Appendix VII through X - Barringer and Associates, Inc.

WASH- 1400(NUREG-75/014)Reactor Safety StudyAn Assessment ofAccident Risks in U.S. CommercialNuclear Power PlantsA.Appendices VII, VIII, IX and X'.4-October 1975United States Nuclear Regulatory Commission

'A7,a

WASH- 1400(NUREG 75/014)aRELEASE OF RADIOACTIVITYinREACTOR ACCIDENTSAPPENDIX VIItoREACTOR SAFETY STUDYbyR. L. RitzmanE. F. AberC. AlexanderM. H. FontanaD. K. LandstromD. L. LessorH. L. McMurryD. L. MorrisonP.C. OwzarskiG. W. ParkerL. F. ParslyM. PobereskinA. K. PostmaH.S. RosenbergW. A. YuillContract W-7405-eng-92Battelle Columbus LaboratoriesU.S. NUCLEAR REGULATORY COMMISSIONOCTOBER 1975

Table of Contents (Continued)SectionPage No.3.1.2 Multicompartment Containment Model (Continued)3.1.2.3 12 Equilibrium ...............................3.1.2.4 Intercompartmental Flow Rates ................3.1.2.5 Design of Corral for PWR Analyses ............3.2 BWR Multicompartment Model ...................................3.2.1 Natural Deposition - Effect of Heat Transferfrom BWR Vessel Walls .................................3.2.2 Pool Scrubbing ........................................3.2.3 Standby Gas Treatment System - (SGTS) .................3.2.4 Natural Deposition in the Drywell Annuar Airspace.....3.2.5 The Computer Code CORRAL .............................3.3 Estimation of Methyl Iodide Fraction Used in PWR and BWRAnalyses With the Corral Code ................................3.3.1 Methyl Iodide Fraction For PWR Analyses ...............3.3.2 Methyl Iodide Fraction For BWR Analyses ...............3.3.3 Fission Product Decontamination During Leakagesto the Atmosphere .....................................3.3.4 Potential Groundwater Contamination in a MeltdownAccident ..............................................VII-25VII-26VII-26VII-27VII-27VII-28VII-28VII-29VII-31VII-33VII-33VII-34VII-34VII-35V3.3.4.13.3.4.23.3.4.33.3.4.43.3.4.5Problem Definition ...........................The Hydraulic and Ion Transport Models .......Results of the Depressurization ReleaseProblem ......................................Results of the Glass Leaching ReleaseProblem ......................................Exposure Reduction Processes andContamination Control Measures ...............VII-35VII-36VII-38VII-38VII-38REFERENCES ...............................................................VII-40APPENDIX AAPPENDIX BAPPENDIX CAPPENDIX DAPPENDIX EAPPENDIX FAPPENDIX GAPPENDIX HAPPENDIX ICALCULATION OF FISSION PRODUCT RELEASE TO FUELROD GAS GAP .................................................RELEASE OF FISSION PRODUCTS FROM PWR AND BWRFUEL PINS ...................................................CALCULATION OF GAP RELEASE OF RADIOACTIVEFISSION PRODUCTS ............................................SURVEY OF EXPERIMENTAL WORK ON FISSIONPRODUCT RELEASE FROM MOLTEN U02 . . . . . . . . . . . . . . . . . . . . . . . . . . . .AN EVALUATION OF FISSION PRODUCT AND FUELCONSTITUENT RELEASE FROM REACTOR FUELSBASED ON A THERMODYNAMIC ANALYSIS OF THECOMPOUND SPECIES PRESENT IN THE FUEL .........................SUMMARY OF DATA ON FISSION PRODUCT RELEASE FROM U0 2DURING OXIDATION IN AIR .....................................ESTIMATIONS OF FISSION PRODUCT RELEASE FROM MELTDURING CONCRETE PENETRATION .................................EXAMINATION OF SOLID FISSION PRODUCT PLATEOUT INPRIMARY SYSTEMS .............................................IODINE DEPOSITION IN PWR AND BWR REACTOR VESSELS ANDASSOCIATED EQUIPMENT .........................................VII-51VII-65VII-85VII-103VII-117VII-127VII-137VII-145VII-157VII-ii

Table of Contents (Continued)SectionPage No.APPENDIX JAPPENDIX KTRANSPORT AND DEPOSITION OF AIRBORNE FISSION PRODUCTSIN CONTAINMENT SYSTEMS OF WATER COOLED REACTORSFOLLOWING POSTULATED ACCIDENTS ..............................DIFFUSION OF RADIOACTIVE FLUID THROUGH SOIL SURROUNDINGA LARGE POWER-REACTOR STATION AFTER A CORE MELTDOWNACCIDENT .....................................................VII-171VII-247List of TablesTablePage No.VII 1-1 Fractions Released to Gap (Total Core) ..........................VII 1-2 Gap Release Component Values ....................................VII 1-3 Meltdown Release Component Values ...............................VII 1-4 Vaporization Release Component Values ...........................VII 1-5 Fission Product Releases During Steam Explosions ................VII 1-6 Fission Product Release Service Summary - BestEstimate Total Core Release Fractions ...........................VII-13/14VII-13/14VII-13/14VII-13/14VII-13/14VII-13/14VII 2-1VII 3-1Estimated Average Temperatures of Upper Structural Material ina BWR Pressure Vessel During Dry Meltdown .......................Fission Product Removal Rate Constants Calculated for aLarge PWR Containment Vessel ....................................VII-19/20VII-43/44VII 3-2 Equilibrium Data for 12 With Boric Acid Sprays ..................VII 3-3 Equilibrium Data for 12 With Caustic Sprays .....................VII 3-4 Input Data and Results for Filter Overbeating Criteria ..........VII 3-5 DF vs Particle Size and Reynolds Number in Drywell AnnularGap..............................................................VII 3-6 Hydraulics Model Parameters .....................................VII-43/44VII-43/44VII-43/44VII-43/44VII-43/44VII 3-7Distribution Coefficients for Various Radionuclides inGroundwater .....................................................VII-43/44VII 3-8 Core Inventory Depressurization Release Fractions ...............VII-43/44VII 3-9VII 3-10Comparison of Calculated Groundwater Effluent Concentrationsto MPC Limits - Depressurization Release Case ...................Comparison of Calculated Groundwater Effluent ConcentrationsTo MPC Limits - Glass Leaching Release Case .....................VII-45/46VII-45/46VII-iii

List of FiguresFigureVII 3-1Drop-Collection Efficiency as a Function of Liquid VolumeSprayed .........................................................Page No.VII-47/48VII 3-2 Schematic of CORRAL-PWR .........................................VII 3-3 Schematic of CORRAL-BWR .........................................VII 3-4 Computer Code CORRAL Flow Diagrams ..............................VII 3-5 Idealized Flow System Used for Analysis .........................VII 3-6Radionuclide Elution Curves for DepressurizationRelease .........................................................VII 3-7 Radionuclide Elution Curves for Glass Leaching Release ..........VII-47/48VII-47/48VII-47/48VII-47/48VII-49/50VII-49/50VII-iv

Appendix VURelease of Radioactivity in Reactor AccidentsIntroductionThis report describes results of theFission Product Source Term Task whichhas been conducted by Battelle'sColumbus Laboratories as part of theU.S. Atomic Energy Commission's ReactorSafety Study. The objective of theReactor Safety Study is the evaluationof postulated accidents in large watercooledpower reactors with respect tothe probability of occurrence and themagnitude of resulting consequences.The primary purpose of the FissionProduct Source Term Task has been tospecify the size of.the fission productsource which would escape the containmentboundary as a function of time forvarious accident conditions defined bythe Reactor Safety Study. In doingthis, specialists in the areas of fissionproduct release, transport, anddeposition met periodically with the ReactorSafety Study to guide the developmentof realistic source term estimates.In addition, key analyses on one or moreportions of the total problem were performedat the different laboratoriesrepresented by each of the specialists.The group was composed of D. L. Morrisonand R. L. Ritzman from Battelle-Columbus, W. A. Yuill from AerojetNuclear Company, A. K. Postma and P. C.Owzarski from Battelle-Northwest, and G.W. Parker and M. H. Fontana from OakRidge National Laboratory.A dominant portion of the effort on thetask concerned fission product behaviorunder reactor core meltdown conditions.Accordingly, considerable interactionoccurred between this task and the ReactorCore Meltdown Task, which was alsoconducted at the Battelle-Columbus Laboratories.The Reactor Core MeltdownTask (Ref. 1) provided data on physicalevents and conditions that wereessential in developing the definitionsand procedures used to specify fissionproduct movement within and loss fromthe containment boundary.This report specifically presents themethodology that was evolved to enablethe performance of calculations of fissionproduct escape to the atmospherefor various accident sequences in alarge PWR or BWR. The essentials of themethodology are described in the sectionswhich follow. Supporting documentationor more detailed treatment ofcertain components can be found in thecollection of papers that are appendedat the rear of the report.VII-1

Section 1Fission Product Release FromReactor Core MaterialFission product release from core materialduring accidents involving meltdownwould probably occur more or less continuouslyuntil the system finallycools. During this period, releaserates should vary over wide limitsdepending on fission product properties,* system temperatures, and the surface tovolume ratio of the molten material.However, it is possible to identify fourconditions or times at which majordriving forces for release exist. Theseperiods of high rates should account formost of the total release. The fourmajor release components are:a. Gap release - fission productrelease which occurs when the claddingsexperience initial rupture.It consists mostly of activity thatwas released to void spaces withinthe fuel rods during normal reactoroperation and rapid depressurizationof contained gases provides thedriving force for escape.b. Meltdown release - fission productrelease which occurs from the fuelwhile it first heats to melting andbecomes molten. High gas flows inthe core during this period sweepthe activity out of the core region.c. Vaporization release - fissionproduct release which occurs afterlarge amounts of molten core materialfall into the reactor cavityfrom the pressure vessel. Turbulencecaused by internal convectionand melt sparging by gaseous decom-* position products of concrete producethe driving forces for escape.d. oxidation release - fission productrelease which occurs just after and- is a result of a steam explosionevent. Finely divided fuel materialis scattered into an oxygen atmosphereand undergoes extensive oxidationwhich liberates specific f issionproducts.By concentrating on the processes andfactors which will control these fourcomponents, it was decided that releaseterms for representative accident sequencescould be developed. Each onepertains to a specific, identifiabletime period in the core meltdown scenarioof either water reactor type. Therefore,PWR and BWR analyses both canutilize the same set of release componentsbut the particular number ofcomponents and the timing of the releaseswill depend on the reactor typeand the particular accident sequencebeing examined. The subsections whichfollow describe each component in moredetail and present the release valuesthat were arrived at on the basis of thecurrent interpretations of fission productbehavior that are contained inappendices to this report. The relevantappendices are cited within the text.1.1 GAP RELEASE COMPONENTFor the purposes of this work, gaprelease is defined as the fission productinventory that is free to escape ingaseous or vapor form from core fuelrods if the cladding ruptures. Thenumber of core fuel rod claddings thatwill rupture depends on the effectivenessof the emergency core coolingsystems. For highly effective emergencycooling no claddings may rupture, whilefor degraded cooling conditions, leadingto fuel rod melting, essentially allcladdings will rupture during thetemperature rise to melting. Betweenthese two extremes there exist a largeseries of core cooling temperature conditionswhich can lead to various percentagesof cladding ruptures.Cladding rupture. temperatures dependupon several factors; rate of temperaturerise, internal gas pressure, andcladding physical and mechanical properties.However, rupture temperatures arelikely to range from about 1400 to2000 F. Fission products which havemigrated to the surface of the fuelpellets or to the interior surface ofthe cladding during normal reactoroperation can potentially be released.The driving force for escape comes fromthe rapid release of inert gases (heliumand fission gas) stored in the plenumand gas gap spaces of the fuel rods.Internal gas pressures ranging from afew hundred psia to 2000 psia can exist.only fission product material whichexists in vapor form should escape duringthe rod depressurization. Therefore,if vaporization from condensedphases or reaction layers within thefuel rod is incomplete when depressuri-VII-3

zation takes place, the potential releasewould not be realized. After thedepressurization, very little drivingforce exists to carry fission productvapors along the narrow annulus to therupture location in the cladding.Consequently, in this work the gaprelease component will be confined to anestimate of the release that occurs atthe time of cladding rupture only. Foraccident conditions which are less severethan core meltdown, the gap releasecomponent values developed here may bereduced according to the percentage ofthe fuel rod claddings that are expectedto experience rupture.Two fractions make up the gap releasecomponent: (1) the release fraction,and (2) the escape fraction. Therelease fraction defines the potentialfor release from U0 2 fuel and is theresult obtained from most fission productrelease models. The escape fractionrepresents an estimate of the degree ofvolatility of the fission product atfuel rod cladding, rupture. The processesand reactions that need to be consideredin estimating the escape fractioninclude physical condensation and reactionswith fuel material, cladding,other fission products and gaseous impuritiesin the rods.1.1.1 THE RELEASE FRACTIONThe three sets of gap release calculationswhich are reported in AppendicesA, B, and C produced the release fractionestimates shown in Table VII 1-1.These results are based on PWR coreproperties but results for a BWR are sosimilar that Table VII 1-1 values willbe used for both reactor types. Thevalues in Table VII 1-1 also representbest estimate releases for the severalspecies and these contain differentuncertainties depending upon the releasemodels used, variations in basic parameters,or differences in methods to computetemperature profiles in operatingfuel. Inspection of results given inAppendices A, B, and C show that uncertaintiesrange from factors of +2 forsome species to factors +10 or mor-e forothers. As a rule, the magnitude of theuncertainty tends to decrease as thedecay half-life increases.The ultimate use of these release fractionsin total accident release calculationsdemands that a single value beassigned for each chemical group; thatis, isotopic dependent release behaviormust be ignored. Consequently, in thelast column of Table VII 1-1, averagerelease values are listed which will beused for all isotopes of each of theapplicable chemical elements. The averagevalues in each case are the simplearithmetic means of the three calculatedreleases for the particular isotope.This isotope was selected on the basisthat it represents the radiologicallyimportant one for the element. Notethat no direct calculations of telluriumreleases were made. Results of out-ofpileexperiments (Ref. 2) indicate thatits release should be similar to iodineand cesium, and on this basis the valueof 0.10 for the principal isotope, Te-132, was selected. Since the isotopesthat were selected to represent eachchemical group have relatively long halflives, the uncertainties in the averagerelease fractions should be nearer thelower end of the uncertainty range asnoted above. This was the basis for theuncertainty factors that are specifiedin Table VII 1-1.The data in Table VII 1-1 represent thebest effort that can be made withcurrent mathematical models in calculatingfission product releases duringreactor operation. Due to lack of basicinformation for some species or parameters,simplifying assumptions are oftenmade which tend to overestimate ratherthan underestimate releases. Therefore,the results should be interpreted ascurrent state-of-the-art and not asabsolute values. Future experimentaland theoretical study may indicate lowerreleases, and when this occurs themodels and results used here can bemodified.1.1.2 THE ESCAPE FRACTIONThe rationale used to select escapefraction values will be discussed on anelement by element basis.1.1.2.1 Noble Gases.The noble gases are gaseous at roomtemperature and are known to be veryunreactive chemically. At claddingrupture the only mechanism which couldretard their escape would be flowrestrictions along the gas gap to thehole or split in the cladding. Sincethis constitutes only a delay process,which depends on very specific detailsof fuel rod structure, no retention ofreleased noble gases in the fuel rod gasspace can be claimed.1.1.2.2 Halogens.Elemental iodine would be entirelygaseous at normal reactor fuel rodoperating temperatures and particularilyat the cladding rupture temperatures.However, iodine readily reacts withVII-4

metals to form iodides which havedifferent volatilities. Possibilitiesinclude zirconium iodides (reaction withcladding), cesium iodine (reaction withfission-product cesium) , and hydrogeniodide (reaction with trace hydrogen orwater vapor).Experimental work by Feuerstein (Ref. 3)has shown that a series of zirconiumiodides can form when iodine in Zircaloycapsules is heated, and several minutesis required to volatilize appreciablefractions of the reaction product at atemperature of 800 C (1472 F) . Indirectevidence of iodine reaction withZircaloy in operating fuel rods has beenobtained by Weidenbaum (Ref. 4).Collins, et al (Ref. 5) performed puncturetests with irradiated Zircaloy-2clad U6 2 in steam at about 1000 C(1832 F) from which it was concludedthat only 10 percent of the iodine thatwas expected to be free within the cansescaped through the puncture hole.Lorenz and Parker (Ref. 6) haveconducted a pair of in-reactor fuel rodfailure transient tests with pressurizedZircaloy-2 clad fuel rods. The fissionproduct release data for the two experimentsindicated 25 percent and 100percent escape of the free iodine relativeto free noble gases. These limitedstudies suggest that iodine retention byZircaloy cladding could limit the escapeof iodine from the gas spaces of fuelrods during rupture. The escape fractionvalue cannot be specified veryaccurately but would be expected to fallwithin the range 0.1 to 1.Contrary to the cladding reaction mechanism,thermodynamic analyses of thefuel-cladding fission product systemconsistently predict that CsI would bethe most stable chemical form for iodineat elevated temperatures (see AppendixE). Since the fission yield for cesiumisotopes is more than ten times that foriodine isotopes, there is sufficientcesium for complete conversion of theiodine. If CsI is the dominant iodinespecies in fuel rods then iodine wouldexhibit a significantly lower volatilityat cladding rupture temperature i.e., avapor fraction in the range of 0.01 to0.1 would be expected. As shown above,experimental escape data do not coincidewith these low values. In additionthere is limited evidence that cesiummay undergo compound formation with U0 2and thus prevent formation of CsI (Refs.7,8). Thermodynamic analyses have notconsidered this reaction. Finally,there appears to be no experimental confirmationof the presence of CsI inirradiated fuel rods. Therefore, thepossibility of CsI being a major chemicalform is not sufficiently establishedto justify consideration in this work.However, additional experimental work inthis area would be useful.The formation of hydrogen iodide, whichmight be significant at high temperatures,also has not been verified byexperimental work on irradiated fuelrods. The existence of HI as a majorspecies would not alter iodine volatilityat cladding rupture temperaturesappreciably. Therefore HI will not beconsidered an important iodine form inthis analysis.In summary the escape fraction foriodine gap release should be based onavailable experimental evidence thatindicates at least partial retention byZircaloy cladding. On the basis of therange indicated a best estimate value of1/3 with an uncertainty factor +3 isappropriate.1.1.2.3 Alkali Metals.The normal boiling point of cesium metalis about 960 K (1270 F) and if thefission product exists in this elementalform, the gap release fraction could becompletely vaporized at cladding rupturetemperatures. On the other hand thethermal transient may be too rapid andincomplete vaporization would haveoccurred at this point. Also, compoundformation with the fuel (noted above) orwith the cladding (possibly with corrosionproducts) could result in significantlyreduced volatility. There isalmost no experimental data related toescape of fission product cesium underthese conditions. The in-pile transienttests of Lorenz and Parker (Ref. 6)provide the only known experimentalestimate. The results of two testsindicate an escape fraction value ofabout 2/3. Because of the approximatenature of these measurements, it wasdecided to use the same escape fractionfor cesium that is used for iodine;i.e., 1/3 with an uncertainty factor of+3.1.1.2.4 Alkaline Earths.Depending upon the oxygen activity inthe system, fission product strontiumwould predominately exist as either themetal or the monoxide. However, neithercondensed phase has appreciable volatilityat clad rupture temperatures. Themetal which exhibits the higher vaporpressure should limit vapor fractions toless than l0-4 of the total strontium.The ANC fission product release modelcalculation for strontium, which considersthermodynamic equilibrium in theVII-5

fuel body, obtained a maximum releasefraction of 4 x 10-6 (Appendix B) . Thein-pile transient test data of Lorenzand Parker (Ref. 6), while quite crudefor strontium, indicat~ escape 6 fractionsranging from about 10- to 10-6. On thebasis of this evidence, it appears thatin conjunction with a release fractionof 0.01, a best estimate value for theescape fraction would be 10-4 with anuncertainty factor of +100.1.1.2.5 Tellurium.Thermodynamic calculations indicate thattellurium can exist as either theelement or an oxide in the fuel. Thestable vapor form at cladding rupturetemperatures is probably TeZ, butseveral experimental studies indicatethat tellurium will react with Zircaloy.Genco, et al. (Ref. 9) demonstratedextensive reaction of tellurium vaporwith zirconium at temperatures above 400C (752 F). The in-pile transient testsof Lorenz and Parker (Ref. 6) while verylimited indicate an escape fraction fortellurium of between 10-1 and 10-5. Acomplex kinetic situation involvingcompetition among vaporization, reactionwith cladding, and escape in the gaseouspuff probably determines the escapefraction. Therefore, the value was setat 10-3 with an uncertainty factor of+100.1.1.2.6 Other Species.The volatilities of other fissionproduct species, besides those which arechemical analogs of the elements discussedabove, are expected to be so lowthat their escape at cladding rupturecan be considered negligible. Therefore,the results of all the gap releasecomponent analysis can now be summarized.Table VII 1-2 presents such asummary showing the gap release fractions,the escape fractions, and theproduct of these two - the gap releasecomponent. The release and escapefractions listed here may be somewhatdifferent from values obtained or recommendedby the individual laboratoriesthat contributed to the analyses. Thisis because current knowledge does notclearly show one analytical approach issuperior to the others.1.2 MELTDOWN RELEASE COMPONENTThe conditions pertaining to this releaseperiod begin with rapid boiloff ofthe water coolant which uncovers thereactor core. Steam, flowing up throughthe heating core, initiates Zr-H 2 0 reactionand this accelerates the rate oftemperature rise. The cladding begins tomelt within one minute and in a few moreminutes fuel-melting temperatures areapproached in the hotter regions. Theprocess spreads throughout the core andwithin 30 minutes to 2 hours (Ref. 1)nearly the whole core is molten attemperatures ranging from roughly 2000to 3000 C. During the later stages ofthis process molten core material canrun through or melt through the gridplate and fall into the bottom of thepressure vessel. If a steam~ explosiondoes not occur when residual water iscontacted in the lower portion of thepressure vessel, partial quenching andtemporary solidification of portions ofthe molten mass can take place. However,the internal heat generationcauses remelting and the inevitabledownward migration continues until thepressure vessel fails, probably bymeltthrough. Pressure vessel failure isexpected to require about 1 hour aftermost of the core has melted (Ref. 1).Prior to this the high internal temperatureshave caused melting of some of thepressure vessel steel and interiorstructural components. The molten ironis not miscible with the core material(oxide phase) although partial conversionto iron oxides could produce somedissolution in and dilution of the corematerial. Nevertheless some fissionproducts (i.e., the noble metals) wouldtend to distribute to the metallic ironphase.Initial fuel melting is expected tooccur in only the center regions ofthe rods on almost a pellet by pelletscale. Thus the melting fuel will offera relatively high surface area forrelease of fission products. As themelting front moves outward, the meltingof the individual pellets may continue,but it is also conceivable that largersections of fuel may collapse. into themolten mass. If this fuel melts withinthe mass rather than at the edge, thenfission product release could beinhibited by the time required fortransport to a free surface. On theother hand, gaseous fission products,present as bubbles in the U02 could risequickly to the surface of the moltenmass and escape. It appears that mostof the fission product release that doesoccur will take place early in themelting period at each core location.Then as the melted fuel mixes with therest of the molten mass and the massincreases in. size, fission productrelease rates will 'become much slower.The melting of structural steel in thepressure vessel during this later periodis expected to produce a layer of molteniron above the molten core materialwhich would offer a further barrier toa.4VII-6

fission product release. Other factorswhich can inhibit release during meltdownin the pressure vessel include thepossibility of crust formation at themelt surface and partial quenching whenmelt runs or falls into water that maybe left in the bottom of the vessel.The atmosphere in the core region andpressure vessel during meltdown isexpected to be a steam-hydrogen mixturewith small concentrations of fissionproduct and core material vapors andaerosols. This may be classified anonoxidizing atmosphere for most fissionproducts, and it, of course, resultsfrom partial consumption of steam bymetal-water reaction, yielding H 2 in thecore region. Although the metal-waterreaction that does occur is steam supplylimited, some steam flow passes throughcooler portions of the core regionwithout complete reaction. It is estimatedthat during the meltdown phaseonly about half the Zircaloy is reactedand other metal-water reaction producesonly about 50 percent more H 2 . Thustotal metal-water reaction is only theequivalent of 75 percent Zr-H 2 0 reaction.Thermal analyses of core meltdown provideonly generalized data on core temperatureprofiles, geometry changes, andmelt behavior versus time. This, combinedwith the uncertainties which existin fission product properties at veryhigh temperatures, argue against constructionof a highly mechanistic modelto calculate fission product releaseduring the meltdown phase. Therefore,in this work, fission product release istreated as being simply proportional tothe fraction of core melted.Mathematically,RCFx(t) = [RFx] - [FCM(t)]where RCFx(t) = Core release fractionfor fission product x as a function oftime (t):RFx = Release fraction of fissionproduct x from melted fuelFCM (t) = Fraction of core melted as afunction of time (t)It is important to note that this approachassigns all release, which isexpected to occur during the time corematerial remains in the pressure vessel,to the early period of first melting.This is consistent with two key observations;(1) the highest steam flows and(2) the highest fuel surface areas areexpected during the early period. Thusthis should be the period of maximumdriving force for fission product escapefrom the core region. The releasefractions (RFx) for the various fissionproducts were estimated by considering:a. The limited datafrom small-scaleU0 2 (Appendix D),that are availableexperiments withandb. The predictions of limited thermodynamicanalyses of fuel-claddingfissionproduct system at high temperature(Appendix E)The results are summarized in Table VII1-3 and a short description of therationale connected with each value isgiven in the following paragraphs.a. Noble Cases (Xe, Kr) - Experimentalwork shows that nearly total releaseof these essentially chemicallyinert gases would be expected duringthe meltdown period if the surfaceto-volumeratio of melting materialremains high. Although this seemslikely and considerable releaseshould occur even before the fuelmelts, some gases could becometrapped as the molten mass enlargesduring the later stages of meltdown.Accordingly, a range of 50 to 100percent release is consideredreasonable but 90 percent should beassumed probable.b. Halogens (I, Br) - Again nearlytotal release is expected due to thehigh volatility, but the rate ofrelease could be limited bytransport in the melt to an externalsurface. Therefore, the same range(50-100 percent release) shouldapply and 90 percent is consideredthe probable value.c. Alkali Metals (Cs, Rb) - Nearly totalrelease would be expected butexperimental data on cesium releasefrom molten U0 2 and thermodynamicstudies show that the alkali metalsare not so highly volatile as thenoble gases or halogens. Therefore,the release rate could be somewhataffected by internal transport inthe melt or by the possible tendencytoward compound formation. Experimentalevidence indicates a range of40 to 90 percent with a probablevalue of 80 percent.d. Tellurium - Simple thermodynamicsindicate that Te should volatilizealmost completely from melting corematerial in the elemental form.However, experimental data indicateVII-7

extensive reaction with unoxidizedZircaloy cladding would tend to holdtellurium in the melt, even thoughmuch of the cladding may oxidizeduring the meltdown period. Thetellurium apparently migrates fartherinto the cladding to react withremaining free metal rather than,diffuse out through the oxide layer.Release of the tellurium from aparticular core region will occurwhen nearly all or all of the claddinghas been oxidized. Since anaverage of 50 percent of the corecladding is expected to become oxidizedduring meltdown, this representsan upper limit for release.However, the oxidation is spreadunevenly over the core and a smalleramount of cladding undergoes completereaction. On this basis telluriumrelease is estimated to rangefrom 5 to 25 percent.e. Alkaline Earths (Sr, Ba) - The chemicalform and the volatility ofthese two fission products are verysensitive to the oxygen partialpressure in the system. Strontiummetal is more volatile than bariummetal but barium oxide is morevolatile than strontium oxide.Thermodynamic analyses produce conflictingestimates of volatility andchemical form due to variations inoxygen activities. Experimentaldata on release from molten fuelmaterial indicate that the two elementswould experience about thesame release. Data obtained withzirconium clad U0 2 showed up to 20percent release of strontium andbarium over several minutes in aneutral atmosphere, while only a fewpercent loss was found for bare orstainless steel clad U0 2 . The lowervolatility of these elements and theprobable existence of unoxidizedcladding suggest that releases inthe range of 2 to 20 percent wouldoccur. Since incomplete claddingoxidation is expected, the probablerelease value should lie above thegeometric mean for this range.Hence, 10 percent is used as theprobable value.f. Noble Metals (Ru, Rh, Pd, Mo, Tc) -Although these elements probablyvolatilize as the oxides, thermodynamiccalculations suggest that thevolatile oxide forms are not verystable at the high temperatures andthe lower oxygen partial pressurethat are expected to be associatedwith the core meltdown. The firstthree elements probably exist in themetallic form in the fuel, while thelatter two are probably in the formof lower oxides. The metallic speciescould partially distribute tothe molten iron phase and be retained.Experimental data on rutheniumrelease from molten U0 2 in anoxygen-deficient atmosphere indicatelow release. Releases in the rangeof 1 to 10 percent are consideredpossible, and 3 percent (the approximategeometric mean) is used as theprobable value.g. Rare Earths (including Y and Np,Pu) - The rare earth elements willgenerally exist in the fuel as thesesquioxides (M 2 0 3 ) while the actinides,neptunium and plutonium,should form the dioxides (MO 2 ). Theoxides characteristically exhibitlow volatility but an estimate ofthe release fraction is difficult tomake. Experimental data for ceriumrelease from small specimens ofmolten UO 2 indicate losses ofseveral tenths of a percent over afew minutes, or about the same asthe U02 vaporization loss. InAppendix H, estimates of fuelvaporization rates indicate lossesin the range of 0.01 to 1 percent.This range can also be used for therare earth species but the probablevalue, 0.3 percent, reflects cautionin selecting a characteristic releasewhen the estimate is soapproximate.h. Refractory Oxides (Zr, Nb) - Theoxides of these elements are sostable and of such low volatilitythat they probably would experienceless release than the rare earths.However, for simplicity the samerelease definitions are used here asfor the rare earths.1.3 VAPORIZATION RELEASE COMPONENTWhen the molten core and iron penetratethe pressure vessel and fall (or run)into the reactor cavity, the materialwill be exposed to oxygen from the containmentatmosphere, steam from contactwith water or vaporized from concrete,and carbon dioxide from thermal decompositionof the limestone aggregate in theconcrete. Passage of steam and carbondioxide through the molten mass willproduce a gas sparging effect. In additionthese gases or their dissociationproducts will create highly oxidizingconditions. One can speculate that theiron phase would be (at least partially)converted to oxide which could thendissolve in the core melt. The meltshould also contain products of theconcrete decomposition; silica, calciumVII-8

Ssilicates, or calcium oxide. The productswould eventually reduce the densityof the oxide phase and the thenheavier iron phase might sink to thebottom of the mass (Ref. 10) . Conversely,incomplete dissolution could leave arelatively pure U0 2 phase which wouldcontinue penetrating downwards (Ref.11). Lower melting mixed oxide phasesforming ahead of the U0 2 would tend torise and cover the U02. Thus, developmentof several immiscible or partiallymiscible phases is conceivable. Internalconvection would promote mixingwithin and exchange between phases.Depending upon their solute properties,fission product oxides could distributeto these phases thereby altering theheat source distribution and the temperatureprofile in the melt system (Ref.10). Vaporization of fuel and structuralmaterials from the upper surface ofthe molten mass should produce denseaerosol clouds (smokes) above the meltand buildup of condensation products onnearby surfaces. ' Agglomeration andgrowth of smoke particles is expected tocause some vaporized material to settleback. Thus much of the vaporized materialshould be retained in the reactorvessel cavity. Vaporized fission products,mixed with the much larger quantitiesof structural material vapors andsmoke, should generally follow the distributionof the bulk material. Exceptionsto this would be the highvolatility species which could escapeinto the upper part of the containment.only highly simplified analyses of thephysical situation just outlined havebeen performed (Refs. 10,11,12). Thereare many unknown details concerning mostof the chemical, physical, thermal, mechanical,and metallurgical propertiesof the complex system. Analytical resultsare dependent on hasic assumptionswhich differ among models. No largescaleexperimental work on the relevantsystem has been performed to guidemodeling. Concrete penetration ratescannot be estimated accurately becauseof uncertainties in heat transfer mechanism,melt interaction effects, and/orboundary limit definitions. Releasecalculations can and have been performed,but the assumptions and data extrapolationsneeded, lead to estimates thatare usually upper limit values (Ref. 11)(See Appendices E and G). The releasemodels are useful in identifying keyrelease processes and species that havea high potential for release by one ormore of these processes.To a first approximation, the extent ofrelease will depend upon species volatilityand the rate of transport in themolten mass to an external surface. Forthe large molten masses that result fromcore meltdown, the latter process shouldcontrol release for all except the verylow volatility species. For example,estimates have shown that pure diffusiontransport would require several hundredhours to achieve significant releasefractions, regardless of volatility(Ref. 13). However, additional estimatesindicated that thermally inducedinternal convection currents might increasemass transfer rates such thatcorresponding releases would occurwithin several hours (Ref. 13). Veryrecent approximations based on gassparging assumptions suggest total releasefor volatile species in fractionsof an hour (Appendix G). The releaserates that would actually occur wouldprobably be some mixture between thelatter two processes. Also, gross vaporizationof the melt could assist therelease by creating a receding surface.The uncertainties of the problem requirethat a simple approach be used tospecify fission product releases forthis portion of the accident.1.3.1 VOLATILE FISSION PRODUCTSThe very volatile fission products willescape the melt if they can reach anexternal surface. The gas spargingprocess offers a mechanism for inducingmixing and creating a large effectiveexternal surface area. The convectivemass transfer process offers an alternativeand usually slower rate fortransport to surfaces where vaporizationcan occur. Both processes would resultin an exponential decrease in the volatilefission product inventory withtime. In each case the half-time forrelease might range from less than onehour to many hours, depending on the gasflow or mass transfer conditions.Other work on the Reactor Safety Studyhas resulted in an estimated time forcore penetration of the concrete base ofabout 18 hours (Ref. 1) . The analysesalso indicate that considerable spallingand decomposition of the concrete wouldoccur within the first half-hour ofcontact. This would be a period ofrapid gas (steam and carbon dioxide)generation during which the spargingprocess could be a dominant drivingforce for escape of volatile fissionproducts. Subsequently, lower rates ofrelease would be expected as gas flowsdecrease but sparging could stillefficiently deplete the melt of volatilespecies (See Appendix G). Thermallyinduced internal convection and alsosurface evaporation might assist in therelease.VII -9

The potential importance of gas spargingto the release of volatiles from themassive melt dictated that treatment ofthe vaporization release componentshould be based on this process.Accordingly, a pseudo-exponential rateexpression was designed which requiredonly a single input parameter, thecharacteristic release half-time.Mathematically,whereVLF(t) = 1 -exp [-0.693t/T]VLF(t) = The vaporization losstion after time (t)frac-T = The characteristic releasehalf-timeIn order to avoid excessive calculationsduring actual accident sequence analyses,a cut-off time for the expression,equivalent to four half-time intervals,was selected. In practice the rateexpression was used exactly as writtenfor the first three half-time internvals,but then complete escape of theremaining activity was compressed intothe fourth interval. Since this laststep involves only 12.5% of the totalvaporization release, the approximationshould produce only slight perturbationsin results.The value of the characteristic releasehalf-time would be a function of thefission product volatility, the gassparging rate, and other kinetic factors.In a rigorous sense, there wouldbe a unique value for each fissionproduct species which would vary con-'tinuously with the sparging conditions.This complexity is not warranted hereand so the half-time value was assignedonly to roughly match the period of highsparge gas flow. This period, as notedearlier, has been estimated to last onthe order of one-half hour. Thus thebest estimate half-time value is consideredto be 30 minutes. Use of thisuniform value for all fission productsprobably produces underestimates of therate of release for the most volatilespecies and overestimates of the rate ofrelease for species of lesser volatility.Also, the sparging process may notbe fully effective in sweeping volatilefission products from the melt eitherbecause some sections of the melt arenot exposed to sparge gas or becausemass transfer limitations inhibit vaporsaturation of sparge gas bubbles. Dueto uncertainties such as these the 30minute half-time value which is used incalculations should be considered uncertainby at least an order of magnitude.The fission products that are sufficientlyvolatile to experience totalloss during this vaporization phase areXe, Kr, I, Br, Cs, Rb, Te, Se, and Sb.The meltdown release of the first six ofthese is expected to be quite large,because of their high volatilities, andso little will be left to contribute tothe vaporization release component.This is not true for the tellurium groupelements, whose meltdown release isconsidered inhibited by compound formationwith Zircaloy cladding. However,by the time the vaporization releasebegins most of the free zirconium, whichexists during core meltdown, should beoxidized (probably from reaction withU0 2 ). If not, then the oxidizing spargegases (steam and C0 2 ) should quicklyeliminate free zirconium so that telluriumand its chemical analogues,selenium and antimony, will become veryvolatile and also experience totalrelease.1.3.2 LOW VOLATILITY FISSION PRODUCTSEssentially all the remaining importantfission product species must be consideredlow volatility components. Thustotal release of these species shouldnot occur. With one exception thisgroup of elements should be present asoxides in the oxide phase. The noblemetals (Ru, Rh, Pd) and to a lesserextent Mo and Tc probably exist as themetals and would be expected topartition into the metallic iron phaseof the molten system (Ref. 14). Lowrelease (less than 1 percent) of thesespecies is expected unless completeoxidation of the iron should occur(Appendix G). Although this is notconsidered likely, localized oxidationcould lead to some release and a valueof 5 percent is considered to be arealistic estimate. This is probablyuncertain by a factor of + 5. The otherfission products (alkaline earth oxidesand rare earth oxides) should be dissolvedin the oxide phase of the melt(Ref. 14). Under the generally oxidizingconditions which persist over thisperiod these species are less volatilethan U02 (Appendix E). The vaporizationof U0 2 and of other oxide materials inthe melt is very difficult to estimateowing to uncertainties in compositionand temperature. High vaporizationrates (high temperatures and oxygenpressures) should produce dense aerosolclouds above the melt which would tendto settle out, carrying condensablefission products back down. Low vaporizationrates (lower temperatures) wouldalso indicate low losses for these fissionproducts. Considering these limitingprocesses it is doubtful that moreapVII-10

athan 1 percent of these fission productspecies could be distributed to theatmosphere in the containment over thecourse of the reactor melt-throughperiod. This estimate is also probablyuncertain by a factor of +5.The rate of release of the low volatilityfission products, for lack of betterdefinition, is assumed here to followthe same exponential function that isused to describe the release rate of thevolatile species. On this basis thepercentage release values given in TableVII 1-4 indicate the amount of eachfission product, remaining in the corematerial after gap release and melt-downrelease have occurred, that will escapeto the containment atmosphere during thevaporization period.1.4 OXIDATION RELEASE COMPONENTA steam explosion event will result inthe scattering of finely divided U0 2(containing fission products) into theatmosphere outside the containment orinto the air-steam atmosphere insidecontainment. In either case, the U0 2particles will cool and undergo reactionwith oxygen to form U 3 0 8 at temperaturesbelow about 1500 C (Ref. 15). Thereaction is exothermic and is accompaniedby release of fission productsthat are volatile under these conditions.Oak Ridge work on measurement offission product release during fueloxidation by air at elevated temperaturesis directly applicable (Ref. 16).These data summarized and discussed inAppendix F show large releases of raregases, iodine, tellurium, and rutheniumduring 10 to 15 minute exposures to airat temperatures of 1100 and 1200 C.Since the data indicate positive temperaturecoefficients for each species,comparable releases should occur in muchshorter times at higher temperatures.On this basis the release percentagegiven in Table VII 1-5 may be treated asessentially instantaneous values. Note,however, that the releases apply only tothe fraction of the U0 2 fuel that isexpected to be dispersed into an air(oxygen)-containing atmosphere.1.5 USE OF RELEASE COMPONENT VALUESThe four release components describedabove would occur more or less sequentiallyduring a reactor meltdown accident.In all cases, the gap componentwould occur first, followed by themeltdown component, and then by thevaporization component. However, steamexplosions could potentially occur anytime after appreciable amounts of thecore have melted. Thus, the oxidationcomponent is somewhat randomly timeoriented.In using the release component valuesfrom Tables VII 1-2 through VII 1-5 tospecify release source terms for a particularaccident sequence, it should beobvious that proper inventory balancesfor each fission product must be maintained.For example, the fraction ofthe total inventory that experiences gaprelease is then not available forrelease by any of the other threeprocesses. Consequently, care must beexercised in setting up release sou=,eterms. To illustrate this point, abasic release source summary is providedin Table VII 1-6. Here individual corerelease fractions are given for eachcomponent and fission product assumingthat, except for the steam explosionfraction, the total core is involved inthe release processes. That is, allfuel rod claddings rupture to give thegap release fraction, total core meltingoccurs, and all of the core meltcontributes to the vaporization releasefraction (unless preceded by a steamexplosion). It is also implicitlyassumed that a steam explosion will notprecede total meltdown release.It is emphasized that the single valueslisted in Table VII 1-6 are based on thebest estimate values taken from TablesVII 1-2 through 1-5. Each of the valuescontains uncertainties as noted in thoseTables and should, therefore, not beconsidered absolute release fractions.VII-11

References1. Carbiener, W. A., et al "Physical Processes in Reactor Meltdown Accidents",WASH-1400, Appendix VIII, October, 1975.2. Parker, G. W., et al., "Out-of-Pile Studies of Fission Product Release fromOverheated Reactor Fuels at ORNL, 1955-1965", ORNL-3981, p 81 (July, 1967).3. Feuerstein, H., "Behavior of Iodine in Zircaloy Capsules", ORNL-4543 (August,1970).4. Weidenbaum, B., et al., "Release of Fission Products from U0 2 Operating at HighPower Rating", CONF-650407, Vol. 2, p 885-904, USAEC (1965).5. Collins, R. D., et al., "Air Cleaning for Reactors with Vented Containments",CONF-660904, Vol. 1, p 419-452 (1967).6. Lorenz, R. A., and G. W. Parker, "Final Report on the Second Fuel Rod FailureTransient Test of a Zircaloy-Clad Fuel Rod Cluster in TREAT", ORNL-4710 (January,1972).7. Chen, H., and P. E. Blackburn, "ANL Reactor Development Program Progress Reportfor September, 1969", ANL-7618, p 117 (October, 1969).8. Crouthamel, C. E. and I. Johnson, "ANL Reactor Development Program Progress Reportfor July, 1971", ANL-7845, p 522 (August, 1971).9. Genco, J. M., et al., "Fission-Product Deposition and Its Enhancement Under ReactorAccident Conditions: Deposition on Primary System Surfaces", BMI-1863,p 38 (March, 1969).10. Jansen, G. and D. D. Stepnewski, "Fast Reactor Fuel Interactions with FloorMaterial After a Hypothetical Core Meltdown", Nucl. Tech. 17, 85-95 (1973).11. Morrison, D. L., et al., "An Evaluation of the Applicability of Existing Data tothe Analytical Description of a Nuclear Reactor Accident-Core Meltdown Evaluation",BMI-1910, Appendix B (July, 1971).12. Ergen, W. K., et al., "Emergency Core Cooling: Report of Advisory Task Force onPower Reactor Emergency Cooling (USAEC).13. Fontana, M. H., "An Estimate of the Enhancement of Fission Product Release fromMolten Fuel by Thermally Induced Internal Circulation", Nucl Appl, 9, 364-375(1970).14. Fischer, J., J. D. Schilb, and M. G. Chasonov, "Investigation of the Distributionof Fission Products Among Molten Fuel and Reactor Phases. Part I - TheDistribution of Fission Products Between Molten Iron and Molten Uranium Dioxide",ANL-7864 (October, 1971).15. Parker, G. W., et al., "Out-of-Pile Studies of Fission Product Release fromOverheated Reactor Fuels at ORNL, 1955-1965", ORNL-3981, p 4 (July, 1967).16. Parker, G. W., et al., "Out-of-Pile Studies of Fission Product Release fromOverheated Reactor Fuels at ORNL, 1955-1965", ORNL-3981, p 85 (July, 1967).VII-12

TABLE VII 1-1 FRACTIONS RELEASED TO GAP (TOTAL CORE)Calculated FractionsAverageFission Product Chemical Release(decay half-life) ANC BCL(b) ORNL Groups FractionXe, Kr (long lived) 0.06 0.10 0.08Xe-133 (5.27 day) 0.04 0.02 0.02 Noble Gases 0 . 0 3 (d)Xe-135 (9.2 hour) 0.0002 0.004 0.004I, Br (long lived) 0.06 0.10 0.141-131 (8.06 day) 0.06 0.03 0.05 Halogens 0.051-132 (2.3 hour) 0.0007 0.005 0.0061-133 (20.8 hour) 0.007 0.01 0.02Cs, Rb (long lived) 0.20 0.05 0.21 0.15(e)Cs-138 (32 minutes) 0.00001 0.0005 0.005 Alkali MetalsSr, Ba (long lived) 0.000004(f) 0.02 0.02 0.01(d)Sr-89 (51 day) - 0.01 0.015Sr-91 (9.7 hour) - 0.002 0.01 Alkaline EarthsTe-132 (78 hour) (estimated value) Tellurium 0.10(d)(a)(b)(c)See Appendix BSee Appendix ASee Appendix C(d) Values can be higher or lower by a factor of 4(e) Value can be higher by a factor of. 2 or lower by a factor of 4(f)This value results from thermodynamic restrictions not considered in the other twmodels. See discussion of the escape fraction for this species.TABLE VII 1-2 GAP RELEASE COMPONENT VALUESFission Gap Gap Total GapProduct Release Escape ReleaseSpecies Fraction Fraction ValueXe, Kr 0 . 0 3(a) 1 0.03I-Br 0 . 0 5 (a) 1 / 3 (c) 0.017Cs, Rb 0 . 1 5 (b) 1 /3(c) 0.05Sr, Ba 0 . 0 1 (a) 1 0 -4(d) 0.000001Te, Se, Sb 0 . 1 0 (a) 1 0 -3(d) 0.0001Others - Negligible(e)(a)(b)(c)(d)(e)Values can be higher or lower by a factor of 4Value can be higher by a factor of 2 or lower by a factor of 4Values can be higher or lower by a factor of 3Values can be higher or lower by a factor of 100While no numerical value was developed for these various species, the numbershould not exceed that used for strontium-barium.

TABLE VII 1-3 MELTDOWN RELEASE COMPONENT VALUESRelease RangeBest EstimateElements (percent) (percent)Xe, Kr 50-100 90I, Br 50-100 9040Cs, Rb 40-90 80Te(a) 5-25 15Ba, Sr 2-20 10Noble Metals (b) 1-10 3Rare Earths(c) .01-1 0.3Zr, Nb .01-1 0.3(a) Includes Se, Sb(b) Includes Ru, Rh, Pd, Mo, Tc(c) Includes Y, La, Ce, Pr, Nd, Pm, Sm, Eu, Np, PuTABLE VII 1-4 VAPORIZATION RELEASE COMPONENT VALUES (a)Fission Product Release, PercentXe, Kr 100I, Br 100aCs, Rb 100Te, Se, Sb 100Ru, Rh, Pd, Mo, Tc 5 (c)Refractory Oxides(b) 1 (c)(a) Releases for the amount that remains after the gap and meltdown releaseshave occurred. The rate is approximated by an exponential function with ahalf-time of 30 min although this value is considered uncertain by an orderof magnitude.(b) Includes Sr, Ba, Y, La, Ce, Nd, Pr, Eu, Pm, Sm, Np, Pu..(c) Values can be higher or lower by a factor of 5.

TABLE VII 1-5 FISSION PRODUCT RELEASES DURING STEAM EXPLOSIONSFission ProductRelease From Oxidation, PercentRangeBest ValueXe,Kr80-100 90I, Br0-100 90Te,Se (Sb)40-80 60Ru (Mo, Tc, Pd, Rh)80-100 90TABLE VII 1-6 FISSION PRODUCT RELEASE SOURCE SUMMARY-BEST ESTIMATE TOTAL CORERELEASE FRACTIONSFission Gap Release Meltdown Release Vaporization Release Steam ExplosionProduct Fraction Fraction Fraction(d) Fraction(e)Xe, Kr 0.030 0.870 0.100 (X) (Y) 0.90I, Br 0.017 0.883 0.100 (X) (Y) 0.90Cs, Rb 0.050 0.760 0.190 --Te(a) 0.0001 0.150 0.850 (X) (Y) (0.60)Sr, Ba 0.000001 0.100 0.010 --Ru(b) -- 0.030 0.050 (X) (Y) (0.90)La(c) -- 0.003 0.010 --(a) Includes Se, Sb(b) Includes Mo, Pd, Rh, Tc(c) Includes Nd, Eu, Y, Ce, Pr, Pm, Sm, Np, Pu, Zr, Nb(d) Exponential loss over 2 hours with halftime of 30 minutes. If a steamexplosion occurs prior to this, only the core fraction not involved in thesteam explosion can experience vaporization.(e) X = Fraction of-co e involved in the steam explosion. Y = Fraction of inventoryremaining for release by oxidation.Table VII 1-1 - Table VII 1-6VII-13/14

Section 2Fission Product Release FromThe Primary Coolant Systemonce fission products escape the coreregion, transport through portions ofthe primary coolant system is requiredto reach the containment atmosphere.Plateout or deposition on internal surfacesin the primary system couldpotentially limit the amount of activitythat leaves the system. The retentionof fission products would depend uponbulk gas flow rates (mass transferrates), temperatures, materials properties(fission products and structuralmaterials), and geometry, all of whichvary with time and position in thesystem. Comprehensive analysis of thisproblem, if all the needed data wereavailable, would require a complexmathematical model and would generateinvolved sets of fission product primarysystem release rates as a function oftime. It was determined that such anapproach was not feasible for thisstudy. Instead, it was decided to usegeneralized bounding calculations offission product behavior to developsimple retention factors for the primarysystem transport step. These factorsare analogous to the gap escape fractionsdescribed earlier and can betermed primary system escape fractions.General calculations relating to fissionproduct retention in the primary systemare described in Appendix I for iodineand in Appendix H for the solid fissionproducts. These analyses concentrate onthe upper region of the pressure vesseland on bulk steam flow conditions whichwould be characteristic of water boiloffduring core meltdown. Therefore, theresults are applicable to all PWR andsome BWR accident sequences. Conclusionsreached from the analyses regardingprimary system escape fractions aresummarized and discussed in the nextsubsection. This is then followed by anassessment of several special caseswhich can be encountered in some accidentsequences where the analyses inAppendices H and I are not applicable.2.1 GENERAL PRIMARY SYSTEM ESCAPEFRACTIONS2.1.1 NOBLE GASESXenon and krypton are gases at roomtemperature and essentially inert chemically.These fission products would becarried directly out of the primarysystem by the bulk gas flow. Therefore,an escape fraction of unity should beused.2.1.2 IODINEThe analysis presented in Appendix Iutilizes experimental deposition data.It indicates that very limited deposition(at most a few percent) would occurin the upper region of either a PWR or aBWR pressure vessel, whether the iodineis assumed to be elemental or HI. Thisamount of retention is insignificantcompared to that which would be carriedout by the bulk gas flow so an escapefraction of unity should be used.2.1.3 SOLID FISSION PRODUCTSThe bounding calculations described inAppendix H indicate that fission products,having normal boiling pointsbelow roughly 1500 F, should experienceonly temporary retention on internalpressure vessel surfaces during a reactorcore meltdown. The slight timedelay in transport out of the vessel isexpected to have no appreciable effecton the subsequent fate of these speciesin the containment volume. The fissionproducts that would generally exhibitthis behavior include the alkali metals(cesium and rubidium) the telluriumgroup (tellurium, selenium, and probablyantimony), and the alkaline earths(strontium and barium). Therefore, anescape fraction of one should be usedfor these species.Conclusions formulated in Appendix Halso indicate that an escape fraction ofone should be used for the noble metals(ruthenium) and rare earth group(yttrium) because of the potential forparticle formation. If ruthenium releasedfrom the core region shouldconvert to the volatile oxide duringtransport in the pressure vessel, thenan escape fraction of unity is even morecertain.It is important to note that the coremeltdown release values given in TableVII 1-3 being used for the solid fissionproducts are based to a large extent onexperimental melting studies. Theexperimental data are usually indicativeof release from a high temperatureregion of the apparatus rather than fromthe peak (melting) temperature location.Since the pressure vessel should reachcomparable temperatures during and fol-VII-15

lowing core meltdown, one would, byanalogy, expect complete escape of themeltdown release fractions. The calculationspertaining to plateout inAppendix H tend to support this conclusionand offer some insight concerningthe reasons and processes that could beinvolved.2.2 SPECIAL CASES OF PRIMARY SYSTEMRETENTION OF FISSION PRODUCTSThree special situations were identifiedin which the use of unit escape fractionsfor released fission products inthe primary coolant system should perhapsbe modified. Modifications weremade in two of these cases but not inthe third.2.2.1 PWR COLD LEG PIPE BREAKFor a cold leg break in a PWR, steam orsteam-hydrogen flow from the core regionwill pass through the upper plenum ofthe pressure vessel just as for a hotleg break. However, instead of exitingdirectly to the containment vessel, thegas and fission product vapors must thentravel through the steam generator sectionin order to reach the containment.The large surface area of the steamgenerator tubes and the relatively lowtemperatures (initially about 500 F)would present favorable conditions forplateout of condensable species. Noretention of the noble gases would occurand the analysis in Appendix I indicatesthat very limited deposition of fissionproduct iodine would be expected.Therefore, an escape fraction of unityfor these two species is still valid.The solid fission products should experiencehigh retention during the earlyphases of core meltdown. Exceptionscould result from penetration of particulatescontaining low volatility fissionproducts. Otherwise, plateout would beexpected to occur at the head end of thesteam generator tubes. The plateoutarea would heat-up due to heat transferfrom the incoming bulk gas and absorptionof decay energy from the condensedfission products. Fission prodtct vaporizationfollowed by re-condensationfarther along the tubes could occurrepetitively and smear the activity out.The eventual loss of bulk gas flow mightleave this fission product materialdeposited in the steam generator. Onthe other hand, steam generator tubetemperatures may rise sufficiently tocause significant fission product escapebefore loss of bulk gas flow.This complex heat and mass transferproblem was not analyzed. Instead, itwas decided to treat both types of pipebreaks identically with respect toprimary system solid fission productretention; i.e., unit escape fractionswere used regardless of break location.This represents worst conditions forrelease of radioactvity to the containmentspace, an assumption which appearsvalid for hot leg breaks but possiblyoverpessimistic for cold leg breaks.2.2.2 BWR CORE MELTDOWN WITH ECCINJECTIONFor a BWR loss of coolant accidentresulting from a recirculation linebreak, it is conceivable that a situationdevelops in which abnormalconditions in the reactor core preventeffective cooling even though normallyadequate emergency coolant pump flow isachieved. Thus, core melting and metalwaterreaction occurs, but either theaffected region is surrounded by ECCwater (due to core spray and lowpressure injection flow) or the openingin the pressure vessel is submergedunder several feet of water (due toflooding to the top of the jet pumps).This presents a unique environment forthe fission products that are liberatedduring fuel melting.In either condition the released fissionproducts, carried by the bulk steamhydrogenmixture, must pass throughseveral feet of water in order to escapefrom the pressure vessel. The steamshould be condensed but the remaininghydrogen will provide the driving forcefor carrying fission product activityout to the primary containment. Theamount of fission product activity whichdoes escape will depend on the trappingcapability of the water and the degreeof deposition on upper vessel structures.Deposition may only delay theescape of activity as vessel temperatureseventually rise, and fissionproducts trapped in the ECC water couldcirculate within the pressure vessel,sometimes encounter the melting coreregion, and undergo vaporization again.Consequently, the determination of aprimary system retention factor is notas straight-forward as it would firstappear.Recognizing that the core meltdown fissionproduct release fractions definedearlier contain an implied primary systemplateout factor, the escape fractionfor this flooded meltdown situationshould consider only the retention effectof the ECC water. Limited work onfission product cleanup in steam suppressionsystems (Ref. 1) offers somebasis for estimating this effect. The4VII-16

data indicate decontamination factors(the reciprocal of escape fraction)ranging from about 101 to 103 for iodineand particle removal from steam-airmixtures passed through a water lute.While the geometry and temperatureconditions are different than indicatedfor the present ptoblem, the fundamentalprocess in both cases is the removal oftrace species from a condensablenoncondensablegas mixture duringpassage through a water column. Becauseof the more severe conditions, the possibilityof ineffective removal must beconsidered.Therefore, escape fractions extendingfrom one (no retention) down to 0.001(high retention) appear feasible. Avalue of 0.1 should be used forrealistic accident calculations. 12.2.3 BWR DRY VESSEL MELTDOWNA second BWR accident situation thatleads to unique conditions for escape ofreleased fission products from the pressurevessel is a dry core heatupstarting with the end-of-blowdown conditions.In this accident there is nowater in the pressure vessel (no ECCdelivery) and consequently no boiloffsteam flow occurs to carry fissionproducts out of the vessel. Therefore,the only gas flow driving fissionproducts out of the vessel is due tosimple gas expansion as internal temperaturesincrease from decay heating.(The fission product noble gas inventory,about 4.5 lb-moles, would notcontribute significantly, to the totalgas content of the vessel.)Thermal analyses (Ref. 2), for dry heatupin a typical BWR indicate steam flowvelocities in the steam separators ofless than a foot per minute. Thus,fission products escaping from the coreshould spend at least on the order ofhalf an hour in the steam separator anddryer region of the pressure vessel. Asthe fission products plate out and decay,much of the fission product decayheat would be absorbed in the structuralmaterial above the core during this timeperiod. Table VII 2-1 lists averagetemperatures of the structural materialabove the core for different accidenttimes. These temperatures were calculatedassuming that all of the decay1 For noble gases a value of unity shouldbe used.heat lost from the core due to fissionproduct release is absorbed in the materialabove the core. The weight of thematerial was 229,000 lb. The resultsindicate that the decay heat of thereleased fission product is insufficientto completely melt all the steel structureabove the core at an accident timeof one hour. However, since core meltdownhas been estimated to require abouttwo hours, it appears that pressurevessel temperatures of about 2800 Fwould be reached sometime during thelatter half of the meltdown period.Therefore, from the beginning to the endof core melting, the absolute temperatureinside the pressure vessel shouldapproximately triple, producing a correspondingexpansion of the containedgases. Accordingly, when the meltdownperiod is completed about two-thirds ofthe original gas will have expanded outof the vessel. Ignoring retention oninternal surfaces because of the hightemperatures, this fraction can be usedto approximate the escape of fissionproducts released from the core duringmeltdown. Thus, at the end of coremeltdown in accident sequences of thistype, 2/3 of the fission product releaseterm is assumed to have escaped thepressure vessel, and the other 1/3 isassumed to remain as vapors inside thevessel.2.3 SUMMARY OF PRIMARY SYSTEMESCAPE FRACTIONS2.3.1 PWR SYSTEMSAn escape fraction offission products is usedtions of PWR accidentspipe break location.2.3.2 BWR SYSTEMSone for allin all calcularegardlessofIn BWR accident sequences where waterboiloff after ECC interruption occurs,an escape fraction of one is used forall fission products.In BWR accident sequences where ECC flowoccurs but with core meltdown, an escapefraction of one is used for noble gasfission products but a value of 0.1 isused for all other fission products.In BWR accident sequences where no ECCdelivery ever occurs, it is assumed thatat the end of core meltdown, 2/3 of allfission products that have been releasedwill have escaped the pressure vessel.VII-17

References1. Diffey, H. R., C. H. Rummary, M. J. S. Smith and R. A. Stinchcombe, "IodineCleanup in a Steam Suppression System", British Report AERE R-4882 (1965).2. Carbiener, W. A.,1400,et alAppendix"Physical VIII, ProcessesOctober, in Reactor1975.Meltdown Accidents", WASH-VII-18

TABLE VII 2-1 ESTIMATED AVERAGE TEMPERATURES OF UPPER STRUCTURAL MATERIAL IN A BWRPRESSURE VESSEL DURING DRY MELTDOWNAccident Time(minutes)Average Temperature,F0 55020 80040 120060 168080 224095 2750Table VII 2-1VII-19/20

Section 3Fission Product Leakage FromThe Containment SystemAnalysis of fission product behavior inthe reactor containment system is thefinal step in specifying the accidentsource term. The amount of radioactivematerial which escapes this barrierconstitutes the source for dispersion inthe natural environment. Fission productsinjected into the containmentspace by one of the four releaseprocesses will undergo removal from theinternal atmosphere by a combination ofmechanisms. The exact combination wouldvary among different plant types andaccident-sequence conditions, but commonto all is the removal that would occuras a result of natural transport anddeposition processes. Various engineeredsafety systems can operate to removefission products from containment atmospheres;i.e., aqueous spray systems andrecirculating filter units in PWRplants, and suppression pool scrubbingand once-through filter units in BWRplants. The size of the fission productsource which escapes to the environmentcritically depends upon how effectivelythe removal processes within containmentcompete with leakage from the containment.Analysis of this problem ofcompeting rate processes required developmentof rate models.3.1 PWR CONTAINMENT MODELSModels developed for PWR accident sequenceanalyses were of two types; (1) asingle-volume hand calculation model,and (2) a multicompartment computercoded model. The single volume modelwas used for preliminary hand calculationsof PWR accident sequences usingsimplified constant rate coefficientvalues. The multicompartment model wasdeveloped and used for all detailedaccident sequence calculations, incorporatingvariable rate coefficients,multivolume capabilities, and timevariable fission product input andcontainment leakage quantities. The twomodels are described below.3.1.1 SINGLE VOLUME CONTAINMENT MODELThe single volume containment modelassumes that the vapor phase consists ofone well-mixed compartment. This assumptionenables one to write the followingsingle differential equation foreach fission product species:C.d-• = - (a X.d3-) C. -1a. C. + R. (t)1(VII 3-1)Initial Condition: Ci = Ci(t') at t = t'C. = airborne moles of component1 i13 = removal rate constantmechanism ja. = leak rate, fraction of1 volume/timeRi (t) = source term (moles/time).viatheThis equation is easily solved for constantXij, ai and Ri to getCi =ZiR. 1 + a.)Ri(Exij + ai)- Ci (t')] exp - (EXij + ai)(t-t')(VII 3-2)The escaped amount during the intervalt-t' is the integralOi = ft'tCiVci dt,where V is the volumeThusR0 V •.0. = (TX E + a)(-'x l-exp [- (Eij + ai) (t-t')]of the vessel.R. 1ij+ a.)r(VII 3-3)The two equations for Ci(t) and Qi(t-t')can be used to calculate airborne fractionsand leakage fractions for variousaccident sequences by hand. The valuesof Ri to be used for various species areVII-21

described in earlier sections of thisreport. The only release of fissionproducts encountered where Ri = constantfor some time period is the melt release.This simplifies most of thecomputations to using an initial conditionfor a time period where the Xij'sand ci's are considered to be constant.If the X's or a's change at some time,t', a new t' should be considered as thebeginning of a new time period ofconstant new X's or a's. The bases forthe various values of Aij are discussedin Appendix J and a list of these coefficientsis given in Table VII 3-1.Values of ai must be specified fromanalyses of containment response for theaccident sequence being calculated.3.1.2 MULTICOMPARTMENT CONTAINMENTMODELTo better simulate containment geometryand time variable removal rates, a modelwas developed to analyze the atmosphericsource from a set of compartments whoseairborne contents are well mixed and arealtered by intercompartmental flow inaddition to deposition rates, leakrates, spray removal rates, etc. Thissystem is described by a set ofequations of the formdCmr/dt =ZHTic(VII 3-4)V. = volume of compartment jJ6 = 1 if i = j6i =0 if i jNote that the flow terms assume uniformmixing within each compartment.For a given release type and materialtype, the equation set can be written inmatrix notationdC_ -d = HC dt (VII 3-5)Where C is the column vector of airbornerelease fractions in the respective compartmentsand H is the matrix of rateconstants governing the evolution. Thesolution of this differential equationfor the column vector C for constantcoefficients in the H matrix is givenexactly byorC(t) = eHt C(o)C(t) = C(t 0 ) + (t-t 0 ) H(VII 3-6)((t-t 0 )H_(t-t 0 )H /whereC(to0).(VII 3-7)= airborne fraction of initial1 release of material m inrelease type r containedcompartment im = (xm +ZGki/Vk) 6iJJ+ (Gji/Vj) (1-Em)X.mr removal coefficient from com-1 partment i by settling, leak,spray removal, and other processesnot involving flow toor from another compartmentG.. = volume flow rate from com-J1 partment j to compartment iinFast computer techniques for generatingthis solution have been developed by B.H. Duane (Ref. 1) and were utilizedhere. By calling the numerical integrationsubroutines developed by Duane ateach time step after calculating the Hmatrix, the accuracy limit becomes thatimposed by the assumption of constant Hmatrix elements within the time step.Numerical error in generating the solutionto the coupled set of equationswith constant coefficients can be madeon the order of 10-6 percent.To integrate the amount leaked withinthe time step, N additional fictitiouscompartments were defined whose functionis to accumulate the leaked materialfrom the N real compartments. The fractionsCi+N, i = 1, 2, "'N, are cumulativeleaks obtained from the solution ofdd- (Ci+N) = aiC.Emrji= filter removal fraction formaterial m from release rwhere ai is one part of Hii. The previouslydefined N x NH matrix wasVII-22

augmented to form a 2N x 2N H matrixaccording toHi+N,j = 'k 6iji, j+N 0Hi+N, j+N 0For i = 1, 2, -, Nj = 1, 2, .- , NIf one takes the material quantity inthe fictitious compartments to be zeroat the start of a time step, the fractionof the initial release materialleaked during the time step will be2NCZ.i=NComputer code CORRAL was written tosolve the set of equations and at thesame time compute each rate parameter asa function of time and/or as a functionof vessel conditions (p, T, humidity).A more detailed description of the variousmodels used to compute rate parametersfollows.The first four parameters enable one tocompute transport coefficients involvedin depletion rate coefficient calculations.The bulk gas viscosities of steam-airmixtures (Nm) are computed according tothe following equations (Ref. 2):•m-whereI 1 AYS1+ - AsYa+IS+ YAYs3•SAPm = viscosity of mixtureTR1 0 .76811 A = 0.0414 L492J. , lb/ft/hr0.003339 (T, R)1.5Ps = 7(TR + 1224.2)v_ = mole fraction of airYS= mole fraction of steam(VII 3-8)Nine time dependent parameters are computedfrom input data at the time ofsolution of the differential equations.These are:1. Compartment pressure, p(k), psig•As = [2 ý2[l + (v)A/]1/AI/22. Compartment temp, T(k), F3. Compartment vapor mole fraction,VAP (k)4. Compartment wall-bulk temperaturedifference, DELTA T(k), F5.-6. Leak rates of molecular iodineand particulates from each compartment,ELI2(k), ELP(k), fractions/hr7. Flow rates between compartments,G(j, k), ft3 /hr8.-9. Filter decontamination efficienciesfor flow between compartments,EFI2(J,K), EFP (j, k),dimensionless.=sA above with subscripts reversed.The diffusivity of 12 in the steam-airmixtures was found using data and equationsfrom Knudsen (Ref. 2).whereD - 112 YA YsDADs(VII 3-9)DA = 2.03 E-05 (T, K) 15/(P, atm)/A WA, cm 2 /secDB = 3.24 E-05 (T, K) 15/(P, atm)/WSWA = 0.7075 + 141.73/T, KWS = 0.7075 + 454.72/T, K.iContainment Of Radionuclides ReleasedAfter LOCA.The diffusivity of 12 in water (spraydrops) was computed using the standardWilke-Chang relationship (Ref. 2), whereVII-23

D£=(7.4 x 10-8) (xM) 1/2 T(K) cm2/sec, 2£ = length of the wallwhereL(VII 3-10)x = degree of solvent association= 2.6 for H 2 0Mt = molecular weight of solventPL = solvent viscosity, cpPL = 100/{2.1484 [(T, (K) -281.6) + (8078.4+ (T,(K)-28l.6) 2 )0.5 ]-120} for H 2 0V = molar volume of diffusingsubstance = 71.5 cm 3 /gmolefor 12.3.1.2.1 Natural Deposition.The mechanism of natural deposition of12 is governed by diffusion with naturalconvection generated by temperature differencesbetween bulk gas and the wall(DELTAT(k)). Knudsen and Hilliard (Ref.3) claim that a mass transfer analogycan be made with correlations predictingnatural convective heat transfer coefficients.In similar manner, the modeluses expressions for Sherwood numbersfor laminar and turbulent flow using athermal Grasholf number,k = mass transfer coefficient.cA combination of the two Sherwood numbersis used to compute the actualSherwood number. Since the turbulencein most cases occurs around 10 ft fromthe top of the wall, a weighted averagewas used.NSh (overall)i-i0- 1£ N (turbulent)9, Sh+ -N +- NSh (laminar).To convert the mass transfer coefficientinto a deposition lambda (Aij) is a simplestep. Thus,A..(k) 13 = k c(k) A(k)/V(k)where A(k) and V(k) are surface area andvolume of compartment k, respectively.Another natural deposition process ofinterest is the settling of particulates.This involves a calculation ofterminal settling velocities, Vs assumingspherical, unit density particles.d2 (pp - PM) g/18 Vm = Vs (VII 3-13)Gr =(Twall -Tbulk)NM/Pm )2 TbulkThus for laminar flow (Gr

ficient. The expression for the spraylambda (s) is given by3F Eh2 V d' (VII 3-15)=--i- exp- kt ed(H + kg/k )(VII 3-14)whereF =d =V =E =spray flow ratespray fall heightspray diametercompartment volumeSpray collection efficiency.whereF = spray flow rateH = equilibrium distribution coefficient(Cg = C k/Hj atequil.)V = spray compartment volumed = spray drop diameterte = drop residence time - heightof fall/terminal velocityand the gas mass transfer coefficient,k , is given by Ranz and Marshall(Ref. 5) askg d {2.0 + 0.60 Re 1 / 2 Scl/JIn CORRAL empirical results from CSEdata are used to predict E. Apparentlythe efficiency is a function of a normalizedliquid volume sprayed (totalvolume sprayed/total compartment volume- Ft/V). Figure VII 3-1 shows the CSEdata, and the curve in this figure wasused to compute drop collection efficienciesin CORRAL. The diffusiophoresiswas subtracted from the efficiencyand the following expressions were fitto the remaining curve. To make theserelationships apply to a spray lambdaFtV0 - 0.002 E = - 15.825 (Ft/V) + .055.002 - .0193 E = .04125 - (.08626E+ 42.68 (Ft/V)] 1 / 2 /21.34.0193 E = .0015andk£ P292D3dD = 12 diffusivity in liquid.The latter is the Griffiths model discussedby Postma (Ref. 6). Incorporatingthe latter is a more conservativeapproach when kg/kk > 5 H. The terminalvelocities of the falling drops arefound by matching the velocity independentdimensionless numberfD Re2 = 4 pM (PL-PM) d 3 g/3jM 2with the appropriate range of Reynoldsnumber (Re). For spray drops the rangeof Re is 10-700. For 10 < Re < 100,fDRe 2 = 15.71 Re 1 - 4 1 7 and for 100 < Re< 700, fDRe 2 = 6.477 Re 1 - 6 0 9 (Ref. 7).Spray lambdas for removal of particlesfollow the equation (Ref. 8):with multiple sprays being used at varioustimes, the quantity Ft/V is now thesum I Fi ti/V for any one release ofparticles. Each release of particleswould have its own spray aging relationship,and at this time no simple meansof tying sequential release togetherinto one relationship seems possible.Only when particle size distributionsare known throughout a spray aging processcan sequential releases be tiedtogether.It should be noted that in CORRAL, nospray cutoff is used at C(t)/C(O) = .02;spray aging is used in its place.3.1.2.3 12 Equilibrium.When airborne molecular iodine is depletedby either sprays or natural deposition,the depletion rate becomes independentof the two above mechanisms whenthe concentration falls below about 1percent of the initial value (a conservativeestimate, i.e., a lower value isless conservative (Ref. 8). At concentrationsbelow this level, an apparentequilibrium situation exists where theconcentrations in liquid and gas phasesare related by an equilibrium distribu-VII-25

tion constant, H = Ci/Cg. H is a functionof time (probably due to slowliquid phase chemical reaction) and hasbeen experimentally determined. Inprogram CORRAL it has been possible toincorporate H = H(t) when equilibriumconditions exist.To get the equilibrium described quantitatively,an equivalent lambda for depletionof gas phase 12 had to be developed,since the value of H increaseswith increasing time, the gas phase isbeing depleted as time goes on. To getthis equivalent lambda, we can write amass balance for 12. If Cgo is the initialairborne concentration, thenCg V C V +C VorC C VCg go C£V g +CgVg g -(VII 3-16)1 1C P + 1 H.•- +1CV _ _9 V9.(VII 3-17)Then for H = H(t), we can write themoval rate of 12 byd Cg_ = _dt 2(Vk ) dHwhere the equivalent lambda isdCg __ 1Cg { -+ 1)gV 9VgdH\dtJdt = -Xdt.(VII 3-18)Data show that H VZ/Vg >> 1 for boricacid and caustic solutions in equilib-I2. so riumwith that1 dHH dt_Typical data for spraysTables VII 3-2 and VII 3-3.(VII 3-19)are shown in3.1.2.4 Intercompartmental Flow Rates.In two cases intercompartmental flowreratescan be calculated. If circulationfans move air throughout the containmentvessel or if boil-off occurs with steamevolution into one compartment, one canuse these flow rates to provide a basisfor estimating intercompartmental flows.In addition, natural convection drivenby wall-bulk gas temperature differencescan be a major contributor to flowrates. These rates can be estimated butwith a high degree of uncertainty. Inthis study, high flow rates have beenused to eliminate mixing as a parameter.Numerically, the high flow rate selectedwas equivalent to 10 PWR containmentvolumes per hour.3.1.2.5 Design of CORRAL for PWRAnalyses.A schematic diagram of the containmentgeometry used in CORRAL-PWR calculationsis given in Fig. VII 3-2. Four internalcompartments of a large PWR containmentvessel are modeled.a. The Primary Cubicle - The smallinternal compartment which housesthe piping, steam generator, andpump of one of the primary coolantloops, specifically the loop whichexperiences the large pipe break.b. The Main Volume - The large spaceabove the reactor cavity inside thepolar crane wall and including thehigh dome. This volume is sprayed.c. The Outer Annulus - The annular volumebetween the polar crane wall andthe containment vessel outer wall.d. The Lower Volume - This correspondsto the basement or the lowest levelinside the containment structure.The gas flow path between the four compartmentsis indicated on the diagram.The interchange between compartments 1and 2 (Qo) is determined by the primarysystem steam generation rate, while thecirculation flow between compartments 2,3, and 4 (QF) is set at a high value (asnoted in the previous section) to simulateefficient interchange between thesevolumes. The PWR contaminant is thusdivided into four well-mixed regionshaving different rates of airborne fissionproduct removal. The removal processesthat are included for each compartmentare listed below the diagramalong with the fission product injectionlocations. Note that all external leakagefrom the containment is assumed tooccur from the outer annulus. Due tothe well-mixed condition that is used,this affects only the amount of depositionthat is calculated to occur in thisVII-26

compartment and not the total inventorythat leaks. An exception to this leakagepath definition occurs when accidentsequence analysis specifies puff releasesfrom containment (i.e., suddenfailure by explosion or overpressure).In such cases the same volume percent ofthe airborne contents of all four compartmentsis included in the puff lossvalue.3.2 BWR MULTICOMPARTMENT MODELThe multicompartment feature of codeCORRAL was essential to estimating atmosphericsource calculations from a BWR.A BWR containment system is not a set ofopenly connected compartments like aPWR, where the whole containment systemcan be usually considered as "wellmixed". The compartments of a BWR areusually closed to one another and flowsbetween them occur in complex waysduring postulated accidents.;As many as six compartments are used insome BWR accident sequences. The firstfive are:a. The Drywell where natural depositioncan occur.b. The Wetwell where pool scrubbing landnatural deposition can occur.c. The Drywell annular gapiwhere naturaldeposition can occur;are made to estimate this temperaturedifference.3.2.1 NATURAL DEPOSITION - EFFECT OFHEAT TRANSFER FROM BWR VESSELWALLSThe mass transfer coefficient for 12removal (see Appendix J) is proportionalto the temperature-difference,Twall-T bulk = ATwraised to the 1/4 or 1/3 power. For anorder of magnitude range in ATw, themass transfer coefficient change's byonly a factor of 2, a relatively insignificantchange. However, during rapidcooling of the drywell (depressurization),ATw can span more than an orderof magnitude for short durations. Duringrapid heating, the condensing heattransfer coefficient is large (h = 1502Btu/hr ft F) and a steady ATw israpidly reached. This is about 0.14 Fin the drywell. In cooling, the heattransfer coefficient from the steel(approximately one inch thick) wall islow (2-5 Btu/hr ft F), and can lagbehind the bulk gas temperature for sometime. Neglecting any temperature gradientin the steel, for a sudden stepchange in bulk gas temperature, ATs, thetemperature-difference, ATw, is given byd. The Reactor Building where naturaldeposition can occur.AT w= AT exp (-ht/. cp)e. The Standby Gas Treatment System (a (VII 3-20)series of filters).IThe sixth compartment used was a fictitiousdumping ground foF' ground levelatmospheric sources / when elevated(stack) sources occurred simultaneously.The schematic diagram of CORRAL for BWRaccidents is shown in Fig. VII 3-3./Natural deposition is the most commonremoval mechanism for fission products,although it is not necessarily the mosteffective mechanism. Natural depositionof particulates occurs on horizontalsurfaces in all large compartments justas in a PWR. Turbulent deposition ofparticulates and iodine in the drywellannular air gap are via different mechanismsto be discussed later. Naturaldeposition of iodine occurs on all surfaceswith the rate controlled by naturalconvection as discussed for the PWRmodel. However, since the wall-bulk gastemperature difference that drives thenatural convection is highly variable ina BWR, transient heat transfer analyseswhereh = heat transfer coefficientt = time after ATsk = wall thicknessp = wall mass densityCp = wall heat capacitymass)If the temperature change isi.e., linear with respect totemperature-difference, ATw,given byATw = ((per unitgradual,time, theis now£cpc/h) P [exp - (ht/kpc) -1](VII 3-21)VII-27

where $ is the linear bulk-gas coolingrate, F/hr. Equations VII 3-20 and VII3-21 are used to compute ATw for varioustime intervals during cooling in thedrywell and wetwell. Values of h dependon gas velocities in the above compartmentsas well as the drywell annular airgap.Little information could be readily obtainedto estimate ATw's in the mainreactor building. These would be highlydependent on positions in the buildingand the outside environment temperatures,as well as gas flow parametersfrom the reactor. To allow for someminimum natural deposition in the mainbuilding, ATw = 0.1 F was used. Thiswould result in a natural depositionrate of X = 0.5 hr-I in the main reactorbuilding for 12. This A is five timesthe gas displacement rate for the buildingunder normal conditions (2,000 cfmthrough the Gas Treatment System).3.2.2 POOL SCRUBBINGIn a number of BWR accident sequences,gas flow occurs through the vent linesfrom the drywell to the wetwell waterpool. The water pool occupies slightlyover one-half the toroidal volume of thewetwell and is approximately 17-feetdeep. Pool scrubbing is a major decontaminationmechanism. Some data existson pool scrubbing, but comprehensiveexperimental studies on pool scrubbingin a BWR wetwell pool that includeinvestigations of all parameters (a widerange of particle sizes, steam quality,pool temperatures, flow rates, 12 concentrations,downcomer L/D ratios,etc...) have not been reported. Applyingsuch data, if available, would nothave refined atmospheric source estimatesgreatly because the available datashow that scrubbing is fairly effectiveon the fission products entering thepool. Also, trial calculations showedthat often 30 percent or more of theavailable fission products in any sequencewould escape the primary containmentby other paths (such as through thedrywell annular gap directly to thesecondary containment system).The best available data appear in a paperby Diffey, Rumary, et al. (Ref. 10),where 12, methyl iodide, and .06 V particlesin a steam-air mixture were poolscrubbed. The typical decontaminationfactors were 100 for 12, 2 for CH 3 I, and50-100 for the .06 p particles with 90percent steam-air mixtures (higher steamfractions give better decontamination).For CORRAL-BWR cases, the values usedare 100 for both 12 and particles, and1.0 for CH 3 I. A decontamination factorof 1.0 is also used for the noble gasfission products.The frequent result of higher than 90percent steam partially justifies theuse of DF = 100 for particles. Eventhough the BWR accident particles areassumed to be 5-15 p, or much moremassive than those studied above, thescrubbing of particles is largely due todiffusiophoresis (condensing steam onthe bubble wall carries particles), andtherefore largely independent of particlesize. However, larger particleswould have more boundary layer penetrationinertia in a rapidly circulatingbubble and this should then furtherjustify the choice of DF = 100 for particles,rather than the lower DF = 50.3.2.3 STANDBY GAS TREATMENT SYSTEM -(SGTS)The SGTS in a typical BWR is a set oftwo parallel filter trains upstream fromthree exhaust fans that releases filteredsecondary containment building air atan elevated level via a stack. However,under accident conditions only one ofthe filter trains would normally beused; the other being held in reserve.The SGTS keeps the building at subatmosphericpressures to minimize groundlevel leaks.Under normal conditions, the flow ratethrough the system is about 2,000 cfmbut the exhaust fans are capable of10,000 cfm. Sources in excess of thismaximum would create positive buildingpressures and produce a ground level,unfiltered source. The filter trains,according to plant Technical Specifications,are routinely tested to insurethe following removal efficiencies atall flow rates up to the 10,000 cfmmaximum:0099% for99% for99% forparticulateselemental iodineorganic iodideThese specifications are derived essentiallyfrom in-place testing criteriapublished by the USAEC in RegulatoryGuide 1.52. The criteria incorporatelimits on bypass flow for the filter andadsorber sections. It is recognizedthat substantially better performancemay be realized for the filter systemduring actual use. This is because thespecifications implicitly anticipatesome deterioration in system effectivenessbetween testing steps which may ormay not occur. However, since filteredleakage will be a minor contributor tooverall accident consequences for meltdownsequences in the BWR, the efficien-VII-28

cies given above were used in all CORRALcalculations. In addition, CORRAL isprogrammed to identify the filter andadsorber activity loadings to discoverpossible overheating conditions, atwhich time the filtration efficiencieswould decrease significantly.The overheating criteria for both theHEPA filters and for the charcoal bedswere based upon typical thermal performancedata obtained from Safety AnalysisReports and the design temperaturelimits for the components; 250 F at 2000cfm for the HEPA filters and 640 F at2000 cfm for the charcoal filters. Theheating rate of the filters (QF) wasassumed proportional to the temperaturedifference between the filter materialand the flowing inlet air.QF a (TFilter - TAir, inlet)(VII 3-22)For the maximum heating rate TFilter =TDesign. The inlet air temperature, bydefinition, isifAT AirTAir,inlet = Air 2(VII 3-23)The change in air temperature is givenby a heat balance,whereATAir -W=Cp=PQFP wCP (VII 3-24)air densityair flow rateair specific heatReference thermal performance data canbe used to obtain an average air temperaturefrom,QF a (TFilter - TAir)(VII 3-25)Using Equations (VII 3-22) through (VII3-25) and the reference data given inTable VII 3-4, the maximum heat loadsfor the HEPA and charcoal filters wereobtained for an air flow rate of 2000cfm. The results are also given inTable VII 3-4.In CORRAL calculations the fraction ofthe core fission product inventory whichis captured by the filter components(all particulate species on the HEPAfilter and elemental and organic iodineon the charcoal filter) is continuouslyrecorded. Data from inventory computationsof fission product decay energyemission rates are used to convert thefractions to heat loads. If the calculatedheat loads equal or exceed thedesign limits the filters are assumed tofail. At this point, the activity whichhas been sorbed by either filter type isassumed to remain fixed, but the filteringefficiency for any further activitywhich passes through the system is setequal to zero. This procedure is basedon the conclusion that the filter mediawill experience a physical degradationdue to the elevated temperature ratherthan chemical combustion. For HEPA'sthe degradation consists of deteriationof fiber binder materials and sealermaterials such that structural damage tothe filter media and bypass flow canoccur. For the charcoal beds it isassumed that the water dousing systemoperates, so that combustion is prevented,but the waterlogged beds will developchannels resulting in bypass flow.The use of zero filter efficienciesunder these conditions is probablyoverly pessimistic but conditions aretoo uncertain to estimate another valuewith reasonable confidence.3.2.4 NATURAL DEPOSITION IN THE DRYWELLANNULAR AIRSPACEThe drywell shell is surrounded by atwo-inch air gap between the steel shelland the concrete shield containing theshell. Under normal leakage, isolationloss leakage, or under certain primarycontainment failure conditions, gas flowfrom the primary containment is assumedto pass through the space and exit atthe operating floor of the reactorbuilding. Therefore, any fission producttransport and deposition whichoccurs in the region of secondary containmentbelow the operating floor areapproximated in CORRAL by behavior inthis annular air space. For cases inwhich leakage occurs directly from thedrywell shell, the wetwell torus, or theconnecting vent pipes this provides anaccurate description of the flow path tothe operating floor level. For cases inwhich primary containment leakage occursin one of the many subcompartments inthe lower region of the secondary containmentbuilding, the model constitutesa simplified approximation of the geometryand flow paths between the variousleakage locations and the operatingfloor level. However, the methodVII-29

described below for calculating fissionproduct transport and deposition in theannular region, is expected to produceoverestimates rather than underestimatesof fission product concentrationsreaching the refueling building underthis alternate leakage path condition.This is because the combination oflonger residence time with deposition inthe lower regions of secondary containmentwould usually be more effectivethan the decontamination factors predictedfor the annulus. Only under highflow conditions should decontaminationin the annulus predict concentrationslower than might be obtained for theother leak path.The annular region cannot be modeledlike the well-mixed compartments withdeposition on the walls and/or floor.It is better described as plug flowalong a cylindrical annulus with masstransfer to the walls. A simple firstorder differential equation defines amass transfer coefficient, k, that canbe estimated from known correlations for12 transfer. For particles, k can beestimated from particle deposition datafrom moving gas streams with more difficultyand uncertainty. The differentialequation for transfer to the walls of anannular slit is:whereTwC = - ( _j)dx, (VII 3-26)C = the concentration of thetransferring substanceC w= wall concentrationU = plug average annual velocityAr = annulus widthx = axial distancek = mass transfer coefficient.If Cw = 0, Equation (VII 3-26) integratesfor 0< x

Rh = Hydraulic radiusflow cross sectional areawetted parameter= Ar/2 for the annulus.The transition region, 2100

(f) Compartment filter decontaminationrates.(g)Fractions of compartmentsreleased during a puff release.2. Variables (with time)(a) Pressure, temperature, andwater vapor content of eachcompartment. Temperaturedifference between bulk gasand walls.(b)(c)(d)Flow rates between compartments.Decontamination factors betweencompartments.Particle sizes.(e) Leak rates to atmosphereand leak DF's (decontaminationfactors).b. Initial Conditions1. All concentrations set equal tozero at t = 0 except gas releaseconcentrations in first compartment.2. Spray set to operate in maincompartment (PWR).3. All amounts released and DRF's(dose reduction factors) setequal to zero. A zero DRF meansthat nothing has been released.c. Computation of PropertiesRemoval Ratesand7. Spray lambdas for particle.8. Terminal spray velocities, gasphase mass transfer coefficient,liquid phase mass transfercoefficient, and spraylambdas.9. 12 equilibrium equivalent lambdas(if needed).10. Overall lambdas.d. Solution of Differential EquationsThe solution of the differentialequations is discussed in the previoussection. To properly age thecontinuous releases, it was necessaryto divide these releases intodiscrete impulses releases. Themelt release was divided into tenequally spaced and sized releases,each independent age wise from theother nine. The vaporization release(released at an exponentiallydecaying rate) was divided into 20impulse releases, each successiverelease at an exponentially lowervalue than the first. The sum ofthe first ten releases equals 1/2the total release, and the remainingten equals the remaining 1/2. Theduration of the period of the firstten is one half life. The durationof the second ten should be threehalf lives for a reasonable approximationof an exponential decay.Thus the total number of differentialequations solved (one each forparticulates, organic iodides and12) for any time step is (N=numberof compartments)1. Pressure, temperature, and watervapor content and T(bulk)-T(wall) by parabolic interpolation.2. Intercompartmental flow ratesand decontamination factors andleak rates and respective DF'sby parabolic interpolation.GAP RELEASEEXPLOSIONRELEASEMELT RELEASE3N equations3N30N3. Particle sizes by linear interpolation.4. Gas phase viscosities and 12diffusivities and Schmidt numbers.5. Mass transfer Grashof numbersand corresponding depositionrates.6. Particles settling velocitiesand their natural deposition.VAPORIZATIONRELEASE60N96N equationsThe accuracy of the output dependson the rate of change of the ratecoefficients, so short time stepswould be desirable immediately aftereach discrete release (aging israpid at first, especially if spraysare on). Long time steps are sufficientfor old releases.VII-32

e. Output Variables1. Airborne contained fractionsrelased at time, t.(a) For each release: 12, organiciodides, particulates.(b) For each compartment foreach release: 12, organiciodides, particulates, attime, t.2. Escaped fractions released (foreach release: 12, organic iodides,particulates, at time,t).3. Escape fractions of the core forany desired isotope.4. Dose reduction factor for eachrelease (12 and particulates) attime, t.5. Overall dose reduction factor(12 and particulates) at time,t.6. Total fraction of core iodineescaped and core particulatesescaped up to time, t.3.3 ESTIMATION OF METHYL IODIDEFRACTION USED IN PWR AND BWRANALYSES WITH THE CORRAL CODEConsiderable experimental evidence existsthat some relatively small portionof the fission product iodine which isreleased from the reactor fuel willappear in the containment volume asorganic iodide compounds. The dominantcompound, methyl iodide, because of itschemical properties would tend to remainin the gas phase rather than deposit onsurfaces, dissolve in water, or reactwith spray solutions. Trapping on charcoalfilters is also less effective fororganic iodides than for inorganic iodinespecies. Therefore, estimates ofthe amount of this difficult-to-removeform that could be produced in anaccident are necessary. This task ishampered by uncertainties in the mechanismof formation and limited data, butconversion values have been obtainedbased on the information that is currentlyavailable. The approach used andthe resulting values for each waterreactor type are described in the twofollowing subsections.3.3.1 METHYL IODIDE FRACTION FOR PWRANALYSESThe fraction of iodine airborne as methyliodide following postulated accidentcases has been estimated from the informationsummarized by Postma andZavadoski (Ref. 13). The total percentageformation was obtained by addingthat formed by radiolysis to that formedby nonradiolytic mechanisms.For nonradiolytic formation, the leastsquares fit of experimental measurementsindicates conversion of 0.056 percent ofiodine. Upper and lower limits indicatedby the measurements are 1 percentand 0.004 percent, respectively. Thesevalues were summarized asNonradiolytic PercentConversion = 0.1% +0.9%-0.1%Radiolytic formation process lead tocombination of elemental iodine andtrace level organic materials in the gasphase. On the basis of available experimentaldata, the maximum percentageconversion for a DBA case is 2.1 percent.The minimum formation due toradiolysis is near zero. Therefore, forno removal, we estimateRadiolytic Conversion For No RemovalCase = 1.1% + 1%.The no-removal case can never be realizedbecause natural processes willdeplete an appreciable fraction of theairborne iodine, making it unavailablefor radiolysis. For natural removal,the percentage conversion was reduced bya factor of 2. The result may be summarizedas follows:Radiolytic Conversion For NaturalRemoval Case = 0.6% + 0.5%.Sprays will remove iodine rapidly fromthe gas phase, further reducing thepotential for organic iodide formation.The effect of spray removal was handledby reducing the conversion by an additionalfactor of 2. The result is summarizedas follows:Radiolytic Conversion For Spray RemovalCase = 0.3% + 0.2%.Total formation of organic iodides willbe the sum of that for radiolysis andthat formed by nonradiolytic mechanisms.For the natural transport case the resultisTotal Conversion For Natural RemovalCase = 0.7% +1.4%-0.6%and for the spray case,VII-33

Total Conversion For Spray RemovalCase = 0.4%+1.1%-0.3%It should be noted that this approachdefines an instantaneous methyl iodidefraction existing near the beginning ofthe accident sequence when most of thecore iodine has been released. A moresophisticated approach would account fornonradiolytic formation as above, butwould account for gas phase radiolysisthrough use of a G value for formation.A continued formation of organic iodides,based on the instantaneous airborneiodine concentration, G value for formation,and radiation dose would be usedin place of the fractional conversionslisted for the two conditions. However,this was considered an unnecessary complicationbecause overall accident consequencesare only partly dependent oniodine releases to the environment.3.3.2 METHYL IODIDE FRACTION FOR BWRANALYSESIt appears there is not enough evidencethat methyl iodide formation in a BWRshould be significantly different thanin a PWR. Although radiolytic formationrates are higher in a BWR due to higherdose, radiolytic destruction is highertoo. The net rate is perhaps much thesame in both reactor types. Naturaldeposition lambdas are nearly identical.The driving force for natural convectionis the bulk gas - wall temperature difference(AT). For the condensing atmosphereof the accidental sequencesAT (PWR) ! 10 AT (BWR)•(BWR)F W = (AT(BWR))1I/ XT(PWR) /A(BWR)/V(BWR)kA(PWR)/V(PWR)!If we compare the drywell A/V with thePWR, then {A(BWR)/V(BWR) }/{A(PWR) /V (PWR) } 1 101/3, thereforeX(BWR)A X(PWR)Since the natural transport case is theonly condition requiring considerationin BWR analyses, a total conversionvalue for organic iodides of%+1-. 0.6%1%can occur during accident sequences foreither reactor system. These can beclassified as follows: (1) low (design)leakage from one or a collection ofsmall undefined paths, (2) isolationloss leakage from a single and relativelylarge open path, and (3) massiveleakage (a puff) from a failed containmentvessel. In PWR systems the keystructure with respect to atmosphericleakage is the steel-lined concrete containmentvessel, while in BWR systemsthe important structure is the concreteand metal panel reactor building, alsoknown as the secondary containment.During leakage from these structures itis possible that some of the fissionproduct vapors and aerosols, which arecarried by the bulk gas flow, woulddeposit on surfaces along the leakpassage. Leakage through a network ofrough-walled cracks would favor depositionwhile little deposition shouldoccur during flow through an orifice.In analyses performed throughout thisstudy no credit has been taken forfission product deposition along theleak path when the leak path leadsdirectly to the atmosphere. For largeleaks (classes (2) and (3) above) thisapproach is probably close to actualexperience. However, it is also consideredan acceptable, although conservative,approximation for small leaksbecause; (1) the generally undefinednature of small leak distributions inhibitsthe ability to predict accuratedecontamination factors, and (2) theaccident sequences which are analyzedinclude containment failure at somepoint in the accident scenario andfission-product leakages beyond thesepoints overwhelm the earlier low leakagevalues. In calculations the assumptionof no decontamination is represented byusing DF = 1.There is one containment leakage situationwhere the release of activity tothe atmosphere is modified by using a DFvalue greater than one. This caseoccurs only in PWR analyses where meltthroughof the concrete base mat isfollowed by a puff release of the containedgas until the internal pressureequalizes with the external pressure. 1Any gases and airborne fission products.A.will be used.3.3.3 FISSION PRODUCT DECONTAMINATIONDURING LEAKAGES TO THE ATMOSPHERESeveral different types of leakage fromreactor containment to the atmosphere1 Since the bottom steel containmentliner is embedded and anchored in concreteall around the reactor cavityarea, no significant gaseous releasefrom the base of containment is expecteduntil the mat is actually penetratedby the melt.VII-34

escaping in this manner would have topass through many feet of ground materialin order to reach the atmosphere.The base of large PWR containments typicallyextend 40 to 60 feet below gradeand part of this height is usuallysaturated with subsurface groundwater.The ground material may be low or highpermeability soils, but next to the containmentwall (perhaps within a cofferdam) the material is likely to consistof layers of gravel, sand, and/or porousconcrete backfill.This combination of external mediashould act as an effective filter forwater soluble and particulate contaminantsin the permeating gas stream.Work reported in Appendix K indicatesmany minutes or hours will be requiredfor active gas to penetrate to theground surface; that is, containmentpressure relief by a single huge puff isnot expected. In the water saturatedzone of this natural filter bed solublevapor species such as elemental iodineshould be absorbed. Even though partialchanneling may occur along fissurescaused by the pressure gradients, theinterfaces between the several layers ofdifferent materials should act as crackarrestors to produce meandering transportpaths. Aerosol particles should beefficiently removed during passagethrough the dry upper layers of thebackfill material. Work on the effectivenessof sand filters for removingaerosols from air streams (Ref. 14,15)indicate that efficiencies in the rangeof 99 percent to 99.99 percent are typicalfor one to three foot deep beds dependingupon gas flow velocity, particlesize, and degree of bed packing.Equivalent or better efficiencies shouldbe expected for the situation of interesthere because dry bed depths of 10 to20 feet are anticipated.Based on the above rationale it was concludedthat fission product escape tothe atmosphere via pressure relief fromthe base of the containment would becharacterized by a large DF value formost species. Preliminary calculationswere performed which indicated that anyDF greater than 1000 would produce atmost a ten percent decrease in the totalatmospheric source term for accidentsequences involving containment meltthrough;that is for DF>1000 the leakageprior to meltthrough accounts for nearlyall of the atmospheric release. Therefore,a conservative value of 1000 wasused in all such calculations. This DFapplies only to elemental iodine and theparticulate fission products. A DF = 1was always used for the fission productnoble gases and the organic iodidefraction. It is also emphasized thatthis procedure is followed only in PWRaccident sequence analyses. In BWR accidentsequences containment failure bysome other process always precedes themeltthrough pathway.3.3.4 POTENTIAL GROUNDWATER CONTAMINA-TION IN A MELTDOWN ACCIDENT 1Most of the reactor meltdown accidentsequences are considered to ultimatelyresult in melthrough of the bottom ofthe containment structure. This eventwill bring the molten mass into contactwith the natural soil system at estimateddepths of 50 to 100 feet. Radioactivecontamination of the local groundwatercan occur by several processesduring this event. In meltthrough accidentswhere sprays have operated, someportion of the spray liquid may bereleased to the soil-water system. Inaccidents where sprays have not operated,airborne activity in the containmentmay be carried into the groundwaterduring containment building depressurization.Finally, groundwater leachingof the core mass itself can provide adelayed but long-term contaminationsource. The radionuclides introduced byany of these processes can be transportedalong the groundwater route to locationsof potential interaction withhuman usage. It is of interest to thepresent study to obtain an estimate ofthe magnitude of this potential contaminationpathway.3.3.4.1 Problem Definition.The analysis assumes a 3200 MW(th) reactor,after operating for .550 days atfull power, experiences an accidentevent which results in core meltdownfollowed by meltthrough of the concretefloor of the containment structure asdescribed in Appendix I. It has beenestimated that containment meltthroughwould occur about one day after theaccident-initiating event (Ref.. 36).Therefore, contamination of-groundwateris unlikely before this time. Two basicmodes were defined for analysis; one toexamine the implications of early radioactivityrelease and the other to examinethe effect of delayed leaching ofradioactivity from the solidified coremass.1 The discussion presented here is based-on the analyses performed by A.E.Reisenaur and R. C. Routson of Battelle'sPacific Northwest Laboratories.VII-35

In the first mode, it is assumed thatcontainment meltthrough at one day isfollowed by rapid depressurization, andthe airborne radioactivity in containmentis delivered to the underlyingsoil-water system. This condition isindicative of a meltthrough accident inwhich containment sprays have notoperated. Since detailed accident sequencecalculations were performed inthe Reactor Safety Study for this typeof accident, core activity release fractiondata are available. The releasefractions range from a fraction of apercent of the core inventory for theless volatile species up to severalpercent of the core inventory for themore volatile species. Exact valuesused will be defined later in thissection. The above contamination sourceterm for the groundwater system isprobably excessive for those accidentsequences in which contaminated spraysolution might leak into the soil becausean equivalent radioactivity releasewould require leakage of severalthousand gallons of solution past themolten core material. It is more likelythe solution would vaporize back intothe containment atmosphere. Accordingly,the depressurization release shouldbe considered a worst case assumption.The analysis further assumes the airborneactivity which is so releaseddissolves rapidly and completely in thereceiving groundwater. This is probablyhighly conservative because much of theradioactivity will be in the form ofoxide aerosol particles which shouldexperience a relatively slow dissolutionrate. In addition, the particulatesource may have to penetrate a "driedout"soil zone around the molten mass inorder to reach the groundwater system.These two effects should delay the appearanceof contamination in thegroundwater and lead to lower peak concentrationsthan will be obtained underthe present groundrules.The second mode considers the effect oflong-term leaching of fission productsfrom the core-soil mass after it hascooled and solidified in the groundunderneath the containment floor. It isexpected that, during penetration intothe ground, the size of the melt willincrease as soil material is melted.Additional oxide and silicate phaseswill probably form which should mix withor partially dissolve in the core material.Fission products would be expectedto distribute among the several maoltenphases under the influence ofconvective transport and chemical reactiondriving forces. During this timethe molten material should be surroundedby a relatively dry soil zone so thatdirect contact with groundwater would beunlikely. Eventually the melt penetrationwould stop because of decreasingdecay-heat generation and the dilutingeffect of the molten soil constituents.Then the melt should gradually coolforming a solidified glass-like mass(Ref. 16). It is expected "-he masswould fracture during the coolingprocess and subsequent return of groundwaterto the surface could begin leachingout fission products. The dimensionsof the core-soil mass and the timedelay for leaching to begin are difticultto predict accurately. Boundingheat-transfer calculations have indicatedmelt sizes may increase to maximumradii of 30 to 50 feet in 1 or 2 yearsin dry soils (Ref. 18). The presence ofgroundwater would tend to lower the numbersbut results of the overall analysishere are not highly sensitive to exactdimensions. Elementary heat balancecalculations also indicate that at about1 year the heat removal capacity of theanticipated water flow (without boiling)would balance the decay-heat generationrate of the mass. The above resultsprovide some guidance for selectingsource dimensions and the starting timefor leaching.In both modes the source is modeled as acylinder 70 feet in diameter and extendingover the assumed 60 feet depth ofthe groundwater system. Radioactivityreleased from this boundary enters thegroundwater where it is transported bythe groundwater flow to a nearby outletwater body such as a river, lake, orbay. Therefore the rate of appearanceof radionuclides at the outlet willdepend on: (1) the release rate, (2)the groundwater flow velocity, and (3)the distribution coefficients for sorptionof the dissolved radionuclides onthe soil material of the groundwatersystem. In other words the soil-watersystem acts as a large chromatographiccolumn to selectively separate and delaythe discharge of the individual radionuclidesin aqueous solution.3.3.4.2 The Hydraulic and Ion TransportModels.The hydraulics model for the two modesdefined above was formulated as follows.An analytic solution to a point sourceand point sink in a uniform flow fieldwas expanded to include a fully penetratingcylindrical source of finiteradius. A diagram of the flow systemused is given in Fig. VIII 3-5. Thedata chosen for the slope of the groundwatersystem, distance to the outletwater body, and soil type were selectedfrom examining several eastern UnitedVII -36

States power reactor site reports andgenerally represent conditions whichwould cause the most rapid dispersion ofthe radionuclides. Table VII 3-6 showsthe input values used in the hydraulicsmodel calculations. The normal groundwaterflow was assumed to be perpendicularto the outlet water body.In performing calculations with themodel, the presence of the cylindricalsource was assumed to disturb the uniformflow system out to twice the sourceradius. Therefore, thecross section ofthe contaminated flow system was 140-feet wide by 60-feet deep. -The volumetricrate of fluid flow in this rectangularchannel was 11,200 ft3 /day, and thelinear flow velocity was 7 feet/day.Accordingly, the travel time for waterfrom the source to the outlet water bodywas 215 days or 0.59 year.The radionuclide transport calculationswere made using a model based on thetheoretical cell concept in which equilibrationbetween the liquid and solidphases is achieved before the fluid isconvected into the next cell (Ref. 17).The equilibrium stage approach has traditionallybeen considered satisfactoryfor groundwater systems because of thespeed of the chemical reactions in relationto the fluid flow rate.Sorption coefficients (Kd) for each isotopeof concern were determined in thefollowing fashiona. Kd's for Sr and Cs in a typical 1groundwater system were determinedfrom direct data.b. The Kd's for all other components inthe given solution were calculatedby assuming that the ratio of Kdx toKdcs or Kdsr reported in otherreferences is constant.Using the best available data, a conservativedistribution coefficient (Kd) waschosen for each specie of concern.These coefficients with appropriate referencesare listed in Table VII 3-7.3.3.4.3 Results of the DepressurizationRelease Problem.The radionuclides chosen for calculationin this mode were selected on the basisiGroundwater having 28 ppt total salinityand having a composition of 0.38 MNa+, 0.008 M K+, 0.008 M Ca++ and 0.04M Mg+.of their half-life, distribution coefficient,and inventory. The fraction ofthe reactor core inventory of each radioactivespecies released to the groundwatersystem was based on calculateddepressurization release fractions for acore meltthrough accident. The valuesused are given in Table VII 3-8. Inexecuting these calculations, the mixedcell transport model utilized 100 columnsegments, and the individual radionuclideinventories (corresponding to therelease fractions in Table VII 3-8) were"injected" to the groundwater flow channelover a two segment interval. Thistranslates to an injection period ofabout 4.3 days which may or may notcorrespond to actual periods. However,final results are not too sensitive tothis factor, particularly for specieswhich experience adsorption and delayduring transport to the outlet waterbody.Resulting activity release rates of selectedradionuclides contained in thegroundwater released to the water bodyhave been plotted and the curves areshown in Fig. VII 3-6. The semi-logplots show the release rates in curies/day as a function of time. Note theanalysis predicts breakthrough of thenon-sorbed Ru-106 at about 0.35 yearwith the release rate peaking at 0.59year (215 days). The indicated axialdispersion is an artifact of the mixedcell model and may overestimate actualconditions. However, less dispersionwould lead to higher peak release ratesbecause the activity balance of the radionuclidemust be maintained. In contrastto the non-sorbed ruthenium-106,the sorbed ruthenium-106 ions breakthrough only after about 7 years and theions leached from the soil have a muchlower and broader elution curve. Technicium99, with a low distribution coefficient(Kd = 0.1) breaks through insignificant concentrations in about 0.7years and peaks before one year.Strontium-90 which has an intermediatedistribution coefficient (Kd = 2) arrivesin significant quantities in 3 to4 years and peaks at about 6 years. TheAntimony-125 and the Cesium-137 withtheir higher Kd values lag behind theother nuclides arriving approximately 22and 30 years after release and reachingtheir peak concentrations at 35 and 51years, respectively.The other radionuclides investigatedwere significantly delayed allowing radioactivedecay to reduce them to innocuouslevels at the time of arrival inthe water body. In order to place theseresults in perspective with respect toradiological impact, the peak effluentVII-37

concentrations discharged from thegroundwater system for each of theradionuclides calculated may be comparedto the maximum permissible concentrationsspecified in 10CFR20 (Ref. 30) forwaters discharged to uncontrolled areas.This comparison along with other informationis given in Table VII 3-9. Itmay be noted that three of the radionuclideeffluent concentrations (Ru-106,Sr-90, and Cs-137) are predicted to bemuch higher than their respective MPCvalues. However, it must be emphasizedthat these are likely to be overestimatesand they do not include theeffects of several processes which wouldact to limit human exposure to thiscontamination source. These processeswill be discussed following presentationof the results for the glass leachingrelease mode.3.3.4.4 Results of the Glass LeachingRelease Problem.In accordance with estimates given earlier,the inventory of radionuclides inthe core mass one year after the initialevent was used to specify the leachingsource strength. The significant speciesand their inventories were determinedfrom the fuel irradiation historyand consideration of prior fission productreleases from the molten core. Allisotopes of the noble gases, halegens,alkali metals, and tellurium group elementswere assumed absent from the coresoilmass.1 From the remaining refractoryelements the following key specieswere selected; Tc-99, Sr-90, Ce-141, Zr-95, Eu-155, Pm-147, Eu-153, Sm-151, U-235, U-238, Pu-241, Pu-239, Pu-240, andPu-242. The relationship used to predictthe rate of leaching for all componentscan be expressed asThey = -5.222 +0 . 3 3 4 xy = logl 0 (fraction leachedfrom time 0 to time T)x = logl0 (elapsed time fromtime 0 to time T)Time is in hours, j is fixed at1 year and T is variable.relationship was derived from data"This assumption is consistent with therelease fractions presented earlier inthis report for the volatile fissionproducts.presented by Saidl and other (Ref.31,32,33). The melt was conservativelyassumed to be basalt glass since it isprobably much more soluble than any.other glass, and the characteristiccomponents were Sr and Cs which are themost leachable of the components ofinterest.The elution curves obtained for severalof the more important nuclides are presentedin Fig. VII 3-7. All the curvescharacteristically rise to a peak andthen slowly decay. The peak effluxrates and concentrations of the radionuclidesfor this case are generally afactor of 100 to. 1000 lower than for thecorresponding nuclides in the depressurizationrelease case. However, the elutioncontinues for extended periods oftime. Each of these features is adirect result of the very slow leachingprocess. In Table VII 3-10, the peakeffluent concentrations discharged fromthe groundwater system for each of theradionuclides calculated are compared tothe maximum permissible concentrationsspecified in 10CFR20 (Ref. 30) forwaters discharged to uncontrolled areas.It is apparent that except for Sr-90 allthe predicted peak effluent concentrationsare well below the MPC limits.3.3.4.5 Exposure Reduction Processesand Contamination ControlMeasures.It should be emphasized that the hydraulicmodel parameters, the radionuclidedistribution coefficients, and the radionuclideleaching rate used in theseanalyses were selected to produce overestimatesfor the rate of appearance ofthe radionuclide sources at the outletwater body. For example, the soilpermeability coefficient is indicativeof well-sorted sands with gravel and offissured limestone formations (Ref. 34),the distribution coefficients are probablylow by factors of 10 or 100 (Ref.35), and the leaching expression assumesa relatively highly soluble glass containingfissures which increase theeffective surface area by a factor of100 or more (Ref. 32). In addition,calculations of human radiation dosefrom use of the receiving water bodyresource would have to include thedilutioneffect that would occur in thewater body beyond the efflux point forthe contaminated groundwater. Thus, thegroundwater contamination problem atmany reactor sites is expected to beless severe than indicated here.Another important factor to consider inevaluating the above results is the significanttimes which are required forVII-38

movement of radionuclides through agroundwater system. Several months andin many cases years should elapse beforecontamination would appear in waterbodies used for the support of a significantpopulation group. This delaywould allow ample time for institutingmonitoring operations and for setting upan effective warning network. More importantly,the time would most likely beused to execute procedures for controllingor even eliminating the spread ofcontamination beyond the reactor site.This would involve drilling wells formonitoring and pumping purposes tocontrol the local groundwater flowgradient. The withdrawn water could bestored temporarily in surface tanks orin sealed holding ponds for subsequenttreatment. After movement of the radionuclidesis under control, it would seemfeasible if it were to be considerednecessary to form a vault-like barrieraround the radioactive zone using acombination of excavation, drilling, andconcrete injection operations.Even without the above engineered mitigatingactions, the basic conclusion ofthis analysis would not be changed.Specifically, the analysis has shown thehydrologic contamination problem occurson a much longer time scale than doesthe atmospheric contamination problemfor a core meltthrough accident. Therefore,warning actions "alone should besufficient to limit population radiationdoses from the hydrologic source to lowlevels compared to the doses received asa result of the atmospheric source.VII-39

References1. Reactor Physics Quarterly Report, April-June 1969, BNWL-1I50, Battelle-Northwest,Richland, Wash., (B. H. Duane, p 2.5-2.7), (1969).2. Knudsen, J. G., "Properties of Air-Steam Mixtures Containing Small Amounts ofIodine", BNWL-1326, Battelle-Northwest (1970).3. Knudsen, J. G., and R. K. Hilliard, "Fission Product Transport by Natural Processesin Containment Vessels", BNWL-943, Battelle-Northwest, Richland, Washington(1969).4. Hilliard, R. K. and L. F. Coleman, "Natural Transport Effects On Fission ProductBehavior in the Containment Systems Experiment", BNWL-1457, Battelle-Northwest,Richland, Washington (1970).5. Bird, R. B., Stewart, W. E., and E. N. Lightfoot, "Transport Phenomena", JohnWiley and Sons, N. Y. (1960).6. Postma, A. K. and W. F. Pasedag, "A Review of Mathematical Models for PredictingSpray Removal of Fission Products in Reactor Containment Vessels", BNWL-B-268,Battelle-Northwest, Richland, Washington (1973).7. Postma, A. K. and R. K. Hilliard, "Absorption of Methyl Iodine by Sodium ThiosulfateSprays", AND Trans. 12, p 898-899 (November, 1969).8. Hilliard, R. K., Postma, A. K., et al., "Removal of Iodine and Particles bySprays in the Containment System Experiment", Nuclear Technology, 10, p 499-519(1971).9. Postma, A. K., L. F. Coleman, and R. K. Hilliard, "Iodine Removal from ContainmentAtmospheres by Boric Acid Spray", Report No. BNP-100, Battelle-Northwest,Richland, Washington (197).10. Diffey, H. R., Fumary, C. H., M. J. S. Smith and R. A. Stinchcombe (AERE,Harwell, England), "Iodine Cleanup in a Steam Suppression System," CONF-650407(Vol. 2), Internation Symposium on Fission Product Release and Transport UnderAccident Conditions, Oak Ridge, Tennessee, p 776-804 (April, 1965).11. Rosenberg, H. S., Genco, J. M., and D. L. Morrison, "Fission-Product Depositionand Its Enhancement Under Reactor Accident Conditions: Deposition on Containment-SystemSurfaces", BMI-1865, Battelle-Columbus (May, 1969).12. Sehmel, G. A.,"Particle Eddy Diffusivities and Deposiition Velocities For IsothermalFlow and Smooth Surfaces", Aerosol Science 4, p 125-139 (1973).13. Postma, A. K., and R. W. Zavadoski, "Review of Organic Iodine Formation UnderAccident Conditions in Water-Cooled Reactors", WASH-1233, (October, 1972).14. Yoder, R. E., and Empson, "The Effectiveness of Land as a Filter Medium", Amer.Ind. Hyg. Assoc. J., 19, 107 (1958).15. McFee, D. R., and Sedlet, J., "Plutonium-Uranium-Molybdenum Fume Characteristicsand Land Filtration", J. Nucl. Energy, 22, 641 (1968).16. Jansen, G. and D. D. Stepnewski, "Fast Reactor Fuel Interactions with Floor MaterialAfter a Hypothetical Core Meltdown,", Nucl Tech, 17, 85 (1973).17. Routson, R. C. and R. J. Serne, "One-Dimensional Model of the Movement of TraceRadioactive Solute Through Soil Columns" The PERCOL Model", BNWL-1718 (1972).18. Ergen, W. K., et al., "Emergency Core Cooling: Report of Advisory Task Force onPower Reactor Emergency Cooling (USAEC).VII-40

19.20.21.22.23.24.25.26.27.28.29.30.31.32.33.34.35.36.Schroeder, M. C., "Laboratory Studies of the Radioactive Contamination of Aquifers",AEC Document UCRL 13074 (1963).Rhodes, D. W., "The Effect of pH on the Update of Radioactive Isotopes from Solutionby a Soil", Soil Sci. Soc. Amer. Proc. 21:389-396 (1957).Pillai, K. C. and T. R. Folson, "The Concentration of Lithium, Potassium, Rubidiumand cesium in some Western American Rivers", Geochim. and Cosmochim. Acta,32:1229-1234 (1968).Barrow, N. J., "Comparison of the Adsorption of Molybdate, Sulfate, and Phosphateby Soil", Soil Sci., 109:282-288 (1970).Geering, H. R., E. E. Gary, L. H. P. Jones and W. H. Allaway, "Solubility andRedox Criteria for the Possible Forms of Selenium in Soil", Soil Sci. Soc. Amer.Proc., 32:35-40 (1968).Pourboug, M., "Atlas of Electrochemical Equilibria in Aqueous Solution", PergamonPress, New York (1966).Touhill, C. J., B. W. Mercer and A. J. Schuckrow, "Treatment of Waste SolidificationCondensates", AEC Document BNWL-723 (1968).Raja, M. E. and K. L. Babcock, "On the Soil Chemistry of Radio-Iodine", SoilSci., 91:1-5 (1961).Price, K. R., Unpublished Data, Battelle-Northwest, Richland, Washington (1973).Wildung, R. E., Unpublished Data, Battelle-Northwest, Richland, Washington(1973).Routson, R. C., Unpublished Data, Battelle-Northwest, Richland, Washington(1973).Code of Federal Regulations, Title 10, Part 20, "Standards for ProtectionAgainst Radiation". USAEC. Appendix B, Table II.Ralkova, J. and J. Saidl, "Solidification of High Level Wastes: Part 3, Diffusionand Rates of Radionuclides Incorporated in Basalts", Kernergie, 10:161(1967).Mendell, J. E., "A Review of Leaching Test Methods and the Leachability of VariousSolid Media Containing Radioactive Wastes", AEC Document BNWL-1765 (1973).Saidl, J. and J. Ralkova, "Immobilization of Strontium and Cesium by Fixation ofRadioactive Waste in Baslat", Collection, Czechoslov. Comun., 31, 871 (1966).Fisher, H. L., "Prediction of the Dosage to Man From Fallout of Nuclear Devices- Transport of Nuclear Debris by Surface and Groundwater" UCRL-50163 Pt. 6(January, 1972).Higgins, G. H. "Evaluation of the Groundwater Contamination Hazard from UndergroundNuclear Explosions", J. Geoph. Res, 64, 1509 (1959).Carbiener, W. A., et al "Physical Processes in Reactor Meltdown Accidents",WASH-1400, Appendix VIII, October, 1975.VII-41

TABLE VII 3-1 FISSION PRODUCT REMOVAL RATE CONSTANTS CALCULATED FORCONTAINMENT VESSELA LARGE PWRFlowFractionRateInitialSpray System gpm Release X, hr-Molecular Iodine1 CSR, Boric Acid 3500 1.0 - .01 3.122 CSR(a), Boric Acid 7000 1.0 - .01 6.24(pH = 5)(Boric Acid -- 10-2 ÷ 1.8 x 10-3 1.15Equilibrium 'Boric Acid -- 1.8 x 10-3 ÷ 3 x 10-5Conditionsi Boric Acid --

TABLE VII 3-3 EQUILIBRIUM DATA FOR 12 WITH CAUSTIC SPRAYS (Ref. 9)Time, min H C /Cgo0-100 Constant H, A=0 .01100-1000 Variable H, k=.095 hr- .011000 7.0 x 104 3.86 x 10-42000 1.5 x 105 1.8 x 10-44000 5 x 105 5.4 x 10-5>7000 1 x 106. 2.7 x 10-5TABLE VII 3-4 INPUT DATA AND RESULTS FOR FILTER OVERBEATING CRITERIAFilter Air Flow Filter Temperature Heat LoadType cfm F WattsHEPA 2000 105 2.2 x 10 3HEPA 2000 112 5.7 x 103HEPA 2000 2 5 0 (a) 5.4 x 104Charcoal 2000 110 1.3 x 10 3Charcoal 2000 126 3.4 x 103Charcoal 2000 6 4 0 (a) 6.4 x 10 4(a)Design LimitTABLE VII 3-5 DF vs PARTICLE SIZE AND REYNOLDS NUMBER IN DRYWELL ANNULAR GAPdpi DF (Re

TABLE VII 3-7 DISTRIBUTION COEFFICIENTS FOR VARIOUS RADIONUCLIDES IN GROUNDWATERRadionuclide Kd (ml/g) Reference1 0.1 25Br 0.1 25Cs 20 18Rb 15 20Te 20 23Se 20 22Sb 15 27Ba 3 18,23Sr 2 181/2 Ru 4 241/2 Ru 0 24Mo 5 21Pd 25 23Rh 25 23Tc 0.1 23,25Y 200 19Zr 200 19Ce 50 28Pr .60 28Pm 60 28Sm 60 28Eu 60 28U 100 26Pu 200 19TABLE VII 3-8 CORE INVENTORY DEPRESSURIZATION RELEASE FRACTIONS(a)ReleaseGroup Chemical Element Fraction1 I, Br 0.0112 Cs, Rb 0.0273 Te, Se, Sb 0.0404 Ba, Sr 0.0295 Ru, Mo, Pd, Rh, Tc 0.00286 Others 0.0005(a)Obtained from CORRAL calculated output for PWR accident sequence ABe.Table VII 3-1 - Table VII 3-8VII-43/44

TABLE VII 3-9 COMPARISON OF CALCULATED GROUNDWATER EFFLUENT CONCENTRATIONS TO MPCLIMITS - DEPRESSURIZATION RELEASE CASENuclideSourceInventory,CuriesTime ofPeak, yearsPeak EffluentConcentrationijCi/cc (a)MPCwCi/ccRu-106 (b)Tc-99Ba-140Sr-89Sr-90Ru-103Ru-106Te-129Te-127mSb-125Cs-134Cs-135Cs-1372.6 x 1042.1 x i004.5 x 1063.1 x 1061.5 x 1052.7 x 1052.6 x 1044.2 x 1055.1 x 1041.1 x 1044.5 x 1043.7 x 10-11.5 x 1050.590.85101124354351513.34.75.66.55.1 x 101-1.4 x 10-47.2 x 10-437.4 x 10-107.0 x 10-11.1 x lo-235.5 x 10-52.9 x 10-711.9 x 10-366.6 x 10-72.0 x 10-92.2 x 10-72.8 x 10-21 x l0-52 x 10- 42 x l0-53 x 10-63 x 10-78 x 10-51 x 10-58 x l0-45 x 10-51 x l0-49 x 10-61 x 10-42 x 10-5(a) Taken at 1500 ft from the point of release(b) Non-sorbed fractionTABLE VII 3-10 COMPARISON OF CALCULATED GROUNDWATER EFFLUENT CONCENTRATIONS TO MPCLIMITS - GLASS LEACHING RELEASE CASESource (a) Peak EffluentInventory, Time of Concentr~t'ýon MPCNuclide Curies Peak, years pCi/cc"cJ wCi/ccTc-99 7.5 x 102 0.9 3.6 x 10-6 2 x 10-4Sr-90 5.1 x 106 5.9 7.1 x 10-4 3 x 10-7Ce-141 7.6 x 104 13 (b) 9 x 10-5Zr-95 3.2 x 106 27 (b) 6 x 10-5Eu-155 6.9 x 104 103 (b) 2 x 10-4Pm-147 1.4 x 107 115 (b) 2 x 10-4Eu-154 5.6 x 104 148 1.3 x 10-8 2 x 10-5Sm-151 5.8 x 103 158 2.3 x 10-7 4 x 10-4U-235 2.3 x 100 267 2.5 x 10-10 3 x l0-5U-238 3.2 x 101 267 3.5 x 10-9 4 x 10-5Pu-241 3.8 x 106 418 (b) 2 x 10-4Pu-239 1.2 x 104 535 8.0 x lo-7 5 x 10-6Pu-240 1.2 x 104 535 8.0 x 10-7 5 x 10-6Pu-242 3.9 x 101 535 2.6 x 10-9 5 x 10-6(a)(b)(c)Time started from beginning of leaching or 1 yearPeak concentration less than 1 x 10-10 pCi/cc.Taken at 1500 ft from the point of release.after the incident.Table VII 3-9 - Table VII 3-10VII-45/46

0.06o Run A-3* Run A-40 Run A-6v Run A-7o Run A-9Id0 "=0.046O.Q0002 --0E for diffusioohoresisI I I I I I I0.002 0004 0.006 0.008 0.010 0.012 0.014Normalized N Volume of Liquid Sprayed, FT V-0.016FIGURE VII 3-1Drop-Collection Efficiencyas a Functionof Liquid VolumeSprayedPoints are plotted at mid-times of spray periods(fresh spray only)Outer Annulus External3 LeakagePrimaryCubicleIQOMain Volume 2OF/ Lower Volume4Compartment234Removal ProcessesNatural depositionNatural depositionSpray absorptionRecirculation filtersNatural depositionLeakageNatural depositionFission Product SourcesGap and melt releasesSteam explosion releasesVaporization releaseFIGURE VII 3-2Schematic of CORRAL-PWR

ElevatedReleaseCompartmentI2345Removal ProcessesNatural depositionExternal leakagePool scrubbingNatural depositionExternal leakageNatural depositionNatural depositionExternal leakageOnce -through filtrationFission Product SourcesGap,melt,steam explosion,and vaporization releasesFIGURE VII 3-3Schematic of CORRAL-BWRSection ISection 2Section 3Section 4Section 5FIGURE VII 3-4Computer Code CORRAL Flow Diagrams

FIGURE VII 3-5Idealized Flow SystemUsed for AnalysisFig. VII 3-1 - Fig. VII 3-5VII-47/48

I . 103I . 021x 1011 x0101010"1I .010,20 2 6 a 10 12 14Y.-FIGURE VII 3-6Radionuclide ElutionCurves for DepressurizationRelease1x 10"1Tc-99-Depressurization -Release-1 x 10.21 x i0, 31.,C31 x 10,4Ix 10" 51 x 1060 0.5 1.0 1.5 2.0 2.5 3.0 3.5Years0 2 4 6 8 10 12 14YearsFIGURE VII 3-6(Continued)

1 x10"3 1 x10"11 x 10-41 x 10001x10-0,I x 10"1U,nC-)1.10-6x 10-21 x 10-71 x 10,3lx10-8YEARS60YEARSFIGURE VII 3-6(Continued)GLASS LEACHING~ RELEASEQ40lx10-60 1 2 ] 4 S 6 IYEARS4 6 a 10 12 14 16YEARSFIGURE VII 3-7Radionuclide Elution Curves for Glass LeachingRelease

400 100 200 300 400 So0 600YEARS100 zo0 300 400 500 600 700 800.YEARSFIGURE VII 3-7(Continued)t300 400 SOo 600 700 800 900 1000YEARSFIGURE VII 3-7(Continued)Fig. VII 3-6 - Fig. VII 3-7VII-49/50

APPENDIX ACALCULATION OF FISSION PRODUCTRELEASE TO FUEL ROD GAS GAPbyR. L. RitzmanBattelle's Columbus LaboratoriesVII-51

Appendix ATable of Contents-9SectionAl. THE REGAP METHOD .................................................A2. RESULTS OF REGAP CALCUALTIONS ....................................REFERENCES ..............................................................Page No.VII-55VII-57VII-59List of TablesTableVII A-1Specified Input Data for Fission Product Gap ReleaseCalculations ....................................................VII A-2 Regap Method Results for Best Estimate Conditions - NobleGases ...........................................................VII A-3 Regap Method Results for Best Estimate Conditions -Halogens ........................................................VII A-4 Regap Method Results for Best Estimate Conditions - AlkaliMetals and Alkaline Earths ......................................VII A-5 Final Gap Release Estimates from the Regap Method ...............Page No.VII-61/62VII-61/62VII-63/64VII-63/64VII-63/64VII-53

Appendix ACalculation of Fission Product Release to Fuel Rod Gas GapbyR. L. RitzmanBattelle's Columbus LaboratoriesOne of the fission product release termsfor a reactor loss-of-coolant accidentis the rapid escape of gaseous activitythat occurs when fuel claddings firstrupture. The rapid depressurization ofthe ruptured rods carries out the fissionproducts that have accumulated inthe gas gap and the gas plenum duringthe normal operating period of the fuelrods. This release term is frequentlyreferred to as the gap release component.Analytical techniques have beendeveloped by different laboratories forestimating the magnitude of the release.These techniques rely on different mathematicalmodels or different interpretationsof experimental release data inmaking gap release estimates. One ofthese techniques is the REGAP computercode developed at Battelle-Columbus.This Appendix describes the results ofapplying the REGAP code to calculate thefractions of various fission productsthat would be released from the fuel ina reactor core under normal operatingconditions. The results obtained by theother two laboratories involved in theexercise are given in Appendicies B andC.All calculations were performed for acommon group of fission product nuclidesand a common set of reactor operationinput specifications. Two typical endof-lifecore conditions were used; onefor a large PWR and the other for alarge BWR. The specified data input arelisted in Table VII A-1. With the inputconditions fixed in this manner therelease results obtained from thedifferent calculation methods shoulddepend only on variations in releasemodels and/or the approach that eachuses to generate fuel temperature profilesfrom the core power distributiondata.Al. THE REGAP METHODIn this method calculations are performedassuming that the fission productsmigrate through the fuel by a diffusionprocess during reactor operation,and are released to the rod gap andplenum spaces by escape from externalsurfaces, surface-connected cracks, orsurface-connected porosity. The equivalentsphere model originally proposed byBooth (Ref. 1)1 and discussed by Lustman(Ref. 2) constitutes the basis for calculatingrelease from the sintered U0 2characteristic of water reactor fuels.The sintered fuel, although near theoreticaldensity, contains some interconnectedporosity which effectivelyincreases the surface area for releaseof fission gases. In the equivalentsphere approach, the fuel body is consideredto be composed of sphericalparticles of uniform size whose surfaceto volume ratio is equivalent to theactual surface area to fuel volume ratioof the sintered material. Solutions forthe appropriate diffusion equations fora sphere in which production and decayof the fission product are taken intoaccount have been obtained by Beck (Ref.3) and these are used directly in theREGAP method.The diffusion equation for the concentrationin a sphere is:ac = _C_, 2(rC) + B -at r ar2with boundary and initialwhereC (a,t) = C (r,O) = 0conditions(VII A-l)C = concentration in the sphere,atoms per cm 3D = diffusion coefficient, cm 2 perseciReferences are listed at end of thisAppendix.VII-55

B = production rate, atoms/ (sec)(cm3)A =decay constant, sec-It = time, secr = radial coordinate in thesphere, cma = equivalent-sphere radius, cm.The rate of release (R) is found fromthe concentration gradient at the outersurface and for a unit volume of thesolid is given byR3DR (32r) r = a (VII A-2)The accumulation of a fission productexternal to the solid can be calculatedby considering the rate at which it isreleased and the rate at which itdecays:dN = R - AN (VII A-3)d-twhere N is the accumulation of undecayedrelease atoms from a unit volume ofsphere (atoms per cm 3 ). Beck (Ref. 3)restated the problem in terms of dimensionlessvariables and obtained thefollowing solution:G = 3 (iL---coth /-•[-exp (-PT)]where6pexP2 (r) Zw n=11 - exp (-n Tr T)n 2 (n 2 2 + P)(VII A-4)NG = BA = the ratio of nondecayedatoms outsidethe sphere at any timeto the total non-decayedatoms in the systemat equilibriumXa 2DtIT = --aD = Do exp (-Q/RT)D = limiting diffusion coefficient,0 cm 2 per secQ = activation energy for diffusion,cal per g-moleR = gas constant, cal/(deg) (g-mole)T = absolute temperature, deg K.For reactor operating times that arelong with respect to the half-life of aradioactive fission product, G approachesthe equilibrium value:G = 3[(i/vj) coth /U - i/i.(VII A-5)For stable fission-product species, therelease is expressed in terms of theratio of the amount released to theamount produced. ThusN 1Bt 15T+ 6 exp(-n 2 7T)2Tr T1 n(VII A-6)The first five terms of the summationare usually sufficient to obtain anaccurate value for G.Equations (VII A-5) and (VII A-6) constitutethe basis of the REGAP method.However, each requires some basic inputdata for the various parameters. Thediffusion coefficients (D) are obtainedfrom the Arrhenius expression as a functionof temperature by specifying thelimiting diffusion coefficient value(DO) and the activation energy for diffusion(Q). Of course, these values aredifferent for each fission product chemicalelement. Extensive experimentalwork has been done over the past twodecades, heavily directed at the noblegas elements krypton and xenon, todetermine the fundamental D and Qvalues for U0 2 fuel. Most o? the workhas involved post-irradiation laboratoryheating experiments although some efforthas been spent on extracting diffusioncoefficients from results of inpile U0 2irradiation studies. Many reviews orsummaries of the field have appearedincluding ones by Lustman (Ref. 2),Childs (Ref. 4), Parker, et al. (Ref. 5),Morrison, et al (Ref. 6), and Carroll(Ref. 7). The basic diffusion parameterdata used in the present calculationswere obtained from fission productrelease results reported by Parker, etal. (Ref. 5) for trace irradiated U02.VII-56

The Do and Q values used for eachfission product are listed in Table 10of Reference 6.The fuel temperatures that are neededfor the calculation of diffusion coefficientsare generated within the computercode that incorporates the REGAP method.The radial temperature profile insidethe fuel is approximated by first dividingthe pellets into five radial segmentsof nearly equal volume. Thetemperature at each radius is computedstarting from the outside surface usinga constant U0 2 thermal conductivity andthe appropriate operating surface heatflux value according to the equation:wherer s 2k 1 -E2 (VII A-7)Tr = temperature at radius r, FTs = temperature at fuel surface, FQ = surface heat flux, Btu/(sec)(in.2 )k = fuel thermal conductivity,Btu/(sec) (in.) (F)a = fuel-pellet outside radius,in.The fuel surface temperature is calculatedby computing the temperature riseacross the rod gas gap and adding thisto the cladding surface temperature.Mathematically,whereTs = Tc = Q/K(VII A-8)Tc = cladding surface temperature,FK = gas gap conductance, Btu/secin 2 ,F.Once the radial temperature profile isobtained a simple average temperature isassigned to each of the five radialzones. This process is carried out forall subregions of the reactor core(axial and radial) that are required.In the present cases, the cores of bothreactor types were divided into 9 axialand 8 radial increments, so temperatureprofiles and fission product releaseswere calculated for 72 separate subregions.The remaining parameter that is criticalto the release calculation is the equivalentsphere radius value. This is relatedto the density of the U0 2 fuel asnoted earlier, and is easily obtainedfrom a knowledge of the percent theoreticaldensity for the fuel pellets usingFigure 9.18 and Equation 9.23 of Reference2.The mechanics of the REGAP method consistof first calculating the fuelpellet temperature profiles as outlinedabove. This step requires the followinginput data: The core average surfaceheat flux, the total power peakingfactor for each subregion, the assumedvalue for fuel rod gap conductance, anda suitable fuel thermal conductivityvalue. As noted, five zone temperaturesare calculated for each core subregion.These temperatures are then used in theArrhenius equation to compute diffusioncoefficient values for each fissionproduct species. Finally the diffusioncoefficient values are used along withthe specified equivalent sphere radiusand the appropriate radioactive decayconstant in Equation (VII A-5) or (VIIA-6) to compute the fractional releaseof each fission product specie from eachzone of each core subregion. Therelease fractions for a particularspecie are summed over the radial zonesand then these values, weighted accordingto the core power distribution, arecollected along the axial direction ofeach radial core region to generate aset of core fractional release values ona full length rod basis. Since thepresent cases use eight radial coreregions, the set contains eight values.These may then be summed to produce atotal core release fraction for each ofthe fission product species.A2. RESULTS OF REGAP CALCULATIONSA series of calculations was performedin which the values of various parameterswere changed to examine the effecton release predictions. As indicated inTable VII A-l, two average gap conductancevalues were used, 500 and 1000Btu/hr,ft 2 ,F. In addition the fuelthermal conductivity was varied from 1.2to 1.5 Btu/hr,ft,F, the equivalentsphere radius was varied from 0.00613 to0.0184 cm, and Do values for eachfission product specie were varied from1 to 100 times the reference values.The ranges of these variations wereconsidered to represent the approximatedegree of uncertainty in the availableinput data although the variation in DoVII-57

values could also be interpreted as anestimate of the possible effect of fuelburnup on the release rates. The bestestimate values for this series ofparameters are considered to be thefollowing:a. Gap Conductance = 1000 Btu/hr, ft2 , Fb. U0 2 Thermal Conductivity = 1.2Btu/hr,ft,Fc. Equivalent Sphere Radius = 0.00613cmd. The Do Parameter = the referencevalue for each fission productelement.The fission product release results obtainedusing the REGAP method for eachof the reactor types (PWR and BWR) arelisted in Tables VII A-2 to VII A-4.The results for the stable fission productspecies are not of great interestexcept that they indicate the magnitudeof release of radioactive species whosehalf-lives are long compared to the corelifetime of about 3 years (108 seconds).This includes species such as Kr-85, I-129, Cs-137, and Sr-90. The results forthe much shorter lived species calculatedhere illustrate the square root ofhalf-life dependence for release thatthe diffusion model contains. Comparingresults for the two reactor types showsvery little difference and so there isno need to separate the release values.The total core release results are thedata of principal interest. The effecton these values of the parametersensitivity examinations can be summarizedas follows:a. The total core release values forthe shorter lived fission productspecies were affected to a greaterdegree than were the values for thestable (or very long lived) species.b. A factor of two change in gap conductanceproduced a factor of 1.4 to4 change in release values.c. A twenty-five percent change in U0 2thermal conductivity produced a factorof 1.4 to 6 change in releasevalues.d. A factor of three change in equivalentsphere radius produced a factorof 2 to 3 change in release values.e. A factor of one hundred change inDo values produced a factor of 2 to10 change in release values.Because of these variations and thesemi-empirical nature of the REGAPmethod itself, it is possible to provideonly a rough estimate of the gap release.A listing of the estimates thatare indicated by this work are given inTable VII A-5. These total core releasevalues are generally applicable towater-cooled power reactors and shouldbe considered the "best estimate" thatis currently possible. Nevertheless thesensitivity calculations suggest thateach is uncertain by a factor of aboutfive, either higher or lower. Finallyit deserves mention that the data inTable VII A-5 indicates potential forrelease from the U0 2 fuel only, thevolatility of released species in thefuel rod gas spaces has not beenconsidered. Therefore, the occurrenceof plateout on or of chemical reactionswith the cladding could cause areduction in the fraction of the gaprelease inventory that is available forrapid escape when the cladding perforates.VII-58

References1. Booth, A. H., "A Method of Calculating Fission Gas Diffusion from U0 2 and ItsApplication to the X-2-f Loop Test", AECL Report CRDC-721 (Sept. 1957).2. Lustman, B., "Irradiation Effects in Uranium Dioxide", Chapter 9 in UraniumDioxide: Properties and Nuclear Applications, edited by J. Belle, USAEC, 1961.3. Beck, S. D., "The Diffusion of Radioactive Fission Products from Porous FuelElements", BMI-1433 (April, 1960).4. Childs, B. G., "Fission Product Effects in Uranium Dioxide:, J. Nucl. Mater, 9,217 (1963).5. Parker, G. W., et al., "Out-of-Pile Studies of Fission Product Release fromOverheated Reactor Fuels at ORNL, 1955-1965", ORNL-3981 (July, 1967).6. Morrison, D. L., et al., "An Evaluation of the Applicability of Existing Data tothe Analytical Description of a Nuclear Reactor Accident", BMI-1779 (August,1966).7. Carroll, R. M., "Fission Gas Behavior in Fuel Materials", Nucl. Safety, 8, 345(1967).VII-59

TABLE VIIA-1SPECIFIED INPUT DATA FOR FISSION PRODUCT GAP RELEASE CALCULATIONSPWRBWRFuel DataFuel pellet diameter, in. 0.367 0.488Cladding O.D., in. 0.422 0.562Cladding thickness, in. 0.024 0.032Fuel density, % T.D. 94 94Thermal DataAverage heat flux, Btu/hr, ft2 191,000 163,0002Peak heat flux, Btu/hr, ft534,000 425,000Maximum peaking factor 2.8 2.6Average cladding surface temp., F 610 560Average gap conductance, Btu/hr, ft, F 1000 1000Average gap conductance, Btu/hr 500 500Average thermal output, Kw/ft 6.2 6.2Maximum thermal output, Kw/ft 17.3 18.3Other Operational DataCore end-of-life burnup, MWD/MTU 30,000 30,000Core end-of-life operating time, sec 108 108Core Segmentation and Power DistributionsAxial - 9 equal length sections - samefor both reactor typesAxial SectionPower Factor1 (top) 0.482 0.803 1.124 1.305 1.486 1.327 1.158 0.829 (bottom) 0.53Radial and Local - 8 unequal volume, full-length regionsfor each reactor but with different power factorsas tabulated below. These regions really representdifferent groups of full length rods with the higherpower regions of the core being segmented in greaterdetail.Radial X Local Fraction of Sum of Radial X Local Power FactorRegion Total Core Rods Fractions PWR BWR1 0.01 0.01 1.85 1.702 0.02 0.03 1.65 1.543 0.02 0.05 1.50 1.444 0.05 0.10 1.45 1.355 0.10 0.20 1.33 1.266 0.15 0.35 1.19 1.147 0.25 0.60 1.05 1.028 0.40 1.00 0.68 0.76

TABLE VII A-i(CONTINUED)Fission Product SpeciesStable or Long-LivedXe, I, Cs, SrRadioactiveXe-133Xe-1351-1311-1321-133Cs-138Sr-89Sr-91Half-Life(5. 27da)(9. 2hr)(8.06da)(2. 3hr)(20. 8hr)(3 2min)(51da)(9. 7hr)TABLE VII A-2 REGAP METHOD RESULTS FOR BEST ESTIMATE CONDITIONS - NOBLE GASESTotal Core Release Percentages for Noble Gas SpeciesRadial Fraction of Stable Xe Xe-133 Xe-135Region Core Volume BWR PWR BWR PWR BWR PWR1 0.01 0.56 0.66 0.15 0.18 0.05 0.052 0.02 0.83 0.90 0.17 0.18 0.05 0.053 0.02 0.63 0.65 0.11 0.10 0.03 0.034 0.05 1.37 1.54 0.21 0.23 0.06 0.075 0.10 1.91 1.98 0.25 0.26 0.07 0.076 0.15 1.96 1.89 0.24 0.23 0.07 0.067 0.25 1.97 1.76 0.24 0.20 0.06 0.068 0.40 0.40 1.29 0.04 0.01 0.01 0.004TotalCore 1.00 9.63 10.67 1.41 1.39 0.40 0.39Table VII A-I - Table VII A-2VII-61/62

TABLE VII A-3 REGAP METHOD RESULTS FOR BEST ESTIMATE CONDITIONS' - HALOGENSRadialRegionFraction ofCore VolumeBWRStable IPWRTotal Core Release Percentages for Halogen Species1-131 1-132BWR PWR BWR PWRBWR1-1331-133PWR120.010.020.67 0.781.01 1.110.33 0.40 0.071 0.0860.44 0.46 0.071 0.0730.17 0. 210.19 0. 2030.020.80 0.830.28 0.28 0.039 0.0380.11 0.1140.051.74 1.950.60 0.69 0.077 0.0880.22 0.2650.102.64 2.750.64 0.63 0.077 0.0740.23 0. 2260.152.65 2.580.61 0.53 0.072 0.0610.21 0.1870.252.44 2.070.54 0.39 0.063 0.0450.19 0.1380.400.18 0.020.02 0.003 0.003 -0.01 0.001TotalCore1.0012.13 12.093.46 3.38 0.47 0.471.33 1.31TABLE VII A-4 REGAP METHOD RESULTS FOR BEST ESTIMATE CONDITIONS - ALKALI METALS AND ALKALINE EARTHSRadialRegionFractionof CoreVolumeStable CsBWRPWRTotal Core Release Percentages for Alkali MetalCs-138Stable SrBWR PWR BWR PWRand Alkaline Earth SpeciesSr-89BWRPWRBWRSr-90PWR10.010.43 0.510.012 0.015 0.34 0.390. 24 0.280.065 0.07820.020.59 0.630.010 0.011 0.42 0.460.26 0.280.042 0.04330.020.42 0.410.005 0.005 0.28 0.240.15 0.120.018 0.01340.050.81 0.910.008 0.009 0.41 0.440.18 0.200.019 0.02050.100.97 0.960.008 0.008 0.43 0.380.15 0.140.015 0.01360.150.75 0.660.006 0.005 0.15 0.130.05 0.040.004 0.00370.250.55 0.420.004 0.003 0.01 0.0080.004 0.00380.400.30 0.005TotalCore1.00 4.55 4.51 0.053 0.056 2.04 2.05 1.03 1.06 0.16 0.17

TABLE VII A-5 FINAL GAP RELEASE ESTIMATES FROM THE REGAP METHODSpeciesKr, Xe (stable-long lived)Total Core Release (a)(Percent)10.0I(stable-long lived)10.0Cs (stable-long lived)Sr (stable-long lived)Xe-133Xe-1351-1311-1321-133Cs-138Sr-89Sr-915.02.02.00.43.00.51.00.051.00.2(a) Values estimated uncertain by about a factor of +5.Table VII A-3 - Table VII A-5VII-63/64

APPENDIX BRELEASE OF FISSION PRODUCTS FROM PWR AND BWR FUEL PINSbyH. L. MeMurry and E. F. AberAerojet Nuclear CompanyVII-65

Appendix BTable of ContentsSectionBI. METHODS ..........................................................B2. APPLICATION TO RELEASE OF LONG-LIVED NOBLE GASESAND IODINES ......................................................B3. APPLICATION TO SHORT-LIVED NOBLE GASES AND IODES .................B4. TREATMENT OF CESIUM ...............................................B5. EQUATIONS FOR THE POWER DISTRIBUTION .............................B5.1 Equation for the Time Dependent Average Concentrationin a Pin ...................................................B5.2 The Release Fraction Fi from a Single Pin inRadial Segment i ...........................................B5.3 Release Fraction for the Entire Core .......................B5.4 Calculations of Tbji and (DT/dr)bji ........................B6. EQUATIONS FOR STRONTIUM ..........................................B7. RESULTS ..........................................................REFERENCES ..............................................................ADDENDUM ................................................................Page No.VII-69VII-70VII-70VII-71VII-71VII-72VII-72VII-73VII-73VII-74VII-76VII-77VII-83List of TablesTableVII B-1Specified Input Data for Fission Product Gap ReleaseCalculations ....................................................Page No.VII-79/80VII B-2 Long Lived Fission Product Species ..............................VII B-3 Short Lived Isotopes ............................................VII B-4 Release Percentages of Xe and I Isotopes At 108Second ..........................................................VII B-5 Cesium Release Percentages at 108 Seconds .......................VII-79/80VII-81/82VII-81/82VII-81/82VII B-6Boundary Temperatures and Heat Ratings for the Hot Pin inthe PWR .........................................................VII-81/82VII-67

Appendix BRelease of Fission Products from PWR and BWR Fuel PinsbyH. L. McMurry and E. F. AberAerojet Nuclear CompanyThis work is in support of the ReactorSafety Study which aims at realisticevaluations of the effects of reactoraccidents. To determine the result ofcracking of the cladding in fuel pins ofoperating reactors, it is important toknow the amount of fission productswhich have migrated to the fuel-claddinginterface during operation. This calculationgives the fractions of variousfission products which have accumulatedin the gap next tothe inner surface ofthe cladding. The isotopes consideredwere the noble gases, their iodine precursorsand cesium and strontium.To give a common base for comparingcalculations from different groups, tworather typical cases were considered.One reactor is a PWR while the other isa BWR. The operating conditions usedare the same as those used by the otherlaboratories. These input conditionsare given in Table VII B-1.BI. METHODSThe methods used here are based on abalance equation for fission productdiffusion which assumes that the drivingforce for diffusion of fission productsis the gradient of the chemical potentialof the diffusing particles (Ref.1). W. Yuill has incorporated this ideainto his models for calculating fissionproduction diffusion (Ref. 2-5).The balance equation in cylindrical coordinatesis (Ref. 6)dc = P(r) - Xc(r,t) + I A r [c dii3dt r dr LRT? jFwhere(VII B-l)c (r,t) concentration at radialposition r at time t ofsome particular fissionproduct. (mols/cc)P(r) = Production of the fissionproduct at r (mols/ccsec.)IaG\A = decay constant for thefission product (sec- 1 )D = Fick's law diffusion constantfor the product(cm 2 /sec)C )T= chemical potential for thefission product. The G isthe Gibbs free energy.[Cal/mol/cc]The expression (Dc/RT)-(dp/dr) whichappears in Equation (VII B-1) involvesapproximation. First it is assumed thatit may be approximated by the idealequation (Ref. 7).1= O (T) + RT ln c (VII B-2)The equation for the diffusion currentis (Ref. 6).J=-'dii _ 'F i'-Dr -D1 [STPc! T.+ !dr 43cO rIf there were no temperature gradientsthis would reduce to Fick's law. Using(DP/Dc)T from Equation (VII B-2) thisyieldsD'- DcRT "This is how the Dc/RTEquation (VII B-1).appearsTo apply Equation (VII B-l) requires aknowledge of (ap/3T)c, D and dT/drthroughout the pin. The temperaturegradient dT/dr can be calculated from aknowledge of the heat generation withinthe pin and the heat conductivity Kc(T)as a function of temperature. This isreadily accomplished if it is assumedinVII-69

that Kc depends only on temperature andis not seriously affected by such structuralchanges as may occur within theU0 2 pellet due to the marked temperaturechanges between the boundary and *theinterior. The methods used for calculatingdT/dr will be discussed in connectionwith the calculation on Cs andSr isotopes where explicit determinationof dT/dr at the boundary has been made.For the noble gases and their iodineprecursors the code used (Ref. 5) calculatesthe boundary temperature from anempirical relation involving the heatrating fK cdT (see Addendum).The diffusion coefficient D is assumedto have a temperature dependence givenbyD = D e-Q/RT0 (VII B-3)Data on Do and Q for noble gases areincorporated into the FPFM code (Ref.5). Data for Cs and Sr are given inReference 4. Data for (Dp/aT)c areavailable only for Cs and Sr (Ref. 4).Because of this the releases of noblegases and their iodine precursors arecalculated using a correlation which ispartly based on Equation (VII B-l) andpartly empirical.G = heat rating = jlO(see below) T T bPz = watts/cm in pinTKcdT Zc 7Kc = heat conductivity (watts/cm-C)A, B, C are constants derived fromanalysis of capsule data.The heat rating and the power per centimeteris used in several places. Theseare related through the heat conductionequation. Thus in cylindrical coordinatesa (rK ) = P - z (r) x rDrc d/ rb 2Where Pz is the average watts/cm andp(r) is a function to take account ofthe variation in production rate withposition r. In this work a flat productionis used and p(r) = 1. The casef(r) = 1.0 is discussed in the Addendumto this Appendix. When f(r) = 1.0,B2. APPLICATION TO RELEASE OF LONG-LIVED NOBLE GASES AND IODINESTbK cdTTT = G. (VII B-5)In calculating releases of long-livednoble gases and iodines Yuill, et al.(Ref. 4) developed equations based onthe steady state form of Equation (VIIB-l) with dc/dt = 0. For stable gases(X = 0) they cast their equations intothe form (Ref. 5).ln (1-F) = -wheree-Q/RTb e[A+BG+CG2 ]b 2 TbF = release fractionb = pellet radius (cm)e(VII B-4)Tb = boundary temperature of pellet(OK)Q = activation energy for Fickiandiffusion as in Equation (VIIB-3)Also,IdT\Tr)bP zb (F(VII B-6)To apply Equation (VII B-4) to an actualreactor, account must be taken of thepower distribution. The FPFM code doesthis from the power distribution specifications.These have been provided forthe PWR and BWR cases (see Table VIIB-l).B3. APPLICATION TO SHORT-LIVED NOBLEGASES AND IODINESYuill has shown that the release fractionsfor short-lived isotopes is proportionalto 1/X. A detailed account ofthis is given in connection with the Csand Sr calculations.Following Yuill's suggestion, the assumptionhas been made in this work thatthe release of 8.05d 1131 is the same asVII-70

for its stable Xe1 3 1 daughter. Then thereleases of all other noble gases andiodines with half lives shorter than8.05d are calculated from (see equationVII B-20)Define1. f =P 1X 131F = F 131 -F 131m Xe A Xemtmti131(VII B-7)In Equation (VII B-7) ti131 andto the half lives of 1 1 3 1 tm referand thespecies m, respectively. The Equationis applied for noble gases or iodinesfor which tm

B5.1 EQUATION FOR THE TIMEDEPENDENT AVERAGECONCENTRATION IN A PINIfdudrtaii_)a+ RT dcc drdTl + /ai P 2c~ = 2jj'\ dTThen Equation (VII B-1) yieldsa- Xa = 2 KDF(VII B-12)B5.2 THE RELEASE FRACTION Fi FROM ASINGLE PIN IN RADIAL SEGMENT iLetN .. = amount ofproduct in thethegapfissionat thej,ilocation (mols)bd I Cr dr dtf 0Define- JbPr dr - X J Cr dr0 0dc+ b[DKc +DTrb(VII B-10)cji(t) JdN .S2idtAh == average concentration offission product in a pinin radial segment i atthe jth axial level.(mols/cc)= 27rthb[DK] bjici - ANgi(VII B-13)length of a segment (9Ah= Z in this problem)(DK)bji = value of DK on the pelletboundary of a pin in ra-b 2 C(t) 2bC (r, t) rdrdial segment iaxial level.at the jth20fbP (r) rdr .The first term in Equation (VII B-13) isthe leakage rate from the pellet, thesecond is the decay of the material.Integrating,0Then N gj(t) -rb 2 Aha..e- tfotcji (t)e'tdtdc (t) -dt-Ac + ý[ DKc + 29ldb(VII B-lI)0L . (DK),j (VII B-14)InEquations (VII B-10) and (VII B-lI)K .(i) kýTc RTT1 dT dr"Using Equation (VII B-12) and writingthe power density in the ji elements asP h.f. z 1 1With the temperature gradients characteristicof these pellets Kc>> dc/dr and(Ddc/dr)b may be neglected in Equation(VII B-11).The further simplification is now madethat Cb(t) - C(t). Then Equation (VIIB-Il) yieldsyieldsN gi (t) = AhPz hjf iaVII-72

Ie-Xt (eji t-1)a i1 (1-At)(VII B-15)Equation (VII B-19) was used to calculaterelease of Cs 1 3 7 . It has thecorrect asymptotic form, namelyF.i 0 as t 0The fractional release fromradial segment i is given byapin atF. ÷ 1 as t +1For short lived isotopes Xt>>l and cji//

The a in Equation (VII B-23) is a heattransfer coefficient calculated from theequation1 Arh=b K C+ 1 rbh bg(VII B-24)hb is a conductance for the water-cladboundary layer. It was estimated by usingthe temperature difference betweenthe average water temperature and theaverage clad boundary temperature.Water temperatures for the typical reactorswere used, to compute hb for thePWR and BWR cases, respectively. TheKc is the clad conductivity and r is theclad thickness. The hg is the gapconductance given in Table VII B-1 andrb/b is the ratio of the clad radius topellet radius. Equation (VII B-24) iswritten so that Tb-Tw = a ý where c isthe heat flux.This isgiven byproducts out of the pin. This is truefor Cs because (Dp/at)c is positive(Ref. 4). For Sr (3p/3t)c is negative1and the thermal gradient alone will keepthe fission products in the pin. However,there must be some release due toevaporative processes. This has beenestimated as follows.Consider an annular segment of thicknessk at the outer boundary of the pellet.At equilibrium the fission products producedwithin the segment per secondequals the outflow. For the j,i locationthis may be writtenP,2TrbAh9 3rb2= 27rbAh(KDC + $]bji7rb22 [KDC + $]b.. 3(VII B-27a)(VII B-27b)NThis leads to\cdr,/b j ( 1T i =4ýjiIn Equations (VII B-27a, 27b), Pji isthe average watts/cm in the pin at J,i.Ah = length of the segmentTb.312=T + - G cjw b ji(VII B-25)k = thickness of the annular regionadjacent to the boundary(cm)2 G.i = flux out of the j,i location.The (dT/dr)b.i is given (see Equation(VII B-6)] bi( drTb2 GKb Kcb(VII B-26)= flux of particle leaving thesegment due to evaporativeprocesses, moles/cm 2 /sec.They go into the gap and areassumed not to return.An equation for ýb is needed. Let Q bethe potential barrier which must beovercome for a particle to emerge intothe gap. Assume, for simplicity, thatthe particles in the boundary layer havea Maxwellian distribution of energiesdefined byKcbji = conductivity on the boundary.This is temperaturedependent and was computedusing Lyon's equation (Ref.8).C (E)27rCb [1TkT] 3/2'_ eE/k = N ~ E/kTK = + CT3 ;A 28.24 °Kc B+T B = 129.4 OK -13C = 4.79 x 10B6. EQUATIONS FOR STRONTIUMEquations (VI B-19,22) assume that thethermal gradient drives the fissionIValues of Do, Q, (ia/aT)c for Sr givenin Table X of Reference 4 areDo= 5.14 x 10-7 cm 2 /sec, andQ = 24,400 cal/mol, (Wa/T)c = -4cal/mol deg. This leads to DK beingpositive when dT/dr is negative. Theresult is an inward driving force fromthe thermal gradient.VII-74

C(E)dE = mols/cc of particles withenergies in dE at E.Assume that particles emerging into thegap have energies E-Q (E>Q) and particleswith E

The speed v may be estimated by writingJ2/,J k T []1/2 e-Q/kT2.2 x 105[ li'3V' T Q e-Q/kT(VII B-34)In Equation (VII B-34) the 2.2 x 105cm/sec is the speed of an atom of oneatomic unit at 293*K. The next factorin brackets converts this to the speedwhich an atom of mass m would have at2temperature T. The [Q/kTIl/ e-Q/kTthen appears as a factor which takesaccount of the barrier. The result is,using m = 88 for Sr,= 2.1 x 104 F TUsing T • 1000, Q/kT 1, 3,ifThis isn6 x 1 04e-Q/kTQ = 24,000 cal/mol.1 cm/secto be compared withDK =D eQ/kT 1 dT tau\e-Q/kTwould seem large. This would lead to afractional release- of the order of4 x 10-4 percent.B7. RESULTSThe PWR and BWR operating conditionsused in the calculations are containedin Table VII B-1.Tables VII B-2 and VII B-3 give data onthe long-lived and short-lived products,respectively.Table VII B-4 gives releases in percentfor xenon and iodine isotopes. Two valuesare given. The higher value is anupper bound to what might be expected.The lower value (1/3 of the higher) is amore probable or "realistic" value. Thefactor 1/3 was suggested by W. A. Yuillas being an appropriate scaling factor.The high values were calculated directlyfrom the FPFM code (Ref. 5) which wasdesigned to give an upper bound to whatwould actually occur. The power distributionsgiven in Table VII B-1 were usedto compute the required heat fluxes atdifferent core locations. Equation (VIIB-l) was used to estimate releases ofshort-lived species relative to theXe 1 31 release. This is certainly conservative.Table VII B-5 lists the releases forCs 1 3 7 (29y) and Cs1 3 8 (32 min). Thesewere calculated using Equation (VIIB-19) for the three. long-lived isotopesCs 1 3 3 (stable), Cs 1 3 5 (2.6 x 106 y) andCs 1 3 7 (29 y). Equation (VII B-20) wasused for the short-lived Cs 1 3 8 (32 min).The data on Do, (3p/pt)c, and Q are fromTable X, Reference 4.anu 5 x 10- 13 cm/sec.D= 8.53 x 10-9 cm2/secIt is seen that v >>DK.Assuming v >>DKandF 29F. b2 1 (VII B-35)F = Fm. * :1(VII B-36)Equation (VII B-35) simply says that theentire production adjacent to the gapleaves by evaporation. The releasedepends on the depth from which Sr atomscan be expected to evaporate. Thiscannot be very large, 100 A or 10-6 cmk•f = 5 cal/mol deg.Q = 6100 cal/molThe methods used here based on Equation(VII B-35) indicate that the release for28 y Sr 9 0 is very small - less than0.0004 percent. It would be completelynegligible for the short lived isotopes.Since the releases are strongly affectedby the calculated boundary temperaturesand temperature gradients, they aregiven in Table VII B-6 for the hot pinin the PWR. " This is the most extremecase.aVII-76

References1. An excellent discussion of the ideas involved is given in "The Thermodynamics ofthe Steady State," by K. G. Denbigh, John Wiley, (1950). Yuill gives additionalreferences to recent applications. A forthcoming report describes the derivationand gives a different result for Equation (VII B-l).2. Baston, V. F., E. G. Good, W. A. Yuill, "Fission Product Release Analysis," ID0-17242, May 1969.3. Yuill W. A., V. F. Baston, J. H. McFadden, "Release of Noble Gases from U02 FuelRods," IN-1346, November 1969.4. Yuill, W. A., V. F. Baston, J. H. McFadden, "An Analytical Model Describing theBehavior of Fission Products in Opefating Fuel Pins," IN-1467, June 1971.5. Baston, V. F., J. H. McFadden and W. A. Yuill, "An Analytical Method for CalculatingSteady-State Fission Gas Release.. .Fission Product Fuel Model (FPFM)Code," ANCR-1010, September 1971.6. The derivation of Equation (VII B-l) is discussed in a forthcoming report (H. L.McMurry, "Theoretical Basis for Calculating Diffusion of Fission Products in FuelPins," CI-1253.) A variant of Equation (VII B-l) replacesdij by QXdT+ 1 ddr T dr c/T crHere QX is the heat of transport. This is the form actually used in thecalculations.7. Denbigh, K. G., "The Principles of Chemical Equilibrium," Cambridge UniversityPress 1964, Chapter 3.8. Lyons, M. F., et al, "U0 2 Pellet Conductivity from Irradiations with CentralMelting," GEAP=4724, July 1964.VII-77

TABLE VII B-1 SPECIFIED INPUT DATA FOR FISSION PRODUCT GAP RELEASE CALCULATIONSPWRBWRFuel DataFuel pellet diameter, in. 0.367 0.488Cladding O.D., in. 0.422 0.562Cladding thickness, in. 0.024 0.032Fuel density, % T.D. 94 94Thermal DataAverage heat flux, Btu/hr, ft2191,000 163,000Peak heat flux, Btu/hr, ft2534,000 425,000Maximum peaking factor 2.8 2.6Average cladding surface temp., F 610 560Average gap conductance, Btu/hr, ft2 ,F 1000 1000Average gap conductance, Btu/hr 500 500Average thermal output, Kw/ft 6.2 7.0Maximum thermal output, Kw/ft 17.3 18.3Other Operational DataCore end-of-life burnup, MWD/MTU 30,000 30,000Core end-of-life operating time, sec 108 108Core Segmentation and Power DistributionsAxial - 9 equal length sections - same for bothreactor typesAxial SectionPower Factor1 (top) 0.482 0.803 1.124 1.305 1.486 1.327 1.158 0.829 (bottom) 0.53Radial and Local - 8 unequal volume, full-length regionsfor each reactor but with different power factorsas tabulated below. These regions really representdifferent groups of full length rods with thehigher power regions of the core being segmentedin greater detail.Radial X Local Fraction of Sum of Radial X Local Power FactorRegion Total Core Rods Fractions PWR BWR1 0.01 0.01 1.85 1.702 0.02 0.03 1.65 1.543 0.02 0.05 1.50 1.444 0.05 0.10 1.45 1.355 0.10 0.20 1.33 1.266 0.15 0.35 1.19 1.147 0.25 0.60 1.05 1.028 0.40 1.00 0.68 0.76

TABLE VII B-1(CONTINUED)Fission Product SpeciesRadioactiveHalf-LifeXe-133(5.27da)Stable or Long-Lived Xe-1351-131(9.2hr)(8.06da)Xe, I, Cs, Sr 1-132 (2.3hr)eJ1-133(20.Shr)Cs-138Sr-89Sr-91(32min)(51da)(9.7hr)TABLE VII B-2LONG LIVED FISSION PRODUCT SPECIES (a)Chain a(b)Yield (%) a barns5 4 Xe131 Sn131(3.4m)-Sb 131(23m) 85> Tea31(24m)-II 3 1 (8.05d)0.992 >Xee31 (stable) 2.9 Xe 131851131 0.754Xe132 Sn132 (2.2m)-Sb 132(2.1m)Te 132(77h)-I 32(2.30h)-Xe132(stable) 4.3 Se132Te 129(72m)_Ii29(l.7x107y) ---- Xe129(stable) 0.9 112928Tota1 5 3 I 0.95 5 Cs133 Sb133(4.1m)0. 7 2 Te 133(63m)-Ii33(20.8h)0.9 7 6 Xe133(5.27d)-Cs133 6.5 x133 1905 5 Cs135 Te135(

TABLE VII B-3 (SHORTabLeVEDISOT2OPES frncs(See Table VII B-2 for references)IsotopesChain Yield, %a'barns5 4 Xe 133(5.27d)Sb 133(4.1m) 7 2 -Te 133(63m)-I133 (20.8h)-Xe 133(5.27d)6.6Xe133 190133531 (20.8h)5 4 Xe 135(9.2h)Te 135(

TABLE VII B-5CESIUM RELEASE PERCENTAGES AT 108 SECONDSCesium-137-135-133Gap Conductance, Btu/hr-ft 2 -FCesium-138Gap Conductance,(32 min)Btu/hr-ft2-FRadialRegionPWR1000 500BWR1000 500PWR1000 500BWR1000 50014461 2440 0.00400.00790.00170.003323854 2034 0.00310.00600.00140.002633349 1931 0.00250.00480.00120.002443147 1627 0.00230.00460.00096 0.001952841 1425 0.00200.00360.00087 0.001762435 1120 0.00160.00280.00067 0.001371929916 0.00120.00220.00056 0.000978813590.00050.00080.00030 0.00050Total Core203110170.00140.00260.00063 0.0012D = D e-Q/RT with D =o 0 8.53 x 10-9 cm 2/sec and Q = 6100 cal/mol * c= 5 cal/mol deg (see Ref. 4).TABLE V1I B-6 BOUNDARY TEMPERATURES AND HEAT RATINGS FOR THE HOT PIN IN THE PWRAxial G Gap Cond = 1000 Btu/hr ft2 F Gap Cond 500 Btu/hr ft2 FRegion (watts/cm) T (OK) (dT/dr (*C/cm) Tb (OK) (dT/dr (oc/cm)1 14.4 726 -1370 835 -15402 24.0 827 -2550 1010 -30003 33.6 927 -3930 1180 -47804 38.9 984 -4790 1280 -59305 44.3 1040 -5720 1375 -71306 39.5 990 -4900 1290 -60507 34.4 937 -4070 1200 -49808 24.6 833 -2630 1020 -30909 15.9 742 -1540 862 -1750Tb calculated from Tb.. = T 2 + a ji PWR3h= 1000 Btu/hr ftg -21 + Rclad + rclad 1 h = 500hfilm kclad b hgap a = 4.16a= 2.39 C/w/cm C2 FFor PWR hfilm = 5788 Btu/hr-ft -F based on average bulk h BWRwater T of 577 F and average cladding F g = 1000T of 560 F. 2 = 2.55Kclad for Zircaloy - 4 = 8.89 Btu/hr-ft -F.c h 4500ga= 4.31Table VII B-3 -Table VIIVII-81/82

AddendumtoAppendix BofAppendix VIIRelation of Heat Rating to Power PerUnit Length of RodFission product releases are oftenexpressed in terms of the heat ratingIn Equation (A2.3) the P is the averagepower densityJTbTK dT(A2.1)bP -2 f P(r)rdr0where Kc is the conductivity and To andTb are the temperatures at the centerlineand outer boundary, respectively.The heat rating is related to the powerper unit length, Pz, in the pin. Therelation is particularly simple when thepower density in the pin is flat.Let P(r) be the heat production (powerdensity) in cal/cc sec or watts/cc.Then if the P depends on r, but not onthe radial angle 0If Equation (A2.3) is used(A2.2) the result isbJo KcdTc d0Tb 24ina n(n -+ 2)Equation2WA. 4)2rrP(r)rdr = -K dT - 2rrfc dr0(A2.1)In this equation the left side gives thetotal heat.production out to the radiusr. Under steady state conditions thismust all flow outward over the boundaryat r by conduction. The right side ofEquation (A2.1) expresses this conduction.Integration of Equation (A2.1) yieldsThe average power density, F(r), ifrelated to the power per unit length inthe rod, Pz, bybPz = 2Tf P(r)rdr=zb20(A2.5)fb 1 f P(r')r'dr' droTf OKcdTTbUsing Equation (A2.5) in Equation (A2.4)yieldsAssume(A2.2)0bKdT = 2 1 +4n n=Oa n 21"(n+2)P(r) = (1 n~o(A2.3)It is seen that the heat rating is simplyrelated to Pz when P = P, i.e.,when the power distribution in the pinis uniform. This approximation is oftenassumed in release analyses.VII-83

APPENDIX CCALCULATION OF GAP RELEASE OFRADIOACTIVE FISSION PRODUCTSbyG. W. ParkerOak Ridge National LaboratoryVII-85

Appendix CTable of ContentsSectionCl. BACKGROUND .......................................................C2. CALCULATION OF PWR AND BWR FUEL ROD GAP RELEASES .................C2.1 Diffusion Coefficient Values ...............................C2.2 Release Computation Procedure ..............................C3. THERMAL ANALYSES OF THE REFERENCE CORES ..........................C3.1 Thermal Analysis of the PWR Fuel (Nine CapsulePositions) .................................................C3.2 Thermal Analysis of the PWR Fuel at the Peak and AverageHeat Fluxes and With Gap Conductances of 1000 &500 BTU/HR.Ft2 -F ...........................................C3.3 Thermal Analysis of the BWR Fuel (Nine CapsulePositions) .................................................C3.4 Thermal Analysis of the BWR Fuel at the Peak and AverageHeat Fluxes and With Gap conductances of 1000 and500 BTU/HR.Ft2 .F ...........................................REFERENCES ..............................................................Page No.VII-89VII-90VII-90VII-91VII-92VII-92VII-94VII-94VII-94VII-95List of TablesTableVII C-1 Values for Constants in the Arrhenius Equation...................VII C-2 Initial Gap Release Results .....................................VII C-3 Final Estimate of Gap Release Values ............................Page No.VII-97/98VII-97/98VII-97/98VII C-4VII C-5Results of the Nine Capsule Position Calculations at PeakRadial Power Position in PWR Core ...............................Results of the Nine Capsule Position Calculations at PeakRadial Power Position in BWR Core ...............................VII-99/100VII-99/100VII-87

List of FiguresFigureVII C-I Diffusion Parameter D' Derived from ORNL - Xenon AnnealingData with Comparison Data of 01 and TAGAKI and AECL .............VII C-2 Computed D' Values for Fission Products .........................Page No.VII-101/102VII-101/102VII C-3Section of PWR Fuel Rod Showing Geometry and Thermal DataUsed for Temperature Analysis ...................................VII-101/102VII C-4 Urania U0 2 Thermal Conductivity .................................VII-101/102VII C-5Section of BWR Fuel Rod Showing Geometry and Thermal DataUsed for Temperature Analysis ...................................VII-101/102VII-88

Appendix CCalculation of Gap Release ofRadioactive Fission ProductsOak Ridgeby.W. ParkerNational LaboratoryC1. BACKGROUNDConventional methods for calculatingfission gas release from U0 2 are basedmainly on the Booth diffusion model concept;however, calculating the releaseof radioiodine is slightly more involvedsince the diffusion parameter, D', foriodine often is reported to exceed thatfor xenon by a factor of at least two.In fact, the results of experimentaldiffusion studies, the annealing type,show several fission products (Cs andTe) with higher diffusion coefficients.The effect of extended fuel irradiation(burnup) is also significant and may accountfor more than an order of magnitudein the equilibrium rate of release.Calculations based on conservative assumptionsfor radioiodine release to theclad gap were given in a recent paper(Ref. 1). An important conclusion isdrawn from this paper in that at heatratings below 10 kw/ft only a few percentof the iodine is released whiledoubling the power increases the releaseby a factor of ten.In addition to the steady-state diffusionprocess, a second contribution offission gases, radioiodine and perhapscesium may come from the heating burstduring the thermal transient phase ofthe accident. In exploratory analysesthe release of iodine (i.e., the gap inventory)has been estimated to increaseonly 10 percent in highly rated fuel butup to 300 percent in fuel at normallylow heat rating.Actual release requires rupture of thesheathing and depends as well upon otherinherent retention processes likely tobe in effect in a fuel element (e.g.,iodine and cesium interaction). Unfortunately,only token credit can be affordedthe retention effect until someadditional experimental work or experienceis actually reported.The assessment of the potential inventoryof volatile fission products (raregases and halogens) may be outlined byuse of an accident sequence and coreresponse process as follows:a. Reactor vessel depressurizationrapid loss-of-coolant.b. Core heatup from sensible heat andradioactive decay.c. Rod rupture beginning at 1200-1400 Faccompanied by prompt release ofmost of the gap inventory of raregases, a fraction of the iodine,cesium, etc.d. Continued core heatup prior to turnaroundby emergency coolant. Thisis accompanied by an additional releaseconsisting of a heating burstand a time dependent, slower, thermaltransient diffusional release.The peak temperature achieved may bebetween 1800 and 2200 F.e. Upon cooling of the core, some additionalgaseous release is incurredby fuel pellet breakup.f. Assuming that fuel rod fragmentationis prevented by the safeguard coolant,the remaining contribution maycome only from aqueous leaching ofthe exposed or expelled fuel includingany incurred by a U02-water reaction.g. If the maximum clad temperaturereached is conservatively taken as2200 F (no cladding will melt);however, 100 percent of the rodswill fail. The time elapse of thetransient before turn-around couldbe 5 minutes.Six critical points seem to be involvedin the justification of this method ofevaluation of the loss-of-coolant promptrelease fission product source term:a. The validity of the application ofclassical diffusion theory or theonVII-89

Booth model (Ref. 2) which correlatesthe experimental release datawith temperature, U0 2 density,length of irradiation' time, etc.with the equilibrium fuel gap radionuclideinventory.b. The validity of the higher releaserate or diffusion parameter, for iodinecompared to xenon from U0 2 tothe fuel voids. We cite mainly tworeferences 3,4 using 70 separate determinationsof which 63 give D' ratiosof iodine over xenon rangingfrom 1.5 to 12. The median ratio is6.25 and the square root of the ratio(2.5) is approximately the increasein diffusion rate of iodineover xenon in reactor grade 94 percentdense U0 2 .c. The correlation of burnup with therate of release of individual fissionproduct elements. ORNL diffusionannealing data (Ref. 5) show asignificant burnup effect; however,one cannot assign a quantitative relationfrom it. Therefore, use ofBMI (Ref. 6) and UK (Ref. 7) analysesof the Westinghouse data (Ref.8) has been proposed. These suggestan increase of one order of magnitudein D' for each 15,000 Mwd/T ofaccumulated burnup. This correction,however, may be over-conservativeaccording to a recent study byBailey, et al. (Ref. 9).d. The significance and magnitude ofthe heating and cooling burst contribution.This effect is recognizedby most experimenters as a departurefrom simple diffusion or graingrowth. It is characterized by arapid adjustment of fission productinventory remaining in the UO, forthe transient temperature change,and may be about one-third of theequilibrium difference inventory betweenthe two temperatures.e. Since radioiodine may be retained inthe ruptured rod, by fuel and claddinginteraction, a release coefficientescape fraction must be assignedto each release process. Atpresent, this value is mainly speculative;however, tentative valuesderived by Collins (Ref. 10) andFeuerstein (Ref. 11) may be used.These range from about 0.10 to 0.65depending upon pressure, rupturetemperature, etc.C2. CALCULATION OF PWR AND BWR FUELROD GAP RELEASESC2.1 DIFFUSION COEFFICIENT VALUESIn order to formulate a direct determinationof gap releases of the variousfission products, extensive use was madeof both the experimentally determinedrelative diffusion rates for differentelements from early ORNL work, as wellas an empirically determined absoluterelation between temperature and thediffusion parameter D! (D Prime, Empirical)by Lorenz at ORNL (Ref. 12) basedon a series of Canadian capsule experimentsin which only xenon and kryptonwere measured. The ORNL data, like mostexperimental data, are derived from socalledannealing type experiments whichthen must be critically evaluated forthe initial burst contribution as wasdone for Fig. VII C-1 by Oi and Tagake(Ref. 13) and by Lorenz and Creek (Ref.14).Percent release vs temperature valueswere taken initially from Fig. 6.3 ofORNL Report 3981 (Ref. 3) and the releasevalues were used to compute D'(D-prime) values according to the originalmethod of Booth and Rymer (Ref. 2).where7T n=ln2 22F = fraction released(VII C-l)t = time in seconds at temperatureReliable approximations of Equation (VIIB-l) are possible.1. When F > 0.77F661 -2. When F < 0.77-7r2 D'tF = 6(D't/7) 1 / 2 - 3D't.(VII C-2)(VII C-3)The computed D' values were then plottedas log D' versus 10 4 /T (see Fig. VIIC-2) and least squares fitted to astraight line by the expression:VII-90

orwherein D' = a + B(10 4/T)1B(04D' = Ae S \T/(VII C-4)(VII C-5)T = absolute temperature and ln(A) = aThe constants are listed in Table VIIC-I.With these constants, an expression forD-primes for each element can be obtainedby the use of the relationship inEquation (VII C-5). For example for Xe& Kr,= 232.3 e -4.00534(10 4/T)(VII C-6)Using the data from Lorenz (Ref. 14)(Fig. VII C-l), the relationship used tocompute D-primes would:= 0.206 e-3. 08778(10 4/T)(VII C-7)The D' values obtained by using Equation(VII C-6), as expected, were inordinatelyhigh due to the fact that the "bursteffect"could not be evaluated from thedata from which the equation was derived.The D-primes for the rare gases,based on the corrected data (Fig. VIIC-1) and expressed by Equation (VII C-7)were assumed to be accurate, so that acorrection factor, dividing Equation(VII C-7) by Equation (VII C-6), wasused to adjust all the D-primes derivedfrom Equation (VII C-6). The value ofthe correction factor obtained was8.8793 x 10-4 e 0 .91756(10 4 /T)(VII C-8)Multiplying each D' expression by thisfactor then uniformly corrects all D-primes for the burst contribution asthough each element responded in thesame way. While this is probably acceptablefor the semivolatile cesium andtellurium, it is most unlikely to becorrect for strontium or ruthenium.The diffusion parameters (D's) were alsoadjusted for burnup by a correction, K,as follows:D'=D x, (MWD/DO expo ~.=k250001 ) = DK(VII C-9)where D; = D calculated from the referenceexperimental data,(MWD/T)5,00015,00025,000K0.1580.3981.000The D's derived from the ORNL experimentaldata were empirically normalized tothose for 25,000 MWD/T burnup, by comparisonof the xenon-krypton data fromnumerous other measurements (Refs. 7, 8,9) of the relation between gaseous releaseand burnup at constant temperature.C2.2 RELEASE COMPUTATION PROCEDUREDuring loading, for both types of reactors,the fuel rods are shifted annuallyin approximately the following manner:(1) fresh rods are loaded into the outer1/3 of the reactor, (2) those rods,formerly in the outer 1/3, are shiftedto the middle third of the reactor, and(3) those rods in the center third, areremoved and the rods formerly in themiddle third, are shifted to the centerthird of the reactor.A simple computer code was written,called DIREL, which summarizes thefractional release of twenty principalnuclides using the thermal profilepredicted by the thermal analysis codedescribed later. In addition considerationis taken of the following approximations:a. The maximum fission product inventorythat a reactor may have shouldreoccur at the end of any equilibriumcycle. In a typical third-yearfuel replacement scheme, the averagetime of irradiation would be two cyclesor about 650 days. This valueaffects mainly the longlived or stablenuclides.b. The average burnups for fuel in theouter, middle, and center thirds ofa reactor core are assumed to beabout 5000, 15,000 and 25,000 megawattdays/ton (MWD/T), respectively,at the end of an equilibrium cycle.VII-91

c. Diffusion coefficients for fissionproducts increase by a factor of 10for every 15,000 MWD/T burnup (Ref.6).d. Thermal conductance across the claddinggap is assumed to range between500 and 1000 Btu/hr/ft 2 . The outerone third and half of the centerone-third are assumed to be unrestructuredand to conform to thelower conductance.e. Six reactor volume fractions wereassigned to the 25,000-Mwd/T region,the 7th to the 15,000 Mwd/Tregion, and the 8th to the 5000-Mwd/T region and burnup correctionsare made on the D's accordingly.In the reference LWR reactor core, eightthermal zones are defined with zones 1to 6 inclusive, being in the center onethirdof the reactor; zone 7 is in themiddle one-third and zone 9 is in theouter one-third.For thermal analysis, a fuel rod is dividedinto nine equal lengths. Seventemperature values are given for sevenpoints along the fuel radius of eachlength. In the gap release computerprogram: (1) a D-prime for each activityis calculated for each of the seventemperatures using Equation (VII C-5);(2) F (fraction release) values, atequilibrium, are computed using the D'values according to:F = 3(WD'/X)I/2(VII C-10)When T < 400 C, F = 0 and W = correctionfactors. If X is less than 2 x 10-8,then the isotope is considered to bestable and F is computed according toF = 4(WD't/7)I/2(VII C-ll)where F = fraction released and t = timein seconds; (3) the seven values of "F"are averaged; (4) steps 1 and 2 are repeatedfor the remaining eight lengthsof the fuel rod; (5) the nine "averageF" values are averaged to produce the"fuel rod release"; and (6) the averagerelease is computed for the eight regionsspecified and thenTotal release =8i=*lx PF. :1% releasei x VFiwhere VF = volume fraction and PF =power factor.Initial results of the gap release cal 7culation for the PWR and the BWR systemsare tabulated in Table VII C-2 as afunction of gap conductance (fuel temperature)and possible variation incomputed D' values. From these resultsthe final release estimates for eachreactor type, including effects of thegap release coefficient escape fractionand the heating burst, were obtained asshown in Table VII C-3.C3. THERMAL ANALYSES OF THE:REFERENCE CORESC3.1 THERMAL ANALYSIS OF THE PWRFUEL (NINE CAPSULE POSITIONS)The "Little Mamu" (Ref. 15) program designedfor the IBM 360-91 or 75 computerwas used to calculate the temperaturedistribution within the fuel pin. Thisis a one dimensional (radial) heattransfer program used in many ORNL reactorexperiments and furnished the fuelzone temperature needed for gap releasegas and halogen inventory calculations.In this problem the 12 foot long fuelpin was divided into nine equal capsulepositions of 16 inches. Capsule position1 is at the bottom. The radial andaxial power factors for a reference reactorwere used and the peak heat flux(534,000 Btu/hr/ft 2 ) was used to calculatea volumetric heat generation in thefuel at the central capsule position No.5. It is assumed that all the heat is.generated in the fuel (none in the cladding,gap or coolant). A peak heat generationrate (capsule position 5) wascalculated to be 2.796 watts/gram withinthe U0 2 fuel. The heat rates at theother capsule positions were calculatedby proportion according to the givenaxial power factor distribution. Forcomputer input the fuel pin was dividedinto eight regions at each of the ninecapsule positions in the geometry shownin Fig. VII C-3. In Fig. VII C-3, Region1 represents the Zircaloy-2 claddingwhich was given; (1) a density ofzero (for no heat generation) and (2) alinear thermal conductivity versus temperaturefunction of K = 6.7 + 0.004T.Region 2 represents the gap or a materialwhose density is zero (for no heatgeneration) and whose thermal cnnductivity(K) is equal to the product of thethermal conductance (1000 or 500 Btu/hr-ft 2 -F) and the gap thickness (X +0.0035 inch). Regions 3 through 8 representsix equal volumes of U0 2 fuelwhich were given a thermal conductivityinitially according to CURVE C of Fig.VII C-4. The coolant was given; (1) an(VII C-12ýVII-92

inlet temperature of 543 F, (2) a densityof zero (for no heat generation), (3)a specific heat, flow area, and flowrate. A thermal heat transfer coefficientwas also calculated and input. Nothermal expansion was considered. Allheat is assumed to flow radially to thecoolant by conduction. Given the abovedata, the computer program first calculatesthe heat generated in capsuleposition 1. Half of this heat is addedto the coolant to increase its temperaturefrom Tc inlet to Tcl at the midlengthof capsule position 1 by theequationwhereTcl - T = QTI/2WCPCinletQTI = totalBtu/hrheat in cap. pos. 1,W = flow rate, lb/hrCp= specific heat coolantBtu/lb, FThe computer then calculates the outsidecladding temperature, To, at the midlengthof Capsule Position 1, Btu/hrwhereT -TQTI/(HTC)AsHTC = heat transfer coefficient,Btu/hr, ft, FAs = area surface outside clad,2ftThe thermal gradient across the Zircaloycladding is then calculated by theequationwhereT1 -To = QTI(In R0/RI)/2TLK,Ro and R, are the outside andradii of the claddingL iscapsule length = 16 inchesinsideK is thermal conductivity of thecladding at clad mean temp. (Obtainedby an iterative process).The thermal gradient across the gapthen calculated by the equationT2 - T1 = QT1 (in Rl/R2)/2iTLKiswhereRl and R2 are outside and inside gapradiiK is thermal conductivity of gap -C(Rl - R2), where C is the thermalconductance (1000 or 500 Btu/hrft2F).The thermal gradient in region 3 (theoutside fuel region) is then calculatedby the equationT 3 - T 2 = (q (R22 -R32)/4K)where-(q1 R32 (ln R2/R3)/2K)+ (Q4- 8 (ln R2/R3)/2rrLK)R2 and R3 are the outside and insideradii of region 3K is the thermal conductivity at themean temperature of region 3 accordingto Figure VII C-4, Curve C,Btu/hr-ft-Fq''' is volumetric heat generation inthe fuel, Btu/hr/in 3Q 4 - 8 is the heat generated in regions4 through 8, Btu/hr.The thermal gradients in the remainingfuel regions 4 through 8 are calculatedin the same manner as above.The computer then calculates the coolanttemperature at the midlength of capsuleposition 2. Half the total heat generatedin capsule positions 1 and 2 isadded to the coolant to increase itstemperature to Tc 2 by the equationwhereTc 2 - Tc 1 = (QTl + QT2)/2WCQT2 is the total heat generated incapsule position 2, Btu/hr.The temperature distribution at the variousregion boundaries at each of thesucceeding capsule positions is calculatedin a like manner. Initial computerruns were made and the coolant flowrate was adjusted to 5686 lb/hr to obtainan outlet coolant temp. of 608.3 F.Also the heat transfer coefficient was2made HTC = 6586 Btu/hr - ft - F toattain a peak outside cladding temperatureof To(peak) = 657 F(347.2 C) atVII-93

capsule position 5. Two final computerruns were made, one case with the gapconductance equal to 1000 Btu/hr, ft2 , Fand the other case with the 5ap conductanceequal to 500 Btu/hr, ft2 , F. TableVII C-4 lists the computer calculatedtemperature distribution for the twocases.C3.2 THERMAL ANALYSIS OF THE PWR FUELAT THE PEAK AND AVERAGE HEATFLUXES AND WITH GAP CONDUCTANCESOF 1000 & 500 BTU/HR • FT 2 FFour computer runs were made in thisanalysis, one run or case each for the1000 and 500 gap conductance for thepeak and the average heat fluxes. Onlyone capsule position of unit length isused for these cases. The input wasmade such that the outside cladding temperaturewas 610 F (321.1 C) for the twocases of average heat flux and 657 F(347.2 C) for the two cases of peak heatflux. Heat generation rates in the fuelwere calculated and input for each conditionof heat flux. The geometry anddata (as applicable) of Fig. VII C-3were input and the calculations for thetemperature distribution were made inthe manner described for the PWR ninecapsule position cases.C3.3 THERMAL ANALYSIS OF THE BWR FUEL(NINE CAPSULE POSITIONS)Two cases were run for the BWR fuel (9capsule positions) - one each for gapconductances of 1000 and 500 Btu/hr,-ft 2 ,F.The manner in which temperatures wereattained is the same as that describedfor the PWR fuel except that U0 2 thermalconductivities were taken from CURVE Bof Fig. VII C-4. Figure VII C-5 showsthe geometry and heat transfer data usedfor the BWR fuel. The coolant flow rateand specific heat were made large sothat the coolant temperature (Tc) wouldremain constant at each of the 9 capsulepositions. The inlet coolant temperaturewas made 546.4 F which is the saturationtemperatureoperating pressure.transfer coefficientthe average heat fluxequationandat the 1000 psigA coolant heatwas calculated atposition from the(Q/A)ave = (HTC) tf, where (Q/A)ave= 163,000 Btu/hr - ftAtf = Toave - Tc = 588 - 546.4= 11.6 F.The HTCave equals 14,069 Btu/hr -F. From the relationAtfave/ Atfpeak =[(Q/Aave)/(Q/Apeak)]I/4a Atf peak was obtained to calculate a(HTC)peak = 28,912 Btu/hr,ft 2 ,F forcomputer input. Table VII C-5 lists thecomputer calculated temperature distributionfor the two cases.C3.4 THERMAL ANALYSIS OF THE BWR FUELAT THE PEAK AND AVERAGE HEATFLUXES AND WITH GAP CONDUCTANCESOF 1000 AND 500 BTU/HR • FT 2 -FSame as PWR except BWR fuel and theoutside cladding temperature was made560 F (293.3 C) for the average heatflux conditions and 565.7 F (296.5 C)for the peak heat flux conditions. Thegeometry and data (as applicable) ofFig. VII C-5 were input for the fourcases.2ft2-VII-94

References1. Parker, G. W., R. A. Lorenz, and G. E. Creek, "The Calculation of FissionProduct Release from Core Coolant Safeguarded Reactors", ANS Trans. 12, No. 2,p. 902-3, San Francisco (Nov. 3, 1969).2. Booth, A.sion XenonCouncil ofH. and G. T. Rymer, "Determination of the Diffusion Constant of FisinU0 2 Crystals and Sintered Compacts", CRDC-720, National ResearchCanada (August, 1958).3. Parker, G. W., G.of-Pile StudiesORNL, 1955-1965",E. Creek, W. J. Martin, C. J. Barton, and R. A. Lorenz, "OutofFission-Product Release from Overheated Reactor Fuels atORNL-3981, Oak Ridge National Laboratory (July, 1967).4. Davies, D., G. Long, and W.Products from Uranium Dioxide",(June, 1963).P. Stanaway, "The Emission of Volatile FissionAERE-R-4342, Atomic Energy Research Estab.5. Parker, G. W., G. E. Creek, R. A. Lorenz, and W. J. Martin, "Parametric Studiesof Fission Product Release from U0 2 Fuels", TID-7641, Third Conf. on Nuclear ReactorChemistry, Gatlinburg, Tenn. (1962).6. Morrison, D. L., W. Carbiener, and R. L. Ritzman, "An Evaluation of the Applicabilityof Existing Data to the Analytical Description of a Nuclear-Reactor Accident",BMI Report 1856 (January, 1969).7. Evans, D. M. and A. W. Shilling, "Parameters for Estimating the Release of FissionGases from U0 2 Fuel", TRG Report 1240 (w) (May, 1966).8. Daniels, R. C., et al.,"Effects of High Burnup on Zircaloy-Clad Bulk U0 2Fuel Samples", WAPD-263, Westinghouse Electric Corp. (September, 1962).Plate9. Baily, W. E., C. N. Spalaris, D. W. Sandusky, and E. L. Zebroski, "Effect ofTemperature and Burnup on Fission Gas Release in Mixed Oxide Fuel", Paper presentedat American Ceramic Society Meeting, to be published (May, 1969).10. Collins, R. D., J. T. Hillary and J. C. Taylor, "Air Cleaning for Reactors withVented Containment", UKAEA Report TRG 1318(w) (August, 1966).11. Feuerstein, Horst, "Behavior of Iodine in Zircaloy Capsules", USAEC Report ORNL-4543 (August 1970).12. Parker, G. W. and R. A. Lorenz, "Prompt Release of Fission Products from Zircaloy-CladU0 2 Fuels", in Nuclear Safety Program Annual Progress Report for PeriodEnding December 31, 1967, ORNL-4228, Oak Ridge National Laboratory (April,1968).13. Oi, N., and J. Takagi, Zeit Naturf 20a, 673 (1965).14. Lorenz, R. A., and G. E. Creek, unpublished work supplied by personal communicationat ORNL (1973).15. For a discussion"PreirradiationORNL-TM-3446, p.of the basis for "Little Mamu" calculations, see: Olsen A. R.,Data for ORNL Series II and BNW Oxide Fuel Tests in EBR II",22, Nov. 1971.VII-95

Appendix CCalculation of Gap Release ofRadioactive Fission ProductsbyG. W. ParkerOak Ridge National LaboratoryC1. BACKGROUNDConventional methods for calculatingfission gas release from U0 2 are basedmainly on the Booth diffusion model concept;however, calculating the releaseof radioiodine is slightly more involvedsince the diffusion parameter, D', foriodine often is reported to exceed thatfor xenon by a factor of at least two.In fact, the results of experimentaldiffusion studies, the annealing type,show several fission products (Cs andTe) with higher diffusion coefficients.The effect of extended fuel irradiation(burnup) is also significant and may accountfor more than an order of magnitudein the equilibrium rate of release.Calculations based on conservative assumptionsfor radioiodine release to theclad gap were given in a recent paper(Ref. 1). An important conclusion isdrawn from this paper in that at heatratings below 10 kw/ft only a few percentof the iodine is released whiledoubling the power increases the releaseby a factor of ten.In addition to the steady-state diffusionprocess, a second contribution offission gases, radioiodine and perhapscesium may come from the heating burstduring the thermal transient phase ofthe accident. In exploratory analysesthe release of iodine (i.e., the gap inventory)has been estimated to increaseonly 10 percent in highly rated fuel butup to 300 percent in fuel at normallylow heat rating.Actual release requires rupture of thesheathing and depends as well upon otherinherent retention processes likely tobe in effect in a fuel element (e.g.,iodine and cesium interaction). Unfortunately,only token credit can be affordedthe retention effect until someadditional experimental work or experienceis actually reported.The assessment of the potential inventoryof volatile fission products (raregases and halogens) may be outlined byuse of an accident sequence and coreresponse process as follows:a. Reactor vessel depressurizationrapid loss-of-coolant.b. Core heatup from sensible heat andradioactive decay.c. Rod rupture beginning at 1200-1400 Faccompanied by prompt release ofmost of the gap inventory of raregases, a fraction of the iodine,cesium, etc.d. Continued core heatup prior to turnaroundby emergency coolant. Thisis accompanied by an additional releaseconsisting of a heating burstand a time dependent, slower, thermaltransient diffusional release.The peak temperature achieved may bebetween 1800 and 2200 F.e. Upon cooling of the core, some additionalgaseous release is incurredby fuel pellet breakup.f. Assuming that fuel rod fragmentationis prevented by the safeguard coolant,the remaining contribution maycome only from aqueous leaching ofthe exposed or expelled fuel includingany incurred by a U02-water reaction.g. If the maximum clad temperaturereached is conservatively taken as2200 F (no cladding will melt);however, 100 percent of the rodswill fail. The time elapse of thetransient before turn-around couldbe 5 minutes.Six critical points seem to be involvedin the justification of this method ofevaluation of the loss-of-coolant promptrelease fission product source term:a. The validity of the application ofclassical diffusion theory or theonVII-89

Booth model (Ref. 2) which correlatesthe experimental release datawith temperature, U0 2 density,length of irradiation' time, etc.with the equilibrium fuel gap radionuclideinventory.b. The validity of the higher releaserate or diffusion parameter, for iodinecompared to xenon from U0 2 tothe fuel voids. We cite mainly tworeferences 3,4 using 70 separate determinationsof which 63 give D' ratiosof iodine over xenon rangingfrom 1.5 to 12. The median ratio is6.25 and the square root of the ratio(2.5) is approximately the increasein diffusion rate of iodineover xenon in reactor grade 94 percentdense U0 2 .c. The correlation of burnup with therate of release of individual fissionproduct elements. ORNL diffusionannealing data (Ref. 5) show asignificant burnup effect; however,one cannot assign a quantitative relationfrom it. Therefore, use ofBMI (Ref. 6) and UK (Ref. 7) analysesof the Westinghouse data (Ref.8) has been proposed. These suggestan increase of one order of magnitudein D' for each 15,000 Mwd/T ofaccumulated burnup. This correction,however, may be over-conservativeaccording to a recent study byBailey, et al. (Ref. 9).d. The significance and magnitude ofthe heating and cooling burst contribution.This effect is recognizedby most experimenters as a departurefrom simple diffusion or graingrowth. It is characterized by arapid adjustment of fission productinventory remaining in the UO forthe transient temperature change,and may be about one-third of theequilibrium difference inventory betweenthe two temperatures.e. Since radioiodine may be retained inthe ruptured rod, by fuel and claddinginteraction, a release coefficientescape fraction must be assignedto each release process. Atpresent, this value is mainly speculative;however, tentative valuesderived by Collins (Ref. 10) andFeuerstein (Ref. 11) may be used.These range from about 0.10 to 0.65depending upon pressure, rupturetemperature, etc.C2. CALCULATION OF PWR AND BWR FUELROD GAP RELEASESC2.1 DIFFUSION COEFFICIENT VALUESIn order to formulate a direct determinationof gap releases of the variousfission products, extensive use was madeof both the experimentally determinedrelative diffusion rates for differentelements from early ORNL work, as wellas an empirically determined absoluterelation between temperature and thediffusion parameter D! (D Prime, Empirical)by Lorenz at ORNL (Ref. 12) basedon a series of Canadian capsule experimentsin which only xenon and kryptonwere measured. The ORNL data, like mostexperimental data, are derived from socalledannealing type experiments whichthen must be critically evaluated forthe initial burst contribution as wasdone for Fig. VII C-I by Oi and Tagake(Ref. 13) and by Lorenz and Creek (Ref.14).Percent release vs temperature valueswere taken initially from Fig. 6.3 ofORNL Report 3981 (Ref. 3) and the releasevalues were used to compute D'(D-prime) values according to the originalmethod of Booth and Rymer (Ref. 2).wherea22F = 1 - E -2 e- n2D'tn=lnF = fraction released(VII C-l)t = time in seconds at temperatureReliable approximations of Equation (VIIB-i) are possible.1. When F > 0.7767T2. When F < 0.77F = 6(D't/T)I/2 -3D't.(VII C-2)(VII C-3)The computed D' values were then plottedas log D' versus 10 4 /T (see Fig. VIIC-2) and least squares fitted to astraight line by the expression:VII-90

orin D' = a + B(10 4/T)(VII C-4)adjusted for burnup by a correction, K,as follows:1B(04D' = Ae S \T/(VII C-5)D' = DO exp [2.303 25000 D KwhereT = absolute temperature and ln(A) = aThe constants are listed in Table VIIC-1.With these constants, an expression forD-primes for each element can be obtainedby the use of the relationship inEquation (VII C-5). For example for Xe& Kr,D' =232.3 e~-4.0 0 5 3 4 (10 4 /T)(VII C-6)Using the data from Lorenz (Ref. 14)(Fig. VII C-l), the relationship used tocompute D-primes would:-3.08778(104/T)D' = 0.206 e(VII C-7)The D' values obtained by using Equation(VII C-6), as expected, were inordinatelyhigh due to the fact that the "bursteffect"could not be evaluated from thedata from which the equation was derived.The D-primes for the rare gases,based on the corrected data (Fig. VIIC-l) and expressed by Equation (VII C-7)were assumed to be accurate, so that acorrection factor, dividing Equation(VII C-7) by Equation (VII C-6), wasused to adjust all the D-primes derivedfrom Equation (VII C-6). The value ofthe correction factor obtained was8.8793 x 10-4 e0.91756(10 4 /T)(VII C-8)Multiplying each D' expression by thisfactor then uniformly corrects all D-primes for the burst contribution asthough each element responded in thesame way. While this is probably acceptablefor the semivolatile cesium andtellurium, it is most unlikely to becorrect for strontium or ruthenium.The diffusion parameters (D'S) were also(VII C-9)where D; = D calculated from the referenceexperimental data,( MWD/T)5,00015,00025,000K0.1580.3981.000The D's derived from the ORNL experimentaldata were empirically normalized tothose for 25,000 MWD/T burnup, by comparisonof the xenon-krypton data fromnumerous other measurements (Refs. 7, 8,9) of the relation between gaseous releaseand burnup at constant temperature.C2.2 RELEASE COMPUTATION PROCEDUREDuring loading, for both types of reactors,the fuel rods are shifted annuallyin approximately the following manner:(C) fresh rods are loaded into the outer1/3 of the reactor, (2) those rods,formerly in the outer 1/3, are shiftedto the middle third of the reactor, and(3) those rods in the center third, areremoved and the rods formerly in themiddle third, are shifted to the centerthird of the reactor.A simple computer code was written,called DIREL, which summarizes thefractional release of twenty principalnuclides using the thermal profilepredicted by the thermal analysis codedescribed later. In addition considerationis taken of the following approximations:a. The maximum fission product inventorythat a reactor may have shouldreoccur at the end of any equilibriumcycle. In a typical third-yearfuel replacement scheme, the averagetime of irradiation would be two cyclesor about 650 days. This valueaffects mainly the longlived or stablenuclides.b. The average burnups for fuel in theouter, middle, and center thirds ofa reactor core are assumed to beabout 5000, 15,000 and 25,000 megawattdays/ton (MWD/T), respectively,at the end of an equilibrium cycle.VII-91

c. Diffusion coefficients for fissionproducts increase by a factor of 10for every 15,000 MWD/T burnup (Ref.6).d. Thermal conductance across the claddinggap is assumed to range between500 and 1000 Btu/hr/ft 2 . The outerone third and half of the centerone-third are assumed to be unrestructuredand to conform to thelower conductance.e. Six reactor volume fractions wereassigned to the 25,000-Mwd/T region,the 7th to the 15,000 Mwd/Tregion, and the 8th to the 5000-Mwd/T region and burnup correctionsare made on the D's accordingly.In the reference LWR reactor core, eightthermal zones are defined with zones 1to 6 inclusive, being in the center onethirdof the reactor; zone 7 is in themiddle one-third and zone 9 is in theouter one-third.For thermal analysis, a fuel rod is dividedinto nine equal lengths. Seventemperature values are given for sevenpoints along the fuel radius of eachlength. In the gap release computerprogram: (1) a D-prime for each activityis calculated for each of the seventemperatures using Equation (VII C-5);(2) F (fraction release) values, atequilibrium, are computed using the D'values according to:F = 3(WD'/X)I/2(VII C-10)When T < 400 C, F = 0 and W = correctionfactors. If X is less than 2 x 10-8,then the isotope is considered to bestable and F is computed according toF = 4(WD't/f) 1/2(VII C-11)where F = fraction released and t = timein seconds; (3) the seven values of "F"are averaged; (4) steps 1 and 2 are -repeatedfor the remaining eight lengthsof the fuel rod; (5) the nine "averageF" values are averaged to produce the"fuel rod release"; and (6) the averagerelease is computed for the eight regionsspecified and thenTotal release =il8% release. 1 x VF. 1where VF = volume fraction and PF =power factor.Initial results of the gap release cal7culation for the PWR and the BWR systemsare tabulated in Table VII C-2 as afunction of gap conductance (fuel temperature)and possible variation incomputed D' values. From these resultsthe final release estimates for eachreactor type, including effects of thegap release coefficient escape fractionand the heating burst, were obtained asshown in Table VII C-3.C3. THERMAL ANALYSES OF THEREFERENCE CORESC3.1 THERMAL ANALYSIS OF THE PWRFUEL (NINE CAPSULE POSITIONS)The "Little Mamu" (Ref. 15) program designedfor the IBM 360-91 or 75 computerwas used to calculate the temperaturedistribution within the fuel pin. Thisis a one dimensional (radial) heattransfer program used in many ORNL reactorexperiments and furnished the fuelzone temperature needed for gap releasegas and halogen inventory calculations.In this problem the 12 foot long fuelpin was divided into nine equal capsulepositions of 16 inches. Capsule position1 is at the bottom. The radial andaxial power factors for a reference reactorwere used and the peak heat flux(534,000 Btu/hr/ft 2 ) was used to calculatea volumetric heat generation in thefuel at the central capsule position No.5. It is assumed that all the heat is.generated in the fuel (none in the cladding,gap or coolant). A peak heat generationrate (capsule position 5) wascalculated to be 2.796 watts/gram withinthe U0 2 fuel. The heat rates at theother capsule positions were calculatedby proportion according to the givenaxial power factor distribution. Forcomputer input the fuel pin was dividedinto eight regions at each of the ninecapsule positions in the geometry shownin Fig. VII C-3. In Fig. VII C-3, Region1 represents the Zircaloy-2 claddingwhich was given; (1) a density ofzero (for no heat generation) and (2) alinear thermal conductivity versus temperaturefunction of K = 6.7 + 0.004T.Region 2 represents the gap or a materialwhose density is zero (for no heatgeneration) and whose thermal cnnductivity(K) is equal to the product of thethermal conductance (.1000 or 500 Btu/hr-ft 2 -F) and the gap thickness (X +0.0035 inch). Regions 3 through 8 representsix-equal volumes of U0 2 fuelwhich were given a thermal conductivityinitially according to CURVE C of Fig.VII C-4. The coolant was given; (1) an4.x PF. 1 (VII C-12,VII-92

inlet temperature of 543 F, (2) a densityof zero (for no heat generation), (3)a specific heat, flow area, and flowrate. A thermal heat transfer coefficientwas also calculated and input. Nothermal expansion was considered. Allheat is assumed to flow radially to thecoolant by conduction. Given the abovedata, the computer program first calculatesthe heat generated in capsuleposition 1. Half of this heat is addedto the coolant to increase its temperaturefrom Tc inlet to Tcl at the midlengthof capsule position 1 by theequationwhereTcl -T ci Cinlet = Q Tl /2WC PQTI = totalBtu/hrheat in cap. pos. 1,W = flow rate, lb/hrCp = specific heat coolantBtu/lb, FThe computer then calculates the outsidecladding temperature, To, at the midlengthof Capsule Position 1, Btu/hrwhereTO -Tc = QTI/(HTC)AsHTC = heat transfer coefficient,Btu/hr, ft, FAs = area surface outside clad,2ftThe thermal gradient across the Zircaloycladding is then calculated by theequationwhereT1 -To = QTI(ln RO/Rl)/27rLK,Ro and R, are the outside andradii of the claddingL iscapsule length = 16 inchesinsideK is thermal conductivity of thecladding at clad mean temp. (Obtainedby an iterative process).The thermal gradient across the gapthen calculated by the equationT2 - T 1 = QTI (in Rl/R2)/2irLKiswhereR1 and R2 are outside and inside gapradiiK is thermal conductivity ofC(Rl - R2), where C is theconductance (1000 or 5002ft F).gap -thermalBtu/hrThe thermal gradient in region 3 (theoutside fuel region) is then calculatedby the equationT 3 - T 2 = (q (R22 -R3 2)/4K)where-(q1 R32 (in R2/R3)/2K)+ (Q4- 8 (in R2/R3)/27rLK)R2 and R3 are the outside and insideradii of region 3K is the thermal conductivity at themean temperature of region 3 accordingto Figure VII C-4, Curve C,Btu/hr-ft-Fq''' is volumetric heat generation inthe fuel, Btu/hr/in 3Q4-8 is the heat generated in regions4 through 8, Btu/hr.The thermal gradients in the remainingfuel regions 4 through 8 are calculatedin the same manner as above.The computer then calculates the coolanttemperature at the midlength of capsuleposition 2. Half the total heat generatedin capsule positions 1 and 2 isadded to the coolant to increase itstemperature to Tc 2 by the equationwhereTc 2 - Tc 1 = (QTl + QT2)/2WCQT2 is the total heat generated incapsule position 2, Btu/hr.The temperature distribution at the variousregion boundaries at each of thesucceeding capsule positions is calculatedin a like manner. Initial computerruns were made and the coolant flowrate was adjusted to 5686 lb/hr to obtainan outlet coolant temp. of 608.3 F.Also the heat transfer coefficient was2made HTC = 6586 Btu/hr - ft - F toattain a peak outside cladding temperatureof To(peak) = 657 F(347.2 C) atVII-93

capsule position 5. Two final computerruns were made, one case with the gapconductance equal to 1000 Btu/hr, ft2 , Fand the other case with the gap conductanceequal to 500 Btu/hr, ft , F. TableVII C-4 lists the computer calculated'temperature distribution for the twocases.C3.2 THERMAL ANALYSIS OF THE PWR FUELAT THE PEAK AND AVERAGE HEATFLUXES AND WITH GAP CONDUCTANCESOF 1000 & 500 BTU/HR • FT 2 FFour computer runs were made in thisanalysis, one run or case each for the1000 and 500 gap conductance for thepeak and the average heat fluxes. Onlyone capsule position of unit length isused for these cases. The input wasmade such that the outside cladding temperaturewas 610 F (321.1 C) for the twocases of average heat flux and 657 F(347.2 C) for the two cases of peak heatflux. Heat generation rates in the fuelwere calculated and input for each conditionof heat flux. The geometry anddata (as applicable) of Fig. VII C-3were input and the calculations for thetemperature distribution were made inthe manner described for the PWR ninecapsule position cases.C3.3 THERMAL ANALYSIS OF THE BWR FUEL(NINE CAPSULE POSITIONS)Two cases were run for the BWR fuel (9capsule positions) - one each for gapconductances of 1000 and 500 Btu/hr,-ft2 ,F.The manner in which temperatures wereattained is the same as that describedfor the PWR fuel except that U0 2 thermalconductivities were taken from CURVE Bof Fig. VII C-4. Figure VII C-5 showsthe geometry and heat transfer data usedfor the BWR fuel. The coolant flow rateand specific heat were made large sothat the coolant temperature (Tc) wouldremain constant at each of the 9 capsulepositions. The inlet coolant temperaturewas made 546.4 F which is the saturationtemperatureoperating pressure.transfer coefficientthe average heat fluxequationandat the 1000 psigA coolant heatwas calculated atposition from the(Q/A) ave = (HTC) tf, where (Q/A) ave= 163,000 Btu/hr - ftAtf = Toave - Tc = 588 - 546.4= 11.6 F.The HTCave equals 14,069 Btu/hr -F. From the relationAtfave/ Atfpeak = [(Q/Aave)/(Q/Apeak)a Atf peak was obtained to calculate a(HTC) eak = 28,912 Btu/hr,ft 2 ,F forcomputer input. Table VII C-5 lists thecomputer calculated temperature distributionfor the two cases.C3.4 THERMAL ANALYSI OF THE BWR FUELAT THE PEAK AND AVERAGE HEATFLUXES AND WITH GAP CONDUCTANCESOF 1000 AND 500 BTU/HR • FT 2 •FSame as PWR except BWR fuel and theoutside cladding temperature was made560 F (293.3 C) for the average heatflux conditions and 565.7 F (296.5 C)for the peak heat flux conditions. Thegeometry and data (as applicable) ofFig. VII C-5 were input for the fourcases.2ft2VII-94

References1. Parker, G. W., R. A. Lorenz, and G. E. Creek, "The Calculation of FissionProduct Release from Core Coolant Safeguarded Reactors", ANS Trans. 12, No. 2,p. 902-3, San Francisco (Nov. 3, 1969).2. Booth, A.sion XenonCouncil ofH. and G. T. Rymer, "Determination of the Diffusion Constant of FisinU0 2 Crystals and Sintered Compacts", CRDC-720, National ResearchCanada (August, 1958).3. Parker, G. W., G.of-Pile StudiesORNL, 1955-1965",E. Creek, W. J. Martin, C. J. Barton, and R. A. Lorenz, "OutofFission-Product Release from Overheated Reactor Fuels atORNL-3981, Oak Ridge National Laboratory (July, 1967).4. Davies, D., G. Long, and W.Products from Uranium Dioxide",(June, 1963).P. Stanaway, "The Emission of Volatile FissionAERE-R-4342, Atomic Energy Research Estab.5. Parker, G. W., G. E. Creek, R. A. Lorenz, and W. J. Martin, "Parametric Studiesof Fission Product Release from U0 2 Fuels", TID-7641, Third Conf. on Nuclear ReactorChemistry, Gatlinburg, Tenn. (1962).6. Morrison, D. L., W. Carbiener, and R. L. Ritzman, "An Evaluation of the Applicabilityof Existing Data to the Analytical Description of a Nuclear-Reactor Accident",BMI Report 1856 (January, 1969).7. Evans, D. M. and A. W. Shilling, "Parameters for Estimating the Release of FissionGases from U0 2 Fuel", TRG Report 1240 (w) (May, 1966).8. Daniels, R. C., et al.,"Effects of High Burnup on Zircaloy-Clad Bulk U02Fuel Samples", WAPD-263, Westinghouse Electric Corp. (September, 1962).Plate9. Baily, W. E., C. N. Spalaris, D. W. Sandusky, and E. L. Zebroski, "Effect ofTemperature and Burnup on Fission Gas Release in Mixed Oxide Fuel", Paper presentedat American Ceramic Society Meeting, to be published (May, 1969).10. Collins, R. D., J. T. Hillary and J. C. Taylor, "Air Cleaning for Reactors withVented Containment", UKAEA Report TRG 1318(w) (August, 1966).11. Feuerstein, Horst, "Behavior of Iodine in Zircaloy Capsules", USAEC Report ORNL-4543 (August 1970).12. Parker, G. W. and R. A. Lorenz, "Prompt Release of Fission Products from Zircaloy-CladU0 2 Fuels", in Nuclear Safety Program Annual Progress Report for PeriodEnding December 31, 1967, ORNL-4228, Oak Ridge National Laboratory (April,1968).13. Oi, N., and J. Takagi, Zeit Naturf 20a, 673 (1965).14. Lorenz, R. A., and G. E. Creek, unpublished work supplied by personal communicationat ORNL (1973).15. For a discussion"PreirradiationORNL-TM-3446, p.of the basis for "Little Mamu" calculations, see: Olsen A. R.,Data for ORNL Series II and BNW Oxide Fuel Tests in EBR II",22, Nov. 1971.VII-95

TABLE V1I C-1VALUES FOR CONSTANTS IN THE ARRHENIUS EQUATIONRadioelementXenon-kryptonIodineCesiumStrontiumRutheniumTelluriumConstant A232.32.9361.3693.37 x 10153.849553.27Constant B-4.00534-2.95177-2.62966-11.2368-7.01379-3.3510TABLE VII C-2INITIAL GAP RELEASE RESULTSReactor PWR PWR PWR PWR BWR BWR BWR BWRGap Conductance 1000 1000 500 500 1000 1000 500 500Conditions (1) (2) (i) (2) (1) (2) (1) (2)Xe-133 0.75 2.0 2.6 5.9 0.62 1.7 2.0 4.7Xe-135 0.21 0.63 0.76 2.2 0.17 0.52 0.58 1.7Kr-85 3.4 6.2 9.4 14.2 2.9 5.4 7.6 11.61-129 7.3 12.8 16.0 23.8 6.5 11.2 13.5 20.11-131 1.8 4.7 4.9 10.8 1.6 4.0 3.9 8.81-132 0.2 0.58 0.54 1.6 0.17 0.5 0.43 1.31-133 0.59 1.75 1.6 4.6 0.51 1.5 1.3 3.7Cs-137 .12.1 19.4 23.3 32.8 10.9 17.2 20.2 28.4Cs-138 0.17 0.49 0.42 1.2 0.15 0,42 0.34 0.96Sr-89 0.42 0.74 2.2 3.1 0.32 0.57 1.6 2.3Sr-90 0.63 0.88 2.6 3.8 0.45 0.74 2.0 2.9Sr-91 0.12 0.24 0.83 1.3 0.08 0.19 0.54 0.96Conditions:(1) Used one tenth (1/10) of computed D-primes.(2) Used burn-up factors on computed D-primes.

TABLE VI1 C-3FINAL ESTIMATE OF GAP RELEASE VALUES(a)(b)Xe 1 3 3Xe 1 3 5Kr 8 5i1 2 9i131i132i133Cs 1 3 7Cs138SrSr8 99 0Sr 9 1Xe 1 3 3Xe1 3 5Kr 8 5i129131132133Cs 1 3 7Cs 1 3 8Sr 8 9SrSr9 09 1Percentin Gap2.30.77.814.44.80.61.721.30.51.51.71.01.90.66.512.44.00.51.417.10.41.11.40.4Noble gases and halogeneffect" for the thermalGap release coefficientthe main text.Gap Release Percent(a) ReleasedCoefficient (b)from RodsSummary of PWR Data1.03.11.01.01.010.50.59.60.53.20.50.40.51.20.511.00.50.250- 41.5 x 10-410-41.7 x 10-4.10- 41.0 x 10-4Summary of BWR Data1.02.51.00.81.08.70.58.30.52.70.50.40.50.90.58.60.50.210-41.1 x 10-410-41.4 x 10-410- 4 0.4 x 10-4releases have been increased by one-third for "burstreleases have been increased by one-third for "bursttransient following loss of coolant.corresponds to the escape fraction terminology used inTable VII C-I - Table VII C-3VII-97/98

TABLE VII C-4RESULTS OF THE NINE CAPSULE POSITION CALCULATIONS AT PEAK RADIAL POWERPOSITION IN PWR COREPosition TC TO Ti T2 T3 T4 T5 T6 T7 T8Gap Conductance -21000 Btu/hr, ft , F1 285.0 301.1 325.9 446.8 508.3 573.3 642.5 716.8 797.6 885.02 287.7 312.6 350.4 537.6 642.8 760.4 894.9 1037.9 1191.3 1357.93 291.6 326.6 378.8 641.2 810.5 1005.8 1220.9 1463.4 1725.1 1975.24 296.6 336.7 396.1 697.2 909.4 1146.3 1415.4 1713.3 1999.8 2254.15 302.3 347.3 413.1 750.8 998.7 1278.7 1601.6 1931.6 2223.5 2471.06 307.9 347.4 405.4 702.0 911.8 1145.2 1409.7 1702.5 1985.7 2237.97 312.7 346.8 396.9 652.5 819.1 1009.9 1219.5 1455.0 1709.0 1954.08 316.6 340.9 377.1 559.6 664.5 781.9 914.2 1054.9 1205.8 1369.59 319.2 333.8 355.7 465.2 521.7 581.1 644.1 711.2 783.6 862.12Gap Conductance = 500 Btu/hr, ft2, F1 285.0 301.1 325.9 567.7 636.5 710.4 790.5 877.7 968.3 1062.92 287.7 312.6 350.4 724.7 854.3 994.6 1144.7 1307.1 1483.9 1669.73 291.6 326.6 378.8 903.6 1107.8 1335.0 1587.4 1847.6 2087.2 2301.84 296.6 336.7 396.1 998.4 1246.4 1530.5 1832.5 2107.1 2348.0 2553.95 302.2 347.3 413.1 1088.4 1383.0 1715.7 2035.2 2314.0 2547.9 2752.36 307.9 347.4 405.4 998.6 1242.6 1522.0 1820.4 2092.6 2332.6 2537.57 312.7 346.8 396.9 908.0 1107.1 1327.9 1573.0 1827.7 2063.6 2275.18 316.6 340.9 377.1 742.2 871.2 1009.1 1156.5 1315.8 1488.6 1669.99 319.2 333.8 355.7 574.8 637.3 704.0 775.7 854.0 935.0 1019.1TC = temperature of coolant, C.TO = temperature of outside surface of Region 1, C.T9 = temperature of inside surface of Region X, C.TABLE VII C-5RESULTS OF THE NINE CAPSULE POSITION CALCULATIONS AT PEAK RADIAL POWERPOSITION IN BWR COREPosition TC TO T1 T2 T3 T4 T5 T6 T7 T8Gap Conductance = 1000 Btu/hr,ft2, F1 285.8 289.6 316.2 412.5 476.1 542.2 611.2 683.5 759.5 840.02 285.8 291.7 332.5 481.5 585.4 696.7 817.3 950.2 1099.8 1267.83 285.8 294.1 350.8 559.8 715.6 890.6 1094.6 1334.6 1610.9 1918.74 285.8 295.3 360.2 600.1 785.7 1001.2 1263.6 1573.7 1925.5 2297.75 285.8 296.5 368.9 637.9 853.9 1114.5 1435.0 1814.7 2230.8 2634.36 285.8 295.2 359.1 595.3 777.3 987.7 1243.7 1546.2 1889.6 2255.778285.8285.8293.9291.5349.2331.4552.7.476.8703.5577.7872.0685.61066.7802.21295.8930.01559.91073.01855.01234.09 285.8 289.3 313.4 400.6 457.7 516.8 578.1 641.9 708.6 778.5Gap Conductance -500 Btu/hr,ft2,F12285.8285.8289.6291.7316.2332.5508.9630.6576.4745.5646.9870.7721.01009.8799.11168.7882.11344.2971.01538.73 285.8 294.1 350.8 768.9 951.7 1168.4 1420.6 1707.8 2024.6 2349.64 285.8 295.3 360.2 840.0 1066.3 1339.9 1660.7 2021.7 2394.8 2747.156285.8285.8296.5295.2268.9359.1906.9831.61181.81052.51518.11319.11908.61631.82326.11983.92723.82351.13099.02700.578285.8.285.8293.9291.5349.2331.4756.3622.2932.2733.51138.5854.41378.6988.01652.61139.31955.71307.12271.51493.69 285.8 289.3 313.4 487.8 548.0 610.6 675.9 744.2 816.0 892.0TC = temperatureTO = temperatureTX = temperatureof coolant, C.of outside surface of Region 1, C.of inside surface of Region X, C.Table VII C-4 - Table VII C-5VII-99/100

Temperature,CaI~ "-- Stable Fission Gas10% in 2.27doys- 90%n 3 Yearszo-79%in3 years0 10% in 22.7 days-50%in 3years~ 30% in3 years- 1-0%in 227daysI in 2.27day,10%.. 2270 days1%/ in 22.7 days-Iin 1%227 days!0,O000/T N°)0 1 31 1-25,000Mwsi/TI/Xe: 2.5131-Iow bansapL/Xe: 2.5FIGURE VII C-1Diffusion ParameterD' Derived fromORNL - Xenon AnnealingData with ComparisonData of 01 andTAGAKI and AECL0.,0000 0TE MP E RATU R EC8 100550.0I010-ttO4 510'/T(K)FIGURE VII C-2Computed D' Valuesfor Fission Products

(0 /A")e = 163,000 Btu/hr-ft 2(Q/A)0 pk=425,000 Btu/hr-ft6To(vel=560 F 1STc uit= 546.4 FGp Cond xx= 0.005 inch•-Regiona-HTC-2t990.5Btu/hr-tftCoolant Flow Rate= 10' lb/hrArea Coolant Channel = 0. 2966 inch 2FFIGURE VII C-3Section of PWR Fuel Rod ShowingGeometry and Thermal Data Usedfor Temperature Analysis22.4 ~ - 4 lO" O t T18* I.32 100 4 e 88,'I ýT 1~'650 CK O'0Cl9 -132 - 0- e ' asO" 02! . 55-,CC

TC(Q/A)ove=I 9 1,000 8tu/hr-ft'0QAPI3 =5,0008Buhr-lTcout=608.3Tcinilzt5 4 3FFToo~)610 311C-Coolant Flow Rate z 5686 lb/hrAreo Coolant Channel =0.1771 inch 2FIGURE VII C-5Section of BWR Fuel Rod ShowingGeometry and Thermal Data Usedfor Temperature AnalysisFig. VII C-1 - Fig. VII C-5VII-101/102

APPENDIX DSURVEY OF EXPERIMENTAL WORK ON FISSION PRODUCTRELEASE FROM MOLTEN U0 2byR. L. RitzmanBattelle, Columbus LaboratoriesandG. W. ParkerOak Ridge National LaboratoryVII-103

Appendix DTable of ContentsSectionDl. OUT-OF-PILE MELTING STUDIES ......................................D1.1 ARC-Image Furnace Experiments ..............................D1.2 Tunsgten Crucible Melting Experiments ......................D1.3 Tungsten Resistor Melting Experiments ......................D1.4 Melting Experiments at the NSPP ............................D1.5 Melting Experiments at the CDE .............................D2. IN-PILE MELTING STUDIES ..........................................D3. SUMMARY AND CONCLUSIONS ..........................................REFERENCES ..............................................................Page No.VII-107VII-107VII-108VII-108VII-109VII-109VII-II0VII-lllVII-112List of TablesTableVII D-1 Fission Product Release from U0 2 Melted in ImpureHelium ..........................................................VII D-2 Fission Product Release from U0 2 Melted in PureHelium by the Tungsten-Crucible Method ..........................Page No.VII-113/114VII-113/114VII D-3Fission Product Release from Trace-Irradiated PWR-TypeU0 2 Melted in a Single Element Tungsten-Resistor FurnaceFilled with Helium ..............................................VII-113/114VII D-4 Summary of Fission Product Releases Observed in NSPP U02Melting Experiments .............................................VII D-5 Fission Product Release from Heated Zircaloy-Clad U02 . . . . . . . . . . .VII-113/114VII-115/116VII D-6Results of ORNL Experiments on Fission-Product Release DuringIn-Pile Melting of Stainless-Steel Clad U02 .....................VII-115/116VII-105

Appendix DSurvey of Experimental Work on -Fission ProductRelease From Molten U0 2by'R. L. RitzinanBattelle, Columbus LaboratoriesandG. W. ParkerOak Ridge National LaboratoryExperimental work on the release of anumber of important fission productsduring melting Of U102 samples has b~eenperformed at several laboratories.Numerous factors which can affect releasemagnitudes or rates have beenexamined to various degrees, but anextensive data base for use in predictingrelease during a reactor core meltdowndoes not exist. This is mainlybecause no moderate or large amounts ofirradiated U102 have been melted, particularlyin a reactor core configuration.The studies that have been done withsmall fuel samples do, however, provideuseful insights to the general releasebehavior of the more important fissionproducts. The purpose of this survey isto summarize the more pertinent findingsof these studies without developing acomprehensive review of all the workthat has been done. Emphasis will beplaced on the general conclusions thatcan be drawn from the data as a wholeregarding release magnitudes and significantinteractions which. should affectthe magnitudes.Dl.' OUT-OF-PILE MELTING STUDIESMost of the experimental work of thistype has been performed at Oak RidgeNational Laboratory (ORNL), but someadditional useful data have been obtainedat the Contamination-DecontaminationExperiment (CDE). Much of the OBNL workconsisted of melting . small samples ofirradiated U0 2 using different furnaceheating techniques. The work has beendescribed in a report by Parker, et al.(Ref. 1). Limited investigation ofparameters which can affect fission productrelease was made including burnup,sample size, time -molten, atmospherecomposition, melting method, and type ofcladding. 'Most of the melting experimentswere performed in helium becauseof reactivity of container or heatermaterials with oxygen, but a few experimentswere performed with CO 2 and air.D1.l ARC-IMAGE FURNACE EXPERIMENTSThe first melting experiments wereconducted with small (0.2 to 0.6 gram)trace-irradiated and unclad 1102 specimensheld in a BeO support tube. Theentire assembly was melted (sometimesincompletely) in an arc-image furnacefor periods of 1-1/2 to 3 minutes.Following melting the release of raregases, iodine, tellurium, cesium, ruthenium,strontium, barium, and rare earthelements was measured by radiochemicalanalysis. Release was defined as thepercentage of each specie that escapedthe U0 2 -BeO assembly. The helium covergas used in the experiment containedsmall amounts of air as an impurity.Table VII D-1 lists the results of thefission product release measurements forthese experiments. The rare gas releasevalues should be considered indicativeof the degree of melting that wasachieved. In general, the data fromthese small specimens indicate two rangesof release values; one set of highreleases which includes rare gases, iodine,tellurium, cesium, and ruthenium,and another set of low releases whichinclude strontium, barium, and rareearths. It is possible that the airimpurity in the helium cover gas influencedsome of the release values.Strontium, the rare earth elements, andto some extent barium, are known to formquite stable oxides with volatilitiessignificantly less than the respectivemetals (Ref. 2 and 3). Such reactions*may have contributed to the low releasesobserved for these fission products. onthe other hand, ruthenium can form volatileoxides (Ref. 2), and this may havecaused the observed high release values.The releases of the remaining fissionproducts would not be expected to beparticularly sensitive to the presenceof air impurities, since at 110? meltingtemperatures both the oxides, if formed,and the elements exhibit high volatility.VII-107

Limited additional experiments of thistype were performed in air and carbondioxide and impure helium atmospheresusing U0 2 having burnup levels up to11,000 Mwd/T. All fission product releasedata were very similar to theresults described above; i.e., eitherhigh or low release for the identicalsets of fission products. A weak positivecorrelation of release with burnuplevel was suggested, but the effect alsocorrelates with differences in samplesizes. No appreciable effect of thedifferent cover gases used on the extentof release was indicated. This wasprobably because each of the atmospherescreated an essentially oxidizing environment.Generally, the data indicatethat burnup level should be consideredno more than a secondary factor and thatair and CO 2 have about the same effecton fission product release from moltenU0 2 .D1.2 TUNGSTEN CRUCIBLE MELTINGEXPERIMENTSThe next series of experiments consistedof trace-irradiated U0 2 specimens thatwere melted by induction in tungstencrucibles. The results of these experimentsare more useful for two reasons:(1) the larger samples, about 29 grams,are comparable to actual reactor fuelpellet sizes, and (2) an essentiallyinert furnace atmosphere was maintainedso that release data are indicative ofnon-oxidizing conditions. The data obtainedfrom these experiments are givenin Table VII D-2. Comparison with TableVII D-1 shows that, generally, the samerelease results were obtained by the twomelting techniques except for the fissionproduct ruthenium. As noted earlier,ruthenium is known to be quite oxygensensitive and it is quite possiblethat its higher release on melting inthe arc-image furnace c-an be attributedto traces of oxygen in the helium supplyand the absence of the good oxygen getter(tungsten) that was available in thelater experiments with larger samples.Note also in Table VII D-2 that the certain(rare earths) release values approximatelycorrespond to the percentU0 2 that vaporized. This observation isconsistent with the known refractorynature of the rare earth oxides. Thetime that the U0 2 remained molten inthese experiments varied to some extentand a weak correlation may exist betweenrelease and time molten for the morevolatile elements. Nevertheless, thedata lead to the conclusion that themelting of bare U0 2 fuel pellets willcause rapid land high release of the morevolatile fission products (rare gases,iodine, tellurium, and cesium). Therelease of ruthenium is very sensitiveto the oxidizing nature of the system.Strontium and rare earth release ratesshould be low and comparable to U0 2vaporization rates while barium releaserates may be somewhat higher (probablydue to the lesser stability or highervolatility of its oxide).D1.3 TUNGSTEN RESISTOR MELTINGEXPERIMENTSThe third and final type of laboratoryscale out-of-pile melting experimentinvolved use of a tungsten rod resistorheating element which was passed throughcored U0 2 pellets. Both clad and uncladelements were employed in these experiments,but they were limited to a heliumatmosphere and complete melting of specimenscould not be accomplished beforethe tungsten rods melted. Nevertheless,the results are quite useful because thehigh interior fuel temperature and coolersurface achieved with this heatingmethod more nearly simulate nuclearheating than any other out-of-pile technique.The release data obtained fromthe melting of trace-irradiated U0 2 aregiven in Table VII D-3. The resultsfrom the unclad specimens, after adjustingfor the fraction melted, are verysimilar to the results of the othermelting methods with respect to raregases, iodine, tellurium, cesium, andcerium (rare earths). The releases ofstrontium and barium are somewhat higherand the release of ruthenium is againlow, which suggests that the oxygenactivity in the system was even lowerthan in the tungsten crucible meltingexperiments. The release from thestainless-steel clad specimen was quitesimilar to that from unclad fuel.However, it is evident that the zirconiumclad specimens gave very differentresults for some fission products. Thereleases of strontium and barium weresignificantly higher while tellurium releaseexperienced a sharp decrease.Post-melting examinations indicated thatthe molten zirconium had wet the U0 2 andspread over the surface. Since zirconiumis known to have a high affinity foroxygen, this behavior probably causedthe clad to serve as a very effectiveoxygen getter. This led to a very lowoxygen activity in the system causingconversion of both strontium and bariumto the more volatile metallic forms.The low tellurium release can apparentlybe explained on the basis of reactionwith and retention by the molten zirconium.This conclusion is supported byindependent results obtained by Genco,et al. (Ref. 4) which shows that tellu--rium vapor reacts extensively with zir-VII-1 08

conium metal at temperatures above about400 C. Since tellurium belongs to thesame periodic group as oxygen, it is nottoo surprising that stable zirconiumtellurides could exist at high temperatures.The same considerations providesome insight regarding the lack of effectof stainless steel cladding onfission product releases. Iron, themajor component of stainless steel, hasa much lower affinity for oxygen (lessnegative free energy of formation) (Ref.5) than does zirconium. Consequently,it should not have a strong effect onthe oxygen activity in the system (whichcan influence strontium, barium, andperhaps ruthenium volatility), and byanalogy should not tend to form thermallystable tellurides either.D1.4 MELTING EXPERIMENTS. AT THE NSPPA different series of U0 2 melting experimentswere performed at ORNL as part ofthe Nuclear Safety Pilot Plant (NSPP)program (Refs. 6-9). Samples of cladU0 2 were passed under a plasma torch indifferent atmospheres to generate anaerosol source for the study of containmentbehavior processes in an adjacentmodel containment vessel. The experimentswere characterized by many differingconditions and material balanceswere poor in some instances but the resultsprovide some useful general observations.The release data for theseexperiments are summarized in Table VIID-4. Two sets of release percentagesfor various fission products are listed;one for release from the immediate fuelregion and another for release from thesteel transfer line to the model containmentvessel. Looking at the firstfive experiments (stainless-steel cladding)the data in general indicate highreleases from molten fuel for iodine,tellurium, and cesium with much lowerreleases for the other species. Rutheniumbehavior was erratic and probablyindicative of the oxygen level in thefurnace. None of the release resultsappear to depend on changes in furnaceatmosphere except ruthenium. The lattertwo experiments with Zircaloy clad highburnup U0 2 are more difficult to interpret.The difference in degree of meltingprobably masks most atmosphere effectsexcept for ruthenium. Consistentwith other data the ruthenium releasewas relatively high in an air atmosphere.The low ruthenium release insteam could be attributed to a morereducing environment resulting fromhydrogen produced by the zirconium-waterreaction that would take place betweencladding and steam. Also consistentwith the out-of-pile melting work describedearlier, iodine and cesiumreleases remain high and apparentlyindependent of atmosphere. The bariumrelease in Experiment 14 and strontium,barium, and cerium releases in Experiment15 appear somewhat anomalous. Theoxygen gettering ability of Zircaloyleading to volatilization of strontiumand barium metal could possibly explainthe high releases in Experiment 15, butcerium should not be strongly affectedand the ruthenium release is not consistentwith this hypothesis. The explanationis unknown and much more informationabout exact conditions in thefurnace during melting would be neededin order to analyze the data further.The second set of release numbers inTable VII D-4 provides an indication ofthe effectiveness with which fissionproducts released from the fuel regionwere transported to the MCV. Eventhough temperatures in the transfer linewere low (probably several hundred degrees)the transit time was short. Ingeneral, plateout factors for all speciesduring this transfer were about thesame and ranged from factors of about 2to factors of about 10. Thus, plateoutin the transfer line did not cause amajor change in the character of thefission product source that was transportedto the model containment vessel.D1.5 MELTING EXPERIMENTS AT THE CDEAnother short series of out-of-pile experimentsin which fission product releasefrom molten U0 2 was measured wasconducted at the CDE by Freeby, et al.(Ref. 10). These consisted of five runsin which Zircaloy clad U0 2 pellets weremelted by induction in steam atmospheres.The fuel weight ranged from 70to 80 grams and molten times of about 30minutes were achieved in each case.Fuel burnups among the five samplesranged from about 500 to 2000 Mwd/T.Some difficulties in achieving materialbalances were encountered, particularly.for iodine and tellurium. However,these experiments are specifically relevantto the conditions that would existduring reactor core melting. The combinedresults of the five experimentsare given in Table VII D-5 along withother pertinent data that investigatorsobtained from reports of related work.The CDE results generally agree withother out-of-pile release data in thathigh volatility is indicated for thenoble gases, iodine, tellurium, and cesium,and relatively low volatility forstrontium, barium, and ruthenium. However,contrary to the ORNL results, therelease data do not indicate a dominanteffect of the Zircaloy cladding on telluriumor strontium and barium release.VII-109

This may be due to the steam atmospherein the CDE experiments which caused extensiveoxidation of the Zircaloy cladding(possibly liberating some retainedtellurium) and perhaps producing a highenough oxygen activity in the system tohold strontium and barium as their oxides.The data listed in the tableunder CMF were the result of a singlemelting experiment under a steam-airatmosphere in that facility at ORNL.The data continue to support the earlierobservations at ORNL concerning the effectof zirconium cladding on telluriumand strontium release. Therefore, therelease behavior of tellurium and strontiumand barium during melting of Zircaloyclad U02 is somewhat uncertain. Acomparison of the data obtained in thedifferent experiments identified inTable VII D-5 shows a possible trendtoward increase in iodine release withincreasing fuel burnup. However, differencesin specimen temperatures ormelting periods probably influence theresults such that the indicated burnupeffect is quite indefinite.In the CDE experiments efforts were madeto record the plateout of released fissionproducts that occurred within ashort stainless steel transfer line to acontainment vessel. The temperature ofthe line was maintained at about 800 Fand transit times ranged from about oneto two seconds. In general, it wasfound that less than 1/3 of the enteringfission products deposited in the lineexcept for cerium which was nearly alldeposited. The results roughly agreewith experience at the NSPP and indicatethat longer residence times orhigher surface areas would be needed topromote efficient plateout. Anothersignificant observation was recordedduring the CDE melting experiments.Even though the small fuel pins werekept molten for 30 minutes or so, mostof the fission product release tookplace within a one- to two-minute periodrelatively early in the experiments.This rapid release coincided with meltthroughof the Zircaloy cladding (pinlying on its side in U02 powder). Subsequentlylarge quantities of particulates(cladding and fuel) were releasedcreating a white smoke in the furnace.A continuous slow release of fissionproducts occurred during the period thefuel was maintained in the molten state.Although unconfirmed, it seems probablethat the early rapid release was due toescape of the highly volatile fissionproducts while the slower continuousrelease was due to the low volatilityspecies. The appearance of a dense aerosolin the furnace is indicative of theease with which released vapors willcondense and agglomerate to form a particulatesource.D2. IN-PILE MELTING STUDIESIn addition to the out-of-pile meltingexperiments, a series of in-pile fuelmelting experiments were conducted atORNL (Ref. 15). These experiments consistedof melting miniature stainlesssteel-cladU0 2 fuel elements (-30 gramsU02) in various atmospheres in the OakRidge Research Reactor and measuring thefission products released. Fission andgamma heat in the reactor raised thetemperature of the miniature fuel elementssufficiently high to melt the U0 2without the use of external heat. Datawere obtained for fission product releasefrom the fuel zone (fuel elementand thoria holder) and from the hightemperature zone (thoria and zirconiainsulation) which had a minimum temperatureof 1000 C. Although over twentyexperiments were run most of the majorresults are contained in the summary ofseven runs that are given in Table VIID-6. The first set of values (releasefrom the fuel zone) show extensive releaseof iodine, tellurium, and cesiumregardless of the type of atmosphere.The other fission products and the U02fuel also experienced relatively largereleases which were reduced somewhat inhigh steam concentration atmospheres.The second set of values in the table(release from the high temperature zone)are more useful to reactor accidentanalysis because they give a better indicationof escape from a core regionwhich is not entirely molten. Thesedata are also more comparable with outof-pilestudies which often achieve incompletemelting and have steep thermalgradients in the sample region of thefurnaces. The release data in Table VIID-6 can, in general, be divided into twogroups, a group having high values andanother group having low values. Iodine,tellurium, and cesium belong tothe high group while ruthenium, strontium-barium,and zirconium - cerium belongto the low group. In Experiments13 and 18 ruthenium release was exceptionallyhigh which was probably due tothe highly oxidizing nature of the atmospherein the system. These observationsare generally consistent with theresults that have been obtained from thevarious out-of-pile melting studies.The in-pile experiments resulted in anumber of other conclusions (Ref. 17)which are pertinent to reactor coremeltdown analysis.a. Stainless steelruthenium and,conditions, toappears to retainunder oxidizinglower the meltingVII-II0

point of the U0 2 . However, retentionof the ruthenium by oxidizedstainless steel was not observed.b. Fissiona weak,factorburnup.product release showed onlyif any, correlation with aof 1,000 increase in fuelc. Electron photomicrographs of particlescollected on filters in the inpileapparatus indicated the sourceconsisted of sub-micron sized particleswhich were generally sphericalin shape.D3. SUMMARYAND CONCLUSIONSThe experimental work that has been doneregarding fission product release fromU0 2 during melting can be generally summarizedas follows:a. Nearly total release of the noblegases, iodine, cesium, telluriumshould be expected within a fewminutes from standard pellet-sizemasses of U0 2 . The release of thesespecies seems largely independent ofwhether oxidizing or reducing conditionsexist during melting. Releaseof the four elements is unaffectedby stainless steel cladding and onlytellurium is influenced by Zircaloycladding. Considerable retention oftellurium, probably through compoundformation with zirconium, can occurbut complete oxidation of the Zircaloywould be expected to liberatesuch reacted tellurium.b. The release behavior of rutheniumcan be quite complex because itforms volatile oxides. The degreeof oxide formation appears quitesensitive to the oxygen activity inthe system. The oxygen activitydepends on both the gaseous atmospherethat exists and the otheroxygen reactive materials that arepresent. Air, carbon dioxide, andalso steam represent oxidizing atmospheresfor ruthenium. However,zirconium (and Zircaloy) is a veryeffective oxygen getter. Consequently,the release of rutheniumfrom Zircaloy clad U0 2 in the aboveatmospheres should be expected todepend on the degree of claddingoxidation. If the cladding is completelyoxidized, then excess air,C0 2 , or steam would be available andfission product ruthenium should experiencelarge release. If the Zircaloyis only partly oxidized, theoxygen activity would be low, rutheniumshould remain in the metallicstate and be expected to experienceonly small release. Since stainlesssteel is less reactive with oxygen,it should have less effect on oxygenactivities than Zircaloy. However,evidence indicates that metallic rutheniumwill partition into stainlesssteel from U0 2 and considerableoxidation of the stainless-steel maybe necessary to cause release of thedissolved ruthenium.c. The release behavior of strontiumand barium should also depend on theoxygen activity in the system butnot to the degree nor in the mannerof ruthenium. Metallic strontiumand barium are not very volatile andtheir oxides are even less volatile.Therefore, the higher releases ofthese fission products should occurunder conditions of low oxygen activity,or, for clad U02, in situationswhere incomplete oxidation ofthe Zircaloy has left some free zirconium.The limited experimentaldata tend to confirm this behaviorand indicate that even under veryreducing conditions, the release ofthese elements will not reach highvalues. The data also show thatstainless-steel has essentially noeffect on strontium or barium releasefrom U02. Thus, stainlesssteel is apparently not effective inreducing the oxygen activity farenough to cause conversion of thealkaline earth oxides to the respectivemetals. If these fission productsexist as oxides then lowrelease from molten U02 is to be expected.d. The release of cerium (a rare earth)and fission product zirconium werefound to be quite low in essentiallyall the melting experiments. Theseelements form very stable nonvolatileoxides and their release fromU0 2 should be nearly insensitive toexternal atmosphere and chemicaleffects of cladding materials. Theexperimental results almost entirelysupport this conclusion.VII-ill

References1. Parker, G. W., et al., "Out-of-Pile Studies of Fission-Product Release from OverheatedReactor Fuels at ORNL, 1955-1965", ORNL-3981 (July, 1967), p. 92-102.2. Bedford, R. G., and Jackson, D. D., "Volatilities of the Fission Product andUranium Oxides", UCRL-12314 (January, 1965).3. Gabelnick, S. D., and Chasanov, M. G., "A Calculational Approach to the Estimationof Fuel and Fission-Product Vapor Pressure and Oxidation States to 6000 K",ANL-7867 (October, 1972).4. Genco, J. M., et al., "Fission-Product Deposition and Its Enhancement Under ReactorAccident Conditions: Deposition on Primary System Surfaces", BMI-1863(March, 1969) p. 38-42.5. Coughlin, J. P., "Contributions to the Data on Theoretical Metallurgy. XIIHeats and Free Energies of Formation of Inorganic Oxides", U.S. Bureau of MinesBulletin 542 (1954).6. Parsly, L. F. and Row, T. H., "Study of Fission Products Released from Trace-Irradiated U0 2 into Steam-Air Atmospheres (Nuclear Safety Pilot Plant Runs 8 and9)", ORNL-TM-1588 (May, 1966).7. Parsly, L. F. and Row, T. H., "Behavior of Fission Products Released From SyntheticHigh Burnup U0 2 in Steam Atmospheres (Nuclear Safety Pilot Plant Runs 10-12)", ORNL-TM-1698 (February, 1967).8. Parsly, L. F. and Row, T. H., "Behavior of Fission Products Released into aSteam-Air Atmosphere From Overheated U0 2 Previously Irradiated to 20,000 Mwd/T(Nuclear Safety Pilot Plant Run No. 14, Part I), ORNL-TM-1908 (September, 1967).9. Parsly, L. F., et al.,"Release and Transport of Fission Products Released FromFuel Pins Irradiated to 20,000 Mwd/T: Summary Report of NSPP Run 15", ORNL-TM-3533 (February, 1972).10. Freeby, W. A., Lakey, L. T., and Black, D. E., "Fission Product Behavior UnderSimulated Loss-of-Coolant Conditions in the Contamination-Decontamination Experiment",IN-1172 (January, 1969).11. Collins, R. D., and Hillary, J. J., "Some Experiments Related to the Behavior orGas-Borne Iodine", TRG-933 (W) (1967), p. 5.12. Collins, R. D., Hillary, J. J., and Taylor, J. C., "Air Cleaning for ReactorsWith Vented Containment", TRG-1318 (W) (August, 1966), p. 18.13. Cottrell, WM., B (prog. dir.), "Nuclear Safety Program Annual Progress Reportfor Period Ending December 31, 1966", ORNL-4071 (March, 1967), p. 76.14. Browning, W. E., Jr., et al, "Release of Fission Products During In-Pile Meltingof U0 2 ", Nucl. Sci. and Eng., 18, 151-162 (1964).15. Cottrell, Wm. B. (prog. dir.), "Nuclear Safety Program Semiannual Progress Reportfor Period Ending June 30, 1964", ORNL-3691 (November, 1964), p. 32.16. Cottrell, Wm. B. (prog. dir.), "Nuclear Safety Program Semiannual Progress Reportfor Period Ending December 31, 1964", ORNL-3776 (March, 1965) p. 113.17. Cottrell, Wm. B. (prog. dir.) "Nuclear Safety Program Semiannual Progress Reportfor Period Ending June 30, 1965", ORNL-3843 (September, 1965), p. 37.18. Fischer, J., Schilb, J. D., and Chasanov, M.G., Investigation of the Distributionof Fission Products Among Molten Fuel and Reactor Phases. (Part I - TheDistribution of Fission Products Between Molten Iron and Molten Uranium Dioxide)",ANL-7864 (October, 1971).VII-112

TABLE VII D-I FISSION PRODUCT RELEASE FROM UO2 (a) MELTED IN IMPURE HELIUM(b)Percentage of Individual Fission Product ReleasedSample Time atRun Weight Temp. RareNo. (g) (sec) Gases I Te Cs Ru Sr Ba TRE(c)1 0.57 120 64 71 60 59 28 0.1i2 0.34 120 91 70 72 25 60 0.07 0.8 0.23 0.56 120 93 84 86 34 32 0.16 0.9 1.14 0.56 180 56 67 63 24 75 0.11 1.3 0.75 0.58 180 63 46 54 12 36 0.11 2.6 0.56 0.37 120 69 51 43 7.1 20 0.26 0.5 0.37 0.18 120 99.4 84 86 90 72 0.20 2.0 0.78 0.25 90 99.6 95 96 93 76 3.9 7.3 3.8(a)(b)(c)Trace-irradiated pellet meltedarc-image furnace.Helium flow rate, 100 cc/min.Total rare earths.simultaneously with BeO support tube inTABLE VII D-2 FISSION PRODUCT RELEASE FROM U0- 2 a) MELTED IN PURE HELIUM b) BY THETUNGSTEN-CRUCIBLE METHODMoltenTimePercentagePercentofU02Individual Fission Product Released(min.) Vaporized Xe-Kr I Te Cs Ru Sr Ba Ce1.0 0.10 93 77 90 63 0.45 0.33 4.8 0.051.5 0.16 98 98 98 66 0.05 0.47 2.6 0.072.0 0.16 99 99 99 60 0.32 0.41 3.0 0.172.5 0.25 99 95 99 72 0.33 0.53 2.4 0.131.5 (c) 99 88 92 80 0.20 0.26 2.6 0,402.5 (c) 99 93 96 89 0.70 0.50 3.6 0.10(a) Sample: 29g PWR U0 2 irradiated at tracer level and preheated in helium for 4.5 to 5.0 min.(b) Atmosphere: purified helium flowing at a rate of 700 cc/min.(c) U0 2 sample had a slightly higher density than the first four samples.

TABLE VII D-3FISSION PRODUCT RELEASE (a) FROM TRACE-IRRADIATED PWR-TYPE U0 2 MELTEDIN A SINGLE ELEMENT TUNGSTEN-RESISTOR FURNACE FILLED WITH HELIUM(b)ElementHeatDuration(min.)U0 2Vaporized(%)Percentage of Individual Fission Product ReleasedXe-Kr I Te Cs Ru Sr Ba Ce/REUO 25.00.86347 56 44 1.6 1.65.30.6UO 24.00.25030 42 41 0.4 0.82.90.5UO 24.40.33425 33 40 0.05 1.24.30.5U0 2(SS clad)4.70.256 52 31 46 0.5 1.04.2 0.3U0 2(Zr clad)7.00.152 24 1.1 280.1 10.1U0 2(Zr clad)6.70.0441 50 0.6 320.2 i0.07.5 0.5(a) Results are not corrected for the fraction of the sample melted which is approximatelyequal to the percent rate gas release. Release is from fuel and cladding.(b)Helium flow rate, 400 cc/minTABLE V11 D-4 SUMMSARY OF FISSION PRODUCT RELEASES OBSERVED IN NSPP U0 2MELTING EXPERIMENTSPercent Release From Fuel MassExpt. Furnace Clad DegreeNo. Atmos . Mater. Melted I Te Cs Sr Ba Ru Ce Zr U8 (a) He-steam SS 5% 9 2 6 0.1 1.0 0.3 0.1 0.01 0.19 (a) He-steam SS 80% 61 30 29 0.1 0.3 16.0 0.7 0.1 1.010(b) A-H 2 SS 70-80% 26 -- 20 0.07 -- 2.0 0.2 -- 0.3ii(b) Air SS 70-80% 96 -- 44 0.05 -- 27.0 0.2 -- 1.0A-H 2 SS 50% 74 32 85 -- 2.0 0.9 0.5 -- 2.014(c) Steam Zry-2 limited 10 -- 6 0.9 64.0 0.4 0.3 ....15(c) He-air Zry-2 high 70 -- 56 41.0 32.0 29.0 10.0 ....Percent Transferred to MCV8 He-steam SS 5% 2 0.3 0.1 0.02 0.2 0.02 0.02 -- 0.019 He-steam SS 80% 35 8.0 10.0 0.08 0.06 2.0 0.2 0.03 0.310 A-H 2 SS 70-80% 24 -- 19.0 0.03 -- 1.0 0.2 -- 0.211 Air SS 70-80% 84 -- 26.0 0.02 -- 12.0 0.008 -- 0.612 A-H 2 SS 50% 9 3.0 4.0 -- 0.02 0.09 0.05 -- 0.0214 Steam Zry-2 limited 2 -- 0.7 0.2 2.0 0.1 0.2 ....15 He-Air Zry-2 high 27 -- 42.0 14.0 12.0 0.4 1.0 ....(a)(b)(c)Clad U0 2 trace irradiated.Simulated high burnup U0 2 -premixed.Irradiated to 20,000 Mwd/T.Table VII D-1 - Table VII D-4VII-113/114

TABLE VI1 D-5 FISSION PRODUCT RELEASE FROM HEATED ZIRCALOY-CLAD UO 2ExperimentFuel Burnup(Mwd/T)CDE460-2000CMF (b)7000Fission Product Releases -Percent of Fuel InventoryXe, Kr 86-90I(a) 15-44 91.0Te 7-22 2.0Cs 3-21 30.0Sr 0.01-0.05 0.2Ba 0.04-0.20 -Mo 2-8 -Ru 0.003-0.7 0.2Zr-Nb 0.0001-0.1(a) Some British work at low burnup has been performed:(b) See Reference 13.Reference Burnup Iodine Release,%11 1-2 14-1512 100 8-35

TABLE VII D-6RESULTS OFMELTING OFORNL EXPERIMENTS ON FISSION-PRODUCT RELEASE DURING IN-PILESTAINLESS-STEEL CLAD UO 2U0 2Expt. Sweep Melted,No. Gas % I Te Cs Ru Sr-Ba Zr-Ce U0 2Material Released from theFuel Zone(b) -%4,5,9 Dry He 100 99+ 99 99 78 77 65 4813,18 Moist air (a) 100 98 97 95 38 33 28 2716 87% steam 100 90 96 96 13 16 9 713% air17 87% steam 100 95 90 89 21 19 5 312% He1% H 2Material Released from theHigh Temperature Zone(C)- %4,5,9 Dry He 100 86 86 71 4 1.4 0.4 0.09(a)13,18 Moist air 100 95 89 83 16 1.4 0.6 0.9016 87% steam 100 67 92 42 0.8 0.8 0.1 0.213% air17 86% steam 100 73 87 30 1.2 3.6 0.5 1.312% He1% H 2(a) The steam concentration in this atmosphere was low, corresponding to airsaturated with water vapor at 23 C.(b)Molten times lasted about 5 minutes.(c) Minimum temperature, 1000 C.Table VII D-5 - Table VII D-6VII-115/116

APPENDIX EAN EVALUATION OF FISSION PRODUCT AND FUEL CONSTITUENTRELEASE FROM REACTOR FUELS BASED ON A THERMODYNAMICANALYSIS OF THE COMPOUND SPECIES PRESENT IN THE FUELbyM. Pobereskin, C. Alexander and R. RitzmanBattelle's Columbus LaboratoriesVII-117

Appendix ETable of ContentsSectionPage No.El. INTRODUCTION ..................................................... VII-121E2. ANALYSIS FOR TEMPERATURES BELOW FUEL MELTING ..................... VII-122E3. ANALYSIS FOR MOLTEN FUEL ......................................... VII-123E4. CONCLUSIONS ...................................................... VII-123REFERENCES ............ .............................................................. VII-124List of TablesTablePage No.VII E-1 Results of Thermodynamic Analysis at T = 800 K .................... VII-125/126VII E-2 Results of Thermodynamic Analysis at T = 1300 K ................... VII-125/126VII E-3Results of Thermodynamic Analysis at T = 1900 K ................... VII-125/126VII E-4 Results of Thermodynamic Analysis at T = 2500 K ................... VII-125/126VII E-5Vapor Species and Maximum Volatilization Rate Estimates,T = 3100K .......................................................VII-125/126VII-119

46Mk

Appendix EAn Evaluation of Fission Product and Fuel Constituent ReleaseFrom Reactor Fuels Based on a Thermodynamic Analysisof the Compound Species Present in the FuelbyM. Pobereskin, C. Alexander and R. RitzmanBattelle's Columbus Laboratories-0El. INTRODUCTIONConsiderable research has been performedaimed at establishing the nature andamount of fission products released fromthe core of a light water nuclear reactorin event of a loss-of-coolant. Experimentaland analytical approacheshave both been attempted (Ref. 1). Generallyspeaking, the experiments havebeen on a laboratory scale, and requireextensive interpretation to reflect reallife conditions. Similarly, the analyticalapproaches have been predicted onexistence of chemical species presumedfrom general knowledge of the chemistryof the element.In this study, the evaluation of fissionproduct behavior is based upon a thermodynamicanalysis of the fission productand fuel species present in the fuelunder pertinent conditions. This ismade possible through the application ofthe EQUICA computer program to theanalysis of the nature, amount, andstate of the species at thermodynamicequilibrium.EQUICA rapidly computes the compositionof a given set of reaction species. Itis especially versatile in its abilityto accept input data which can incorporatethermal and entropy effects otherthan those of the tabulated values givenin handbooks and the JANAF Tables.EQUICA consists of a basis which maycontain 20 species in which all elementspresent in the computation are presentin either elemental or some combinedform. An additional 40 species may beincluded in the computation. There areno other stipulations applied to thebasis other than it can contain no morethan 20 different elements. As a matterof fact, the program has provision forshifting the basis so that the speciesrichest in a certain element would formthe basis for that element in all speciesin which it appeared, and allequilibrium constants would be based onits equilibrium value in that particularcomposition. For instance, in a computationat very high temperature gaseouscarbon may form the basis for carbon,under pyrolysis conditions elementalcarbon may form the basis, and undercombustion conditions carbon in carbondioxide may form the basis for carbon.This ability to shift basis greatlyimproves the convergence for complexsystems when the final composition differsmarkedly from the initial estimate.Unlike most thermochemical programsthere is no limit on the number ofsolids in either the basis or nonbasissets, and one can have the same chemicalsubstance present as both a gas or vaporand a condensed phase or as a componentin a solution. Under these conditionsEQUICA will correctly determine the saturationvapor pressure of the vaporspecies. EQUICA will also allow for thevanishing of a condensed phase if it isestablished thermodynamically that thephase should not exist. This is inkeeping with the phase rule and ensuresthat the proper degrees of freedom aremaintained.The program utilizes a compositionmatrix made up of N molecular speciescomprised of S chemical elements. Acomposition matrix relates the speciesto the elements. If one considers thenthat the composition Ni of the Ithspecies be changed by AC, then to preservestoichiometry the followingchanges in the composition of the basisare needed:1Ni = Ni + Ac.It is possible that for some minor speciesN. could go negative, and since NientersI the program for the form of (lnNi), then this relation becomes undefinedas Ni - 0. For this reason asubroutine FUDGE has been written andVII-121

incorporated to determine if the currentAC produces non-positive Ni values. Ifit is found that negative compositionsexist then the subroutine computes a newAE value which allows all Ni's to remainconstant. The input data include initialestimates of composition and Gibbsfree energies of all species at thedesired temperature and pressure. Theoutput consists of the following:a. The equilibrium compositions of allspecies.b. The equilibrium constants of allspecies. (Although equilibrium constantsfor the basis species are notdefined, the program sets theirvalue to one. Thus a one in theequilibrium constant column indicatesthat this species currently isa basis species.)c. A defined byThis AvolumeoccupyT.PA =nEyiN iivalue is thenthe gaseousat pressure Prelated to theproducts wouldand temperatured. The convergence vector, C, whichdefines the extent of convergence ofall major and minor species.e. Partial pressure for each species.f. Total pressure (sum of partial pressures).E2. ANALYSIS FOR TEMPERATURESBELOW FUEL MELTINGIn the present analysis, it was necessary,of course, to include uranium,oxygen, zirconium, hydrogen, and heliumamong.the element inputs to the program.This left room in the program forfifteen elements to be selected from themajor fission products. It was foundthat selecting these elements was not aslimiting as was the condition that only60 compound species could be considered.In addition to the elements listedabove, Cs, Ba, La, Sm, Mo, Ru, Te, I,Br, and Sr were included in the analysis.This choice of elements was basednot only on abundance in fission yieldand representation of the major chemicalgroups, but also on uniqueness of thermochemicalproperties. For example, thelanthanum oxide would satisfactorilyrepresent the sesquioxides except thatelemental samarium was considered to besufficiently volatile that in the firstanalysis samarium needed to be consideredapart from the other rare earths.Two computer runs were made based onapproximate preaccident and postaccidentconditions. The initial run was designedto provide some prediction of thechemical species that could exist infuel rods just before the accident, andto indicate fission product volatilityas a function of temperature for periodsafter this. The effect of several parameterswas examined.a. Fuelusedsionburnup - 1 and 4 percent wereto produce variations in fisproductconcentration.b. Temperature - equilibrium calculationswere made for 800, 1300, 1900,and 2500K, corresponding roughly totemperature regions in operatingfuel rods and to accident temperaturesfor an overheating reactorcore.c. Pressure - 1000 and 2000 psia wereused to simulate internal rod pressures.d. Steam concentration - 30 ppm waterin the fuel and unlimited steamsupply were used to simulate normalconditions and more oxidizing conditionsfor the fuel-fission productcladdingsystem.Elemental compositions were predicted ona nominal fuel rod containing 10 poundsof U0 2 and 2 pounds of zirconium. Solutioneffects of oxides in the parentfuel were not included because, for themost part, the oxides have very limitedsolid solubility in U0 2 .The results of the parametric analysisto 2500K are summarized in Tables VIIE-1 through VII E-4, which for each temperaturegive the predicted dominantcondensed phase and vapor phase formsand the approximate mole ratio of materialin the vapor phase to material inthe condensed phase. This ratio indicatesthe volatility of the fission productspecies. Several conclusions weredrawn from the data.a. The degree of burnup and the totalpressure had negligible effects onthe results.b. The volatility of some species (generallyBa, La, Sm, Zr, and Sr) isdepressed by excess steam (more oxidizingconditions) with the effectmore evident at lower temperatures.VII-122

c. The volatility of some speciesenhanced by excess steam - Ru andat higher temperatures and Te,Br, and Cs at lower temperatures.isUI,d. The most volatile fission productspecies are Cs, Te, and CsI. Athigher temperatures, particularlyunder more reducing conditions, Baand Sr begin to approach the abovethree in volatility.Generally speaking, the results confirmedwhat one might have predictedfrom a general knowledge of thechemistry of the involved elements. Anexception was the indication that iodineand bromine react with the excess cesiumin the fuel. In the subsequent runs,with HI as a species in the matrix, theCsI was still the indicated major iodinecarrying species, even at temperaturesto 3100K. Furthermore, CsI will notoxidize in air to Cs20 and 12 nor inwater vapor to Cs20 + 2HI. Therefore,based on these calculations it appearsthat CsI may be present in the releasedfission product vapors. It should benoted though that no experimental confirmationof this prediction is available,and other reactions that cesiummight undergo could alter the conclusion.E3. ANALYSIS FOR MOLTEN FUELIn this computer run, several new specieswere introduced; e.g., stainlesssteel constituents Fe, Cr, Ni (and theiroxides), and Zr(g), Mo(g), Ru(g), andHI(g). To make room for these species anumber of the nonreactive species fromthe first run were dropped from thematrix. Conditions were fixed at 3100K,1000 psia, 1 percent burnup, and unlimitedwater. For this run the free energieswere adjusted and the oxides wereconsidered to form an ideal solution.The results of the run showed that theFe, Cr, Ni compound species did notaffect the reaction patterns establishedin the systems at temperatures to 2500K.The overall situation with excess waterpresent was quite oxidizing in that thepartial pressure of U0 3 was more thantwo orders of magnitude greater thanU02 in all cases.The data were used to calculate volatilizationrates utilizing the Knudsen-Mayer equation:Z = 44P p M/T(VII E-l)where Z is the volatilization rate ingrams per second per square centimeter,P is the vapor pressure in atmospheres,M the molecular weight, and T is thetemperature in degrees Kelvin. Althoughthis relationship holds for molecularflow in vacuum, Fonda (Ref. 2) hasindicated that in the presence of aninert cover gas, the same relationshipapplies but the rate is effectivelyreduced by a factor of 1/80 to 1/100.For practical purposes, at temperaturesnear 3000K, one may utilize therelation:Z % 0.01 P i--M- g/sec/cm 2 (VII E-2)as long as the rate-controlling step ofdiffusion through a boundary layer filmEquation (VII E-2) is expected to beapplicable. Under conditions tending tobe more static, the evolution rate wouldbe even lower. Thus Equation (VII E-2)may be considered as an upper limit forvaporization loss. Equation (VII E-2)appears to be only a slowly changingfunction of pressure at atmosphericpressure or greater. Thus at pressuresas high as 50 psia Equation (VII E-2)should closely approximate actual releaseconditions.The vapor species expected to be releasedand the maximum-volatilizationrates obtained for the molten fuel caseare given in Table VII E-5.It is assumed that convective forces areoperative in the gas phase which willlift the volatiles along with the gaseouscurrents and these species will beentrained in the gas at least until thespecies are contacted by a surface wherethey may condense. They could be carriedas aerosol particles however.These results indicate that uranium isabout as volatile as anything else andits volatilization should assist theloss of volatile fission products fromthe receding surface. Care must be usedin attempting to extrapolate the volatilizationrates to any real core meltdownsituation. The values given inTable VII E-5 should be considered onlyindicative of the relative volatility ofthe different species under the preciseconditions selected for the calculation;i.e., excess water and 3100K. In thiscontext the data show the potentialvolatile nature of species in moltenU02--even some species which are normallyconsidered refractory.E4. CONCLUSIONSIn order to provide a basis for evaluatingthe dynamics of release, an analysisVII-123

was made of the degree of volatilizationof the fission product species as afunction of temperature. These data aresummarized in Tables VII E-1 to VII E-4.Assuming cladding rupture at 800 K,three major phases in the release processare recognized. These are identifiedbelow and estimates of releasemagnitudes are provided.a. Release at Time of RuptureCs -Te -Several percent of inventoryLess than 1 percent of inventory12 - Less than 1 percent of inventoryb. Release During Heat-up to MeltingCs - remainder, mostly as Cs vapor(early in period)Te -remainder as Te vapor (early inperiod)I - remainder, as CsI vapor (laterin period)Ba - a large percent as Ba or BaOvaporSr -a few percent as Sr vaporc. Release from Molten FuelRu -Sr -Mo -potentialRuO3potentialSrpotentialMoO 3for rapid release asfor rapid release asfor rapid release asLa -Zr -potential for slow release asLaOpotential for very slow releaseas ZrO 2U - potential for rapid vaporizationas U0 3Certainly, the release characteristicsof the fission products indicated aboveare restricted to the thermodynamic systems,components, and conditions thatwere analyzed. Kinetic limitations andan exhaustive treatment of all potentialcompound species could not be included.Nevertheless the predictions in manycases tend to agree with other methodsused to estimate fission product releases.Exceptions appear to betellurium and iodine release during fuelheatup to melting, but here possibletellurium reaction with the cladding wasnot included and cesium iodide stabilitywas considered only with respect toformation of cesium oxide or formationof hydrogen iodide. The results indicatedfor molten fuel are based on arather extreme condition; i.e., puremolten fuel in equilibrium with excesswater. In fact fuel will probably bediluted with lower melting oxides ofstructural materials, the mass is likelyto be non-isothermal, and the availabilityof water should be limited. Consequently,the releases from molten fuelhave only a potential for occurring andacutally represent upper limit predictionsat this time. Considerable experimentalwork is needed in the area ofcore meltdown and materials interactionsbefore much improved projections offission product escape can be made.Thermodynamic analyses such as these arevery useful in revealing the importantproblems which require investigation.References1. Morrison, D. L., et al., "An Evaluation of the Applicability of Existing Data tothe Analytical Description of a Nuclear Reactor Accident", BMI-1779 (August,1966).2. Fonda, G. R., "Evaporation of Tungsten Under Various Pressures of Argon", Phys.Rev., 31, 260 (1928).VII-124

TABLE VII E-1 RESULTS OF THERMODYNAMIC ANALYSIS AT T = 800 KFissionProductDominantCondensed FormDominantVapor FormApproximateComposition RatioVapor/CondensedLimited SteamCSBaLaSmZrMoRuTeIBrSrUCsBaLa203Sm 2 0 3ZrMoO 2RuTeCsICsBrSrOU0 2CsBaLaOSmN.C.N.C.N.C.Te12Br2SrN.C.2211714xxxxxxxx10-i0- 9i0-2510- 1410-3io-8i0-i0-18CsBaLaSmZrMoRuTeIBrSrUunlimited SteamCsBaOLa 2 03Sm 2 0 3ZrO 2N.C.RuTeCsICsBrSrOU0 2CsBaON.C.N.C.N.C.N.C.N.C.TeCsICsBrSrN.C.5 xixlxlx5x3 xi0- i110-2010-210-i0- 310-23N.C. = Not certain because convergence not achieved in the iteration limit.TABLE VII E-2 RESULTS OF THERMODYNAMIC ANALYSIS AT T = 1300 KApproximateFission Dominant Dominant Composition RatioProduct Condensed Form Vapor-Form Vapor/CondensedLimited SteamCs Cs 2 0 Cs 1 x 105Ba Ba Ba 2 x 1013La La 2 0 3 LaO 2 x 1071Sm Sm 2 0 3 Sm 2 x 1018Zr Zr ZrO 1 x 10-18Mo MoO 2 N.C.Ru Ru N.C. 0Te Te Te 4 x 10 0 2I CsI CsI 3 x 10-1Br CsBr CsBr 2 x 10-1Sr SrO Sr 5 x 10- 1U U0 2 U0 5 x 10-IUnlimited SteamCs CS 2 0 Cs 1 x 1095Ba BaO BaO 1 x 1016La La 2 0 3 LaO 1 x 10_19Sm Sm 2 0 3 Sm 2 x 10 3Zr ZrO 2 ZrO 2 1 x 10-2Mo N.C. N.C. -0_27Ru Ru RuO 3 1 x 10 2Te TeO 2 Te 3 x 1041I CsI CsI 7 x 103Br CsBr CsBr 4 x 10_11Sr SrO Sr 1 x 10_17U U02 U02 1 x 10N.C. = Not certain because convergence not achieved in the iteration limit.

TABLE VII E-3 RESULTS OF THERMODYNAMIC ANALYSIS AT T = 1900 KApproximateFission Dominant Dominant Composition RatioProduct Condensed Form Vapor Form Vapor/CondensedLimited SteamCs None Cs >1 x 1083Ba Ba Ba 2 x 10 7La La 2 0 3 LaO 2 x 104Sm Sm 2 0 3 Sm 7 x 10-11Zr Zr ZrO 4 x 10Mo N.C. N.C. -Ru Ru RuO 3 2 10Te TeO 2 Te 3 x 104I CsI CsI >8 x 10 3Br CsBr CsBr >4 x 10 3Sr SrO Sr 7 x 10•2U U0 2 UO 1 x 10-10Unlimited SteamCs None Cs >1 x 1048Ba BaO BaO 2 x 10 -4La La 2 0 3 LaO 7 x 109Sm Sm 2 0 3 Sm 2 x 10-13Zr ZrO 2 ZrO 2 1 x 1013Mo N.C. N.C. -Ru Ru RuO 3 lx 10-16Te None Te >2 x 103I CsI CsI >8 x 103Br CsBr CsBr >4 x 10JSr SrO Sr 5 x 10-6U U0 2 U0 3 2 x 10-9N.C. = Not certain because convergence not achieved in the iteration limit.TABLE VII E-4 RESULTS OF THERMODYNAMIC ANALYSIS AT T = 2500 KApproximateFission Dominant Dominant Composition RatioProduct Condensed Form Vapor Form Vapor/CondensedLimited SteamCs None Cs >1 x 1082Ba Ba Ba 3 x 1024La La 2 0 LaO 3 x l0o-Sm Sm 2 o0 Sm 3 x 10_7Zr Zr ZrO 3 x 10Mo N.C. N.C. -23Ru Ru RuO3 2 x 10-7Te None Te >2 x 103I CsI CsI >8 x 10•Br CsBr CsBr >4 x 103Sr SrO Sr 3 x 1047_U U0 2 U0 2 x 10Unlimited SteamCsBaNoneBaOCsBaO>1 x 1088 x 10-2La La20 3 LaO 1 x 10-4Sm Sm 2 0 3 Sm 7 x 10-7Zr ZrO 2 ZrO 2 1 x 10-8Mo N.C. N.C. -Ru Ru RuO 3 6 x 1013Te None Te >2 x 107I CsI CsI 7 x 103Br CsBr CsBr 3 x 1032Sr SrOU U02SrU021 x 10-61 x 10N.C. = Not certain because convergence not achieved in the iteration limit.

TABLE VII E-5VAPOR SPECIES AND MAXIMUM VOLATILIZATION RATE ESTIMATES,3T = 100KSpeciesEquilibriumVapor FractionEquilibriumPartial Pressure,atmLimitingVolatilizationRate,g/cm 2 /secCs1.0BaOTe, TeOCsIRuO 3SrMoO 3LaOZrO2UO 31.01.01.00.470.0340.0240.0170.0090.5118182xxxxx710-210 410310-510-61812(a)(a)(a)(a)x 105x i0-b*104~x 105x i0-J* 0i(a) These are too highly volatile to estimate rates. No condensed phase existsat 3100 K.Table VII E-1 -- Table VII E-5VII-125/126

APPENDIX FSUMMARY OF DATA ON FISSION PRODUCT RELEASEFROM U0 2 DURING OXIDATION IN ARbyR. L. RitzmanBattelle's Columbus LaboratoriesandG. W. ParkerOak Ridge National LaboratoryVII-127

Appendix FTable of ContentsSectionSUMMARY OF DATA ON FISSION PRODUCT RELEASE FROM U0 2 DURINGOXIDATION IN AIR ........................................................REFERENCES ...............................................................Page No.VII-131VII-132List of TablesTableVII F-IFission Product Release from U0 2 Oxidized inAir Sample: .....................................................Page No.VII-133/134VII F-2 Fission-Product Relase from PWR-Type U0 2 - Irradiated to4000 Mwd/T ......................................................VII F-3 Fission-Product Release from PWR-Type U0 2 - Irradiated to7000 Mwd/T ......................................................VII-133/134VII-133/134List of FiguresFigureVII F-I Fission Product Release by the Oxidation of U0 2 to U 3 0 8in Air, Showing Discontinuity between 600 and 900 C ..............Page No.VII-135/136VII F-2Effect of Burnup on Fission-Product Release by Oxidation at1000 and 1200 C ................................................. VII-135/136VII-129

Appendix FSummary of Data on Fission Product ReleaseFrom U0 2 During Oxidation in AirbyR. L. RitzmanBattelle's Columbus LaboratoriesandG. W. ParkerOak Ridge National LaboratoryWhen uranium dioxide is heated in air attemperatures below about 1550 C, theUOý is oxidized to U 3 0 8 with an accompanyingexpansion of the solid lattice(Ref. 1). The U308 surface layer cracksupon accumulation of sufficient stressand exposes the underlying U0 2 to continuedoxidation (Ref. 2). The oxidationreaction is exothermic and ignitionor burning of U0 2 has been observed(Ref. 3). These conditions--increasedsurface area, phase change with latticeexpansion, and elevated temperatures--are conductive to fission product releaseduring the oxidation process.Parker et al.(Ref. 2) first studied releasefrom trace-irradiated PWR-typeU0 2 pellets of 94 percent theoreticaldensity and later (Ref. 4) from the sametype of fuel irradiated to differentlevels of burnup to a maximum of 7000Mwd/T. Approximately one-gram sampleswere heated for different periods in airflowing at 100 cc/minute at temperaturesranging from 500 C to 1400 C.The pertinent fission product releasedata that were obtained in these seriesof experiments at ORNL are summarized inFig. VII F-1 and in Tables VII F-l, VIIF-2, and VII F-3. Release values determinedwith specimens irradiated at tracelevel (nl Mwd/T) are plotted in Fig. VIIF-1. Results show that release is not asimple function of temperature, butabove 900 C releases tend to increasewith temperature.The effect of varying the heating timeat different temperatures is shown inTables VII F-2 and VII F-3. Specimensemployed to obtain these data were PWRtypematerial with a density of 93 to 94percent of theoretical, irradiated to aburnup of 1000 Mwd/T and 4000 Mwd/T,respectively. Increasing exposure timein air in the range investigated seemedto have no significant effct on fissionproductrelease below 800 C, but at thistemperature and above, increasing releaseof some isotopes with increasingexposure was observed.The effect of burnup on oxidation releaseof the more volatile fission productsis shown graphically in Fig. VIIF-2 for two temperatures. The largesteffect in the case of iodine and rutheniumcame in the first 1000 Mwd/T ofburnup and this was also true of therare gases at 1200 C. The release oftellurium appeared to increase more orless regularly with increasing burnup inthe range tested. The release of cesium,even at 1200 C, was too low to establishan unequivocal correlation butthe results obtained indicate a slightincrease in release with increasingburnup.The outstanding feature of these results,particularly at the higher temperatures,is the high release of raregases, iodine, ruthenium, and tellurium.Tellurium release is not quite as largeas the other three but still much greaterthan cesium, strontium, or barium.Ruthenium release is probably largelydue to formation of a volatile oxide.Tellurium oxide is also reasonably volatileat the higher temperatures. Thevery low releases of cesium, strontium,and barium are indicative of the lowervolatility of their oxides. In general,the results also indicate that only afew minutes exposure to air at temperaturesof 1200 C or above would be necessaryto cause large releases of raregas, iodine, ruthenium, and telluriumfission products.VII-131

References1. Belle, J., Editor, Uranium Dioxide: Properties and Nuclear Applications, USAEC,1961, Chapter 8.2. Parker, G. W., G. E. Creek, and W. J. Martin in "Chemistry Division AnnualProgress Report for Period Ending June 30, 1961", ORNL-3176 (Sept 1961), p 68.3. Leitnaker, J. M., M. L. Smith, and C. M. Fitzpatrick, "Conversion of UraniumNitrate to Ceramic Grade Oxide for the Light Water Breeder Reactor: ProcessDevelopment", ORNL-4755 (April 197.2).4. Parker, G. W., et al., in "Nuclear Safety Program Semiannual Progress Report forPeriod Ending June 30, 1962", ORNL-3319 (Aug 1962), p 11.VII-132

TABLE VII F-1 FISSION PRODUCT RELEASE FROM U0 2 OXIDIZED IN AIRSample: Intermediate density PWR U0 2 (93-94%)Irradiation: 1000 Mwd/tonAir flow:. 100 cc/minTime atTemperature Percentage of Individual Fission Products ReleasedTemp (min) Rare(C) He Air Gases I Te Cs Ru Sr Ba U500 9 4.4 4.6

TABLE VII F-2 FISSION-PRODUCT RELEASE FROM PWR-TYPEU0(a)2 (a)Irradiated to 4000 Mwd/T and Heated in AirTime atTemperatureTemp (min.)(c) He AirPercentage of Individual Fission Products ReleasedCGases I Te Cs Ru Sr Ba U500 16.0 23.0 1.518.0 90.0 2.93.6

Temperoture, C8000C) o .ooO o 80 0 0.22 0ý 0 00oII-2-- x I04FIGURE VII F-1Fission Product Release by the Oxidationof U0 2 to U 3 0 8 in Air, ShowingDiscontinuity between 600 and 900 C100 -908070* 60• 50a- 403020t0-FIGURE VII F-2Effect of Burnup on Fission-ProductRelease by Oxidation at 100 and 1200 CFig. VII F-i - Fig. VII F-2VII-135/136

APPENDIX GESTIMATIONS OF FISSION PRODUCT RELEASEFROM MELT DURING CONCRETE PENETRATIONbyL. F. Parsly, Jr., and M. H. FontanaOak Ridge National LaboratoryVII-137

Appendix GTable of ContentsSectionESTIMATIONS OF FISSION PRODUCT RELEASE FROM MELT DURING CONCRETEPENETRATION .............................................................Release During Gas Sparging ......................................REFERENCES ..............................................................Page No.VII-141VII-141VII-142List of TablesTableVII G-1Removal of Fission Products from Melt by Sparging with CO 2from Concrete Decomposition .....................................Page No.VII-143/144VII-139

Appendix GEstimations of Fission Product ReleaseFrom Melt During Concrete PenetrationbyL. F. Parsly, Jr, and M. H. FontanaOak Ridge National LaboratoryConditions are created which could promotefission product release from themolten core material during its penetrationof the containment concrete base.Carbon dioxide generated from decompositionof limestone aggregate, and steamliberated from solid hydrates or includedwater, may pass through the melt producinga gas sparging effect. Analysesperformed in the core meltdown task indicatethat high-gas generation rateswould occur during the first half-hourafter the molten core contacts the concretebase. Subsequently, the gas generationrates should decrease as concretepenetration slows, but thecumulative gas generation during thisperiod would exceed that of the earlyperiod. In order to obtain an estimateof the effectiveness of this process inpromoting fission product release, simplevolatilization calculations wereperformed.RELEASE DURING GAS SPARGINGThe assumption was made that concretewould be decomposed and that a significantdecomposition product would be C02.The C02 at 3000 K and 1 atm pressure wasassumed to bubble through the moltencore, and it was assumed that equilibriumdistribution of fission productsbetween the melt and the gas bubbleswould be established. Distribution coefficientsfor the fission products betweenthe melt and the C02 were calculatedfrom the data given in Gabelnick andChassnov (Ref. 1) for 3000 K, 5 percentburnup and 5 g/cc smear density. Thesewere assumed to be independent of concentration.Based on the assumed equilibrium,a differential material balancecan be written. From this we find that:whereH VeGC = CO0 exp ( VL ), (VII G-l)C = final concentrationCO = initial concentrationH = distribution coefficientconcentration in gas/concentrationin liquidVG = volume of gas which has bubbledthrough the meltVL = volume of melt.Thus, the fraction retained isCo= exp - v L (VII G-2)and the fraction removed isC -C0C 0=l - exp (VII G-3)The calculated distribution coefficientsand the fraction removed values obtainedfrom Equation (VII G-3) are given inTable VII G-1. The results are based on3a melt volume (VL) of 932 ft and atotal C02 volume (VG) of 1.25 x 106 ft3 .This volume of CO 2 corresponds to decompositionof approximately 5.8 x 104 lbof concrete (Ref. 2). Complete decompositionof this mass may not occur and/orthe generated C02 may partially bypassthe melt. Thus, two columns of fractionremoved values are given in Table VIIG-l; one assuming only 20 percent of thetotal C02 volume sparges the melt andanother which assumes 100 percent. TheVG/VL values in Equation (VII G-3)applying to these two conditions arerespectively 268 and 1340%The results of these calculations indicatethat the fission products can begrouped into four classes as follows:a. Highly volatile - essentially completelyremoved by even limitedsparging. These include Ag, Cd, Cs,In, I, Rb, Sb, Se, Sn, and Te.VII-141

. Refractory - almost completely retainedafter sparging. These includeCe, Nb, Rh, and Zr.c. Moderately volatile - extent of removaldepends on amount of sparging.These include Eu, La, Mo, Nd, andPd.d. Uncertain - has one volatile and onenon-volatile form. These includeBa, Gd, Pm, Pr, Ru, Sm, Sr, Tc, andY. However, as noted in the table,the volatile form of each of thesefission products is the minor speciesin the thermodynamic system analyzed.The sparging process wouldcreate highly oxidizing conditions.On this basis, Ba, Gd, Pm, Pr, Sm,Sr, and Y should exist as oxides andcan legitimately be considered refractory.Under the same conditionsRu and Tc should form the volatileoxides but the presence of metalliciron in the real melt system wouldinterfere. Therefore, only the volatilityclassification for these twofission products remains questionable.It must be cautioned that the volatilizationfractions listed in Table VII G-lwere obtained using a simple rateexpression based on convenient assumptionsregarding chemical and physicalequilibrium. While very useful indeveloping the qualitative conclusionsjust presented, the data should not beconsidered a quantitative prediction ofactual releases in a containment meltthroughsituation.References1. Gabelnick, S. D. and Chasanov, M. G., "A Calculational Approach to the Estimationof Fuel and Fission Product Vapor Pressures and Oxidation States to 6000 K", ANL-7867 (October, 1972).2. Reimers, R. A., and Seidenfeld, A., "Investigation of the Consequences of aNuclear Reactor Core Melt-through Accident, ORNL-MIT-62 (October, 1968).VII-142

TABLE VII G-1REMOVAL OF FISSION PRODUCTS FROM MELT BY SPARGING WITHCO 2 FROM CONCRETE DECOMPOSITIONFission ProductDistributionb)CoefficientPercent Sparge Gas Volume20 100Fraction Removed from MeltAgAg 2 0Ba (a)BaOCdCdOCeCe 2 03CsCs 2 0EuEu 2 03Gd (a)Gd203IInIn 2 0 3La (a)La 2 03MoMoO 2NbNbO 2Nd (a)Nd 2 03PdPdOpm(a)Pm 2 03pr(a)Pr 2 03RbRb 2 0RhRh 2 0RuRuO 2(a)SbSb2033.261.629.354.094.933.454.141.256.511.532.732.433.335.684.78406.02.003.673.351.212.472.929.651.276..123.808.843.251.031.387.534.082.067.931.116.124.408.80xxxxxxxxxxxx10-210110-210510 210-5 --2105101-5310110-110-361076x 10-210-410510 810-710-410-610-610-410-410-410-410-610- 610-310-110- i0-0-60- 510-6110- 2>0.999>0.999>0.9990.01>0.999>0.9990.010.999>0.999>0.9990.480.590.999>0.999>0.9990.050.010.9990.045>0.999>0.9990.0450.017>0.999>0.999>0.9990.96>0.9990.999>0.999>0.9990.230.0510.9990.030.010.02>0.999>0.999>0.999

TABLE VII G-1(Continued)Percent Sparge Gas VolumeDistributionb) 20 100Fission Product Coefficient Fraction Removed from MeltSeSeO 2Sm (a)Sm 2 03SnSnOSr (a)SrOTcTcO 2(a)TeTeO2(a)Y203ZrZrO2U0 2Pu024.49 x 10-14.103.24 x 10- 15.45 x 10-54.02 x 10-33.62 x i022.45 x 10-18.18 x 10-73.69 x 10-71.57 x 10-30.62 x 10-13.40 x 101-3.45 x 1041.77 x 10-55.2 x 1071.59 x 10- 84.71 x 102.64 x 10-76>0.999>0.999>0.9990.010.66>0.999>0 .9990.9990.0560.010 .9990.070.995>0.999>0.9990.0010.999>0.9990.3840.02

APPENDIX HEXAMINATION OF SOLID FISSION PRODUCT PLATEOUTIN PRIMARY SYSTEMSbyR. L. RitzmanBattelle's Columbus Laboratories.VII-145

Appendix HTable of ContentsSectionHI.ESTIMATED VAPOR CONCENTRATIONS OF FISSION PRODUCTS ANDCONDENSATION CONDITIONS ..........................................Page No.VII-149H2. ESTIMATION OF THE THERMAL HISTORY OF UPPER PRESSURE VESSELINTERNALS ........................................................H3. ESTIMATED VAPORIZATION OF CORE STRUCTURALS AND BEHAVIOR IN UPPERPLENUM ...........................................................H4. CONCLUSIONS ......................................................REFERENCES ..............................................................VII-150VII-151VII-152VII-153List of TablesTableVII H-IVII H-2Estimated Solid Fission Product Partial Pressures as a Functionof Gas Temperature after Leaving Core Region ....................Approximate Supersaturation Ratio for the Solid Fission Productsas a Function of Gas Temperature ................................Page No.VII-155/156VII-155/156VII H-3 Average Exit Gas Temperatures from Core Heatup Analyses .........VII-155/156VII H-4Upper Plenum Structural Temperatures Versus Time for Heatingby Steam Only ...................................................VII-155/156VII H-5 Materials Data and Potential Vaporization Loss Rates ............VII H-6 Saturation Vaporization Rates and Concentrations ................VII-155/156VII-155/156VII-147

Appendix HExamination of Solid Fission Product Plateoutin Primary SystemsbyR. L. RitzmanBattelle's Columbus LaboratoriesFission products that are released fromthe reactor fuel before and during coremelting must travel through the rest ofthe primary system in order to escapeinto the containment atmosphere. Thefirst and principal component of thesystem is the pressure vessel, and it ispossible that some of the releasedfission products could be retained oninternal structures inside this vessel.Specifically fission product vapors thatare swept out of the core region willencounter cooler temperatures in theupper plenum. The vapors may condenseon the cooler structural surfaces inthis region thus preventing their escapeinto the containment. However, it isalso conceivable that the vapors mightundergo gas phase condensation withother vaporized materials, forming anaerosol which would be swept out of thevessel. The following sections of thisdocument present the results of severalcalculations that were performed toprovide some insight into the fate ofsolid fission products in the reactorvessel. Examination of iodine depositionin the primary system is covered inanother document.HII. ESTIMATED VAPOR CONCENTRATIONSOF FISSION PRODUCTS ANDCONDENSATION CONDITIONSIn order to obtain an estimate of thepotential for solid fission productplateout in the upper plenum region ofthe pressure vessel, a series of volatilitycalculations were performed.Fission product vapor concentrations inthe bulk gas flowing through the upperplenum were estimated using the meltdownrelease rates and bulk steam-hydrogenflow rates that would be characteristicof the meltdown period.The estimated length of the meltdownperiod for a PWR or BWR varies somewhat(Ref. 1) but 45 minutes is a reasonablevalue. In this analysis it is assumedthat the solid fission product meltdowncomponent releases are released at aconstant rate over the 45 minute period.Combining the fraction released valueswith calculated core fission productinventory data, an estimate of theindividual mass release rates can bemade. Then knowing the bulk gas flowrate from the core region one cancalculate the fission product gas phasepartial pressures. The bulk gas (steamhydrogen)flow originates from waterboiloff inside the pressure vessel andthe rate is estimated to range fromabout 5000 to 10,000 cfm (Ref. 1). Thesmaller rate will be used to maximizethe estimated partial pressures. Theresults of performing this exercise fora series of gas temperatures are summarizedin Table VII H-1. Note that thepartial pressures are calculated withoutregard to vapor saturation restrictions.The result of considering saturationconditions for each of the fissionproduct elements is presented in TableVII H-2. Here the ratio of the calculatedpartial pressures given in Table VIIH-1 to the respective saturation pressures(Ref. 2) for each fission productelement and temperature are tabulated.This ratio indicates the degree of supersaturationthat would occur in eachcase if the hot gas mixture leaving thecore region were suddenly cooled to thegiven temperatures. Values greater thanone indicate that condensation of thefission product would be favored undersuch conditions while values less thanone tend to preclude any condensation.Although the supersaturation ratios inTable VII H-2 are only approximations,several useful observations can be made.For example, at gas or system temperaturesabove about 1000 F, no stable condensedcesium or tellurium phase shouldexist. Above about 1500 F the same isexpected for strontium. However, at alltemperatures considered here, condensationof ruthenium and yttrium. is stronglyfavored, and at the lowest temperature(530 F) strontium, tellurium, andprobably cesium should also condense.These results indicate the importancethat thermal conditions will have on thebehavior of solid fission products inprimary systems.VIT-149

H2. ESTIMATION OF THE THERMALHISTORY OF UPPER PRESSUREVESSEL INTERNALSThe upper plenum of a PWR pressurevessel (above the upper core plate)contains a collection of support columnsand control rod guide tubes. When thereactor core begins to heat up, afterthe LOCA, the temperatures in thisregion would be about 500 to 600 F.During the core meltdown period, hotsteam and hydrogen flowing from the corewould heat the internal masses. Thisheatup process would also be assisted bydecay heat from fission products thatare carried into the upper plenum.From thermal analysis of core heatup(Ref. 3) a typical water boiloff ratefor the meltdown period is 0.01 lb/minper channel or about 400 lb/min for theentire core. Also from core thermalanalyses the average temperature of thissteam as it entered the upper plenum canbe obtained. These data, listed inTable VII H-3, indicate average gastemperatures of 2000 to 3000 F duringthe first half of core meltdown andincreasing temperatures afterwards. Thepoint of interest here pertains to thefirst half of core meltdown; specificallyhow rapidly the steam would heat thestructural internals of the upper plenumto temperatures in the range of 1000 to1500 F. Therefore, a simple model wasformulated to solve the heat transferproblem.The plenum was considered to be a wellmixed chamber with steam entering atF lb/hr and at constant temperature Tiand leaving at the same rate but at alower temperature TS. Inside the chamberthe structural masses are at temperatureTMO initially and heat to increasingtemperatures denoted by TM as afunction of time. On this basis thefollowing equations can be written:dTMMMCM dt hA - (T TM)dTST-TVpsCs dt - FCS i - ) - HA (Ts - TM)whereA hA1 MMCMF + hA 2 + X3 Vp-- VpsCS 2 4Here,FhA2 =Vp s- VsCS p4MM = mass of structural masses, lbCM = heat capacity of the solid,Btu/lb, Fps = steam density, lb/ftCS = heat capacity of steam, Btu/lb, FV = volume of chamber, ftF = steam delivery rate, lb/hrA = surface area of masses, fth = heat transferBtu/hr, ft2 , F.The solution for TM as atime for this problem is:I LTMO (X 3 -a 2 )TM Ti +a1-a2+332coefficient,functionexp (-a 2 t)X 1 Ti 012-a) 1a 1 (a 2 -a 1 ) exp (-a 1 t)oforwheredTM=t 1 S 1 M(a 1 +X 3 )2+ (X 1 +X 3 ) 2+ 4 - l 2dT S= 2 Ti - 3 TS + 4 TM(A 1 +X 3 )a2 2( X 1 + X 3 ) 24 11 2 .VII-150

The heat transfer coefficient, h, wasestimated from a dimensionless equation(Ref. 4) for gases flowing normal tobanks of staggered tubes for DoGmax/Pffrom 2000 to 32,000. ThuswhereZP =M =T =vaporization rate in g/cm 2 /secvapor pressure in atmospheresmolecular weight in gramstemperature in degrees Kelvin.hD 0 = .3[p 0.33 FDo~max]0.6The value obtained for h at a steamdelivery rate of 400 lb/min or 24,000lb/hr was 26.8 Btu/hr, ft, F. Valuesof TM as a function of time (t) werecalculated for the case; TMO = 500 F attime zero and an inlet steam temperatureTi = 2000 F. The results are given inTable VII H-4 along with the values usedfor the various parameters. It can beseen that upper plenum structural temperaturesare predicted to exceed 1000 Fwithin 15 min and 1500 F within 45 min.Considering these results in conjunctionwith the data shown in Table VII H-3, itappears that structural surfaces in theupper plenum would be well above 1000 Fand probably exceed 1500 F before halfthe core has melted. This is becauseaverage steam temperatures continue torise as melting occurs and heating priorto the onset of melting would raise thematerials temperatures above the 500 Fvalue used in the heat-transfer calculation.H3. ESTIMATED VAPORIZATION OF CORESTRUCTURALS AND BEHAVIOR INUPPER PLENUMThe structural components in and nearthe molten core region will vaporizeduring the meltdown period. In factthese vapors may comprise most of thetotal mass which is carried by steamflow into the upper region of thepressure vessel. Cooling of the hotvapor-steam mixture as it mixes withsteam that has passed through unmeltedsections of the core will create supersaturatedvapor conditions. Assumingthis causes particles to form it isnecessary to estimate the particulateconcentrations that could occur and todetermine how the resulting particulatesurface area compares with the upperplenum system surface area.The first step in the analysis requiresan estimate of the core vaporizationrates for important components. Thiswas made by using the Knudsen-Mayerequation modified for the presence of anexternal atmosphere (Ref. 5) . Thus,Z = 0.44 PMIT,Two vaporization temperatures (T) wereselected; 2800 C (the approximate meltingpoint of U0 2 and also close to theboiling point of iron) and 2300 C. Thelatter temperature takes into accountmelting point depression of the fuelthrough possible eutectic reactions andit is also well into the liquid regionfor molten iron (melting point about1550 C). In order to obtain a corevaporization rate the above equationmust be multiplied by an effectivesurface area. A value of 4 x 105 cm 2was used since this is about 1 percentof the total fuel rods surface area in alarge reactor core and it is also only afew times larger than the cross sectionalarea of the reactor pressure vessel.On this basis the potential vaporizationloss rates and other pertinent data forimportant materials were obtained asshown in Table VII H-5. It can be seenthat under the assumed conditions,nearly all the vaporized material willbe made up of tin and iron at both temperatures.Fuel vaporization contributesminor amounts and zirconium lossis insignificant. Therefore, zirconiumvaporization may be ignored during therest of the analysis.The potential vaporization loss rates inTable VII H-5 would not be realized ifinsufficient gas flow is available tocarry away vapors from the vaporizingsurfaces. Therefore, the rates must bemodified to take into account gas phasesaturation limitations. This requiresspecification of the gas flow rate inthe core, and since boiloff calculationsindicate steam flows in the range of5000 to 10,000 cfm (Ref. 1), the highervalue was used to correspond with thehigh core temperatures that were selected.Assuming ideal gas behavior, thepotential vaporization rates from TableVII H-5 would exceed saturation concentrationsat this bulk gas flow rate;that is, the 10,000 cfm steam flow istoo low to achieve the potential vaporizationrates. The rates that could beachieved at each temperature are tabulatedin Table VII H-6 along with thegas phase vapor saturation concentrationthese represent. It can be seen thattin still exhibits the highest massrelease rate, with iron a close second,and then U0 2 somewhat less. However,there is insufficient tin in a largereactor core (about 2.6 x 10 5 g) to sustainthese rates for the entire meltdownVII-151

period (about 4 5 minutes). Since ironand fuel vaporization are not so limited,the remainder of this analysis isbased only on these two materials.The vapor concentration values in TableVII H-6 can now be used to arrive atestimates of particle concentrations inthe upper plenum of a pressure vesselduring meltdown. However, one furtherfactor must be considered in developingthis estimate. In the early stages ofcore meltdown only a fraction of thetotal steam flow will pass through theregion of the core that is melting.Therefore, only a fraction of the vaporconcentrations given in Table VII H-6will occur in the upper plenum as thissteam mixes with the remaining steamthat has passed through cooler coreregions. On this basis a rather widerange of vapor concentrations ispossible for the upper plenum whichpractically could extend from about 2 x10-4 g/l up to about 2 x 10-1 g/l. Thisrange was used to estimate particleconcentrations, the lower value beingindicative of very early core meltdownand the higher one indicative of verylate core meltdown.According to experimental studies, themean size of particles formed when bothclad and unclad U02 are melted in gasstreams ranges from about 0.02 to 0.1 pam(Ref. 6). Assuming all the vaporizedmaterial condenses in the gas phaseforming 0.1 pam diameter particles, usinga particle density of 5 g/cc, andconsidering the effect of reduced temperaturein the plenum region on specificvolume (roughly a factor of two), thevapor mass concentration range givenabove indicates a concentration rangeof 0.1 Pam particles from 1.5 x 1011particles/l to 1.5 x 1014 particles/l.These concentrations can now be used toestimate particulate surface areas forthe upper plenum volume even thoughagglomeration would be expected tooccur. This is because in forming agglomerates,the surface area representedby the primary particles is probablyonly slightly reduced (less than afactor of two). Electron photomicrographsof particle agglomerates usuallyshow a very loose, open structure inwhich most of the primary particlesurface area is exposed to the gas (Ref.7). Therefore, assuming spherical particles,the specific particulate surfaceareas corresponding to the above concentrationscould range from about 4.5 x101 cm 2 /l up to 4.5 x 104 cm 2 /l. Thesevalues translate into a total particulatesurface area for the upper plenumvolume (about 1850 ft 3 ) of from 2600 sqft up to 2.6 x 106 sq ft. Since thestructural surface area of the plenumregion is estimated at about 3000 sq ft,the calculations indicate a particulatesurface area roughly equivalent to thestructural surface area at the beginningof core meltdown and tens or hundredstimes greater during following parts ofthe meltdown period.The above analysis of course assumesthat all the material vaporized in thecore region condenses to form particlesas the hot gas mixture cools in theupper plenum rather than condensing onthe cooler surfaces. This probably isnot the case but the process of coolingthe hot gas-vapor mixture through mixingwith cooler gases will rapidly producehighly supersaturated vapor conditionsdirectly in the gas phase. Such conditionsare favorable for particlenucleation and it can be shown fromcollision theory that collision frequencies,between species in the gas phasewill exceed collision frequencies withsolid surfaces. Therefore, it seemsthat particulate aerosols would bepresent and the surface area of theaerosol could exceed the surface area ofstructural components in the upperplenum region during much of themeltdown period. During the late stagesof meltdown, this condition is lessprobable because of the predicted highgas temperatures (see Table VII H-3).However, the steel structurals mightvery well have melted by this time.114. CONCLUSIONSThe three sets of analyses that havebeen presented lead to several expectationsregarding solid fission productbehavior in primary systems during coremeltdown. For example, it is doubtfulthat either cesium or tellurium would,through condensation, become associatedwith the particulate aerosol. By thetime appreciable aerosol concentrationsoccur the expected gas temperatures, andhence particle temperatures, are toohigh (over 1000 F) to allow cesium ortellurium condensation. However, in theearly stages of core meltdown cesium andtellurium vapors would probably plateouton the relatively cool (below 1000 F)structure in the upper plenum. Thenwithin a few minutes, as the structureheats above 1000 F, this condensedmaterial would vaporize and be swept outto the containment atmosphere. Henceplateout in the primary system isexpected to offer only a temporarydelaying effect in the escape ofreleased cesium and tellurium to thecontainment volume.VII-152

The behavior of strontium should besimilar to that of cesium and telluriumwith escape to containment delayed somewhatlonger by the need for higher upperplenum temperatures to vaporize the portionthat would experience earlyplateout. Some of the early condensationof strontium vapor could occur onaerosol particles however. This materialwould then be carried directly to thecontainment without significant delay.Materials like ruthenium (noble metals)or yttrium (rare earths) have sufficientlylow volatilities that condensationshould be expected in the plenumregion throughout most of the core meltdownperiod. Competition may occurbetween condensation on structuralsurfaces and condensation on or with theparticulate aerosol. The gas phase isexpected to be so highly supersaturatedthat co-condensation with structuralvapors from the core should be verycompetitive. In fact, these materialsmay enter the plenum region not asvapors but as very small particulatesand act there as condensation nuclei forthe structural vapors. In either case,it seems inappropriate to assumesignificant retention of the meltdownrelease fractions for these species inthe primary vessel.References1. Carbiener, W. A., et al., "Investigations of Reactor MeltdownReactors" Report on the Reactor Safety Study (Nov. 1973).inLight Water2. Nesmeyanov, A. N., "Vapor Pressure of the Elements", Translated and Edited byJ. I. Corasso, Academic Press, New York (1963).3. Carbiener, W. A., et al., "Investigation of Reactor Meltdown in LightReactors", Report on the Reactor Safety Study, Appendix A (Nov. 1973).Water4. Colburn, A. P., Trans Am Inst. Chem. Engnrs., 29, 174-210 (1933).5. Fonda, G. R., "Evaporation of Tungsten Under Various Pressures of Argon", PhysRev, 31, 260 (1928).6. Morrison, D. L., et al., "An Evaluation of the Applicability of Existing Data tothe Analytical Description of a Nuclear Reactor Accident", BMI-1779, p. 147-150(Aug. 12, 1966).7. Baurmash, L., C. T. Nelson and R. L. Koontz, "The Characterization of Aerosols byAerodynamic Parameters" in Treatment of Airborne Radioactive Wastes, IAEA,Vienna, p. 63-75 (1968).VII-153

TABLE VII H-1ESTIMATED SOLID FISSION PRODUCT PARTIAL PRESSURES AS AFUNCTION OF GAS TEMPERATURE AFTER LEAVING CORE REGIONEstimated(atm) forEstimated MolecularPatlIncte GsMeltdown Release Weight Partial Indicated GasRelease Rate Ib/ib- Pressure(f) TemperaturesSpecies Fraction lb/min mole 530F 980F 1520F 2060FCs (a)Te (b)Sr (c)Ru (d)(e)0.81 6.00.15 0.20.10 0.30.03 0.150.003 0.0051 33 6.5 x 10-3 9.5 x 10-3 1.3 x 10-2345 8.4 x 10-5 1.2 x 10-4 1.7 x 10-588 4.9 x 10-4 7.2 x 10-4 9.9 x 10-4103 2.1 x 10-4 3.0 x 10- 4.2 x 10-489 8.1 x 10-6 1.2 x 10-5 1.6 x 10-51.72.11.35.32.1xxxxx10-210-410-310-410- 5(a)(b)(c)(d)(e)(f)IndicatorIndicatorIndicatorIndicatorIndicatorEstimatedspecies for alkali metalsspecies for Te, Se, Sbspecies for alkaline earthsspecies for noble metal groupspecies for lanthanides and actinidesfrom ideal gas equation ignoring saturation restrictions.TABLE VII H-2 APPROXIMATE SUPERSATURATION RATIO FOR THE SOLID FISSIONPRODUCTS AS A FUNCTION OF GAS TEMPERATURESupersaturation Ratio - Partial Pressure InGas At Assumed Release RateSSaturation Pressure of the ElementSpecies 530 F 980 F 1520 F 2060 FCs 3.8 0.049 4.2 x 107 >5.3 x 10 7y >8.1 x 105 >1.2 x 10 6 1.5 x 106 310TABLE VII H-3: (a)AVERAGE EXIT GAS TEMPERATURES FROM CORE HEATUP ANALYSESAccident Approximate ApproximateTime (Min) T (gas) (OF) PercentCore Melt30 1400 040 2000 550 2600 2560 3300-3600 50-6070 4200-4700 60-9080 4600-5000 70-100(a)Range of values for higher core melt fractions due to differences in core heatupmodel assumptions regarding steam flow through melted regions.

TABLE VII H-4 UPPER PLENUM STRUCTURAL TEMPERATURES VERSUSBY STEAM ONLYTIME FOR HEATINGTime (min)Tm(OF)0615304560MM = 35, 130 lbCM = 0.17 Btu/lb, FPS = 0.0433 lb/ft 3CS = 0.5 Btu/lb, F50073210111349156917143V = 1850 ftF = 24000 lb/hr2A = 2186 fth = 26.8 Btu/hr,2ft , FTABLE VII H-5MATERIALS DATA AND POTENTIAL VAPORIZATION LOSS RATESPotential VaporizationMolecular Vapor Pressure, atm Loss Rate, g/secMaterial Weight 2300 C 2800 C 2300 C 2800 CU0 2 270 4 x 10-4 6 x 10-2 23 3100Iron 56 8.5 x 10-2 >1 2200 >24000Zirconium 91 5 x 10-6 4 x 10- 0.17 12Tin 119 1.7 x 10- 1 >1 6400 >35000TABLE VII H-6SATURATION VAPORIZATION RATES AND CONCENTRATIONSVaporization Rate, g/secVapor Concentrations, g/lMaterial 2300 C 2800 C 2300 C 2800 CU0 22.3 300 5 x 10-4 6.4 x 10-2Iron 110 >1000 2.3 x 10-2 >2.2 x 10-1Tin 460 >2100 9.6 x 10-2 >4.7 x 10-1Table VII H-1 - Table VII H-6VII-155/156

APPENDIX IIODINE DEPOSITION IN PWR AND BWR REACTOR VESSELSAND ASSOCIATED EQUIPMENTbyH. S.. Rosenberg and D. K. LandstromBattelge's Columbus LaboratoriesVII-157

Appendix ITable of ContentsSectionIi. CONSTANT RELEASE VERSUS PUFF RELEASE .............................I1.1 Constant Release Rate ......................................11.2 Puff Release ...............................................12. DETERMINATION OF THE OVERALL DEPOSIITION COEFFICIENT, Kt,AND CALCULATION OF DEPOSITION FRACTIONS ..........................12.1 PWR Reactor, Upper Internals ...............................12.2 PWR Reactor, Steam Generator ...............................12.3 BWR Reactor, Steam Separators ..............................12.4 BWR Reactor Steam Dryers ...................................13. SUMMARY OF RESULTS ...............................................REFERENCES ..............................................................Page No.VII-161VII-161VII-161VII-162VII-162VII-163VII-163VII-164VII-164VII-164List of TablesTableVII I-i Values of Ka, Kw, Kt, Md and fd for Various Conditions ofIodine Depo ition PWR Upper Internals ...........................VII 1-2 Values of Kg, Kw, Kt, Md, fd, Md (Total) and fd for VariousConditions of Iodine Deposition PWR Steam Generators ............VII 1-3 Values of Kg, Kw, Kt, Md and fd for Various Conditions of IodineDeposition BWR, Steam Separators ................................VII 1-4 Values of Kg, Kw, Kt, Md, fd, Md (Total) and fd Total forVarious Conditions of Iodine Deposition ..........................VII 1-5Definition of Terms and Units Used to Calculate IodineDeposition ......................................................VII 1-6 Summary of Typical PWR Surface Areas ............................VII 1-7 Summary of Typical BWR Surafce Areas ............................Page No.VII-165/166VII-165/166VII-165/166VII-165/166VII-167/168VII-167/168VII-169/170VII-159

Appendix IIodine Deposition in PWR and BWR Reactor Vesselsand Associated EquipmentH. S. Rosenberg and D. K. LandstromBattelle's Columbus LaboratoriesA model to predict the deposition offission-product iodine on the internalsand associated equipment of PWR and BWRreactors was developed and calculationswere made as a function of temperature,concentration, gas flow rate, gas composition,and puff versus constant releaserate. It can be shown that the amountof iodine deposited is independent ofthe release model and depends only onthe steam boil-off rate (Q), the surfacearea available for deposition (A), andthe overall deposition coefficient,(Kt). Further, it can be shown that thedeposition of iodine is controlled bythe wall deposition velocity (Kw), andfor all cases of interest Kt 2 Kw.I1. CONSTANT RELEASE VERSUS PUFFRELEASE 111.1 COMTS-ANT RELEASE RATEFor a constant release over a timeperiod, the concentration of iodine isgiven by the following equation:Cg CgU = W [1 -e-(Q/V)t](VIi I-1)where, Cg = concentration of 12 in gas,W = mass release rate of 12,Q = volumetric flow rate, V = volume ofchamber, and t = time period.The amount of iodine deposited is givenby:tMd = f C K tAdtg(VII 1-2)M Md- =WtA Q [t It + Q (eQ/V) t- 1 )]Now,by definition(VII 1-3)W = M /t (VII 1-4)where Mo = the total mass of 12, andsince the product of Q and t is largecompared to V, Equation (VII 1-3) reducestoMd = Mo K tA/Q (VII 1-5)11.2 PUFF RELEASEFor the case of a puff release of a mass(Mo) of iodine into a chamber of volumeV, the concentration is given by:Co = Mo/V (VII 1-6)Steam being producedremoves iodine atTherefore:anddC'at a rate (Q)a rate C'gQ.dC'dt6Cgdt % Cg (VII 1-7)dt Qc,/V. (VII 1-8)where, Kt = overall deposition coefficient,A = deposition area and Cg is asdefined in Equation (VII I-1).Inserting and integrating gives:Therefore:Co = Coe-Qt/Vg(VII 1-9)where t is the deposition time.1 See Table VII 1-5 for units and definitionsof terms.there-The amount of iodine deposited isfore:VII-161

Md =Md = KtACo0 CK tAdtf te-Qt/VdtM Md = KtACoV Q [e-Qt/VSince the product of Q andwith respect to V, thisreduce to:Md = K tACoV/Q.(VII 1-10)(VII I-ll)(VII 1-12)t is largeequation will(VII 1-13)Since from Equation (VII 1-6) Co = Mo/V,Equation (VII 1-13) further reduces toMd = Mo K tA/Q. (VII 1-14)Equation (VII 1-14) is the same asEquation (VII 1-5) for the constantrelease case so that Md is independentof the release model and depends only onthe steam boil-off rate (Q), the surfacearea available for deposition (A), andthe overall deposition coefficient, Kt.12. DETERMINATION OF THE OVERALLDEPOSITION COEFFICIENT, Kt,AND CALCULATION OF DEPOSITIONFRACTIONSAs will be described later, it was determinedfor both the PWR and BWRreactors that the deposition of iodineis essentially controlled by the walldeposition velocity, Kw, since the calculatedvalue of the mass-transportcoefficient, Kg is large with respect toKw. The following relationship existsbetween the coefficients:1 1 +Kt Kg w(VII 1-15)If K 9 is large with respect to Kw,Equation (VII 1-13) will reduce to asimple equivalent between Kt and Kw,i.e.:Kt = Kw (VII 1-16)12.1 PWR REACTOR, UPPER INTERNALSIn order to prove the validity of usingEquation (VII 1-16) it was necessary tocalculate values of Kg and compare themto previously determined values of Kw(Ref. 1). For the PWR reactor upperinternals this was done by using theChilton-Colburn analogy between heat andmass transfer to determine Kg, since anempirical equation is available to calculatethe heat transfer coefficient,hm, for fluid flowing normal to banks ofstaggered tubes. The heat transfer coefficientequation is (Ref. 2)hmDo= 0.33[Ckr [D1/3 [D 10ma-x]0.6(VII 1-18)where Do = outside diameter of tubes,Cp fluid heat capacity, p = viscosityoT =fluid, kf = thermal conductivity offluid, Gmal = mass velocity of fluidthrough minimum cross section.Using the definitions of ChiltonColburn (Ref. 3) for jH and JD:J - h mC pPUaveand[ 7Cp 1 2/3 (VII 1-19)jDýý [ p D v1 2/3(VII 1-20)JD =gmGmax vwhere hm, C, P, kf, Gmax are previouslydefined, K• = gas film coefficientPBm = logari~hmic mean of the partialpressure, p = density, Dv = diffusivityof iodine in fluid, Uave = mean fluidvelocity.Assuming jH = JD,K' g such thatK'U PKg = g ave Bmmaxdefiningadjusting the units ofIf the above is true then Equations (VII1-3) and (VII 1-12) can be written asfollows:Md = M oKwA/Q . (VII 1-17)C Pf= Prandtl Number = N and V= Schmidt Number = Nsc' thenVII-162

Thush Cp 2/3C LprJl] 2/3= g Nsc](VII 1-21)12.2 PWR REACTOR, STEAM GENERATORFor the deposition of iodine in thesteam generator, Gilliland's Equation(VII 1-7) which represents mass transferinside wetted-wall columns was used toobtain JD as follows:hr ]2/3 sc]2/3g PLprjmNI(VII 1-22)JD= 0.023 (Re)-0.17 (Nsc)0.11(VII 1-24)hm is obtained from Equation (VII 1-16)and the diffusivity term, Dv, in theSchmidt number is evaluated using thefollowing equation (Ref. 4)whereRe = Reynolds Number -1.858 x 10-3 T3/ 2 [ml + m 2 ]/2where= hydraulic radiusD DI1=2 12 ~ P GI1 2 QD D(VII 1-23)where T = temperature (K), ml and m 2 arethe molecular weights of the twocomponents, 2 DGI = Lennard-Jones forceconstant, and D = Collision integralfor diffusion.Values of Gl, and %D can be obtainedfrom tables listed in Reference 4 .Solving Equation (VII 1-22) for Kg atvarious conditions produce the valueslisted in Table VII I-1, which alsoshows the values of Kw taken from Reference1. Table VII 1-1 also shows theamount of iodine deposited (Md) fromEquation (VII 1-17) and the fraction depositedf= Md/Mo for various conditions.1i1 calculations assume an iodinemass release of 22 lbs (10kg),deposition surfaces of pre-filmed 304stainless steel, the atmosphere in thepressure vessel is well-mixed, and steamboil-off rates of 12 lbs/sec (3 x 106cc/sec) and 1.2 lbs/sec (3 x 105cc/sec). The area available for depositionwas estimated from scale drawingsand is approximately 3000 sq ft. for PWRupper internals (see Table VII 1-6).Calculations were performed for temperaturesof 250 F, 500 F, 1000 F, and2000 F. At the two higher temperatures,two iodine species, elemental iodine andhydrogen iodide, were considered alongwith different fluid compositions becauseiodine could conceivably reactwith hydrogen in the core region to formHI.and Nsc is the Schmidt number as previouslydefined. Therefore combiningEquations (VII 1-20) and (VII 1-21)0.023 Gmax -0 17 -056K; PBm (Re) (Nsc)By adjusting unitsft/hr. we have:Kg(VII 1-25)to obtain K g0.023 Gmax (Re)p0.17(Nsc)0.56in(VII 1-26)Solving Equation (VII 1-26) for Kg atvarious conditions produces the valueslisted in Table 1-2. Kw is taken fromReference 1 and Md was calculatedusing Equation (VII 1-17). All calculationsassume the same conditions usedfor the upper internal calculations exceptthat the deposition area is approximately51,500 sq. ft. Table VII 1-2also lists the Md (total) and fd (total)by summing the deposition from the reactorupper internals and the steamgenerators.12.3 BWR REACTOR, STEAM SEPARATORSA procedure identical to that used tocalculate deposition in a PWR steamgenerator was used to calculate theiodine deposition in the steam separatorsof a BWR. Gilliland's Equation(Ref. 5) was used with a suitableVII-163

modification of the Reynolds Number tothe steam separator geometry. The samesteam generation conditions as notedpreviously were used with a depositionarea of approximately 9,850 sq. ft.Since the results were expected to besimilar to the previous cases only thedepositions at 500 and 1000 F werecalculated. Table VII 1-3 lists thevalues of K 9 , Kw, Kt, Md, and fd forthese conditions.12.4 BWR REACTOR STEAM DRYERSA procedure similar to that used to calculatedeposition in the BWR steamseparators was used to calculate theiodine deposition in the steam dryers ofa BWR. Gilliland's Equation (Ref. 5)was again used with a modification ofthe Reynolds Number to the steam dryergeometry. The same steam generatingconditions as noted previously were usedwith a deposition area of 32,400 sq. ft.Again, deposition was calculated only at500 and 1000 F. Table VII 1-4 lists thevalues of Kg, Kw, Kt, Md, Fd (total),and fd (total) for these conditions.13. SUMMARY OF RESULTSIn all the cases examined for both thePWR and BWR primary systems, comparisonof Kg and I values show that iodinedeposition would be surface rate controlledeven at relatively low bulk gasflows. The results in Table VII 1-2also show that minimal deposition ofelemental iodifte should be expected inthe PWR primary system under LOCA conditions;that is, at temperatures above500 F and at bulk steam generation ratesin the range of 5 to 10 lb per second.In addition, considerably less than halfthe HI (if it should form) would beexpected to deposit under these conditionseither.Deposition predictions for iodine in theupper regions of a BWR vessel (TablesVII 1-3 and VII 1-4) show results thatare very similar to those obtained forthe PWR cases. Therefore, primary systemdeposition should not cause significantretention of iodine during itstransport to the containment region ineither reactor system.References1. Genco, J. M., W. E. Berry, H. S. Rosenberg, D. L. Morrison, "Fission-ProductDeposition and its Enhancement under Reactor Accident Conditions: Deposition onPrimary-System Surfaces", BMI Report No. 1863 (March, 1969).2. McAdams, W. H., Heat Transmission, McGraw Hill Book Co., pg. 272 (1954).3. Sherwood, T. K., R. L. Pigford, "Absorption and Extraction", McGraw-Hill BookCo., Inc.,*pg. 61 (1952).4. Reid, R. C., T. K. Sherwood, "The Properties of Gases and Liquids", McGraw HillBook Co., Incl, pg 523 (1966).5. Sherwood and Pigford, op. cit., p 78.VII-164

TABLE VII I-1 VALUES OF Kg, Kw, Kt, Md AND fd FOR VARIOUS CONDITIONS OFIODINE DEPOSITIONPWR UPPER INTERNALS (a)TemperatureF250250IodineSpecies1212FluidComposition100% H 20100% H 2 0SteamBoil-offRatelbs/sec12.01.2Kgft/hr339.085.2Kw(b)ft/hr0.3540.354Ktft/hr0.3540.352Mdlbs.6.13xi0-2fd2.79xi-036.10x10 1 - 2.77x10- 25005001000100010001000200020002000200012121212HIHI1212HIHI100% H2 0 12.050% H 21.250%100% H 2 0100% H201.212.0H20100%1.2H 2050% H2 12.050%250% H 21.2100% H 2 0100% H 2012.01.250%50%H2H 2 012.0233.7 0.023658.7 0.0236421.8106.0906.731.53x10-1.53xi0- 30.590227.7 0.590844.8212.11807.72. 01xl0- 42. 01xl0- 40.3540.02360.023631.53xi0-1.53xi0- 30.5900.5882. 01x10- 42. 01xlO- 40.3544. 09xi0- 31 .86x10- 44. 09xi0- 231.86xi0-2.65xi0-4 1. 20x10- 542. 65xi0-3 1.20x10~0.1024 .64x10-31.02 0.04633.48x10- 561.58xi03.48x10- 451.58x10"6. 13xi0- 2 2.79xi0- 3454.1 0.354 0.3546.13xl10 1 2.79x10 -2(a) Based on Deposition area of approximately 3000(b) Values obtained from Reference 1.sq. ft. - See Table VII 1-6.TABLE VII 1-2 VALUES OF Kg' Kw, Kt, Md, fd' Md (TOTAL) ANDCONDITIONS OF IODINE DEPOSITIONPWR STEAM GENERATORS (a)fd FOR VARIOUSSteamBoil-off Kw(d) Kt Md MdfdTemperatureF SpeciesIodinelbs/secRateft/hrgft/hr ft/hr lbs fdTotallbs Total2 5 0 (b) 12 12.0 520.0 0.354 0.354 1.05 4.78x10- 2 1.11 0.0512 5 0 (b) 12 1.2 76.9 0.354 0.354 10.52 4.78x10-I 1.09 0.51500(b) 12 12.0 327.0 0.0236 0.0236 7.02xi0- 2 3.19x10- 3 7.28xi0- 3 3.38x10-3500(b) 12 1.2 48.4 0.0236 0.0236 7.02x10 1 - 3.19x10- 2 0.0729 0.03.381000(b) 12 12.0 545.0 1.53xi0- 3 1.53xi0- 3 4.55xi0- 3 2.07x10- 3 4.82x10- 3 2.08xi0- 31 0 0 0 (b) 12 1.2 80.6 1.53xi0- 3 1.53xi0- 3 4.55xi0- 2 2.07x0.0- 2 0.023 0.0211 0 0 0 (c) HI 12.0 1142.0 0.590 0.590 1.754 7.97xl0- 2 1.855 0.08431000(c) HI 1.2 169.0 0.590 0.590 17.54 7.97xi0- 18.55 0.8442 0 0 0 (b) 12 12.0 972.0 2.01x10- 4 2.01xi0- 4 5.98x10- 4 2.72xi0- 5 6.33xi0- 4 2.88xl0-52000(b) 12 1.2 143.0 2.01xlO- 4 2.01x10- 4 5.98x10- 3 2.72x10- 4 6.33x10- 3 2.88xlo-42 0 0 0 (c) HI 12.0 2057.0 0.354 0.354 1.05 4.77x10- 2 1.11 0.0512000(c) HI 1.2 304.0 0.354 0.354 10.52 4.77xi0-I 11.1 0.506(a) Based on deposition area of approximately 51,500 sq. ft. - See Table VII 1-6.(b) Fluid composition, 100% H 2 0.(c) Fluid composition, 50% H 2 0 - 50% H 2 .(d) Values obtained from Reference 1.

TABLE VII 1-3 VALUES OF Kg, Kw, Kt, Md AND fd FOR VARIOUS CONDITIONS OFIODINE DEPOSITIONBWR, STEAM SEPARATORS (a)TemperatureF50050010001000IodineSpecies12121212FluidComposition100% H 2 0100% H 2 0100% H 20100% H 2 0SteamBoil-offRatelbs/sec12.01.212.01.2K gft/hr8.981,3314.342.12ft/hr(b)0.02360.023631.53x10-31.53x10-Xtft/hr0.02350.023231.53xi0-31.53xi0-Mdlbs.0.01340.1338.7xi0-438.7xi0-fd6. 90xi0- 46. 02xi0- 33 .96xi0- 53. 96xi0- 4(a) Based on deposition area of approximately 9,850(b) Values obtained from Reference i.sq. ft. - See Table VII 1-7.TABLE VII 1-4 VALUES OF Kg, Kw, Kt, Md, fd' Md (TOTAL) AND fd TOTAL FOR VARIOUSCONDITIONS OF IODINE DEPOSITIONBWR, STEAM DRYERS (a)SteamITemperatureF SpeciesIodinelbs/secRateft/hrgft/hr ft/hrKt12 12.0 2.131 0.0236 0.0233Mdlbs0.043fd1. 95x0.0-3MdTotallbs0.0564fdTotal2. 56xi0- 35 0 0 (b) 12 1.2 0.3156 0.0236 0.02201 0 0 0 (b) 12 12.0 4.088 1.53xi0- 3 1.53x10- 31 0 0 0 (b) 12 1.2 0.504 1.53xi0- 3 1.52xi0- 30.4122. 86x10 30.02860.0191. 3x10- 41.3x10-30.5453 .73x10- 30.03730.0251. 69xi0- 431.69x10-(a) Based on deposition area of approximately 32,400 sq. ft. - See(b) Fluid Composition, 100% H 2 0.(c) Values obtained from Reference 1.Table VII 1-7.Table VII I-i - Table VII 1-4VII-165/166

TABLE VII 1-5DEFINITION OF TERMS AND UNITS USED TO CALCULATE IODINEDEPOSITIONACC, gCoCpDoD12Dvf dGmaxGI12hmkfk gK;gKtKwMdM 0m1Im 2N prN scPPBmQReRHTtVUavepDPDeposition Area, ft2Concentration of iodine in gas (constant release), lbs/ft 3Iodine concentration (puff release), lbs/ft 3Initial concentration of iodine in gas (puff release), lbs/ft 3Fluid heat capav-ity, Btu/(lb) (IF)Outside diameter of tubes, ft.Diffusion coefficient (diffusivity of species1 through a mixture of 1 and 2) cm /sec.2Molecular diffusivity of iodine in fluid, ft /hrFraction of iodine deposited, dimensionlessMass velocity of fluid through minimum cross section, lbs/(hr) (ftLennard-Jones force constantheat transfer coefficient, Btu/(hr) (ft 2)thermal conductivity of fluid, Btu/(hr) (ft)mass transport coefficient, ft/hrIndividual or gas-film coefficient, lb-moles/(hr) (ft ) (atm)Overall deposition coefficient, ft/hrWall deposition velocity, ft/hrAmount of iodine deposited, lbsInitialmass of iodine, lbsMolecular weight of species 1Molecular weight of species.2Prandtl Number,Schmidt Number,Pressure, atmdimensionlessdimensionlessLogarithmic means of the partial pressure, atm.Volumetric flow rate (steam boil off rate) ft 3/hrReynolds Number,Hydraulic radius, ftTemperature,(*K)time period, hr.dimensionlessMean fluid velocity (volumetric flow rate, Q divided by cross section),ft/hr3Volume, ftCollision integral for diffusionfluid density, lbs/ft 3fluid viscosity, lbs/(ft)(hr)(F)(F)2

TABLE VII 1-6 SUMMARY OF TYPICAL PWR SURFACE AREASThermal ShieldSurface Area Inside) = 7.75 x 104 in2 = 538.2 ft2Surface Area (outside) = 8.0 x 10 4 in 2 = 555.8 ft 2Core Barrel (Lower)Surface Area (Inside) = 6.99 x 104 in2 = 485.5 ftSurface Area (Outside) = 7.20 x 104 in2 = 500.0 ft22Core Barrel (Upper)Surface Area (Inside) = 4.1 x 104 in = 284.8 ftSurface Area (outside) = 4.2 x 104 in2 = 294.1 ftCoolant opening area (3764 in2 = 26 ft)2 subtracted from above.22Top Support PlateUnder Surface Area = 17,066 in2 = 118.5 ft2 (hole area subtracted)Volume (minus holes) = 54,612 in 3 = 31.6 ft 3Mass = 15,834 lbs (for 304 SS)Upper Structural Support ColumnsExternal Surface area (per column) = 15 ft 2For 14 Support Columns = 210 ft2For 25 Support Columns = 375 ft2Volume of each Support Column = 4951 in3 = 2.87 ft3For 14 Columns (o40.1 ft 3For 25 Columns = 71.6 ftn3Mass of each Support Column = 1438 tbsFor 14 Columns = 20,130 fFor 25 Columns = 35,950 fControl Rod Guide TubesSurface Area (per tube) = 37.3 ft 253 Control rods = 1,976 ft2Mass (per tube) = 283 lbs53 Tubes = 15,000 lbs (total)Pressure VesselInside Surface Area(Upper Core Plate to 2Upper Support plate)364 ftAreas of water inlets and outlets subtractedBottom Support CastingAssuming approximately 50% open area3Volume = 43 ftMass = 21,650 lbsInstrument Guide TubesInformation not available to estimate total number of tubes.Estimated weight for each tube = 144.7 lbs.bsbs

TABLE VII 1-6(Continued)Steam Generator (Heat Exchanger U tubes) Single UnitInternal Surface Area = 51,500 ft2 (3338 U tubes)I.D. of tubes = 0.775 inLength of single tube = 914 in 76 feetSurface Area/tube = 2222 in 2 = 15.4 ft2 /tubeReactor Coolant PumpsInternal Surface Area (approx.) = 61.5 ft2 per pump (3 total)Piping (Single Loop)Approx. total length of pipingApprox. total Internal Surface AreaApprox. length Reactor to Steam GeneratorApprox. length Steam Generator to PumpApprox. length Pump to Reactor= 81 ft= 600 ft= 25 ft (29" I.D.)2= 22 ft (31" I.D.)= 34 ft (27.5" I.D.)Table VII 1-5 - Table VII 1-6VII-167/168

TABLE VII 1-7 SUMMARY OF TYPICAL BWR SURFACE AREASSteam DomeVolumeSurface Area= 1,560 ft 3-515 ft2CoreEmpty Volume (without bypass)Bypass VolumeFuel Rod VolumeFuel Assembly VolumeNet Core VolumeCladding Surface AreaChannel Surface Area= 2,420 ft 3-267 ft3= 221 ft= 99.1 ft= 2,100 ft 3= 59,100 ft= 29,800 ft322Core ShroudInside and Outside Surface AreaMass= 2,160= 84,122ftlb2Lower Plenum and Control Rod DrivesEmpty VolumeControl Rod Guide Tube Displaced VolumeNet VolumeGuide Tube Outside Surface AreaTotal Surface Area= 2,910= 426= 2,482= 3,410= 4,439ftftftftft33322Downcomer and Jet PumpsEmpty VolumeJet Pump VolumeNet VolumeJet Pump Outside Surface AreaTotal Surface Area= 1,510= 182= 1,328= 874= 3,244ftftftftft33322Jet Pumps Total Mass (20 Pumps)= 20,000lbSteam SeparatorsEmpty VolumeSeparator Displaced VolumeSeparator VolumeNet VolumeSeparator Outside Surface AreaTotal Surface Area4,3002,5002241,8008,9009,850ftftftftftft333322

TABLE VI1 1-7 (Continued)Steam DriersEmpty VolumeDrier VolumeNet VolumeDrier Surface AreaTotal Surface Area (approx.)= 2,970= 1322,83831,700= 32,400ftftftftft33322Recirculation PipingLength (approx.)Inside VolumeInside Surface AreaInlets (approx. length)Inlets (volume)Inlets (inside surface area)Suction Lines and Pumps Approx. LengthSuction Lines and Pumps Inside VolumeSuction Lines and Pumps Inside Surface Area= 80- 64= 254= 15= 75= 375- 100= 340- 920ftftftftftft3232ft each (2 total)ftft32Table VII 1-7VII-169/170

APPENDIX JTRANSPORT AND DEPOSITION OF AIRBORNEFISSION PRODUCTS IN CONTAINMENT SYSTEMSOF WATER COOLED REACTORSFOLLOWING POSTULATED ACCIDENTSbyA. K. Postma, P. C. Owzarski,and D. L. LessorBattelle Pacific Northwest LaboratoriesVII-171

Appendix JTable of ContentsSectionJil.MECHANISMS WHICH CONTROL THE REMOVAL OF AIRBORNE FISSION PRODUCTSFROM CONTAINMENT ATMOSPHERES .....................................J1.l Physical and Chemical Forms of Airborne FissionProducts ...................................................J1.2 Mechansims Expected to Control Depletion from the GasPhase ......................................................J1.2.1 Noble Gases ........................................J1.2.2 Methyl Iodide ......................................J1.2.3 Elemental Iodine ...................................J1.2.4 Aerosol Particles ..................................Page No.VII-175VII-175VII-175VII-175VII-175VII-176VII-176J2. NUMERICAL EVALUATION OF REMOVAL RATES ............................ VII-177J2.1 Noble Gases ................................................ VII-177J2.2 Methyl Iodide .............................................. VII-177J2.3 Elemental Iodine ........................................... VII-178J2.4 Aerosol Particles .......................................... VII-180J3. CONTAINMENT MODELS ............................................... VII-180J3.1 Single Volume Containment Model ............................ VII-180J3.2 Multicompartment Containment Model - PWR ....................... VII-181J3.2.1 Natural Deposition .................................J3.2.2 Spray Removal ......................................J3.2.3 12 Equilibrium ......................................J3.2.4 Intercompartmental Flow Rates ..........................VII-183VII-184VII-185VII-185J3.3 Multicompartment Containment Model - BWR ....................... VII-186J3.3.1 Natural Deposition - Effect of Heat Transferfrom BWR Vessel Walls ..............................VII-187J3.3.2 Pool Scrubbing .....................................VII-187J3.3.3 Standby Gas Treatment System - (SGTS) ................. VII-188J3.3.4 Natural Deposition in the Drywell AnnularAirspace ...........................................VII-188J3.4 Computer Code Corral ....................................... VII-190REFERENCES ..............................................................ADDENDUM ................................................................VII-192VII-197VII-173

List of TablesTableVII J-lFission Product Removal Rate Constants Calculated for a LargePWR Containment Vessel ...........................................Page No.VII-193/194VII J-2 Equilibrium Data for 12 with Boric Acid Sprays ..................VII J-3 Equilibrium Data for 12 with Caustic Sprays .....................VII J-4 DF Versus Particle Size and Reynolds Number in Drywell AnnularGap..............................................................VII J-5 Fission Product Releases ........................................VII-193/194VII-193/194VII-193/194VII-193/194List of FiguresFigureVII J-1 Partitioning of Elemental Iodine by Recirculated Sprays .........Page No.VII-195/196VII J-2VII J-3VII J-4Schematic of Quantity of Airborne-Contained Fission ProductsVersus Time .....................................................Schematic of Quantity of Fission Products Released to AtmosphereEVersus Time .....................................................Drop-Collection Efficiency as a Function of Liquid VolumeSprayed .........................................................VII-195/196VII-195/196VII-195/196VII J-5 Gas Flow for PWR Cases ..........................................VII J-6 Flow Schematic for a BWR - (Possible Atmospheric Source) ........VII J-7 Computer Code CORRAL Flow Diagram ...............................VII-195/196VII-195/196VII-195/196VII J-8Typical Sequence of Spike Fission Product Releases forPostulated Accidents ............................................VII-195/196VII-174

Appendix JTransport and Deposition of AirborneFission Products in Containment SystemsOf Water Cooled ReactorsFollowing Postulated AccidentsbyA. K. Postma, P. C. Owzarski,and D. L. LessorBattelle Pacific Northwest LaboratoriesIn this report, fission product transportand deposition in containmentsystems of water cooled reactors is analyzed.The scope of the work involvesprediction of fission product behaviorfrom the point of release from the primarycoolant system to the entry to theearth's atmosphere.This study is one element in the FissionProduct Source Term Task of the USAECReactor Safety Study. The goal of theReactor Safety Study is to realisticallyevaluate the risk to the public of postulatedcatastrophic accidents in watercooled nuclear reactors.This report contains a description ofremoval mechanisms which will depletefission products airborne in containmentsystems, a listing of the calculationalmethods used to predict removal rates,and a description of a computer code forcalculating atmospheric release of airbornesubstances from a multicompartmentedcontainment system.JI. MECHANISMS WHICH CONTROL THEREMOVAL OF AIRBORNE FISSIONPRODUCTS FROM CONTAINMENTATMOSPHERESJ1.1 PHYSICAL AND CHEMICAL FORMSOF AIRBORNE FISSION PRODUCTSThe mechanisms which control depletionof a specific fission product specie dependon its physical and chemical form.For example, gaseous fission productforms such as noble gases, organic iodidesand elemental iodine are removedat widely different rates. because ofgreat differences in equilibrium solubilityin water. Except for the noblegases and halogens, fission productswill be present as solid particles.Fission product forms considered inpresent study include the following:thee elemental iodineo methyl iodide* noble gases& aerosol particles.These categories are expected to be sufficientlybroad to include all fissionproducts of importance.J1.2 MECHANSIMS EXPECTED TO CONTROLDEPLETION FROM THE GAS PHASE 1J1.2.1 NOBLE GASESFor the accident cases considered, noremoval mechanisms of importance werepresent. Therefore, removal of noblegases from the gas phase was consideredto be negligible. It is recognized thatseveral means for collecting noble gaseshave been proposed. These include clathratecompound formation, trapping bysurface active agents such as refrigeratedcharcoal, separation by selectivemembranes, and absorption by fluorocarbonsolvents. If any of these processeswere considered operable following postulatedaccidents, a removal term wouldbe added to the present formulations.Jil.2.2 METHYL IODIDEMethyl iodide and other alkyl halidesare relatively unreactive and are onlyslightly soluble in water. Organic iodidesare removed by hydrolysis inwater, by adsorption on special reactivepaints, by solution reaction with reactivespray additives such as sodiumthiosulfate, and by activated charcoalfilters.iThe discussion here is oriented mainlytoward PWR systems but portions applyalso to BWR systems. These plus additionalBWR processes are describedlater in the report.VII-175

For accident cases where engineeredsafety systems do not function, removalwill be governed by hydrolysis in waterand by adsorption on paint. The removalrate due to both of these mechanisms isslow, being equivalent to a removalhalf-time of many hours. Thus methyliodide removal may be neglected forthese accident cases.Removal of methyl iodide by both causticand boric acid spray is primarily by hydrolysisof methyl iodide in the spraysolution. This hydrolysis rate is slow,and, like the case for no spray, the removalrate is usually negligible.Addition of 1 percent by weight of sodiumthiosulfate to the spray solutionmarkedly increases the methyl iodideremoval rate compared to that for boricacid or caustic spray. For the thiosulfatespray, methyl iodide removal is theresult of adsorption with simultaneouschemical reaction into wall films andinto falling spray drops.Activated charcoal filters such as thoseused in a recirculating air cleaningsystem remove methyl iodide at appreciableefficiencies. For the saturatedair-steam atmospheres encountered underaccident conditions in PWR's, removalefficiencies in the neighborhood of 70percent are anticipated. Thus the removalrate will depend primarily on theair flow rate through the filters.J1.2.3ELEMENTAL IODINEElemental iodine is reactive in aqueoussolution and may be rapidly absorbed bysprays. The absorption rate for boricacid is limited by the gas-liquid equilibriumpartition coefficient. Forcaustic sprays the absorption rate iscontrolled more by gas phase mass transferresistance than by equilibrium.Liquid phase mass transfer resistance isnegligible for sprays containing 1 percentby weight of sodium thiosulfate.Activated charcoal filters are quite efficientfor removal of elemental iodine.A removal efficiency of 99 percent couldbe expected for a charcoal filter installedin a recirculating filter loop.The removal rate would be controlledprimarily by the air flow rates.natural convection boundary layer arerelatively small, so the removal rate bynatural deposition is small compared tothat which can be achieved by reactivesprays.J1.2.4AEROSOL PARTICLESElemental iodine is also removed by naturalprocesses if engineered safety systemsdo not work. The removal rate maybe estimated from models in which iodinemass transfer is limited by diffusionthrough a gas boundary layer which flowsalong the walls of the containment vessel.Mass transfer coefficients for theAll non-gaseous fission products whichbecome airborne will be present in theform of aerosol particles. Due to therelatively high mass concentration ofairborne particles, the particles willconsist of numerous primary nuclei agglomeratedinto larger particles. Theagglomerate particles would be expectedto contain primary particles of the manyfission product nuclides present. Thusit is expected that the transport characteristicsof all solid fission productswould follow the behavior of theaerosol particles.Particles may be removed by sprays, byfilters, by deposition onto surfaces,and by gravitational settling.Mechanisms which cause collection ofparticles by spray drops include inertialimpaction, interception, diffusiophoresis,and diffusion. Prediction ofthe collection rate requires knowledgeof the single drop collection efficiency.In the present study, the singledrop collection efficiency was calculatedfrom large scale spray experimentsin which particle removal was measuredunder simulated accident conditions.This means for estimating the collectionefficiency is expected to result in anunder-prediction, because in the postulatedaccidents higher aerosol concentrationsfavor formation of larger,easier-to-remove particles.Particulate filters used in recirculatingfilter systems are highly efficientin removing airborne particles of allsizes. Therefore, recirculating filterssystems, if included in a containmentsystem design, could be relied on toremove more than 99 percent of airborneparticles per pass.Under natural transport conditions,gravity settling of particles on horizontalsurfaces, and turbulent depositionon vertical surfaces represent theprimary removal mechanisms. Experimentshave demonstrated that gravity settlingwill dominate in the post accident containmentatmosphere. In the presentstudy, the particle gravity settlingvelocity, and hence the aerodynamic diameter,were obtained from experimentaldata obtained in large scale containmentvessels.VII-176

J2. NUMERICAL EVALUATION OFREMOVAL RATESJ2.1 NOBLE GASESFor the accident cases which have beenstudied to date, no specific means havebeen included to remove noble gases.Since these gases would not be removedto a significant extent by sprays, filters,or natural deposition, the removalrate from the total system has been setequal to zero. Removal from a specificcompartment of the containment vesselmay occur by convective flow of gas, butthis output term would represent an inputterm to the connected compartment.J2.2 METHYL IODIDEMethyl iodide is removed only very slowlyunder conditions of natural depositionor where sprays of boric acid orcaustic are used. Removal half-timesare many hours for these three cases,hence methyl iodide removal may be neglectedwithout excessive error.Appreciable removal of methyl iodidewould occur if a recirculating charcoalfilter system were incorporated in thecontainment system. Results obtained inthe Containment Systems Experiment (Ref.1) have demonstrated that a removal efficiencyof approximately 80 percent isachieved for methyl iodide when saturatedsteam-air atmospheres pass throughcharcoal filters impregnated with 5 percentiodine. The removal rate constantsystem may be express-for such a filtered aswhere(dc)f= Cfc = FE CfC(dC)= removal rate due to filtersystem,,(VII J-l)X = removal rate constant forfilter system,F = gas flow rate through filter,E = fractional removal efficiency,C = airborne concentration ofremoved species.Based on the CSE tests, E should be chosenas 0.8 for methyl iodide, and equalto 1.0 for both elemental iodine andaerosol particles.Methyl iodide is removed at an appreciablerate by sprays containing sodiumthiosulfate at a level of one weightpercent. The removal rate constant forreactive drops may be predicted (Ref. 2)bywhereXdropsFH (V+kte)Xdrops = methyl iodidespray drops,vF = spray flow rate,(VII J-2)removalV = volume of containedphase,bygasH = equilibrium partition coefficient,k = first order solution reactionrate,te =drop exposure time.This equation applies for a well mixeddrop model, and is expected to realisticallyrepresent the situation prevailingfor containment sprays.In addition to removal by spray drops,methyl iodide will also be removed byspray liquid which has wet the wall ofthe containment vessel. If the wallfilm is assumed to be stagnant, themethyl iodide absorption rate is givenbywhere(wall)- AH Ak_ tanh [ (k/D)6JV(VII J-3)A(wall) = removal rate constant forwall film,A = surface area wet by wallfilm,H = equilibrium partition coefficient,D = diffusivity of methyl iodidein spray liquid,= wall film thickness,VII-177

V = volume of containedgases.The net removal rate will be the sumthat for wall film and spray drops:FH (l+kte) AH /kEtanh [(k/-•)61V + Vof(VII J-4)Predictions based on equation (VII J-4)are in good agreement with large scaleexperiments (Ref. 2).J2.3 ELEMENTAL IODINEElemental iodine is removed by naturaldeposition at a rate which is controlledby mass transfer through the gas boundarylayer at the walls of the containmentvessel. The natural depositionremoval rate constant is simplywherek A(NT) V gV(VII J-5)k = 0.13 (Gr Sc) 1/3D V(VII J-7)Laminar flow persists until the Grashofnumber exceeds a transition value. Accordingto an analysis presented byHales (Ref. 4) transition from laminarto turbulent flow occurs for lengthswhich correspond to Grashof numbers between1.5 and 10-8 and 1.5 x 1010. Forcontainment atmospheres, this transitionwould occur at distances 5 to 10 ft downthe vertical wall. (Ref. 5) Thus Equation(VII J-6) would apply for the first10 ft, and Equation (VII J-7) wouldapply for the remainder of the distance.Hilliard and Coleman (Ref. 5) have providedestimates of iodine removal bynatural deposition based on use ofEquations (VII J-6) and (VII J-7). Theairborne concentration is predicted tobeC_ g-= 0.96e- 1.42t + 0.04e-0.0143tC go(VII J-8)JLiX(NT) = removal rate constant fornatural deposition,A = surface area for masstransfer,V = volume of contained gases.whereC airborne concentration,gC = airborne concentrationgo time zero,atKnudsen and Hilliard (Ref. 3) have discussedmethods for calculating k forcontainment vessels. The value of ýg isgoverned primarily by the Grashov numberfor natural convection. For laminarflow, the mass transfer coefficient isgiven bywherek 9,cD= 0.59 (Gr Sc)1/ 4 (VII J-6)kc = mass transfer coefficient,k =length along surface,D = diffusivity of iodine in gasv phase,Gr = Grashov no. for wall boundarylayer,Sc = Schmidt no. for iodine insteam.For turbulent flowt = time in hours.The first term in Equation (VII J-8)represents removal by mass transferthrough the wall boundary layer. Thesecond term accounts for an approach topseudo-equilibrium between the gas andliquid phases. The initial removal halflife for a large vessel as predicted byEquation (VII J-8) is approximately 30minutes.For spray removal of elemental iodine,the removal rate may be predicted from amodel in which the spray is consideredto be an assemblage of noninteractingsingle drops. The overall drop absorptionprocess includes the followingsteps:" mass transfer across the gas film" equilibrium dissolution at the gasliquidinterface" diffusion into the drop" reaction within the liquid phase.VII-178

0A conservative model for spray absorptionmay be formulated under the assumptionthat the drop consists of an outerstagnant film and a well-mixed interior.For this model, the absorption efficiency(defined as fractional saturationachieved by a drop falling through thecontainment vessel) is given bywhere6kE=l1- exp- -kt ekL+(VII J-9)the value attained in a few seconds.This time specification is necessarybecause iodine dissolution is greatlyinfluenced by chemical reactions. Onlythose reactions which are fast enough toconvert elemental iodine to non-volatileforms within a time period shorter thanthe drop exposure time will aid sprayabsorption.On the basis of available information,Postma and Pasedag (Ref. 6) have concludedthat applicable H values are asfollows:* for caustic, pH 9.5, H = 5000a for boric acid, pH = 5, H = 200E = drop absorption efficiency,0 for basicH = 100,000.sodiumthiosulfate,3k D (2 + 0.6 Re0"5 Sc0.3 )g d= gas phase mass transfercoef.,2 2 DLkL = 3d - liquid phase masstransfer coefficient,t e= drop exposure time,H = equilibrium partition coefficient.The basis and assumptions involved informulating Equation (VII J-9) are discussedin detail by Postma and Pasedag(Ref. 6).The spray removal rate constant may berelated to the spray flow rate and dropabsorption efficiency byThe first order removal rate predictedby Equation (VII J-10) will not continueto apply for an indefinite time becauseof equilibrium between the gas and liquidphase. For the first two hours followingspray initiation, equilibrium maybe accounted for by limiting spray removalto 1 percent of the release concentration(Ref. 6). Beyond the timeneeded to remove 99 percent of airborneiodine, the concentration should be assumedto remain constant at 1 percent ofthe release concentration.For times longer than two hours, the gasconcentration may be allowed to decreasefurther as a result of chemical reactionsin the liquid phase. For equilibriumconditions, which would be attainedfor times longer than a few hours, theairborne iodine concentration is givenbyC 9C gowhere1 VL1 + H V --g(VII J-ll)=F H Es V (VII J-10) C = equilibrium airborne concengtrationwhere C = initial airborne concentrationfor puff release,goX = spray removal rate constant,H = equilibrium partition coefficient,E = drop absorption efficiency,V = volume of contained gas.The partition coefficient appearing inEquations (VII J-9) and (VII J-10) isH = equilibrium partition coefficient,VL = volume liquid phase,V g = volume of gas phase.The partition coefficient, H, in Equation(VII J-ll) increases with time dueto kinetically slow liquid phase reac-VII-179

tions. In this study, the variation ofH with time was taken from experimentalresults reported by Postma, et al. (Ref.7). Experimental results for causticand boric acid solutions are shown inFig. VII J-l. The data presented inFig. VII J-l were used to conjunctionwith Equation (VII J-ll) to predict longterm behavior of elemental iodine incaustic and boric acid spray.For spray containing 1 percent sodiumthiosulfate, partition coefficients willbe large. As discussed by Postma andPasedag, (Ref. 6) the cutoff concentrationfor sodium thiosulfate sprays maybe taken as 0.1 percent of the releaseconcentration.J2.4 AEROSOL PARTICLESParticles are removed by natural transportprimarily by gravitational settling(Ref. 5). The removal rate constant isrelated to the particle settling velocitybywhere= Ut AcsV= removaltling,constant forA = cross sectionalcs vessel,(VII J-12)areaV = volume of contained gases,set-Ut = terminal settling velocity ofparticles.Results from CSE experiments (Ref. 5)indicate that soon after fission productrelease, particle diameter was 15 microns.A few hours later, average particlesize decreased to about 5 microns.The CSE results indicate that aerosolparticles will be removed slowly bynatural processes. Because of higheraerosol mass concentrations in postulatedcore meltdown accidents, the CSEresults will tend to underpredict removalfor the meltdown accidents. We havenot attempted to account for this differencein the present study.For recirculating filter systems, particleremoval would be described by Equation(VII J-l) with an efficiency fractionof unity.Spray removal of aerosol particles maybe predicted from a model in which thespray is considered to be an assemblageofof non-interacting single drops. Theoverall removal rate is the sum of removalrates for individual drops. For awell-mixed gas space, the removal rateconstant is expressible aswhere3 hF E2T V(VII J-13)= spray removal rate constant forparticles,h = drop fall height,F = spray flow rate,E = drop collection efficiency,d = mean drop size,V = volume of gas phase.The magnitude of the drop collectionefficiency, E, depends on a number offactors as discussed earlier. Particlesize of the aerosol is the most importantsingle parameter, and size wasestimated from CSE spray experiments.In applicable CSE tests (Ref. 8) theeffective particle collection efficiencyvaried from a maximum value (E = 0.06)achieved early to a minimum value(E = 0.0015) reached after most of theaerosol was removed. In single volume,hand calculation models an average valueof 0.02 may be used. After the airborneconcentration reduces to 2 percentof the release concentration, the dropcollection efficiency used in the calculationshould be lowered to 0.0015. Inmulticompartment containment models, Eshould be allowed to vary with the degreeof removal as was observed in theCSE experiments.Examples of removal rates calculated fora large PWR containment vessel are shownin Table VII J-l. These values wereused in preliminary hand calculationswith a single volume containment model.J3. CONTAINMENT MODELSJ3.1 SINGLE VOLUME CONTAINMENTMODELThe single volume containment model assumesthat the vapor phase consists ofone well-mixed compartment. This assumptionenables one to write the followingsingle differential equation foreach species:VII-180

dt = - LijCi - N C + Ri(t)Initial Condition: C. = C. (t') at1 1t = t'(VII J-14)C. = Airborne moles of component1 iX.. = Removal rate constant via1J mechanism ja . = leak rate, fraction of the1 volume/timeRi(t) =source term (moles/time)This equation is easily solved for constantXij, ai and Ri to getR.i= (Zij + i)+R.(Exij + i)summarized in Table VII J-l. Typicalcurves describing the moles airborne andescaped are shown in Fig. VII J-2 andVII J-3.The only release of fission productsencountered where R. = constant for sometime period is the melt release. Thissimplifies most of the computations tousing an initial condition for a timeperiod where the Xij's and ai's are consideredto be constant. If the X's ora's changed at some time, t', a new t'can be considered as the beginning of anew time period of constant new X's orals.J3.2 MULTICOMPARTMENT CONTAINMENTMODEL - PWRTo better simulate containment geometryand time variable removal rates, a modelwas developed.to analyze the atmosphericsource from a set of compartments whoseairborn contents are well mixed and arealtered by intercompartmental flow inaddition to deposition rates, leakrates, spray removal rates, etc. Thissystem is described by a set of equationsof the form- Ci(t')] exp - (ZXij + ai) (t-t')dh dCTr/d 1 rey= Hm •imr(VII J-17)(VII J-15)whereThe escaped amount duringt-t' is the integralQi =dtthe intervalC 1 = air borne fraction of initialrelease of material m inrelease type r contained incompartment iwhere V is the volume of the vessel.Thus= (Xmr + E Gki/Vk)Siji k k/k6]er 1j+ (Gji/Vj) l-Em )Ri VaiQi (E i lij + ai)F R.[(7Tij + ai)- Ci(tX ij + ix l-exp [-(EXij + ai)(t-t')](VII J-16)The two equations for Ci(t) and Qi(t-t')can be used to calculate airborne fractionsand leakage fractions for variousaccident sequences by hand. The valuesof Xij, used for various species arediscussed earlier in this report and,Tr= removal coefficient from compartmenti by settling, leak,spray removal, and otherprocesses not involving flowto or from another compartmentG.. = volume flow rate from com-3. partment/compartment j toE.V.13compartment i= filter removal fraction formaterial m from release r= volume of compartment j= 1if i = j, 6. = 0 if i 3j1JVII-181

Note that the flow terms assumemixing within each compartment.uniformFor a given release type and materialtype, the equation set can be written inmatrix notationdCd-- = HC(VII J-18)Where C is the column vector of airbornrelease fractions in the respectivecompartments and H is the matrix of rateconstants governing the evolution. Thesolution of this differential equationfor the column vector C for constantcoefficients in the H matrix is givenexactly byorC(t) = eHt C(o)(VII J-19)/e(t-to) HiC(t) = C(t ) + (t-t )H ((tIto)H C(to).0 0 Ft t 0 )H / 0(VII J-20)Fast computer techniques for generatingthis solution have been developed by B.H. Duane (Ref. 9) and were utilizedhere. By calling the numerical integrationsubroutines developed by Duane ateach time step after calculating the Hmatrix, the accuracy limit becomes thatimposed by the assumption of constant Hmatrix elements within the time step.Numerical error in generating the solutionto the coupled set of equationswith constant coefficients can be madeon the order of 10-6 percent.To integrate the amount leaked withinthe time step, N additional fictitiouscompartments were defined whose functionis to accumulate the leaked materialfrom the N real compartments. The fractionsCi+N, i = 1, 2, -.- N, are cumulativeleaks obtained from the solution ofdd-t(Ci+N) = aiCiwhere ai is the leak rate coefficientfrom compartment i. Note -ai is onepart of Hii. The previously defined N xN H matrix was augmented to form a 2N x2N H matrix according toH i+N,j = 06ijHi, j+N 0Hi+N, j+N = 0For i = 1,2, ",Nj = 1, 2, ", NIf one takes the material quantity inthe fictitous compartments to be zero atthe start of a time step, the fractionof the initial release material leakedduring the time step will be2Ni=NComputer code CORRAL1 was written tosolve the set of equations and at thesame time compute each rate parameter asa function of time and/or as a functionof vessel conditions (p, T, humidity).A more detailed description of the variousmodels used to compute rate parametersfollows.Nine time dependent parameters are computedfrom input data at the time ofsolution of the differential equations.These are:1. Compartment pressure, p(k), psig2. Compartment temp, T(k), F3. Compartment vapor mole fraction,VAP(k)4. Compartment wall-bulk temperaturedifference, DELTA T(k), F5.-6. Leak rates of molecular iodineand particulates from each compartment,ELI2(k), ELP(k),fractions/hr.7. Flow rates betweenG(j, k), ft3 /hr.compartments,8.-9. Filter decontamination efficienciesfor flow between compartments,EFI2(J,KO, EFP (j, k),dimensionless.The first four parameters enable one tocompute transport coefficients involvedin depletion rate coefficient calculations.1 Containment Of Radionuclides ReleasedAfter LOCA.VII-182

a-The bulk gas viscosities of steam-airmixtures (vm) are computed according tothe following equations (Ref. 10):whereU1A11 m y1 + Y sAsaIS+ sY As= viscosity of mixture1A =0.0414 [TR] , lb/ft/hr0.003339 (T, R)1"5Is = (TR + 1224.2)Ys== mole fraction of airmole fraction of steamVs/ (q 11[1 A 1/2 Ps1/4] 2OAs - / 1/22/7 +ýsA(VII J-21)above with subscripts reversedThe diffusivity of 12 in the steam-airmixtures were found using data andequations from Knudsen (Ref. 10).whereD = 1D12 YA YsDADsDA = 2.03 x 10-5 (T,K) 1.cm2/secDB = 3.24 x 10-5 (T, K) 1.5/(VII J-22)5 /P, atm) WA,DWilke-Chang relationship (Ref. 10),where= (7.4 x 10 8)(xM1)I/2 T(K) 2D V0.6 ,cm/secwhere(VII J-23)x = degree of solvent association= 2.6 for H 2 0M, = molecular weight of solventpL =solvent viscosity, cpVL = 100/{2.1484 [(T,(K)-281.6)+(8078.4 + (T, (k)-281.6) 2 )0.-1201 for H 2 0J3.2.1V = molar volume of diffusing substance= 71.5 cm 3 /gmole for12-NATURAL DEPOSITIONThe mechanism of natural deposition of12 is governed by diffusion with naturalconvection generated by temperature differencesbetween bulk gas and the wall(DELTAT(k)). Knudsen and Hilliard (Ref.3) claim that a mass transfer analogycan be made with correlations predictingnatural convective heat transfer coefficients.In similar manner, the modeluses expressions for Sherwood numbersfor laminar and turbulent flow using athermal Grashof number,Gr = Z3 (Twall-Tbulk)im /Pm)2 T bulkThus for laminar flow (Gr < 109), theSherwood number isSh - kc - 0.59 (Gr Sc) 1 /4D 12g5 ](VII J-24)(P, atm)/WS and for turbulent flow (109 < Gr < 1012)WA = 0.7075 + 141.73/T, KWS = 0.7075 + 454.72/T, KSh - k C z - 0.13 (Gr Sc) 1 1 3(VII J-25)The diffusivity of 12 in water (spraydrops) was computed using the standardwhere Sc = Schmidt number =IM/PM D 1 2VII-183

andZ =length of the wallX = F- 1 - exp - d(H + kg/k))k = mass transfer coefficient.cA combination of the two Sherwood numbersis used to compute the actualSherwood number. Since the turbulencein most cases occurs around 10 ft fromthe top of the wall, a weighted averagewas used.Sh(overall) =:£-I0 Sh(turbulent)+ 1 Sh(laminar).To convert the mass transfer coefficientinto a deposition lambda (Xij) is asimple step. Thus,Xij (k) = kc (k) A(k)/V(k)where A(k) and V(k) are suface area andvolume of compartment k, respectively.Another natural deposition process ofinterest is the settling of particulates.This involves a calculation ofterminal settling velocities, Vs, assumingspherical, unit density particles.d 2 (Pp- Pm) g/18 Pm = Vswhere/H at equilibriumF = spray flow rate(VII J-27)H = equilibrium distribution coefficient(Cg = C k /Hj at equil.)V = spray compartment volumed = spray drop diametert = drop residence time - heighte of fall/terminal velocityand the gas mass transfer coefficient,ka, is given by Ranz and Marshall (Ref.if) asandkg 1 {2.0 + 0.60 ReI/ 2 Scl/3,a gn(VII J-26)Particle diameters used were consideredas functions of time. Hilliard andColeman (Ref. 5) report that the settlingvelocities decrease with time afterrelease. In CORRAL it was decidedto use their data and assign an earlyparticle diameter (15p) and a late particlediameter (50) ("several hours"later = 4 hrs) and linearly interpolatebetween them. After 4 hrs., the particlesize was kept constant at the latevalue. To get the natural depositionlambda for particulates useX= A(floor area)Vs V(compartment volume)J3.2.2 SPRAY REMOVALThe spray removal model of I by boricacid and caustic sprays used in CORRALcombines a gas phase mass transfer coefficient,a drop-gas interfacial equilibriumdistribution coefficient and astagnant liquid film mass transfer coefficiet.The expression for the spraylambda is given by27'r 2k -Ef---, D, Z 12 diffusivity inliquid.The latter is the Griffiths model discussedby Postma (Ref. 6). Incorporatingthe latter is a more conservativeapproach when kg/k£ >5H. The terminalvelocities of the -falling drops arefound by matching the velocity independentdimensionless numberfD Re 2 = 4 pM (P -0 M) d 3 g/3M 2with the appropriate range of Reynoldsnumber Re. For spray drops the range 2 ofRe is 10-700. For 10

whereFhdVEspray flow ratespray fall heightspray diametercompartment volumespray collection efficiency.In CORRAL empirical results from CSEdata are used to predict E. Apparentlythe efficiency is a function of a normalizedliquid volume sprayed (totalvolume sprayed/total compartment volume- Ft/V). Figure VII J-4 shows the CSEdata, and the curve in this figure wasused to compute drop collection efficienciesin CORRAL. The diffusiophoresiswas subtracted from the efficiencyand the following expressions were fitto the remaining curve. To make theserelationships apply to a spray lambdaFt/V0 - 0.002 E = - 15.825 (Ft/V) + 0.055Ebeen experimentally determined. Inprogram CORRAL it has been possible toincorporate H = H(t) when equilibriumconditions exist.To get the equilibrium described quantitatively,an equivalent lambda for depletionof gas phase 12 had to be developed.Since the value of H increaseswith increasing time, the gas phase isbeing depleted as time goes on. To getthis equivalent lambda, we can write amass balance for 12. If Cgo is the initialairborn concentration, thenorCggo Vg = C + CgVg (VII J-29)CgVgC CgVo + CgVg CYVZq +g gV g1(VII J-30)0.002 - 0.0193 E = 0.04125 - [0.08626+ 42.68 (Ft/V) 1 2 /21.34Then for H = H(t), weremoval rate of 12 bycanwrite the0.0193 E = 0.0015with multiple sprays being used at varioustimes, the quantity ft/V is now thesum E Fi ti/V for any one release ofparticles. Each release of particleswould have its own spray aging relationship,and at this time no simple meansof typing sequential release togetherinto one relationship seems possible.Only when particle size distributionsare known throughout a spray aging processwill sequential releases be able tobe tied together.It should be noted that inspray cutoff is used at0.02; spray aging is used inJ3.2.312 EQUILIBRIUMC /Ctdt1 dH/ dg9where the equivalent lambda isC d at2= -VVgdHdtdt = - AdtCORRAL, no(VII J-31)C(t)/C(O) =its place. Data shows that H V£/Vg> >1 for boricacid and caustic solutions in equilibriumwithI2, so thatWhen airborn molecular iodine is depletedby either sprays or natural deposition,the depletion rate becomes independentof the two above mechanisms whenthe corpentration falls below about 1percent of the initial value (a conservativeestimate, i.e., a lower value isless conservative (Ref. 8). At concentrationsbelow this level, an apparentequilibrium situation exists where theconcentrations in liquid and gas phasesare related by an equilibrium distributionconstant, H = CZ/Cg. H is a functionof time (probably due to slow liquidphase chemical reaction) and has1 dHTypical data for spraysTables VII J-2 and VII J-3.J3.2.4(VII J-32)are shown inINTERCOMPARTMENTAL FLOW RATESIn two cases intercompartmental flowrates can be calculated. If circulationfans move air throughout the containmentvessel or if boil-off occurs with steamevolution into one compartment, one canuse these flow rates to provide a basisfor estimating intercompartmental flows.VII-185

In addition, natural convection drivenby wall-bulk gas temperature differencescan be a major contributor to flowrates. These rates can be estimated butwith a high degree of uncertainty.In this study, high flow rates have beenused to eliminate mixing as a parameter.The flows for PWR cases follow the schematicin Fig. VII J-5. Flow rates usedwere Q1 = blow down steam rate and Q 210 PWR containment volumes per hour.J3.3 MULTICOMPARTMENT CONTAINMENTMODEL - BWRThe multicompartment feature of codeCORRAL was essential to estimating atmosphericsource calculations from aBWR. A BWR containment system is not aset of openly connected compartmentslike a PWR, where the whole containmentsystem can be usually considered as"well mixed". The compartments of a BWRare usually closed to one-another andflows between them occur in complex waysduring postulated accidents.As many as six compartments are used insome BWR accident sequences. The firstfive are:a. The Drywell where natural depositioncan occur.b. The Wetwell where pool scrubbing andnatural deposition can occur.c. The Drywell annular gap where naturaldeposition can occur.d. The Reactor Building where naturaldeposition.can occur.e. The Standby Gas Treatment System (aseries of filters).The sixth compartment used was a fictitiousdumping ground for ground levelatmospheric sources when elevated(stack) sources occurred simultaneously.The schematic flow diagram for BWR accidentsis shown in Fig. VII J-6.Although the BWR containment model israther specific, it still must be consideredan approximate treatment ofactual conditions. For example, poolscrubbing of vapor-phase fission productsin the wetwell is modeled using asimple decontamination factor ratherthan a rigorous solution of the dynamictransport and absorption problem. Inaddition, the drywell annular air gap isused to simulate the entire lower regionof the secondary containment building.For situations in which leakage from theprimary containment boundary occurs directlyfrom the drywell or wetwell steelshells, this provides an accurate descriptionof the geometry and flow pathto the refueling floor level of the secondarycontainment. For cases in whichthe leakage site is in one of the severalsublevels of secondary containment(e.g., isolation valve failure), themodel represents an approximation of thegeometries and flow paths between thesevarious leakage locations and the refuelingfloor level. However, the methoddescribed later for calculating fissionproduct transport and deposition inthe annular region is expected to overestimaterather than underestimate fissionproduct concentrations that wouldreach the refueling building under suchconditions. This is because the combinationof longer residence time and depositionin the lower regions of secondarycontainment would usually be moreeffective than the decontamination factorspredicted for the annulus.Note in Fig. VII J-6 that potential releasesof fission products directly tothe atmosphere are indicated for allcompartments. Release from the standbygas treatment system is the normal mode,and this would be an elevated releasevia the plant stack. All other potentialreleases would be at ground level.Direct releases from the main reactor(refueling) building could occur becauseof failure to isolate the building orbecause gas flow into the building duringa particular accident might exceedthe exhaust capacity of the standby gastreatment system blowers. Direct releasesfrom the annulus (and/or lowersecondary containment) could occur ifhigh volume gas flow, caused by isolationfailure of a large penetration inprimary containment, would result instructural failure of the reactor buildingwall panels. This could also resultfrom rupture of primary containment,provided the event does not cause failureof an outside wall in the sublevelsof secondary containment. The alternative(i.e., structural failure of outsidewalls) could produce the directrelease from either the drywell or wetwellcompartments. In calculating suchdirect release cases, internal fissionproduct removal is considered only forcompartments up to and including the onefrom which the direct release occurs.Natural deposition is the most commonremoval mechanism for fission products,although it is not necessarily the mosteffective mechanism. Natural depositionof particulates occurs on horizontalsurfaces in all large compartments justas in a PWR. Turbulent deposition ofparticulates and iodine in the drywellannular air gap occur by different mech-VII-186

anisms to be discussed later. Naturaldeposition of iodine occurs on all surfaceswith the rate controlled by naturalconvection discussed in paragraphJ3.2.1. However, since the wall-bulkgas temperature-difference that drivesthe natural convection is highly variablein a BWR, transient heat transferanalyses are made to estimate thistemperature-difference.J3.3.1 NATURAL DEPOSITION - EFFECT OFHEAT TRANSFER FROM BWR VESSELWALLSThe mass transfer coefficient for 12 removal(see Equations VII J-6 and VIIJ-7), is proportional to the temperature-difference,Twall - Tbulk = ATw (VII J-33)raised to the 1/4 or 1/3 power. For anorder of magnitude range in ATw, themass transfer coefficient changes byonly a factor of 2, a relatively insignificantchange. However, during rapidcooling of the drywell (depressurization),ATw can span more than an orderof magnitude for short durations. Duringrapid heating, the condensing heattransfer coefficient is large (H & 150Btu/hr,ft 2 ,F) and a steady ATw is rapidlyreached. This is about 0.14F in thedrywell. In cooling, the heat transfercoefficient from the steel (approximatelyone-inch thick) wall is low (2-5Btu/hr, ft, F), and can lag behind thebulk gas temperature for some time. Neglectingany temperature gradient in thesteel, for a sudden step change in bulkgas temperature, ATs, the temperaturedifference,ATw, is given bywhereATw = ATs exp (-ht/9.pcph = heat transfer coefficientt = time after ATsZ =wall thicknessp = wall mass density(VII J-34)Cp = wall heat capacity (per unitmass).If the temperature change is gradual,i.e., linear with respect to time. Thetemperature-difference, ATw, is nowgiven byAT = ( £Pc p/h) [exp - (ht/kpcp) - 1](VII J-35)where 8 is the linear bulk-gas coolingrate, F/hr. Equations (VII J-34) and(VII J-35) are used to compute AT. forvarious time intervals during cooling inthe drywell and wetwell. Values of hdepend on gas velocities in the abovecompartments as well as the drywell annularair gap.Little information could be readily obtainedto estimate ATw's in the main reactorbuilding. These would be highlydependent on positions in the buildingand the outside environment temperatures,as well as gas flow parametersfrom the reactor. To allow for someminimum natural deposition in the mainbuilding, ATw = 0.1 F was used. Thiswould result in a natural depositionrate of A A 0.5 hr- 1 in the main reactorbuilding for 12. This X is five timesthe gas displacement rate for the buildingunder normal conditions (2,000 cfmthrough the Gas Treatment System).J3.3.2 POOL SCRUBBINGIn a number of BWR accident sequencesgas flow occurs through the vent linesfrom the drywell to the wetwell waterpool. The water pool occupies slightlyover one-half the toroidal volume of thewetwell and is approximately 17-feetdeep. Pool scrubbing is a major decontaminationmechanism. Some data existson pool scrubbing, but comprehensiveexperimental studies on pool scrubbingin a BWR wetwell pool that include investigationsof all parameters (a widerange of particle sizes, steam quality,pool temperatures, flow rates, 12 concentrations,downcomer L/D ratios,etc...) have not been reported. Applyingsuch data, if available, would nothave refined atmospheric source estimatesgreatly, because the availabledata show that scrubbing is fairly effectiveon the fission products enteringthe pool. Also, trial calculationsshowed that often 30 percent or more ofthe available fission products in anysequence would escape the primary containmentby other paths (such as throughthe drywell annular gap directly to thesecondary containment system).The best available data appears in a paperby Diffey, Rumary, et al., (Ref. 12)where 12, methyl iodide, and .06 p particlesin a steam-air mixture were poolscrubbed. The typical decontaminationfactors were 100 for 12, 2 for CH 3 I, and50-100 for the 0.06 p particles with 90percent steam-air mixtures (higher steamfractions give better decontamination).For CORRAL-BWR cases, the values usedare 100 for both 12 and particles, providedthe bulk gas is mostly steam andcondensation is expected in the pool. AVII-187

decontamination factor of 1.0 is alwaysused for methyl iodide and the noblegases.Even though the BWR accident particlesare assumed to be 5-15 p, or much moremassive than those studied above, thescrubbing of particles is largely due todiffusiophoresis (condensing steam onthe bubble wall carries particles), andtherefore largely independent of particlesize. However, larger particleswould have more boundary layer penetrationinertia in a rapidly circulatingbubble and this should then further justifythe choice of DF=100 for particles,rather than the lower DF of 50.J3.3.3 STANDBY GAS TREATMENT SYSTEM -(SGTS)The SGTS in a typical BWR is a set oftwo parallel filter trains upstream fromthree exhaust fans that releases filteredsecondary containment building air atan elevated level via a stack. However,under accident conditions only one ofthe filter trains would normally beused, the other being held in reserve.The SGTS keeps the building at subatmosphericpressures to minimize groundlevel leaks. Under normal conditions,the flow rate through the system isabout 2,000 cfm but the exhaust fans arecapable of 10,000 cfm. Sources in excessof this maximum would create positivebuilding pressures and produce aground level, unfiltered source. Thefilter trains, according to plant TechnicalSpecifications, are routinelytested to ensure the following removalefficiencies at all flow rates up to10,000 cfm maximum:* 99%* 99%* 99%for particulatesfor elemental iodinefor organic iodideThese specifications are derived essentiallyfrom in-place testing criteriapublished by the USAEC in RegulatoryGuide 1.52. The criteria incorporatelimits on bypass flow for the filter andadsorber sections. It is recognizedthat substantially better performancemay be realized for the filter systemduring actual use. This is because thespecifications implicitly anticipatesome deterioration in system effectivenessbetween testing steps which may ormay not occur. However, since filteredleakage will be a minor contributor tooverall accident consequences for meltdownsequences in the BWR, the efficienciesgiven above were used in all CORRALcalculations. In addition, CORRAL isprogrammed to identify the filter andadsorber activity loadings to discoverpossible overheating conditions, atwhich time the filtration efficiencieswould decrease significantly.J3.3.4NATURAL DEPOSITION IN THEDRYWELL ANNULAR AIRSPACEThe drywell shell is surrounded by atwo-inch air gap between the steel shelland the concrete shield containing theshell. Normal leakage across the shellor rupture of the shell can result intransport through the air gap with anexit into the secondary containment.For analysis purposes, the drywell leakagepoint is considered to occur nearthe base, in the vicinity of the eightvent lines leading to the wetwell. Theexit point is at the top flange assembly.The flow path of gases from theshell rupture is thus around the sphericalpart of the shell, longitudinallythrough the cylindrical upper annulusand out the flange.The annular region cannot be modeledlike the well mixed compartments withdeposition on the walls and/or floor.It is better described as plug flowalong a cylindrical annulus with masstransfer to the walls. A simple firstorder differential equation defines amass transfer coefficient, k, that canbe estimated from known correlations for12 transfer. For particles, k can beestimated from particle deposition datafrom moving gas streams with more difficultyand uncertainty. The differentialequation for transfer to the walls of anannular slit is:wheredC (2k dx(C-C) w - dx (VII J-36)C = the concentration oftransferring substanceCw = wall concentrationU = axial plug flow velocityAr = annulus widthx = axial distancek = mass transfer coefficient.If Cw = 0, Equation (VII J-36) igrates for 0< x

assuming-- AQavan average axial velocity, U, flow cross sectional areawetted parameter= Ar/2 for the annuluswhereThe transition region, 2100 0.1 p). For these inertialparticles, Brownian diffusion isnil, so k = o has been assigned. (DF =1 for Re < 2100 for all particles).Only the largest of the 5-15 p particleshave a significant turbulent depositionvelocity.An emperical fit of Sehmel's data ispossible for k = k (U, dp), but he hasonly two U values for vertical wall deposition.For this reason, only a firstorder approximation can be made for k.This also eliminates major overhaul ofCORRAL's matrix computations to incorporatea new set of variables. The ratiok/U = 1/300 = constant for all k andU for T particle size midway between 10and 15 1. Table VII J-4 shows the widerange of particle DF's versus Reynoldsnumber.To be conservative, an upper limit of DF= 100 seems more reasonable to assign tod = 15 p (for k/U = 1/300, DF = 27).With this upper limit, the followingVII-189

emperical equation fits the particlerange:DF =1.0 + 0.1dp-595(VII J-42)for 10

7. Spray lambdas for particles.8. Terminal spray velocities, gasphase mass transfer coefficient,liquid phase mass transfercoefficient, and spraylambdas.9. 12 equilibrium equivalent lambdas(ifneeded).10. Overall lambdas.d. Solution of Differential EquationsThe solution of the differentialequations is discussed in the previoussection. To properly age thecontinuous releases, it was necessaryto divide these releases intodiscrete impulse releases. The meltrelease was divided into ten equallyspaced and sized releases, each independentage wise from the othernine. The vaporization release (releasedat an exponentially decayingrate) was divided into 20 impulsereleases, each successive release atan exponentially lower value thanthe first. The sum of the first tenreleases equals 1/2 the total release,and the remaining ten equalsthe remaining 1/2. The duration ofthe period of the first ten is onehalf-life. The duration of the secondten should be three half-livesfor a reasonable approximation of anexponential decay.Thus the total number of differentialequations solved (one each forparticulates, organic iodides, and12) for any time step is (N = numberof compartments).GAP RELEASEEXPLOSION RELEASEMELT RELEASEVAPORIZATION RELEASE3N equations3N30N60N6N equationsThe accuracy of the output dependson the rate of change of the ratecoefficients, so short time stepswould be desirable immediately aftereach discrete release (aging is rapidat first, especially if spraysare on). Long time steps are sufficientfor old releases.e. Output Variables1. Airborne contained fractions releasedat time, t.(a)For each release:ganic iodides,lates.12, or-particu-(b) For each compartment foreach release: 12, organiciodides, particulates, attime, t.2. Escaped fractions released (foreach release: 12, organic iodides,particulates, at time,t).3. Escape fractions of the core forany desired iosotope.4. Dose reduction factor for eachrelease (12 and particulates) attime, t.5. Overall dose reduction factor(12 and particulates) at time,t.6. Total fraction ofescaped and coreescaped up to time,core iodineparticulatest.For more complete details on actual operationof the CORRAL code refer to theAddendum to this Appendix which providesa user's guide for the code.VII-191

References1. McCormack, J. D., R. K. Hilliard, and A. K. Postma, "Removal of Airborne FissionProducts by Recirculating Filter Systems in the Containment Systems Experiment",BNWL-1587, Battelle-Northwest, Richland, Washington (June 1971).2. Postma, A., and R. K. Hilliard, "Absorption of Methyl Iodide by Sodium ThiosulfateSprays", ANS Trans. 12, p 898-899 (November 1969).3. Knudsen, J. G., and R. K. Hilliard, "Fission Product Transport by Natural Processesin Containment Vessels", BNWL-943, Battelle-Northwest, Richland, Washington(1969).4. Hales, J. M., T. W. Horst, and L. C. Schwendiman, "Aerosol Transport In a Condensing-SteamBoundary Layer", BNWL-1125, Battelle-Northwest, Richland, Washington(1970).5. Hilliard, R., and L. F. Coleman, "Natural Transport Effects on Fission ProductBehavior in the Containment Systems Experiment", BNWL-1457, Battelle-Northwest,Richland, Washington (1970).6. Postma, A. K., and W. F. Pasedag, "A Review of Mathematical Models for PredictingSpray Removal of Fission Products in Reactor Containment Vessels", BNWL-B-268, Battelle-Northwest, Richland, Washington (1973).7. Postma, A. K., L. F. Coleman, and R. K. Hilliard, "Iodine Removal from ContainmentAtmospheres by Boric Acid Spray", Report No. BNP-100, Battelle-Northwest,Richland, Washington (1970).8. Hilliard, R. K., A. K. Postma, et al., "Removal of Iodine and Particles bySprays in the Containment System Experiment", Nuclear Technology, 10, pp 499-519(1971).9. Reactor Physics Quarterly Report, April-June 1969, BNWL-1150, Battelle-Northwest,Richland, Washington (1969) (B. H. Duane, pp 2.5-2.7).10. Knudsen, J. G., "Properties of Air-Steam Mixtures Containing Small Amounts ofIodine", BNWL-1326, Battelle-Northwest (1970).11. Bird, R. B., W. E. Stewart, and E. N. Lightfoot, "Transport Phenomena", JohnWiley and Sons, N. Y. (1960).12. Diffey, H. R., C. H. Rumary, M. J. S. Smith and R. A. Stinchcombe (AERE, Harwell,England), "Iodine Cleanup in a Steam Suppression System", CONF-650407 (Vol2), International Symposium on Fission Product Release and Transport under AccidentConditions, Oak Ridge, Tenn, pp 776-804 (April 1965).13. Rosenberg, H. S., J. M. Genco, and D. L. Morrison, "Fission-Product Depositionand its Enhancement under Reactor Accident Conditions: Deposition on Containment-SystemSurfaces", BMI-1865, Battelle-Columbus (May 1969).14. Sehmel, G. A., "Particle Eddy Diffusivities and Deposition Velocities forIsothermal Flow and Smooth Surfaces", Aerosol Science 4, pp 125-139 (1973).VII-192

TABLE VII J-1 FISSION PRODUCT REMOVAL RATE CONSTANTS CALCULATED FOR ALARGE PWR CONTAINMENT VESSELFlowFractionRate InitialSpray System gpm Release A, hr- 1Molecular Iodine1 CSR, Boric Acid 3500 1.0 + .01 3.12(a)2 CSR a Boric Acid 7000 1.0 .01 6.24(pH = 5)E Boric Acid -- 10-2 -1.8 x 10-3 1.15Equilibrium Boric Acid -- 1.8 x 10-3 - 3 x 10-5C Boric Acid --

TABLE VIIJ-3 EQUILIBRIUM DATA FOR 12 WITH CAUSTIC SPRAYS (a)Time, min H C /Cgo(a) See Reference 70-100 Constant H, X=0 0.01100-1000 Variable H, X=.095 hr-I1000 7.0 x 10 4 3.86 x 10-42000 1.5 x 10 5 1.8 x 10-44000 5 x 105 5.4 x 10-57000 1 x 106 2.7 x 10-5TABLE VIIJ-4 DF VERSUS PARTICLE SIZEANNULAR GAPAND REYNOLDS NUMBER IN DRYWELLd DF(Re

106Run' ARun A-10 (c.-,ResultsCis fov ! I0 COO0 20003000 4000FIGURE VII J-1Partitioning of ElementalIodine by RecirculatedSprays1.016-tiTimeFIGURE VII J-2Schematic of Quantity of Airborne-Contained Fission Products VersusTime

67OvtITimeFIGURE VII J-3Schematic of Quantity of Fission ProductsReleased to Atmospheres Versus Time0.08o Run A-3* Run A-4* Run A-6* Run A-70.06 1-0 Run A-9wC0 0.04Points are plotted at mid-times ofspray periods (fresh spray only).oCiLi0.02 I--E for diffusiophoresisI I i i I I I IOh0 0.002 0.004 0.006 0.008 0.010 0.012 0.014 0.016Normolized Volume of Liquid Sprayed, T VFIGURE VII J-4Drop-Collection Efficiency asa Function of Liquid VolumeSprayed

QI 01AtmosphericSourceFIGURE VII J-5Gas Flow for PWR CasesFIGURE VII J-6 Flow Schematic for aBWR - (*PossibleAtmospheric Source)

Section ISection 2Section 3Section 4Section 5FIGURE VII J-7Computer Code CORRAL Flow DiagramGapReleaseEx St eamE xplosion1.00.750.50z*0.25Melt Release!= 1.0Vaporization ReleaseY = 0.5 Y .T3T11111i{~11 1 1 1 1 1 I I I0 1 2 3Time, HoursFIGURE VII J-8Typical Sequence of Spike Fission ProductReleases for Postulated AccidentsFig. VII J-1 - Fig. VII J-8VII-195/196

ADDENDUM TO APPENDIX JCORRAL CODE USER'S GUIDEbyP. C. Ouzarski and A. K. PostmaBattelle Pacific Northwest LaboratoriesVII-197

Addendum to Appendix JTable of ContentsSectionPage No.Jl. INTRODUCTION ..................................................... VII-201J2. INPUT DATA ....................................................... VII-201J2.1 Fission Product Release Data ............................... VII-201J2.2 Multicompartment Input Data .................................... VII-202J2.3 Spray Input Data ........................................... VII-202J2.4 Other Input Data ........................................... VII-203J3. OUTPUT DESCRIPTION ............................................... VII-203J4. EXAMPLE PWR CASE ................................................. VII-204J5. EXAMPLE BWR CASE ................................................. VII-205J6. COMPUTER CODE CORRAL LISTING FOR BWR ACCIDENT SEQUENCES ............. VII-211J7. COMPUTER CODE CORRAL LISTING FORPWR ACCIDENT SEQUENCES ...........................................VII-226J8. SUBROUTINES IN CORRAL ............................................ VII-241VII-199

Addendum to Appendix Jof Appendix VIICORRAL Code User's GuidebyP. C. Owzarski and A. K. PostmaBattellePacific Northwest LaboratoriesJ1. INTRODUCTIONThis addendum to Appendix J is intendedto provide a workable tool to anyone desiringto do further containment analysesof particulates, molecular iodineand noble gases. When the realizationthat a multicompartment containment modelwas needed to analyze the complexaccident sequences, we hoped that afairly generalized computer code couldbe developed. We think that with thetwo versions of. CORRAL (PWR and BWR)described herein are.quite versatile.CORRAL was always under constant revision.The final version of CORRAL includesrevisions required to handle themany accident sequences and core releasesthat were defined. Built intothe two versions are four types of fissionproduct releases, capability ofspraying in one compartment with boricacid or caustic solutions, and naturaldeposition in as many as nine compartments.These rooms can be connected inany way with filtration within or betweenthem.The user of CORRAL must input the containmentthermodynamics (pressure, temperature,vapor composition) the intercompartmentalflow rates and leak ratesas a function of time. CORRAL interpolatesbetween input data and computesall natural deposition rates and sprayremoval rates. It solves a large arrayof simultaneous differential equations,and keeps an inventory on the airbornefission products. CORRAL outputs thecumulative amount leaked of the variouscore radionuclides based on the inputdata of the core fraction airborne witheach release. CORRAL can compute thelosses for one explosive event persequence where the containment pressureis rapidly reduced to atmospheric byflow through a large hole.This users manual in conjunction withthe rest of Appendix J describes thecode input requirements and output. Asimple PWR example is shown followed byan elaborate BWR accident sequence example.Both versions of CORRAL are listedalong with subroutines.J2. INPUT DATACORRAL was coded to use the very informalNAMELIST method of input. Dependingon the type of computer, the use ofNAMELIST permits input simply by namingthe variable, followed by an equal sign,the value, and then a comma. For CORRALon the Univac 1108, the procedure is toenter beginning in column 2 on the firstdata card $BREAK N=5, NDATA=12, etc. forall input variables. The last datum isfollowed by $. Two complete lists ofinput data are shown in section J4 ofthis addendum.J2.1 FISSION PRODUCT RELEASE DATABoth PWR and BWR accident sequencestypes of releases are discussed elsewherein this study. The first releaseis the Gap Release that is at t=O. Thisis a single spike release occurring inthe PWR primary cubical (Compartment 2).The other releases occur at later times.Figure VII J-8 shows a typical sequenceof events. Table VII J-5 lists theother releases and pertinent variablesas programmed into CORRAL. All releasesin BWR accidents occur in the drywell(Compartment 1). All times designatedas a variable name in Table VII J-5 mustbe entered in the NAMELIST.The continuous Melt and VaporizationReleases were programmed as discreetspikes so each release could be "aged"properly. Aging is necessary for settlingparticles as well as sprayedparticles. Molecular iodine equilibriumis also time dependent. The aging ofnonsprayed particles is controlled bythree input constants:DPE -= 15. (micron, initialdiameter)DPL = 5. (micron, finalsize)averageaverageVII-201

TD= 4. (hrs, time to go from DPEto DPL).These are discussed in paragraphJ3.2.2.volume ofAF(I), ft2 ,and ceilingmum N isprogram.each V(I), ft3 , floorand wall area, AW(I),height HT(I), ft. The9 as dimensioned inarea,ft2 ,maxitheSince the total of each of the fourreleases is normalized to 1.0, to getthe amount of each isotope escaped tothe atmosphere, an appropriate isotopecore fraction per release must beentered in the input data. Eight basicgroups of isotopes were identified ashaving similar "release" behavior. Theeight can be grouped into three classesaccording to their airborne character:molecular iodine (12-Br group), inertgas (organic iodides and Xe-Kr) andparticulates (Cs-Rb, Te, Ba-Sr, Ru, Lagroups). The input array in fractionsof core airborne per release isCFR(I,J):1 Xe-Kr2 Org-I3 1 2 -Br4 Cs-Rb5 Te6 Ba-Sr7 Ru8 La1, GAP REL.2, MELT REL.For each accident sequence, the user ofCORRAL has a choice of up to 20 sequentialevent changes. For our analysesthese were matched primarily to changesin leak rates from the containment system.However, changes in other parametercan be programmed into the accident.The time input variable is TI(I), I=l,NDATA (NDATA=I-20). The time dependentvariables are:TMY(J,I),J + 1-NI = I-NDATAPI (J,I)Compartment temperature, OFCompartment pressure,psigVAPI(J,I) Compartment water vapormole fractionDELTTI (J, I) Compartment wall-bulk gastemperature difference, OF.(para. J3.2.1 and J3.4.1 ofAppendix J)GI (J,K, I)Intercompartmental gas flowrate (J to K), ACFH.EP(J,K,I) Corresponding particulatedecontamination for J to Kflow. (fraction removed) 1El2 (J,K, Z)ELKP (J, I)ELK12 (J,I)ELK0I (J,I)Same for molecular iodine. 1Fractional loss per hourof compartment J particulates,iodine or organiciodide leaked from containmentto environment. Normallythese would be identicalat any time J.3, STEAM EXPL. FDP (I)4, VAP. REL.CORRAL computes a cumulative leak amountfor each isotope, e.g.,412 leaked = CFR (1,3) x Fraction12=of release 12 leaked)It does not correct for radioactivedecay loss.FDI2 (I)Removal rate of particulateswithin compartmentby filters, ice condenser,etc. Fractions of vol I/Hr.SameJ2.3 SPRAY INPUT DATAfor molecular iodine.Sprays to remove particulates and iodineare important in the postulated PWR accidentsequences. Two types of sprays,the containment system injection spray(CSI), and the recirculation spraysJ2.2 MULTICOMPARTMENT INPUT DATABasic input data is the number of containmentsystem compartments, N, thel(e.g., refer to para. J3.4.2 for poolscrubbing EP(I,2,K)=0.99 for a decontaminationfactor of 100)VII-202

(CSR) could be operating in the main PWRvolume (Compartment 1). The input variablesfor these are:HCSI, height of CSI spray abovefloor, ftHCSR, height of CSR spray aboveDCSI,DCSR,floor, ftCSI droplet diameter, cmCSR droplet diameter, cmCSI, flow rate, ft 3/hr.CSRI,flow rate for firstsprays, ft3 /hrthetheset of CSRCSR2, flow rate for second set ofspraysHO,CSRequilibrium partition coefficientfor 1-2H0 = 5000 for caustic spray= 200 for HBO 3 sprayCUT0V,TCSI,fraction of an 12 release belowwhich airborne concentrations arein equilibrium with the sprayliquid. CUT0V = 0.01 for thecaustic and HB0 3 sprays.TCSIE, endingbeginning time of CSI sprays, hrs.time of CSI sprays, hrs.TCSRlTCSRIE~TCSR2| same for CSR spraysTCSR2ETCSR2E JThe 12 equilibrium calculations inCORRAL are carried out for the sprayedvolume only (Compartment 1). A subroutineELAM must be compiled along withCORRAL for the type of spray used-Three versions of ELAM exist:ELAM/CAUS,ELAM/NONE,ELAM, for HBO 3 spraysfor caustic spraysfor sequences not havingany sprayingThe first two versions use the data inTables VII J-2 and VII J-3 and EquationVII J-32 to calculate an equivalentremoval A for iodine when the concentrationfalls below CUTOV = 0.01. ELAM/NONE puts X=0. See last section forlistings of these subroutines.J2.4 OTHER INPUT DATAIf the containment vessel is breachedwith rapid loss of containment atmosphere,such as in a steam explosion orfloor meltthrough, then the variablesnecessary are: TPUFF, in hours, whenthe puff occurs. XPUFF, the fraction ofcontainment atmospheric still contained,and DFPP, DFP0I and DFPI2, the respectiveparticulate, organic iodide andiodine decontamination factors. A puffrelease takes an equal fraction (l.-XPUFF) from each PWR compartment andadds it to the total atmospheric sourceterm. Since N=4 was the BWR SGTS CharcoalFilter compartment, CORRAL-BWRavoids dumping (l.-XPUFF) of compartment4 into the source term. Even if thereis no puff in the accident sequence, anumber greater than zero must be enteredfor DFPP, DFP0I, DFPI2 and XPUFF. Also,enter TPUFF > TEND to avoid a puff.(See below). This is also true foravoiding TEXl, TMR, and TVR1 and TVR2.The last of the input data necessary forthe execution of CORRAL are the timeinput controls. The first is DT, inhours, the time between computations.DT must be entered for each case ifmultiple cases are being run, since itgets altered in the execution. Thevalue of DT will control the spacing ofoutput between the gap and the nextrelease, and will control the spacingafter the last release is completeduntil TJUMPI which triggers the newspacing level DT2, both in hours. Thecalculations progress until TJUMP2 whichtriggers a jump to TEND where the lastset of calculations in the execution iscompleted.CORRAL-BWR has two additional inputintegers. These are MCOMP and MANN. IfMANN is any integer other than 1, gasflow can proceed through the drywellannulus from compartment number MCOMP (1or 2). Some accident sequences had flowfrom the drywell (compartment 1) towetwell (compartment 2) and then backthrough the annulus.J3. OUTPUT DESCRIPTIONThe first output is the input data withlabeled descriptions. The same unitsare in this output as in the input. Seethe example cases for a comparison.The calculational output follows. Foreach time value TT = TT+DT, CORRALwrites:a. The inventory of airborne particles,iodine and organic iodides (or noblegases) in each compartment for eachrelease. The values are expressedin fractions of the normalized release(1.0).VII-203

. The inventory summed over the Ncompartments.c. The cumulative fraction of particles,iodine and organic iodides ofeach release escaped to the outsideatmosphere.d. The dose reduction factor of particlesand iodine for each release.This number is the cumulative amountof isotope leaked divided into thecumulative amount leaked if therewere no deposition or washout in thecontainment system. This was donesimply by dividing the cumulativefractions leaked into the cumulativefraction of organic iodides leaked,since no credit was given for washoutof organic iodides.e. The cumulative fractions of coreleaked for each of the eight isotopegroups described earlier.The last item constituted the bulk ofthe data used for determining populationdoses.J4. EXAMPLE PWR CASEThe case selected to represent PWRinput/output data was a typical DesignBase Accident (Case A) where only a gaprelease occurred after the large pipebreak. This accident sequence is verysimple and illustrates how CORRAL can beused to study one of many parametersthat have some influence on the fissionproduct loss to the environment. Forexample, the flow rate between the outerannulus, the main sprayed compartment,and the lower volume was slowed down bya factor of 10 in this run compared toCase A examined in the Reactor SafetyStudy. This slowdown lowered the fissionproduct losses in the first fewminutes after the gap release, but hadno long term effect. Figure VII J-5 isthe schematic of the gas flow. Dischargeto the atmosphere was fromCompartment 3, the outer annulus. Thesprayed main volume is Compartment 1,the primary cubicle is Compartment 2,and the lower volume is Compartment 4.In all PWR cases with the 10 containmentvolumes/hour circulating from Compartment1 to 3 to 4 to 1 again the containmentwas essentially "well-mixed."There was rapid interchange betweencompartments 1 and 2.INPUT CARDSPwaP SA"PLE CAtF - nF IGN RAq• ACC(11NT WITH LOW INTERNAL CIRCdLATIONTrI( I.1..4.3..1..5.G I (10, 1)1=1 2+4 (, 1(4.2, 1)=O.. r, 1 (2.. , 1 1.2+*GT(1,7, 2)=1.2-06. '114,7, 2)1.,, P1(4,1, 2-1,2-06,4T( 1 '( 2 . n I ( 4 7 , . 1 = n*(2LO, FL"1T(l= 1, 4).. (•1 2 ,, . 4•TITVI1.')4=1..•A PIl .' .1'.. ,P c( *., 4):.'+•*G I(] I. 51)=1.61÷6. , I 1(9, 4. a;) = 1.61-6, 1 (4.] 1 . =1.61÷6.TYY(1,.(4*)47.P÷6 I 2,Or)l.644 .'. yAI)2), 4(I(1,1. 6)= .,1.61. T( 1.4 , 6 '=.61+6, 61!4,1,=}.?+4,(,)=1.61 6. ,

hLT . .00, OU -OM ýý;'SPWR CASE A - SAMPLE FISSION PRODUCT OUTPUT DATACO"I-ARTMEjIf AIHLO ," . FR4CTI )'4S COQTAINE',C GAP ,Ap, GAP "ELTU RLLtQ A E , '1E- LA•A RELFAS" EL V A',EM PAT IC•.LS 12 'I PA•,TICLLSI 9.iS.9-03 4. (3 _,-,3 q*4 7-r~ - .)n," 9-9.1b 5 -01 9.6.7;,-('1 ?.O,700. I 0.00(53 .Z659-06 42, b.,7"-11, 4.25•63-0,, O.OnOCMELT MELT EXPLOSION EXPLOSION EXPLOSION VAPOR VAPOROrLEASE OELEASE RELFAqE RIELEASE RELEASE RELEASE RELEASEI2 01 PARTICLLS 72 01 PARTICLE% 120.00000.000"0.0n00O00n00Oono00.00ou0.00000.00tU0.00000.00000.00000.00000.00000.00000.0000O.OOO00.00000.00000.00000.00000.00000.0000•.OOOVAPORRELEASE010.00000.00000.00000.0000IUTA. AIRIBON.E F1 AC'lu rlS CO'!TATNErGAl,AoRLLEA.E RELEAS.PA 0 1IL.LS 129.9do.-ul ,.o2. I _0'0.. .G•r "FLI MELT MELT EXPLoSIoN EXPLnSION EXPLOSION VAPOR VAPORSELFASE PEU EASE WrLEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE.TI PAI-!CLL0. TP 01 PARTICLE9 1? 01 PARTICLES 120.10n'in.OSVAPORRELEASE01.OOT .0 0 0.00"0 0.O0ou 0.00O0 0.0000 0.0000 0.0000OSCA,'E FRACTIU!:S Or EACH RELEASEGAP";APPARTICLES '2GPP "ELT MELT MELT EAPLOSION EXPL"SION EXPLOSION VAPOR VAPOR,LLTASF "EI EAS RSLEASE RELEASE RELFASL RELEASE RELEASE RELEASE RELEASE01 PAROICLES I .01 PARTICLES 12 01 pARTICLEs 121.Uu5b-12 i.U.75-12 1.1611-I.? Q,0O0n 0.000n 0.00n0 0.0000 0.0000 0.0000 0.0000 0.0000VAPORRELEASE010.0000UOSE RE1'"CTICN FACT 7ýS OF FAC,HFL-ASEGAP ýAP GAPT ELT "MELT I ELT EXPLOSION EXPLOSION EXPLOSION VAPOR V"POp..LLAýE RELEASL SELFASE FELEA..L RELEASE RELEASE RELFAE RELEASE RELEASE RELEASE: RELEASEPAKT1CLES 12 01 PARTICLES 12 01 PARTICLES 12 01 PARTICLES 121,0097400 1.0015'AO 0.0n00 0.0001 0.000o 0.0000 0.0000 0.0000VAPORRELEASE01FRACTIONS OF COREINVE!ITORY LEAKFOr,R-Yg 01 I20-P CS-B TE BA-SR RU3.e072-14 7.2696-17 I.P147-14 5.2940-14 1.05A8-16 1.0588-18 O.u00OLA0.0000J5. EXAMPLE BWR CASEIn contrast to the simple PWR example,one of the most complex BWR sequences ispresented as an example. SequenceAGJ-6 had a number of features that madeit difficult to fit into CORRAL's usualcapabilities. CORRAL normally calculatesone lumped atmospheric source. Thissequence had:a. An elevated stack source through theGas Treatment System (SGTS, Compartment4).b. A ground level source.c. Charcoal filter beds whose trappingcapacity for 12 and organic iodidecould fail if too much 12 collectedon them.d. Flow from the drywell (Compartment1) to the wetwell (Compartment 2)with pool scrubbing.e. Return flow to the drywell.f. Flow from the drywell through thedrywell annulus to the Main ReactorBuilding (Compartment 5).A Compartment 3 was used as an artificialdumping ground for the drywellannulus. To segregate the ground levelsource from the elevated source, anotherartificial Compartment 6 was created.Flow to it was considered as the groundsource and core fractions in it had tobe hand calculated from the setsCFR(I,J) and CG$I(6,2) etc.The elevated source, from filtered airwas easy to handle as a simple low levelleak rate from Compartment 5 for particulatesand 12 only. This meant that fornormal filter operation, the total flowof particulates from Compartment 5 was1.001 times larger than it should havebeen. The flow of 12 was 1.001 timestoo high. This was negligible. Thenormal organic iodide leak rate was 150times larger than the 12 leak ratecreating a 1.15 overdepletion rate. Toavoid this, no organic iodide leakedfrom Compartment 5, but rather wasallowed to build up on the filter(Compartment 4) and 15 percent of it washand calculated to add to the totalatmospheric source.The SGTS overheat problem was easilyhandled by checking on the 12 buildup.When the charcoal beds did overheat, itwas necessary to hand calculate all the12 and organic iodide as a differencebetween the buildup of new 12 andorganic iodide in Compartment 4 and thebuildup level at the time of overheat.VII-205

No real problems arose but the Univac1108 chugged 1512 seconds for the 24-hour accident sequence. This is largelydue to the use of N=6 compartments. Thetime for matrix solution of the differentialequations increases dramaticallyas N increases.1WR EXAMPLF INP,(T DATA FO)R ACCIDENT ýFQUENCE AGJ-OELTASQQPFA:,N=6.NfOATA=2n1 .COMP=I. MAN'!=2,T((111 .,.116..117.1.5.1.5194.15,4.151,4.483.5.483.5.6595.651,6.1596.151,6 .483 .6 .484 ,111. 15, 10. 151,1 3.15, 13. 151, 24..HT(1 =1110. .29 .9 ,.I,*1,56..APIiI(3526..(083?.,).,11..1.96+4.AW(1(=14810.,163011..1..11..3.14+4,V(II (1.59+ý5,1.19+5,27.8,27.8,1.1+6.TMY( I, ITMY ( 1.20=27.,P 0( 1.1 1=24.1 , VAPT(, (I )=.91, 'WLT=1 8,1. PTIr I *? 1=7.1n1. VADI(, I (';>).99, DFLT' T1(I,2 TI(1.1 1=5n.. ):Sn..TMY( 1.4 (=181).. P 1(1,4 =4.1100. IAOI(1 .4 )=.99. DFLT TI(1.4 )=3.3.TMY(l.5 )=18n1., PT(1.95 (=4.311. VAPU(1.5 )=.99. DELT TI(1.5 1=3.3,TNY(1,6 )=190.. P1(1.6 )=.300. VAPI((,6 )=.99, DFLT' T111.6 )=.14.TMY(1,7 )=190., P1(1,7 )=.300, VAPI(1,7 )=.99. DELT TI(I,7 )=.14,TMY(1.A )=158., PT(1.8 (=19.3, VAPI(1,8 )=.99. DFLT TI(I.8 )=.14,TVY(1.9 (=3119*, PT(I.9 )=61.1, VAP((1,9 )=.19. r3ýLTTMY(1,101=329., PT( 1,11711=87.1. VAPt(1 ,111(=.99, DELT'TMY(3.11l(=172.. PT(I.111=87.1, VAPTIl,11(=.0Q, ')FLTTT(I .111=1.11,TMY(1.31'297..TIPT11,111=49.3. VAPI(1.13(=.911, nFLT TI(I.13(=.14.(1,12=1=4.,TMY(1,14(.322.. PT(1,141=77.1. ('6P1(1.141=.911, DELTT%'Y(1.''(.27?., PI(1,35(1=770. VAPI(1,19)=.90, (WrLTTI(I,15)=.14.JFLOWTMY(1.16('328., PI(l164=85.3, VAPI(1,16(=.90, DFLT TI(1.16)=.14,TMY(I1.7)=328., Pfll,17(=85.3, VAP1(1,17((=.90. DELT TI(1.17)=.62,T'1Y(1.18(=216.. PTI1,18(1=1.311, VAP((1.18)=.90, DELT TI(I.18)=.62.TMY(1.19(=216., PI(I.19(=1.30. VAP1(1,191=.90. DELT TI(1,19)=.14,TMY(1.10(=?16.. P1(1.201=1(.20. VAPI((1201=.90. DFLT T F(L.P0)=.L .BAF(2)=11.,AW(2(=.0.AFI6(=11.,AW(6(=11.,V(6(.10011.,G1(1.?,7 (=1.94+6. EP412.,7 )-.99, E12(1,2.7 )-.99.TO WETIWALIL. WITH POOL SCRUBBING0,1(1,?.8 1=9.66+1,s EPI1.2.8 )=.Q9. E12(1,298 )-.99.r1(I.7.l01=6.:14+r,, FP(l.;2.10(=.O.19 2121(.9CM1(,.141=4.81+5. EP(I,2.14)=.99. F12(1,2,14)=.99,G1(1.2.15(=3.55+4. EP(1.2.15)=.99, E12(1.2,15(..99,01(2,1 *11)=8.87+4,11?(?,1,12 (=8,87+4,GI(2.1,17)=3.32+4,GI(2,1,18)=3.32+4,61(10,.1 )=2.31+5.GT(I.3.2 )=2.31+5,r, ( 1.10• )=7,11+1.0111.3,4 (=2.31+5,GT0.13,5 I=1.17+5,1(M1.3,6 1=1.17+,,Gy0(1 r (I .3.31=2.18.5'A. Q I =1 . 72+5.GIl1,3,12)=1.64+5,GI(1.3.13)=1.69+5.01(1.33,14)=2.69+5,0I(1,3.15)=1.34+5.rI(1.31]6)=2.(145.rI(I.3 .39=1.84+5.r1( .3.01=1.187+5.rTI(]0.IO)=j.87*5.GI1(%A,7 (=7.58+5,T.1(5,6.8 (:2.58+5.GI(5.6,Q )(2.58+5,GI(';,6.10'=2.58+5,GI(5.6.15) 3.60+5.GI(5,6.161=3.60+5.FLSP( 5 ,1 )-2.29-5,FLKP(9.2 )=2.29-5.FLKPI5.3 )=2.29-5.cLvP(q.4 )-P.29-5.FLKP(•5•, =1.11-5.FLKP(9.6 1=1.10-%,FLKPIS.7 )=5.45-S.FLKP(S.A )=5.45-5.FLKPI,9.)=5.45-5,FLKP(%5.0)=5.45-5.FLKP(5,11)=5.45-5.I RETURN FLOW FROM WETWELL6T(3,5,1 )-2.31+5, GI(5,4.1 )=2.52+50GI(3,5.2 )=2.31+5. G1(5,4.2 )=2.52+5.cI((.,5,3 (=?°31÷50G1(5,4,3 )=2.52+5.GI(3,5,4 )=2.31+5,r.1(1.5.5 )=1.17+5.GI(5,4.4 )=2.52+5.GI(5.4.,5 )=1.20+5#rI(3..,6 (=I.17+5, G1(, 4.6 )(1.20+5#01(3.,7 (4. ~177O(r,.4.7 (=6.111+5,111(3.5,T ( S37+,1(,4,8 I(=6. 111+5,6T(3.5.9 (=2.08,5, GI(,.4,) )(6.n0+5,1T('G5,11)=2.1 . 01(5.4. 1 ()=6.(I()+1.01(=,5,11(=1.64+5, 01I(5,4.11)=6.OC+5.01(3,5.12)=1.64+5. GI(5.4.12)=6.00+5.0I(1,5.13)=I.69+5. GI(5,4.1)3=6.00+5,01(3,5.14)=1.69-5. GI(R.4.I4)=6.00+5,GT(3.5.15)=2.18+5. G01(54.15)=6.n0+5%01I(3,5,]6)=2.1÷.5, GI(05,4.16)=6.00+5,r,11 (1 ,1 =1.34,5+i. G1(.4,17)=5.10+5,AT(!.5.(9)=1.P7+5. 111I(5.4.191=2.7(++,1(3,5.20(=l.8'÷.5IFLOW0I(5.4,20)=2.70+5.AS A GROUND LEVEL SOURCEELKI2(5,1 )=2.29-4,FLKI2IS.2 )=2.29-4,ELKT2(5,3 )=2.29-4,FLKT2(5,4 )=2.79-4,FLK(2(5.5, (=1.111-4,FLKI2(5.6 3=1.111-4.ELKI2(5.7 )=5.4,-4,ELKI2(5,8 1-5.45-4,ELKI?(5,9 )(5.45-4.ELK 12(15. 10 5. 45-4.FLK(2(5.11 (.5.45i-4.1ST COLUMN IS FLOW THROUGHDRYWELL ANNULUS2ND COLUMN IS FLOW INTOMAIN REACTOR BUILDING3RD COLUMN IS FLOW INTOTHE FILTERVII-206

FLKP(".12)=5.45-5, ELKI2(5,12)=5.45-4.FLKPI5.131=5.45-5. ELK12(5.]3)=,.45-4.FLtK'(5.)4 (=5*45-5, ELKT2(5.14lx5.4S-4.rLKP(".15) 5.4-5. FLKI2(5,151=5.45-4.FLKPlS.16)=5.45-5. FLK1215.]6)=5.45-4.FLKP(5,17)=4.64-•, FLKt (5,171=4.64-4,FLKP(R.18)=4.64-•t ELKT?(5.18)=4.64-4,FLKPI5.)9 (=2.45-s. FLKT?(5,J9)=2.&.5-4,FLKP( 5.20 (=2.45-5, PLK! 20(520(=2.45-4,CFP( 1,1 )=.03..87,0.,.1,CFR(R,4)=.O05,.7 6 ,0.,.19,(PR) 1,5)=1.-5..15,O. .. 85.(PR( 1,7(=O. ..03.0,.,05.CFR( 1.R)=O...O03,.t1, .01Tr)=4. 00F=19. r)PL=5.,CST=1. . CSR=.,CP2=1..rC.'S=1.,(•R=1. VCLI. F(CCI=I., DUMMYTrSRI1=.+4÷T(SR2=1.+4. TrSI=1.44,TCS1Icf.+4.TC.SRlF=I.+4,T(FR2FI.+,4.HCSR=I..H0=1..HCS=I..CUTOV=.01,TPX]1I.+4,TPJFF=1.+4,XPUFF=1.,DFPP=1.,DFPO!=1.,DFP 121.,T)0=4.48R3,TMF=5.48a.TVRl=6.1;.TV02=6.65,TVRF=8.15,rT=.44 Ar)T2=. ,TJIJMP1=8.2 TJUMP?=I 1 ?.2TND=24.,.DA[A AAV ASSUMPTIUNSNO. OF COMPANTME17S 1= 6OUTPUT - BWR SEQUENCE AGI-6FISSION PRODUCT RELEASE AND CLEANUPCOMPARTMENT PRESSURES((PI(I.J),I=I, 0),J=1,20)=2.830000+01 0.0oU0o0 0.000000 0.000000 0.000000 0.0000007.300j)00+()0 0.000000 0.000000 0.000000 0.000000 0.0000007.300000+00 0.O•oooo 0.000000 0.000000 0.000000 0.0000004.U00000+00 0.0OU0ou 0.000000 0.000000 0.000000 0.0000004.000000+90 0.000300 0.000000 3.000000 0.000000 0.0000003.000u00-91 0.000000 0.000000 0.000000 0.000000 0.000000s.uOoouuO-9I 0.00000i 0.000000 0.000000 0.000000 0.0000001.930000+01 0.00o0n0 0.000000 0.000000 0.000000 0.000000b.1301100+UI 0.000000 0.000000 0.000000 0.000000 0.0000008.730000+91 0.000000 0.000000 0.000000 0.000000 0.000000b.730000+01 0.00)000 0.000000 0.000000 0.000000 0.0000004.9301300+01 O.u0000 0.000000 3.000000 0.000000 0.0000004.930000+91 0.000000 0.000000 0.000000 0.000000 0.0000007.730000+01 0.000000 0.000000 0.000000 0.000000 0.0000007.730300+01 0.000300 0.000000 0.000000 0.000000 0.0000008.030000+91 0.000000 0.000000 0.000000 0.000000 0.0000006.5300100+91 0.000000 0.000000 0.090000 0.000000 0.0000001.300000+90 0.000000 0.000000 0.00co00 0.000000 0.0000001.300000+90 0.000000 0.000000 0.000000 0.000000 0.0000001.300000+00 0.000000 0.000000 0.000000 0.000000 0.000000COMPARTMENT TEMPERATURES(CTHY(I.J),I:1, o),J=I,20)=2.720000+92 7.500300+C1 7.500000401 7.500000.01 7.500000+01 7.500000+011.800000+02 7.5nUO0o+C 7.500000+01 7.50o000001 7.500000+01 7.500000+011.600000+92 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+011.800000+02 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+011.800000+02 7.500000+C0 7.500000+01 7.500000+01 7.500000+01 7.500000+011.900000+02 7.500000+C1 7.500000+01 7.500000÷01 7.500000+01 7.500000+011.900000+92 7.500000+01 7.500000+01 7.500000.01 7.500000+01 7.500000+011.580000432 7.500000+01 7.500000+01 7.50000+01 7.000 7.500000+013.080000*02 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+013.290000+02 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+013.290000*02 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+012.970000+02 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+012.970100+02 7.500000+01 7.500000.01 7.500000+01 7.500000+01 7.500000+013.220000+02 7.500000+01 7.500000+11 7.500000+01 7.500000+01 7.500000+013.220000+02 7.500000+01 7.500000+61 7.500000+01 7.500000+01 7.500000+013.280000t02 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+013.280100+92 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+012.160U00+(12 7.5000•0+01 7.500000+01 7.500000+01 7.500000+01 7.500000+012.160000+02 7.500000+01 7.500000+01 7.5n0000+01 7.500000+01 7.500000+012.160u00+02 7.500000+01 7.500000+01 7.500000+01 7.500000+01 7.500000+01IEMPERATURE OIFFERENCES TBULK-TWALL(CUELTTI(IJ).I=I. b) ,J=1,20)=I.1ooJOO+00 1.000000-01 1.000000-01 1.000000-Cl 1.000000-01 1.000000-015.000000+01 1.0noooo-Cl 1.000000-01 1.000000-01 1.000000-01 1.000000-015.000000+-)01 1.00000-lo 1.000000-01 1.00c000-01 1.000000-01 1.000000-013.300lJ000(+0 1.00f0306--01 1.000000-01 1.000000-01 1.000000-01 1.000000-013.300000+00 1.000000-Cl 1.000000-01 1.000000-01 1.000000-01 1.000000-011.400000-01 1.000000-Cl 1.00000-01 1.000000-01 1.000000-01 1.000000-011.400000-Oi 1.000000-C1 1.000000-01 1.00v000-01 1.000000-01 1.000000-011.400000-01 1.000000-Cl 1.000000-01 1.000000-01 l.0ooo0n-oi 1.000000-011.400000-01 .onuoo0-Cl 1.000000-01 1.000000-01 1.000000-01 1.000000-011.400000-01 1.000000-Cl 1.000000-01 1.000000-01 1.000000-01 1.000000-011.000000+Oo 1.000000-01 1.000000-01 1.000000-01 1.000000-01 1.000000-014.800000+00 1.000000-Cl 1.000000-01 1.000000-01 1.000000-01 1.000000-01VII-207

1.4"0000-01 1.O0U000-C1 1.000000-01 1.000000-01 1.000000-01 1.000000-011.400000-l1 1.000000-01 1.000000-01 1.000000-01 1.000000-01 1.000000-011.400000-01 1.000000-01 1.000000-01 1.00C000-O1 1.000000-01 1.000000-011.400000-01 1.00000,-C0 1.000000-01 1.OOCO00-01 1.000000-01 1.000000-016.20O00-O1 1.000000-01 1.000000-01 1.00CO00-01 1.000000-01 1.000000-01b.200000-O1 1.000000-Cl 1.000000-01 1.000000-01 1.000000-01 1.000000-011.400000-01 1.0n0o00-01 1.000000-01 1.000000-01 1.000000-01 1.000000-011.000000-01 1.0000oo-el 1.000000-01 1.o0oo0o-01 1.000000-01 1.000000-01WATER VAPOR MOLE FRACTIONSl(VAPI'ItJ) *I=I, 6) ,J=1.20)=9.900000-019.900000-013.000000-023.UN0000-023.000000-023.000000-023.000000-023.00C000-023.000000-023.000000-023.000000-023.000000-029.900000-91 3.00oOOO-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.OOUOOO-C2 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.000000-C2 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.000000-C2 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.000000-C2 3.000000-02 3.000000-02 3.000000-02 3.000000-029.900000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-CZ 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-c2 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-C2 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000U00-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-02 3.000000-02 3.000000-02 3.000000-02 3.000000-029.000000-01 3.000000-C2 3.000000-02 3.000000-02 3.000000-02 3.000000-02FRACTION OF 12 LEAKED PER HOUR( 0ELKL2(IJ),0=0 , 00-J=0,2C)00.000000 0.000000 0.000000 0.000000 2.290000-04 0.0000000.0OOOoo U.00000000000 00000000 2.290000-04 0.0000000.U00U00 0.000000 0.000000 0.000000 2.290000-04 0.0000000.000000 0.000000 0.000000 0.00c000 2.290000-04 0.0000000.000000 0.000000 0.000000 0.000000 1.100000-04 0.0000000.000000 0.000000 0.000000 0.000000 1.100000-04 0.0000000.000000 0.000000 0.000000 0.00OO00 5.450000-04 0.0000000.u0000 0.000000 0.000000 0.000000 b.450000-04 0.0000000.000000 U.000000 0.0000u0 0.000000 b.450000-04 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-04 0.0000000.000000 0.000000 0.000000 0.000000 b.450000-04 0.0000000.000u00 U.000000 0.000000 0.000000 5.450000-04 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-04 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-04 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-04 0.0000000.000000 0.000000 0.000000 0.000000 5.650000-04 0.0000000.000000 0.000003 0.000000 0.000000 4.640000-04 0.0000000.o00000 0.000o0u 0.000000 0.000000 4.640000-04 0.000000U.OOOOOc0.00000c0.000000G.0000030.0000000.0000000.0000000.0000002.450000-042.450000-040.0000000.000000FRACTION OF PARTICULATES LEAKED PER HOUR((ELKP(IJ),I=1, b6)J=I120)=0.000000 0.000000 0.000000 0.000000 2.290000-05 0.0000000.000000 0.000000 0.000000 0.000000 2.2q0000-05 0.0000000.000000 0.000000 0.000000 0.000000 2.290000-05 0.0000000.000000 0.000000 0.000000 0.000000 2.290000-05 0.0000000.000000 0.000000 0.000000 0.000000 1.100000-05 0.0000000.000000 0.000000 0.000000 0.000000 1.100000-05 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-05 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-05 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-05 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-05 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-05 0.0000000.000000 0.000000 0.000000 0.000000 5.450000-05 0.0000000.000000 0.000000 0.000000 0.000000 b.450000-05 0.0000000.000000 O.00O000 0.000000 0.000000 5.450000-05 0.000000o.000000 0.000000 0.000000 0.000000 5.450000-05 0.000000o.oobOo0 0.000000 0.000000 0.000000 5.450000-05 0.0000000.000000 0.000000 0.000000 0.000000 4.640000-05 0.0000000.000000 0.000000 0.000000 0.000000 4.640000-05 0.0000000.000000 0.000000 0.000000 0.000000 2.450000-05 0.000000.0.000000 0.000000 0.000000 0.000000 2.450000-05 0.000000FRACTION OF ORGANIC IODIDES LEAKED PER HOUR FROM COMPARTMENT I((ELKOI(I.J).*=I. 6) J=12C)=0.000000 0.000000 0.000000 0.000000 0.000000 0.0000000.000000 0.000000 0.000000 0.000000 0.000000 0.0000000.000000 0.00U000 0.000000 0.000000 0.000000 0.0000000.000000 0.000000 0.000000 0.000000 0.000000 0.0000000.000000 o.on0000 0.000000 0.000000 0.000000 0.0000000.000000 0.000000 0.000000 0.000000 0.000000 0.0000000.000000 0.000000 0.000000 0.000000 0.000000 0.0000000.000U00 O.Ono000 0.000000 0.000000 0.000000 0.0000000.000000 O.uou00) 0.000000 0.000000 0.000000 0.0000000.300000 0.000003 0.000000 0.090000 0.000000 0.000000O.uo0000 0.00000:) 0.000000 0.00cooo 0.000000 0.0000000.000000 0.000000 0.000000 0.000000 0.000000 0.0000000.000000 0.000000 0.000000 0.0OC000 0.000000 0.000000VII-208

0.000,000 0 .00100o 0.0000000.90 o(00 0.0000o0 0.00000011o.00000C O.0010rou 0.0000000. 00000 0.000000 ,0.00000011.000000 0.000000 0.0000000.000000 0.000000 0.0000000.000000 0.000000 0.000000PI

TCSHI=TIME CONTAINMLNT SYSTEM RECIRCULATING SPRAY STARTSICSRIE=TIME CUNTAINMENT SYSTEM RECIRCULATING SPRAY ENDST"X1:TIME OF EXPLOSION NO. I1PJFF=TI,.A OF PUFF RELEASETJUMP1=TIME rU SI1TCH TO TIML STEP DT2TýjUAP2=TIME TO STEP TO ENDTIb 4.00009+00 0Tz 4.48300-01 TVHI= 6.150O+00 TVR2: 6.65000+00 TVRE= 8.5000000+00 TMR= 4.48300000+00TMF: 5.48300•00+00 TCSI= 1.0000000+04 TCSIE: 1.00000000+04 TECCI: 0.00000000 TCSRI: 1.00000000+04TCSRIE= 1.0O•9UC0GO4 TCSR2: 1.000DO00+04 TCSR2E= 1.00000000404 TEXX: 1.O00000000lO4 TPUFF= 1.00000000+04Uf2= 5.00O0COOO-Ul TJUAPI0 8.19999990+00 TjUMP2= 1.31999999+01 TEND: 2.40000000+01REACTORCOMPART.IEiT LATACOMPART;AENT WALL AREAiAW(I),I1I. 6):1.481000+9.O4 0.000000 0.000000 0.000000 3.140000+04 0.000000COMPARTMENT FLOOR AREA(AF(l),I=I, 6)=3.1)26000+03 0.000000 0.000000 0.000000 1.960000+04 0.000000COMPART74ENT iEISHT1.000o00+92 2.990000+01 1.0ofOOOO-01 1.000000-01 5.600000+01 0.000000COMPARTMENT 4OLUME (CUBIC FEET)(V(L).l1, 6)=1.!l 0t)OO÷05 1.19000÷05 2.780000+01 2.780000+01 1.100000+06 1.000000+03LOi.TAINd,.EC4TAih) CLEANUP SPECIFICATION4S•I=CONTAI&•TiET SYSTEM IN;JECTION FLOW IN CUBIC FEET/HR.CSRI:CONTAIl4lAEl4T SYSTEM NO.1 RECIRCULATING FLOW IN CUBIC FEET/HR.HCSI,r4CSN4:HEI(HT THROUGH WHICH CONTAINMENT SYSTEM SPRAYS FALLOCSI,UCS:=UIAMETER OF CONTAINMENT SYSTEM SPRAY DROPLETSCLITOV:COIJCECITRATIOrJ OF 12 [3ELOW WHICH SPRAY REMOVAL IS INEFFECTIVEvLL=VULUME OF COOLANT LIQUIUHU=RATIO OF IODINE IN SOLUTION IN WATER TO THAT IN VAPOR AT EQUILIBRIUMELCI:EMERGENCY CORE COOLANT FLOW RATEuFPP,DFPuI.DFPI2=PUFF RELEASE DECONTAMINATION FACTORSXFPUFF=FRACTION OF AIRRORNE MATERIAL RETAINED AFTER PUFF RELEASE6PE,Jl'L=EFFECTIVL DIAMETER OF PARTICLES IN MICRONS AT EARLY AND LATE TIMESCSIZ l.uooe(o000400 CSRI 1.00000000+00 HCSI: 1.00000000+00 HCSR: 1.00000000+00 DCSI: 1.00000000+00DCSM= 1.000CO000+00 CUTOV= 9.99999990-03 VCL= 1.00000000+00 No= 1.00000000+00 ECCI= 1.00000000+00LJFPP= 1.uOOCflOO+00 FPOI= 1.0000000+00 DFPI2= 1.0000000+00 CSR2= 1.0000000+00 XPUFF= 1.0000000+00OPE= l.bO0OUO+01 OPL= 5.0000000+00CORE FRACTIuLiS RELEASED BYMETHOD J IN ISOTOPE GROUP I CFR(JI)GAP RNEEASE MELT RELEASE STEAM EXPLOSION VAPORIZATIONKR-AE 3.00C00-02 8.7C990-01 0.00000 1.00UO0-OL01 0.00000 7.00000-03 u.00000 0.0000012-UH 1.70C00-03 8.83900-01 U.00000 1.00000-01CS-Rd 5.OOC9u-03 7.6C000-01 0.0OG0O 1.90000-01TL I.O1,00•-05 L.50900-01 0.00000 8.50000-01OA-SR 1.000-07 1.O0000-01 0.0000 1.00000-02RU 0.00000 3.00900-02 0.00000 5.00000-02LA 0.00COO 3.00000-03 0.00000 1.00000-02SAMPLE OF FISSION PRODUCT OUTPUT DATATIME: 5.Ai13+Oa0HNSCOMPARTMý11T AlRd-•NIE FRACTIONiS CONTAINEDL 6AP GAP GAP MELT MELT MELT EXPLOSION EXPLOSION EXPLOSION VAPOR VAPOR VAPORU RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASEM PARrICLES 12 01 PARTICLES 12 O0 PARTICLES 12 or PARTICLES 12 01I ob6.12-09 ?.2949-03 9.2327-C7 8.8233-02 1.2575-01 1.3296-01 0.0000 0.0l00 0.0000 0.0000 0.0000 0.00001. "0-0, .6499-0o 1.4684-02 5.0654-03 6.2838-03 6.5028-01 0.0000 0.0000 0.0000 0.0000 0.00003 i.0b-L2 2.979b-Ie 1.ol63-1' 1.b542-05 1.6279-05 2.3264-05 0.0000 0.0000 0.0000 0.0000 0,000 00;0004 1.5)3-02 2.44o2-02 6.5264-C1 1,2950-92 2.3502-02 4.4345-02 0.0000 0.0000 0.0000 0.0000 0.0000 0.00005 T.7j.i-03 .,OdjO-03 ;.0437-Cl 5.o998-02 8.5738-02 1.5332-01 0.0000 0.0000 0.0000 0.0oo0 0.0000 0.0000U 1.cl-d3 .6.3-3-03 1.2831-01 S.5bhb-03 8.08158-03 1,90o8-02 0.0000 0.0000 0.0000 0.0001 0.0000 0.0000TUTAL AIR=iORi4EFRACTIONS CONTAINED.AP OAP GA" MELT MELT MELT EXPLOSION EXPLOSION EXPLOSION VAPOR VAPOR VAPORRLLASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASEPAITICLE. Ile 01 PARTICLES I? O0 PARTICLES 12 O1 PARTICLES 12 011.5353-02 :.2939-U2 1.000+09 1.7003-01 2.4711-01 1.000÷+00 0.0000 0.0000 0.0000 0.0000 0.0000 0.0000VII-210

LýCAPC FRACTIvrS JF EACH RELEASE.AP ,AP GAP MELT ,MELT MELT EXPLOSION EXPLOSIWN EXPLOSION VAPOR VAPOR VAPORREýLASE kELLASE RELEASE RLLEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASEiARIIC.ES £2 O0 PARTICLES 12 01 PARTICLES 12 oI PARTICLES I2 Ov1.5.L,4-Oo Z.4b514-05 .uD003 1.293 9 -Ob 2.SRS-0S 0.OOO 0.0000 0.0000 0.0000 0.0000 U.0000 0.0000UGSE NEOUCTIOj FACTORS OF LACH RELEASE.AP VAP GAP MELT MELT MELT EXPLOSION EXPLOSION EXPLOSION VAPOR VAPOR VAPORRULEASE RELEASL RiLEASE RLLEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASE RELEASEeARIICLES 0• 00 PARTICLES 12 01 PARTICLES 12 01 PARTICLES I 01O.Ou0O O.0Q00 0.0000 G.0100 0.0000 0.000 0.0000 0.0000FRACTIONS OF COREINVENTORY LEAKEDKR-XL 01 12-GR CS-RB 7E BA-SR RU LAO.Ouol 0.0000 1.0130-05 9.9108-07 1.9411-07 1.2940-07 3.8818-08 3.B818-09J6. COMPUTER CODE CORRAL LISTING FOR BWR ACCIDENT SEQUENCES1* C2* C THI-i VERSIOtN OF COcRAl. IC FOý G-p tCr'XlEIT SEutiFiCES3* Cw* .i) I-r•.SToý! PI(9,2n},Tl(kfl).Glfqlq. n).VAPI~q.20),V(9),5* IvAP(1I.P(qI,TMRP(.),rELTTI(q,?0I.OELTeT(q).FLKP(q.0)I,E•-(9,9y2n),•* IEP~q,9o2O)-FFI;)(qeqilHT{G)pE. lI(cl)oEIPI9)-r(1)1),CMP(9.I,2fI)7* ILLK12{9,2O),CKG(-,),rYP(lr)-()-Atin)}0.B(In).* XICAOI(9,102,2.io*ICB12(9.00.2),!a* 14* ICSP;(9,20,),ICA12(9,10,?),7* ItCFP(9,2)01.5* LI'"FSIO' CcG•( ,2},ir*ýo*i•sICGO2(9,2):ICMI2(9,10,24.17 CSI?(19,21.1CGOI(9.21'1;* ICS P(9.2).1C IMtP(g~.,!* ICMTT2(9),k3* ICMTOI(9)..ý0 ITMI3(IO),t,*ITMV 1TM•flo), 110).,7. lOKICIA(10)3* LUIMFNSIOE' nFTAPG(9),.0 IdETAPS(9)g51* 1AW(q).AF(9),•" 1SK&'(9),1SKCI(9,10).34* 1FOp(9),j5* 1FD1219),jh*IHTA12SIS)131* IBTAI2M(9,1I).j8* IFTAI2G(9),jq" lCVTOI(I9tCVTI2(91)40* IGKNC I(9,10 1 .41* loKt!CG(9),42* IK;;CS(9)4 . 1,oKICA( 0):46* 1OK'.CA(9,10).47* iAT1IA2A(9,1n)I,4b* 11; T A Te b9, 1, ..49*•{*IAMCA*IP 11) .IAMLAIM(i0),b IAMr: AI A(IA),b4* lCID\ 0I9IO)ý,3"IFLoI(9),lLK~j~q,20),54* IOK,-!Ct3 9,10),Ub*ICATlI(q)b7* ITAI(lU),t•d*I1(lo).tO9ITA3(10)#ul* ITA(110),U2* *TAu(IU),)3* 1 IIEI5T19ý5* IvA!°(9).VII-211

1'b IST'9)."7# IPAS(9).t.9"I VM (q),70* lwA1q),71* IWS(9)./2* 1,S (Q) .73* IDA (9),",4*75*1012(9),1SC(g])76 LI.SIOh A.•,L,APA( In ),A'fPB( 10) 1 8TAP'PA(9,.0),BTAPVB(pjj1)'CAP(9,17* 110,2) .CBP(9,10,2) .CATp(9) ,C3TP(9),CVTP(9),DvA(10),DVO(IU)1 EFA(lO)7m* 2,EF, 101) SKGA(9.10.).SKGe(9'tln)'TVAI1.i).TVA(1n)194 DImIFNSIOl GR(g)t•o* 1SKGI (9),01l'ICOV(10),1CO({10}),o3'ICOA(IO),b4$ I ELAMr4(1f0),o5*Lb*1t•LA"A(I0),I CL J'A.L ( 10),bT,* IF.Fv (10I)o8" IDM(') ,9* 1TMY(9.20),TCiAlG6(20 )90* C CFR(JELISOT)zCO1F FRACTION REI.EASEr' FOo MITEPTAL I-;U1 UY91* C ET;0O JNEL.2•2' F•I.J.'rISIOLt CF(•I(,0),;'JAMF (10),CORLFK(IO):,0.6i* L(.,) VALE•CE(TA 4(I),TVAul ), (TR'.(1),T, i6(l)I944 C THt ISOTOPE ý,ROUPS WILL "E TAKE.! A' '-R-XE A.'n 01 bORA -,C I.lDInF5* C L1,K).,12-010(12 LIKr)*A14U CS-!:BTE.eA-SRPI,.PU(PART ICdL;.!E LIKE)'95* LATA NAMrS(1)/6HKR-XE /,ý!AMF.S(2)/6HOT /.NAMES(j()/bHi)bP 197* I1 1 A!'ES(4)/6HCS-Pb /.N'ApI1;(5)/iHTF /,NAmFF 16)/6HbA-sH /ý1* 2r..A'"FS(7)/6HP'H /.NAM•-';()/IHLa /1-,0* LOSTCAL FWITCHlul#• CO•"*1Oh•//A ( 5.n}, AbSMA X, AFS•'5, F'(5(}) ,F7,O ) ,FAP,ý (50 e 50) ,VF[ U,!t))1U2* 1,H(MO50) , II,If(U,.IRFL,LL,LOST*"APS."ARK,NTTNA, H.,NE.XI ,NT5Iu3$2•,PL"T( ,2,LOL) .,(5P,50),PEL•AX'RELPMoRtJP'(l(O) ,T(5O) ,TA;% 50)..TYPrlu4.$ 3.Ut.ITFUrNITT,W(50) .ZCSU,50),ZN(RO.,S0),ZNII(50)1L.5slub*shrA.rLIST/3PES',K/h,flflATATT!.DEt.TTODT. T2.TJUJVPFI.TJL•PP.iII .I ,EPET2, Al-, AFFLK 12,ELKi-,H'rTtH-Nr-,TE~n, TVRIlu7* 2.TVr,ý.,TVFEGK.rPEF.)PL,VAPI VFOP.FoD1,,TC"1,ELKOI1Mk, ,.F,TCý,iiFJ.Iud# 3CS12,TCSP2',HCS1, TC T-XI ,TCSIF CSIT ioTCSpIE.CSRJ.19' 4U)CSP.HCSFCUTUV,DCEI,TECCIVCL TMY.TCSP ,.ECCI,DF1v* 5, Tpi;FF, DFPP,.•FPO(,D FPI•,CFR.YPUJFFYC"MP, 'ANNil1* c :;CnV7' COMP fLnv, F~om MC-4P THRII A"N,WLuJS112' ,#A!t;= 1 ' nypAS ANNI'LoS11:* [iA/F.LIST/TRI JIA/ý1?.UR.SC.CK('.PTMP.'NST.VMPTAI,ý-•*,l'&u-EL12114* ..115"AVELIST/MIX/P, TMP0tDTVAIR.VSY.PASPSAVI;)ATP•. ABSZ" A) I 1, -s• /, AbSP-S1! E-R I,RFLmAX / .E-8 ,R-R•IF•lib*1 COP.TINUE11i7* !z h A {5,9fEAK, E•'!D=g9n)1±05 ,RITE(b.20n1)N119* 2J0 I F'ORIIAT'f(I,//I,32X,•btAFISý,ZOt• PP'ODUCT RELf.'ASE Aý4C CLc.A,..oP,//,1•[]* 15X,?0HOATA A.!ND ASSI.'mpTIOP'S./-6X,22HN(,. OF COMPARTMENrS iv= *12)11"* wRITL(6.?002)N.NOATA1Ž;*1,:3*2u02 FORHAT(IHO16x,P1HCnmPAHTt'FNT PRESSIRrS,/.6X.II4( (pT( IJ), 1:1,, 2,6H) ,J=I-,, 1 2#4}))1i4* LO 15O Jj1,UATAle•" wRjTE {6,2003) (PI ( I,J).,I:1 ,N)lut* 1500 CONTINUE1,7* 2003 FOPWAT(6X,1Pý3Et4.6)1..* i..R!TE(6,2004)NNDATA1.q* 204 FO!.'AT(l110.6X,24HCOMrIATt"ENT TE'pERAIUQES,/,6X.15H((TM.(I.J).:1I='1.0* 112,'6,f) *dJ,1-2.2W=)1.11 00 1505 J=I,SOATA1.2* wRITE(6.2003)(TMY(1.J).I=1N)1.13* 1505 CONTINUE1.,4* WRjITE-6,2006)N,NrATA1,5*1...bO2006 FOOVA) )H',6'5.-5ETEY'EHt'TRE OIFFEPEICES TIuLK-Tv;ALL,/.X,l' Ibk( [.ELTT l( l,J),•I=! ,, 12, H) ,OJ= ,, 12,2H): )I.a 7O1510 JI,.[0ATA1.)8* 1 .)96ý,PITE(6,2003) CDELTTI( IJ),I=IN)1510 CO-:TINLIE140* •,v T•(6,?00C)NiiC.A T A11* 2UO6 FOF!AT(IH0,A..2bIoWATER VAPOR MOLE rRoCTIOS ./,6X,142* l1bd1 (,AP!C(fd) l:1,,I,.') ,Jzl,, I,2L)1.3*14,4*00 1521 J1 ,::0AT,"ý.R ITE.(6,2001I (VAF'I1(lI j),1=1,1e!)145,5 1520I COt.ITINUEl.u*v.;RzTE L(6. 01 )N.!NA'A1,7* 2iG10 FOR'.,AT({1!;(C6, ,OvHFPACTIOI OF 12 LEAKWD PF) i'OUR,/,6X,1,9* :;0 1525 J I,':r/ATAiLuO#•.R Tt.(6,2003) {FLK1l•(l.J) ,I=I,NI]t1h* 1525 COlTINIJE1,2* ',HITL(6-201()WNr. A'AI',3* 2o 12 FOR'.'AT(IHrl,h*.b OtiF"ACTIO' OF PARTTCULATEc LEAKED PER Hj,.j, ./.6X.lb•* i 1l"P( (LLKF(I .. ~).,I-,,) ,!) Jl 22)='LO 1'530 J=I i.ATAIL.,-* ;.RPITL(6.P20 3)(FLKP(TJ} A1:.7- 1:,3(1 COI:T IIJL)E15t5. ,,RjTL(6-2016)N-fr.ATATIlt)VII-212

It,9-. 2016 FOPVAT(1If".n6)A,,214FPACTIOP' OF ORtRAN1C IODIDF; LEAKED jEo mOIIR FRn.IU* 1r.Or'ARTMEt:T I ,/T6vI7((FLK0I(1,J('1=I1,6H),JZI.,E?.2H)loi*'O 11530 J=1,NPATAI•24 11530 ýRIT• (b,•0(3) (FLKOT (I,J).1 =1 JINIo3*o;RITE(6,2014)trnATA, (TI(J),J=1 .NlATA)Io4* 2014 FORWAT(1H Ax,Ilh0HDE-CL9T!jG rOUALY I.•OFXF.) 1XPRESSIO',jS .,LRE INrEAElv5, 10 6I! COMPARTMENT (I) ANP TIMF (J). ThE INPUT TIME ARpAf IS ri.1,7* ",. ,RITE (6, 2?O1V TCP•A1G1.6* 22615 FORA7AT( 10-0 6X.72HT-IMES TCHANG SIGNAL RtIfITEpPOLATION OF FLOW RATESl,9* * AIrX THERMODVNAMIC DATA ./,6x, 7 TWTCMA'G'-"/,(6X,1PBE14.6))110svGRIT-(6,2018)N171* 2018 FORAAT(1H(f.6X,77HFPACTIOm OF PARTICUI.ATES RFMOVED PER ..oýUR UY ImTE172* IIUOR FILTER IN COMPARTMENT I ,6X,1,H(FPP(1),I1.'I2,AtH)173* 4,RITF(6,2003) (FDOP(T), 1,N)1y4* ARTT{(6,2020)N1"5* 2020 FORMAT(1HO,6x.76tiFPACTIOr OF IODINE (12) REP.OVEO PER HouR BY INTER176*1177

2$ t,.2 * ATTEf6,2003)(V(I)1T=l,Id)2.)* RTHL(6,PO~4P)ký4* 2u 4 ,3 FOr'AT(1O O6,•Iý.tICONTAII)'ENT AN[ CLEANUP SPFCIFICATIuN,.,//,oX.5S2:ý5*lIICST=CONTAIt-"El:T SYSTF)" rjJECTItr, PL'W It' CUBIC FEET/H,.,/6X,6Li2tbt,$ IHCSf;;=COlJTAJtAFNT 6YSTL' -10.1 RECIPCLAT-IjG FLOW IN CU.IC FEET/HR.2L.75 l./tX,61,*ICSI.HCSR=HEI(HT THROU604 WHTCH CONTAINMENT SYSTEM SPRAYSLlt•$i* if-ALI. ,/,(,XlHrC'IorCcR=r'AMrTER Or e ONTTN"ENT SISTtw SPRAY OR'^PL2ETS ,/.,6Y,6-,.CUTOV=COILFtITRATIO?) OF 12 BrLO;. wHICH SPRa4 RFMOVAI3S I!rJFFECTIV. ,/,6Y.ŽeHVCL=VnLUk4F OF COOLANT LIQUID ./.2Ai* IbX,7IHH0=RATIO OF M131i.E IN SOLUTION Itl WATER TO THAT L,, VAPOR .j2o2* ILOuILIBRTUM./,6X,3 7 HECCI=EMEoGENCY C'RRF COOLANT FLOW RAIE,/,oX,2 u3* 153IfFPPrFPOI,.FPI2=PUFF RELFASF DFC'NTAWINATION FACTO,.,./,6X,2,ý4$I*b3k.xPuFF=FPACTT3', 'F AI'PORNIF MATEOIAL RFTAINED AFTEI? pFF RELASE:2•5$ 1 ./-,,t,7,mDPE,0PL:EFFECTTVE lIAmErER OF rARTICLES 1:, MICRONS ATew)* I EA'LY AN:D LATE TIvES2u7$ ,RITE (6,2050)CSI .CPItHCSI HCSRtDCqIDCSP#CUTOV,VCLiOLCCI,2.C* lIbFpZDFPOI,)'FPI2,CSP2,XPUFF~nPEnPL2.4* 2Ub0 FOFPAT(lHO,b',IjHCSI=.IPEi-.8,3X.5HCSolr.]PE4.A,3X,5HHC.•,I=.PEIt,.82iU* 1,3k,5,HCSR=,IPF14-.n,3X.5H(CSI:IPEI4.Rb/,6(.SHDCSR=,IPF i4.A3X,2711 I(,HCI TOV=,'IPF 14 F.,3y,'4HVCL= -IRE I u. 8, 33HHI=. IPE14.8. 3X.:)IE~cCI=.2*1* IIPF14.c,3X'/.6X.bHDFPP:.,IPE13.7.3X.6HDFPOI=.IPEl3.7, 1 X.21.5* 18HDFPI2=,1PEI13.7,3X,SHCSP?=,IPEIT.?,IX.6HXPUFF=,PE13.7.//214# 17XHUDPE:tPE17,7X,4X•HCPL=.IPE13.7 )275* wRITE(6,2052)276' I'ISO=NOI+NI2+NrAP217* ARITE(6,2094)((NAMFS(J).(CFR(IIJ).T=1'4))J=:,l•NISO)?d8 20 5 2 fOR'.AT(I40,6.X.63HC"RE FRACTT(NS RELE!SeD 1Y METHOU J I,, ISOOPF 6R2i9* loUp I CFR(J,I),//,12XIIPGAP RELEASE,4X.1,?HMELT RELEASE.WXoLb0$215HSTEAM EXPLOSION.IX.12F1VAPORIZATIOC)2ol* 2t(54 V~1fA*N, I*1~~.-' E~m..~ ~*I PI2u2* C INITIAL CONTDITIOIS cF Cn,,ulTrr VARIAr'LFS2ý3* VTOT= ,2.,4 C vTrf=TOTfL CONTAT'EN'JT VWLUr.FTR2o5* LIO I IZ~l,'2at,* CGoT(l, 1)=2o7* CGP(I, 1): =.2ka* CSpI, I)= n.2V•0C6)(I, 2):o2•1$ LG7(I. 2)= 0.2,ý2 CSP(I, 2)= fl.29j3* .SCT(), 2):2-)4* CMTP(I)IO.2-i5' CMTTi:(I)O.29u,*297"CMT0I(1)ZO.CMT%,I)(1):.298* CVTI2(1)=O.2•4$CATf'I(1)=O.3uOi* CBTOI(I)=O.lulsCATn(1):C.352$ CBTr(I):=n.5.3- CG,2(1' I): =3U4* CS1,111, J)= (i.3,,b* CG: (I, 2): 2 .3u6'3.7*CSI2(I, P)= 6.vTCT:vTOT 4 V(I)3,8* 00 J=:I,103uq* C SET 1, CONCF.:T"ATIONS NO--ZE.O FIR C .ToV Tr';T ONLY310$ CMI2(IJ.I):I.511*31i,*CAIP(I.J.1):I.Cbl2t],J-l)=t-343* CACT(I, J, ))= 0.314* c'CBti(. J, 1): 0.31uo rAP(I. J, 1): A.3.7* CMP(1. j. 1): n.3184 CMOIll, J, 1)= 0.314* CACT(I, J, 2)= 0.4u* LHd'I{I, J, 2): 0.351'31ý*CBP(I,CAP(I,J. 2): '.J. 2)= 0.3•3* CMp(I, J, 2)= 0.3,4' CMOT(I, J. 2): 0.se5* CAI)(Ip J, 2)= 0.3,6$ C8121I1, J, 2): 0.3,7$ CMIT(I, J, 2): 0.31$' 8 CONTINUE329* TEX:TEXl3.51 uO I0 ,J:1 i,]3.>.e$ TM3(J)=O"333* TMU ).)TmPR-FI.OAT(J-1)$1T':F-T'R)/ 103,44 TA1(J): 0.35* TAI)J): 0.3-7$ TA3{J):o,3.;1* T83(J0:0.,i'ITAu(J)=TVRI-r'L(AT(J-t)*(TVR2-TVP)I/i"IUu)J}=TVH2+rFLOAT(JI)*(TVRE-TVP2)il)"341o 10 Cot.TINUE3-42$ TA: 0.343* TSA7 u.53i. CWLVP:0,VII-214

3", b* CSLKP=O,Zý6*CGLKP=O.3!.7* CMLW"2=0.3.d.*CVLK,=0 .349* CSL.KI2=0.3tO*CGLVI2=0.3=,1* CVLVI:e=O.31ý2* ~ ctMLK01=03 Z3.* CSLleQI =A3ýý4*CGLK"OIO 03)b*CVLKOI=D..•bb'COKFnG1=O03:8* OKEOG2=0.3,9* TG2=U.3oU*3,•15OKEQS1=O.GKISI=O,1S1:0.3.2* CKE1S2=0.3.3* TS2=0.34u4* FOOF:1.3o5* C FEGIN CALCULATIONS.o7* 7TT O.3c0b' C SAVr LVT3o9* i;T1U137'/*ICHANG=I371* 6 COK.TINUE372* IF(TT.LT.TMP)OT=AMINI(OT1,TMP)3730 IF(TT.GE.TMR)DT=.0 *(TMF-TMR)3'1*IF(TT.GE.TVRI )rT=..n5*(TVP2-TVRI)37b* IF(TT.GE.TVPI .AND.TT.LTTMF)P.T=.fS5*A.•I NI (TMF-TMR.TVR2TVHRI)3ib*IF(TT.GE.TVR2)DfT=.I*(TVRF-TVR2)377* IF(TT.GE.TVPE r)T=DTI137d* IF(TT.6T.TJLIVPI )DT=OT2319* TF(TT.LT.IPUrF.AfrO.TT+UT.GT.TpUr')rT-TPUIFF-TT+1 .E-830J* IF( T.LT.TCFI.ANDý.TT4IT1.T.TCSI)OrTZTISI-TT+1.E-A3)* I F ( TT.LT.TCSýu2.MAI0.TT+UT.OT.TCSR?)rT:TCSP?-TT+1.E-83o3* IF(T1.GST.TJU!PP2)rT=TENU-1T3o4* TT=TT÷OT3.b C rYDA-,ýS PNAREATFrf RFCOP0IfATInN IF TH.R.AOryNAMICS IS tNC.tANGEU.3oh* C CHA'IGES OCCUr AT T!MES TCHANr3,o7*,W'ITCH=.FALSE.Lob* C IF(-'.uT.,.A!o:;.,T.LT.TCHA!rG(Nr7HAý!C,))Gl TO 13!33LU9* 1JCH .G=NCHAr . +1390* SWITCH:.THF'E.oo" C CO.'r':UTF rI.),T(r),fELrAT(,],AP{J).EpI.F.1(J) FO,, ,r,t PARTICU-392*393*C LAR TIME TT.hN= 13 J14 * O 11 .J=7,ýnr''Tt3'o z) * IF(A T)(TI (J.TT).L.AroS(T (NNi)-TT) )N,=J3v,.*197411 CO.TItjLIENN=-.AXO (2,M~i.O (N1,r),ATA-I)398*399*Al:TI(NN-1)A2= rI (tjt)L4uu*A3=TI(NN+1)4Ui*4vZ*w,:O 13 K#t#iPWK)=0INT ( A1 A2, A-'-TT.PI (K ,N•--I) pIP (KNIN) ,PI (K-rtj+I÷))4u3* TmPr(K)=QINT (A A2A3TT.TTY(K.NN-1) v MY0(INt,) TMY(KiNti))U0* ;. ELTAT (K )=') I':T (A1. ^•.A3.TT ,nFLTTj ( K., N-1) . OELTTI (K tN!) ,uELTIIl(K.(,• ub*1,JN+1 ))IvAP (K) =Q1t.IT ( •I ,A2. Al. TT , JAPT ( K* ."V-1) ,VAPI (K, M ) , VAPI (K,-.4+147. ) T ,,A,,EL1W(.AtAA 3.17 ELKI2 (wNrl- ,ELK2 ) (K rp) , ELK Il (K,4W+1)4 L'* ELP(IK):INT(AlA2,A3,TT,L-LKP (K.NN-1),ELKfP(KNN),ELKP(K,idNNI))4u9* 1LOTtK) =:INT(A ,A2,A3,TTELKI(VfNN-I) .ELKOI(KNN)(ELKUIC(KNN+IJ I4 O* C CO.''uTE G(I,,JIEFIP(I,,J),EFIP(IoJ) FmR ANY TIME IT4.if* DO0 13 L=1-f! .W•Z* ~~~C- (K I, L) =(Iq4T ( •A ,Al 3TTKI{,'.+o ),I( ,,N G • • ,N I4,A3* EF12(K,L)=(ýIt,T(A1,A2-A3,TTEY2(lfLIN-,-I) ,FI2(KL,iiN) ,EI,kK,LPNt:+I414* 1))415* Ep (K.L):IlJTC(Al.A2.A3,TTEP(K(L.NN-m)eEP(K.LNN),EP(K..N.+I))4j1* 13 CONTINUE417* C1333 COrTINUE4•I4*4.9* C COMPUTE 3AS PARAI.ETERS.T;.ANS!'ORT CoErFICIENTS. OIMENýL,,LESS NO.54,0* C4el* -;O 14 K=1,N4••2*4,3*CALL ERR(3,1FHFRACASIF(.tNoT.SWITCH)GO TONO 1411024e4* P,(VI.P1K)+14.74i0' TR=TMPC(K)*460. RANKINE4e'b* (iNT (K )nK p() /( 10.73*TR)4aý7* ONST(K)=DNST(K)*VAP(K).1.+C(I.-VAPIK))*DNST(K)*29.46* C ABfVL IN LDSI/FT3.,.:9 * T RI5TR*S0R (TR)A* 0'-1*tVST(K):.no3339*TPF+5/(TRt 12?2.2)vAT" (K) =.04 1'te fTP/49P. )*.. °76A14e... IF(VAP(K).EO..)C0 TO 7-463* IF(VAP(K),EO.O.)GO TQ 7734-4* .'AS( =1.*SCRT.( ()(V^ IN(I)/VST(K) )*SRT(18./2. ) ) **2/..,-7U45* ISA (K)=(. +SRTL((VST(KL)/VZAIR/(K)*SART(29. /18) ))**2/3.uOaVII-215

4.;35 #OW TO 7794, 9t.* 7 VM(., )=VAT P(K )-.,0,1 ý0 TO 77944, * 773 VWf)K)=VST(K).41 '7;; co* -1,.LIE.43*4'.,.l•C ,HI"VL IJ LU/'T/LRý,-A(K}--O,707ý.÷141,7ý/iTtJ/1,8)1$5* wSc,) =3 7n0754-•4 7.7/( TR/1 .8),4,t,b C -A .)=,Ot2 ,•35(TR/l1,P*iI.5"•.lQ61/( P(K)/14°7)*4".43,*.•.WACK})),..9l' C 1*Wf;)I.)) t;" * ( ( I'6')fR * I,•~2 3 / I )! ° )3g I *cU rjS 1 0 %.?0Op , I -I T (KI R p ,+ / I Fk12 )!/. •. ÷ % R ,) ( )* 0'4b~l C AOTE DIFFUISIVITIES IN Ctp/SFCS C I? (rK) I'! FT /I"Rr'CCC0' ILl 'O= I cIsCOSTTy/,(f:F-ISTY*r'IFrUr ,):C( .)4jb,',C(GC)I=V A (K) / (uL9ST (K)*[.L2 (K) )7* c ,PA-i.UF rUn POP NATil-AL CoNVrCTTO.NCG=MT CrnEFF IN Fl/, 4 i4:,9* ;HR{) =•731£+HTTAT(K)g**DT K K**2 LTAT)K)/CTI)V1(K1)*2)'JuO• iF CGrSC-LE.I. +r•) CI"O(d)z.•4.{'3R?(SrOHT(0R*0Y:})J2(d)/rlT(')4L lIFV(kSC.CGT.] .4+') CYG (K)=.13*SBRTGPSi)*DI7(K)/HT(K),.l2* 11U2 COrT IN JE4.3* C CO'.'PUTF ,FTTLII'G vrLOCITIES foR PAr•TTCLFC.,* C ,,Av- •AR)TICF1 rn; 1! SO•PCES *OAP RrLFASEr.n TIMr I,jT).KVALS OFL,* TiIF AELT RFLEASE + 7 FAPLOSIMNS + ;0 VAPnRI7ATION RELEASES.¢ u C L'e•. 0 r xI O X ,IV V (14,,7* )G= l;-iAX I ( DPF + I( PL -PPE ) )T7/T$ )*O)PL 114.6* i;EX I=Af'AX 1 ( t)iE+ ( rPL-DpL ) - ( TT-TEX 1 ) /Tr* o nPL ),Ug$ Ir(TT.LT.TFXI) r)EY l=o.'. 71#IF(((TT.(:T.TCST.A•nf.TT.LT.TC•IEI.OP.(TT.,T.TCSR1.AND.TT.LT.ITCSPIE).O•R.(TT.GT.TCSR- .PND.TT.LT.TCýR2E)).ANU.K.L•Q,1)SPRAYzl.47.)* IF(TT.LT.TMR)GC TO 1955,14, ? * O 15 IrI:l , to475" 4M( IL)ZAMAXI (D"L, DPe+ (UPL-DOE $TTT-T (IN) )/TO)476* IF(TT.LT.TM4(IP))D'(ID)=0.4177 C t'ARTICLE T)IA:ETEPS A"F it AICrPONS4'9*,oO*1r CO,_T INUE1055 C1) t ITMUE4•,.S1 •K,•S (i.) = ( I. PrPAY) 0( I 7*n(;**,'/Vw ) )*t-F ( )/V (K)402* !ýKGI (K)=(1.-$F'PAYI) * (.p15.T7*EXI*2/V(K) ) *Ar(K)/V(K)q,•o3' IF(TI.LT.TMF)G3C TO 3010,44* DO P7015 T) 1,l1 1=.$ .'IVA ( D) ) =A 4 ' AA C PPL* P9F+ 1 I.-."'r'E)*( TT-'VA( Tr) )/TO)4 ut)* IF (I TLT. TVA f In) r'A I U) 7:04,t 7- VLU ( Lb) -ArýAXI(roPL - 110F+ (rl'L-OrE) 0 (TT-'VP ( Tr) )/TW)4od 4 IF (TT .LT. TV" (I1 ) )fVp( IL) .0.4ý4* '.,K r. ( K ' [')=:( l -SPRAy )*$( . 015 ?*ID A ( I ) $*2/V• ( K) )*AF (K )/ , K£)4',0* %SK-"(KIf')=(I.-SPRtY)*(.rI595'*ODVflI•O*)*2/Vf(l)}*AF(K)/vCK)wj I * e ul, CO-T I;4IJE4V2* 3 u 1 ),r cO !T I71LIE4'3* C sK--=:J•A rJAL ,:EPOSITIG;T I_-O6 I" W4-1,4Yb5*114 C0,T II, I4Wi),* IF(VAI`N.EG.1) rO TO 3t)O4I'7*F ".1 ) .OT.o.L) TCr)",P' -44"•,9" IFC(ci.5).LE..flI.AC43.GC1,3).LE.E.O1) "-0 TO %4')f4:=Soo*VEL=G(MCf.VP,•)/225.7ci, LRY .C .ELL A'1r:rLLIS RryrJoLPS NO.- REA.,N%O?0REANNO .333*DM$TM(I*VEL/VM(I)503$ IF(?LArr,.LF.,lfnG.) ,(r, TO .O13IF( LAlIN.GT.,.tVU.) G,) TI ;•P0¢t-•5* BOUb !• FP (XVIMP , 3)-- 0.,5u" kID•••'P3=.'v(lI O* I()*CPEAtUN*OC(1)*oOO0,j 4 .*.•/VEL) •5u7# ,O TQ 6n3 S35!)V* 8 U I-P( C'COMP, 3) =. P2LF17 (r!COvaP. A1=1 .- EVP (-15( . *DT2(1 )SpE'NP**.8,SC (1) *.3/vEL)5r0* 8OU3 IF(rFI2(vCOMf,.3).Gl.o9) EFTF21(-Mjr..)=:.abls 8u04 CO) TINLUC512*I C CARDS FOLLOWIIIG TO 35420 TEST CHARCOAL FILTERS FOR OVERLOAD'513* C OF RAUXOZOOINE5i4* C4I2=CMTI2(4)*CFR(2.3)+CVT12t4)*CFR(IC.31+CGI2 4.1)*CFR(l131515s FILCAP:.000007*G(5#4)*POOF/6G. "516. XSI2=C4I2-FILCAP517* IFtG(5,4).NE.O..AIIU.XSI,.GT.I.E-8) GO TO 35421518* GO TO 35420 uo -c519* 35421 POOF=O..520* WRITL(6.35422)-521* 35422 FORmAT(1I8 FILTER OVERHEATED)IZ=• 0522s 35420 CONTInUE523* C524s C525* C5uLu C SPnlY LAWMBfA9 P'EED TO LE COMI4UTFOi ý'O"+b,,7* Cel, C pA:,TICILJTE ,PPAY I.A"PFA IN rOMflARTMNT I5•9* CDjO* C IRsT ESTARLISH W0U(FT/V1)) AIM• IMXPACT FFrICIrnCrVII-216

t,.; I$ F I =AV A X I ( 0 - AM T N I('r-I C S I -A%4 T N1 t nT /2. , ITT -T•I7 /2, ) L 1• E- T CS I502* 1*CST").5* F R z XFRIAfmAXI (i . .AI .TTT-I.TCYR2-PMTN (fT /2, TT -TCSR )/2 ) ICSR1.E-"'Jo*•7*5J4" ~1TCSrP1 ) T*CSRTffob* ~FR•.n-A•AX 0. *r~ oAI:TT..1CtP2-P MIt; If /' ,2 o (TT-TCSR2) /2,) , 1L•SR2E-ITCS-)))*CSR2fA=(AFRHI+FR?+F I)/VT()I F (rA .GE 0• . A1l . FA .LT. b. 02 ) Ft 15. n2 'F A rIF (cA. GT. • I0 E=.,lr)a 1555 C THEL: tF.FICI~F:CIES AifC Or ,M GAP PELF-A;,E ONLY543* C IN ThE EVEMT THAT TWF SPI'AYS ARF wflRING AFTER AllY EiP. ýSION AN;?b4* C SOLQ-CE RELEASE S,Tý'c EFFICIECY CF R-MOVAL KUST BE RECALCULATED54," C FOP RE-YOVINC, S.THE PEMOVAL EFFICIEýty IS FFX5.o*' C5-, l* F 1Z." 'AX 1 (0{1•,ýIH•NT 4(TT-TEW 1-Au Till 'T/2..( TT-TE 1 )/2, ) 'TCI;I{.-'•,6"ITEX )t)*CSIF1ZV.AXU1 T (0 , I JIT 1 (TT.-TfX 1-AMTNI (lT/2., (TT-TEX1 1/2. ICxlE-.be, u* 11ExI }) *CSRI5i.)i1TE~t))*CSRI1F2--A,:AX I(0.. ^•:TNI (TT-TEX I-APv rII (DT/2, , ITT-TFYI)/2, ) eTC•,,2L-5t.2#ITEX l ) ) *C'1?2ES•z=I+F24FT.,4.* cES-VrS/V(1)IFT(F .GE.O..,':,r).FS.LT.U.002) EEY=-1 5.825*F.6IF 'S. GE. .•f l ArJ• -.ES. LT • .0 103) FEW=. 04625- 0862b6+4e d.o*E%) **. /25•h* ~~IF I,-'S. GT • .0) "3)E --0n.5 9* C CALCULArF EcF. FoR PAPTIClLATES FROM MFLT PrLEASE FO< M-Ch OF T,.s•O* C 10 OýIVISIONS, CALL IT EFV(J)5oi'*562*IF(TT.LT.TMPTGO TO00O 17 J=1-1017775ý30 -AMt.•.'P ( J ) 0.!,o", IF(TT.LT.TMQ4(Jlf)CO To 175,,5* FMI=AMAXT(0.,AtI'"Il TT-P'T(JT]-AMI;jI {OT/2 ,(TT-TT4(JT))/,.).Sun•*507*ITCSI.-TMt(Jd)))CSIFftl=AMAXt 1(0 , A•IP; ( TT-Tý', (i)-AMI~t (D'f/2.,,(TT-TM (i) )/2. )boB8* 1TCSR1E-TM4(j)) ) *CSP5o9* FM -AMAXT(O.,4,'ZINl(TT-TMl(fJ)-AMINI(OT/2.,(TT-TM4(J))/2I,5 i0* I TCr'ZL-T!4 (J) ) I *CSP2571* iT" FvI +F"4Fr:2512*b73*ETmET/VTl)IF ( ET.GE .O .AlIr.,.F T .LT. 0.o2)rFM(J)=-15.825*ET, ,.06514* IF(7T.GE.,O02.ANV.ET.LT.O.0I93) EF-,(.j):.04625-(.Od626+4..68*ET)**s575" b*/21.34b l* IFfpl .01. .••E;3) FF'(J)=.nn15t 17* C COWULTF51 * IF(TT.•'.TCSLo. Nr.TT .L1.CSI') ^'rAM.(J)=AMý)AMP(J)4I.5bIoCSI*CST*LF5 790' 1:1ý(J)T/(DCSI1*V(1)/30-48 )5SL0" IF(TT.GT.TCSI'T .Ar'.,TI.LT.TCS, IE) AYDýMP(J):$AMDAIP(J)÷I.,*HCR*C¢(i* 1777 COT'IWUE56* Z 'cLLT RFLEASE PARTMIrLATF LAMnAS AR- OOMPLFTF5.7* C CO"rUTE OTHEIý PAPTICULATE LAk'DAS NfW5o8* AMDAF'G=O.hUg*AMNOPS=9.'ju* IFrTT.GT.TCST .ANr',Tt .L' .TCSIF) Alr!AP;.4OAPc+1.5*dCSI*'-.I*E/(DCSI*b•L0 tv(11/!n.48)592* IF(rT.GT.TCSi.oNP•.TT.LT.ICSIF)A-'rAnS:AMnAPS+IT. .HCSI.Ci*kLEX/(PrS5,3* lT*v(1)/30).4FT594* IF(TT.GT.TCSPl.AND.Tr.LT.TCSPIE) APDAPGAMOAPG+ 1. I*HCS.,*CSRI*E/tUC5,b* ISR•'y{1/30,4T)5,0,* IF('-T.CT.TC'1.,*.A'!D.TT.LT.TCS"IE) AMDAPS=AVOAPS+I.,*HCSw.CSPI*EFx/(:3-74 IuC•p*V( 1)/3p ,4n)51-iti*IF("T.GT.TCSr.2*.A !D.TTr.LT.TCSo2E) A!DIPS=A-'DAPS+] .5*Hc5S,.CSP2*EE"v/I599* 1jCSq*V fl)/3q.4A)6_04 [}i, IF(I I.GT.TCS:.'2..Al.D.TT.LT.TCS2?E) AMýD APG= A•.P0•+T, *HcS,."CSR2*E/fL)CholtI SR*v(1)/50.',)i-2* C CO',PUIE vAPOrI7ATT'4 pAvlTICULATF RFM-VAL RY SPRAYS,,O 217 J=11il14-AMDAPA(Jd)=O,"U5* AMAPB(J)=fl.6ý6*•bT*IFtTI.LT.AmI }(TCSI.TCSPITC',R2).Or.TT.GT.A.-AXI(TCSIEISRIE_1TCSPZE))GO Tn 217Ubt* IF(TT.LT.TA{.4(J)3GO To ;e17bd9* FAI=AMAXI(0..AMIlt•( TT-TA-[J)T-AMIII1( DT/2..fTT-TA4(J))12).)610* ITCSIE -TA4(J)T)*CST611* FAI=ApAX1(O.,AVItIl(TT-TA4•tJ)-AMTIJI(O,/2.,(TT-TA4(J))12.),b!2* ITCSRiE-TAL4T(J) ))*CSR1bij*' F A2AMAXITU..,AIh YT-iT,',dJ)-AMT'I IOT/2.,•TT-TA4(J))/2.),140 11CStnkE-TA4(J) ] *CSP2615s FBI=AMAXI(O..AMtM.I(TT-T[4(J)-AMI•i ot/2.,tT1-TB4(J))/2.)I6bi*ITCSIE -TRt.(J)))*CSYb,7* FB3-AMAX1(0..AMINI(.T-T•R4(J)-AMIN1(Dr/2..(TT-TBh4())/2.),6A8* ITCSI•IE-TR4(J)))*CSPI6i9* F82=AMAX1I(n.,At'IF1(lTT-TR4(J)-AMN1( DT/2..)TT-TR4(J))/2.),.?0* ITCSP2E-TB4(J) T )*CSP?6dla*EA=(FAI+FA1,FA2)/V)1)622* EB:(FHI*+FHI+F82)/V(1)6"*• IF(FA.LT..0193)GO To 1750VII-217

..: 4 * [.FAfj^t le OIn0-2%* GO TO 48b066b* 475n IF (i:A.LT.o.Ot21GO TO 47Fn.,7* t: F I (,J -- 0 4W2', - S PT 09ebP 8 +4 .,6 8 E A I 13 . u'd"s'L-,O TC 4B50O6Z9'* 47b0 EFA(J)=-15.325*EA+.O06D.)U 4L,5U C O0 :T I N UEt,), I*0.)2*IF(rE.LT..n1'3)GC TO 5750;.'F[ ( J) = . E 0 1.56.$3* 60 lo ",ý0(.:4.* 5750 IF (E3.LT. .wl2)G]G TO 5761oj6$ 60 TO 58q.O6,7$ 5763 F'Fri)J) =-15 .A5*Efl+.0 6bib* 5.50 COý. T tIiE0.J.9* IF{ T.LT.TCSI.0i•.TT.GT.TrSIE)GO TO 5,54t•,4* AMC.IPB(J)=1 .-5*"CSI*CSl*EFO(J) /lnCSi*'!,(1l /30.48)6.+e* 5054 IF(TT.LT.TCSi;I.Or,.TT.G1.,CSR1E)C-0 TO 5356IMCAPA(UJ=AA,,APA (J) 41.5b*IJCSR*CSrt*FFý (J)/ (rCSR*V I)/jO .,8).45* 5.55b iF(TIoLT.TCSP2.0:.TT.G1.TCSR•E)GO TO 217t)ý7*A•I.M APA(.J)A'=AP A(J)4 1.b*I;CSR*CSP2*EF" (J)/(DCSR*Vh()/jO. d)f64t* Z17 CO,;TTI%-JEo"9" C ALL rARTICIIL4TF 5P,0AY LA.tr)AS COUPLFTr(,)U* Ct,'* C IO'i,;E'L'. LMOVAL HY S(rAY rALVIJLAT-D NO.qb0e * Cbt3* C SP•1 LA; .'t COr"'P!TEn Nr. FoR 12(AM'AIT,.AMh;AISeA.,DAIM(K))b,4* C6ýt)* C CO'PUTE TERmINaL VFLOCITY OF SPRAy tlOPSt)*Jb* C6-,?. FDPrIz(4./3.J*F2.Ž4*r.s;'T(])e(r CSt13.,,)*•.17312UOO.lv,4(1)e"L6t5* FDREK:FnrrE I*(C"SR/%CSIJ**3tb"g" IFFLFORFI.LT.10700.) RE•=(FDRE I/15.7)**,7027b64* IF(FOREI.GE.10700.) REI:(FDRI1 /6.477)**.621qt-,• I *IF(PLRER.LT.17O00.)RFN=( Fl)RFR/l5.71)**.7n27;.,•.2 * IF(rORER.GE.o10"00.) RERNFDRFR/6.47j#*.,215,.o3* $jT=NRLT*VMI()/(0t.ST(I)*DCSI/30.UP)6.4* uTR:RR VMM(I)/(OtIST(1)*DCSR/30•'• ')Q5b* c TI,UTR IN FT/HR6oo*Go7*C MASS xFR COFFF TO THESe •, ROPý CM-PuiTFD RFLO. CGGIGG6)GGI=LI2(1) * (ý.+.6*REI**.' *SC(1 )'*.33)/{OCSl/30,48)IN FT/HRDot* GR12(fl'(2.-.fA$V4R**. .SC( 1'*e.331/(OCRR/3U.W4)* c (,As FHASF CO,ýPtETE.,ow C'"PUTE Lloiil" PHASE M'ASS AFR CILFF.GLIIGLR6706/'1*C UL:L1,UI.C VTSCOSITY.Cp,TI=(TMP(1)+460, )/I.8G.L:OIFFUSIVITY OF 12 IN H2#O.T2/HR672* uIIO./12.14B2. -673* IrLt.5137E'07*( T I/IIL).74*,I GL l=(, 5R*r) / (0("SI/ltn .4bl5$ CLpu.° 53 L/(OCS 0 /'xf .4b)T7b* c STA.;,JAT DRO" MODEL. UEIJD,FFFICIFI;Cv:,-EXr(-;,.KGOT/S(-i+ýo/KL))D77# ECsI=1 °-FXP(-6.*eGlt*(HCSI/UTi )/{ (DCST/30,48)* (HO+6GI/r-Li) ))U16* ECSP=! ,-EXP(-6.*GGR. (HCSV/uTP) / ((0C5"',/•0.4%3) * O+GGR/GL14)) )09* C SP,01 LAtfIOAS=F*wO'E/V.IF C/,(1)MGJ.r(EMITTED)/(',UM v(I})*CUTOFFw.41* C EQUILIRRItIM LAfItWA , ISL-) AT tOWn•? COI.C.upi* C IN%7rT A Sr'rCTAL DFCK T0 ýiAN'LE F• 0 IRI• !.' CALCjiLATIO., FOR FACHb62* C TYPr OF SPPAYo0(3 COG=LGI2( I,1) *VTOT/V(I)644* COS:CSI2(1,1)*VTOT/V(1).ýQ5$b,.,bo*UO 14 J=1l10C~k, tJ) =V• 12(1 -J, I VTO7/-'(1) /.Iti,7*o0d:*CO[{)(J)C =I2 (IJ.CO• (o.) C ,AIP'( !.J,I *VToT/'J (I)/( n75,?*EXP(-. -I *FLOA r( -J-1fl)I) * v T v ( 1 ) / ( •75:?8.EXP( -. 1*FLOAT ( -1 ) )049$ 1. COtJTINUEovO$('KG=C."'i I*OKS:6.6,1* IF (.OG. G' .CUTOV I OKG=I.t1j*uV'+e*IF(COS.GT.C[iTOV60 1I• J=1.10) UKS:1.'15* 10•*0 K"( J) =0.CKt(J)=O.bj7 *OKp(J)0O.69* • IF(Cijm(J) .rT.ClUTOVIOKv(J)=l..-)•jq*IF(C-JA(J) .GT.C''TCV).KA(JL1I74uU IF(Fl (J1 .GT.CIHTOV)OKP (J)=1.7,j7a1*i! *19 COv:TINUEIF( T I.LT. TCSIE.At:r).TT.GE .TCSI ) f'KC-,=.I-7(:3* IF(TT.GE.TCSIE.OR.TT.LT .TCSl) OKCsIO.7u44IF(TT.LT.TCSPlE.AfJfr.TT.GF.TCýRl) OKCRI=I.7.5* IF(TT.GE.TCI,,E.JFR.TT.LT.TCSPl) CKCSo1fl.106* IF(TT.LT.TCS,;2F.ANP.TT.GF.TCSR 2 ) O)KCSR2=1.7w7* IF ( TT.GE.TCS,'IF.CR.TT.LT.TCSP2) OKCSr,2=O.7(,9* 1CSF?*ECSPeOKG*OKCSP?/V(1)710sAMýA I1% (HO*O* SiVr 1) ) *(CST*ECr I*0KCSI÷CSRI *ECSR*OKCSR i+cSR2*ECSR*OK7.L14 1CS 2I)7A.,!'LO '0 J=-l ln7iJ* tMLIm( JV)(Ho*OKv.(J)/V(1 i)*(CSICECSI.OKCSI+CSRI*ECSH O,•;SRI+CSR2*7j 4 *1ECSF.OKCSR2)715* AM.'AIA(J) (Ho0*fKA(J)/V(1) )*(t-SI* FCSIOKCST+CSRI*ECSR*OKCSR÷+CsP?*7,6'1ECSR*OKCSR2)VII-218

717* AMDAIIA(J)=(HC*trKR(J)/V(i ))*C(SI*ECFI.OKCSI CSRI*ECSR.O,•CSRI÷C5P2*71.1* IECSP*OKCSR2)719. 20 CONTINUE7eU* C SUMP LIQUID vOL. AND 12 CONC. IN THE LIQ.CONTROL EOUILIURIUMPVSUMP721* C EQ ECCI÷VCL#CSI*(TT-TCSI)7 '* IF(TT.GE.TECCI) OKFCCI-1.71.3* IFITT.LT.TFCCI) OKrCCI=O.7.,4* IF(TT.GT.TCSIE) 0KCSIE=l.7,•5"IF(TT.LE.TCSIE) eKCS IF.I)7i;o* 0Su:.PZVCL+ECCI*OECC2+7:7* C VSuMP=TOTALTCSIE-TCSI)SUMP LIOUIU.FT3CCI*OKCSIE÷(TT-TCI)*CLSI*OKCS17,.* C CALCULATION rF LA'4OAS FOP EQI'ILITBRLU, PROCEED PELOW. T,-.SL ARE7Ž9. C F'LA%ý,ELARS,ZELAr.'(T).FACH SPVAY TYPE REG'rIPFS A SPECIA, UECK TO7.0* C CO.'FOuTF THESE+. FOR THE CASE OF Mo ,P,'AY. uSE H-00 DECK.7A'I* C7.*2 C tiOrIC ACI[O.HY'•0• DEC- FvP COwUTTNG E.UILTR1IUM COCCE,,7.,iTIOwS7,L3* C LDESj,,'ATE TGt Aii[: TG: AS TIPE 00I'Tj- wHFrI C.GI2(1,K) E(.EEOS ANO7.54* C IS LT THF EOUIL. rUTOFF.CUTOV=.0l FAR ALL SPRAYS EXCEPT THIOSULFA705* C TE,%0O* IF(1.G.GT.CITOV)G. TO b51127.?7* IFT.).OT.U(TT.GF.TCSI.AN.rTT.F.FTCSJE).OR.rTT.GF.TCSRI .•,,.U.TT.L-.7.•h* 21C5u•iE).OR.(TT.GE.'rCSRp .AND.TT.LE.TCSR2E)I) GO TO 65427,.)9* TA=TADT7•.0*LA'bG=ELAM(TA)741* 60 TO 6S457,2" b4,1 CO,"TICJE714.* FLIAfG=O7ý,*T A=I.7.5* 6545 CONT114UE740* C F.LAVG IN I/;HR747. IF(COS.GT.CUTOV.OR.TT.LT.TEXI)G' TO ,5A374.t3 IF .uOT.((TT.GF.TC'TI.AND.TT.LE.'rCSIE,.OR.(T.GE.TCSR1.,.,.U.rTTLE.7-9* 21CS"•1•).R. (TT.GE .CSP2.AMiD.TT1.-E.TCSR2E)1) GO TO 6b5•37:,O*TSA=TSA+nT7t5s * ELA'4S=ELAMITSA)7Z2* GO 10 65P57L3. 6ý,83 COTI.ITINUE7t4ELA"!=O.755* TSA=O.7-6* b685 C0,4TINUE7,7* C ELt.f'S IN HP-I7z,8* IF(TT.LT.TMR)Gf TO 9OoL7b9* Do 1000 jIl.07.U* IF(COMtJ).GI.CI'TOV.OR.TT.LT.TM4(J) )Gr. TO '1237ol* IF(.;,oT., (TT.Gr.TCSI.AND.TT.LETCSITE) OR.(TTrGF.TCSRI.. ,',U.TT.LE.7uL'6 2TCS;0lE).OR,(rT.GE.VCSR2.AND.TTLETCý,R2E)II GO TO 71237b3* TMI(.J)=T(.IJ)+DT7•4*ELANvJ)=ELAC(TM(I J))705* GO TO 80007ubo 7123 COI,4I NUE7tO7ELAMM(J)=07o8oTMI(J)=O.7o9* 8000 CONTINUE770* 8001 CONTINUE7710 IF(TT.LT.TVfI)GO Tn goUl7/2*773*O 'o0o J=1.1OIFICOA(J).0T.CI'TV.OR.TT.LT.TA 4 fj)GlG TO V; 4 4774* IF(.i0o'r.((TT.GF.TCSI.Ahr..TT.IE.TCSIE).nR.(TT.GE.TCSRI.•I.D.TT.LF.775* 2TCSPIE).OR.(TT.GE.TCSRŽ.AND.TT.LF.TC.R2E))) GO TO 36.,776* TA(,rJ) =T1c(J)4prT777* ELAVA(J)-ELApn(TAI(,I))77.1* GO TO 9n007"19* 3t. 4 4 CO? TiTlUE7;0 * 'LA'?A(J)=0.7ul* TAIJ)=O.7t,2" 9000 CONTINUE7.5* 9001 CONTItUE7o•4*IFfTT.LT.TVR2)r-O TP 5go±l7ib*7"A.,*Lo 3000 J=l.O0IFtCG{p(J).GT.CLiTOV.Ouu.Tt.LT.TB 4 (J))G," TO 4R1l7.7* IF(.1UT. C (TT.GF.TCSI.A, .TT.[E*rCStE).OR. CTT.GE.TCSR. 1 .L.U.Tf.LF.TL6* 2TCSI'E)I.OR.(TT.GE.TCSRp.tND.TT.LE.TC-RpE)I) rO TO 46217091* TiI C,.)=Tnl(J)4*OT7'0* ELAv I(J) ZEL At(TH I (J))7,1* ;,0 TO 5OO07v'e* 4i.21 CLOT I rTIJE79ý3* E.L AV•. ( J) ZO •7 .ý4,* THI.(J)=O.7 ;b* 500 COtiTINUE79I,* 5001 CO•:T IiUE797* C I:LAI':-(JI I'l H l7b-d*7v~9*CAT THIS POI,:. Ft O'l IRP.11" LAkl;64; 4Rý COOPPL7TE!uoJ* C rIFrEruE,'TIAL EONS (AS UIFFEq'FMICfr EON-) AýV. TO PE SOLr--,I NOW:;ui' CTnu2* C FST43LISH RELEASF t4PiITS AS INITIAL rONCFNTrATIOnS F)R cxPLOSIOt,6.3* C IOELrASF AMrP EtT RFLEASE.FOr r.AP RrLASE, ICS ARE ESiAIHEDt)04* C INCLLDE VAPn;H PELEAsESt; *8,b,C EXPLOSIONJ RFLEASEIF(TT.GE.TEXI.Arid.TXQ.r.o. )TX=TTRU7* IF(TX.EO.TT) OK ICSrI.Alu8*o.9* CIFTIX.tIE.TT) OKICS=O..FELT kELEASEVII-219

,.0 00 3 J=-lfln,'4il' IF (TT.GE.TK.4 (J) . A NCI.7m.5(,J) E0.0,.) TMý(J)=TTIZ*1 IF(TA43(J).FQ.TT) OKTCm(J)=.l6,304 IF fTMV3 (d } NE. TT" O) ICM(J)=0.dl4* IFITT.GE.TA4(J) .ANO.1A3(J) .EO.O.) TAN(J)=TT015s IF(T.GE.TB4(J) .ANr.TB3(J) .E('.0.) T83(J)=TTo~bo IF( TA3(JI.EQ.TT) nKICA(JQ=.C.752AR*XO(-.1 *FLOAT(J-1))R17* IF(t Tf3(J).Eo.TT) 0(1•1B(J)=.204'.*E4Pf-.I*FLoAT(J+9))81a* IF( TA3(,J).NE.TT) ()KICA(J)=).b19* IF( 53(CJ) .NE.TT) OKICb(J)=O.6eo0 23 CO?-TINUEk341* CU%,2 C .0 ITE. DIFFFLiCE EOi;S IFO,: GAr P CONC. IN COMPARTME.,;T i TO N683* CALL [PR(3,1.3HOPRAL .f!. 23C,,4*0O 190 IVAT=1,3•5*8,:b*C SET ALCE'Y FIXED I'PUT VARIAnLES0•=16.7* I1=2*N6-46* LL-1lj.:9"GO T0(30.60.;2,),VI.ATA)g* 30 CONTINUE6JI* C SET OFF PIAC-NAL AAT,TIX FLEMIENTS FOR P BRANCHO0 24 IC(M=I.Nli.>3*H { '~txO •+N. ICo,,,)=El P(ICoM)8.)4s DO P4 JCrM=O.Nt)ýb f 24 ?{C o r.', I cO•s) =G(f I C 0tj CU vp (I .- E F 1(C MJ r 0.V )V(I COr.,c.ýb"ý)O WO IREL:I,bB,,7* GO TO (31.41,51,1061,1071).IREL8.)8* C PRECLDINrJ PRANCH I- To GAP RFLEAqEMrLT PELEASESTEA,., Ai-LOSION,k.J,4' C VAPP'4.IZATIOt, RFLEASE eHA.4E A. %)APeRIZATTON RELEASE ..HrE SR40* 31 COITINUE6,1* C !:ELEASE IN Pa.TICIJI ATFS PDANCH6 ,3w T(I)= DT6ý40 CO )7 TC0VNS,Nt'.5* HQ COm, ICON I=o.t]3*7*IET { ICO7=OiSKG1) (TCOM)+=LP(TCOP-'().IOICDv)e47*IF(ICON'. EG.* lFT APP IC Ic I ETAPC,( ICO,) +A%'0Ar-G3.o F (TF ) =CGP6( C jrO.1)9ý#5tlU*rF ICOIO+N)=O.IF (,.%EQ. IANo. ICOM.Fl, 1 ) F(IC(OM):I,tlz•1"00 1.9 JCOM=I,NbO2*s"•3#317 i4(Ir:.,ICOM)=H(ICOIICo})-G(TCOm,Jro.O)/VUTCoM)37 1 ( I • ,", TCOk ) •i ( IC0'%ICO ) -8FTAP( (leO".)854* CALL ALCEMYst5bO00 38 ICOM=I,N5,•6sCGLKP=CGLKP+FT( ICO4t4N,1)a7 # 30 CGp(ICOM.2) =FTfICO".l)P.b* C GAr r.ELEASF pARTCVLATLS COMPLETE FO" PPE"ENT TIME SrE,.e,49* (60 TO 506uO' 41 CONTINUE8u1* C IF "ELT RELEASE FiAf NoT STARTED. 9YPA.SSpok*, IF(TT.LT.TMR)GO TO 538o3* DO 45 JREL=I.10bu4' IF(TT.LT.TM4(JREL))GO TO 456ub*I (I)=AMINI(rTTTTil#4(.JHEL))8o0'DO 4.7 ICOM:I,NBo7#FQCXOM)=CMP(ICOM,JPFLl)8b8* IF(ICOM.EO.1.AND.OKICM(JREL).NEo.) F(ICOM)=OKICM(JREL)€•,.,9.F (IC£,yN) :0.67()* H(CnI~'tk 71* f.ETAi TCO.,.j,FL, (I =KGC (ICr MJu'ELI +ELP hICO")+FDP (ICOe)ii.72#IF (!COFQ.EIG)HETAPV(ICuV,.JREI )=PETAP'"(ICC•"•.JPEL)+AMU,,M,4(JREL)i.13"ro N49 JCOtMliI,N674* 49 ý!( I,ý .ICOM)=H(ICOM,!Ccv)-G(TCOMJCO-,)/V(IC,7M)%, 75* 47 (CQr:"ICOM)=HC1COM.ZC)-RETAP4( ICO..JRFLI6/'*t,77*CALL ALCEMYk',0 4j ICCM=1.,14/8* CMLKP=C.LKP+t" T ( IrOv+,q,1 IP,19* 45 CMFtICO•MJREL,)=FT(COt"'I. )&Os 45 COT ItUE4t'l* C "ELT iELFASf PARTIrULATF; COvPLFTE F-R PRESFNT TIME STii-,.,2*jr3*60 TO 5051 COIVTINuE STEAM ExAý4* C Ir :TEA" F.Yi'L"STOPI HAS t'oT 0CCIIPRFD,BYPASS.ot). IF(TT.LT.TEXI)GO TO bniBu6* 7I(I)=A•NIN(DT,TT-Trxi),3.1'uO 'j7 ICOMI.Nf.:o,6* h(ICOmICOL n.ý,,,* F ( ICo;C )=CSP( lCPl. I)3:='1"F(ICUVN+N) 0.~IFI [0 .CKEP, 1 AH•O. OI CSI- * O, 0.( IECO"-) =OKICSi892* HETfPS(ICO1) SKG1( ICM)+-LP(TCO")+rDp (ICPm)W'3. IFI(CoM.FO.1)HFTAPS,(ICui) BETAPS ICO'")+At'DAPS6ý'•*;3v5*,O 59 JCOM=IN5' (* ,.'IO ) H IO ,TO.)-G ( TCOO'sro ,)/v (I¢r")_•b* 57 ( !CW =HO.zH ICOM., C(f,')-BFTAPI (1?O ':3U7*CALL ALCEMY698* TF'r'(-1,L").NE.n)3oTO 504.6P 9* 5086 COt IMUE'uUt)* C ENr. OF AC 1(..C ýRITTF9ulo O0 .o ICm:lImI,3.j2* CSL`1lCSLKP+FT ( ICOll4'. 1)vII-220

9U3* 58 CSP(ICOM 2)=FT(ICOMlI9u4* C S7EA!- EXPLOSIOt' PARTICULATES COuPLFTE FOP PRESENT TI,4E STEP9u5* GO TO 50966* 1081 COiwTINUE VAPREL A9'7* C IF vAPORIZATION HAS NoT STARTED, BYPASS.9d6* IF(TT.LT.TVRIGO TO 509u9* GO 1065 JREL=I-10910 IFITT.LT.TVAtJRELI)GO 1O 1W6r911* TCI)=AMIN1(DT.TT-TVA(JREL))9124 DO 1067 ICOMZlN913* FICOMX=CAP(ICOMJREL.1)914* IFPtCOM.EG.I.AND.OKICA(JREL).NE.,.IF(ICOMI=oKICAIJREL)915* FCICOM÷N)=.9i6* H(ICOMPICO4)sO.917* BTAPVA(ICOM@JREL):SKGA(ICOM'JREL)4ELP(ICOM),FDP(ICOM)916* IE(ICoM.EO.1)BTAPVA(IcO!V.JREI.)=RTAPVp(TCr)MeJREL)+AiAKpIJJREL)919*9,0*DO 1069 ,JCOM=1,NIUb9 H(IcorICOM)=H(ICOM,ICrký).G(TCOJco..,)/VIICcM)9%L* 1067 if(ICOMICO•)=H(ICO",ICO•).BTAPVA(ICO--,JRFL)9.L2*91ý3*CALL ALCCMYUO 1068 ICOM=I.N9Zi4CVLKP=CVLKP+FT(ICOM+N,I)9,5* 1068 CAptICOM-JREL.P)=FT(IC0V,.I)9.e6* lb5 CONTINUEC VAPP;,I7ATI'N RLLFASE PAOTIrUiATES CnmPLETL F.)R 9Z7* PHA-E A IHIS TI"F STFP9dd* GO TO 509£9* lu7l COhTIjIJUE0.0* C IF. vAPORIZATION RELEASE rHASr B HAS r OT STA4TEr), UYPAS;VAPREL 89.51* IF(IT.LT.TVR2)GO TO 509,2* -ý.O 1075 JREL=1,1O9)3* IF(TT.LT.TVP(JPEL))GO TO 107'.944* T(l)ZANIN1l(T,TT-TV8RJREL))9g5* 0O 1077 TCOM=t.NQ396* F(ICoM):CBP(ICehmJPELI)9.17*93(3*IF(TCOM.EC.I.AND.OKICB(JPEL).NE.O.)F(ICOM)=oKICR(JRELJF (IC w,.+N) =n.9.9* H(ICOWICOM):U.940* NTAPVU(ICOM.jRFL)=SKGB(ICOM.JREL)÷ELD(ICOM)+FOP(ICOM)941* IFITCOM.,O.I)BAPVP(ICOMJREL)RPTAnVI(ICOMJREL)+AMDAP,{JRFL)142ý00 1079 JCOM=I,N9 4 3* it)79 H(ICiMICOM)=H(ICOA',ICO").G(TCOV.JCO")/V(ICoM)944* lu77 H(IOM.ICOM)=H(ICOM,ICOM).BTAPVA(ICO..,JPEL)9ý!5*CALL ALCFMY946* DO IC78 ICO,.=IN947* CVLKP=CVLKP+FT(ICOP+N,I)948* 1U7l CBP(ICOMJREL.2)=FT(iCOM,I)9"9* 1075 CONTINUE9ý0* C PHASE R VAPOfIZATIfN RLLFASE CO"PLET7D FOR THIS TIME STLP,951* 5U COt!TINUE9t2*9t3*GO TO 19060 CONTINUE ORGANICI954* C SET OFF EIAroIIAL MATRIX ELEMENTS FOR ORGANIC IODIDE HRArNCH9tb*DO 600 ICOM:I.N1%lb*H(XCGM÷N°ICO'I=ELOT(ICOV)9LP7 DO 600 JCOM=1.N943* Of0 HtJCOMIC0)=G(ICOM,JCOM)/V(ICOM)959* 0O A0 IREL=1,596$* GO TO(61,71,81,91o101),IPEL961* C FIVE wAY BRAa4CH 1S TO EXPLOSION# GAP, MELT, PHASE A vApo•lZATION,9b2* C AND PHASE R vAPOPI7ATION9630 61 CONTINUE EXPLODE9o4* C IF STEAM EXPLOSION HAS NOT OCCURRED, BYPASS965* IFCTT.LT.TEXI)GO TO 809Jb6* T(l:AMrNIIOT,TT-TWXI)967*908*nO 6? TCOM=I,NH(I¢OmIcOM}=O.9u9* F(IrCN)=CSOI(ICH.)l)970* FIICON+N)=n.971* IF(ICON;.EG.I.AI!D.OKTCS.tiJr.O.)F(ICOvi=OKICS91•$DO '•9 JCOM:I,N973* 9735 6 bg ý!(ICU 0 69Jt)=. ,IICO•t)=H(ICO'l.IrGM)-G(TCOMtJCO-.)/V(IC,)M)974* b7 HIIUM,ICOM)=H(ICOhTCUM)-ELOIICOV)q75* CALL ALCFMY97O* t'O (c ICOMh•:1,977* CSLKOI=CSLKor+FT(ICOf+NII)978* b8 CSOI(ICOK,Z)=FT(TCPO,I)9/9* C ORGANIC IODIDE EXPICSIOnI RELEASE COMn'LETED FOR THIS TImL STEP9o0* GO TO 80901* 71 CONTiNUE902* 7(11=UTGAP REL9403* DO 77 ICOM=I1N9b4* H(ICOpICOM)=o.9c5sF ICOp.)=CGOI(ICOx',I)9o6* F(ICOMN)=n.9o7* IFIICOM.Fo.1.APDO.V. E.I) F(I toM)=I.qOd*00 79 JCOM=I,N909' 74 •(ICO•.•.CO•}:H(1COYICQfA}-G(TCO",JCO•)/V(IC•)9-0* 77 H(ICOUICOM)=H(ICOk.TCU!J)-ELrI(ICOr.)991* CALL ALCFMY992* O0 7a3 ICOM=I1N9103* CGLKOI:CGLKOI+rT(IfIf÷+NI I994* 7g CGnI(ICOQ,2)ZFT(ICnW,1)995* C ORGA;:IC IODIDE RELEASE CjrNE FOR THIS TIME STEPVII-221

Y96* GO TO 809"7* 81 COI'TINUE MELT REL9*-e* C IF VELT RELEASE HAS NOT STARTED, RY-'ASS1910 IF("T.LT.T"1P)G0 TO AnlouO*[;o P.5 JRFL=Z.1nloul* IF(TT.LT.T,.•JREL))GO TO 85i U.2$T(1)=AMr!I(•l7.TT-TM4(CJEL))IOu3* 00 A7 ICOM=I,41,u4441lo)5*HIICOM. ICOM)=0.F {ICU"P)=Cm•OI({ICO"',JDFL, 1)lIOubF(ICOM+N)=0.loi 7 * IFF(ICOM.EQ.1.AND.OKICM(JREL).NE..I) r(ICOm)=OKICm{JREL)10)85* O A9 JCOM=I.T.)t09' 81 HCZ(4I4.)ICOMI=H(ICOP?, ICO1)-G(TCOUMJCO.)/V(ICoM)1010* 87 H(ICOM.ICOM)NH(ICOAICOM)-ELOI(ICOM)loll*CALL ALCEMYlui2* O0 va3 ICOMOI,N1013sCMLKOI=CMLKOI+FT(ICO,•÷NI)1014* 88 CMOT(ICOM.JREL-2)FT.(ICO1,1 )o0155 85 CONTIrjUElol6* C ORriIjIC IODILDE MELT RELFASE COMPLETE-l FOR T';IS TIME STFpl0•7. GO TO 80Injd* 91 CO0TINUE VAPREL A1019' C IF VAPOP rQELFASF "tAS NOT STARTED. BYPASS110-:1*,00 1F( TT .LT.TVR .)GO ,TO 14000 Qb JREL=I,ln104z* IF(rT.LT.TA4(JPEL))rmO TO 9541 ,?.3÷ T(t) AMItn (.T.TT-TA4(jREI) IIU•.* DO '7 IC(P=INIUZ5* 1;( IcoM, ICOM}=0.h'l..;o* F ( ICON);CAO! ( IC~t, PL I,lr~~7( IrCM+N)=0.IF ( ICOM.FO•1. .,V!D.OrC)(JkEL) .NE.O.)F ( ICO"V) =rKICA4(JREL)Iu,.9410.•0.(1;0 '49 JCOM=I.N94 11( ICL;•* ICO-V) =H ( ICOm, TCýý'] -G( TCO1', J(O%ý)/V( ICrM)1ý131*• 97 .,( le:!'.* ICh•} =H( IrOl', jCUw' )-EL1I ( ICOP)IU.)2*LALL ALCFMY13335 DO qb ICOM=INfiL)4sluý)5*CVLvIoGCVLKOTIFT(I4O÷+N i)9H CAoI(ICOwJRE[L,2)=FT(ICOV,1)Iojo* 95 C OT1" I t4UElo.)7' C ;1F(Gt;,IC IOCI_,F VAPOrP ILrbSE P 4 ,SE A COMPLFTED FO;( T14I,< TI'E STFP10.8b* GO O 501o.)9' 101 COqTINUE VAPREL 81n-.U* C IF OHASE [ý'APOR RFLLASF PAS NOT STA!-TEO, BYPASSlul,IF(TT.LT.Tv!)I)GO TO ý.jlU.$2* 1'0 1l5b JREL:j -10I(0,3* IF (T.LT.TRL+(JPEL))GO 10 1051U64* T 11)=AN44I4tl(nT,TT-T4 (IJRE.) II1,•5*00 Ii7 ICOP=lN104,* 0C If:O., ICOM)=0.IU7*F(ICOW)zCRO1hICON, JRELl)10.8* F( ICOM+N)=0.1,!49* IF(IuLO".EO.I.A?!D.OVICH(JrEL)°NE.O.)F(ICO!'l=oKICo(JREL)11),O*00 lug JsCOmIjNlful' i0'9 H(IC.ICOM)=,I( IO(ICOVICQý)-G(ICOM.JCO-.)/V(ICn!)10L)2 107 HI( ICOMICOM)=H(ICOM.ICOM)-ELOI(ICOM)lu4,3* CALL ALCEMY1()4* .;00 108 ICOm:j.Nli)5.* CCVLkOI CVLK0I+FT(IC0O+NI )1O:zb# 108 ,UCI(ICOM,JREL,2)rFT{ICOP-.1Ilot,7* 105 COjT IN4UEl)bbs C ORGAh.IC IODIDE vAPro RELEFASE PHASE B COMPLETED FOR ThI, TIME STEPsl1jt,9* 80 CONTINUEluU$* GO TO 1901OuI' 120 CONTINUE 12BRANCH10."' C iýATUPAL CONVECTION LAMBDA WILL Br ASSUMED ZERO I; ANY CoRPArIj.lou 3 * C MENT WHFRE CI2/V.LT.CUTOV/VTOT.104o* uO 1037 I=lij10c,5*KSpPPAY=(Iub* IF(T.EQ.1•A',. ( (TT.4rT.TCSRI•ANO.TT.L?.TCSRtE).OR.(TT.Gi.TCSR2.A."011iu 7 * *.TT.LT.TCS•?L).OP. TT.GT.TCSI.ANO.TT.LT.TCSIF)))KSPRAY=ILku,*KC KPC( I=FLfl4Tt(I-KSPRAY)1Ou9*OKr.ýCS( I )1FLLAT(1-KSPHAY)lf70* IF(CSI2(I, II/V(I).LT.CUTnV/VTOT)OKNCS (I)= 0.I071* IF(CGI2(TI1/VII).LT.cUTOV/VTOT)OKpCG(1)=n.1012* 4ý0 1037 J=1,101U'/3SCALE=•.07521.FYP(-.I.FLOAT(J-1)I1 /4* 0K.'CA(I.J)=FLOAT(I-KSpHAY)1075" OK1t:rCL(I.J)=FLOAT(i-KSphAY)Iulb* OK . (I..J) =FLOAT (1-KSPRAY 1j1'77*IF CAI2( I *d, 1)/V(I1/SCAUT.LT.CUlGV/V"OT)OKNCA(IJ)-0.1f'178* IFCCLI2(IJI)/V(I)/ScALF.LT.CUTOV/V.OT)OKNCB(IJ)=O.11/9- jF(C'CIIP(T.J. .)/V(I)/.i.LT.CUIOV!VT"TIOKt"(I•j)=O.10•ou 1.37 COr:TI,4JEb0ol* C .SET CFF DIAG;'NAL ;'AT;ZIX FLELUETS FoR 12 e.RA:JCHlr,2*10o5$JDO 10 I20 IC:INH{ ICOv'+., ICO"-)=EL I:•ICO11)1OV4W1'), 5*'C',0 1200 JCO%.=I.N1,.; O'1 ý ( j C 0. , I ('ý0"•) = (i ( I C 0 , JC•• I l -F E 1 I ,-Orj , JC o. })I/ V ( I CO 0,lO 170 IPEL=I.•1u.,7*O TO(121,131.141,'151,161) ,IPELIL,4* C FIVF ,.AY PPA]'CH IS Tn GAP RELEACE,STrAU FXPLOSION-MELT ,t:LEASE,VII-222

1069* C VAP3;OIZATION PHASE A. ANF VAPORTZATIIIN PHASE B.1090* 121 CONTINUE GAP REL1091* T(I)=DT1092* O0 127 ICOM=1.,1093* H(IcOMICOM)=o.1U94* F(ICUM)=CGIP(ICOM'1)1045. F(ICOM+N):O.109"* IF0M.EQ.I) F(ICOM):nKRON(ICO-!,I)'1097* vTA•G (ICOM)=ELI!(]CUM)+FDIPIICI)'+OvNrGtIC•,) *CKG( .* CnF10J99**)4Ar(ICOM))/VITCom)1o1* IF{ICOM.EG.1)BTAI2G(ICOM)=BTAI2G(ICO"1)48",AIG+ELAMGIlol*00 129 JCOM=1,NIIU2* 129 H(ICOM.ICOMI=H(ICOM.ICOW.)-G(1COMJCOm)/V(ICoM)2flu3* 127 H(IC6W,ICOM)=H(ICOM.ICOJM)-BTAI 8(ICO.)11u4* CALL ALCEMY -I105. UO 128 ICOM=1,N1166* CGLKI2=CGLKI2+FT(ICOM+N,1)lI0o* 128 CGI2(ICOM ,2)=FT(IC ,o1.j1108* C 12 GAP RELEASE COMPLETLO FOR THIS TIME STEP1169* GO TO 170ST EXPLO1110* 131 CONTINUE1111* C IF STEAM EXPLOSION HAS NCT OCCURREO.OYPASS1112* IF(TT.LT.TEX1)GO TO 170II13oT(V)=AMINI(DT#TT-TEX1)l1,4*00 137 ICOM=IP'1115* H(ICOM.TCOM)=n.1116* F(ICOM)=CSI2(ICOfI)1117* F(ICOC+N)fl.1lia*,1F(OKICS.GT.A:.)F(ICO:2)=DoRON(ICOM,I)1119* FITA'2S(ICOM)=ELI2{ICOM)+FD12(ICOm)1140* IF(OKNCS(ICOr).NE.O.)BTAy2S(1COM)=PTAIPS(ICo*)+CK3(ico:•)11a1* 1*(AC(ICOr)+AF(TCOM))/v(ICOM)11,2* IF(TCOM.FO.I)HTA12S(ICO•):BTAI25(ICO..)+AvrAIS÷ELAMS113* LUO 139 JCOV=1.NllI.4* 139 H(ICo0,ICOM)=H(ICO".ICO•')-G(TCOmJCO,..)/V(UCcM)lIm* 137 ,(ICCM-ICOM)=H(IC-OICuC)uBTAI2SI ICO,)14t)* CALL ALCFMY1147* DO 138 ICOM=I H11zo*Iie9"CSLKI2:CcLKI?+FT(Iro.i+N,])138i CSIý(ICOV-,2)=FT(ICnM,I)11.U* C 12 STEAM ExPLOrION COMPLETED FOR ThI's TIP'F STEP111.* GO TO 17n1132* 141 COP:TINLEMELT REL1133* C IF MELT PELEASF HAI tj(oT STARTEO.RYPArS114* JIF(TT.LT.TMF1GM TO 17011,5* DO 145 JPELO1.1011.)o*11:57oIF(TT.LT.TU4(JPEL))GO 10 145Til)=Amlt:I{rT.TT-TM4(JkE1.))11.38*11•9q*flU 147 ICOW=1.Nf4QCOpIcOv)=O.11ý0" F (I co p.) =r mI ? (I COf',JP EL 1I1141* F(ICOM+N)=P.11142* IF(Or,ICM(JPEL).GT.0.) F(ICO4 )= DKROr(ICOM.0)*OKICM(JRHL)1143* PTAI1-( ICOM.JRFL)=rLI (ICOM)+FO!2( TCýM +pKrM(ICo,4.jEL)*11,4* 1(CKG(ICOM)*(Aw(ICOM)+AF(ICO2))/V(ICO,:))1145* IFtTCOM.FO.I)BTAT2-(ICuM.JREI)=RTA12.(ICnIA,,REL)*AMDAI.+(JREL)+l14b* *ELA'•A(JRfL)11.7* UO 149 JCOM=1.N1194* 149 HZCJ,%.-ICOV)=H(ICO"IICOM)-G(TCO•,JCO!,)/V(ICOV)1.I9* 147 H(IC.r4,ICOM)VH(ICOM.ICOM).BTAI2M(ICO.,,JRFL)11•0*CALL ALCFMY11:4* 00 148 ICOM=,1Ml152* CMLvI2=CLKI2+FT(ICOM+N,11l13* 148 CMIPilCO':lJRFL.22=rT(ICOP,,1)145 CONTINUE12 'LLT PELFASE COMPLETED FOr THIS TIME CTEP1134*11•5* C116* 60 TO 170I1•7* 151 CONTItUEIID"* C IF 'APOR RELEASE ttAS NOT STAPTEPRYP•SSVAP REL11t9* 1F(TT.LT.TVR1)GO TO 170lAI*1Ioi*DO 155 JREL=1.10IF(TT.LT.TAU(JPEL))GO TO 155llio2*T(I)=AMINI(DT,TT-TA4(JREL))11.3* DO 157 !(OV1IN11o5* F(IC0M)=CAI2(ICO",JRELI),IUD$FQICOm+N)=0.11o7* IF(CICA(JRELj.GT.n.)F(TCOM)=DKRONJ(IrCOm,)*rKICA(JREL)flu8*I109*F.TAT2ACICOJRVL)=FL1I2(ICOM)+FDO!(IC,:M)+rKNCA(ICCA,,JI(CKG•(ICOM)*(AW(ICO-)+AF(ICOM))/VCIIO-,)I4 EL)*1110* IF(tICOM.Fr.1)0TAI2A(ICOMJRE[.):RTAT21,(TCnr..JREL)+AMaAý(CJHEL)11/1* I+ELAIA(JPEL)1172* no 159 JCOM=I,N1113s lb9 (I.rl,.ICOM)zH(ICO0.ICUO•;-G(CCOvJCO..)/V(ICoM)1174* 157 C;(I• ,ICOW)=HhICOM,ICOM)-BTAI2A(ICOo•,JRFL)1175. CALL ALCEMY1176* 00 158 ICO"=IN1177* CVLK 12=CVLKTI+FT(ICOM+pJr1)1118A* lE CA!? ( ICO'1oJf'EL2)=FT( ILO". 1)1179* 155 COý:TINUE11;0* C 12 vAPORIZATIOr r'ELFASE PjiASF A COMPI ETED FcR THIS TIMF STEPllbl* GO T0 17n11o2* 161 CONTINUEVAPREL uVII-223

11,3* C IF 01hAE n v;POIý RFIEASE HAS NOT STATEn, RYPASS11o4* IF{TT.LT.TVP2)GO TO 170lleb*CO 165 JREL=1,1011b'6 IF(TT.LT.TP4(JPEL))GO TO 16511o7* T(1)=AmINl('lT,TT-Tn4(JKEL))1168* 00 167 Ir.OP=l.N11"9*1190"*H(F1C0vCO")=rl.F(If:O)=C I?( ICO•,JPELI1llvl* F( ICOm*14) =0,1192 IFjnAJC9(jqEj I.GT.n.)F(TCOM)=DKgtoN{•lOm, Il*oKLC9(,IREL)19j*TrýAiZbI COM,-.FL1)=FLI2(Cn+M)+FO!?(1CoM)+OKN(1I(ICOM.,JEI.)*1194* *CKG(ICOM)*(AW(ICOM)IAF(ICOM))/V(IcOM)llIj5* IF(1!G.4.FO.IIiTAI2B{ICO,*JREI.)=RTAI2-(ICOY,JPEL)+11b* *AMDAIH(JPEL)+ELAl'B(JREL)1197* 10 169 JCOM=I.N111Ib 1b69 H(ICONI,ICOM)H(ICOM.ICOM)-G(TCOM,PJCO,)/V(ICoMII199* 167 H(ICu•.ICOM)=H(ICOM.ICOM)-BTAI25(ICO•,,JREL)12uO*12u1"CALL ALCEMY0O !.t,8 ICOW=I,ý!12u2* CVLKI2=CVLKII+FT(IoOM+N,1)1263* 186 CBI!(ICONIJREL, 2)=FT(ILOr',l112U4* 1b5 CONTINUE1205* C 12 vAPORIZATION RELEASE PHASr B COMPL.ETED FOR THIS TIMF STEP12U60 170 CONTINUE1207* 190 CONTINUE12ud6 IF(AbS(TT-TPUFF).GT.3.L-Fl]GO TO 20001269* C MULTIPLY ALL AIRBORNE FRACTIONS EXCEPT THOSE FROM STEA,• EXPLOSIONS HY1210* C XPUFF. TNCREASF LEAvED FRACTION FOR THESE BY (1.-XPUFF)I*AIRBORNE1211* C FRACTION PRESENT PRIOR TO PUFF-1212' XPC=t.-XPUFF1213* XPCSZXPC/XPIJFF1214* 0o 21001 ICO.,1=1,W1215* IFiTCOH.F0.I) rO Tn P1001121o* CGLKPZCGLKP+xPC*CG6(ICUV:,2)/nFPP1217' CGp(IC0M,2):IPtIFF*CP(IC•,2)1216* CSLKPrCSLKP+PrS*C'CP(ICO'.,2)/flFPP12i9* CGLKI=CGLKOI+XPC*CGCI(ICOMP)/nFPnI12.0* L(.OI(ICOv.2):xPUFF*CGoI(ICOM,2)12,1. CSLKIT=CSL.I+XpCS*C;oI(ICO",2)/rFpO,1@2,:CGLKI12CGLKI;I+XPC*CCI2(ICOM,2)/rFPI212:3* CGI2(ICO'I,2)=XPUFF*CGI2(ICOO.21IZ"* CSLKIZ=CSLKI+XPCS*CSI2IICOM,2)/nFPI?12•6* 00 '1000 JPIL=IIO122b* CML'F=CIPLKP÷xPC.*CM0P(ICOn'.JREI . 2 )/oFP-,12g7* CMP(.CO•,JRFL, )=Xr,)FF*C;•Pf COMJRFL,2)12,60 CVLKP=CVLKP+XPC*(CAP(ICOtN.JIFL2 )+CBfl(ICOM,,REL,Ž2))/ IFpp12ý9* CAp(ICOM,JREL,2)=XPUFF*CýP(ICOMIJRFL,2I12.30* C6P(ICOMJREL, I:XPHFF*CPP(ICOM-JREL,2)12.ýI*CMLwOI=CFLKOI+YPC*rMOI(ICoMjREL.2)/•FPOI12.2* C1OIICOt. JRF'L.21=YPFF*CsOT(1C0M.JR'L.2112.3* CVLKOI:CVLKOI+YPC*CCAoI(TCOM,JPELP)+CgOT(ICOMJRL,2))/L)FP6 I12.,4* CAOTIICO.Jý,LL.2)=XPUFF*CAOI(ICOM,JRrL,2)125.ýSCBOI(ICO',JRFL.2)=YPUFF*CF0OI(ICM.JRFL,211236' (:MLKI2:C'LKTý+YPC*CMI2(I(OMJREL.,2)/nFPI?1237' CMI;( ICnfYJPEL,2)=yP!JFF*CMI2(ICn.JRýL.2)1230* CV/KI2=CVLK•+Ypc*(CAlk(TCOvJRFL,?)+CRI•(ICOMJR•L,,}•/DFP1212.9* CAI'IlCO"'JPEL,2)IXPuFFCAI12(ICOM.JRrL.2)12.0* CiI?(ICo.*:JrEL.2)XP()FF*¢CRI2(ICeM.JR,#L.21241* 21000 CONTINUE1242' 21001 CONTINUE1243* 20000 CONTINUE12.4* C IOTAL VIPOPI1ATTO•' ArId "FLT RELEASE COMPONFNTS PER Cu.PARTMEM-1245* O t150 I:1=N124o* SUVA:u.1247* SUMn:O.1248' SUWC=O.1249* SUwrO.1250" ýSU'F=0.1251' SUP:H=O.1262* UO 44 J:1.1012ý3'SUMA=SUMA+CMp(I.J,2)12540 SUMn=SUMR+CKI2(I,J,2)1265* SUMC=SU%4C÷CMOI(I.J.2I12tj6* SUPn=SUMP+CAI?( IJ,2)÷CP'( IJ-2I12•75 SUPr=SUME4CAOI(I.J,2lC~riI(I.J,2)12tbd*SUPIý=SUMH÷iCAp(T#J,2)÷CbP(IJ,2)1259' 44 CONTINUE1260* CMTP(I):SUMA1201* CMTri(I):SU!:C12Ž3* CVTlll (I )SUrl12u4* CVTt2(1):SII•012.b*CVTPII):SUMP12.b* 4bO CONTINUE1267' CTMnF=O.12,A*CTsnPzO.12(6* CT,.11Žzfl.12i1'CTS1I2=O.1272* CT0-I1Z:.1273* CTVql2=0.12i4* CTY{wI0.1275' CTSRoI'=0.VII-224

12764 CTe•OI=O .1277* CTv'ZO1ZO.127d* CTvRP.=o12-9 " ,O0 46 I=lr:121U'CTV'rP=CT'#PP+C•TP(I)12.1* CTS*ýfT=CTSRP+÷SP(T.;)1 2 2c * CIcrrP=CTDRP*CGP(II2)1203* CTS7OI=CTSP0IfCSOI(I2)12.4* CTGO'I:CTGPOI*CGOIIT.2)1235*120u*CT.,QOI=CTMPOI+CMTOTfI)CTVROI=CTVRnl+CVTOI(1)1237* CTvQI?=CTVRI?+CVTI?(1)12eon* cTSil2CTSTRIý+CSI2(I,2)lýý0* CTGfI2=CTGRI,?+CG12I 12)1290* CTvRI2=CTMRI2+CMTTP(1)1291* CTVPP=CTvRP+CVTP(I)1292* 46 COPITINUE1293s C AIR'loHNE FRA:1!O):S (TOTA!.) CnPIPITE', %0OVF1 2 '4* C CO•'PUTF APO0'.TS LEAKED12ý5* C CO'VrUTE DOSE RFOLOCTION FACTOPS12-)b$URFYI2=O.1297* DRFSI=O.1298* ORF'P=O.12993* ORFV12=0.13UOODRFGI2=O.13uI*URFGP=O.13;02* oRFSP=O.13u3$ IF(CrLKP.GT.o,. rRFmP=CMLKOI/CMLKFP13UL*IFfCSLKP.GT.n.) ORFSP=CSLKOI/CSLKP1305v13oh*IF(CCLKP.GT.n.) ORrGP=CGLKOI/CGLVP.IF(C-ALKI2.GT.0.) DPFMI2=CMLKnl/CPLKII13U7* IF(CSLKI2.GT.O.) DRFSI2=CSLKOI/CSLKI213U8'IF(CGLKT2.GT.O.1 OrFGr2:CGLKVI/CGLKI213g* IF(CVLKI2.GT.D.) ODFVI2=CVLKOI/CVLKI?1310* IF(CVLKP.GT.O.)iP•-"V-P=cVLVOI/CVLKP13i1* .=MI1312* C IF YOL. NANT TO BYPASS ORITINr. PUT A GO TO nq8 HERE.13i3* wRITE(6,300O)TT"1314* 3000 FOR 'AT(I I,/,SXYHTI)E.,IPEII.5,3X.3-1HRS)1315* 0R1TE(6,AOO1)1310* wRITE(6e,3O02)1317* 3101 FORMAT(I1HO,5X.4OCHC(.PAHTwENT AIRROPNr FRACTIONS CONTAILO ./1131j* 3002 FORY!AT(1X,IHC,11H GAP .tlN rAP 111H GAP1319* 111,4 VELT .11H MELT 114H MELT IIH EXPLULON13iu$ 211H LxPL0SIt; .111H EXpLOSION .114H VAPOP .0OH VApOk ,10H VA13i1" 3FOp P/,I1,IHI RELEASE '11H RELEASE .11H RELEASE I13,2* 411H RELFASE IIH RELEASE ,ll1 R.LFASE .11H RELEýý,13.3* 5i1i ;(ELEASE .11H 1 FLEASE ,t111 RELEASF ,OP1 RELEA.,L13,4* 610m RELEASE ,/,IXH41114H . RIICLvS 114H 121345* 711H oI13ý6* 711H PARTICLES ,I1n I; .111H of *11H PARTICLES1327* 811H 12 IIH O0 .I1H PA-TICLES 10OH 1213d* 910H 01 ./)l•JO* CMTnI(I),CSPtl,2)'rSI2(1,P),rSOI(I,2i,CVTP(T)-CVT12(i).LVTI01 I))I13,i1.2I=1,N•)1,.* 3003 FOR'AT(12,1PlAl).UIPlOIl.4,IPEIO.u)1333* wRITL(6,004)13,14* 3004 FOF•AT(IHn,//.rX,3'tHTO1AL AIRBORNE FrACTIONS CONTAINED #/)13j5* .RITE(6.3007113*,s •RITL(6,'nO6)CTGRPCrnKI,'CTGROTCYMP,CTIIP?.CTM!-ROICilxPI13-7* ICTSIl12,CTSROI.CTVRPCTVPI2.CTVROI13.38* WRITL(6.3005)13.19* 3605 FORXAT(Iwn,/,95),32HFSCAPE FRACTICN, rF EACH RELEASE ,/)1340* .RITE(6,3007)13414 .vRITE(6,3006)COLKP-CrLKIPCGtKOT-CML•LCMLKI2,CMLK01,C;LKPo1342* 1CSLwI2,CSLKOIrVLKPCVLKT2,CVLKmT1343* 3lOb FOPVAT(2XelPEIP.4.IPIOEIt4,.IPEIO, 4 )1,44* 3007 FORt.!ATIIXIH .11h GAP 1114 rAP .111 GAP1345* I11w MELT 114H MELT .11p, MELT IIH EXPLO'SION1346* 211H ExPLOSIO: .114H ExpLOSION 11H vjAPOR 10OH VAPOR1347* 310w VAPOR ./,IXeIH .IH PELEASE .11w RELEASE1348* 411H RELEASE I1349* 4111H RELEASE 111H RELEASE .11H RrLEASE 114H RELE.SE ,13ýo* 511H RELEASE .11H RELEASE 011H RELEASE .0OH RELEAc .- OH REL135l* 6EASF ,/,1XlH ,111i PARTICLE' '1IH 12 .11H 011332* 711H PARTICLES ,11H 12 .111H O .11H PARTICLES1353* 811H 12 .111 01 111H PATICLES .10H 1213,4. 9,10,4 0 ./)1355* wRITE(6,3010)13!eb* 3010 FORMATflHO./.SY.3 8 4Ht 0sL REDUCTIMN FACTORS OF EACH RELEAS ./)13•7*wRITE(6.3007)1358* 4RITE(6,3015)oRFGPPORFGI2,ORFMP.ORFM12tDRFSp.DRFS12,DRFVPORFV1213ý)9* 3019 FOR'AAT(2X,lPF~lq.4,,PE11.q,I~ly,lO2Ell.4,11•°•P2Ell.4,Ilx,131(0* 11PF11...,IPE1'.. )13ui* C AD ;KýC qPITF Tr SEF IF CEFFFTCIFrTq ARE PFTý:S CALCULAII.U CfH14ECTLY.13v2* C PUT NEW VALIIES lI'T' NO. I 61ICKETS.13o3* UO Q780 J=lZi 41314* CGP(J.I)=CGP(J,2)1-•5" CGI?(j,)=C6I2(J'2)13.6* CGOTlJ,])=CRoI(J.2)13.7* IF(TT.LT.TFY)2 O Tn 177513.8s* CSP{J,l]=CS'(J.2)VII-225

13.9" CSI'(J,1)=CS12(dJ2)13/* CSo.-(J, I I C soI (J,;>)1311* 9775 CONT rIUE1372* 1O n7-0 K=1.ld1.3/3. [FiTI.LT.T'4.(K))6O TO 977R13/14*1375*CMP(JK,I)=C:-P(J,K.2).Ii,,)•2JK2137tb* CM.j(JK.1)=CMOI(JK,2) 91-177 9779 IF(TT.LT.TA4(K))6O TO 7bO1373* CAP(J,K,1)=CAP(J-KP)15/9* CAj?(J-K,1)=CAT2(J.K.2)I ,,ý1) * CAlýTtJK,l)=CAoI(J-K,2)13i1. 1F(TT.LT.TH41K))GO TO 97fn1.312* CBr(J,K-1)=Cý,P(JK-2)cljZI(JK.I)=C3T2(J'K-211-5.4* C83T(J,KI)=CHOI(JlKl2)13o5* 9763 CO'T IJUE1:n* C fEFjif 3 TYP.S OF rA:Torn:;LIý.E mEPoS7TTO-i Ar,:D TPA,4SPLJHT TYPES'137$ C t:R6A,.IC 1O0,1E LIKE, IOjT.E LIKE. LkNr PAPTICULATE LLE13o8*ljo9*00 6b1l I:1= '.OT6b0l CORL-EKII)=CG3LKAIl*CFR(JI'I+CMtKOI*C;:R(2'I)÷CsLKO'I*CFR(3,1)*XPUFF13. *13913t1+CVLLKOI*CFR(*-1)!jOjpj=NOI+÷11312s,I;2r;JD=NOI+".'113,3* iOo k&02 l=Nl0IP1 rJeF•I2n1394* boO 2 CONLIK(I)=CGLKI2*CFR(lI()+CMIKI2*CFR(2f*)+CSLKI2*CFR(3,1)*XPUFF115'1* ICVLKI2*CFPt4,I)1396* liSTAkT=N12EFr7+l1397* t. ISl=tOI4t•JINPAR1,'t O 253 I=NSTAPT.NTSO11511 6tU3 COLEK(I(=CGLKr*CFP(1.1),CMLKP*CFR(2,1)+CSLKP*CFR(3.1)*APUFF14U,*I+CVLKP*CFR(4,I)1.j1*jRITL(6.3032)(NAN.E6(I),1=1.N1SO)14o2* 3j32 FORktAT(lH0AX.74HFPACTlONS OF COPE I"VENTORy LEAKED,//.'OXUlO(A6,bX14U3*14.4*1))NRITE(b-.'03'1$)(CoP•LFK(I),TI~leHISM) 14Ods GýO TO .11405* 3034 FORMAT(IH ,5X0l041PE1o.4,WX)) * 1 4u9* 999 CO'.TIJNLE14ub* lF(rT.LT.TENC-1.-8)GO 70 6 1410" STOP140T* CALL ERR(3,ti.HCORRAL S.N. 30.4 1 1411* (NOJ7. COMPUTER CODE CORRAL LISTING FOR PWR ACCIDENT SEQUENCES1* DIMENWION PI(9.20)eTI(20),GI(.9g,20),VAPI (920),VC9),2* IVAP(9),P(9),TMP(9),DELrTTt9,2l),DELTAT(9),ELKP(9t20),El2(9,9,2D),3* IEP(9,9,20),EFI2(9,9),HT(9),ELI2(9),ELP(9),G(9,9).CMP(9,10,2),4* 1ELK12(9,20),LKG(9),OKM(IO)OKA(IO).OKBIIO),D* ICA 1( . 10,2),6.7.ICdOI(9,10,2),1C412(Y.10,2),8* 1CAI2(9,.1.2),9* IEFP(9,9)10* UIMENSION CGP(9,2).114 lCSP(9.2).I-,* IC612(9,2),15* 1C.412(9i10.2).14* ICSI2(9,2).15* ICGOI(9,2).lb* ICAOI(9,10.2),17*id*1CSOI(9,8),IC%;TP(9),19* 1C.1TIk(9),o* JICm;TOI(9),21. ITAI(1f).I T,;3(10).23.24*1T.,A4(1O)f1OKIM•ID)46- VIAENSION [ETAPG(9),2b* 1BETAPH(9,10)t27* IJETAPS(9),2d* 1AW(91,AF(9),29*30*ISKGG(9),1'ýKG{9,101,31* 1FOP(9),3.e* 1FUI2'(9),33. IBTA12S(9).34* IbTAI2M(9g0),3b* b1TAI2G(9).3b* lCVuII(9).CVT12(9),57. IOKNCM(9.IO),38* IOKNCG(9).359* IUKNCS(9)404 DIMENSIOtN41* IOKICA(IO),4i* IOKICB(IO),.3* lUKNCA(9.10),44* IBTAI2A(9,10),15* IBTA123(9,10),4b* 1AMOAMP(11),47* IAMUAIM(IO),4b- IA;ADAIA(19),VII-226

SO* IELOI(9).ELKOI(9,20).51* IOKNCB(9.10),ae* ICATOI(9)53* DIMENSION CHTOI(9).54*55*1TAI(1O),IT13 (in) p56* ITA3(10).57* 1TB3(10),50* 1TU4(10),59* *TA4(10)60* 01MENSIONo1* IDIIST(9)02* IVAIR(9):63* IVST(9gp64* IPAS(9),65* IPSA(9)oGo.*lV(9),61* lWA(9),od .s()69* IDS(9)170* lDA(9),71* 1012(9)p72* 1SC(9)73. DIMENSION AMDAPA(10).AMDAPB(10).BTAPVA(9t10)PBTAPVB(9,l0),CAP49,74* 110,2) PCP(9,IO12),CATP(9) CBTP(9),CVTP(9),DVA(IOI)DVI(10)pEFA(IO)7a* 2,EFB(1O),SKGA(9,1QIPSKGB(9glO).TVA(IOITVO'tO)70* DIMENSION GR(9),77. ISKG1(9)v78* ECOM(1O)p79* 1COB(IO).aO*1COA(IO),81* 1 ELAMA(IO)p82* IELAMA(IO)p63* IELAMB8iO),84* tEFM(10),85* IDM(9),do* ITMY(9,20),TCHANG(20)67* C CFR(JREL.ISOT)=CORE FRACTION RELEASED FOR MATERIAL ISOT BYa*. C METHOD JREL.89. DIMENSION CFR(4,10),NAMES(10),CORLEK(IO)90* EQUIVALENCE(TA4(1),TVA(I)) (TB4(1)HTVB(1))91* C THE ISOTOPE GROUPS WILL BE TAKEN AS KR-XE AND 0I (ORGANIC IODIDE92* C LIKE).I2-HR(12 LIKE),AND CS-RB.TEBA-SRoRUPU(PARTICULATE LIKE)95* DATA 4AMES(I)/6HKR-XE /,NAMES12)/6HOI /,NAMES(3)/bHI2-BR /,94* lNAMES(4)/6HCS-RB /PNAMES(5)/6HTE /.NAMES(b)/6HBA-SR I,9S* 2NAMES(7)/6HRU /,NAMES(8I/6HLA /9o* DATA NOI/2/,H12/1/,NPAR/5/97* LOGICAL SWITCH98* COMMUN//A(SC)eASMAX.ARSRMSR(50).F(50) FABZ(50,5O0)FT(50,50)99. liI(5JSDU) ,II.INOINFLLL,LOST,MAPS,MARKNTT,NA,Nr*,NEXTNT100* 2,PLOT(5.2,lOO),Q(50,SO),RELMAXRELRMS.RtJN(1O),T(50).TAG(5OITYPE101* 3,'JNITF,UNITTw(5O),Z(5Oi50),ZN(50,50).ZNII(5O)102* NAMELIST/JREAK/N#NDATATIDELTTI.DTtDT72TJUMPlTJUMP2,103* 1PI,GI,EP,EI2,AWAF.ELKI2,ELKP,HT.TCHANGTENO,TVR1104* 2,TVR2,TVRE,GK,DPEDPL,VAPI,V.FDP,FDI2,TCSI,ELKOITMRTMFTCSR2E,105. 3CSR2.TCSR2,HCSITD.TEXleTCSIE,CSIiO,TCSRIE,CSRI,loo* 4DCSRHCSR.CUTOVDCSI.TECCIVCL,TMY.TCSRIECCIDF107* 5,TPUFF,OFPP,DFPOI.OFPI2,CFR.XPUFF10* NAMELIST/TRIVIA/DI2,GR.SCCKGPTMPDNSTVMUTAI2G.OKNCGELI2109* ;NAME-IST/AIX/PTMPDNSTtVAIR,VSTPASPSA,VM110* DATA ABSMAX/1.E-8 /.ABSkMS/1.E-8 /,RELMAX/I.E-8 /,RELRMS/I.E-8/111* I CONTINUE CONTINUE TO NEXT CASE112* REAO(0,8REAK.END=999) READ INPUT DATA113* WRITE(b,2001)N114* 2001 FORMAT(1HI,///e32X,35HFISSION PRODUCT RELEASE AND CLEANUPP//v11* 15X#20GllATA AND ASSUMPTIONS,/,bXP22HNO. OF COMPARTMENTS N= .12)I1o*WRITE(6,20021NNDATA117* 2002 FORMAT(1HO36X,21HCOMPARTMENT PRESSURES,/t6X,l1a* 114H((PI(IJ),I=I,,12,bH),J=I,,I2,2H)lJ119* DO 1500 J=I,NDATA120* WRITE(6.2O03)(PI(I#J)rI=lN)121* IaOU CONTINUE122* 2003 FORMAT(6X.IP8L4.6)123. WRITE162004)N.NDATA124* 2004 FORMAT(tH0,6X,24HCOMPARTMENT TEMPERATURES,/,bX,15H((TMY(IJ) It=,,120* DO 1505 J=1,NUATA127. WRITE(6,23O3T(TMY(IJ),I=IN)128* 1595 CONTINUE129* WRITE(b,2006TN,NDATA130* 2006 FORMAT(iHO,6X,35HTEMPERATURE DIFFERENCES TBULK-TWALLp/t6X#131s 1168-I((DELTTI(I,J),I:llI2,6H).J=1,,I2,2H)=)132* DO 1510 J=INDATA113* WRITE(G.2003)(DELTTI(IJ),I=:1N)134* 1510 CONTINUE135* WRITE(bt2OO8TNPNDATA136* 2008 FORMATI1HO,6X,26HWATER VAPOR MOLE FRACTIONS P/,6X,'137* 116H((VAPI(I,J).IzI,,12,6H),J=lI2,2H)=Ii1,5*JO 1520 J=I,NDATA1139* WRITE(6,2003)(VAPI(IJ),PIIN)14i2 1520 CONTINUE141* WRITE(6,201C)N,NDATA142* 2CI0 FORMAT(IHUNX,3OHFRACTION OF 12 LEAKED PER HOUR,/.6X,VII-227

143- 117H((ELK12(I,J)#I=I,,12,bH),J=le,I2,2H)=)144* 00 1525 J=I,NUATA145# WRITEI6,2003I(ELKI2(I.J),I=I.N)146* I52b CONTINUE147* WR ITE(6,2U12)N,NDATA148*149#2012 FORMAT(1HOt6X,4OHFRACTION OF PARTICULATES LEAKED PER HOURllbH((ELKP(I,J),I=lp,I2,bH).J=I,ol2o2H)=)*/,6X,150*151l*O0 1530 J=I.NUATAWRITL(6,20O3J(ELKP(I,J)vI=l.N)lb2* 1530 CONTINUE153* WRITE(6,2016) N,NUATA154s 2016 FORMAT(IHO,6X,62HFRACTION OF ORGANIC IODIDES LEAKED PER HOUR FROM155# 1COMPARIMENT I ./,6X,17H((ELKOI(IJ)jIl=lI2,bH),J=I,,I2,2H)Z)156# DO 11530 J=INDATA157. 11530 WRITE(6,20O3)(ELKOI(I,J).I=IN)156* tIRITL(b,2014)rO)ATA.(TI(J),J=.,NDATA)159* 2014 FORMAT(IH ,6X,I1OHPRECEDING DOUBLY INDEXED EXPRESSIONS WERE INDEXE160*161*1D ON COMPARTMENT (1) AND TIME (J). THE INPUT TIME ARRAY2,/,6X,11N(TI(J),J=I,,I2,2H)=.I,(6XPlPBEt4.6))IS TI.162* WRITE(6,22015)TCHANG163* 22015 FORMAT(IHOp6X,72HTIMES TCHANG SIGNAL REINTERPOLATION OF FLOW RATESlo4* * AND THERMODYNAMIC DATA ,/p6X.7HTCHAN6=,/P(6X,1PBE14.6))105* WRITE(6.2018)N16*b 2016 FORMAT(1HO,6X.77HFRACTION OF PARTICULATES REMOVED PER HOUR BY INTElol7Ius*IRIOR FILTER IN COMPARTMENT IWRITEtb.2OO3)(FDP(I),I=I#Ni/,I6Xv12H(FDP(I),I=.,eI2p2H)=)169* WkITE(6,2020)N170* 2C20 FORMAT(IHO,6X,76HFRACTION OF IODINE (12) REMOVED PER HOUR BY INTER171* IIOR FILTER IN COMPARTMENT I /,/bX,13H(FDr12(l)I=l.leI2o2H)=)172* WAITEIb,2003)1 FDI2(I),I=lN)173* WHITE(6,2021)174* 2021 FORMAT(1HO,/,bX,41HINTER-COMPARTMENTAL TRANSFER COEFFICIENTS ./)175* WRITE(6.2022)176* 2022 FORMAT(1HO,6X,84HCUBIC FEET OF FLOW PER HOUR FROM COMPARTMENT I TO177. 1 COMPARTMENT K AT DATA TIME ENTRY J178* D0 1540 J=1,NOATA179* DO 1540 I=IN180* WRITE(b.2024)IoJ#N181* WRITE(b,2OO3)(GI(IKJ),K=IN)182* 1540 CONTINUE163* 2024 FORMAT(IH ,6X 9 4H(GI(,I2.3HuKu,I2,bH),K=l1,I2e2H)=)184* WRITE(6,2026)185* 2026 FORMAT(1HO,6XIO7HFRACTION OF IODINE (12) REMOVED BY FILTER IN FLIda* 1OW FROM COMPARTMENT I TO COMPARTMENT K AT DATA TIME ENTRY J )I17* O0 1545 J=INUATA188* DO 1545 I=1,N189* WRITE(6.2028)I.JrN190* WRITE(6,2OO3)(EI2(I.K,J),K=1,N)191* 154! CONTINUE192* 2028 FORMAT(IH .6X,5H(EI2(,12.3HKo,I2,6H)tK=l,,I2,2H)=)193* WRITEi(,2030)194* DO 1550 J=1,NOATA19* 2030 FORMAT(IHO,6XIO7HFKACTION OF PARTICULATES REMOVED BY FILTER IN FL196* lUW FROM COMPARTMENT I TO COMPARTMENT K AT DATA TIME ENTRY J197* UO 1550 I:I,N198* WRITE(6,2032)I#JJN199* wRITE(b,2003)(EP(IKeJ)vK=l.N)200* 1550 CONTINUE201* 203e FORMAT(iH ,6X,4H(LP(,I2.3h.K,.I2,6H)PK=I#,I2,214)=)202* WRITE(b.2034)203* 2034 FORMAT(1HO.bX.2OHINPUT TIME VARIABLES ,//.6XI08HTD=TIME INTERVAL204* lOVER WHICH EFFECTIVE PARTICLE DIAMETER CHANGES LINEARLY FROM DPE I205* 2DIAMETER PARTICLES EARLY) t/.6X.33HTO DPL (DIAMETER PARTICLES LATE20b* 3). ,/.6X,51HDTDT2=TIME STEPfHRS) BEFORE MELT AND AFTER TJUMP1207* A/,6X,39HTVRI=TIME OF FIRST VAPORIZATION RELEASE208* 4/,bX,36HTVR2=TIME OF VAPORIZATION RELEASE 2. ./,209* 5bX,37f1TVRE=TIME OF VAPORIZATION RELEASE END I/t210* 6bX,24HTMR=TIME OF MELT RELEASE ,/l211* 76X,31HTMF=TIME OF MELT RELEASE FINISH ./p21Z* 8bX,5oHTCSI=TIME CONTAINMENT SYSTEM INJECTION SPRAY PUMP STARTS,/,215* 36X,5OHTCSIE=TIME CONTAINMENT SYSTEM INJECTION SPRAY ENDS ./.214* 96XP40HTECCIZTIME OF EMERGENCY CORE COOLANT INJECTION #/,21"* O6X,5bHTCSRI=TIME CONTAINMENT SYSTEM RECIRCULATING SPRAY STARTS ,/o216* 16X,55HTCSR1E=TIME CONTAINMENT SYSTEM RECIRCULATING SPRAY ENDS ./v217* 26Xv28HTEXI=TIME OF EXPLOSION NO. I P/,21d* 36X,26HTPUFF:TIME OF PUFF RELEASE t/0219* 46X,30HTJUAP1=TIME TO S4ITCH TO TIME STEP OT2 #/,220. 5bX,26HTJUiP2:TIME TO STEP TO ENID v/1221* WRITE(6 9 2036)TTDDTTVRITVR2,TVRETMRvTMrTCSI.TCSIETECCI TCSRIv222* *TCSRIETCSR2,TCSR2E.TEXITPUFFDT2,TJUMPI,TJUMP2,TEND223* 2036 FORMAT(IH .6X,3HTU=,IPE12.S,3X.3HDT=vlPE12.5,3Xo5HTVRI=,IPE12.5,2X224* A.5HTVR2=,IPEI2.5.2X225* 1.5HTVRE=vlPE14.8,3X,4dTMR=,1PEI4.8,/,6X,4HTMF=,lPE14.8,3X,5HTCSI=,22o* 11PE14.8,3XbHTCSIE=,lPE14.8,3X#6HTECCI=,lPE14.8,3X,6HTCSRI=lPE14227* I.8,/,6X,7HTCSRlE=IPE14.8,3X.6HTCSR2=eIPE14.8.3Xv7HTCSR2E=v1PE14.8,22d* 13A.5HTEXI:=,PE14.8.3Xo6HTPUFF=,IPE14.8v/,229* 16Xv4HDT2=,iPE14.B,3X,7HTJUMP1:IPE14.8v3X,7HTJUMP2=,lPE14.8v3Xv230* 15HTEND=,IPE14.8)231* WRITE(6.2038)232* 2038 FORMAT(IHOv6X,24HREACTOR COMPARTMENT DATA ,/I233* WRITE(ov2040)N234* 2040 FORMAT(IH ,6X,21HCOMPARTMENT WALL AREA#/#6XvlLH(AW)I)vI=lv#I2pVII-228

235* 124)=)23o* WR ITE(o#20O3)(AW(1).I=lN)237* WRITE(6.2042)N238* 2C42 FORMAT(IH o6X.22HCOMPARTMENT FLOOR AREAv/@6XolIH(AF(I),I=I,,12t239* 12H)=)240* aRITE(6,20C3) (AF(I),1=11J)241$ WRITE(6,2O44IN242*243.2C44 FORMAT(IH .bX,1BHCOMPARTMENTWRITEt6,20O3) (HT(1),I=lN)HEIGHT,/.6X.IIH(HT(I).IZp,2.2H)=)244* WHITE(b,2046)N245* 2046 FORMAT(IH .6X*31HCOMPARTMENT VOLUME (CUBIC FEET),/.6X,24b* 11OH(V(I),1=z, I2,2H)=)247* WRITE(v200O37(V(I)pI=1,N)246* WRITE(b,2048)249* 2048 FORMAT(IHOp6XP38HCONTAINMENT AND CLEANUP SPECIFICATIONSo//PbXt55250* 1HCSI=CONTAINMENT SYSTEM INJECTION FLOW IN CUBIC FEET/HR.P/eoX,652bl* IHCSRI=CONTAINMENT SYSTEM NO.1 RLCIRCULATING FLOW IN CUBIC FEET/HR.252* 1./,6X,61HHCSIHCSR=IIEIGHT THROUGH WHICH CONTAINMENT SYSTEM SPRAYS253* IFALL ,/.6Xtb5HDCSIPDCSR=DIAMETER OF CONTAINMENT SYSTEM SPRAY DROPL254* 2ETS ,/p6XF66HCUTOV=CONCENTRATION OF 12 BELOW WHICH SPRAY REMOVAL I255* 3S INEFFECTIVE ./e6XP28HVCL=VOLUME OF COOLANT LIQUID P/t25o* Il6X,71HIIO=RATIO OF IODINE IN SOLUTION IN WATER TO THAT IN VAPOR AT257* IEOUILIBRIUM,/,6X,37HECCI:EMEROENCY CORE COOLANT FLOW RATEp/,bX#2!d* 153HOFPP,DFPOI,DFPI2=PUFF RELEASE DECONTAMINATION FACTORS,/PbXo259* 1b3HXPUFF=FRACTION OF AIRBORNE MATERIAL RETAINED AFTER PUFF RELEASE260* 1 /,#6X.74HuPEDPL=EFFECTIVE DIAMETER OF PARTICLES IN MICRONS AT2bl* I EARLY AND LATE TIMES262* WRITE(b,2D50)CSICSRIeHCSI,3iCSR,DCSIDCSRCUTOV,VCL.HOECCIP263* IDFPPOFPOIDFI'12.CSR2,XPUFFDPEDPL264* 20O" FORNAT(1HOV6X,4HCSI:,IPE14.8,3X.SNCSRI=,IPE14.8.3X.5HHCSI=:IPEI4.8265* 1,3X,5riHCSR=,IPE014.e3X,SHDCSI=PIPE14.8./p6X,5HOCSR=.IPE14.8,3X.288* 1bHCUTOV=:.PE14.8,3X,4HVCL=,IPE14.8.3X.3HHO=,IPE14.8,3XSHECCI=,2o7* 11PE14.8,3X,/,bX,SHDFPP=, PE13.7e3X,6HDFPOI=:1PE13.7,3X,288* 16HDFPI2=,lPE13.7,3X.5HCSR2=.lPE13.7,3X,6HXPUFF=,lPE13.7./,269* 17X,4HOPE=,IPE13.7.3X,4HDPL=:tPE13.7270* WRITE(b,2052)271* NISO=NOI+,.JI2+NPAR272* WRITE(b.2054) (NAMES(J),(CFR(I.J).I=I:1.),J=l.NISO)273* 20:2 FORMAT(1HO#6X,63HCORE FRACTIONS RELEASED BY METHOD J IN ISOTOPE .GR274* IOUP I CFR(J,I).//.12X.IIHGAP RELEASEp4Xe12HMELT RELEASE,4X,275* 215HSTEAM EXPLOSIONIX,12HVAPORIZATION )27o* d0t4 FORMAT(6X.A6,IX.IPEIO.5,6XlPE-tOo.56XIPEIO5,bXIPEIO.S)277* C INITIAL CONDITIONS OF COMPUTED VARIABLES279* VTOT=O.279* C VTOT=TOTAL CONTAINMENT VOLUMEPFT3280* DO 8 I=I,N281* COOI(I, 1)= 0.282* CGP(Ip 1)= 0.283* CSP(I, 1)= 0.2b4* CSOI(I. 1)= 0.285* CGOI(I, 2)= 5.266* CGP(I, 2)= C.267* CSP(I, 2)= 0.288* CSOI(I. 2)= 9.289* C.ATP(I)=O.290* CMTI2(I)=U.291* CMTOI(I)=O.292* CATI2(I)=O.293* CVTI2(I)=O.294* CATOI(I)=O.295* CJTOI(I)=O.296* CATP(I)=O.297- CBTP(I)=O.296* CGI2(I, 1)= 0.299* C512(I 1)= 0.300* CGI21I, 2)= 9.301* CSI2(I, 2)= 0.302* VTOT=VTOT÷V(Il303* 00 8 J=1,10304* C SET 12 CONCENTRATIONS NON-ZERO FOR CUTOV TEST ONLY305* CMI2(lJ.I)=I.306* CAI2t1.J,1)=I.307* CU12(1.,Jl)=I.308* CAOI(I. J, 1)= 0.309* CtOI(I., J, 1): 0.310* CdP(I, Jv 1)= 0.311* CAP(I, J. 1)= 0.312. CMP(I. J, 1)= 0.343V COI(I. Je 1)= 0.314* CAOI(It J, 2)= 0.315. CROI(If J. 2)= 0.31o. CBP(I, J, 2)= 0.317* CAP(I, J, 2)= 0.318. CMP(I, J, 2)= 0.319* CAOI(I, J. 2)= 0.320* CAI2CIP J, 2): 0.321. C8I2(I. J, 2)= 0.322* CMI2(I, J. 2)z 0.323. 8 CONTINUE324* TEX=TEXi325* TX=O.32b* DO 10 J=1.10327* TM3(J)=O.VII-229

328*329*330*331*332*333*334*33•533b*357*338*339*340*341.342.34*3.344*345.346*347.34,3.349.350.351.352*353*35b4*3b6*35t,*357*358.359*360*361*3u2.3b6*3ob43b5*36o*367*368*369.370*371*37i*373*374*375*376.377*378*379*3b0*38L*3824383*3tsS-305*381*369*390*391-39e*3933s39 *39b*390*397* 401*398* 395.399*40U* 394.307.412*406*410*415*414.410*412*4171419.420*TM4(J)=TMR+FLOAT(J-1TMIIJ)= 0.TAI()= 0.TdI(j)= 0.TA3IJ)=O.*(TMF-TMR)/10.T63(J)=0.TA4(j)=TVRI+FLOAT(J-I)*(TVR2-TVRI)/10.T64(J)=TV;2+FLOAT(J-I)*(TVRE-TVR2)/10.1U CONTINUETA= 0.TSA= 0.CMLKP=O.CSLKP=O.CGLKP=O.Ci4LKI2=O.CVLKP=U.CSLKI2=O.CGLKI2=0.CVLKI2=0.CMLKOI=O.CSLKUI=O.CGLKOI=O.CVLKOI=O.OKEjGI=O.TGI:O.OKEOUZ=0.T32=0.OKEQSI1O.TSI:u. -OKE0S2=0.TS2=j.BEGIN CALCULATIONSM=ITT=O.SAVE OTNCHANG=I6 CONTINUE CONTINUE TO NEXT TIME STOPIF(TT.LT.TMR)UT=AMINI(OTITMR)IF(TT.GE. TMRIUT=.O0,.TMF-TMR,)IF(TT.GETVR )DT=.O5.(TVR2-T R1)IF(TT.GE.TVRI.AND.TT.LT.TMF)DT=.O5*AMINI(TMF-TMRTVR2-TVRI)IF(TT.GE.TVR2)DT=.I*(TVRE-TVR2)ADJUST DTIF(TT.GE.TVRE)DT=DTIADJUSTPERIF(TT.GT.TJUMPI)DT=OT2TO PROPERIF(TT.LT.TPUFF.AND.TT+DT.GT.TPUFF)DT=TPUFF-TT÷1.E-8SIZEIF(TT.LT.TCSLANDoTTtOT.GT°TCSI DT=TCSI-TT÷I.E-8IF(TT.LT.TCSRIAND.TT+OT.GT.TCSR1)DT:TCSR1-TTI.E-8IF(TT.LTTCSR2.AND.TT÷DT.61TTCSR2)DT=TCS•2-TT4I.E-8IF(TT.GT.TJUMP2)DT=TEND-TT NEW TIME TT, gRSTT=TT+DTBYPASS PARAMETER RECOMPUTATION IF THERMODYNAMICS IS UNCHANGED.CHANGES OCCUR AT TIMES TCHANGSOITCH:.FALSE.C IF(M.GT.I.AtND.TT.LT.TCHANG(NC(iANG))GO TO 1333 TRIS TEST CAN BE USED TOWCHANG=NCHANG+1 BYPA55 IO 1333 7! NO RHANGES OCCURSWITCH=.TRUE.C COMPUTL P(N),T(N).DELTAT(N',VAP(J),ELI2(jI)ELP(J) FOR THE PARTICU-C LAR TIME TT.DO 11 J=2.NDATAIF(AbSiTI(JI-TT).LT.At•SiTI(NN)-TT))t4N=J11 CONTINUE4;'I=.AXO(2,M I IN IN. NDAIA-I)AI=TI (NN-I)A2=T1 (SWA3TTI(NN+I)UO 13 K=INP(K)=UINT(Al, A2, A3, TT,PI(K,NN4-1) ePI(KNN) ,PIU(KNN+I))TA-P(r)=QL IT( Al, A2, A3, TT. TMY(K,tlN-1), TMY(K,NN) , TMY (K.NN÷1)pELTAT(K)=QINT(AI,A2,A3,TTOELTTI(KNN-1),DELTTI(K.N N)PDELTTIC(K,VAP)K)=5IINT(A1.A2.A3,TTVAPI(K.NN-1),VAPI(K.NN).VAPI(K.NN+I))ELI2(K):0[NTCAIA2eA3,tT.ELKI2(K.NN-1).ELKI2(KNN),ELKI2(KNN+1)) INTERPOLATIONELP(K)=OI'4T(A1.A2.A3,rTELKP(K,NN-1),ELKP(K,NN),ELKP(KNN+I ) OF INPUT DATA"OI)K)=QINT(AI,A2,A3,TT,ELKOI(K,IIN-1),ELKOI(KNN),ELKOI(K,NNI1))TO GET VALUE AT TT.C COMPUTE G(I,J),EFI2(I.J),EFI2(I,J) FOR ANY TIME TT00 13 L:ltlG(K,L)=QINT(A1.A2,A3* TTGIC(KL,NNJ-1) GI(K.LNN),GI(KLNN+I)EFI2(K,L)=UINT(AltA2,A3,TT#EI2(K,Lt4N-I)#EI2(KrL,NN)PE12(KeLNN÷I1)) EFP{vk,L}=lJ~t~l(Al,A2,A3.TT-EP(K,L.NN-1) EPtK.L,N14) EP(K#L#NN÷I))13 CONTINUEC1a33 CONTI11UECC COMPUTE GAS PARAMLTERSTRANSPORT COEFFICIENTS, DIMENSIONLESS NO.sCDo 14 K=1,NCALL LRR(3,181HFRACAS D0 14IF(.NOT.Sv•ITCH)GO TO 1102P(K)=P(K)+14.7TR=TMP(K)+469.0NST(K)=P(K)/(lO.73*TR)PSIARANKINEVII-230

421*422*423*424*42b*426*'40*429*430*431*43z2433*434*435*437*438*439*440*441*449*443*444*445*446*447*44d*449*450*452.452*454*457*4S7*458*459*4,,0*461*462*463*464*4tb5*466*467*468*469*470*471*472*473**474*475*47o*477*478*479*480*481*462*483*484 *486*417*488*4L$9*490*491*492*493*494*495*49b*497**496*499*500*501*502*503*504*505*506*507*508*509*510*511*512*5134JriST(K)=DNST(K)*VAP(K)*18.+(1.-VAP(K J)*DNST(K}*29.D'ENSITYC ABOVE IN LBS/FT3TRIP5=TR*SGRT(TR)VST(K)=.003339*TRIP5/(rR+1224.2)VISCOSITII IS WATER VAPORVAIR(K)=.0414*(TR/492.)**.768AIRIF(VAP(K).EO.l.)GO TO 7IF(VAP(K).EO.O.)GO TO 773PAS(K):(I.+SQRT(iVAIR(K)/VST(K))*SORT(I8./29.)))**2/4.5704PSA(K)=(I.+SQRT((VSTtK)/VAIR(K))*SQRT(29./1B.)))**2/3.6008RMOL=VAP(K)/(I.-VAP(K))VM(K)=VAIR(K)/(I.+RMOL*PAS(K) )VST(K)/(I.+PSA(K)/RMOL) EQUATI ON (VII J-21)GO TO 7787 VM(K)=VAIR(K)GO TO 778773 V,.(K)=VST(K)778 CONTINUEC AuOVE IN LB/FT/HRWA(K):O.7075414I.73/(TR/I.8)WS(K)0.7075*454.72/(TR/I.8)12 DIFFUSIVITYC DA(KJ=.0020435*(TR/1.8)**1.5*.19b1/((P(K)/14.7)*4.43b**2*WA(K))JA(KJ=.8432578E-5*TRIP5/(P(K)/14.7*WA(K)IC OS(K)=.0020201*((T(K)+460.)/1.8)**1.5*.2438/((P(K)/14.7)*3.901**2(EQUATION VII J-22)C I*WS(K))OS(K)=.134012b9E-4*TRIP5/( P(K)/14.7*WS(K)}C AbOVE DIFFUSIVITIES IN CM2/SECDI2(K)I./(VAP(K)/DS(K)+(I.-VAP(K))/DA(K))DI2(K)=DI2(K)*3.875C 012(K) IN FT2/HRC SCHMIDT NO= VISCOSITY/IDENSITY*DIFFUS.)=SCtK)SC(K)=VM(K)/(0NST(K)*OI2(K))C GPASHOF NO. FOR NATURAL CONVECTIONCKG=MT COEFF IN FT/HRGN(K)=4.17312E+8*HT(K)**3*DNST(K)**2*DELTAT(K)/(TR*VM(K)**2)MASS TRANSFERIF(HT(K).LE.10.)CKG(K)=.59*DI2(K)*SGRT(SORT(GR(K)*SC(K))) /HT(K) 1lf(HT(K).8T.10.)CKG(K)zOI2(K)*(S.9*SQRT(SORT0RI(K)*SC(K)))/HT K)] COEFF, EQNs1+.13*(HT(K)-1O.)*CBRT(GR(K)*SC(K))/HT(K))/HT(K)VII J-24, J-251102 CONTINUEC COMPUTE SETTLING VELOCITIES FOR PARTICLESC HAVE PARTICES FROM 13 SOURCES *GAP RELEASE,10. TIME INTERVALS OFC THE MELT RELEASE + 2 EXPLOSIONS * 20 VAPORIZATION RELEASESC DG.OM(IO).DEXIDEX2,DVA(IO) 0V8110)OG=AMAXI (UPE+IDPL-DPE)*(TT/TD),DPL)DEXI=AMAXI(DPE+(DPL-DPE)*(TT-TEX1)/TDDPL)IF(TT.LT.TEX1) DEX1=O.SPRAY=O.IF(((T.T.GT.TCSI.AND.TT.LT.TCSIE).OR.(TT.GI.TCSR1.AND.TT.LT.1TCSRIE).OR.(TT.GT.TCSR2.AND.TT.LT.TCSR2E)).AND.K.EO.1)SPRAY=l.IF(TT.LT.TMRJGO TO 1555DO 15 IODel10DM1IDI=AMAXI(DPLDPE+(DPL-OPE)*(TT-TM4(ID))/TD)IF(TT.LT.TM4(ID))UM(I)=O.C PARTICLE DIAMETERS ARE IN MICRONSSKG(KID)=(1.-SPRAY)*(.01557*DM(ID)**2/VM(K))*AF(K)/V(K)1ý CONTINUElb55 CONTINUESKGG(K)=(i.-SPRAY(*(.01557*DG**2/VM(K))*AF(K)/V(K)SKG1(KI)(1.-SPRAY)*(.01557*DEXI**2/VMIK))*AF(K)/V(K)IF(TT.LT.TMFTGO TO 3016DO 2015 ID = 1.10DVA(ID)=AMAXi(DPLDPE+(OPL-DPE)*(TT-TVA 10))/TD)IF(TT.LT.TVA(IO))0VA(ID)=O.DVB(IO)=AMAXI(OPLDPE,(DPL-DPEl*(TT-TV8(IDI /TD)IF(TT.LT.TVB(ID)lDVB(ID)=0.SKGA(KID)=(1.-SPRAY)*(.O1557*DVA(ID)**2/VM(K))*AF(K)/V(K)SKGB(KIDIO(L.-SPRAYI*(.O1557*DVB(IDI**2/VM(K)I*AF(K)/V(K)2015 CONTINUE3016 CONTINUEC SKG:NATURAL DEPOSITION LAMBDA IN HR-I (PARTICULATES)CCCCCCCCCCC14 CONTINUEREMOVAL COEFF FOR 12.AWD +PARTICULATES COMPLETE (W/O SPRAYS)SPRAY LAMdDAS NEED TO BE COMPUTED NOWPARTICULATE SPRAY LAMBDAIN COMPARTMENT IFIRST ESTABLISH SUM(FT/V(1)) AND IMPACT EFFICIENCYFI=AMAXL(O..AMIN1(TT-rCSI-AMIN1(DT/2.#(TT-TCSI)/2.).TCSIE-TCSI))l*CSIFRI:AiAAXI(0.,AMINI(TT-TCSRI-AMINI(DT/2..(TT-TCSRI)/2.),TCSRIE-1TCSRI))*CSR1FR2=AMAXI(OAMINI.(TT-TCSRZ-AMINL(DT/2.,(TT-TCSR2)/2.).TCSR2E-ITCSR2})*CSR2FA=CFRI+FR2+FI)/Vt1)IF(FA.GE.O..AND.FA.LT.O.002)L=-15.825*FA+.06IF(FA.GE.0.-02.AND.FA.LT.O.193)E=.04625-SORTI.08626+42.68*FA)/21..14IF(FA.GT..0I93)E=.O015TESE EFFICIENCIES ARE FOR GAP RELEASE ONLYIN THE EVENT THAT THE SPRAYS ARE WORKING AFTER ANY EXPLOSION ANDSOURCE RELEASE S#THE EFFICIENCY OF REMOVAL MUST 8E RECALCULATEDFOR REMOVING S.THE REMOVAL EFFICIENCY IS EEXFI=AMAXI(O.,AMINI(TT-TEXI-AMINI(DT/2..(TT-TEXI)/2.)PTCSIE-ITEXII)*CSI;EQN VII J-28 andFIGURE VII J-4VII-231

514* FI=AMAXI(O.,AMINI(TT-TEXI-AMINI(DT/2.,(TT-TEX1)/2.),TCSRiE-515* 1TEXI))*CSRI516* F2=AMAXI(O.,AMINI(TT-TEX1-AMINI(DT/2.,(TT-TEX1)/2.),TCSR2E-517* 1TEXI))*CSR2516* ES=FI+F2+FI519* ES=ES/V(l)520* IF(ES.GE.O..AND.ES.LT.O.002) EEX=-15.825*ES+.O6521* IF(ES.GE..OO2.ANO.ES.LT.O.0193) EEX=.04625-(.08626+42.68*ES)**.5/2522* 11.34523* IF(ES.GT..0193)EEX=.0O15524* C CALCULATE EFF. FOR PARTICULATES FROM MELT RELEASE FOR EACH OF THF525* C 10 DIVISIONS* CALL IT EFM(J)52b* IF(TT.LT.TMRIGO TO 1777527* DO 17 J=1 9 1O528* AMDAMP(J)=O.549* IF(TT.LT.TM4(.)))GO TO 17530* FM4I=AMAX1(O.,AMINI(TT-TM4(J)-AMINI(DT/2.,(TT-TM4(J))/2.).531* ITCSIE-1M4(J)))*CSI532* F4I=AMAXI(O.,AMIN1(TT-1M41()-AMINI(DT/2..(TT-TM4(J))/2.)p53a* ITCSRlE-TMU(J)))*CSRI534* Fl2=AMAXI(O.,AMINI(TT-TM41J)-AMIII(DT/2..(TT-TM4(J))/2.)o535* 1TCSR2E-TM4(Jl))*CSR253b* ET= FMI+FMI+FM2537* ET=ET/V(I)538P IF(ET.GE.I..AND.ET.LT.O.O2)EFM( J)=-15.825*ET+.Ob539* IF(ET.GE..GC2.AND.ET.LT.O.0I931 EFN(J):.O4625-(.08626+42.68*ET)**.54U* 15/21.34541* IF(ET.GT..019J) EFM(J)=.O015ý42* C COMPUTE543* IF(TT.GT.TCSI.AND.TT.LT.TCSIE) AMDAMP(J)=AMDAMP(J)+1.5*HCSI*CSI*EF)44* IM(J)/IuCsI*V(1)/30.48)545, IF(TI.GT.TCSFI.ANU.TT.LT.TCSRIE) AMDAMP(J)=AMOAMP(J)+1.5*HCSR*CSRI54b* l*EF4(J)/(DCSH*V(I)/3U.48)547* IF(TT.GT.TCSR2.AND.TT.LT.TCSR2E) AMDAMP(J)=AMDAMP(J)+I.S*HCSR*CSR2546* I*EFM(J)/[OCSR*V(1)/30.48)549* 17 CONTINUE550* 1777 CONTINUE5nl* C MELT RELEASE PARTICULATE LAMDAS ARE COMPLETE552* C COMPUTE OTHER PARTICULATE LAMDAS NOW553* AMDAPG=O.554* AMDAPS=0.555* IF(TT.GT.TCSI.AND.TT.LT.JCSIE) AMDAPG=AMDAPG+1.5*HCSI*CSI*E/(DCSI*556* IV(1)/.U.46)5ti7*IF(TT.GT.TCS1.AND.TT.LT.TCSIE)AMDAPS=AMDAI'S+1.5*ICSI*CSI*EEX/(DCS556* ll*V(1)/30.481559* IF(TT.GT.TCSR1.AND.TT.LT.TCSRIE) A'IDAPG=AMDAPG*I.5*HCSR*CSRI*E/(DC5O* 1SR*V(11/30.48)501* IF(TT.GT.rCSH1.AND.TT.LT.TCSRIE) A!*4DAPS=A,.DAPS+1.5*HCSR*CSRI*EEX/(5o 2 * IOCSR*V(t)/3C.48)5,3* IF(TT.GT.FCSRŽ.ANU.TT.LT.TCSR2E) A4DAPS=AMDAPS+1.5*HCSR*CSR2*EEX/( I5o4* IDCSR*V(1)/6C.48)565* IF(TT.GT.TCSR2.AND.TT.LT.TCSR2E) AMOAPG=AMDAPG÷1.5*HCSR*CSR2*E/(DC5Oh* 1SR*V(I)/30.48)567* C COMPUTE VAPORIZATION PARTICULATE REMOVAL BY SPRAYS5bb*DO 217 J=1#19569* AMUAPA(J)=O.570* AMUAPU(J)=O.571* IF(TT.LT.AMINI(TCSITCSRi.TCSR2).OR.TT.GT.AMAXI(TCSIEPTCSRIE.572* 1TCSR2E))GO 10 217573* IF(TT.LT.TA4(J))60 TO 217574* FAI=AMAX1(O.,AMINI(TT-TA4(C)-AMINI(OT/2.,(TT-TA4IJ))/2.).575* ITCSIE -TA4(J)))*CSI576* FAI=A*IAXI(O.,AMINI(TT-TA4(J)-AMIIJ1(DT/2.odTT-TA4(J))/2.).577* 1TCSRIE-TA4(JT))*CSRI570. FA2=AMAXI(O.,AMINI:(TT-TA4(J)-A'"INI(DT/2.,(TT-TA4(J))/2.).579* 1TCSR2E-TA4(J)))*CSR25600 FIA.1AXI(O..AMINI(TF--TB4(J)-AMIHI1(DT/2.,(TT-TU4(J))/2.2.hb1*Sode*ITCSIL -TB4(J)))*CSIF1I•=A;.4AX I(O..A! I I JI(TT-TR4 (J)-A;A I N1(OT/2.. (TT-TB4 (J) )12.)5•* ITCSRIL-TF4(J) ) CbR1564' Fj2=AMAAXI(O.,AMIrhI(TT-TR4(J)-A'4IIlOT/2.,(TT-Tb4IJI)/2.),58ýo ITCSR2Lt-T4(IJ)))*CSR25oosS•=(FAI+FAI+FA2)/V(1)56t7- E =(F6[+FiI+FL.2)/i(1)5o6* LF(EA.LT..O19j)GO TO 4750btY*cFA(J)z.'JJ155,O TO 4851591* 475b IF (EA.LT..C;)GO0 TO 47o0592* EFA())=,.0425-SORl(.06b26+42.08*LA)/21.3451J. PGO TO 485i.)94. 4 7 Ll EPA (J) =-ID. tl2*EA+. U,59!* 4o1O CuNTINIJE59o* Iý(Eb.LT-.0193)GO TO 575U597* EFU(Jd:.Oul5596* :0 TO 58•5(s596o)It 3 *5750 IF (Lu.LT..C92)GO TO n7601) d()Z.(%62()-SNT(.dhu20+42.o8*EL))/21.34bl)l* GO To 535U602* 5?o 1 L.Fb4J)=-,5.A2t5*Eb+.0Ob03* 5d00, CQI ITil4)UE6(04* IF(TT.LT.TCSI.OR.IT.GT.TCSIE}GO TO 585460a- AiDUAPA(Jd)1.5*HCSI*CSI*EFA(J)/(DCSI*V(1)/30.48)((lb* .A;iuAad(J)=I.,*tHCSi*CSI*EF(J)/ iDCS I*V( 1) / 504)VII-232

07*60do609*610*615*tojb*617*old*620(*621*622*623*624*625*62b*627*628*629*b30*t310632*633*634*635*63b*oil-638-639*640*641*642*b43*644*646*647*648*649s650*651*652*653*654*655*65b*657*658*659*660*661*662*663*664*665*666*6o7*668*6o9*670*671s672*673*674*675*bbP4 IF(TT.LT.TC'|RI.OR.TT.GT.TCSRlE)GO TO 595btAi-AOAPA(J)=AMDAIPA(J)41.5*HCSR*CSRI*EFA(J)/(DCSN*V(1)/30.48)AUAP(J)?ANJAP8b1J)1+.5*HCSR*CSRI*EFBIJ)I/(DCSk*V(1)/3D.48)56bu IF(TT.LT.TCSM2.OR.TT.GT.TCSR2E)GO TO 217L.tDIUAPA(J)ZAMDAPA(j)+1. 5*HCSR*CSR2*EFA(J)/(DCSR*V(1)/30.48)AýIDAPU(J)=AMUAPf3(J)+i.b*HCSR*CSR2*EFO (J)/(DCSO*V(I)/30.48)21ý7 C1H1 I1UEC ALL PARTICULATE SPRAY LAMDAS COMPLETECC IUUINEtI2. REMOVAL BY SPRAY CALCULATED NOwCC SPRAY LAMUDAS COMPLiTEO 1406 FOR 12(AMDAIG,AMDAIS.AMDAIM(K))CC COMPUTE TERMINAL VELOCITY OF SPRAY DROPSCFUREIC{4./3.)*62.4*DNST(1)*(DCSI/3C.48)**3*417312000./VM(I)**2F'ORER=FDREI*(UCSR/DCSI)**3IF(FDRLI.LT.1I700.) REI=(FDREI/15.71)**.7027IF(FUREI.GE.I6700.) REI=(FDREI/6.477)**.6215IF(FDRER.LT.i0700.J RER=(FDRER/15.71)**.7027IF(FDRER.GE.10700.) RER=(FDRER/6.477)**.6215UTI=REI*VM(I)/(DNST(I)*OCSI/30.48)UTR=RER*VM(1)/(DNST(1)*DCSR/30.48)C UTI.UTR I- FT/HRC MASS XFR COEFF TO THESE DROPS COMPUTED BELOW (GGIuGGR) IN FT/HRGGI=DI2(1)*(2.+.6*REI**.5*SC(1)**.33)/(DCSI/30.48)GGR:DI2(1)*(2.+,.6*RER**.5*SC(1)**.33)/(DCSR/30.48)C GAS PHASE COMPLETEtNO4 COMPUTE LIQUID PHASE MASS XFR COEFF.GLItGLRC UL=LIQUID VISCOSITY.CP, DLZDIFFUSIVITY OF 12 IN H20vFT2/HRTl=(TMP(I)+460.)/1.8UL=100./(2.1482*((T1-281.6).(8078.4+(TI-281.b)**2)**.5)-120.)DL=1.5137E-07*(Ti/UL)GLI=b.58*DL/(UCSI/30.48)GLR=6.58*OL/(DCSR/30.48)C STAGNANT DROP MODEL USEDEFFICIENCY=I-EXP(-6.KGOT/S(H÷KG/KL))ECSI:±.-EXPL-o.*OGI*(HCSI/UTI)/gIDCSI/30.48)*(HO+GGI/GLI)))ECSR=t.-EXP(-6.*GGR*(HCSR/UTR)/((DCSR/30.48,*(HO+GGR/GLR)))C SPRAY LAMBDAS=F*HO*E/V.IF C/V(I).GJ.C(EMITTEO)/(SUM V(I))*CUTOFFC EQUILIBRIUM LAMBDASUSED AT LOWER CONC.C COMPILE PROPER ELANC SUBROUTINECOG=CGI2(1, 1*VTOT/V(I)COS=CS12(1.1)*VTOT/V(I)DO Ib J=1,IOCaM(J)=CMI2(j.j,1)*VTOT/V(1)/.1CB0(J)ZCBI2(IJ.I)*VTOT/V(1)/(.O7528*EXP(-.1*FLOAT(J-1)))C0Alj)=CAI2(IJ,1)*VTOT/V l)/(.07528*EXPI-.I*FLOAT(J-I)))ib CONTINUEOKG=U.6G GAP RELEASEOKS=O. S- STEAM EXPLOSIONIF(COG.GT.CUToV I OKG=I.IF(COS.GT.CUTOV ) OKS=I.DO 19 J=1.@1OKM()=O.MN= MELT RELEASEOKA(J)=O. AI 1st HALEIOF-VAPORZATION RELEASEOKB(J)=O. B- 2nd HALF OF VAPORIZATION RELEASEIF(COM(J).GT.CUTOV)OKM(J):1.IF(COA(J).GT.CUTOV)OKA(J)=I.IF(COB(J).GT.CUTOV)OKB(J)=I.19 CONTINUEIF(TT.LT.TCSIE.AND.TT.GE.TCSI) OKCSI=I.IF(TT.GE.TCSIE.OR.TT.LT.TCSI) OKCSI=O.IF(TT.LT.TCSRIE.AND.TT.GE.TCSRI) OKCSRIZI.IF(TT.6E.TCSRIE.OR.TT.LT.TCSRI) OKC-SRl:O.IF(TT.LT.TCSR2E.AND.TT.GE.TCSR2) OKCSR2=I.IF(TT.GE.TCSR2E.OR.TT.LT.TCSR2) OKCSR2:0.AMDAIG=CSI*HO*ECSI*OKG*OKCSI/V(13)CSRI*HO*ECSR*OKG*OKCSRI/V(I)÷HO*ICSR2*ECSR*OKG*OKCSR2/V(I)AMDAIS=(HO*OKS/V(i))* CSI*CSliOKCSI÷CSRI*EC~SOKCSR•+CSR2*ECSR*OKHEQN VII J-27 ANDON PARA. J3.2.2ZHHHH02 Z U))11~ IN>4 r00FOLLOWINGb7T*677*678*679*680*661*682*683*684*665*668*689*690*691*b92*693*694*695*696*b97*b9d*699*ICSR2)DO 20 J=1,1CAMDAIM(J)=(IIO*OKM(J)/V(1)I*(CSI*ECS5*OKCSI+CSRI*ECSR*OKCSRI+CSR2*IECSR*OKCSR2)AMUAIA(J)=(HO*OKA(J)/V{1))*fCSI*ECSI*OKCSI÷CSRI*ECSR*OKCSRICSR2*IECSR*OKCSR2)AMDAIi3(J)=(HO*OKB(J)/V(I))*(CSI*ECSI*OKCSI+CSRI*ECSR*OKCSR1+CSR2*1ECSR*OKCSR2)20 CONTINUEC SUMP LIQUID VOL. AND 12 CONC. IN THE LIG.CONTROL EGUILIBRIUM,VSUMPC Eo ECCI+VCL+CSI*(TT-TCSI)IF(TT.GE.TECCI) OKECCI=I.IF(TT.LT.TECCI) OKECCI:O.IF(TT.GT.TCSIE) OKCSIE=I.IF(T7.LE.TCSIE) OKCSIE=O.VSUMP=VCL+ECCI*OKECCI+(TCSIE-TCSI)*CSI*OKCSIE+(TT-TCSI)*CSI*OKCSIC VSUMP=TOTAL SUMP LIQUID.FT3C CALCULATION OF LAMOAS FOR EQUILIBRIUM PROCEED BELOW. THESE AREC E6iAG.ELAMSELAMM(I).EACH SPRAY TYPE REQUIRES A SPECIAL DECK TOC CaMPUTE THESE. FOR THE CASE OF NO SPRAY# USE H803 DECK.CC BJRIC ACIO,HbU3 DECK FOR COMPUTING EQUILIBRIUM CONCENTRATIONSC DESIGNATE TGI AND TG2 AS TIME POINTS WHEN CGI2(lK) EXCEEDS ANDC IS LT THE EQUIL. CUTOFFCUTOV=.Ol FOR ALL SPRAYS EXCEPT THIOSULFA;0rA E4ý0H 40c0VII-233

700* C TE#701* IFtCOG.GT.CUTOV)GO TO 6542702w IF(.NOT.((TT.GE.TCSI.AND. TT.LE.TCSIE).OR. TT.GE.TCSRI.AND.TT.LE.71)3* 2TCSRIE).OR.CTI.GE.TCSR2.AND.TT.LE.TCSR2E))) O0 TO 6542704* TA=TA+DT70"5. ELAMG=ELAM(TA)70** GO TO b545707* 6542 CONTINUE708* ELAMG=O.709* TA=O.710* 6645 CONTINUE711* C ELA.46 IN I/HR712* IF(COS.GT.CUTOV.OR.TT.LT.TEX1)GO TO 6583713* IF(.NOT.((TT.GE.TCSI.AND.TT.LE.TCSIE).OR.(TT.GE.TCSRI.AND.TT.LE.714* 2TCSRIE).OR.(TT.GE.TCSR2.ANO.TT;LE.TCSR2E))) 60 TO 65837iZ•*TSA:TSA+OT716* ELAMS=ELAM(TSA)717* GO TO b585718* 6z83 CONTINUE719* ELAMS=O.720* TSA=O.721* 6585 CONTINUE722* C ELAMS IN HR-I723* IF(TT.LT.TMR)GO TO 8001724* DO 8000 J=IplO725* IF(COM(J).GT.CUTOV.OR.TT.LT.TM4(J))60 TO 71237.6* IF(.NOT.((TT.,E.TCSI.AND.TT.LE.TCSIE).OR.(TT.4E.TCSRI.AND.TT.LE.727* 2TCSRIE).OR.(TT.GE.TCSR2.AND.TT.LE.TCSR2E))) Go TO 712372t* TMItJ)=TMI(J)+OT729* ELAMM(J)=ELAM(TMI(J))730* 60 TO 8000731* 7123 CONTINUE732* ELAMM(J)=O.7330 TMI(J)=O.734* 8000 CONTINUE73:5 8001 CONTINUE73io* IF(TT.LT.TVRI)GO TO 9001737* DO 9000 J=1.10738* IF(COA(J).GT.CUTOV.OR.IT.LT.TA4(J))GO TO 3644739* IF(.NOT.((TT.GE.TCSI.AND.TT.LE.TCSIE).OR.(TT.6E.TCSR1.AND.TT.LE.74•0* 2TCSRIE).OR.(TT.GE.TCS,42.AND.TT.LE.TCSR2E))) GO TO 3644741* TA1tJ)=TAIfJ)÷OT742* ELAAMA(J)=ELAM(TA1(J))743* GO TO 9000744* 3o44 CONTINUE745* ELAMA(J)=O.746* TAI(J)=O.747* 9000 CONTINUE748* 9001 CONTINUE749* IF(TT.LT.TVR2)GO TO 5001750* 00 5UOO J=1ll751* IF(COB(J).GT.CUTOV.OR.TT.LT.TB4(J))GO TO 4821752* IF(.NOT.((TT.GE.TCSI.ANO.TT.LE.TCSIE).OR.(TT.GE.TCSRI.AND.TT.LE753* 2TCSRIE).OR.(TT.GE.TCSR2.AND.TT.LE.TCSR2E))) GO TO 4821754* Tdl(J)=T81C(J)+UT755* ELAMB(J)=ELAM(TBI(J))756* GO TO 5000757* 4dda CONTINUE758* ELAMB.(J)Uo.759* T31(J)=O.7b0* 5C00 CONTINUE761* 5001 CONTINUE7o2* C ELAMM(J) IN HR**-1763* C AT THIS POINT EQUILIBRIUM LAMBDAS ARE COMPLETE7b4* C765* C DIFFERENTIAL EONS ARE TO BE SOLVED NOW76o* C7b7* C ESTAdLISH RELEASE INPUTS AS INITIAL CONCENTRATIONS FOR EXPLOSION7686* C RELEASE AND MELT RELEASE.FOR GAP RELEASEP ICS ARE ESTABLISHED769* C INCLUDE VAPOR RELEASES770* C EXPLOSION RELEASE771* IF(TT.GE.TEXI.AND.TX.EQ.O.)TX=TT772* IF(TX.EQ.TT) OKICSZI.773* IF(TX.NE.TT) OKICSZO.774* C MELT RELEASE775* 00 23 jd1.1C776* IF(TT.GE.TM4(J).AND.TM3(J).EG.O.) TM3(J)=TT777* IF(TM3(J).EQ.TT) OKICM(J)=.l77d* IF(TMS(J).NE.TT) OKICM(J)=O.779* IF(TT.GE.TA4(J).AND.TA3(J).EQO.0.) TA3(J)=TT760* IF(TT.GE.TB4(J).AND.TB3(J).EG.O.) TB3(J)=TT781* IF( TA3iJI.EG.TT) OKICA(J)=.O7528*EXP(-.l*FLOAT(J-1))78•* IF( TB3(J).EQ.TT) OKICB(J):.2045*EXP(-.I*FLOAT(J+9))7o3* IF( TA3(J).NE.TT) OKICA(J)=0.784* IF( Tb3(J).NE.TT) OKICB(J)=0.785* 23 CONTINUE786* C787* C SOLVE MULTICOMPARTMENT DIFFENTIAL EQUATIONS (EQN VII J-17) BY SUBROUTINE788* CALL ERR(35.IHCORHAL S.N. 23 1 ALCEMY (SEE PARA. J3-2)789* DO 190 IMAT=1,3790* C SET ALCEMY FIxED INPUT VARIABLES791* NT:I792* 11=2*NVII-234

793* LL=O794* GO TO(30.60.120)PIMAT795* 30 CONTINUE79b* C SET OFF DIAGONAL MATRIX ELEMENTS FOR P BRANCti797* 00 24 ICO'4•1.N798* H(ICOM+N, ICOM)=ELP(ICOM)799* 00 24 JCOM=I#N800. 24 HIJCOArICOM)=G(ICOMJCOM)*(1.-EFP(ICOMJCOM))/V(ICOM)801* DO 50 IREL1.5802* GO TO (31,.4151,1061.IOT71,IREL803s C PRECEDING BRANCH IS TO GAP RELEASEMELT AELEASE.STEAM EXPLOSION,804* C VAPORIZATION RELEASE PHASE A, VAPORIZATiON RELEASE PHASE B805* 31 CONTINUEBOb* C RELEASE 1I PARTICULATES BRANCH807* c8nft*T1})=JT809* DO 37 ICOM=I,N810* H(ICON. ICOM)=O.811* BETAPG(ICOM)=SKGG(ICOM)+ELP(ICOM)+FDP(ICOM)812* IF(ICOM.EQ.IIBETAPG(ICOM)=BETAPG(ICOM)+AMDAPG813* F(ICO.4)=CGP(ICOM,1)814* F(ICOM÷NIzO.815. IF(M.EG.l.ANO.ICDM.EQ.2)FIICOM)1=.816* D0O 39 JCOM=IN817* 39 H(ICOMPICOM)=H(ICOMICOMl-G(ICOMPJCOM)/V(ICOM)818* 37 H(ICOM#ICOMIZH(ICOM.ICOMI-BETAPG(ZCOMI8194 CALL ALCEMY820* DO 38 ICOM=IpN821* CGLKP:CGLKP+FT(ICOM+N,)822* 38 CGP(ICOM.2)=FT(ICOMPI)823. C GAP RELEASE PARTICULATES COMPLETE FOR PRESENT TIME STEP.824* GO TO 50825* 41 CONTINUE MELT R82b* C IF MELT RELEASE HAS NOT STARTED. BYPASS827* IFITT.LT.TMR)GO TO 508s8* DO 45 JREL=I,10829. IF(TT.LT.TM4(JREL))GO TO 45830* T(I)=AMINItDT.TT-TM4(JREL))831* 00 47 ICOM=I*N832* F(ICOP4)=CMP(ICOMJRELI)833* IF(ICOM.EO.2.AND.OKICM(JREL).NE.O.)F(ICOM)=OKICM(JREL)834* F(ICOM+N)=O.835. H(ICOM.ICOM)=O,836. dETAPM(ICOM.JMEL)=SKG(ICOM.JREL)+ELP(ICOM)+FDP(ICOM)837* IFIICOM.EG.1)BETAPM(ICOMJREL)=BETAPM(ICOMJREL)+AMOAMP(JREL)83d* DO 49 JCOM=I.N839* 49 H(ICOM.ICOM)}H(ICOM.ICOM)-G(ICOM,JCOM)/V(ICOM)840* 47 H(ICOM.ICOM)=H(ICOMICOM)-BETAPM(ICOMPJREL)841* CALL ALCEMY842* DO 48 ICOM=1.N843* CMLKP=CMLKPeFT(ICOM÷N.1)944. 48 CMP(ICOMJREL,2)=FT(ICOMI)845, 45 CONTINUE84u* C MELT RELEASE PARTICULATES COMPLETE FOR PRESENT TIME STEP847* 0O TO 5084d* 51 CONTINUE STEAM EX849* C IF STEAM EXPLOSION HAS NOT OCCURRED.BYPASS850* IF(TT.LT.TEXI)GO TO 50851* T(I)=AMIN1(DT.TT-TEXI)852* DO 57 ICO4=1,N853. H(ICOMICOM):o.854* F(ICOM)=CSP(ICOM.i)855* F(ICOM÷N)=O.850* IF(ICOM.EG.I.ANO.OKICS.NE.O.)F(ICOM)=OK;CS857* BETAPS(ICOM)=SKG1I(ICOM)+ELP(ICOM)÷FDP(ICOM)85d8IF(ICOM.EQ.IGETAPS(ICOMN=BETAPS(ICOM)+AMDAPS859. 00 59 JCOMI.N860* 59 H(ICOMICOM}=H(ICOMICOM)-G(ICOMPJCOM)/V(ICOM861* 57 H(ICOM, ICOM)=HIICOMICOM)-BETAPS(iCOM)802* CALL ALCEMY863* IF(MODIM-1.10).NE.O)GO TO 5086864* 5036 CONTINUE865* C EIlO OF AD HOC WRITE8bb.00 58 ICON=k1N867* CSLKP=CSLKP+FT(ICOM+N.l)808* 58 CSP(ICOMo2)=FT(ICOM.1)869. C STEAM EXPLOSION PARTICULATES COMPLETE FOR PRESENT TIME STEP870* GO TO 50871* 1061 CONTINUE VAPREL A872. C IF VAPORIZATION HAS NOT STARTED, BYPASS.873. LF(TT.6T.TVR1)GO TO 50874* OU 1065 JRCL=L,1IO875* IF(TT.LT.TVA(JREL))GO TO 106587b* TII)=Av.INi(DTTT-TVA(JREL))877* 00 10o7 ICOM=L,N878* F(ICOH)=CAP(ICOM.JRELI)879* IFIICOM.Ea.1.AND.OKICAIJREL).NE.O.)F(ICOM)=OSICA(JREL)880*ddl*FIICOM+N)=O.H(ICQ:4*IC0M)=0..862# (iTAPVA[ICOM,JRELI=SK.A(ICOMJREL)÷ELP(ICOr4I+FOP(ICOM)863* IF(ILOA.EO.lI}TAPVA(ICOMjREL)=BTAPVA(ICO,4,JNEL)+AMDAPA(JREL)804* DO l0•9 JCOM=I.N8Bo* 1C69 II{ICOMICOM)=IIl1COM.ICOM)-G(ICOM.JCOM)/V(ICOM)VII-235

80b* l0u7 H(CICO~1,ICOM)=H(ICOM.ICOM)-BTAPVA(ICOMJREL)887* CALL ALCEMYdOd*00 1066 ICOM=M1N889* CVLKP=CVLKP÷FT(ICOM*N,I)890* 108 CAP(ICOMJREL,2Z=FT(ICOMel)891* 1065 CONTINUE892* C PHASE A VAPORIZATION RELEASE PARTICULATES COMPLETE FOR THIS TIME STEP89.5* &0 TU 53894* 1071 CUtTItNUE VAPREL H893* C IF VAPORIZATION RELEASE PHASE R HAS NOT STARTED. NYPASS890* IF(TT.LT.TVR2)GO To 50897* 00 1075 JqEL=1O1089%* IF(TT.LT.TV0(jREL))GO TO 1075899* T(1)=AMINIDITTT-TVB(JREL})9UO*DO 1U77 ICOM=IN9014 F(ICOMICdP(ICOMJREL.,)902* IF(ICOM.EO.1.AND.OKICS(JREL).NE.O.IF(ICOM)=OKICB(JREL)903* F(ICOM+NI=O.904* H(ICOMICOM)=O.905* 8TAPV8(ICOM,JNEL):SKG6(ICOM.JREL)+ELP(ICOI4)*FDP(ICOM)906* IF(ICOM.EJ.1)UTAPVB(ICOi4lJREL)=BTAPVB(I•OMJREL)+AMDAPB(JREL)907* UO 1079 JCOM=IN906* 1079 H(ICOr4,1COM)=H(ICOM,ICOMI-GIICOM.JCOM)/V(ICOM)909* 1077 H(ICOi4,1COM)=H(ICOM, ICOM)-BTAPVB(ICOM.JREL)910* CALL ALCEAY911* 00 1070 ICOMNIN912. CVLKP=CVLKP+Fr(ICOM+N,I)913* 107d CdPIICO.A.JREL,2)=FT(ICOMtI)91j* 1075 CONTINUE915* C PHASL 8 VAPORIZATION RELEASE COMPLETED FOR THIS TIME STEP.916s 50 CONTINUE917* 60 TO 190918. 60 CONTINUE ORGANICI919* C SET OFF DIAGONAL MATRIX ELEMENTS FOR ORGANIC IODIDE BRANCH920* 00 600 ICOM=,pN921* ýl(ICOM.N, ICOf4)=ELOI(ICJMJ922* D0 600 JCOM=I,N923* 600 H(JCUMICOM)=b(ICOM.JCOM)/V(ICOM)94* DO S0 IREL=I.5925* GO TU(bl,71,81,91,101),IREL92u* C FIVE WAY BRANCH 15 TO EXPLOSION, GAR., MELTv PHASE A VAPORIZATION,927* C ANO PHASE B VAPORIZATION928* 61 CONTINUE EXPLODE929* C IF STEAM EXPLOSION HAS NOT OCCURRED, BYPASS930* IF(Tr.LT.TEXI)GO TO 80931* T(I)=AMINI(DTTT•1EXI)932* 00 67 ICOM=lNr933* rHICO4,ICuM)=u.9.4. F(ICO•)=CSOIIICOM,1)935* F(ICO.4+N)=O.930* IF(ICOM.Eu.I.AND.UKICS.NE.O.)F(ICOI)=OKICS937* 00 69 UCO(=l.,i9383* o 9 H(ICO,4,ICOMI)H(ICUNM,ICOH)-G(ICOM.JCOM)/V(ICOM)939. 07 H(ICOM,ICOMI)=(ICONICOM)-ELOI(ICOM)940* CALL ALCEMY941* LU 68 ICOM=1,IJ-942. CSLKOI=CSLKOI+FT(ICOM*!4,1)943* o8 CSOI(ICOM,2)ZFT(ICOM.1)9414* C ORGANIC IODIDE EXPLOSION RELEASE COMPLETEO FOR THIS TIME STEP945* GO TO 8094t*947.71 CONTINUET(1)=DlGAP REL948* DO 77 ICO1.Mi949* i(ICOm,'ICoM)=p.950* F(ICOp)DCGOI(ICOMI)951* F(ICUM+N)=O.952* IF(ICOM.EQ.2.ANO.M.E8.IJF(ICOM)1l.953* UD 79 JCO:41,N954- 79 HtICU.,IlCQ )=HIICOMtICOMI-G(ICOMJCOMI/V(ICOM)9-5* 77 IIIICOM,ICOM}:H{ICOM,ICOM)-ELOI(ICOM)95o* CALL ALCEi4Y957* DO 78 ICO'4=1,N9bd*C5LKOI=CGLKOI+FTtICOM+NI)959* 78 CGOIIICONP2)=FT(ICOM,I)9o0* C ORGANIC IODIUE RELEASE DONE FOR THIS TIME STEP9b1* GO TO 80962* dW CONTINUE MELT REL963* C IF MELT RELEASE HAS NOT STARTED. BYPASS964* IF(TT.LT.TMR)GO TO 809b5* 00 85 JRELI1O109u•* IF(TT.LT.TM4(JREL))0O TO 8597- T(I1)=AMINIIOT,TT-TM4(JREL))908* 00 87 ICOM=I.N969* HIICOMICOM)=O.97U* F(ICOA)=CAOI(ICOM,JREL,1)971* F(ICOA+N)=O.97.* IF(ICO9.EO.2.AND.uKICi(JHEL).ihE.O.)F(ICOM)=OKICM(JREL)973* U0 89 JCOI-:1.1974*97b*9 titlCO.l,ICOM)I=ICICO9,IC04)-GtICOM,JCOM4/V(ICOM)o7 ii(ICO,•,.ICOA)=Ii(ICOM, ICOM)-ELOI(ICO.')97b* CALL ALCEMY977* Ou lid ICO=IN978* CMLKOI=CALKOI+FT(ICOM+N.I)979* ob CHIOIKICO\,JREL,2)=FT(ICOM,1)VII-236

980* o5 CONTINUE91* C ORGANIC IOLIUL MELT RELEASE COMPLETED FOR THIS TIME STEP902* GO TO do9b5* 91 CONTIJUE VAPREL A964* C IF VAPOR RELEASE HAS NOT STARTED. BYPASS965* IF(TT.LT.TVRIfGO TO 8;9 1o* OU 95 JREL=IU9d7* IF(Ti.,T. rA4(JRELJ)6O ro 9598o* T(1)=A4IN1(DTTT-TA4(JREL))9o9* JJ 97 ICO=I1.pN990* H(ICO!4, ICOM)=1.991* F(ICOA)=CAOI(ICOM.JRELI)992* F(ICOM+N)=O.993. IF(ICO9. E-.1AIJD.OKICA(JRLL).IIE.O.)F(ICOM)=OKICA(JRELI994* Do 99 JCOM=1.N990* 99 HIICOMICOM)=H(ICOMICOM)-G(ICOM.JCOM)/V(ICOM)990* 97 f(ICUý,ICOM)Ii(ICOM.ICOM)-ELOI{ICOM)997* CALL ALCEM1Y998* UJ 98 ICOq=I,rj999. CVLKUI=CVLKOI+FT(ICOM+N.1)1000* 98 CAOI(ICOMJREL,2)=FT(ICOM.1)1001* 95 CONTINUE1002* C ORGANIC IODIDE VAPOR RELEASE PHASE A COMPLETED FOR THIS TIME STEP1003* Go TO ol1004* 1,11 CONTI1UE VAPREL ii0It0* C IF PHASE U VAPOR RELEASE IAS NOT STARTED. BYPASS1006* IF(TT.LT.TVR2)GO TO 801007* 00 105 JREL=1.10109d* IF(TT.LT.TB4(JRELI)GO TO 1051009* T(1)=AAIN1I(DTTT-0B4(jREL))1010* 00 107 ICOm=IN1u11* H(ICOMI,ICOM)=O.101* F(ICOM)=C;iOIiICOM,JREL.I)101,* F(ICO,4N)=O.1014* IFCICO'4.EO.I.AND.OKICB(JREL).NE.O.)F(ICOMi=OKICB(OREL)1015* DO 109 JCOM=1.N1016* 199 H(ICOiIGOM)=H(ICOMICOM)-G(ICOMJCOM)/V(ICOM)1017* .107 H(ICOM. ICOM)=HIICOM.ICOMI-ELOI(ICOM)1018* CALL ALCEMY1019* 0 10a ICOM=E,N1020* CVLKOI=CVLKOI+FT(ICOM+I4.I)1021* 190 C3O3(IC04,.JREL.2)zFT(ICOMI)1022* 105 CONT INUE1025* C ORGANIC IODIL VAPOR RELEASE PHASE B COMPLETED FOR THIS TIME STEP.104* do CONTINUE1025* GO TO 1901026* 120 CONTINUE 12URANCH1027* C NATURAL CONVECTION LAMBDA 4ILL BE ASSJMED ZERO IN ANY COMPART-102d* C MENT .HEHL CI2/V.LT.CUTOV/VTOT.1029* 00 1037 I=1,Niu3U*KSPRAY=I1051* IF(I.EG.I.AHOD.(TT.GT.TCSRI.ArJD.TT.LT.TCSNIE).OR.(TT.GT.TCSR2.AND1032* *.rT.LT.TCSH2EI.OR.(TT.GT.TCSI.AND.TT.LT.TCSIE)))KSPRAY=I1035* OKNCG(I)=FLOAT(t-KSPRAy)lUJý*OKfNCS(I)=FLOA1(I-KSPRAY)1031* IF(CSI2(I,I)/V(I).LT.CLJTOV/VTOT)OKNCS(I)=U.0.o6* IF(CGi2(II)/V(I).LI.CJTOV/VTOT)OKi'CG(I)=u.1037* DO 1037 J=1,101038* SCALE=.07528*LXP(-.l*FLOAT(J-1))1039* OKNCAIIJ)=FLDAT(I-KSPRAY)104•*UKWCb(I,J)=FLuAT(I-KSPRAY)1041* UKNCM(I1,J)=FLOAT(I-KSPRAY)104.* IF(CAI2(I,J.1,/V(I)/SCALE.LT.CUTOV/VTOT)OKNCA(I,J)=O.1043* IF(Cbl2(I.JIO/V(I)/SCALE.LT.CNTOV/VTOT)OKNC8(I,J)=O.1044* I-(CEMIlI IJ,1I/V(I)/.I.LT.CUIOV/VTOT)OKNCH(I.{J)=O.L04!* iC57 CONTINUEIO4b* C SET OFF fIAGOI4AL MIATRIX ELEMENTS FOR 12 bRANCH1U47* DO l0o ICOM=IN104d* H(ICOM+e, ICOM)=ELI2(ICON)104Y9* UO 1200 JCOM=IN1050* 1200 H(JCOMICOM)=(GICOM.JCOM)*(I.-EFI2(ICOMJCOM))/V(ICOMI1051* DO 170 IREL:1.51052* GU TU(121,131,141,151,161),IREL1053* C FIVE 4AY dRANCII IS TO GAP RELEASESTEAM EXPLOSIONMELT RELEASE,1054* C VAPORI2ATION PHASE A, AND VAPORIZATION PHASE B.10b* 121 CONTINUE GAP REL1056* T(1)=DT1057* DO 127 ICOM=IN1050* H(ICO;.1,ICOm)=n.1059* F(ICUjA)=CGI2(ICOM,1)1060* F(ICOM+N)=O.1061* IF(M.EG.1)F(ICON)ZDKRON(ICOM,2)1b0* BTAI2G(ICON)=ELI2(ICO9,*FDI2(ICOM)+OKNCG(ICOM)*CKG(ICOM)*.AW(ICOM1063* *)+AF(ICOM))/V(ICOM)1ub1*1065* IF(ICoM.Eq.1)uTAI1G(ICOM)=BTAI2G(ICOM)+AMDAIG+ELAMG1i6)*DO 129 JCOM=I,4N1007* 129 H(ICO,4,1COM)=H(ICOMICOM)-G(ICOMJCOM)/V(ICOM)1068* 127 H(ICOQ;4ICOM)=Il(ICOR,ICOM)-NTAI2G(ICOM)1009. CALL ALCEMY1070* 00 120 ICOM=IN1071* CGLKI2=CGLKi2+FT(ICO.÷iN,1)lul2* 12s C612AICO'4,2)FFT(ICOM, I)Vii-237

1075* C 12 GAP RELEASE COMPLETED FOR THIS TIME STEP1074* GO TO 1701075* 131 CONTINUE ST EXPLO107b* C IF STCAM EXPLOSIOtN HAS NOT OCCURREO.BYPASS1077* IF(TT.LT.TEXI)GO TO 1701U07* T(I)=AMINI(OTTT-TEX1)1079* O0 137 ICOMr=1,1000, H(ICOA,ICOM)=0.o1001 F(ICOM)=CSI2(ICON,i)lObi*FElCOM+N)=O.10o.*IF(OKICS.GT.O.)F(CCOM)=OKROtE(lCOMI)10d4* dTAL2S(ICui4)=LL12(ICOA)4FD12(ICOM)1885. IF(OKIECS(ICOM).NE.O.)dTAI2S(ICOM)=BTA12SE1COA1)+CKG(ICOM)lo0s* 1-(A4(ICOM)+AF(ICOM))/V(ICOM)jOdT*IF(ICOM.EO.1i)TAI2S(LCOM)=BTA12S(ICOM)+AM0AIS÷ELAMS1018* O0 139 JCOM=1.N10419. 139 H(ICOH,ICOM:)ZI(ICOMICOM)-G(ICOM.JCOM)/V(ICOM)1090. 107 r(ICOlrlICOM)=HIICUM, ICOM)-HTAI2S(ICOM)1091* CALL ALCEMY109.* DO 1.5 ICOM=1.NI093* CSLKI2=CSLKI2+FT(ICOM÷N+11)1094-* i3) CS12(1COM,2)=FTtICOM,I)C EXPLOSION CO;4PLETED FOR THIS TIME STEP1095. 12 STEAA109i)- 00 TO 1701097* 141 C3wTINUE040y. C IF MELT RELEASE HAS NOT STARTEDAYPASSMELT REL1099*1100*IF(TT.LT.TMP)GO TO 170Do 145 JREL=1,101101* IF(TT.LT.rM4EJREL))bO TO 14511U2* l(1)=AMII41(OTTT-TM4(JREL))1103. UO 147 ICOMI.I41104* H(ICO.4,ICOM)=fl.1111t* F(ICOM)C%1I2(ECOMJREL.1)11004 FEICOM+N)EU.1107* IF(OKICM(JREL).GT.O.)F(ICOM)=OKRON(ICOM.2)*OKICM(JREL)11003 UTAI24(ICOM,JREL)=ELI2(ICOM)*FD12(ICOM)÷OKNCM(ICOM,JREL)*11119* I(CKG(ICOM)*(AN(ICOM)*AF(ICOM))/V(ICOM))1110* IF(ICOM.EQ.IBTA12M(ICOM.JREL)=BTA12MEICOMJREL)÷AMDAIM(JREL)+1111* *ELAMMIJREL)1112* U0 149 JC0M=I.N111ý* 149 1E(ICOMICOM)=H(ICOM.ICOM)-G(ICOMJCOM)/V(ICOM)1114* 147 H(ICOM. ICOMIZ=iE(ICOM. ICOM)-BTAI2M(ICOMJREL)111b* CALL ALCEMY1116* 00 148 ICOM=1011117* CHLKL2zCMLKI2+FT(ICOR+t4,1)111* 148 CMI2(ICOM.JHEL,2)=FT(ICOMI)1119* 14! CONTINUE11?0* C 12 MELT RELEASL COMPLETED FOR TIllS TIME STEP1121* GU TO 1701122* 151 CONTINUE VAP REL1123* C IF VAPOR RELEASE tlAS NOT STARTED.hYPASS11I4* IF(TT.LT.TVR1)GO TO 170II•5.O00 155 JREL1.1012o*. IF(TT.LT.TA4(JREL))GO TO 1551127* T(1)=AMINI(D1.TT-TA4(JREL))1128* 00 157 IC-M=Lf41129* jE(ICu0, ICOM)=n.1150* FIICOM))CAI211CONJREL.1)1131* F(ICO.A+N)=U.113?* IF(ONICA(JREL,.GT.O.)F(ICOM)=DERON(ICOM1I)*OKICA(JREL)113.* HTAI2A(ICOM.,JiEL)=ELI2(ICGE)+FUI2(ICOM)+OKNCA(ICOM,JRiEL)*1154* I(CKG(ICOM)*(A*(ICOM)÷AF(ILOM))/V(ICOM))1135* IF(ICoM.EU.1)BTAI2A(ICUM,JREL)=RTAI2A(ICOMJREL)+AMDAIA(JREL)11536 I+ELAMA(JREL)1137* ;)U 159 JCOm=lR I1166* 159 ii(ICO.4,ICOR)=tl(ICOM,ICOM)-G(ICOM,JCOM)/V(ICON)1139* 157 ,( I .CO. I I)=I,(ICOM. ICOM) -BTAI2A(ICOM, JREL)11ts*CALL ALCEIY1141* O0 150 ICO4=1,N1142* CVLKI2=CVLK12+FT(ICOM4+lJ,)i143* l118 CA12(ICOM.JRLL2)=FT(ICOM.1)1144* lbb EONTiNUE114b* L 12 VAPORIZATION( KELEASE PHASE A COAIPLETEE FOR THIS TIME STEP114b* W) TO 1701147* i10 CUItTINUEli+b* C IF PHASE 6 VAPOR RELEASE HEAS 10T STARTED. BYPASSVAPREL1149*11•0-iFtTT.LT.TVR2)GO TO 170Uo lu5 JREL=1,1O11j1* IF(TT.LT.rb4(ORELE)GO TO 16b11524 T(I)="IrNI(DTTT-Tb4(JREL))1153* UO Io7 ICOM=1,NLlZ)*H(ICU,4,ICuM)fl.l155* F(ICOM)=C8I2(ICOMJREL.1)limb*F(ICO,4+N)=O.l1117* IF(OKLICB(JREL).GT.O.)F(ICOM)=DKRON(ICOM,1)*OKICB(uREL)1158* 3TAI1U(ICOMJREL):ELI2EICOMS÷FDI2(ICOM)+OKNCBIICOM.JREL)*1159* *CKG(ICOM)*(AW(ICOM)+AF(ICOM))/V(ICOM)11.U*IF(ICOM.EO.I)iITAI28iICOMJREL)=BTA12B(ICOM.JREL)÷1161' *Av.IAIb(oJREL)*ELAMO(JREL)110o* UO 169 JCOM=I.EJ11o05. 19 litICOAElICOM):H(ICOM. ICOM)-G(ICOMJCOM)/V( ICOM)1114* Ib7 EII(CO., ICOM)=iE(ICOM.ICOM)-BTAI2B(ICOM.JREL)llb5*CALL ALCEMYVII-238

lloo* DO 168 ICOM:IN1167* CVLKI2=CVLKI2+FT(ICOM+1,1)1168s 1b8 CdI2(ICOM,JREL.2)=FT(ICOM,.)1169* lot CONTINUE1170* C 12 VAPORIZATION1 RELEASE PHASE B COMPLETED FOR TRIS TIME STEP1171* 170 CONTINUE117Z* 190 CONTINUE117.* IF(ABS(TT-TPUFF).GT.3.E-8)GO TO 200001174* C MULTIPLY ALL AIRBORNE FRACTIONS EXCEPT THOSE FROM STEAM EXVLOSIONS BY1175* C APUFF. I14CREASE LEAKED FRACTION FOR THESE BY '(I.-XPUFF)*AIRUORNEL11T* C FRACTION PRESE14T PRIOR TO PUFF.1177. XPC=1.-XPUFF1178* XPCS=XPC/XPUFF1179. iO 21000 ICOM=INlldO*CGLKP=CGLKP+XPC*CGP(ICOM,2)/UFPP1181* CCP(ICOM,2)=AiUFF*CGP(ICOM,2)1184* CSLKP=CSLKP+XPCS*CSP(ICOM,2)/DFPP11615. CGLKOI=CGLKOI+XPC*CGOI(ICOM,2)/DFPOI11I4* CGOI(ICOM.2)=XPUFF*CGOI(ICOM,2)11ti.* CSLKOI=CSLKOI+XPCS*CSOI(ICOM,2)/DFPOI11ob*CLKI2=CGLX12+XPC*CG12(ICOM,21/OFPI21167* CG12(ICOM,2)=XPUFF*CG12(ICOM,2)1188* CSLKI2=CSLKI2+XPCS*CSI2(ICOM,2)/UFPI21109. 03 21000 JREL=I.IO1190* CMLKP=CMLKP+XPC*CMP(ICOMJREL,2)/DFPP1191* CHP(ICOM,JREL.2)=XPUFF*CMP(ICOM.JREL,2)1192* CVLKP=CVLKP*XPC*(CAP(ICOM,JREL.21+CBP(ICOM.JREL.2)/)OFPP119.5* CAP.(ICOMJREL,2)=XPUFF*CAP(ICOM.,JREL,2)1194* CLiP(ICOM,JREL,2)=xPUFF*CBP(ICOM,JNEL,2)1195* CiLKOI=CNLKOI+XPC*CMOIfICOMJREL,2)/DFPOI1196* CMOI(ICOMJREL,2)=XPUFF*CMOI(ICOMJREL,2)1197* CVLKOI=CVLKOI+XPC*(CAOI(ICOMJREL.2)÷CBOI(ICOM.JREL.2))/DFPOI119d* CAOI(ICOM,JRLL,2)=XPUFF*CAOI(ICOMJREL,2)1199* CBOI(ICOM.JREL,2)=XPUFF*CcOI(ICOM,JREL,2)1200* LRLKI2=CMLKI2+XPC*CMI2(ICOMJREL,2)/OFPI21201* CMI2(ICOM.JREL.2)=XPUFF*CMI2 ICOMJREL.2)120,2* CVLKI2=CVLKI2+XPC*(CA12(ICOM#JRELe2)+CBI2(ICOMeJREL.2))/DFPI21203. CAI2(ICOMJREL,2)=XPUFF*CAI2(ICOM,JREL.2)1204sC612(ICOMJREL,2)=XPUFF*CBI2(ICOMJREL,2)1205* 21000 CONTINUE1206* 20C00 CONTINUE1207. C TOTAL VAPORIZATION.AND MELT RELEASE COMPONENTS PER COMPARTMENT1206* 00 450 I1IN1209* SUMA=O.1210* SUMB=O.1211* SUMC=O.121z* SUMDZO.1213* SUME=O.1214* SUMHZO.1215* DO 44 J=1.1O1216* SUMA=SUMA+CMP(I.J.2)1217* SUHM=SUMR4CMI2(I.J,2)1216* SUMC=SUMC+CMOI(I.J,2)1219* SUMD=SUMU÷CA12(I,J,2)+CRI2(I1J,2)1220. SUME=SUME÷CAOifI.J,2)+CBOI(I,J,2)1221* SOMI=SIJMH*CAP(I.J,2)÷CBP(IJp2)1222* 44 CONTINUE1223* CMTP(I)=SUMA1224* CMTI2(I)=SUMB1220.i 2;1b*CNITOI(I)=SUMCCVT0I(I)=SI.IME1227* CVTI2(l)=SUMD122b* CITP(I):SUMH12z9* *50 CONTINUE12.0. CfMRP=O.12.1) CrSRPZU.1232* CfGRPz('.12.5:* CTMRI2=O.12.54* CTSRI2=O.1235* CTGRL2=0.1236* CTVRL2=O.1237* CIMROI=3.,L238*CCSROI=O.12-39* CTGROI=O.1240* CTVRUI=3.1241* CTVRP=U.1242. 00 46 I=1,N124.)* CTMRP=CTMNP+CMTP(1)1244* CTSRP=CTSRP+CSP(I.2)1245. CIGRP=CTGRP+CGP(I#2)1246* CTSROI=CTSROI.CSOI(I.2)1247* C1GROI=CTGROI+CGOI(I.,)1248* CIMROI=CTMROI+CMTOI(I)1249* CrVROI=CTVROI+CVTOI(1)1250O* CTVRI2=CTVRI2+CVTI2(1)1251* CTSRI2=CTSRI2+CS12(IP2)1252* CTGRI2=CT&RI2+CGI2(I,2)120.5* CTMRI2=CTMRI2÷CMT12(I)12540* CTVRP=CTVRP+CVTP(I)125b* 46 CONTINUE125o* c AIRBORNE FRACTIONSL(TOTAL) COMPUTED ABOVE1257* C COMPUTE AMOUNTS LEAKED125d* C COMPUTE DOSE REDUCTION FACTORSVII-239

1259eD0FMI2=0.12oO*DAFSI2=0.121,1 DHFNP=O.120fURFdI2=0.12630 ORFGI2=O.12o4* DOFGP=G."12"#* DRFSP=O.12b0* IF(CMLKP.GT.O.) OHFMP=CMLKOI/CMLKP12s7TLF(CSLKP.GT.O.) ORFSP=CSLKOI/CSLKP12b01* IF(CGLKP.GT.9.) DRFGP=CGLKOI/CGLKP1209* IF(CMLKI2.GT.O.) ORFMI2=CMLKOI/CMLKI21270# IF(CSLKI2.bT.O.) ORFSI2=CSLKOI/CSLKI21271. IF(COLKI2.GT.0.) DRFGI2=CGLKOI/CGLKI21272. IF(CVLKI2.OT.t'.) URFVI2=CVLKOI/CVLKI2127i$ IF(CVL.P.GT.l.)DRFVP=CVLKOI/CVLKP1274* M=M+1127bT C IF YOU WAiT TO BYPASS WRITING, PUT A GO TO 998 HERE.127b* WRITE(cr30UCJTT1277. 3000 FORMAT(1HI,/bX,S5HTIME.PIPE11.5,3X.3tIHRS)127ooWRITElb.3UO1)1279s WRITE(o.30f2)12d03 iC91 FUR,.AT(IHO,5X,4OHCOMPARTMENT AIRBORNE FRACTIONS CONTAINED ./)12b1$ 3C02 FURMAT(lXp1lHC,'11H GAP 11lH GAP .1111 GAP128zv 111H MELT 1111 MELT .11H MELT .1H EXPLOSION1263$ 211H EXPLOSION .111i EXPLOSION .l1 VAPOR oI0H VAPOR .101 VA12b4* 3PUR ./,IX,klU.llH RELEASE .11H RELEASE .11H RELEASE126b* 41111 RELEASE .11H RELEASE ,11H RELEASE .11H RELEASE12b.# 51114 RELEASE .111, RELEASE .11H RELEASE .10H RELEASE12070 61011 RELEASE ,/,1X,1IIMH113l PARTICLES #11H 12120b* 71111 011239* 71111 PAhTICLES ,1111 12 .11H 01 11H PARTICLES v1290* 811H 12 ,11H 01 .11i PARTICLES IOH 121291$ 91011 01 ./)1292* WN•ITLEu.3003)(IeCGP(I.2).CGI2(I,2)PCGOI(I,2}hCMTP(I)eCMTI2(1).129.* 1C:TOII.),CSP(1,2),CSI2II.2)eCSOICI.2).CVTP(I)*CVJI2(I)PCVTOI(I)),1294* 21=ltt4)1295$ 3004 FOIRMAT(12-1PEIO.4,IPIOE11.4.LPEIO.4)12960 WkITE(b.3004)1297* 3CC4 FORMAT4UHU,//.5X.54HTOTAL AIRBORNE FRACtIONS CONTAINED P/)129d* wRITL(o,3007)1299kWRITLio.306)CTGRP.CTGRI2.CTGROI.CTMRP.CTNRI2.CTMROICTSRP,1300$ ICTSR12,CTSROI.CTVRP.CTVRI2,CTVROI1301* wIRITE(6.3005)1302$ 3005 FORMATI1IIO./,SX.32HESCAPE FRACTIONS OF EACH RELEASE ./11306* WRITZ(b#3007)1304* WRITEfb,3O0bICGLKP.CGLKI2.CGLKOI#CNLKP.CMLKI2.CMLKOI.CSLKPv130b* ICSLKI2.CSLKOI.CVLKP.CVLKI2.CVLKOI1306$ 3096 FORMAT(2X,1PEIO•.41PLOE11.4.IPElO.4)1307$ 3007 FURMAT(IX.iH 114H GAP .11H GAP .l1H GAP p130* 11114 MELT .11|t MELT .11H MELT .11H EXPLOSION ,1309$ 211H EXPLOSION #11H EAPLOSION .11H VAPOR .1OH VAPOR1310$ 310H VAPOR */,IXtlH .11H RELEASE .11H RELEASE @1311$ 411H RELEASE p1312* 411H RELEASE 11.H RELEASE .11H RELEASE .11H RELEASE *131.1 511H RELEASE ,11H RELEASE *1111 RELEASE 10OH RELEASE .I1H REL1314* bEASE ,/,1X.ll .1111 PARTICLES 114H 12 .111H 01131tý* 71111 PARTICLES 114H 12 .11H 01 PI1H PARTICLES o131o* 111i 12 .11H 01 .11H PARTICLES 10OH 121317$ 9,3OH 01 P/I131d* WRITL(n,3010)1319$ 3V10I FORNAT(lH0./p5X,3c11DOSE REDUCTION FACTORS OF EACH RELEASE ./)1320$ WRITE(6.3007)1321$ WI(ITE(6,3015)IHFGPDRFGI2.ORFMPDRFMI2.DkFSPDRFSI2,DRFVPDRFVI21322* 3C15 FURMAr(2X.IPEIO.4.iPEII.4.llX.1P2ELI.4*.IX.1P2E11.4.I1X.1323$ 1IPE11.4,IPE19.4)13ý4* C1325* C PUT NEW VALUES INTO A0. 1 BUCKETS.13•o*DO 9780 J=.1N13•7*CUPIJv1)=CGP(Jp2)1328* CG!2(1J.l=CGI2(J.21329* COOI1J.l)=CGOi(J#2)1330$ IFtTT.LT.TEXI)GO TO 97751331$ CSP(Jp.)=CSP(J,2113324 CS12(oe1)=CS12(J.2)13.33 CSOI(J.1)=CSOI(J.2)1334* 9775 CONTINUE13:b4 00 9780 K=1.1..13J6* IF(TT.LT.TM4(K))G0 TO 97781367* CMP(.,•KI)=CMP(JPK2)13381 CMI2(J.K.1)=CMI2hJKI e)1339$ C.AOI(jpK.1)=CMO1(Cj#K.2)1340* 9778 IF(TT.L7.tA4(K))GU TO 97801341* CAP(.JK, )=CAP(J.Ko2)1342* CAI2(J.K.1)=CAI2(J#K.1)134ý* CAOI(JKPL)=CAOI(J.KP2)1344* IF(TT.LT.Te4(K))6O TO 97801345* CbP(iJKeL):CBP(J#K,2)1346$ CoI2(JvK.1)=CUI2(1J.K2)13474 C*OIij.Ksl)=CLOI(J.K,2I1346* 9760 CONTINUE1349* C UEFINE 3 rYPES OF RADIONUCLIDE DEPOSITION AND TRANSPORT TYPES.1350b* C ORGANIC IODIDE LIKE, IODINE LIKE. AND PARTICULATE LIKE1351. 00 6001 I=,NOIVII-240

132a* bOl CORLEK( )=CGLKOI*CF I II)4CNLKOI*CFR 2pt)+CS"0O*CFR(3o )XPUFF13b3* I+CVLI.OI*CFR(4#. I)1354* NOIPIZI401+113bb.NU2ENDONOI+14X213Db*.435?*00 6602 I=NOIPI.N12Ek)b602 CORLLK ( I =CGLK12*CFR( I P 11 +CMA12*CJR(2e -CSLK12*CFRt 3o 1) *XPUFF13btl* 1÷CVLK12*CFR(4oI)13b9sNSTAHT=N12Etfl+11,3bO*13bl*NISO=NOIO.N12+t4pAR00 6603 Z=NSTART.NISO1362Z- 6o,13 CURLEK ( IC) ZCQLP*CFR 1. Ill +CLKP*CFR(2.*11CSL.KP*CVH3P I *XPUFF130.4* 1 +CVL•,P*CFR (4 1 )lab"* WHITEW3H 32)2 AMACSIZ,3P1 =NIEOI1300 3('32 F;Hik4AT(Il10vbxe3%KFRACTION5. OF CORE INVENT'IY LEAKEDei/,9Xr10(A6P5X131)0# !) )13u7s WATP g 3r• 34 ItCORLEK I I) u -IINISOIa36s J30AC FORMAT(MH .bX#1,OQPE1O.A.IX1I13b94 1F(ITT+1.E-8l.LT.TENOD)O TO 6 *NEWjL370* CALL LRR(3e18i4CO0RAL S.N. 3034 -1371. GO TO 11,572*137301999 CQ4TI!4UESTIOP137qeEN.OJ8. SUBROUTINES IN CORRALThe followiqg listings are suoroutiues used in CORRAL Or in thesubroutines. These were programmed for the UNIVAC 1108, and mightnot be universally adaptable to other macbhies.Sr(CHECK FOR DIVISION BY ZERO OR OVERFLOW)Or- roýATTfN -ATC *~ TTMý 1117?~)ni-i-n I týT(MA*JYI-ALL O)VýQcl (m"7-4I IJNIVAr>',FT GonhrII I*,n TO• Inl, •r T TN11o.. I1ECK FORK IVISION BY ZERO OR OVERFLOW'I" .OTLI., ' nlvln•lnahr I V IT RUN RAn. nIA1-40AF "Q0'PLFS Orc 1*ATTIM -. ATe - " J" j0 . lIv 1 .7.LnTA l TyPer / r*.6ALI OVIUQFI tM'ICH) UNIVAC IYEANOP' ill M111'4 (41-1. Ii NTVAr 1vi'kNe'.II. i .. 7?F:(INO.M•T.e|N9 Lnccn:• M[VFRGFMnrINrHP+N, IIL+LnrC,.FrPA I'OI Tn 2nrl GO0'LMCT=•MAXtfh.STL0PeT.+.1 1NO ERASET1rrI~nT.r'r.i .T;T) An Tn fnl TrO RAnn• PF y O r Mnifm• V O r NI LLP IY + NTT ,L. INK1*'( T7 ,AiMA (T2,1i.t I- ~III~~ ~~VPC/IIYA) 7)~p~e 1 Tyurgfl.)1 LAOI.hf~lnAO1 tADojn^^) mrt~mncerr - (GENERATES Zri(e•--1 WHERE Z IS A SQUARE MATRIX)C.'aQn-11ry 4 V FAP.7Mr Or;l= COVATTnt IATV - ? JIjl 7'. T)W 1g .17.7+*.Mr.O IhI;.*- IP)9.I0M 1.TP1Ap . l.,'),• . r• iIm i QL..Y4O-L"..+ ,•,,• 4 • 9"N r-n RT.A , S 7 T4anF.) © . • TAG, pT •t)TYP0V11-241

- ~ ~ ,JPh'- wT~w cArT{nRAL(N1 mVC vOV(Npl((.b-hAX RýLtýA A-S-iAIS RI'LRI NA NAl T1 1T1- A(N).I4-1.NA ROU.,N=1.NAwV0O119nFT~li k4&VTO O'n.FO CCICC AniO~Onr vTr -1-nY m IRnnFAirO fl.ArN V.,TN7VV INO' ,~ TN. I.Im &n m AI NAQ" !R-.? PAC 10T UO %FRIIFc tITFRM 7N ANr) -.CPlvi SLIM FABZTOl FA)(% i l.1)'1 1.1r.n TA I A(* CCT 1,1 FRCOn MOACtIIDPr F"IRA NcXT TIPRM. (JIM CRC.MrA-;IIRF CHRORnn 14n TIýi~ttnnI 14InI 12-.11 /Al 2.1I'1[Y-AmA~j I DLF -DN)110 ON.&AFISZ4(T7.111lINr Ic fl-'VF- cEZ!FC IJIr1FY 1CIIN'1-INNP) CMNVFQCFD* CANvVFRCD, OIVFPGFD- :~1-MAT~r CLLTOADM r-CK -OR ';PILL ANn PIVSTON QY ZPRO*ýr(ALl-ckv/ PO (SOLVES SET OF SIMULTANEOUS 1ST ORDER DIFFERENTIAL EQUATIONS)'I,'-Qr)UTI* aLj-um47 TIUc-VAq1ITT IjZOTOPIC TRANqMIITALTTON IINW PROGRAM ALCIHFMY-1! (CHOP)P- DC CTI~fI 4 ITF - 7 JUN 714. TIMO 13.17.2810*QLF 0!crtc.?CIj OCYDcr nAT& PnwcR2/'( 3 7740r10rI910f/.TCP/88.n296919/ UINIV.ACU)ATA POWFQ2 ,,7764e*(flfl Torl'IonnItI, TOP /74.1.6674632/C tV MIATRIX HYDCRr.POMvTRTC FUNCTION Tn $:X-ONFNTAL INTFGRA1 (F*'21I)/ZMA V%I Fm-cR TIMF Lfy)P. -T09c TII4v ?NmC NY? IN COMMONUNTT=Mje IFIT).=c(.1) + Ftn*4H*IVC*@T*H-1)/(T*4N1)*T17 ccr 13P lCCcRen cTcLD AerC~vIAL IN Anl In "M1.I NTFR LAPLAiC TRANeffOR( LOOP %TORF LAPLACF T'4rwFX L IN COMMONVII-242

LL.L=LL+lmO 220 L1.-1,LLLI-IAF3SIL1-1C ASSFMPLI7 T*fH-'.) IN 7 WITH ~IIN OF SGOIARES OF FLEMENT5S CALED UNDER ONEno I 11-1.170 2021!70 1?'? 1 -W7L,+1)~ralr v RYpeWcQs MP TWn. 9nIn R-Q+7(?2.III**2TM.! (NJ )R.ARS(TN4lORT PR))FAST IIPJSCALTNfl40 P-bNnt0D,PmWPQ21TN.TN/Qsnno 6" 17-.1?* VMMAYT7TX IJv0FD090MMTPIf qg:?fg~l IF**Z-1 1/7 YN FANZrALL PAVZm* AS-M~ EPONFNTTALLY-SCAL:Fn SWIPCF FIFLO Q*(F**Z-1)fl IN ZNI! BUFFEROn 7t% 12-1.11ZN? 1(12)-A.n0 7M r1-i,?1Tr 'ZrAL~n, ASSFUNIF UM~CALr 4rcmeQa?0Q mqrR~c INzI7r..-?7 ZNIV) 9A 72-1.11700 'MN!172.1..ZNl T1,TJ Iq(1,1*ZT*3C UNS,7ALc LOOP *I)0 ml 117I T7s1,1,In 17n l T11,?f7 -;T UP~ NrXT PASS170 7(17,!' -ZN4T2.I11r unf)?cY ANn TRANSFvQ RIJPFPRpn cUM?nn 130 12=1#IT130 ZN11(I2)-.5*(AlI2)+ZNTI(12)1llu)c'76Lc LofV) vX!T- mIAflc U'.4S-.tiNr- C.W CAT0R MATRIX744111, 1111,o.no 130T.1?CONT!NUF UNSrALTNGr-o To 11In- ~ rMS~4LF r.F**f*W(LI1~n cTW=-WL*T(Nt17' vAUZ~o PrTPiRNS SORTI' IN'WX AS C0f)VF~fFNCF-STATUS INnIcATOR170) VTW..,OTW?~n=V~AXA(ITNO,TNF+I)DIVERGEOC IF 5 0 ILL9 DIVIDE/Ov OR ID!VERGFNCE, AFET RUN SAO* DIAGNOSE TROUBLE.180 CALL ERROR DIVERGEDC. Ooi FXP RvTURNS MINUS SIrN AS FRROR, TNnCATOR100 T 0 (FTWI 170.200-.200?n0 eTW.vTh.TgNi1CAcF?48LP ANDf ACf7RUF SOURfF FIFLfl OIL,..e*T4N-0*.T*WhL,,/(-W(L,)no 71f) 11.1,??C7' rT $10 NvXT PASS710 A(12)'=0fI7,LllWL.--wLl)IF SPILL, rlYVDE/09 OR fDIvCRGFNCF* SET RUN BAD, DIAGNC".E TROUBLECALL ERRORPilo LAPLArF TRANSFO0RM LOOP74n 7'ONTINI- 0C Cmm '?MAc 1000^I707 I ?1.IT I0OVARTANT VFCrflq rl-IO T IFT1N.lTT7PlN7 'oItT*IT 57TAT7?S T-nT'aT IN 1 rT LncT.9fl. 74Om-7cP0).4nnM^0, QAAOVII-243

CnCe ' TN (QUADRATIC OR LINEAR INTERPOLATION SCHEME)FINCTION OTNT(XIjoX2,X3.)X4.VeYIVV3l0:. r e.R4tVoN MArr - 2'. JffL 71. T(Mw* 21.5el.35n T..FN STAN X(•,1, Y;.)C PUNfTTON GONT RFT!JRN•P IAORATICALLY INTFRPOLAT~n VALUE Y4r rr.RRcspnflrNýN TM X4. -TVrN TNr)PPrN.FNT VARIARLE VALUFý Xl.XZ.(X3A~ rO~Q~It~rNmof '?PPFNnNT V/ARTARLP VALI'-- VJ.y?.Vl.(11 .Xj'ftl) X7v:4;=XhYMAX.AMAXIIARS(X!).ARý-(Xý).An•(XI)IY (4)=X'fA XAAr9'(YVAX.CQ.MoIrO To 21,10 lnO T=I101nO Y(I)=Y(tl/YwAXAn wo0 qYP4cs r r3uAnRATTf IN FAVOR OF LINPAR TNTrRPOLATTONIc(YMAX.GTenetGO TO SOMIF'n.40.40 TO I*=(Y(11*X(11-y(II)*l:X(•IY1a*Y(7|-Y(1)*X(11*YI!)X()-y{?)*Xll*(l|)+X1)*x(lj*Y(11-X(?j*X(ý-j*Y(j)j/n=ICY(I I*Xt ?oxI tl•*cX( 7)-X( q 1 +Y(71*X(I I*x(1 )*(X(%t-xII1) ÷y(•I.)X( I!*XT(A*IX{I4--X(4)+RX)/•fV) TM, 3AI 4 TTX(4))MAlf Yl~u&I RrSORT TO LIN94R INTJ RVOLATPONO O4nit T NUF T TOmNTJ-1mcunm - XfJl - Y(21tr CO0 TO 721)tT - Y)J) Y I(y(Ji- (XCI 1 - XtJ )/-NOMY(P4177 ^NTAYfJy*I-4XrM Tn AM4000 N0TNI|F17ALL PRR(I*AHIINTI0TvT.rQ1J4T4YUAXltntrK nKRtON t (l EERATES IRONECKER DELTA)91INCTION IKRONITJJ)C • YRFTtON n4TF - J AU' 73. TNMF 13.25.39vo^NcrKcR ncLTA CUNrTrnNacrt#RN0'?~TT]A'.n 1-CTIO FtAWUf" EQUILIBRIUM LAMBDA, NO 'SPRAY)P•fP rQATION nATF - 11 JUL 73, TIMF 6°2Q.18,RPTIIRNtI0'?4•r:÷OCITA-l.JN'T/OelnHT-- =• L•" (QUILIBRIUM LAMBDA. HB03. SPRAY)S Prc ep¢AttnN mATc -- 77 J111N 739 TIC .r,26,45HC=27A#ERTATrEHERTGO -in O I N10 'cfTAeC7T .O . -/60.)- TO 3O 23r- TO TAS0 •¢TMe TOM•./OIOO?0 !0rTAerTejnA0./6Oe)•I To inTO 36M'-4.NC On4 Ft(TTA-IOl. ./60.)An ý LMBD. BO3 SPA10 T CQlAfTON TM JAVII-244

100 VL6LtA S"ST/HSarTURN*O CL rLk(I;J$ (EQUIL.BRIUM LAMBDA, NaOH SPRAY)cIJ*CTI3N •LAM(TA1f PCr CQFATt'N MATV - 11 £l'I 71. TTMF *FRe...44r CMCST!0• €M W T^ PCTLION OnItILRRTRIUM QATV C0NuTAMTS FOR CA&IITIr'?(YA..ITf1.ne/&0.O TO in1CI.1An Tn inn10 IfITA.GT.000./60.IGO TO 20O)Hc0Th.flnlW:=l.rr TO 1n20 lVCTA*.0T*7Onlf6'*10fl Tr) 3nH14S7.F+04*f)IISnT*l TA-lOInfl./6n.10 T~fTA.C-T.a4l00./6Q.)16 Tn 40r-0 TO 10040 ,FITA.GT.7Cono./60.)G0 TO 50'1#4cr)T-l.F+n4Hc~=5.E+05+nHOSIT*( TA-4 000./6".)M0 TO 1n01n0 Vj-AM=OHCnfT/b4qQ.T1'RVII-245

APPENDIX KDIFFUSION OF RADIOACTIVE FLUID THROUGH SOILSURROUNDING A LARGE POWER-REACTOR STATIONAFTER A CORE MELTDOWN ACCIDENTbyJ. H. Pitts, B. R. Bowman, R. W. Martin, and J. P. McKayLawrence Livermore LaboratoryVII-247

Appendix KTable of ContentsSectionKl. INTRODUCTION .....................................................K2. METHOD OF ANALYSIS ...............................................K2.1K2.2One-Dimensional Ideal Gas ..................................One-Dimensional Multiphase Flow ............................K2.3 Two-Dimensional Ideal Gas ..................................K2.4 Heat Transfer ..............................................K3. RESULTS ..........................................................K3.1 One-Dimensional Ideal Gas ..................................K3.2 One-Dimensional Multiphase Flow ............................K3.3 Two-Dimensional Ideal Gas ..................................K3.4 Steam Generation by Molten Pool ............................K4. ADDITIONAL CONSIDERATIONS ........................................K5. CONCLUSIONS ......................................................ACKNOWLEDGMENT ..........................................................REFERENCES ..............................................................Page No.VII-251VII-252VII-252VII-253VII-254VII-254VII-255VII-255VII-256VII-257VII-257VII-258VII-258VII-259VII-259List of FiguresFigureVII K-1Schematic of Two Possible Configurations After a Core MeltdownAccident in a Large Power-Reactor Station .......................Page No.VII-261/262VII K-2 Configuration for the Two-Dimensional, Single-Phase Calculationin Cylindrical Coordinates ......................................VII-261/262VII K-3VII K-4VII K-5VII K-6VII K-7Schematic of the Element Mesh for the Steam GenerationCalculation .....................................................The Effect of Time and Pressure on the Normalized Active MassEfflux Calculated in Cartesian Coordinates for an Ideal Gasand for a Value po of 14.7 psia .................................Normalized Time at which the Active Containment-Shell Gas isReleased at Radius Ro; Calculated in Cylindrical Coordinatesfor an Ideal Gas and a Value of po of 14.7 psia .................Normalized Time at which Active Containment-Shell Gas isReleased at Radius Ro for Varying Radius Ratios .................Containment-Shell-Pressure Decay and Active Mass Efflux for CoreMeltdown Analysis ...............................................VII-261/262VII-261/262VII-261/262VII-261/262VII-261/262VII-249

List of Figures (Continued)Page No.FigureCore After-Heat of 3.5 x 10-:7 BTU/H and a Containment-ShellVolume of 2 x 106 Ft 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . VII-263/264.VII K-8 Water Saturation, Normalized Pressure, and TemperatureDistribution for pl = 100 psia ..................................VII K-9 Water Saturation, Normalized Pressure, and TemperatureDistribution for P, = 50 psia ...................................VII-263/264VII-263/264VII K-10 Water Saturation, Normalized Pressure, and TemperatureDistribution for pl = 25 psia ...................................VII-263/264VII K-Il Part (A) Shows Water Saturation, Normalized Pressure, andNormalized Temperature Distribution for Pressure Decay forP1 = 1 P, = 100 psia. Part (B) Shows Containment-ShellPressure as a Function of Time at Pl = 100 PSIA .................VII K-12 Pressure Profile of a Two-Dimensional Calculation att = 2.9 H .......................................................VII K-13 Pressure Profile of a Two-Dimensional Calculation att = 6.2 .H .......................................................VII K-14 Pressure Profile of a Two-Dimensional Calculation att = 14.7 H ......................................................VII-263/264VII-263/264VII-263/264VII-263/264VII K-15 Steam Generation for Core Melt in Contact with Wet Soil for aVII-250

Appendix KDiffusion of Radioactive Fluid Through SoilSurrounding a Large Power-Reactor StationAfter a Core Meltdown AccidentbyJI H. Pitts, B. R. Bowman, R. W. Martin, and J. P. McKayLawrence Livermore LaboratoryThe f low of f luids through the soilsurrounding a large power-reactor stationafter a core meltdown accident wasanalyzed. Fracture of, or penetrationthrough, a portion of the containmentfloor was assumed so that radioactivegases would be driven through the soilby pressure within the containmentshell.Results for both one-dimensional idealgas and multiphase flow were obtainedusing dimensionless variables so thatresults are applicable for all soil conditions.Two-dimensional results areincluded for a specific case where permeabilityof the soil varies spatially.These calculations show that the timerequired for radioactive gases to permeateto the ground surface are years forsilty-type clays and about 10 h forsand-type soil.Heat transfer calculations that estimatethe amount of steam generated should thecore melt penetrate the containmentfloor and contact wet soil show thatcooling systems must be in operationwithin four or five days in order toprevent overpressurization of the containmentshell. Fluid motion existingin the region immediately around thecontainment shell and also caused byfissure propagation is mentioned.K1. INTRODUCTIONA reactor accident which causes coremeltdown will probably never occur in alarge power-reactor station. However,because the effects of such an unlikelyaccident could be serious, a study ofconditions following core meltdownshould be included as part of any safetyanalysis report. Such studies are usefulin the design of future power stations.Also they help to minimize oreliminate adverse effects should an accidentoccur in an existing power plant.We examined one means by which radioactivematter could reach the ground surfaceand be released into the atmospherefollowing a core meltdown accident.Consider that the melted core interactswith the containment floor so that eitherthe molten pool penetrates thefloor, or the floor is fractured to suchan extent that it no longer can be consideredas a containment barrier. Radioactivefluids would then be releasedinto the soil beneath the containmentshell and permeate in all directionsthrough the soil (see Fig. VII K-1).From a calculational standpoint, theanalysis is approximately the same forboth cases since active material isreleased by essentially the same, pointsource. The driving pressure is thatpresent in the containment shell.Changes in the concentration of radioactivematter due to filtering interactionwith the soil or adsorption are notincluded.We have completed a four-part analyticalstudy that predicts the extent of penetrationof radioactive fluid into thesoil and possible release into the atmosphereusing methods similar to thosedeveloped by Morrison (Ref. 1,2). Theproblem was first bounded using our onedimensionalprograms with a noncondensiblegas. Geometries included applicablecases in Cartesian, cylindrical, andspherical coordinates.In the first study, we considered a Cartesianmodel of a constant-area columnwith a length equal to the distance fromthe expected source through the soil tothe ground surface. Soil properties(e.g., permeability and porosity) wereassumed constant. The results of thiscalculation are used as a bound on thetime for release of active gas to theatmosphere and for comparison with condensiblefluid calculations describedlater. Use of dimensionless parameterspermits application of the results tonumerous cases with different gases andhaving different soil properties. Amodel using cylindrical coordinates simulatespossible radial flow of noncondensiblegases in a single layer ofsoil. For example, consider the eartharound the reactor to consist of a horizontallayer of sand with clay above andbelow. The permeability in the sandregion would be probably several ordersVII -251

of magnitude above that of the clay.Should gases reach this relatively highpermeability region, vertical flow couldbe neglected in comparison to radialflow. Spherical coordinates may be usedto determine the extent of flow belowthe containment floor shown in Fig. VIIK-1. This model is accurate until apressure response is felt at the outerradius of the cofferdam. Beyond thistime, gases would be able to permeateboth vertically toward the ground surfaceand radially outward in a twodimensionalfashion. If the outer radiusof the one-dimensional spherical coordinatesum ofproblem is taken equal to thethe distance from the source tothe outer cofferdam radius and thevertical distance to the ground surface,the predicted time of arrival would beshorter than that actually occurring.A second set of calculations using ourone-dimensional, Cartesian coordinate,multiphase flow model was completed sothat effects of steam condensing withinthe pores of the soil could be included.In all calculations, a mass balance wasmaintained so that the pressure withinthe containment shell decreased as flowinto the soil occurred. We found thatthis pressure decay due to flow out ofthe containment shell was small for mostcases. (A more rapid decay would occurif the permeability was high or if thegas source volume was reduced in size.)As a result, we also studied effects ata constant containment-shell pressuresince this permitted further generalizationof results and the use of similaritysolutions.The third part of the study utilizedBowman's (Ref. 3) two-dimensional extensionof the noncondensible program.This permitted us to closely model thegeometry of the containment shell shownin Fig. VII K-l, and to include variouslayers of soil, each with appropriateproperties. The extent of radioactivegas penetration with time for condensibleflow may be estimated by comparisonof changes found for a noncondensiblegas on a one-dimensional basis.The fourth and last series of calculationswas aimed at bounding the extentof heat dissipation from the outer wallof the containment shell and the steamgeneration caused by contact of the coremelt with surrounding ground watershould the molten pool penetrate thecontainment floor. These results wereused to justify the use of a constantpressure (25, 50, or 100 psia) insidethe containment shell. The last twovalues correspond to what might be thedesign pressure and rupture pressure ofthe containment shell.K2. METHOD OF ANALYSISK2.1 ONE-DIMENSIONAL IDEAL GASIn the first part of the study we consideredone-dimensional, isothermal flowof an ideal gas. The restriction ofisothermal flow is reasonable since theheat capacity of the soil is much largerthan the energy present in the gas flowingthrough the pores of the soil. Thecontinuity equation for compressibleflow through porous media with constantporosity may be written asV - (P) + E: 5T 0 (VII K-l)where p is the fluid density, E is theporosity, t is time, and S is an apparentvelocity equal to the volume flowrate per unit area normal to the flow.Conservation of momentumusing Darcy's law (Ref. 4)uVpissatisfied(VII K-2)which is a constitutive equation forlow-Reynolds-number flow through porousmedia. Here p is the pressure in thefluid, p is the fluid viscosity, and kis the permeability of the solid. Usingan ideal gas equation of statep = pRT ,(VII K-3)where R is the gas constant and T is thefluid temperature which we take to beconstant, the governing equation becomesrl1-n D rn-i pr - aP(rn Pr ;20)= k D t'(VII K-4)where r is the spatial variable and n isset equal to 1 for Cartesian coordinates,2 for cylindrical coordinates,and 3 for spherical coordinates.Equation (VII K-4) is placed in dimensionlessform by definingX=r0where Ro is the outerporous medium,p =Pl-Po(VII K-5)radius of the(VII K-6.)where po is the initial fluid pressurein the pores of the solid and p is theinitial driving pressure and 1VII-252

where T isk(pl -po)R2S p R2dimensionless time and(VII K-7)The total mass having left the poroussolid isM M=)n l Xon x ddX0i 0N = pl/Po . (VII K-8)( ,2 + 2P \IXldIn terms of these dimensionless quantities,Equation (VII K-4) becomesa 2 2 + + _n-_ +_2P_Xax ()+x 2TR(P N - )=2 a ;T(VII K-9)Boundary conditions used are:P = 0 at T = 0P = 0 at X = 1 for all TP = 1 at X = X 0 at T > 0The value Xo is the ratio of inner toouter radius and is set equal to zero ifCartesian coordinates are used. A timedependentdriving pressure at X = Xo,instead of a constant value, may be incorporatedif desired. The time dependencyof pressure decay may be calculatedto correspond to the mass flow rate intothe porous media or any other mathematicalrelationship.Ifwe define a dimensionless massmm 0(VII K-10)where m is the total mass that has leftthe outer radius of the solid, and mo isthe mass of gas originally in the poresof the solid, we may write the continuityrelation where A is the flow area asd(m/mO) ÷dM==puAdt dt m (VII K-11)0(VII K-13)When M equals unity, the mass havingleft the solid will just equal the originalmass of gas in the pores. Gasoriginally in the containment shellwould start to be released at this time.Equation (VII K-9) is solved using a finitedifference technique with Equations(VII K-12) and (VII K-13) used to determinethe rate and amount of mass thathas left the solid. Use of dimensionlessparameters is very advantageoussince it makes the results applicablefor all values of permeability, porosity,and gas viscosity.K2.2 ONE-DIMENSIONAL MULTIPHASEFLOWIn the second series of calculations weconsidered one-dimensional flow of steamand air in a Cartesian-coordinate system.Even though pure steam was consideredto be injected into the soil, airor other noncondensible gases are presentunder ambient conditions in thepores of the solid so that the solutionmust consider two species as well as twophases. For each species we write acontinuity equation of the forma (p) aT-X (pu) + E t (pS) = 0 , (VII K-14)where S is the saturation of the speciesor the fraction of the pore volume accessibleto the component species. If£ represents liquid water, m the mixtureof air and water vapor, a the noncondensibleair alone, and v the water vaporalone; we may write the two necessarycontinuity equations asa- (-E + s (PaSm) = 0 (air)When combined with Darcy's law and thedefinitions of dimensionless parameters,this becomesdM n N - 1 dd-- (1 - Xo)n 2 Txanda- (p u£ + Pvum) + e -L (P S£ + PS)(VII K-15)(P ,2 + RN2P)I x =(VII K-12)= 0 (H 2 0 liquid and vapor).(VII K-16)VII-253

All species in gaseous phase are assumedto move at the velocity of the gaseousmixture. Each phase obeys Darcy's lawwith a relative permeability included toaccount for the interaction of the twophases;U - -- k -Lk 11Z rm axU k k a.M V rm ax(VII K-17)(VII K-18)The relative permeabilities, kr£ andkrm, are taken to be functions of watersaturation in the pores of the solid andare the ratio of effective permeabilityof a phase to the absolute permeabilityof the solid.For an elemental volume of the columnthe energy equation is written asax-- (ptuh + Pvumhv + Pa u mha)+ 1-- (p S e + pvSme + pSme)at Z , viny a maaes7 tPs 0 (VII K-19)where h is enthalpy and e is internalenergy. Equations (VII K-15) through(VII K-19) are combined using dimensionlessvariables as shown in Reference 2with equations of state for steam, water,vapor, and air, and then solvedusing finite difference techniques. Byneglecting gravitational effects andaxial heat conduction, and considering asemi-infinite medium, a similarity solutionexists where water saturation,pressure, and temperature may be determinedas a function of the similarityvariable, 0, where_ "a2 kplt (VII K-20)K2.3 TWO-DIMENSIONAL IDEAL GASA third set of calculations modeled thegeometry of a typical, large, powerreactor in a two-dimensional ideal gascalculation (see Fig. VII K-2). We usedthe equations of the one-dimensionalideal gas calculation and extended themto two dimensions. That is, the transientflow is governed by Equations (VIIK-l) through (VII K-3) except that nowthe permeability takes on the form of atensor. The equations are solved forthe pressure in cylindrical coordinatesusing an explicit, finite-differenceapproximation with a constant temperature.Variations of soil permeability betweenlayers of sand and clay are permitted asshown in Fig. VII K-2. The permeabilityis assumed to be isotropic within eachsoil layer and does not vary with time.Porosity values are typical for the gasfilledporosity in alluvial clays andsand (Ref. 5). This porosity representsthe fraction of interconnected void volumeavailable for flow of gas in theporous media. It is assumed that theonly effect of liquid in the pores is toreduce the porosity.Pressure within the containment shell isassumed to decay with time as a resultof mass flow from the shell into thesurrounding soil; however, this decaywas again found to be small except forcases of reduced gas source volume orhigh soil permeability. The containmentshell is assumed to be impermeable exceptfor a 10.3-ft-radius cylinder representingthe volume occupied by themass of the molten core.K2.4 HEAT TRANSFERWe estimated the dissipation of heatgenerated after core meltdown by consideringconduction through the containment-shellwall. If the core does notpenetrate the containment floor and allcooling systems are inoperative, a simpleheat balance may be used to estimatethe inside surface temperature of thecontainment shell. Taking an averageheat generation rate of 3.5 x 107 Btu/hover the first 30 h after core meltdown,the temperature drop through a 2.5-ftthickconcrete containment shell exceeds3000 F. This temperature is in excessof what concrete will withstand. Withoutan operative cooling system or theaddition of material whose heat capacityis very large, failure of the containmentshell would be certain.Should the core melt penetrate the containmentfloor, the energy generatedwould vaporize the water present in thepores of the soil adjacent to the moltenpool. We examined the problem of heatconduction in an infinite, homogeneousmedium capable of undergoing phasechange in order to determine the amountof steam which could be generated by thecore under these conditions. The internalboundary was a sphere of 10.3-ftradius with a constant internal heatgeneration rate of 3.5 x 107 Btu/h (seeFig. VII K-3).The solution followed the finite elementmethod of Wilson and Nickell (Ref. 6).VII-254

The molten pool was assumed to be stationaryand the thermal properties ofthe soil during phase change were takento be an average of the saturated anddry conditions. Steam was assumed toflow from the soil into the containmentshell with negligible pressure losssince any substantial buildup of pressurewould fracture the soil.The basic solution technique uses thematrix equationswith[K] (8) = (Q)(VII K-21)[K) = 1K) + IC] (VII K-22)where K and C are matrices which includethe thermal conductivity and heat capacity,6 is the temperature differencevector, and Q is a thermal force vectorwhich includes heat flux across the elementsurface and internal heat generation.The solution begins at time zerowith given initial conditions, and proceedsuntil the average temperature ofthe first spherical shell of soil equalsthe saturation temperature specified.When saturation temperature is attained,the thermal properties of the soil inthat shell are changed to those representingthe average moisture conditionduring phase change and the appropriatethermal matrices are reevaluated. Whenthe added energy is equal to that requiredto transform the water to steam,the phase change is complete. Soilproperties are then modified again toreflect the dry condition of the soil.or gas viscosity. Three values of containmentshell pressures are shown,which should bracket expected values formost foreseeable accidents.In order to explain the use of dimensionlessparameters, we shall considerthe following example of axial flow inthe annular space between the outersurface of the containment shell and thecofferdam (see Fig. VII K-l). Since theflow is axial, we use the Cartesian coordinateresults in Fig. VII K-4.Length = 40 ftArea = fr/4 x (1502 - 1352) = 3360 ftPorosity = 0.1Pore volume = 0.1 x (40 x 3360)= 13,400 ftPl - Po = (50 - 14.7) = 35.3 lb/in. 2Density at ambient conditions= 0.075 lb/ft 3Gas viscosity = 5.0 x 10-7 lb - sPermeabilityE 11 R 2ot ~ 0 Tk(P - Po)30.1 Darcy= [0.1 x 5.0 x 10-7 (lb • s/ft2ft22K3. RESULTSx 402 ft2x T]/[0.l Darcy x 1.06K3.1 ONE-DIMENSIONAL IDEAL GASResults for the one-dimensional, idealgascalculations are shown in Figs. VIIK-4 to VII K-7 using dimensionless variables.In Fig. VII K-4, the normalizedactive mass effluxI, Mo, is defined asthe amount of active mass released tothe atmosphere divided by the amount ofinactive gas initially contained withinthe pores of the earth in the column.In conjunction with the dimensionlesstime, T, defined previously in Equation(VII K-7), the results are universallyapplicable for any area or length of thecolumn and any porosity, permeability,x lo-11 (ft 2/Darcy) x 35.3 (lb/in.2x 144 (in 2/ft 2)]= 14,800 T (t in seconds)= 4.12 T (t in hours)The initial mass of gas in the pores ofthe solid, calculated from the densityat ambient conditions, is1 Active mass is defined as the mass ofmaterial initially contained in thecontainment shell.mi = 0.075 lb x 13,400 ftm = 1,000 x M o (m in lb).3_ 1,000 1-VII-255

We conclude for this example that noactive gas is released until T = 0.70(from Fig. VII K-4) or t = 2.88 h. AtT= 1.0, or t =4.12 h, the amount ofactive gas released would be N 0 = 0.65or m = 650 lb. The average flow rateover this time period is 524 lb/h.Results for cylindrical coordinateswhere the dimensionless time for initialrelease of active gas to the atmosphereis plotted against the ratio of inner toouter radius are presented in Fig. VIIK-5. The radius ratio is a parameterwhich must be included in cylindricaland spherical coordinates since the flowarea is a function of radius. As in thecase of Cartesian coordinates, there isa period of time when no active gas isreleased into the atmosphere. Duringthis delay time, active gas is permeatingthrough the solid and displacing thegas initially in the pores. The flowrate increases gradually from zero;however, by the time active gas startsto be released the flow rate is normallysubstantial. Use of results from thecylindrical coordinate case would beappropriate, for example, if active gaswas released into a sand layer justbeyond the radius of the cofferdam (seeFig. VII K-1). Since the clay above andbelow the sand layer would prevent appreciablevertical flow, gas would permeateradially outward to a place whereit could be released such as a riverchannel.The spherical coordinate results of Fig.VII K-6 may be applied to the analysisof the flow of pressurized gas afterpenetration of the containment floor.From a calculational standpoint bothcases shown in Fig. VII K-1 are nearlyidentical.In Fig. VII K-4 to VII K-6, results arepresented for a constant driving pressure.Figure VII K-7 shows a specificcase in spherical coordinates when themass leaving the containment shell andpermeating into the soil was calculated.The driving pressure in the containmentshell would then decay as mass was lostto the soil. At the time active gasstarts to be released from the outerboundary (T = 1.0) the driving pressurehas decayed by only 5%. This pressuredecay would have been less if cylindricalor Cartesian coordinates were used.We felt that the use of a constantpressure, in lieu of the more accuratepressure decay, is justified since theeffect of mass loss is usually small andother factors such as operation of theemergency containment cooling system areprobably more dominant.K3.2 ONE-DIMENSIONAL MULTIPHASEFLOWTwo phase flow with condensation wasstudied and the results are presented inFigs. VII K-8 through K-11. Containmentshell gas is taken to be pure saturatedsteam at the indicated pressure. If thecontainment shell pressure is held constantfor example at 100 psia, the pressure,temperature, and water saturationdistributions may be found for anyposition and time using the similarityvariable e (see Fig. VII K-8). Thissimilarity variable gives spatial variationfor a given time where a smallvalue of 0 corresponds to positionsclose to the pressure source. It alsogives temporal variation for a givenspatial position with time increasingfrom right to left in the direction ofdecreasing 0.If we consider a given time, the interfacebetween the soil and the cavityfilled with steam is at e = 0. Here,.the temperature and pressure are equalto the containment shell conditions sothat the normalized values are unity.Also the steam quality is 100% yieldinga water saturation that is zero. As wefocus on increasing values of X (or 0),the pressure decreases as a result ofmomentum loss in the pores of the solid.In this region, 0 < 6 < 0.04, the soilhas been heated to the saturation temperatureof the steam. The slightlynegative slope corresponds to a decreasein saturation temperature with decreasingpressure. Steam had previously condensedto water in raising the temperatureof the soil, but no furthercondensation is taking place in thisregion. Some of this water remains inthe pores resulting in an increase inwater saturation.As 0 increases just beyond 0.04, a pointis reached where the soil has not yetreached the steam saturation temperature.Condensation here results in anabrupt increase in water saturation.This region is very narrow and thetemperature decreases abruptly to ambienttemperature.Beyond 60 0.04, the water saturationremains higher for some distance becausethe previously condensed water has beenpushed ahead of the condensation region.Water and gas flow are nearly incompressiblein this region and thepressure gradient is almost linear. ForO > 0.2, flow is essentially that of thegas phase alone.The active fluid front is determined byintegrating the velocity of the gaseousphase with time. For all three valuesVII -256

of pressure examined, this interfaceoccurred at a value of e = 0.05.In comparing the ideal gas calculationto the two-phase, condensible flow calculationin order to determine theextent of radioactive penetration, wemust consider that while the total porosityof the soil is probably close to30%, the liquid fills most of the voidspace. If we choose the initial saturationof 70%, then we should compare anideal gas calculation with porosity ofabout 10% to the two-phase calculationof total porosity equal to 30%. Comparingresults of Figs. VII K-4 and VII K-9for 50 psia, we have an interface timeT and similarity variable 8 of 0.07 and0.05, respectively. Converting to adimensional time for the ideal gas andtwo-phase flow cases, we havev. pR 2ti = 1(Pl _ 0P) T (ideal gas)2 2tt- X - 2 (two-phase flow)4kpIe(VII K-23)(VII K-24)where as before X = x/R 0 . We areinterested in the downstream face of thecolumn where X = 1, so thatt t = t(pl -po)t i4EiiP1 e20.3 x (50 - 14.7) psia4 x 0.1 x 50 psia x 0.052 x 0.70= 303 (VII K-25)In other words, the estimated time forrelease of active material in the twophase,flow calculation is about 300times longer than in the ideal gasapproximation. These two sets of results(containment-shell gas consistingof 100% steam or ideal gas) are boundsfor release of radioactive material.Cases where mixtures of steam and idealgas are forced through the soil wouldgive results in between these limits.Figures VII K-9 and VII K-10 showsimilar results for containment-shellpressures of 50 and 25 psia. Figure VIIK-Il shows containment-shell pressuredecay as fluid is lost into the soil.The initial containment-shell pressuredoes not change appreciably for thiscase so that results are similar tothose of Fig. VII K-8. The pressurecurve is normalized using the cavitypressure in all cases, but in Fig. VIIK-11 this pressure is a function oftime. Temperature curves are normalizedin a similar fashion.K3.3 TWO-DIMENSIONAL IDEAL GASWe performed a two-dimensional calculationwith the specific geometry shown inFig. VII K-2 to show effects of spatialvariation in soil permeability and aspecific power-station geometry. Thesecalculations were found to be timeconsuming. Results in Figs. VII K-12 toVII K-14 show the pressure distributionand the location of the active gas frontat three different times. At the latertimes, a deviation from the onedimensionalspherical case occurs whichdelays the propagation of the, active gasfront. An initial value of containmentshell pressure equal to 100 psia waschosen and allowed to decay as massdiffused into the soil. The change indriving pressure was negligible for thiscase.Note that the permeability of 0.1 Darcy,which corresponds to that expected forsand, was chosen for the lower region.Had a lower value of permeability beenused, results would have been similarexcept that propagation of the frontwould have taken a longer time.K3.4 STEAM GENERATION BY MOLTENPOOLIf the core melt has penetrated thecontainment floor (second case in Fig.VII K-i), the after-heat from thereactor would vaporize the water presentin the pores of the adjacent soil. Thissteam generation is described in Fig.VII K-15 along with values of thequantity of steam required to pressurizea 2 x 10 6 3-ft containment shell to 50and 100 psia. Although the heat generationrate was held constant at3.5 x 107 Btu/h in the molten pool, thesteam generation rate decreases withtime since heat must penetrate more andmore soil before reaching an area wherewater saturation is still significant.Some steam is lost by flow into thesoil. This loss of steam is normally aninsignificant percentage of the steamgenerated and only becomes important ifrelease of radioactive gases occurswithin a few hours after the accident.Heat dissipation through the containmentshell into the atmosphere was found tobe negligible. Some cooling system mustbe functioning to keep the pressurewithin the containment shell below itsVII-257

selected design or burst pressure of 50or 100 psia, respectively.K4. ADDITIONAL CONSIDERATIONSNo study would be complete without mentioningperipheral areas of interest.Although we studied permeation of fluidsthrough soils, some attention should begiven to the annular region-between thecofferdam and the containment shell as apossible area from which some radioactivematerial might be released to theatmosphere. our experience indicatesadequate containment can be achievedwith proper attention to the materialsused in filling this region. It may bedesirable to experimentally obtain aneffective permeability through regionsconsidered t~o be nearly gas tight seals.Results presented in this and futurestudies could be used to estimate theseverity of any leakage and to specifyany corrective action needed.Another subject which requires specialattention is the propagation of fissuresthrough the soil outside the cofferdam.These -fissures (Ref. 7) may-6pen shouldpressures below the containment shellreach values near the hydrostatic overburdenpressure exerted by the soil.The overburden pressure is of the orderof 1 psi per ft of depth which generallyexceeds the design pressure of thecontainment shell. From a standpointof fissure propagation, the most susceptiblesoils are those with low permeability.one would desire a highpermeability to eliminate fissure propagationand a low permeability to preventrelease of radioactive debris bydiffusion through the soil. Backfillcan be selected to achieve optimumconditions.K5. CONCLUSIONSa. Results for constant driving pressuresof 100, 50, or 25 psia showthat the time required for ideal radioactivegases to reach the surfaceof the ground are of the order ofyears if soil permeabilities arenear those anticipated for siltytype clays (0.0001 Darcy). Theserelease times are shortened to about10 h if permeabilities expected forsand type materials (0.1 Darcy)exist.b. Decay of the driving pressure in atypical containment shell due toleakage of gas into the soil isinsignificant for most cases becauseof the' large volume of the shell.Incorporation of a pressure decay isimportant for small gas sourcevolumes or if the soil permeabilityis large. Other factors such asoperation of various cooling systemsare probably more dominant in affectingthe driving pressure.c. An estimate of the release time formultiphase flow (injection of allsteam from the containment shell)indicates that it may be over twoorders of magnitude longer than thatpredicted with injection of idealgas. Cases where mixtures of steamand ideal gas are forced through thesoil would give results falling betweenthese bounds.d. Steam generation, which results fromcore melt penetrating the containmentfloor, is large enough thatsome cooling systems must, be inoperation within 4 or 5 days afterthe accident in order to preventoverpressurization of the containmentshell. Heat dissipationthrough the containment shell to theatmosphere is negligible.e. Examination of the permeabilitybetween a typical cofferdam andcontainment shell is suggested sincethis annular region is an area fromwhich some radioactive materialmight possibly be released to theatmosphere. Results depend upondetailed knowledge of materialproperties and exact constructiondetails.f. This analysis and future studies maybe used to specify materials at reactorsites yet to be built. Backfillnear the reactor may bespecified for optimum conditions andmaterials adjacent to the containmentshell may be chosen to assureadequate confinement of radioactivematerials.VII-258

AcknowledgmentsA study such as this is most efficient if several people combine their differenttalents. Calculations of flow through the soil were performed by B. R. Bowman andJ. H. Pitts. J. P. McKay calculated dissipation of heat through the containmentshell wall to the atmosphere, organized most of the input numbers used in thecalculations, and performed some of the computer work. R. W. Martin performed thesteam generation calculation. J. E. Mellor helped with a portion of the computerwork.V. N. Karpenko, C. E. Walter, and D. M. Norris offered suggestions for the study andreviewed the writing of the report. H. Lentzner edited this report and made numerousimprovements to the writing.R. Ritzman of Battelle Memorial Institute was responsible for the instigation of thisstudy, formulation of the problem, and determination of the basic input parameters.He also guided our work as we progressed so that results would be of generalapplicability to safety analysis problems.References1. F. A. Morrison, Jr., Transient Gas Flow in a Porous Column, Lawrence LivermoreLaboratory, Rept. UCRL-52929 (1971).2. F. A. Morrison, Jr., Transient Multiphase Multicomponent Flow in Porous Media,Lawrence Livermore Laboratory, Rept. UCRL-74327, Rev. 1 (1972).3. B. R. Bowman, TWIG, A Computer Code for the Calculation of Two-Dimensional IdealGas Flow in Porous Media, Lawrence Livermore Laboratory, Rept. UCID-16320 (1972).4. R. J. M. Dewiest, Flow Through Porous Media (Academic Press, 1969) pp. 1-6.5. Ibid, pp. 78-82.6. E. L. Wilson and R. E. Nickell, "Application of the Finite Element Method to HeatConduction Analysis," Nuclear Engineering and Design 4, 276 (1966).7. J. H. Pitts, Gas Initiated Crack Propagation in a Porous Solid - A ProgressReport, Lawrence Livermore Laboratory, Rept. UCID-16236 (1973). a40 a40VII-259

Crack. area in containment floorLegendConfiguration ASConcreteSandCi ayBackfillConfiguration BFIGURE VII K-1Schematic of Two Possible ConfigurationsAfter a Core Meltdown Accident in a LargePower-Reactor Station

DarcysFIGURE VII K-2Configuration for theTwo-Dimensional,Single-Phase Calculationin CylindricalCoordinates20Radial distance - ft50Condit-onCore reeltSoilWet Average Dry,c Bto/ft r. Ft 10.0 1.0 0.75 0.0C. Stu/Ib R 1.12 044 035 0.26p. IWr 3 270 125 118 112No. of diaoe-s n each regionRange. ft No. of radial divisions0.0 ro 10.3 a10.3 to 13,0 1013.0 to 20.0 1420.0 to 50.0 30FIGURE VII K-3Schematic of the ElementMesh for theSteam Generation Calculation

Mi(from containmentshell)(toatmosphere)ý- R. -- Ia. Model flow through aconstant-area porous column0x0ZFIGURE VII K-4IDimensionless time -- rb. ResultsThe Effect of Time and Pressureon the Normalized ActiveMass Efflux Calculated in CartesianCoordinates for anIdeal Gas and for a Value poof 14.7 PSIAMoR R 02-a. Model-radial flow througha porous cylindrical diskE0Z.0.1 0.2Ratio of inner to outer radius -b. ResultsRi/RoFIGURE VII K-5Normalized Time at which theActive Containment-Shell Gasis Released at Radius Ro; Calculatedin Cylindrical Coordinatesfor an Ideal Gas and aValue of Po of 14.7 PSIA

RoRia. Model-radial flowthrough a porous sphereWEZFIGURE VII K-6b. ResultsNormalized Time at which ActiveContainment-Shell Gas isReleased at Radius Ro forVarying Radius Ratios: Calculatedfor a po Value of 14.7PSIA0.06 r- .O..000.05 I-Ecu0.03 FE0.041-0.021-0z0.01010 I .2Dimensionless time - 73FIGURE VII K-7Containment-Shell-Pressure Decay andActive Mass Efflux for Core MeltdownAnalysis: Calculated in SphericalCoordinates for an Ideal Gas and theFollowing Parameters: pl = 100 PSIA,E = 0.08, Ri =b15 Ft, Ro = 120 Ft,= 5 x 10-7 Lb S/Ft 2 , and t = 1.3T/K, where t is the Time in HoursFig. VII K-I - Fig. VII K-7VII-261/262

0 0.1 0.2 0.36ala]Similarity variable, 0 - - 1½20.3Similarity variable, a - x -Ela 1 /2"2 Lk-PltjIFIGURE VII K-8Water Saturation,Normalized Pressure,and Temperature Distributionfor p 1100 PSIAFIGURE VII K-10Water Saturation,Normalized Pressure,and Temperature Distributionfor p, =25 PSIAIiiI I I I I 111111 'II .0O0.8Sw0.60.4PC -Po0.2n'a=k.! m0•0 0.1 0.2 0.3TSimilarity variable,.0 -2ý L kP 1 lt JIa.C.CLSimilarity variable, 0 - • [ t]or t I I-J I I I I I I I I I I I 1 1 ,,-0 0.005 O.OIO 0.015rb.FIGURE VII K-9Water Saturation,Normalized Pressure,and Temperature Distributionfor pi = 50PSIAFIGURE VII K-11Part (A) Shows Water Saturation,Normalized Pressure,and Normalized TemperatureDistribution forPressure Decay for p 1 = 1pl = 100 PSIA. Part (B)Shows Containment-ShellPressure as a Function ofTime at p 1 = 100 PSIA

Distance, ftDistonce, ftC.)C•5LOFIGURE VII K-12Pressure Profile ofa Two-DimensionalCalculation att = 2.9 HFIGURE VII K-14Pressure Profile ofa Two-DimensionalCalculation att = 14.7 HzDistance, ft- R_Quantity of steam for 100 psia incontainment shellQuantity of steam for 50 psia incontainment shell- 0 5SCDCC-wU,30raE4ccýi I0(I,103 0 40 80 120Time - h160 2000I I IFIGURE VII K-13 Pressure Profile ofa Two-DimensionalCalculation att = 6.2 HFIGURE VII K-15Steam Generation forCore Melt in Contactwith Wet Soil for aCore After-Heat of3.5 x l0-7 BTU/H anda Containment-ShellVolume of 2 x 106Ft 3Fig. VII K-8 - Fig. VII K-15VII-.263/264

WASH- 1400(NUREG 75/014)PHYSICAL PROCESSESinREACTOR MELTDOWN ACCIDENTSAPPENDIX VIHtoREACTOR SAFETY STUDYbyW. A. CarbienerP. CybulskisR. S. DenningJ.M. GencoN. E. MillerR. L. RitzmanR. SimonR. 0. WootonContract W-7405-eng-92Battelle Columbus LaboratoriesU.S. NUCLEAR REGULATORY COMMISSIONOCTOBER 1975

Appendix VIIITable of ContentsSectionPage No.1. INTRODUCTION .................................................. VIII-11.1 Objectives ............................................... VIII-21.2 Interactions ............................................. VIII-21.3 Approach ................................................. VIII-31.4 Limitations .............................................. VIII-31.5 Definitions .............................................. VIII-31.5.1 Sequence ......................................... VIII-31.5.2 Subsequence ...................................... VIII-32. PWR ANALYSES .................................................. VIII-92.1 Accident Sequences Resulting in Core Meltdown ............... VIII-92.2. Analysis of Degraded Accident Behavior ......................... VIII-102.2.1 Basic Assumptions ................................ VIII-102.2.2 Accident Time Scale .................................... VIII-II2.2.3 Core Meltdown .................................... VIII-II2.2.4 Primary Vessel Meltthrough ............................ VIII-122.2.5 Containment Vessel Penetration ......................... VIII-122.2.6 Steam Explosions ................................. VIII-132.2.7 Hydrogen Combustion .................................... VIII-142.2.8 Containment Response ................................... VIII-142.2.9 Leak Rate from Containment ............................ VIII-162.3 Subsequent Probabilities ................................. VIII-162.3.1 Definitions ...................................... VIII-172.3.2 Containment Failure Resulting from aSteam Explosion ............... ................... VIII-172.3.3 Containment Meltthrough with ContainmentIsolation Failure ................................VIII-192.3.4 Hydrogen Combustion .................................... VIII-192.3.5 Containment Overpressurization ......................... VIII-192.3.6 Containment Meltthrough ................................ VIII-203. BWR ANALYSES .................................................. VIII-253.1 Accident Sequence Resulting in Core Meltdown ................. VIII-253.1.1 Electric Power System (EPS) ........................... VIII-253.1.2 Vapor Suppression System (VSS) ........................ VIII-253.1.33.1.4Containment Leakage (CL) ...............................Emergency Cooling Injection (ECI) .....................VIII-25VIII-253.1.5 Emergency Cooling Function (ECF) ..................... VIII-253.1.6 Scram ............................................ VIII-253.1.7 Long-Term Cooling (LTC) ................................ VIII-263.2 Analysis of Degraded Accident Behavior ........................ VIII-263.2.1 Basic Assumptions ................................ VIII-263.2.2 Accident Time Scale .................................... VIII-263.2.3 Core Meltdown .................................... VIII-263.2.4 Pressure Vessel Meltthrough ........................... VIII-273.2.5 Steam Explosions ................................. VIII-273.2.6 Containment Vessel Penetration ......................... VIII-283.2.7 Hydrogen Combustion .................................... VIII-283.2.8 Containment Response ................................... VIII-28VIII-i

Table of Contents (Continued)SectionPage No.3.2.9 Containment Isolation Failures ....................... VIII-303.2.10 Secondary Containment ............................ VIII-303.3 Subsequence Probabilities ................................ VIII-31APPENDIX A THERMAL ANALYSES ....................................... VIII-37APPENDIX B PHYSICAL EXPLOSIONS RESULTING FROMCONTACT OF MOLTEN MATERIALS AND WATER ......................VIII-69APPDNDIX C MODEL STUDY OF PHYSICAL EXPLOSIONS ......................... VIII-87APPENDIX D NONCONDENSABLE GASES ................................... VIII-lllAPPENDIX E CONTAINMENT FAILURE MODES EVALUATION ....................... VIII-129VIII-ii

Appendix VIIList of TablesTableVIII 1-1 PWR Data ........................................................VIII 1-2 BWR Data ........................................................Page No.VIII-5/6VIII-7/8List of FiguresFigureVIII 2-1VIII 2-2VIII 2-3VIII 2-4VIII 2-5VIII 2-6VIII 2-7VIII 2-8VIII 2-9VIII 2-10VIII 2-11VIII 3-1Time Required for 80 Percent Core Melting from the Onsetof Melting ......................................................PWR Containment Pressure Response During the DesignBasis Accident and with the Assumed Failure of theLow Pressure Recirculation System ...............................PWR Containment Pressure Response with Assumed Failuresof the Containment Spray Injection System and LowPressure Recirculation System ...................................PWR Containment Pressure Response with Assumed Failuresof the Containment Spray Recirculation System and LowPressure Recirculation System ...................................PWR Containment Pressure Response with Assumed Failuresof Containment Heat Removal System and Low PressureRecirculation System ............................................PWR Containment Pressure Response with Assumed Failuresof Containment Spray Injection System, ContainmentSpray Recirculation System, and Low Pressure RecirculationSystem ...................................................PWR Containment Pressure Response with Assumed Failureof the Electric Power System and the Attendant Loss ofAll Engineered Safeguards .......................................PWR Containment Pressure Response with Assumed Failuresof Safety Injection System, Containment Spray InjectionSystem, and Containment Spray Recirculation System ..............PWR Containment Pressure Response with Assumed Failuresof Safety Injection System, Containment Spray RecirculationSystem, and Containment Heat Removal System ..............PWR Containment Pressure as a Function of Leak Rate foran Assumed LOCA with No Containment Safeguards ..................PWR Containment Event Tree ............... .......................PWR Containment Pressure as a Function of Compositionand Temperature .................................................Page No.VIII-21/22VIII-21/22VIII-21/22VIII-21/22VIII-21/22VIII-21/22VIII-23/24VIII-23/24VIII-23/24VIII-23/24VIII-23/24VIII-35/36VIII-iii

List of Figures (Continued)FigureVIII 3-2VIII 3-3VIII 3-4VIII 3-5BWR Containment Pressure Response with Assumed Failureof Emergency Core Cooling System Function ..........................BWR Containment Pressure Response with Assumed Failureof Emergency Core Cooling System Operation ......................BWR Containment Pressure Response as a Function of LeakRate with Assumed Loss of Long-Term Cooling and All theNoncondensables in the Suppression Chamber ......................BWR Containment Pressure as a Function of Leak Rate withAssumed Loss of Long-Term Cooling and NoncondensablesEqualized Between the Drywell and Suppression Chamber ...........Page No.VIII-35/36VIII-35/36VIII-35/36VIII-35/36VIII 3-6 BWR Containment Event Tree ...................................... VIII-35/36VIII-iv

This report describes the results of theReactor Meltdown Task conducted at Battelle'sColumbus Laboratories as a partof a larger effort being undertaken bythe U. S. Atomic Energy Commission toevaluate the probabilities and consequencesof postulated accidents in largelight water power reactors. The detailsof the investigations performed underthis task are presented in the appendicesto this report. The body of thereport gives the results and conclusionsthat are required as input to the othertasks of the overall study.1.1 OBJECTIVESThe objective of this study has been todescribe the course of events that wouldbe expected to occur during various hypotheticalreactor meltdown accidents.The key parameters and physical processesconsidered included:a. Times for the initiation and completionof core melting*b. Steam generation rates duringmeltdownSection 1Introductioncorec. Rate and extent of reaction betweenZircaloy cladding and water (steam)d. Likelihood and potential consequencesof hydrogen burning or explodingin the containment buildinge. Probability and magnitude of steamexplosions in the reactor vessel dueto the interaction of the moltencore with water; probability of containmentfailure as a result of suchsteam explosionsf. Time at which the molten core penetratesthe reactor vesseli. Interaction of the molten core withthe concrete foundation mat of thecontainment, including the possibilityand timing of containment meltthrough.In the cases analyzed, core degradationwas a result of a loss-of-coolant accident(LOCA) accompanied by the assumedfailure of various combinations of engineeredsafety features. 1 For the pressurizedwater reactors (PWR) the initiatingevent was a complete severance of alarge cold-leg or hot-leg pipe in theprimary system. For the boiling waterreactors (BWR) severance of a recirculationline was assumed. It is expectedthat the consequences of these largepipe-break accidents will bound the consequencesof other potential accidentsin which the principal difference is anextension of the time scale.1.2 INTERACTIONSAs one facet of the AEC's investigationof reactor accidents, this Task interactedclosely with other areas of AEC'soverall study. The combinations of engineeredsafety features which are assumedto fail in the various LOCA sequencesleading to core meltdown were furnishedto Battelle by the Reactor Safety Studylocated at the AEC. The event trees andfault trees that define the various accidentsequences and failures of engineeredsafety features involved weredeveloped by the AEC. Further, the ReactorMeltdown Task has been closely associatedwith the Fission Product SourceTerm Task, also performed at Battelle'sColumbus Laboratories,2 which involvedthe definition of the size of the fis-g. Potential for steam explosions asthe molten core drops to the floorof the reactor cavity and the probabilityof containment failure due tosuch explosionsh. Pressure-time histories within thecontainment building, including thepotential and timing of containmentfailure due to internal over-presurizationiSubsequent analyses have been performedfor small-break and transient accidentsbased upon the methods described inthis Appendix. These sequences arediscussed in Appendices V and XI. Theresults of these analyses are includedin Appendix V.2With assistance from Battelle NorthwestLaboratory, Oak Ridge National Laboratory,and Aerojet Nuclear Company.VIII-I

sion product source term that would escapethe containment boundary as a functionof time for the various accidentsequences delineated by the AEC ReactorSafety Study. The results of the ReactorMeltdown Task are a necessary inputfor the evaluation of the release ofradioactivity from the containment.1.3 APPROACHThe Reactor Meltdown Task was undertakenin the realization that it may not bepossible to describe with certainty someof the complex physical phenomena thatmay occur during a meltdown accident.In those instances where data or understandingof a particular aspect were incomplete,bounding calculations wereperformed to establish the limits of uncertainty.Typically, the accident sequencewas divided into a few discretetime intervals and the events in eachinterval were modeled to the extent possible.Basic assumptions of, and inputsinto the models were then varied toestablish the bounds of uncertainty forthe course and consequences of theevents of interest.1.4 LIMITATIONSThe PWR design utilized is a three-loopplant with subatmospheric containment;the BWR design assumed a steel drywelland toroidal suppression pool containment.1 Since reactor design clearlyaffects the probabilities and consequencesof accidents, it may not be appropriateto generalize the conclusionsdrawn from these studies to all PWR'sand BWR's. Although the specific casesanalyzed were based on design basis pipebreaks, it is expected that the resultscan be used to approximate or bound theconsequences of other accidents, such assmaller pipe breaks, involving the samefailures of-engineered safety features.However, care must be exercised in extendingthese results to other accidentsto ensure that the same physical phenomenaoccur at the same relative time inthe meltdown accident as those that havebeen evaluated. The methods discussedin this report should be applicable tomeltdown accidents resulting from initiatingevents other than those specificallyassumed here.As noted above, the Reactor MeltdownTask is one portion of a larger study.The output of this task has been direct-1Basic data for the PWR and BWR investigatedare presented in Tables VIII 1-1and VIII 1-2, respectively.ed at supplying the needs of the overallprogram and a concerted effort has beenmade to keep the assumptions in thistask consistent with those of the broaderstudy. Therefore care should be exercisedin excerpting results or indrawing conclusions from the resultspresented in this report divorced fromthe greater context. In particular, theseriousness of the consequences of reactoraccidents cannot be evaluated separatelyfrom the probabilities of theiroccurrence.In reporting times for events, magnitudesof various effects, and probabilitiesof the occurrence of particularevents, a best estimate for the value isaccompanied by an error band. The erroror uncertainty bands have been derivedfrom sensitivity studies and physicalconstraints on the system and should encompassthe actual value of any givenparameter. Thus the best-estimate valuesreported here should be consideredtogether with the associated uncertaintybands.1.5 DEFINITIONSThe use of event trees to describe thecombinations of events that might occurin a reactor accident has been extremelyhelpful in studying the accidents and inproviding a common basis for the exchangeof information between the differenttasks. However, since the eventtrees were subject to change during theinvestigation, it was felt that the resultsof this report should not be relatedto the format of a specific tree.In the realization that the results ofthe report will be employed in conjunctionwith event trees the following definitionare used to distinguish betweenbranches which represent failure of engineeredsafety features and the brancheswhich represent alternative paths tocontainment failure in a meltdown accident.1. 5.1 SEQUENCEAn accident sequence involves a specifiedcombination of failures of engineeredsafety features accompanying aLOCA. If A, B, and C represent engineeredsafety features in an event tree,then Sequence ABC represents the accidentwhich involves the combined failureof the three safety features. The probabilityof the occurrence of ABC is determinedby fault tree analysis.1.5.2 SUBSEQUENCEA subsequence is defined as a path tocontainmbnt failure within a meltdownVIII-2

accident. Thus for Sequence ABC possiblesubsequences might involve meltthroughof the containment floor,overpressurization failure of the containment,or breaching of the containmentby missiles developed in a steamexplosion.In the terminology of Appendix I "EventTrees" of the Reactor Safety Study report,a sequence is a branch of an accidentevent tree and a subsequence is abranch of a containment event tree.Since engineered safety systems are generallycomposed of diverse and redundantcomponents, the inoperability of a safetysystem implies multiple failures.When safety features do operate it isnecessary to establish the level or capacityat which they perform. Sincesafety systems are overdesigned to assurea margin of safety in their performance,partial operation of systemsis often adequate to provide the requiredfunction. Therefore, a set ofminimum safeguards has been determinedto set clearly defined limits for functioningor nonfunctioning of systems.In the analysis of meltdown accidents,minimum safeguards are assumed for thosecases in which safety system performanceis indicated. The times at which systemsbegin to operate and/or fail to operatewill also have a bearing on the consequencesof a given accident sequence.For the present analyses it is assumed asystem failure, if it occurs, will takeplace at the earliest time that the particularsystem is required to function.VIII-3

TABLE ViII 1-1 PWR DATANominal powerInternal energy of waterSensible heat in the coreTotal water in the systemAug.Pressuretemperature (Excl. pres.)Reactor coolant system volumePressurizer volume, totalwatersteamThree accumulators, total volumewater volumepressuretemperatureContainment recirculation spray2 systems, flow eachContainment free volumeInitial temperatureInitial pressureDew pointPrimary system hot metalTemperature2,441 Mwt8,331 x 106246.9 x 10616.35 x 106423,200 lb571.8 F2280 psig8,387 ft 31,336 ft 38165204,350 ft 32,775 ft 3675 psig120 F3,500 gpm1.8 x 10 6 ft105 F10 psia80 F1,686,285 lb572 FBtu/hrBtuBtu3Containment Heat SinksThickness, ftArea, ft2Walls inside containmentWalls inside containmentWalls inside containmentWalls inside containmentWalls inside containmentContainment wallDomeFloor above foundation matFoundation matContainment linerWallsDomeFloorMiscellaneous metal -1,200,000 lb1.02.03.04.06.54.52.52.010.00.38 in.0.50. in.0.253,32027,60019,4005,0002,10046,74725,00011,25011,25046,74725,00011,250CoreEquivalent diameterActive heightL/DTotal cross sectional area119.7 in.144.0 in.1.20278.3 ft 2

TABLE VIII 1-1(Continued)Core (Continued)No.of fuel assembliesRods per assemblyPitchAssembly dimensionsFuel rod diameterClad (Zr-4) thicknessTotal number of fuel rodsCore weightU0 2ZircaloyMisc.Fuel pellet diameter, Region 12 and 3Fuel pellet lengthDiametral gap, Region 12 and 3Fuel density, Region 123Fuel enrichment, Region 123No.of grid spacersNeutron adsorberCladClad thicknessNo.of control assembliesFull lengthPart lengthRods per assemblyBurnable poison rodsNo.per assemblyNo. of assembliesMaterialO.D.I.D.CladBoron (natural) loadingReactor vesselI.D. of shellBelt line thickness (w/o clad)Head thicknessClad thicknessOverall height1572040.563 in.8.426 in. square0.422 in.0.0243 in.32,028226,200 lb175,60036,30014,3000.3669 in.0.3659 in.0.6 in,0.0065 in.0.0075 in.94%92911.85 w/o2.553.107Ag-In-Cd304 ss0.024 in.,53485208161268Borosilicate glass0.4395 in.0.2365 in.304 ss0.0429 g/cm157 in.7.875 in.5.0 in.0.125 in.40 ft-5 in.

TABLE V11I 1-i (Continued)Reactor vessel (Continued)InletnozzlesOutlet nozzlesWater volume with core and internalsin placeCore barrel I.D.O.D.Thermal shield I.D.O.D.Safety Injection Charging PumpsNumberDesign pressure, dischargeDesign pressure, suctionDesign temperatureDesign flowMaximum flowDesign headLow Head Safety Injection PumpsNumberDesign pressure, dischargeDesign temperatureDesign flowDesign headMaximum flowContainment Spray PumpsNumberDesign flowDesign headDesign pressureRecirculation Spray Pumps InsidcNumberDesign flowDesign headRecirculation Spray Pumps OutsidLe ContainmentNumberDesign flowDesign headRecirculation Spray CoolersNumberDesign duty, eachRefueling Water Storage TankVolumeBoron concentrationDesign pressureDesign temperatureWater temperatureContainment27.5 in.tapered to 35.4 in.29 in.3,718ft 3133.9 in.137.9 in.142.6 in.148.0 in.32750 psig250 psig250 F150 gpm600 gpm5800 ft2300 psig300 F3000 gpm225 ft4000 gpm23,200 gpm225 ft250 psig23,500 gpm230 gpm23,500 gpm249 ft455,534,520 Btu/hr350,000 gal2,500 ppmHydraulic head150 F45 FTable VIII i-1VIII-5/6

TABLE Vill 1-2BWR DATA-aRated powerSteam flow rateCore coolant flowFeedwater flowFeedwater temperatureSteam pressure in domeCoolant enthalpy at inletCore Avg.qualityRecirculation pump flowNormal primary system operating temperaturePrimary System Coolant Inventory During OperationReactor VesselSubcooled liquidSaturated liquidSteamPipingRecirculationFeedwaterSteamVolume in Reactor Vessel up to Jet Pump InletControl rods fully inControl rods fully outReactor VesselInside diameterInside heightDesign pressureDesign temperatureThickness (with clad)Vessel WeightsBottom headVessel shellVessel flangeSupport skirtOther componentsTop headTotal vesselVessel InternalsCore shroudShroud head-steam separatorCore supportTop guideFuel support piecesControl rod guide tubesJet pumpsSteam dryersCore spray spargerTotal Weight of Internal's (excluding fuel,control rods, feedwater spargers, vesselhead cooling nozzles, in-core guide tubes,start-up sources, temporary control contains)3,293 Mwt1.12456 x 1010 Btu/hr13.381 x 106 lb/hr102.5 x 106 lb/hr13.3 x 10l6 b/hr376.1 F1020 psig521.2 Btu/lb13.2% steam45,200 gpm547 F7,847.4 ft4,004.7 ft8,813.5 ft1,227.7 ft815.9 ft3,125.6 ft4,100 ft4,000 ft251 in.72 ft-ll-i/8 in.331250 psig575 F6-5/16 in.207,500 lb842,300105,80028,20065,000252,2001,501,000 lb116,900 lb139,600- 20,50015,20011,30046,35022,70090,0004,317462,000 lb333333

TABLE VIII 1-2(Continued)Primary System WeightsRecirculation pipesRecirculation pumps and motorsRecirculation valvesSteam linesSteam valvesSafety relief valvesFeedwater pipeFeedwater valvesConcrete Heat Sinks2 ft thick3 ft thick6 ft thick12 ft thickMiscellaneous metal in drywellDrywellSpherical section, diameterSpherical section, heightSpherical section, wallKnuckle section, diameterKnuckle section, heightKnuckle section, wallCylindrical section, diameterCylindrical section, heightCylindrical section, wallTop head (2:1 ell.), diameterheightVent PipesNumberInternal diameterDowncomer PipesNumberInternal diameterSubmergence, nominalwallPressure Suppression ChamberChamber inner diameterTorus major diameterDrywell Free VolumePressure Suppression Chamber Free VolumeDrywell steelNPSH RequirementsRHR pumpsHPCICore Spray124,612 lb140,00050,800164,00052,00013,00094,00025,0001,361,000 lb585,300622,0502,669,2508 x 106 lb67 ft58 ft-4-3/8 in.3/4 - 1-1/4 in.Varies5 ft. - 10-3/8 in.2-7/8 in.38 ft - 6 in.34 ft - 2-1/4 in.3/4 - 1-1/2 in.32 ft - 4 in.15 ft - 1-1/2 in.1-1/2 in.86 ft. - 9 in.962 ft.4 ft.31 ft111 ft - 6 in.3159,000 ft119,000 ft 31,550,980 lb35 ft at 11,900 gpm25 10,00018 4,00028 3,85027 3,200Table VIII 1-2VIII-7/8

Section 2PWR Analyses2.1 ACCIDENT SEQUENCES RESULTING INCORE MELTDOWNFailures of the following engineeredsafety features have been considered invarious combinations as they mightaffect the course of reactor meltdownfollowing a LOCA: electric power, containmentleakage, containment sprayinjection, containment heat removal,emergency core cooling injection andemergency core cooling recirculation.Definitions of the failures of thesesystems and their effect on the accidentsequences are given below.ELECTRIC POWER SYSTEM(EPS)Insufficient a-c or d-c power availableto the emergency buses to operate minimumdesign basis engineered safety featureequipment. Since the accumulatorsare a passive system, they are assumedto discharge in the loss-of-power caseand deliver at least part of their waterinventory to the reactor vessel. Theaccumulators by themselves are not, however,adequate to prevent reactor meltdown.CONTAINMENT LEAKAGE(CL)Inability to close one or more of thepenetrations that serve as potentialleak paths from the containment atmospherein the event of a loss-of-coolantaccident. Inability to isolate the containmentwill not in itself lead to reactormeltdown. Loss of isolation inthe event of a reactor meltdown will,however, provide a path for fission productrelease and may affect the pressuretransient in the containment.CONTAINMENT SPRAY INJECTIONSYSTEM (CSIS)Delivery of borated water to the containmentatmosphere through spray nozzlesless than the equivalent of thecapacity of one containment spray pump.The operation of the CSIS in a reactormeltdown accident affects both thepressure transient in containment andthe scrubbing of fission products fromthe containment atmosphere.CONTAINMENT SPRAY RECIRCULATIONSYSTEM (CSRS)Delivery of recirculation spray waterthrough spray nozzles less than theequivalent of the output of two recirculationspray pumps for the first 24hours or less than the equivalent of theoutput of one recirculation spray pumpthereafter. In addition to containmentpressure reduction and fission productscrubbing, the CSRS provides circulationto the hot side of the CSHX; thus, failureof the CSRS implies failure of CHRS.SODIUM HYDROXIDE ADDITION SYSTEM(SHA)NaOH injection less than a rate proportionalto the total rate of depletion ofthe refueling water storage tank. NaOHthat does not go directly to the CSIS isassumed to go into Emergency Coolant InjectionSystem (ECI) water and then tothe CSRS due to ECI overflow. Thus successfuloperation of CSIS is not a prerequisitefor successful NaOH injection.The primary role of the NaOH is theremoval of radioactive iodine from thecontainment atmosphere.CONTAINMENT HEAT REMOVAL SYSTEM(CHRS)Delivery of service water to the shellside of at least two of four containmentheat exchangers corresponding to operatingCSRS loops. Failure of CSHX constitutesloss of heat sink for the overallsystem and will lead to eventualcore melting and containment failureeven if the other safeguards are operating.EMERGENCY COOLANT INJECTION(ECI)Failure is defined as (1) delivery ofless borated water than the equivalentto that contained in two accumulators(ACC) to the primary system cold leg immediatelyfollowing a large pipe break,and (2) delivery of borated water at arate less than the equivalent to the designoutput of one Low Pressure InjectionSystem (LPIS) pump to the reactorsystem cold leg starting at about 30VIII-9

EMERGENCY COOLANT RECIRCULATION(ECR)seconds and continuing until the refuelingwater storage tank is effectivelyemptied. Operation of the High PressureInjection System (HPIS) is not requiredfor control of large pipe break accidents.Delivery of water to the reactor coolantsystem cold legs from the containmentsump at a rate less than the equivalentdischarge from one LPIS pump. This modeof operation is designed to start at theend of ECI and requires the manual realignmentof the pump suction from therefueling water storage tank to the containmentsump.2.2 ANALYSIS OF DEGRADED ACCIDENTBEHAVIOR2.2.1 BASIC ASSUMPTIONSThe initiating event for the assessmentof consequences has been assumed to be adouble-ended cold-leg break, takingplace instantaneously while the reactoris operating at full power with a corecontaining an equilibrium concentrationof fission products. This type of accidenthas been found to result in themost rapid depressurization of the primarycoolant system and leads to themost demanding requirements on the performanceof the Emergency Core CoolingSystem (ECCS). The assumption of adouble-ended cold-leg break implies thatECCS in the broken loop will not be ableto perform their function and any emergencycoolant supplied to that loop willspill out the break. If the primarysystem rupture were assumed in a hot legof the system, then, since ECC injectionis into the cold-leg piping, there wouldbe a significantly higher probabilitythat all of the emergency core coolantwould perform its intended function.A complete spectrum of accident conditionscan be conjectured which involvethe failure or partial performance ofeach of the redundant components of theengineered safety features. In performingdetailed analyses of accident sequences,it is necessary to restrict thecombinations of system performance to beevaluated to a finite tractable number.The results of the cases studied arethen considered to be representative ofa broader range of accidents that differonly slightly in detail. For the purposesof analysis, therefore, the failureand success of systems were assigneddefinitions which imply a precise levelof performance rather than the range ofperformance that could occur in an accident.In the analysis of the accidentconsequences the potential availabilityand function of only the minimum designbasis engineered safeguards were considered;i.e., if a component and/orsystem were assumed to function, onlythe minimum required capacity or capabilitywas considered. For example, fora system consisting of two componentswith 100 percent redundancy, functioningwould be taken as the operability of oneof the two components at full capacity,although conceptually the same effectcould be achieved by the operation ofboth components at some combination ofreduced capacities.For the PWR studied the minimum engineeredsafeguards would be:" Two of three accumulators" One of three high-pressure injectionpumps" One of two low-pressure injectionpumps* One of two containment spray injectionpumps" Two of four containment spray recirculationpumps" Two of four containment heat exchangers.For the purposes of these analyses, theemergency core cooling systems, if operablefor any given accident sequence,are assumed to refill the reactor vesseland keep it filled to the level of theprimary piping connections; excess capacityof the systems is assumed tospill to the containment sump. TheLPIS, HPIS, and CSIS pumps, if initiallyoperating, are assumed to function aslong as a supply of water is availableto each. Low Pressure RecirculationSystem (LPRS), if operable, starts assoon as ECI stops and continues to functionas long as there is water in thecontainment sump, barring pump cavitationdue to sudden containment depressurization.The CSRS pumps, if assumedto operate, start on demand and continueto function as long as there is water inthe containment sump, again barring cavitation.The CHRS, if assumed operating,begin to function at the same timeas the CSRS pumps and operate at theirdesign capacity as long as the CSRS operates.For those accident sequences inwhich ECI was assumed to be inoperative,potential operation of ECR was not considered.VIII-IO

Since the accumulators comprise essentiallya passive system, they wereassumed to discharge in all cases anddeliver at least part of their waterinventory to the reactor vessel. This-assumption tacitly implies that theprobability of no emergency core coolingwater being delivered to the vessel issignificantly less than the probabilityof the delivery of some water by eitherthe active or passive systems. Coremeltdown sequences in which water ispresent in the vessel following blowdownentail significant metal-water reactionsand the potential for steam explosionswhen the molten core comes into contactwith the residual water in the bottom ofthe reactor vessel. If the accumulatorswere to fail to deliver any water to thereactor vessel at the same time that thepumped ECI failed, core melting wouldtake place without significant metalwaterreaction and there would be nopossibility of steam explosions in thereactor vessel. The sequences of ECIfailure which include accumulator failureare therefore expected to be ofsignificantly lower probability and tohave lower potential consequences thanthe cases which have been investigated.If the accumulators fail but the pumpinjection operates, the consequences ofthe meltdown that may ensue would not beexpected to be significantly differentfrom those considered.2.2.2 ACCIDENT TIME SCALEFor an instantaneous double-ended coldlegbreak, primary system depressurizationwould be essentially complete in10-11 seconds. The accumulator dischargepressure of 650 lb/in. 2 would bereached in 6-7 seconds after the break,and accumulator discharge would be completedat about 30 seconds into the accident.The safety injection chargingpumps and the low-head safety injectionpumps will be activated by any of a numberof signals such as low pressurizerpressure and water level, high containmentpressure. The time at which thesesystems can be considered to be operablewill depend on the availability of offsitepower. With off-site power thepumped injection can be activated extremelyrapidly; in the absence of offsitepower the pumped injection will bepowered by the emergency diesels andwill require on the order of half a minuteto start. For the purposes of thisevaluation, both the pumped injectionand CSIS are assumed to start 30 secondsafter the break. Given the assumptionthat pumped injection and/or CSIS operate,uncertainty of a few minutes as totheir exact starting time is of littleconsequence since these systems will operatefor time periods of the order ofan hour.Both the pumped injection and CSIS systemstake their suction from the refuelingwater storage tank whose total capacityis 350,000 gal; 50,000 gal ofthis capacity is reserved for the exclusiveuse of the ECI. With both ECI andCSIS systems operating at their minimumdesign levels, i.e., one safety injectioncharging pump at 150 gpm, one lowheadsafety injection pump at 3000 gpm,and one containment spray injection pumpat 3200 gpm, the containment spray injectionwill end at 47 minutes and coreinjection will last until 63 minutes.These times are based on utilizing theentire capacity of the refueling waterstorage tank. For those cases where theCSIS is assumed to fail, ECI will lastuntil 111 minutes; if the ECI is assumedto fail, CSIS will last until 94 minutes.In the latter case 50,000 galwill remain in the refueling water storagetank. The operation of more thanthe minimum safeguards or at effectivepump capacities other than the minimumvalues assumed would, of course, alterthe time intervals.The CSRS system is slated for activationin sequence, with half the capacity beingavailable as early as 90 seconds andthe other half starting as late as 5minutes into the accident. For the purposesof the containment pressure analyses,operation of only two of the fourcontainment recirculation spray pumps of3500 gpm each was assumed, with a startingtime of 200 seconds. Here again,since the total accident times of interestare much longer than the possibleuncertainty in CSRS starting time, smallvariations in starting time would be ofno consequence to the overall results.The ECCS is designed to be switched manuallyfrom the injection (ECI) to therecirculation (ECR) mode of operationupon the receipt of a low water levelsignal from the refueling water storagetank. For the present analyses theLPRS, if operating, was assumed to startat the end of ECI as determined aboveand, barring pump cavitation due to suddencontainment depressurization, continueas long as there is water in thecontainment sump.2.2.3 CORE MELTDOWNCore meltdown in a LOCA is assumed to bethe result of failure of ECI, ECR, CSRS,CHRS, or EPS. In each case the expectedcourse of reactor meltdown is quite similar,varying primarily in time scale.Although containment pressure would beVIII-II

expected to have some effect on theboiloff of water from the core, thiseffect is small and within the bounds ofuncertainties on the progress of boiloffand meltdown. The time at which theflow of emergency cooling water stopsand boiloff begins is therefore the cnntrollingparameter that influences thetime scale of meltdown for the differentaccident sequences. Since the decayheat decreases with time from the LOCAevent, the later the time at which boiloffbegins, the more extended the meltdowntime scale. Although the boiloffwhich occurs for ECI failure begins withonly a partially flooded core ratherthan fully covered, the remainder of theboiloff and meltdown follow the samecourse as for the other accident sequences.In Appendix A, the computer code whichwas used to investigate core meltdownand the results obtained with variousmeltdown models are described. Due tothe many uncertainties associated withthe phenomena of the movement of moltenU0 2 and its interactions with other materialsit was necessary to utilize approximatemodels to describe the processesinvolved. Bounding calculationswere performed to evaluate the importantparameters in the meltdown process;among these are: fraction of zirconiumreacted, time at which melting begins,rate at which melting proceeds, and thetime at which the molten core leaves thenormal confines of the fuel region. Byvarying the assumptions and inputs inthe models a reasonable picture of themeltdown process is believed to havebeen developed.As discussed in Appendix A, it is believedthat during the period in whichcore melting is occurring, the fuel doesnot drop to the core support plate orthe lower plenum of the vessel. Rathera molten pool is formed and is supportedby a frozen crust which proceeds outwardand downward with time as more of thecore becomes molten.The analyses described in the appendixindicate that when a major fraction ofthe core is molten, -80 percent, it willno longer be possible to retain the corein the normal fuel region. Melting ofthe core support plate or failure of theupper core barrel will permit the moltenfuel to move down and contact water inthe lower plenum. The accident willthen enter a new phase with the moltencore proceeding into the reactor vesselbottom head where either a steam explosionor vessel meltthrough would occur.The time at which melting begins can beobtained from the curve marked 0% meltingin Fig. VIII A-6. In Fig. VIII 2-1the additional time required to go from0% to 80% melting of the core is plottedversus time after the LOCA at whichboiloff is initiated. The estimatedvalues and uncertainty bands were derivedfrom the results of computationspresented in Appendix A. Although theremaining 20 percent of the core willeventually melt, the melting may be extendedover an appreciable time becauseof the low peaking factor in these regions.Since this fraction of the corecontains a small fraction of the fission-productinventory, it is reasonableto assume that the principal period offission-productcore melting isrelease due to initialfrom the time of meltinginitiation to 80 percent molten. Thehydrogen generated in the meltdown accidentis estimated to be the equivalentof 75 + 25 percent of the zirconium inthe core reacted. This band providesfor potential hydrogen sources fromother reactions such as iron-water andU0 2 -steam during the time period of coremeltdown and primary vessel meltthrough.Total hydrogen release is relativelyinsensitive to the time scale of themeltdown accident.2.2.4 PRIMARY VESSEL MELTTHROUGHAppendix A describes the models used toanalyze the time required to penetratethe lower head cf the pressure vessel.The principal effect of the time requiredto melt through the pressure vesselis to delay contact of the moltencore with the floor of the containmentbuilding. Since the estimated time forprimary vessel meltthrough is much lessthan the band of: uncertainty involved incontainment meltthrough, the uncertaintyin the time for vessel meltthrough isnot considered particularly significant.If the molten core enters the lower plenumin 1-2 hours, the time required topenetrate the vessel is 60 (+40, -20)minutes. For accident sequences inwhich loss of containment heat removalcapability is the cause of reactor meltdownthe time of reactor meltdown is delayeda number of hours. In these casescontainment failure precedes meltdownand the time of pressure vessel meltthroughis not particularly significant.2.2.5 CONTAINMENT VESSEL PENETRATIONWhen the molten core (mixed with moltenzirconium, zirconium oxide, steel, andiron oxides) falls on the concretefloor, the vaporization of free waterbelow the surface of the concrete willVIII-12

cause spalling of the concrete and avery rapid penetration rate of the meltinto the concrete. As the products ofconcrete decomposition are dissolved inthe melt, the melt temperature will decreaseuntil the mixture becomes viscousor the constituents of the mixture beginto precipitate. From this point onwardin time, the progress of themelt through the concrete is controlledby the rate of decay heat generation.While there are some sequences in whichmeltdown is delayed for many hours, theyusually involve containment failure precedingmeltdown and the time required tomelt through the concrete pad does notaffect the atmospheric consequences.For the subsequences in which the meltthroughfailure mode provides the majoratmospheric release, the molten fuelcontacts the concrete at approximately2-3 hours after the LOCA. For thesecases, containment meltthrough is estimatedto require an additional 18 (+10,-5) hours. Although the molten corepenetrates the steel liner shortly afterfalling on the containment floor, substantialpressure relief is not expecteduntil a significant fraction of the reinforcedconcrete foundation mat hasbeen penetrated. As discussed in AppendixA, the molten region is expected toproceed an additonal 10-50 ft into theearth before arresting.For cases in which ECI or ECR failure iscombined with CSRS or CHRS failure, arace can occur between overpressurefailure of the containment and meltthrough.During the process of decompositionof the concrete, carbon dioxideand water vapor are generated and add tocontainment pressure. At the time ofcontainment meltthrough, the C0 2 wouldtypically contribute about 13 psi to thecontainment pressure. The effect ofthese components on containment pressureis included in Figs. VIII 2-2 throughVIII 2-6.Release of radioactive gases through theground to the atmosphere and the interactionof the molten core with groundwaterfollowing containment meltthroughis discussed in Appendix VII.2.2.6 STEAM EXPLOSIONSDuring the time in which the core ismelting down, it is possible for smallor moderate amounts of molten U0 2 tointeract with water with resultantpressure generation. Unless the amountOf U0 2 is approximately 20 percent ofthe core or more, however, the analysisdescribed in Appendix C indicates thatcontainment integrity would not bethreatened. If, during meltdown themolten material remains in a pool withinthe normal confines of the core until amajor fraction of the fuel has melted(as has been indicated in the presentevaluation), then, at the time the'lowergrid plate fails, large quantities ofmolten fuel and water can interact.Thus when a large fraction of the corehas melted, the potential for a steamexplosion in the primary vessel leadingto containment failure must be considered.The consequences of a steamexplosion which results in containmentfailure are serious not only becausecontainment is breached at an early timein the accident but also because dispersalof the core into the containmentatmosphere will result in oxidation ofthe U0 2 and an enhanced fission-productrelease.There is the possibility that missilesgenerated in a steam explosion in thereactor vessel could damage the containmentspray systems. In the design considered,the CSIS consists of two redundantfull-capacity subsystems and theCSRS consists of four redundant halfcapacitysubsystems. Since damage ofall the redundant subsystems is unlikelyand since a small margin is believed toexist between explosions with the capacityto damage the containment sprays andthose that will breach containment, thismode of spray failure was not includedin the present evaluation.A potential for steam explosions canalso occur at other times in the meltdownsequence, e.g., when the core meltsthrough the lower head of the primaryvessel. Steam explosions in the reactorcavity are not expected to threaten containmentintegrity since the resultingpressures within the containment freevolume are predicted to be small comparedto the design levels and becauseinternal structures and equipment provideeffective shielding for the containmentshell for any missiles originatingfrom such steam explosions. Sincethe occurrence of steam explosions isconsidered to have low probability, thepotential effect on fission-product releaseof steam explosions which do notresult in containment failure is ignored.Dispersal of the core withincontainment in a steam explosion couldalso result in a core configuration thatcan be cooled and contained. It isunlikely, however, that the entire corecould be dispersed into such a configuration.The posssibility of being ableto contain part or all of the molten butdispersed core is recognized but is consideredrelatively unlikely.VIII-13

Localized interactions can also occur atthe surface of the melt when it comes incontact with concrete or when it contactswater-laden gravel after penetratingthe floor of the containment building.These interactions do not have thepotential for the rapid dispersal of a'large quantity of molten material intowater or a resulting large energy release.The consequences of these reactionswill therefore be limited to thevicinity of the melt.2.2.7 HYDROGEN COMBUSTIONThe hydrogen generated in the meltdownaccident is estimated to be the equivalentof 75 + 25 percent of the zirconiumin the core reacted. As discussed inAppendix D, the hydrogen concentrationresulting from the reaction of 75 percentof the zirconium with water wouldbe outside the detonation limits for theresulting hydrogen-air-steam mixture.With 100 percent of the zirconium reacted(the upper limit of the estimatedextent of reaction), the mixture couldbe within the detonation range if thecontainment temperature is sufficientlylow. In Appendix D it is shown that theimpulse achieved in the shock wave fromhydrogen detonation is not sufficient tofail the containment.The effects on containment of a rapidself-propagating hydrogen deflagrationhave also been considered. For initialpressures higher than 32 psia, the steamconcentration in containment is abovethat at which a self-propagating reactionwill occur. At lower initial pressures,the incremental pressure from thereaction will not be sufficient to rupturecontainment.The contribution of the energy source ofhydrogen burning as it is evolved intothe containment, in combination withhigh steam pressures, can be sufficientto increase the containment pressure atthe end of the meltdown period to apoint at which containment is threatened.This effect is included in theevaluation of the containment pressureresponse.2.2.8 CONTAINMENT RESPONSEFigures VIII 2-2 through VIII 2-9illustrate the containment pressureresponses for various assumed combinationsof failures of engineered safetyfeatures; a typical design basis accidentpressure-time history is includedfor comparison. The assumptions andmethodology utilized in determining thecontainment pressure responses are discussedin Appendix A; the salient featuresof the results are discussedbelow.During a design basis accident the containmentis pressurized rapidly to apeak of 54 psia as a result of primarysystem blowdown. The CSIS and CSRSsystems reduce the pressure to subatmosphericin approximately 30 minutes.After the delivery of chilled spraywater by the CSIS stops, there is aslight increase in containment pressureto a level of about 14 psia where it ismaintained by the combined action of theCSRS and CHRS.With the failure of ECR and the othersafeguards operating at their minimumdesign levels, the initial portion ofthe containment pressure transient willnot be changed from the design basisaccident case. The generation of hydrogenby the zirconium-water reaction willlead to a containment pressure of about16 psia at the end of core meltdown.After the core melts through the bottomhead of the reactor vessel, C02 andsteam will be generated from the decompositionof the concrete by the moltencore. The steam will be condensed bythe containment sprays, but the C02 willraise the pressure to about 25 psia bythe time of containment meltthrough.With the operation of the containmentsprays the burning of hydrogen will haveno appreciable effect on the pressuretransient. This case is illustrated inFig. VIII 2-2 together with the designbasis accident.The failure of CSIS alone would not, ofcourse, lead to core melting; it wouldaffect the rate of containment depressurizationfollowing primary systemblowdown. In conjunction with the failureof ECR, the inoperability of CSISwould extend the time for which ECI isoperable and thus delay core melting,although the overall effect would not beexpected to be significant. Again, inthe presence of containment sprays, hydrogenburning would not have a significanteffect on the transient. The containmentpressure responses for thesetwo cases are given in Fig. VIII 2-3.The containment pressure responses withassumed failures of CSRS and ECR areillustrated in Fig. VIII 2-4. With theinitial failure of CSRS and theattendant failure of the CHRS, but withthe ECR operating, the containment pressurewill be decreased to about 25 psiaby the CSIS system. After the CSISstops, the pressure will increase continuously,due to the generation ofVIII-14

steam by the core decay heat untilcontainment failure by overpressurization.At the time of containmentfailure the LPRS pumps would be expectedto cavitate and core melting wouldfollow. If the LPRS fails independentlyof CSRS, core melting would take placewith the containment intact. The pressurein the containment will rise due tothe steam generated by the boiloff inthe reactor vessel. This pressure risewill stop when all the water in thereactor vessel has been vaporized.During pressure vessel meltthrough thecontainment pressure will decline due tosteam condensation on cold surfaces.After the core melts through the pressurevessel a peak containment pressureof about 80 psia would be reached as aresult of the evaporation of the waterin the reactor cavity and the initialdecomposition of the concrete. Themagnitude of this pressure is limited bythe quantity of water expected in thereactor cavity. As the molten coreprogresses through the containment foundation,the pressure would be expectedto decrease since the rate of CO 2 andsteam generation would be counteractedby steam condensation on the varioussurfaces within containment. If, in thecase of LPRS failure independent ofCSRS, the hydrogen from the zirconiumwaterreaction is assumed to burn as itis generated, higher pressures willresult and containment failure by overpressurizationcan be expected afterpressure vessel meltthrough.In the event of CHRS failure, the pressureresulting from primary systemdepressurization will be quenched,largely by the chilled CSIS water. Withcontinued steam generation from coredecay heat, the water inventory withincontainment will heat and the pressurewill increase until the containmentfails due to overpressurization. Containmentfailure will be accompanied byLPRS pump cavitation and subsequent coremelting. If LPRS fails independently ofCHRS, core melting will take place withthe containment intact; with the contributionof the hydrogen generated duringcore melting the containment pressure atthe completion of core meltdown would beabout 16 psia. When the molten corepenetrates the reactor vessel bottomhead it will come into contact with thewater in the reactor cavity and the CSRSwater. Vaporization of the water willlead to a continuous increase in containmentpressure; at the same time themolten core will be attacking the containmentfoundation mat. As shown inFig. VIII 2-5, in this case containmentfailure by overpressurization or meltthroughis predicted to occur at aboutthe same time. Hydrogen burning, if ittakes place as hydrogen is generated,would not have an appreciable effect onthe pressure-time history.Containment pressure response in theabsence of containment sprays is shownin Fig. VIII 2-6. With both CSIS andCSRS failed but LPRS operating, thepressure would increase continuously dueto steam generation from the decay heatuntil containment failure by overpressurization.Containment failure wouldbe accompanied by LPRS pump cavitationand subsequent core meltdown. If theLPRS fails independently of thecontainment sprays, core melting wouldstart while the containment is stillintact; by the end of core meltdown,however, boiloff of the water in thereactor vessel would have raised thepressure sufficiently to failcontainment. Should the containment beintact at the end of core meltdown, thepressure would decline somewhat due tosteam condensation during pressurevessel meltthrough; it would probablyfail after reactor vessel meltthroughwith the pressure increase from theinitial rapid decompositon of concrete.If, in the absence of containmentsprays, LPRS failure is accompanied byhydrogen burning, the containmentfailure pressure would be exceeded evenbefore the end of core meltdown.Figure VIII 2-7 illustrates the containmentpressure response with an assumedloss of electric power and the attendantloss of all active engineered safeguards.As would be expected, there isno reduction in the post-blowdown pressure,and the boiloff of the waterinjected from the accumulators raisesthe pressure to a peak of about 75 psia.The peak pressure is limited by theamount of water available for vaporization.As the core melts through thebottom of the reactor vessel there is adecrease in containment pressure due tocondensation of steam by the variousstructures within containment. Afterthe molten core falls to the bottom ofthe reactor cavity, pressure increasesdue to the initial rapid rate ofconcrete decomposition. Subsequentlythe pressure would be expected todecline as the melt works its waythrough the concrete foundation mat.Upon containment meltthrough, the pressurewould be controlled by the statichead of the ground water. If (in theabsence of all active safeguards) coremelting is accompanied by hydrogen burning,the containment pressure wouldreach about 100 psia at the end of coremeltdown. At that level containmentfailure could be expected.VIII-15

The effects of ECI failure in combinationwith various containment safeguardfailures on containment pressure responseare illustrated in Figs. VIII 2-8and VIII 2-9. If ECI is the onlyfailure, the containment sprays wouldreduce the pressure as in the designbasis accident. Hydrogen generationduring the meltdown in combination withthe noncondensables initially presentwould limit the minimum pressure attainableto about 16 psia. The generationof CO 2 from the concrete after pressurevessel meltthrough would raise thepressure to about 25 psia; after containmentmeltthrough the pressure wouldbe controlled by the external hydrostatichead. If, in the case of ECIfailure, the hydrogen were to burn, theenergy release would be accommodated bythe sprays and the net result would be aminimum pressure of about 13 psia. WithCSIS as well as ECI failing, the steamreleased by the primary system blowdownwould be condensed by the CSRS acting inconjunction with CHRS, although thepressure would be at an elevated levellonger than for the previous case. Thepressure prior to containment meltthroughwould again be controlled by thenoncondensables. The containment pressureresponse with ECI, CSIS, and CSRSfailure would be identical to that forcomplete loss of electric power whichwas discussed previously.If the LOCA is accompanied by thefailure of ECI and CHRS, the containmentpressure will be reduced rapidly by theoperation of CSIS and CSRS to a level ofabout 16 psia at the end of coremeltdown. After the CSIS stops and themolten core has penetrated the bottom ofthe reactor vessel, the water inventorywithin containment will be heated tosaturation and partially vaporized,leading to a rise in pressure. As themolten core is boiling off water, it isalso attacking the concrete foundationmat; containment failure by the twocompeting mechanisms is predicted tooccur at about the same time. Theburning of hydrogen as it is generatedwould not have a significant effect onthis pressure transient since the energyreleased would be spread over asubstantial water inventory. In theevent of the failure of the CSRS inconjunction with ECI, the blowdownpressure would be effectively suppressedby the CSIS, though not as. quickly as inthe previous case. After the CSIS stopsand the molten core drops to the floorof the reactor cavity, the pressurewould again increase due to the evaporationof water and the decomposition ofconcrete. In this case the quantity ofwater in the reactor cavity would berelatively small and the pressureresulting from its vaporization wouldnot be sufficient to fail containment.Thus the expected failure mechanismwould be containment meltthrough. Hydrogenburning, since it would beexpected to take place while the CSIS isoperating, would again have littleeffect on the overall pressuretransient.Figure VIII 2-10 illustrates the effectof leak rate on containment pressuretimeresponse in the absence of anycontainment sprays. It can be seen thatleak rates of about 200 v/o per day andgreater will preclude containment failureby overpressurization. A leak rateof 200 v/o per day would require a hole3.6 inches in diameter or its equivalent.Within this context the designbasis leakage of 0.1 v/o per day can beconsidered negligible.2.2.9 LEAK RATE FROM CONTAINMENTThe release of fission products to theenvironment is controlled by the leakrate from the containment. In theperiod of time preceding containmentfailure, the leakage is determined bythe internal pressure of the containmentand the hole sizes associated witheither containment isolation success orcontainment isolation failure. At thetime of containment failure by steamexplosion, overpressurization, or meltthrough,the fraction of the contents ofthe containment that are releaseddepends upon the containment pressure.After containment failure has occurredby one of these mechanisms, the leakrate from the containment will be determinedby the generation rate ofnoncondensing gases. If sprays areoperating with subcooled water, steamgenerated within containment will becondensed and the generation of hydrogenand C02 will provide the driving forcefor leakage. When sprays are notfunctioning, the steam generated withincontainment will also result in leakage.The assumptions used in calculating leakrate from containment for the PWRsubsequences are described inAppendix A.2.3 SUBSEQUENCE PROBABILITIESThe event tree shown in Fig. VIII 2-11illustrates the five paths to containmentfailure that are postulated for ameltdown accident in a PWR: steamexplosion in primary vessel, containmentisolation failure, overpressurization ofthe containment resulting from hydrogenburning, overpressurization of contain-VIII-16

ment from steam generation and carbondioxide production, and containmentfloor meltthrough. It is expected thatcontainment floor meltthrough wouldaccompany each of the other containmentfailure events. However, meltthroughwould not significantly add to theatmospheric fission-product source termin the cases where other containmentfailure modes have preceded it. It isassumed that the different paths tocontainment failure are exclusive. Forexample, if a steam explosion introducesa large hole in containment, it is notnecessary to consider containment overpressurization.For a given sequence some of these pathsto containment failure may not becredible. These subsequences areassigned a probability of zero for thatsequence. In the following paragraphsthe method of obtaining the probabilitiesof the five subsequences (Pa, P1,Py, P 6 , P.) for the different accidentsequences are described.For the purposes of estimating andcombining uncertainties in the predictedvalues of the various parameters, fourtypes of statistical distributions areconsidered: normal, log normal, skewed,and skewed log. The normal distributionis characterized by a mean value and astandard deviation which can be used toevaluate the probability of exceedingany given value by means of normaldistribution tables. The log normaldistribution is distinguished by a meanvalue and a standard deviation expressedas powers to the base ten. The skeweddistributions are characterized by amean value and two standard deviations,one representing the distribution forvalues greater than the mean and theother for values less than the mean.2.3.1 DEFINITIONSPPyP 3P4P2 =P3P4Irelative probabilities ofeach of the containment failuremechanismsprobabilityeach decisioncontainmenttreeof failure atpoint in thefailure eventPfw = probability of a large massof molten U02 (>20 percentcore) contacting a similarmass of water during themeltdown processasxsprobability of a dispersalevent occurring when moltenU02 falls into water (effectiveparticle size

P1 = Pfwasxsa cf" (VIII 2-1)Calculations discussed in Appendix Aindicate that molten fuel cannot travelfar through cooler regions of the corewithout refreezing. The accumulation ofa large pool of molten fuel is thereforequite likely. The uncertainties in Pfwinclude potential for extensive vaporizationof the fuel and the possibilityof no water in the lower plenum at thetime when the molten fuel and waterwould interact. The latter two factorsare expected to be small for the PWRaccidents considered. For these reasonsthe probability of Pfw is estimated tobe near unity with a small uncertainty.Pfw = 10-0.05 (±0.05)Experience with explosive interactionsbetween molten materials and water isdescribed in Appendix A. Insufficientdata exist for molten U0 2 in water topredict under any given conditionswhether or not an explosive interactionwill occur. From the observed behaviorof other molten materials and water,particularly with regard to thesignificance of subcooling to knowndispersal mechanisms, it is felt thatthe likelihood of steam explosionsoccurring under the given conditions inthe primary vessel is small. Because ofthe lack of understanding of the U02-water interaction, however, the possibilityof an explosion event must berecognized. The range of probabilitywhich has been assigned to asxs isintended to represent the probability ofa dispersal event which is expected tobe small but could be moderate. Thenumerical value assigned to thisprobability is largely a matter ofjudgment based on the history ofphysical explosions and discussions withresearchers in this and related areas.coefficients. The results of thesestudies indicate that if the dispersalmechanism results in effective particlesizes less than 4000 v', containment willbe breached for insertion rates of fuelinto the water of 200 percent core massper second or greater. The minimumquantity of U0 2 that must be involved inthe interaction to lead to containmentfailure is about 20 percent of the core.Although it would be unlikely to observedropping rates of U0 2 into water of themagnitude of 200 percent per second, twophenomena enhance the likelihood ofobtaining such effective insertionrates. Delay times are often observedin steam explosions between the time atwhich the hot fluid is introduced andthe time of the explosion. During theperiod of the delay a large mass of hotfluid can be introduced into the water.Also, as the first amount of U0 2interacts with the water in the lowerplenum, the water surface will swellrapidly engulfing the remainder of themolten fuel. This fuel may then be ableto participate in the explosion event.Other uncertainties in the probabilityof failing containment in an explosionevent involve the effective quantity ofwater with which the molten fuelinteracts and the manner in which therapidly expanding mixture interacts withthe upper head. The estimated value foracf reflects results which indicate thatthe conditions required to rupturecontainment are difficult to achieve butmight be expected with a moderately lowprobability. The numerical value ofthis probability is derived on the basisof the calculational results discussedin Appendix C.acf = 10-1 (+0.5, -1)The combined and rounded probability ofthis path to containment failure isP Z 10-2 (+1, -2)asxs - 10-1 (+0.5, -1)Appendix C describes a series of calculationsperformed to determine theconditions under which containmentfailure would result from a steamexplosion. Key parameters consideredare effective surface area produced bythe dispersal mechanism, time requiredto introduce the molten core into thewater, and magnitude of heat-transferAlthough small differences would existin the consequences of a steam explosionwith and without containment isolation,for the purposes of this analysis steamexplosion consequences are considered tobe independent of the condition ofcontainment isolation.P = P (VIII 2-2)VIII-18

2. 3. 3 CONTAINMENT MELTTHROUGH WITHCONTAINMENT ISOLATION FAILUREAs illustrated in Fig. VIII 2-10, leakrates of 200 v/o per day (3.6-in.-diamhole) and greater would preclude containmentfailure by overpressurizationeven in the absence of any containmentsprays. This value was therefore takenas a convenient dividing point betweenleak rates considered as containmentisolation success and containment isolationfailure. The design leak rate forthe type of containment considered is0.1 v/a per day, with temporaryincreases up to 1 v/o per day consideredpermissible. Considering all possibleleak rates up to 200 v/o per day, theprobability of a leak greater than thenormal operating range would be expectedto be much less than the probability ofa leak within the operating range. Thusa representative leak size for thisregion should be within the operatingrange. The upper limit on the normaloperating range of leakage, or -1 v/oper day, was, therefore, chosen tocharacterize containment isolationsuccess.For hole sizes greater than the 3.6-in.equivalent diameter, the average size isin the range 8-10 in., corresponding toa leak rate of about 1000 v/o per day.For leak rates of this magnitude, theexact size of the hole in containmenthas little effect on the consequences ofa meltdown accident. The driving forcefor fission-product leakage from containmentconsists of the steam andnoncondensables that are generatedduring the course of the accident.Except for short time periods, theproduction rate of gases and vaporsavailable for leakage is less than thechoked flow rate through a 8-10-in. diamhole. Thus the leak rate is effectivelydetermined by the net production ofgases and vapors and is independent ofthe exact hole size. Accident sequencesinvolving containment isolation failurehave accordingly been represented by aleak rate of 1000 v/o per day.The probability of containment isolationfailure will then be given bywhere2 = 7• for Li >3.6 in., (VIII 2-3)1'Ii = probability of leak size Li.The probability that loss of isolationwill be the principal mode of containmentfailure will bePa = (1 - P 1 ) P 2 . (VIII 2-4)2.3.4 HYDROGEN COMBUSTIONThe containment pressure at the end ofmeltdown is calculated including theburning of hydrogen to evaluate thepotential for overpressure failure. Thecontribution of hydrogen from 75 percentequivalent zirconium-water reaction isshown in Figs. VIII 2-4, VIII 2-6, andVIII 2-7. The probability of failure,Phf, is assessed assuming a normaldistribution of failure pressures with astandard deviation of 15 psi around theestimated failure pressure of 100 psia.For the pressure transients in whichmore than one peak is predicted, themaximum value of the pressure is used.The likelihood of hydrogen burning as itis evolved from the primary vessel, Phb,must also be considered. The locationof the break (hot leg or cold leg) willaffect the temperature at which thehydrogen leaves the vessel. For a coldlegbreak, the hydrogen will passthrough the unbroken loops and be cooledto the temperature of the steam generatorsbefore entering the containmentatmosphere. Combustion will require theavailability of oxygen as well asignition source. In a hot-leg break,the hydrogen will probably enter thecontainment at a temperature above theautoignition temperature. Whether thehydrogen burns or not will then dependupon the availability of oxygen at thebreak. If the hydrogen is transported asignificant distance before encounteringoxygen, its temperature will decreaseand an ignition source may again berequired for combustion to take place.Another possibility is a condition ofonly partial burning. An overall valueof Phb = 0.25 (+0.25) has beenemployed.P 3 = hfPhb,P = P 3 (l - P 2 ) (1 - Pl)-(VIII 2-5)(VIII 2-6)2.3.5 CONTAINMENT OVERPRESSURIZATIONIn some cases the rise in containmentpressure is not water-limited and theVIII-19

internal pressure would reach extremelyhigh levels if the containment were toremain intact. In other cases theinternal pressure literally runs out ofsteam at a point near the band offailure pressure. In the formersequences, the failure of containment isconsidered a certainty. In the lattersequences, the probability of failure,P... is assessed assuming a normaldistribution of failure pressure. Againa standard deviation of 15 psi and afailure pressure of 100 psia are used.The calculated peak pressure includesthat from the CO 2 and the steamgenerated fiom interaction of the moltenfuel with the concrete floor.Since overpressure of the containmentand containment meltthrough can occur atapproximately the same time, it isnecessary to consider the competitionbetween the two events. Since theuncertainty in containment meltthroughtime is great, this time has beentreated as having a skewed distributionabout the estimated time, 18 (+10,-5 hours). The probability ofcontainment overpressurization occurringprior to the time at which containmentmeltthrough can relieve pressure, Pom,is determined by using normal distributiontables.P4op omP6 =P 4 (1 - P 3 ) (i - P 2 ) (1 - P1).2.3.6 CONTAINMENT MELTTHROUGH(VIII 2-7)(VIII 2-8)If the other modes of containment failureare avoided, containment meltthroughis treated as a certainty.PC =( - P 4 ) (1 - P(3)1( - P2 (1 - Pl).(VIII 2-9)VIII-20

400300,-42001-400U41000Time at Start of Boiloff, hrFIGURE VIII 2-1Time Required for 80 PercentCore Melting from the Onsetof Melting140-- Design Basis Accident........... LPRS failure1201008004020. Containment.. me1tthrough10FIGURE VIII 2-2PWR Containment Pressure Response During theDesign Basis Accident and with the AssumedFailure of the Low Pressure RecirculationSystem

140 '- CSIS failure.......... CSIS and LPRS failure120100_________________________Faiqlure pressure80 -r60 -44020Containmentmeltthrough10a a I ri.jaaa.~ aL I . ý I -'- ýAccident Time, mini..1 4IFIGURE VIII 2-3PWR Containment Pressure Response with AssumedFailures of the Containment Spray Injection Systemand Low Pressure Recirculation System140 1 - CSRS failure. CSRS and LPRS failure120....... CSRS and LPRS failure with H2 combustion1000.60U .4020Accident Time, min.FIGURE ViII 2-4PWR Containment Pressure Response with AssumedFailures of the Containment Spray RecirculationSystem and Low Pressure Recirculation System

140- CHRS failure1201Go4,.S8060o402010Accident Time, minFIGURE VIII 2-5PWR Containment Pressure Response with AssumedFailures of Containment Heat Removal System andLow Pressure Recirculation System140120- CSIS and CSRS failure........... CSIS, CSRS, and LPRS failureCSIS, CSRS, and LPRS failure withH2 combustionoverpressurizationContainment10080I60~40ContainmentIme tthoh2010 I,,,,A T 102Accident Time, minI., I I I ' I i - 4FIGURE VIIi 2-6PWR Containment Pressure Response with AssumedFailures of Containment Spray Injection System,Ccntainment Spray Recirculation System, and LowPressure Recirculation SystemFig. VIII 2-1 - Fig. VIII 2-6VIII-21/22

140 "EPS failure-. -. - . EPS failure with H 2 combustion120100 -80Failure nressure/,\ .*i~/ \~'%~........................... \:60S""Containment-. meltthrough0UC4020t0FIGURE VIII 2-70"2Accident Time, minPWR Containment Pressure Response with AssumedFailure of the Electric Power System and theAttendant Loss of All Engineered Safeguards104140120too880- 60•0Accident Time, minFIGURE VIII 2-8PWR Containment Pressure Response with AssumedFailures of Safety Injection System, ContainmentSpray Injection System, and Containment SprayRecirculation System

140120S100'0S806060402010 r?_________Accident Time, minFIGURE, Vill 2-9PWR Containment Pressure Response with AssumedFailures of Safety Injection System, ContainmentSpray Recirculation System, and Containment HeatRemoval Systemr.Accident Time, minFIGURE VIII 2-10PWR Containment Pressure as a Function of LeakRate for an Assumed LOCA with No ContainmentSafeguards

CRV5SECL5,CR- BY,CR- Or'6CR-MTC6P =2IP (I - P1)P (I P I)(I P F 2 )(1P )PC (1 1)1 )( 23)(1- P4FIGURE VIII 2-11PWR Containment Event Tree4LFig. VIII 2-7 - Fig. VIII 2-11VIII-23/24

Section 3BWR Analyses3.1 ACCIDENT SEQUENCE RESULTING INCORE MELTDOWNFailures of the following engineeredsafety features have been considered invarious combinations as they might affectthe development of degraded coreconditions in the event of a recirculationline break in a BWR: electricpower, loss of vapor suppression, lossof containment isolation, emergency corecooling system operation, emergency corecooling system function, scram, andlong-term cooling. The definitions ofthe failures of these systems and theirinfluence on the course of the accidentare given below.ELECTRIC POWER SYSTEM (EPS)Failure to provide a-c and d-c power forthe operation of the engineered safetyfeatures required to mitigate the initiatingevent.VAPOR SUPPRESSION SYSTEM (VSS)Failure of the vapor suppression systemto condense an adequate quantity ofsteam discharged from the drywell to thewetwell to prevent an overpressurizationof the primary containment sufficient toresult in its rupture. Loss of vaporsuppression with the attendant failureof containment may not of itself lead tocore meltdown since the other safeguardscould still be operable, assuming failuredoes not lead to the loss of waterfrom the torus.CONTAINMENT LEAKAGE(CL)Three cases of containment isolationsfailures are considered:a. Minor failures of Primary ContainmentSystem (PCS) isolation whichresult in leakage to the SecondaryContainment System (SCS) of lessthan 100 v/o per day; this is equivalentto approximately a l-in.-diamhole.b. Failures of the PCS which result inleakage to the SCS in excess of 100v/o per day but less than 3600 v/oper day. With these leak rates andthe loss of long-term cooling of theECCS pumps will fail due to the lossof the required Net Positive SuctionHead (NPSH); however, these leakagesare sufficiently high to preventprimary containment failure by overpressurization.c. Leakage greater than 3600 v/o perday (approximately a 6-in.-diamopening). This will result insecondary containment failure duringthe blowdown phase of the accident.EMERGENCY COOLING INJECTION(ECI)Two definitions for emergency core coolingoperational failures have beendeveloped:a. Failure of (a) any Core Spray InjectionSystem (CSIS) pump with no LowPressure Coolant Injection System(LPCIS) operational or (b) more thantwo CSIS pumps and one LPCI pump; aswell as their associated controls,piping, etc.b. Failure of (a) two CSIS pumps withno LPCIS operational or (b) threeLPCI pumps with no core sprayoperational or (c) more than threeLPCI pumps and more than three corespray pumps; as well as their associatedcontrols, piping, etc.For success, the minimum design ECCS(Emergency Core Cooling System) mustoperate until the core is reflooded atwhich time one ECCS pump is sufficientto maintain the water level in thevessel.EMERGENCY COOLING FUNCTION(ECF)Failure to provide the required quantityof water to the reactor core to preventcore melt, due to structural failuressuch as: core shroud failure, jet pumpfailure, or other damage which mayresult from initial blowdown.SCRAMThe failure to insert control rods priorto reflooding the core with ECCS inorder to ensure subcriticality of thereactor.VIII-25

LONG-TERM COOLING(LTC)Failure to provide core cooling throughat least one Residual Heat Removal (RHR)loop consisting of one pump, one heatexchanger, one high-pressure servicewater pump, one emergency service waterpump and the associated controls,valves, piping, etc., or by use of onecore spray pump in conjunction with theabove. Loss of long-term cooling willlead to overheating of the suppressionpool water, overpressurization of thecontainment, failure of the ECCS, andcore meltdown. The particular sequenceof events may vary if failures of othersafeguards accompany loss of long-termcooling.3.2 ANALYSIS OF DEGRADED ACCIDENTBEHAVIOR3.2.1 BASIC ASSUMPTIONSThe governing accident case for BWRsafety evaluations is generally taken tobe a double-ended rupture of a primarycoolant recirculation line. This is thesituation selected as the starting pointfor the present degraded accident consequenceassessment. The design of BWRsystems normally provides a diversityand redundancy of standby core coolingsystems as protection against the entirespectrum of potential LOCA's. For thepurpose of this study, however, theoperability of only the minimum designbasis safeguards has been assumed. Specifically,for the double-ended recirculationline break, the minimum requiredsafeguards were assumed to consist oftwo Low Pressure Coolant Injection(LPCI) pumps, one Core Spray Loop whichcontains two Core Spray Injection (CSI)pumps, and one Residual Heat Removal(RHR) loop. No High Pressure CoolantInjection System (HPCIS) or ContainmentSpray operability were considered. Itis recognized that several other combinationsof the total available safeguardscould provide the desired levelof protection.For purposes of degraded accident evaluationthe ECCS, if operable, are assumedto refill the reactor vessel and keep itfilled to the level of the jet pumps;excess capacity is assumed to spill tothe floor of the drywell and flow backto the suppression pool. Failure ofemergency core cooling operation isinterpreted as no water injected intothe reactor vessel. Loss of emergencycore cooling function is taken as deliveryof the requisite quantity of waterto the reactor vessel but failure toprovide the desired cooling due to sometype of structural failure. LTC, ifavailable, is assumed to be manuallyactivated at 600 seconds into the accident.3.2.2 ACCIDENT TIME SCALEFor the above accident and safeguardsoperation, the expected sequence ofevents would be as follows:EventDouble-ended recirculationline break with attendantloss of normal auxiliarypowerReactor vessel low waterlevel initiates main steamisolation valves closureReactor scram signal fromhigh drywell pressureDiesel generators signaledto start. Primary containmentisolated. Corespray and low pressurecoolant injection systemssignaled to startPrimary system depressurizationcomplete. Corespray system operatingLow-pressure coolantinjection system operatingSignal for emergency servicewater pumps to startEmergency service waterpumps at rated speedCore effectively refloodedManual activation ofresidual heat removalsystemMaximum suppression pooltemperature reachedTime,sec0.00.51.03.030.043.055.060.095.0600.04.6 x 104The above sequence of events shouldbe representative of most large BWRsystems.3.2.3 CORE MELTDOWNAccident sequences involving the failureof the ECCS to operate or the loss ofelectric power lead to dry meltdowns inwhich the core heats essentially adiabatically.Since essentially all theprimary system water is lost during theblowdown, there is no significant quantityof steam for the zirconium-waterVIII-26

eaction in these cases. There is alsolittle driving force to remove fissionproducts that have been released to theprimary vessel. As in the PWR cases,the core is expected to remain above thegrid plate until a large fraction ofthe core has melted. From Fig. VIII A-7the time after the LOCA has occurred tothat at which a large fraction of themolten core would be expected to movedown into the lower head is estimated tobe 150 + 30 minutes.Core degradation due to loss of emergencycore cooling function is particularlydifficult to evaluate and would be expectedto depend on the particular natureand extent of damage encountered.The present analyses indicate that coremelting in the presence of emergencycoolant whose heat-removal effectivenesshas been impaired would take place intwo phases. In the hottest portions ofthe core, decreased cooling effectivenesswould result in heat generation dueto the zirconium-water reaction and thusaccelerate the heatup and melting ofthose regions. The cooler portions ofthe core would take some time to reachtemperatures where zirconium-water reactionswould be significant and thuswould reach melting at significantlylater times. The time after LOCA forthe hotter half of the core to melt isestimated to be 20 + 10 minutes and forthe cooler half it is 150 + 30 minutes.The cases in which LTC failure is thecause of reactor meltdown are similar tothe PWR accidents analyzed. The corehas been quenched and is partiallyflooded at the time the emergency coolingpumps fail. The water then boilsoff and the core melts. The time requiredto achieve 80 percent core meltingis quite similar for the BWR and PWRboiloff meltdown cases. This time cantherefore be estimated from Fig. VII A-6as a function of time at which boiloffbegins. For the BWR wet meltdowns thefraction of zirconium reacted is estimatedto be 50 + 15 percent of the coreinventory.3.2.4 PRESSURE VESSEL MELTTHROUGHThe lower plenum of the BWR is partiallyfilled with control rod guide tubes andthe lower head contains a number of controlrod drive penetrations. When themolten core leaves the fuel region itmay partially refreeze on control rodguide tubes or penetrate these tubesprior to failure of the bottom head. Ifthe time required to heat the entirevessel to 2400 F with the decay heat isconsidered an upper bound to the timerequired to penetrate the primary vessel,then penetration must occur in lessthan 72 minutes. The best estimate forthe meltthrough time after the coredrops into the lower plenum is 30 + 20minutes for the dry meltdown cases. Forthe wet meltdown cases, an additional 30minutes will be required for vaporizationof the water in the bottom headand, for later onset of core melting,the meltthrough time will be furtherextended due to the decrease in decayheating.3.2.5 STEAM EXPLOSIONSIf during the course of a meltdown accidentlarge quantities of molten corematerials can come into contact withwater, the possibility of steam explosionsmust be considered. The analysesdescribed in Appendix C indicate that if20 or more percent of the core whenmolten comes in contact with water,steam explosions in the pressure vesselwith the potential to fail containmentcould result. For meltdowns which are aconsequence of loss of electric power orfailure of the ECCS to operate therewould be little or no water in the reactorvessel and hence little likelihoodof steam explosions within the reactorvessel. If, however, meltdown were toresult from failure of the ECCS to functioneffectively, failure of the reactorto scram, or emergency coolant pump cavitation,there would be the possibilityof steam explosions as the molten corecomes in contact with the water in thereactor vessel. If core meltdown takesplace with the containment alreadyfailed, as would be the case with theloss of vapor suppression or loss ofLTC, the occurrence of a steam explosionwould not be expected to have, a majoreffect on the accident consequences.There is also the possibility of steamexplosions upon the meltthrough of thereactor vessel bottom head as the massof molten material falls into the wateron the drywell floor. The configurationof the drywell to wetwell vents wouldpermit approximately a 2-foot depth ofwater to remain on the drywell floor;the temperature of this water couldrange from subcooled to saturated, dependingon the particular accidentsequence and point in time of interest.If the molten materials were to fallinto the water in relatively small quantitiesand over a period of time, theywould simply be quenched and the steamgenerated in the process would be condensedin the suppression pool. If, onthe other hand, a significant fractionof the core were to fall into the watercoherently, there would be the potentialVIII-27

for a violent interaction. Such violentinteractions are of concern if they havethe potential to fail containment and ifthey occur while the containment isstill intact.3.2.6 CONTAINMENT VESSEL PENETRATIONWhen the molten core together with someof the reactor internals and part of thevessel bottom head fall to the floor ofthe drywell, vaporization of the waterthere as well as decomposition of theconcrete will ensue. Any water on thedrywell floor will be vaporized by themelt and in the absence of drywell failurewill be forced into the suppressionpool and condensed. The penetration ofthe concrete proceeds in a manner similarto that previously described for thePWR's. Initially the vaporization offree water below the surface of the concretewill cause spalling and a rapidrate of penetration of the concrete bythe melt. As the products of concretedecomposition are dissolved in the melt,the melt temperature will decrease tothe point where some of the constituentsof the mixture begin to precipitate.From this point onward in time the progress,of the melt through the concrete iscontrolled by the rate of decay heatgeneration. The initial rapid spallingis estimated to penetrate a 1.5-ft depthof concrete in about 20 minutes, thesubsequent meltthrough proceeds at arate of about 0.5 ft per hour. The CO 2and water vapor generated by the decompositionof concrete will be forced intothe suppression pool where the latterwill condense. The CO 2 together withthe other noncondensables will collectin the suppression chamber gas space andlead to containment failure by overpressurization.Figure VIII 3-1 gives thepartial contributions as well as combinedpressures of the various constituentswithin the containment. The CO 2pressure contribution is based on thedecomposition of approximately a 20-ftdiamx 8-ft-high cylinder of concretebelow the reactor vessel within thedrywell vessel. The quantity of CO 2together with the other noncondensableswhen confined in the suppression poolgas space will result in a pressure morethan sufficient to rupture the containment.If the melt progresses radiallyoutward as well as down, as might beexpected, a larger quantity of C02 wouldbe generated prior to the time that themelt reaches the steel vessel. Further,since the bottom of the drywell vesselis embedded in concrete, meltthrough ofthe steel vessel would not necessarilyimply a breach of containment. Thus,with less than about 100 percent per dayleakage, containment would fail by overpressurerather than meltthrough of theconcrete foundation.Interaction of the molten core withgroundwater following containment meltthroughis discussed in Appendix VII.3.2.7 HYDROGEN COMBUSTIONBWR containment atmospheres are inertedby purging with nitrogen until theoxygen content is below 5 percent. Theoxygen is maintained at or below thislevel during normal operation. At thisoxygen concentration the containmentatmosphere is outside hydrogen flammabilitylimits, regardless of the quantityof hydrogen in the atmosphere. Radiolyticdecomposition of water during thecourse of the accident would add oxygenas well as hydrogen to the atmosphere.During the accident times of interesthere, i.e., of the order of a day orless, radiolytic decomposition of waterwould not result in a flammable mixture.For longer time periods, the containmentwould have failed by other mechanismsand hydrogen combustion would not be ofconcern. If the hydrogen generated bythe zirconium-water reaction is includedin the consideration, the atmospherewill be outside the flammability rangeeven if oxygen in the containment approachesthe equilibrium value achievableby radiolysis.Thus it is concluded that hydrogen combustioncan be eliminated as a mechanismof containment failure in the BWR degradedaccident sequences. Some of thebases for this conclusion are furtherdiscussed in Appendix D.3.2.8 CONTAINMENT RESPONSEFunctioning of the vapor suppressioncontainment as designed leads to relativelylow accident pressures within thecontainment except for the early portionsof the blowdown. If, however,condensation of the steam generated byprimary system depressurization does nottake place, for whatever reason, pressuressignificantly in excess of designlevels would result. Assuming thatsteam condensation does not take placeand that the entire free volume of thecontainment is available, primary systemdepressurization would result in a pressureof about 250 psia within the reactorcontainment design assumed for thisstudy. As this is well above the 175 +25 psia failure pressure determined iiiAppendix E, primary containment rupturewould be expected. If this failure didnot lead to loss of suppression poolwater, but only in venting of the steamfrom the containment, it may still beVIII-28

possible for the standby core coolingsystems to operate and thus prevent coremelting.The failure of long-term cooling withall the other safeguards systems workingwould lead to a gradual increase in thesuppression pool water temperature andattendant gradual containment pressureincrease. With a design basis containmentleak rate, the loss of LTC willlead to containment failure by overpressurizationin about 25 hour. With largerleak rates the time to failure couldbe extended; inloss of standbysuch cases, however,core cooling systempumps due to cavitation must be considered.This aspect is discussed in alater section of the report.The minimum safeguards for the purposeof this evaluation were assumed to betwo LPCI pumps, two CSI pumps, and oneRHR loop. This combination is sufficientto effectively reflood the coreand more than sufficient to maintain thecore water inventory. With the combinationof bottom flooding and core spray,decay heat would be accommodated withlittle or no net steam generation.Under degraded ECCS operation, there maybe continued generation of steam fromthe core; this could lead to the drywellpressure remaining about 3 psi higherthan that in the suppression chamberwith no equalization of the noncondensables.Under such conditions the containmentwould remain at an elevatedpressure for an extended time.The location of the noncondensable gaseswill have a significant effect on theoverall BWR containment pressures followingan accident. Figure VIII 3-1gives the partial as well as combinedpressures of the various constituents ofthe containment atmosphere as a functionof containment temperature and dispositionof the noncondensables. The pressurecontribution due to hydrogen isbased on the reaction with water of 50percent of the zirconium in the core(equivalent of 83 percent of the fuelcladding); the carbon dioxide contributionassumes the decomposition of a 20-ft-diam by 8-ft cylinder of concrete atthe bottom of the drywell. It may benoted that with minimum design basissafeguards operating, the containmenttemperature range of interest is 130 -190 F.Figure VIII 3-2 gives the expected containmentpressure response for an accidentsequence involving failure ofemergency core cooling function. Theinitial rapid pressure decrease isassociated with the condensation ofsteam by the suppression pool and theequilibration of the noncondensables.As the core overheats and melts, thecontainment pressure increases due tothe generation of hydrogen by thezirconium-water reaction. The reactionof the cladding with water has beenassumed to have run its course by theend of core meltdown, i.e., 150 minutes;thus there is little pressure changeduring the time of pressure vessel meltthrough.Assuming no damaging steam explosionas the molten mass falls to thedrywell floor, pressure vessel meltthroughwill be followed by vaporizationof the water at the bottom of the drywelland the rapid attack of the concretefloor. During this period, thepressure within the system will increaserapidly and lead to large flow rates ofsteam and noncondensables from the drywellto the suppression pool, where theformer will be condensed. The rapidpressure increase occurring between 210-230 minutes in Fig. VIII 3-2 is associatedwith the transport of all the noncondensablesto the suppression chambergas space. After 230 minutes the pressurewill continue to increase but at asignificantly lower rate due to the continuedevolution of CO 2 from the decompositionof concrete. This will continueuntil the containment fails.Containment pressure response for failureof emergency core cooling operationis shown in Fig. VIII 3-3. In this casethere would not be enough water availablein the reactor vessel for an appreciableextent of zirconium-water reactionand the containment pressure wouldremain at a low level during the coremelting and pressure vessel meltthroughphases of the accident. When the moltenmass falls to the drywell floor it willcome into contact with water, and thereaction of the molten zirconium withwater can be expected. However, sincethe geometry in this case will not be asfavorable for the reaction as during theinitial core meltdown, the extent of reactionis estimated to be 25 percent ofthe total zirconium in the core. Againbarring a damaging steam explosion,pressure vessel meltthrough will be followedby a rapid pressure increase andthe flow of steam and noncondensablesinto the suppression pool. The rapidpressure increase from 180-210 minutesin Fig. VIII 3-3 reflects the transportof the noncondensables to the suppressionchamber. The slower pressure increaseafter that is due to the continuedevolution of C02 from concretedecomposition; this continues untilcontainment failure.VIII-29

3.2.9 CONTAINMENT ISOLATION FAILURESThe LPCI and CSI pumps take their suctionfrom the suppression pool and haveNPSH requirements of 25 to 28 ft. Theyare not always designed to operate inthe absence of positive pressure in thecontainment. Thus the possibility ofpump cavitation must be considered inthe evaluation of the various accidentsequences, particularly those involvingloss of containment isolation. Theheat-removal capacity of a single RHRloop, the minimum design basis, is lessthan the decay power early in the accident.This imbalance is accommodated byan increase in the suppression poolwater temperature until about 4.6 x 104seconds; after that time the heatremovalcapacity exceeds the decay heatgeneration and the pool temperature decreasesgradually with time. With morethan the minimum design basis LTC availablethere would not be this temporaryimbalance between heat generation andheat removal. If containment integrityis maintained, i.e., leakage from thecontainment does not exceed the designleak rate, pump cavitation will not occurregardless of pool water temperature.In the event of loss of LTC thepool temperature will increase continuouslyand the containment will fail fromoverpressure. Containment failure willthen be followed by pump cavitation.Failure of containment isolation, i.e.,excessive leakage rather than the totalloss of containment integrity, wouldlead to the loss of noncondensables fromthe containment. This would not resultin pump cavitation if the RHR systemwere functioning. With loss of containmentisolation and failure of the RHRsystem, cavitation of the standby corecooling system pumps would be expected.Figures VIII 3-4 and VIII 3-5 illustratecontainment pressure-time histories withloss of RHR and several containment leakrates with the noncondensables in thesuppression chamber and equalized,respectively. It is seen that for leakrates in excess of 100 percent per day,pump cavitation will occur beforecontainment overpressurization. Basedon critical flow of steam with adischarge coefficient of 0.6, a leakrate equal to 100 percent of the freevolume of the containment corresponds toan orifice size of 1 inch; a 1000percent per day leak corresponds to a3.2-in.-diam hole, etc. With the containmentat atmospheric pressure, pumpcavitation would be expected to occurwhen the suppression pool water temperaturereaches 190 F, or about 1.5 x 10seconds into the accident.If the containment fails due to thefailure of vapor suppression but withoutloss of water from the suppression pool,the standby core cooling systems couldstill operate. If loss of vapor suppressionis accompanied by the loss ofresidual heat removal capability, pumpcavitttion would take place at about 2.4x 10 seconds. In this situation, thepool water would not have absorbed theenergy of the flashing primary coolantand thus would require longer to attainthe temperature at which cavitationwould occur.The location as well as the size of containmentisolation failures can have abearing on the overall accident consequences.If the containment isolationfailure takes place in the drywell, fissionproducts released from the core mayleak directly to the secondary containment.In the event of secondary containmentfailure this may result in alarge release of radioactivity to theenvironment, even if gross failure ofthe primary containment is avoided. Forisolation failures in the suppressionchamber gas space, on the other hand,the leakage of fission products will bethrough the pool and then out to thesecondary containment. The water in thepool would be expected to retain most ofthe particulates and a large fraction ofthe iodine. Thus even in the event ofsecondary containment failure therewould be a significant reduction in thefission-product release for these cases.3.2.10 SECONDARY CONTAINMENTThe secondary containment building forBWR's houses reactor auxiliary systems,radwaste equipment, refueling operations,etc. It is a controlled leakagestructure which is normally maintainedat a slightly subatmospheric pressure byventilation fans which exhaust throughthe stack. In the event of an accident,the normal ventilation system is isolatedand the building atmosphere ispassed through the Standby Gas TreatmentSystem (SBGTS) before being releasedthrough the stack. Thus the consequencesof any activity release from theprimary containment could be greatlyreduced by the presence of the secondarycontainment. For the degraded accidentanalyses considered here, however, thepossibility of secondary containmentfailure must also be considered.The secondary containment assumed inthis evaluation has a free volume of 2.53x 106 ft and is designed for an internalpressure of 0.25 psig. The designleak rate is 100 v/o per day at a differentialpressure of 0.25 in. of water.VIII-30

Due to the low pressure capability ofthe secondary containment, any catastrophicfailure of the primary containmentis expected to fail the secondaryalso. Thus steam explosions which failthe drywell would at the same timebreach the secondary containment building;similarly, overpressurization failureof primary containment will fail thesecondary. In addition to the above,loss of primary containment isolationcould result in leakage to the secondarycontainment greater than the capacity ofthe SBGTS and the leakage out.By considering the flow into the secondarycontainment through the containmentisolation leak, flow to the SBGTS andbuilding leakage as a function pressure,it is possible to determine the size ofisolation failure at which the secondarycontainment building pressure will exceed0.25 psig following blowdown intothe drywell. It has been determinedthat loss of. containment isolationequivalent to about 5 to 6-in.-diam holewould lead to the failure of the secondarycontainment in the event of a largeLOCA.The nature and location of secondarycontainment failure can in itself havean influence on the accident consequences.A steam explosion in the reactorvessel which fails the primary containmentwould be expected to breach thesecondary containment above the reactorrefueling floor. Thus any ensuing activityrelease would be directly to theenvironment. An overpressure failure,on the other hand, could take place anywherealong the primary containment envelope.If a failure of the latter typewere to occur in the lower portions ofthe structure, the activity could bereleased into the various portions ofthe secondary containment where considerableholdup and decontamination couldtake place before the release reachesthe environment. There are parts of thesecondary containment structures whosefailure would lead to a direct releaseof activity to the environment. Ingeneral, however, it would be expectedthat the activity released as a consequenceof an overpressure failure of theprimary containment would undergo someholdup and decontamination in passingthrough the secondary containment, eventhough the latter is also failed.3.3 SUBSEQUENCE PROBABILITIESFigure VIII 3-6 illustrates the variouspaths to containment failure that arepostulated following meltdown accidentsin BWR's: reactor vessel steam explosion,containment steam explosion, overpressurization,containment leakage fromthe drywell, containment leakage fromthe wetwell, secondary containment failure,and standby gas treatment systemfailure. It is expected that containmentfloor meltthrough would accompanyeach of the other containment failuremodes. However, since meltthrough isnot the primary mode of containmentfailure in any of the cases and since itwould not add appreciably to the atmosphericfission-product source term, ithas been omitted from the core meltevent tree. The violent modes of containmentfailure, i.e., steam explosionsand overpressurization, are treated asexclusive, e.g., if a steam explosionresults in gross containment failurethere is no need to consider the potentialoverpressurization. Also, violentprimary containment failures as well aslarge containment isolation failureswill be accompanied by failure of thesecondary containment.For a given accident sequence, i.e.,LOCA in combination with various engineeredsafeguards failures, some of thepaths to containment failure may not berelevant or credible. In such casesthese paths are assigned a probabilityof zero for that particular accidentsequence. The method of determining thevarious probabilities of containmentfailure for the accident sequences ofinterest are described below.In Fig. VIII 3-6 the numerically subscriptedfactors, PI, P 2 , etc., representthe probability of failure at eachdecision point in the containment eventtree. They are defined as follows:P - probability of primary containmentfailure due to areactor vessel steam explosionP 2 - probability of primary containmentfailure due to a containmentsteam explosionP 3 - probability of primarytainment failure dueoverpressurecontoP 4 - probability of a containmentleak in the wetwellP 5 - probability that a leak in thedrywell is greater than 2400percent per dayP 6 - probabilitywetwell ispercent perthat a leak in thegreater than 2400dayVIII-31

P 7 - probability of secondary containmentfailureP8 - probability of standby gastreatment system failure.The probabilities of containment leakage,P 4 , P 5 , and P 6 , probability ofsecondary containment failure with primarycontainment intact, P 7 , and independentprobability of SBGTS filterfailure, PS, are not dependent on thecore melting sequence. The numericalvalues of these probabilities have to bedetermined from other studies such asfault tree analyses of the particularsystems. They are included in the presentanalysis since they can affect thecourse and consequences of the overallaccident sequence. The other probabilitiesare dependent on the particularaccident sequence in question, and theirevaluation has been a significant partof the present study.The relative probabilities of each ofthe branches of the BWR containmentevent tree for any given accident sethenquence canbe determined asfollows.P( = P1PB = (1 -P)2This latter set of probabilities takesinto account the exclusive as well asdependent nature of the various failuremodes for each accident sequence.3.3.1 CONTAINMENT FAILURES RESULTINGFROM A STEAM EXPLOSIONAs discussed previously, the only casesin which primary vessel steam explosionsare a concern are those in which meltdowntakes place with water in the primaryvessel and the containment intact.These would include sequences entailingfailure of emergency core cooling function,loss of long-term cooling, andfailure to scram. The probability ofcontainment failure due to a steam explosionis given byP1 = Pfwasxsacf" (VIII 3-1)The terms in this equation have beenpreviously defined. The probabilitythat a significant fraction of the corewould accumulate in the molten state andbe available to interact with the water,Pfw, is believed to be high. The basesfor this conclusion are discussed inAppendix A. The numerical value assignedto this factor is the same asestimated for the PWR cases,Py = (1 - P 1 ) (1 - P2)P3- P= (1- )2 ) (l - P 3 )P 4 P 6= (i1- P(P E - P 2 ) (1 - P3)P4(i - p 6 )P 7E8 = (i - P 1) (I - P2 ) (i - P3 ) P4(l - P 6 ) (1 - P 7 )P 8P = (i - P 1 ) (1 - P2)(I - P3)P4(I - P 6 ) (1 - P 7 ) (I - P 8 )(1 - 1) (I(l - P4 )p 5- P 2 ) (1 - P 3 )- (1 - P 1 ) (1 - P 2 ) (1 - P 3 )(1 - P 4 )(l -P5)P7P = (i - )( - 2) (1 - P 3 ) (l - P 4 )(1 - P 5 ) (1 - P 7 )P 8P = ( - )((1 - P 5 ) (1- P (i - P 3 ) (i - P 4 )- P 7 ) (I - P 8 )Pfw = 10-0.05 (±0.05)The probability that a mass of moltencore material upon contact with waterwill disperse finely enough for a violentinteraction, asxs, is believed tobe depending on the temperature of thewater. If the water is subcooled, theprobability of dispersal is significantlyhigher than if the water is at ornear saturation conditions, as is discussedin Appendix B. For those degradedaccident sequences in which thewater in the reactor vessel can be expectedto be at or near saturation,e.g., loss of LTC, the probability of adispersal event is estimated to be thesame as that for the PWR cases consideredpreviously:sxs = 10-1 (+0.5, -1)In those cases where the water in thebottom of the reactor vessel may besubcooled, e.g., failure of emergencycore cooling function, a higher probabilityof dispersal is utilized,aVIII-32

asxs = 10-0.3(+0.25, -1)Appendix C describes calculations performedto determine the conditions underwhich drywell failure might result froma steam explosion in the pressure vessel.These results indicate that thepotential for containment failure fromsteam explosions in the reactor vesselin BWR's is comparable to that in PWR's.While the PWR containment shell is protectedby a control rod drive shield andby some distance, the internal equipmentin the upper portion of the BWR pressurevessel is effective in absorbing some ofthe energy of the expanding mixture ofhot (molten) core materials and steam.Thus the probability of a reactor vesselsteam explosion failing containment isestimated on the basis of these calculationsto beacf = 10-1 (+0.5, -1)Combining the individual factors assuminga skewed log-normal distribution,the probability of a reactor vesselsteam explosion leading to containmentfailure with saturated water isP 1 = 10-2 (+1, -2)and with subcooled water itP1 = 10-1.35 (+0.8, -2)Steam explosions are also possible uponpressure vessel meltthrough as the massof molten materials falls into the wateron the drywell floor. As is the casewith reactor vessel steam explosions, asignificant quantity of molten materialmust interact coherently with water toproduce a damaging explosion. Hereagain the probability that a large quantityof molten material would be availableto interact with the water, Pcfw,is believed to be large. Accordingly,Pcfw = 10-0.05 (±0.05).The probability that the molten materialis finely dispersed upon interactionwith the water is taken to be identicalto that for reactor vessel steam explosions,being 10-1(+0.5, -1) and10-3 l25,-l) for saturated and subcooledwater, respectively.isThe containment steam explosion wouldtaken place within the 20-ft-ID reactorvessel support cylinder which is steellined,reinforced concrete approximately3 ft thick. The explosive vaporizationof a significant fraction of the approximately2 ft of water at the bottom ofthis cylinder will produce pressureswell above the strength of the cylinder,the latter being a few hundred psi.Although there is an opening in the sideof this cylinder for an access door andthe space between the reactor vessel andthe sacrificial shield provides partialopening at the top, the time scale of asteam explosion would be too short topermit significant pressure reliefthrough these vent areas. The rise timeof the steam explosion would be on theorder of 30 msec, much larger than thesonic transit time through the concretewall which is calculated to be about 0.3msec. Thus spalling of the concretewould not be expected, and failure ofthe support cylinder will take place byan essentially static overpressurization.The steam will then be released to thedrywell. If the pressure rise withinthe drywell is sufficiently rapid, thedrywell-to-wetwell vents may not be ableto clear in time to avert overpressurizationof the drywell. The time requiredto clear the vents for a pressurerise of the type considered here hasbeen calculated to be 0.1 sec; an additional0.1 sec would be required to flow10 percent of the steam into the suppressionpool. Thus, if a major fractionof the heat transfer between themolten core and water takes place in 0.2sec or less, venting of the steam to thesuppression pool will not be possible.The effective particle diameter associatedwith this heat transfer time isabout 600 microns, based on the analysesin Appendix C. This particle size iswell within the range of sizes observedin steam explosions.If the water within the reactor vesselsupport structure were vaporized fasterthan it could be transported to thesuppression pool, the incremental increasein the static pressure in thedrywell would be 112 psi. The resultingpotential for drywell failure from overpressurewould depend upon the preexistingpressure in the drywell at the timeof the event. Thus the potential forfailure due to a containment steam explosionwill have to be considered separatelyfor each accident sequence inquestion. If drywell failure does notoccur, the steam will subsequently flowinto the suppression pool and condense.VIII-33

In addition to the quasiequilibriumpressure of steam, missile generationcould be a potential mechanism of failurein the event of a containment steamexplosion. As discussed above, the risetime of steam explosions is not expectedto be rapid enough to produce spallingon the outside of the reinforced concretecylinder. The overpressure failuremay produce concrete fragments, butthese would not be expected to have ahigh velocity. After pressure relief,the reinforcing steel together withother supporting structures would stillbe expected to provide adequate supportto the shell of the pressure vessel.The sheet metal door could be a potentialmissile, but one that is unlikelyto penetrate the drywell. Further,after a 2-in. air gap, the drywell vesselis backed up by concrete; thus thesteel shell could deform locally andtransfer the missile loading to theconcrete. Thus it is concluded thatfailure due to missiles in this case hasa significantly lower probability thanfailure due to overpressure.3.3.2 CONTAINMENT OVERPRESSURIZATIONFor meltdown accidents resulting fromloss of LTC or failure of vapor suppression,containment failure due to overpressurizationprecedes core meltdownand is accompanied by failure of thesecondary containment. Thus, alternatepaths to containment failure need not beconsidered for these cases. In casesinvolving failure of emergency corecooling operation or function, if thecontainment is still intact after initiationof meltthrough of the drywellfloor, failure due to overpressurizationis considered a certainty. As notedpreviously, the C02 generated by thedecomposition of concrete together withthe other noncondensables will raise thecontainment pressure above its failurelevel before the melt reaches the drywellvessel.Small containment isolation failures,i.e., equivalent to l-in.-diam holes orless, will not prelude containment failureby overpressurization.3.3.3 SECONDARY CONTAINMENT FAILUREAs noted previously, violent failures ofthe primary containment would be expectedto result in failure of the secondarycontainment. If large containment isolationfailures existed during the blowdown,secondary containment would beexpected to fail. Small containmentisolation failures would not result insecondary containment failure directly,although the latter would occur when theprimary containment fails due to overpressureor other mechanisms. Thuseffectively the only degraded accidentsequences which benefit from the secondarycontainment are those in which intermediatecontainment isolation failuresare present. These failures arelarge enough to prevent overpressurizationof the containment but not so largeas to overpressurize the secondary containmentbuilding.3.3.4 UNFILTERED ELEVATED RELEASEThe SBGTS, consisting of demisters,high-efficiency particulate filters, andactivated-charcoal filters, is designedto clean the secondary containment atmosphereand exhaust through the stack.Thus the consequences of any radioactivityrelease from the primary containmentcould be greatly reduced by theSBGTS operation. The degraded accidentanalyses have also considered the possibilityof failure of the SBGTS. Thelatter can occur in several ways:a. Failure of the secondary containment,discussed aboveb. Failure of the components of thetreatment train, e.g., leakagethrough or around the filters,improper installation. (Thisprobability requires separatestudies on the reliability of gastreatment systems.)c. Failure of some or all of the componentsof the train due to overloadingwhen subjected to off-designconditions, as may be the case inmany of the accidents consideredhere. (This possibility is beingevaluated in the Fission ProductSource Term Task.)The SBGTS can mitigate the consequencesof radioactivity release even with someof its components failed by leading toan elevated rather than a ground levelrelease.3.3.5 HYDROGEN COMBUSTIONDue to inerting of the containment atmosphere,the potential for hydrogencombustion in BWR's is small and itsinfluence on the consequences of meltdownaccidents would not be very significant.Accordingly this possibilityhas not been included in the containmentfailure evaluation.aVIII-34

-- A i -, 0 O 0.-- 40 C'2160140120.Loa806040200N- All noncondensables in suppression chamberAoncondensables equalized18o 14o iso 220 260 30 34o 0 38oTemeerature, FFIGURE VIII 3-1 PWR Containment Pressure asa Function of Compositionand Temperature200180160140pressureK...............120Cd100,Cn802003010 3.Time, minFIGURE VIII 3-2 BWR Containment Pressure Response withAssumed Failure of Emergency Core CoolingSystem Function

200 ý180e Fresst±re160140.............. ....... -120-LO100U)Q)80'60.40.20n0 50 100 150 200 250 360 350 4d0Time, min450 50O 5ý0 660FIGURE VIII 3-3 BWR Containment Pressure Response withAssumed Failure of Emergency Core CoolingSystem Operation100.8016 T ,Accident Time, hrFIGURE VIII 3-4 BWR Containment Pressure Response as a Functionof Leak Rate with Assumed Loss of Long-TermCooling and All the Noncondensables in theSuppression Chamber

180140 Conti........... Pump12010080604020....................0 .0 5FIGURE VIII 3-5Accident Time, hrBWR Containment Pressure as a Function of LeakRate with Assumed Loss of Long-Term Coolingand Noncondensables Equalized Between theDrywell and Suppression ChamberCore MeltDW vs WWVSE CSE O Lek LCL SCF SGTSa 8 y 6/c 4 0Sequence6Y8P 1FIGURE VIII 3-6 BWR Containment Event TreeFIG. VIII 3-1 - FIG. VIII 3-6VIII-35/36

APPENDIX ATHERMAL ANALYSISVIII-37

Appendix ATable of ContentsSectionPage No.Al. CALCULATIONS ..................................................... VIII-41Al.I Core Heatup Calculations ...................................A1.1.1 Computer Program BOIL ..............................A1.1.2 BOIL Calculations ..................................VIII-41VIII-41VIII-41A1.1.2.1 Heat Transfer ............................ VIII-41A1.1.2.2 Water Boiloff ............................ VIII-46A1.1.2.3 Core Meltdown ............................ VIII-49A1.2 Results of BOIL Calculations ............................... VIII-51A1.2.1 PWR Results ........................................ VIII-51A1.2.2 BWR Results ........................................ VIII-53A1.2.3 Bottom-Flooding .................................... VIII-54A2. CONTAINMENT HEAT TRANSFER ........................................ VIII-54A2.1 PWR Analyses ............................................... VIII-54A2.1.1 Energy Sources ..................................... VIII-54A2.1.2 Containment Heat Sinks ............................. VIII-54A2.1.3 Containment Atmosphere ............................. VIII-55A2.1.4 Containment Pressure-Time Evaluation .................. VIII-55A2.1.5 Containment Leakage ................................ VIII-55A2.2 BWR Analyses ............................................... VIII-56A2.2.1 Energy Sources ..................................... VIII-56A2.2.2 Containment Heat Sinks ............................. VIII-56A2.2.3 Containment Response ............................... VIII-57A3. PRESSURE VESSEL MELTTHROUGH ...................................... VIII-57A3.1 PWR Analyses ............................................... VIII-57A3.1.1 Realistic Failure Time ............................. VIII-58A3.1.2 Maximum Failure Time ............................... VIII-59A3.1.3 Minimum Failure Time ............................... VIII-59A3.2 BWR Analyses ............................................... VIII-59A4. CONTAINMENT VESSEL MELTTHROUGH ................................... VIII-59A4.1 PWR Analyses ............................................... VIII-59A4.2 BWR Analyses ............................................... VIII-61REFERENCES ..............................................................VIII-63VIII-39

List of FiguresFigureVIII A-1VIII A-2VIII A-3Characteristics Curves For Release of Volatile FissionProducts From Overheated Fuel ...................................Mixture Void Fraction and the Locations of the MixtureLevel and Molten Fuel Regions in a PWR for MeltdownModels A and B as a Function of Time for Boiloff Startingat 1.0 Hr .......................................................Steam Generation Rate During a Boiloff Accident Startingat 1.0 Hr in a PWR as a Function of Time for MeltdownModels A and B ..................................................Page No.VIII-55/56VIII-55/56VIII-55/56VIII A-4 Fuel Rod Temperatures in the Central (Maximum Power) CoreZone of a PWR as a Function of Time for Boiloff Startingat 1.0 Hr for Meltdown Model A ..................................VIII-55/56VIII A-5VIII A-6VIII A-7VIII A-8Effect of Meltdown Model on the Percent of a PWR CoreMelted as Function of Time for Boiloff Starting at1.0 Hr ..........................................................Effect of Meltdown Model on the Percent of the CladdingReacted in a PWR Core as a Function of Time for a BoiloffAccident Starting at 1.0 Hr .....................................Effect of the Time When a Boiloff Accident Starts (DecayTime) on the Core Meltdown Sequences in a PWR for MeltdownModel A ....................................................Percent of a BWR Core Melted as a Function of Time for(1) a Dry Core Heatup Starting with End-of-Blowdown Conditionsat 25 Sec and (2) a Boiloff Accident Starting at100 Hours with the Core Flooded to the 8 Foot Level .............VIII-55/56VIII-55/56VIII-57/58VIII-57/58VIII-40

Appendix AThermal AnalysesAl. CALCULATIONSThe primary purpose of the calculationsdescribed in this appendix is to provide,as quantitatively as possible, adescription of the physical conditionsexisting during the course of the accidentand the time of occurrence of steamexplosions, hydrogen explosions, containmentmeltthrough, the size of thefission-product source, and the size ofthe heat load on the containment. Thatis, the calculations in this appendixdefine the time spectrum during whichconditions exist which may eventuallylead to breach of the containment. Forexample, the core heatup calculationsdefine the amounts of molten UO? andwater available for a steam explosion asa function of time. These calculationsalso predict the quantity of Zircaloycladding that has reacted with steam toproduce hydrogen, which may possiblyexplode or burn to breach the containment.AI.l CORE HEATUP CALCULATIONSThis section describes core heatup calculationsperformed for several degradedcore cooling situations. The calculationswere performed for two specificreactors, representative of a large PWRand BWR, respectively. The models inthe computer program and many of theresults have general applicability tothe accidents studied, although some ofthe results apply only to the specificdesigns considered. The computer programBOIL written for this evaluationis described and the results of the calculationsare presented. Some of thecalculated results appear elsewherethroughout the report, and are notalways repeated or given further documentationin this appendix.The situation of principal concern inthis discussion is the "boiloff" accident.In the "boiloff" accident, as theterm is used here, circulation of waterthrough the core is assumed to stop.Because the core continues to generatedecay heat and no further water is beingadded, the water in the core begins toboil off, and the water level decreases.The uncovered portions of the core heatup and eventually melt. The computerprogram BOIL was written to describethis sequence. The program may also beused in some cases to calculate: (1)core heatup for (very) low bottomfloodingrates which lead to core meltingand (2) dry heatup of the core inthe complete absence of water.In the work that follows, the computerprogram BOIL is discussed, and theresults of calculations with BOIL arepresented.A1.1.1COMPUTER PROGRAM BOILThe computer program BOIL was written tocalculate core heatup in an accidentwhere the fission-product-decay heatboils the water out of the pressure vesseland uncovers the core. The approachused is to divide the core into smallvolumes or nodes and (1) calculate theheat produced in each node and performheat balances between the fuel and coolantnodes, (2) calculate the water-steammixture level in the core and the steamboiloffrate, and (3) perform a meltdowncalculation when the temperature ofa node exceeds the melting point of U0 2 .The models in each of the three parts ofthe program are discussed below.Al.l.2 BOIL CALCULATIONSA1.1.2.1Heat Transfer.This section of BOIL (1) calculates thefission-product-decay heat as a functionof time, (2) accounts for the reductionin the heat source due to fissionproductvolatilization, (3) calculatesthe heat produced by the zirconium-steamreaction in the cladding, (4) calculatesconvection heat transfer between thefuel rods and the steam or water coolant,(5) calculates radiation heattransferlosses from the ends of thecore, and (6) performs heat balances onthe fuel and coolant.A1.1.2.1.1 Core Nodalization. In theBOIL calculations, the core is dividedinto a maximum of 10 radial zones composedof fuel rods and their associatedflow channels. The radial zones aresectioned into a maximum of 50 axialnodes. Most of the calculations havebeen performed for 5 radial and 24 axialnodes. The sizes of the radial zonesare arbitrary, and are normally chosenin a manner which conveniently describesthe radial power distribution of thecore. In the heat-transfer calculationsVIII-41

the only coupling between radial zonesbefore core melting occurs is by a uniformsteam flow rate (lb/hr ft2) at theinlet to the steam-covered region. Atthe channel exit, the steam concentrationdepends on the amount of metalwaterreaction in that particular channel.There is normally no coupling ofradial zones by conduction, convection,or radiation heat transfer. After coremelting occurs, radial zones may (programoption) be coupled by steam crossflow and mixing within the pool of moltenfuel in the core. The axial fuelrodnodes within a given radial zone arenormally coupled in the heat-transfercalculations only through their connectionto a common steam channel. Duringcore meltdown, the axial nodes are assumedto be coupled within the moltenregion by convective mixing. The axialnodes are not coupled by conduction heattransfer. The axial nodes in the flowchannel are connected by the axial flowof steam.A1.1.2.1.2 Decay Heat. The fissionproduct decay heat is calculated fromthe Proposed ANS Standard for infiniteirradiation time:P/Po~ 0.076t0181P/P= 0.0766 t , 10 < t < 150 sec,P/Po =0.130 t-0.283(VIII A-la)150 < t < 4 x 106 sec, (VIII A-lb)where P/P is the fraction of operatingpower and t is the decay time in seconds(Ref. 1). For most of the time periodof interest, the Proposed ANS Standardis estimated to give decay heats between10 percent too low and 20 percent toohigh. The decay heat of a given corenode as a function of time is obtainedusing Equation (VII A-l) with axial andradial power-peaking factors applied tothe average core power density.A1.1.2.1.3 Fission-Product Release. Forfuel-rod temperatures below 1500 F, theheat-transfer calculations assume thatno fission products are lost. For rodtemperatures between 1500 F and 2000 F,the fission-product loss is primarilycomposed of a puff release (RP), whichdepends only on the operating power historyof the fuel node. Above 2000 F,the total release depends on both therod power and the current temperature ofthe fuel node. The fission-productreleasefraction programmed in BOIL fortemperatures between 1500 F and 2000 FisRP = 0.13(F - 1.35), (VIII A-2)where F is the local power-peaking factor.For local peaking factors lessthan 1.35, RP is taken to be zero. Fortemperatures above 2000 F the releasefraction isRT = 0.471 (1 - F/3.65) (TR/3000 - 0.67),(VIII A-3)where TR is the temperature (F) of thefuel-rod node. In the BOIL calculations,the fission-product-decay heat ofa fuel node is reduced by the fractionRP or RT to account for loss of fissionproducts. The maximum fractionalreduction permitted in the program is0.30. It is assumed the fission productsare lost from the core and are notredeposited elsewhere in the core.The above expressions for fission productrelease were derived from thecurves shown in Fig. VIII A-1. Thesecurves were constructed on the basis ofcalculations of fission-gas release presentedin BMI-1885 (Ref. 2). The calculationswere performed for the case of aloss-of-coolant accident with no emergencycore cooling. The preaccidentdiffusion release of fission gas to thefuel rod gas gap was found to vary withthe power peaking factor as shown by theinitial (1500 F) values of the curves inFig. VIII A-1. The preaccident gap contentswere assumed to be available forrelease immediately upon rod rupture.The rod rupture temperature was taken as1500 F, though it recognized that fuelrod failures could occur over a range oftemperatures; the release values wouldnot be sensitive to this parameter atthe high core heatup rates associatedwith meltdown accidents. The releasecurves at temperatures above 1500 Fgiven in Fig. VIII A-1 were developedfrom the postaccident fission gas releasevalues in Table 1 of BMI-1885 andthe core heatup rate data from Fig. 1 ofBMI-1885. To accomplish this, an approximaterelationship between fissiongas release and fuel temperature wasconstructed. This approximate relationshipis given by the bottom curve inFig. VIII A-l, and it represents thethermally activated release of fissiongas remaining in the U02 itself. Theappropriate fractional releases from theU0 2 are added to the preaccident gap releasesto generate the total releaseVIII-42

curves for the various powerfactors shown in Fig. VIII A-i.curves were extrapolated to 100release at a temperature of 5000peakingAll thepercentF.The release versus temperature curves,while specifically derived using noblegas diffusion coefficient values, wereassumed to also apply to the other highlyvolatile fission products, namelyiodine, cesium, and tellurium. Whilethere is some experimental data thatgenerally supports this assumption, itis clearly an approximation. The fourgroups of fission products classifiedhere as volatile contribute about 30percent of the total fission productdecay energy at shutdown times of aboutan hour.A1.1.2.1.4 Cladding - Water Reaction.The zirconium-steam reaction rate in theZircaloy cladding is calculated from theminimum predicted by (1) Baker's ratelaw or (2) a gas phase diffusion model(Ref. 3). The model is the same as thatused in NURLOC (Ref. 3, Eqs. 75-82), andthe equations are not repeated here.The reaction is stopped at a node if allof the steam in a flow channel has beenconsumed or if all of the cladding hasbeen consumed. The cladding-steam reactionis also stopped after a fuel-rodnode is completely melted, on the assumptionthat steam will not penetrate amolten node. When core melting occurs,the user of the program may select variousoptions which scope the effects ofchannel blockage on the metal-water reactionin the unmelted core zones. Theprogram user may (1) let the steam flowcontinue through a channel as if noblockage occurs, (2) assume the channelplugs and stop the metal-water reactionin a given channel above the lowestmelted node, (3) let the steam flow bypassmelted (plugged) channels and,therefore,. increase the steam flow ratein unplugged channels. Options (2) and(3) can significantly reduce the fractionof the cladding in the core whichreacts with steam. In general, it isbelieved that option (1) overestimatesthe amount of cladding reacted, particularlyin BWR's with shrouded fuel ele-~ments. Options (2) and (3) probably underestimatethe metal-water reaction inboth open-lattice and shrouded elements.A1.1.2.1.5 Thermal Radiation. Radiationheat transfer from the uncoveredpart of the core to (1) the internalstructure above the core and (2) thewater inside or below the core is evaluatedin the BOIL calculations. The heattransferred by radiation from a nodewith emitting area A R is0 rad 0.173 F A R [(T R/100) 4 - (T 0/100) 41whereQrad = radiationBtu/hrheatA = radiating area, ft 2(VIII A-4)transfer,T R =temperature of radiatingnode, RTo= temperature ofnode, RreceiverF =radiation interchange factor.The interchange factor is calculatedassuming the radiating and receivingsurfaces are parallel planes;I thus,whereF =L_( C R 0 - 1)(VIII A-5)E R = emissivity of radiating nodeE0= emissivity of receiver node.The area, AR, is computed in a planeperpendicular to the axis of the core.The total heat radiated is calculated bysumming the Qrad over the total crosssection of the core. For radiation heattransfer from the core to the water, theTR are evaluated in the plane of thefirst uncovered node above the water.For heat transfer to the internal structureabove the core, the TR are evaluatedfor nodes in a plane at the top ofthe core. It is assumed in the calculationsthat the maximum amount of heatthat can be radiated from a (solid orunmelted) node is the decay heat of thenode. This restriction is placed on theradiation heat transfer because themechanisms for transfer of heat (by conductionand radiation) in the axial directionare normally limited. Since thecore was divided into 24 axial nodes,the maximum amount of heat lost from the(unmelted) core by radiation is 2/24 ofthe decay heat. When the top of thecore is melted, the radiation heat lossfrom the top of the core i~s limited (inmeltdown Model B) to the total decayheat in the molten pool. The mechanismfor moving this heat to the surface ofthe molten pool is internal natural convection.VIII-43

The heat fluxes calculated from Equation(VIII A-4) are used to calculate thetemperature of the structure above thecore. In these calculations, the structureis divided into pieces of knownmass and heat capacity. If one of thepieces melts, it is assumed to fall awayand expose another piece to the radiationheat flux.The radiation heat flux to the watercontributes to the boiloff of the waterand is included in the water massbalance equations. The boiloff ratefrom radiation heat transfer isgenerally small in comparison with theother mechanisms, unless the core isnearly uncovered.A1.1.2.1.6 Convection Heat Transfer.The core is divided into two convectionheat-transfer regions, a steam-coveredregion and a region covered with awater-steam mixture. In the regionabove the mixture, the rod-steam heattransfercoefficient, h, is calculatedfrom a simplified Dittus-Boelter correlation(Ref. 4),where08 0.2h = 0.0144 C G /D (VIII A-6)Cp= specific heat of steam, Btu/lbFGsteam flow rate, lb/hr ftD = equivalent diameter of channel,ft.Steam properties are used in the calculations,and the properties are notchanged to account for changes in thesteam-hydrogen mixture ratio due tometal-water reaction. For the steamflow rates obtained in the boiloff calculations,the convection heat loss froma fuel-rod node is generally a smallfraction of the decay heat. That is, theheatup rates are nearly adiabatic attemperatures below 2200 F where thecladding-steam reaction is small, andthe gas temperatures closely follow therod temperatures. In view of its smalleffect, a more detailed treatment ofconvective heat transfer to steam and/orhydrogen does not appear warranted.In the water- or mixture-covered regionsof the core, the rod temperatures areassumed to be in a steady-state equilibriumwith boiling water. The temperatureof a fuel-rod node in the mixturecoveredregion is2whereTR = Tw + Q D,R/AhB,(VIII A-7)TW = saturated water temperature= 300 FQD,R = decay heat of node, Btu/hrA = node heat-transfer area,ft2hB = boiling heat-transfer coefficient- 350 Btu/hr ft2 F.According to Equation (VIII A-7), all ofthe decay heat produced in a watercoveredfuel-rod node goes into boilingwater. The fuel-rod temperatures givenby Equation (VIII A-7) are generallyless than 10 - 20 F greater than thewater temperature. Thus a more sophisticatedtreatment, taking into accounttime-dependent changes in saturationtemperature, heat transfer coefficient,etc., is not necessary. At the end of atime-step in the BOIL calculations it ispossible that a fuel-node temperature inthe mixture-covered region may be greaterthan that given by Equation (VIIIA-7). This occurs in two situations:(1) when the mixture level increasesduring a time-step due to bottom floodingand covers a fuel node which waspreviously dry, and (2) when molten fueldrops into the mixture-covered region.When this occurs, the fuel node is assumedto be quenched (in one time-step)to the node temperature given by Equation(VIII A-7). The assumption ofquenching in one time-step is justifiedby transient-conduction analysis. Thesecalculations indicate that for a (typical)1-minute time step, a fuel rod willlose about 90 percent of its storedheat.A1.1.2.1.7 Node Heat Balances. Usingthe nodalization treatment, heatsources, and heat-loss mechanisms describedabove, BOIL performs heat balancesand calculates the fuel-rod andflow-channel node temperatures. In theBOIL nodalization treatment, a fuel rodhas only one radial node. Thus, theconduction heat-transfer problem is notsolved, and the heat balances involveonly changes in stored heat and the heatsources and losses. The use of only oneradial node for a rod is a satisfactorytreatment for the low, nearly adiabaticheatup rates (few hundred F/minute) encounteredin the boiloff situations; atdecay power levels the temperature dif-VIII-44

ference between the fuel center andsurface would be only the order of 100F.In the mixture-covered region, the fuelto-waterheat balance isR --+ Fmelt XPVR = QD,R +MW+ Qmelt - Qrad - hA (TR - TG),(VIII A-10)QD,R = hBA (TR - TW) = mRhfg"wherehfg(VIII A-8)boiloff rate per node, lb/minheat of vaporization, Btu/lb.and the other notation is the same as inEquation (VIII A-7). Note that the rodtemperature TR changes very slowly (withthe decay heat) in the water-covered region.The total contribution of the decayheat to the boiloff rate is obtainedby summing the ffIR for each mixturecoveredfuel-rod node. If at the end ofa time step the temperature of a fuelrodnode in the mixture-covered regionis greater than that obtained from Equation(VIII A-7) (because the node wasnot water-covered during the previoustime step), the node is quenched accordingto the relation,whereQMWQmelt= heat from metal-waterreaction, Btu/hr= heat added to node fromslumping during meltdown,Btu/hrradiation heat-transferlosses, Btu/hiFmelt = fraction of node melted= heat of fusion of node,Btu/lbh = Equation (VIII A-6)TG = gas or steam node temperature,F.The increase in the steam temperaturealong the length of the channel is givenbyQmelt = PCVR (TR - TR) R,QA t hfg,C ýT G = h p (T Tp -TT R - G(VIII A-11)where(VIII A-9)p = density of fuel rod, lb/ftC = heat capacity of rod, Btu/lb FVR = volume or node, ftAt = time step in BOIL calculation,hr= boiloff rate from quenching,QR,Qlb/hrTRIT R= Equation (VIII A-7)= node temperaturequenching, F.33beforeFor a fuel-rod node in the steam-coveredregion, the node temperature increasesbecause the steam flow is insufficientto carry off the decay heat. The generalizedfuel-rod-node heat balance iswhere= total steam flow rate, lb/hrZ = distance measured up the channel,ftp= fuel rod circumference, ftand the other notation was given previously.In Equation (VIII A-10), Qmeltis calculated assuming the molten poolmixes and has a uniform temperature.During node melting, the node temperatureis constant. (That is, @TR/3t = 0when 0 < Fmelt < 1). Qrad is zero exceptfor the node at the top of the coreand immediately above the mixture level.QMW is always zero when Fmq1tt= 1, andmay also be zero (optional) i the nodeis above a molten region.Equations (VIII A-10) and (VIII A-11)are programmed in finite-difference formfor solution in BOIL. The finitedifferenceforms of the equations areVIII-45

T R (t + At) = T R(t) + qAt/pCV R and the flow out the break in V 2 is(VIII A-10a)TG(Z + Az) = TR(t + At) -(TR(t + At)1fl -n v2ddt 2A heat balance on volume V 2gives(VIII A-13)- TG(z))e (\Cp ),(VIII A-lla)dT 2P 2 v 2 cp i 1CpT 1 - 2-hpvApV(T 2 - TpV),where q is the sum of all the heatlosses and gains in Equation (VIIIA-10). In most of the calculations, atime-step size of At = 1.0 minute and anode length of Az = 6.0 inches was used.The calculation is numerically stablefor all mesh sizes; however, accuracyappears to suffer if the temperaturechange is much greater than a few hundreddegrees (F) per time step. Experienceindicates that the primary effectof too large a time step is to overestimatethe metal-water reaction rate and,consequently, the core heatup rate.AI.1.2.1.8 Dry Heatup. BOIL permitsany initial water level in the pressurevessel, including zero. Consequently,BOIL may also be used to calculate dryheatup of the core in the absence of anywater.An item of interest to the determinationof fission-product release during dryheatup is the efflux of steam from thepressure vessel due to simple gas expansion.since there is no steam boiloffto carry out fission products, gas expansionprovides the principal drivingforce for fission-product transport.BOIL includes a simple gas-expansionmodel for calculating efflux during dryheatup. The model assumes the core isin volume Vl, and the remaining part ofthe pressure vessel is represented byvolume V2. The gas in volume V 1 followsthe core temperature, expands, and flowsinto volume V2. Volume V 2 contains aleak which vents to the containment.Heat transfer is permitted in volume V 2 .The two volumes are assumed to be atapproximately the same pressure withnegligible flow resistance through theleak paths. The steam flow from V 1 intoV 2 isdo -V 1 d-t- 1(VIII A-12)wherep = steam density, lb/ft 3T = gas temperature, F(VIII A-14)hPV = heat-transfer coefficientfrom V 2 tQ structures in V 2 ,Btu/hr ft2 FAPVTPV= heat-transfer area2structures in V2, ft= temperature of APV, F.These equations, along with the perfectgaslaw (Po = constant = pRT), are programmedin BOIL to obtain the steam flowM2) out the break in V2. The parametersApV and TPV are input constants,and hpV = 0.2 AT 1 73 is used to estimatethe convection heat-transfer losses(Ref. 3) from V 2 .A1.1.2.2Water Boiloff.In this section of BOIL, the boiloff ofthe water in the pressure vessel is calculated.The components of the boiloffcalculation discussed below include calculationsof (1) the heat input to themixture region, (2) the steam-generationrate, (3) the water mass, (4) the mixturelevel, and (5) limitations on theuse of the model to calculate bottomflooding.Al.l..2.2.1 Heat Input to Water. Theheat input to the water includes (1) allof the decay heat in the fuel-rod nodescovered by the swollen mixture of waterand steam, (2) radiation heat transferfrom the first steam-covered node immediatelyabove the mixture level, and(3) the stored heat from the quenchingof fuel nodes which were dry during theprevious time step. These components ofofVIII-46

the heat input to the water are given inEquations (VIII A-8), (VIII A-5), and(VIII A-9). When molten fuel drops intothe water, the decay heat and the storedheat are included in the heat input tothe water, but metal-water reaction isneglected. It is also implicitly assumedthat steam explosions do notoccur, and the only effect of moltenfuel is to boil water.A1.1.2.2.2 Steam Flow Rate. It isassumed that the water in the pressurevessel is at the saturation temperature.The steam generation rate isýn= Q total /h fg,(VIII A-15)sidered. However, level swell also increasesthe steam-generation rate anddecreases the water mass more quickly.The net effect of level swell is toretard or delay core heatup by severalminutes. A simple level-swell modelwhich assumes a linear variation of thevoid fraction with height inside thecore is incorporated in BOIL. For amixture level, Y, the water mass isM = (1 -0[1 O(4)].= AtotYPL (VIII A-17)where Qtotal is the sum of all the heatinputs to the water as discussed above.All of this steam is assumed to flowinto the steam-covered part of the core;that is, bypass flow is neglected. Whencore melting occurs, steam flow may(optional) be diverted around channelsassumed to be plugged.Al.1.2.2.3 Water Mass. The water massis calculated by adding the changes inthe water mass to the value in theprevious time step. The mass balance iswhereM(t + At) = M(t) - &At + PLACVF t,M = water mass, lb(VIII A-16)PL = density of saturated water,lb/ftAC = flow area of core, ftVF = "bottom flooding" rate, ft/hr= Equation (VIII A-15).A1.1.2.2.4 Mixture Level. The water inthe pressure vessel is actually a mixtureof water and steam. The volume ofthe mixture is larger than the equivalentmass of water; and, likewise, theswollen mixture level is greater thanthat of water. It is assumed in theBOIL calculations that core nodes coveredby the mixture are cooled. Consequently,a given mass of water will coolmore of the core if level swell is con-2whereAtot = total cross-sectional areain pressure vessel occupied2by mixture, ftY = mixture level above bottomof core, ft(X = void fraction at the top ofthe mixture, and= OTZ/Y and is usedevaluate the integral.The void fraction at the top of themixture is related to the boiloff rateaccording to the relationwhereQDK = r hfg = PsUTaThtothfg,to(VIII A-18)QDK = total decay heat in themixture-covered fuel nodes,Btu/hrPs density of saturated steam,lb/ftUT steam separation velocity atthe top of the mixture,ft/hr.The possible contributions of heatinputs from radiation heat transfer andthe quenching of molten fuel are neglectedin the evaluation of the mixturevoid fraction. In the level-swell calculations,Equation (VIII A-18) is firstused to calculate aT, and the Equation(VIII A-17) is used to calculate a newvalue of the mixture level Y. The voidfraction in the water above and belowVIII-47

the core is assumed to be zero. If Y isgreater than the core height, the totalmixture height isY = H o(M/PLAtotHo + aH2),(VIII A-19)where H0 = active fuel length, ft. If Yis at a level under the core, aT = 0 inEquation (VIII A-19); and negative valuesof Y are calculated in BOIL. Thelevel-swell model in BOIL is similar inconcept to those used in RELAP (Ref. 5)and by General Electric (Ref. 6). However,RELAP assumes a linear variationof the mixture density, G. E. assumes alinear variation of the mixture quality,and BOIL assumes a linear variation ofthe void-fraction. RELAP and BOIL requirethe program user to specify (inputconstant) values of the steam separationvelocity UT. G. E. calculates values ofUT based on the data of Wilson (Ref. 7).Wilson's experiments did not mock up theaxial variation of the steam void fraction.G. E. assumes that Wilson's constantor uniform void fraction is equivalentto the average void fraction inthe volume. RELAP required use of aconstant separation velocity of UT - 3 0ft/sec to reproduce experimental blowdownin a one-volume vessel (Ref. 4).G. E. obtained experimental confirmationof its level-swell model in a 12-in.-diameter x 14-ft-high pressure vesselusing (Wilson's) values of UT between1.0 and about 6 ft/sec. (G. E. assumesin the calculations the minimum value ofUT is 1.0 ft/sec.) Wilson's correlationat a saturation temperature of 300 Fgives (assuming aave = 1/2 aT)U = 34T ý 60244(a T)1.283 ft/sec,(VIII A-20)where D is the hydraulic diameter of theregion in inches (Ref. 6). It is notclear, however, how D should be interpretedfor a core composed of tube bundles.Wallis (Ref. 8, p. 287) indicatesthat D should be the bundle-housing diameterrather than the flow-channelequivalent diameter. In a typical BWRcore the maximum value of UT from Equation(VIII A-20) is 3.47 ft/sec if D isthe equivalent diameter, and the maximumUT is about 5.9 ft/sec if D is the housingdiameter. In the BOIL calculationsa constant value of UT = 4.5 ft/sec wasused. Use of this separation velocityin Equation (VIII A-18) results in amaximum value of the void fraction (allcore nodes covered) of aT = 0.75 for adecay time of 1.0 hr in a typical core.For decay time less than about 25 min,Equation (VIII A-17) predicts maximumvalues of cT greater than 1.0; and BOILreduces a T to 1.0 in that case.The BOIL calculations do not include an"entrainment" model. "Entrainment" iscaused by steam being produced so rapidlyin a channel that it "entrains" orsuspends water droplets due to dragforces, and carries the water dropletsup through the channel. This phenomenonis observed, for example; in the FLECHTbottom-flooding experiments. It is notbelieved, however, that the entrainmentphenomenon occurs in the boiloff accidentstudied here. Entrainment isbelieved to require (1) large forcedcoolant-injection rates and (2) initiallyhot (%1600 F) rods. Neither factoris present in the boiloff calculations.A1.1.2.2.5 Bottom Flooding. BOIL maybe used, in some limited cases, for thestudy of core heatup with degraded "bottomflooding". It is assumed in theBOIL calculations that coolant is addedat a very low rate to the water inventoryat the bottom of the pressure vessel.BOIL calculates the response of the coreto this bottom-flooding rate using exactlythe same model assumptions (withregard to level swell, fuel-node quenching,steam generation, etc.) as are usedin the other boiloff calculations. Thelimitations on the use of BOIL to calculatebottom flooding include (1) saturatedwater is injected, (2) the water isadded to the inventory at the bottom ofthe pressure vessel, (3) the floodingrate should not result in covering morethan about one rod node per minute. Thelatter restriction is primarily a consequenceof the assumption in BOIL that apreviously dry fuel-rod node can bequenched in one 2 time step. Since thetime constant (pcr /k) of a fuel rod isapproximately 1 minute, the quenchingassumption is questionable if more thanone previously dry node per minute iscovered. For the nodes considered inthe BOIL calculations, a practical upperlimit on the bottom-flooding rate isabout 0.2 in./sec. It is also expectedthat for large flooding rates, liquidentrainment will become significant,particularly when the fuel rods are atinitially high temperatures. A floodingrate of 0.2 in./sec of saturated wateris just marginally able to remove thetotal core-decay heat at the end ofblowdown assuming complete coolant vaporization.Flooding rates of thisorder will not prevent core heatup ifVIII-48

the fuel rods are initially at elevatedtemperatures and liquid entrainment isencountered.A1.1.2.3 Core Meltdown.To scope the effects of core meltdown oncore heatup, three core meltdown modelswere developed for incorporation inBOIL. The models are not phenomenologicalin the sense that slumping is notbased on calculations of stress levels,'creep rates, or flow rates of molten materials.Fuel slumping is triggeredwhen a fuel node reaches the meltingncjint Of U02 and absorbs additional energyequal to the latent heat of fusion.Two of the meltdown models (Models A andB) assume the molten fuel is retained inthe core as a continuous region, and onemodel (Model C) assumes the molten fuelto fall out of the core as it melts.Calculations indicate that the differentmodel assumptions can significantly affectthe course of core heatup, primarilybecause of the influence of the coremeltdown on the boiloff rate and thecladding-water reaction. This sectionof Appendix A, thus, contains a discussionof the core meltdown and a descriptionof the meltdown models developedfor BOIL.A1.1.2.3.1 Core-Meltdown Behavior. Thebehavior of a core during a meltdownaccident is uncertain. No cores havebeen melted. Experiments involving morethan a cupful of molten U02 are still inthe planning stages. Some properties ofmolten U0 2 are known: melting point,boiling point, and heat of fusion. How'-ever, little is known about the viscosity,internal thermal convection, surfacetension, or the metallurgical effects ofvarious diluents. Nevertheless, considerableinsight has been developed bypeople working in the field about thepossible course of core meltdown.Because the core power tends to peak(about 2.5 times average) at the centerof the core and because the claddingsteamreaction increases the heatup rate* of the hotter regions, core meltingstarts at the center of the core. Becauseof the power peaking and the presenceof water in the bottom of thecore, the core temperatures a foot awayfrom the melted region are frequentlycalculated to be more than 1000 F belowthe melting point of the fuel. In theserelatively cool regions, the U0 2 wouldremain solid although the cladding couldbe melted. Because the fuel rods in thecore are relatively closely packed,there is not room for solid fuel pelletsto fall out of the core nor for grossdistortion of the solid portions of thecore. In this situation, it is believeda region of solid rubble would form underthe molten fuel, and the molten fuelwould tend to be retained in the core.However, since the rubble continues togenerate heat, it will eventually melt,and the increasingly larger molten regionwill move downward. If the poolmoves downward fast enough, it will interceptthe water that is boiling out ofthe bottom of the core. When this happenseither (1) steam explosions willoccur or (2) the boiloff rate and,therefore, the cladding-steam reactionrate will increase. when the moltenregion grows to include 50 to 80 percentof the core, it becomes questionablewhether or not the molten region can beretained inside the core. At this time,the molten pool will be 3 to 4 ft thick,and will presumably be held up by a layerof rubble. When large fractions ofthe core are molten, the core-supportplates and shrouds are also exposed tohigh thermal loadings. Failure of thesemajor structural members would releasethe molten pool and either (1) the restof the water boils out of the pressurevessel or (2) a steam explosion results.It is assumed in the development ofmeltdown models A and B that a largepool of molten U0 2 can be retained inthe core. It should be noted that aconsequence of this assumption is thatdamaging steam explosions become moreprobable. For a holdup mechanism toexist, at least two prerequisites arenecessary: (l) there must be a mechanismfor removing the decay heat fromthe molten pool of fuel and (2) theremust be a solid crust or rubble layer atthe bottom of the pool to retain theliquid U0 2 'The possible existence of a crust whichsupports or tends to restrain the motionof the molten pool Of U0 2 has been investigatedanalytically. Attempts weremade to show (by calculation) that (1) ahot liquid film Of U0 2 flowing down acold, solid, fuel rod or (2) a dropletof molten U0 2 falling in air through acold array of rods would quickly resolidify.However, the calculations wereinconclusive because of the difficultyof justifying the model assumptions andfinding experimental data on the viscosityand surface tension of liquid U0 2 ,Neglecting surface-tension effects,these calculations indicate that smalldroplets and films 0'u50 mil) will resolidify.However, films and droplets aslarge as the rod-to-rod spacing may notresolidify before dropping out of thecore. Another calculation including theeffects of surface tension indicatesVIII-49

that droplets may have to weigh morethan 5 grams before they will start tofall. It is concluded that the existenceof a holdup mechanism for moltenU0 2 is largely based on experience withother materials and, to some extent, intuition;based on present knowledge, thecalculational evidence is inconclusive.The work of Hesson indicates that internalnatural convection will be adequatein most cases to remove the decay heatfrom the interior of a pool of moltenU0 2 retained within the core (Ref. 9).Hesson's experiments were performed usinginternally heated saltwater solutions,and the data were correlated interms of the usual dimensionless heattransferparameters. With thermophysicalproperty data appropriate to moltenU0 2 , the correlations are used to estimatethe maximum heat fluxes obtainablefrom the surfaces of a molten pool ofU02 without boiling off the U0 2 . Inthese correlations boiling inside thepool is assumed, but there is no netloss of U0 2 from the pool by vaporization.The boiling inside the pooldrives the natural circulation, and thevapor is condensed inside the pool.Values obtained for the maximum downward(qD) and horizontal (qH) heat fluxes areandqD = 133,000 Btu/hr ftqH = 382,000 Btu/hr ft22(VIII A-21a)(VIII A-21b)Hesson does not give a value for theupward heat flux. Since the naturalconvection heat transfer coefficientsfor horizontal and vertical plates areapproximately equal (Ref. 4), it isassumed that the upward heat flux (qT)equals the horizontal flux. Assumingthe pool is a cylinder and using theseheat fluxes, the total heat lost fromthe pool by convection can be calculatedand compared with the decay heat. Theresults of the calculations indicatethat natural convection will transportthe decay heat from a completely meltedcore to the outside surface if themeltdown occurs after about 6 minutes inthe BWR and after 17 minutes in the PWR.Since complete core meltdown is notlikely to occur this early in the accident,it is concluded that a pool ofmolten UO 2 will not boil away during thecore-meltdown process.Although UO2 boiloff is not expected tobe a significant heat loss mechanism,some U0 2 vaporization may occur. Adense smoke, which is apparently U0 2vapor, has been observed experimentallyabove molten samples of U0 2 . This smokecould affect some core heat-transferprocesses by inhibiting thermal-radiationheat transfer. In the BOIL calculationsthe heatup of the support structuresabove the core is controlled byradiation heat transfer. Thus, the existenceof the U0 2 smoke could reducethe heatup rate of these structures asdetermined by radiation; this would becompensated by the heatup due to thecondensation of the U0 2 smoke on thesame surfaces. The UO 2 smoke would notsignificantly affect the core heatuprate, however.A1.1.2.3.2 Meltdown Model A. In meltdownModel A, it is assumed that theheat in the molten pool is transferreddownward. There is no convection ofheat to the top and sides of the pool.This model maximizes the downward movementof the molten pool. According tothe model assumptions in BOIL, if themolten region moved downward any faster,it would resolidify. Hesson's work indicatesthat the convection heat fluxesfrom a molten pool are in the ratioqD:qH:qT = 1:2.9:2.9. Thus, the downwardheat flux in Model A exceeds Hesson'spure convection value. For ModelA to be physically consistent, it isnecessary for the molten region to actuallypenetrate the solid regions belowthe molten pool. Thus, Model A is physicallyconsistent with a meltdown situationin which the molten region tends tocover and mix (downward) with the solidregion at such a rate that the homogenizedmolten region remains just at themelting temperature. It is assumed thatno solid core material falls into themolten pool from above. Such an occurrencewould tend to keep the temperatureof the solid-liquid mixture at or belowthe melting point.Al.I.2.3.3 Meltdown Model B. In meltdownModel B, it is assumed that theheat in the molten pool is transferredupward, and none is transferred down.Within the molten region, heat may betransferred radially if the averagetemperature of a radial power regionexceeds the melting temperature. Theheat transferred upward is used in theBOIL calculations to melt solid corematerial, which is assumed to fall intothe molten pool. The amount of solidmaterial falling into the molten pool issufficient to keep the homogenized temperatureof the molten pool at the melt-VIII-50

ing point of U0 2 . When the top nodes inthe core are melted, it is assumed thisheat may be radiated to the supportstructures above the core. Model B isphysically more consistent than Model Awith the pure convection heat fluxesgiven by Hesson's work, where the downwardheat flux was comparatively small.Model A, on the other hand, considersthe downward motion of the molten fuel.Whether Model A or Model B is a betterdescription of an actual core meltdowncannot be stated definitely. The twomodels yield very similar results forcore-meltdown fractions of up to about50 percent. However, for larger meltdownfractions, Model A results infaster core heatup. In Model A, themore rapid downward progression of themolten region results in increasedmetal-water reaction when the moltenregion intercepts the water level. Ifit is assumed that at any time duringthe meltdown, as predicted by Model B, asmall part of the molten core (,I percentper time step) falls into thewater, the results in the two models aresimilar.A1.1.2.3.4 Meltdown Model C. In ModelC it is assumed that when a fuel nodemelts, it immediately falls to the bottomof the pressure vessel. The fuelnode is quenched in one time step, andthe decay heat of the node is added tothe water. The large boiloff ratesobtained under these assumptions resultin very high heatup rates, due to thecladding-steam reaction, between thetimes when core melting first starts andall of the water in the pressure vesselis boiled off. Model C is not believedto give a realistic picture of coremeltdown. It was developed to illustratethe effect of molten fuel droppingout of the core rather than being retainedin a molten zone within the coreregion.A1.2 RESULTS OF BOIL CALCULATIONSThe results of calculations arepresented in this section of the report.Although the details of the resultsdepend on the particular reactor beingstudied, many of the conclusions andtrends seen in the BOIL results havesome general applicability. The mostimportant variables in the calculationsappear to be the time after shutdownthat the boiloff sequence starts and theinitial pressure-vessel water inventory.These variables control the core-decayheat, the water boiloff rate, and thesubsequent core heatup rate. After coremelting starts, the meltdown or slumpingmodel used in the calculations can alsosignificantly influence the core heatup.In the calculations described in thisappendix, the time when the boiloff partof the accident starts and the sequenceof events preceding it are assumed to beknown. These sequences and their probabilityare discussed elsewhere.A1.2.1PWR RESULTSFor most of the BOIL calculations on thePWR, the initial water inventory in thepressure vessel was equivalent to anunswollen water level at the top of thecore. For these cases, the swollenmixturelevel is actually above thecore. This is representative of situationsin which the emergency core coolingsystem has functioned in the injectionmode. Many of the results in thissection are plotted as a function ofaccident time, measured from some predeterminedinitial condition in the BOILcalculations. As used in this sectionof the report, the "time when the boiloffcalculations starts" is not the timewhen the LOCA accident starts. It isthe time when the water inventory in thepressure vessel decreases to a knownunswollen level, usually the top of thecore. The time when the boiloff calculationstarts is discussed elsewhere; itdepends on the particular accident sequencebeing considered, e.g., the timewhen the emergency core cooling recirculationsystem fails. In addition toresults that illustrate the expectedbehavior of the reactor, the sectionincludes results that illustrate effectsof model assumptions in the BOIL program.Figure VIII A-2 shows how the swollenmixturelevel and void fraction at thetop of the core decrease for a boiloffaccident starting at 1.0 hr decay time.The initial unswollen water level was atthe top of the core (12 ft from the bottomof the core). Note that after 30min the average mixture void fraction(aT/2 ) is less than 7 percent, and theeffect of level swell is not significant.The most important effect oflevel swell is to keep the core completelycovered (and cooled) for almostthe first 10 minutes after boiloffbegins. Thus, the level swell retardsthe initial core heatup by about 10minutes for this accident. Figure VIIIA-1 also illustrates the growth of themolten-fuel zone in the core. The axiallocation of the melted fuel nodes in thehighest (radial) power zone are shown.(The core is divided into 5 radial and24 axial regions or nodes.) Note thatsignificant melting occurs while thereis still water in the bottom of thecore. -In meltdown Model A, the mixturelevel decreases rapidly after about 55VIII-51

minutes as molten fuel intercepts themixture level, and the core-meltdownrate increases due to increased metalwaterreaction. In Model B the moltenzone remains above the mixture forconsiderably longer, assuming no fuelfalls out of the molten zone. In theModel A meltdown, the molten zonereaches the bottom of the core in about60 minutes.Figure VIII A-3 shows the water-boiloffor steam-production rates for the samesituation considered in Fig. VIII A-2.The boiloff rate remains approximatelyconstant, at a comparatively high value,for the first 10 minutes because thecore is completely covered due to levelswell. The boiloff rate decreases asthe mixture level goes down and less ofthe core-decay heat is absorbed in thewater. At 60 minutes in meltdown ModelA, the boiloff rate increases rapidlyfrom quenching molten fuel entering thewater.Figure VIII A-4 illustrates typicaltemperature transients for fuel-rodnodes (in the hottest radial zone) atthe top, bottom, and midplane of thecore. The accident sequence is the sameas considered in Fig. VIII A-2. The topof the core uncovers first, and initiallyheats up at a rate dependent on thefuel-rod-decay heat. The midplane uncoverslater, but heats up faster due toits larger decay heat. When the nodesreach temperatures above 2500 F, theyshow typical heatup rates produced bythe zirconium-steam reaction. The rapidheatup of the bottom of the core after55 minutes is primarily due to the moltenfuel slumping in meltdown Model A.Figure VIII A-5 shows, for the sameaccident considered in Fig. VIII A-l,the fraction of the core melted as afunction of time for the three meltdownmodels discussed previously. Models Aand B both predict that 50 percent coremelting is reached in about 60 min.However, at about this time Model Apredicts that molten fuel contacts thewater, and the increased zirconium-steamreaction rate causes the results todiverge for larger meltdown fractions.A calculation is also shown, labeledModel B', which is the same as Model Bexcept that, after 60 minutes, itassumes 1.0 percent per minute of themolten core falls into the water. Theresulting increased steam flow andmetal-water reaction produces a resultsimilar to that obtained in Model A.Thus, if a small part of the moltenregion falls into the water, the Model Aand B results are similar. Since thisseems likely to occur, it is concludedthat meltdown Model A probably gives amore realistic picture of core meltdownthan Model B, particularly for verylarge (')80 percent) fractions of thecore melted.In the calculations above, it is assumedthat core melting does not prevent steamflow and stop the metal-water reactionin the core regions above the moltenfuelzone. If it is assumed that coremelting (channel plugging) stops themetal-water reaction in a given channelabove the melted zone, the result (labeledModel A') looks similar to thatobtained for Model B. Channel pluggingis probably more likely in a BWR corewith shrouded fuel elements than in anopen-lattice PWR core.Acronyms (labeled OZM, MZAB, UCBM, UCPM)indicate situations where massive structuralfailure of the core is possible.These situations are noted but the BOILcalculation is continued as if thefailure did not occur. If damagingsteam explosions occur, they are likelyto follow these massive structural failures.When the molten region reachesthe outside of the core (OZM), the corebarrel may be exposed to molten fuel.Since the core barrel supports the core,the core could drop into the bottom ofthe pressure vessel at this time. Inmeltdown Model B there is a large thermalradiation loading on the upper coreplate and core barrel (after the uppercore plate melts) above the core. Theupper core barrel is predicted to melt(UCBM) at 110 minutes in Model B. InModel A, the molten zone reaches thebottom of the core (MZAB) in about 61minutes, and the core would presumablyfall to the bottom of the pressurevessel shortly after. The metal-waterreaction with the molten iron is notconsidered in the BOIL calculation.Model C shows a very rapid heatup aftercore melting starts due to the largesteam flow rates and metal-water reaction.However, after about 40 minutes,all of the water in the pressure vesselis boiled off and the metal-water reactionstops. The results of the calculationsindicate that all of the corenodes hot enough (over 2500 F) to have alarge metal-water reaction rate quicklymelted, and the nodes that were below2000 F remained relatively cool. Thus,the calculated core condition showsabout half of the core in the bottom ofthe pressure vessel at an average temperatureof about 4000 F. The originalcore has a hole melted out of it, butthe remaining fuel is relatively cool(%1500 F). This core-meltdown model isnot believed to be realistic; however,VIII-52

it illustrates the significance of themetal-water reaction in the core heatup.Figure VIII A-6 shows the fraction ofthe cladding reacted for the sameaccident shown in Fig. VIII A-2. Forthe reasons discussed above (Fig. VIIIA-5), the results are sensitive to themeltdown-model and channel-plugging assumptionswhen more than about 30 percentof the cladding is reacted. Theresults of these calculations indicateit is highly likely that more than 30percent of the cladding will react, andthat much higher fractions are possible.Figure VIII A-7 shows how the results inFigs. VIII A-2 to VIII A-6 can be generalizedin terms of the time when theboiloff part of the accident starts.These results indicate that core meltdownin a boiloff accident can be expectedin a few hours even for accidentsstarting several days after core shutdown,assuming no further preventativeaction is taken. A calculation is alsoshown in Fig. VIII A-7 for an accidentstarting at 60 sec with an initial pressurevessel water inventory equivalentto an unswollen water level of 6 ft(core midplane).A1.2.2BWR RESULTSFor a boiloff accident starting at thesame decay time, the BWR and PWR resultsare quite similar. Differences in theboiloff calculations are primarily attributableto differences in the corepower density, pressure-vessel water inventoryat the start of the accident,and the fuel-element design (open latticeversus shrouded). At the start ofthe BWR boiloff calculations, the initialmixture level inside the coreshroud is 8 ft above the bottom of thecore. The presence of the fuel-bundleshroud has two effects: (1)-the Zircaloyin the shroud adds to the potentialfor metal water reaction, and (2) shroudprobably makes channel blockages morelikely and steam bypass of the pluggedchannels less likely. Thus, there arecompensating effects on the amount ofzirconium reacted. A third considerationis that the BOIL programming doesnot perform a specific calculation ofthe shroud temperature. The shroud istreated approximately by including it inthe fuel-rod cladding.In addition to the boiloff calculations,BOIL calculations were also performedfor a dry heatup (with no steam flow)and for a degraded bottom-floodingsituation starting with the end-ofblowdownconditions.Figure VIII A-8 shows the core-meltdownfraction as a function of time for twoBWR accident situations: (1) a dryheatup starting with end-of-blowdownconditions, and (2) a boiloff accidentstarting at 100 hr decay time with thecore covered to the jet pump nozzle(8 ft from the bottom of the core). Thecore temperatures at the end-of-blowdownwere given by the relation (Ref. 10)TR = 923 (F - 0.332) + 540(VIII A-22)where F is the local power-peaking factor(Ref. 10). This relation gives anaverage core temperature of 1157 F and apeak temperature (F = 2.1) of 2172 F atthe end of blowdown. In the boiloffcalculation it was assumed that thespray system was operative until theboiloff accident started at 100 hr.Thus, the whole core was assumed to beinitially cooled even though the mixturelevel was only at an elevation of 8 ft.For the dry heatup case in Fig. VIIIA-8, meltdown Models A and B predict thesame curve of fraction core melted versustime. However, the molten zonemigrates to different axial positions inthe two models. At 60 min in Model A,the molten zone has reached the bottomof the core. In Model B, the moltenzone is still 2.5 ft from the bottom ofthe core at an accident time of 60 min.After about 80 min in Model B, the topof the core is melted, and the radiationheat transfer from the surface of themolten pool begins to melt the steelstructure above the core.For the boiloff accident starting at 100hr, the core-meltdown behavior is similarto that obtained for the PWR. Forboiloff starting at 100 hr, both Fig.VIII A-7 for the PWR and Fig. VIII A-8for the BWR predict 80 percent coremelting with the molten zone at thebottom of the core (Model A) in about225 min. (Note that the boiloff in thetwo reactors starts with different waterelevations in the core.) The above calculationsassume there is steam flowingto unmelted core zones above the water.If it is assumed that melting causeschannel blockage so that no Zircaloycladding reacts above a melted fuel-rodnode, then the meltdown results forModel A look approximately the same asthose for Model B.For the boiloff accident shown inVIII A-8, about 25 percent of theZircaloy (cladding plus shrouds)Fig.corehasVIII-53

eacted at the time when 50 percent ofthe core is melted. This may increaseto 35 to 60 percent Zircaloy reacted,depending on the meltdown model selected,when 80 percent of *the core ismelted.Al. 2.3BOTTOM-FLOODINGBOIL calculations were performed forseveral degraded bottom-flooding rates.It was assumed that the core temperaturesand decay heat were given by theend-of-bJlowdown conditions at 25 sec.The water level in the pressure vesselwas assumed to be at the bottom of thecore, and saturated water was added tothe vessel inventory at rates equivalentto bottom-flooding rates between 0.02and 0.2 in./sec. In these calculations,metal-water reaction with the fuelbundleshroud was neglected. Theresults indicate that the worst floodingrate, in terms of producing the largestamount of zirconium-steam reaction, isabout 0.1 in./sec. For higher or lowerflooding rates, the metal-water reactionis reduced. For a flooding rate of 0.1in./sec, the amount of cladding reactedis between 20 and 70 percent, dependingon the assumptions made relative tochannel blockage caused by fuel melting.For flooding rates less than 0.2in./sec, the BOIL calculations predictmore than 20 percent core melting. Formeltdown Model A and a flooding rate of0.02 in./sec, the molten zone is predictedto be at the bottom of the core atan accident time of 230 min.A2. CONTAINMENT HEAT TRANSFERA2.1 PWR ANALYSESA2.1.1ENERGY SOURCESThe reactor was assumed to be operatingat full power at the time of the accident.The integrated power generationfor the first 10 sec after the break wasdetermined from a typical power shutdowncurve for large breaks. From 10 sec on,power generation was determined by theproposed ANS standard (Ref. 1) for decayheat assuming infinite irradiation time.The assumption of infinite irradiationtime, on the one hand, will overestimatedecay heat generation; on the otherhand, it should compensate for potentiallynonconservative uncertainties inshutdown heating.The sensible stored heat of the reactorvessel and internals was assumed to begiven up to the cold injection water.It was found that the stored heat wasapproximately equal to that required toraise the injection water in the refloodedvessel to saturation. The heatstored in the primary piping as well asin secondary coolant in the steam generatorswas given up to the steam generatedby core-decay heat. The rate of heattransfer was found to be limited by theheat capacity of the steam.The energy generated by the zirconiumwaterreaction was not input into thecontainment pressure calculations directly,but was assumed to go into coreheating. It is recognized that thegases (i.e., steam and hydrogen) leavingthe core which is undergoing significantmetal-water reaction will be at veryhigh temperatures; the influence of thishigh-temperature gas on the containmentpressure behavior is problematical, however.For a cold-leg break, the hotgases will have to pass through theprimary-system piping as well as thesteam generators to reach the containmentatmosphere; in such situations thegases will be cooled to the temperatureof the piping and steam generators. Fora hot-leg break the heated gases willhave a direct path to the containment;but, at extremely high temperatures, thehydrogen would be expected to igniteupon exposure to the containment atmosphere.The energy generated by hydrogenburning would far exceed the sensibleheat of the gas stream. Energy releasesdue to hydrogen burning, where applicable,were included in containment energybalances.A2.1.2CONTAINMENT HEAT SINKSThe available passive heat sinks withincontainment consist of the containmentliner, miscellaneous metal, and concretestructures. The quantities and geometricalparameters of each were taken orinferred from the FSAR. A steam-condensingcoefficient of 150 Btu/hr-ft 2 -Fwas assumed for all "cold" surfaceswithin containment: this value is consistentwith the results obtained in theCVTR containment experiments (Ref. 13).The rate of heat absorption by thevarious sinks was evaluated by means oftransient-conduction nomograms, as givenby McAdamns (Ref. 4) and others. Withthe above condensing coefficient, thecontainment liner would be heated to thetemperature of the steam in the atmospherein less than an hour. Since thespecific dimensions and quantities ofmiscellaneous metal in the containmentwere not available, the latter wasassumed to heat up at the same rate asthe containment liner. The concretewalls within containment were heated onboth sides; the external walls andfoundation, on one side only. HeatVIII-54

absorption by the concrete was found tobe governed by conduction within theconcrete itself and largely insensitiveto the condensing coefficient utilized.A2.1.3 CONTAINMENT ATMOSPHEREThe temperatures and pressures in thecontainment atmosphere are assumed to behomogeneous. The containment spraywater, if available, is heated to thecondensing temperature of the steam inits passage through the atmosphere. Theenergy due to hydrogen burning, if applicable,is distributed uniformly overthe contents of the containment atmosphere.A2.1..4CONTAINMENT PRESSURE-TIMEEVALUATIONThe evaluation of the containmentpressure-time history, taking into accountthe above described cons iderations,is carried out as follows. Atany point in time, an energy balance inthe containment atmosphere is calculatedby summing the energy in the steampresent at the previous time, 'the energyadded by the steam from the core, thewater added by the sprays, the heatremoved by the heat exchangers, lossesby condensation on the containment heatsinks, and leakage from the containment.This gives the total mass and averageenthalpy of the water in the containmentatmosphere. Using this enthalpy and anassumed pressure (which is based on thepressure at the previous iteration) aquality, and hence the amount of steamin the containment atmosphere, is determined.From this quantity of steam andthe known volume of the containmentbuilding a resulting partial pressure ofsteam is determined. If the sum of thispartial pressure of steam and the partialpressure of the noncondensablesdoes not agree with the pressure assumedfor the computation of the quality, thecalculation is repeated with an adjustedpressure until a consistent solution isreached. The steam and water propertiesutilized are adjusted in accord with thepredicted temperatures and pressures.The steam that condenses and the spraywater are returned to the containmentsump and mixed with the water there.Mass and energy balances are computedfor the sump water at each point intime.After the conditions in the containmentatmosphere and in the sump have beenevaluated, time is advanced and theprocedure is repeated as many times asrequired to cover the accident sequencein question. As time is advanced, appropriateadjustments are made to thedecay power level, safeguards operation,rates of heat loss, etc. The results ofthe containment pressure-time evaluationfor the various accident sequences aregiven in the body of this report.A2.1.5CONTAINMENT LEAKAGEThe release of fission products to theenvironment is controlled by the leakrate from the containment. Prior togross containment failure, the leakageis determined by the operational leakrate or by the characteristic hole sizefor containment-isolation failure andthe containment pressure. If the internalpressure exceeds about 25 psia, theleakage flow through an assumed orificewill be choked and the leak rate will be4.2 x 10-2 ybo/hr with the containmentisolated and 42 ybo/hr with containmentisolationfailure. The leak rate isassumed to vary linearly from zero tothe choked flow rate in the pressurerange 15-25 psia. If the containmentpressure is subatmospheric, as would bethe case with all safeguards functioning,the leakage is considered to bezero.In the event of a containment-isolationfailure, the pressure would approach oneatmosphere rapidly after the blowdown.Leakage will then depend upon theavailability of a driving force. Withthe containment sprays and containmentheat exchangers operating, any steamgenerated will be condensed and the onlycontributors to leakage will be thehydrogen from the zirconium-water reactionand the carbon dioxide from thedecomposition of concrete during meltthrough.In sequences where the con-.tainment sprays are not operating, steamgeneration could be the principaldriving force for leakage. Similarly,if the sprays are working but the containmentheat exchanger has failed,condensation will not take place afterthe spray water has reached saturation.If the effective generation rate ofgases and vapors is less than 42 v/o/hrwith containment-isolation failed, theleak rate is equal to the generationrate. If the generation rate exceeds 42v/o/hr, the leak rate is held at thislevel and the pressure is allowed torise as required to maintain consistencybetween production and loss of gases andvapors.Containment failure by steam explosion,overpressure, or meltthrough is accompaniedby a puff release of a fraction ofthe containment atmosphere; the magnitudeof the puff release will depend onthe containment pressure at the time ofVIII-55

the failure and the location of thefailure. Failures due to steam explosionsand overpressure are assumed tooccur above ground and vent to theatmosphere. Meltthrough failures ventto the atmosphere against the statichead of the groundwater; in some casesthe latter may exceed the pressurewithin containment.After containment failure due to a steamexplosion or overpressure, the breach incontainment is assumed to be too largeto limit flow and the leak rates aretaken equal to the net generation ratesof gases and vapors. For sequences inwhich sprays are operating and a steamexplosion fails containment, the spraysare assumed to continue operating, eventhough there is a significant probabilitythat the steam explosion may fail thesprays also. For over-pressure failurethe spray pumps would cavitate if theywere operating at the time. If the primarymode of containment failure ismeltthrough, the subsequent pressurewithin the building will be controlledby the static head of groundwater surroundingthe building foundation. Anysteam generation would be condensed asit flows through the breach in thefoundation mat. Noncondensables willleak out through the ground as well asthrough normal leakage paths.During the period of core melting thesteam generation rate is approximatelyequal to the available decay heat, i.e.,total decay heat reduced by the escapeof the volatile fission products. Allthe hydrogen production is assumed totake place during initial core meltdown.After the molten core drops to the lowerplenum, the water there is evaporated in15 minutes by partial quenching of thecore. During the remainder of pressurevessel meltthrough there is no furthergas or vapor generation. After the meltdrops into the vessel cavity it willstart to vaporize any water there aswell as attack the concrete. If neitherthe ECI or CSI operates, there will belittle or no water in the vessel cavity.With both ECI and CSI operating, 1.83 x105 lb of water will be available toproduce steam. If the CSR is operating,a continuing supply of water will beavailable for steam generation in thecavity, although the steam may be condensedsubsequently. Meltthrough of theconcrete generates 0.2 lb of steam and0.264 lb of CO2 per lb of concretedecomposed.The uncertainties in the calculated leakrates are believed to be relativelylarge. Many of the assumptions, particularlythose associated with steam explosionsand containment meltthrough,are intended to be representative of thephenomena that could exist, rather thanattempts at an exact delineation ofevents. The range' of possible leakrates is estimated to be within a factorof two of the calculated values. It isbelieved that in many sequences theconsequences of fission product releasewill be dominated by the puff releaseand thus will be relatively insensitiveto the uncertainties in the timedependentleak rates.A2.2 BWR ANALYSESA2.2.1ENERGY SOURCESThe reactor was assumed to be operatingat full power at the time of the accident.Reactor scram, though not necessaryfor shutdown, was assumed to takeplace early in the accident. Decaypower was determined from the proposedANS standard for decay heat assuming aninfinite irradiation time (Ref. 1). Inthe absence of a control-rod scram, thereactor would be shut down by core voidsand the integrated power during shutdownwould not differ appreciably from thatof the scram case. The absence of scramor control rod run-in, however, couldlead to recriticality and subsequentpower generation on core reflooding bythe emergency core-cooling system. Asbefore, the assumption of an infiniteirradiation time will overestimatedecay-heat generation which could compensatefor other, nonconservative uncertainties.The sensible stored heat of the reactorvessel, internals, and primary pipinginside the dry well were considered tobe given up to the cold core-coolingwater. The energy generated by thezirconium-water reaction was assumed togo into core heating.A2.2.2CONTAINMENT HEAT SINKSThe available passive heat sinks withincontainment in addition to the suppressionpool water are the dry-well vessel,miscellaneous metal within the dry well,and the concrete in the dry well. Sincethe dry well is separated from itssurrounding structures by a 2-in. airgap, the only heat-transfer path out ofthe dry well is by conduction throughthe concrete base. The dry-well vesseland the concrete base were assumed to beheated on one side, the concrete wallswithin the dry well were assumed to beheated on both sides. The rate of heatabsorption by these sinks was evaluatedby means of transient-conduction nomo-VIII-56

grams, such as given by McAdams (Ref. 4)and others. A steam condensing coefficientof 150 Btu/hr-ft 2 -F was assumedfor "cold" surfaces within the dry well.Within the above guidelines the dry-wellvessel was found to be heated to thetemperature of the steam in less than anhour. Heat absorption by the concretewas found to be governed by conductionwithin the concrete and insensitive tothe condensing coefficient utilized.The temperatures of the suppressionchambercomponents were assumed to be atthe temperature of the water.A2.2.3CONTAINMENT RESPONSEPrimary-system depressurization leads toa rapid rise in dry-well pressure, whichin turn leads to the flow of the releasedsteam and water into the suppressionpool where the steam is condensed.During the blowdown process, essentiallyall the air initially present in the drywell will be transported to the gasspace in the suppression chamber. Atthe end of the blowdown, with both coresprayand core-flooding systems operating,the pressure in the dry well decreasesas the relatively small quantityof steam remaining there begins to condense;the pressure in the suppressionchamber, being largely controlled by thenoncondensables, stays relatively constant.When the pressure in the drywell drops below that in the suppressionchamber, vacuum breakers open and allowthe noncondensables in the suppressionchamber gas space to flow back into thedry well. This results in a decrease inthe pressure level within containment asthe noncondensables equilibrate betweenthe dry well and wet well and the steamin the dry well continues to condenseand move toward equilibrium with thepool. Under design conditions, thepressures in the two portions of containmentremain equilized indefinitely.Under design-basis accident conditionsthere is little or no net steam generationfrom the core during reflooding.As a consequence, what steam is generatedwill be condensed by the suppressionpool, and the noncondensables will beequalized between the dry-well and wetwellgas spaces. With degraded performanceof the standby core-cooling system,significant steam generation by thecore may continue during the refloodingphases of the accident; this wouldresult in the dry-well pressure remainingabout 3 psi higher than that in thesuppression chamber with no equalizationof the noncondensables. Under such conditionsthe containment would remain atan elevated pressure for an extendedtime.Functioning of the vapor-suppressioncontainment as designed leads to relativelylow accident pressures within thecontainment except for the early portionsof the blowdown. If, however,condensation of the steam generated byprimary-system depressurization does nottake place, for whatever reason, pressuressignificantly in excess of designlevels would result. Assuming thatsteam condensation does not take placeand that the entire free volume of thecontainment is available, primary-systemdepressurization would result in a pressureof about 250 psia within thereactor-containment design assumed forthis study.The failure of long-term cooling withall the other safeguards functioningwould lead to a gradual increase in thesuppression-pool water temperature and acorresponding gradual increase in pressure.With a design-basis containmentleakrate the loss of long-term coolingwill eventually lead to containmentfailure by overpressurization. For containment-leakrates in excess of about100 percent per day, overpressurizationfailures would not be expected.The containment pressure-time historiesfor the specific accident sequences consideredare given in the body of thisreport.A3. PRESSURE VESSEL MELTTHROUGHA3.1 PWR ANALYSESIt is assumed that, as the core temperaturerises, most of the core remainsabove the grid plate until a majorfraction is molten (\80 percent). Atthis time the molten fuel drops rapidlyto the bottom of the vessel where it ispartially quenched as it boils away thewater in the lower plenum. The timerequired for the core to break throughthe pressure vessel serves as a delayperiod before the melt can begin to workon the containment floor. The time atwhich the molten fuel enters the lowerplenum in the following analyses is -2hours.Using J. C. Hesson's model (Ref. 11),the heat convected downward from themolten pool through a solid layer ofU0 2 into the steel head isQt =2k (T 2 - T 1 ) q + Q2(VIII A-23)VIII-57

whereQt = heat flux into steelk = thermal conductivity in solidUO 2T2 = melting point of UO 2T1 = melting point of steelq = volumetric heat generation incoreQ = heat flux from liquid to solidUO 2 -plies the driving force for convection.The core will therefore absorb additionalheat as its temperature is increasedto an average between melting and boilingpoints, %5475 F.As the stress in the bottom head fromthe weight of the core is relatively low(< 1000 psi), failure of the head is notexpected until it becomes quite weak athigher temperatures. By extrapolatinglower temperature data, the tensilestrength of vessel steel decreased to1000 psi at about 2400 F. From this, anaverage steel temperature of 2400 F isassumed to be the failure point.Q isevaluated fromA3.1.1REALISTIC FAILURE TIMEQ = \ k2) ) At,where C, m, and n are constants(VIII A-24)S = vapor fraction at differentlevelsL = pool depthAt = temperature difference in poolabove the boiling point ofU0 2 .Hesson found Q independent of L,m = 1/3, and n between 0 and 1/3 inexperiments with internally heated saltwaterbaths. Allowing S = 1 for maximumconvection and using Cp = 0.124 cal/g-C,= 0.46 poise, p = 8.7 g/cm 3 , andk = 0.005 cal/sec-cm-C values of Q of8.52 cal/sec-cm 2 for n = 0, 10.43 forn = 1/4, and 10.22 for n = 1/3determined.Using the value 10.22 cal/sec-cm 2 C(136,000 Btu/ft 2 hr F) and solving forthe heat flux into the steel, we findthat Qt = 1.39 x 105 Btu/ft 2 -hr F forour case.Since this total heat flux is onlyslightly greater than the convectioncomponent, the internal heating componentin the solid layer is small andtherefore the solid U0 2 layer is thin.Thus the quenched core will absorb heatuntil nearly the entire core remelts,or, ignoring this thin solid layer, heatequal to that removed by the evaporatingwater during quenching will be absorbedto remelt the core. In addition, themodel assumes the temperature gradientin the melt to vary from its meltingpoint to its boiling point which sup-The time from the core falling into thebottom of the vessel until the bottomhead fails is the time for the decayheat to heat the core to its equilibriumtemperature, plus the time to heat thebottom head to its failure temperatureat the above heat flux Qt.tI =AQ water + (MC AT) coreQdecay(VIII A-25)where AQ is the heat to evaporate thewater; MCpAT is the mass, specific heat,and temperature difference in the moltencore; and Qdecay is the decay heat.3.13 x 107 + 0.62 x 107t6.00 x 10 = 0.625 hr,t 2and(MCp AT) steelQt= 0.406 hr,t + t 2 = 0.625 + 0.406= 1.03 hr = 62 min.Mass is expressed in lb/ftsteel head.(VIII A-26)224* x 0.12 x 21001.39 x 105(VIII A-27)(VIII A-28)2of theamVIII-58

It is unrealistic to assume that theseevents occur consecutively. However,when bounding estimates are made on thetime to failure, more detailedcalculations do not appear to bewarranted.A3.1.2MAXIMUM FAILURE TIMEIf it is assumed that only the longerlived (tl/2 > 3 min) noncondensablefission gases are lost from the vesselso that 80 percent of the decay heatremains in the vessel, the time requiredto heat the whole vessel and internalsto 2400 F isA3.1.3AQ water + (MC pAT) steelQdecay= 3.13 x 107 + (6.855 x 105 x 0.113x 2.100 x 103)/6.86 x 107= 2.67 hr = 160 min. (VIII A-29)MINIMUM FAILURE TIMESome of Hesson's experiments showed horizontalheat flow approaching 4 x 105Btu/ft 2 -hr F. If it is assumed thisheat flux is directed into the steelnear the surface of the molten coret 5.65 x 10 + 0.625 = 0.766 hr4.0 x 105= 46 min. (VIII A-30)A3.2 BWR ANALYSESThe lower head of a BWR pressure vesselis partially filled with control-rodguide tubes and contains a number ofcontrol-rod-drive penetrations. Whenthe molten core drops into the bottomhead, the drive penetrations may permitsome of the molten material to escapeprior to the failure of the bottom headas a whole. In other respects, analysesof BWR vessel meltthrough follows thesame reasoning as given above for PWRs.If the time required to heat the entirevessel to 2400 F with decay heat isconsidered an upper bound to the timerequired for penetration, then penetrationof the bottom head in the absenceof water will occur in about 72 min. Anadditional 30 min will be required ifthere is water in the bottom head. Thebest estimate for the meltthrough timeafter the core drops into the lowerplenum is 30 + 20 min for the drymeltdown case and 60 + 30 min if thereis water in the bottom head. Theseresults are based on core meltdownstaking place early in the accident; forthe later onset of core melting, themeltthrough time will be extended due tothe decrease in decay heating.A4. CONTAINMENT VESSEL MELTTHROUGHA4.1 PWR ANALYSESThe processes by which the molten coreinteracts with the concrete floor of thecontainment building are very complexand not fully understood. In the absenceof definitive experimental informationit is only possible to estimateapproximately the time required forpenetration of the containment floor bythe molten core. When the molten fuel(together with molten zirconium, zirconiumoxide, steel, iron oxide, etc.)falls onto the concrete, vaporization offree water below the surface will causespalling of the concrete and result in avery rapid penetration rate of the meltinto the concrete. Based on the vaporizationof the free water, initial spallingrates are calculated to be 15-30ft/hr. As the concrete heats up it willgive up its water of hydration at about900 F; then at 1400-1600 F the limestonewill decompose into CO 2 and CaO. It isexpected that the water vapor and carbondioxide would escape from the melt andbe released to the containment atmosphere,but the calcium and siliconoxides would stay with the melt. As theproducts of concrete decomposition areabsorbed and/or dissolved by the melt,the melt temperature will decrease untilconstituents of the mixture begin toprecipitate. It is estimated that bythe time the melt has penetrated through1-1/2 feet of concrete, U0 2 would, beginto precipitate from the multicomponentmixture. From this point on the progressof the melt through the concreteis controlled by the rate of decay-heatgeneration, i.e., the quantity of concretepenetrated is directly proportionalto the energy available to decomposethe concrete and raise the temperatureof the products to the temperature ofthe melt.The temperature required to maintain themulticomponent mixture in a fluid stateis not known. The principal componentsof the melt and their approximateVIII-59

melting points are: U02 (5200 F) , ZrO2(4900 F), Zr (3350 F), stainless steel(2550 F), Fe (2790 F), FeO (2590 F),Fe 2 0 3 (2850 F), CaO (4680 F) and SiO 2(3100 F). The melting point ofCaO.SiO 2 is 2750 F; however, in ordinaryconcrete there is almost twice as muchcalcia as silica. A melt temperature of4000 F has been assumed for the presentanalysis.After the initial rapid penetration ofconcrete by spalling, the melt is expectedto remain relatively viscous,decomposing and dissolving concrete at arate compatible with decay-heat generation.Because the silica and calciathat are introduced at the lower surfaceare less dense than the body of themelt, convection currents that should beestablished will prevent the center ofthe melt from reaching very high temperatures.Also, the carbon dioxide andwater vapor released by the concrete,and/or their further decomposition products,will provide additional agitationas they rise through the melt.The upper surface of the melt is likelyto be covered with a solid crust and tobe radiating heat to the remains of thepressure vessel and to the walls of thereactor cavity. If there is water inthe cavity it will be vaporized by themelt, but even the continued addition ofwater would not avert containment meltthroughsince the geometry of the moltenmass is not favorable for effectivecooling. If the structures above themelt reach elevated temperatures, theycould fall into the melt.That at least part of the mass of fuel,structural, and other material remainmolten is required by heat-transfer considerations.If this mass should solidifyat some point during the accident,then the only mechanism for dissipatingdecay heat would be conduction. Forconduction to transfer the decay heatout. of the largely low conductivity massof material involved, temperatures atthe center of the mass would have toexceed boiling points of some andmelting points of all the constituents.Hence the conclusion that at least someof the mass will remain in a fluid statefor considerable time, with convectionwithin the melt tending to maintain themelt temperature near the effectivemelting point of the mixture. It isrecognized that as the size of the meltincreases and the heat-generation rateper unit volume decreases due todilution, solidification must eventuallyoccur. Solidification is not, however,expected to take place prior to thepenetration of the containment foundationmat.In estimating the time required to penetratethe 10-ft-thick foundation matthree different configurations wereconsidered for the quantity of concretedecomposed by the molten core: (1) a15-ft-diameter, 10-ft-high cylinder, (2)a 10-ft-radius hemisphere, and (3) a 32-ft-diameter, 10-ft-high cylinder. Thefirst case assumed that the melt progresseddownward through the concretefaster than it did horizontally; thesecond case assumed equal rates of progressof the melt downward and horizontally.The third case assumed that themolten core materials spread within theconfines of the reactor cavity and thenattacked the concrete at equal rates inthe downward and horizontal directions.Because the carbon dioxide and watervapor produced by the decomposition ofconcrete would tend to sparge the meltof fission products, the decay heatutilized for this analysis correspondedto 60 percent of that at the time ofLOCA. This implies complete loss of thevolatile fission products and fractionalremoval of some of the less volatilespecies from the melt. Assuming all thedecay heat goes into the concrete, thetimes required to penetrate the concreteand bring the entire melt to 4000 F willbe 7, 9, and 36 hours for the threecases considered. These times are basedon contact of the concrete by the moltencore at 2-3 hr after the start of theaccident. Making an allowance for heatlosses from the upper surface of themelt, the best estimate for the timerequired to penetrate the containmentfoundation mat is 18 hours.More rapid meltthrough of the concretethan calculated could occur if the spalledpieces of concrete were floated tothe surface of the melt without undergoingdissolution and being elevated tothe temperature of the melt. In choosingthe distribution for the time ofcontainment meltthrough as 18 (+10, -5)hours from within the broad range ofuncertainty, attention was paid to thepotential competition between containmentmeltthrough and containment overpressurizationas failure modes. Whereasthe consequences of meltthrough arerelatively small and insensitive tomeltthrough time, it was considered essentialnot to underpredict the meltthroughtime in such a manner as topreclude the possibility of failure byoverpressurization if it would haveoccurred at a later time in the absenceof meltthrough. The distribution forcontainment meltthrough time overlapsthe expected time of containment over-UVIII-60

pressurization in such a manner as togive a high probability of overpressurizationfailure in those accidents inwhich it is a potential failure mode.The effect of the steel liner and rebarsteel have not been included in theabove analyses; the effect would not beexpected to be large, the rebar being ofthe order of 5 percent of the concreteby volume.An evaluation of the penetration of amolten core through basaltic concretehas been performed by Jansen andStepnewski (Ref. 12). Their analysisassumed a basaltic concrete compositionwith the water removed; the meltingpoint of the basalt was taken to be,2000 F; 80 percent of. fission productswere assumed to stay with the melt.This material is considerably differentfrom the ordinary concrete utilized inthe present analysis. Their results fora 3000-Mwt reactor indicate penetrationof 10 ft of concrete in approximately 5hours. The assumptions made in thisanalysis would tend to under-predictmeltthrough time for the present study.In calculating the response of the containmentpressure to the release ofgases during the time the core ismelting through the concrete, the rateof carbon dioxide generation has beencalculated for a typical concrete composition.It should be recognized thatnot all concretes contain carbonatesthat would result in carbon dioxideevolution. The water released from theconcrete has been assumed to escapewithout interacting with constituents ofthe melt to produce hydrogen. Since alarge amount of molten steel couldaccompany the molten core, hydrogenproduction would be expected to occur ifthe steam were to bubble up evenlythrough the melt rather than to bypassit. If a mass of steel equal to thelower head of the pressure vessel werereacted, enough hydrogen would beproduced to increase the containmentpressure by - 15-20 psi for accidents inwhich containment sprays operate. Thepressure at the time of meltthroughwould still be well below the predictedfailure pressure, however. For accidentsin which containment sprays orcontainment heat removal are inoperativein the meltthrough period, water vaporand hydrogen would have equivalenteffects on containment pressure. Thepressure transient would therefore beessentially unaffected by whether or notthe reaction occurred.A4.2 BWR ANALYSESThe phenomena associated with the penetrationof the concrete foundation bythe molten core in a BWR would be thesame as described above for the PWR.However, because of differences in containmentconfiguration and volume, theconsequences of containment meltthroughin a BWR are expected to be quite differentthan in a PWR.When the molten core together with someof the reactor internals and part of thevessel bottom head fall to the floor ofthe dry well, vaporization of the waterthere as well as decomposition of theconcrete will ensue. The C02 and watervapor will be forced into the suppressionpool where the latter will condense.The C02 together with the othernoncondensables will collect in thesuppression-chamber gas space and, inthe absence of pressure relief, lead tofailure by overpressurization. FigureVIII 3-2 in the body of this reportshows the partial contributions as wellas combined pressures of the variousconstituents of the containment atmosphere.The partial pressure of C02 wasbased on the decomposition of approximatelya 20-ft-diameter by 8-ft-highcylinder of concrete below the reactorvessel, implying that the lateral spreadof the melt was fixed by the reactorvessel-supportcylinder and the progressof the melt was straight down. Thequantity of C02 released from the abovemass of concrete together with the othernoncondensables when confined to thesuppression-pool gas space will resultin a pressure more than sufficient torupture containment. If the melt progressesradially outward as well asdown, as might be expected, a largerquantity of CO 2 would be generated priorto the time that the melt reaches thedry-well vessel. The time of the overpressurefailure will depend on the particularaccident sequence of interest.The melt is estimated to penetrate thedry-well shell in about 12 hours afterthe core falls on the concrete.If some of the water produced by decompositionof the concrete were to reactwith molten steel to produce hydrogenrather than to bypass the melt, the timeof overpressure failure of the drywellcould be reduced. The uncertainty infailure time associated with the rate atwhich concrete is decomposed is muchgreater than the uncertainty associatedwith the potential for reaction betweenthe iron and water, however.Since the bottom of the dry-well vesselis embedded in concrete, initial melt-VIII-61

through of the steel shell would notnecessarily mean gross failure of containment.If a large hole were meltedthrough the bottom of the shell withoutit having failed previously, the ultimatestrength of the dry-well vesselcould be reduced significantly.After the melt has penetrated the drywellshell, it will continue to progressthrough the concrete foundation underneaththe containment. An additional 18hours are estimated to be required forcomplete meltthrough.After the molten material has penetratedthe concrete floor, the melt front willproceed into the underlying gravel andpossibly into the earth. The ultimateextent to which the molten zone can growdepends upon the heat removal processesat the upper and lower surfaces and thechemical and physical processes withinthe melt. Estimates have been made ofthe ultimate extent of the growth ofthis region. Assuming that heat removalat the surface is limited to conduction,the maximum radius of molten sphere hasbeen calculated to be 30 ft and 50 ftfor growth into media of limestone anddry sand, respectively (Ref. 14). Theanalyses of Jansen and Stepnewski (Ref.12) for basaltic concrete indicated amaximum radius of 38 ft for a moltenhemisphere. Since the ground underneathcontainment is well below the level ofthe water table, conduction heat transferat the surface of the melt should beaugmented by steam generation and convection.It is therefore likely thatthe melt will not proceed more than 10-50 ft below the bottom of the containmentbuilding. Greater depths couldonly be achieved if the core materialwere able to melt through the underlyingmaterial without mixing and beingdiluted by the products of decomposition.Although it has been predictedthat small pellets of solid U0 2 couldtravel some distance before being dissolvedinto the molten products of themedium being penetrated (Ref. 14), goodmixing should occur between molten U02and the products of decomposition ofconcrete or soil in the configurationsexpected in the meltdown accident.VIII-62

References1. "Proposed ANS Standard, Decay Energy Release Rates Following Shutdown ofUranium-Fueled Thermal Reactors", Subcommittee ANS-5, ANS Standards Committee(October, 1971).2. Carbiener, W. A., and Ritzman, R. L., "An Evaluation of the Applicability ofExisting Data to the Analytical Description of a Nuclear Reactor Accient,Quarterly Report for April through June, 1970", BMI-1885 (July, 1970).3. Walters, C. T., and Genco, J. M., "NURLOC-1.0, A Digital Computer Program forThermal Analysis of a Nuclear-Reactor Loss-of-Coolant Accident", BMI-1807 (July6, 1967).4. McAdams, W. H., Heat Transmission, 3rd Ed., McGraw-Hill Book Co. (1954).5. Rettig, W. H., et al., "RELAP-3, A Computer Program for Reactor BlowdownAnalysis", IN-1321 (June, 1970).6. Slifer, B. C., and Hench, J. E., "Loss-of-Coolant Accident and Emergency CoreCooling Models for General Electric Boiling Water Reactor", NEDO-10329 (April,1971).7. Wilson, J. F., et al., "The Velocity of Rising Steam in a Bubbling Two PhaseMixture", ANS Transactions, 5, (1), p. 151 (June, 1962).8. Wallis, G. B., One-Dimensional Two-Phase Flow, McGraw-Hill Book Co., p. 287(1969).9. Baker, L., Jr., and Hesson, J. C., "Heat-Transfer Model for Internal HeatGeneration in Pools", ANL-RDP-7, p. 9.25 (July, 1972).10. Morrison, D. L., et al., "An Evaluation of the Applicability of Existing Data tothe Analytical Description of a Nuclear-Reactor Accident, Core MeltdownEvaluation", BMI-1910, p. A-1 (July, 1971).11. Hesson, J. C., ANL-RDP-2, p. 8.37.12. Jansen, G., and Stepnewski, D. 0., "Fast Reactor Fuel Interactions with FloorMaterial After a Hypothetical Core Meltdown", Nuclear Technology, 17 (1), 85-95(January, 1973).13. Schmitt, R. C., Bingham, G. E., and Norberg, J. A., "Simulated Design BasisAccident Tests of the Carolinas Virginia Tube Reactor Containment - FinalReport", IN-1403 (December, 1970).14. Ergen, W. K., et al., "Emergency Core Cooling, Report of Advisory Task Force onPower Reactor Emergency Cooling", U.S. Atomic Energy Commission.15. Private communication with G. Jansen, Battelle's Northwest Laboratories.VIII-63

100a -16V0.1 L ,1500 2=FIGURE VIII A-12500 3000 350 4000 4500 5000Fuel Temeature, FCharacteristics CurvesFor Release of VolatileFission Products FromOverheated Fuel14012C:0108LevelMelted Fuel Nodes in Central Core Zone,Meltdown Models A and B0.j64230 40 5D 60 70 80Time After Boiloff Accident Gtarts, ruin90 100FIGURE VIII A-2Mixture Void Fraction and the Locations of theMixture Level and Molten Fuel Regions in a PWRfor Meltdown Models A and B as a Function ofTime for Boiloff Starting at 1.0 Hr

C,*35,325 Channels in Total Care'5A and Bcr0E010 20 30 40 50 60 70 80 90 100Time After Boiloff Accident Starts, minFIGURE VIII A-3Steam Generation Rate During a BoiloffAccident Starting at 1.0 Hr in a PWR asa Function of Time for Meltdown ModelsA and BC0onCaon60005000C'42-.C-,C40003000Axial Midplaneof Core (b ft).Ca1.0 ft Above Bottom20000~BCa*00on1000Ca0'4-0~0 10 20 30 40 50 60 70 80 90 100 110Time After Boiloff Accident Starts, minFIGURE VIII A-4Fuel Rod Temperatures in the Central (Maximum Power)Core Zone of a PWR as a Function of Time for BoiloffStarting at 1.0 Hr for Meltdown Model A

8070Meltdown Model-. A,B,CBMZAB60 1 2000eg A'5004030COZM - Outer Zone MeltsUCPM = Upper Core Plate MeltsUCBM = Upper Core Barrel MeltsMZAB = Molten Zone at Bottom of Core201000 10 20 30 40 50 60 70 80 90 100 110Time After Boiloff Accident Starts, min80FIGURE VIII A-5Effect of MeltdownPWR Core Melted asStarting at 1.O HrModel on the Percent of aFunction of Time for BoiloffA706050A'-, 40B= 30201000 10 !U 30 40 50 60 70 80 90 100 110Time After Boiloff Accident Starts, minFIGURE VIII A-6Effect of Meltdown Model on the Percent of theCladding Reacted in a PWR Core as a Function ofTime for a Boiloff Accident Starting at 1.0 HrFig. VIII A-1 - Fig. VIII A-6VIII-65/66

1000100-Molten zone onlower core plate,-80/. core melted00 10VVMixcure levelat core midplanecore meltedcore meltedcore meltedW-Note: Initial unswollen levelat 6 ft. this case only.01 .I11.0 10.Time When Boiloff Accident Starts, Decay Time,hr100.FIGURE VIII A-780Effect of the Time When a Boiloff(Decay Time) on the Core MeltdownPWR for Meltdown Model A- A-MZABAccident StartsSequences in aB-USHM -706060 B"UGPM ý-'"-B-UGPMS50OZMa 400A-MZABa 304DryS20 --A-OZMBoiloff accidents(t = 100 hr)1000 1I0 50 150 200 250 300Time After Boiloff Starts, min350 400 450FIGURE VIII A-8Percent of a BWR Core Melted as a Function ofTime for (1) a Dry Core Heatup Starting withEnd-of-Blowdown Conditions at 25 Sec and (2) aBoiloff Accident Starting at 100 Hours with theCore Flooded to the 8 Foot LevelFig. VIII A-7 - Fig. VIII A-8VIII-67/68

APPENDIX BPHYSICAL EXPLOSIONS RESULTING FROMCONTACT OF MOLTEN MATERIALS AND WATERVIII-69

Appendix BTable of ContentsSectionPage No.BI. REVIEW OF LITERATURE ............................................. VIII-73B1.1 Introduction ............................................... VIII-73B1.2 Representative Incidents Involving Steam Explosions ........ VIII-73BI.2.1 Metal Industry .....................................VIII-73B1.2.1.1 Mallory-Sharon Incident(Ref. 35) ................................VIII-73B1.2.1.2 Reynolds Aluminum Incident(Ref. 35) ................................VIII-73B1.2.1.3 Quebec Foundry IncidentB1.2.1.4(Ref. 4) .................................Western Foundries IncidentVIII-74(Ref. 36) ................................VIII-74B1.2.1.5 Armco Steel Incident(Refs. 37,38) ............................VIII-74Bl.2.1.6 East German Slag Incident(Ref. 2) ................................. VIII-74B1.2.1.7 British Slag Incident(Ref. 3) .................................VIII-74BI.2.2 Paper Industry .....................................VIII-74B1.2.3 Nuclear Reactor Industry ........................... VIII-74B1.2.3.1 Canadian NRX ReactorB1.2.3.2(Ref. 35) ................................Borax I ReactorVIII-75(Ref. 35) ................................ VIII-75B1.2.3.3 SPERT 1-D Reactor ......................... VIII-75B1.2.3.4 SL-I Reactor ............................. VIII-75B1.3 Metal Fragmentation and the Initiationof Physical Explosions .....................................VIII-75Bl.4 Interaction of U0 2 with Water .............................. VIII-77B1.5 Experiments with Metals in Water ........................... VIII-78B2. CONCLUSIONS ...................................................... VIII-79REFERENCES ..............................................................VIII-80VIII-71

List of Tables-TableVIII B-IVIII B-2Methods Leading to Metal Fragmentation and the Initiationof Physical Explosions ..........................................Results of Molten U0 2 Being Dropped Into Water atVarious Temperatures (Ref. 25) ..................................Page No.VIII-83/84VIII-83/84List of FiguresFigurePage No.VIII B-i Particle Size Distribution of Quenched U0 2 Samples .......VIII-85/86VIII-72

Appendix BPhysical Explosions Resulting fromContact of Molten Materials and WaterBII. REVIEW OF LITERATURE11l.1 INTRODUCTIONWhen molten metal comes into contactwith a quenching fluid, a violent explosioncan occur. This so-called "vaporor physical explosion" is a well-known,but little understood, phenomenon. Avapor explosion is usually characterizedby an initiating event that leads tofragmentation and by the sudden conversionof thermal energy to mechanicalenergy due to very rapid heat transferaccompanied by a subsequent pressurewave. The violence, or magnitude, ofsuch an explosion depends upon the quantityand rate of energy release. Numerousincidents have been reported in theliterature (Refs. 1-12). Such explosionshave occurred in the steel (Refs.1-4), aluminum (Refs. 5,6), coppersmelting (Ref. 5), paper (Refs. 7,B),and nuclear industries (Refs. 9,10).The mechanism that triggers or initiatesthe explosion is not known; however, twobasic facts have been established.First, the causative mechanism is notdue to chemical reaction (Ref. 13), andsecond, fragmentation of the sample materialis usually involved. Both experimentalresults and analyses (Ref. 14)have shown that the heat-transfer ratesrequired to release the observed energyfrom a smooth metal sample are severalorders of magnitude higher than the maximumrates that can be obtained in laboratorystudies. Thus it has been concludedby numerous investigators thatfragmentation of the metal to generatelarge surface areas is required toobtain the observed explosion violence.Several review articles have consideredthe potential problem of steam explosionswith reference to the safety ofnuclear reactors (Refs. 12,15). Thearticles by Brauer, et al. (Ref. 16),Ergen (Ref. 17), Witte, et al. (Ref. 12),and Flory, et al. (Ref. 18) describe thevarious factors thought to be involvedin steam explosions, while Pearlman(Ref. 19) reviewed chemical reactionswhich occur between water and certainmetals at elevated temperature. Epsteinsummarized many ~of the early developmentsregarding both physical and chemical-metal-water explosions (Ref. 15).A variety of materials have beenexamined including molten aluminum inwater (Refs. 5,6,20) , copper mat inwater (Refs. 21,22), iron in water(Refs. 23,24), smelt in water (Refs.7,8), and, more recently, molten U0 2 inwater (Refs. 25,26) and U0 2 in sodium(Refs. 26,27,28,29). An intensive investigationinto the fundamental causesof physical explosions is being pursuedat the University of Houston (Refs5.11,30-34).BL.2REPRESENTATIVE INCIDENTSINVOLVING STEAM EXPLOSIONSExplosive incidents periodically occurringin the paper and metal industrieshave been reported. Several such incidentshave been summarized (Ref. 11) andare cited to demonstrate the magnitudeof the destructive forces present andthe physical circumstances leading tothe incidents.B1.2.1 METAL INDUSTRYExplosivethe metaloccur, theaccidents areindustry butdestruction isinfrequent inwhen they dosevere.B1.2.1.1 Mallory-Sharon Incident(Ref. 35).In 1954, a titanium arc-melting furnace,which was water-cooled, exploded at aplant in Ohio. Nine injuries includedfour fatalities and property damage was$30,000. The explosion was believed toresult from water entering the meltingcrucible.B1.2.1.2Reynolds Aluminum Incident(Ref. 35).In 1958, an aluminum-water explosionoccurred in Illinois involving some 46injuries, 6 fatalities and approximately$1,000,000 in property damage. The explosion"rocked a 25 mile" area. Wetscrap metal was being loaded into a furnacewhen the explosion was triggered.VIII-73

B1.2.1.3Quebec Foundry Incident(Ref. 4).The accident occurred in a foundrybuilding approximately 18 million cubicfootvolume. One hundred pounds of moltensteel fell into a shallow troughcontaining about 78 gallons of water.The resulting explosion injured millpersonnel (one fatally) and caused$150,000 damage to the foundry buildingincluding cracking a 20-inch concretefloor, breaking 6000 panes of glass, andstructural damage to the walls and ceilings.Damage was also incurred byanother structure separated some 75yards from the foundry building. Thisaccident is one of the better documentedincidents.B1.2.1.4 Western Foundries Incident(Ref. 36).In 1966, while 3000 pounds of moltensteel was being poured from an electricfurnace into a tile-lined ladle, a cablebroke and the hot steel dropped into awater filled pit. The result was aviolent explosion that injured threeworkers and tore a 600-square-foot holein the roof of a building of some12,000-square-foot floor area. The explosionwas heard some 3 miles from thefoundry.B1.2.1.5Armco Steel Incident(Refs. 37,38).In 1967, an explosion occurred whenmolten steel fell on "damp" ground. Aladle containing some 30 tons of moltensteel had been elevated some 40 feetwhen the ladle fell. Injuries sustainedby some 30 workers included 6 fatalities.Evidently, sufficient moisturewas present in the porous ground totrigger small-scale explosions thatshowered molten steel over a wide area.Although the injuries were attributedprimarily to burns, an explosion accompaniedthe incident.B1.2.1.6East German Slag Incident(Ref. 2).Appearing in 1959, an East German articlediscusses a number of slag-water explosionsthat have occurred in Germanopen-hearth steel mills. Two accidentswere discussed in which explosionsresulted from spraying water on moltenslag in open slag pits. One of theexplosions resulted in a fatality and anumber of other injuries. Severestructural damage was also noted. Thesecond explosion was less severe. Bothexplosions were attributed to excesswater on the slag passing down into thecracks to the hot molten material below.A third instance resulting in an explosionoccurred when a slag pot was placedon a slag bed that had been previouslysprayed with water. An explosion occurred,killing one man. The explosionwas attributed to the heavy slag potcausing cracks in the surface of the hotslag bed and excess water on the surfacegetting into these cracks. Other explosionsbriefly described include rainwaterleaking through an unsealed roofover a slag bed, resulting in an explosion,and two instances of explosionsresulting when molten slag was pouredinto dump cars that had small amounts ofwater in the bottom.B1.2.1.7 British Slag Incident(Ref. 3).In 1964, an explosion occurred in aBritish steel mill when a ladle beingused to tap a blast furnace was sprayedwith lime water and returned to service.When it was next used, the ladle explodedwhen it was about three-fourths fullof slag (12 to 14 tons). Damage to thestructure and injuries to personnel werereported.B1.2.2PAPER INDUSTRYThe paper industry experiences explosionssimilar to those in the metalindustry more frequently but they areless destructive. Explosions occur whenpaper smelt (mostly fused sodium carbonatewith a few percent of sodium sulfide,'sodium chloride, and minor ingredients)is quenched in large containersof "green liquor" (Refs. 7,8). Also,explosions frequently occur when boilertubes in waste-heat boilers fueled by"black liquor" fail, and water is injectedinto hot molten smelt and blackliquor. These explosions occur withconsiderable destruction to the furnaceand plant facilities.B1.2.3NUCLEAR REACTOR INDUSTRYExplosive vapor formation when hot,molten core materials have come in contactwith water have also been observedin nuclear reactors.B1.2.3.1 Canadian NRX Reactor (Ref. 35).In 1952, at Chalk River, Ontario, duringa low-power experiment, a nuclear excursionwas experienced. Although theduration of the incident was less than62 seconds, the damage was sufficient toresult in contamination of the facility.The reaction between uranium and steam(or water) was the principal cause ofdamage.A:IVIII-74

B1.2.3.2 Borax I Reactor (Ref. 35).In 1954, at the National Reactor Testingstation in Idaho, the Borax I reactorwas deliberately subjected to a potentiallydamaging power excursion in reactorsafety studies. A power excursionlasting approximately 30 millisecondsproduced a peak power of 19,000 megawattswith a total energy release of 135megawatt-seconds. The power excursionmelted most of the fuel elements. Thereactor tank (1/2-inch steel) was rupturedby the pressure (probably inexcess of 10,000 psi) resulting from thereaction between the molten metal andthe water. The sound of the explosionat the control station (1/2 mile away)was comparable to that from 1-2 poundsof 40 percent dynamite.B1.2.3.3 SPERT 1-D Reactor.During the final test of the destructivetest program with the SPERT 1-D core,damaging pressure generation was observed.Pressure transducers recordedthe generation of a pressure pulse largerthan 3000 psi which caused the destructionof the core. The pressurepulse occurred some 15 millisecondsafter initiation of the power excursion.The power excursion rapidly overheatedthe fuel plates; the increased temperaturemelted the metal and the claddingof the fuel plates. After the transient,much of the fuel that had beenmolten was found dispersed in thecoolant.BI.2.3.4 SL-I Reactor.In January, 1961, a nuclear excursionoccurred in the SL-I reactor in Idaho.The total energy released in the excursionwas approximately 130 Mw-sec (Ref.51). Of this, 50 Mw-sec was produced inthe outer fuel elements in the core.This portion of the energy was slowlytransferred to the water coolant over 2sec period, and no melting (uraniumaluminumalloy) of the outer fuel elementsoccurred. About 50-60 Mw-sec ofthe total energy release was promptlyreleased by 12 heavily damaged innerfuel elements to the water coolant inless than 30 msec. This prompt energyrelease resulted in rapid steam formationin the core which accelerated thewater above the core and produced awater hammer that hit the pressurevessel lid. The vessel, weighing about30,000 lbs with its internals, shearedits connecting piping and was liftedapproximately 9 feet into the air by themomentum transferred from the waterhammer. Calculations of the mechanicaldeformation of the vessel indicate thatabout 12 percent of the prompt energyrelease or 4.7 percent of the totalnuclear release was converted into mechanicalenergy (Ref. 52).In each instance, under differing circumstances,a hot molten material fell,dropped, or spewed into a mass of coolerliquid and destructive pressure generationresulted. The complex mechanismstriggering this type of reaction are notcompletely understood.It may be noted that all of the abovereactor tests and incidents involvedplate-type fuel elements consisting ofuranium-aluminum alloy fuel clad inaluminum. These are substantiallydifferent from the uranium oxide fuel,Zircaloy-clad rods used in power reactors.B1.3 METAL FRAGMENTATION AND THEINITIATION OF PHYSICALEXPLOSIONSThe nature of the initiating phenomenafor physical explosions is not clearlyunderstood. However, fragmentation orparticle dispersion is thought to be arequirement for such events. As indicatedby Witte, Cox, and Bouvier (Refs.11,12), there appear to be several possiblemechanisms for initiating physicalexplosions (Table VIII B-1). Allinitiation mechanisms result in thespontaneous subdivision of the moltenmetal in intimate contact with water.After exhaustive experimentation withaluminum-water explosions, Long postulatedthat trapping or encapsulation ofcold water at a solid surface undermolten metal can lead to its rapid subdivision(Refs. 27,28). By his theory,the trapped water is rapidly vaporizedand overcomes the surface tension forcesof the metal with the result that themetal becomes finely divided.Sallack (Ref. 7) and Nelson and Kennedy(Ref. 8) have proposed a surface crackingmechanism to explain explosions thatoccur within the bulk of the liquid. Ithas also been applied by Tetzner to explosionsoccurring at a solid interface(Ref. 2). The hypothesis is that as anunshattered ball of molten material (forexample, salt or slag) begins to freezein an environment of cold water, surfacecracking occurs due to thermal stress,some water enters the bulk of the moltenmaterial and is trapped within fissuresand pores, and the ball of semimoltenmaterial shatters as the water is convertedto steam in the interior of thematerial.VIII-75

Another mechanism, termed violent boiling,is predicated upon the fact thatwhen a mass of molten metal isintroduced into a liquid, a quenchingprocess occurs (Refs. 39,40,53). In thetransition region between film and nucleate,the boiling process can becomevery violent -hydrodynamically with theresult that large mixing forces aregenerated. Under such conditions thesteam-metal interface is ruptured withthe rocket-like acceleration of volumesof liquid water and molten material intothe steam interface between the twophases. This dispersion can be comparedto the Leidenfrost phenomena when watercontacts a hot surface or to water dropletsin hot oil. Swift and Baker (Ref.53) suggest that fragmentation may onlyoccur when the materials pass throughthe range of temperature from the boilingpoint to critical point of the fluid(in which violent boiling can occur)prior to freezing of the melt. This restrictionwould not permit fragmentationin U0 2 -H 2 0 systems.The shell or encapsulation theory hasbeen proposed by researchers at theUniversity of Kansas (Refs. 16,18).These researchers performed a series ofexperiments using high-speed photographyto study the fragmentation of moltenmetals in water. Experiments were performedby dropping a number of hotmetals (Pb, Sn, Bi, Zn, Cu, Al, Hg andWood's metal) into water and photographicallyrecording the ensuing reaction.Brauer suggests that molten globules areblown apart by an internally generatedpressure (Ref. 16). This theory hypothesizesthat somehow liquid is trappedinside the molten-metal globules, theliquid rapidly vaporizes, and the resultingpressure increase fragments themetal. Brauer's hypothesis coincides toa great extent with the surface crackinghypothesis.Molten-metal fragmentation may also resultwhen inertial forces acting on themolten metal (as it is passing throughthe liquid) exceed the surface tensionforces of the metal (Ref. 41). Such aninitiation mechanism has been proposedby Witte and co-workers (Ref. 11). Theratio of the inertial forces to thesurface-tension forces of a molten-metalglobule is expressed as the Webernumber, Nw = prU 2 /y. If the Webernumber exceeds a critical value, about5-10 for molten-metal drop tests withHg, Pb, Bi, and Sn, the inertia of theparticle overcomes the surface-tensionenergy and the globule fragments intosmaller, more stable sizes.Forced initiation refers to any eventbrought about by an external drivingforce which leads to the rapid dispersalof molten metal in water. Occurrencesof this type have been observed in foundrypractice. Also, Higgins broughtabout physical explosions by usingblasting caps and spray injection techniquesto disperse various molten metals(Zr, Zircaloy-2, Al, U, NaK, and stainlesssteel) in water (Refs. 42,43). Thespray-injection experiments (performedin an explosion dynamometer) permittedmeasurement of explosive power, reactiontime, reaction efficiency and the effectsof temperature and droplet size(Ref. 43). Peak overpressures were ofthe order of 20 to 200 psig and pressure-riserates were approximately 5000to 120,000 psig/sec. Explosion periodsor times required for the pressure risewere of the order of milliseconds.The superheat theory used by Katz andothers to explain the occurrence ofvapor explosions in the LNG industry(Refs. 44,45,46) it has been extended byFauske (Ref. 29) to explain the mechanismOf U0 2 -Na vapor explosions. Contactbetween LNG and water at first resultsin the establishment of film boiling ofthe LNG. The most volatile component ofthe LNG, methane, is preferentiallydepleted in time, and an increase in theboiling point of the LNG results. Thestable vapor blanket becomes thinner, iseventually eliminated, and liquid-liquidwetting contact between the LNG and thewater is substituted. Unlike that insolid-liquid contact, unstable nucleateboiling cannot commence because nucleationsites are lacking. A thin layer ofLNG along this contact surface explodesbecause the instantaneous contact temperatureis above the limit of superheattemperature of the enriched LNG, forwhich homogenous nucleation occurs.This exploding layer presumably fragmentsand disperses some or all of theremaining LNG into the water, therebytriggering the main explosion.For U0 2 and Na, on the other hand,Fauske points out that the developmentof film boiling along the liquid-liquidinterface would require a U0 2 temperaturemuch higher than the melting pointOf U0 2 , hence wetting contact occursimmediately. However, the instantaneouswetting contact temperature is well belowthe superheat limit for Na, and thetype of triggering event described abovefor LNG does not occur. Instead, anyglobule of liquid Na that might becomeentrained in the U0 2 would heat uprelatively slowly, depending upon itsVIII-76

size, until a substantial portion or allof the globule became superheated, inthe absence of induced nucleation sites.When the interface temperature finallydoes reach the limit of superheat, theentire superheated portion of the globulewould explode violently. There isno need for a prior fragmentation anddispersion mechanism, although whateverfragmentation and dispersion this explosioncould produce in the rest of themixture would contribute to theintensity of the explosion. In theactual reactor situation, the probabilityof the occurrence of such an explosionwould depend upon the density ofnucleation sites such as fission fragmentsand gas bubbles that could initiateprior boiling of the liquid Na.The collapsing-vapor-bubble theory recentlyadvanced by researchers at theUniversity of Houston is based on studyof the way in which the vapor film surroundinga heated silver sphere wasdestabilized, i.e., the sequence ofevents that causes the boiling behaviorto become transitional between the filmand nucleate regimes. Two distinctivemechanisms for destabilization wereidentified: one, a precipitous collapseand the other, progressive instability.The precipitous collapse occurs veryrapidly, on the order of 0.250 millisecond,while the progressive instabilityrequires from 50 to 100 milliseconds forcompletion. The progressive instabilitywas triggered by small bubble-like irregularitiesthat move on the liquidvaporinterface and it occurs only insaturated (or nearly-saturated) water.Precipitous collapse occurs in subcooledwater. There appeared to be a directrelationship between the two types ofvapor film collapse and observed fragmentationbehavior for molten materialsbeing rapidly quenched. Extensive fragmentationgenerally occurs only in subcooledsystems; here, the vapor filmcollapse is very rapid and energetic.Little fragmentation occurs in saturatedsystems; here, the collapse is relativelyslow and nonenergetic.The superheat-limit theory has been combinedwith the collapsing-vapor-bubbletheory to formulate a jet-action theoryfor triggering the usual type of vaporexplosion - one that requires rapidfragmentation and dispersion of oneliquid into the other (Ref. 47). In thejet-action theory, the exploding thinfilm degenerates into a multiplicity ofvapor bubbles or globules as it expands.The expansion is greater toward the lessdense liquid. When these globulescollapse, the rectilinear motion towardthe less dense liquid accelerates, as isnecessary to conserve linear momentum.Shape instabilities, developing as theglobules collapse, eventually result inthe formation of jets of the more denseliquid that project through the globulesinto the less dense liquid (Ref. 48).Experimental evidence has been presentedon the role played by collapsingcavities and the accompanying formationof microjets in initiating detonationsin slow-burning liquid films (Ref. 49).Violent high-frequency oscillations ofthe vapor film have been observed undersome conditions following pulsed-laserbeamtransient heating of a metal filmimmersed in water (Ref. 50). The resultantperiodic liquid-liquid contact resultsin a greatly increased heat fluxover that ordinarily obtained in thefilm-boiling regime. The authors postulatethat the pressure pulses developedduring the collapse phases of the oscillationsmay be important in the dispersalof the hot molten material in vaporexplosions. This dispersion could alsobe augmented by the formation of jets,as previously described, during the collapsephases of the oscillations.B1.4 INTERACTION OF U0 2 WITH WATERVirtually no data could be found whichwere directly applicable to the lightwater-reactormeltdown situations consideredin this program, i.e., largeamounts of molten U0 2 falling intosaturated or near-saturated water (93 Cfor the cases considered). However, twoexperimental studies directed at investigatingthe reaction of molten U0 2 withwater are applicable to some extent(Refs. 25,26).At Battelle's Northwest Laboratory, thereaction of molten U0 2 with water wasinvestigated in a high-pressure furnaceby allowing melted U0 2 samples (1 to 10grams) to fall into water heated to varioustemperatures (see Table VIII B-2).The tested U0 2 samples were analyzed forthe degree of oxidation by coulimetrictitration. Oxygen-to-metal ratios increasedfrom an average initial value of2.002 to values between 2.06 to 2.16,depending on the initial water temperature(30 to 250 C). Water temperatureswere arbitrarily selected to determinethe effects of water temperature on thefinal results and encompass the temperaturerange expected in the case of amajor accident. The oxygen-to-metalratio values correspond to a maximumhydrogen generation of 12 ml (STP)/gU0 2 . The oxidation product, as deter-VIII-77

mined from Guinier X-ray powder patterns,*was U 4 0 9 .Attempts to measure pressure transients,produced by evaporation and/or boilingof water in contact with the-molten U0 2 ,were generally unsuccessful. Such measurements,if successful, would havebeen directly applicable to possiblephysical explosions Of U0 2 in water, Ofsignificance, however, was the fact thatrapid injection of molten U0 2 into waterdid not produce a powdered end product.Rather, the quenched U0 2 was very friableand could be readily reduced to acoarse powder. Tyler screen analyses onthree different powdered samples gaveessentially identical average particlediameters of 620 microns (0.062 cm) (seeFig. VIII B-1).In two experiments, IAmblard, et al.dropped approximately 1 gram each ofmolten U0 2 at 3000 and 3100 C, respectively,into 10 grams each of water at20 C and atmospheric pressure (Ref. 26).In both experiments, long hollow tubesOf U0 2 that formed, imprisoned thecoolant. In the first experiment (3000C), the particles were about 10 to 12 mmlong and 3 to 6 mm wide. In the secondexperiment (3100 C), the particlelengths ranged from about 6 to 27 mm andwidth from about 1.5 to 2.5 mm. In athird experiment, a cylindrical sampleOf U0 2 about 5.7 mm in diameter by 55 mmlong (i.e., about 15 grams) clad instainless steel was melted in a pressurevessel capable of withstanding a pressureof 850 bars (12,325 psi). Thestainless steel cladding was cooled byhot water adjacent to the clad underpressure from a cover gas. Upon heating,the clad was pierced by the moltentJ0 2 and the fuel was able to disperse inthe water which was at 230 C (446 F).The clad fusion and U0 2 -water contactwere photographed. Some pulverizationof the U0 2 did occur, but was much lessthan experienced in similar experimentsperformed with U0 2 in liquid sodium.The pressure in the vessel was recordedas a function of time. This trace didnot indicate appreciable dynamic overpressurebut rather a slow pressureincrease of 15 bars (218 psi), i.e.,from 150 bars (2175 psi) absolute to 165bars (2393 psi) absolute in approximately10 seconds. The static pressureincrease of water and cover gas wasthought to correspond to an enthalpyincrease resulting from 50 percent ofthe U0 2 in the sample being molten.B1.5 EXPERIMENTS WITH METALS INWATERNumerous experiments have been conductedinvolving the potential for explosionswhen molten metals come into contactwith water (Refs. 6,16,18).Long conducted some of the initial experimentson explosive vapor formationby means of aluminum (m.p. = 1200 F)drop tests (Ref. 5). More than 880 experimentswere conducted under conditionsthat simulated sudden accidentalpouring of molten aluminum into a poolof water. Fifty pounds of commerciallypure aluminum in the molten state wassuspended in a crucible above a containerpartially filled with water. Whenthe molten metal had obtained the desiredtemperature, a remote-controlledplug was removed from the crucible andthe molten metal flowed through a taphole into the water. Long investigatedabout ten variables and noted that themolten aluminum when at 1382 F fallinginto the water tank did not produce anexplosion if the water temperature wasat least 140 F or above,' 1provided thewater was not highly ionic. Under highlyionic conditions (e.g., addition ofNaCl to produce a 15 percent salt solution)explosions could be obtained attemperatures up to 172 F but not in saturatedwater.Brauer, Green, and Mesler conducted astudy in which aluminum (MP = 1220 F)and lead (MP = 621 F) were heated totemperatures well above the meltingpoint and dropped into liquid water withthe objective of obtaining fragmentationof the metal (Ref. 16). A high-speedmotion picture camera recorded macroscopicobservations. In general, it wasnoted that fragmentation did not occurfor metal temperatures just slightlyabove the metal's melting point (i.e.,the higher the metal temperature aboveits melting point the greater the fragmentation)or for cooling-water temperaturesgreater than 60 C (140 F). For agiven metal, fragmentation if itoccurred, was more violent for the lowerwater temperatures.Flory, Paoli, and Mesler (Ref. 18) carriedout numerous experiments at KansasUniversity. Heating was accomplished byan induction coil and the hot moltenmetal was dropped into a Plexiglas tankcontaining water. Pressure disturbanceswere monitored and the interactions werephotographed with a high-speed camera.Tests included a wide range of metal and1Actually, Long's data indicate that heobserved explosions up to 122 F (seeTable 7 of Reference 5).VIII-78

ath temperatures. Metals studied includedlead (MP = 621 F), tin (MP = 449F), bismuth (MP = 520 F), zinc(MP = 787.1 F), copper (MP = 1981 F),aluminum (MP = 1220 F), mercury (MP =37F), Wood's metal, and Cerro Bend. Numerousexperimental observations weremade concerning fragmentation mechanisms.The temperature of the quenchwater was varied from -u8 C to 100 C(17.6 F to 212 F). "Near the freezingpoint of water the violence of fragmentationis not a function of bath temperature,while above 25 C (77 F) theviolence and extent of fragmentationdecreased rapidly to zero at 90 to 100 C(194 to 212 F). Tests with severalmetals dropped into liquid nitrogen atits boiling point gave no fragmentationwhatever, giving supporting evidence tothe results of the water experiments,namely that metals will not fragment ina saturated liquid."B2. CONCLUSIONSThe single most-pertinent question thatmust be resolved is whether molten U0 2can be fragmented to give the high surfaceareas required to initiate a physicalexplosion when contacted with water.Although pertinent, the above describedexperiments with U0 2 and water are by nomeans definitive, *and, in fact, no completelydefinitive experiments could befound in the literature.it is questionable whether the experimentalobservations with regard to droppingmodest amounts of "relatively lowtemperature" molten metals into watershould be extrapolated to the case oflarge amounts of molten U0 2 . The meltingpoint of the U0 2 is SO much higherthan that of the other materials investigatedthat entirely different phenomenacould govern the process. The quantityof molten material that could interactwith water in a meltdown accidentis orders of magnitude greater than hasbeen investigated in any controlledexperiments7 such an extrapolation insize would be fraught with uncertaintieseven if the phenomenology were understood.The experiments with the variousmolten metals indicate that the probabilityof obtaining a high degree offragmentation decreases as the temperatureof the water approaches saturation.Taking into account the history of steamexplosion in industry observed under awide variety of conditions, the modestamount of work available with U0 -~waterinteractions, and the considerable experimentalwork on various metals inwater, it is concluded that the probabilityof initiating U0 2 -water explosionsthat require a high degree offragmentation is very low. Such wouldbe the most violent types of explosionswhich require effective average particlediameters of less than about 2000 microns,as described in Appendix C. Forthe less violent explosions, also capableof causing containment failure, theaverage particle diameter can be betweenabout 2000 to 5000 microns (2 to 5 mm)and the probability must be consideredsomewhat greater.The overall probability of occurrence ofan event in a PWR or BWR reactor vesselthat would produce an effective surfacearea equivalent to 2- to 5-mm diameterparticles of U0 2 and disperse themthrough a significant part of the volumeof water is considered to be low.VIII-79

References1. Anonymous, "A Method for Handling Steelmaking Slag", Iron & Steel Eng., 28 (10),85-89 (October, 1951).2. Tetzner, Herman, "Slag Explosions in Open-Hearth Steelworks", Neue Hutte, 6 (4),352-359 (June, 1959).3. Anonymous, "Water was Energy Source in Ladle Explosion", Steel Times, pp. 56-57(January, 1964).4. Lipsett, S. B., "Explosions from Molten Material and Water", Fire Technology, 2(2), 118-126 (May, 1966).5. Long, G., Stroup, P. T., and Ennor, W. T., "Explosions of Aluminum and Water",Report No. 2-50-33 (NP-5471), Process Metallurgy Division, Aluminum ResearchLaboratories, Aluminum Company of America (1950); Long, G., "Explosions of Aluminumand Water - Cause and Prevention." Metal Progr., pp. 107-112 (May, 1957).6. Rengstorff, G. W. P., Lemmon, A. W., Jr., and Hoffman, A. 0., "Review ofKnowledge on Explosions Between Molten Aluminum and Water to the AluminumAssociation", Battelle's Columbus Laboratories, p. 53 (April, 1969).7. Sallack, J. A., "An Investigation of Explosions in the Soda Smelt DissolvingOperation", Pulp and Paper Mag. (Canada), 56 (9), 114-118 (1955).8. Nelson, W., and Kennedy, E. H., "What Causes Kraft Dissolving TankExplosions - I. Laboratory Experiments", Paper Trade J., 140 (7), 50-56 (July,1956).9. Ivins, R. 0., and Baker, L., "Chemical Engineering Division Semiannual Report,January-June, 1964", ANL-6900, pp. 270-280 (August, 1964).10. Mally, C. T., et al., "Kinetic Studies of Heterogenous Water Reactors. AnnualSummary Report, 1964", STL-372-12 (December 30, 1964).11. Bouvier, J. E., Witte, L. C., and Cox, J. E., "The Vapor Explosion - HeatTransfer and Fragmentation. I. Introduction to Explosive Vapor Formation",Technical Report No. ORO-3936-I, Department of Mechanical Engineering,University of Houston, 21 pp. (November, 1969).12. Witte, L. C., Cox, J. E., and Bouvier, J. E., "The Vapor Explosion", Journal ofMetals, 22 (2) 39-44 (February, 1970).13. Epstein, L. F., "Analytical Formulation for the Reaction Rate", Metal WaterReactions, 6 37-39 (September, 1959).14. Higgins, H. M., "The Reaction of Molten Uranium and Zirconium Alloys WithWater", Aerojet General Report No. AGC-AE-7, p. 18 (April, 1965).15. Epstein, L. F., "Recent Development in the Study of Metal-Water Reactions",Progr. Nucl. Energy, Ser. IV, 4, 461-483 (1961).16. Brauer, S. D., Green, N. W., and Masler, R. B., "Metal/Water Explosions", Nucl.Sci. Eng., 31 (3), 551-554 (1968).17. Ergen, W. K., et al., "Emergency Core Cooling - Report of Advisory Task Force onPower Reactor Emergency Cooling", TID-24226 (1966).18. Flory, K., Paoli, R., and Mesler, R., "Molten Metal-Water Explosions", ChemicalEngineering Progress, 65 (12), 50-54 (December, 1969).19. Pearlman, H., "Chemical Reactions Between Water and Certain Metals at ElevatedTemperatures", U.S. Atomic Energy Commission, NAA-SR-Memo-858, p. 22 (1954).VIII-80

20. Hess, P. D., and Brondyke, K. J., "Causes of Molten Aluminum-Water Explosionsand Their Prevention", Metal Prog., (4), pp. 93-199 (April, 1969). (Presentedat the Metals Congress, Detroit, October, 1968.)21. Azuma, Kiyoshi, Goto, Sakichi, Kametani, Hiroshi, "The Explosion of Mat byWater. I. Nature of the Explosion", Nippon Kogyo Kaishi, 73, 295-300 (1957).22. Azuma, Kiyoshi, Goto, Sakichi, Kametani, Hiroshi, "The Explosion of Mat byWater. II. Chemical Reaction Between the Molten Copper Mat and Steam", NipponKogyo Kaishi 73, 449-452 (1957).23. Murty, G. V. L. N., Subba Rao, D., and Ramaswami, T. R., "Highly Water-ReactiveForms of Transitional Elements of the Iron Groups", Current Sci. (India) 25,197-188 (1956).24. Gasik, M. I., and Khitrik, S. I., "Interaction of Liquid Iron with a Steam-Hydrogen Mixture", Met. i Koksokhim., Mezhvedonstv. Resp. Nauchn.-Tekhn. Sb. 3,5-16 (1965).25. Gilby, R. L., "Reaction of Molten UO2 With Water", BNWL-362 (1967).26. Amblard, M., Semeria, R., Vernier, P., Gouzy, A., and Najuc, M., "Contact EffectBetween Molten UO 2 and Sodium and Molten UO 2 and Water", EURFNR-811 (March 3,1970).27. Anderson, R. P., and Armstrong, D. R., "Laboratory Tests of Molten-Fuel-CoolantInteractions", Trans. Amer. Nucl. Soc., 15 (1), 313 (June, 1972).28. Armstrong, D. R., Testa, F. J., and Raridon, D., Jr., "Interaction of SodiumWith Molten UO 2 and Stainless Steel Using a Dropping Mode of Contact", ANL-7890(December, 1971).29. Fauske, Hans K., "On the Mechanism of UO 2 -Na Vapor Explosions", Nuclear ScienceEngineering, 51, 95-101 (1973).30. Stevens, J. W., Bullock, R. L., Witte, L. C., and Cox, J. E., "Transition BoilingFrom Spheres to Water", AEC Report ORO-3936-3 (April, 1970).31. Hsico, K. H., Cox, J. E., Hedgecox, P. G., and Witte, L. C., "Pressurization ofa Solidifying Sphere", AEC Report ORO 3936-4 (April, 1970).32. Witte, L. C., et alo, "Rapid Quenching of Molten Metal", AEC Report ORO-3936-6(August, 1971).33. Bradley, R. H., et al., "Investigation of the Vapor Explosion Phenomenon Using aMolten-Metal Jet Injected into Distilled Water", AEC Report ORO-3936-7 (October,1971).34. Witte, L. C., et al., "Transition Film and Transition Boiling from a Sphere",AEC Report ORO-3936-9 (September, 1972).35. Epstein, L. F., "Reactor Safety Aspects of Metal-Water Reactions", Metal-WaterReactions VII, GEAP 3335, p. 2 (January 31, 1960).36. Thomas, P., "Three Hurt in Foundry Blast", The Denver Post (February 14, 1966).37. Anonymous, "Hot Metal Spilled at Plant", Houston Chronicle, 67 (66), Section 1,p. 1 (December 18, 1967).38. Anonymous, "6th Worker Dies in Armco Tragedy", The Houston Post, No. 30, 211,Section 1, p. 10 (December 21, 1967).39. Vogel, et al., "Reactor Development Program Progress Report for February 1969",ANL-7553 p. 118 (March 25, 1969); "Reactor Development Program Progress Report",ANL-7478, p. 128 (July, 1968).VIII-81

40. Firstenberg, A. F., et al., "Kinetic Studies of Heterogenous Water Reactors.AEC Research and Development Report Annual Summary Report,1966", STL-372-50(December, 1966).41. Duffield, R. B., and Lawroski, S., Argonne National Laboratory, Reactor DevelopmentProgress Report, ANL-7399, pp. 162-167 (November, 1968).42. Higgins, H. M., "A Study of the Reactions of Metals and Water - Interim Report",USAEC Report AECD-3664 (April, 1955).43. Higgins, H. M., and Schultz, R. D., "The Reaction of Metals with Water andOxidizing Gases at High Temperatures", USAEC Report IDO-28000, Aerojet-GeneralCorp. (April 30, 1957).44. Nakanishi, E., and Reid, R. C., "Liquid Natural Gas-Water Reactions", Chem. Eng.Prog., 67 (12), 36 (1971).45. Katz, D. L., and Sliepcevich, C. M., "LNG/Water Explosions: Cause and Effect",Hydrocarbon Processing (November, 1971).46. Enger, T., and Hartman, D. E., "Explosive Boiling of Liquified Gases on Water",paper presented at the National Research Council Conference on LNG Importationand Terminal Safety, Boston, Mass. (June 13, 1972).47. Simon, R., "Computer Model Study of the Smelt-Water Physical Explosion", paperto be presented at the Atlanta Heat Transfer Conference, A.I.Ch.E. (August,1973),48. Chincholle, L., "Bubbles and the Rocket Effect", J. Appl. Phys., 41, 4532(1970).49. Bowden, F. P., and Chaudri, M. M., "Initiation of Explosion of AgN 3 and aPbN 6Single Crystals by a Collapsing Bubble", Nature, 220, 690 (1968).50. Board, S. J., Clare, A. J., Duffey, R. B., Hall, R. S., and Poole, D. H., "AnExperimental Study of Energy Transfer Processes Relevant to Thermal Explosions",Int. J. Heat Mass Transfer, 14, 1631 (1971).51. USAEC Idaho Operations Office, "IDO Report on the Nuclear Incident at the SL-IReactor, January 3, 1961, at the Nation Reactor Testing Station", IDO-19302(1962).52. Proctor, J. F., "Adequacy of Explosion-Response Data in Estimating Reactor-Vessel Damage", Nuclear Safety, 8 (6), 565 (November, 1967).53. D. L. Swift and L. Baker, "Experimental Studies of the High-Temperature Interactionof Fuel and Cladding Materials with Liquid Sodium", ANL-7120, pp 839-847(1965).VIII-82

TABLE VIII B-1METHODS LEADING TO METAL FRAGMENTATIONPHYSICAL EXPLOSIONSAND THE INITIATION OFInitiation MethodTrapping WaterFissures, Surface CrackingViolent BoilingShell TheoryInertial Forces(Weber Number Effects)Reference27, 282, 7, 839, 40, 5316, 1811, 41Forced InitiationSuperheat TheoryCollapsing Vapor BubbleJet FormationOscillating Vapor Blanket42,29,3447504344, 45, 46TABLE VIII B-2RESULTS OF MOLTEN U02 BEINGTEMPERATURES (Ref. 25)DROPPED INTO WATER AT VARIOUSMelted UO2Weight,gWaterTemperature,CT,cRatioOxygen/UraniumBeforeAfterCalculatedH2 Generation,ml (STP)/g U020.7562752.002 2.0584.59.3728732.005 2.0806.20.57110052.002 2.0776.21.126105152.004 2.0826.57.705110612.012 (a) 2.1289.69.25100802.002 2.0877.18.49210702. 012 (a) 2.15511.911.381255492.029 (a) 2.1469.79.13260402.006 2.14211.3(a)Oxidation resulting from leakage of water into furnace chamber from water vessel.Table VIII B-I - Table VIII B-2VIII-83/84

50001oo ]Z1E.2100100 20 40 6080 100Weight Percent FinerFIGURE Vill B-1 Particle Size Distribution of Quenched UO Samples2Fig. VIII B-IVIII-85/86

APPENDIX CMODEL STUDY OF PHYSICAL EXPLOSIONSVIII-87

Appendix CTable of ContentsSectionPage No.Cl. MODEL GEOMETRY ................................................... VIII-91C2. EXPLODING MIXTURE ................................................ VIII-92C3. BASIC MODEL RELATIONSHIP ......................................... VIII-93C3.1 Mathematical Equations ..................................... VIII-93C3.2 Heat-Transfer Rate ......................................... VIII-93C3.3 Additional Assumptions ..................................... VIII-94C3.4 Chemical Reactions ......................................... VIII-95C3.5 Vessel - Wall Stresses ..................................... VIII-95C4. RESULTS OF COMPUTATIONS .......................................... VIII-96C5. MATHEMATICS OF THE MODEL ......................................... VIII-98C5.1 Basic Model Equations ...................................... VIII-98C5.1.1 Conservation of Energy ............................. VIII-98C5.1.2 Conservation of Momentum ........................... VIII-99C5.1.3 Conservation of Mass ............................... VIII-100C5.1.4 Equations of State .................................VIII-100C5.2 Mechanical Response Time ................................... VIII-100C5.3 Dynamic Stresses in Vessels Walls .......................... VIII-101C6. HEAT-TRANSFER-RATE CONSIDERATIONS ................................ VIII-101C6.1 Heat-Transfer Coefficient .................................. VIII-101C6.1.1 Radiactive Heat Transfer ........................... VIII-102C6.1.2 Conductive Heat Transfer ........................... VIII-102C6.2 Convective Heat Transfer ................................... . VIII-102C6.3 Temperature Gradient in Melt Particles ......................... VIII-103REFERENCES ..............................................................VIII-105VIII-89

List of TablesTableVIII C-1Values of Parameters Used For Massive Grid-Plate Failure,Steam-Explosion Computations ...................................Page No.VIII-107/108List of FiguresFigureVill C-IVIII C-2VIII C-3VIII C-4VIII C-5Basic Computational Model Geometry ..............................Typical Computer Results ........................................PWR Results .....................................................BWR Results .....................................................Dimensionless Presentation of Calculational Results .............Page No.VII -109/110VIII-109/l10VIII-109/110VIII-109/110VIII-109/ll0VIII-90

Appendix CModel Study of Physical ExplosionsThis appendix describes the analyticalmodel used to provide information regardingboth the probability and damagepotential of rapid energy exchanges betweenlarge quantities of molten materialsand water. These exchanges, termed"physical" or "steam explosions" in thisdescription, are pertinent to the overallstudy as they could occur followinga significant core meltdown during apostulated accident. For the purposesof this discussion the term "explosion"is broadly defined as any heat-transferevent during which sufficient mechanicalenergy could be developed in forms thatcould result in appreciable damage ordeformation to surrounding structures.The phenomenological model, though necessarilysimplified, includes treatmentof the basic physical phenomena. Usingthis model in parametric and sensitivitystudies permitted the development ofregion mappings which provided guidanceas to the probability and damage potentialof the various events in additionto the identification of the more significantvariables. The model proved tobe adequate for the purposes intendedand a more sophisticated treatment wasnot considered warranted. It should bereiterated that for the purposes of thisevaluation the meltdown event was assumedto occur; only the potentialensuing events are addressed here.It should be recognized that the simplemodel that has been used in this studyprobably yields an overestimate of theconversion of thermal energy to mechanicalenergy than would occur in anexplosive event. The results of themodel indicate conversion efficienciesof 10 percent or greater. Conversionefficiencies observed in experiments oraccidents are typically lower.C1. MODEL GEOMETRYFigure VIII C-1 showsthe primary reactorthe geometry ofvessel assumed inThis appendix presents the variousaspects of the model along with theresults, detailed mathematical and heattransfer aspects, and pertinent references.the model after all or a portion of themolten core has dropped into the waterbelow it, at some time after thebeginning of a meltdown accident. Thisgeometry consists basically of a steelcylinder, closed on both ends with flatsteel plates, and with the mixture initiallyoccupying the lower portion ofthe cylinder and expanding upwards asthe explosion proceeds.The hemispherical top head of the actualvessel is replaced in the model with acylindrical end cap of equal volume.The bottom plate of the model representsthe top of the grid plate for thecomputations involving dropping of themolten portion of the core of the PWRinto the water remaining above the gridplate. For the computations involvingmassive grid-plate failure, in which allof the core and the molten grid platedrop into the water reservoir below thegrid plate, the hemispherical bottomhead is also replaced by a cylindricalend cap of equivalent volume.As shown in Fig. VIII C-1, the modelalso provides for orifices of a giventotal area located at a given distanceabove the bottom plate. The steam andhydrogen above the mixture leak out ofthese orifices, at a rate given by thestandard orifice formula, as long as thetop of the mixture is below the orificelevel. When the mixture rises above theorifice level, it leaks out instead.The actual geometry of the vessel, ofcourse, could be simulated more closelyby the model. The actual shapes of theend caps and changes in effective crosssection as a function of vertical positionin the vessel could be taken intoaccount, including the effects of variousstructures within the vessel. Suchrefinements would add to the complexityof the computations, but would not beexpected to have any major effect on thefinal results for the PWR. However, forthe BWR calculations the effects ofmajor internal structures above thereactor core cannot be neglected. Howthese effects are taken into account isdescribed later.VIII-91

The model provides for the possibleexistence of a reservoir of a givenvolume above the top of a vessel inorder to simulate steam explosions thatmay occur at the bottom of the reactorcavity if the bottom head should meltthrough.C2. EXPLODJING MIXTUREThe "exploding" mixture consists of thevolume of water, steam, and hydrogen gasbelow the melted-down portion of thereactor core plus the molten and solidifiedportion of the core that hasdropped into this volume. The corematerials dropping into the water are,assumed to become uniformly dispersedthroughout the volume.Transfer of heat from the hot core componentsto the colder water-steam-gascomponents of the exploding mixturecauses the mixture to expand, primarilydue to the generation of steam. Expansionis essentially confined to theupward direction by the vessel walls andis impeded initially by the self-inertiaof the exploding mixture and later bythe compression of the steam and hydrogengas above the top of this mixture.The pressure developed in the mixture bythe inertial impedance to free expansionmay become great enough to rupture thesides and bottom of the reactor vessel.It may also propel upward a substantialportion of the exploding mixture withsufficient kinetic energy to rupture thetop head of the vessel, due to a combinationof rapid compression of the gasand vapor above the mixture and impactof the mixture on the top head. It maythereafter impart considerable kineticenergy in the form of upward motion tothe top head and to any shielding orother structures above it. The resultant"1projectile" may reach the dome ofthe containment building and causefailure. The direct effect of pressurepotentially rupturing the vessel wallsas well as the later effect of thedevelopment of the "projectile" phenomenonare considered in the computationalmodel.As discussed in the introduction, theterm "explosion" is broadly defined asany heat-transfer event in which sufficientmechanical energy is developed informs that can result in an appreciablelevel of deformation or damage to surroundingstructures. These forms mayconsist of potential energy (pressure)and/or kinetic energy (directed motionof mass). This broad definition ofexplosion also encompasses the ordinarydefinition of the rapid development of ahigh-pressure pulse that leads to thegeneration and propagation of a shockwave in a medium.The exact procedure for computing thedynamic response of the exploding mixture,i.e., the pressure buildup andmotion of the mixture as a function ofvertical position, is to divide the mixture,or medium, into a number of horizontallayer pairs. Each layer pairconsists of a lower layer that incorporatesthe expansion and compliant propertiesof the medium at the verticallocation of the layer pair and an upperlayer that incorporates its inertialproperties. The resultant equations andtheir stepwise numerical integration intime, beginning with the initial conditions,simulate the solution of the nonlinearacoustic wave equation for themedium, with pressure generating sourcesdistributed throughout the medium.The simplified model shown in Fig. VIIIc-l consists on only one such layerpair. The use of a single layer pairresults in simplification of the computations;the model then essentiallysolves a set of simultaneous ordinarydifferential equations rather than a setof partial differential equations withmixed boundary conditions. This simplificationpermitted the application of apreviously developed computer code forsmelt-water explosions (Ref. 1) to thereactor vessel case, an adaptation whichinvolved conversion from spherical torectilinear geometry and several othermodifications.If the upward motion of the top surfaceof the exploding mixture can be assumedto result from a uniform expansion ofthe mixture, the single layer pair isbest approximated by considering theupper layer to be a "piston" containinghalf the mass of the mixture. Uniformexpansion of a homogeneous medium withthe bottom end fixed in position entailsan acceleration, velocity, displacement,and negative of the pressure gradientthat vary linearly from zero at the bottomto maximum values at the top surface.The pressure itself would varyquadratically from a maximum value atthe bottom to a value corresponding tothe pressure in the gas and vapor abovethe mixture at the top.For acoustic wave propagation effects,the uniform expansion approximation isgood for wavelengths greater than about6 times the depth of the medium. Thetime scales of the pressure eventsobtained from the results of the computations*show this wavelength condition4VIII-92

to be met under most circumstances ofinterest. However, it does break downunder one important circumstance: whenthe gas above the top surface of themedium becomes rapidly compressed as,this surface approaches the top head ofthe vessel, a compression wave alsopropagates downward into the medium.The Taylor-type of instabilities (Ref.2) that can develop in the shape of thissurface when its upward motion israpidly decelerated may also contributeto the inaccuracy of the uniform-expansionassumption at this stage. Anotherfactor that tends to question the uniform-expansionassumption to some extentis the pressure difference that developsbetween the bottom and the top of themedium. This precludes the uniformityof generation and expansion of steamthroughout the medium which is inherentin the assumption of uniform expansion.C3. BASIC MODEL RELATIONSHIPC3.1 MATHEMATICAL EQUATIONSThe mathematical equations of the model,which are presented and discussed indetail in the Mathematics Section, arebased on four fundamental laws.a. Conservation of energy: The rate ofheat input to the colder portion ofthe exploding mixture equals the sumof the rates of change in internalenergy of the liquid water, steam,and gas, the rates of work done intheir expansion, and the latent heatof vaporization of water multipliedby the mass rate of change of liquidwater to steam. At the same timethis must equal the rate of sensibleheat loss from the various componentsof the melt plus the sum ofthe products of heats of fusiontimes mass rate of conversion fromliquid to solid for the various meltcomponents.b. Conservation of momentum: The netforce on the inertial layer equalsthe rate of change of its momentum(which includes effects of changesin its mass).c. Conservation of mass: For each component,the sum of the products ofvolume and density for its twophases equals the original mass ofthe component less the loss of massby flow through the orifices and bychemical reactions.d. Equations of state: The density ofeach phase is a unique function oftemperature and pressure.C3.2 HEAT - TRANSFER RATEThe rate at which heat is transferredfrom the hot melt to the colder componentsof the exploding mixture is ofprimary importance in the model relationships.Heat-transfer considerationsare discussed in detail in a latersection. In brief, the rate of heattransfer is given by the product of aheat-transfer coefficient, the totalsurface area of the melt exposed to thecolder components, and the temperaturedifference between the two.The heat-transfer coefficient was computedby assuming that some fraction ofthe assumed spherical melt particles issurrounded by an essentially infinitebody of steam, while the remainder issurrounded by a blanket of steam, as instable film boiling, beyond which liesan essentially infinite body of liquidwater. This fraction is equal to thevolume fraction of the water that is inthe form of steam. For the former, heattransfer is essentially by conductioninto the steam. For the latter, heattransfer is essentially that required toreplenish the steam film against loss bynatural convection, as given byBronley's equation (Ref. 3). Forcedconvectionheat transfer due to turbulentrelative velocities between themelt particles and the water-steam-gasis neglected, since these velocities arenot known. The average value of therelative velocity is zero, in accordancewith one of the special assumptionsdiscussed later. The radiative heattransfercoefficient, which is significantat the tempeature of the melt inquestion (initial value, 2850 C - justabove the melting point of U0 2 ), isadded to the conductive and/or convectivecoefficients.The surface area in the heat-transferrateequation is computed on the basisof the assumed average particle diameterfor the dispersed melt. The surfacetemperature of the melt was initiallytaken to be the same as the averagetemperature of the melt particle (isothermalmelt assumption). This is agood assumption for melt particles of100 microns and smaller, but not for thelarger particles. An approximate analysisdeveloped to take into account thefinite rate of heat transfer within themelt particles is described in themathematics section.VI1-9 3

The rate of transfer of heat from themelt to the water-steam per unit mass ofwater depends also on the speed that themelt drops into the water. The modelassumes that the melt drops into thewater at a uniform rate over a finitetime interval, which is specified in theinput information, and the heat-transfermathematics takes this finite drop rateinto account. Its effects on the resultsare discussed later.C3.3 ADDITIONAL ASSUMPTIONSCertain assumptions are needed in additionto the basic equations and the heattransfer rate equations in order toobtain definitive solutions with themodel analyses. These are:a. The liquid water, steam, and gas inthe exploding mixture are at alltimes at the same uniform temperaturethroughout the mixture (exceptfor the steep temperature gradientin the region adjacent to the surfaceof a melt particle). The existenceof turbulent convective mixingwould be necessary for thisassumption to be valid, since thethermal diffusivities of liquidwater and steam are both much toosmall for heat conduction to maintaintemperature equilibrium and homogeneityover the time and distancescales involved. As long as somewater exists in the liquid phase, arelationship between the temperatureand the pressure is needed in orderto be able to solve the problem.The most convenient and often usedone is the vapor pressure-temperaturerelationship along the vaporliquidequilibrium curve.b. The exploding mixture is at a uniformpressure equal to that at thebottom of 'the mixture. This assumptionis necessitated by the use of asingle layer pair to represent themechanical properties of the medium.The expansion and compliant propertiesof the medium are functions ofboth pressure and temperature aswell as composition; they are evaluatedat the single assumed uniformpressure and single uniform temperature.The entire pressure differentialbetween the top and bottom ofthe mixture is assumed to appearacross the upper inertial layer.Because of the parabolic variationin pressure which results from theuniform expansion assumption, thepressure of most of the medium isactually closer to that at its bottomthan that at its top; the meanpressure is 2/3 of that at thebottom.C. The melt is at a uniform temperatureat all times, i.e., there is a morerapid interchange of -heat betweenthe various portions of the meltthan between the melt and the water.Clearly, this assumption is not realisticfor separated melt particlesof different sizes, since the smallerones would cool down faster thanthe larger ones and there is noopportunity for direct interchangeof heat among the various particles.Furthermore, for melt dropped intothe water over a finite period oftime, the later additions would beat the original melt temperaturewhereas the particles dropped earlierwould have cooled. There is alsothe difference between the averageand surface temperature of a particlethat depends upon its size, aspreviously mentioned. Despite thequestionability of the details ofthe assumption of melt temperaturehomogeneity, this assumption shouldyield an average rate of heat exchangethroughout the medium that isapproximately correct.d. As soon as any core material fallsinto the medium it becomes fragmentedinto the given average particlesize and becomes uniformly dispersedthroughout the medium. Furthermore,this particle size value and uniformityof dispersion remain unchangedas the mixture generatessteam and expands. This assumptionentails a very high relative velocitybetween the melt particles andthe medium for a very short time,until their dispersion in the mediumis completed. Thereafter the averagerelative velocity is zero, butthe average square of the relativevelocity is not zero, because ofturbulence. If dispersion of themelt by one or another of the possibleexplosion-triggering mechanismsis delayed until some time after themelt had started to drop into themedium, the resultant explosionwould be more severe, since thissituation would correspond approximatelyto an instantaneous drop anddispersion of that portion of thecore which had already reached themedium when the triggering eventoccurred. If the incoming melt weredispersed over only a portion of themedium when the explosion was triggered,the result may be more severelocal deformation and damage of theYessel but less energy expended inacceleration of the mixture.VIII-94

C3.4 CHEMICAL REACTIONSOnly one chemical reaction, that betweenthe zirconium and the water, is incorporatedin the model., This is consideredto be the principal reaction for theconditions of interest (Ref. 4). Thediffusion-limited parabolic reaction lawas given by Baker was employed in themodel (Ref. 5). The surface area ofzirconium in contact with the water wastaken to be the same as if the mass ofzirconium were broken up into particlesof the same average diameter as thatspecified for the entire melt.The parabolic reaction law yields aninfinite reaction rate at zero time.The computat~ional difficulties with aninfinite rate were avoided by assumingthat the zirconium already had a thinfilm of oxide on its surface when themelt was dispersed in the water. Theamount of hydrogen assumed to existinitially in the mixture was taken toresult from the oxidation of this initialfilm. Although the initial rate ofreaction is sensitive to the thicknessof this initial oxide film, later valuesof the amount reacted and of the heatgenerated as functions of time arelittle affected by this assumption.The liquid water and the steam werepresumed to participate equally in thechemical reaction; that is, in proportionto their mass fractions. The heatgenerated by the chemical reactionturned out to be 1 percent or less ofthe heat transferred from the melt tothe water in the short time intervalinvolved in most cases of interest. Theamount of hydrogen gas generated wasalso small, of the order of 10-5 that ofthe mass of the water. The chemicalreaction could have been safely ignoredin this phase of the computations.C3.5 VESSEL-WALL STRESSESProper computation of the vessel-wallstresses and comparison of thesestresses with the strength properties ofthe material are important considerationssince the main consequences of asteam explosion depend upon whether ornot the integrity of the vessel ispreserved.The distribution of stress in an engineeringstructure such as a reactorvessel is very complex. The vessel issubjected to many different types ofloadings, some of which are alleviatedand others intensified under the conditionsof a melt-down accident. The mainconcern in the model computations is inthe magnitude of the gross circumferentialand longitudinal stresses in thevessel walls developed by the transientinternal-pressure excursion resultingfrom the steam explosion. The onset ofdeformation and rupture should dependupon the combination of the new and thepreexisting stresses. However, the latterare ignored in the model since theyare expected to be small in comparisonwith the former.An excess of pressure inside the vesselover that outside at the same verticallocation produces a circumferential(hoop) tensile stress in the vessel wallequal to the excess pressure times theratio of radius to wall thickness. ifthe vessel is closed on both ends thereis also a longitudinal tensile stressequal to half the hoop stress. Thehemispherical heads are subjected tomembrane tensile stresses (equal in alldirections) given by the same formula asfor the longitudinal stress (Ref. 6).Superimposed upon these are bendingstresses due to restraints to radialexpansion at supports and flanges, pressuregradients in the mixture itself,and stress concentrations at tube entrances,at abrupt changes in shape orcross section, and at points of attachmentof internal and external supportstructures. Failures of certain portionsof the vessel structure might thusoccur for exploding mixture pressures of1/2 to 1/4 of those necessary for producinggross longitudinal ruptures fromexcessive hoop stress. To identify thepossibility of such occurrences, a. longitudinal"flaw" (of given length at agiven distance from the bottom of thevessel with a stress-concentration factorof 4) was introduced into the computerprogram.The simple direct proportionalitybetween excess internal pressure andwall stress applies strictly only tostatic or slowly changing internalpressures. For pressures that changesignificantly in a time comparable tothe natural period of radial vibrationsof the cylinder, inertial effects preventthe wall stresses from followingpressure changes instantaneously, sincestresses are accompanied by strains and,hence, radial displacements of thewalls. For pressure pulses that areextremely short compared to the radialvibration period, the maximum wallstress will be developed after the pulseis over and will be proportional to thepressure impulse, the integral of thepressure pulse over time, rather than tothe maximum pressure value itself.VIII-95

For some of the preliminary computationswith the PWR, a differential equationwas incorporated in the computer programfor determining the dynamic hoop stressesfrom the computed pressure variationsin the exploding mixture. With suitabledamping incorporated into the equationto avoid "ringing" effects, the dynamicstress was found to follow the staticstress variations rather faithfullyexcept for a time delay of about 1/4 ofthe natural period for radial vibrationsof 3 milliseconds. This is because thepressure changes for these computationsoccurred on a scale of tens of millisecondsfor the assumed particle size (400microns). As they had been shown to beunimportant to the situations of interest,computations of the dynamic stresseswere then eliminated from thecomputer program.In the computer program it is assumedthat the wall would rupture wherever andwhenever the hoop stress exceeds thetensile strength of the steel, or 1/4thof this value at the assumed flaw location.These cracks could not open fastenough to have any significant effect onthe course of the remainder of theexplosion. Computations were continueduntil the top head was assumed to rupturedue to excessive longitudinalstress just below the flange. Cracksshould also begin to form in the hemisphericalshell of the top head at aboutthis time, but only the development of avery large opening could relieve theinternal pressure buildup fast enoughfor the cases involving high rates ofheat transfer.In the above considerations the responseof the material to the applied stress isgreatly simplified. Failure is assumedto take place when the vessel-wallstress exceeds the tensile strength ofthe material. The energy that could beabsorbed by plastic deformation is notconsidered. For the more energeticinteractions this is a good approximationsince the kinetic energy of themoving mass by far exceeds the energyrequired to fail the vessel. In less-,energetic interactions, where failure ofthe vessel is marginal, this approach isoverly conservative.C4. RESULTS OF COMPUTATIONSA number of cases were computed forwhich it was assumed that the moltenportion of the core at a given timefollowing the initiation of a meltdownaccident dropped suddenly into the waterremaining above the grid plate at thattime. PWR reactor specifications wereused and the cases ranged from 3.9percent of the core dropping into 37,400pounds of water to 88.3 percent droppinginto 4500 pounds of water. The averageparticle diameter of the dispersed meltwas assumed to be 400 microns, inaccordance with the findings of Higginsand Schultz (Ref. 7). In all cases thetop head of the reactor vessel wasruptured and, except for the 3.9 percentcase, sufficient energy remained afterinelastic collision with the concreteshield above the vessel to breach thedome of the containment building.It is considered unlikely that a partiallymolten core would drop down intothe water above the grid plate; ratherit would tend to resolidify as it progresseddown the cooler core structurebelow it. Consequently, All subsequentcomputations were limited to the case ofmassive grid-plate failure, for whichall of the molten core and the moltengrid plate would drop into the bottomheadwater reservoir.Figure VIII C-2 shows the plotted resultsof a typical computer output for adispersal of 400-micron-diameter meltparticles. The various dependent variablesshown as functions of time are:P - Pressure in exploding mixture,psi/lOOQ - Pressure in gas and steamexploding mixture, psi/lOGaboveT - Temperature of steam-water, deg C/10U -Temperature of melt, deg C/1OOX - Position of the top of explodingmixture, ftS -H -Mass fraction of steam, percentHeat input to steam-water, cal/g/1OE - Conversion efficiency, heatkinetic energy, percent.The maximum pressure developed is notenough to rupture the bottom of the vesselfor this particular case. (The vesseldoes rupture at its bottom for anaverage particle diameter of 320microns). The pressure above the top ofthe mixture is seen to increase veryrapidly as the model "piston" approachesthe top head. Rupture of the top headoccurs because of this pressure buildupin combination with a hammer effect assome of the exploding mixture penetratesthe gas pocket above it to impact thetop head. (The Taylor instabilitydeveloped at the top of the mixture astoVIII-96

it rapidly decelerates explains thepenetration.)As discussed in Appendix B, it is notknown whether there would be any operativemechanism to disperse molten corematerials dropped into saturated water.In view of the possibility that finefragmentation and dispersion would notoccur with any great likelihood, aseries of computations were made withlarger average particle diameters,ranging up to 100,000 microns (10 cm).Melt-drop times were taken to vary from0.001 second to 3.2 seconds. The valuesfor the various input parameters usedfor these calculations are listed inTable VIII C-1.The results of these computations areshown in Fig. VIII C-3 for PWR reactors.Computations were made for all combinationsof values of APD, the averageparticle diameter, and TAU, the meltdroptime, for which a capital letterdesignator appears on the log-log grid.The number to the side of each letter,except for E, is the height in feet towhich the top head and control-rodshield would rise against gravity ifunimpeded by other forces or constraints.The significance of theletter designates is as follows:A - The vessel wall ruptures near thebottom and also later at the assumedflaw position running from 6 to 8feet above the bottomn with a stressconcentrationfactor of 4. The tophead also subsequently ruptures,hits the control-rod shield block,and the combined projectile penetratesthe dome of the containmentbuilding, 95 feet above the shieldblock.B - Same as A except that the vesseldoes not rupture near the bottom.* C - Same as A except that neither thebottom nor the assumed flaw locationruptures.D - Same as C except that the projectile* assembly does not have enough energyto reach the dome of the containmentbuilding.E -No rupture of the vessel occurs.The lines drawn on Fig. VIII C-3 delineatethe approximate boundaries betweenthe areas on the grid characterized bythe above-lettered combination ofevents. The second line from the rightis particularly significant since it isthe dividing line between the set ofconditions for which no containmentfailure is predicted to occur and thosefor which the containment might fail.Figure VIII C-4 shows the results ofsimilar computations for a BWR, for thevalues of the parameters listed in TableViIl C-1. The BWR reactor vessel containsa considerable amount of internalstructure above the shroud head thatcovers the reactor core. The movingmass (the hypothetical piston that formsthe top layer of the exploding mixture)is presumed to collide inelasticallywith the shroud head and to tear itloose from its mountings. Energy lossesare suffered in flowing past the internalstructures in the vessel both in theform of viscous drag resistance to flowand inelastic collisions when the shroudhead is torn loose and as it is plasticallydeformed. The entire movingassembly then collides inelasticallywith the top head and ruptures it if ithas enough energy to do so. The assembledmass then continues on upwardtoward the roof of the containmentbuilding. The BWR has no shield blockabove the vessel.Despite the differences between the PWRand the BWR and the differences in initial*and boundary conditions, the use ofimproved functions for the properties ofliquid water near the critical point forthe latter, and a first-order correctionfor the differences between the averageand surface temperatures of melt particles,of various sizes for the latter,the sets of curves for the PWR and theBWR are quite similar.The ordinate scales of Figs. VIII C-3and VIII C-4 can also be plotted on atime-scale basis, like the abscissascales, by using the thermal relaxationtime, TRT, rather than the averageparticle diameter. The thermal relaxationtime is the time required for theheat content of the melt to decrease tol/e of its initial value. In order tocompute the heat content of the melt itis first necessary to find the finalequilibrium temperature of the mixture,assuming a slow, constant pressure heattransfer and hence no dynamic effects.The equilibrium temperature turns out tobe 1710 C for the PWR and 1150 C forthe BWR (smaller melt to water ratio).The heat content for the former is thus217 cal/g and for the latter 300 cal/g.The value of TRT is simply the heatcontent divided by the initial rate ofheat transfer per gram of melt. Thecomputer prints out the value of thisinitial rate of heat transfer. However,it is readily computable from the massof melt, the total surface of the melt,VIII-97

the initial temperature difference, andthe sum of the radiative and filmboilingconvective heat-transfer coefficients.For the purpose of thiscomputation the average melt temperatureis used as its surface temperature.The TRT and the TAU scales can both berendered dimensionless in a significantway by dividing them by TMR, the mechanicalresponse time for the system. TMRis a measure of the time required forthe piston mass equivalent of the inertiaof the system to respond appreciablyto changes in forces imposed on it. Aformula for TMR is derived in theMathematics Section. This formula isthe same as that for about 1/6 of thenatural period of oscillation of a massequal to the piston mass supported on anair spring with the same cross-sectionalarea and initial depth as that of theexploding mixture, with the air springpressurized above ambient pressure justsufficiently to support the weight ofthe piston against gravity. The valuesof TMR, the mechanical repsonse time,turn out to be nearly the same for thePWR and the BWR, 0.31 second for theformer and 0.32 second for the latter.The dimensionless plots of TRT/TMR versusTAU/TMR are given in Fig. VIII C-5for the PWR and BWR. The curves differslightly in shape and in where theyintersect the coordinate axes. However,the similarities between the two sets ofcurves are more striking than the differences.They would not be expected tobe exactly alike since there are manyother dimensionless parameters involvedin the behaviors of the two systems.However, TRT/TMR and TAU/TMR appear tobe the primary parameters. Also significantis the fact that the change insystem behavior from a nonexplosive oneto a very explosive one is characterizedby a dividing line corresponding to thevalue of unity for the vector sum(square root of sum of squares) forthese two parameters. This dividingline is approximately that betweenRegions B and C of Fig. VIII C-5. Novessel ruptures occur for values ofTRT/TMR of about 10 to 12 or greater orfor values of TAU/TMR of 4 to 6 orgreater, regardless of the value of theother parameter.Maximum efficiency of conversion of heatenergy to mechanical energy occurs whenthe heat-exchange rate is so great thatessentially no expansion can take placeuntil all of the water has been convertedto very high pressure steam (constantvolume boiling). The system would thenbecome insensitive to heat-exchangerate, i.e., greater values of heatexchangerate would be of no consequence.This condition would correspondto both of the parameters TRT/TMR andTAU/TMR being much less than 0.1, whichwould require average particle diametersof less than 100 microns and melt droptimes of less than a millisecond.The question of the likelihood ofoccurrence of a steam explosion of givenseverity in a reactor vessel seems toreduce itself to the question of thelikelihood of occurrences of the variouscombinations of TRT/TMR and TAU/TMRalong one of the actual curves or alongan interpolated curve of Fig. VIII C-5.It is seen that if TAU is much less thanTRT, the phenomenon depends only on thesingle parameter TRT/TMR. On the otherhand, if TAU is much greater than TRT,the rate of heat exchange depends primarilyon the rate at which melt dropsinto the water and not on how rapidlyheat is exchanged between the water andan individual particle of the melt.Hence the value of TAU/TMR becomes thesignificant quantity.C5. MATHEMATICS OF THE MODELC5.1 BASIC MODEL EQUATIONSC5.l.l CONSERVATION OF ENERGYwhereM=wM +Ms + Mq +QM w,M s,Mgqw'qs'qgsLA(VIII C-l)= The rate of heat inputto the colder componentsof the mixture from themelt,= Instantaneous values ofthe masses of the liquidwater, steam, and gas,respectively,= Rates of heat input perunit mass to the water,steam, and gas, respectively,andL wvw = Latent heat of vaporizationtimes mass rate ofIn accordancethermodynamics,conversionsteam.of water towith the first law ofdQ = dH - VdP = Cps dT - T (DV/aT)p dP.VIII-98

Hence, for the steam, for example,qs = C ps T - T OV s /ýT) p P stwhere(VIII C-2)Cps = Constant pressure specificheat as a function of Ps andT,V 5 = Volume per unit mass of steamas a function of P and T,P = Partial pressure of thesteam, andT = Absolute temperature.Equations analogous to Equation (VIII C-2) are also set up for the water and forthe gas (hydrogen). Since the hydrogenis assumed to be a perfect gas,T (3V g/T)p = Vg.For the melt,4= - • (C~iM£ + CM.)Tmi C il ki +C si Msi) Mfusion of Component j and Msj is themass rate of increase of its solidphase. Equation (VIII C-3b) continuesto apply until all of the liquid phaseof Component j has been converted tosolid, then control reverts to Equation(VIII C-3a). The various components ofthe melt are thus assumed to be in separatephases at thermal equilibrium: thepossibility of formation of liquid orsolid solutions is not considered.The additive term on the right of bothequations represents the rate at whichthermal energy is added to the meltwithin the exploding mixture by thecontinual dropping into the mixture ofadditional melt from the core at itsoriginal temperature Tmo over a timeperiod T. Here Wki is the total mass ofliquid over this time period. This secondterm is deleted from the two equationsupon completion of the droppingprocess; i.e., when the time begins toexceed T.C5.1.2CONSERVATION OF MOMENTUMA(P - P ) -Wg = Wk + (W%/T) x,(VIII C-4)4+iL C 9iW i (T - mi= 1wherex = Position of the topexploding mixture,oftheor(VIII C-3a)W = Effective mass (half of instantaneousmass of melt pluswater),6 = -LjMsj4+il C~iWi (Tmo - Tm )/T.A = Cross-section area,P = Pressure at the bottom of themixture,(VIII C-3b)Equation (VIII C-3a) applies while theinstantaneous melt temperature Tm is notat any of the melting points of the fourcomponents, UO 2 , ZrO 2 , Zr, and Fe. Q isthe same quantity as on the left side ofEquation (VIII C-l). C.i and Mii arethe specific heat and instantaneous masswithin the exploding mixture, respectively,of the liquid phase of Componenti, and Csi and respective Msi are thesame respective quantities for its solidphase.Whenever the melt temperature reachesthe melting point Tmi of the j-th component,Equation (VIII C-3b) begins toapply, wherein L is the latent heat ofP = Pressure in the space abovex the mixture (filled with steamand hydrogen),g = Acceleration of gravity, andW m= Total mass of the melt.Since T is the melt drop time, Wm/T representsthe rate of increase of melt inthe mixture.Px is computed as a function of time byapplying an equation similar to that of(VIII C-l) to the steam and hydrogen inthe upper part of the vessel with Q = 0.The steam in this part of the vessel isnot quite saturated initially and becomeseven less so as it is compressedadiabatically.VIII-99

C5.1.3CONSERVATION OF MASSxA - V wt (1 - awg)Ww [FVs + (1 - F) Vw].(VIII C-5)Equation (VIII C-5) states that thevolume of the exploding mixture at anytime, less the volume occupied by themelt components, equals the volume occupiedby the steam plus the volume occupiedby the water. Ww is the originalmass of water; F is the mass fraction inthe form of steam, and Vs is the volumeper unit mass of steam; Vw is the volumeper unit mass of water; wt is the massfraction of the mixture remaining afterthe escape of some through the orifice;and 1 - awA is the mass fraction of waterremaining after chemical generationof a mass of hydrogen equal to wgWwIwhere a is the ratio of molecularweights of water to hydrogen. Thestandard orifice equations yields wt,from which wt is obtained by numericalintegration. Likewise, *g comes fromthe chemical reaction rate equation.Since the hydrogen gas occupies the samevolume as the steam, we have, furthermore,(1 - awg) FVs = WgVg. (VIII C-6)The time derivatives of EquationsC-5) and (VIII C-6) furnish neededtionships for the analysis. Theof Mws in Equation (VIII C-l) canbe expressed in terms of F and theparameters.C5.1.4EQUATIONS OF STATE(VIIIrelavaluethenotherFor steam, the analytic expression quotedby Justi for Vs(Ps,T) was employed(Ref. 8). Tabulated density values forboth saturated and super-heated steamcould be reproduced to within an accuracyof 1 percent by using this expression,except in the neighborhood of thecritical point. The needed pressure andtemperature partial derivatives of Vswere also obtained from this expression.The density of saturated liquid water asa function of temperature was obtainedfrom a mathematical expression quoted byDorsey (Ref. 9). This expression wasmodified to apply to subcooled water byadding a term proportional to the excesspressure over the saturation value, withthe coefficient of proportionality afunction of temperature, so as to approximatethe tabulated values quoted byDorsey (Ref. 10).The specific heat at constant pressurefor the steam as a function of Ps and Twas obtained from an expression for theentropy, S, of steam as quoted by Justi(Ref. 8), using the thermodynamic relationshipCps = T(aS/9T) . The specificheat of saturated watep was obtained bymatching a polynomial to published data(Ref. 11). In addition, analytic expressionswere used for the vapor pressureof water as a function of temperature(Ref. 12) and for the latent heatof vaporization of water as a functionof temperature (Ref. 13).The ideal gas law was used for theequation of state and the specific heatof the hydrogen gas. This was deemedsatisfactory because the hydrogen is aminor component with a low partial pressure.The densities, specific heats, heats offusion, and melting point temperaturesof the various components of the meltwere considered to be constants at averagevalues over the temperature range ofinterest (Ref. 14). The effects of thevariations in the density of the meltwith temperature and pressure were neglected.C5.2 MECHANICAL RESPONSE TIMEEquation (VIII C-4) may be rendered dimensionlessby dividing through by APo,where P 0 is the initial pressure at thebottom of the mixture, obtainingWx 0 d 2 (X/X)WAP0 dt2 AP= (P 0 -; X-O) 0WmXoAPTd (x/xo)dt(VIII C-7)wherein x. is the initial depth of theexploding mixture.The coefficient of the term on the lefthas the dimensions of time squared. Itssquare root, given bytmr = WXo/APO, (VIII C-8)VIII-IO0

is precisely the factor needed to makethe time scale of the equation dimensionlessthroughout; e.g., the left sidebecomes d 2 (x/xo)/d(t/tmr) . SincePo = Pa + P-' the ambient pressure plusthe additiogal pressure at the bottom ofthe mixture due to gravity, equals(W/A)/g, Equation (VIII C-8) can also beexpressed in the formtmt = V(xo/g)/(1 + Pa/Pg)(VIII C-9)Equation (VIII C-9) for tmr is the sameas the equation for 1/27 times theperiod of a pendulum of length xo /(l +P a/Pg ).Equation (VIII C-8) is the same as thatl/2ff times the period of small (isothermal)oscillations of a mass W mounted onan air spring of cross-section area A,length xo, and filled to a pressure P.,just enough above ambient pressure tosupport W. The differential equationfor a small displacement, x, of W mountedon the air spring would bedt2x(_O) xP 0A,(VIII C-10)since -(x/x 0 Po) is the isothermal pressurechange produced by the displacementx (from Boyle's law). This equationintegrates to simple harmonicmotion with a period of27rT WXo/APo-C5.3 DYNAMIC STRESSES IN VESSEL WALLSThe static circumferential (hoop) tensilestress in a cylinder of radius Rand wall thickness tw

C6.1.IRADIATIVE HEAT TRANSFERRadiative heat transfer is a majorcontribution at the melt temperature ofinterest. The net rate of radiation ofheat per unit surface area away from asurface of emissivity ems and temperatureTms toward a surface of emissivityc and temperature T is given bysteady-state solution of the heat-conductionequation in spherical coordinatesyields the value of the temperaturegradient in the medium at thesurface of the sphere and hence the rateof heat flow into the medium from thesphere. From this heat-flow rate theheat-transfer coefficient for conductionbecomesa(Tms 4 T4)q T= - ) (VIII C-14)Hcond = 2K/D,(VIII C-17)CmsYwhich corresponds to a value of 2the Nusselt number, HD/K.forwhere a is the Stefan-Boltzmann constant(Ref. 15). If the melt particles arefar enough apart and of such shape thatone portion of a melt particle surfacecannot receive radiation directly fromanother portion, and if the value of cfor the water-steam-gas is effectivelyunity (perfect absorber and radiator),then Equation (VIII C-14) becomesq = m (Tms 4 - T4)" (VIII C-15)Each melt particle is considered in themodel to consist of the same mixture ofthe two oxides and the two metals. Inpractice, the proportions of these mayvary greatly from one particle to another.However, the possibility thatany metallic component, with a low emissivity,can form an appreciable fractionof the melt surface is ruled out, sincea thin film of oxide would form almostimmediately upon contact with steam orliquid water. The available informationon the emissivities of U0 2 and ZrO 2 attemperatures in the neighborhood of2800 C indicate that 0.8 would be anappropriate value for 6ms (Ref. 14).Hence the contribution of radiation tothe heat transfer coefficient is givenbyHrad = 0.80 (Tms + T)Tms+ Tms T + T2),(VIII C-16)where Tms and T are in degrees absolute.C6.1.2CONDUCTIVE HEAT TRANSFERIf a hot sphere of surface temperatureTms and diameter D is immersed in ahomogeneous medium of ambient temperatureT and thermal conductivity K, theEquation (VIII C-17) is employed in themodel for the conductivity contributionto the heat-transfer coefficient whenthe fluid is all steam. If Vol is thevolume fraction of steam before all ofthe liquid has vaporized, then the expressionincorporated in the model isHco = ( - Vol)Hbr + Vol Hcond,(VIII C-18)where Hbr is the Bromley heat transfercoefficient applicable to stable filmboiling, discussed later. Equation(VIII C-18) is tantamount to assumingthat a Vol fraction of the melt particlesis surrounded primarily by steamand the remainder is surrounded by afilm of steam only, beyond which isliquid water.It is recognized that Equation (VIII C-17) represents an approximate evaluationof the contribution of heat conductionto the heat-transfer coefficient. Sinceconduction to steam would be expected tobe a lesser contributor to overall heattransfer than natural convection andradiation to liquid water, this approximationis deemed reasonable.C6.2 CONVECTIVE HEAT TRANSFERNo account was taken in the model of thecontribution to the heat-transfer coefficientof a melt particle in steam dueto either natural or forced convection,although the latter could be significant.Much turbulence no doubt wouldexist in practice, but the value of theReynold's number, needed to compute thecontribution due to forced convection,is unknown. As previously pointed out,the average relative velocity betweenthe melt particles and the steam is presumedto be zero. The contribution tothe heat transferred during the timethat the melt particles are being rapidlydispersed throughout the medium isVIII-102

neglected, by assuming this time to berelatively very small.An important consideration is the rateof heat transfer in the initial stagesof the explosion when the cold portionof the exploding mixture is almost entirelyliquid water. The theory of heattransfer in the stable film boiling regimewas developed by Bromley (Ref. 3).He assumed that the vapor blanket arounda hot, horizontal, stationary tube rosecontinuously under the action of buoyantforces, at a rate limited by viscousdrag on the tube. The heat transferredis that needed to replenish the vaporblanket by vaporization of the surroundingliquid. This heat is transportedthrough the film primarily by conductionand radiation. The resultant contributionto the heat transfer coefficient isH br K=v0.62 ( Pv (Pk- pv)g En)l/4 ,Driv(Tms -T)(VIII C-19)where Kv, pv and n. are the thermal conductivity,density, and viscosity of thevapor, respectively; pk is the densityof the liquid; and En is the differencein enthalpy between the vapor blanket atits mean temperature and the liquid.Analytic expressions were used for theviscosity (Ref. 16) and the enthalpy(Ref. 8) of steam as functions of temperatureand pressure in Equation (VIIIC-19). The thermal conductivity was obtainedfrom the viscosity and the constant-pressurespecific heat (previouslydescribed) by using a Prandtl numbervalue of 1/1.15 (Ref. 11). The analyticexpressions for the liquid and vapordensities have been discussed.A considerable amount of theoretical andexperimental work has been reported onheat transfer to liquids from movingspheres, the case of forced convection(Refs. 17, 18). Heat-transfer rates ofabout one order of magnitude greaterthan those characteristic of naturalconvection can thus be attained. Forcedconvection effects are neglected in themodel since the relative velocity betweenthe melt particles and the waterdue to turbulence is not known. Aspointed out previously, the average velocitydifference is assumed to be zeroand the heat transfer during dispersionis neglected.Radiative heat transfer and natural convection(Bromley's correlation) makeabout equal contributions to the initialmagnitude of the heat-transfer coefficientfor 100-micron-size particles.However, as the particle size increases,the radiation contribution become predominant.Thus the overall heat-transferrate predicted is not overly sensitiveto the particular/film-boilingcorrelation utilized. These results ofthe present model computations are inaccord with the meager amount of difficult-to-obtainexperimental data on heattransfer at such high temperatures. Inexperiments on film boiling with anelectrically heated carbon rod in water,Oki and Nabemoto measured heat-transfercoefficients for rod temperatures rangingfrom somewhat under 1500 C to somewhatover 3000 C above the saturatedwater temperature (Ref. 19). They claimtheir experimental results to be inagreement with radiative heat transferand Bromley's equation. Attempts toobtain the individual magnitudes of thetwo terms from pairs of points alongtheir curve proved to be futile fortemperatures above 2000 C. The curvedrawn through their data points agreeswell with the assumption of only radiativeheat exchange above this temperature,for an emissivity value close tounity.C6.3 TEMPERATURE GRADIENT IN MELTPARTICLESIn all preliminary computations it wasassumed that the surface temperature ofeach melt particle was the same as itsaverage temperature. This assumptioncan be justified only for particles ofabout 60 microns or less in diameter onthe time scale of events of interest, ifthermal conduction is the only mechanismof internal heat transfer in the particle.(The thermal-diffusivit value formolten U0 2 is about 3.8 x i0-ý cm 2 /sec.)Other mechanisms such as internal convectionwould have to be invoked to justitythe isothermal assumption for thelarger particles. However, since theassumed initial temperature of the meltis just above the melting point of U0 2 ,a solidification front would begin toprogress from the surface inward.as soonas the melt is dispersed throughout thewater. The solid crystallites of U0 2and ZrO 2 might form as a porous matrixthat would hinder the convection of thelower-melting-point liquid metals.An approximate procedure was devised tocorrect for the fact that the melt particlemay not be isothermal. The temperaturedistribution within the meltparticle is assumed to be parabolic,with a maximum temperature at the centerof the assumed spherical particle and amaximum (negative) temperature gradientVIII-103

at its surface. The value of this gradientis determined simply from the rateof heat transfer in the medium adjacentto the particle surface and the thermalconductivity of the particle (weightedaverage over the four components).Since the average temperature is known,the assumed parabolic distributionenables the surface temperature to bedetermined. The result is the followingformula for the surface temperature ofthe melt:Tm+ Htc D)Tis TsKm( : 1TA solution exists for the transientheat-transferequation for the case ofthe cooling of an initially isothermalsphere immersed in a medium characterizedby a fixed value of Ht_ at the surfaceof the sphere (Ref. 20). Althoughthis solution applies only to the assumptionof a constant ambient tempera-approximate one proposed above. Thetemperature gradient at the surface ofthe sphere for this case is initiallymuch greater than that which would begiven by a parabolic-temperature-distributionlaw. This is to be expected,H tc D (VIII C-20) since the sphere is assumed initiallywhere H c is the heat transfer coefficienttor the medium adjacent to thesurface and Km is the thermal conductivityof the melt particle. The quantityT. is the average temperature of themelt particle and T is the average watertemperature away from the particle surface.In applying Equation (VIII C-20) to determinethe value of Tms by computer,the value of Tms is first assumed to bethe same as Tm. This enables the valueof Htc to be computed in accordance withthe formulas previously presented inthis Appendix. Equation (VIII C-20) isthen applied to obtain an improved valueof Tms, and Htc is recomputed. The processis repeated until successive valuesof Tms are in good agreement.ture for the surrounding medium, it isinstructive to compare it with theisothermal. However, an approximatelyparabolic distribution is attained afteronly a few percent of the total time ittakes for the sphere to lose about 98percent of its original temperaturedifference from the medium. Toward theend of the cooling process the parabolicdistribution begins to give too high avalue for the (negative) gradient at thesurface. (The parabolic distribution isthe exact steady-state solution for asphere in a cooling medium with heatsources uniformly distributed throughoutits volume (Ref. 21).The computed initial differences betweenTm and Tins ranged from only a few degreesfor a particle size of 100 micronsto about 600 C for a particle size of10,000 microns. The uniform-particletemperatureassumption would thus appearto be quite reasonable for all but thelargest particle sizes considered.VIII-104

References1. Simon, R., "Computer Model Study of the Smelt-Water Physical Explosion", Presentedat the Atlanta Heat Transfer Conference, A.I.Ch.E. (August, 1973).2. Taylor, G. I., "The Instability of Liquid Surfaces When Accelerated in a DirectionPerpendicular to Their Planes", Proc. Roy. Soc. London, A201, p. 290(1950).3. Bromley, L. A., "Heat Transfer in Stable Film Boiling", Chem. Eng. Progress, 46,p. 221 (1950).4. Morrison, D. C., Ritzman, R. L., Wooton, R. 0., Genco, J. M., Barnes, R. H.,Fackelmann, J. M., and Carbiener, W. A., "An Evaluation of the Applicability ofExisting Data to the Analytical Description of a Nuclear-Reactor Accident - CoreMeltdown Evaluation", BMI-1910, Topical Report, Task 18, Battelle-Columbus Laboratoriesto AEC, p. C-7 (July, 1971).5. Baker, W. L., and Ivins, R. 0., "Analyzing the Effects of a Zirconium-Water Reaction",Nucleonics, 23, pp. 70-74 (July, 1965).6. Jensen, A., Applied Strength of Materials, McGraw Hill, New York (1957), pp. 58-59.7. Higgins, H. M., and Schultz, R. H., "The Reaction of Metals with Water and OxidizingGases at High Temperatures", Report No. IDO-28000 from Aerojet GeneralCorporation to AEC (April 30, 1957).8. Justi, E., Spezifische Warme, Enthalpie, Entropie and Dissoziation Technischer-se, Springer Verlag, Berlin (1938), p. 30.9. Dorsey, E. N., Properties of Ordinary Water-Substance, Reinhold PublishingCorp., New York (1940), p. 584.10. Dorsey, E. N., Ibid., p. 240.11. Meyer, C. A., McClintock, R. B., Silvestri, G. J., and Spencer, R. C., Jr.,Thermodynamic and Transport Properties of Steam, ASME Publication (1967).12. Dorsey, E. N., Ibid., p. 574.13. Dorsey, E. N., Ibid., p. 613.14. Brassfield, H. C., White J. F., Sjodahl, L., and Bittel, J. T., "RecommendedProperty and Reaction Kinetics Data for Use in Evaluating a Light-Water-CooledReactor Loss-of-Coolant Incident Involving Zircalloy-4 or 304-SS-Clad UO21,'Report GEMP-482, General Electric Co., (April, 1968).15. McAdmas, W. H., Heat Transmission, Third Edition, McGraw-hill, New York, (1954),p. 63.16. Dorsey, N. E., Ibid, p. 64.17. Stevens, J. W., Witte, L. C., and Cox, J. E., "The Vapor Explosion - Heat Transferand Fragmentation. VI. Transient Film Boiling and Transition Boiling froma Sphere", Technical Report ORO-3936-9, University of Houston to AEC (September,1972).18. Vliet, G. C., and Leppert, G., "Forced Convection Heat Transfer from an IsothermalSphere to Water", J. Heat Transfer, ASME Transactions, p. 163 (May, 1961).19. Oki, H., and Nabemoto, A., "Film Boiling at a High Temperature of about 3000 C",Report presented at Int. Symposium of Japanese Society of Mechanical Engineers,September 1967, - Abstract #13 in Heat Transfer Newsletter No. 2, Committee ofReactor Safety Technology, European Nuclear Energy Agency, CREST-NL-2 (December,1967).VIII-105

20. Carslaw, H. S., and Jaeger, J. C., Conduction of Heat in Solids, Second Edition,Clarendon Press, Oxford (1959) , p. 234.21. Carslaw, H. S., and Jaeger, J. C., Ibid., p. 242.VIII-106

TABLE VIII C-i VALUES OF PARAMETERS USED FOR MASSIVE GRID-PLATE FAILURE,STEAM-EXPLOSION COMPUTATIONSPWRBWR(a) More representative values would be 70,000-80,000 psi. The energy imparted toMasses, lbUO2ZrO 2ZrH 2 0H 2Top head, flange, andcontrol-rod shield,PWRShroud head and internalstructure175,60030,0006,30037,4000.374323,5000369,800105,20022,100149,5001.495358,000299,500Initial ConditionsPressure, psiaMelt temperature, CWater temperature, CVessel ParametersRadius, internal, ftHeight, internal, ftWall thickness, in.Wall tensile strength,Flaw position, ft fromFlaw strength, psipsi (a)bottom13.22,850907.1834.607.875130,0006 to 832,50014.72,8509310.4665.956.313130,00012 to 1632,500(a) More representative values would be 70,000-80,000 psi. The energy imparted tothe pressure vessel head is quite insensitive to the assumed tensile strength.Table VIII C-1VIII-107/108

.rq.rqmP Tx X Steam andGasOrificesaInertialLayerExpandingLayerx tn}Mixture of Hot.............................................................................................4.X~Molten Metals andOxides with Water,PTX Steam, and GasesXv- ..-.- -Cylindrical VesselFIGURE VIII C-1Basic Computational Model Geometry100n"7550-, 50C 5o 2500 10 20 30 40 50 60IfTime,rmsecO Mass fraction of steam, % 0 Melt temperature, C/100* Pressure in exploding mixture, 0 Position of exploding mixture,psi/100ftO Temperature of coolant, C/1O 0 Heat to kinetic energy6 Heat input to coolant, conversion, %cal/g/lO0 Pressure above exploding mixturepsi/lO0FIGURE VIII C-2Typical Computer Results

E E E1 0 10(19) 6)(1D9) 0(1F) 5(19)T(122) .C(122) C (122) :C(F2) 2(6X)103 I B(582) B(582) B(579) B(56 (58)D(EA(1837) A(1837) A(1616) .A(1541\ B(795 C(283) D EI .'..---.'-..-.i-'--..•_ __° 2tS) A~5 9 '- A (8---" '.~....... Y.. .....2~~ AL5) ~ A 19) A 10) A80 (419) C(547) D(49)110 3 10- 2 t0- L 100 10Melt Drop Time (TAU),secFIGURE VIII C-3PWR Results° I0 04 DQ7) (7 mEEC(177) C(174) *C(143) 120)1O 3 B. B(997) *B(958) C(466\ (7 )I J 19) E10 v'. . .. ........................... '"' ..... I10 I0" 2 2 I0" 1 1000 10 1Melt Drop Time (TAU), secFIGURE VIII C-4 BWR Results

(a) PWR (b) BWR100100 -10E10NE0c \AA \N..\ \.1B \II.I0.10. 1FIGURE VIII C-51TAU/MR0.110 I0.11.0 10TAU/TMRDimensionless Presentation of CalculationalResultsFig. VIII C-1 - Fig. VIII C-5VIII-1.09/110

APPENDIX DNONCONDENSABLE GASESVIII-Ill

'aS

Appendix DTable of ContentsSectionDl. HYDROGEN IN A PWR ................................................Page No.VIII-11501.101.2Dl. 3Production .................................................Transport to Containment ...................................Events in the Containment Atmosphere .......................Dl.3.1Dl.3.2Dl.3.3Flammability Limits ................................Detonation Limits ..................................Combustion of Hydrogen at Entry toContainment ........................................VIII-II5VIII-115VIII-115VIII-115VIII-116VIII-116D2. EXAMINATION OF DETONATION EVENTS .................................D2.1 The Mechanism of Detonation ................................D2.2 Calculations of Detonation Parameters ......................D2.3 Duration of the Shock Waves ................................D2.3.1 Application to a Large Vessel ......................D2.3.2 Hydrogen in a BWR ..................................D2.4 Carbon Dioxide Generation in PWR and BWR ....................REFERENCES ..............................................................VIII-117VIII-117VIII-118VIII-119VIII-120VIII-121VIII-122VIII-123List of TablesTableVIII D-1 PWR Hydrogen Sources ...........................................Page No.VIII-125/126VIII D-2Calculated Detonation Parameters for LimitingCompositions ...................................................VIII-125/126List of FiguresFigureVIII D-1VIII D-2Flammability Limits of Hydrogen-Air-Steam Mixtures(Ref. 1) .......................................................Composition Isotherms for Containment Atmosphere atSaturated Steam Conditions .....................................Page No.VIII-127/128VIII-127/128VIII- 113

Appendix DNoncondensable GasesThere are two major sources of noncondensable-gasgeneration in a reactoraccident in which core meltdown occurs.Metal-water reactions in the pressurevessel will produce hydrogen gas (H 2 )shortly before and during the meltdownprocess. Later in the accident, carbondioxide (C0 2 ) will be generated as themolten core material causes thermaldecomposition of limestone aggregate inthe concrete base mat of the containmentstructure. The production of thesegases presents two potential threats tocontainment integrity. Both gases willcause a buildup in internal gas pressurein the system. Hydrogen generation canalso lead to combustible mixtures withthe oxygen already present in thecontainment atmosphere. Ignition of thehydrogen-oxygen mixtures can produce anexothermic chemical reaction which, dependingupon conditions, might developinto a detonation. The introduction ofadditional thermal energy into thecontainment atmosphere will cause a risein pressure, perhaps coupled with ashock-wave loading on the containmentwalls if the detonation occurs. Theimportant question in all of thesesituations is whether, or under whatconditions, would they likely result incontainment failure by overpressurization.The hydrogen-generation problemis examined first for each type ofwater-reactor system; an analysis of thecarbon dioxide generation problem follows.D 1.HYDROGEN IN A PWRD1.1 PRODUCTIONSeveral sources of hydrogen exist for aPWR system during the course of a coremeltdownaccident. The principalsources are identified in Table VIII D-1along with an estimate of the rate ofhydrogen production and the potentialyield for each over a selected timeperiod. It is apparent that the metalwaterreaction is the most importantsource of hydrogen, both in amount andrate of generation. While the othersources could become significant atsufficiently long times, the metal-waterreaction generated hydrogen will dominateduring the important first severalhours after the accident. The bestestimate of the amount of hydrogen generatedby this source is 600 lb-moleswhich corresponds to 75 percent zirconiumreaction in the reactor-core regionduring the period of core meltdown. Thevalue is based on parametric heattransferanalyses of the meltdown processwhich takes into account steamavailabilityand core-slumping effects.More or less metal-water reaction can bepredicted depending on assumed conditions,but it is shown that the finaleffects are not to sensitive to theseuncertainties.D1.2 TRANSPORT TO CONTAINMENTThe hydrogen-steam mixture will exitfrom the core at high temperatures,probably 2000 F or higher, and the pathsfrom this location to the containmentdepends upon the type of pipe break.For a hot-leg break the gas mixture willexit from the primary system directlyfrom the upper-plenum region of thepressure vessel without much cooling.For a cold-leg break the gas mixturemust travel from the upper plenumthrough the steam-generator tubes beforeexiting to the containment. This willresult in cooling of the gas mixture totemperatures of 500-600 F. Shapiro andMoffette have noted that the spontaneous-ignitiontemperatures of hydrogensteammixtures can range from about 1000to 1500 F depending upon exit velocity,steam concentration, and geometry of thesystems (Ref. 1). Thus the temperatureof the steam-hydrogen mixture that exitsfrom the broken pipe into thecontainment should be above thespontaneous ignition temperature in ahot-leg break but not for a cold-legbreak.D1.3 EVENTS IN THE CONTAINMENTATMOSPHERED1.3.1FLAMMABILITY LIMITSShapiro and Moffette reviewed this subjectabout 15 years ago and developedthe limit composition curves shown inFig. VIII D-1 (Ref. i). In spite of theconsiderable amount of work on hydrogenoxygencombustion since then there hasbeen no adequate additional study of theeffect of water vapor (steam) on thelimits. However, the limit lines givenin Fig. VIII D-1 probably represent theminimum compositions for flame propagation,i.e., the flammability region mayVIII-115

actually be smaller than indicated.Several factors may cause this reduction,such as downward rather thanupward flame propagation and combustionin smaller vessels which would result inenergy losses to the vessel walls. Noteon the figure that increased temperatureor pressure have a relatively smalleffect on the flammability limits.Therefore, the flammability regiondefined in the figure for a temperatureof 300 F and a pressure of 100 psig isused as the basis for determininghydrogen-combustion conditions in thisstudy.D1.3.2DETONATION LIMITSShapiro and Noffette state that inhydrogen-air mixtures (no steam) as thehydrogen concentration increases beyond9 v/o, not only does the flame propagatein all directions but the rate ofpropagation increases rapidly (Ref. 1).Then when the hydrogen in air reaches 19v/o, the mixture detonates. This conditionpersists until the hydrogen in airexceeds 57 v/o. The detonation occursin a shock wave which may travel atvelocities of several thousand metersper second. Very small ignition sourcessuch as sparks can trigger detonationsby initiating local combustion. Thecombustion advances quickly and producesturbulence and shock waves in and aheadof the flame front which rapidly transformthe deflagration into a detonation.The detonation shock wave is sustainedby the energy of the chemical reactionwhich is itself initiated by thetemperature and pressure of the wave.The detonation compositions noted abovefor hydrogeni-air mixtures constitute thelower and upper limits of the assumeddetonation limit boundary in Fig. VIIID-1. In constructing the boundary itappears that Shapiro and Moffette simplyassumed parallel behavior (to the lowtemperature and pressure flammabilitylimit boundary) at increasing steamconcentrations. Thus, the assumed detonationlimits must be considered quiteuncertain, particularly in view of thevery irregular detonation limits thathave been observed in other threecomponentsystems (Ref. 2). For thepurposes of this study, as it is shownthat results of hydrogen generation willnot be sensitive to detonation-limitpredictions, no effort was made toobtain better estimates.D1.3.3 COMBUSTION OF HYDROGEN AT ENTRYTO CONTAINMENTThe hydrogen-steam mixture from thereactor primary system will enter thelarge containment volume via therelatively small room which contains thebroken primary coolant loop. Theserooms are typically only a few percentof the total containment free volume, sofor expected hydrogen-steam flows in therange of 5000-10,000 cfm the residencetime of hydrogen in the room would beonly a few minutes. It is doubtful thatany significant hydrogen burning wouldoccur in this room for several reasons.First, the reactor-coolant blowdowriwould be expected to drive nearly allthe original air out of the room, thuseliminating the source of oxygen for thereaction. In some accident sequencescontinued steam flow from water boiloffin the core would prevent return of airexcept by back diffusion. This wouldrequire many hours. In other accidentsequences, containment spray waterfalling through upper gratings in theroom could bring entrained air back morerapidly. The sprays would also reducethe steam partial pressure to the rangewhere theoretically flammable hydrogenair-steamconcentrations could develop.Although some combustion might begin forthe hot-leg break where the hydrogensteammixture could enter the room abovethe spontaneous ignition temperature,the oxygen supply probably could not bemaintained and the reaction would sooncease. on this basis it is concludedthat hydrogen combustion in the smallloopcompartment can be ignored.The next point at which the hydrogensteammixture might ignite and burn isat the floor grating to the large uppercontainment space (i.e., as the gasflows up out of the small-loop compartment).To maintain continued combustionwould require a large ignitionsource and a continuous supply of air(oxygen). These conditions could not bemet for a cold-leg break, but for a hotlegbreak it is possible that the hydrogen-steammixture would come through thegrating at or above the spontaneousignitiontemperature. Once the combustionstarted, the heat liberated wouldfavor continued reaction provided oxygencould be supplied at the necessary rate.Complete combustion of 600 lb-moles ofH (Table VIII D-1) would require 3001i-moles Of 02 or about one-half thecontainment air supply. Assuming thehydrogen is generated at a uniform rateover 3/4 hour this fraction of thecontainment atmosphere would have tocirculate to the combustion site withinthis time period. The required rate ofcirculation would steadily increase asoxygen was consumed by the reaction.Under these conditions it is doubtfulhow long the burning period would last.VIII-116

However, if complete combustion of thehydrogen is assumed, the rate ofthermal-energy addition to thecontainment atmosphere can be estimated.Since the heat of combustion of the hydrogen-oxygenreaction is 1.04 x 105 Btuper, lb-mole H 2 consumed, the rate ofenergy addition (assuming uniform H 2production) would be 8.3 x 107 Btu/hrfor the 3/4 hour period. In someaccident sequences this rate may besignificant in terms of the internalpressure rise that would result.If the hydrogen does not burn at the-floor grating in the large upper containmentspace then it will tend to riseand accumulate in the upper regions ofthis space. Internal convection currentsin the containment atmosphere willgradually mix the hydrogen throughoutthe volume, but after appreciablehydrogen has been generated, explosivemixtures can develop. For a typical PWRcontainment, 75 percent zirconium-waterreaction will lead to the followinghydrogen concentrations in the containmentbuilding assuming complete mixingand steam saturated conditions:Temp,230130FH 2 , v/o714If only the upper containment space isconsidered, the hydrogen concentrationswould be about twice these values. InFig. VIII D-2 it can be seen thatregardless of the hydrogen concentrationin a steam-saturated atmosphere at temperaturesabove about 230 F, the compositionwill lie outside the flammabilityregion. Therefore, a propagating combustionwave cannot occur. In accidentsequences where containment sprays arenot operable, the atmosphere must coolto 230 F or below before any explosiveH 2 -0 2 reaction is possible. However, insequences where sprays operate, theatmosphere is cooled to 230 F in amatter of minutes and then down to therange of 130 F within an hour. Thebuildup of hydrogen under these conditionscreates gas mixtures well withinthe flammability limits. The appearanceof a small spark in such mixtures isoften considered to present relativelyhigh potential for initiating a detonationevent. A detonation will producetwo coupled effects on the containmentstructure. First the detonation shockwave will deliver an impulse loading tothe containment wall. This dynamicloading will quickly decay to a somewhatmore sustained pressure pulse from theexpanding gas that has undergone adia-batic heating. Heat transfer from thegas to the wall and to internal structureswill cause decay of this secondarypressure pulse. Because of the importanceof these loadings to containmentintegrity, the detonation problem wasanalyzed in more detail. The results ofthis analysis are given in the nextsection.D2. EXAMINATION OF DETONATIONEVENTSD2.1 THE MECHANISM OF DETONATIONA gaseous detonation may be initiated bya shock of sufficient intensity, or itmay develop more slowly from a laminarflame. In the latter case the combustionwill produce an increase in pressurewhich will start weak pressurewaves propagating into the unburned gas.Such small disturbances, especially asthey interact with the walls, promotethe formation of turbulence in the flamefront. Stronger shock waves are generatedas a result of turbulent burning,especially as little pockets of gasexplode in the flame front. These shockwaves are propagated into the unburnedgas, accompanied by local compressionand temperature rise. When the shockwave is strong enough, the chemicalreaction begins in advance of the flamefront. The velocity of sound is greaterat higher temperatures. Therefore successiveshock waves overtake one anotheruntil a detonation results. The detonationis characterized by an exothermicreaction which is initiated by the shockwave and which in turn drives the shockwave. The velocity of the detonation isdetermined by the velocity of sound inthe burned gas, and is usually severaltimes the velocity of sound in theoriginal unburned mixture.By writing down equations expressing theconservation of mass, of momentum, andof energy in the gas as a shock wavepasses through, one can describe therelationship between the pressure, p,and the specific volume, v, for any gas.This relationship is known as theHugoniot curve. When no chemical reactionis involved, the original pressureand volume, p, and vl, will be transformedinto a pressure and volume in theshock wave, p 5 and v , which depend onthe intensity of theSshock.Such a shock wave will die out as itpasses through an unreacting gas, thevelocity will decay until it becomesequal to the velocity of sound. For achemical reaction that proceeds to completion,the energy generated in the re-VIII-117

action maintains the velocity constant.This additional constraint on thesystems allows for the calculation of aunique detonation velocity, togetherwith a specific volume and a temperaturedepending only on the composition of thegas and the original conditions ofpressure, volume, and temperature. Thiscalculable condition after complete reaction,is known as the Chapman-Jouquetstate. This is not yet the final stateof the gas; pressure equilibration hasnot yet taken place, and no heat hasbeen exchanged with the surroundings.According to this greatly simplifiedpicture, the events transpiring as thedetonation wave passes are approximatelyas follows: a shock wave, traveling atseveral times the velocity of sound,causes a very brief and very sharp increasein pressure, density, and temperature.It is caused by the net forwardmovement of molecules in the shockfront. This net movement of moleculesresults in an impulse greater than thatrepresented by the very considerablepressure spike. Immediately behind theshock front is a region of hot gasundergoing a prereaction "induction"period. During this period the concentrationof reaction-causing free radicalsis building up, but little energyis being released. At the end of theinduction period the reaction rate increasesrapidly and proceeds toward thecondition of chemical equilibrium. Finally,the wave passes and the reactionproducts assume a final equilibriumstate that marks the beginning of heatexchange with the vessel.The simple theory sketched out above isnow known to be entirely inadequate asan explanation of what happens in thedetonation process. The detonation waveis not a simple plane discontinuity. Itis, in fact, highly turbulent and isassociated with transverse waves, thestructure of which depends on the shapeand size of the vessel. Neverthelessthe older theory has been successful inpredicting detonation velocities andother paramters of the system. It isused here to give at least a reliableorder-of-magnitude estimate of the effectof a detonation under assumed conditions.D2.2 CALCULATIONS OF DETONATIONPARAMETERSThe first step in the calculation, asoutlined, for example, by Lewis and vonElbe (Ref. 3), is the determination ofthe conditions at the Chapman-Jouquetplane. In addition to the specificationof the starting conditions of composition,pressure, specific volume and temperature,this step requires that thefollowing be known: the average specificheat of the final gas through thetemperature range, the ratio of specificheats (specific heat at constantpressure divided by the specific heat atconstant volume) of the final gas at thefinal temperature, and the heat of reaction.All of these, but especially thelast, depend on the final composition,taking note of the fact that at elevatedtemperatures the partial pressures of H,0, and OH will not be zero. Aniterative process is required; here onlya few iterations are needed sincetemperatures were not extreme and highaccuracy was not required. There issome question whether the time availableis sufficient to fully excite the vibrationalmodes. It was assumed thatsufficient time is available. The specificheats were therefore taken as thestandard values without modification(Ref. 4).Having determined the velocity of thedetonation, it is possible to calculatethe conditions in the initial shockplane, as outlined, for example, byKistiakowsky and Kydd (Ref. 5). Thesecalculations require the knowledge ofthe ratio of specific heats of theunreacted gas averaged over thetemperature range. Again the questionof the excitation of energy levelsarises. The time is too short forvibrational modes to be excited. It wasassumed, however, that rotational levelsare excited. Therefore, the specificheatratio was taken as 1.4 for thediatomic molecules, and 1.33 for H 2 0.The final value for the impulse musttake into account the net forward motionof molecules in the shock wave. Themomentum of these molecules passing unitarea in unit time is w2/v 5 , where w isthe net forward velocity (particlevelocity) and v is the specific volumein the shock 2ront. According tohydrodynamic theory, w 2 is calculated by2w = (v 1 - vs) (ps - pl), (VIII D-1)from the specific volume and pressure ofthe original and the shock-front gas.Therefore the pressure equivalent, thatis, the rate of momentum transfer, isw2 v 1 -v(Ps - P (P - Pl)(VIII D-2)VIII-118

If thisincrease,becomesv 1is added to the pressureps - Pl, the total impulse(Ps PS - pl)" (VIII D-3)From this value, the effect of theinitial shock can be estimated.Finally the gas reaches a state ofequilibrium, which for an adiabaticprocess, is found from the heat ofreaction, the average specific heat atconstant volume, and the gas laws. Inthis calculation the equilibriumdecomposition of water was neglected.The results of these calculations forthree assumed compositions are shown inTable VIII D-2. The starting compositionswere selected by adding to air at10 psia enough water vapor to saturateat 130 F and enough H 2 to approximateShapiro and Moffette's assumed leandetonation limit (Case I), the stoichiometriccomposition (Case II), andShapiro and Moffette's assumed richdetonation limit (Case III). Theresults are computed for idealizedlimiting circumstances which do notnecessarily correspond to the conditionsexpected during an accident. Turbulenceand transverse shock waves will resultin many variations in temperature,pressure, and density. Obstructions inthe vessel will cause perturbations. Ifa low-speed flame were to persist forsome time before the transition todetonation, a considerable pressureincrease could occur in the unburned gasbefore detonation, thus leading to amore severe shock (Ref. 6). The fullcalculated impulse is delivered to thewall only if the direction of propagationis normal to the wall.The core-meltdown analyses described inAppendix A predict the extent ofzirconium-water reaction to be 75 + 25percent of Zircaloy in the PWR cladding.Case I corresponds to approximately 100percent Zircaloy reacted, or the upperlimit of the Appendix A calculations.Case II, the stoichiometric composition,would require the equivalent of a 150percent Zircaloy-water reaction. Whilethis may be physically realizable fromthe reaction of all the cladding incombination with contributions fromiron-water, U02-water, and other reactions,it is considered quite unlikely.Further, the reaction of the hydrogengenerated in Case II would require theconsumption of all the oxygen initiallypresent in the containment building.Case III, the assumed rich detonationlimit, is not believed attainable underthe reactor-accident conditions ofinterest here.The quasiequilibrium pressure resultingfrom the detonation or rapid selfpropagatingcombustion of the hydrogenfrom an equivalent 75 percent claddingreaction would produce peak pressures of71-77 psia within the PWR containmentbuilding. At these levels, containmentfailure would not be expected. Theeffect of the initial impulse of theshock wave is more difficult to assess.It will depend on the duration as wellas the magnitude of the impulse. In thenext section the time scale is examined.D2.3 DURATION OF THE SHOCK WAVESAn examination of the literature onhydrogen-oxygen detonations shows thatan enormous research effort has beenexpendedprocess.on various aspects of thisAlmost all of it has been donein long narrow tubes at reduced pressure,diluted, if at all, with inertgas. It has not been possible to findspecific information on the time requiredfor the detonation process underconditions similar to those of the presentstudy. Since detonation velocitiesare of the order of thousands of metersper second, the resolution of measuringinstruments must be of the highestorder. It is for this reason that lowpressures and high dilutions have beenso widely used. Nevertheless, with somejudgment and some extrapolation, reasonableestimates can be made at leastof the order of magnitude of the timescale.Bradley reports some determinations ofthe thickness of shock fronts leading toan estimate of about 6Lo, where Lo isthe mean free path in the undisturbedgas (Ref. 7). The thickness of theshock front would then be about 4 x10-7 m for Case I, and the duration ofthe initial shock, about 3 x 10-10 sec.The average molecule would remain in thefront somewhat longer than that becauseit is swept forward with the shockfront. One estimates, for this casethen, that the average molecule wouldundergo approximately 100 collisions asthe shock front passes by. This lengthof time is of the same order ofmagnitude as that required forrotational excitation, but is too shortfor the vibrational levels. This seemsmore reasonable than the 2 or 3 meanfree paths calculated by Thomas (Ref.VIII-119

8). Experimental evidence on weakshocks in N 2 at higher pressures alsoindicated thinner fronts, although notso thin as indicated by the Thomascalculation (Ref. 9).It is concluded that a reasonable valuefor the duration of the original shockwould be 3 x 10-10 sec.The detonation wave is being drivenforward by the formation of hot gas inthe reaction immediately following. Thepressure in the shock front decays notto the original ambient pressure, but tothe pressure corresponding to the burnedgas, that is, to the Chapman-Jouquetcondition. The time duration of thisperiod, during which the reaction goesto completion, was estimated from themeasurements of the reaction rates inthe hydrogen-oxygen system.The reaction normally proceeds by way ofan initiation step in which the freeradicals, H and OH, are formed, followedby a chain-branching step in which thenumber of these radicals is multiplied.In this part of the reaction, known asthe induction period, little H20 isformed and little heat is produced.When sufficient chain carriers areformed, they begin to react to form H 2 0,thus raising the temperature and acceleratingthe reaction. Throughout theinduction period and the main reactionperiod chain-stopping reactions involvingthird bodies such as N 2 or H 2 0 serveto slow the process. The chain-stoppingstep involves the formation of a metastableradical, HO?. At high temperatures,the induction period is probablynegligible since the H 2 0 originallypresent will decompose and form H and OHand since the chain-stopping HO 2 becomesentirely unstable. Near the combustionlimits and under marginal conditions,however, the length of the inductionperiod may determine whether or not thedetonation is propagated.The induction period has been studied inshock tubes at relatively low pressures,and with inert gas, but not H 2 0, as adiluent. The results are usually shownin straight-line plots of log 102] T. vsT 1 , where (021 is oxygen concentration,Ti is the induction time and T is thetemperature. Using Case I as an example,these results indicate 0.8 to 1.6-usec induction times (Refs. 7,10,11).These values are determined by a considerableextrapolation, however, since thework was done at low pressures, whereTi was considerably greater.Kistiakowsky and Kydd (Ref. 5) measuredthe density changes during the detonationreaction and interpreted these interms of the time for 0.50 or 0.75 ofthe reaction to be completed. Theyfound no evidence of an inductionperiod. Their work was also done at lowpressures. Extrapolating their resultsto atmospheric pressure gives 0.26 l.secfor a stoichiometric mixture, 1.7 psecfor a mixture containing 14 percent H 2and 30 percent 02. Dove and Tribbeckdevised a computational method forstudying the duration of the reaction(Ref. 12). They give reaction times,for the undiluted mixture at 1 atm of0.6 Psec for 16 percent H 2 , 0.02 psecfor a stoichiometric mixture, and 0.2Psec for 92 percent H 2 . They reportthat experimental reaction times atlower pressures are 5-10 times longer.In summary, reaction times in the range0.1 to 10 3isec seem reasonable. Interms of its effect on the reactiontime, water will probably act only as adiluent (Ref. 5). Because of the likelihoodof compositions somewhat removedfrom stoichiometric, together with thepresence of N 2 and H 2 0 diluents, thetimes are likely to be closer to theupper end of this range.D2.3.]APPLICATION TO A LARGE VESSELMost of the experimental work which underliesthe discussion above was done inshock tubes. Pressures were generallyof the order of 0.1 atm or less, and thediluents, if any, were N 2 , Ar, or He.Since the present case differs in severalrespects, some attention may be givento each aspect of the difference.Detonation in a tube represents, atfirst glance, a one-dimension process.The process is too rapid for any appreciableheat-transfer to the walls. Itis now known, however, that the structureof transverse waves which accompanythe longitudinal wave is dependent onthe tube walls. This complex system ofshock waves probably plays an importantpart in the course of the chemicalreaction, especially near the detonationlimits. In a freely expanding detonation,the transverse-wave structure willbe considerably different. Struck andReichenbach observed no turbulencebehind the combustion zone, no cellularstructure in the wave front, and notransition from deflagration to detonationexcept by interaction with the wall(Ref. 13). Soloukhin and Ragland observedmore evidence of structure, butalso noted the difficulty of establishingthe detonation in a freely expandingshock wave (Ref. 14).VIII-120

It has been necessary to use data obtainedat subatmospheric pressure oftenwith no inert ingredients (scarcely anyin the presence of added water vapor) toevaluate detonations at pressures aboveatmospheric, with N 2 and H 2 0 present.With respect to the properties of thedetonation wave, its velocity, pressure,and duration, these extrapolations arereasonably reliable. The theory onwhich calculations were based has notproved adequate in a descriptive sense,but it has been successful in predictingdetonation parameters. More uncertaintyis involved with respect to the limitsof detonation. The presence of watervapor may serve to quench the incipientreaction by providing an efficientchain-stopping mechanism. Neither theorynor experimental evidence is adequateto establish the detonation limits veryprecisely under the free-expansion conditionsthat exist in the containmentvessel. Furthermore, the conditionsnecessary for initiating a detonationunder these circumstances is not wellknown. It seems likely that there is aconsiderable composition range in whichdetonation is possible, but detonationis highly unlikely because of theseverity of the conditions needed forits initiation.Nevertheless, if it is assumed that theinitiation occurs and the detonationproceeds, the effect on the containmentstructure can be estimated as followsusing the data in Table VIII D-2 and theestimated shock-duration times. Thenatural period of vibration of typicalcontainment structures is much greaterthan the duration of a detonation shockwave, so the loadings can be approximatedby an impulse loading. Newmark hasdeveloped a convenient empirical formulafor such use which gives answers to + 5percent accuracy (Ref. 15). For a brittlematerial (the reinforced concretewall) the formula is,pt = 0.32 RT,(VIII D-4)where p is the maximum pressure in theimpulse, t is the time of thedetonation, R is the loading thatproduces the maximum elastic deflectionfor the structure, and T is the naturalperiod of vibration. Letting R = 120psia (i.e., twice the design pressurefor a typical PWR (containment vessel)and T = 0.02 second, the estimatednatural period for a typical reinforcedconcrete PWR containment vessel (Ref.16), the right side of the aboveequation equals 0.77. Therefore, if theproduct, pt, is less than this value thestructure will not fail under theimpulse loading of the detonation shockwave. Using the pressure impulse valuesgiven for the three composition cases inTable VIII D-2 along with the estimatedupper limit value of 10 Usec for theduration of the impulse, the followingpt values are obtained:Case p t, secIIIIII12001331172710-510-510-5pt0.0120.0130.017The pt values are well below the allowablelimit of 0.77 so failure of thecontainment due to the detonation shockwave is not expected.The above shock-wave analyses were performedfor the same idealized limitingcompositions as considered previouslyfor the evaluation of the quasiequilibriumpressures. For the actualquantity of hydrogen predicted to beavailable, the possibility of containmentfailure would be even less.Since neither the shock wave nor theensuing quasiequilibrium pressure of thedetonated gas mixture approach thelevels required for failure, the probabilitythat a hydrogen detonation willresult in PWR containment failure can beconsidered negligible.D2.3.2HYDROGEN IN A BWRThe influence of hydrogen production inaccident sequences involving core meltdownin a large BWR is quite differentfrom that in a large PWR. In sequenceswhere the ECC systems do not operate,the primary coolant blowdown leaves thepressure vessel essentially dry so noappreciable metal-water reaction hydrogenwill be produced during initial coremeltdown. As the core melts through thepressure vessel and falls to thecontainment floor, it will come intocontact with water and appreciable hydrogengeneration may result. Forsequences where the core meltdown occurswith water in the vessel, hydrogenproduction will be similar to PWR cases.If the containment fails prior to coremelting, the subsequent generation ofhydrogen will not be of concern.BWR containment atmospheres are inertedby purging with nitrogen until theoxygen content is below 5 percent. Theoxygen is maintained at or below thislevel during normal operation. At thisoxygen concentration, the containmentVIII-121

atmosphere is outside hydrogen flammabilitylimits, regardless of thequantity of hydrogen in the atmosphere(Ref. 18). Under these circumstances,the primary significance of hydrogen isas a noncondensable medium for containmentpressurization. Since the freevolume of BWR containment is muchsmaller than that of PWR's, the relativesignificance of hydrogen as a pressuresource is much greater in BWR's.Radiolytic decomposition of water duringthe course of the accident would addoxygen as well as hydrogen to the atmosphere.Maximum accident times ofinterest here are of the order of a dayor less. Using an integrated fissionproduct decay energy over one day or 14x 1030 ev, an energy absorption factorof 0.1 in the core region, and the upperlimit G(H 2 ) value of 0.44, the maximumyield of hydrogen at the end of thisperiod would be about 8100 SCF or 23 lbmoles.The radiolytic oxygen yieldwould be a factor of two lower.Neglecting any hydrogen from the zirconium-waterreaction, the above wouldresult in hydrogen and oxygen concentrationsof 2.8 and 6.5 percent, respectively.This composition is outside theflammability range. If the hydrogengenerated by the zirconium-waterreaction is included, the atmospherewill be hydrogen rich and outside theflammable range.Thus it is concluded that hydrogencombustion can be eliminated as amechanism of containment failure in BWRdegraded accident sequences.D2.4 CARBON DIOXIDE GENERATION INPWR AND BWRCarbon dioxide can be generated in areactor-meltdown accident through thermaldecomposition of concrete containinglimestone aggregate according to thereaction,CaCO 3 + heat - CaO + C0 2 +. (VIII D-5)pressure vessel and falls onto theconcrete floor and base mat of thecontainment structure. The CO 2 generationby itself probably would not besufficient to cause overpressurizationof either a PWR or a BWR containment,since other failure mechanisms areeffective competitors. However, as CO 2can contribute to overpressure failureunder certain conditions, the generationcannot be ignored. The containment responseanalysis takes all the processesinto account; thus, only the basicinformation needed to consider the CO 2generation mechanism is provided here.Reimer and Seidenfeld have presented auseful documentation of the problem andthe pertinent basic data given below istaken from their work (Ref. 17). Thecomposition of the concrete that hasbeen used in this study is presentedbelow. Although this characterizationdoes not represent a typical concrete,it provides an upper bound to thequantities of gases (H 2 0 and C0 2 ) thatmight be released during thedecomposition of concrete.MaterialCaCO 3H 2 0SiO 2weight Percent602020Concrete apparently will begin torapidly disintegrate and spall at about900 F as the free water expands, rupturesthe surface, and escapes. Theyield of CO 2 when the CaCO 3 decomposescan be given in various units as followsbased on a concrete density of 150 lb/cuft.COYield per Unit Amount of Concrete3325 ft CO 2 (STP) per ft concrete30.905 lb-mole CO 2 per ft concrete0.264 lb CO 2 per lb concrete3Standard texts indicate that thedecomposition of CaCO 3 becomes rapid ata temperature of about 800 C (1470 F).Therefore, this represents a latersource of gas in the accident when themolten core material penetrates theThe energy needed to decompose calciumcarbonate is given as 42,500 cal/g-moleCaCO 3 . This converts to a value of 460Btu/lb concrete. Other energy requirementsto heat the concrete materials tothe decomposition temperature and aboveVIII-122

can be determined with the following fusion of these two materials is givenlist of heat capacities. as:Material Heat Capacity Material Heat of FusionCO 2 0.26 Btu/lb/F SiO 2 64.5 Btu/lbH20 0.58 Btu/lb/F CaO 578 Btu/lb'CaO + Si0 2 0.22 Btu/ib/FContinued energy input after the calcium These physical parameters can be used tocarbonate is decomposed can result in evaluate the gas yield and the thermalmelting of first SiO 2 (MP = 3100 F) and energy requirements of the concretethenCaO (MP = 4400 F). The heat of decomposition problem.References1. Shapiro, Z. M., and Moffette, T. R., "Hydrogen Flammability Data and Applicationto PWR Loss-of-Coolant Accident", WAPD-SC-545 (September, 1957).2. Miles, J. E. P., Munday, G., and Ubbelohde, A. R., "Effects of Additives onMarginal Detonation in Gases", Proc. Roy. Soc., 269, 165-179 (1962).3. Lewis, Bernard, and Van Elbe, Guenther, "Combustion, Flames and Explosions ofGases", Second Edition, Academic Press, p. 524 (1961).4. JANAF Thermochemical Tables, Second Edition, NSRDS-NBS 37 (1971).5. Kistiakowsky, G. B., and Kydd, P. H., J. Chem. Phys., 25, 824-835 (1956).6. Brinkley, S. R., Jr., and Lewis, B., Symp. Combust, p. 807 (1958).7. Bradley, John, N., "Shock Waves in Chemistry and Physics", John Wiley & Sons(1962), Chapter VII.8. Thomas, L. H., J. Chem. Phys., 12, 449 (1944).9. Cowman, George R., and Harnig, Donald F., J. Chem. Phys., 18, 1008-1018 (1950).10. Schott, G. L., and Kinsey, J. L., J. Chem. Phys., 29, 1177 (1958).11. Mullaney, G. J., Kw, Peh S., and Botch, W. D., AIAA Inl., 3, 873 (1965).12. Dove, J. E., and Tribbeck, T. D., Astronautica Acta, 15, 387-397 (1970).13. Struck, W. G., and Reichenback, H. W., Symp. Combust., II, 677-82 (1966).14. Soloukhin, R. I., and Ragland, K. W., Comb. and Flame, 13, 295-302 (1969).15. Newmark, N. M., "An Engineering Approach to Blast Resistant Design", Proc ASCESeparate No. 306, Oct. 1953.16. Morrison, D. L., et al., "An Evaluation of the Applicability of Existing Data tothe Analytical Description of a Nuclear Reactor Accident - Core MeltdownEvaluation", BMI-1910, p. C-19, Table C-5 (July, 1971).17. Reimer, R. A., and Seidenfeld, A., "Investigation of the Consequences of aNuclear Reactor Core Melt-Through Accident", ORNL-MIT-62 (October, 1968).18. Coward, H. F., and Jones, G. W., "Limits of Flammability of Gases and Vapors",Bulletin 503, U. S. Bureau of Mines (1952).VIII-123

TABLE VIII D-1PWR HYDROGEN SOURCESEstimate of H, Period of EstimatedHydrogen Generation Rate, Generation, Yield in Period,Source lb-moles/hr hr lb-molesCore-metal/ (a)water reaction 800 0.75 600Radiolysis ofwater(5) 1.9 24 46Corrosion ofmetals by spraysolutions(c 0.15 48 7(a)(b)Based on core zirconium content and thermal analysis of meltdown period.Simple average based on decay integrated over one day, 20 percent absorbed, andG(H 2 ) = 0.44. The rate decreases with time.(c) Based on corrosion of aluminum paint by alkaline spray. Rate will decrease bymore than ten for longer times.TABLE VIlI D-2CALCULATED DETONATION PARAMETERS FOR LIMITING COMPOSITIONSComposition, Specific Energy of Detonationvolume percent Pressure impulse Temperature Volume, Combustion, VelocityCase H 2 0 H 2 02 N 2 OH H/m 2 psia N/0 2 psia K F m 3 /kg J/kg m/secI (a) 15.0 18.0 14.1 52.9 1.029 x 105 14.9 327.6 130 1.182(b) 1.906 x 106 276 8.30 9 10 6 1200 1360.0 1990 0.269(c) 36.1 5.5 58.2 0.31 1.072 x 106 156 2175.0 3454 0.684 1.9325 x 106 1650(d) 36.2 5.6 58.2 6.30 x 105 91 2008.0 3156 1.182II (a) 13.4 25.6 12.8 48.2 1.130 x 105 16.4 327.6 130 1.024(b) 2.050 x 106 297 9.19 x 106 1331 1398.0 2930 0.216(c) 44.4 0.1 55.3 0.32 1.469 x 106 213 2784.0 4550 0.583 2.998 x 106 1797(d) 44.7 55.3 9.06 x 105 131 2625.0 4275 1.024III (a). 13.0 40.0 10.3 38.7 1.407 x 105 20.4 327.6 130 1.026(b) 2.690 x 106 390 1.190 x 101 1727 1340.0 1952 0.219(c) 35.1 21.6 43.2 0.10 1.670 x 106 242 2497.0 4035 0.590 2.926 x 106 1925(d) 35.2 21.6 43.2 9.75 x 105 142 2269.0 3625 1.026(a)(b)(c)(d)Original gasShock frontC-J planeFinal gasTable VIII D-1 - Table D-2VIII-125/126

Flammability Limits75 F - 0 psig---- 0---- 300 F - 0 psig-- &--- 300 F -100 psig10% reaction containerambient temperature 216 F100/o reaction container100 '/o reaction ambient 280 Fambient 216 F 00 %l reaction container40 ambient 280 F• "d e t o a m m ai l i t m i _ t s _0000 80 60 40 20 0 0Hydrogen, v/oFIGURE ViIl D-1 Flammability Limits of Hydrogen-Air-Steam Mixtures (Ref. 1)k

3.000 Hydrogen Concentration for75 Percent Zr ReactionStoichiometric;ombustion and)etonation Limitsrom Figure D-1270 F010080 60 40 20 0Hydrogen, v/o100FIGURE Viii D-2Composition Isotherms for Containment Atmosphere at Saturated SteamConditionsFig. VIII D-1 - Fig. VIII D-2VIII-127/128

APPENDIX ECONTAINMENT FAILURE MODES EVALUATIONVIII-129

Appendix ETable of ContentsSEectionEl. SUMMARY AND CONCLUSIONS .........................................E2. DISCUSSION: PWR CONTAINMENT ....................................E3. DISCUSSION: BWR CONTAINMENT ....................................Page No.VIII-133VIII-133VIII-136Paart I ................................................................1. GENERAL .........................................................1.1 Loading Conditions at Accident .............................1.2 Design Basis ...............................................2. EFFECT OF INCREASING INTERNAL PRESSURE ..........................2.1 General ....................................................2.2 Conditions at the Yield Point of Reinforcement .............2.2.1 Summary for Conditions when Reinforcementat Yield ............................................2.3 Conditions at the Ultimate Strength of theReinforcement ..............................................2.4 Ultimate Strength of the Concrete in Shear .................2.5 Ultimate Strength of the Liner .............................3. EFFECT OF RISE IN INTERNAL TEMPERATURE ..........................4. CONCLUSION ......................................................LIST OF REFERENCES .....................................................LIST OF FIGURES ........................................................VIII-139VIII-141VIII-141VIII-141VIII-141VIII-141VIII-142VIII-143VIII-143VIII-145VIII-147VIII-148VIII-148VIII-150VIII-150Paart II ................................................................1. DESCRIPTION OF CONTAINMENT FAILURE ANALYSES .....................1.1 Reinforced-Concrete Containment Structure ..................1.2 Steel Containment Structure ................................1.2.1 Computer Calculation for AxisymmetricPressure Vesses .....................................1.2.2 Burst Pressure Calculations for IndividualComponents ..........................................2. DISCUSSION AND RECOMMENDATIONS ..................................2.1 Reinforced-Concrete Containment Structure ..................2.2 Steel Containment Structure ................................REFERENCES .............................................................VIII-153VIII-156VIII-156VIII-157VIII-157VIII-158VIII-158VIII-158VIII-159VIII-160VIII-131

List of TablesTableVIII E-l Estimated Containment-Building Failure Pressures ................VIII E-2 Distortion Energy Failure Estimates .............................VIII E-3 Summary of Calculated Stresses in Typical Vapor-Suppression Containment for Boiling-Water Reactors ..............Page No.VIII-161/162VIII-161/162VIII-161/162VIII E-4VIII E-5Summary of Estimated Burst Pressures for Components ofa Typical Vapor-Suppression Containment for Boiling-WaterReactors ........................................................Summary of Computer Analysis Used for Vapor-SuppressionContainment for Boiling-Water Reactors ..........................VIII-161/162VIII-161/162List of FiguresFigurePage No.VIII E-1 Axisymnietric Model for Steel Shell Computer Analysis ......VIII-136/164VIII-132

Appendix EContainment Failure Modes EvaluationE.1 SUMMARY AND CONCLUSIONThe objective of this subtask was todetermine the thresholds and modes offailure for reactor containment buildingswhen subjected to internal pressuresand temperature greater than thedesign values. Containment structuresrepresentative of boiling as well aspressurized water reactors were consideredwith the intent of determining the,conditions at which failure would occur,identifying the probable failure locations,and defining the nature and sizeof the most likely failures.Subatmospheric containment of reinforced-concreteconstruction was consideredfor the PWR, and the light bulb andtorus vapor-suppression arrangement wastaken to be representative of BWR containments.A specific containmentstructure of each type was selected foruse in this study in order to have ameaningful basis for evaluation. To theextent possible, the extension of theresults to other containments of similardesign has been indicated.Of necessity, the extent and depth ofthis evaluation were limited, withemphasis on the principal features,design criteria, and materials utilized.Nevertheless, a significant number ofaspects were considered for each of thetwo containment concepts in arriving atthe conclusions presented here. Torefine the conclusions, considerationwould have to be given to the as-builtdetails of specific containment structuresas well as to the results of testsand nondestructive examinations. Severalconsultants, both at BCL as well asexternal, were utilized to guide thisevaluation; their reports are presentedin Parts I and II of this Appendix. Theconclusions presented here are based onthe recommendations of the consultantsas well as consideration of other factorsand substantial additional analyses.Subject to the limitations and conditionsdiscussed below, the conclusionsof this study are as follows:a. A reinforced-concrete containmentbuilding designed for an internalpressure of 60 psia (45 psid) whensubjected to slow overpressurizationby steam would be expected to failat a pressure of 100 + 15 psia. Thefailure would be initiated by yieldingof the reinforcing steel and/orcrumbling of the concrete, and leadto a gross failure of the structure.Such a falure would take place inthose parts of the structure withoutany radial reinforcement of the concrete,e.g., upper portions of thecylindrical shell and/or the hemisphericaldome.b. Vapor-suppression containment of thesteel drywell and torus arrangementdesigned for an internal pressure of71 psia (56 psid) when subjected tooverpressurization would be expectedto fail at a pressure of 175 + 25psia. Typically, the weakest portionin such a structure would be inthe upper part of the toroidal suppressionchamber; other highlystressed locations are the toroidalknuckle between the spherical andcylindrical parts of the drywell,the thinner cylindrical parts of thedrywell, the thin part of thesphere, and the expansion joints onthe drywell-wetwell vents. Thefailure would be expected to startat a weld or a discontinuity such asthe attachment of a structuralsupport, penetration, or embedment.As applied in this study the failurepressure is not a single discretevalue, but a continuous variablewith a normal distribution about thenominal value. This approach recognizesthat the probability of structuralfailure is small at loadsslightly above design, but increaseswith increasing loading. By definition,the probability of failure atthe nominal failure pressure is 0.5;it approaches unity as the loadingapproaches the ultimate strength ofthe structure.E.2 DISCUSSION: PWR CONTAINMENTSubatmospheric containment of reinforced-concreteconstruction was assumedto be representative of pressurizedwater reactor practice. Typically, sucha structure has a flat base mat, avertical cylindrical shell, and a hemi-VIII-133

spherical dome; a thin steel liner onthe inside provides a leaktight barrier.The design was assumed to be in accordancewith the ACI code for reinforcedconcretestructures; the materialsutilized are representative of currentpractice. In the prediction of failurethresholds the materials propertiesutilized are expected values rather thanthe minima specified by the variouscodes.The determination of the failure thresholdfor reinforced concrete reactorcontainment buildings requires consideringthe complex interaction among theliner, concrete, and the reinforcingsteel. The strength of the liner islow, but its integrity must be maintainedif the structure as a whole is tofunction. Should the liner lose itsintegrity, the reinforced concrete,though structurally intact, will not beable to maintain internal pressures dueto porosity and cracking of the concrete.The concrete, though it may becracked, serves to transmit the internalpressure loads to the reinforcing steel.Deterioration of the concrete willdefeat the function of the reinforcingsteel and lead to failure of the structure.The reinforcing steel is the mainload-bearing element in the structureand its failure is synonymous with thefailure of the structure as a whole.The determination of the failure thresholdof a reinforced-concrete containmentbuilding requires the definition ofthe conditions at which the three mainelements of the structure cease to functionas a whole.The steel liner serves as the leak-tightbarrier for the containment. It has arelatively low strength in comparisonwith that of the reinforced concrete anddepends on the latter for support. Noallowance is given to the strength ofthe liner in the design of the structure.The liner material is typicallymuch more ductile than the reinforcingsteel; thus in a gross sense theconcrete behind the liner would have tofail first. Careful attention, however,must be paid to the details of the linerinstallation and anchoring, particularlyat discontinuities (such as the buildingwall-to-foundation joint and penetrations)to ensure that the liner candeform without failure under all designloadings. In the designs *considered,this objective appears to have beenachieved.The ideal ultimate strength of the containmentstructure considered is about140 psia, based on the tensile strengthof the reinforcing steel and liner. Atthis loading, failure of the structurewould be a certainty. It is recognized,however, that the achievement of theultimate strength in all parts of thestructure may not be possible. Accordingly,the possibility of failure atless than the ultimate loading must beconsidered.The design of reinforced-concrete containmentstructures is based on a safetyfactor of 1.5 on the yield strength ofthe r~einforcing steel with an additional10 percent reduction in working stressesto allow for nonuniformity in materialsand workmanship. This results in anominal safety factor of 1.67 on theyield strength of the reinforcing steelat the design loading. No allowance isgiven to the strength of the steel linerin determining the quantity of reinforcing,although the latter should be includedin the evaluation of the ultimatestrength of a structure. In Part I,Mast suggests that the ultimate strengthof reinforced-concrete containment willbe reached when the reinforcing steel isat its yield point. The large strainsassociated with stresses beyond yieldwill result in the separation of thesteel from the concrete and loss of anystrength from aggregate interlock as aresult of cracks in the concrete. Thefailure would be expected to take placein those regions which lack radialreinforcement - typically the upper portionsof the cylindrical shell as wellas the hemispherical dome. For a structuredesigned for 45 psid this criterionleads to an ultimate strength of 75psid. Adding to this the yield strengthof the liner yields a maximum loading of92 psid. Based on the criterion suggestedby Mast, failure of the structurecould thus be expected at an internalpressure of 107 psia.The evaluation given in Part II arguesthat due to the overreinforced characterof typical reactor containment structuresthe concrete will fail while thesteel is at relatively low stress levels.That is, the differences in stresslevels between the inner and outerlayers of reinforcing due to the thermaleffects and geometrical constraints willbe greater than can be accommodated bythe concrete as the loading exceedsdesign levels. If the concrete cannoteffectively distribute the internalpressure loads among the various layersof reinforcing, most of the load will betaken up by the innermost layers ofreinforcing and the containment liner.Using this line of argument and assumingthat failure will take place approximatelyhalfway between the yield andultimate strength of the load-bearingVIII-134

liner- reinforcing steel combination, afailure load of 75 psid corresponding toan internal pressure of 90 psia isestimated.The value of 90 psia for the failurepressure must be considered conservativesince it accounts for only part of thereinforcing provided. The 107 psiafailure pressure should be more realistic.Both these values are considerablybelow the ideal ultimate strength of thestructure noted above. In view of thephysical bases for these two values, thenominal failure loading is taken as theapproximate mean of the two.In addition to the strength of thestructure as a whole, the effects ofdiscontinuities, penetrations, and otherfactors have been considered. Smallpenetrations are normally placed so asnot to interfere with the pattern ofreinforcing steel and thus require aminimum of effort to achieve a sounddesign. The evaluation of the failurepotential of all the various smallpenetrations would require a detailedlook at each and is beyond the scope ofthis study. major penetrations such aspersonnel- and equipment-access hatchesrepresent significant perturbations thatcould result in potentially weak pointsin the system. Their net effect ishighly dependent on design details.Substantial additional reinforcing steelis provided to compensate for reinforcementinterrupted by these large openings;the resultant structure, intheory, is at least as strong as theunperturbed structure. Such additionalreinforcing steel, however, makes itextremely difficult to achieve soundconcrete pours and good bond strengthbetween the steel and concrete. In theabsence of a sound concrete matrix thestrength of the steel reinforcing cannotbe effectively developed, and weakeningof the structure will result. Suchconsiderations are difficult to quantify,at best, and cannot be evaluated inthe absence of specific details ondesign, quality of workmanship, andresults of nondestructuve examinations.Other factors bearing on the failurepotential of major penetrations includethe manner in which they are anchored tothe concrete and the reinforcing steel,the reinforcing of the liner around suchpenetrations, and the attachment of theliner to the penetrations. The significanceof these factors will obviouslyvary with the details of each design.A potential low-threshold-failure locationexists in some reinforced-concretecontainments near the apex of thehemispherical dome where the reinforcingbars end in a cylindrical ring. In suchdesigns the final closure is achieved bybolting a flat plate across the openingand utilizing a gasket seal. While thethickness of the containment liner willtypically be increased near such anopening, neither the flat plate nor theliner are necessarily backed by concreteor other rigid structures. Thus atpressures and temperatures appreciablyabove nominal design levels the gasketedsurfaces may distort and lead to leakagepast the gasket. Also, the gasket materialcould deteriorate at temperaturesand pressures above design levels. Suchfactors could possibly lead to substantialleakage at pressures lower thanthose required for gross structuraldamage. Again, the potential failurethreshold will be highly design dependent.As noted above, the design of reinforcedconcrete containments incorporates anominal safety factor of 1.67 on theyield strength of reinforcing steel.However, reductions of 10 percent in thestrength of individual components, e.g.,rebar splices are commonly permitted.Further, the large amount of reinforcingincluded in these structures often leadsto difficulties in obtaining sound concretepours. Thus it is possible thatlocally the strength of the structurecan be lower than would be predicted onthe basis of nominal design parameters.It can be expected, however, that highlylocalized weaknesqes, should they occur,can be accommodated by local yieldingand/or deformation without affecting theoverall integrity of the structure.As is typical practice with reactor containments,this structure has beenstrength and leak tested at loadingssomewhat in excess of design levels.Successful performance during such testingindicates that the design objectiveshave been met in practice.On the basis of the evaluations discussedabove as well as the input of theconsultants whose reports are given inParts I and II of this Appendix, it isconcluded that a typical reinforcedconcretecontainment building designedfor an internal pressure of 60 psia canbe expected to fail at 100 + 15 psia.The error band should encompassuncertainties as to the mechanism andthreshold of failure as well as expectedvariations in materials, properties, andworkmanship. Structural failure wouldbe expected in the upper portion of thecylindrical shell or hemispherical domeas a result of yielding of the reinforcingsteel and/or shear failure of theVIII-135

concrete matrix. The eventual size ofthe failure is difficult to evaluate,but it would be expected to be largeenough to defeat the primary function ofthe containment, i.e., retention of thepressurized contents of the containmentatmosphere. Major penetrations and discontinuitiesin the structure can bepotential weak points; with careful attentionto design detail they can beeliminated.E.3 DISCUSSION: BWR CONTAINMENTBoiling water reactor containment wasassumed to be typified by a steel drywelland torus vapor-suppression arrangement.The nominal design pressurewas taken to be 56 psid and the designassumed in accordance with Section IIIof the ASME Pressure Vessel Code.Materials selected are representative ofindustry practice; materials propertiesutilized in failure evaluation areexpected values rather than codespecifiedminima.A computer stress analysis on a typicaldrywell vessel was performed at BCL andthe results compared with typicalresults presented in reactor-safetyanalysisreports. This analysis indicatedthe primary stress distribution inthe structure as a whole as well aspinpointed locations of high secondarystresses. The results of this analysisand comparison are discussed in Part II.The computer stress analysis did notinclude consideration of the suppression-chambertorus, the drywell top headand its closure, or the expansion jointson the drywell to wetwell vent pipes.Accordingly, additional detailed evaluationswere performed on these componentsand other selected details of the BWRcontainment structure to provide a consistentbasis for the prediction of afailure threshold..The above combination of analyses indicatedthe most highly stressed region tobe the inside upper part of thesuppression-chamber torus. This is aconsequence of the typically reducedwall thickness in this area combinedwith the geometrical configuration.Other highly stressed areas are thetoroidal knuckle between the sphericaland cylindrical parts of the drywell andthe upper part of the drywell sphere.In addition to the above areas havinghigh primary stresses, substantial bendingstresses are predicted for the followinglocations: the attachment of thedrywell to its support skirt, the embedmentof the drywell in the concretefoundation, and the conical transitioncylinders of different diame-betweenters.A number of factors have been consideredin developing a failure criterion forBWR containment of the type consideredin this study. These include the materialsof construction, expected qualitycontrol, type of loading, and constraintson the system. The materialsof construction are generally mediumcarbonsteels which are quite ductileand thus can undergo substantial plasticdeformation before failure. Thesestructures are designed and fabricatedin accordance with Section III of theASME Pressure Vessel Code, thus implyinga high degree of quality control in materialsand workmanship. The numerouspenetrations through the structure, bythe above code requirements, will bereinforced so as not to detract from theultimate strength of the structure as awhole. These considerations support theselection of the ultimate strength ofthe base material as the failure criterion;this is the criterion selected inthe evaluation of Part II to this Appendix.However, the conclusion that thestructure can maintain its integrityuntil the ultimate strength of the basematerial is reached implies that all thewelds have a strength and ductilityequal to those of the base material. Itfurther assumes that the structure caneverywhere deform as required to accommodatesecondary stresses such as localbending. It does not seem prudent toassume that the large number of fieldwelds required in the fabrication of thedrywell and wetwell structures can beaccomplished without any flaws or imperfections.Also, at internal pressuresabove the design levels the drywell vesselwill contact the concrete shieldingand other structures surrounding it. Onthe one hand, the concrete may tend tosupport the steel vessel; on the other,it will impede free expansion and thusmay lead to additional stresses and/orstress concentrations. Further, a numberof internal structural supports areattached to the walls of the drywell;these will impede the free expansion andmay act as stress raisers. It shouldalso be noted that the varying metalthicknesses and changes in geometry andoverall dimensions will result in nonuniformdeformations in various parts ofthe structure; this in turn will lead tovarying degrees of interaction with thesurrounding concrete and possibly producebuckling-type loads on parts of thestructure.If the ultimate strength of the structurecan be developed, the failure pressurewould be 250 psia as indicated in9VIII-136

Part II of this Appendix. As discussedabove, however, the potential for failureat lesser loadings must be recognized.It is concluded that a reasonablefailure criterion for *the containmentstructure in the present study is astress level in the base materialhalfway between the yield and ultimatestrength of the metal. Thus the bestestimate for the failure pressure for atypical drywell and torus vapor suppressioncontainment, designed for 56 psid,is 175 + 25 psia, with the failuretaking place in the top half of thetoroidal suppression chamber. Such arupture would not be expected to lead tostructural failure of the containment asa whole, but would be expected to besufficiently large to rapidly depressurizethe system. The mechanical propertiesutilized in this evaluation havebeen the expected properties of thematerials rather than the minima specifiedin the applicable codes. Whereappropriate, account has been taken ofthe expected changes in materialsproperties with changes in temperature.Potential weak points in the type ofcontainment considered here are theexpansion joints on the vent pipes fromdrywell to suppression pool, unlessspecific measures are taken to ensuretheir integrity at pressures significantlyin excess of design values. Suchmeasures would consist of limiting thetotal axial motion of the joint andconfining the flexible elements in sucha way as to maintain their prescribedgeometry. With such precautions, acommon practice in the design of reactorcontainment, the ultimate strength ofthe expansion joints should be comparableto that of the rest of the structure.In their absence, the expansionjoints could fail at about twice thedesign pressure, or at an internalpressure of about 125 psia within thecontext of the present discussion.The top closure of the drywell iseffected by a bolted head with sealinggaskets. Conceptually the head, flanges,and gaskets could be weak points inthe system. The head, since it is notbacked by concrete, is made extra strongso as to be able to resist localized jetforces. The bolted flanges are ofnecessity made quite rigid, and thecylindrical portion immediatey below theflange is typically heavier than otherparts of the structure. Thus internalpressure stresses near the gasketed topclosure are generally substantiallylower than in other parts of the system.It is thus concluded that the top heador its seal do not generally constitutea structural weak link in the system.At temperatures and pressures abovedesign levels, leakage past the gasketshowever, may be possible before failureof the structure.A steel drywell and torus vapor suppressioncontainment structure designed for56 psid could be expected to fail at aninternal pressure of 175 + 25 psia. Theerror band should encompass uncertaintiesas to the failure threshold as wellas expected variations in materials,properties, and quality of workmanship.The most likely location for the failureis in the inner top half of thesuppression chamber torus; additionalhighly stressed areas are the toroidalknuckle between the spherical andcylindrical parts of the drywell and theupper part of the drywell sphere.VIII-137

APPENDIX EPART IULTIMATE STRENGTH OF CONTAINMENT STRUCTUREbyPAUL E. MASTVIII-139

j/) f~)f/)/JENGINEERING BUILDING, SUITE 1819f'./ 205 WEST WACKER DRIVE~ 8~a~CHICAGO. ILLINOIS 60606BUS.: (312) 641-3675June 25, 1973RES.: 1312) 251-0450ULTIMATE STRENGTH OF CONTAINMENT STRUCTURE1. GENERAL1.1 LOADING CONDITIONS AT ACCIDENTIt is assumed that after the hypothetical accident thetemperature and the pressure will rise to 3400 and to120 PSI, respectively, both within approximately1000 minutes or 17 hours. The purpose of this studyis to predict the point of failure within thisperiod and to describe the failure mode.1.2 DESIGN BASISThe structure has been designed for a design pressureof 45 PSI associated with an internal temperature of1500. For these loading conditions, common reinforcedconcrete ultimate strength design principles wereapplied. According to the latter, the variouscomponents and materials of the structure at designlevel are stressed to approximately 60 per cent oftheir yield strength.2. EFFECT OF INCREASING INTERNAL PRESSURE2.1 GENERALDisregarding discontinuities, the structure isessentially a shell in a membrane state of stress.Therefore, increasing the internal pressure meansincreasing the membrane stresses, except the membraneshear stress. However, there will be an increase inradial shear stresses at discontinuities, commensuratewith the increase of all other stresses.VIII-141

Page - 2 -2.2 CONDITIONS AT THE YIELD POINT OF REINFORCEMENTThe stress in the reinforcing steel at design pressure(45 PSI) is approximately0.9 x 50 = 30 KSI.Therefore, with a yield stress of the main reinforcementof 50 KSI, the internal pressure at yieldamounts toP 1 = 45 50 - 75 PSI.30-The elastic strain in the reinforcement at this point isso-30,000 0.0017.The corresponding tension stress in the concretesurrounding the reinforcing steel at this point is waybeyond its tensile strength, so that the concrete mustbe assumed cracked. To estimate the crack width atthat stage, one can use the data obtained from thepressure test (Reference #2) which revealed cracksat about 18 inches on center. If one multiplies thestrain in the reinforcing bars with the crack spacing,one arrives at a crack width ofd =18 x 0.0017 = 0.03 inches.This is the crack width for which concrete shearcapacity due to aggregate interlock (surface roughness)was tested at MIT and Cornell University(Reference #1). Unfortunately, the tests wereperformed only for membrane shear strength but notfor radial shear strength. The difference lies inthe behavior of the reinforcement acting as dowels.Concerning membrane shear, practically all reinforcementperpendicular to the crack is engaged in fulldowel action.Concerning radial shear, dowel action of reinforcementnear the faces of the concrete walls is nil, becausethe reinforcing bars will cause the thin concrete coverto spall (see Figure #1).VIII-142

Page - 3 -2.2.1 SUMMARY FOR CONDITIONS WHEN REINFORCEMENT AT YIELDThe knowledge about reinforced concrete at the yieldpoint of conventional reinforcing steel is wellfounded. The mechanics of the failure mechanism arethe basis for the Ultimate Strength Design Method onwhich all modern Concrete Building Codes are based.The crack width in the tension zone is in theneighborhood of the above described 0.03 inches,which assures conventional radial shear capacitydue to aggregate interlock. Concerning in-plane(membrane) shear capacity, conditions aresignificently better because of the dowel actionof the reinforcement. Only reinforcement near thecenter of a section (see Figure #2) can be consideredeffective to develoo some radial shear capacity bydowel action.In view of the above, it ;s concluded that the subjectstructure can withstand an internal pressure of75 PSI with a uniform safety factor of 1.0 throughout.Based on the assumption that all components ofthe structure were equally designed by conventionalultimate strength principles, there should be no"weak link" up to this point.2.3 CONDITIONS AT THE ULTIMATE STRENGTH OF THE REINFORCEMENTThe reinforcement consists of ASTM A615-50 steel whichhas a yield strength of 50 KSI and an ultimate strengthof 80 KSI.It would be wrong to assume from the above figuresthat the vessel can withstand an ultimate internalpressure ofPu = 75 x 50 = 120 PSI,although it was stated before that at a pressure of75 PSI the reinforcement would just reach its yieldpoint (50 KSI).The reason why the above correlation is not true liesin the strains associates with steel stresses abovethe yield point. For the reinforcing steel specified,it may be assumed that the following strains prevail:VIII-143

Page - 4 -1. At the beginning of strain hardening: = 0.0152. At the end of strain hardening: = 0.0303. At ultimate strength (fu = 80 KSI): = 0.154. At fracture: = 0.20Looking at the above strains, one realizes that atthe end of strain hardening, the strains in thereinforcing and, thereby, the cracks in the concreteare roughly 18 times biqqer than at yield, but thatthe stress in the steel and thereby the resistanceof the reinforcing cage have really not increased.In other words, there is a state of transitionafter the reinforcement has reached its yield pointat which there is no increase in membrane tensilecapacity, but at which the concrete undergoesextensive cracking. The above established crackwidth of o.o3inches at the end of strain hardeningcould easily reachd = 0.03 x 18 = 1/2 inch +.Needless to say that no shear capacity due toaggregate interlock can be developed across a crackwidth. Therefore, to assure the integrety of thestructure, and the applicability of reinforcedconcrete design principles, one has to dependentirely on the radial shear capacity of thereinforcement, be it diagonal shear reinforcement,if any, or be it dowel action from bars perpendicularto the crack and positioned near the center of asection.In other words, at this stage of progressivecracking, the formation of weak links will be inareas of high radial shears with insufficientreinforcement to absorb these shear forces. It wasstated before that membrane shear in this investigationis no problem because of the radial symmetry of theloading conditions, and because of the diagonalearthquake-shear reinforcement, and because of theability of all orthogonal reinforcement to provideshear resistance by dowel action. However, concerningradial shear, the magnitude of the latter will increasein direct proportion to the internal pressure applied.VIII-144

Page - 5 -Furthermore, only the region near the base hasradial diagonal shear reinforcement, and thedowel shear capacity of the main reinforcementis neglible, because it is placed close to thefaces of the concrete walls.Reference #1 (page 32) shows that the slipresistance decreases with increasing crack width.In other words, the horizontal slip due to a shearforce applied to a precracked specimen increasedby 40 per cent when the pre-set crack width wasincreased from 0.02 in. to 0.03 in. It appearsjustified to extrapolate this result to theanticipated crack width of 1/2 inch at the end ofstrain hardening, and to evaluate the formation ofweak links on the basis of large cracks, associatedby large relative horizontal slip-type displacementsof concrete chunks. The structure, in this stage,can be visualized as a cage of reinforcing barsbulging in regions of high radial shear, such asthe vicinity of large penetrations (equipmenthatch: 14.5 feet diameter).Concerning public safety, it will be the linerintegrety which becomes critical in these areas atthis stage, before the structural integrety isjeopardized. Lateral liner deflections underdesign loads are usually limited to 0.01 inches.The strain imposed on a liner at wide cracks in theconcrete, together with the bending stresses due torelative transverse slip of the back-up concrete,most likely will cause the liner to tear. This, inall probability, will be the ultimate failure modeof the structure. The design documents (ReferenceNo. 1) described the liner to be of ASTM A-442,Grade 60 material, modified to A-300, with a yieldstrength of 32 KSI. No information concerningductility or ultimate strain at rupture was included,which would waive the possibility of the abovedescribed failure mode.2.4 ULTIMATE STRENGTH OF THE CONCRETE IN SHEARIn case the liner material should be capable ofenduring the above described longitudinal andtransverse (slip-type) strains, another possibilityof the formation of a weak link lies in the structuralfailure of the concrete in radial shear, per se.VIII-145

Page - 6 -Radial shear prevails at shell discontinuities,such as the boundaries at the top and at the bottomof the cylinder. Furthermore, radial shear prevailsat the larger penetrations. While the latter increasesin direct proportion to the internal pressureapplied, radial shears at the other discontinuities(top and bottom of cylinder), are relieved by theoverall yielding of the structure.Some radial shear reinforcement was providedat the top and bottom boundaries of the cylinder.Therefore, and because of the aforementioned selfrelievingbehavior, radial shear failure around thelarge penetration appears to be more imminent thanat the other boundaries.One should not be mislead by the fact that nocracks appeared around the penetrations during theloadtest. The observed behavior was underconditions well below the yield strength, and therebyirrelevant to the anticipated behavior afterextensive cracking of the concrete occurs. Ifthe design provided reinforcement around thepenetrations such that it would not yield beforeany other part of the structure breaks down, onlythen could one assume the penetrations to be nopotential hazard concerning failure in radial shear.It has become common practice in the design ofconcrete containment structures and concrete pressurevessels to provide the penetration sleeves withshear lugs. This was considered adequate toprevent a blow-out. In general, no welding to thereinforcing cage appeared necessary. With theinternal pressure acting on the penetration, the shearlugs will bear on the surrounding concrete and theywill prevent the sleeves from horizontal slip. Thelugs, however, are ineffective in preventing thesurrounding concrete from cracking under high strainsand thereby from failing in radial shear and beingblown out together with the penetration.To really appreciate the condition under yieldingreinforcing steel, the surrounding concrete at thisstage of failure must be considered to consist ofindividual chunks held by a grid of reinforcing bars.The steel strain of the reinforcing bars times theVIII-146

Page - 7assumed crack spacing would result in ultimate crackwidths ofdu=0.15 x 18 = 2.7 inches.A~t these strains, bond between reinforcing bars andconcrete has ceased to exist, so that the concreteis held by the reinforcing bars just like coarsegravjel in a sievie. The ineffectiveness of shearlugs on the penetration sleevies in that stage ofultimate strain becomes obviious. One could avoidthis potential weak link by a positive connectionbetween the penetration sleeve and the surroundingreinforcing cage, such as welded straps or hookedanchors.2.5 ULTIIMATE STRENGTH OF THE LINERIt has been pointed out before that a high-ductileliner material may prevent the liner to tear beforestructural failure occurs.Beneficial in this respect is. the rise in internaltemperature during the accident. The linerexpansion due to a temperature rise of about 2700is equivalent to a strain of onlyeS = 0.0017. Thisis by far not enough to compensate for the largestrains imposed on the liner by the yieldingreinforcing bars. The increased ductility of the linerunder high temperatures, howevier, may suffice toprevient its ovjer-all failure. Therefore, theductility of the liner under high temperature shouldbe specified by the manufacturer and be evaluated bythe designer, to check whether the liner will holduntil structural failure of the viessel occurs.The point of greatest weakness of the liner is itsanchorage to the rigid base (see Figure 15.5.1.8-1and Figure 15.5.1.1-1 of Reference No. 2). TheV'-611 long vjertical anchors rigidly tie the linerto the base, at a point where large deformationsmust be anticipated under the hypothetical accident.The curvjed transition at the base joint is beneficial.Whether it will suffice to allow the liner to followrotations and translations under gross deformations,can only be evaluated by a limit-design type failuremechanism analysis.VIII-147

Page - 8 -3. EFFECT OF RISE IN INTERNAL TEMPERATUREThe maximum temperature associated with thehypothetical accident is 340o. It is assumedthat this temperature prevails over the entire17 hours after the accident, until the structuralfailure occurs.The structure is built of 3000 PSI concrete. Suchconcrete is expected to withstand temperatures upto 2500 for an indefinite period of time withoutsignificent strength reduction. To provide asafety factor, most codes on concrete containmentvessels specify a maximum operational temperatureof 150 degree (Reference No. 3, paragraph 2.5.1.(a)).Over a short period of time, such as this hypotheticalaccident, an exposure to 3400 may be considered tohave no consequences because the other potentialfailure modes, as described above, are more seriousand likely to occur sooner. Under short-durationaccident conditions, exposure of the interiorsurface of the concrete shell to less than 350 deg.would even be permissible under common designpractices (Reference No. 3, paragraph 2.5.1.(b)).Exposure of the containment to 3400 from theinside will result in an average temperature of theconcrete walls of less than 150 degree. This isassurance enough that the very important shearcapacity of the concrete will not be impaired(Reference No. 4). If the hypothetical accidentwould be associated with temperatures of over350 degree, for this low-strength concrete acertain loss in shear strength would have to be takeninto account. The corresponding loss in modulusof elasticity, in compression strength, and intension strength would be of no significance forthis investigation4. CONCLUSION4.1. It is predicted that the vessel will fail at aninternal pressure of 80 PSI.At that stage, the yielding reinforcement hasundergone strain hardening and the concrete willshow 1/2 inch wide cracks. The liner will tear atthese cracks, and fission material will be released.VIII-148

Page - 9 -4.2. If the liner should be ductile enough to absorbhuge longitudinal and transverse strains, the containmentat this point uwill fail structurally in radial shearwhere no radial (diagonal) shear reinforcement isprovided to absorb the latter. Especially in viewof the lou-strength concrete (:3000 PSI), the concretesurrounding large penetrations is susceptible tosuch shear failure, This will cause a blow-out ofthe penetration, unless the penetration sleeve iswelded or otherwise anchored to the reinforcingcage (Figure No. 3).4.3. Should the penetration be designed to prevent thereinforcement in that area from yielding, and therebyprevent the concrete from cracking, under any overpressure,the next weak link will be the junctionof the cylindrical shell with the rigid base.The bending stresses at this point are very high,even under design pressure. Yielding under an appliedoverpressure will cause the formation of a partialhinge and, thereby, partially relieve these stresses.However, the strains encountered in the reinforcingsteel beyond the strain hardening area (6- = 0.03)are so big that one has to disregard the capacityof the concrete completely, both in shear, flexureand in compression.Therefore, since the base is not designed asa true hinge, it will fail by horizontal translation(see Figure #4). However, if the liner should becapable to absorb these translations, the point offailure may be close to the ultimate strength ofthe reinforcing bars, i.e., close to an internalpressure ofPult. - 57' x 120 PSIas derived before. A~t that stage, one may visualizethe structure to act like a gas-filled balloon, withthe reinforcing cage in pure membrane tension toprovide the only system of resistance.VIII-149

Page - 10 -LIST OFREFERENCES1. "Behavior of Pre-Cracked Concrete, Subjectedto Reversing Shearing Stresses",by M. J. Holley (MIT) and R. H. White (CornellUniversity), published by Stone & Webster,Boston, 19692. Summary Description of Containment Structurefurnished to Battelle by client.3. "Criteria for Reinforced Concrete NuclearPower Containment Structures",by ACI Committee 349, published Journal of theAmerican Concrete Institute, January 1972,Detroit.4. Effects of High Temperature Exposure on Concrete",by H. S. Davis, Proceedings, Meeting AmericanNuclear Society, June 20 - 24, 1965.LIST OFFIGURESFig. #1: Dowel Action of Reinforcing Bars near the FaceFig. #2: Dowel Action of Reinforcing Bars at the CenterFig. #3: Radial Shear Failure near PenetrationFig. #4: Formation of Hinge at the BaseVIII-150

Page - 11 -eN .l-0aI NE rF ECTXI V EN E--s OFZEýINT=Oft-fNG 33AP-SAS RADIAL ScHEA7R-,D 0W FLS),I~ F PL-AC r- -c-o-UO.&e -ro-rii &VR.FACE--Ft~. ~ZSh e-cLr-4RýEUAJFRC.INa E9AP-S LCC- L 0 S G TOT++ cNIERG C .l-OfA &EC-T-IOM AR-ff S-7fFFC-T1L"67AS VOW:L-S FOR- JZAPIA- cifG!AFLVIII-151

Page - 12 -RADIAL SHE7AP- FIAILiJ,/R.!LJ.E 70 HIGH TEN.SiLeCRTACK INa C-AV.SD e.I ZY I ErU1L. I N r "P-E I F N T 1.CI N G.\",CO N C F. ET 7-E A N• r>Pe7N c--'r-FZ-- •LE7E AR-E N0"r POSI-T1V ELy A '-A Cf4 e:l> Orfi~ JR-SPIFOIZCIN15 CAG7E.-PU N EThATI ON,F=A I L U R.eCOMB ED2 50EAR 4FL- X 0 P•A L__ALVitEVIII-152

APPENDIX EPART HCONTAINMENT FAILURE MODES STUDYbyS. G. Sampath and M. F. Kanninen and P. E. KordaVIII-153

PART HCONTAENMENT FAILURE MODES STUDYbyS. G. Sampath and M. F. Kanninenl and P. E. Korda 2This study was premised upon an accidentin which the pressure within a givencontainment building increases steadily,but relatively slowly, without limit.The objective was to determine both thepressure level at which the structurefails and the most probable mode offailure. The results for the twocontainment buildings considered in thisstudy are given in Table VIII E-1.The analysis of the steel structure(BWR) was much the simpler of the twocontainment types considered. The calculationswere performed by the authorsafter consultation with Mr. E. C.Rodabaugh 3 , a recognized expert in thestress analysis of pressure vessels.The reinforced-concrete structure, however,represented a more formidableundertaking. Here, separate analyseswere performed by the authors and by Mr.Paul Mast 4 , an authority on reinforcedconcrete design. Interestingly, the twoindependent approaches predict failurepressures that are close enough that thesingle entry in Table VIII E-1 suffices.The failure mode for the reinforcedconcretestructure is believed to arisein the following manner. First, as thepressure rises above the design value,the concrete between the two layers ofreinforcing bars will fail in shear.1 Applied Mathematics and MechanicsSection, Battelle's Columbus Laboratories.2 Korda and Associated, Ltd., Columbus,Ohio.3 Applied Solid Mechanics Section,Battelle's Columbus Laboratories.4 Consulting Engineer, Chicago, Illinois.This will lead to a complete separationat the planes of reinforcing where theconcrete is weakest. The structure willstill retain its integrity, however,because the inner layer of reinforcingbars and the liner will remain intact.The point at which the latter sectionwill fail is not certain, but it can bebounded. The lower bound, which correspondsto the pressure at which theliner and the reinforcing bars are justat yield, is 63.6 psig. The upper boundcorresponds to the pressure at whichthey are stressed to their ultimatestrength. This is 87.5 psig. The entryin Table VIII E-1 is approximately theaverage of the two bounds; it representsthe best estimate presently obtainablein the absence of details of constructionand workmanship.The analysis of the reinforced-concretestructure performed by Paul Mast isfully described in Part I. He considersthat an increase in pressure will give aproportional increase in the membranestresses in the structure. Becauseunder design conditions the stress inthe steel reinforcing bars is 60 percentof the yield stress, the pressure atyield can be determined to be 75 psig.The bulk of his analysis consists ofshowing that failure must inevitablyfollow when this condition is reached.In essence, Mast's argument is that thestrain associated with yielding in steelwill cause cracks in the concrete thatare wide enough to tear the linerwhereupon fission material will bereleased at a catastrophic rate.The analysis of the steel structureproceeded in two stages. The first wasa computer stress-analysis computationdesigned to investigate the effect ofstress concentrations at junction pointsin the complete structure. The secondstage involved calculations of the burstpressure of individual components. Theresult was the finding that one sec-VIII-155

tion - the toroidal knuckle - can besingled out as the region in which afailure is most likely to occur. Theburst pressure, as given in Table VIIIE-1, is estimated as 250 psig.it should be emphasized that the conclusionsreported here must be consideredas tentative. The reason is that, dueto severe constraints on both time andthe lack of details, many areas receivedonly a cursory examination or were notpursued at all. Several of these aspectswere suggested by the literaturereferences contained in the Bibliography,others from the experience of theconsultants. Some of the more seriousareas of concern are given in theDiscussion and Recommendation section.Eli-i. DESCRIPTION OF CONTAINMENTFAILURE ANALYSESAs noted above, the analyses of the twocontainment types considered were performedindependently. Accordingly,these are described separately asfollows. Note that the following doesnot include the analysis of Paul Mast.His work, performed in the capacity of aconsultant, is described in Part I ofthis Appendix.Ei- 1.1 REINFORCED -CONCRETE:CONTAINMENT STRUCTUREA number of different failure mechanismswere considered and, so *far as possible,quantitative estimates made for each.The accuracy associated with theestimated failure pressures was limitedby the lack of specific design details.It is felt that the uncertainty in thepredicted results could be reduced witha more detailed examination. Specificareas that may warrant additionalattention are noted in the concludingsection of Part II.A qualitative description of the mostlikely failure mode is as follows. Asthe pressure increases beyond theordinary working levels, the stress inthe outer layer of reinforcing bars willbecome significantly less than that inthe inner layer. The reason is that thecombined shear and tension imposed onthe concrete by the hoop stress and thestrain difference between the two layersquickly exceeds the strength of theconcrete. This is in accord with theeffects on the concrete of the thermalenvironment. As stated in a typicaldesign evaluation:"The thermal operating load in thecontainment concrete wall combined withincident condition loadings produces astress difference of approximately 6,000psi between the reinforcing steel adjacentto the inside face of the wall andthe reinforcing steel adjacent to theoutside face of the wall. Thisdifference exists in both the longitudinalsteel and the hoop reinforcingsteel."This conclusion was obtained from acalculation based on the assumption thatthe concrete will constrain the reinforcementto deform in accordance withthe "1plane sections remain plane"hypothesis. If this is the case, thethermal operating load alone would causea shear stress equal to almost 4.5 ksi.This is approximately 40 times theultimate capacity of the concrete.obviously, strain compatibility willcease long before the theoretical stresspattern indicated by the computeranalysis could develop.With increasing internal pressure, thefirst major area of distress, then, willbe a massive shear failure in theconcrete. There will probably becomplete separation at the planes of thereinforcing where the concrete is initiallyweakened. Concrete will crumbleand crack away from the reinforcingbars. The outer layer of reinforcingwill be effective for a time due to thephysical restraint exercised over theenclosed volume. Eventually, chunks ofconcrete will spall off the outside.The outer layer of reinforcing will thenbe virtually relieved of all loadbearingcapacity.The liner and the reinforcing near theinside face are only inches apart. Thiscircumstance probably required verycareful workmanship during constructionto achieve proper consolidation of theconcrete and its integrity is somewhatsuspect. However, assuming that thissection is properly constructed, thecontinuous liner on the inside and theapproximately 4 inches of concrete coveron the outside will probably result insufficient confinement that the linerand the inner reinforcing will worktogether through the critical- phases.it is this steel whose load-carryingcapacity will determine the pressure atfailure.The reinforcing rods will yield at 50ksi while the steel liner will yield at32 ksi. The large strains accompanyingyielding will accelerate the crumblingand pulverizing of the concrete, bothoutside the inner layer of reinforcing4VIII-156

and, eventually, between it and theliner. At some point, strain hardeningwill commence to accommodate furtherincrease in the internal pressure. Itis not possible to predict the preciseextent to which this will happen priorto ultimate failure. But, while thestress level in the steel is between theyield and the ultimate strength, localdeformations will cause the concretefailure to extend all the way to theliner. When this happens over a largeenough area, the combined tension andbending will cause a blowout with thepossibility of the crack propagatingseveral feet before the sudden releaseof the internal pressure will cause itto stop. Because a catastrophic amountof containment will be released by thisprocess, this is deemed a failure mode.In the state described above, the hoopstress is twice the longitudinal stress,and the circumferential reinforcingavailable is twice that available in thelongitudinal direction. Hence, thegoverning condition can be taken to bein the direction of the hoop stress.The reinforcing bars are in uniaxialtension with a minimum guaranteed yieldstrength of 50 ksi and a minimum tensilestrength of 70 ksi. The steel liner isessentially in a state of biaxialtension with the hoop stress approximatelyequal to twice the longitudinalstress. Under these conditions, themaximum distortion energy failurecriterion results in an increase by afactor of 2/Vr = 1.155. The guaranteedminimum yield strength is 32 ksi and theminimum tensile strength, 60 ksi.Estimates of unit stress and hoop forceare shown in Table VIII E-2. The insideradius of the containment structurebeing 63 ft, these correspond to a lowerbound pressure of 63.6 psig and an upperbound of 87.5 psig.In summary, therefore, failure willoccur no sooner than when the steelliner and the inner layer of reinforcingare stressed to yield and no later thanwhen they are subjected to theirultimate tensile strength. This givesrise to an upper and lower bound for thefailure pressure. One would mostcertainly anticipate a significantpressure increment to failure beyond thelower bound (the extent will primarilybe a function of the uniformity ofworkmanship both in welding steel and inplacing concrete). A good estimate of alikely single value would be somewherenear the middle of the range. Thisgives the value of 75 psig quoted inTable VII E-1.E!I-1.2 STEEL CONTAINMENT STRUCTUREThe analysis of the steel containmentstructure consisted of two phases: acomputer analysis of the entire structureas an axisymmetric pressure vesseland failure calculations based on thestress concentrations existing in theindividual components. Local stressconcentrations arising from accesses,pipes and the like, of course, areignored in an axisymmetric analysis.Hence, the computer calculation isintended to evaluate only the potentialfor failure at positions with discontinuities,e.g., changes of thickness. Inturn, these effects are ignored in theevaluation of local stress concentrationswhich follow, it being reasonableto assume that the interaction betweenthose effects can be neglected.EII-I.2.1COMPUTER CALCULATION FORAXISYMMETRIC PRESSURE VESSELA Battelle computer program (MONSAS) wasused to calculate the stresses in theprimary containment shell under thedesign basis accident conditions. Inessence, the program performs a nonlinear(large deformation) analysis ofelastic shells of revolution using amultisegment, direct-integration techniquecoupled with a Newtonian-typeiterative scheme. The shell geometrywas modeled as shown in Fig. VIII E-1.Because tuberances, accesses, piping andvents in the shell were not taken intoaccount so that the problem could betreated as axisymmetric. The profile ofthe shell consisted of five parts, asshown in Fig. VIII E-1. Because precisedetails of the structure were notavailable, the wall thickness had to beestimated. The assumed thickness distributionused in the computation isgiven in Table VIII E-3. Note thatbetween any two points within a singlepart, the program assumes a linearlyvarying thickness.At the concrete embedment at elevation119'11", a fully clamped condition wasassumed for the shell. The model wasterminated at the flange connection atelevation 210'4", with a simple supportand no inplane constraint. However, astatically equivalent in-plane load(6000 lb/in.) was prescribed at thispoint to account for the longitudinalstress that arises from the internalpressure. A value of 62 psig for thepressure was utilized with an assumeduniform temperature of 281 F. Theelastic constants for the material ofthe shell were those specified for ASTM-VIII-157

A5l6, Grade 70, Fire Box Quality, USSSteel.The calculated principal stresses atvarious locations are given in TableVIII E-3. For comparison, the designvalues for a typical configuration arealso given. The overall agreementbetween the two results appears fair;the variation in the thickness distributionassumed in the two cases probablyaccounts in large part for the specificdifferences. The discontinuity in thestress at elevation 119111" due tobending can be expected to disappear ifaccount is taken of the sand bedprovided between elevation 117'1-5/8"and elevation 119111". A high value forstress was also obtained at the junctionbetween the conical section and thecylindrical part (see Fig. VIII E-1,elevation 202'). A large bending stresscan also be expected at the flangeconnection at elevation 210'4". However,due to a lack of available detailof the joint and sealing arrangement,the present analysis did not take thisinto account.Other factors which need to be consideredfor a detailed study of the failuremode of the shell are discussed in theconcluding section.EII-l. 2.2BURST-PRESSURE CALCULATIONSFOR INDIVIDUAL COMPONENTSAs mentioned above, the complexitiesintroduced by the fittings and othernonaxisymmetric features of the containmentstructure precluded an exactanalysis. Instead, guidance from designexperience as incorporated into the ASMENuclear Pressure Vessel Code was used toestimate the burst. pressure for eachpart of the containment structuretreated individually. These calculationswere carried out under thefollowing restrictions:a. A single application of the loadneed be considered.b. All welds are of uniformly highquality and are completely ductile.C. The stiffness and strength contributedby the girders to the toroidalsuppression chamber is neglected.d. No large defects are present in thevessel which would cause catastrophicfracture by unstable crackpropagation.e. Flanged joints are adequately sealedto prevent leakage prior to burstingof vessel.f. Reinforcements at nozzle openingsare equivalent to at least 100 percentarea replacement (as requiredby the ASME Code).Given the above conditions, burstpressures for each of the major parts ofthe containment vessel can be calculatedby the simple strength of materialsrelationships. The results are shown inTable VIII E-4.Having determined from the computeranalysis that the overall design of thecontainment structure is basically correct,it is expected that failure wouldoccur at a local stress concentrationwithin an individual part. Hence, fromthe results given in Table VIII E-4, itis concluded that failure would occur inthe toroidal knuckle section of thetypical configuration. The calculatedvalue for this possibility given inTable VIII E-4 corresponds to the entrygiven in Table VIII E-1 of this report.EHI-2. DISCUSSION ANDRECOMMENDATIONSIn this section, the principal areas ofuncertainty in the analyses performedare discussed. The results given in thepreceding section should be interpretedin the context of these uncertainties.E11-2.1 REINFORCED -CONCRETECONTAINMENT STRUCTUREThe primary uncertainties in theanalysis of the reinforced-concrete containmentvessel are deemed to be asfollows.a. The structure has an excess of steel(in the neighborhood of 5 percentwhereas it is usual to find 1-2percent steel in typical welldesignedconcrete structures).Proper compaction of concrete undersuch conditions would be difficultto achieve.b. The currently accepted ACI Code 318-71 is more demanding than the 318-63Code under which the structure wasdesigned with regard to adequatedevelopmental requirement for deformedbars. This is based onrecent theoretical research and onfield experiences. A calculation ofthe development length for Number 18reinforcing bars indicates that alength of about 100 inches isrequired. This is greatly in excessof the lengths available indesign in local areas.theVIII-158

c. Seismic loads are assumed to appearnonconcurrent with an accidentcondition. No account of theseismic load has been taken into anyof our discussions.d. In the personnel- and access-hatchdesigns, the steel ring could carrythe shearing forces as hoop tension.In this case, shear cracking ofconcrete and the consequent loss oflamination between the outer-innerreinforcement is not credible exceptas exists in the overall structure.e. The liner is in a biaxial state ofstress so that its yield stress isunderestimated by uniaxial data.While some account was taken of thisfact, the actual failure pressuremay be sensitive to the exact value.EU- 2.2 STEEL CONTAINMENT STRUCTUREIn essence, the primary uncertainty inthe analysis of the steel containmentstructure stems from a lack of somespecific design details. To put thecurrent effort into perspective, thecomputer analysis performed here issummarized in Table VIII E-5. It can beseen from this and the discussion in thepreceding sections that many details ofthe design have not been considered. Ifa more precise prediction of failurepressure were desired, analyses shouldbe performed to account for details ofthe reinforcing around pipeand access holes.penetrationA reactor building encloses the primarycontainment and also serves the additionalpurpose of limiting deflectionsof the steel shell under the action ofconcentrated forces to prevent failure.Although permanent deformations of theprimary containment and the suppressionchamber are not considered as failureper se, possible reactive pressures(arising due to the concrete shield wallas a result of the overall deflectionsof the primary chamber exceeding the gapprovided between the two walls) warrantsconsideration.The elastic computer analysis wasperformed with the pressure load assumedas 62 psig. The results gave shelldeflections normal to the surface ofabout 1 inch. It is to be expected thatreactive pressures would arise on theshell surface long before the hypothesizedburst pressures (calculatedwithout taking this point into consideration)were reached. The contour andthe stiffness of the concrete shield notbeing known, it appears that unequalconstraints will tend to alter thestress distribution calculated along thesteel-shell contour under incipientburst conditions. Intuitively, theknuckle portion of the drywell and thedouble-gasketed flange, due to bendingand subsequent leaking appear morevulnerable than had been stated earlier.An inclusive analysis should beworthwhile.VIII-159

References1. Merkle, J. G., "The Strength of Steel Containment Shells", Nuclear Safety, 7(3), 346-353 (1966).2. Doyle, J.M., and Chu, S.L., "Some Structural Considerations in the Design ofNuclear Containment Liners", Nuclear Engineering and Design, 16, 294-300 (1971).3. Browne, R. D., and Blundell, R., "The Behavior of Concrete in PrestressedConcrete Pressure Vessels", Nuclear Engineering and Design, 20, 429-475 (1972).4. Tan, Chen Pang, "A Review of the Technology of Prestressed Concrete ReactorPressure Vessels", Nuclear Safety, 11 (1), 25-33 (1970).5. Costantino, C. J., "The Strength of Thin-Walled Cylinders Subjected to DynamicInternal Pressures", Trans. ASME, Jour. Appl. Mech., 32, 104-106 (1965).6. Haydl, H. M., and Sherbourne, A. N., "Effect of Transverse Shear on Limit Loadof Cylindrical Shells", Nuclear Engineering & Design, 22, 290-295 (1972).7. Rashid, Y. R., "Ultimate Strength Analysis of Prestressed Concrete PressureVessels", Nuclear Engineering & Design, 7, 334-344 (1968).8. Halligan, D. W., "Structural Design Criteria for Secondary ContainmentStructures", Nuclear Engineering and Design, 8, 427-434 (1968).9. Whitman, G. D., "Technology of Steel Pressure Vessels for Water-Cooled NuclearReactors", Nuclear Safety, 8 (5), 429-442 (1967).10. Denkins, F. R., and Northrup, T. E., "Concrete Containment Structures", NuclearSafety, 6 (2), 194-211 (1964-65).11. Langer, B. F., "Rupture of Cylindrical Shells", PVRC Interpretative Report ofPressure Vessel Research, Welding Research Council, 95 (August, 1964). Section I-Design Consideration.VIII-160

TABLE VIII E-1 ESTIMATED CONTAINMENT-BUILDING FAILURE PRESSURESBuildingFailureClassification Wall Structure Pressure, psigSub-atmosphericcontainment forpressurized water Reinforcedreactors concrete 75Vapor suppressioncontainment forboiling waterreactors Steel 250TABLE VIII E-2-DISTORTION. ENERGY FAILURE ESTIMATESUnit Stress,Hoop Force,Area, ksi kip/lin. ft2in. / Lower Upper Lower Upper1in. ft Bound Bound Bound Bound2 Number 18 bars on 12-in.centers 8.00 50 70.0 400 5603/8-in. steel liner 4.50 39 51.9 176 234576 794TABLE VIII E-3 SUMMARY OF CALCULATED STRESSES IN TYPICAL VAPOR-SUPPRESSIONCONTAINMENT FOR BOILING-WATER REACTORSPrincipal Stresses at*Design Basis Accident.Conditions(a), psi CalculatedAssumed WallDeflections,Elevation Thickness, inch Calculated (MONSAS) Design inch119'11" 1.25 (b) 11,802 0.00126' 1.25 10,400 10,679 1.041~132'8y" 1.25 9,977 10,147 1.06106137'71" 1.25 9,975 16,327 1.08154'-9" 1.07 11,650 15,900 1.14170' 315- 1.93 12,500 15,592 1.21171'2i" 2.8125 16,895 14,200 1.26172'8 1 2.8125 19,700 16,712 1.18175'91" 2.8125 11,284 9,714 0.65178'5" 0.75 14,299 -- 0.66179'3" 0.78 18,178 0.69183'8" 0.985 15,728 -- 0.68200'3" 1.25 15,747 - -- 0.65202' 1.25 5 5 , 2 7 8 (c) -- 0.44204'11.6" 1.39 12,179 1.56207'4" 1.5 33,481 -- 1.70210'4" 1.5 4,000 -- 0.00(a) Pressure =-62 psig, temperature - 281 F.(b) Discontinuity stress, clamped boundary condition.(c) High bending stress.

TABLE VIII E-4SUMMARY OF ESTIMATED BURST PRESSURES FOR COMPONENTS OF A TYPICALVAPOR-SUPPRESSION CONTAINMENT FOR BOILING-WATER REACTORSCalcuiatedLower Upper Burst Pressure,Part Model Elevation Elevation psig1 Spheroidal 119ll" 171'21"83002 Toroidal knuckle 1712-- 175'91 2503 Cylinder 1759 202 2604 Conical 202' 207 4" 3755 Cylinder 20714" 210'4" 620iTABLE VIII E-5SUMMARY OF COMPUTER ANALYSIS USED FOR VAPOR-SUPPRESSION CONTAINMENTFOR BOILING-WATER REACTORSAccounted for orAssumed in AnalysisNot Accounted forin Analysis1. Program MONSAS: Thin-shell analysisCode Used-2. Analysis Linear elastic, large deforma- Plasticitytions3. Geometry Shell of revolution; axisym- Effect of protrusions,metric (see Fig. VIII E-1)vents, accesses, sundrypipingModeled from elevation 119'11"and terminated at flange atelevation 210'4"Ellipsoidal head above210'4"; discontinuityflange (details of flangeconnection not availableyet)4. Thickness variable thickness was assumed(see Table VIII E-l). The programDetailed thickness distributionunavailableassumes a linear distribution from drawingbetween prescribed points.5. Load Internal pressure = 62 psig and Dead load of structure,uniform temperature - 281 F seismic loads (verticaland horizontal) hydrostaticload from floodedcondition6. Boundary At elevation 119'11", the edge Air gap between elevationCondition was assumed clamped; i.e., nor- 1 " e amal displacement w, meridional 117'o" and elevationdisplacement u, and rotation 11911'iwas set = aAt elevation 210'4", the edge Discontinuity and bendingwas free with an axial tensile stresses at flange andforce being applied to account the effect of the ellipforthe longitudinal stress due soidal head. This wouldto pressure. This force was cause compressivecalculated from statics; i.e., stresses.force/unit circumference = P.7. Mechanical E 29 x 10 6 psi, 0 = 0.27, a a'ultProperties a = 8.4 x 10-6 in./in./FTable VIII E-1 - Table VIII E-5VIII-161/162

PbPort 5Part 4PartNo.I23Part NameSpheroidalToroidalCylindricalainches402.0303.0231.0b ,inches72.0314.5bPart 3- a -~-4Conical231.074.05Cylindrical194.035.914a -~b2Integration/aPart IInitialFIGURE VIII E-1Axisymmetric Model for Steel Shell Computer AnalysisFig. VIII E-1VIII-163/164

WASH--1400(NUREG 75/014)SAFETY DESIGN RATIONALEforNUCLEAR POWER PLANTSAPPENDIX IXtoREACTOR SAFETY STUDYU.S. NUCLEAR REGULATORY COMMISSIONOCTOBER 1975

APPENDIX IXTable of ContentsSectionPage No.1. INTRODUCTION AND SUMMARY .......................................... IX-12. PRIMARY BARRIERS TO RELEASE OF RADIOACTIVE MATERIAL ................... IX-52.1 Fuel Rods .................................................... IX-52.2 Reactor Coolant System and Containment ............................ IX-53. GENERAL BASIS FOR SAFETY DESIGN ................................... IX-93.1 Normal Operation at Power .................................... IX-93.1.1 Confinement of Radioactivity ......................... IX-93.1.2 Cooling of Fuel Rods ................................. IX-93.2 Anticipated Transients and Accidents ......................... IX-104. PROTECTION AGAINST OVERPOWER ...................................... IX-134.1 Reduction in Temperature of Reactor Coolant ...................... IX-134.1.1 Temperature and Flow of Feedwater ........................ IX-134.1.2 Increase in Steam Demand ............................. IX-144.1.3 Rupture of a Main Steam Line in a PWR .................... IX-144.2 Increase in Flow of Coolant to Reactor Core ...................... IX-164.3 Increase in Pressure in a BWR ................................ IX-174.4 Uncontrolled Movement of Control Absorber ........................ IX-174.4.1 Withdrawal of Control Assembly by DriveMechanisms ...........................................IX-174.4.2 Dilution of Boron in PWR Coolant ......................... IX-184.4.3 Insertion of a Control Assembly ........................... IX-184.4.4 Ejection of a Control Assembly ............................ IX-195. PROTECTION AGAINST A REDUCTION IN CORE COOLING ........................ IX-215.1 Reduction in Reactor Coolant Flow ............................ IX-215.2 Mismatch Between Feedwater and Steam Flows ....................... IX-215.3 Reduction in Steam Demand .................................... IX-225.4 Reduction in Pressure in Reactor Coolant System .................. IX-236. PROTECTION AGAINST LOSS OF REACTOR COOLANT ............................. IX-256.1 Cooling of Fuel Rods ......................................... IX-256.1.1 Small-Break Accidents ................................ IX-256.1.2 Large-Break Accidents ................................ IX-266.1.3 Breaks of Intermediate Size .......................... IX-276.2 Containment of Radioactivity During Accident ..................... IX-286.3 Post-Accident Heat Removal and Confinement of FissionProducts .....................................................IX-296.3.1 Cooling of Fuel Rods ................................. IX-296.3.2 Cooling of Containment and Retention of FissionProducts .............................................IX-296.3.3 Control of Hydrogen Concentration in Containment ..... IX-296.4 Special Containment Situations ............................... IX-30IX-i

Table of Contents (Continued)SectionPage No.7. MEASURES TO ENSURE PERFORMANCE OF SAFETY FUNCTIONS .................... IX-337.1 Quality Assurance ............................................ IX-337.2 Protection Against Effects of Severe Natural Phenomena ....... IX-337.3 Reliable Power Supply ........................................ IX-347.4 Redundancy in Safety Systems ................................. IX-347.5 Testability of Safety Systems ....................................... IX-358. CONCLUSION ........................................................ IX-37List of FiguresFigurePage No.IX 2-1 Typical Reactor Core Fuel Rod ................................... IX-7/8IX 2-2 Typical PWR Reactor Coolant System .............................. IX-7/8IX 2-3 Typical BWR Reactor Coolant System .............................. IX-7/8IX 3-1 Heat Production of Decay Fission Products ............................ IX-11/12IX 6-1 Typical PWR Containment ......................................... IX-31/32IX 6-2 Typical BWR Containment ......................................... IX-31/32IX-ii

Section 1Introduction and SummaryAs indicated elsewhere in the ReactorSafety Study, the AEC's regulatorybodies have put into use a well developedset of safety design requirementsfor pressurized-water (PWR) and boilingwater(BWR) nuclear power plants. Theprimary purpose of this appendix is todescribe, for the benefit of those notwell-schooled in reactor safety, thebasic logic for most of these safetyrequirements, unimpeded by the stylizedlanguage of regulations, licensingapplications, and technical codes andstandards. To avoid interfering withthe logical development, a minimum ofdescriptive material is presented. Thereader is presumed to have some understandingof the fission process and theoperation of nuclear power reactors.For those without this understanding,many references are available. 1Another purpose of this appendix is toprovide a bridge between the traditional"qualitative" approach to reactor safetyand the "quantitative" approach taken inthe Reactor Safety Study. This appendix,in essence, describes the. conventionalapproach to reactor safety ascurrently used by the AEC's regulatorybodies and the industry. In the Study,this approach to safety is analyzedquantitatively and on a more realisticbasis than is usually employed to predictthe probabilities and consequencesof reactor accidents. Although theStudy is not limited to consideration ofthose accidents defined in the conventionalapproach, it has started withthem and then expanded the scope of thesafety investigation to include manyother accidents of risk significance.Nuclear power plants, like fossil-fueledplants, use heat produced by reaction ofa fuel to generate steam to drive aturbine-generator to produce electricity.Both types of plants must bedesigned to protect against the hazardsof handling large volumes of water and1 Nuclear reactors and the fission processare described briefly and simplyin the booklets entitled "Nuclear PowerPlants" (IB-505) and "Nuclear Reactors"(IB-507) of the "Understanding theAtom" series, prepared by the U.S.Atomic Energy Commission, Office ofInformation Services.steam in large equipment and piping athigh temperature and pressure. In thenuclear plant, however, the radioactivefission products that are the byproductof the heat-producing fission reactionin the uranium fuel constitute an additionalhazard. A plant generatingelectricity at a rate of 1 thousandmegawatts (1000 MWe) contains about10 billion curies of radioactivity. 1Much of this radioactivity persists forlong times after the fission reactionhas been turned off. If it were possibleto disperse the radioactive productsfrom a large plant, people might receivesevere radiation exposures and use ofland could be restricted over manysquare miles in the vicinity of theplant.The real business of reactor safety isprevention of the release of radioactivityfrom the reactor fuel. Although thefission process creates large amounts ofradioactivity in the fuel rods, theceramic pellets of uranium dioxide fuelretain more than 98% of this radioactivity.About 2% of the radioactivity,chiefly the gaseous krypton, xenon, andiodine, diffuses into the gas plenumbetween the fuel pellets and the sealedZircaloy cladding.Small amounts of radioactive gas mayleak through the cladding in normaloperation (the core of a large reactorcontains roughly 40,000 fuel rods andsome will surely leak), but this is ofminor consequence. The leakage isconfined by two additional barriers tothe release of radioactivity to theenvironment - the piping and vessels ofthe reactor coolant system and thecontainment. Waste, treatment systemsare provided to dispose of this radioactivity,and, although the leakageslightly' complicates the operation ofthe plant, it is not a hazard to thepublic.If the fuel rod cladding is heated to atemperature several hundred degreesabove the normal operating temperature,it can fail and release the radioactivityin the gas space plus a small quantityof volatile solid fission products1 This excludes very rapidly decayingisotopes.IX-I

located at the surface of the ceramicpellets. This small fraction of thetotal radioactivity is easily confinedby the reactor coolant system and/or thecontainment. The only way that a largefraction of the radioactivity can bereleased is to severely overheat or, inessence, melt the uranium dioxide pellets.if uranium dioxide remainsmolten, about 15% of the radioactivityis volatilized in a few minutes andabout 25% is released in a few hours.Melting of the fuel in a reactor corewould also lead to failure of the otherbarriers as they are now designed.These facts have been known for manyyears. Consequently, in the design,engineering, construction, and operationof nuclear power plants, much effort hasbeen expended on preventing damage tothe cladding and melting of the fuel.In principle, there is only one way forthe fuel to melt: it must generate moreheat than is being removed from it.When the reactor is operating, the fuelis cooled by water flowing at high velocityover the surfaces of the fuelrods. When the. reactor is shut down,cooling must be continued, but at a significantlylower rate, to remove theheat due to radioactive decay of thefission products. Three kinds of eventscan cause the heat generation in thefuel to exceed the cooling capability ofthe water: the heat generation canrise, the cooling capability candecrease, or the coolant can be lostfrom the reactor core. Such eventsmight be caused by equipment failures inthe plant due to faulty design, construction,or operation. Externalforces exerted on the plant by earthquakesand other severe natural phenomenaand by airplane crashes are also potentialcauses.Measures are taken to ensure highquality in design, construction, andoperation of the plant so that thenumber and severity of events caused byinternal failures will be minimal.Abundant cooling is provided to removeheat from the fuel and allow flexibilityin operation and margin for accommodatingupsets. Control systems keep thereactor power and coolant conditionswithin prescribed limits when the plantis operating. The plant is designed tooperate or be shut down and maintainedin a safe condition in the event of anearthquake, flood, tornado, or hurricaneof a magnitude not expected to occurduring the lifetime of the plant.In spite of these measures, it isassumed that operational upsets andaccidents will occur, and measures areprovided to cope with them. A reactorprotection system is provided to shutdown the plant and maintain it in a safecondition when limits on operating conditionsare reached. Engineered safetyfeatures are incorporated to cool thefuel and augment confinement of theradioactivity by the containment underabnormal and accident conditions. Thesesystems are designed to terminate anticipatedtransients - events that arelikely to occur during the lifetime ofthe plant - without damage to the fuelrods or the coolant system. Severeaccidents that have a lower probabilityof occurrence are also postulated andexamined. Some damage to the fuel rodsis considered acceptable in theseevents, but the safety features mustprevent the rods from melting or radioactivematerials from escaping from theplant in sufficient quantity to causeserious exposures to the public.An unintended rise in power in the reactorcore could be produced by:a. A reduction in temperature of thecoolant to the core. Such a reductionwould increase the density ofthe coolant, increasing its abilityto slow neutrons. This would leadto an increase in neutrons availablefor fissioning with a resultantincrease in power. An increase inflow or decrease in temperature ofthe feedwater being returned to thereactor coolant system or, in a PWR,an increase in steam demand or ruptureof a steam line will lower thecoolant temperature.b. An increase in flow of coolant tothe core. Start-up of an inactivereactor coolant recirculation loop,or abnormal operation of the reactorcoolant flow controller in a BWR,can increase the flow and in someinstances lower the temperature ofthe coolant.C. Uncontrolled removal of control absorberfrom the core. Uncontrolledwithdrawal of control assemblies bythe drive mechanisms, ejection of acontrol assembly by rupture of acontrol rod housing in a PWR or bygravity forces in a BWR, and dilutionof the boron in the coolant ofa PWR remove control absorber fromthe core.d. An increase in pressure in a BWR.Abnormal operation of the pressurecontroller or rapid closing ofvalves in the main steam lines, as1ýIX- 2

in a turbine trip, increase thepressure in the core.e. Insertion of a control assembly. Ifother control assemblies are withdrawnto maintain the total corepower, the power will rise in someregions of the core to compensatefor the reduction in the vicinity ofthe inserted assembly.A reduction in cooling capability wouldresult from:a. A reduction in reactor coolant flow.Stoppage of one or more of the reactorcoolant pumps or abnormal operationof the reactor coolant flowcontroller in a BWR will reduce thecoolant flow.b. A reduction in pressure of the coolantin a PWR. Release of steam fromthe pressurizer or an increase insteam demand can reduce the pressure.c. A reduction in the inventory ofwater in the steam generators of aPWR or the reactor vessel of a BWR.A mismatch between the feedwaterflow and the steam flow caused byabnormal operation of pumps andvalves or breaks in pipes will leadto a reduction in inventory and incapacity to cool the core in theevent of loss of feedwater flow.d. A reduction in steam demand. Closingof valves in the main steamlines stops the flow of steam andthe removal of heat from the reactorcoolant.Loss of reactor coolant would resultfrom the opening of a valve or a breakin a pipe that allowed water or steam tobe discharged from the reactor coolantsystem in an uncontrolled manner. Theopening can range in size from a smallcrack to a hole produced by severance ofthe largest pipe in the reactor coolantsystem.Because of these considerations, theplants are designed so that fuel rodsare likely to be slightly damaged onlyin the more severe accidents. Normalconditions are restored by the controlsystems without shutting down the reactor,or the protective system shuts downthe power to prevent the fuel from overheatingin most of the potential powerriseand reduction-in-cooling incidents.Auxiliary feedwater systems and highpressure coolant injection systems providewater for the steam generators andthe reactor coolant system, and steamrelief systems provide a means for dissipatingthe decay heat. The reactorprotection system is designed to preventthe start-up of an inactive loop in aPWR under conditions that could causefuel damage. Only the most severecontrol-assembly-ejection accidents arelikely to damage fuel rods. Such accidentshave a low probability of occurrenceand the damage would be limited toperforation of the cladding of a smallfraction of the rods.Loss-of-coolant accidents fall into twogeneral categories - small-break andlarge-break accidents. High pressurecoolant injection systems are providedto maintain sufficient water in the coreto keep the fuel rods from being damagedduring the time required to depressurizeand cool the reactor system in smallbreakaccidents. In large-break accidentswater and steam are dischargedfrom the reactor coolant system sorapidly that no means has been found toprevent the core from being voided ofliquid during the depressurization.Systems are provided to inject waterinto the reactor vessel at intermediateand low pressure to cool and reflood thecore rapidly. Because a large-breakloss-of-coolant accident is an event oflow probability, the possibility of somedamage to fuel rods is acceptable. Thecooling systems are designed to ensurethat the cladding will not melt or beseriously oxidized and embrittled duringthe temperature transient.Depressurization of the reactor coolantsystem in the larger loss-of-coolantaccidents results in the discharge of alarge amount of water and steam into thecontainment and a consequent rise intemperature and pressure in the containment.Sprays are provided in most PWRcontainments to condense-the steam andcool the containment and to removeiodine and other radioactive particlesfrom the containment atmosphere. Thepressure suppression system serves tocondense steam in the BWR containment.Water recirculation systems are providedfor removing the fission product decayheat from the reactor core and the containmentfor a long time after an accident.Recombiners in the PWR containmentand a nitrogen atmosphere in theBWR containment prevent the hydrogen,produced chiefly from radiolysis of thewater, from reaching a flammable concentration.High quality of design, construction,and operation and redundancy and testabilityof equipment and systems areimportant factors in ensuring that theIX-3

safety features will function whencalled upon. If the containment andsafety features perform as the designersexpect, few, if any, of the fuel rodswould be damaged, little radioactivitywould be released, and little hazard tothe public would result from any of theaccidents.IX-4

Section 2Primary Barriers to Releaseof Radioactive MaterialThe large power reactors now operatingor under construction in the U. S. arepressurized water and boiling waterreactors (PWR and BWR) that use waterfor cooling the uranium fuel. Althoughthese reactors differ in some importantfeatures, the design approach to radiologicalsafety for both is to provide adefense in depth against the release ofradioactivity from the plant under allcircumstances considered credible. Thisdefense in depth consists of physicalbarriers augmented by the auxiliary systems,engineered safety features, procedures,and administrative controlsnecessary to make the barriers highlyreliable and to maintain their effectiveness.2.1 FUEL RODSThe first and most important physicalbarrier is the fuel rod. The fuel isuranium dioxide powder that has beenformed into small cylinders and fired athigh temperature to produce ceramic pellets.Groups of pellets are inserted inmetal tubes made of a zirconium alloy,'Zircaloy, and the ends of the tubes aresealed by welding to form fuel rods, asshown in Fig. IX 2-1, for use in thereactor core.Most of the heat produced in the reactoris generated by fissioning of uraniumand plutonium in the pellets of the fuelrods. Most of the radioactive materialsproduced - those of greatest health andsafety concern - are the fission productsleft in the fuel pellets by thefission reaction. Some of the fissionproducts are gasses which will diffusefrom the pellets into the space aroundthem. The principal gaseous productsare krypton, xenon, and iodine. Typically,the fuel rods in a l000-Mwe reactorcontain about 110 million curies ofkrypton, 250 million curies of xenon,and 700 million curies of iodine, ofthese amounts, 0.4 million curies ofkrypton, 3.6 million curies of xenon,and 9.5 million curies of iodine areestimated to be present in the gasspace. About 98% of the gaseous fissionproducts and essentially all the solidfission products remain in the ceramicpellets.At normal operating temperatures thefuel pellets are well below the meltingtemperature, the cladding is strong andductile, and the fuel rod is an effectivebarrier to the release of fissionproducts. However, abnormal or accidentsituations can be envisioned in whichthe temperature of the cladding wouldrise several hundred degrees, and thecladding would fail and release thefission products contained in the gasspace.If the temperature were to rise aboveabout 48001F, the uranium dioxide pelletswould melt. All the krypton andxenon and about 90% of the iodine wouldbe released quickly from the fuel, andsome of the solid fission products wouldbe dispersed as an aerosol. Release ofthis radioactivity from the plant couldhave serious, widespread effects.2.2 REACTOR COOLANT SYSTEM ANDCONTAINMENTTwo physical barriers beyond the fuelrod provide protection against the releaseof fission products during normaloperation and in accident situations.The first of these is the reactor coolantsystem (Figs. IX 2-2 and IX 2-3).The fuel rods, grouped into fuel elementsand installed in the reactorvessel, form the reactor core, which iscooled by the water circulating in thereactor coolant system. Fission productsreleased by the fuel rods are confinedby the piping and vessels of thereactor coolant system as long as theyremain intact. The structure thathouses the reactor coolant system - thecontainment (see Figs. IX 6-1 and IX6-2) - is designed to confine radioactivematerials that escape from thecoolant system.Preferably each barrier - fuel rod, reactorcoolant system, and containment -should be independent and self sufficient.That is, failure of one of twobarriers should not affect the integrityof the remaining barrier(s). Failuresof the containment that are consideredcredible would not affect the integrityof the reactor coolant system or thefuel rods. The components and piping ofthe reactor coolant system are designedand supported and missile shields areprovided so that failures of the reactorcoolant system would not directly affectIX- 5

the integrity of the containment. However,in the absence of special safetyfeatures, some failures of the reactorcoolant system would lead to extensivemelting of the fuel rods. Extensivemelting, whatever the cause, would befollowed by breaching of the reactorvessel and containment and the releaseof a large amount of fission productsfrom the plant. Consequently, the foremostsafety requirement in the design ofa nuclear power plant is to provide reliablecooling for the fuel rods to preventthem from melting.IX-6

Typical Fuel DataPWRBWROverall length, in. 149.7 "- 164Outside diam., in. 0.422 0.563Metal wall thickness, in. 0.0243 0.037Pellet diam., in. 0.366 0.477Pellet length, in. 0.600 0.5Pellet stack height, in. 144 144Plenum length, in. 4.3 16Fuel rods in fuel assembly 204 49Fuel rod pitch, in. 0.563 0.738Fuel assemblies in core 193 764FIGURE IX 2-1Typical Reactor Core Fuel Rod

Steam Outlet (to turbine)Steam Outlet(to turbine).Steam GeneratorFeedwater Inlet(from condenser)FeedwaterInlet(fromMain Coolant PumpPressurizerAccumulatorReactorVesselFIGURE IX 2-2Typical PWR Reactor Coolant System

FeedwaterInletCoolantRecirculationInletCoolantRecirculationOutletDischarge Valve -RecirculationPumpBypass ValveFIGURE IX 2-3Typical BWR Reactor Coolant SystemFig. IX 2-1 - Fig. IX 2-3IX-7/8

,is

Section 3General Basis for Safety Design3.1 NORMAL OPERATION AT POWER3.1.1 CONFINEMENT OF RADIOACTIVITYSuccessful operation of the plant requiresthat almost all the fission productsremain in the fuel rods. Should,for instance, the cladding of too manyfuel rods contain imperfections thatrelease products from the gas space, thelevel of radioactivity in the plantwould be high enough to interfere seriouslywith operation and maintenance.Furthermore, it would be difficult tokeep the radioactivity in routine effluentsat the low level prescribed to protectthe health of the public. Thus, asa part of normal plant operation, careis required to keep small the amounts offission products and other long-livedradioactive materials in the coolingwater.This entails, first, the manufacture offuel pellets and cladding of very highquality so that the number of imperfectionswill be small. However, a reactorcore contains almost 40,000 fuel rods,and some do leak. Water from the reactorprimary system is processed continuouslyby evaporation and ion exchangefor several purposes, including the removalof radioactive fission productsand corrosion products. Waste liquids,gases, and solids from this processingare stored until their radioactivitydecays to a level safe enough for releaseto the surroundings; alternativelythe wastes are packaged and shipped to adisposal facility.Excessive leakage of water from the reactorcoolant system into the containmentwould make operation and maintenancedifficult, so the coolant systemmust be of highest quality. The reactorcoolant system operates at about 1050psi in a BWR and 2250 psi in a PWR, anda small amount of radioactive gas andwater passes through seals on valves,pumps, and other equipment into the containment.The air in the containmentand the liquid collected in sumps at thebottom of the containment are processedto remove and confine their radioactivity.3.1.2 COOLING OF FUEL RODSIn addition to the requirement forfuelrods of high integrity for normal operation,the fuel must be prevented fromreaching temperatures that could meltthe uranium oxide pellets (>4800"F) orthe cladding (3300*F). Because Zircaloyreacts exothermally with steam, thecladding is surrounded by water, causingclad oxidation and embrittlement, it isdesirable to keep the cladding below1800 0 F, the temperature at which thisreaction begins to be significant. Forthis reason, the perturbations thatcould cause excessive fuel or claddingtemperatures are examined and means areprovided to control them.All the processes that could cause thetemperatures to become excessive dependon the same fundamental physical condition.The fuel is a source of heat andthe heat must be removed at about thesame rate that it is generated or thefuel pellets and the cladding will overheat.This is important not only whenthe reactor is operating but also afterthe fission reaction has been stopped.At full power the uranium oxide in thefuel rods is at a much higher temperaturethan the cladding - above 40001F inthe center of some pellets. The storedenergy in the pellets is equivalent toabout 4 sec of full power operation inregions of maximum heat generation, andthe cladding must be kept at an acceptabletemperature until the excess heathas been conducted to the coolant.Removal of this stored energy takes afew seconds, and thereafter only theheat produced by radioactive decay ofthe fission products establishes thecooling requirement. The heat productionrate from this fission productsource is about 7% of full power at theinstant of shutdown and decreases asshown in Fig. IX 3-1. Fuel from a 1000-MWe reactor would still produce 6000kilowatts of heat 30 days after shutdown.Heat is removed from the fuel in normaloperation by circulating water over thecladding at high velocity. A certainmass flow rate of water at the designtemperature and pressure is required tocool the core adequately at it's designpower; substantially more than the requiredflow is provided. The heat productionis not uniform throughout thecore, and this must be taken into consideration.The maximum heat generationrate in the fuel and the heat flux fromIX-9

the cladding to the coolant in normaloperation are 1.5 to 2 times the averagefor all the fuel in the core. A valuewell above 2 is assumed for design purposesto allow flexibility in operationof the reactor. On the basis of thishigher value the coolant flow, temperature,and pressure are specified so thatthe critical heat flux 1 is greater by afactor of 1.8 or more than the localheat flux anywhere in the core. Thisdoes not mean that the reactor powercould be increased by a factor of 1.8without damage to the fuel rods. Itdoes, however, provide considerablemargin to bring upsets under controlbefore the fuel rods are damaged.3.2 ANTICIPATED TRANSIENTS ANDACCIDENTSIf the power level were increased beyondthe capacity of the water flow to removethe heat; if the water flow, temperature,or pressure were changed in such away as to reduce the effectiveness ofthe coolant; or if the water were lostfrom the reactor core, the fuel rodswould overheat. Various events have thepotential for producing such effects.Some involve changes in operating conditions,failure of equipment within theplant, or errors by the operators. Othersinvolve external forces, such asearthquakes and other severe naturalphenomena that can affect the plantequipment. Several levels of safety areprovided in the plant to protect againstand mitigate the effects of such events.1The heat flux at which steam begins toblanket the surface and the claddingtemperature begins to rise rapidly todamaging levels because the steam impedestransfer of the* heat from thefuel to the bulk of the coolant.The plants are designed conservativelyand built, tested, operated, and maintainedin accordance with high qualitystandards to reduce equipment faults andoperating errors to a minimum. Specialrequirements are imposed on the designand operation to prevent external forcesfrom causing failures. Instrumentationand control systems are provided to keepthe plant operating steadily within prescribedlimits.In spite of these measures, some operationalupsets must be expected to occurduring the service life of the plant.The design of the plant is examined inlight of the incidents believed mostlikely to occur and measures are providedto cope with them. One of theimportant measures is a reactor protectionsystem that sounds alarms and holdsor reduces the reactor power when operatinglimits are approached and thentrips the reactor (stops the fissionreaction) and establishes safe shutdownconditions if these limits are reached.A design requirement is that the safetyfeatures should terminate these antici-~pated transients without damage to thefuel rods or the reactor coolant systems.Finally, additional systems and marginsare included in the design to protectthe public in the event that certainhighly unlikely but very severe accidentsoccur. Major failures of plantcomponents and systems are assumed andengineered safety systems are designedto control the accidents that wouldensue. Fuel rods may be damaged andradioactivity may be released from themin these design-basis accidents, but thedesign of the plant and its engineeredsafety features must provide assurancethat the fuel rods will not melt andthat radioactive materials will notescape from the plant in sufficientquantity to cause serious exposures tothe public.IX-10

0C-U_CLC_0.1 1 10 102 103 104 105 106 107 108 109Time (Sec)FIGURE IX 3-1Heat Production of Decay of Fission ProductsFIG. IX 3-1IX-11/12

V.'--

Section 4Protection Against OverpowerSeveral types of events can cause thepower generated in a reactor core torise so that the control or protectionsystem must act to prevent the fuel rodsfrom overheating. A reduction in temperatureor a rise in flow of coolantcauses the average temperature of bothfuel and coolant in the core of a PWR tobegin to fall, and the nuclear characteristicsof the core cause the powerto rise in an effort to hold the averagetemperature near its initial value. Areduction in temperature or an increasein flow of coolant or in system pressurecauses the volume of steam in the coreof a BWR to begin to decrease and thepower rises in an effort to maintain thesteam volume fraction. The fissionreaction is controlled by assemblies ofneutron absorbing material which aremoved in and out of channels in thereactor core. Uncontrolled withdrawalof control assemblies would cause apower transient and increases in fueland coolant temperatures in the core ofa PWR,volumeor in fuel temperature and steamfraction in the core of a BWR.PWR's also use boron dissolved in thecoolant for control purposes. Dilutingthe boron has an effect similar to thewithdrawal of control rods.Power reactors are designed to produceheat to satisfy the demand for steam bythe turbine generator. The core poweris permitted to vary with demand, andthe control systems function to keep thecoolant conditions within limits relatedto the power. Because an excess ofcooling is provided, the core can bepermitted to exceed design full power by10 to 20% before being tripped by theprotection system. Neutron detectorsthat measure the fission power directlyand nearly instantaneously from radiationemitted by the core are the primarysensors for initiating protective actionagainst overpower. The rise in temperatureof the water on passing through thecore is used as a backup indicator ofreactor power to initiate protectiveaction in a PWR.4.1 REDUCTION IN TEMPERATURE OFREACTOR COOLANTThe temperature of the water enteringthe core of a PWR depends on the amountthat the water was cooled in passingthrough the steam generators in thecoolant loops outside the reactor vessel.Reducing the temperature orincreasing the flow of feedwater to thesteam generators or increasing the rateat which steam is withdrawn from themlowers the temperature of the waterentering the reactor core. In a BWR thereactor core is the steam generator.The feedwater is mixed with the coolantas it recirculates to the core, and thetemperature of the coolant is influenceddirectly by the feedwater temperatureand flow.4.1.1 TEMPERATURE AND FLOW OF FEEDWATERSteam that is generated in the steamgenerators of a PWR and in the core of aBWR is passed through a turbine whereabout one-third of the energy is used inturning a generator to produce electricity.The remaining energy is removedin condensing the steam in a condenserat the outlet of the turbine. Thiscondensate is then recycled as feedwaterto the steam generators of the PWR andthe reactor vessel of the BWR. Thethermal efficiency of the steam cycle isoptimized by withdrawing part of thesteam from several points in the turbineand using that steam to preheat thefeedwater as it is being returned to thereactor system. The temperature of thefeedwater delivered to the reactorsystem can be lowered below its normaltemperature by reducing the flow ofsteam to the feedwater heaters or bybypassing some of them. The characteristicsof the feedwater systems are suchthat changes in feedwater temperaturefrom these causes would occur slowly orwould be small. The greatest effect isexpected when the reactor is at fullpower. Analyses indicate that the powerwould rise and stabilize below theoverpower trip level in most feedwatercooling incidents.The rate of feedwater flow to a BWRreactor coolant system is controlled byseveral control valves and/or feedwaterpumps operating in parallel. Malfunctionof feedwater controllers couldcause the flow to increase rapidly.With the reactor at full power, theamount of flow increase in many caseswould not be sufficient to raise thepower to the overpower limit. If, however,the flow were increased to thefull output of the feedwater pumps, anoverpower trip would ensue.IX-13

A rapid rise in feedwater flow to a ratewell below full flow with the reactor atvery low power would result in a powertransient with a peak that could beseveral times the full power level. Insuch an incident a reactor trip would beinitiated as the power passed through 10to 25% of full power (the range of tripsettings when the reactor is at lowpower). The reactivity of the uraniumoxide fuel decreases with increasingtemperature, and the rapid heating ofthe fuel would stop the rise in powerand cause it to fall back rapidly.Insertion of the control assemblies in asecond or two would end the incident.The amount that the feedwater flow couldreasonably increase is limited by thesystem design. Within those limits thetotal energy released during a transientwould not.be enough to damage the fuelrods.A BWR has systems other than the feedwatersystem that supply cold water tothe reactor vessel in normal or emergencysituations. They include a highpressure coolant injection system andshutdown cooling systems. Addition ofwater from any of these sources wouldcause the reactor vessel water levelcontrollers to compensate by reducingthe feedwater flow. The effect would beto lower the temperature of the net flowof water to the vessel, and the resultingtransient would be less severe thanone produced by interrupting the flow ofsteam to the feedwater heaters.4.1.2 INCREASE IN STEAM DEMANDThe demand for steam from the reactorsystem can be increased by opening morefully the turbine control valves, byopening valves that release steam fromthe main steam lines, or by rupture of asteam line. If any of these events wereto occur in a PWR at full power, thereactor core power would rise abovedesign full power, but this could alsohappen with the plant initially at lowerpower. In bringing the plant from lowpower to full power the reactor coremust provide the heat for warming thewater and equipment in the plant systemsas well as for generating the steam forthe turbine. If the load were increasedtoo rapidly, the reactor core powerwould rise above the design full powerin satisfying the total demand.The immediate response to a sudden releaseof steam in a BWR is a decrease inpower. The steam pressure falls, thesteam in the reactor core expands, whichincreases the steam volume fraction anddecreases the reactivity, and the powerfalls. The control system responds byincreasing the coolant flow to the coreto maintain the water level within theoperating range and raises the powerlevel, which restores the pressure.The PWR and BWR plants are designed toaccommodate, without a reactor trip,rates of increase in demand over therange from low power to full power thatare well above those normally imposed bythe electrical transmission system. Theturbines are equipped with load-limitingand/or speed-control devices to preventthe demand for steam from exceeding fullrated flow by more than a few percent.Any malfunction or misoperation thatcaused the steam demand and reactorpower to rise excessively would resultin a reactor trip.The steam systems contain turbine bypassvalves that discharge steam from themain steam lines directly to the turbinecondenser, power-operated relief valves,and code safety valves. These valvesrelease steam under particular conditionsto cool the reactor, control thereactor coolant system pressure in aBWR, and protect the steam system fromexcessive pressure in a PWR. Misoperationor malfunction of one or more ofthese valves would increase the demandfor steam from the reactor coolant system.Numerous valves are used in thesesystems and they operate separately.Opening of one or two valves usually canbe accommodated by the reactor controlsystem without a reactor trip. Theprotection system would trip the reactorif the steam demand raised the powerexcessively.4.1.3 RUPTURE OF A MAIN STEAM LINE INA PWRRupture of a main steam line would causethe largest and most rapid increase inflow of steam of any of the increase-insteam-demandincidents. This is one ofthe low probability accidents for whichmeasures are taken in the design to protectthe fuel rods from melting and thepublic from serious exposure to radiationfrom radioactivity inIt is an overpower incidentthe steam.in a PWRplant. It is a loss-of-coolant accident,and is discussed later as such, ina BWR.A steam line in a PWR might rupture whenthe reactor is operating at any power upto full power, or when the control assembliesare fully inserted and thereactor is shut down but near operatingtemperature. With the reactor operating,the power would rise very rapidlyand the reactor would be tripped by ther"ýIX-14

protection system on indication ofoverpower by the nuclear power sensors.Shutdown would occur quickly enough toavoid damage to the fuel rods.Simply inserting the control assembliesinto the core would not be sufficient tokeep the reactor shutdown under all circumstances.The control assemblies in aPWR are designed to stop the fissionreaction quickly and to keep the reactorshut down at the operating temperature.Boron must be added to the water if thereactor is to remain shut down when thecoolant temperature is lowered. Continuedgeneration of steam in the steamgenerators would lower the reactor coolanttemperature rapidly, and hence thepossibility and consequence of startingthe fission reaction again must beconsidered.Several measures are taken to limit thecooling and to bring the reactor to asecurely shutdown condition.a. Block valves in each main steam linejust outside the containment closeautomatically on indication of abreak in a steam line. If the breakwere outside the containment closingthe valves would in most instancesstop the steam flow. If a blockvalve failed to close, however, connectionsbetween lines would allowsteam to be discharged to theatmosphere until all the water hadboiled out of one steam generator.If the break were inside the containment,the affected steam generatorwould discharge its steam insidethe containment until dry. In anyevent the release would be limitedto one steam generator.b. The flow of feedwater to the steamgenerators is stopped until theaccident is brought under control.C. A concentrated boric acid solutionis injected into the reactor coolantby high pressure pumps to minimizethe rise in fission reaction.The maximum rate, magnitude, and durationof cooling would occur if thereactor were at hot shut down with thecontrol rods fully inserted and a steampipe were to rupture at a steamgenerator. Safety analyses for a typicalplant indicate that the coolingwould allow the fission reaction tobegin within 15 sec after a steam linerupture. The nuclear power would riserapidly and stabilize at about 30% offull power. Concentrated boron solutionwould begin to enter the reactor coolantloops in about 90 seconds, about thetime that the steam generator would boildry, and the reactor would be shut downa few seconds later.In the steam-pipe-rupture incident thepower reached by the reactor on coolingwould be well below full power. However,if the control assembly of greatestworth were to stick in the fullywithdrawn position during the initialshutdown, the power distribution in thesubsequent transient with all hut onecontrol assembly in the core would beseverely peaked. The maximum heat generationrate in the fuel near theposition of the withdrawn assembly atthe lower power might exceed the maximumat full power allowed for in the design.In studies of this situation, it hasbeen found that there is no danger ofthe fuel rods overheating even with thereactor coolant pumps stopped and thewater circulating by thermal convection.With the reactor shut down at operatingtemperature, the sticking open of apressure-relief valve or a turbinebypass valve would cause a coolingincident similar to the steam line ruptureincident. Although the fuel rodswould not be damaged, it is not desirableto have the reactor resume operationin such events involving a largeblowdown which are likely to happen morefrequently than main steam line breaks.By use of a large number of smallvalves, the r~ate of cooling caused bythe sticking open of one valve is reducedand the effect can be compensatedby the boron injection system.Steam would be discharged into the containmentin some of these steam-releaseincidents and to the atmosphere in allof them. 1 If, during normal operation,reactor cooling water leaks into thesteam system through flaws in the tubingor tube sheets of the steam generators,the steam discharged in an incidentwould contain some radioactivity. Whenthe break is inside the containment,limiting the blowdown to the contents ofone steam generator ensures that thepeak pressure in the containment wouldbe below the design pressure and theradioactivity would be confined, as inthe loss-of-coolant accidents describedin section 6. when the break is outsidethe containment, all the steam releasedwould pass into the atmosphere. The1Steam is released to the atmosphere tocool the reactor in cases where thesteam line block valves close.IX-15

potential radiation doses to the publicare small because the concentration ofradioactivity in the reactor coolant,the rate of leakage of reactor coolantinto the steam system, and the amount ofradioactivity in the steam system arerequired to be at low levels.4.2 INCREASE IN FLOW OF COOLANT TOREACTOR COREThe reactor coolant pumps in a PWR runat a fixed speed and the coolant loopshave no flow control valves; hence theflow cannot be increased in normal operation.However, PWR's have multiplecoolant loops and 3 or 4 coolant pumpsand are permitted to operate at reducedpower with one pump out of service.Since uncontrolled startup of an inactiveloop could, in some circumstances,cause a substantial rise in reactorpower level, protection against such anevent is included in the design. TheBWR makes use of a variable speed pumpin each of two loops to control the rateof recirculation of coolant through thecore and, thereby, the reactor power.Increasing the flow lowers the steamvolume fraction in the core and thepower rises to restore it. A malfunctionof the pump speed controls thatincreased the flow or startup of aninactive loop would produce a powertransient.Two types of startup of an inactive loopin a PWR must be considered - one inwhich a pump has been idle but the loophas no isolation valves or the valveshave been left open, and the other inwhich the loop has been isolated byclosing valves in the inlet and outletlines. The latter is of greatest concernbecause it presents the possibilityof adding rapidly, to the bulk of thecoolant, water that is cool and low inboron content.With the pump stopped but the loop opento the remainder of the reactor coolantsystem, water flows in the reverse directionthrough the inactive loop andthe water in the steam generator tubesis cooled to a temperature below that ofwater being delivered to the reactorcore by the operating loops. To bringan inactive loop into operation withouttripping the reactor, operating proceduresrequire that the plant be broughtto low power in order to equalize thetemperatures before starting the pump.Without this precaution, the rise inreactivity of the core caused first bythe increase in flow and then by thedecrease in coolant inlet temperaturewould be too rapid to be compensated bythe control system. The nuclear powerwould rise to the trip point in lessthan a second and the protective systemwould trip the reactor. The peak powerwould be well above full design power,but the heat flux from the cladding tothe coolant would not reach the fullpowerheat flux during the transient andthe fuel rods would not be damaged.In the instance of the isolated loop,simply opening the isolation valves andstarting the pump could, under the mostadverse conditions, cause a very largerise in core power and possibly damagethe fuel. Therefore administrative controlsrequire the power to be reduced tozero and the temperature and the boronconcentration of the water in theisolated loop to be near those in theactive loops before the isolation valvesare opened. Because these administrativecontrols are judged to give insufficientprotection, interlocks are alsoprovided in the reactor protection systemto ensure a low probability ofstartup of an isolated loop containingwater at low temperature and boronconcentration. These interlocks preventthe valve in the line from the dischargeof the isolated coolant pump to thereactor vessel from being opened unlessthe water in the isolated loop has beenrecirculated and mixed slowly, through arelief line, with the water in theremainder of the reactor coolant system.Because of the protective interlocks,the most drastic misoperation likely tooccur would be to begin the startup inviolation of administrative controlswith the isolated loop containing waterat low temperature and boron concentration.The flow through the relief lineis limited by the line's size, and thuswater from the isolated loop would mixslowly with the bulk of the reactorcoolant. If the water in the isolatedloop contained no boron, several minuteswould be required before the shutdownmargin of the control assemblies wouldbe surpassed and the fission reactionwould begin. The reduction in shutdownmargin would be indicated as it occurredby the neutron flux monitors in thecontrol room. The operators would haveample time to take corrective actionwhich would include injection of concentratedboric acid to regain the shutdownmargin.In the BWR, a malfunction of equipmentthat controls the pump speed to increasethe speed at the maximum rate has agreater effect than does startup of anidle loop. In either case, the mostsevere transient is found to occur whenthe initial power is 1,70% of full powerIX-16

and the flow is \,50% of full flow. Inthe case of malfunction of speed controls,the reactor would be tripped onindication of overpower by the neutronsensors. The thermal neutron flux wouldpeak at a value above 200% of full powerbut the transient would be so short thatthe heat flux from the cladding to thecoolant would at no time exceed thefull-power heat flux. In the case ofstartup of an idle loop, the characteristicsof the recirculation pump andpower supply and of the jet pumps thatit drives combine to make the incident a'-V, minor one. The reactor core power wouldrise briefly to only a few percent abovethe full-power value before settling ata new steady-state condition and thereactor would not be tripped.4.3 INCREASE IN PRESSURE IN A BWRAn increase in reactor system pressurecompresses the steam in the core of aBWR and the power rises rapidly in aneffort to restore the steam volume fraction.The reactor is equipped with apressure regulating system, but thissystem cannot reasonably be providedwith the capacity and extra featuresneeded to prevent the pressure fromrising when the turbine control valvesclose rapidly, or the turbine blockvalves or the isolation valves in themain steam lines close. The turbinecontrol valves close rapidly on rejectionof a large percentage of the loadby the electrical transmission system.The turbine block valves close on anysignal that calls for a turbine orreactor trip. The main steam line isolationvalves close to prevent steamfrom being discharged to the atmosphereon indication of certain potentiallyunsafe situations in the reactor, thecontainment, or the steam system. Misoperationor malfunctions can also causethe valves to close.Rapid closing of the turbine controlvalves or stop valves interrupts thesteam flow in a fraction of a second. Areactor trip is initiated when thesevalves close, and normally the pressurecontrol system opens turbine bypassvalves to release steam to the turbinecondenser. The pressure rises sufficientlyin a few seconds to cause reliefvalves to open briefly to limit the presurerise, but the control assembliesinsert neutron absorbing material rapidlyenough to keep small, any increase inpower beyond rated full power.The incident is more severe if the turbinebypass valves cannot open due tothe condenser being out of service.Relief valves open as the pressure risesfrom the nominal operating pressure of1050 psi to a peak just above 1200 psi.The core neutron level rises briefly toabout double that at full power but theheat flux rises only a few percentbefore the power transient is terminated.Closing the main steam line isolationvalves produces a transient similar tothat produced when the turbine blockvalves close and the turbine bypass systemis inoperative. The transient isless severe, however, because the isolationvalves close less rapidly. Thepressure in a transient produced by theclosing of one turbine block valve orsteam line isolation valve is controlledby the turbine bypass valves alone, butthe power increases enough to cause thereactor to trip from overpower.4.4 UNCONTROLLED MOVEMENT OFCONTROL ABSORBERAlthough the control systems are designedto prevent the heat flux in thereactor core from rising to excessivelevels, malfunctions or accidents involvingthe control systems can causesuch a rise. Three types of situationare considered in the reactor design:(1) uncontrolled withdrawal of neutronabsorber by the control systems, (2)insertion of a control assembly into thecore with the reactor at power, and (3)ejection of a control assembly from thecore under pressure or gravity forces.The first and third result in increasesin power, the second in a change in thepower distribution and an increase inthe local heat flux in the core. Theconsequences of these incidents are limitedby limiting the worth of eachcontrol assembly and/or the rate ofmovement and using a large number ofassemblies to satisfy the total neutroncontrol requirements of the reactor.4.4.1 WITHDRAWAL OF CONTROL ASSEMBLY BYDRIVE MECHANISMSThe uncontrolled withdrawal of a controlassembly might occur with the reactorshut down or at power. Several provisionsof the design, limit the rate atwhich the reactivity of the core can beincreased in such an incident. Thespeed of the drive mechanism is limitedby the design, and the number of assembliesthat can be moved at one time islimited by interlocks in the controlsystem and by the sequencing of power tothe drives.IX- 17

With the reactor initially at zero power,continuous withdrawal of a controlassembly could cause the core power torise very rapidly. The incident wouldbe terminated automatically by reactortrips initiated at low, intermediate,and high power by signals from theneutron sensors. Under the most adverseconditions of the highest worth controlassembly being withdrawn and time delaysin the transmission of signals and operationof trip mechanisms, the neutronlevel might rise to almost 10 times thefull design power, but only for afraction of a second. The power risewould be stopped and the power would hemomentarily reduced by the inherentnegative reactivity effect from heatingof the fuel. Insertion of the controlassemblies, which takes 1 to 2 secIwould shut down the reactor. Because ofthe short duration of the transient, thetemperatures of the fuel and claddingand the maximum heat flux from thecladding to the water would not reachthe full-power values.The effects of uncontrolled withdrawalof a control assembly have been studiedover the full range of power for a widerange of withdrawal rates and assemblyworths. Under most circumstances theincident is minor and the fuel rods areprotected by a reactor trip on indicationof overpower by the neutron sensors.This overpower trip does not,however, provide the desired degree ofprotection in a PWR when the reactivityis increased slowly. The steam controlvalves to the turbine limit the rate ofgeneration of steam by the steam generatorsand thereby the cooling of thewater in the reactor primary system.The water temperature rises as the powerincreases and the critical heat fluxdecreases with rising water temperature.A trip based on coolant temperature andreactor power is included in the reactorprotection circuits for this situation.A reactor trip is initiated when thetemperature rise of the reactor coolantin passing through the core (a measureof the reactor power) exceeds the temperaturerise allowable for the prevailingcoolant temperature in the reactor.4.4.2 DILUTION OF BORON IN PWR COOLANTDuring operation of a PWR, water isrecycled at a small rate but continuouslyfrom the reactor coolant system,which is at high pressure, to a chemicaland volume control system, which operatesat low pressure. The boron concentrationin the reactor coolant can bereduced, and the reactivity of the corethus increased, by returning unboratedwater to the reactor coolant system.Boron dilution is a manual operationthat is conducted under administrativecontrols. Valves must be opened andpumps started by the operator in orderto add unborated water to the reactorcoolant system. The chemical and volumecontrol system is designed to limit themaximum rate of dilution to a valuethat, after indication through alarmsand instrumentation, provides the operatorwith several minutes to correct thesituation before the reactor is trippedby the protection system.After a trip, the operators have manyminutes to stop the dilution before thereduction in boron concentration causesthe fission reaction to begin again withthe control assemblies fully inserted.The procedures made erroneous dilutionunlikely, numerous alarms and indicationsalert the operators to the condition,ample time is available to correctthe situation, and therefore no otherprotection against a dilution incidentis provided.Dilution during startup is similar tothe uncontrolled withdrawal of controlassemblies, except that reactivity isadded more slowly. The reactor operatorswould be alerted to the changingboron concentration by a graduallyincreasing count rate on the nuclearinstruments in ample time to takecorrective action.4.4.3 INSERTION OF A CONTROL ASSEMBLYDriving (or, in a PWR, dropping) one ormore control assemblies into the corewith the reactor at power, or leavingone in when the reactor is brought topower, reduces the heat generation ratein the fuel in the vicinity of theinserted assembly. The heat generationrates in other parts of the core mustincrease in order to hold the totalpower constant and thus the heat generationrate increases in these otherregions. The same effect is obtained ifthe control assemblies are not keptproperly positioned relative to eachother during operation.During operation, control assemblies aremoved in a preselected sequence to maintainproper alignment. The positionsare shown by indicators and lights.Alarms sound when assemblies deviatefrom prescribed positions by more than afew percent. In some designs, movementof control assemblies is inhibited whenalignment limits are reached. Accidentalinsertion of a control assemblywould shut down the reactor or would beindicated (1) by a sudden drop in power,IX- 18

(2) by an asymmetric power distributionsensed by the neutron power detectors(also by thermocouples in the waterleaving the core in a PWR) , and (3) bythe various position indicators andalarms.The effect of control assembly misalignmentis a factor considered in establishingthe value for the maximum heatflux allowed by the reactor design. Ifa reactor were operated at full powerwith one control assembly fully insertedand the others aligned within thelimits, the maximum heat flux from thecladding to the coolant would still bewell below the critical heat flux.Since studies indicate that serious situationsare not likely to arise fromcontrol-assembly-insertion incidents,the operator is largely depended on totake any corrective measures that nightbe needed.4.4.4 EJECTION OF A CONTROL ASSEMBLYThe drive mechanisms for the controlassemblies are mounted on the top headof the reactor vessel in a PWR and aresupported from the bottom head in a 'BWR.If a drive mechanism housing were torupture, the pressure inside the reactorvessel would tend to expel the driveshaft from the vessel. The attachedcontrol assembly would be ejected fromthe core very-rapidly, with consequencesto the fuel rods and reactor coolantsystem that would depend on the reactivityworth of the assembly.Use of boron in the reactor coolant forpart of the neutron control and use ofmany control assemblies makes the fullworth of an individual assembly small ina PWR. The housings are manufactured tohigh standards of quality so that theprobability of a drastic failure sufficientto eject a control assembly issmall. Furthermore, the results of arod ejection accident in the PWR ispredicted to result in limited ornegligible fuel rod damage. Because ofthese considerations, no mechanical featuresare provided in the design of aPWR to prevent rupture of a housing fromcausing a control assembly to be ejectedfrom the core.Although many control assemblies areused in a BWR the full worth of individualassemblies is greater than in a PWR.Structural members are thus incorporatedin the BWR design to prevent the driveshaft from moving more than a shortdistance in the event of a housingrupture. Several types of failurecould, however, occur in a drivemechanism that would permit a controlassembly to drop out of the core. Thecontrol assembly incorporates a devicethat limits the velocity of fall inorder to reduce the consequences of suchan accident.In most instances, a control assemblyejection incident would not produce aserious power transient. When the reactoris at power, the control assembliesare partly to fully withdrawn. In manyinstances, the power would not rise tothe overpower trip level, in others theprotective system would trip the reactoron indication of overpower and thetransient would be terminated with thefuel rods undamaged. When the reactoris shut down with the control assembliesfully inserted, ejection of one wouldnot produce a rise in power.Under some circumstances, ejection of acontrol assembly could result in adamaging transient. Because the probabilityof such an accident is very low,the risk of limited damage to some fuelrods is accepted provided that only asmall percentage of the fuel rods wouldbe likely to fail and the failures wouldnot cause a damaging pressure surge inthe reactor coolant system.By subjecting fuel rods to rapid powertransients in test reactors, it has beenshown that the damage sustained by afuel rod can be related to the totalenergy generated in the fuel pelletsduring a transient. These data are usedin assessing the consequences of controlassembly ejection accidents in a PWR andcontrol assembly drop accidents in aBWR. Analyses of plants of acceptabledesign show that less than 10% of thefuel rods would be perforated during themost severe accidents. The pressure inthe coolant system would rise moderatelybut would not impose appreciable additionalloads on the coolant system.In the case of rupture of a control rodhousing, reactor coolant would bedischarged into the containment and theaccident would also be a small-breakloss-of-coolant accident. Measures takento prevent the fuel rods from beingdamaged further and to protect thepublic from radioactivity in such accidentsare discussed in section 6. Inthe case of a control rod drop accidentin a BWR, no coolant would be lost andthe fission products would be confinedinitially to the reactor coolant systemand the steam system. Isolation valvesin the main steam lines would close in afew seconds on indication of a highlevel of radioactivity in the steam andIX-19

thus limit the amount of radioactivityin the steam system. The steam systemalso would be isolated by closing valvesto ensure that onlythe radioactiveiodine could escapea small fraction ofkryton, xenon, andto the surroundings.IX-20

Section 5Protection Against a Reductionin Core CoolingThe cooling effectiveness of the reactorcooling water is reduced by changes incooling conditions that lower the criticalheat flux. A reduction in flow, arise in temperature, or a reduction inpressure of the coolant are suchchanges. The cooling capacity, particularlyin some types of abnormal occurrence,depends on the inventory of waterin the steam generators of a PWR and inthe reactor vessel of a BWR. Changes inoperation that reduce the inventory andthus the cooling capacity include reductionin flow of feedwater to or increasein flow of steam from the steam generatorsin a PWR or the reactor vessel in aBWR.The design approach to protectingagainst reductions in cooling effectivenessand capacity is to:a. Design the plant so that such eventswill be infrequent and the magnitudeand rate of reduction of coolingcapability will be within limits.b. Provide control and protection systemsthat will, on indication of areduction in cooling effectivenessor capacity, reduce the power ortrip the reactor, depending on themagnitude of the upset.C. Ensure that the cooling capacityafter a reactor trip will be sufficientto dissipate the energy storedin the fuel at high temperature andthe fission-product decay heat.5.1 REDUCTION IN REACTORCOOLANT FLOWLoss of power to one or more of thereactor coolant recirculation pumpmotors, seizure of a pump rotor, orshaft breakage would cause the flow ofcoolant to the reactor core to decrease.The reduction in flow results in areduction in the critical heat flux andthe core power must be reduced to preventthe fuel rods from overheating.The pumps of a PWR are equipped withflywheels to increase the inertia of therotary elements so that the pumps coastdown slowly and the flow decreases graduallyover many seconds. The combinedinertia of the pumps and the motor generatorsets used to control the pumpspeed serves the same purpose in a BW1R.Seizure of a pump rotor or shaft breakagewould decrease the flow in onecoolant loop abruptly, but, because ofthe number of loops in a PWR and thecharacteristics of the jet pumps in theBWR, the flow through the core would notfall below about 60% of full flow.In a PWR, as the flow decreased the coretemperature would begin to rise and thepower to fall. Until the average temperatureand temperature rise of thecoolant in passing through the corereached the operating limits, the reactorcontrol system would withdraw controlrods to hold the power constant.This would aggravate the situation, sothe protective system is arranged totrip the reactor on indication of lossof power to one or more pumps or lowflow in one or more coolant loops.In a BWR, as the flow decreased thevolume fraction of steam in the corewould begin to rise and the power andthe rate of steam production would beginto fall. The pressure regulating system,in order to hold the pressuresteady, would gradually restrict theflow of steam to the turbine, the reactorwould not be tripped, and the powerwould stabilize at a level appropriateto the reduced flow. The density differenceresulting from the production ofsteam in the reactor core induces asubstantial circulation of coolant whenthe pumps are stopped.Stoppage of all the recirculation pumpswould cause the greatest reduction inflow, but the initial rapid reductionwhile the reactor is still at full powermakes seizure of a pump rotor or shaftbreakage the most severe reduction-incoolant-flowincident. Seizure of apump rotor and shaft breakage are eventswith such low probabilities of occurrencethat damage to a small percentageof the fuel rods could be tolerated.However, analyses show that the criticalheat flux in decreasing with thedecreasing flow would approach but stillexceed the maximum heat flux in the coreand thus the fuel rods are not likely tobe damaged.5.2 MISMATCH BETWEEN FEEDWATER ANDSTEAM FLOWSA reduction in feedwatersteam generators of a PWR,flow to theor to theIX-21

eactor vessel in a BWR that is notcompensated by a reduction in flow ofsteam to the turbine, results in adecrease in the inventory of water inthe respective steam generators andreactor vessel. An increase in steamflow uncompensated by an increase infeedwater flow has the same effect.Conversion to steam of inventory in thePWR steam generators and in the BWRreactor vessel provides the cooling forthe reactor until water can be suppliedfrom other sources when the normal flowof feedwater has been interrupted forany reason. A reduction in inventoryrepresents a reduction in coolingcapacity.Events that can cause a reduction infeedwater flow include malfunctions offeedwater pumps, valves, and controllers,interruption of power to the pump,and breaks in feedwater lines. Malfunctionsof valves and controllers andbreaks in lines can cause the steam flowto increase. Signals to the reactorprotection system that indicate, directlyor indirectly, excessive steam flowor excessive steam flow relative tofeedwater flow and low water level inthe steam generators or reactor vesselare taken as indicators of reduction ofthe water inventory toward the minimumacceptable volume. The protective systemtrips the reactor and the turbine toreduce the steam flow and heat generationrate. The main feedwater pumps maycontinue to function and supply water tothe reactor systems, but an auxiliaryfeedwater system in a PWR and a highpressure coolant injection system in theBWR are provided to supply water automaticallywhen necessary.Some potential incidents are not terminatedquickly by these actions andadditional measures are required tobring them under control. A break in afeedwater line inside the containment orin a steam line inside or outside thecontainment could result, in the case ofa PWR, in the release of all the waterand steam in one steam generator. Theother steam generators contain enoughwater to cool the reactor, and otheraspects of such an incident were discussedin section 4.1.3.In a BWR, a break in a feedwater line ora steam line inside the containmentgives rise to a loss-of-coolant accidentthat is discussed in section 6. If thebreak occurs outside the containment,valves in the feedwater lines and isolationvalves in the main steam linesclose in a few seconds to stop the flowof water and steam. Some of the radioactivityin the fluids is released inthe turbine building and thence to theatmosphere. The potential doses to thepublic are limited by using valves thatclose rapidly and limiting the amount ofradioactivity in the reactor coolant.5.3 REDUCTION IN STEAM DEMANDWhen a reactor is operating, the fuelrods are cooled, directly in a BWR andindirectly in a PWR, by release of steamto the turbine. Throttling the flow bythe turbine control valves to satisfy areduction in demand for electricity, orclosing the turbine control valves, turbineblock valves, or steam line isolationvalves, for whatever cause, reducesthe flow of steam and the cooling availableto the core. If the reduction indemand for steam is partial and slow,the reactor control system can simplyreduce the fission rate to compensate.At times, however, the steam demand mustbe reduced more rapidly than the controlsystem can reduce the fission rate withouttripping the reactor. Also, theenergy stored in the fuel and fissionproducts and reactor systems at highpowertends to sustain the high heatgeneration rate when the fission rate isreduced. Therefore, a system is providedto maintain sufficient steam flowto cool the fuel and keep the pressurein the steam and reactor systems fromrising excessively while new steadystateconditions are being established.This is necessary whether the power isreduced partially or the reactor istripped.The steam relief system consists ofturbine-bypass valves, power-operatedrelief valves, and ASME code safetyvalves on each main steam line. In theevent of a rapid reduction in steamdemand, the bypass valves are opened bythe controllers and release steam to theturbine condenser at the rate requiredto hold the steam pressure in a BWR andthe steam pressure and reactor coolanttemperature in a PWR within the operatinglimits while the reactor power isreduced to match the demand of the turbine.The bypass valves have a capacityof 25 to 50% of full steam flow so thatthe plant can reduce load, without areactor trip, at rates well above thosenormally required to satisfy the demandsof the electrical transmission system.A reactor trip is initiated automaticallyby a turbine trip from any powerlevel above the design capacity of thebypass system, and a turbine trip isinitiated by any reactor trip. After atrip, the turbine bypass system usuallyrelieves the steam necessary to cool theIX-22

eactor to the shutdown condition. Whenthe pressure in the turbine condenser istoo high or the steam line isolationvalves are closed, the bypass systemcannot be used, and in some transientsthe bypass valves do not have enoughcapacity to hold the pressure within thedesign limit. Under these conditions,the power-operated relief valves andcode safety valves discharge steam tothe atmosphere in a PWR and the suppressionpool or containment in a BWR tolimit the pressure rise. Discharge tothe atmosphere is acceptable in the caseof the PWR, because the steam is notgenerated from the reactor coolant andcontains much less radioactivity. Whensteam is discharged through the reliefand safety valves, feedwater is drawnfrom a large supply of storedcondensate. In the BWR, water can alsobe supplied from the suppression pool.The steam pressure control and reliefsystem acts directly to prevent thepressure from becoming excessive in thereactor coolant system of a BWR. However,it affects the pressure in thereactor coolant system of a PWR onlyindirectly, through its effects on thecoolant temperature. Provisions arerequired in a PWR to protect the coolantsystem from being overpressured by expansionof the water as its temperaturerises in the more severe transients thatcan follow a turbine trip. Sprays inthe pressurizer condense steam from thevapor space to hold the pressure nearthe normal level. If the pressure risesabove the set points of the poweroperatedrelief valves and code safetyvalves, they discharge vapor from thepressurizer to reduce the pressure.Reactor trips on high-pressurizer pressureand high-pressurizer level areincluded in the protective system asbackup to the automatic reactor trip ona turbine trip. They also ensure thatthe reactor coolant system pressure willnot be excessive and that there will beample free volume in the pressurizer atthe time of a turbine trip.5.4 REDUCTION IN PRESSURE IN REACTORCOOLANT SYSTEMA reduction in pressure in the reactorcoolant system reduces the effectivenessof the coolant by lowering the criticalheat flux and also portends a reductionin inventory and cooling capacity. In aPWR, pressure reductions are associatedwith faults in the pressurizer such asbreaks in pipes connected to the vaporspace or the sticking open of pressurizerrelief valves. In such an event,the fuel rods are protected initially bya reactor trip on indication of a lowerthan-allowablepressure. If caused by apipe break or an unclosable relief valvethe incident evolves into a loss-ofcoolantaccident requiring the additionalactions described in section 6. In aBWR, pressure reductions are ordinarilyassociated with breaks in steam lines orwith malfunctions of the pressure controlsystem or steam bypass and reliefvalves, all in the direction of increasingsteam flow. This type of incidentwas discussed in section 5.2.IX-23

Section 6Protection Against Loss ofReactor Coolant4coolantLoss of reactor coolant can be a seriousevent. It could result from misoperation,malfunction, or failure of equipment,such as a break in a reactor coolantpipe, that allowed uncontrolleddischarge of water or steam from thereactor coolant system. It constitutesa failure of a primary barrier to theescape of radioactivity from the plant.The escaping fluid must be confined.More important, loss of all coolant fromthe reactor core for minutes could leadto melting of fuel rods followed by destructionof the reactor vessel and penetrationof the containment. Because ofthis, many provisions are included inthe design to prevent loss-of-coolantincidents and to mitigate the consequencesof any that do occur.Measures taken include the following:a. High standards are specified for thedesign, construction and operationof the reactor so that the number ofmisoperations, malfunctions, andfailures of equipment in the reactorcoolant system will be minimal.b. Means are provided for detectingsmall rates of leakage from the reactorcoolant system so that in mostinstances the plant can be shut downto repair the equipment before theleakage becomes serious.c. The reactor coolant system, its auxiliariesand associated engineeredsafeguards are designed so that thefuel rods are not likely to be damagedin the most probable loss-ofincidentsand would be cooledacceptably in the most seriouspostulated accidents.d. Features incorporated in the reactorcoolant system and the containmentand its related engineered safetyfeatures are designed to ensureconfinement of all but a smallamount of radioactivity in all typesof loss-of-coolant accidents.As suggested by (c) and (d) above, thereare two general areas of concern in theinitial phase of a loss-of-coolant accident.The fuel rods must be cooled sufficientlyto prevent them from meltingand also to prevent the cladding frombecoming so embrittled by the reactionof Zircaloy with steam that it mightshatter on cooling and permit the fuelpellets to fall into an uncoolable mass.The water, steam, and radioactivity dischargedfrom the reactor coolant systemmust be confined. After the initialphase, the fuel rods and the containmentmust be cooled to remove the heat generatedby the radioactive decay of thefission products and measures must betaken to ensure continued confinement ofthe fission products.6.1 COOLING OF FUEL RODS6.1.1 SMALL-BREAK ACCIDENTSLoss-of-coolant accidents are classifiedaccording to the size of opening throughwhich the fluid is discharged. In asmall-break accident the water and steamwould be expelled from the reactor coolantsystem relatively slowly and thesystem would remain at a pressure ofseveral hundred psi until almost empty.The fuel rods would be uncovered and thecladding would melt before water couldbe added at low pressure, so a highpressure coolant injection system isprovided to maintain sufficient inventoryin the reactor vessel to cool thefuel rods until the system is depressurizedand at low temperature.The high pressure coolant injectionsystem is designed to ensure that thefuel rods will be covered by water, orby water and the froth produced byboiling, when the water in the reactorvessel reaches its lowest level duringthe incident. Small-break incidents aremore severe when the break is in a waterline, because a larger rate of coolantloss occurs. A high pressure pumpingsystem of reasonable capacity can maintainthe normal water level when theopening is smaller than the equivalentof a hole about 3/4 in. in diameter andcan maintain the required coverage ofthe fuel rods for breaks up to about 6in. in diameter. The duration of theaccident - the time during which thesystem depressurizes and, in the largerbreaks where the water level must bereturned to above the top of the core -may vary from about 15 min. for a 4-in.break to several hours for a very smallbreak. Since the fuel rods are wellcooled, they are expected to survive asmall-break accident undamaged.IX-25

In some PWR's the high pressure coolantinjection system makes use of the normalcharging pumps operated up to full capacityto inject borated water into thereactor coolant system. In others, separatepumping systems are provided. Ifthe break is quite small, (< 3/4 in. indiameter) the pumps can maintain theoperating level in the pressurizer andthe operators can shut the reactor downand cool it by normal procedures. Forlarger breaks, the reactor trips onindication of low pressure and low levelin the pressurizer, and the cooling isconducted in accordance with emergencyprocedures. In either event, steam isgenerated in the steam generators andreleased through the turbine bypassvalves or the relief valves to cool thereactor coolant until the temperatureand pressure are reduced to the level atwhich the reactor shutdown cooling systemcan function. Feedwater is suppliedto the steam generators by the mainfeedwater system initially and then bythe auxiliary feedwater system. Therefueling water storage tank provides alarge reserve of borated water for injectioninto the reactor coolant system.Measures taken in the BWR are similar.If the break is small enough that thereactor core isolation cooling systemcan maintain the water level in thereactor vessel, the plant is shut downand depressurized by normal procedures.For larger breaks or those requiring themain steam line isolation valves to beclosed, the high pressure coolant injectionsystem automatically supplieswater. Depending on the circumstances,steam is released through the turbinebypass valves or the relief valves-atthe rate required to cool the fuel rods.Water added to the vessel is drawn froma large supply of condensate or from thecontainment.6.1.2 LARGE-BREAK ACCIDENTSIn a large-break accident, all the waterand steam in the reactor coolant systemwould be discharged into the containmentin a time from a few seconds to about 2min. No way of preventing the core frombeing voided of liquid has been devised;so systems are provided to restore coolingbefore the fuel rods are damagedexcessively. One of these is a lowpressure coolant injection system thatpumps water into the reactor coolantsystem at a high rate. It is not practicalto provide a pumping system withthe capacity to refill the reactor vesseland reflood the core in the timeallowed under some postulated accidentconditions, so additional measurrs aretaken.PWR's are equipped with an accumulatorsystem consisting of several tankspartly filled with cool borated water,pressurized with nitrogen, and connectedby separate lines to the reactor coolantsystem. Water from the accumulatorspasses into the coolant system throughcheck valves when the pressure falls tothe range of 600 to 300 psi, dependingon the particular design, and the accumulatorsdischarge until empty. Thequantity of water is sufficient to refillthe reactor vessel to a level atleast high enough in the core so thatthe fuel rods will be cooled adequatelyby the low pressure coolant injectionsystem as it completes the reflooding.In BWR's a core spray system sprays alarge volume of water onto the top ofthe fuel elements to cool the fuel rodsuntil the core is reflooded by the lowpressure coolant injection system.Large breaks range in size between theequivalent of about a 6 in. diameteropening and the two openings produced byseverance of the largest pipe (about 30in. diam) in the reactor coolant system.The instantaneous severance of a largepipe in the reactor coolant system -generally referred to as the designbasis loss-of-coolant accident - is themost serious accident for which protectionis provided in a nuclear powerplant. Rupture of the reactor vesselwould be more serious but extra measuresare taken in design, manufacture, andtesting to produce vessels of highestquality and in operation to ensure thatthe vessels will be operated well withinthe design specifications. These measuresmake the probability of rupture ofa reactor vessel so small that protectionagainst such an event is deemedunnecessary.The severance of a large pipe is also ahighly improbable event but by design ofthe safety features to protect againstsuch a drastic event, the pipes are providedwith considerable margin over thecapacity required to cope with the moreprobable smaller breaks. To be certainof this, the performance of the safetyfeatures is analyzed for a wide range ofpossible break sizes and locations inwater and steam lines.The most severe accident is that resultingfrom a large break in the pipe betweenthe discharge of a reactor coolantrecirculation pump and the inlet to thereactor vessel. The sequence of eventsin such an accident in a PWR is describedbelow with reference to Fig. IX2-2. Figure IX 2-2 shows the configurationof a PWR reactor coolant system andIX-26

containment and the location of thebreak.In a fraction of a second followingseverance of the pipe, the rush of waterthrough the break and into the containmentcauses the pressure and level inthe pressurizer to fall to the setpoints. This initiates a reactor tripand actuates the coolant injection systems.Actually, the pressure decreasesso rapidly that boiling begins in thereactor core in less than a second. Thenegative reactivity effect of liquidbeing displaced by vapor in the corestops the fission reaction, and insertionof the control rods augments thiseffect.D~uring depressurization, or blowdown, ofthe reactor coolant system, water andsteam flow from the reactor vessel tothe break through inlet and outletlines. The direction and rate of flowthrough the core vary with time, and thecladding temperature is calculated torise and fall as the flow rate and qualityof the coolant change in the core.The critical heat flux falls below themaximum heat flux generated early in thetransient, so the temperature of thefuel rod cladding rises well above normaloperating temperature. Since thefission reaction has been stopped, themajor concern is that the cooling besufficient to remove the heat stored inthe fuel pellets and the fission productdecay heat without the cladding exceedinga safe temperature.Analyses indicate that in less thanabout 20 sec after the break all but asmall amount of water is discharged fromthe reactor coolant system into the containmentso that the pressure in thecontainment and that inside the coolantsystem are about equal. The accumulatorsbegin to discharge into the coolantsystem at about 10 sec, but it isassumed that the accumulator water alsoleaves through the break until the endof the blowdown.Following the blowdown phase there is arefill period of about 15 sec duringwhich the liquid from the accumulatorsfills the reactor vessel to the level ofthe bottom of the fuel rods. Assumingloss of off-site power and delays inbringing the emergency supply into operation,the low pressure coolant injectionpumps begin to deliver boratedwater from the refueling water storagetank to the reactor coolant systemsomewhat later in the refill period.The fuel rods are essentially uncooledduring this period and the temperaturerises about adiabatically.Reflooding of the core begins when thewater reaches the bottom of the fuelelements. As it rises into the core, itis heated by the lower ends of the fuelrods and begins to boil. When the waterlevel is about 1 ft above the bottom ofthe core, the boiling becomes vigorousenough to entrain droplets of water inthe steam. As the water level rises,the mixture of droplets and steam providesincreasingly effective cooling ofthe fuel rod surfaces above the waterlevel. The accumulators are empty atabout somewhat less than one minute, andat about the same time the cooling atthe hot spot becomes sufficient to terminatethe temperature rise. The pumpscontinue to inject water, the peakcladding temperature falls gradually,and the water level rises to the top ofthe core completing flooding within afew minutes after the break. The reactorsystem continues to fill until waterflows out the break in the pipe. Provisionsfor long-term cooling are broughtinto operation after most of the waterin the refueling water storage tank hasbeen pumped into the reactor coolantsystem and the containment.The scenario for a design basis accidentin a BWR is similar. Because of differencesin design, the coolant flow isupward through the core throughout blowdownand depressurization takes aboutone minute. The water level drops belowthe bottom of the core in less thanabout 30 sec. The core sprays and thelow pressure coolant injection systemsbegin to add water from the suppressionpool at about 30 sec. The core is refloodedand the temperature rise terminatedwithin a few minutes after thebreak.Because the probability of a large-breakloss-of-coolant accident is low, somedamage to the fuel rods is acceptable.The cooling is designed to prevent thecladding from melting or being seriouslyembrittled during the transient. Thefuel rods and the piping and equipmentare designed so that movement producedby the blowdown forces cannot impedeoperation of the safety systems; thusthe fuel rods should remain in a coolableconfiguration throughout the accident.Analyses indicate that only asmall fraction of the fuel rods wouldreach a temperature at which the claddingwould swell, rupture, and releasethe fission products from the gas space.6.1.3 BREAKS OF INTERMEDIATE SIZEProtection must also be provided againstbreaks of intermediate size. In the PWRthe pressure at which the accumulatorsIX-27

egin to discharge is selected so thatthe high pressure coolant injection systemand the accumulators combine to keepwater in the core. In the BWR an automaticdepressurization system operatesthe relief valves on the steam lines toreduce the pressure in the reactor coolantsystem rapidly if the high pressurecoolant injection system is unable tomaintain the water level above the corein the reactor vessel. This provides abackup to the high pressure injectionsystem for breaks of small and intermediatesize. With the reactor depressurized,the core spray and low pressurecoolant injection systems cool the core.6.2 CONTAINMENT OF RADIOACTIVITYDURING ACCIDENTA large-break loss-of-coolant accidentpresents the greatest challenge to thereactor containment. Practically allthe water and steam in the reactorcoolant system is expelled into thecontainment in less than a minute, andthe stored energy in the fluids and heatreleased by the fuel flash part of thewater into a large volume of steam. Thecontainment must be sealed tightly toprevent the escape of radioactivity. Itmust be capable of withstanding a moderatepressure and must have a large volumeor the capacity to condense much ofthe steam as it is released.Two general types of containment haveevolved; one characteristic of PWRs, theother of BWRs and they are shown inFigs. IX 6-1 and IX 6-2. The typicalPWR containment is a cylindrical buildingwith a domed roof, all constructedof heavily reinforced concrete. Concreteis difficult to keep gas tight sothe containment has a welded steel linerattached to the inside face of the concrete.Entrance to the containment whenfuel is in the reactor is through anairlock. Electrical wiring and pipingpass through special sleeves embedded inthe concrete and welded to the liner.The thick concrete structure providesstrength to withstand the pressure producedby the flashing of reactor coolantinto steam and it shields persons outsidefrom radiation emitted by fissionproducts that are dispersed in the containmentatmosphere. In a typical containmentthe internal pressure in alarge break accident would rise to about40 psig. The containment is designed towithstand about 45 psig and is testedabove 50 psig.The BWR containment consists of a smallerstructure, called a drywell, aroundthe reactor coolant system that iscoupled by vent pipes to a pressuresuppression chamber. The fluids areexpelled from the reactor coolant systeminto the dry well in a loss-of-coolantaccident and the rise in presssure inthe dry well causes the steam to flowthrough the vent pipes into the pressuresuppression chamber. The pressure suppressionchamber contains about one milliongallons of water; the vent pipesterminate several feet below the waterlevel; the steam disperses in bubblesand condenses as it rises through thewater. The large volume of water alsoserves as one of the sinks for heatafter the blowdown in a loss-of-coolantaccident. Steam is discharged directlyinto the containment by the safetyvalves and into the suppression pool bythe relief valves in incidents such asoverpower and reduction-in-cooling discussedpreviously.Typically, the dry well and the pressuresuppression chamber, the latter a torus,have been constructed as steel tanksinside a reinforced concrete structure.The pressure in the dry well and thatabove the water level in the pressuresuppression chamber are estimated torise to less than 45 and 30 psig, respectively,in a design basis loss-ofcoolantaccident. The containment isdesigned for a pressure above 50 psigand is tested above 60 psig.The BWR containment is inside a reactorbuilding that provides an additionalbarrier to the escape of radioactivityin an accident. A standby gas treatmentsystem is put into operation on indicationof radioactivity in the building.Air from the building is passed throughfilters to remove radioactive particlesand iodine before it is discharged up astack to the atmosphere.Many steam lines, water lines, and otherpipes penetrate the containment andprovide a multitude of potential pathsfor radioactive gases and liquids topass outside. The lines are equippedwith isolation valves. Valves in linesthat have no safety function in anaccident close automatically on signalsthat call certain of the safety systemsinto operation.Making a large containment with its manypenetrations completely gas tight isvery difficult, if not impossible.Proving that one is gas tight might beequally difficult, so a reasonablyattainable and demonstrable limit isplaced on the leakage. Generally, theleakage at full design pressure must beless than 0.1 to 0.5% per day of thecontainment volume, depending on theIX-28

effectiveness of other provisions forconfining the radioactivity. Conformancewith this requirement must bedemonstrated prior to operation andperiodically during the life of theplant.6.3 POST- ACCIDENT HEAT REMOVAL ANDCONFINEMENT OF FISSION PRODUCTSThe loss-of-coolant accident is stabilizedand under control when the fuelrods are reflooded and cooled belowabout 300OF and the reactor coolant systernand the containment are at about thesame low pressure. Cooling of the fuelrods must be continued after the accidentto remove the fission product decayheat at the rates indicated in Fig. IX3-1. Although the containments are designedto withstand the accident pressurefor long times, good sense dictatesthat the potential hazard should be reducedby lowering the pressure and thusreducing the mobility and leakage of theradioactive iodine, one of the principalhealth hazards.6.3.1 COOLING OF FUEL RODSDuring the accident and early in thepost-accident period, the fuel rods arecooled by water injected by the accumulator,core spray, and low pressurecoolant injection systems. In the PWRthe principal source of water is therefueling water storage tank, whichcontains about 300,000 gallons of coolborated water. in the BWR the water isdrawn largely from the pressure suppressionpool.After the injection phase, systems forlong-term cooling are put into operation.Basically, these systems recirculatewater from the containment - thesumps in the PWR and the pressure suppressionpool in the BWR - through heatexchangers and back to the containment.Part of the cooled water is pumped intothe reactor coolant system and returnsto the containment through the break. Areliable supply of cooling water fromthe heat exchangers is provided by theplant service water and high pressureservice water systems. The residualheat removal system~, which cools thefuel during normal shutdown operation,and other pumping systems are combinedto provide a reliable post-accident heatremoval system. The way in which thisis accomplished differs considerablybetween plants.6.3.2 COOLING OF CONTAINMENT ANDRETENTION OF FISSION PRODUCTSSpray systems are provided in the PWRcontainments to condense the steam,absorb radioactive iodine, and washradioactive particles from the containmentatmosphere. When the containmentpressure reaches the range of 10 to 25psig, these systems go into operationautomatically and rain droplets downthrough the containment atmosphere fromnozzles in the dome. Initially, thewater is supplied from the refuelingwater storage tank. As it leaves thetank, sodium hydroxide or alkalinesodium thiosulphate is added to reactwith iodine and hold it in solution.Later, part of the water recycled by thepost-accident heat removal systemreturns to the containment through thespray systems.In some designs the spray system aloneis depended on to cool the containment,condense the steam, and lower the pressure.In others the air circulation andcooling system that conditions the airin the containment during normal operationis designed to provide part of thecapacity. This system may also containcharcoal filters for absorbing iodine.The containment spray and air circulationsystems do not have enough capacity,nor can they go into operationrapidly enough if only emergency poweris available, to affect the initialpressure rise in a large pipe breakaccident. They function to return thepressure to near the normal operatingpressure and terminate or greatly reduceany leakage of radioactive gases in anhour or less.Sprays are installed in BWR containmentsto cool the gases in the drywell and inthe space over the pool in the pressuresuppression chamber. Because transferof heat to the large surface of water inthe pressure suppression chamber and inthe drywell after an accident is expectedto provide adequate cooling, use ofthe sprays is optional. Water is suppliedto the sprays from the pressuresuppression pool. It contains no additivesto immobilize iodine, since thelarge volume of water is somewhat effectivein holding the iodine. Gases thatleak from the containment pass into thereactor building and can be processed(filtered) to remove iodine before beingreleased from the plant.6.3.3 CONTROL OF HYDROGEN CONCENTRATIONIN CONTAINMENTDuring the temperature transient in aIX- 29

loss-of-coolant accident, steam isexpected to react with some of theZircaloy cladding of the fuel rods toproduce zirconium dioxide and hydrogen.In the PWR, the alkaline spray solutionwill react with aluminum in the containmentto produce aluminum oxides, hydroxides,and hydrogen. Radiation emittedby the fission products will decomposewater in the core and containment regionsproducing hydrogen and oxygen. Ifthe hydrogen were permitted to accumulatein air in the containment, it couldignite after the concentration reachedabove 4 vol %. Recombiners are providedin PWR containments to prevent the hydrogenconcentration from rising to theflammable limit. The BWR containmenthas a nitrogen atmosphere when the reactoris operating so a flammable mixturecannot be produced immediately followingan accident.6.4 SPECIAL CONTAINMENT SITUATIONSThe containment and its engineeredsafety features are designed to hold thewater, steam, and radioactivity thatwould be released in a break, large orsmall, inside the containment in anywater line, reactor coolant line, steamline, or piece of equipment except amajor break of the reactor vessel. Thenumber of safety features that mustoperate in an accident depends on thesize of the break, the magnitude of therise in containment pressure, and theneed for emergency cooling. The role ofthe steam line block valves in limitingthe escape of radioactivity in the eventof water-line and steam-line breaks outsidethe containment has been mentionedpreviously. Incidents associated withrupture of a tube in a PWR steam generator,damage to a fuel element in handlingduring refueling or storage, andrupture of storage tanks containingwaste liquids or gases must also beconsidered in the design.Leakage of reactor coolant throughcracks developing in steam generatortubes during operation has occurredfrequently in PWR plants. In mostinstances the rate of leakage was smalland the plant could continue to operatesafely; the leaking tubes were pluggedduring a scheduled maintenance period.The probability of complete tube ruptureis very small, but rupture of one tubein one steam generator is considered inthe design of the plant. This is asmall-break accident and the reactorcoolant is discharged into the steamsystem, also a high pressure system.Many signals indicate the beginning ofthe accident, trip the reactor and theturbine, and bring the safety systemsinto operation. Among these is indicationof a high radiation level in thenoncondensable gases drawn from theturbine condenser by its vacuum systemand in the normal liquid blowdown fromthe steam generators. The gaseous dischargecan be diverted to the containmentin some PWR designs and the liquidblowdown is interrupted by block valves.This limits the amount of radioactivegases released.The next step is to cool the reactorcoolant system and bring its pressuredown to the level of the pressure of thesteam system in order to stop the transferof liquid. Normally, this is doneby releasing steam through the turbinebypass valves to the condenser, and theradioactivity is confined. If the bypassvalves cannot be used, the steammust be discharged to the atmosphere.After this initial reduction in pressure,which takes about 1/2 hour, theaffected steam generator can be isolatedand cooling of the reactor and steamsystems can be completed by use of theother steam generators and cooling systems.Potential radiation doses to thepublic are kept small in this incident,as in other steam release incidents,by limiting the amount of radioactiveiodine in the reactor coolant.If an assembly of fuel rods were droppedduring refueling, the cladding of someof the fuel rods might be damaged enoughto release fission products from the gas;pace. The refueling of a PWR is done.nside the containment and the fuel isstored in a pool in an adjacent building.The containment would confine anyradioactivity released there and theventilation system of the fuel storage.building is provided with filters andabsorbers to reduce the release of iodineto the atmosphere. The BWR is opento the reactor building during refuelingand the fuel is stored in the reactorbuilding. Protection is provided by thestandby gas treatment system in thereactor building.Radioactive liquids are processed andstored in tanks in the reactor buildingor in well-ventilated auxiliary buildings.The buildings are designed sothat liquid leakage from the processingequipment and storage tanks drains tosumps from which it can be pumped toother tanks; alternatively the tanks arein pits designed to hold any leakage.Any gaseous leakage is discharged fromthe plant through the ventilation system.Potential doses to the public arekept small by limiting the amount ofradioactivity in the gas in any tank.IX-30

125 Ton Crane LIE - T.V. CameraPrimaryCoolantPumpRemovableFIGURE IX 6-1Typical PWR Containment

ReactorBuildingSprayDrywel(PrimaryContainment)ReactorVesselSpray -VentPipePressure- SuppressionPoolFIGURE IX 6-2Typical BWR ContainmentFig. IX 6-1 - Fig. IX 6-2IX-31/32

Section 7Measures to Ensure Performance of Safety FunctionsHaving provided special features in thedesign of the plant for cooling the fuelrods and containing the radioactivity,there must be assurance that the equipmentand systems will function whencalled upon. High quality of design,construction, and operation and redundancyand testability of equipment andsystems are important factors in providingthis assurance.7.1 QUALITY ASSURANCEThe dependence on high quality has beenmentioned several times in previous sections.Use of appropriate standards andcriteria helps to ensure that theplants are designed carefully and conservatively.The minimum requirementsfor the principal design criteria arespecified by the ABC in Appendix A of 10CFR Part 501 of its regulations.Criteria for quality assurance programsfor the design, construction, and operationof each nuclear facility arespecified in Appendix B of 10 CFR Part50. AEC Regulatory Guides describe acceptablemethods for implementing specificparts of the regulations.The technical societies, industry, andthe ABC and its laboratories have beenworking for several years to developappropriate codes and standards fornuclear power plants. Results of theirefforts are embodied in the ASME 2 Boilerand Pressure Vessel Code Section III,"Nuclear Power Plant Components," ANSI 3B31.7 "Nuclear Piping," IEEE 4 Standard279 "Criteria for Protection Systems forNuclear Power Generating Stations," andin many other codes and standards thatgovern the design, construction, and1Title 10 - Atomic Energy - of the Codeof Federal Regulations, Part 50 Licensingof Production and utilization Facilities.3American National Standards Institute.4Institute of Electrical & ElectronicsEngineers.2American Society of Mechanical Engineers.operation of nuclear facilities. Aspart of the design, specifications areprepared to ensure that appropriate materialsand procedures will be used inconstructing the plant. The design criteria,particularly of equipment andsystems important to safety, are reviewedand judged to be adequate by theABC regulatory staff and the AdvisoryCommittee on Reactor Safeguards before aconstruction permit or an operatinglicense is authorized for a plant.7.2 PROTECTION AGAINST EFFECTS OFSEVERE NATURAL PHENOMENACareful consideration of the effects ofsevere natural phenomena on the safetyof the plant is important in providing aconservative design. Studies are madeof local and regional geology, hydrology,meterology, and seismology and ofhistorical records. On the basis ofthese studies the plant structures andequipment of importance to safety aredesigned to withstand the effects ofearthquakes, tornados, hurricanes,floods, tsunamis, and seiches withoutloss of capability to perform theirsafety functions.Generally this means that the structuresmust be capable of withstanding tornadicwinds having a tangential velocity of300 mph and a translational velocity of60 mph. Protection must be providedagainst missiles, such as heavy plankstraveling at high velocity and automobilestraveling at low velocity, thatmight be propelled by such winds. Dependingon the site, the structures andequipment must also be capable of withstandingthe effects of the winds ofhurricane force.The plant must be built on ground sufficientlyhigh so that equipment andstructures would not be reached by thewaters of a flood of the greatest magnitudethat can reasonably be projectedfor the site or barriers against floodingmust be incorporated in the design.Wave action generated by hurricane orstorm winds at the time of a maximumflood and the effects of seiches andtsunamis, where appropriate, are importantfactors in establishing the maximumwater level and the protection that mustbe provided.IX-33

Because earthquakes occur quickly andwithout warning and can develop forceslarge enough to destroy piping systemsand structures, much attention is givento designing nuclear power plants towithstand the shaking caused by earthquakes.Earthquakes of two magnitudesare considered: the Safe ShutdownEarthquake and the Operating BasisEarthquake. The Safe Shutdown Earthquakeis the most severe earthquake thatcould reasonably be conceived to occurat the plant site considering the regionaland local geology and seismologyand the characteristics of the localsubsurface material. Some damage wouldoccur, but certain parts of the plantmust be designed to remain functionalwhen subjected to the vibratory groundmotion produced by such a severe earthquake.They are the structures, systems,and components necessary to ensurea. The integrity of the reactor coolantpiping and vessels,b. The capability to shut down thereactor and maintain it in a safeshutdown condition, andC. The capability to prevent or mitigatethe consequences of accidentsthat might result in appreciableradiation exposures to the public.The Operating Basis Earthquake is onethat could reasonably be expected toaffect the site during the operatinglife of the plant. The vibratory motionfrom such an earthquake might cause theplant to shut down, but the reactorplant must be designed to remain in acondition such that operation can becontinued safely.7.3 RELIABLE POWER SUPPLYThe safety of a nuclear power plantdepends on the availability of directcurrent electricity for operating the,instrumentation and control systems andalternating current for operating pumpsin some of the safety systems. Twoseparate battery and distribution systemsare installed to provide a reliablesource of direct current. Operation ofeither provides sufficient capacity foressential loads.Several sources are provided to ensure asupply of alternating current, and detailsof accomplishing this may differbetween plants. Each plant is connectedto a regional power grid through atleast two separate l "ines. On shutdownýof the plant turbine generator, powerwould normally be supplied from the gridto continue the operation of essentialequipment. However, loss of off-sitepower can occur, so emergency suppliesare provided on-site. Usually the onsitesource is two or more diesel generatorsets, but at some plants gasturbine-drivengenerators or nearbyhydrogenerators provide diverse sourcesof emergency power. At least twosystems are installed for distributingthe alternating current electricitywithin the plant. Functioning of onedistribution system and part of theemergency generators provides sufficientpower for operation of essentialequipment. The on-site power sourcesand distribution systems are designed tosafety standards that make them capableof delivering power during floods,storms, earthquakes and accidents.7.4 REDUNDANCY IN SAFETY SYSTEMSA basic assumption in the plant designis that some of the equipment in thesafety systems will not function or willoperate at reduced capacity in anemergency; so extra, or redundant,systems are incorporated. Measurestaken in the design of the electricalpower supplies is one example. Thedesigns of engineered safety featuresfor the containments and the emergencycore cooling systems show otherexamples.In some PWR containments sprays aloneare depended on to cool the containmentand remove fission products from theatmosphere after an accident. Two spraysystems are installed, each with morethan enough capacity to handle the totalload. Each spray system has one or morespray headers in the containment domeand its own pumps, valves, and piping tosupply borated water to the sprays. Thetwo spray systems may be interconnectedto increase the number of flow paths andpumps available to each system. Twovalves in parallel may be installed atcritical locations in the piping so thatfailure of one valve to open will notblock the flow. In other PWR containmentsonly one spray system is used, andthe air circulation system that coolsthe containment in normal operation isequipped with charcoal filters to absorbfission products and designed tofunction in a loss-of-coolant accidentenvironment. This combination has theredundance of two 'spray systems andgreater diversity, thus reducing theprobability of a common mode failurecausing the simultaneous loss of bothcooling systems.Redundancy of cooling forin the reactor core in anprovided in several ways.the fuel rodsemergency isIn a typicalIX-34

BWR, the High Pressure Coolant InjectionSystem goes into operation on indicationof low level in the reactor vessel orhigh drywell pressure. Only one pumpand piping system is provided, but thepump is driven by a turbine thatoperates on steam produced by theresidual heat in the reactor andrequires only direct current power tofunction. Redundancy is supplied by theAutomatic Depressurization System, whichlowers the pressure in the reactor bydischarging steam to the suppressionpool if the High Pressure CoolantInjection System cannot maintain thelevel. Five relief valves are used andfour valves give amiple capacity. Afterthe pressure is reduced, the Core SpraySystem and the Low Pressure CoolantInjection System can each provide morethan the required Capacity of waterdelivery and they go into operationsimultaneously. The Core Spray Systemconsists of two separate pumping systems,each of which has the requiredpumping capacity. Each system has 2-50%capacity pumps, and its own piping,valves, instrumentation, and sprayheader in the reactor vessel. The LowPressure Coolant Injection System hastwo separate pumping systems, each discharginginto a recirculation loop ofthe reactor coolant system. Each systemhas two pumps, and operation of three ofthe four pumps will provide the requiredcapacity.The PWR's do not have an AutomaticDepressurization System so a typicalreactor will contain two separate highpressure injection systems with sufficientpumps so that partial operation ofthe systems' pumps delivers the flowrequired for small break loss-of-coolantaccidents. The accumulators requireonly the decrease in reactor coolantsystem pressure that accompanies alarger break to bring them intooperation. Since the location of thebreak could prevent the contents of oneaccumulator from reaching the reactorvessel, they are designed so that one oftwo, two of three, or three of fourinstalled units will have the requiredcapacity. Redundant low pressure pumps,each having the full design capacity,are provided for injecting and recirculatingcooling water after the pressurehas fallen. They may also function toprovide coolant delivery to the containmentspray system.Any of several relevant parameters beingout of limits may automatically initiateoperation of safety system, e.g., lowliquid level in the reactor vessel orhigh pressure in the drywell in the BWRstarts the High Pressure CoolantInjection System. The reactor operatorcan start any of the systems manually.At least, three sensing circuits areprovided to monitor each importantparameter. Out-of-limits indication bytwo of three sensors initiates safetyaction. This permits a circuit to betested without interrupting the reactoroperation, but the protection system isdesigned so that a faulted circuit orone under test gives one of the twosignals required to start a safetysystem.7.5 TESTABILITY OF SAFETY SYSTEMSAs a final measure, the safety systemsand their instrumentation and controlsmust be testable. Some tests, such as ahigh pressure test of the containment,need be conducted only a few timesduring the life of the plant. Overalltests of most of the engineered safetyfeatures are conducted during therefueling interval which occurs annuallyin most plants, but some systems aretested semiannually. The operability ofmany of the components of the safetysystems, such as pumps and valves, istested at least monthly and some of theinstrumentation is checked daily. Thetests are designed to show, as near aspractical, that the safety featureswould be brought into action and wouldfunction properly if a situation requiredtheir use.IX-35

Section 8ConclusionNuclear Power plants are equipped withmany special safety features to preventexcessive amounts of radioactive materialfrom being released in normal operationand in abnormal situations. Ifthese features perform as the designersexpect, few, if any, of the fuel rodswould be damaged and little radioactivitywould be released during reactortransients or accidents. The risk tothe public would be very small.However, in assessing the potentialhazard it must be assumed that only onsiteemergency power will be available,that it will be delayed in starting, andthat the safety features will functionat reduced capacity in an accident.Although the designers must show thatthe fuel rods will not melt under thesecircumstances, the performance of thecontainment features in the design basisloss-of-coolant accident is evaluated onthe basis of a release of fission productsfrom the fuel that could beproduced only by melting. In order forthe measures taken in the design of theplant to be considered satisfactory, thedesigners must show that the probableradiation doses to the most heavilyexposed members of the public would bewell within AEC guidelines.For incidents that might occur occasionallyin the lifetime of a plant, thepotential doses must be far below levelsthat are considered to be harmful. Forserious accidents, which have a considerablylower probability of occurrence,the guideline doses are 25 rem to thewhole body and 300 rem to the adultthyroid apply. Calculations of thesedoses are based on conservative estimatesof fission product release fromthe fuel and removal of airborne fissionproducts by natural phenomena and engineeredsafety features so that in realitythe average doses received by amember of the general public would befar below the potential dose to the mostexposed individuals located nearer theplant site boundary.IX-37

WASH- 1400(NUREG 75/014)DESIGN ADEQUACYAPPENDIX XtoREACTOR SAFETY STUDYU.S. NUCLEAR REGULATORY COMMISSIONOCTOBER 1975

Appendix XTable of ContentsSection1.2.3.INTRODUCTION ....................................................RESULTS .........................................................SUMMARY .........................................................Page No.X-1X-3X-15APPENDIXADESIGN ADEQUACY EVALUATION OF SELECTED SAFETY RELATEDEQUIPMENT AND STRUCTURES IN NUCLEAR POWER GENERATINGPLANTS ......................................................X-17List of TablesTablePage No.X 2-1X 2-2X 2-3X 2-4X 2-5PWR Component Review - Seismic Design AdequacySummary .........................................................PWR Component Review - Environmental Qualification Summary ......BWR Component Review - Seismic Design Adequacy Summary ..........BWR Component Review - Environment Qualification Summary ........Summary of Bases for Reduced Safety Margin Determination ........X-5/6X-7/8X-9/10X-11/12X-13/14X2-6Summary of Items for Which Design Adequacy Could Not beAssessed ........................................................X-13/14X-i

Section 1IntroductionThe quantitative assessments of failureprobabilities performed in tbis studyare based on component failure rate dataobtained from various sources, as discussedin Appendix III. However, thisdata base generally does not includeinformation pertinent to some of thespecific conditions under which certainnuclear components important to safetyare required to operate. These conditionsinclude loadings due to externalphenomena such as earthquakes and tornadoes,and the adverse environmental conditionsthat certain equipment locatedinside the reactor containment buildingsmay be subjected to following potentialloss of coolant or transient-causedaccidents (e.g., coincident high temperature,high pressure, high humidity andhigh radiation levels). Since thesephenomena are not included in thegeneric data base and could potentiallylead to common mode failures that couldaffect the probability of occurrence ofvarious accidents, assessments were madeof the ability of specific componentsand structures to perform as designedwhen subjected to these phenomena. Thisdesign adequacy assessment was done todetermine whether: (1) these phenomenarepresent a common mode failure potential;and (2) if dependencies do exist,they are factored into the probabilitypredictions.In addition to the above, the genericdata base did not include componentdesigns built under the quality assurancerequirements presently applicableto nuclear power plants. As indicatedin Appendix II the liziated infoxmationcurrently available om nuclear •comp•nentfailure rates seims to indicate Tlttledifference between nuclear ,and monnuclearcomponent failure rates. Thus,this study has found little basis lorusing failure rates for n-mclear :compunentsthat are lower than those fr •n~nuclearcomponents. The exception ttothis rule is that of the probability ofgross reactor vessel failure. As -discussedin Appendix V, the data on thelikelihood of failure of conventibonalpressure vessels has been reduced by afactor of 10 for nuclear pressure resselsdue to the high standards and attentiongiven in the design, fabrication,inspection and testing of nuclearvessels.The design adequacy assessment was performedfor the Reactor Safety Study bythe Franklin Institute Research Laboratories(FIRL). The general approachused was to take a sample of thecomponents, systems and structures thatare engineered safety features or areclosely associated with them and tocheck the adequacy of the implementationof safety design and specialized environmentaltesting requirements. Some ofthese samples were checked for seismicdesign adequacy and others to see ifthey have been properly designed andqualified for operation in the unusualpost-accident environments. In addition,the structures housing the dieselgenerators were examined with regard totornado design.X-1

Section 2ResultsTables X 2-1, X 2-2, X 2-3 and X 2-4summarize the results of the evaluations.The detailed report of the FIRLinvestigation is attached as Appendix A.As shown in Table X 2-1, 30 PWR itemswere examined with regard to seismic design.Of these, 25 were found to beadequate (83%). Design adequacy was notdemonstrated for five items (17%) (reactorcoolant pump nozzles, low head safetyinjection system instrumentation,recirculation spray pump outside containment,the diesel generator day tank,and the AC and DC switchgear), becausesufficient information was not availableto permit an assessment of adequacy tobe made. For three items (the containmentcrane, the low head safety injectionpumps, and the reactor protectionsystem), it was found that the designwas adequate in that failure is not expectedunder seismic excitation. However,the margin to failure was found tobe less than that normally expected consideringapplicable code and qualificationrequirements because either: (1)errors were found in assumptions used incalculating stresses; or (2) seismicqualification tests were not sufficientlycomprehensive or were not performed.The adequacy of the tornado and tornadomissile design was evaluated for thestructure housing the diesel generators,the doors to the structure, and fueltank soil cover. It was found that thestructural design and the soil coverwere adequate. However, in the event ofimpact of a tornado missile on the doorof the structure, it was found thatwhile the door should survive the impact,the outward rebound of the doormay fail the door pins and cause thedoor to open. This potentially couldexpose a diesel-generator to adverse environmentalconditions (e.g., rain,small missiles, hail, etc.). The probabilityof diesel failure under theseconditions has not been assessed, but itshould be considerably less than unity.Furthermore, the loss of a single dieseldoes not create an immediate safetyproblem since: (1) the remaining dieselis available should off-site power beinterrupted by a tornado; and (2) thesteam-driven portions of the auxiliaryfeedwater system would also be availableto remove decay heat even if all ACpower is lost.Table X 2-2 indicates that nine PWRitems were examined with regard to environmentalqualification. All werefound to be adequate although five items(reactor coolant piping snubbers, lowhead safety injection system snubbers,the low head safety injection ssystempumps, the containment recirculatingspray pump located outside containment,and the reactor protection system) weretested under conditions which were notas comprehensive as tests performed topresent standards.BWR components were also evaluated.Thirty-two BWR items were evaluated withregard to seismic design as shown inTable X 2-3. Of these, 28 (88%) werefound to be acceptable. For two ofthese items (6% of the total evaluated)(the recirculation lines and the reactorprotection system), it was found thatalthough the design was adequate, thesafety margin is less than normallyexpected. Adequacy could not beassessed for four items (13%) (themissile barriers inside containment,reactor pressure relief valves, corespray system instrumentation, and the480-Volt load centers) because sufficientinformation was not available topermit an assessment to be made.The BWR diesel-generator housing wasalso analyzed with regard to tornado andtornado missile design. The design ofthe structure was found to be adequate.Insufficient information was availableto permit an assessment of the adequacyof the door to withstand the reboundfollowing missile impact or the impactof a missile on the corrugated metalbarrier attachments. With regard toenvironmental qualification, as shown inTable X 2-4, eight items were examined.Six (75%) were found to be acceptable.Of these, the reactor coolant systemsnubbers were found to be acceptablewith reduced safety margin becausetesting had not been performed. Insufficientinformation was available on theremaining two items (25%) (reactorpressure relief valves, and the electricalcables, terminations and connectors)to permit an assessment of theiradequacy.For both plants, 62 items were examinedwith regard to seismic design. Thirtysix(58%) were found to be acceptable byx-3

satisfying code and qualificationre---quiriements, 12 (i1%), were found'- accept_-a-bl-e based on engineering- judgment-, andfive.% (•)% were found to be a-aceptabl:ewith rediuced margin. Thus, 53: items:(85%V of. those analyzed for: seismic:des:igýY were found to be- acaeptable-.Design:adequacy was not demonstrated for9 items (15% of the total examined) dueto lack of sufficient information.With:-. regard to tornado design, thediesel-generator housings were studiedon both plants. Both structures werefound-to be adequate. The PWR buildingexterior door was found to be likely-tofail whentrebounding after impact of atornado missile. Insufficient informationwas available to assess the adequacyof the BWR door in this regard, orof the corrugated metal barrier attachmentsfor the BWR diesel generatorbuilding.Sevenrteen items were evaluated concern--ing.envi~ronmental qualification on- bothplanrts. Six: (35%) were found-to meetqualification requirements, three (18.%)we-reý found acceptable on the basis- ofengineering judgement, and four (23%)were- foundi to be adequate with-reducedmargin mainly because the tests were: notsufficiently' comprehensive. Thus., 13('7T6V o~fthe total evaluated) were deemedto-- be- acceptable. Design adequacy wasnot demonstrated for four items (24%- ofthe-tOtal)- because of lack of sufficientinf6rnration- to permit an assessment tobe- made.Tables-,X 2ý75. and X 2-6 indicate thebasesfbr determining adequacy with reducedmargin and insufficient informationto assess- adequacy, respectively.X-4

6TABLE X 2-1 PWR COMPONENT REVIEW - SEISMIC(a) DESIGN ADEQUACY SUMMARYReview Data Basiso o • Design Criteria-H H4 2 Satisfied(b)Adequate (Failure Not Expected)Assessment of Design AdequacyNot Demonstrated (Failure Possible)W 01 Analysis Design Criteria Insufficient Information ReviewedSubsection V) W - a and/or Engineering Not Completely Information to Indicates that_4 ') (a IVin Text Component Description > 0 a u E. Test Judgment Satisfied (c) Assess Adequacy Failure is Likely NotesA6.3.1 Reactor BuildingA6.3.1.1 Soil-Structure Interaction Model / /A6.3.1.2 Containment Internal Structure / / (d)A6.3.1.4 CraneA6.3.2 Reactor Coolant System (RCS)A6.3.2.1 Reactor Coolant System Loop /A6.3.2.2 Steam Generator and Pump Supports V /A6.3.2.3 Reactor Coolant Pump Nozzles / (e)A6.3.2.4 Pipe Whip Restraints:Main Steam LineFeedwater Line//// (d)A6.3.2.5 Snubbers MA6.3.3 Low Head Safety Injection System (LHSIS)A6.3.3.1 Piping / / (d)A6.3.3.2 Pumps and Drives /A6.3.3.3 ValvesA6.3.3.4 Motor Operators /A6.3.3.5 Snubbers and Hangers (f)A6.3.3.6 InstrumentationA6.3.4 High Head Safety Injection Systems (HHSIS)A6.3.4.1 Accumulator Tank Nozzle / / (d)A6.3.4.2 Accumulator Piping Connection to RCSA6.3.4.3 Charging Pumps and Drives /A6.3.5 Containment Recirculation Spray SystemA6.3.5.1 Pump and Motor Located Outside Containment:Pump / /Motor / / /A6.3.5.2 Motor Drives Inside Containment /A6.3.6 On-Site Electric Power SystemsA6.3.6.1 Diesel Generator Housing:(a)Walls and Roof / I /Door / /Soil Cover for Fuel Tanks /A6.3.6.2 Diesel Generators /A6.3.6.3 Day Tanks /A6.3.6.4 Air Bottle Supports /A6.3.6.5 Batteries and Battery Supports /A6.3.7 Electric Power Distribution SystemsA6.3.7.2 Electrical ContainmentPenetrations and Connectors / /A6.3.7.3 Cable TraysA6.3.7.4 AC and DC SwitchgearA6.3.8 Reactor and Engineered SafeguardsProtection Systems Sensors andLogic CabinetsA6.3.9 Intake Canal// /

(a)Diesel Generator Building was evaluated only for tornado resistance.(b) Design Criteria as stated in FSAR; or current criteria, if original criteria were not explicitor found to have insufficient margin.(c)More sophisticated analysis methods may in some cases show that the design criteria have infact been satisfied; in all cases however, loss of function is not expected.(d) Deficiencies in original analysis have been either revised or investigated further. Revisedresults have been evaluated and design is found to be adequate.(e)(f)Bijlaard formulae not applicable.Periodic seal replacement in the snubbers is recommended.Table X 2-1X-5/6

TABLE X 2-2 PWR COMPONENT REVIEW - ENVIRONMENTAL QUALIFICATION SUMMARYReview Data BasisAdequate (Failure Not Expected)C: 4J0 0 W Design Criteria-H 04 V4 o Satisfied(a)Assessment of Design AdequacyNot Demonstrated (Failure Possible)( 4 p 0 Analysis Design Criteria Insufficient Information ReviewedSubsection ( and/or Engineering Not Completely Information to Indicates that.1. a) 0 et Jdmnin Text Component Description > oSatisfied (b)0 • Test Judgment St Assess Adequacy Failure is Likely NotesA6.3.1 Reactor BuildingA6.3.1.3 Paint Coatings within Containment /A6.3.2 Reactor Coolant System (RCS)A6.3.2.5 Snubbers / / (c)A6.3.3A6.3.3.2Low Head Safety Injection System (LHSIS)Pumps and Drives /(d)A6.3.3.5 Snubbers / / (c)A6.3.4 High Head Safety Injection Systems (HHSIS) (d)A6.3.5 Containment Recirculation Spray SystemA6.3.5.1 Pump Located Outside Containment / /A6.3.5.2 Motor Drives Inside ContainmentA6.3.6 On-Site Electric Power Systems/ / /(d)A6.3.7 Electric Power Distribution SystemA6.3.7.1 Electrical Cables and Terminations " / /A6.3.7.2 Penetrations and Connectors / /A6.3.8 Reactor Protection System / /(a) Design Criteria as stated in FSAR; or current criteria, if original criteria were not explicit or found to have insufficient margin.(b)More sophisticated analysis methods may in some cases showis not expected.that the design criteria have in fact been satisfied; in all cases however, loss of function(c) Periodic seal replacement in the snubbers is recommended.(d)Active components of this system that are outside of containment and not subject to LOCA environment have not been included.Table X 2-2X-7/8

ifA

TABLE X 2-3 BWR COMPONENT REVIEW - SEISMIC(a) DESIGN ADEQUACY SUMMARYReview Data Basis0 o Design Criteria0 Satisfied(b)Adequate (Failure Not. Expected)Assessment of Design AdequacyNot Demonstrated (Failure Possible)Analysis Design Criteria Insufficient Information ReviewedSubsection and/or Engineering Not Completely Information to Indicates thatin Tet Component Description Test Judgment Satisfied (c) Assess Adequacy Failure is Likely NotesA6.4,1 Containment StructuresA6.4.1.1 Primary Containment Structure " /A6.4.1.2 Reactor Building (Secondary Containment) / /A6.4.1.3 Containment Piping Penetrations ' /A6.4.1.4 Suppression-Chamber-to-Drywell VacuumBreaker Valves /A6.4.1.5 Missile Barriers Inside Containment /A6.4.2 Reactor Coolant SystemA6.4.2.1 Reactor Pressure Vessel and Internals / VA6.4.2.2 Reactor Pressure Vessel Nozzles / /A6.4,2.3 Reactor Pressure Vessel Skirt / /A6.4.2.4 Recirculation Lines V VA6.4.2.5 Main Steam Lines /A6.4.2.6 Main Steam Line Isolation Valves / / /A6.4.2.7 Pipe Whip Restraints for RecirculationLines /A6.4.2.8 Reactor Pressure Relief Valves /A6.4.3 Core Spray SystemA6.4.3.1 Piping /A6.4.3.2 Hangers and Snubbers:HangersSnubbersA6.4,3.3 Pumps and Drives /A6.4.3.4 Valves /A6.4.3.5 Valve Motor Operators /A6.4.3.6 Instrumentation /A6.4.4 HPCIS Turbine / / /A6.4,5 RHR Pumps and Drives " / /A6.4,6 On-Site Electric Power SystemsA6.4.6.1 Diesel Generator Buildings:(a)Structure and Door / /Door HingesCorrugated Barrier AttachmentsA6.4.6.2 Diesel Generators / VA6.4,6.3 Batteries and Battery Racks:Batteries /Battery Racks / /A6.4.7 Electric Power Distribution SystemsA6.4.7.2 Electrical Containment Penetrations / /A6.4.7.3 4 kV Switchgear " / /A6.4.7.4 480 V Load Centers /A6.4.7.5 480 V Motor Control Centers / /A6.4.7.6 DC Distribution Panels and Fuse Boxes /A6.4.7.7 Cable Trays / /A6.4.8 Reactor Protection System /(d)(a)Diesel Generator Building was evaluated for both seismic and tornado loads.(b) Design Criteria as stated in FSAR; or current criteria, if original criteria were not explicitor found to have insufficient margin.(c) More sophisticated analysis methods may in some cases show that the design criteria have in factbeen satisfied; in all cases however, loss of function is not expected.(d) Periodic seal replacement in the snubbers is recommended.Table X 2-3X-9/10

aITABLE X 2-4 BWR COMPONENT REVIEW - ENVIRONMENT QUALIFICATION SUMMARYReview Data BasisAdequate (Failure Not Expected)Assessment of Design AdequacyNot Demonstrated (Failure Possible)Subsectionin TextComponent Descriptiono 0 W.H. -4 04J'W> 0 0Q U CDesign CriteriaSatisfied(a)Analysisand/orTextEngineeringJudgmentDesign CriteriaNot CompletelySatisfied(b)InsufficientInformation toAssess AdequacyInformation ReviewedIndicates thatFailure is Likely NotesA6.4.1A6.4.1.4A6. 4. 1. 6A6.4.2A6.4.2.6A6.4.2.8A6. 4. 3A6.4 .3. 2AG.4.5A6. 4. 7A6.4.7.IA6. 4.7. 2ContainmentSuppression-Chamber-to-Drywell VacuumBreaker ValvesPaint Coatings Within ContainmentReactor Coolant SystemMain Steam Line Isolation ValvesReactor Pressure Relief ValvesCore Spray SystemSnubbersRHR Pumps and DrivesElectric Power Distribution SystemsElectrical Cables Terminations andConnectorsElectrical Containment PenetrationsVWI*1/'IWI/ /W' I II//'/(a) Design Criteria as stated in FSAR; or current criteria, iforiginal criteria were not explicit or found to have insufficient margin.(b) More sophisticated analysis methods may in some cases show that the design criteria have in fact been satisfied; in all cases however, loss of functionis not expected.Table X 2-4x-ll/12

0vf

TABLE X 2-5 SUMMARY OF BASES FOR REDUCED SAFETY MARGIN DETERMINATIONComponentA. PWR Seismic1. Containment Crane2. Low Head Safety Injection Pump3. Reactor Protection SystemB. PWR Environmental1. Reactor Coolant System Snubbers2. Low Head Safety InjectionSystem Snubbers3. Reactor Protection System4. Low Head Safety InjectionSystem Pumps5. Containment Recirculation SprayPumpC. BWR Seismic1. Recirculation Lines2. Reactor Protection SystemD. BWR Environmental1. Reactor Coolant System SnubbersBasisThe charging floor response spectra was used.Use of response spectra- at the crane elevationresults in calculated stresses in theend ties in excess of the yield strength.No seismic qualification test performed.Inadequate seismic qualification testingperformed.Inadequate environmental qualificationtesting performed.Inadequate environmental qualificationtesting performed.Inadequate environmental qualificationtesting performed.Inadequate environmental qualificationtesting performed.Inadequate environmental qualificationtesting performed.Analysis performed assumed a constantspectral acceleration for the verticalcomponent of the earthquakes considered.Use of the vertical response spectraindicates stresses will exceed code allowablebut yielding will not occur.Inadequate seismic qualification testingperformed.There is no evidence that environmentalqualification tests were performed.

TABLE X 2-6 SUMMARY OF ITEMS FOR WHICH DESIGN ADEQUACY COULD NOT BE ASSESSEDA. PWR SeismicComponent1. Reactor Coolant Pump Nozzles2. Low Head Safety Injection SystemInstrumentation3. Containment recirculation spraypumps and drive.4. Diesel generator day tank.S. AC and DC SwitchgearB. BWR Seismic1. Missle barriers insidecontainment.2. Reactor Pressure Relief Valves3. Core Spray System Instrumentation4. 480V load CentersC. BWR Environmental1. Reactor Pressure Relief Valves2. Electrical Cables, Termination andConnectorsBasisImproper Use of Bijlaard Formulae.Revised calculations have not beenperformed.No evidence of seismic qualificationtests.No evidence of pump seismicqualification tests.No evidence of seismic design havingbeen performed.No evidence of seismic qualificationtests.Improper penetration formulae used.No evidence of seismic qualificationtests.No evidence of seismic qualificationtests.No evidence of seismic qualificationtests.No evidence of environmentalqualification tests.Testing performed not sufficientlycomprehensive.Table X 2-5 - Table X 2-6X-13/14

Section 3SunnmaryThe majority of the items examined werefound to be acceptable (seismic design -85%, environmental qualification - 76%,tornado design - 50%). Nine percent ofthe seismic items considered acceptableand 31% of the environmental qualificationitems considered acceptable have amargin to failure somewhat smaller thanthat which could be expected consideringapplicable code and qualification requirements.The only item (PWR dieselgenerator building doors) for whichfailure was considered likely underdesign loadings has a negligible impacton the risk assessments performed inthis study. In addition, there wereitems for which design adequacy couldnot be assessed because sufficientinformation was not available (15% ofthose items considered for seismicevaluation, and 24% of those consideredfor environmental qualification). Theseinadequacies occurred because either:(1) improper assumptions were made indesign calculations or the calculationswere not performed; or (2) seismic orenvironmental qualification tests apparentlywere not performed. The increasedemphasis on qualification testing andquality assurance programs which hasdeveloped since the two specific plantsunder study were designed may haveimproved the situation.X-15

APPENDIX ADESIGN ADEQUACY EVALUATION OF SELECTEDSAFETY RELATED EQUIPMENT AND STRUCTURESIN NUCLEAR POWER GENERATING PLANTSX-17

Appendix ATable of ContentsSectionPage No.Al. INTRODUCTION ...................................................... X-27A1.1 General Background .......................................... X-27A1.2 Objectives of the Design-Adequacy Study ........................ X-27A1.3 Implementation of the Objectives ............................ X-27A2. DESIGN BASIS EVENTS ............................................... X-28A2.1 Earthquakes ................................................. X-28A2.1.1 Seismicity of PWR and BWR Plant Environs .............. X-28A2.1.1.1 PWR Plant ................................. X-28A2.1.1.2 BWR Plant ................................. X-29A2.1.2Response of Structures to Earthquake GroundMotion .............................................. X-29A2.2 Tornado Considerations ...................................... X-30A2.3 Accident Considerations ..................................... X-31A3. SYSTEMS AND COMPONENTS EXAMINED ................................... X-31A3.1 PWR Item Selection Basis .................................... X-32A3.1.1 Reactor Building .................................... X-32A3.1.2 Reactor Coolant System .............................. X-32A3.1.3 Low Head Safety Injection System (LHSIS) .............. X-32A3.1.4 High Head Safety Injection Systems (HHSIS) ............. X-32A3.1.5 Containment Recirculation Spray System ................. X-33A3.1.6 On-Site Electric Power Systems .......................... X-33A3.1.7 Electric Power Distribution Systems .................... X-33A3.1.8 Reactor and Engineered Safeguards ProtectionSystems Sensors and Logic Cabinets ..................... X-33A3.1.9 Intake Canal ........................................ X-33A3.2 BWR Item Selection Basis .................................... X-33A3.2.1 Containment Structures .............................. X-33A3.2.2 Reactor Coolant System .............................. X-33A3.2.3 Core Spray System ................................... X-33A3.2.4High Pressure Coolant Injection System (HPCIS)Turbine ............................................. X-33A3.2.5 Residual Heat Removal (RHR) Pumps and Drives ........ X-33A3.2.6 On-Site Electric Power Systems .......................... X-34A3.2.7 Electric Power Distribution Systems .................... X-34A3.2.8 Reactor Protection System ........................... X-34A4. DESIGN CRITERIA AND PRACTICES ..................................... X-34A5. METHOD OF EVALUATION .............................................. X-35A5.1 Information Acquisition ..................................... X-35A5.1.1 Final Safety Analysis Report Review .................... X-35A5.1.2 Site Visits ......................................... X-36A5.1.3 Visits to Architect/Engineers, Utilities,and Suppliers ....................................... X-36A5.1.4 Review Data Sheets .................................. X-36A5.1.5 Other Documents Reviewed ............................ X-36X-19

Table of Contents (Continued)SectionPage No.A5.2 Evaluation of Information ................................... X-36A5.2.4 Codes and Standards ................................. X-36A5.2.2 Appropriateness of Mathematical Models ................ X-36A5.2.3 Adequacy of Load Definitions ........................ X-37A5.2.4 Adequacy of Boundary Conditions ........................ X-37A5.2.5 Adequacy of Computer Code Numerical Techniques ...... X-37A5.2.6 Effect of Variation on Material Properties ............ X-37A5.2.7 Adequacy of Qualification Test Programs ............... X-37A6. EVALUATION OF COMPONENTS .......................................... X-38A6.1 Seismic Effects ............................................. X-38A6.1.1 PWR Plant ........................................... X-38A6.1.1.1 Seismic Loads ............................. X-38A6.1.1.2 Damping ................................... X-40A6.1.2 BWR Plant ........................................... X-40A6.1.2.1 Seismic Loads ............................. X-40A6.1.2.2 Damping ................................... X-41A6.2 Tornado Effects ............................................. X-41A6.2.1 PWR Plant ........................................... X-41A6.2.2 BWR Plant ........................................... X-42A6.3 Design Adequacy of PWR Plant Components ........................ X-42A6.3.1 Reactor Building .................................... X-42A6.3.1.l Soil-Structure Interaction Model ............ X-42A6.3.1.2 Containment Internal Structure .............. X-43A6.3.1.3 Paint Coatings Within Containment ......... X-44A6.3.1.4 Crane ..................................... X-45A6.3.2 Reactor Coolant System (RCS) ........................ X-46A6.3.2.1 Reactor Coolant System Loops ................ X-46A6.3.2.2 Steam-Generator and Pump Supports ........... X-48A6.3.2.3 Reactor Coolant Pump Nozzles ................ X-49A6.3.2.4 Pipe-Whip Restraints .......................... X-50A6.3.2.5 Snubbers .................................. X-51A6.3.3 Low Head Safety Injection System (LHSIS) .............. X-51A6.3.3.1 Piping .................................... X-51A6.3.3.2 Pumps and Drives .......................... X-53A6.3.3.3 Valves .................................... X-54A6.3.3.4 Valve Motor Operators ........................ X-54A6.3.3.5 Snubbers and Hangers .......................... X-55A6.3.3.6 Instrumentation ........................... X-55A6.3.3.7 Commentary and Conclusion .................... X-56A6.3.4 High Head Safety Injection Systems (HHSIS) .......... X-56A6.3.4.1 Accumulator Tank Nozzle ...................... X-56A6.3.4.2 Accumulator Piping Connection to RCS ...... X-57A6.3.4.3 Charging Pumps and Drives .................... X-57A6.3.4.4 Commentary and.Conclusion .................... X-57X-20

Table of Contents (Continued)SectionPage No.A6.3.5 Containment Recirculation Spray System (CRSS) ....... X-57A6.3.5.1 Pump and Motor Located OutsideA6.3.5.2Containment ...............................Motor Drives ..............................X-58X-58A6.3.6 On-Site Electric Power Systems .......................... X-59A6.3.6.1 Diesel Generator Building .................... X-59A6.3.6.2 Diesel Generators ......................... X-61A6.3.6.3 Day Tanks ................................. X-62A6.3.6.4 Air Bottle Support ............................ X-62A6.3.6.5 Batteries and Battery Support ................ X-62A6.3.7 Electric Power Distribution Systems .................... X-63A6.3.7.1 Electrical Cables and Terminations ........ X-63A6.3.7.2 Electrical Containment Penetrationsand Connectors ............................. X-63A6.3.7.3 Cable Trays ............................... X-64A6.3.7.4 A-C and D-C Switchgear ....................... X-65A6.3.8Reactor and Engineered Safeguards ProtectionSystems Sensors and Logic Cabinets ..................... X-65A6.3.8.1 Environmental Qualification .................. X-65A6.3.8.2 Seismic Qualification ......................... X-65A6.3.8.3 Commentary ................................ X-66A6.3.8.4 Conclusion ................................ X-66A6.3.9 Intake Canal ........................................ X-66A6.4 Design Adequacy of BWR Plant Components ......................... X-67A6.4.1 Containment Structure ............................... X-67A6.4.1.1 Primary Containment Structure ................ X-67A6.4.1.2 Reactor Building (SecondaryContainment) .............................. X-67A6.4.1.3A6.4.1.4Containment Piping Penetrations .............Suppression-Chamber-to-Drywell VacuumX-68Breaker Valves ............................ X-69A6.4.1.5 Missile Barriers Inside Containment ....... X-70A6.4.1.6 Paint Coatings Within Containment ......... X-71A6.4.2 Reactor Coolant System .............................. X-72A6.4.2.1 Reactor Pressure Vessel and Internals ..... X-72A6.4.2.2A6.4.2.3Reactor Pressure Vessel Nozzles .............Reactor Pressure Vessel Skirt ................X-73X-73A6.4.2.4 Recirculation Lines ....................... X-74A6.4.2.5 Main Steam Lines .......................... X-75A6.4.2.6 Main Steam Line Isolation Valves ............ X-75A6.4.2.7 Pipe Whip Restraint for RecirculationLines ..................................... X-78A6.4.1.8 Reactor Pressure Relief Valves ............. X-78A6.4.3 Core Spray System ................................... X-79A6.4.3.1 Piping .................................... X-79A6.4.3.2 Hangers and Snubbers .......................... X-79X-21

Table of Contents (Continued)SectionPage No.A6.4.3.3 Pumps and Drives .......................... X-80A6.4.3.4 Valves .................................... X-81A6.4.3.5 Valve Motor Operators ......................... X-81A6.4.3.6 Instrumentation ............................ - X-82A6.4.4 HPCIS Turbine ....................................... X-82A6.4.4.1 Commentary ................................ X-83A6.4.4.2 Conclusion ................................ X-83A6.4.5 RHR Pumps and Drives ................................ X-83A6.4.5.1 Seismic Qualification ........................ X-83A6.4.5.2 Environmental Qualification .................. X-84A6.4.5.3 Commentary ................................ X-84A6.4.5.4 Conclusion ................................ X-85A6.4.6 On-Site Electric Power Systems ........................ X-85A6.4.6.1 Diesel Generator Building .................... X-86A6.4.6.2 Diesel Generators ......................... X-88A6.4.6.3 Batteries and Battery Racks .................. X-89A6.4.7 Electric Power Distribution Systems .................... X-90A6.4.7.1 Electric Cables, Terminations, andConnectors ................................ X-90A6.4.7.2 Electrical Containment Penetrations ....... X-91A6.4.7.3 4-kV Switchgear ........................... X-91A6.4.7.4 480-V Load Centers ........................ X-92A6.4.7.5 480-V Motor Control Centers .................. X-92A6.4.7.6 D-C Distribution Panels and Fuse Boxes.... X-93A6.4.7.7 Cable Trays ............................... X-93A6.4.8 Reactor Protection System ........................... X-94A6.4.8.1 Commentary ................................ X-94A6.4.8.2 Conclusion ................................ X-95REFERENCES ............................................................... X-95X-22

List of TablesTableX A-IX A-2X A-3X A-4X A-5X A-6X A-7X A-8X A-9X A-10X A-lIX A-12X A-13X A-14X A-15X A-16X A-17X A-18X A- 19X A-20X A-21X A-2 2X A-23X A-24X A-25X A-26X A-27Design Basis Environments - PWR Plant ...........................Design Basis Environments - BWR Plant ...........................PWR Components ..................................................BWR Components ..................................................PWR Component Review - Seismic Design Adequacy Summary ..........PWR Component Review - Environmental Qualification Summary ......BWR Component Review - Seismic Design Adequacy Summary ..........BWR Component Reviw - Environment Qualification Summary .........Comparison of Damping Values Used in PWR Plant With ThoseCurrently Required ..............................................Comparison of Damping Values Used inCurrently Required ..............................................BWR Plant With ThoseComparison of Missile Characteristics...... ......................Comparison of Computed Frequencies ..............................Estimate of Modal Damping .......................................Summary of Moments and Stresses in Crane (Trolley at CenterLine) ...........................................................Critical Stress Summary (psi) ...................................Possible fe/fs Values ...........................................Steam Generator Supports - Maximum Member Loads - PipeRupture Plus Seismic ............................................Reactor Coolant Pump Supports - Maximum Member Loads - PipeRupture Plus Seismic ............................................Summary of Critical Stresses (psi) in LHSIS .....................Summary of Accumulator Tank Nozzle Pipe Loads ...................Load Condition and Stress Limits - Code Case No. 1607 ...........Accumulator Piping Connection Nozzle Loads ......................Reactor Building Natural Frequencies ............................Computed Stress Intensities in BWR Vessel Nozzles ...............Reactor Pressure Vessel Elastic Analysis ........................Natural Periods and Spectral Acceleration .......................Critical Stresses (psi) .........................................Page No.X-99/100X-99/100X-101/102X-101/102X-103/104X-I05/106X-107/108X-109/II0X-109/I10X-11/112X-1II/112X-III/112X-l11/112X-113/I14X-113/I14X-113/114X-i13/114X-115/116X-115/116X-115/116X-I17/1I18X-117/118X-117/118X-119/120X-119/120X-119/120X-119/120X-23

List of Tables (Continued)TablePage No.X A-28X A-2 9X A-30X A-31X A-32X A-33Periods of Vibration and Spectral Accelerations for MainSteam Line ......................................................Critical Stresses (psi) for Main Steam Line .....................Summary of Loads on Pipe Whip Constraints .......................Critical Stresses (psi) .........................................Results of Dynamic Analysis of Battery Rackss ....................Hanger Attachment Types .........................................X-121/122X-121/122X-121/122X-121/122X-121/122X-121/122List of FiguresF i gureX A-1X A-2X A-3Design Adequacy Review Data Sheet ....................................Supplementary Review Data Sheet......................................Example of a Multi-Element Structure Connected to RigidStructures at Both Ends ...........................................X A-4 Ground Response Spectra (OBE) - PWR ...............................X A-5 Ground Response Spectra (DBE) - PWR ...............................X A-6 Response Spectra (DE) - BWR .......................................X A-7 Response Spectra (MCE) - BWR ......................................X A-8 Cross Section Through Reactor Building ............................X A-9 Soil-Structure Interaction Model ..................................X A-10 Containment Internal Structure - Cross Section of LowerPortion ...........................................................X A-lb Reactor Coolant System - Seismic Analysis Model Used bySupplier ..........................................................X A-12X A-13Horizontal OBE Floor Response Spectrum Used for ReactorBuilding ..........................................................A/E Mathematical Model for Dynamic Analysis of PrimaryCoolant Loop ......................................................X A-14 Typical Reactor Coolant Pump ......................................X A-15 Pipe-Whip Restraint - Main Steam Line (not to scale) ..............X A-16 Accumulator Tank Nozzle ...........................................Page No.X-123/124X-123/124X-125/126X-125/126X-125/126X-125/126X-125/126X-127/128X-127/128X-127/128X-127/128X-127/128X-129/130X-129/130X-129/130X-129/130X-24

List of Figures (Continued)X A-17X A-18Horizontal DBE Floor Response Spectrum Used for DieselGenerator Building ................................................Vertical DBE Floor Response Spectrum Used for Diesel GeneratorBuilding ..........................................................Page No.X-131/132X-131/132X A-19 Day Tank ..........................................................X A-20 Air Bottle Support (Not to Scale) .................................X A-21 Elevation View Through BWR Plant ..................................X A-22 Reactor Building Mathematical Model ...............................X A-23 Response Spectra Comparison, Design Earthquake, 5% g, 2%Damping ...........................................................X A-24 Vacuum Breaker Valve ..............................................X A-25 Missile Barriers Inside Drywell ...................................X A-26 Safety Valve ......................................................X A-27 Recirculation Line ................................................X A-28 Main Steam Isolation Valve ........................................X A-29 Pressure Relief Valve .............................................X-131/132X-131/132X-131/132X-133/134X-133/134X-133/134X-133/134X-133/134X-135/136X-135/136X-135/136X A-30A/E Mathematical Model for Dynamic Analysis of Core Spray Piping..X-135/136X A-31 Core Spray Pump and Drive - Lower Portion Showing PumpMounting ..........................................................X A-32 Core Spray Pump and Drive - Upper Portion Showing Motor andCoupling to Pump ..................................................X A-33 Typical Valve Motor Operator Installation .........................X A-34 HPCIS Pump and Interstage Piping ..................................X A-35 HPCIS Turbine Drive ................................................X A-36 Watertight Entryway to RHR Pump Room ..............................X A-37 RHR Pump and Drive ................................................X A-38 RHR Pump and Motor ................................................X-135/136X-135/136X-135/136X-135/136X-137/138X-137/138X-137/138X-137/138X A-39Selected Displacement Records from RHR Pump Vibration TestsShowing Cyclic Displacement Peaks .................................X-137/138X A-40 Diesel Generator Building - Side Facing River .....................X A-41 Diesel Generator Building - Opposite Side .........................X-137/138X-137/138X A-42Lumped Mass Model for Diesel Generator Building SeismicAnalysis ..........................................................X-137/138X A-43 Diesel Generator Unit - View from Generator End ...................X A-44 View of Diesel Generator Unit Showing Turbo-Charger ...............X-139/140X-139/140X-25

List of Figures (Continued)FigurePage No.A-45A-46A-47A-48A-49A-50A-51A-52A-53A-54.A-55A-56A-57A-58Diesel Generator Dynamic Model ....................................Batteries and Battery Racks - 125/250-V System ....................Batteries and Battery Racks - 24-V System .........................Lumped-Mass Model for Seismic Analysis of 58-Cell FTC-21Battery Rack. (Typical of Model Scheme for All BatteryRacks) . ............................................................Electric Power and Control Cables Installed in Cable Trays ........4-kV Switchgear Unit ..............................................4-kV Switchgear Racks .............................................480-V Load Center .................................................480-V Motor Control Centers .......................................D-C Distribution Panels ...........................................RPS Instrumentation Racks (Front. View) ............................RPS Instrumentation Racks (Rear View) .............................Reactor Protection System Logic Cabinet ...........................Control Room Panels for Reactor Safety Systems ....................X-139/140X-139/140X-139/140X-139/140X-139/140X-139/140X-141/142X-141/142X-141/142X-141/142X-141/142X-141/142X-141/142X-141/142X-26

Appendix ADesign Adequacy Evaluation of SelectedSafety Related Equipment and Structuresin Nuclear Power Generating PlantsA.1 INTRODUCTIONA1.1 GENERAL BACKGROUNDThis attachment to Appendix X of the ReactorSafety Study assesses the adequacyof the design of a sample of safetyrelatednuclear equipment and structuresfrom the viewpoint of resistance tocommon-mode failures. An earthquake,for example, has the potential for causinga common failure in the piping ofboth the reactor coolant and the emergencycore cooling systems, if thesesystems are not properly designed. Similarly,a tornado has the potential forcausing a common failure of both offsiteand emergency on-site power due towind damage to off-site transmissionlines and missile penetration of thediesel generator building if the buildingis inadequately designed. These areexamples of system failures that canresult from a single cause. Failure ofone system as the result of the failureof another system due to inadequatedesign would also be classified as acommon-mode failure. A break in a recirculationline, for example, couldresult in a whipping pipe, which could,in turn, cause the failure of engineeredsafety features, such as the containmentbarrier or emergency core cooling systems,if pipewhip restraints are inadequatelydesigned.The Atomic Energy Commission, throughits rules and regulations 1 and guides,The AEC Rules and Regulations are publishedin the Code of Federal Regulations,Title 10, Chapter I (10 CFR) andhave the force of law.2 Regulatory Guides are aimed at providingacceptable methods for implementingspecific parts of the AEC regulationsbut are not mandatory.establishes safety design requirementsfor nuclear power plants to avoid thesefailures. The objective of the designadequacystudy was to determine whetherthe AEC requirements have been properlyimplemented in the samples studied.Two nuclear power plants of recent design,licensed and presently in operation,were selected for study. Oneplant is a boiling-water reactor (BWR);the other is a pressurized-water reactor(PWR).The design-adequacy study was performedby the Mechanical and Nuclear EngineeringDepartment of the Franklin InstituteResearch Laboratories (FIRL) with supportfrom the Atomic Energy Commissionand cooperation from the owners of thetwo nuclear plants that were studied aswell as the architect/engineers (A/E)and nuclear steam supply system (NSSS)suppliers. The A/E, in particular, notonly cooperated with the study but spenta considerable amount of time with theFIRL review team.A1.2 OBJECTIVES OF THE DESIGN-ADEQUACY STUDYThe principal objective of the FIRLstudy was to assess the design adequacyof a sample of nuclear power plant componentsand systems, with detailedconsideration given to the followingdesign basis events that have thepotential for causing common-mode failures:" Earthquake* Tornado" Loss-of-coolant accident (LOCA)A1.3 IMPLEMENTATION OFTHE OBJECTIVESFirst, we examined design basis eventsin general and in particular as theypertain to PWR and BWR plants (sectionA2 contains a discussion of theseevents).X-27

Next, we selected the systems and componentsto be examined. Lists of thesesystems and components and the bases fortheir selection are presented in sectionA3.Finally, we evaluated the methods andthe results of the analyses and testsused in qualifying the equipment andcompared this information with the designcriteria in effect at the time theplants were licensed. To provide arealistic assessment of design adequacy,we also evaluated the equipment inaccordance with current criteria, whichincorporate the latest knowledge. Thesecriteria, as reflected in pertinentcodes, standards, regulations, andstate-of-the-art knowledge, are brieflydiscussed in section A4, as are qualificationmethods anticipated for thefuture.In section A5 we describe the procedurewe used to acquire and evaluate thedesign calculations and the tests thatwere used to qualify the equipment.The results of our evaluation, with detaileddiscussions on specific components,are presented in section A6.'Tables X A-5 and X A-6 provide a summaryof the components reviewed and an assessmentof their design adequacy forthe PWR and BWR plants, respectively.A2. DESIGN BASIS EVENTSAppendix A of Title 10 of the Code ofFederal Regulations, Part 50, contains55 principal design criteria for lightwater nuclear reactor plants. Of thesethe following two describe design basisevents that must be considered in anassessment of the design adequacy of anuclear power plant:a. "Criterion 2 - Design bases forprotection against natural phenome-'na. Structures, systems, and componientsimportant to safety shall bedesigned to withstand the effects ofnatural phenomena such asearthquakes, tornadoes, hurricanes,floods, tsunamis, and seiches withoutloss of capability to performtheir safety functions. The designbases for these structures, systems,and components shall reflect: (1)Appropriate consideration of the-most severe of the natural phenomenathat have been historically reportedfor the site and surrounding area,with sufficient margin for thelimited accuracy, quantity, andperiod of time in which thehistorical data have been accumulated,(2) Appropriate combinations ofthe effects of normal and accidentconditions with the effects of thenatural phenomena and (3) Theimportance of the safety function tobe performed."b. "Criterion 4 - Environmental andmissile design bases. Structures,systems, and-components important tosafety shall be designed to accommodatethe effects of and to becompatible with the environmentalconditions associated with normaloperation, maintenance, testing, andpostulated accidents including lossof-coolantaccidents. These structures,systems, and components shallbe appropriately protected againstdynamic effects, including theeffects of missiles, pipe whipping,and discharging fluids, that mayresult from equipment failures andfrom events and conditions outsidethe nuclear power unit."Criterion 2 describes natural phenomenathat originate outside the plant; Criterion4 describes events that originatelargely inside the plant. Our assessmentof the design adequancy of BWR andPWR plants has concentrated on earthquakes,tornadoes, and accident conditionsassociated with equipment failure.The essential features of thesephenomena will be outlined to promote anunderstanding of their characteristics.A2.1 EARTHQUAKESA2.1.1 SEISMICITY OF PWR AND BWR PLANTENVIRONSA2.1.1.1PWR Plant.The PWR site, located in the Atlantic.Coastal Plain Province, is bounded onýthe east by the Atlantic ocean and onthe west by the fall line and thePiedmont Province. The crystallinebasement rock crops out near the fallzone about 50 miles west of the site.The basement surface slopes gently tothe southeast and is overlain by Cretaceousand Tertiary sediments, which areabout 50 miles west of the site. Apostulated fault trending northwestsoutheastin the basement rock beneaththe James River has been discounted.Eight earthquakes with epicentral intensitiesof V through VII on the modifiedMercalli scale have been reported within100 miles of the site since the late18th century. The closest major earthquakes(estimated to have been about 7*X- 28

on the Richter magnitude scale, withdamage to buildings and bridges) haveoccurred at Charleston, S.C., withepicenters about 350 miles southwest ofthe site.on the basis of the seismic history andthe geological structure of the area, itis estimated that the site could experiencean earthquake with horizontalground acceleration equal to 0.07g and amaximum credible earthquake (MCE) equalto 0.15g in the horizontal direction.The vertical component of the earthquakehas been estimated to be two-thirds ofthese values.A2.1.1.2BWR Plant.The BWR site, located within the Piedmontupland section of the PiedmontProvince of the Appalachian Province, isunderlain by metamorphosed sedimentaryand crystalline rocks of the Paleozoicand Precambrian eras. The fall zone,which represents the physiographicboundary between the Piedmont Provinceand the Atlantic Coastal Plain Province,is located about 20 miles southeast ofthe site at its closest approach.Two major Paleozoic and older faultsystems are prevalent in the region asare faults of the Triassic period.These faults do not involve youngerMesozoic or Cenozoic strata and arecompletely healed. There is a Paleozoicfault associated with a syncline thatpasses about 1 mile south of the BWRsite, but this fault has been inactivefor 140 to 200 million years.The BWR plant is built on bedrock overlainby 10 ft of sediment consistingprimarily of fine sand and silt withoccasional clayey zones.On the basis of records dating back tothe early 18th century, it is estimatedthat the BWR region has not experiencedan earthquake of intensity greater thanVII. Two earthquakes estimated to be ofintensity VII have been recorded: oneoccurred in October 1871, 40 miles fromthe site, and the other occurred inFebruary 1954, 100 miles from the site.on the 'basis of the seismic history andgeological structure of the area, it isestimated that the site could experiencean earthquake with horizontal groundacceleration equal to 0.05g and an MCEwith horizontal ground accelerationequal to 0.12g. The vertical componenthas been estimated to be two-thirds ofthese values.A2.1.2RESPONSE OF STRUCTURES TOEARTHQUAKE GROUND MOTIONNuclear reactors are housed in massiveconcrete structures with fundamentalfrequencies (lowest natural frequency)that generally fall within the peakacceleration response range of earthquakes,i.e., 2.5 to 9 cycles per second(cps).The response of a building to an earthquakewill depend on the dynamiccharacteristics of the building (stiffness,mass, and damping) as well as onits foundation and on the earthquakeitself. If a building is founded onrock (as is the BWR), then the freefieldseismic ground motion will betransmitted directly to the base of thebuilding. If, however, the building isfounded on soil (as is the PWR), thenthere will be dynamic interactionbetween the building and the soil. Inthis case, a single mathematical modelrepresenting both the building and thesoil will be required to determine thedynamic response.The response of structures is usuallyobtained by a method of analysis inwhich the mass of the structure isconcentrated at discrete locations andstiffness properties are computed on theassumption that the structure willbehave like a beam that deflects owingto both shear and bending loads. Sincenuclear structures have relatively smallaspect ratios (height/width), sheardeflection predominates.The dynamic response of the structuremay be obtained by modal analysismethods 1 that use either a real or anartificial time history earthquakerecord or a response spectrum curve fordetermining the response.A singl e point on a response spectrumcurve for a given percent of criticaldamping n is obtained by determining thetime history response of an oscillatorwith damping n and a given natural frequencyto base motion defined by theaccelerogram. The maximum response ofthis oscillator then determines theordinate of the response spectrum, andthe natural period of vibration of theoscillator is the abscissa. The processis then repeated for other assumedperiods. The resulting response spectrumcurve is quite jagged, and the1 See Ref. 1 for a discussion of modalanalysis methods and techniques.X- 29

process may be repeated for severalearthquake records to obtain an average.The Housner curve is a smoothed averageresponse curve that is in common use.The dynamic stresses in the structuremay be obtained from accelerationresponse spectra by first determiningthe natural frequencies and correspondingmode shapes of the structure as wellas modal participation factors. Themodal acceleration is then equal to themode shape times the modal participationfactor times the acceleration valuetaken from the acceleration responsespectrum curve (for the assumed modaldamping), which corresponds to the modeshape frequency. Modal stresses maythen be obtained by applying theinertial forces (corresponding to themodal accelerations) to the structure.The total response should be computed bycombining these stresses in accordancewith the square root of the sum of thesquares of the modal stresses except formodes with closely spaced frequencies,as will be explained subsequently. Thedisplacements can be similarly determined.The response spectrum approach is aconvenient method of obtaining displacements,forces, and stresses in thestructure; however, since we are dealingonly with maximum values in each mode,we do not know the phase relationshipbetween the modal responses. It iscustomary, therefore, to determine thetotal response by taking the square rootof the sum of the squares (SRSS) of themodal stresses. * The response of modeswith closely spaced frequencies may,however, be nearly in phase. Recentpractice has been to combine the directsum of the absolute value of these modalresponses with the remaining modes inaccordance with the SRSS.The response spectrum method gives totaldisplacement at a given point in thestructure, which is neither time- norfrequency-dependent; so it does notallow an analysis of the dynamic responseof equipment that is mounted atthe various floor levels in thestructure. For this purpose, a floorresponse spectrum curve must be generatedby starting with a time history inputto the base of the structure and thendetermining the time history response atthe floor of interest. The floorresponse spectrum may then be generatedfrom the floor time history response.An artificial time history record at thebase of the structure must first begenerated since it must be compatiblewith the averaged design ground responsespectrum curve. This is usually done bystarting with a real time historyrecord, such as the Taft earthquakerecord, which has been normalized to thedefined maximum ground acceleration forthe operating basis earthquake (OBE) ordesign basis earthquake (DBE), and thenmodifying it so that its correspondingresponse spectrum curve approximates andenvelops the design ground responsespectrum.At the time the PWR and BWR plants werelicensed, it was generally assumed thatonly one horizontal component of theearthquake (e.g., X or Z) would actsimultaneously with the vertical component.Assume for the moment that thestress at a given point in the structuredue to the X, Y, or Z earthquakecomponents is "a." Some analysts wouldassume that the X and Y earthquakeswould be in phase; so the total responsewould be 2a. Other analysts wouldassume that the components would be outof phase; so the total response would begiven by V- 0 a l .4c. Recent evidenceindicates that all three components ofthe earthquake act simultaneously butout of phase; thus the total response inthe assumed example would be VTaz1.7a. Thus, in this example, theanalysts who had assumed the X and Zcomponents to act simultaneously but inphase would give results that would beacceptable for the three simultaneousearthquake components which are out ofphase. Likewise in more realisticexamples, this would be generally, butnot universally true. However, theanalysts who had previously assumed twocomponents out of phase would have toreexamine their results.A2.2 TORNADO CONSIDERATIONSA tornado can develop the greatest concentrationof natural atmospheric poweron earth. It is the most violent of allstorms. Although a tornado covers avery restricted area and is of shortduration, it can cause devastating destructionof homes and other commonstructures and place human life in greatjeopardy from blast and flying debris.Nuclear reactor structures housingsafety equipment should be resistant tothe following 'tornado effects:a. External wind forces.b. Differential pressure between theinside and outside of fully enclosedbuildings.C. Impact from flying missiles.WX-30

A2.3 ACCIDENT CONSIDERATIONSSpecific definitions of tornado effectsare given in sections A6.2.1 and A6.2.2for the PWR and BWR plants, respectively.Although extensive measures are takenduring the design and operation of nuclearpower plants to prevent accidents'from occurring, the AEC requires thatthe designer postulate severe accidents,using unrealistically severe assumptions.The effects of the various typesof hypothetical accidents are investigadoted, considering a variety of plantconditions, to determine situations thatcould result in the off-site release ofradioactive material. The followingtypes of accidents are generally considered:a. Rupture of any single pipe up to andincluding complete severance of thelargest pipe in the reactor coolantsystem.b. Rapid power increase beyond designlimits caused by reactor controlsystemfailure.C. Rupture of fuel rod cladding in afuel assembly or spent-fuel transportcask that is dropped during refuelingoperations.d. Rupture of a gaseous-radioactivewastestorage tank.Each of these types of accidents has thepotential for releasing radioactive materialto the environment. An analysisof each type of accident is performed todetermine whether there is assurancethat adequate safety features have beenengineered into the plant -- in the formof passive barriers or active systems --to limit the release of radioactivematerials and to show that the maximumV radiation dose to an individual at theplant's site boundary and beyond wouldnot exceed the allowable dose even whenhighly pessimistic assumptions are used.Accidents with the largest potential foroff-site radiation exposures for a givenset of assumptions are designated designbasis accidents and are analyzed indetail.Structures, engineered safeguards systems,and the components thereof must bedesigned to perform adequately under allservice conditions including normal operation,maintenance, and testing andunder the conditions of postulateddesign basis accidents. They must alsobe appropriately protected againstflooding and dynamic effects (includingthe effects of accident-produced missiles,pipe whipping, and dischargingfluids) which can result from equipmentor piping failures.Additional conservatism is incorporatedinto the accident analyses by assumingthat an additional, unrelated, unspecifiedfault in some active component orpiece of equipment in the engineeredsafeguards systems also occurs. Thisfault is assumed to result in the malfunctionof a device that is intended tomitigate the consequences of the accident.The assumed result of such an unspecifiedfault is restricted to suchrelatively common events as an electricalfailure, instrument error, motorstall, breaker freeze-in, or valvemalfunction. Highly improbable failures,such as additional pipe breaks,are not assumed to occur coincident withthe assumed accident. The additionalfailure(s) to be considered are inaddition to all possible failures(common-mode or otherwise) caused by theaccident itself. For the most severetype of design basis accident, a LOCA,this means that the worst possiblesingle failure must be assumed to occurin one of the systems designed to preventa LOCA from leading to extremeoverheating of the fuel rods (andthereby melting or otherwise releasinglarge quantities of radioactive materials).The accident-produced environmental conditionspostulated for the BWR and PWRplants are summarized in Tables X A-1and X A-2, respectively. The numericalvalues under headings I and II are thedesign bases for the equipment. Notethat the LOCA conditions under headingII are established under the assumptionthat the worst possible single failurein the engineered safeguard systems hasoccurred coincident with the accident.The tabulated information has beendeveloped from a number of sources, includingthe IEEE standard on equipmentqualification (Ref. 2) and reports ofqualification tests conducted on individualcomponents.A3. SYSTEMS AND COMPONENTSEXAMINEDThe components and systems sampled inthis review were selected on the basisof a combination of factors:a. Safety function.x-31

. Portions of structures which areknown to be vulnerable to highstresses.c. Component types that have experiencedmalfunction during normal operationand/or testing in existingnuclear power plants.d. Potential for common-mode failure.Some major components, such as the reactorpressure vessel, were deliberatelyexcluded from our review 1 because theyrequire a formal stress report and aresubjected to a careful review by the AECand other organizations, which providesan added measure of assurance as totheir adequacy.Some items, such as the reactor building,were reviewed primarily from theviewpoint of adequate dynamic mathematicalmodeling to ensure that computedfloor response spectra curves used foranalyzing equipment were correct.Our sampling process was biased towardpoints of potential weakness.Thirty-one PWR equipment groups (Table XA-3) and thirty-three BWR equipmentgroups (Table X A-4) were examined bythe review team. In the tables eachgroup has been assigned a number thatcorresponds to the subsection in sectionA6 where the group is described andevaluated. For each plant type, a completesystem was selected for reviewplus certain components or structures inother systems. The components wereselected for examination because theymay all be subject to seismic loads andthus have the potential for common-modefailure. In addition, these componentseither perform a safety-related functionor their failure could prevent the safeshutdown of the reactor or cause abreach in containment, as described insections A3.1 and A3.2.A3.1 PWR ITEM SELECTION BASISThe groups shown in Table X A-3 formnine systems whose safety-related functionsare described briefly as follows:A3.1.1REACTOR BUILDINGThe containment structure is the ultimatebarrier between any postulated1 The BWR reactor was considered onlyfrom the viewpoint of imposing seismicdisplacements on connecting piping.large release of radioactivity and theplant environs. Soil-structure interactionmust be considered because the PWRis founded on 1300 ft of soil, whichaffects the response of structures andequipment to seismic events. The internalstructure (also referred to as thecrane wall) supports the nuclear steamsupply system, the crane, and portionsof safety-related equipment, such as theemergency core injection piping. Thecontainment paint coating has thepotential for peeling under the severeenvironment caused by a LOCA and therebyimpairing the effectiveness of some ofthe emergency core cooling and containmentcooling systems. Catastrophicfailure of the crane or its supportcould cause a break in the containmentspray headers, which are directly belowthe crane.A3.1.2REACTOR COOLANT SYSTEMItems in this group are necessary tomaintain the integrity of the reactorcoolant pressure boundary, except thepipe-whip restraints whose failure couldcause a breach in the containment. Componentsor portions thereof were selectedon the basis of the criteria discussedpreviously.A3.1.3LOW HEAD SAFETY INJECTION SYSTEM(LHSIS)This system injects borated water intothe cold legs of the reactor coolingsystem loops as a supplement to theaccumulator discharge of borated waterin the event of a large LOCA. It alsoinjects water (which collects in thecontainment sump) during the recirculationphase. The system is necessary toprovide for the safe shutdown of thereactor in the event of a LOCA. Thesnubbers provide seismic restraints onpipe motion and the hangers support thepiping weight. Both snubbers andhangers should be designed to limitexcessive stress in the piping.Instrumentation controls the LHSISoperation. For this group, all items inthe system were considered in ourreview.,A3.1.4HIGH HEAD SAFETY INJECTIONSYSTEMS (HHSIS)These systems inject borated water intothe cold legs of the reactor in theevent of a LOCA. The accumulator dischargesborated water when there is alarge break in the reactor coolant system(RCS). The nozzle of the accumulatorand the connection to the RCS pipingx-32

were examined because these are criticalregions.A3. 1.5 CONTAINMENT RE CIRCULATION SPRAYSYSTEMThis system is necessary to depressurizethe containment in the event of a LOCA.The key components -- the pumps andtheir drives -- were considered in ourreview.A3.1.6ON-SITE ELECTRIC POWER SYSTEMSEmergency on-site power is provided bythe diesel generator. The system isnecessary to power engineered safeguardssystems in the event of loss of off-sitepower. The day tank provides the initialsource of diesel fuel for the system.The air bottle stores compressed airused in starting the diesel engine. Thebatteries provide emergency power forcontrol systems necessary for the safeshutdown of the reactor.A3.1.7ELECTRIC POWER DISTRIBUTIONSYSTEMSThese systems distribute power to thepump motors, motor-operate~d valves, andcontrol-rod drives which are necessaryfor safe shutdown.A3 .1. 8 REACTOR AND ENGINEEREDSAFEGUARDS PROTECTION SYSTEMSSENSORS AND LOGIC CABINETSThe protection systems sense reactorwater level, reactor pressure, and containmentpressure. Through their logicsystems, they interpret conditions thatcall for immediate shutdown, initiatethis action, and activate containmentisolation and reactor core safety injectionsystems.e A3.1.9 INTAKE CANALThe intake canal carries water pumpedfrom the river, which is the only sourceof cooling for the entire plant.A3.2 BWR ITEM SELECTION BASISThe components shown in Table X A-4 fallinto eight groups whose safety-relatedfunctions are described briefly as follows:A3.2.1 CONTAINMENT STRUCTURESThe drywell and suppression chamber providethe containment pressure boundaryand passive suppression of pressureincrease in the event of a LOCA. Thisstructure is housed in the reactorbuilding, which functions as a secondarycontainment barrier against the releaseof radioactivity off-site should therebe an accident. Containment piping penetrationsare intended to provide forthe safe transfer of material throughthe containment boundary so that a pipebreak will not cause a breach in containmentintegrity. The missile barriersprovide shielding inside the drywellagainst a whipping pipe. As in the PWRcase, excessive peeling of paint insidethe containment could impair the properfunctioning of the emergency core coolingsystems.A3.2.2 REACTOR COOLANT SYSTEMItems in this group maintain the integrityof the reactor coolant pressureboundary as well as that of theradioactive steam produced by the BWR.Pipe lines and nozzle connections to thereactor must be capable of resistingseismic loads. The main steam isolationvalves must prevent the large release ofradioactive steam from within thecontainment. The pipe-whip restraintsare~ intended to prevent a break in therecirculation line from causing asecondary failure in containment or engineeredfeatures. The reactor pressurerelief valves are intended to preventexcessive pressure within the nuclearsteam supply system.A3.2.3CORE SPRAY SYSTEMThis system provides emergency corecooling in the event of a large LOCAafter the pressure in the reactor isreduced. All items in this group mustbe capable of resisting seismic events(which have the potential for causing aLOCA) and must be capable of continuedoperation in the environment following aLOCA. For this system, all items wereanalyzed.A3.2.4HIGH PRESSURE COOLANT INJECTIONSYSTEM (HPCIS) TURBINEThe HPCIS provides initial injection ofemergency core cooling in the event of aLOCA.A3.2.5RESIDUAL HEAT REMOVAL (RHR)PUMPS AND DRIVESThe system limits the temperature insidethe containment after a LOCA by circu-X- 33

lating water f rom the suppression pooiand cooling it via RHR heat exchangers.A3.2.6ON-SITE ELECTRIC POWER SYSTEMSEmergency on-site power is provided bythe diesel generator. The system isnecessary to power engineered safeguardssystems in the event of loss of off-sitepower. The day tank provides the initialsource of diesel fuel for the system.The air bottle stores compressed airused in starting the diesel engine.Batteries provide emergency power forcontrol systems necessary for the safeshutdown of the reactor.A3.2.7 ELECTRIC POWER DISTRIBUTIONSYSTEMSThese systems distribute power to thepum~p motors, motor-operated valves, andcontrol-rod drives which are necessaryfor safe shutdown.A3.2.8REACTOR PROTECTION SYSTEMThe reactor protection system sensesreactor water level, reactor pressure,and containment pressure. Through itslogic system, it interprets conditionsthat call for immediate shutdown, initiatesthis action, and activates containmentisolation and reactor core safetyinjection systems.A4. DESIGN CRITERIA AND PRACTICEScodes and standards for the design, construction,and operation of nuclearpower plants, like codes and standardsfor conventional power plants and otherprocessing plants, are under continualreview and revision. The codes andstandards in existence at the time thetwo plants used in this study wereapproved have undergone considerablerevision in the 6 to 8 years sinceinitial review. For this reason, theplants studied fail to meet all thecodes and standards that are invoked forplants currently being licensed. On theother hand, the PWR and BWR plants weredesigned and constructed to somewhatmore rigid requirements than were thepreviously designed fossil and nuclearplants that provide the data base forthe component failure rates used in theRSS. It is clear, therefore, that, inprinciple, the reliability of the PWRand BWR plants should be at least asgood as the reliability of those plantsfrom which data have been derived. Bythe same token, future plants, includingthose now under construction, must meetmore stringent design, construction, andoperational codes and standards than dothe plants used in this study. Thus,the reliability values used in the RSSare probably somewhat conservative, and,although the plants used in the studymay not meet all the requirements ofpresent codes and standards, their designcriteria requirements are at leastas good as those of the facilities thathave provided basic failure rate data.Our study centers primarily on safetycomponents (i.e., those necessary forsafe operation and safe shutdown) andthe ability of the engineered safeguardsystems to function as required.Each applicant for a nuclear power plantconstruction permit or operating licenseis required to provide assurance thatall safety-related equipment will performadequately under all its serviceconditions, including normal operation,maintenance, and testing (both of thecomponent itself and of other parts ofthe plant which may subject the componentto unusual service conditions), anddesign basis events, such as earthquakesand accidents.Section III, "Design Control," of AppendixB, "Quality Assurance Criteria forNuclear Power Plants and Fuel ReprocessingPlants," to 10 CFR Part 50,"Licensing of Production and utilizationFacilities," requires that a test programused to verify the adequacy of aspecific design feature include suitablequalification testing of a prototypeunit under the most adverse design conditions.The overall quality assuranceprogram for nuclear power plant equipmentincludes, but is not limited to,design, qualification, production qualitycontrol, shipping and storage,installation, maintenance, and periodictesting. The FIPL design adequacy studyincluded a review and evaluation of thequalification portion of the overallquality assurance program.The primary role of qualification is toensure that, for each item of safetyrelatedequipment, the design and manufacturingprocesses are such that thereis a high degree of confidence in thereliability of installed equipment. Theother steps in the quality assuranceprogram require strict control to ensurethat the equipment installed in theplant is the same type as that which wasqualified and that the production itemsare suitably applied, installed, maintained,and periodically tested.Essentially, a component is assumed tobe functional only if it meets estab-X-34

lished codes and regulations asevidenced by the SAR and by specificsupporting documentation of designadequacy which is kept on file. It isrecognized that the evidence presentedto establish design adequacy can takemany forms.Equipment qualification requirements andmethods are embodied in criteria, standardsand guides (Ref s. 1 through 8)prepared by committees of the Instituteof Electrical and Electronics Engineers(IEEE), American Society of Chemical*Engineers (AIChE) , and other professionalsocieties. These documents are thenforwarded to the American NationalStandards Institute (ANSI), which eitherendorses (adopts) them or returns themfor revision. The AEC then eitheraccepts this document as its requirement,or issues a Regulatory Guidecontaining detailed requirements, sometimessupplementing those in theindustry guides. It should be notedthat AEC personnel hold membership onmost of the professional society committeeswhich prepare industry standardsdocuments. In this way the AEC viewpointusually becomes incorporated intothe documents. Meetings of the committeesare extremely valuable because theyprovide a relatively informal setting inwhich industry and government representativescan explore each other's point ofview and basis for desiring that certainfeatures be incorporated into the documents.Misunderstandings can be eliminated,or at least minimized, and atechnically strong document developed.Qualification requirements are constantlychanging, becoming more stringent asindustry, the ABC, and the professionalsocieties concerned with the preparationof test standards gain knowledge aboutthe procedures needed to ensure thereliability of safety-related equipment.When the BWR and PWR plants consideredin this study were designed and theequipment was purchased, qualificationrequirements were very general. Thissituation placed a great burden on thedesign engineers who had to interpretthe generalities in terms of specificprocedures for each type of equipment.For example, seismic qualification ofcomponents was not required initially,and even after seismic qualificationbecame a requirement, it took a longtime for the AEC and the industry totranslate the general requirements intospecific requirements. This process isstill under way. Part of the problemhas been that the equipment necessary toconduct many of the tests is notavailable and, because the tests arehighly specialized and the testrequirements have been changing constantly,test facilities have onlyslowly come into existence. The greatestpart of the problem, however, isthat in other fields standards have beendeveloped after practices have becomeaccepted in the industry. In thenuclear industry, the preparation ofstandards is required before diversepractices are established, and thusaspects of the standards must deal withmatters which have not been studied andfor which the state of technical knowledgemay be rather limited.The need for a more intensive effort indeveloping standards for the nuclearindustry has been recognized, and thelevel of activity has increased dramaticallyin the past two years.Type testing of actual equipment usingsimulated service conditions is usuallythe preferred method of qualifyingequipment. However, a type test alonesatisfies qualification only if theequipment to be tested is properly aged,subjected to all important environmentalinfluences, and operated under posteventconditions to provide assurancethat all such equipment will be able toperform its intended function for atleast the required operating time. Partialtype tests may be augmented bytests of components when size, applications,time, or other test limitationspreclude the use of a full type test onthe complete equipment package.A5. METHOD OF EVALUATIONThe following sections describe the proceduresby which the review teamacquired and evaluated the informationused to assess the design adequacy ofthe systems and components noted in sectionA3.A5.1 INFORMATION ACQUISITIONA5.1.1 FINAL SAFETY ANALYSISREPORT REVIEWBackground material and preliminary informationwere acquired by reviewing thefinal safety analysis reports (FSAR)which describe the physical aspects ofthe PWR and BWR sites, including theirgeology and seismology, design bases,functional aspects, operational requirements,'design evaluation, and test andinspection requirements for the plantsand their various systems.The FSAR is a large documenttypically of 15 sections,specifications, supplementsconsistingtechnicaldescribingX-35

questions by the AEC and answers by theowner, appendices, and supplementaryreports.A5.1.2SITE VISITSThe review team visited the PWR and BWRsites and toured the plants, with theexception of high radiation areas. ThePWR was in operation at the time of thevisit, and the BWR was preparing forinitial start-up. The PWR safety-relatedpiping, motors, pumps, cable trays,tanks, diesel generator, controls, batteryrooms, logic cabinets, and intakecanal were inspected. The tours wereconducted by plant operating personnel.Special attention was given to theinstallation, support, and attachment ofequipment.A5.1.3VISITS TO ARCHITECT/ENGINEERS,UTILITIES, AND SUPPLIERSThe FSAR primarily describe the designbases and functional requirements forthe plant systems and components; theydo not present the design calculationsand test procedures in sufficient detailfor an assessment of design adequacy.The detailed calculations and testreports are retained by the A/E and bythe various equipment suppliers. Inmany cases this information is consideredproprietary; so the A/E and theNSSS suppliers would not permit removalof this information from their premises.However, because these firms wanted tocooperate with the RSS, they permittedthe review team to examine the materialon the premises. The cognizant engineer(if available) responded to our questions,and we were able to make notesand copies of pertinent data.We also visited the utility that ownsthe BWR.A5.1.4REVIEW DATA SHEETSThe review team used review data sheetsand supplementary review data sheets(Figs. X A-1 and X A-2) during themeetings with the A/E and suppliers. Weused these sheets as guides in carryingout the review and also as a means ofrecording pertinent information. Theinformation on these sheets was thenevaluated as described in section A5.2.The design adequacy assessment is reportedin section A6.A5.1.5OTHER DOCUMENTS REVIEWEDIn addition to the FSAR, design calculations,and test reports, we reviewed thesafety evaluation reports prepared bythe AEC for each plant, design specifications,and pertinent portions ofgoverning codes and standards, such asthe B31.1 USA Standard Code for PressurePiping. We also reviewed pertinent AECregulatory guides, which supplement thecodes and standards with the latestmethods and techniques acceptable to theAEC. These are referenced in appropriatesections of this report.A5.2 EVALUATION OF INFORMATIONWe evaluated the information acquiredwith the objective of drawing conclu'--sions about design adequacy, as detailedin section A1.2. Evaluations were basedon the considerations described in thissection.Time did not permit extensive spot-checkindependent analysis. We did, however,perform an independent analysis of theBWR core spray piping (section A6.4.3.1)to compare the results of evaluating thepiping in accordance with the currentlyrequired NB-3600 of the ASME Code, SectionIII, with the results obtainedusing USAS B31.1, which governed thedesign at the time of construction.In addition, if our review uncoveredcalculations or assumptions that seemedunreasonable, we made very conservativesimplifying assumptions to estimate anupper bound on stresses. If the resultswere within the allowables, the designwas considered adequate. If, however,the resulting stress values exceeded theallowables, we informed the A/E or NSSSsupplier that the original calculationswould have to be revised.A5.2.1CODES AND STANDARDSPertinent portions of the codes andstandards cited in the FSAR as governingthe design were reviewed to verify thatcomputed stress levels were within allowablevalues. An assessment was alsomade to determine whether the equipmentwould meet current requirements.A5.2.2APPROPRIATENESS OF MATHEMATICALMODELSPerhaps the most important step in theanalysis process is the development ofthe mathematical model as an idealizationof the real system. When theengineer creates a mathematical representationof a real system, he makes ajudgment about what kind of phenomenaare significant and therefore should beincluded in the model and what phenomenaare insignificant and therefore shouldbe excluded. The review team checkedX-36

the appropriateness of thein the calculations.models usedA5.2.5 ADEQUACY OF COMPUTER CODENUMERICAL TECHNIQUESThe most common modeling error we encounteredduring the course of thisreview was the implicit assumption of ashear connection between multi-elementstructures having a common connection toa fairly rigid structure at either end,such as the yoke that separates thevalve operator from the valve (Fig.X A-3). The analyst assumes that themulti-element structure acts as asingle-element cantilevered beam oflength L with a cross-sectional momentof inertia described by:This incorrect modeling may give rise tocomputed stiffnesses and associatedfrequencies which are far in excess oftheir correct value and which could leadto unconservatively low seismic loads.Dynamic interaction between the structureand heavy equipment, such as theNSS System, should also be considered inmodeling. Guidelines for using decoupledmodels rather than a single, morecomplex model are given in Ref. 1.A5.2.3 ADEQUACY OF LOAD DEFINITIONSour review included an evaluation of theseismic loads used in the analysisand/or testing of equipment and a comparisonof these loads with the expectedseismic loads. Since the seismic loadsare obtained from the frequency-dependentresponse spectra curves, an error inthe computation of natural frequenciesmay lead to an error in seismic loads.Seismic loads are sometimes developed byone group or company and used by anothergroup or company for analysis and evaluationof equipment. This presentspotential interface problems that couldlead to significant errors. During ourevaluation, we looked for these situations.A5.2.4 ADEQUACY OF BOUNDARY CONDITIONSBoundary conditions implicit in theanalysis of any system that is assumedseparated from its surroundings forpurposes of study were evaluated to seeif the assumed boundary conditions werereasonable. For example, hydraulicsnubbers provide a "spring" (albeitgenerally stiff) support but are oftenassumed to provide a rigid constraint onpipe displacement. We considered theeffect of this departure from the realcondition.We did not make an extensive check oncomputer codes since verification ofcomputer codes with experimental results,or by obtaining the same resultfor a given problem with two independentcodes, is generally required before AECacceptance.A5.2.6EFFECT OF VARIATION ONMATERIAL PROPERTIESonly the variation in shear modulus forthe soil used in' the soil-structureinteraction model (see section A6.3.1.1)was considered in our evaluation. Variationin Young's modulus for concrete isgenerally accounted for in the peakbroadening of the floor response spectracurves.The modulus of elasticity for varioussteels for nuclear applications is givenin Table 1-6.0 of the ASME Code, SectionIII. There is negligible variation ofYoung's modulus, B, for steel of a givencomposition; there is a slight decreasewith increasing temperature, however.The value of B for austenitic stainlesssteel, for example, decreases from 28.3x 106 psi at 70OF to 27.1 x 106 psi at3001F. This represents a decrease of 4%and would give rise to a decrease innatural frequency of 2%, a change whichis not considered significant.A5.2.7 ADEQUACY OF QUALIFICATIONTEST PROGRAMSThe procedure for evaluating the adequacyof equipment that had been qualified,wholly or partially, by test includedthe following salient features. First,we made an effort to establish thesafety-related function of the equipmentand the conditions under which it isexpected to function. Then we evaluatedthe available reports containing informationon qualification tests withreference to applicable ABC regulatoryguides and industry standards (principallyIEEE and ANSI equipment qualificationguides). This included an evaluationof the applicability of, and theadequacy of proof for, satisfactoryperformance under the following conditions:a. Degradation occurring during theinstalled life under normal conditions,including thermal aging, radiationaging, and normal wearaging.b. Seismic disturbance.x- 37

c. Loss-of-coolant accident (LOCA), includingexposure to nuclear radiation,steam, and sprays of demineralizedwater and chemical solutions.Although we were able to note whetheraging had been simulated, we were generallyunable to evaluate the adequacy ofthe simulation. Not only is the theoryof accelerated thermal aging to simulateexposures of up to 40 years not wellestablished but data on the bases forthe selection of the temperature and theduration of thermal aging exposures forparticular equipments were not availableto us. Since the radiation exposuresrequired to simulate normal and accidentconditions have been established, theadequacy of radiation resistance wasrelatively easy to evaluate. Theevaluation of seismic qualificationinvolved a comparison of the accelerationlevels, and frequency responsespectrum anticipated at the equipmentlocation with those used in the testsand, in some cases, independent qualitativeanalysis of the vulnerability ofequipment to seismic disturbances.Evaluation of equipment performanceunder LOCA conditions included acomparison of the predicted in-plantenvironmental exposure profiles with theexposure profiles used in the testprogram.A6. EVALUATION OF COMPONENTSThis section summarizes the results ofthe review of PWR and BWR plant components.Section A6.1 describes the overall seismiceffects, with particular emphasis onthe definition of seismic loads for PWRand BWR plants.Section A6.2 describes the tornado effectson the diesel generator buildingswhich house the source of emergencypower.Sections A6.3 and A6.4 present the resultsof the component reviews for thePWR and BWR plants, respectively, togetherwith commentary and conclusions.Tables X A-5 and X A-6 (PWR) and TablesX A-7 and X A-8 (BWR) provide a summaryof the components reviewed andan assessment of their design adequacy.Tables X A-5 and X A-7 give the resultsfor components that were reviewed fromthe viewpoint of seismic design adequacy.An assessment of the tornado resistanceof the diesel generator buildingis also included in Tables XA-5 andX A-7. Tables X A-6 and X A-8 give theresults for components that were reviewedfrom the viewpoint of resistance tonormal and LOCA environments.The tables indicate whether the designof a component is such that failure isnot expected and the component is thereforejudged "adequate" or whether failureis possible so that adequacy is "notdemonstratated." The assessment ofdesign adequacy is indicated by a checkmark in one of five columns as follows:consid-Failure Not Expected (Design isered adequate)i. Design criteria, as explicitly statedin FSAR (or in accordance withcurrent criteria and practice if notexplicitly stated in the FSAR), havebeen satisfied by analysis and/ortest.2. Design criteria have been satisfiedon the basis of engineering judgment.3. Although the design criteria havenot been completely satisfied, lossof function is not expected. Insome cases more sophisticated analysismethods may show that the designcriteria have been satisfied.Failure Is Possible (Design adequacy hasnot been demonstrated)4. There is insufficient information toassess adequacy.5. The information reviewed indicatesthat failure is likely.The tables generally summarize our evaluationof design adequacy in accordancewith the criteria stated in the FSAR.We made exceptions only when currentknowledge indicated that the originalcriteria were in error (for example, theoriginal design temperature for the BWRof 286OF was superseded by a higherdesign temperature of 340 0 F) or whenexplicit criteria were not stated inFSAR.Current criteria and practice are discussedand compared with the originalcriteria in the "commentary" portions ofthe text.A6A SEISMIC EFFECTSA6.1.1A6.1.1.1PWR PLANTSeismic Loads.The site of the PWR is founded on 1300ft of soil above bedrock. On the basisX-38

of tectonic considerations and the seismichistory of the area within a 350-mile radius of the site, the maximum accelerationat bedrock is expected to beno higher than 0.07g. With soil dampingequal to 10% of the critical value,soil-surface acceleration is calculatedto be 0.14g. This figure does not includesoil-structure interaction.consid-The following seismic loads areered for this plant:a. operating basis earthquake (OBE)0.07g horizontal2/3 (0.07)g verticalb. Design basis earthquake (DBE)0.15g horizontal2/3 (0.15)g verticalThe adequacy of Class I structures isevaluated on the basis of the groundresponse spectra shown on Fig.. X A-4(OBE) and Fig. X A-5 (DBE). These curvesare taken directly from the FSAR for theBWR plant and (although somewhat atypicalin appearance) were derived fromHousner' s average response spectra normalizedto the maximum expected groundacceleration for this plant.The adequacy of Class I components isevaluated on the basis of floor spectracalculated from Taft's earthquake timehistory record normalized to the groundacceleration levels of this site. Initialanalysis of components was based onground acceleration spectra. Later thiswas modified to include amplificationdue to structural response.A6.1.1.1.1 Commentary. The maximumground acceleration experienced within a100-mile radius of the site since the18th century is estimated at 0.05g onthe basis of the estimated seismicity ofthe region. The tectonics of the regionappear to have been carefully considered.The ground accelerations of 0.07gfor the OBE and 0.15g for the DBE thereforeappear to be reasonable values forthe purpose of design. The designground response spectra currently requiredby the AEC (Ref. 1) generallygive higher responses than those derivedfrom the average Housner spectra. Thedamping values currently required by theAEC (Ref. 2) are generally higher thanthose acceptable at the time the plantwas designed. For example, 5% dampingfor both the DBE and OBE was used forthe reinforced-concrete containmentstructure. As shown in Table X A-9, thisvalue is lower than the currently acceptablevalue of 7% for the DBE but ishigher than the current 4% for the OBE.The 5% response spectra curve used inthe design of the BWR (curve 5) for theOBE may be compared with the ground responsespectra curve currently required(curve 4) in Fig. X A-4. -Note that thecurrent design curve gives a maximum accelerationof 0.22g at 2.5 cps comparedto the PWR design value of 0.llg between2 and 4 cps.Figure X A-5 shows a comparison betweenthe response spectra used in the designof the PWR for the DBE (curve 5) andthat currently required (curve 7) . Thecurrent design curve gives a maximum accelerationof 0.4g at 2.5 cps ascompared to the PWR value of 0.22gbetween 2 and 4 cps.It is not surprising that the currentdesign response spectra given inRegulatory Guide 1.60 are higher thanthose used in the design of the PWR.The PWR curves were based on the Housnerspectra, which represent average values,whereas the current design responsespectra represent 1 positive standarddeviation from the average values.It should be emphasized here that RegulatoryGuide 1.60 indicates that thedesign response spectra curves may notbe applicable for sites that "have physicalcharacteristics that could significantlyaffect the spectral combinationof input motion." For such sites thedesign response spectra should be developedindividually according to the sitecharacteristics. Because the PWR is on1300 ft of soil over bedrock, individualspectra would be more appropriate.The vertical component of the earthquakefor the PWR design is two-thirds of thehorizontal component; the current designrequirement (as indicated in RegulatoryGuide 1.60) is as follows:Frequency (cps)3.5HorizontalVertical2/32/3 to 11The first two frequencies for the seismicanalysis model of the PWR reactorbuilding were computed to be 2.34 and5.10 cps, which fall on either side of3.5 cps, implying that the requiredseismic response spectra in the verticaldirection are approximately the same asin the horizontal direction.X- 39

The seismic loads used in the design ofthe PWR were in accordance with acceptedpractice at the time the plant wasbuilt; however, the Class I structureswould be subject to seismic loads higherby a factor of approximately 2 if currentdesign spectra and damping valueswere imposed. On the other hand, speciallydeveloped spectra that account forthe 1300 ft of soil over bedrock mayshow the factor to be less than 2.Floor response spectra at the charginglevel of the reactor building were developedon the basis of the Taft earthquaketime history record. A dampingvalue of 0.5% critical was used to generatethese curves. Table X A-9 showsthat currently acceptable damping valuesare 2% critical or higher. It is estimatedthat the higher damping valueswould lead to a seismic load reductionof 30%. However, since the seismicloads on structures may increase by afactor of 2, the net result may be anincrease of 40% in the seismic loads onequipment.A6.1.1.1.2 Conclusion. We concludethat:a. The seismic loads used in the designof the PWR were in accordance withaccepted practice at the time theplant was built.b. The use of current design responsespectra curves and current dampingvalues could lead to:1. An increase in seismic loads onPWR structures by a factor of upto 2.2. An increase inPWR equipment(Note that wethe effectloads.)seismic loads onof up to 40%.do not considerof these higherA6.1.2A6.1.2.1BWR PLANTSeismic Loads.Two Earthquakes with maximum groundacceleration have been considered asfollows:a. Design earthquake (DE)0.05g horizontal2/3 (0.05)g verticalb. Maximum credible earthquake (MCE)0.12g horizontal2/3 (0.12)g verticalClass I structures have been analyzed onthe basis of the ground response spectrashown in Fig. X A-6 (DE) and Fig. X A-7(MCE). These curves are based onHousner's average response spectranormalized to the maximum expectedground acceleration.Class I BWR structures (reinforced concrete)have been assigned 2% criticaldamping for the DE and 5% critical forthe MCE; currently acceptable values are4% and 7% (Table X A-10).Figure XA-6 shows the current RegulatoryGuide 1.60 design response spectra for4% critical damping (curve 4) as comparedto the 2% (curve 2) used in the designof the BWR structures for the DE.These curves show a current peak accelerationof 0.18g at 2.5 cps as comparedto a peak value of 0.15g at 3.5 cps forthe DE. Similarly, Fig. X A-7 shows acurrent peak acceleration of 0.27g at2.5 cps as compared to 0.23 g at 2.5 cpsfor the MCE.Note that the vertical stiffness of theClass I structure was high; thus amaximum vertical ground acceleration oftwo-thirds of the horizontal groundacceleration was used in the design.Regulatory Guide 1.60 indicates that forfrequencies greater than 3.5 cps the124Wc. A specially prepared site spectraaccounting for local soil characteristicsmay show that the increasesare not as large as indicated initem b.A6.1.1.2Damping.Table X A-9 shows the damping valuesused in determining the dynamic responseof the PWR and the values currently requiredby the AEC as described in Ref. 2.1 The DE for the BWR is the same as theOBE for the PWR, i.e., an earthquakefor which safety-related equipmentshould remain operational during andafter the event.2 The MCE for the BWR is the same as theDBE for the PWR, i.e., an earthquakefor which safety-related equipmentshould remain operational so that asafe and orderly shutdown of the plantcan be achieved and maintained.X-40

vertical acceleration is equal to thehorizontal acceleration.Equipment was designed using floor responsespectra generated on the basis ofa Taft earthquake time history recordmodified so that its associated responsespectra are equal to or greater than thesite ground response spectra.Table X A-l0 shows that currently acceptabledamping values are generally 2%critical or higher as compared to the0.5% used in the BWR for equipment suchas piping. This difference could leadto a seismic load reduction of 30% onequipment, which, when factored with theincreased seismic loads on structures,leads to a slight net reduction in seismicloads on equipment.A6.1.2.1.1 Commentary.a. The DE and MCE maximum ground accelerationsappear to be reasonable andappropriate in view of the seismologicalcharacteristics of the areaas described in section A2.1.2.b. Since Class I structures are builtdirectly on bedrock (which wasoverlain by only 10 ft of soil),there is no need to consider theproblems and uncertainties associatedwith soil-structure interaction(see section A6.3.1.1). Thus, thebasic seismic input load characteristicscan be viewed with a highlevel of confidence.C. The design response spectra curvescurrently required by the AEC (Ref.1) could lead to an increase inseismic loads of up to 20% for theDE and 17% for the MCE in the horizontaldirection and an increase of50% in the vertical direction forClass I structures.A6 .1. 2.1. 2that:Conclusion.We concludea. The seismic loads used in the designof the BWR were in accordance withaccepted practice at the time theplant was built.b. Class I structures could be subjectto seismic loads up to 20% in excessof those used for the horizontalearthquake and 50% higher for thevertical earthquake if current designresponse spectra were imposed.our review considered the designadequacy of buildings only from theviewpoint of mathematical modelingto obtain seismic inputs for equipment.The effect of higher seismicstresses on structures due to thesehigher loads has- not been evaluatedtherefore.C. Seismicslightlyspectraused.A6.1.2.2 Damping.loads on equipment could bereduced if current responseand damping values wereTable X A-10 shows the damping valuesused in determining the dynamic responseof the BWR and the values currentlyrequired by the AEC as described inReference 2.The damping values used in analyzing forthe DE are all less than or equal tocurrently accepted values. Similarly,the values used for the MCE are lessthan the currently accepted values, withthe exception of those for steel-framestructures and bolted and riveted assemblies,which are somewhat higher thanthe current values.A6.2 TORNADO EFFECTSA6.2.1PWR PLANTThe tornado design criteria governingthe diesel generator building are givenin Reference 3 and summarized asfollows:a. External wind forces resulting froma tornado having a rotational velocityof 300 mph and a translationalvelocity of 60 mph for a combinedmaximum velocity of 360 mph. Thediameter of the eye of the tornadois 1000 ft, and the maximum windsextend a radius of 200 ft.b. Atmosphericin 3 sec.relieve ordrop.pressure dropVenting canreduce thisof 3 psibe used topressurec. Missiles equivalent to a 40-ft-longby 12-in.-diameter 50 lb/ft 3 utilitypole with a velocity of 150 mph andimpacting on horizontal or verticalbuilding surfaces and a 1-ton autowith a velocity of 150 mph.In addition, the criteria incorporate aprocedure for calculating the windforces, as set forth in ASCE Paper No.3269, Transactions, 1961, and state thatwind forces and the effects of missiles.shall be considered simultaneously. Thefollowing allowable stresses are alsoestablished:a. Rolled steel plates and shapes - 90%of minimum specified yield stress.X- 41

. Steel-reinforcing bars - the minimumyield stress multiplied by the capacityreduction factors as specifiedin Section 1504 of ACI 318-63.c. Concrete - 75% of the ultimatestrength.CommentaryThe criteria for tornado winds and missilesare similar to those consideredtoday and are considered adequate. Theallowable stress limits for the steelreinforcing bars appear to have no conservativemargin.Section A6.3.6.1 presents a detaileddiscussion of the structural resistanceof the diesel generator building totornado loads.A comparison of the characteristicslisted in Table X A-ll shows that thecurrently considered auto missile with avelocity of 147 ft/sec (100 mph) hasabout twice the momentum of the autospecified for the BWR under review. Thesliding panel under current design requirementshas a kinetic energy per unitimpact area (a measure of penetrationcapability) more than four times asgreat as the wood plank considered inthe BWR design.Section A6.4.6.1 presents a detaileddiscussion of the structural resistanceof the diesel generator building totornado loads.A6.3 DESIGN ADEQUACY OF PWR PLANTCOMPONENTSA6.2.2BWR PLANTA6.3.1REACTOR BUILDINGThe diesel generator building is requiredto meet the design limits under thefollowing loading conditions:a. External wind forces resulting froma tornado having a horizontal peripheraltangential velocity of 300mph, which includes the tangentialand translational components.b. Differential pressure of 3 psi betweenthe inside and outside offully enclosed areas. Blowoutpanels are to be included wherenecessary in the design of thestructure to limit pressuredifferentials.c. Missiles equivalent to a 4-in.-thickby 12-in.-wide by 12-ft-long woodplank traveling end-on at 300 mph,or a 4000-lb passenger auto, with acontact area of 20 ft2 , flyingthrough the air at 50 mph not morethan 25 ft above ground.d. A torsional moment resulting fromapplying the wind specified in itema, above, acting on one-half thelength of the building.CommentaryGeneral practice at present is to considera tornado with a combined horizontalperipheral tangential velocity of360 mph.More recent plant design criteria considera greater variety of more severemissiles. A comparison of the missilesused in the BWR design with those usedin more recent plants is given in TableX A-11.The reactor building in this plant is areinforced-concrete containment structure.As shown in Fig. X A-8, thestructure is basically a cylinder with ahemispherical dome and a flat base. Theinternal height is 185 ft and the ID is126 ft. The base mat is 10 ft thick,and the cylindrical walls are approximately5 ft thick. During normal plantoperation the containment is maintainedat a subatmospheric pressure of 9.0 to11.0 psia.A6.3.1.1Soil-Structure InteractionModel.The seismic response of the PWR reactorbuilding structure to the horizontalcomponent of the earthquake was originallyobtained on the basis of a 4-degree-of-freedom mass-spring model asshown in Fig. X A-9. Here kI and k 4represent the static horizontal androcking soil stiffnesses, respectively.A subsequent analysis was performedusing a 14-degree-of-freedom model. Inthis model the external and internalstructures are represented by 7 and 5masses, respectively, as compared to 1mass each in the 4-degree-of-freedommodel. For both models each mass isassumed to have a horizontal translationaldegree of freedom. In addition,the structure is assumed to be capableof rotating as a rigid body in a singlerocking mode.An independent check was made to determineif significant rocking modes otherthan those permitted by the A/E modelcould exist. In particular, it wasnoted that the internal structure fromthe base mat to a height of 25 ft abovethe base mat consists of a blocked-outX-42

structure (columns uniformly distributedon the outside diameter of the structure).A rough calculation indicatesthat the rotational stiffness of theinternal structure supported on theblocked-out section is 7.7 x 109 ftkip/radas compared to 2.7 x 109 ftkip/radrotational stiffness of thesoil. The total rotatory inertia participatingin the rigid-body rockingmode is estimated at 20 x 106 kip-sec 2 -ft, while the rotatory inertia associatedwith the internal structure is only0.8 x 106 kip-sec 2 -ft. It follows thenthat the frequency associated with therocking of the internal structure isconsiderably higher than that of theoverall rigid-body rocking mode. Thisimplies that the dynamic response of theinternal structure rocking mode is of nosignificance compared to the overallrigid-body rocking mode.The frequencies computed using the 4-degree-of-freedom model are comparedwith those computed using the 14-degreeof-freedommodel in Table X A-12.The results shown in the table are basedon a soil shear modulus of G = 18,000psi (2592 ksf). The A/E estimates thatthis value may in fact vary as much as+15% or -30%. The sensitivity to thisvariation in soil property was consideredby the A/E in the 14-degree-offreedommodel but not in the 4-degreeof-freedommodel.The 4-degree-of-freedom model was usedto compute the seismic stresses in thereactor building on the basis of theground response spectra, whereas the 14-degree-of-freedom model with a timehistory record was used to determine thefloor response spectra at the chargingfloor elevation.The A/E investigated the possibility ofsoil liquefaction (loss of shear resistance)under seismic vibratory loads andconcluded that there is an adequatemargin of safety against liquefactionfor a hypothetical ground accelerationof up to 0.25g.The response of the system was determinedfrom the ground response spectracurves corresponding to 5% of criticaldamping for the OBE and 10% of criticaldamping for the DBE. The A/E performedcalculations to estimate strain energydissipated in each mode and came up withaverage values of percent criticaldamping for each mode as shown in TableX A-13.A6.3.1.1.1 Commentary.stiffness calculationThe soilisbased onelastic half-space loaded with a rigiddisk (representing the base mat), withformulae taken from Reference 4. Theanalysis, however, did not consider theeffective mass of the soil, which isestimated to be approximately 25% of thebase mat mass or less than 10% of thetotal mass of the reactor building andhence of minor significance.Since the motion of the structure isdominated by the first three modes,Table X A-13 indicates that the 5% and10% damping values used to determine thedynamic response are reasonable.The model does not account for embedmentof the containment structure in the soilnor does it account for the differingshear moduli of the various soil layers.Re-examination of the soil-structureinteraction problem using finite elementtechniques could account for theseeffects.A6.3.1.1.2 Conclusion. We concludethat the soil-structure interaction modelgives reasonably conservative seismicloads on the structure.A6.3.1.2Containment InternalStructure.The containment internal structureconsists of a reactor support pedestal(13 1/2 ft radius, 4 1/2 ft thick) and25-ft-tall columns arranged around a103-ft-diameter circle (see Fig. X A-8).The cross sections of these columns areapproximately 8 ft in the circumferentialdirection and 2 ft in the radialdirection. The lower portion of thereactor support pedestal (approximately14 ft above the base mat) has a part cutout of its cross section (Fig. X A-10).The internal structure was reviewed onlyfrom the viewpoint of mathematicalmodeling to obtain seismic inputs toequipment such as the reactor coolantsystem loop. The structure itself wasnot evaluated from the viewpoint ofseismic resistance.A6.3.1.2.1 Commentary. Review of thedesign calculations indicated that themathematical modeling of the lowerportion of the internal structure (thefirst 25 ft above the base mat) does notproperly account for the considerablereduction in the lateral stiffness dueto the blocked-out structure. The A/Ereplaced the individual circumferentialcolumn cross sections by an "equivalent"annular ring with a reduced thickness.The cross-sectional moment of inertia ofthis ring, together with that of thecore, was computed, and the sum of theseX-43

values, together with the sum of thecross-sectional shear areas, wasassigned to the beam properties of thestructure between the base mat and themass 1 of the 14-degree-of-freedommodel.Independent calculations indicate that astiffness value equal to 2,812,000kip/ft would result for the lower portionof the internal structure with thedimensions indicated for the equivalentsolved cylinder representing theblocked-out internal structure.This stiffness value, however, is unrealisticin that the A/E's mathematicalmodeling implies that all the columnshave some shear connection throughouttheir 25 ft length when, in fact, theyare connected only at the top (via floorslabs) and at the base mat. The appropriatemethod of computing the lateralstiffness of the lower portion of theinternal structure is by summation ofthe lateral stiffness of the individualcolumns, under proper consideration ofindividual column orientations, and ofthe core. Such an independent ca: -ulationresults in a lateral stiffness of1,262,041 kip/ft for the reactor supportand 185,000 kip/ft for the peripheralcolumns or a total of 1,447,016 kip/ft.In accordance with the above, the actualstiffness of the internal structure isabout 51% of that used by the A/E. Inaddition, since the lateral stiffness ofthe lower portion of the internalstructure is only a fraction of that ofthe upper 96 ft, more of the internalstructure mass will participate in thefirst flexural mode (the third mode ofthe soil-structure interaction model)than is accounted for in the A/E'smodel. It is estimated that twice asmuch of the internal structure mass willparticipate in the flexural mode than iscalculated by the A/E. This means thatboth the modal participation factor(which is a measure of the participationof a particular mode to the total forcingfunction) and the modal effectivemass (which is a measure of the modalreaction at the base of the structure)will be greater for the flexural modethan is implied by the A/E model.As a result of our investigation, theA/E revised his calculations using thecorrect stiffness properties and recomputedfrequencies, mode shapes, modalparticipation factors, and seismicresponse. A comparison of the originalcalculations with the revised calculationsshows very small differences ofless than 10%. The reason for thissmall difference, in spite of adifference of almost 100% in an elementstiffness value, can be explained byreferring to Fig. X A-9. Note that theinternal structure (m 3 ) mass is 310 kipsec2 /ft as compared to a total mass of3185 kip-sec 2 /ft, that is, less than10%. Apparently the system dynamics aredominated by the rigid-body motionassociated with the first two modes asrevealed by the modal participationfactors shown in the following table.Mode1234Modal Participation Factor1.563-0.7080.140-0.012A6.3.1.2.2 Conclusion. In spite of thedifferences noted in the commentary, thesignificance of the error in computingthe stiffness properties of the internalstructure is negligible since theseismic response is dominated by therigid-body modes (rocking and translation)of the system. We conclude,therefore, that the seismic input toequipment resulting from the seismicanalysis of the internal structure isadequate.A6.3.1.3Paint Coatings WithinContainment.The internal surface of the containmentliner, internal concrete structure, andmetallic components are coated withspecial paint coatings. Aside fromaesthetic purposes, coatings are used toprevent corrosion of metal surfaces andto seal the concrete. The coatings mustbe able to withstand normal and accidentenvironments without becoming separatedfrom the walls in any way, such as byruptured blisters, flaking, and cracking.Excessive delamination of coatingfilms could cause the strainers in thecontainment spray and safety injectionsystems to clog and would interfere withtheir proper operation.Two specimens of each of the coatingsystems were subjected to a qualificationtest involving simultaneous exposureto gamma radiation, steam, andchemical spray for a period of 7 days.The exposure to simulated LOCA conditionsbegan with a rapid injection ofsteam, followed by a 40-min dwell at280°F and a 20-min drop to 1401F, wherethe temperature was held for theremainder of the test. The coatingspecimens were sprayed with threedifferent chemical solutions during thetest. The total gamma-radiation doseX-44

eceived by the coatings during the testwas 100 megarads.The specimenstest showed noof any type.at the conclusion of thesignificant degradationA6.3.1.3.1 Commentary. Comparison ofthe qualification test conditions withthose postulated under maximum LOCAconditions shows that during the firsthour, the test temperature profile exceedsthe calculated containment temperatureprofile by a substantial margin,the smallest difference being about 10'Fat the peak containment temperature.The margin between the two temperaturesdecreased during the second hour of thetest; thereafter, the test temperaturewas the same as the calculated designbasis accident (DBA) temperature. Thegamma-radiation exposure and chemicalsprays used in the test were essentiallythe same as those expected in the plant.Based on present qualification philosophyand ignoring, for the moment, considerationsregarding the number ofspecimens tested, the coatings may beregarded as being qualified for somewhatless than a week following a LOCA event.The utility has reported that this is anadequate period for the coatings toretain their integrity. A few daysafter a LOCA event, cooling requirementsare not critical, and any degradation ofcooling system performance caused bycoating failure would not have seriousconsequences.There remains the question of whethertwo specimens constitute an adequatesample of each coating system. Analysisof this subject has shown that testsmust be made of a relatively large numberof specimens (about 50) with nocoating failures to yield modest confidence(about 90% confidence level) inthe adequacy of the coating system. Inaddition, experience has shown that itis difficult to establish that thecoatings on test samples are fullyequivalent to the coating actuallyapplied because the application techniquesand local climatic conditionshave a strong influence on the abilityof the coating to adhere properly underLOCA conditions. In this respect thequalification test is somewhat deficientwhen judged in terms of the evolvingqualification philosophy. 1 At the timeiAt this time, a formal, accepted standardis not in existence, and there isno explicit requirement regarding samplesize.the plant was constructed, however, thequalification requirements were relativelyrudimentary; thus the resultsobtained were regarded as being adequateto demonstrate satisfactory performance.It should also be stated that thequalification test was more realisticthan many that are conducted in thatnuclear radiation, steam, and chemicalspray were applied simultaneously. Mosttests that have been conducted haveinvolved sequential exposures to theseenvironmental conditions.A6.3.1.3.2 Conclusion. The qualificationtest established that there is areasonably good probability that thecoating systems will not experiencelarge-scale failure (flaking and/or delamination)should a LOCA occur.A6.3.1.4Crane.The crane is at elevation 99 ft 1 1/2in. directly over the containment sprayheaders. A dynamic analysis of thecrane was performed by the A/E. Thefundamental frequencies of the cranewith the hoist trolley at mid-span withno hoisted load on the hook were foundto be 4.2 cps in the vertical directionand 2.27 cps in the horizontal direction.The A/E determined the seismicaccelerations from the 5% equipmentdamping response spectra curves at thecharging floor elevation to be as follows:EarthquakeVertical-OBEVertical-DBEHorizontal-OBEHorizontal-DBEAcceleration (g)0.190.300.620.84The A/E stresses computed on the basisof the above accelerations are shown inTable X A-14. Note that the maximumstress was found to be 24.12 ksi; itoccurs in the end ties (which connectthe two girders that span the cranewall) and is due to the horizontal DBE.This stress value is compared to 90% ofthe yield strength = 32.4 ksi.A6.3.1.4.1 Commentary. The seismicevaluation of the crane, under the assumptionthat there is no lifted load onthe hook, is reasonable since this wouldbe the normal condition when the reactoris in operation.The acceleration values considered bythe A/E were based on the response spec-X-45

trum atassumingtne charging f loor elevation,equipment damping of 5% ofcritical. This high damping value isnot in agreement with the values givenin the FSAR. Table X A-9, which summarizesthe FSAR values, shows a 2% dampingvalue for mechanical equipment forboth the OBE and the DBE. The currentlyaccepted values, as given in RegulatoryGuide 1.61, allow 2% for the OBE and 3%for the DBE. In our judgment, if thesedamping values are used, the higherdesign ground response spectra given inRegulatory Guide 1.48 should be used.If we adhere to the 2% values given inthe FSAR for the PWR, then the peak accelerationsin the horizontal directionwould be 1.425g for the DBE and 1.05gfor the OBE, which occur over the frequencybandwidth of 1 to 8 cps, encompassingthe 2.27 cps natural frequencyof the crane in the horizontal direction.In addition, the response spectrum atthe charging floor (elevation 47 ft 4in.) would not reflect the higher accelerationsat the crane elevation, becauseof the rocking mode of the containmentstructure. An estimate of the relativevalues of acceleration at the two elevationsmay be obtained by comparing theproduct of the modal participation factortimes the mode shape coefficient foreach elevation. Examination of theA/E's computer output for the seismicanalysis of the containment structurereveals that this product for the rockingmode is 1.274 for the craneelevation as compared to 0.9033 for thecharging floor elevation.In light of the above it is estimatedthat the maximum computed stress in theend ties for the horizontal DBE wouldbe:1.425 x1.274 x 24.1 = 57.7 ksi0.8-4 0._9033Similarly for the OBE, theputed stress would be:maximum com-1.05 1.-274 x1.1=4. s0._62 x 0.9033x7812ksIt is apparent that the OBE will inducesome plastic flow and that the DBE willinduce plastic flow. The plate is approximately3 ft deep and 3/8 in. thick;thus buckling on one side under theseloads is likely, thereby further increasingplastic flow on the tensionside.A6.3.1.4.2 Conclusion. Although theestimated stress levels exceed the 90%yield strength criterion established bythe FSAR, it is our judgment that thereis sufficient reserve strength (theultimate tensile strength is estimatedto be twice the yield strength) toprevent catastrophic failure.It is our judgment that catastrophicfailure of the crane, which in turncould damage the recirculation sprayheaders (Fig. X A-8), is not expected.A6.3.2REACTOR COOLANT SYSTEM (RCS)The reactor coolant system consists ofthe reactor pressure vessel (containingthe fuel), the pressurizer, three essentiallyidentical coolant loops, a pressurizerrelief tank, connecting piping,and instrumentation. Each reactor coolantloop contains a steam generator, areactor coolant pump, two loop isolationvalves, and interconnecting piping.The reactor coolant piping, steam generators,and reactor coolant pumps areprovided with snubbers, which accommodatethe thermal expansion from thefixed or anchored reactor vessel with noresistance, but which do resist earthquakeand pipe-whip forces. Steamgeneratorrestraints are provided nearthe bottom of each steam generator toresist lateral and vertical loads resultingfrom seismic and pipe-ruptureforces. Additional restraints are providedat a higher elevation to resistlateral loads resulting from seismic andpipe-rupture forces.Portions of the system are discussed insection A6.3.2.1 through A6.3.2.5.A6.3.2.1 Reactor Coolant System Loops.The mathematical model for one of thereactor coolant systems loops, as developedby the supplier, is shown in Fig.' XA-11. Joint 33 represents the supportpoint for the pump. Joints 14 and 18represent support points for the steamgenerator. The loop is estimated toweigh 1340 kip.The mathematical modeling is characterizedby the following features:a. Local flexibility of the nozzleshelljuncture is not accounted for.b. The actual pump and steam generatorsupport members are not representedin the model; instead, their stiffnesscharacteristics are determinedin a separate calculation and assignedto the support points.ax-46

c . The main steam lines and feedwaterlines are not represented in themodel.d. Dynamic interaction between the loopand the reactor building internalstructure (which is estimated toweigh 27,000 kip) is not considered.The supplier has determined that thefundamental frequency of the loop is5.12 cps and that the pump motion dominatesthe fundamental mode. The nexttwo modes in which steam-generator motionpredominates have frequencies of6.57 cps and 7.08 cps.The OBE floor response spectrum used indetermining the seismic response of theloop is shown in Fig. X A-12. Note thatthe peak acceleration value 1.68g occursbetween natural periods of 0.125 sec and0.9 sec corresponding to frequencies of8 and 1.1 cps.Critical stresses were computed for the400 elbow in the upper crossover legnear the steam generator outlet nozzleas shown in Table X A-15.A6.3.2.1.1 Commentary. The A/E computeda fundamental frequency of 8.32 cpsras noted in section A6.3.2.2. As notedin the commentary, this higher value isprobably correct. If the fundamentalfrequency is indeed higher than 8 cps,then Fig. X A-12 indicates that thespectral acceleration values for the OBEwould be between 0.14g and 0.21g, whichis 1/8 the peak value of 1.68g. Theactual seismic loads may be significantlyless than indicated in the analysisperformed by the supplier.The fundamental frequency of 5.12 cpscomputed by the supplier for the loop is97% of the computed third-mode frequency(5.3 cps) of the containment, which isdominated by flexure of the reactorbuilding internal structure. Since theloop is braced to the internal structurevia steam generator and pump supports,there could be significant dynamic interactionbetween the structure and theloop.RDT Standard F9-2T (tentative), p. 20,(Ref. 5) indicates that a single modelwhich couples both the equipment (theloop in our case) and the support (theinternal structure in our case) shouldbe used if:0.5 !ý f /f < 2.0unless;where:Me/M s< 0.0001f e= natural frequency of supportedsystem (the loop)f S= natural frequency of support(internal structure)M =total mass of supported systemMs = total mass of supportIn this case Me/Ms = [3(1340)1/27,0000.15 since there are three loops.The possible values for fe/fs for eachof 2 configurations are given in Table XA-16.Note that both configurations indicatethe need for a coupled model. The sup--plier has indicated that in the designof recent reactor plants a coupled modelwas used and this has resulted in adecrease in seismic stresses in the loopby 30%; so, in this special configura--tion, inclusion of coupling appears tohave value only for obtaining a moreaccurate assessment of stress in theinternal structure.A6.3.2.1.2 Conclusion. It appears thatthe computed stiffness matrix for thereactor coolant pump support does notreflect the as-built condition; so thecomputed seismic stresses shown in TableX A-15 are probably higher than theywould be for the defined DBE and OBEfloor response spectra.Dynamic coupling between the loop andthe internal structure may provide additionallowering of seismic stresses inthe loop. We did not evaluate the seismicresistance of the internal structure;thus we cannot judge the significanceof the effect that inclusion ofcoupling would have on seismic stress inthe internal structure.We conclude that the seismic resistanceof the reactor coolant system appears tobe adequate.x- 47

A6.3.2.2 Steam-Generator and PumpSupports.Each primary loop is composed of a steamgenerator, a primary coolant pump, andpiping that connects these units to eachother and to the reactor vessel. Inthis particular plant the A/E, not thesupplier, is responsible for the designof the component support. Both the A/Eand the supplier have performed independentdynamic seismic analyses of theprimary loop. The A/E states in Ref. 6that the steam generator and reactorcoolant pump were modeled as rigidmembers. Figure X A-13 shows the mathematicalmodeling used by the A/E. Thesupplier accounts for the flexibility inthese members. The A/E provided thesupplier with the design details for thesupport of the pump and steam generatorto be included in the mathematical dynamicmodel of the loop.The design of the support structure forthe steam generator and the primarycoolant pump is governed by loads resultingfrom postulated circumferentialand longitudinal pipe breaks. Theseloads are far in excess of the computedseismic loads. Support loads due to 10postulated pipe ruptures were determinedon the basis of a time history dynamicanalysis using the first 30 modes of themathematical model shown in Fig. X A-13under the following assumptions:a. 0% critical damping.b. 0.015-sec rise time to reach maximumjet force at pipe break.C. Longitudinal break assumed to occurover two pipe diameters.d. section of longitudinal break assumedto maintain full axial loadcapabilities but cross-sectional momentof inertia reduced to 1b% ofthat of unbroken pipe.Table X A-17 summarizes the computedsteam-generator support member loadsresulting from seismic and pipe-ruptureevents. Table X A-18 summarizes thecomputed loads in the members supportingthe primary coolant pump. The originalsupport design of the pump had no horizontalsnubbers.Member loads were evaluated in accordancewith the following criteria:WLoad (a)W + DBEW +DBE +R(a) WDBEERFaAllowable (a)AISC code allowable1.33 x (AISC codeallowable)Tension -Shear -F y0.7 F yBending -F yCompression -1.6 Fanormal operating conditions;design basis earthquake;pipe rupture;specified yield point; andallowable compressive stress..The A/E computed a fundamental frequencyof 8.32 cps for the loop; the suppliercomputed 5.12 cps. The supplier alsodetermined second- and third-mode frequenciesof 6.57 cps and 7.08 cps,respectively. The fundamental frequencymode is dominated by pump motion, whereasthe second and third modes are dominatedby the steam generator.A6.3.2.2.1 Commentary. The 8.32-cpsfundamental frequency computed by theA/E gives a horizontal seismic spectralacceleration of 0.21g; the 5.12-cps fundamentalfrequency computed by the suppliergives 1.68g. This could lead toan increase in seismic loads in thesupport members by a factor of 8. Webelieve, however, that the A/E's computationsreflect the addition of the horizontalsnubbers for the pump, indicatingthat the higher frequency valuerepresents the lower bound for the actualstructure, so that 0.21g accelerationis applicable.In addition, the assumption of 0% criticaldamping in determining the dynamicresponse from pipe-rupture loads is extremelyconservative; Regulatory Guide1.61 permits 3%. In our judgment anysupport member increase in seismic loadsdue to a lowered fundamental frequencywould be more than offset by a decreaseddynamic response to pipe rupture loadswith 3% critical damping.A6.3.2.2.2 Conclusion. The support designfor the pump and generator shouldbe capable of withstanding loads inducedby combined seismic and pipe ruptureevents. The design appears to beadequate.Vx-48

A6.3.2.3Reactor Coolant Pump Nozzles.Each of the three reactor coolant loopscontains a vertically mounted coolantpump and drive motor assembly (Fig. X A-14). The overall height of the unit is25 ft 3 in.This single-stage centrifugal pump isdesigned to deliver 88,500 gpm against a280-ft head. It is driven by a 6000-hp,single-speed, air-cooled induction motormounted directly above it.Coolant returning to the reactor fromthe steam generator is drawn up throughthe 31-in.-ID suction nozzle into theimpeller and discharged through the27.5-in.-ID nozzle in the side of thepump casing.As a part of the primary coolant loop,the pump casing and nozzles are classifiedas Class I seismic components.Table 4.1.6-1 of the FSAR (Ref. 3)indicates that although no specific codeprovision for pumps was in effect at thetime the pump was designed, the reactorcoolant pump casing was designed perArticle 4 of Section III of the ASMEBoiler and Pressure vessel Code (i.e.,to the same rules used for the reactorvessel design).The analysis of the PWR reactor coolantpump nozzles is contained in a genericreport prepared by the major equipmentsupplier (Ref. 7).In general, this major equipment supplierseeks to qualify standard componentsof the steam supply system on the basisof a master one-time analysis applicableto all reactor plants in which thesystem is installed.To accomplish this objective, the supplierperforms the master analysis for aset of thermal and mechanical loadingschosen to represent a conservative envelopeof the most severe loading casesexpected in any specific application.The loading cases used in the masteranalysis are referred to as the umbrellaloads. In the master analysis thiscomponent is shown to meet the requirementsof the applicable structural codeunder the umbrella loads. If the actualloadings in a given plant installationare subsequently shown (either by loadinglimitations established in the specificationsor by analysis) to be lesssevere than the umbrella loads, then thestructural adequacy of the component isconsidered to have been demonstrated forthe specific application.Stresses arising from a variety of loadingsources are required for structuralevaluation of the nozzles. In theanalysis of the 93A casing nozzles (Ref.7), the following procedures are used toobtain these stresses or to qualifycomponents:a. Determine stresses in the nozzlewalls from pipe reactions using beamtype formulae. These equations areincorporated in computer programNUMBRA used for this analysis.b. In the pump casing wall at the junctionwith the nozzles, determinestresses from pipe reactions usingthe formulae of Welding ResearchBulletin 107 (Ref. 8) . TheseBijlaard equations are also incorporatedin computer program NUMBRA.C. Determine thermal stresses (for normaland upset conditions only) inthe nozzle wall due to radial temperaturegradients, using simplecylindrical models of correspondingdiameter and wall thickness.d. Qualify the discharge nozzle for thefaulted condition with a specialfinite element model in conjunctionwith a limit analysis using computerprogram ANSYS. The model used is afull 3600 model with a refined gridof triangular flat-plate elementswith plastic properties. In thisanalysis the boundary at the junctionwith the pump casing was takenas fixed.e. Qualify the casing and nozzles forfatigue by invoking the exemptionprovisions of paragraph NB-3222.4d,Section III,,ASME Code.A6.3.2.3.1 Comamentary. When the methodof a master analysis using umbrellaloads is used to qualify a component fora specific application, it should bedemonstrated that actual loads do infact fall within the limits of the umbrella.The equipment supplier had, atour request, prepared such a comparisonfor the working session. However, thiscomparison was confined solely to thepipe reactions at the nozzles due toseismic loads. Although this limitedcomparison showed actual loads to bebelow the corresponding umbrella loads,it did not include other simultaneousloadings required by the code to becombined with those of seismic origin.The equipment supplier recently indicatedthat other loads have been combinedwith the seismic loads and that theumbrella loads have not been exceeded.x-49

For the discharge nozzle, the stressesin the pump casing wall at the junctionwith the nozzle are assessed usingBijlaard's method of analysis. Thisapproach is of doubtful value in thisapplication since the conditions forvalid application of Bijlaard's methodare not present.Bijlaard's analysis may be used to determinethe stresses in a shell at itsjunction with a smaller cylinder (nozzle),provided the shell is eitherspherical or cylindrical, the intersectingcylinder is normal to the surface,and its diameter is less than one-thirdthe diameter of the shell. In addition,since the analysis is based upon thedeformation characteristics of theshell, the shell is presumed to be thinand uniform.In this application, the following conditionsinvalidate the approach:a. The nozzle diameter exceeds onethirdthe diameter of the pump casing.b. The casing wall certainly cannot beconsidered uniform in the vicinityof the discharge nozzle. Along itsupper periphery, this nozzle abuts athick flange.c. The required condition of uniformityfor the shell is also violated inboth cases by the mutual proximityof the two nozzles; interactionstresses may also occur in the regionbetween the nozzles.The pump casing was not specificallymodeled so that the general stress distributionin the pump casing wall underfaulted conditions was not established.For the discharge nozzle to qualify, alimit analysis using finite elementswith plastic properties was required.This suggests that stresses in the pumpcasing for the faulted condition may bequite high and that a correspondingfinite element analysis of this regionis needed.A6.3.2.3.2 Conclusion. There is insufficientinformation to assess designadequacy. In our judgment a finite elementanalysis is needed before designadequacy can be established.A6.3.2.4Pipe-Whip Restraints.Both the main steam line and the feedwaterlines that connect to the top andside of the steam generator, respectively,are provided with pipe-whip restraintsdesigned to prevent a rupturedpipe from damaging the containmentstructure. Restraints for the mainsteam line, for example, are installedon 5-ft centers, which is equal to theradial distance between the pipe lineand the inside of the containment wall.Figure X A-15 shows the restraint designconcept for the main steam line. Notethat the 6-in. space between the outsideof the pipe and the inside of therestraint should allow unrestrained motionof the pipe under thermal and seismicloads. The pipe comes in contactwith the restraint only if a pipe breakcauses jet forces that induce very largepipe movements.The restraints for the main steam linehave been designed to resist a piperuptureload equal to:whereP = 1.25 p Ap = the design pressure, andA = the flow area.The restraints have been designed todevelop a plastic hinge when the thrustload P is applied at the point of contactbetween the pipe OD and the insideof the restraint.The same procedure (including the use ofthe formulae for calculating P) was appliedin designing the restraints forthe feedwater line. Material for thefeedwater line pipe-whip restraint isAlSl 1144 bar stock with a minimum yieldstrength of 53,000 psi.A6.3.2.4.1 Commentary. Analyses wereperformed to establish the thrust coefficientappropriate to the PWR plantfeedwater line. When the steam generatorwater level drops below the feedring, high pressure steam will enter thefeedline and accelerate an ever diminishingwater mass down the pipe to thebreak.Analysis of a break near the check valveof a 170 ft feedline shows that withconservative mathematical approximations,an impulse with peak force 1.35pA will occur. The analysis is based onthe water in the steam generator beingsubcooled by the amount corresponding tonormal operation. It is anticipatedthat a more refined calculation wouldreduce the coefficient below 1.25 forthis case. The refined calculationshould also be carried out for theV-0x-50

minimum amount of subcoolingoccur in the steam generator.that mayA6.3.2.4.2 Conclusion. The pipe-whiprestraint design for the main steam lineappears to satisfy the design criteria.Present analysis for the thrust coefficientin the feedwater line predicts athrust coefficient of 1.35 existing fora very short time. However, the analysishas some conservatism. In our opinionthe integrity of pipe-whip restraintsfor the feedwater line (basedon a design thrust of 1.25 pA) isunlikely to be compromised; therefore,these restraints also may be judged assatisfying the design criteria.A6.3.2.5 Snubbers.The hydraulic snubbers are designed tooffer no resistance to slowly varyingpipe displacements which, for example,accompany thermal transients but "lock"under the rapid pipe motions induced bydynamic loads such as earthquakes.The snubbers contain a hydraulic fluidand a movable piston; hence seals mustbe provided to prevent the loss of fluidaround the shaft and at certain otherlocations. At the time the snubbers forthe PWR plant were being designed,polyurethane was recommended by the sealindustry on the basis of limited laboratorytesting. Additional testing sincethat time, plus the fact that the sealsin a large number of snubbers built by amanufacturer different from the one whobuilt those for the PWR plant havefailed after installation in nucleargenerating plants, has resulted in theindustry being more keenly aware of theeffects of the hydraulic fluid, ambienttemperature and nuclear radiation on theseal materials. We have been providedwith a copy of the report of testsperformed by the supplier of the sealmaterial used in the snubbers at the PWRplant, and a copy of the specificationsheet for the hydraulic fluid (Ref. 9).A6.3.2.5.1 Commentary. Opinion is dividedabout whether polyurethane, ethylenepropylene, or some other material isbest, but, from a review of the availableinformation, it appears that thereare formulations of either compound thatare adequate for at least the period oftime between refueling outages. A formalqualification program to demonstratethe ability of snubbers to maintainproper sealing of the hydraulic fluidunder the operating environment (includingsevere seismic disturbance followingtemperature and radiation aging, and aLOCA environment) should be conducted.A6.3.2.5.2 Conclusion. Our evaluationof the information available regardingthe hydraulic snubbers installed at theplant leads to the conclusion that theyare probably adequate for the period oftime between refuelinq outages.A6.3.3LOW HEAD SAFETY INJECTION SYSTEM(LHSIS)The LHSIS consists of two pumps, which,in the event of a large-break LOCA,initially draw borated water from therefueling water storage tank and injectit into the RCS hot legs. After theinitial injection phase, system valvesare realigned, and these pumps recirculateto the RCS the spilled reactorcoolant and injected borated water.The low head safety injection pumps arelocated in the safeguards area alongsidethe containment building; their pumpimpellers are actually located within anextension of the containment boundary.The pump casing is connected by a horizontallength of piping to the containmentsump.The vertical centrifugal pumps have adesign discharge pressure of 300 psigand a design flow rate of 3000 gpm each.An unusual feature of these pumps is therelatively long shaft necessary to connectthe impeller, at elevation -32 ft,to the motor, at elevation +12 ft.A6.3.3.1Piping.The LHSIS piping system, consisting of10-in. and 12-in. Schedule 140 piping,is extensive with lines running inside,outside, and around the containmentstructure; it also extends in elevation(in some sections) between the accumulatortank supported at the base mat atelevation -27 ft. 7 in. to a maximumelevation of 24 ft. 6 in. and finallyconnects to the reactor vessel primarypiping at elevation 14 ft.The A/E determined seismic stresses forsections of the piping between anchorpoints in accordance with the followingprocedure:a. Masses are lumped at short intervalsalong the pipe line, and each massis assigned 3 translational degreesof freedom.b. Frequencies, mode shapes, and participationfactors are computed foreach mode.c. Spectral acceleration values areselected for each mode for earth-X-51

quakes occurring in each of theprincipal directions.d. If the anchor points are at twodifferent elevations, then the maximumresponse spectrum for the mostsevere point is used.e. Modal forces on the basis of items athrough d are then computed.f. Seismic stresses at various locationsin the piping are computed byselecting the highest modal forcefor each principal direction at eachmass point and combining it with thesquare root of the sum of thesquares of the remaining modalforces. The resulting force valuesfor each principal direction arethen applied to the mass points asan equivalent static load on thestructure, and seismic stress valuesare computed.g. Stresses due to seismically induceddifferential anchor motion are computedby applying the root meansquare of the seismic displacementsat the anchor position in oppositedirections.h. Seismic stresses generated in itemsf and g are combined with longitudinalpressure and dead weight stressesand evaluated in accordance withthe following criteria:whereSLP + SDL + SOBESLP + SDL + SDBE-l-' Sh18 ShSLp = longitudinal pressure stressSDL = dead weight stressOBE = operational basis imearthquake stress Primary StressesDBE = design basis earthquakestress )S = allowable stress at operatingtemperaturei. In addition, non-seismic stressesare evaluated in accordance with thefollowing criteria:SLP + SDL -STH

C. The method of applying the seismicinput simultaneously in all threeprincipal directions in effect meansthat the earthquake components actsimultaneously and in phase, whichis conservative.In connection with item c, at the timethe PWR was built, the AEC generallyaccepted seismic inputs in only onehorizontal direction and one verticaldirection, which, for example, couldlead to a stress value of 1.45 or 2.05depending on whether the A/E assumedthat the earthquake components were outof phase or in phase. on the basis offurther evaluation of earthquake records,the AEC is now requiring that allthree components of the earthquake beconsidered to act simultaneously but outof phase; this could lead to a stressvalue of 1.7S. The approach used in theanalysis of the LHSIS piping could lead,in effect, to a stress value of 3.OS andis therefore considered to be conservative.The evaluation considered primarystresses and thermal expansion stressesas required by the B31.1 piping code.Current practice would require analysisand evaluation in accordance with theNuclear Code (NB 3600 of ASZ4E CodeSection III), which considers evaluationof secondary stresses due to thermalgradients and cyclic loads. Since theLHSIS piping is not subject to significantthermal gradients and less than4,000 cycles of operation, secondarystresses and cyclic loading conditionsare judged acceptable.The sensitization of portions of thepiping appears to have been properlyhandled and should present no compromiseto design adequacy provided the surveillanceprogram outlined in Ref. 10 ismaintained..A6.3.3.1.2 Conclusion. The design ofthe LHSIS piping appears to be adequate.A6.3.3.2Pumps and Drives.The motor drive is mounted at elevation12 ft. 0 in.; the pump end is approximately50 ft. below that level. Theshaft that drives the pump is laterallybraced inside a pipe, which in turn isencased in concrete at elevation -27 ft.6 in. and laterally braced by a 2 3/8-in. -thick plate at elevation 12 ft. 0in.The supplier performedwhich indicate that thefrequency is 3.55 cps.calculationsfundamentalSeismic loads were taken as 3.0g. Theanalysis indicated that the shaft casingwould deflect 1.15 in. at the pump. Amaximum stress of 9519 psi was computedfor the inner casing at the lowerseismic support.The LHSIS pump supplier prepared staticg-load (3.0 g) calculations to investigatethe ability of LHSIS pump tooperate during and after MCE deflectionsof 1.15 inches. These calculationsincluded:a. Examination of pump shaft deflectionsand stator-rotor clearancesunder conditions of maximum deflectiondue to lateral seismicacceleration.b. Computationloads dueacceleration.of pump shaft bearingto lateral seismicC. Computation of simultaneous additionalbearing loads due togyroscopic torque during seismicvibration.The LHSIS pumpshow:supplier's computationsa. Stator-rotor clearance is not functionallyimpaired.b. The maximum bearing loads to beexpected are small compared to therated load capacity of the bearings.The pump internals (gaskets, seals,etc.) can be exposed to high levels ofnuclear radiation during the recirculationphase following a LOCA, and the A/Estates that these items can withstandthis environment. Documentation madeavailable to substantiate this assertionare Refs. 11, 12, and 13.We understand that the motor drives areidentical to those used on the containmerntrecirculation spray pumps, forwhich a qualification program wasconducted (see section A6.3.5.2).A6. 3.3.2.1 Commentary. The analysisindicates that stresses in the pump arewithin acceptable limits. In our judgmentthe analysis provided to verify theability of the pump to operate underdeflections as large as 1.15 inchesprovides reasonable engineering assurancethat operability will not be impairedby deflections of this magnitude.However, these computations were staticnot dynamic analyses and are not theequivalent of a seismic vibration testX- 53

adequately following a severe earthquakeat the PWR plant.The other aspects of the test program,as reported in Ref. 15, demonstrate thatthe VMO design is rugged and can be reasonablyexpected to function properly inthe plant. This conclusion isreinforced by the fact that the VM'O arelocated separate from any high energypiping systems.A6.3.3.4.2 Conclusion. We judge thedesign of the VMO in the LHSIS to beadequate for the service conditions towhich they may be subjected.A6.3.3.5Snubbers and Hangers.The LHSIS piping is provided at variouspoints with the following types ofsupports:a. Hydraulic Snubbers. These units aredesigned to offer no resistance toslowly varying pipe displacementsthat, for example, accompany thermaltransients, but they "lock" underthe rapid pipe motions induced bydynamic loads, such as earthquakes.b. Hangers. Spring hangers support thepipe weight only, and offer negligibleresistance to pipe displacements.Rod hangers also support thepipe weight but offer resistance todownward pipe motion (since the rodis in tension) and negligibleresistance to pipe motion in otherdirections (since the rod readilybuckles).C. Restraints. These units are designedtrestrain pipe motion at aparticular point in one or twodirections.d. Anchors. These units are designedto restrain pipe displacements at aparticular point in all directions.The piping drawings tabulate the calculateddisplacements at each support aswell as the calculated forces that thesupport should be capable of resistingfor the various load conditions. Thisinformation is used by the supplier todesign the support. Support design calculationswere not reviewed.The design features of the hydraulicsnubbers and the environmental testresults described in section A6.3.2.5are applicable to the LHSIS snubbers.A6 . 3. 3. 5 . 1been usedmany years.Commentary. Hangers haveextensively in industry forIn addition, since they donot resist seismic loads, we need notinvestigate their design adequacy. Restraintsand anchors have also been usedextensively, and their design is suchthat they present no special maintenanceproblems. Since anchors and restraintsresist thermal loads during normal plantoperation, routine inspection would showany weaknesses to resist seismic loads.The hydraulic snubbers, however, onlyresist dynamic loads and are not"exercised'' by thermal loads duringnormal plant operation. As indicated insection A6.3.2.5 opinion is divided asto whether polyurethane, ethylene propylene,or some other material is best,but, from a review of the availableinformation, it appears that there areformulations of either compound that areadequate for at least the period of timebetween refueling outages. A formalqualification program to demonstrate theability of snubbers to maintain propersealing of the hydraulic fluid under theoperating and accident environments (includingsevere seismic disturbance followingtemperature and radiation aging,and a LOCA environment) should beconducted.A6.3.3.5.2 Conclusion. We judge anchors,restraints, and hangers to beadequate on the basis of experience andnormal plant operation. As indicated insection A6.3.2.5, in our judgment, thehydraulic snubbers are probably adequatefor the period of time between refuelingoutages.A6. 3.3.6 Instrumentation.The instrumentation and logic circuitswhich determine that a LOCA event hasoccurred and which then initiate functioningof the LHSIS are part of theplant protection system. None of theinstrumentation within the LHSIS appearsto be vital to the successful operationof the system during the injectionphase, but rather it serves to informthe operator (via indicating lights andgauges in the control room) that thesystem is performing properly.Two highly important instrumentationitems are the level sensors and thelevel alarms on the Refueling WaterStorage Tank (RWST); these are essentialto inform the operator when to changethe valve alignment to change from theinjection phase to the recirculationphase. Because the operator can obtaincomparable information from othersources (i.e., the clock, which advisesthe duration of the injection phase andhence approximately how much water couldhave been withdrawn from the RWST, andthe containment sump level indicators,SX-54

which advise whether there is sufficientwater there to supply the LHSIS), failureof the RWST level instrumentationdoes not automatically lead to a failureof the LHSIS. It must be recognized,however, that failure of this levelinstrumentation could result in maloperationof the system unless theoperator is alert to the situation andtakes the proper action based upon thesecondary instrumentation cited above.The RWST level instrumentation sensorsare located outside the plant and hencecannot be exposed to LOCA conditions.The control room instruments are withinthe plant, but are in an area protectedfrom adverse environmental conditions.Both portions- of the instrumentationsystem are potentially vulnerable todamage as a result of a severe seismicdisturbance, however. Because the PWRplant was constructed before a formalqualification program became an explicitAEC licensing requirement, we have notbeen able to obtain documentation ofsuch qualification in order to make anevaluation of design adequacy.A6.3.3.7Commentary and Conclusion.Because of the absence of evidence thatthe RWST level instrumentation sensorsand control room instruments will functionproperly following a severe seismicevent, we cannot make a determination ofdesign adequacy.A6.3.4HIGH HEAD SAFETY INJECTIONSYSTEMS (HHSIS)There are two independent HHSIS at thisplant: (1) the three 150-gpm chargingpumps (which are normally used to providemakeup flow to the RCS) andassociated valves and piping to enablethese pumps to be used to inject boratedwater from the boron injection tank andthe refueling water storage tank intothe RCS loop piping and (2) threeaccumulators and associated valves andpiping which can supply their containedborated water into the RCS loop pipingcold legs.The safety injection charging pumps arelocated in the auxiliary building, andthe accumulators are mounted on thefloor of the containment. The accumulatorsare passive safety elements, prepressurizedwith nitrogen. Coolantflows from these storage tanks as soonas RCS pressure falls below accumulatorpressure (650 psig). Our review focusedon (1) the main piping nozzles of theaccumulators, (2) the piping connectionsto the RCS, and (3) the charging pumps.A6.3.4.1Accumulator Tank NozzleFigure X A-16 shows the details of theaccumulator piping nozzle connection tothe hemispherical head of the accumulatortank. The ratio of nozzle to shellis 8.25/69 = 0.12. The DBE pipe loadson the nozzle were determined by theA/E's piping flexibility analysis (seeTable X A-20).The FSAR indicates that the tank isclassified as Class C (ASME Code SectionIII), which requires that pipe connectionsto the vessel shall "not overstressthe vessel wall" without settingspecific limits; however, stresses inthe nozzle region of the tank hemisphericalhead would currently be evaluatedin accordance with ASME Code Case No.1607, "(Special Ruling) Stress Criteriafor Section III, Class 2 and 3 VesselsSubject to Upset, Emergency and FaultedOperating Conditions" - Approved byCouncil November 5, 1973. Load conditionand stress limits for this codecase are given in Table X A-21.The stresses in the vessel wall due toDBE pipe loads for System No. 2 weredetermined on the basis of a Bijlaardanalysis. These stresses were combinedwith the membrane stresses due tointernal pressure and are compared withthe stress limits for a faultedcondition as follows:Stress (psi)Membrane = 19,150Membrane+ bending = 32,450Stress Limit (psi)2.0 S = 34,5802.4 S = 41,496A6.3.4.1.1 Commentary. In the event ofa large LOCA, the accumulator tank providesan initial high flow rate ofinjection water to the cold legs of thereactor via a check valve. This issupplemented later by water from theinjection pumps. Since the accumulatortank is not called upon to provide itsfunction during any normal reactoroperation, Safety Guide 26 (Ref. 18)would currently require it to bedesigned, constructed, and stamped inaccordance with the rules of SubsectionB. This option, for example, wouldrequire more stringent weld inspectionprocedures than required for Class Cvessels.X-55

under conditions that simulate pumpsupport, as installed. Because of itsunconventional support seismic testingmay be particularly significant for thisunit. However, there is no evidence ofseismic qualification tests.A6.3.3.2.2 Conclusion. The ability ofthe pump to function after pump casingdeflections of 1.15 has been demonstratedby static analyses. Actual seismictests have not been made; consequently,design adequacy has not been rigorouslydemonstrated. However, results of theanalysis undertaken enhance the probabilitythat the pump would qualify ifseismic testing were performed. Thereforein our judgment, for purposes ofthis study only, the pump and drivedesign may be considered adequate,although design criteria are not completelysatisfied.A6.3.3.3Valves.The valves on the suction side of theLHSIS are 12-in. motor-operated gatevalves with body rated at 150 lb/ANSI-B16.5. The disk is guided duringclosing and opening.The drawings (Ref. 14) show that thedesign of the valves is compact; theirnatural frequency is estimated to begreater than 10 cps.Some of the valves have an extended 11/2-in.-diameter stem, approximately 40ft between the disk and the motor operator.The stem is protected by a 2 1/2-in. pipe sleeve.A6.3.3.3.1 Commentary. In the absenceof a formal qualification program wecould not make a rigorous assessment ofthe design adequacy of the valves. Onthe basis of our experience in performingstress analyses of similar valves,we conclude that there is a goodprobability that the valves willfunction properly under an accidentsituation.We examined extended stem details.There is a 0.11-in. radial gap betweenthe OD of the coupling that connects thestem sections (approximatey 7 ft. long)and the ID of the 2 1/2-in, pipe sleeve.Assuming that the stem ends at the valveand the motor operator are fixed, weestimate that a side load of 0.1 lb/ftwould close the gap at mid-height andthe associated end moments would be13.33 ft-lb. In our judgment this endmoment can readily be reacted by thevalve bonnet.A6.3.3.3.2 Conclusion. We concludethat the design of the valve isadequate.A6.3.3.4Valve Motor Operators.All active valves in the LHSIS arelocated outside the containment structure;hence the motor operators forthese valves cannot be exposed to thehigh radiation level, high-pressuresteam, and chemical-spray environmentthat characterizes the LOCA conditionswithin the containment. The VMO werenevertheless subjected to a rigorousqualification testing program, whichconsisted of aging, nuclear radiation,simulated seismic conditions, andsimulated LOCA exposures. The simulatedseismic test (Ref. 15) consisted of twoparts: (1) a resonance scan between 5and 35 cps (acceleration level notstated) and (2) five 2-min vibrationexposures at 3g acceleration, 35 cps,with a 1-min pause (no vibration)between individual exposures. No resonanceswere "observed" during part 1.Tests to demonstrate the ability of VMOto function after exposure to high temperatures,pressure, and radiation wereconducted in two groups (Ref.15). Inthe first group, a VMO was aged in anoven and then exposed for 9 hr to bothhigh temperature and high pressures(300*F/90 psig peak); these conditionsexceed the peak environmental conditionswithin the containment. The VMO wascycled every 30-min throughout theexposure under simulated valve operatingloads. In the second group, the motorfrom the valve operators was exposed tomore than 150 megarads of gamma radiationin approximately 1 month. Thisexposure is more than the VMO wouldexperience after 1 year following a LOCAif it were installed within thecuntainment.A6.3.3.4.1 Commentary. From the informationprovided to us about the seismictest program, it appears that theprogram was typical of those beingconducted a few years ago. Certainaspects of the test do not meet currentseismic test requirements: the lowestfrequency is rather high; the method ofdetermining resonances is not stated(there is a large number of individualpieces in the unit, inside the case, andit is doubtful that resonances of thesewould be detected by sensors on theoutside of the case); and a bi-axialexcitation was not used. Nevertheless,from the test conducted and from ourexperience in seismic qualification, wewould expect that the unit would performIX-56

We note that the stresses in the vesselwall due to pipe loads on the nozzlehave been computed for System No. 2.Table X A-20 shows that the DBE pipeloads for System No. 3 are significantlyhigher.The supplier of the accumulator tankperformed an additional analysis forSystem No. 3 (after we indicated thathigher pipe loads should be used forthis system) with the following result:Membrane StressPressure (psi)Due to radial load 880Due to moment load 9,530Subtotal 10,410Due to internalpressure 16,780Total 27,190In the analysis for the System 3 nozzleall bending at the nozzle is classifiedas secondary bending (as permitted byTable NB-3217-1 of Section III of theASME Code), thus 27,190 psi is also thePM + PB stress for this nozzle.This classification for bending stresswas not invoked in the analysis of theSystem No. 2 nozzle; instead, as conservatism,all bending at the shell-nozzlejunction was there taken as primary. Itresults between nozzles are to becompared on a consistent basis thisconservatism may be removed; thenPL + PB for System No. 2 becomes 19,150psi.Additional information, data and resultsfor the accumulator head at the nozzlejunction was furnished us in Ref. 19 andwas carefully reviewed.A6.3.4.1.2 Conclusion. Based on clarificationand data provided by Ref. 19 itis our judgment that the analysis nowmeets the requirements of the ASME Code,Section III (including code case 1607)for Class 3 vessels.A6.3.4.2Accumulator Piping Connectionto RCS.The 12-in. accumulator piping connectsto the 27 1/2-in, piping cold leg of thereactor via a nozzle. Pipe loads on thenozzle were determined by the A/E. SystemNo. 2 was found to be the mostseverely loaded and therefore wasanalyzed. The NSSS vendor performed ananalysis to determine and evaluatestress at the nozzle juncture by twomethods as follows:a. Using the rules set forth in paragraph119.6.4 of the B31.1 Code, butwith the stress indices from theNuclear Piping Code (B31.7).b. Using the Bijlaard method as outlinedin the Welding Research CouncilBulletin No. 107 to determinestresses resulting from pipe loadscombined with the B31.1 Code todetermine stresses due to internalpressure and loop loadings.The resulting stresses were then comparedwith the allowables in accordancewith B31.1. The results are given inTable X A-22.A6.3.4.2.1 Commentary. The two methodsgive results that are in reasonableagreement for the earthquake and pressureloads, although thermal-expansionstress results do not agree as well.The supplier asserts that the stressesdue to the seismic response of the NSSSloop are included in the results.A6.3.4.2.2 Conclusion. We concludethat the nozzle connection of theaccumulator piping to the reactor coldleg satisfies the design criteria and isadequate.A6.3.4.3Charging Pumps and Drives.All three of the charging pump and driveunits are installed in the safeguardarea outside containment.The charging pumps and drives arecompact and are connected via horizontalshafts. The supplier has computed thefundamental frequency to 55 cps and hasdetermined that the maximum seismicstress is 1229 psi.A6.3.4.4Commentary and Conclusion.The charging pumps and drives areoutside containment; LOCA environmentsneed not be considered therefore.The seismic analysis results appear tobe reasonable and we conclude that theunits satisfy design criteria and areadequate.A6.3.5 CONTAINMENT RECIRCULATIONSPRAY SYSTEM (CRSS)This system recirculates water from thecontainment sumps through heat exchangersto spray headers mounted high withinthe containment. The water thus circulatedremoves heat from the containment,aids in lowering the temperature andpressure within the containment, andX-57

maintains the containment at subatmosphericconditions once it is depressurized.There are four pumps andassociated drives in the CRSS; two arelocated outside the containment and twowithin it. All four pumps take suctionfrom the containment sump. The CRSSpumps are electric-motor driven fromeither normal or emergency powersources, and are the vertical centrifugaltype with a 3500-gpm flow rating.The pressure head rating for the insidecontainmentpumps is 230 ft; that forthe outside-containment pumps is 249 ft.The former have a throttle bushing seal,and the latter have a tandem mechanicalseal. The inside-containment units arerelatively short and have their motordrive closely coupled. The outsidecontainmentunits are mounted in a mannervery similar to that of the LHSISpumps, the pump being located approximately45 ft below the drive and connectedby a long shaft.A6.3.5.1Pump and Motor LocatedOutside Containment.The installation of the CRSS pump andmotor is similar to that of the LHSIS asdescribed in section A6.3.3.2. The CRSSmotor weighs 2000 lb, the pump 3107 lb,and the interconnecting piping 3893 lb.The A/E performed a dynamic analysiswhich indicates that the installationhas a natural frequency of 1.9cps inthe horizontal direction and 63 cps inthe vertical direction.The pump manufacturer was given the followingseismic load requirements:a. Must be able to operate during andafter experiencing1. 0.17g horizontal and2. 0.llg vertical.b. Must be able to function with noloss of structural integrity afterexperiencing1. 0.34g horizontal2. 0.22g verticalThe A/E's evaluation of the installationwas based on the frequency computationsand corresponding acceleration responsespectra values:a. 1.68g horizontal0.19g vertical Ib. 2.48g horizontal0.3g verticalOBE$DBEThe A/E's analysis indicates that theinstallation is adequate for the aboveloads. However, no evidence was presentedto indicate whether or not thepump manufacturer qualified the pumpoperation for the higher seismic loads.The motor manufacturer performed a staticanalysis for the motor and itsinstallation assuming that the maximumhorizontal g load would be 1.45g andobtained maximum stresses of 12,170 psiin the motor mounting bolts; the allowablestress is 13,850 psi. A comprehensivequalification test of the motor wasconducted and is discussed in the nextsubsection.A6.3.5.1.1 Commentary. The motormounting bolts would be stressed to17,848 psi under the higher seismic loadof 2.48g, which would exceed the allowable.In our judgment, however, themotor would vibrate in the higher modesof the installation so that itsparticipation in the lower mode would besmall.It is not clear, however, whether or notthe pump would be capable of continuedfunctioning under the higher seismicloads.The basis (Refs. 11,12 and 13) for establishingthe radiation resistance ofthe pump internals leaves some questionsunanswered, but it shows that the pumpinternals probably have adequate radiationresistance.A6.3.5.1.2 Conclusion. The seismic designof the motor and motor mountingappear to be adequate.There is insufficient information tojudge the ability of the pump to continuefunctioning after exposure to seismicenvironments. Design adequacy of thepump has not been demonstrated.A6.3.5.2Motor Drives.All four motors are rated at 300 hp, 460V, 3-phase, 60 cps. They are fully enclosed,sealed motors with encapsulatedinsulation. The inside containmentunits are located at approximately elevation-20 ft; the outside containmentunits are at elevation +16 ft.A topical report (Ref. 20) describes acomprehensive qualification test programconsisting of four parts conducted onthe same motor in the following sequence(l)thermal aging, (2) seismictest, (3) nuclear-radiation exposure,and (4) steam/chemical-spray exposure.It should be noted that only the insider!X-58

containment units are subject tosteam/chemical-spray and high-levelnuclear-radiation exposures.Thermal aging consisted of exposing thestator to an average temperature of203 0 C for 155 hr. The report statesthat the aging exposure is based on IEEE275 and IEEE 384, Sec. 4.3.1, but supportingdata are not presented. Theaging exposure is reported to beequivalent to 18,000 hr at 130*C, themaximum operating temperature under"normal" (non-accident) conditions,which is equivalent to 450 hr per yearfor 40 years. (In the plant, the motorsare not expected to be operated morethan 15 hr per year.)The motor was exposed to 2 x 108 rads ofgamma radiation at a maximum rate of 0.5megarad per hour. This dose is the integrateddose that the motor could beexpected to receive during 40 years ofnormal life plus the exposure for a yearfollowing a LOCA.The seismic qualification test consistedof vibrating the motor sequentiallyalong three mutually perpendicular axes(one vertical and two horizontal) at aninput acceleration amplitude of 0.4gwithin the frequency range 4 to 70 cps.Because the motors in the plant are locatedbelow grade, an analysis performedby the A/E shows that qualification canbe demonstrated by vibration at anacceleration amplitude of twice themaximum ground acceleration level, whichis 0.15g.The steam/chemical-spray exposure includedfive cyclic transient exposuresto 45 psig/270 0 F, with dwells of 30 minat these conditions, followed by 7 daysat 0 psig/150 0 F. Except for briefinterruptions for electrical measurementsand to check the ability of themotor to restart, the motor was operatedat full load during the high temperatureportion of each of the five cycles andthroughout the seven-day exposure at150 0 F. The A/E's qualification testspecification states that the motor isrequired to start and operate at ratedload for 30 min during each transientand to be operable for a minimum of 24hr at 150 0 F + 5 0 F and 14.7 to 10.0 psia.A6.3.5.2.1 Commentary. The gamma-radiationdose used in the qualificationtest was more than adequate; and, sinceit preceded the seismic test and thesteam/chemical-spray exposure, the testwas conducted in a conservative manner.During the steam/chemical-spray exposuretest, the winding temperature neverexceeded the rating, 180 0 C (356 0 F), ofthe Class H insulation used in themotor. While present qualification requirementscall for testing at a temperatureexceeding the expected servicetemperature by a margin of 15 0 F, thefact that five transients were conductedwould appear, to provide a margin thattends to compensate for the lack of atemperature margin.A6.3.5.2.2 Conclusion. The qualificationtest program clearly demonstratesthat the motor satisfies the design criteriaand thus is adequate.A6.3.6ON-SITE ELECTRIC POWER SYSTEMSThe electrical power systems at the PWRplant are designed to provide a diversityof dependable power sources which arephysically isolated so that any failureaffecting one source of supply will notpropagate to other sources.The plant is connected to six separateoff-site substations with 230-kV transmissionlines and to two substationswith 500-kV transmission lines. In theevent of total loss of power from offsitesources, auxiliary power is suppliedfrom diesel generators locatedwithin the plant. Station batteries areprovided as a reliable source of controlpower for specific engineered safeguardsand for other functions required whena-c power is not available.A6.3.6.1Diesel Generator Building.The diesel generators are housed in cubiclesthat are part of the servicebuilding. These cubicles are protectedfrom tornado effects as described insection A6.2.1. A discussion of and anevaluation of the housing's resistanceto tornado effects follows:A6.3.6.1.1 Resistancetration. Three typesused to protect thesystem:to Missile Peneofbarriers arediesel generatora. Monolithic concrete walls - 2 ftthick, reinforced in both faces withNo. 11 deformed steel bars on 10-in.centers.b. Precast reinforcedby 2-ft by 12.5-ftbuild up "doors."c. Six feet of compacted soil.concrete - 2-ftbeams used toNo penetration calculations were madeavailable for review. According to theA/E, tests on the 2-ft-thick concretesection are currently under way.X-59

A6.3.6.l.l.l Concrete Walls. On thebasis of experience with other structuresand different missiles, we concludedthat neither of the two missileswill penetrate the concrete wall. (Calculationsindicate a penetration of only1 in. for the utility pole.) Penetrationcalculations for the auto missileare not considered necessary since theimpact area is so large.A6.3.6.1.1.2 Precast Beams Used for"Door." same comments as for concretewalls.A6.3.6.1.1.3 Compacted Soil. On thebasis of the work of Kornhauser, a penetrationof about 20 ft of sand would beexpected for the utility pole. Thepenetration in the compacted soil overthe fuel storage tanks could be the sameor less depending on the colloidalcontent of the particular material andthe degree of compaction achieved.Under the stated criteria it appearsthat missile contact with the fuel tankscould occur. Details of the fuel linelayout and manifold were not madeavailable; therefore, we could not makea final estimate of the damage potential.However, it should be noted that6 ft of compacted soil is a significantcover, and most missiles would thereforeprobably not be able to do any damage.A6.3.6.1.2 Overall Response Calculations.No overall response calculationshad been made by the A/E.A6.3.6.1.2.1 Monolithic Concrete Wallsand Roof. No independent overall responsecalculations were made owing totime limitations. However, the similarityof this structure to others forwhich calculations have been made (butwith different missiles) leads us tobelieve that the structure is adequate.A6.3.6.1.2.2 Precast Reinforced ConcreteBeams. Unlike the walls discussedabove, the beams appear to have lesspotential for withstanding the effect ofthe utility pole impact because thistype of beam cannot distribute theimpact load to a wider portion of thestructure (as occurs in a slab) untilthe load reaches the door jamb. In addition,only a single l-in.-diameterbolt at each end of the beam is availableto resist rebound forces afterimpact. However, for the missiles specifiedthe door beams are considered adequate.A6.3.6.1.3 Wind Loads. In spite of thedesign criterion of a 360-mph peak windvelocity, a velocity of 300 mph was usedto produce a dynamic pressure of 230psf. The wall openings were computed at23.6% of the building envelope area,which led to the conclusion that the 3-psi differential pressure in 3 sec wouldnot lead to any significant loading.The dynamic pressure of 230 psf was modified by the pressure coefficients asrecommended in ASCE Paper No. 3269,i.e., 0.9 on the windward side and -0.4on the leeward side, before being appliedto the walls. For the roof load afactor of -1.4 was used. This resultsin a 300-psf uplift force, which, however,is completely neutralized by thedead weight of the roof.we did not examinedoor structure,believes this loadcal.windbutwouldloads on thethe reviewernot be criti-One important point noted is that asteel brace from the turbine building isanchored on the top of one of the internaldividing walls of the emergencygenerator building. The brace introducesa load, which has horizontal aswell as vertical components, at thispoint of the diesel generator building.Calculations based on the design windloading for the turbine building, whichis lower than the tornado loading, weremade for the braced load. This wasjustified on the basis that blowoutpanels, etc., will partially relieve thewind loading on the turbine building.older calculations based on a differentpreliminary design (braces of a differentdesign) were made using the ultimateload-carrying capability of the brace asthe load applied to the diesel gene~ratorbuilding. New calculations made by theA/E at our request indicate that the newanchor design can take the ultimateload-carrying capacity of the brace withsome plastic straining but no rupture.No calculations were made to combinewind loads with missile loads.A6.3.6.1.4 Commentary. Only wind loadswere considered in any depth in the designcalculations reviewed, and thesewere based on 300-mph winds rather thanthe 360-mph winds called for in thespecifications. However, the reviewer 'sopinion is that wind loads will not becritical for this design. The followingitems, while not considered critical,are of some concern:a. Capability of the ciompactedcover to protect thie fuelagainst damage by a utilitymissile.b. The effects of fuelignition in the eventimpact.soiltankpolespillage andof an auto4X- 60

c . Analysis of the overall impact andwind-load response of the removableconcrete beam doors with an evaluationof the rebound forces for thedoor. (The A/E indicates therebound force will fail the doorpins.)Note: The A/E made additional calculat-ionswhich indicate that the door pinswould fail owing to the rebound forceafter missile impact. The A/E did notconsider this to be a problem, however,since the chances of an additional missileentering the diesel generator areawould be very small. The A/E furtherpoints out that there are obstructionsin front of the door that would preventthe direct impact of missiles; however,a 450 impact is possible. In ourjudgment, if the door rebound forces dueto missile impact were to fail the pinsand expose the diesel generator to thedriving rain and hail often associatedwith a tornado, this could cause aninterruption of emergency power.A6.3.6.1.5 Conclusion. We conclude thefollowing:a. The walls and roof satisfy the designcriteria and are adequate onthe basis of experience and judgment.b. The doors may fail to perform theirfunction; failure of the door pinsin particular is likely. Adequacyhas not been demonstrated.C. The soil cover for the fuel tanks isadequate in our judgment.A6.3.6.2Diesel GeneratorsEach diesel generator weighs 54 tons, isanchored to its base by ten 1 1/2-in. -diameter bolts, and is subject to DBEhorizontal and vertical spectral accelerations,as shown in Figs. X A-17 and XA-18. Note that the peak horizontalacceleration equal to 2.475g occursbetween 1.33 and 6.67 cps, with accelerationsof between 0.3 and 0.225g forfrequencies greater than 6.67 cps. Thevertical peak of 3.6g occurs between1.25 and 3.33 cps and then drops to 0.3gfor frequencies greater than 3.33 cps.The A/E performed calculations to checkthe adequacy of the anchor bolts andfound that they would not be strongenough to resist the peak accelerationsbut would be strong enough to resist atransverse horizontal acceleration of0.3g, conservatively neglecting the 0.7gnet restoring force in the vertical direction.The A/E computed a tensilestress of 3080 psi and a shear stress of2310 psi in the bolt, whose yieldstrength is 36,000 psi. The A/E justifiedthe use of the lower accelerationson the assumption that the fundamentalfrequency of the diesel generator isgreater than 6.67 cps.The supplier of the diesel generatorprovided test results which indicatethat the units frequently experiencelongitudinal (horizontal) accelerationsof lg and vertical accelerations of between0.25 and 0.5g during locomotiveoperation with "no detrimental resultson the operation of this equipment."When the locomotive is coupled with therail cars, the equipment experienceslongitudinal accelerations of 4g andhigher.A6.3.6.2.1 Commentary. The assumptionthat the fundamental frequency of thediesel generator is greater than 6.67cps is reasonable. Section A6.4.6.2,for example, indicates that the computedfundamental frequency for the BWR dieselgenerator is 9.63 cps, which isdominated by the motion of the turbochargers.The A/E's analysis of the anchor boltsassumes that they equally resist theshear load, a situation that could occuronly if all bolts are fitted. In alllikelihood, however, the shear load =0.3 x 108,000 = 32,400 lb would beresisted by friction between the unitand its base.It is estimated that without any preloadin the bolts, there would be an availablefriction force equal to 0.7 x108,000 x 0.3 = 22,680 lb., assuming acoefficient of friction of 0.3. Theadditional required shear resistancewould be provided by the bolt preloadforce.The information supplied by the dieselgenerator vendor is convincing evidencethat the unit can continue to operateafter experiencing 0.3g in the horizontallongitudinal direction together with0.3g in the vertical direction. Althoughno mention is made of transverseg-load measurements, the rocking loadsin normal diesel engine service couldgive rise to non-trivial transverse gloads. It is our judgment that on thebasis of in-service experience thediesel generator should be able towithstand a transverse acceleration of0.3g.A6.3,.6.2.2 Conclusion. Wethat the diesel generator isconcludeadequateX- 61

and c apable of withstanding the DBEseismic loads without loss of function.A6.3.6.3 Day Tanks.Fuel for each diesel generator comesfrom the day tank, which is housed inthe diesel generator cubicle. Each daytank, in turn, is supplied with fuelfrom the storage tanks, which are buriedin the ground outside the building. Theday tank is supported in a horizontalposition by four unbraced posts (approximately3 in. in diameter) as shownschematically in Fig. X A-19.No seismic analysis was available forour review. However, visual inspectionof the as-built unit during a site visitto the PWR plant revealed that one ofthe support posts was noticeably out ofplumb.A6.3.6.3.1 Comunentary. The seismic resistanceof the day tank to lateralloads with unbraced posts must be derivedfrom post bending stiffness. Thisstiffness, however, is normally lowerthan that obtained from supports withbraces acting as truss members.A6.3.6.3.2 Conclusion. A seismic analysisshould be performed to demonstratethe ability of the day tank supports 1 toresist seismic loads. Design adequacyhas not been demonstrated.A6.3.6.4Air Bottle Support.The air bottles provide compressed airfor starting the diesel generator. Thebottles are supported in the verticalposition by a back frame (Fig. X A-20).Each bottle is clamped to the back frameby an upper and a lower strap so thatthe vertical inertia force on the bottleis resisted by friction induced by theclamping action of the strap.No seismic analysis of the air bottlesupport was made available for our review.However, the installation wasvisually inspected during a site visitto the PWR plant.A6.3.6.4.1 Commentary. The ability ofthe air bottle support to resist verticalseismic loads depends on properinstallation to ensure that the strapsprovide sufficient clamping action.There is obviously sufficient clampingto support the weight (1g) of the botltisour understanding that the daytank support is being provided withbracing as a result of the review.tles. On the basis of inspection, thefundamental frequency is greater than3.33 cps; so the maximum expectedseismic g load in the vertical directionwould be 0.3g (see Fig. X A-18), whichrepresents a 30% increase over the deadweight.A6.3.6.4.2 Conclusion. Inment, the design of thesupport is adequate. Thisassumes proper installationcient clamping action of theA6.3.6.5Batteries and BatterySupport.our j udgairbottleconclusionwith suffistraps.Station batteries are provided as areliable source of control power forspecific engineered safeguards and forother functions required when a-c poweris not available (e.g., to operate circuitbreaker controls, turbine shutdownoil pumps, instrumentation, emergencylighting, and plant protection circuits).Two separate 125-V d-c systemsare installed, with an additional systemto supply essential services at theremote circulating-water intake structure.The batteries for each system arelocated in separate ventilated batteryrooms. Each 125-V battery consists of60 individual sealed cells with sealedcovers. Each battery is served by aseparate charger connected to "floatcharge" the battery. The chargers arecapable of carrying the normal d-csystem load and at the same timesupplying sufficient current to keep thebatteries in a fully charged condition.The chargers are supplied from separate480-V a-c motor control centers, eachconnected to an independent emergencya-c bus.The batteries were subjected to a seismicqualification test as follows:1.6g @ 11.50.7g @ 19.30.4g @ 28.70.7g @ 19.3Test No. 1(front to back)cps horizontalcps veticalTest No. 2(side to side)cps horizontalcps veticalSimultaneousSimultaneousAfter completion of each test the batterieswere inspected; there was novisible damage. In addition, a dischargetest subsequent to testing provedthat 100% capacity of the cell was stillavailable.X- 62

For this plant the batteries are supporteddirectly on massive, poured inplace, concrete blocks. Batteries aretied to the concrete block to preventsliding or overturning should a seismicevent occur.Commentary and ConclusionVisual inspection shows the batterysupport to be capable of resistingseismic loads. The testing of thebatteries demonstrates the adequacy ofthe batteries. We conclude that thebatteries and supports satisfy thedesign criteria and are adequate forseismic loads.A6.3.7ELECTRIC POWER DISTRIBUTIONSYSTEMSThe electric power distribution systemsconsist of transformers, switchgear andmotor control centers, emergency buses,containment penetrations, and associatedelectric power and control cables andtheir supporting cable trays.A6.3.7.1Electrical Cables andTerminations.Because safety-related electric powerand control cables are installed invirtually every area of the plant, bothwithin and outside the primary containment,they will be required to functionunder severe environmental conditionsshould an accident occur in the plant.Their ability to withstand such conditionswas established by subjectingsamples of the cables and cable splicesused in the plant to a qualificationtesting program (Ref. 21). In general,these programs consisted in thermal andnuclear radiation aging and steam/chemical-spray exposure. The steam/chemical test of electrical cablesinvolved a temperature in excess of00*F for a duration of approximately 3weeks. The scope of the tests, the testparameters, and the procedures followedwere typical of those used at the timethe test programs were conducted. Fireretardantconstruction was used in thecables to reduce the fire hazard.A6.3.7.1.1 Commentary. Because qualificationrequirements have become moredefinitive during the 4 years since thequalification tests were performed, mostof the tests would probably not beconsidered adequate by a strict interpretationof today's qualificationstandards (Ref. 22). For example, notall the cables were energized during thefull period of the test, and the testduration did not exceed the post-LOCAperiod during which some of the cablesshould be functional. The fact that thecables did perform adequately for 3weeks confirms that their design isadequate for the most crucial periodfollowing a LOCA. A cable failure at alater time, should one occur, may causea relatively minor operational problembut is not expected to result in asafety problem.Because the cables are not structuralmembers and are presumably properly supportedby cable trays, seismic qualificationis not explicitly required. Inprinciple, there is concern that cableinsulation may become embrittled as aresult of aging (exposures to temperatureand nuclear radiation during theirinstalled life) and that the insulationmay therefore be subject to failure bycracking during a severe seismic event.The current qualification philosophy isto cover this possibility by requiring abend test (wrapping the cable specimensbeing tested around a mandrel) after thesimulated LOCA exposure. The qualificationtest that was performed evidentlydid not include this aspect of thedesign requirements. This omission isnot regarded as serious, however,because the nuclear-radiation doselevels in the plant are not high enoughto produce embrittlement of the insulatingmaterials used in cables installedin safety-related systems.A6.3.7.1.2 Conclusion. Although certainaspects of the qualification testprogram were not as comprehensive ascurrent requirements, we neverthelessbelieve that the program meets the originaldesign criteria and does providereasonable assurance that the electricalcables and terminations will performadequately during and following designbasis events at the plant. This conclusionis tempered by the reservationthat, because the state of knowledgewith regard to long-term thermal,humidity, and nuclear-radiation agingeffects is somewhate limited, periodicsurveillance testing of the installedcables should be conducted.A6.3.7.2Electrical Containment Penetrationsand Connectors.The electrical containment penetrationsare components that provide pressurewithstanding,leak-tight, electricallyinsulated paths through the wall of theprimary containment structure for transmittingelectrical signals and electricpower to motors and solenoids. Generallyspeaking, current specifications requirethat penetration assemblies employtwo seals in series to permit monitoringand testing in situ. Both seals mustfail to permit leakage from containment.X-63

The design and qualifications of theelectrical mating connectors are basedupon the requirements of military specificationnumber MIL-C-5015. Connectordesign is such that silastic componentsare provided in the connector to feedthrough the interface. The A/E believesthat this specification is equivalent tospecifications for nuclear power plantapplication, and that this type ofinterface has been proven adequate tomeet environmental requirement. Additionalcapability to withstand elevatedtemperatures is provided in the siliconematerial used for the sealing members.Tests conducted by the manufacturer consistedof the following (Ref. 3):"Connectors installed in theflanges normally operate at ambientconditions of 105'F and 9.75 psiaand were tested for leak rate andtagged for integrity before shipmentto the job. A test facilitywas set up by the manufacturersuitable for 50 psig with provisionsfor thermocycling from 32°Fto 300 0 F. A thermocycle run of atleast three cycles was made on oneof each type flange. A timeinterval of 30 minutes was allowedbetween the thermocycies. The leakrate test after thermocycling wasmade at 50 psig and 3001F. Eachcompleted flange hag a leak rate ofless than 1 x 0- cc/sec perassembled flange. All flanges wereleak tested at 50 psig and 300 0 F.Helium gas was used in the testfacility."Reference 3 also reports that for thetriaxial cable penetrations a more detailedprocedure for the thermocycletest was followed in shop test, asfollows:"The type sample consisted of acontainment side flange disk withhermetic assemblies welded inplace. A thermocouple was installedto monitor disk temperature.The disk was stabilized at32 0 F and then placed in an ovenheated previously to 280 0 F. Onentrance of the disk, the oventemperature was reduced, straightline, to 150°F over a 60 minuteperiod. The disk was removed andcooled to 100*F, while the oven wasreheated to 2801F. The disk wasthen returned to the oven and theoven temperature reduced to 150 0 Fas before. The highest metal temperaturereached during this cyclewas recorded and a 50 psig heliumleak test was conducted at thismetal temperature for allthis type."disks ofEach penetration assembly, without externalcable mating connectors, has beentested in the factory to demonstrateinsulation resistance of at least 1000megohms at 1000 V d-c. In addition,each penetration has passed a highpotentialtest. After installation,each penetration with external cablesconnected was tested at 1000 V d-c for 5min.We understand that an environmental testwhich would simulate pressure, temperature,moisture, and radiation associatedwith a LOCA was scheduled, but theresults have not been made available toUS.Both an analysis and a test were performedto qualify the electricalcontainment penetrations for seismicenvironments. The fundamental frequencywas found to be 3094 cps. Maximumcomputed seismic stress was found to be56 psi in the mounting bolt. A shaketest was performed over a frequencyrange up to 10 cps. The accelerationwas 5g at 10 cps. No visible damage andno loss of electrical function wasobserved after the test.ConclusionFrom our review of the qualificationprogram for the electrical containmentpenetrations and connectors, we concludethat their design is adequate, assumingthat the results of the LOCA environmenttest are positive.A6.3.7.3Cable Trays.The cable trays, which provide supportfor electrical cables, are covered bycorrugated sheets and consist of twoside rails connected by rungs, the centersof which are typically spaced 9 in.apart. The trays are supported fromabove approximately every 8 ft. Noseismic analyses of the trays wereprovided for our review. The trayswere, however, visually inspected duringa site visit to the PWR plant. Thetrays appeared to be well constructedand supported. Natural frequencies areestimated to be above the peak spectralacceleration range.Commentary and ConclusionFrom our inspection of the trays, weconclude that their design is adequate.X-64

A6.3.7.4A-C and D-C Switchgear.No seismic or environmental qualif icationinformation has been made availablefor review and evaluation. We cannottherefore make an assessment of designadequacy of those components.A6.3. 8REACTOR AND ENGINEERED SAFE-GUARDS PROTECTION SYSTEMSSENSORS AND LOGIC CABINETS'The reactor protection system monitorsnuclear, thernial, and hydraulic parameterassociated with the reactor core,determines whether conditions are suchthat the reactor should be shut downimmediately, and, if so, initiates ashutdown. The engineered safeguards instrumentationmonitors parameters todetect failures in the reactor coolantsystem and to initiate containment isolationand engineered safeguards equipmentoperation.The only portions of the plant protectionsystems that were considered in thedesign adequacy review were:a. Pressure and differential pressuretransmitters.b. Nuclear instrumentation system cabinets.c. Radiation monitoring system cabinets.d. Signal conditioning equipment (modulesand racks) for monitoringpressurizer level and pressure; containmentpressure; reactor coolantflow, pressure, and temperature; andsteam-generator water level andfeedwater pressure.e. Safeguards actuation logic racks.Instrument piping is connected to theplant components and routed to instrumentationracks. The instruments thatmonitor reactor, reactor coolant system,and steam-generator process parametersare located within the containmentstructure; the others are located externalto the containment. These rackscontain the pressure transmitters anddifferential pressure transmitters whichtransform hydraulic pressure signalinputs into electrical signals. Theelectrical signals are conveyed byelectrical control cables to instrumentsand to cabinets containing signalconditioning equipment and componentsthat provide the 2-out-of-4 logicfunctions. other electrical cables thencarry the processed signals, whichinitiate reactor shutdown, containingisolation, and ECCS initiation, to thecontrol rod drive system and to thevarious motor control centers thatinitiate valve movements and motor startUPS. Additional signals are generatedin the various systems and carried byelectrical cables to the control room,where gauges and lights are displayed toprovide to the plant operator suchinformation as motor current, systemflow and pump discharge pressure, andvalve positions. The control room alsocontains switches to permit the manualoperation of motors and valves shouldtheir automatic functioning fail tooccur.A6. 3. 8.1 Environmental Qualification.As noted above, some of the pressure anddifferential pressure transmitters arelocated within the containment andtherefore can be subjected to extremeenvironmental conditions should a LOCAoccur. Environmental qualfication testshave been conducted on these components(Ref. 15).A6.3.8.2Seismic Qualification.Evaluation of the seismic qualificationof the PWR protection system was basedprimarily on a review of References 23through 25. The rationale of the testprocedure is presented in References 26and 27. The seismic tests of electricaland control equipment described in thesereports cover essentially three seriesof tests, one for low and two for highseismic disturbances. The more severeset of test criteria (high seismic) weredeveloped to simulate the effect of disturbancesat the base of equipmentlocated on floors of plants with a DBEhaving horizontal ground accelerationsin the range of 0.2 to 0.4g and higher.Since it was found that mechanical deformations(e.g., slippage at boltedjoints) caused the equipment resonantfrequencies at high input accelerationsto be different from those observedduring resonance scans at low inputlevels, the tests were conducted for theýmost part at discrete frequenciesseparated by intervals that wereintended to be small enough so that anyequipment resonance in the range between1 and .35 Hz would be accounted for.Tests were conducted sequentially alongthree orthogonal axes, two horizontaland one vertical. The amplitude of thehorizontal, sine-beat input motion atthe base of the equipment tested was afunction of frequency, having a plateauof 1.5g between 5 and 10 Hz and droppingto 0.5g at 1 Hz and 25 Hz; the amplitudeof vertical input motion was two-thirdsof the horizontal input motion.X-65

The control board was designed to withstandearthquake conditions. An analysiswas performed to verify the adequacyof the seismic design, and this wascomplemented by tests performed on sectionsof the main control board withsome safety related instruments in position.A6.3.8.3 Commentary.A6 .3. 8.3.1 Environmental Qualification.A review of the information presented inRef. 15 discloses that two pressuretransmitters and one differential pressuretransmitter were subjected tosimulated LOCA environments that weremore severe than the worst environmentexpected within the plant under LOCAconditions. The durations of the exposureswere 2 to 3 hr, which is very muchlonger than the time (a few minutes) thesensors that detect abnormal conditionsand initiate plant shutdown, etc., mustfunction. On the other hand, it ishighly desirable that instrumentation beavailable to monitor the conditionswithin the plant for a long period(several weeks) following a LOCA.Although the equipment samples that weresubjected to the environmental qualificationtests were not first subjected toan aging exposure, which is a currentrequirement, there are several obviatingfactors. According to the manufacturer,the pressure transmitters are parts ofrepairable systems which are required tobe tested at one-month intervals duringoperation with the reactor at power.Any degradation of performance resultingfrom normal aging would be detectedduring such tests, and the componentwould be replaced. Also, to the extentthat the duration of the simulated LOCAenvironment exceeded the minimum requiredfor qualification, it may beregarded as partly compensating for thelack of prior aging. We have not seenany qualification information that wouldpermit an evaluation of the capabilityof these components to function properlyunder the following conditions:a. Envelopment in a steam cloud followinga severe seismic event.b. Heavy wetting by water spray followinga severe seismic event.A6.3.8.3.2 Seismic Qualification. Theseismic qualification tests were conductedin accordance with proceduresthat were considered acceptable at thetime the tests were conducted. However,the information available was not adequateto permit a determination of theadequacy of the tests with respect tocurrent seismic qualification requirements.In particular, it presently mustbe demonstrated that multi-frequencytesting, involving simultaneous vibrationin horizontal and vertical directions,is not essential for the specificequipment tested. The blanket comparisonof earlier test procedures with newrequirements in Reference 27 does notaccomplish this.Other reservations regarding seismicqualification include the following:The information available did not permitus to determine that the equipmenttested was the same as, or adequatelysimilar to, the equipment that wasinstalled in the plants. Also, in thosecases where equipment failed during atest and a modification was made, thetest was usually continued from thepoint of failure. Presumably, it wasassumed that the "improved" equipmentwould not fail during a repetition ofthat part of the test that had beencompleted before any change was made;while this is probably usually true,there may be cases in which the changemight cause an unforeseen degradation ofperformance under conditions notretested. Furthermore, in those caseswhere a component failure was judged tobe an " isolated incident," and thereforenot an indication that the test unit hadfailed, the rationale was questionable.A6.3.8.4Conclusion.The results of the extensive qualificationtests conducted on the protectionsystems sensor and logic cabinets leadsus to conclude that they would probablyfunction satisfactorily. However, becauseof the reservations expressed inthe commentary, the design criteria havenot been completely satisfied.A6.3.9INTAKE CANALThe intake canal, which channels waterfrom the reservoir to the plant, ishigher in elevation than the dischargecanal; so it provides a passive pressurehead that ensures a flow of coolingwater to safety systems.The intake canal is lined with concreteand is supported by earth below andaround the entire canal. No seismicanalysis was provided for review, butthe entire canal was visually inspectedfrom the pump house where water from thereservoir is pumped into the intakecanal to the screens at the plant.Commentary and Conclusionour visual inspection of the intakecanal indicates that there is no dangeraiX- 66

of the canal losing its inventory as aresult of the liner cracking during aseismic event, since the surroundingearth will prevent rapid loss of water.We conclude that the seismic designadequacy of the intake canal is ade-•quate.A6.4 DESIGN ADEQUACY OF BWRPLANT COMPONENTSA6.4.1 CONTAINMENT STRUCTUREThe containment system at the BWR plantconsists of a multi-barrier design with:a. A primary barrier consisting of thesteel primary containment with itspressure-suppression feature.b. A secondary barrier consisting ofthe reactor building with a systemto limit the ground level release ofairborne radioactive material fromthe secondary containment.Figure X A-21 shows a cross section ofthe containment structure. The containmentstructure was reviewed only fromthe viewpoint of mathematical modelingand the method of obtaining seismic inputsto equipment that is mounted on thecontainment structure.A6.4.1.1Primary Containment Structure.The primary containment is an enclosurefor the reactor vessel, the reactorcoolant recirculation system, and otherbranch connections of the reactor coolantsystem. Structurally, the containmentconsists of a drywell and apressure-suppression chamber connectedby eight large vent pipes. The drywellis a steel pressure vessel in the shapeof a light bulb, and the pressuresuppressionchamber is a torus-shapedsteel pressure vessel located below andencircling the drywell (see Fig. XA-21). The suppression chamber is halffilledwith water.The A/E indicated that preliminary studiesshowed that the drywell does notdynamically interact with the system;thus its mass was lumped with the reac-* tor building in the mathematical model.A seismic analysis of the pressure suppressionchamber (torus) was performedby the A/E under the assumption that thetorus is fully flooded with water. Thewater was assumed to behave like a frozenmass. The fundamental frequencyunder these assumptions was found to be4.75 cps. The A/E found that the maximumdisplacement of the torus under theassumed conditions was 0.064 in.The torus was originally installed withbaffle plates, but these were removedafter installation.A6.4.1.1.1 Commentary. The assumptionof a fully flooded torus appears conservativefrom the viewpoint of loads onthe seismic restraints. The floodedcondition is not considered to be a realisticone. Normally, the torus ishalf-filled with water.Sloshing modes could be excited by seismicevents since the baffles in thetorus have been removed. The A/E hasrecently supplied us with informationwhich indicates that the natural frequencyof the suppression chamber poolis on the order of 0.1 cps. The groundresponse spectrum curve (5% damping) forthe BWR maximum credible earthquake(Fig. X A-7) shows a horizontal displacementof 4 in. at 0.1 cps, which isnot considered to be significant.A6.4.1.1.2 Conclusion. From our reviewof the information we have received fromthe A/E, we conclude that the modelingof the primary containment structure forproviding seismic inputs to componentsmounted on the primary containmentstructure is adequate.A6.4.1.2Reactor Building (SecondaryContainment).The building is a cast-in-place reinforcedconcrete structure from itsfoundation floor at elevation 91 ft 6in. to its refueling floor at elevation234 ft. Above this floor, the buildingsuperstructure consists of metal sidingand decking supported on a structuralsteel framework. The building isdesigned to be nominally 150 ft by 150ft below elevation 135 ft and is 150 ftby 120 ft above this level. The foundationof the building consists of a monolithicconcrete mat supported on soundrock. The mat also supports the primarycontainment and its internal components,including the reactor vessel pedestal.The exterior and some interior walls ofthe building above the foundation arecast-in-place concrete. Other interiorwalls are normal-weight concrete blockwall-s. The floor slabs of the buildingsare of composite construction with castin-placeconcrete over structural steelbeams and metal deck. The thicknessesof walls and slabs were governed bystructural or nuclear-radiation shieldingrequirements.Each steel-framedcross-braced toearthquake forcessuperstructure iswithstand wind andand supports metalX-67

siding and metal deck for the roof. Theroof consists of built-up roofing overthe metal deck. The frame also supportsa runway for the 125-ton travelingbridge crane.The configuration of the reactor buildingand adjacent buildings is shown inFig. X A-21.The mathematical model used for determiningthe seismic loads on the reactorbuilding, together with the interactionwith the reactor vessel and sacrificialshield, is shown in Fig. X A-22. Thefirst three natural frequencies in thehorizontal directions were computed asshown in Table X A-23.Modal inertia forces were found from theground response spectra curves shown inFig. X A-6 (2% critical damping designearthquake) and Fig. X A-7 (5% criticaldamping MCE). The associated modalshears, moments, and deflections in eachof the building structural elements werethen computed. Next, the modal loads inthe building structural elements werecombined on an absolute sum basis.No dynamic amplification of the verticalground motion was assumed to occur inthe building because of its stiffness inthe vertical direction and because thebuilding is founded directly on rock.Floor response spectra curves were generatedby imposing 'a modified timehistory Taft earthquake at the base ofthe structure. The time history responseat various elevations was thenused to construct floor response spectrafor designing equipment. Figure X A-23shows a comparison between the responsespectrum for 2% critical damping, whichcorresponds to the modified Taft earthquake,and the design ground responsespectrum for the DE. Note that thedesign spectrum peaks at 3.8 cpsIwhereas the Taft spectrum peaks at 2.8cps. Also note that the Taft spectrumexceeds the design spectrum at all frequenciesabove 0.6 cps.A6 .4.*1.2.1 Commentary. The mathematicalmodel appears to be a reasonablerepresentation of the actual structure.The method of combining modal forces inthe structural elements is conservative.Figure X A-23 shows thatTaft earthquake used infloor response spectra isthe modifiedgenerating theconservative.A6.4.1.2.2 Conclusion. The method ofdetermining the seismic response of thereactor building is more than adequate,and the generated floor response spectrafor designing equipment may be used withconfidence.A6.4.1.3Containment PipingPenetrations.Two general types of pipe penetrationsare provided: (1) those which must accommodatethermal movement and (2) thosewhich experience relatively little thermalstress.The piping penetrations that accommodatethermal movement are the high-temperaturelines, such as the steam lines,feedwater lines, and other reactor auxiliarysystem lines. The drywell nozzlepasses through the concrete shield andis attached to a bellows expansion jointwhich, in turn, is attached to thepenetration adapter to form a containmentpressure boundary. The processline which passes through the penetration,is attached to the penetrationadapter and is free to move axially. Aguard pipe immediately surrounds theprocess line and is designed to protectthe bellows and containment boundaryshould the process pipe fail within thepenetration. Thermal insulation is installedin the annular space between theguard pipe and the process pipe.The bellows assembly accommodates thethermal expansion of the process pipeand drywell. The hot process pipe isanchored at the penetration adapter externalto the drywell and guided at theother end of the penetration to allowthermal movement of the pipe parallel tothe penetration. Two isolation valvesare provided, one outside the drywelland the other inside the drywell. Thesevalves are located as close to the drywellpenetration as practical.The bellows expansion joints are of twoplyconstruction and permit leak testingof these penetrations at pressures up tothe primary containment design pressure.The expansion joint assembly is designed,constructed, and tested in accordancewith the previously specifiedrequirements for the primary containmentand code case interpretations, includingCode Cases 1177-5 and 1330-1. The bellowsare fabricated from stainlesssteel. Non-destructive tests of the assembliesinclude radiography and liquidpenetrant tests of welds and pneumaticpressure tests of bellows.Penetration adapters are one-piece forgingswith integral flues made of thesame material as the process pipe. Theadapters are designed, fabricated, andtested in accordance .with the previouslyAIX-68

specified requirements for thecontainment.primaryThe design of the penetrations takesinto account the simultaneous stressesassociated with normal thermal expansion,live and dead loads, seismicloads, and loads associated with a LOCAwithin the drywell. For all these conditions,including combinations of theseloads, the resultant stresses in thepipe and penetration components do notexceed the code allowable design limits.The design also takes into account thejet force loading resulting from thefailure of the steam piping in additionto the other loadings given. The resultantstresses in the pipe and penetrationfor this condition do not exceed90% of the yield stresses of the material.Cold piping and ventilation ducts arewelded directly to the drywell penetrations.Bellows and guard pipes are notnecessary since the thermal stresses aresmall and are accounted for in thedesign of the weld joints.The A/E performed design calculations todetermine stresses in the framework thatanchors the adapters to the reactorbuilding structure. The final design ofthe framework was then analyzed withcomputer techniques.Commentary and ConclusionThe calculations appear to be adequate.There is an inherent check between handcalculations and the computer results.We conclude that the piping penetrationssatisfy design criteria and are adequate.A6.4.1.4Suppression-Chamber-to-DrywellVacuum Breaker Valves.Twelve self-actuating valves connect thedrywell to the suppression chamber toprevent excessive negative pressure inthe drywell. These valves are locatedwithin the suppression chamber at anapproximate elevation of 116 ft on theeight large-diameter vent lines (four ofthese have two valves each, and theremaining four have one each). Thesevalves are exercised by auxiliary airactuators operable at local controlstations external to the containment.Figure X A-24 is a photograph of a valvesimilar to those installed within thesuppression chamber.It is important that the drywell-to-suppression-chambervacuum breaker valvesremain in a fully closed position untilthey are called upon to open to preventthe pressure in the suppression chamberfrom exceeding that in the drywell. Ifthe valves were not closed, a path wouldexist for steam resulting from a smallleak in the primary system (which doesnot entail a significant pressure buildupto help ensure that the valve is heldclosed) to pass directly into the vaporspace in the suppression chamber.At the time of the review, the vacuumbreaker design for the prototype BWRplant was under active review and modification.The tentative new designexisted on the A/E's drawing (Ref. 28)as marked up for proposed revisions.The original design is identical to thatof the breakers for an earlier BWRplant. Modifications have been effectedin the electrical circuitry that providesvalve-position signals in the controlroom and also in the auxiliaryvalve operator. Possible modificationto the valve plate was also under consideration,but no change in the valvebody design is contemplated.The original specification for the valvewas contained in the A/E's purchasespecification (Ref. 29) which did notimpose seismic requirements on the valvedesign. It has since been supplementedwith the A/E's general seismic specification(Ref. 30) which requires a dynamicseismic analysis for all Class Icomponents.The only analysis that the A/E had onhand was that initially provided for theBWR plant, primarily concerning boltstresses. The A/E has marked up thisdocument with comments on omissions andsuggested procedures. The A/E has requestedthat a more comprehensive analysisbe furnished by the vendor and thatthe analysis be updated to reflect thesupplementary seismic specification.A6.4.1.4.1 Commentary. The seismicanalysis of the design has not beenprovided; however, it is felt unlikelythat the valve body (a short, largediameterhollow cylinder) would be foundto be highly stressed by seismicloadings.Moreover, the A/E states that the newdesign (which has been installed) willavoid a feature of the original designthat might cause impairment of the valvefunction. The original design employs amicroswitch recessed into the rim of thevalve seat to initiate a signal indicatingthat the valve is in the closedposition. The operating button isX-69

depressed by the valve rim when thevalve is seated (like a door latch).However, a 9-lb force is needed todepress the switch. As a consequence,the valve plate must close with sufficientmomentum to cam the switch. Itwas observed that in the event of anearthquake the valve plate might bejiggled slightly open. The switch mightthen act as a detent to prevent fullclosure, defeating one of the principalpurposes of the valve.No environmental qualification testshave been performed. Although such atest would be desirable to provide assurancethat the position switches functionproperly, the absence of such materialsas electrical insulation in theportions of the valve which are essentialto its vital functions leads us tobelieve that it will function adequatelyin the event of either a small-leak orlarge-leak LOCA.A6.4.1.4.2 Conclusion. There is no apparentreason to believe that thesevalves will fail to function under postulatedseismic loads. The design criteriaappear to be satisfied on thebasis of engineering judgment.A6.4.1.5Missile Barriers InsideContainment.The recirculation line is provided withpipe-whip constraints but the main steamand feedwater lines are not. The A/Emade a thorough examination of the mainsteam and feedwater line configurations,postulating instantaneous circumferentialpipe breaks at butt welds andselecting regions on the inside surfaceof the drywell that would be vulnerableto impact by the broken pipe whippingabout a plastic hinge. These vulnerableregions of the drywell (which are quiteextensive) have been provided withreinforcing shielding missile barrierplates, as described on pages Q.4.9.2-1through Q.4.9.2-5 of the FSAR (Ref. 31).Figure X A-25, taken from Ref. 31, showsthe location and "thickness" of thereinforcing plates.The A/E's calculations indicate that theimpact energy, EI, to be absorbed by thebarrier is derived from the relationship:EI = ET -EHwhere ET is equal to the thrust load(1.2 x operating pressure x break areafor the main steam line) times the arclength of travel and EH is the energydissipated in forming the plastic hinge.EI is then compared with U, the energyrequired to penetrate a plate, using anempirical formula derived by StanfordResearch Institute (SRI):where:U = D a (0.344T2 + 0.00806WT)m uDm = width parameter of penetratingWauTmissile= height parameter ofpenetrating missile= ultimate strength of targetmaterial= thickness of target material.One part of the calculation indicatesthat a 3 1/2 in. plate thickness wouldnot be sufficient to withstand an impactenergy of 11.71 x 106 ft-lb but that theadditional thickness of the drywellplate (3/4 in. thick) would provide sufficientprotection. The installationdrawings indicate that the final designconsists of two plates, each 1 5/8 in.thick, which, along with the drywell3/4-in.-thick plate, gives a total of 4in. The A/E's calculations state thatthe empirical SRI formula for U may bemodified for multiple thickness barriersby replacing T with ZTi.A6.4.1.5.1 Commentary. The assumptionthat T can be replaced by ZTi in the SRIformula is incorrect since the relationshiphas a square term in T as well as alinear term. The energy required topenetrate multiple-plate barriers shouldbe calculated on the basis of ZUi, whereUi is the energy required to penetrateeach plate of thickness Ti. A roughcalculation using the above applicationof the formula indicates that the multiplebarriers could provide penetrationresistance to roughly 8.8 x 106 ft-lbimpact energy as opposed to the required11.71 x 106 ft-lb, implying a negativemargin of 33%.Note: The A/E has at our requestreviewed the original design and nowfinds that the pipe will impact themissile barriers at an angle of 410;this reduces the impact velocity byapproximately 25%. The A/E's revisedcalculations predict a 17% margin ofsafety. A reduction of impact velocityby 25% reduces the impact energy by 44%;however, the effective penetration areamay also be very significantly reduced,and the impact energy per unit area mayX-70

actually increase if the missile strikesat an angle.Additional analysis or penetration testdata for multiple plates is requiredbefore the design adequacy can beassessed. More refined calculations,which would account for the effectiveincrease in penetration area as eachsuccessive plate is penetrated, may show,that the design is adequate. Someadditional margin may be obtained in thecomputation of the energy dissipated atthe plastic hinge if strain hardening isconsidered.A6.4.1.5.2 Conclusion. In our judgmentthere is insufficient information toassess the adequacy of the missilebarriers. Design adequacy has not beendemonstrated.A6.4.1.6Paint Coatings WithinContainment.The internal surface of the containmentliner, internal concrete structures, andmetallic components are coated with specialpaint coatings. Aside from aestheticpurposes, coatings are used toprevent corrosion of metal surfaces andto seal the concrete. The coatings mustbe able to withstand normal and accidentenvironments without becoming separatedfrom the walls in any way, such as byruptured blisters, flaking, and cracking.Excessive delamination of coatingfilms could cause the clogging ofstrainers on the suction lines of thecore spray and safety injection systemsand interfere with their proper operation.The coating system was applied in twostages. In the first stage an inorganiczinc prime -coat was applied on theindividual shop-fabricated assemblagesby the steel fabricator. The finishphenolic coating was applied after thecontainment had been assembled and afterdamaged (e.g., scratched) areas andregions affected by welding had beentouched up.In March 1971, during a routine inspectionof the torus (suppression chamber)portion of the containment, failure ofthe finish coating was found in thebelow-the-waterline regions (loss ofadhesion over a significant area) andlocalized failures were found in theupper air-filled region. The followingrepair program was carried out in 1972:a. Below waterline: Sandwhiteapplyofzinc.blast tometal andone coatinorganicb. waterline level:c. Vapor zone(above waterline):Sand blast towhite metal in2-ft band andapply one coatof inorganiczinc and one topcoat of a phenoliccoating.Repair defectiveareas only.The top coating in the vapor zone wasretained because the condition of thecoating was essentially good. Coatingadhesion tests performed during the recoatinginvestigation substantiated thisdecision.During preoperational testing of theplant systems in May 1973, the torus wassubjected to a relatively severe dynamicsteam test (demineralized water heatedto 180 0 F, held at this temperature forapproximately 48 hours, followed by afast cooldown). Following this test,blister formation was observed in thevapor zone. The total area of phenoliccoating failure was estimated to coverless than 1% of the surface area. Theinorganic zinc base coat in failureareas was reported by the utility to bein excellent condition. In general, theblistering or delamination occurred inthe areas where a two-coat phenolicsystem existed because of unavoidableoverlapping during the repair work.This determination was made by measuringthe thickness of paint chips and byvisual examination of exposed areaswhere the phenolic finish coating stillexisted. The examination also showedthat the bonding between inorganic zincand the metal surface was still intact.In June 1973, those portions of thephenolic finish coat which exhibitedpoor adherence were removed. A testingprogram to determine the integrity ofthe remaining phenolic top coat was theninitiated. This program included steamdynamic tests simulating the conditionsexperienced during the preoperationaltest. The utility reported that theresults of inspections made after theconclusion of these tests clearly indicatedthat the inorganic zinc primer wasperforming its corrosion resistancefunction and that the vast majority ofthe remaining phenolic finish coat wasfirmly adhered and was satisfactory froma performance standpoint.In view of the foregoing, the toruscoating was then repaired a second time:X-71

a. Below waterline:b. Vapor zone:Sand blast towhite metal andapply one coat ofinorganic zinc.Repair defectiveareas in presentinorganic zinccoating.This procedure provided an inorganiczinc coating on the entire internal surfacesof the torus.A6.4.1.6.1 Commentary. No documentationpertaining to a qualification programwas provided to the reviewers, butwe were informed that the manufacturerof the coating materials and the A/E hadboth conducted extensive testing programswhich established that the coatingsystem can be expected to performadequately should a LOCA occur. Withinthe limits of our familiarity with thetests, this seems to be a valid conclusion.As is noted in section A6.3.1.3,however, it is extremely difficult todevise a qualification test for paintcoatings which has a significant statisticalbasis and is not challengeable onthe ground that the samples tested arenot accurately representative of thecoating actually applied in the field.The utility has also presented ananalysis which demonstrates that thereprobably will not be a safety problemeven if the phenolic top coat shouldfail during a LOCA. The point is madethat this material is thermosetting andhas a melting point of about 1100 0 F.Should it come loose in an accident, thematerial will harden, become morebrittle, and will not react with thedeionized water. Also, since thephenolic is heavier than water, it willsettle to the bottom of the torus.Plugging of strainers is not likelybecause of the configuration of thetorus (segmented by structural ringgirders) and because the core spray andHPCI strainers are connected to 16-in.pipe, the residual heat removal straineris connected to 24-in. pipe, and thestrainers are truncated cones with areasof 4.7 and 7 ft2 , respectively. Thelatter are mounted 1 ft above the torusinvert and have 1/8-in, holes. Further,the strainer flow area is conservativelydesigned with twice the area required topass the design flow at design pressuredrop.A6.4.1.6.2 Conclusion. The qualificationprograms conducted by the paintmanufacturer and for the A/E, in conjunctionwith the testing performed onthe actual coatings in the plant, providereasonable assurance that thecoating systems will not experiencelarge-scale failure (flaking and/ordelamination) should a LOCA occur. Inaddition, the construction features ofthe plant safety systems are such that,even if the containment coating shouldexperience a large-scale failure, it isextremely unlikely that the performanceof these systems will be significantlydegraded.The paint coatings within the containmentsatisfy the design criteria and areadequate.A6.4.2REACTOR COOLANT SYSTEMThe principal components of the reactorcoolant system considered in this studyincluded the reactor vessel and internals,the reactor vessel support, thetwo recirculation lines, the main steamlines and isolation valves, and thepressure relief system. The reactorvessel is supported by a cylindricalskirt that rests on a ring girder atop aconcrete and steel support pedestal.The pedestal is an integral part of thebuilding foundation (see Fig. X A-21).The reactor pressure relief systemconsists of 11 relief and 2 safetyvalves mounted on the main steam lineswithin the drywell. The safety valves(Fig. X A-26) are spring-loaded valvesset at 1230 psig. Each relief valve isself-actuating at set points rangingbetween 1080 and 1100 psig. The reliefvalves can also be actuated by remotecontrol.A6.4.2.1 Reactor Pressure Vessel andInternals.Design calculations prepared by the reactorvendor on this unit were reviewedprimarily from the viewpoint of dynamicmathematical modeling. These calculationsprovide an independent check onthe A/E's model shown in Fig. X A-22.The reactor building frequencies determinedby the A/E and those found by thereactor vendor are as follows:ReactorBuildingMode1212DirectionN-SN-SE-WE-WNaturalFrequency (cps)RCSA/E Supplier5.056.133.645.495.056.213.635.56I0X-72

The maximum shear and moment at thereactor pressure vessel skirt base dueto the seismic response of the systemwas found to be 205 kips and 6320 kipfeet,respectively; the allowables were2396 kips and 96,276 kip-feet.A6.4.2.1.1 Commentary. There appearsto be excellent agreement between thebuildings natural frequencies computedby the A/E and the RCS supplier, andthis provides a high level of confidencein these values.A6.4.2.1.2 Conclusion. The mathematicalmodeling for determining the seismicresponse of the vessel and internals isadequate.A6.4.2.2Reactor Pressure VesselNozzles.The analyses of the structural integrityof a number of the nozzles at penetrationsinto the BWR vessel were reviewed.These include:a. Main Steam Outlet Nozzle, No. 14.b. Recirculation Line Inlet Nozzle, No.7.c. Recirculation Line Outlet Nozzle,No. 8.d. Feedwater Nozzle, No. 10.e. Core Spray Nozzle, No. 11.f. 2-in. Instrumentation Nozzle, No.12.g. Control Rod Drive Hydraulic SystemReturn Nozzle, No. 13.h. Jet Pump Instrumentation Nozzle, No.19.i. 2-in. Drain Nozzle, No. 22.j. Vent Nozzle, No. 204.k. Head Spray Nozzle, No. 206.The analyses for these nozzles wereprepared by the reactor vessel manufacturerto the reactor vendor's specificationsand are outlined in References 32and 33.The same analytical technique was usedin evaluating all these nozzles. ABijlaard analysis, computerized forcylinder-to-cylinder junctions, was usedto obtain stresses in the vessel wall.Corresponding stresses for the nozzleswere computed from beam-type formulae.The membrane stresses due to internalpressure are also considered. In addition,the provisions of article NB-3222.4(d) (Components Not RequiringAnalysis for Cyclic Operation) of SectionIII of the ASME Code were invokedto qualify most of the nozzles infatigue.Table X A-24 presents the stress intensitiesreported for the nozzles.Nozzles that qualified in fatigue underArticle NB-3222.4(d) are indicated.A6.4.2.2.1 Commentary. There is no evidencein the documents offered for ourreview that the discontinuity stressesin the nozzles due to pressure loadingsare accounted for in this analysis.Moreover, none of the documents submittedfor out review contained a thermalanalysis. However, without an analyticaldemonstration or stress analysis,Reference 33 states that "...these nozzlesare not subjected to severe thermaltransient conditions, therefore primaryand secondary stresses will not be critical."We believe, however, that an analyticaldemonstration of this point forat least one of the nozzles, preferablythe 24-in. outlet nozzle, should be inthe documentary record.A6.4.2.2.2 Conclusion. Although theseanalyses do not entirely conform toanalytical standards and procedures ineffect today, we believe that in view ofthe adequate margins exhibited under theloadings considered, the structural integrityof these nozzles can be consideredadequate.A6.4.2.3Reactor Pressure Vessel Skirt.The analysis of the reactor vessel supportskirt was performed by the vesselmanufacturer (Ref. 34) using a finiteelement model of the skirt vessel junctionand computerized techniques ofstructural and thermal analysis. Boththermal transient and mechanical loadingconditions were considered, includingthe effects of the maximum seismic eventand jet loads.An elastic analysis was performed firstwith the results given in Table X A-25.As can be seen the allowable stressintensity range computed on an elasticbasis exceeds the code allowable. Thecritical location was found to occur atthe end of the filleted section nearestthe skirt.A simplified elastic-plastic analysiswas therefore performed. The methodused was subsequently incorporated intoSection III of the ASME Code under para-X-73

graph NB-3228.3. Although the currentcode requires demonstration that somecollateral conditions be satisfied,these were not mandatory at the time theanalysis was made. The analysis for thevessel skirt demonstrated that theelastic-plastic usage factor was 0.55

final evaluation of the adequacy of therecirculation lines.The reactor vendor has provided theresults of a stress analysis of a mainsteam line (for a power plant of morerecent design) which does take intoaccount the dynamic amplification of thevertical component of the earthquake.The results show that stresses arewithin allowables. We have reviewed theanalysis and note the following differencesbetween the design which is thesubject of this study and designanalyzed:a. The recirculationshock suppressorsdesign compared todesign.pump has fouron the analyzedtwo on the studyb. There are a total of ten shock suppressorson each loop of the analyzeddesign compared to a total offour on the study design.c. The analyzedand one nearsor compareddesign.design has one verticalvertical shock supprestonone for the studyIt is apparent that the analyzed designhas significantly more seismic restraintsthan the design we are studying;therefore the seismic analysisresults of the other design are notapplicable.We conclude that the recirculation linefrc7- the viewpoint of seismic resistancedoes noz completely satisfy the designcriteria. Nevertheless, estimatedstress levels indicate that failure isnot ex ected and the design is consideredadequate.A6.4.2.5Main Steam 'ines.The i,±n steam lines extend betweeneleva-,->ns 138 ft 2 i.. and 203 ft 1 1/2in. Thi piping system. was seismicallyqualified by the supplier on the basisof an analysis performed for a similarplant at another site. Only the horizontalresponse spectra were used here,with the vertical response computed onthe basis of constant 0.033g and 0.08gspectral accelerations in the same wayas it was applied to the recirculationlines (see section A6.4.2.4). A summaryof the natural periods and spectral accelerationvalues used in the analysisfor tfh, First nine modes is shown inTabl. '. %-28.The spectral accelevix ion values used bythe analyst were exactly as found on thefloor response spectrum curve withoutpeak broadening. For example, the computedperiod for mode 2 is 0.187 sec,for which the analyst picked the exactspectral acceleration value of 1.81g.The same curve indicates, however, thatfor a natural period of 0.19 sec, (anincrease of 1.6%), the spectral accelerationvalue would be 3.0g, i.e., anincrease of 66% over the actual valueused. Critical stresses are summarizedin Table X A-29.A6.4.2.5.1 Commentary. General commentsregarding the dynamic response dueto the vertical component of the earthquakeare identical to those given insection A6.4.2.4. In addition, since itis not unreasonable to expect that theactual second-mode frequency could behigher than the computed value by 1.6%,it would be more appropriate to use the3.Og value rather than 1.81g.A6.4.2.5.2 Conclusion. The reactorvendor has provided the results of ananalysis of a more recent design whichaccounts for the dynamic amplificationof the vertical component of the earthquakeand shows that streses are withinthe allowable. A comparison between themore recent design and the design whichis the subject of this study shows thatthe degree of seismic restraint issimilar.On the basis of this additional evidence,we conclude that the design ofthe main steam line is adequate.A6.4.2.6Main Steam Line IsolationValves.Because they represent a potentiallarge-diameter path for the escape ofradioactivity to the environment shoulda steam line break outside the containmentstructure, the main steam lines aregiven special isolation consideration.Two 26-in. automatic isolation valves,each powered by both air pressure andspring force, are provided in each ofthe four main steam lines (see Fig.X A-28). One valve is located outside.the primary containment and the otherinside, approximately at elevation 135ft. Each valve is provided with ahydraulic damper to prevent excessivelyfast closing rates that might damage thevalve or piping system.The valves have a Y-pattern body and acylindrical main plug that moves on acenter line 450 from vertical. The aircylinder operator is supported from thevalve body by four rods that also serveas guide shafts for springs which providea closure mechanism with sufficientX-75

stored energy to seat the valve shouldthe air supply be lost.Normal closure is powered by a solenoidtriggeredair cylinder that drives acentral stemn connected to the main plug.A spring and dashpot are incorporatedinto the stem to help control closurerates.The analytical document prepared for theseismic analysis of the BWR MSIV (fromwhich data used in the FSAR were taken)was originally presented for our review(Ref. 35). It contains errors as notedin our commentary.During our initial examination of thisdocument, we questioned some of its assumptionsand methods. Consequently wewere shown Reference 36, an analysisprepared for the seismic qualificationof a MSIV for another plant. The valveis of the same nominal size, is suppliedby the same manufacturer, but differs insome dimensions. However, the dimensionaldifferences are not sufficient toinvalidate the applicability of the calculationsto the BWR valve; we thereforealso reviewed this second analysis.These calculations were prepared to theBWR NSSS supplier's specification thatseismic coefficients of 0.6g verticaland 1.5g horizontal be applied as constantshock acceleration spectra from0.25 to 33 cps.The valve operator weighs about 3/4 tonand is supported from the valve body onfour 3-in.-diameter rods about 4 ftlong. During an earthquake the operatormass will be set into vibratory motion.Items of concern are the consequentstresses in the rods and the ability ofthe valve to remain functional throughoutthe seismic event.Consequently, a 39-node 21-mass dynamicmodel of the actuator, the support rods,and their appurtenances was prepared.Using the STARDYNE computer program,mode shapes and corresponding frequencieswere computed. The naturalfrequency was found to be 13.3 cps andis the only mode within the range ofinterest (0.25 to 33 cps). Because ofthe symmetry of the support rod arrangement,the structure is equally stiff inall transverse directions. It willrespond to horizontal accelerations inthe plane of the excitation and exhibitthe same natural frequency in alldirections.The maximum stress was found to be20,130 psi; the allowable value is48,000 psi.Experimental and analytical evidencebearing on the ability of the main steamisolation valve to function under a widevariety of accident flow conditions iscontained in the NSSS supplier's topicalreport (Ref. 37) . This report describesan extensive test demonstration programin which a commercial main steam isolationvalve was installed in a full-scaleflow test facility and operated underhigh flow conditions closely simulatingthose anticipated during a plant accident.Forty tests encompassing bothsingle-phase and two-phase flow wereconducted, a number of these being atthe highest flow rates postulated for adesign basis accident. These testsincluded:SteamWaterTest TypeTwo Phase (variousqualities, 0.17 to 0.45)Surge (quality 0.1 to0.33)Flow Range(lb/sec)50 to 1080240 to 34901530 to 3860520 to 2970The report concludes that the MSIV willclose under reactor accident flow conditionsas evidenced by their closureduring experimentally simulated reactoraccident flow conditions.In another test reported in an appendixto Reference 36, a valve similar to thatinstalled in the BWR plant was loadedperpendicular to the valve stem centerline by a dead weight applied through apulley. The valve was then operatedunder the side thrust load. The loadingproduced transverse deflections equal tothose expected under dead weight plus1.5g seismic acceleration.Closing time was insignificantly differentfrom that when the valve was notloaded.A6.4.2.6.1, Commentary.A6.4.2.6.1.1 Seismic Analysis of Reference35. The author, although properlyidentifying the fundamental character ofthe analytical problem as that of aframe, nevertheless subsequently treatsthe multi-rod structure as a single cantileverbeam. This implies a shear connectionbetween the rods and results inX-76

Ia misevaluation of the stiffness of theoperator structure. The computed stiffnessis too high by approximately 25times and results in an estimate of thenatural frequency of 74.1 cps, which isabout 5 times too high. Stress estimatesare about 25% of those which thestructure may actually experience duringan earthquake.This computation bears the major equipmentsupplier's stamp attesting to thefact that the document was approved andno further action was required.A6.4.2.6.1.2 Seismic Analysis of Reference36. Although the computationsheets in this report are signed off aschecked by a member of the consultingfirm that prepared them:a. The weight of the modal masses doesnot total to the weight of thestructural components (due to anerror in weight allocation on page17 of Reference 36.b. A spot check indicates that thestructure is not in equilibrium,even for simple dead weight loading.The reactions reported on page 14for dead weight (e = 0O) do notbalance the applied loads (i.e., theset of weight loads actually used inthe model) . Neither the forcesalong the rod axis nor the overturningmoment is equilibrated bythe reported reactions at the baseof the structure. We suspect an inputerror.The error in assigning masses to nodeshas only a small effect on results.Much more disturbing is the fact thatthe structure does not appear to be inequilibrium under its own weight (ormore precisely, under the weight loadsit was taken to have). Failure tosatisfy this kind of a check oftensignals significant errors.However, we believe that in this analysisthe error falls to the conservativeside. It appears to lower the naturalfrequency (which we estimate to be about15 cps) and to overstate stresses.Moreover, there are subsequent conservatismsdeliberately introduced into theanalysis:a. Stress resultants are computed bysumming absolute values of seismicand dead weight stresses withoutregard to sign.b. The analysis considers the possibilitythat the valve might be installedwith the actuator axis tilted4Q0 out of plane (i.e., rotated 400off the vertical, about the pipeaxis). This condition producessignificantly higher stresses thanwould upright installation. In thesummary, this worst-case stress isreported. since one expects verticalinstallation to be the usualpractice, the reported maximumstress is conservative for the majorityof actual installations.The functional flow tests were conductedusing a 20-in, valve supplied by amanufacturer different from the one whosupplied the MSIV for the BWR plant.The A/E contends that in light of thesimilarity of design among valves ofthis type supplied by various manufacturersfor this application the resultshave general relevance for all suchvalves. The report implies that theeffects of such differences as may existbetween designs can be analyticallyaccounted for and are not sufficient toalter the main results. This contentiondoes not seem unreasonable.Although these tests were conducted withsome flow rates in excess of those predictedfor major accidents, the predictedfluid pressures could not be fullyduplicated. This limitation on the testwas imposed by the engineers of the(conventional) steam generating plantfurnishing the steam and water used asthe test fluid. They feared structuraldamage to plant equipment from suchpressures. Since MSIV have been trippedmany times under normal flow conditionsat full reactor pressure, we do not feelthat the inability to duplicate pressuresinvalidates the conclusionsreached in Reference 37.A6.4.2.6.1.3 Environmental Qualification.Evidently, a formal qualificationprogram has not been conducted to demonstrateby test that the electropneumaticcontrol assembly on the MSIVwill be functional when exposed to theambient environmental conditions withinthe primary containment following aLOCA. Discussions with the NSSS supplier'srepresentatives disclosed thefollowing information pertinent to environmentalqualification of the two componentsof this assembly which aresubject to degradation:a. Solenoid Actuators: Qualificationis not needed because the design isfail-safe. The electrical signal isneeded to hold the valve open; itsremoval for any reason- will resultin valve closure unless movement ofthe pneumatic control valve control-x- 77

led by the solenoid is impeded(i.e., by excessive binding of theseals).b. Seals on Pneumatic Valves: Analysisshows that these will not degradeappreciably during 5 years of normalplant service (aging due to exposureto the temperature, humidity, andradiation environments). The sealswill be replaced every 2 years inthe preventive maintenance program.These discussions also disclosed that aseismic qualification of the assemblywas not conducted because of the failsafedesign and lack of a mechanismwhereby an earthquake could cause ahang-up in the proper operation of thepneumatic valves.The previous discussion gives reasonableassurance of the adequacy of the electromagneticcontrol assembly. A formalqualification program would, however,provide additional assurance.A6.4.2.6.2 Conclusion. The ability ofthe i..in steam isolation valve to closeunder the maximum flow rates expectedfrom a major plant accident has beendemonstrated in full-scale tests. Seismicresistance has also been demonstrat-'ed. The design is adequate.A6.4.2.7Pipe Whip Restraint forRecirculation Lines.The pipe whip restraints for the recirculationline are designed to constrainthe piping only in -the event of verylarge pipe movement due to a largebreak. This is accomplished by having alarge radial clearance between the OD ofthe pipe and the ID of the restraint.The restraints are anchored to thestructural framing surrounding thereactor vessel. Maximum spacing of therestraints is determined by assuming apipe break and finding the moment armthat will just develop a plastic hingein the pipe at the restraint. Thethrust load is taken as the systempressure times the pipe flow area. Therestraints are designed to have amaximum stress of 90% of the yieldstrength. They have been designed towithstand the following load cases:a. Case I: Load is along the restraintsymmetry line; it has only onecomponent, T.b. Case II: Load is transverse to theconstraint symmetry line; it has atransverse component, S, and givesrise to a moment on the anchorbolts, M.C. Case III: Load is at 450 to theco-nstraint; it has components T andS and gives rise to a moment, M, atthe anchor bolts.A summary of the actual and allowableloads for the three cases is given inTable X A-30.A6 .4.2.7.1 Commentary. Although therecirculation line carries water, thedesign pressure and temperature are 1250psig and 575 0 F, respectively; so, ifthere is a pipe break, the water willflash into steam. A thrust coefficientof 12.5 is therefore appropriate. Sincethe operating pressure is actually 1000psig, use of a pressure of 1250 psigdoes, in effect, give the thrust coefficient.The calculations appear to bereasonable, and the results show thatactual expected loads are within theallowables.A6.4.2.7.2 Conclusion. The piperestraints for the recirculationsatisfy the design criteria andadequate.A6.4.2.8Reactor Pressure ReliefValves.whiplineareThe 11 pressure relief valves are mountedon the main steam line headers withinthe drywell portion of the containmentat elevation 160 ft. Each relief valve(Fig. X A-29) is self-actuating at thepresent relief pressure (set-pointranges between 1080 and 1100 psig forthe various valves) but may also beactuated by a solenoid-operated airvalve to permit remote manual orautomatic opening at lower pressures.Each valve consists of a main valve diskand piston, operated by a second-stagedisk and piston displaced by either apressure sensing pilot or a pneumaticallyoperated mechanical push rod.The automatic depressurization system(ADS), a subsystem of the steam supplypressure relief system, serves as abackup to the high pressure coolantinjection system (HPCIS) under smallleakLOCA conditions. Five of therelief valves are connected to providefor automatic depressurization of thereactor in this type of accident. Thesevalves are equipped with accumulatorsand check valves arranged to ensure thatthe valves can be opened and held openeven if the air supply should fail. Theaccumulators are sized for a minimum offive valve operations. The reliefvalves also can be operated by remotecontrols from the main control room.IX-78

A6.4.2.8.l Commentary. Discussionswith representatives of the reactor vendor(who was responsible for supplyingthe valves) disclosed that no formalqualification program has been conducted.Because these valves must function*properly at the time of peak temperatureconditions following a small-break LOCA,because they contain some nonmetallicmaterials (e.g., seals and electricalinsulation) which are known to be adverselyaffected by thermal and radiationaging and by LOCA conditions, andbecause the valves are subjected to highinternal and ambient temperatures duringtheir installed life, we strongly believethat a proper, comprehensivequalification program (i.e., one includingaging simulation, seismic testing,and LOCA exposure) should be conducted.We note that Reference 31 states, "Asolenoid control valve, identical tothose associated with the [BWR plant]safety/relief valve was tested for 10hours at temperature in excess of 300'Fand a pressure of 62 psig. Results ofthis test show that the solenoid valveperformed satisfactorily and there wasno indication that temperature limitswere approached." Since an ambienttemperature of 340OF can reasonably beexpected to occur at the time when thesevalves are required to operate, thisstatement is inconclusive. It is particularlyimportant that the electricsolenoid valves that provide for bothADS and manual actuation be fullyqualified.A6.4.2.8.2 Conclusion. We cannot concludethat the design of this componentis adequate on the basis of availableinformation. Design adequacy has notbeen demonstrated.A6.4.3 CORE SPRAY SYSTEMThe core spray system consists of twoindependent, identical loops, each ofwhich is capable of preventing excessivefuel cladding temperatures by coolingthe fuel in the core should a LOCAoccur. Each loop has two 50% capacitycentrifugal pumps driven by electricmotors, a spray sparger in the reactorvessel above the core, piping and valvesto carry water from the suppression poolto the sparger, and associated controlsand instrumentation.The subsequent sections present detaileddescriptions and design adequacy assessmentsof the following components of thecore spray system: piping, piping hang-*ers and snubbers, pumps and drives andpump and drive mounting, valves, valvemotor operators, and instrumentation.A6.4.3.1 Piping.An isometric drawing of a portion of thecore spray piping inside the containmentis shown in Fig. X A-30. The piping isanchored to the drywell shield wall atelevation 143 ft 6 in. and connects tothe reactor vessel nozzle at elevation188 ft 7 1/2 in.The fundamental frequency was computedto be 7.5 cps. Seismic loads were basedon an average of the response spectrafor the upper and lower anchor pointsassuming 0.5% critical damping values.Table X A-31 summarizes the criticalstresses. Modal stresses were combinedby the square root of the sum of thesquares method. The highest of thehorizontal responses was added to thevertical response. The piping was evaluatedin accordance with the B31.1Piping Code.A6.4 .3.1.1 Commentary. An independentcomputer analysis was made, and theresults were evaluated in accordancewith the rules of B31.1 as well as therules of the current NB-3600 portion ofthe ASME Code, Section III.The independent analysis resultsgood agreement with the A/E result.showDiscussion with the A/E indicates thatthe jump discontinuity in the temperaturefrom 150 to 5750F shown on theisometric drawing CFig. X A-30) is. notrealistic since the piping is exposed tothe ambient temperature. It is expectedthat the actual thermal gradient alongthe pipe length is a gradual one.Evaluation of computed stresses in accordancewith the current code requirementsshowed the stress levels to bewithin the allowables if we neglect theunrealistic gross thermal discontinuityshown on the isometric drawing.A6.4.3.1.2 Conclusion. The core spraypiping inside the containment satisfiesthe design criteria and is adequate.A6.4.3.2Hangers and Snubbers.As with other seismic Class I pipingsystems, the core spray system piping issupported by both conventional pipinghangers and special shock absorberstermed snubbers. The locations of thesecomponents are shown in Fig. X A-30.The hangers are designed to avoid placingtwisting forces on the piping systemand to provide a constant supportingforce even when the piping moves as ax-79

esult of thermal expansion as the temnperatureof the contained f luid changes.A6.4.3.2.1 Commentary. No separatedocumentation pertaining to hanger designis available. The hangers arerugged devices, do not contain componentsthat can be degraded as a resultof the nuclear radiation and temperatureconditions present in the plant or whichmay occur in an an accident,,and havebeen extensively used in both fossil andnuclear power plants for many years.The snubbers contain a hydraulic fluidand a movable piston. Seals must beprovided to prevent the loss of fluidaround the shaft. The seals used insnubbers installed in several powerplants have deteriorated within a periodof only several months, resulting in aloss of fluid and consequent loss ofeffectiveness. However, the snubbersinstalled in the BWR plant were producedby a manufacturer other than the onethat produced the malfunctioning snubbers.The hydraulic fluid and the sealsused in the snubbers that were installedin the plant were reported to be radiation-resistantmaterials.No reports on qualification programsconducted for the snubbers used in theBWR plant were made available to thereviewers. It is our understanding,however, that the snubbers at the BWRplant have been modified to use the bestmaterials now available (similar to themodifications made at the PWR plant).A6.4.3.2.2 Conclusion. On the basis ofengineering judgment, we consider thedesign of the hangers and snubbers to beadequate. In view of the fact that sealfailure and loss of snubber fluid haveoccurred in other installations and thata proper qualification program has notbeen conducted, it is essential that aperiodic surveillance of the snubbers bemade, and that the seals and other materialsthat are susceptible degradationas a result of exposure to temperature,humidity, radiation, etc., be replacedbefore properties are seriously altered.There is insufficient information toassess adequacy *of the seal design ofthe snubbers. Resistance to seismicloads may be negated due to loss ofhydraulic fluid.A6.4.3.3 Pumps and Drives.There are four core spray (CS) pumps andmotor units located in four separateflood-protected rooms, two in the southeastcorner and two in the northeastcorner of the reactor building. Thewatertight entryway is identical to thatshown in Fig. X A-36 for the residualheat removal (RHR) pump rooms.Each pump and motor is vertically mounted,and the motor is supported on abracket that bolts to the .pump casing(Fig. X A-31). The pump itself isbolted to a foundation pad on the floor(Fig. X A-32). Pump mountings areoriented in three different directions:two of the pumps are parallel and theother two are oriented at 450 to thefirst pair and 900 to each other. Eachpump can deliver water to the reactorcore at a rate of 3125 gpm when a 105-psi pressure difference exists betweenthe reactor vessel and primary containment.Each CS motor is rated at .600 hpand operates at 482 kW.The weights and heightsgravity for the coremotors are:unitMotorMountingBracketPumpWeight (lb)520053003000of the center ofspray pumps andHeight of CGAbove MountingPad (ft)6.513.941.92The A/E's original seismic calculationsfor the anchor positioned the entirehorizontal inertia force of the motorand pump at the pump CG, thus reducingthe overturning moment due to the higherposition of the motor CG. This waspointed out to the A/E during the courseof the review. The A/E then computed thefundamental frequency of the pump-motor,which permitted the use of lower accelerationvalues from the floor responsespectra curves. The revised resultsshowed a small positive margin. Nevertheless,the A/E has strengthened theanchor installation to provide additionalmargin.There are no piping systems in the CSpump rooms other than those associatedwith the respective pumps.The A/E has performed an analysis whichshows that the pump seals will not besignificantly degraded by exposure tonuclear radiation emanating from the radioactivewater pumped under accidentconditions.A6.4.3.3.1 Commentary. A formal qualificationprogram for the pumps anddrives has not been conducted. However,the reactor vendor has functionallyjX-80

tested a similar motor in a saturatedsteam environment at atmospheric pressure(Ref. 38).From the viewpoint of seismic protection,the pumps are strategically locatedand oriented as follows:a. They are at or near ground level,where amplification of the groundresponse spectrum is minimal.b. They are located in separate roomsin different areas of the building.C. Their mountings are oriented inthree different directions.Thus, for a seismic event, only twopumps (at most) are likely to experiencethe same excitations and consequentlyexhibit similar responses.A static analysis (using 0.6g horizontaland 0.07g vertical) was performed. Themaximum g load from the floor responsespectra is 0.3g horizontal for the DEand 0.72g horizontal for the MCE. Independentcalculations performed by theA/E show that the fundamental frequencyof the unit is 12 cps. The spectralacceleration value for this frequency is0.14g for the DE and 0.336g for the MCE.The 0.6g value used in the analysis istherefore conservative.The most critical stress was found tooccur in the driver stand bolting(16,500 psi computed, compared to thestated allowable stress of 20,000 psi).Margins above allowable stresses aresubstantially larger for all othersections examined.As discussed in section A6.4.5, a pseudoseismic qualification test of the RHRpump has been conducted in which a pump,motor and coupling were set up and runin a test loop. The RHR and CS pumpsare made by the same manufacturer. TheRI-R pump is about twice the size of theCS pump; so its natural frequency wouldbe lower. Since the pumps are similarlymounted, it is reasonable to expect thatthe CS pump would successfully pass aqualification test under a maximum loadof 0.34g. An actual test of the CS pumpwould offer more assurance, however.A6.4.3.3.2 Conclusion. We believe that(1) the seismic computations justify theassumption that the structural supportof the core spray pumps and motors iscapable of meeting the seismic requirementsof the specifications, (2) on thebasis of physical isolation considerationsand the test conducted on asimilar motor, there is no reason toexpect that the functioning of the pumpsand drives will be impaired as a resultof a pipe break, and (3) it is reasonableto expect that the pump and motorwill function adequately following asafe shutdown earthquake, based upon theperformance demonstrated in the vibrationtest of the RHR pump. The unitssatisfy the design criteria and areadequate.A6.4.3.4Valves.Several valves are associated with thecore spray system. Specifications forseveral of the valves were reviewed, andthe procedures that were followed in thedesign of the valves were discussed withthe A/E. We were informed that a qualificationprogram for the various valveswas under way but that no data were yetavailable.A6.4.3.4.1 Commentary. The specificationsdo not require that the ability ofthe valves to function properly underseismic events be demonstrated but ratherthat they not fail as pressure containingmembers. Some relatively minorinconsistencies among the specificationswere also noted (e.g., the requirementto withstand aging and nuclear-radiationexposures under LOCA and post-LOCAconditions).A6.4.3.4.2 Conclusion. In the absenceof a formal qualification program, wecannot make a rigorous assessment of thedesign adequacy of the valves. On thebasis of the reviewer's experience inperforming stress analyses of similarvalves, we can conclude that there is agood probability that the valves can beexpected to function properly under anaccident situation. In our judgment thevalves satisfy the seismic criteria.A6.4.3.5 Valve Motor Operators.Figure X A-33 shows the valve motor operators(VMO) appended to and supportedby the valves they control. A largenumber of VMO, of varying size butsimilar in design, are used in theplant. The VMO units used inside theprimary containment have Class H insulation;the units installed outside haveClass F insulation. None-of the VMO inthe CSS are installed within the primarycontainment.A6.4.3.5.1 Commentary. The VMO installedin the BWR plant were subjected to arigorous qualification testing programthat consisted of aging, nuclear radiation,simulated seismic conditions, andsimulated LOCA exposures in the simulatedseismic test (Refs. 39 and 40). AnX- 81

aged operator was mounted on a vibrationtable and subjected to vibration of 20cps at ig for 2 min on and 1 min off,repeated five times along both thevertical and horizontal axes of theactuator. This test does provide considerablereassurance that the VMO willfunction adequately even though the testdoes not conform to the current requirements.A6.4.3.5.2 Conclusion. We believe thatthe qualificaton testing program thatwas performed, although not fully equivalentto the current requirements, doesnevertheless demonstrate design adequacy.A6.4.3.6Instrumentation.The instrumentation and controls whichsense that a LOCA event has occurred andwhich initiate functioning of the corespray system are part of the reactorprotection system. The CSS instrumentationincludes pressure switches on thedischarge line from each set of corespray pumps to indicate the successfulstart-up of the respective pumps, theflow-measuring instrumentation in eachof the pump discharge lines, the motorcurrent, and the valve-positionindicatingswitches (included as part ofthe valve operators). Flow rates, motorcurrents, and valve positions are displayedin the control room.A6.4.3.6.1 Commentary. Specific informationrelating to the qualification ofthese instruments is not available;hence we cannot assess design adequacy.However, the failure of the instrumentationcannot completely negate the functioningon the system. The dischargeflow rate would be lowered if the pumpdischarge pressure signal were notpresent to cause the bypass valves to beclosed and thereby enable all the flowto be directed to the core.A6.4.3.6.2 Conclusion. Either the corespray instrumentation should be qualifiedor the effect of reduced deliveryto the reactor in the event of a LOCAshould be evaluated. There is insufficientinformation to assess adequacy ofthe CS instrumentation.A6.4.4HPCIS TURBINEThe high pressure coolant injection system(HPCIS) consists of a steam turbinedriving a constant-flow pump, systempiping, valves, controls, and instrumentation.The pump has a design flow rateof 5000 gpm at 1120 to 150 psid (poundsper square inch differential betweenreactor vessel and primary containment).The HPCIS is capable ofdesign flow rate withinreceipt of anthe automaticappropriatecontrols.reaching25 secsignalthefromfromThe HPCIS is installed in the reactorbuilding at elevation 88 ft. Suction isfrom the condensate storage tank and thesuppression pool. Injection water issupplied to the reactor feedwater pipingat a T-connection. Steam supply for theturbine is piped from a main steamheader within the primary containment.This piping is provided with an isolationvalve on each side of the drywellwall. Exhaust steam from the HPCISturbine is discharged to the suppressionpool. Remote controls for valve andturbine operation are provided in themain control room.As is shown in Fig. X A-34, the HPCIpump and turbine are mounted on a massiveconcrete foundation block, which inturn is an integral part of the reactorbuilding foundation. A speed reducer isprovided between the two stages of thepump (booster pump and main pump). Thelarge pipe in the foreground of thephotograph connects the pump stages.Figure X A-35 shows the turbine end ofthe unit.The turbine has two devices for controllingpower: (1) a speed governor, whichlimits turbine speed to- its maximumoperating level, and (2) a controlgovernor, which is positioned by a demandsignal from a flow controller, tomaintain constant flow over the pressurerange of HPCIS operation.A consulting organization performedseismic calculations (Ref. 41) for theturbine.The shafts for the turbine and pump arehorizontal and are connected via a flexiblecoupling. Equivalent static loadingsof 1.5g horizontal and 0.48g verticalwere applied.Seismic stresses were found in the turbinecasing, shaft, hold-down bolts andconnections, pedestals, base plate, andtie-down to the foundation.The analysis determines stresses anddisplacements in 10 critical areas includingbolting, turbine shaft, pedestals,the stop valve assembly and thebase plate connection.The analysis of the yoke that connectsthe cylinder to the stop valve makes theimplicit assumption that there is ashear connection between the two legs ofthe yoke.a4X-82

A6.4.4.1 Commentary.The implicit assumption that there is ashear connection between the two legs ofthe yoke is incorrect. We estimate thatthe correct stresses in the yoke wouldbe approximately 20,000 psi rather thanthe 13,100 psi computed by the vendor.This value is still within the limit of33,000 psi.Otherthan the above error the analysisis thorough and adequate.A6.4.4.2Conclusion.We conclude that the HPCIS turbine iscapable of resisting seismic environments.Design criteria are satisfied.Seismic design adequacy has been established.A6.4.5RHR PUMPS AND DRIVESThe plant has four identical residualheat removal pumps. These are the mainpumps for the residual heat removal(RHR) system. This System can be assignedseveral functions:a. Removal of residual core heatingduring and after plant shutdown (thesystem's normal operating function).b. Re-flooding of the core after anaccident in which watering the coreis temporarily lost. The RHR systemis one of three pumping systemswhich can be assigned this functionand is activated when reactor pressureis reduced and rapid delivery.of a large volume of water isrequired.C. Post-accident containment cooling.Heat is removed from inside thecontainment by continual cooling ofsuppression pool water by circulationthrough the RHR heat exchangers.d. Reduction of primary containmentpressure. After an accident the.RHRsystem can be used to spray waterinto the primary containment (condensingsteam and thereby reducingpressure) simultaneously with thefunction described in item c.Each pump is housed in a separate floodprotectedroom of the reactor building.Figure X A-36 shows the watertight entrywayto one of the rooms. One pair ofrooms is located in the northwest cornerof the building, the other in thesouthwest corner. In each room thepumps are mounted on the floor atelevation 91 ft 6 in. (see Fig. X A-37).In addition to the dispersion of pumplocations, the orientations of the pumpmountings and connected piping are alsodisparate. Two of the pump dischargeconnections are parallel, and each ofthe remaining pair of pumps is orientedat 450 to the first pair and at 9Q0 toeach other.A6.4.5.1Seismic Qualification.Each pump and its motor drive are verticallymounted as a unit, and the motoris supported on a stand that bolts tothe pump casing. The casing, in turn,is mounted on a footing that bolts to afoundation pad. The piping also providesa measure of lateral restraint tothe pump casing. The configuration isshown in Fig. X A-38.The weightsgravity forbelow:UnitMotorMountingBracketPumpand heights of the center ofthis equipment are tabulatedWeight (lb)14, 0002,2005,600Height of CGAbove MountingPad (ft)9.83.671.5The structural adequacy of the pump andmotor mountings was assessed with astatic g-load analysis using 0.6g forhorizontal and 0.07g for vertical loading.This original seismic analysis is partof the supplier's design computations(Ref. 42). These computations coveronly the sizing of structural memberssupporting the RHR pump and motormasses.The effects of the seismic static gloads were examined at five criticalcross sections:a. The motor mounting bolts.b. The cross section of the driverstand at bottom of the cutouts.c. The driver stand mounting bolts.d. The base plate welds.e. The foundation bolts.The results of these computations showthat design stresses are well below theallowable stresses.X-83

A vibration test of the RHR pump hasalso been conducted; results are reportedin Reference 43. In this test apump, motor, and coupling were set upand run in a test loop. Vibrations wereproduced by suppressing the net positivesuction head of the pump, i.e., byintroducing cavitation at the intake.Horizontal accelerations and displacementsalong two perpendicular axes wererecorded for stations at three differentelevations on the PUMP and motorassembly.The maximum horizontal accelerationoccured at the top of the motor and wasapproximately 3.5g.Twenty tests, lasting at least 30 seceach, were, run for a number of netpositive suction head (NPSH) conditionswhile the pump was operating in theloop. The pump operated successfullyunder these vibratory conditions.A comparative inspection of bearingseal, wear ring, and shaft runout wasmade before and after the set of testruns. No visible damage occurred.The report concludes in para. A6.4.5.3.3that the natural period of vibration ofthe pump is about 0.033 sec.In addition, Ref. 43 includes a sectionin which further static g-design computationsof the type described in Ref. 42are reported. These computations generallyconfirm the results of the previousanalysis, and the report concludes thatthe structure is adequate under horizontalg loadings of at least 1.5g.A6 .4.5.*2Environmental Qualification.A formal qualification program for thepumps and drives has not been conducted.However, the reactor vendor has functionallytested a similar motor while itwas subjected to a saturated steam environmentat atmospheric pressure (Ref.38) . in addition, as in the case of thecore spray pumps, the A/E performed ananalysis which shows that for 6 monthsfollowing a LOCA the pump seals will notbe significantly degraded by exposure tonuclear radiation emanating from thepumping of radioactive water. A formalqualification program has not beenconducted, however.A6.4.5.~3Commentary.A6.4.5.3.1 Location, Arrangement, andHousing. The physical arrangement ofthese pumps and the manner in which theyare housed have important implicationsfor the seismic and environmentalintegrity of the units.From the viewpoint of seismic protection,the pumps are strategically locatedand oriented as follows:a. They are at or near ground level,where building amplification of theground response spectrum is nonexistentor minimal.b. They are located in separate roomsin different areas of the building.c. Their mountings are orientedthree different directions.Thus, for a seismic event, only twopumps (at most) are likely to experiencethe same excitations and consequentlyexhibit similar responses.From the viewpoint of the ability of thepumps to perform under accident conditions,the manner in which the pumps arehoused adds a substantial measure ofassurance that they will operate despiteadverse environmental conditions in theplant.The individual pump compartments containno other piping than those required forRHR pump operation (i.e., pump loop piping,RHR heat exchanger piping, servicewater lines for pump cooling, and theroom cooling system). The pump roomhatchway doors (which protect the pumpfrom floods) also isolate the pump fromspray or steam that could result fromrupture of other pipe systems outsidethe pump room.A6.4.5.3.2 Seismic Analysis. The seismicdesign computations of Ref. 42present credible evidence that the mainstructural members supporting the RHRpump and motor have been designed tocode requirements for the specifiedearthquake loadings based on the followingconsiderations:a. The static g loading of 0.6g horizontaland 0.07g vertical is adequatewhen compared to the peakfloor response levels of 0.5g and0.07g vertical.b. The computations employ standardstrength-of-material formulas.* Althoughthese are first-order designapproximations involving simplificationsand although they may notaccount for all stresses in thestructure, they do reflect the majoreffects of the design loads. Moreover,the designer preserved sub-in4X-84

stantial margins over and above theallowable stresses at every sectionbe examined.c. The formulas used are appropriate tothe situations considered. Forexample, when the stresses in themotor mount at sections with cutoutswas examined, a formula was selectedwhich takes into account the reductionin the moment-carrying capacityof such sections due to the lack ofshear connection.d. For the structural members supportingthe pump and motor (the mainthrust of this computation) , theresults are conservative since theytake no credit for the additionalrestraint provided by the piping.e. The computations show evidence ofhaving been numerically checked and(although some numerical disparitiesare indicated) were found to have nonumerical errors of engineering significance.A6.4.5.3.3 Seismic Tests. The naturalperiod of the pump and motor is reportedin Reference 43 to be about 0.033 sec;that is, the natural frequency is 30cps. In our judgment, what we observedwas that the shaft speed was: 30rev/sec x 60 sec/mmn = 1800 rpm; ratherthan the natural frequency.Considering the method of excitation -inducing cavitation by starving thesuction head - one should not be surprisedto find the shaft frequencystrongly represented in the data sincecavitations (although a complex process)can be expected to provide a form ofunbalanced rotor.Moreover, if it were thought that thereported frequency was actually thenatural frequency of this structure, aserious design deficiency should havebeen declared, since the structure couldthen be excited by the motor's rotation.Nevertheless, the test record may supplyan indication of the structural response.Figure X A-39 shows two of thetest records. These are traces of thehorizontal displacement vs. time recordedat the top of the motor. The tracesexhibit detectable repetitive scallopsat 6 to 7 cps which may have been producedby the structure's response. Thisevidence is too meager to support asubstantive conclusion. However, thefrequency of 6 to 7 cps is close to thenatural frequency we predict for thisstructure on the basis of a simple 1independent calcula-degree-of-freedomntion.It is evident that the excitation methodused in these tests fails to modelearthquake-induced vibrations. However,this unit did experience repeated severevibrations over time periods at least aslong as those an earthquake would produceand also experienced g loadings (atleast at the motor mass, which is boththe largest system mass and the onefarthest from the supports) significantlylarger than those expected from thedesign earthquake. On this basis, onecan reasonably infer structural seismicadequacy.A6.4.5.3.4 Environmental Qualification.The residual heat removal system standsout among the engineered safeguards featuresas one whose postaccident functioncalls for heavy-duty operation over aprolonged period. Therefore, the environmentalqualification of this systemis particularly significant and shouldbe reported in a unified document.Nevertheless, it does appear, on thebasis of physical isolation considerations,tests conducted on a similar pumpdrive, and analysis of possible radiationdegradation of the pump seals, thatthere is no reason to expect that thefunctioning of the pump or motor will beimpaired as a result of pipe breaks inother systems or ambient temperatures,steam environments, or radiation exposurefollowing a LOCA.A6.4.5.4Conclusion.Seismic adequacy of the RHRdrive has been demonstratedanalysis and may be inferredvibration test.A6.4.6pump andby staticfrom theON-SITE ELECTRIC POWER SYSTEMSThe electrical power systems at theplant are designed to provide a diversityof dependable power sources whichare physically isolated so that anyfailure affecting one source of supplywill not propagate to other sources.The plant receives power from two separateoff-site sources. In the event ofa total loss of power from off-sitesources, auxiliary power is suppliedfrom diesel generators located on thesite. Each power source, up to thepoint of its connection to a 4-kyemergency auxiliary power bus, iscapable of complete and rapid electricalisolation from any other source. Loadsimportant to plant safety are split andX-85

diversified between auxiliary bus sections,and means are provided for rapidisolation of system faults. Stationbatteries are provided as a reliablesource of control power for specificengineered safeguards and for otherfunctions required when a-c power is notavailable. The diesel generators arehoused in a separate building, and thebatteries are housed within the complexof main plant buildings.A6.4.6.1 Diesel Generator Building.The diesel generators are housed in areinforced-concrete seismic Class Istructure with the floor at elevation127 ft. Each of the. diesel generatorunits is lccated in an individual room.The building is watertight to the designflood level of elevation 135 ft, and isalso intended to provide protectionagainst other natural phenomena, such astornados, lightning, rain, ice, andsnow. The building, which measuresapproximately 62 ft deep, 134 ft wide,and 33 ft high, is located approximately80 ft south of the turbine building.Figures X A-40 and X A-41 show two sidesof the building. The large-equipmentaccesses are provided with missileprotecteddoors; the air intakes on thesecond floor are also designed to preventthe entry of tornado-generated missiles.The diesel generator building is supportedby concrete shear walls and H-beam piles which are carried down torock. The elevation of the rock is wellbelow the floor slab and varies as thegeneral terrain falls off toward theriver. Consequently, the depth to whichthe shear walls extend below the floorslab differs significantly at either endof the building. The floor slab isintermediately supported on sets of H-beam piles embedded in the rock.A6.4.6.1.1 Seismic Resistance. Seismicdesign was carried out in accordancewith the rules of the ACI code and iscontained in the design report of Ref.44. A dynamic analysis was performed todetermine the building response to seismicexcitation. The model for thisanalysis is shown in Fig. X A-42.The 'model departs slightly from theconventional equivalent beam model usedby the A/E to represent other buildingstructures. The slab is represented asa rigid beam supported at either end bymassless springs assigned lateral shearstiffnesses based upon the shear area ofthe individual walls at either end ofthe building. This representation wasused to take into account the unequalheights of the shear walls under theslab. In the model the floor slab isassigned near-rigid stiffness sectionalproperties.With these assumptions the lowest fundamentalfrequency was found to be 6.9 cps(north-south direction) and 10.5 cps(east-west direction).Assuming 2% damping for the designearthquake (DE) and 5% damping for themaximum credible earthquake (MCE) andusing the site response spectra, themaximum deflections of the building weredetermined. These were found to occurat the roof and were:a. 0.00262 ft = 0.031 in. (N-S, DE)b. 0.00453 ft = 0.054 in. (N-S, MCE)The most critically loaded members werefound to be rebars at the bottom of thesouth wall. Computed stresses and theircorresponding allowable values are tabulatedbelow:DEMCERebar SeismicStress13,500 psi24,200 psiAllowable20,000 psi36,000 psiThe A/E stated that these results wereconfirmed by another independent analysis.This model was also used to generatefloor response spectra. For this purposethe time-history input (modifiedTaft 1952 record) used for the reactorbuilding was employed.A6.4.6.1.2 Commentary. The assumptions,modeling methods, and computationalprocedures used for this analysisare judged to be appropriate to showcode compliance for this structure.It is concluded that the diesel generatorbuilding is designed with adequatemargin for seismic loads.A6.4.6.1.3 Tornado Resistance. Thediesel generator building was subject totornado loads as described in sectionA6.2.2.A6.4.6.1.3.1 Penetration Calculations.Four types of barriers are used toprotect the generator system:a. monolithic concrete walls: 2-ftthick, reinforced in both faces with0.44 in. 2 /ft steel rods.4X-86

. Steel doors: 2.5-in, thick.C. Double-thickness corrugatedbarriers with a single flatbetween, i.e., three layersgauge high-strength steel.d. 3.5 ft of compacted soil.metalsheetof 12-Kinetic energy per unit area of impactof the 4-in, by 12-in, by 12-ft plank isgreater than that of the auto; hence,through the barriers provided, itgoverns the design against missile penetration.Concrete walls were originally sized toprotect against turbine missiles. At alater time the turbine missile was ruledout because the orientation of the generatorbuilding made the probability ofsuch a missile striking the buildingunlikely. The initially calculated 2-ftthickness of the walls was, however,maintained. Calculations using thePetry formula indicate that a thicknessof 1 ft is required to prevent scabbingby the plank.Calculations of the penetration resistanceof steel door 's, using the Stanfordformula, showed that the required minimumthickness is 1-3/16 in.The resistance of corrugated metalbarriers to penetration was demonstratedby tests carried out by the vendor inconjunction with a testing laboratory.Penetration of the cover of compactedsoil over the fuel tank was not considered.However, since a manifold systemconnecting all tanks is provided, penetrationhere would not represent asafety problem.An approximate calculation made by thereviewer indicates that the plank wouldpenetrate about 8 ft of sand.A6.4.6.1.3.2 overall Response Caculations.The type of failure associatedwith overall response may occur awayfrom the immediate impact area and giverise to collapse of the missile barrierin much the same manner as would beproduced by an excessively large load ofany other origin.The areas considered for penetration,except for the soil cover, are againexamined for the overall response.A6.4.6.1.3.2.1 Concrete Walls. An approximateanalysis based on an estimateof the energy to be absorbed by thestructures was performed. The procedure,as used, appears to be generallyconservative and the adequacy of theconcrete was demonstrated.A6.4.6.1.3.2.2 Steel doors. An overallresponse analysis was performed by avendor and checked by the A/E. Theapproach used was that of the energybalance. Part of the missile kineticenergy was assumed to be absorbed by thedoor, the remaining energy being dissipatedby plastic deformation of themissile. The energy-absorbing capacityof the door was computed by the limitload method. This method gives an upperbound (conservative) estimate for theenergy absorbed by the door. The methodalso neglects the momentum of the plate,which is a conservative assumption, andthus the net result appears conservative.The analysis indicated that thedoors must absorb 50% of the impactenergy.Only the plank missile was considered.However, since the kinetic energies ofthe plank and auto are about the same,the only question is related to theassumption of a 50% reduction in thekinetic energy of the missile due to theplastic behavior of the missile duringthe impact. An examination *of thisquestion indicates that about 95% of thekinetic energy of the plank missile andabout 50% of the kinetic energy of thecar missile are dissipated through theplastic behavior of the missile at theimpact. Therefore, it appears that theanalysis made was reasonable for theauto and conservative for the plankmissile.No calculations were made to demonstratethe ability of the steel door fastenings(hinges, lugs, etc.) to withstand thedoor rebound force after the initialimpact.A6.4.6.1.3.2.3 Corrugated Metal Barriers.No overall structural responsecalculations were made. It is not clearfrom the vendor's test report how thepanel was supported during the penetrationtest. While of little consequencefor the penetration resistance, themethod of support is important withregard to overall behavior. Forexample, the two braces supporting thetop of the panel are secured to thefloor with only four 3/4-in.-diameterbolts each. An approximate calculationindicates that a strike by the plankmissile near the top of the barrierwould most likely shear these bolts.A6.4.6.1.3.3 Wind Loads. The windloads on concrete walls were dismissedby comparison with seismic loads. Thiscomparison was based on the differentialX-87

pressure induced by 300-mph winds andnot on the 3-psi differential specifiedin the criteria. This, in turn, wasbased on the assumption that adequateventing is provided to prevent realizationof the full 3-psi differential inthe emergency generator building.No calculations were made for the steeldoors. An approximate analysis indicatesthat these doors would have to besubjected to a pressure well in excessof the full 3-psi differential specifiedbefore collapse would occur. However,this load acts outward so that fulldifferential pressure load would bereacted by the hinges, lugs, etc., forwhich an analysis has not been done.The criteria regarding the torsionalmoment (see item 4 in section A6.2.2)were of no consequence for a building assmall and as compact as the diesel generatorbuilding.A6.4.6.1.4 Conclusion. We concludethat design criteria for seismic resistanceof the diesel generator structurehave been satisfied.Considering the governing criteria, thisreview indicates that the emergencydiesel generator building structure anddoor adequately resist tornado effects.There is insufficient information toassess the tornado resistance of thefollowing items:a. Corrugated metal barrier attachments.b. Steel door supports, i.e., theiradequacy to resist the differentialpressure and the elastic reboundforces caused by missiles.A6.4.6.2Diesel Generators.Four diesel generator units are provided,each consisting of a dieselengine, a generator (3250 kW at continuousoperation), and the associatedauxiliaries mounted on a common base.Figure X A-43 shows the diesel generatorfrom the generator end, and Fig. X A-44shows the turbo-charger.The seismic analysis for the dieselgenerators was performed by the vendor.Although the analysis (Ref. 45) was performedfor units to be installed atanother nuclear power plant and does notform an official part of the documentationof the BWR plant, it has beenreviewed by the A/E and approved assuitable (with minor modification) todemonstrate the seismic integrity of thediesel generators at the BWR plant.Figure X A-45 shows the model used forthe dynamic analysis. The unit, includingtwo turbo-chargers, weighs about 40tons; the mass is distributed to fivelumped-mass points.Two of the masses represent the engineand are placed symmetrically about theengine center of gravity to simulate therotational inertia of the engine. Theengine block, in which the masses areimnbedded, was assumed effectively rigid.The generator rotor and stator are consideredseparate masses and are joinedby a spring simulating the bendingstiffness of the rotor shaft. Theeffective stiffness of the stator supportbeams is also represented. Theturbo-chargers (which are mounted on abracket bolted to the front of theengine) comprise a fifth mass with thestiffness of their support bracketmodeled by the spring between the turbochargersand the engine block. In addition,a flexibility matrix was developedto represent the effects of the skidthat supports the diesel generators.This model is used to determine theresponse of the unit to transverse horizontalexcitations. Under vertical accelerations,the generator stator can beexcited in a rocking motion, fore andaft. The model for vertical excitationis, therefore, slightly modified todetect and display such motions.With this model, modal frequenciesfound by computer to be:Mode123wereNatural Frequencies (czps)HorizontalTransverse9.6313.5238.60vertical20.2145.44Examination of the displacement amplitudesshows that the turbo-charger unitsare the only masses exhibiting significantresponse to the lowest naturalfrequency (9.63 cps). Stresses in theturbo-charger supporting bracket and itsbolts were therefore computed. In thiscomputation, modal forces were combinedon an absolute basis (conservative practice),and vertical and horizontalreactions were assumed to occur simultaneously.The bracket is retained by eight bolts.Although these are only 5/8-in, bolts44X-88

(about minimum engineering practice forheavy equipment), the most severelystressed bolt is under 19,400 psi tensionvs. a proof load equivalent to52,000 psi. The most critical sectionof the bracket was also examined andfound to have a maximum principal stressof 19,400 psi vs. a material yieldstrength of 33,000 psi.Stator-rotor clearances under seismicexcitations were also checked and foundadequate.A6.4.6.2.1 Commentary. The modelingappears to have been carefully considered,the computations are clearly presented,and the analysis is judged to beof high quality.For the BWR installation, the dieselgenerator building floor response spectum(elevation 125-ft, 0-in.) indicatesthat a slightly greater acceleration isappropriate for the safe shutdown earthquakethan was used in the analysisreviewed - 0162g at the BWR site vs.0.525g used in the analysis. However,considering the conservatism in theanalytical procedures and the margin ofsafety shown in the calculations, theanalysis is regarded as satisfactoryevidence that the design also complieswith the structural requirements for theBWR application.A6.4.6.2.2 Conclusion. We judge theseismic design adequacy of the dieselgenerator units to have been welldemonstrated.A6.4.6.3Batteries and Battery Racks.Station batteries are provided as areliable source of control power forspecific engineered safeguards and forother functions required when a-c poweris not available.The battery racks were furnished by thebattery vendor. Racks come in 6-, 12-,and 17-ft lengths. General constructionis in the form of a two-step two-levelframe as shown in Figs. X A-46 andX A-47. Batteries at each of the twolevels rest on a pair of longitudinalbottom rails attached to the frame.Diagonal cross-brancing of the frame isprovided in the longitudinal directionby use of cross ties between every otherpair of rear uprights (i.e., the generalscheme is to cross-brace one bay, leavethe next unbraced, cross-brace thethird, and so on). No other bracing isused. Appended to the frame are siderails used to fence-in the batteries sothey will not slide. These rails aresupported by L-shaped brackets that boltto the bottom rails on which thebatteries rest.The general configurations for theseracks are shown on vendor drawings inRef. 46 (12- and 17-ft racks) and Ref.47.The battery vendor also supplied aseismic analysis for the battery racksto the A/E's specification (Ref. 48).This specification (written before theappropriate floor response spectra wereavailable) required that the batteryrack design satisfy both of the followingseismic conditions:a. 0.lg horizontal and 0.033g verticalaccelerations with the resultingstressed combined with other applicablestresses not to exceed normaldesign stresses.b. 0.24g horizontal and 0.08g verticalaccelerations with the resultingstresses combined with other applicablestresses not to exceed 90% ofthe yield point of the materials.Subsequent to the original purchaseorder for the batteries and their racks,the A/E introduced a general specificationgoverning the seismic design of allClass I equipment, instrumentation, andcomponents. This document (Ref. 30) requiresthat Class I components be dynamicallymodeled as a multi-degree-offreedomlumped-mass system and thattheir natural frequencies and modeshapes be determined.Furthermore, if any of the natural frequenciesso calculated are such thatthey fall within the range in whichamplification of the floor accelerationsare predicted by the floor responsespectra, a dynamic analysis must beperformed. If the lowest natural frequencyexceeds the range for whichamplification of the floor accelerationis predicted, a static analysis may bemade.Consequently, the seismic analysis ofthe battery racks was reexamined usingdynamic methods of analysis, as requiredby this specification. The analysis wascarried out by the A/E and is shown inthe A/E's design report.A six-bar two-mass model, as shown inFig. X A-48, was used for this analysis.The model represents an end-on view ofthe frame and emplaced batteries and wasselected because the lowest natural frequenciesfor the structure are expectedto develop when seismic excitationX-89

occurs in a direction paralleldepth of the structure. Theassumes that the batteriesconsidered rigidly attachedframe.to themodelingcan beto theThe response of the battery racks tovertical seismic excitation was alsoexamined. This analysis employed thesame dynamic models as the horizontalanalysis except that battery masses ateach level were distributed to the nodesat the end of vertical bars, i.e., wereassigned to nodes 1, 3, 4 and 6.The computations for the dynamic analysisfound in Ref. 32 were originallycarried out by hand. The A/E statedthat this analysis was subsequentlychecked by running thýe problem on thecomputer.major results of the dynamic analysisfor the battery racks are shown in TableX A-32.As can be seen from the table, resultsof the dynamic analysis show that thecomputed maximum accelerations exceedthe equivalent g loads originally specifiedfor battery rack design.To check design adequacy, the A/E reviewedthe vendor's stress report andapplied correction factors to the originallycomputed stresses. These factorswere taken as the ratio of theaccelerations for the MCE, as shown inTable X A-32, to those originallyspecified.The A/E stated that the stresses sodetermined fall within allowable limits.A6.4.6.3.1 Commentary. The A/E's procedurefor determining the dynamicloading of the battery racks appears tobe proper. The agreement between independenthand computations and computerresults lends credence to the resultsobtained.The stress correction scheme includes aconservatism, since the ratio used forthe correction factor to be applied toMCE stresses does not take credit foradditional damping customarily allowedin the case of the MCE. (That is, thestress ratios are based on 2% damping,whereas credit could have been taken for5% damping. Substantial reductions instress levels would then be expected ifthe higher damping ratio had beenassumed.)No qualification program was conductedfor the batteries, and therefore arigorous assessment of their adequacy towithstand a severe seismic disturbancecannot be made. Based on the successfuluse of batteries of similar design inautomotive, truck, and submarine applications,where they are subjected tosevere shocks and vibrations for longperiods of time, it can be expected thatthey will probably perform satisfactorilyfollowing an earthquake. Without theresults of a formal qualification program,however, we cannot make a moredefinitive statement.A6.4.6.3.2 Conclusion. we consider thebatteries adequate on the basis of engineeringjudgment. The battery rackssatisfy the design criteria on the basisof the analysis.A6.4.7ELECTRIC POWER DISTRIBUTIONSYSTEMSThe electric power distribution systemsconsist of transformers switchgear andmotor control centers, emergency buses,containment penetrations, and associatedelectric power and control cables andtheir supporting cable trays. For thisplant all 13- and 4-ky cables are installedin conduits. Cables of 440-Vpower circuits and 125-V control circuitsare installed in conduits andmetal trays (see Fig. X A-49).A6.4.7.1 Electric Cables, Terminations,and Connectors.Because the electrical cables are installedboth within and outside theprimary containment, they will berequired to function under severe environmentalconditions should an accidentoccur in the plant. The same type ofcable was used in both locations toeliminate the possibility that cablesdesigned to withstand only the lesssevere conditions outside the containmentmight be inadvertently installedwithin the containment.A6.4.7.1.1 Commentary. A qualificationtest program for the 480-V electricalcables was conducted in March 1970(Refs. 49 and 50). Since that time, thequalification test requirements havebeen made considerably more severe. Thecables tested were exposed to 200 megaradsof gamma radiation and were given a7-1/2-day steam/chemical exposure (including12 hr at 298'F and 7 days at160 0 F). Today, in recognition of thefact that it is possible that temperatureswithin the containment could reach340OF under the small-break LOCA, thetemperature profile suggested for thesteam/chemical exposure includes dwellsX- 90

of several hours at 340°F and a totalexposure of 30 days, with the temperaturenever going below 200 0 F. Theexposure to LOCA conditions must bepreceded by an aging exposure to inducewhatever degradation may occur duringthe installed life of the cables priorto the postulated accident. Similarly,the IPCEA horizontal flame test specifiedfor these cables has been replacedby a more severe vertical flame test.Although the electrical performance ofthe cables was satisfactory, it is notpossible to predict whether they wouldqualify under the current requirements.Documentation of a qualification programfor the 5-kV cables was not available tothe reviewers; however, this omission isnot regarded as serious because (1) thecables within the primary containment donot serve safety-related functions and(2) the safety-related cables outsidethe containment are installed withinrigid metal conduits.Because the cables are not structuralmembers and are supported by cabletrays, seismic qualification is not explicitlyrequired. In electrical systemscritical to safety, redundantcables are provided. Cable redundancysubstantially reduces the probability ofsystem inoperability due to cable failure.In principle, there is concernthat cable insulation may becomeembrittled as a result of aging (exposuresto temperature and nuclear radiationduring their installed life) andthat the insulation may therefore besubject to failure by cracking during asevere seismic event. The current qualificationphilosophy is to cover thispossibility by requiring a bend test(wrapping the cable specimens beingtested around a mandrel) after the agingexposure and before the LOCA exposure.The qualification test that was performedincluded a total nuclear radiationexposure in excess of the totalexposure to which the cables would besubjected in the plant, considering bothnormal operating and accident conditions.There is no evidence that eitherthermal aging or a bend test wasperformed, however.No qualification program was conductedfor terminations and connectors; therefore,an evaluation of design adequacycannot be made.A6.4.7.1.2 Conclusion. The informationavailable is insufficient for us to concludethat the cables, connectors, andterminations will perform properly underall conceivable accident conditions.There is reasonable assurance, however,that the electric cables will performadequately for at least the first fewweeks following a "large break" LOCA.A6.4.7.2Electrical ContainmentPenetrations.The electrical containment penetrationsare components that provide pressurewithstanding,leak-tight, electricallyinsulated paths through the wall of theprimary containment structure for thepurpose of transmitting electrical signalsand electric power to motors andsolenoids. The penetrations are providedwith both inner and outerinsulating seals made of an epoxymaterial. No qualification program wasconducted for these components, althoughthe A/E did perform a brief analysiswhich indicated that, although thestrength of the inner seal may becomemarginal as a result of the cumulativeeffects of nuclear irradiation severalmonths after a major LOCA, the outerseal would maintain its integrity.A6.4.7.2.1 Commentary. Although noseismic or thermal testing or analysiswas performed, our experience indicatesthat this should present no problems.Other penetration units have been testedextensively without mechanical or electricalfailures occurring.A6.4.7.2.2 Conclusion. We conclude onthe basis of engineering judgment thatthe penetrations are adequate for expectedseismic, normal and LOCA environments.A6.4.7.34-kV Switchgear.The 4-kV switchgear units control theconnection of electric power sources(either of the two off-site supplies or,should these be unavailable, the dieselgenerator) to the four 4-kV emergencyswitchgear buses. These units alsosupply power to safety-related motorslarger than 200 hp. One of the switchgearunits is shown in Fig. X A-50. Arow of switchgear cabinets is shown inFig. X A-51. This equipment is segregatedfrom non-safety-related electricalequipment in a separate room located atelevation 135 ft near the center of theplant buildings.Because the switchgear cabinets arelocated in a shielded room that isisolated from nonelectrical equipment,this equipment is not expected to besubjected to the nuclear radiation,steam, or temperature environments thatcould result in the event of a high-X-91

energy pipe break outside the primarycontainment. Any steam that may seepinto the room around the door was expectedto be removed by the roomventilation system. Therefore, the onlysignificant design basis event for thisequipment was the safe shutdown earthquake.For newer plants AEC Regulatory Guide1.70 (Rev. 1, Oct 1972) in section3.11.4 entitled "Loss of Ventilation"requires that the base should be providedwhich assures that loss of airconditioning and/or ventilation systemwill not adversely affect the operabilityof safety related control andelectrical equipment located in thecontrol room and other areas.A seismic analysis of the 4-kV switchgearwas performed by the vendor andreported in Ref. 51. We note that theinformation contained in Ref. 51 reportsthe results of analysis and test and theinspection result of a switchgear likethe one installed in the BWR which hadbeen exposed to an actual earthquakeenvironment. Maximum calculatedstresses are 9379 psi in the jackingscrew, compared to an AISC allowable of20,000 psi. A fundamental frequency of19 cps was computed, compared to 20 cpsdetermined by test. Functional adequacywas verified by test. The inspection ofswitchgear that had experienced anactual earthquake revealed no deficiencies.A6.4.7.3.1 Commentary. The data describedabove offer convincing evidencethat the switchgear will resist seismicenvironments.A6.4.7.3.2 Conclusion. Thethe switchgear is adequateexpected seismic conditions.A6.4.7.4480-V Load Centers.design offor theThe 500-kVA 480-V emergency auxiliaryload centers consist of a combination ofa transformer, which reduces the voltagefrom 4-kV to 480-V, and a group of 480-kV switchgear units. One of the twoload centers is shown in Fig. X A-52.Each load center is located at a partialenclosure in the reactor building atelevation 165 ft.A6.4.7.4.1 Commentary. Although we understandtype qualification data is nowgenerally available for such equipmentfrom manufacturers, no qualificationdata has been submitted for our reviewand apparently does not exist in thedocumentary record of this particularplant.A6.4.7.4.2 Conclusion. We cannotan assessment of design adequacy.A6.4.7.5480-V Motor Control Centers.makeA typical rack containing several 480-Vmotor control centers (MCC) is illustratedin Fig. X A-53. Each MCC controlsa component, such as a valve motoroperator. This equipment is situated atseveral locations throughout the plantin the vicinity of the motor beingcontrolled. A seismic test program of a480-V motor control center is reportedin Ref. 52.A6.4.7.5.1 Commentary. In the seismictest program, a steel frame structurewith angle braces was used to simulate"attachment of the cabinet top to abuilding wall or a rigid floor-mountedstructure, and was used as top bracingin all three axes." This limits thevalidity of the test results to applicationswhere the cabinet is similarlysupported at the top, such as at the BWRplant.The frequency range of the tests was 2to 35 cps. While it is generally desirableto extend the range down to 1 cps,for this plant application the 2-cpslower limit is probably adequate becauselower frequencies are highly attenuatedby the building structure.Functional performance was evaluated by"monitor circuits that detected interruptionsof 1 millisecond or longer inthe open and closed circuits." Theadequacy of this evaluation procedure isnot documented. Therefore, lackingother information, we must comment thatthe maximum acceleration levels forsatisfactory operation listed in thereport are applicable only if interruptionsof less than 1 msec are acceptable.A 2-pole switch that intermittentlyfailed to make contact in the closedposition was replaced with a new switch,which functioned correctly. This practiceis highly questionable. In theabsence of any justification for thereplacement, it must be concluded that aswitch of the same type could fail ifsubjected to a seismic disturbance.The report states that "snap-on terminalstrips were modified ... to keep themfrom falling off during seismic evaluation"and that a "plastic insulationcovering the buss bars ... was replaced... with a stronger insulation material.... " The test results are thereforeapplicable only to motor control centerswith equivalent modifications.X-92

During front-to-back vibration, thecabinet doors swung open. The doorlatches were adjusted in an unsuccessfuleffort to prevent this. It must beassumed that the doors could swing openduring a seismic disturbance. Althoughnot a likely occurrence, it is possiblethat the impacts of the doors againstadjacent cabinets might cause a malfunction.Interpretation of the test results wascomplicated by resonances that occurredin the test apparatus within the frequencyrange of the test. For example,this made it impractical to determinethe damping characteristics of thecabinet in the front-to-back axis.A6.4.7.5.2 Conclusion. On the basis ofthe information in the report reviewed,and assuming that the deficienciesidentified in the test program werecorrected in the units installed in theplant, the seismic design of the 480-VMCC is judged to be adequate for theirintended application in this plant.A6.4.7.6D-C Distribution Panels andFuse Boxes.The d-c distribution panels (Fig. X A-54) and fuse boxes are located outsidethe primary containment in a separateroom and hence are not subject todeleterious nuclear-radiation, steam, ortemperature environments. The smallquantity of steam which may seep intothe room was expected to be removed bythe room ventilation system. Therefore,the only significant design basis eventfor this equipment item was the safeshutdown earthquake. For newer plantsAEC Regulatory Guide 1.70 (Rev. 1, Oct.1972), Section 3.11.4 would, in addition,require a demonstration that lossof ventilation would not adversely affectthe operability of this equipment.The manufacturer of the equipment conducteda qualification testing programand reported the results in Ref. 53.The switchboard and components weretested on a machine that permittedsimultaneous (in-phase) vibration in thehorizontal and vertical directions. Theequipment was vibrated in two perpendicularvertical planes. The test programincluded resonance searches, followed bycontinuous sine, sine-beat, and randomvibration.No malfunctions were observed except ontwo relays. The report implies that therelays are not necessarily to be consideredunqualified to withstand seismicdisturbances because of the severity ofthe test, provided certain precautionsand protective measures are taken. Therelays that malfunctioned during thevibration test reported in Ref. 53 wereremoved from the panelboard and placedin their own enclosure and are notincluded in the seismic qualification.A6.4.7.6.1 Commentary. From the referencedqualification test reports andfrom conversations with their principalauthor, we know that the supplierregards the individual components (relays,switches) of the control boards tobe essentially single-degree-of-freedomsystems for which single-frequency testingis adequate. The components doappear to' be single-degree-of-freedomdevices, based on physical inspection.Therefore, *the tests conducted on thesecomponents satisfy the current regulatoryrequirements in this regard. However,the assemblies were not rotated1800 for a repeat of the test in a givenplant as currently required. Althoughit is possible that the omission of suchrepeat tests can be justified analyticallyfor the assemblies tested, no suchanalysis was offered because the testswere conducted prior to the establishmentof the current AEC regulatoryposition.A6.4.7.6.2 Conclusion. The d-c distributionpanels and fuse boxes satisfy thedesign criteria and are considered adequate.A6.4.7.7Cable Trays.The cable trays for the BWR were designedto act like a low frequencypendulum so that they would be subjectto seismic spectral accelerations whichare low.The cable tray hanger attachment to theoverhead structure is either a swiveltype of fixed, depending on the lengthof the support and the number of traysthat are hung from the support, as shownin Table X A-33.A6.4.7.7.1 Commentary. The design approachdescribed above seems to be areasonable method of minimizing theseismic loads. Very low frequencystructures, however, are subject tolarge seismically induced displacements.Sufficient cable slack is normally providedto allow for tray movement withoutcable damage. A large swing of thecable trays might possibly cause impactwith adjacent structures. Adequacy ofcable slack and the possibility ofimpact could be easily checked by visualinspection and by simply pushing thetrays laterally in critical areas andobserving the results: such checks wereX-93

made by the FIP.L review team at thesite.A6.4.7.7. 2 Conclusion. The cable traydesign appears to be adequate for seismicloads.A6.4.8REACTOR PROTECTION SYSTEMThe function of the reactor protectionsystem (RP1S) is to (1) accurately andreliably determine whether conditions inthe plant are such that the reactorshould be immediately shut down, thecontainment isolated, and the emergencycore cooling system (ECCS) activated and(2) if so, to initiate these actions(the ECCS is to be initiated within0.050 sec) . The RPS can also be manuallyinitiated by means of scram pushbuttonsin the control room.The only-portions of the RPS that wereconsidered in the design adequacy reviewwere:a. Pressure transmitters.b. Pressure switches.c. Monitoring racks.d6. Electrical cables. 1e .f .Logic cabinets containing relays.Control room instrumentationand controls.Instrumentation piping is connected tothe reactor vessel and is routed througha special penetration in the drywellwall and isolation valves to instrumentationracks (see Figs. X A-55 and XA-56) located at elevation 165 ft in thereactor building. These racks containthe pressure transmitters that transformhydraulic pressure signal inputs intoelectrical signals proportional to reactorpressure and reactor water level andpressure switches that initiate an electricalsignal when certain preselectedconditions are reached. other pipelines, transmitters, and switches similarlyprovide information concerningpressure within the drywell. Four linesand transmitters are provided for eachphysical parameter. Adjacent panelscontaining circuits of redundant systemshaving less than 3-ft separation havesteel barriers installed between panels(Fig. X A-55).1 Covered in section A6.4.7.1.The electrical signals are conveyed byelectrical control cables to instrumentsand to cabinets containing relays (similarto those in Fig. X A-57) which providethe logic functions. These componentsare located in the control room atelevation 165 ft. Other electricalcables carry the signals that initiatereactor shutdown, containment isolation,and ECCS functioning to the control roddrive system and to the various motorcontrol centers that activate valveclosures and motor start-ups. A portionof the control room panels pertaining tothe reactor safety functions is shown inFig. X A-58. These panels contain instrumentsthat provide such data asmotor current, system flow, and pumpdischarge pressure to the operator,lights which inform him of valve positions,and control switches to permitthe manual operation of motors andvalves.A6.4.8.1 Commentary.The RPS components considered in thereview are located outside of the primarycontainment, and hence are not exposedto severe environmental conditionsin the event of a LOCA. The pressuretransmitters and switches installed invarious plants have had a tendency toexperience set-point drift under normalplant conditions. We have not seen anyqualification information that wouldpermit an evaluation of the stability ofthese components with regard to setpointdrift under the normal plant operatingconditions, and the consequencesof such drift as may occur.The seismic qualification was reviewedprimarily on the basis of information ina report (Ref. 53) that summarized qualificationfor instrument racks andpanels and various electronic devices.The testing was conducted in accordancewith the requirements of IEEE 344-1971(Ref. 54), except that there is no indicationthat the units were pre-agedbefore being subjected to the simulatedseismic exposure. The latter involved aresonance scan at 0.2g followed by continuoussinusoidal vibration at resonantfrequencies to determine malfunctionlimits, the procedure being repeatedalong three orthogonal axes. Because oftest facility limitations, the testingdid not cover frequencies below 5 cps.Although conditions below 5 cps wereanalyzed, it is difficult to demonstrateseismic qualification of complex devicesby analysis; and, in the absence of thedetailed calculations, a definitiveevaluation is not possible.X-94

Present seismic qualification requirementsemphasize multi-axis testing andthe use of random vibration. Past, orfuture, qualification testing should bereviewed particularly with respect tomulti-axis effects.A6.4.8.2Conclusion.Qualification of this equipment forseismic exposures in the low frequencyrange could not be evaluated completelywith the information available. Thequalification programs that were conductedto demonstrate that the equipmentdoes function satisfactorily when exposedto severe conditions which can beconsidered as segments in the type ofmore comprehensive qualification programthat is presently regarded as necessaryto conclusively demonstrate design adequacy.Although the design criteria are notcompletely satisfied, the testing thatwas performed indicates that failure isnot expected. The RPS components examinedare adequate.Section A2References1. Hurty and Rubenstein, Dynamics of Structures, Prentice-Hall, Englewood Cliffs,N.J., 1964.2. IEEE Standard 323-1974, "IEEE Standard for Qualifying Class IE Equipment forNuclear Power Generating Stations."Section A4References1. IEEE Standard 323-1974, "IEEE Standard for Qualifying Class IE Equipment forNuclear Power Generating Stations."2. IEEE Standard 279-1971, "Criteria for Protection Systems for Nuclear Power GeneratingStations."3. IEEE Standard 308, "Criteria for Class IE Electric Systems for Nuclear PowerGenerating Stations."4. ANSI N45-3.4 and IEEE 336-1971, "Installation, Inspection, and Testing Requirementsfor Instrumentation and Electric Equipment During the Construction ofNuclear Power Plants."5. IEEE Standard 344-1971, "Guide for Seismic Qualification of Class I ElectricalEquipment for Nuclear Power Generating Stations,"6. IEEE Standard 334/1971, "Guide for Type Tests of Continuous Duty Class I MotorsInstalled Within the Containment of Nuclear Power Generating Stations."7. IEEE Standard 382-1972 (ANSI N41.6), "IEEE Trial-Use Guide for Type Test of ClassIE Electric Valve Operators for Nuclear Power Generating Stations."8. IEEE Standard 383-1974 (Draft), "IEEE Trial-Use Guide for Type Test of Class IEElectric Cables, Field Splices and Connections for Nuclear Power GeneratingStations."X-95

Section A5References1. RDT Standard F9-2T, "Seismic Requirements for Design of Nuclear Power Plants andTest Facilities," Oak Ridge National Laboratory, January 1974.Section A6References1. U.S.A.E.C. Regulatory Guide 1.60, "Design Response Spectra for Seismic Design ofNuclear Power Plants."2. U.S.A.E.C. Regulatory Guide 1.61, "Damping Values for Seismic Design of NuclearPower Plants," Oct. 1973.3. Final Safety Analysis Report for Surry Power Station Units 1 and 2, 1970.4. Vibration in Civil Engineering - Proceedings of Symposium Organized by theBritish National Section of the International Association for Earthquake Engineering- Section V: "The Dynamic Behaviour of Foundations," by Robert V.Whitman, Butterworths, London W.C.-2, England, 1965.5. RDT Standard F9-2T, "Seismic Requirements for Design of Nuclear Power Plants andTest Facilities," Oak Ridge National Laboratory, January 1974.6. Virginia Electric and Power Co., "Seismic Design Review and Piping - Surry PowerStation," June 30, 1971, Rev. Sept. 15, 1971.7. Westinghouse Electric Engineering Memorandum 4528, "Analysis of the 93A CasingNozzles Using Umbrella Loads," January 7, 1974.8. Welding Research Council Bulletin 107, "Local Stresses in Spherical and CylyndricalShells Due to External Loadings," August 1965, Rev. July 1970.9. Letter from ITT-Grinnel Corp. to FIRL dated 4/19/74 with enclosure (Parker SealCo. Report No. K10, 151-466, dated Nov. 30, 1973).10. Stone and Webster Engineering Corporation, Supplement B to Report No. SW-MER-lA,"Sensitized Stainless Steel in Pressurized Water Reactor Applications, SurryPower Station," prepared for Virginia Electric and Power Co., November 30, 1971.11. C. G. Collins, et al., "Estimated Radiation Stability of Aircraft Components,"APEX 357, General Electric (no date available).12. Crane Packing Company Bulletin No. 3472.13. Crane Packing Company Bulletin No. 100.1.14. ALOYCO Drawing No. E-45467, April 19, 1968.15. Topical Report, "Environmental Testing of Engineered Safety Features RelatedEquipment," WCAP-7744, Volume I.16. Stone and Webster Engineering Corporation, Piping Flexibility Analysis SummarySheets 1 through 4, Drawings numbered 11448-MSK-122 D1/5, D2, D3/6, D4/1.17. Stone and Webster Engineering Corporation, Piping Flexibility Analysis SummarySheets 1 through 4, Drawings numbered 11448-MSK-122 Al/5, A2/2, A3/5, A4/1.18. U.S.A.E.C. Safety Guide 26, March 23, 1972.X-96

19. Westinghouse Electric Corporation Letter with Attachments to Franklin InstituteResearch Laboratories, dated May 23, 1975 (Paddleford to Sailwell).20. M.W. Sheets, "Topical Report on GE Vertical Induction Motors Inside ContainmentRecirculation Spray Pump Motors - Surry Power Station," General Electric, June12, 1973.21. Topical Report, "Environmental Testing of Engineered Safety Features RelatedEquipment," WCAP-7744, Volume II.22. IEEE Std. 383-1974, "IEEE Standard for Type Test of Class IE Electric Cables,Field Splices, and Connections for Nuclear Power Generating Station."23. Topical Report, "Seismic Testing of Electrical and Control Equipment," WCAP-7817and Supplements 1 through 3, December 1971.24. L. M. Potochnik, "Seismic Testing of Electrical and Control Equipment (HighSeismic Plants)," WCAP-7821 and Supplements 1 through 4, Westinghouse ElectricCorporation, October 1971.25. L. M. Potochnik, "Seismic Testing of Electrical and Control Equipment (PG&EPlants)," Westinghouse Electric Corporation, May 1973.26. A. Morrone, "Seismic Vibration Testing with Sine Beats," WCAP-7558, WestinghouseElectric Corporation, October 1971.27. E. G. Fischer and S. J. Jarecki, "Qualification of Westinghouse Seismic TestingProcedure for Electrical Eqiupment Tested Prior to May 1974." WestinghouseElectric Corporation, August 1974.28. Bechtel Drawing LF-240-121.29. Bechtel Specification 6280-M-120.30. Bechtel Specification 6280-G-14 (Rev. 1), "General Project Class I Sesimic Requirementsfor Equipment, Instrumentation, Systems and Components for PeachBottom Atomic Power Station Units 2 and 3 for the Philadelphia ElectricCompany," May 1973.31. Final Safety Analysis Report for Peach Bottom Power Station Units 2 and 3, 1971.32. Chicago Bridge and Iron Co., Report No. 6, "Stress Analysis of MiscellaneousNozzles, Peach Bottom II," Prepared for General Electric Co., Nuclear EnergyDivision.33. Chicago Bridge and Iron Company, Report No. 7, "Piping Reactions for PeachBottom II," Prepared for General Electric Co., Nuclear Energy Division.34. Chicago Bridge and Iron Co., Report No. 8, "Stress Analysis of Support Skirt forGeneral Electric Co., Nuclear Energy Division, Peach Bottom II."35. General Electric Company, "Sesimic Analysis for Atwood and Morrill 26 'Size y'Type Main Isolation Valve, Ref. Dwg. 20851-H," Document No. 2258-44-1, March 26,1970.36. General Electric Company, "Seismic Analysis for Atwood and Morrill 26 'Y' TypeMSIV, Ref. Dwg. 21150-H," Document No. 296418-2, October 11, 1973.37. General Electric Company, "Design and Performance of G.E.-BWR-MSIV," Report No.APED-5750, by Rockwell and Zylstra, March 1969.38. P. J. Thiemann, "Environmental Qualification Test of Vertical Induction Motorfor ECCS Service in Nuclear Power Plants," NEDM-10672, 72NED93, General ElectricCompany, San Jose, California, August 1972.X-97

39. Philadelphia Gear Corp., "Test of Limitorque Valve Operator to Meet GeneralRequirements of an Electric Valve Actuator in Nuclear Reactor ContainmentEnvirorunent," Jan. 2, 1969.40. Addendum #1 dated April 29, 1969 to preceding report.41. Keith, Feibusch Associates, Engineers, "Seismic Capability of HPC1 Turbines (CSand CCS)," Document No. 2763-43-1, prepared for General Electric Company, April30, 1970.42. General Electric Company, "Design Calculations for RHR Pump," Document 21A5790A5(Rev. 5), 1969.43. Keith, Feibusch Associates, Engineers, "Seismic Capability of the Residual HeatRemoval Pump," prepared for General Electric Company, San Jose, California, July31, 1970.44. Bechtel Report on Diesel Generator Building Structure, April 11, 1971.45. Development Engineering Department, Fairbanks Morse Power Systems Division, "ASeismic Analysis for the Diesel Generators to be Installed at Baltimore Gas &Electric's Calvert Cliffs Nuclear Power Station," March 1971.46. Exide Drawing No. PH27004/08, Rev. 0, 2/27/70.47. Exide Drawing No. PH27006, Rev. 0, 2/27/70.48. Bechtel Specification 6280-E-13, "Specification for Batteries, Battery Racks andChargers for the Peach Bottom Atomic Power Station Units 2 and 3 for thePhiladelphia Electric Company," July 1969.49. L. E. Witcher and W. H. Steigelmann, "Test of Electrical Cables Under SimulatedPost-Accident Reactor Containment Service," Addendum to FIRL Final Report F-C2750, Franklin Institute Research Laboratories, June 1970.50. J. G. Stone and L. S. Chapman, Supplement to Ref. 48, Technical Report No. NP-03, Cerro Wire & Cable Co., July 1970.51. General Electric Requisition 320-67245, "Seismic Analysis of 5kV M26Switchgear," Millstone Point Nuclear Unit 2, March 1972.52. D. W. Caba and R. H. Prause, "Seismic Evaluation of an Old Style Unitrol MotorControl Center with G. E. Components and Top Support," Battelle ColumbusLaboratories, Dec. 22, 1972.53. L. D. Test, "Seismic Qualification of Class 1 Electric Equipment," GeneralElectric Licensing Topical Report NEDO-10678, Nov. 1972.54. IEEE Standard 344-1971, "Guide for Seismic Qualification of Class I ElectricalEquipment for Nuclear Power Generating Stations."X-98

TABLE X A-1 DESIGN BASIS ENVIRONMENTS - PWR PLANTRad. Dose Intergrated(a)Temp. Press. Rel. Hum. Rate Rad. DoseCondition and Location (°F) (psig) (%) (rad/hr) (rad) DurationI. Normal (No Accident):A. Inside Containment 80-105 -4 40-100 55 20 x 106 40 yrB. Outside Containment(b) 60-120 %0 40-90 0.001-0.015 2 40 yrII. LOCA (Safety-relatedsystems functioningexcept for worst singlefailure):A. Inside Containment(c) Incr. to 2B0 Incr. to 45 100 7.2 x 10 -- 10 sec280 45 100 6.2 x 106 3.3 x 106 Next 30 minDecr. to 150 Decr. to -1 100 5.8 x 106 6.3 x 106 Next 30 min150 -1 to -2 Decr. to 50 Decr. to 0.1 1.5 x 108 Next 100 days (d)1.4 x 105B. Outside Containment Incr. to 212 0 100 0.001-0.3 -- 30 sec(Localized)(b) 212 0 100 0.001-0.3 0.006-0.02 Next 6 hr150 0 Decr. to 50 0.001-0.2 2-50 Next 100 days(d)(a)(b)(c)(d)Based upon extremely conservative assumptions; actual expected dose rates and doses about 100 timessmaller.Conditions may be higher in closed cubicles and adjacent to piping and equipment containing radioactivematerials.Plus a continuous spray vertically downward with a recirculating aqueous solution containing 3,000 ppmboron adjusted to a pH of 8.0 with sodium hydroxide, at a rate of 0.15 (gal/min)/(ft 2 in a horizontalplane) for a period of 7 days. Effectiveness of spray in decreasing radiation dose rate (approximatelya factor of 2 reduction) is not included.Assumes that accident conditions have essentially disappeared after 100 days.

TABLE X A-2 DESIGN BASIS ENVIRONMENTS - BWR PLANTRad. Dose Integrated(a)Temp. Press. Rel. Hum. Rate Rad. DoseCondition and Location (OF) (psig) (%) (rad/hr) (rad) DurationI. Normal (No Accident):A. Inside Primary 135-150 %o 40-100 25 7 x 106 40 yrContainmentB. Outside Primary 60-120 %0 40-90 0.001-0.015 250-3000 40 yrContainment (b-UII. LOCA (Safety-relatedsystems functioningexcept for worst singlefailure):A. Inside Drywell(c) Incr. to 280 Incr. to 62 100 3.3 x 106 9000 10 secIncr. to 3 4 0 (d) 6 2 (d) 100 3.3 106 0.3 106 Next 5 min340(d) 35 (d) 100 2.3 106 7.5 x 106 Next hr32 0 (d) 3 5 (d) 100 1.1 x 106 12 x 106 Next 3 hr250(d) 25(d) 100 0.3 x 106 18 x 106 Next 18 hr2 0 0 (d) 2 0 (d) 100 8 x 104 25 x 106 Next 3 daysDecr. to 110 Decr. to 0 Decr. to 50 Decr. to 200 32 x 106 Next 96 days(e)B. Inside Supres- Incr. to 150 Incr. to 25 100 0.12 x 10 6 -- 30 secsion Chamber 150 25 100 1.0 x 106 4 x 105 Next 20 minIncr. to 200 Decr. to 15 100 0.3 x 106 3 x 106 Next 10 hr200 Decr. to 10 Decr. to 50 Decr. to 2000 100 x 106 Next 100 daysC. Outside PrimContainmentl Incr. to 212 '0 100 0.001-0.03 -- 30 sec212 o0 100 0.001-0.03 0.006-0.2 Next 6 hr150 .0 Decr. to 50 0.001-0.02 2-50 Next 100 days(e)(a)(b)(c)(d)(e)Based upon extremely conservative assumptions; actual expectedsmaller.dose rates and doses about 100 timesConditions may be higher in closed cubicles and adjacent to piping and equipment containing radioactivematerial.Plus a continuous spray vertically downward with demineralized water at a rate of 0.15 (gal/min) (ftin a horizontal plane) for a period of 24 hr.These temperatures and pressure conditions would not occur simultaneously.Assumes that accident conditions have essentially disappeared after 100 days.2Table X A-I - Table X A-2X-99/100

TABLE X A-3PWR COMPONENTSA 6.3.1A 6.3.1.1A 6.3.1.2A 6.3.1.3A 6.3.1.4A 6.3.2A 6.3.2.1A 6.3.2.2A 6.3.2.3A 6.3.2.4A 6.3.2.5A 6.3.3A 6.3.3.1A 6.3.3.2A 6.3.3.3A 6.3.3.4A 6.3.3.5A 6.3.3.6A 6.3.4A 6.3.4.1A 6.3.4.2A 6.3.4.3A 6.3.5A 6.3.5.1A 6.3.5.2A 6.3.6A 6.3.6.1A 6.3.6.2A, 6.3.6.3A 6.3.6.4A 6.3.6.5A 6.3.7A 6.3.7.1A 6.3.7.2A 6.3.7.3A 6.3.7.4A 6.3.8A 6.3.9Reactor BuildingSoil-Structure Interaction ModelContainment Internal StructurePaint Coatings Within ContainmentCraneReactor Coolant System (RCS)Reactor Coolant System LoopsSteam Generator and Pump SupportsReactor Coolant Pump NozzlesPipe-Whip RestraintsSnubbersLow Head Safety Injection System (LHSIS)PipingPumps and DrivesValvesValve Motor OperatorsSnubbers and HangersInstrumentationHigh Head Safety Injection System (HHSIS)Accumulator Tank NozzleAccumulator Piping Connection to RCSCharging Pumps and DrivesContainment Recirculation Spray SystemPump and Motor Located Outside ContainmentMotor Drives Inside ContainmentOn-Site Electric Power SystemDiesel Generator HousingDiesel GeneratorsDay TanksAir Bottle SupportsBatteries and Battery SupportsElectric Power Distribution SystemsElectrical Cables and TerminationsElectrical Containment Penetrations and ConnectorsCable TraysAC and DC SwitchgearReactor and Engineered Safeguards Protection Systems Sensors andLogic CabinetsIntake Canal

TABLE X A-.4 BWR COMPONENTSA 6.4.1A 6.4.1.1A 6.4.1.2A 6.4.1.3A 6.4.1.4A 6.4.1.5A 6.4.1.6A 6.4.2A 6.4.2.1A 6.4.2.2A 6.4.2.3A 6.4.2.4A 6.4.2.5A 6.4.2.6A 6.4.2.7A 6.4.2.8A 6.4.3A 6.4.3.1A 6.4.3.2A 6.4.3.3A 6.4.3.4A 6.4.3.5A 6.4.3.6Containment StructuresPrimary Containment StructureReactor Building (Secondary Containment)Containment Piping PenetrationsSuppression-Chamber-to-Drywell Vacuum Breaker ValvesMissile Barriers Inside ContainmentPaint Coatings Within ContainmentReactor Coolant SystemReactor Pressure Vessel and InternalsReactor Pressure Vessel NozzlesReactor Pressure Vessel Support SkirtRecirculation LinesMain Steam LinesMain Steam Line Isolation ValvesPipe-Whip Restraints for Recirculation LinesReactor Pressure Relief ValvesCore Spray SystemPipingHangers and SnubbersPumps and DrivesValvesValve Motor OperatorsInstrumentation1WA 6.4.4 High Pressure Coolant Injection System (HPCIS) TurbineA 6.4.5 Residual Heat Removal (RHR) Pumps and DrivesA 6.4.6A 6.4.6.1A 6.4.6.2A 6.4.6.3A 6.4.7A 6.4.7.1A 6.4.7.2A 6.4.7.3A 6.4.7.4A 6.4.7.5A 6.4.7.6A 6.4.7.7A 6.4.8On-Site Electric Power SystemsDiesel Generator BuildingDiesel GeneratorsBatteries and Battery RacksElectric Power Distribution SystemsElectrical Cables, Terminations, and ConnectorsElectrical Containment Penetrations4-kV Switchgear480-V Load Centers480-V Motor Control CentersDC Distribution Panels and Fuse BoxesCable TraysReactor Protection SystemaTable X A-3 - Table X A-4X-101/102

61OLTABLE X A-5 PWR COMPONENT REVIEW - SEISMIC(a) DESIGN ADEQUACY SUMMARYReview Data Basis0 o W Design Criteria1 4 Satisfied(b)Adequate (Failure Not Expected)Assessment of Design AdequacyNot Demonstrated (Failure Possible)M 4 P Analysis Design Criteria Insufficient Information ReviewedSubsection30 Usubsectin s -1U and/or Engineering Not Completely Information to Indicates thatin Text Component Description > 0 0 0 Test Judgment Satisfied(c) Assess Adequacy Failure is Likely NotesA6.3.1A6.3.1.1A6.3.1.2A6.3.1.4A6.3.2A6.3.2.1A6.3.2.2A6.3.2.3A6.3.2.4A6.3.2.5A6.3.3A6.3.3.1A6.3.3.2A6.3.3.3A6.3.3.4A6.3.3.5A6.3.3.6A6.3.4A6.3.4.1A6.3.4.2A6.3.4.3A6.3.5A6.3.5.1A6.3.5.2A6.3.6A6.3.6.1A6.3.6.2A6.3.6.3A6.3.6.4A6.3.6.5A6.3.7A6.3.7.2A6.3.7.3k6.3.7.4&6.3.8A6.3.9Reactor BuildingSoil-Structure Interaction ModelContainment Internal StructureCraneReactor Coolant System (RCS)Reactor Coolant System LoopSteam Generator and Pump SupportsReactor Coolant Pump NozzlesPipe Whip Restraints:Main Steam LineFeedwater LineSnubbersLow Head Safety Injection System (LHSIS)PipingPumps and DrivesValvesMotor OperatorsSnubbers and HangersInstrumentationHigh Head Safety Injection Systems (EHSIS)Accumulator Tank NozzleAccumulator Piping Connection to RCSCharging Pumps and DrivesContainment Recirculation Spray SystemPump and Motor Located Outside Containment:PumpMotorMotor Drives Inside ContainmentOn-Site Electric Power SystemsDiesel Generator Housing: (a)Walls and RoofDoorSoil Cover for Fuel TanksDiesel GeneratorsDay TanksAir Bottle SupportsBatteries and Battery SupportsElectric Power Distribution SystemsElectrical ContainmentPenetrations and ConnectorsCable TraysAC and DC SwitchgearReactor and Engineered SafeguardsProtection Systems Sensors andLogic CabinetsIntake CanalVI //VIVIVIVIVIVIVIVIVIVIVIVIVIVI VI VIVIVIVIVIVIVIVIVIVIVIVIVIVIVIVIVIVIVIVIVVIVIVIVVIVIVIVIVVIVIVII/VIVI(d)(e)(d)(f)(d)(f)(d)

(a)TABLE X A-5 FOOTNOTESDiesel Generator Building was evaluated only for tornado resistance.(b) Design Criteria as stated in FSAR; or current criteria, if original criteria were not explicitor found to have insufficient margin.(c)More sophisticated analysis methods may in some cases show that the design criteria have infact been satisfied; in all cases however, loss of function is not expected.(d) Deficiencies in original analysis have been either revised or investigated further. Revisedresults have been evaluated and design is found to be adequate.0(e)(f)Bijlaard formulae not applicable.Periodic seal replacement in the snubbers is recommended.Table X A-5X-103/104

k.TABLE X A-6 PWR COMPONENT REVIEW - ENVIRONMENTAL QUALIFICATICReview Data BasisWr. 4J0 0 $4Adequate (Failure Not Expected)Design CriteriaSatisfied(a)Assessment of Design AdequacyNot Demonstrated (Failure Possible)-4 > - Analysis Design Criteria Insufficient Information ReviewedSubsection U) U and/or Engineering Not Completely Information to Indicates thatin Text Component Description A o o 4 ) Test Judgment Satisfied (b) Assess Adequacy Failure is LikelyA6.3.1 Reactor BuildingA6.3.1.3 Paint Coatings within ContainmentA6.3.2 Reactor Coolant System (RCS)A6.3.2.5 Snubbers / /A6.3.3 Low Head Safety Injection System (LHSIS)A6.3.3.2 Pumps and Drives / /A6.3.3.5 Snubbers / /A6.3.4 High Head Safety Injection Systems (HHSIS)A6.3.5 Containment Recirculation Spray SystemA6.3.5.1 Pump Located Outside Containment V /A6.3.5.2 Motor Drives Inside Containment / /A6.3.6 On-Site Electric Power SystemsA6.3.7 Electric Power Distribution SystemA6.3.7.1 Electrical Cables and Terminations VA6.3.7.2 Penetrations and Connectors /A6.3.9 Reactor Protection System / /Notes(c)(d)(c)(d)(d)(a) Design Criteria as stated in FSAR; or current criteria, if original criteria were not explicit or found to have insufficient margin.(b)More sophisticated analysis methods may in some cases show that the design criteria have in fact been satisfied; in all cases however, loss of functionis not expected.(c) Periodic seal replacement in the snubbers is recommended.(d)Active components of this system that are outside of containment and not subject to LOCA environment have not been included.Table X A-6X-105/106

AibTABLE X A-7 BWR COMPONENT REVIEW - SEISMIC(') DESIGN ADEQUACY SUMMARYReview Data BasisAdequate (Failure Not Expected)• o • U~ Design Satisfied(b} CriteriaAssessment of Design AdequacyNot Demonstrated (Failure Possible).6 . H Analysis Design Criteria Insufficient Information ReviewedSubsection W . tq and/or Engineering Not Completely Information toin Text Component Description > o U" P" Test Judgment Satisfied(c) Assess Adequacy Failure Is Likely NotesA6.4.1 Containment StructuresA6.4.1.1 Primary Containment Structure / /A6-4.1.2 Reactor Building (Secondary Containment) V //A6.4.1.3 Containment Piping Penetrations / /A6.4.1.4 Suppression-Chaisber-to-Drywell VacuumBreaker Valves /A6.4.1.5 Missile Barriers Inside Containment / /A6.4.2 Reactor Coolant SystemA6.4.2.1 Reactor Pressure Vessel and Internals VA6.4.2.2 Reactor Pressure Vessel Nozzles / /A6.4.2.3 Reactor Pressure Vessel Skirt / VA6.4.2.4 Recirculation Lines / /A6.4.2.5 Main Steam Lines /A6.4.2.6 Main Steam Line Isolation Valves V / /A6.4.2.7 Pipe Whip Restraints for RecirculationLines V VA6.4.2.8 Reactor Pressure Relief ValvesA6. 4 .3 Core Spray SystemV VA6.4.3.1 Piping V VA6.4.3.2 Hangers and SnubbersýHangersSnubbersA6.4.3.3 Pumps and Drives V /A6.4.3.4 Valves VA6.4.3.5 Valve Motor Operators V VA6.4.3.6 Instrumentation VA6.4.4 HPCIS TurbineA6.4.5 RNR Pumps and DrivesVVVV VVVA6.4.6 On-Site Electric Power SystemsA6. .6.1 4 Diesel Generator Buildings:(a)Structure and Door V VDoor HingesCorrugated Barrier AttachmentsA6.4.6.2 Diesel Generators V V VA6.4.6.3 Batteries and Battery Racks:Batteries V VBattery Racks4V V VA6. .7 Electric Power Distribution SystemsA6.4.7.2 Electrical Containment PenetrationsA6.4.7.3 4 kV Switchgear V V VA6.4.?.4 480 V Load CentersA6.4.7.5 480 V Motor Control Centers V V VA6.4. 7 .6 DC Distribution Panels and Fuse Boxes V VA6.4.7.7 Cable Trays V V VA6.4.8 Reactor Protection System V V((d)(a)Diesel Generator Building was evaluated for both seismic and tornado loads.(b) Design Criteria as stated in FSARj or current criteria, if original criteria were not explicitor found to have insufficient margin.(c) More sophisticated analysis methods may in some cases show that the design criteria have in factbeen satisfied; in all cases however, loss of function is not expected.(d) Periodic seal replacement in the snubbers is recommended.Table X A-7X-107/108

a36

14tTABLE X A-8 BWR COMPONENT REVIEW - ENVIRONMENT QUALIFICATION SUMMARYReview Data BasisAdequate (Failure Not Expected)Assessment of Design AdequacyNot Demonstrated (Failure Possible)o o V Design Criteria4 8 Satisfied (a)(d Analysis Design Criteria Insufficient Information ReviewedSubsection Wo HU U 4.) s and/or Engineering Not Completely Information nomtint to Indicates nictsta thatin Text Component Description > .r , E ' Text Judgment Satisfied(b) Assess Adequacy Failure is Likely NotesA6.4.1 ContainmentA6.4.1.4 Suppression-Chamber-to-Drywell VacuumBreaker Valves /A6.4.1.6 Paint Coatings Within Containment V /A6.4.2 Reactor Coolant SystemA6.4.2.6 Main Steam Line Isolation Valves / / VA6.4.2.8 Reactor Pressure Relief ValvesA6.4.3 Core Spray SystemA6.4.3.2 Snubbers / /A6.4.5 RHR Pumps and Drives / /A6.4.7 Electric Power Distribution SystemsA6.4.7.1 Electrical Cables Terminations andConnectors / /A6.4.7.2 Electrical Containment Penetrations /(a)Design Criteria as stated in FSAR; or current criteria, if original criteria were not explicit or found to have insufficient margin.(b) More sophisticated analysis methods may in some cases show that the design criteria have in fact been satisfied; in all cases however, loss of functionis not expected.TABLE X A-9COMPARISON OF DAMPING VALUES USED IN PWR PLANT WITH THOSE CURRENTLYREQUIREDDamping Values (% Critical)OBEDBEStructure or Component PWR Current PWR CurrentReactor-Vessel InternalsWelded Components and Assemblies 1.0 2.0 1.0 3.0Bolted Components and Assemblies 2.0 2.0 2.0 3.0Reinforced Concrete 5.0 4.0 5.0 7.0Steel-Frame StructuresWelded 1.0 2.0 1.0 4.0Bolted 2.5 4.0 2.5 7.0Mechanical Equipment 2.0 2.0 2.0 3.0Piping>12.0 in. dia. 0.5 2.0 0.5 3.0

TABLE X A-10 COMPARISON OF DAMPING VALUES USED IN BWR PLANT WITH THOSE CURRENTLYREQUIREDDamping Values (% Critical)Structure or ComponentDEMCEBWR Current BWR CurrentReinforced-Concrete Structures 2.0 4.0 5.0 7.0Steel-Frame Structures 2.0 2.0 5.0 4.0Welded-Steel Assemblies(a) 1.0 2.0 2.0 3.0Bolted and Riveted Assemblies(a) 2.0 2.0 5.0 3.0Piping>12.0 in. dia. 0.5 2.0 0.5 3.0

TABLE X A-12COMPARISON OF COMPUTED FREQUENCIESFrequency (cps)4-Degree-of-14-Degree-of-Mode Freedom Model Freedom Model1 1.49 1.522 2.87 3.653 4.68 5.3074 6.03 10.0TABLE X A-13ESTIMATE OF MODAL DAMPINGMode Type % Critical Damping1 Rocking 10.42 Translation 36.83 Flexure 2.54 Translation 3.8&Table X A-10 - Table X A-13X-111/112

TABLE X A-14 SUMMARY OF MOMENTS AND STRESSES IN CRANE (TROLLEY AT CENTER LINE)Moment Section Modulus Stress Max.Vert. Horiz. Total AllowVert. Horiz. 3 3 Vert. Horiz. Stress StressLoading Section (kip-ft) (kip-ft) (in ) (in3) (ksi) (ksi) (ksi) (ksi)Main GirderDead LoadLive LoadImpact @ 15%Dead LoadLive LoadDBE10% DBEDead LoadLive LoadOBE10% OBE1781 -- S = 8433c1550 -- ST = 7612233 --35641781 -- S = 84331550 -- ST = 7612.999 152999 1534429 16821781 -- S = 8433c1550 -- ST = 7612633 112863 1134027 12415.072231 5.626.302231 9.705.732231 6.35-- 5.62 < 17.60.49 cy = 17.60.9 c = 32.4Y9.05 16.03 < 32.40.9 y6.68 13.03 < 32.4End TieDead LoadLive LoadDBE10% DBEE'-E-- -- S = 1065.0c..-- ST = 670.6-- 1012-- 101553.7 --0.9 ay24.12 24.12 < 32.41113Dead LoadLive LoadDBE10% DBE..-- S = 588.6c.-- . T = 510.7-- 445-- 45548.5 --0.9 ay.10.72 10.72 < 32.4490Dead LoadLive LoadOBE10% OBEE'6E..-- Sc = 1065.0.. ST = 670.6747-- 75822553.7 --0.9 oy17.81 17.81 < 32.4Dead LoadLive LoadFý,F..-- Sc = 588.6.-- ST = 510.7548.5 --0.9 ay7.92 7.92 < 32.4OBE-- 32910% OBE-- 33362

TABLE X A-15 CRITICAL STRESS SUMMARY (psi)Load (a) Seismic Total Allowable MarginStress Stress Stress SafetyP + W + OBE 10,090 16,343 1 8 , 7 5 0P + W + DBE 14,863 21,116 2 8 , 1 2 5(b) 0.15(c) 0.33(a)(b)(c)P = Pressure, W = Dead Weight1.2 Sh'1.8, Sh.TABLE X A-16POSSIBLE fe/fs VALUESf(a) f (a) fe/fsConfiguration fe s e s1 5.12 5.3 0.972 8.31 5.3 1.57(a)cpsTABLE X A-17 STEAM GENERATOR SUPPORTS - MAXIMUM MEMBER LOADS - PIPE RUPTURE PLUSSEISMICRupture andPipe Rupture - Seismic - Seismic -Member Maximum Loads Maximum Loads Maximum LoadsNo. (kip) (kip) (kip)21-47 1609.6 -98.5 -1708.122-48 242.7 +48.0 290.723-49 -242.7 -48.0 -290.724-50 -1603.1 -98.5 -1701.625-51 412.0 46.0 458.027-52 -412.0 -46.0 -458.031-57 549.7 115.2 664.916-58 -635.4 -116.0 -751.433-59 613.3 115.4 728.732-56 1526.0 25.6 1551.630-55 1555.9 25.5 1581.430-54 2183.3 64.6 2247.929-53 -1444.0 -30.7 -1474.7Table X A-14 - Table X A-17X-113/11.4

TABLE X A-18 REACTOR COOLANT PUMP SUPPORTS - MAXIMUM MEMBER LOADS - PIPE RUPTUREPLUS SEISMICRupturePipe Rupture - Seismic - SeismicMember Maximum Loads Maximum Loads Maximum LNo. (kip) (kip) (kip)11-40 493.9 22.0 515.911-60 -114.9 -12.3 -127.212-60 -696.8 -20.7 -717.513-41 743.3 10.3 733.613-44 -680.2 -17.3 -697.514-42 538.9 20.7 559.614-43 577.0 13.9 590.914-44 -117.2 -11.5 -128.740-42 -212.7 -3.4 -216.140-60 -210.2 -3.6 -213.840-64 -29.7 -5.5 -35.240-65 1243.6 21.9 1265.542-44 -212.8 -3.4 -216.242-64 1257.7 20.2 1277.944-60 134.3 1.9 136.244-64 313.6 1.9 315.544-67 -963.1 -11.4 -974.560-65 348.6 2.1 350.760-67 -18.6 -2.4 -21.060-69 -1041.0 -14.4 -1055.4

ATABLE X A-19 SUMMARY OF CRITICAL STRESSES (psi) IN LHSISLoad Seismic Total Allowable MarginLoad Seismic Total Allowable MarginCombination Stress Stress Stress SafetyP + W + OBE9,38914,69816,860Piping Section0.23 Accu] mulator PipingP + W + DBE13,94419,24725,2900.43 Acctumulator PipingP + W + OBE7,19114,94119,3200.61 LHSILeg- Pump- Loopto Cold1P + W + DBE10,707 18,45128,9800.98 LHSILeg- Pump- Loopto Cold1P + W + OBE8,641 16,36819,3200.34 LHSILeg- Pump- Loopto Cold2P + W + DBE12,834 20,55228,9800.65 LHSILeg- Pump- Loopto Cold2P + W + OBE8,447 15,96819,3200.39 LHSILeg- Pump- Loopto Cold3P + W + DBE12,445 19,98828,9800.72 LHSILeg- Pump- Loopto Cold3TABLE X A-20SUMMARY OF ACCUMULATOR TANK NOZZLE PIPE LOADSPipe (a)LoadDBESystem No.2 (b)ThermalDBESystem No.3 (c)ThermalF x1,767 lb1,615 lb26,647 lb 252 lbF y5,670 lb165 lb4,442 lb2,153 lbF z9,903 lb1,214 lb4,046 lb3,921 lbM x14,337 ft-lb9,924 ft-lb13,223 ft-lb 11,759 ft-lb4M y26,302 ft-lb22,307 ft-lb35,226 ft-lb 50,043 ft-lbM z42,119 ft-lb14,932 ft-lb160,654 ft-lb 30,775 ft-lb(a) x, y, z refer to global coordinates for piping analysis; F isforce, and M is moment.(b) Ref. 16.(c) Ref. 17.Table X A-18 - Table X A-20X-115/116

TABLE X A-21 LOAD CONDITION AND STRESSLIMITS - CODE CASE NO. 1607Load ConditionStress LimitsUpset am < 1.1 S(am or aL) + ab < 1.65 SEmergency am < 1.5 s(a or aL) + ab< 1.8 SFaulted am < 2.0 S(am or a]) + ab < 2.4 S(a)am = general membrane stress (nodiscontinuity effects) producedonly be mechanicalloads.aL = local membrane stress(including discontinuityeffects) due to mechanicalloads only.ab = bending stress due to mechanicalloads only.S = allowabie stress value givenin Tables 1-7.1, 1-7.2,1-7.3 of Appendix I of currentASME Code Section III.

TABLE X A-22 ACCUMULATOR PIPING CONNECTION NOZZLE LOADSMethodLoadMaximum Stress(psi)Code Allowable Stress(psi)1. B31.1onlyThermal expansionOBE + pressure13,05012,67025,79018,800 (1.2Sh),aDBE + pressure16,38028,200 (1.SSh)2. BijlaardandB31. 1ThermalOBE + pressure7,53012,73025,79018,800(1.2S h)DBE + pressure15,79028,200(1.8S h)TABLE X A-23REACTOR BUILDING NATURAL FREQUENCIESMode North-South (cps) East-West (cps)1 5.05 3.642 6.13 5.493 8.15 8.08Table X A-21 - Table X A-23X-117/118

TABLE X A-24 COMPUTED STRESS INTENSITIES IN BWR VESSEL NOZZLESPrimary Membrane Allowable Secondary Plus AllowablePlus Pressure Stress Stress Primary Stress StressNozzle Mark Max. Stress Intensity 1.5 Sm Max. Stress Intensity 3.0 SmDescription No. (ksi) (ksi) (ksi) (ksi)Recirculation outlet 8 27.09 40.0 30.20 80.0Recirculation inlet 7 26.33 40.0 27.77 80.0Steam outlet (a) 14 26.79 40.0 27.01 80.0Feedwater 10 26.41 40.0 27.39 80.0Core spray 11 26.33 40.0 27.35 80.0Jet pump 19 26.29 40.0 26.46 80.0instrumentation(a)CRD hydraulic 13 26.29 40.0 26.43 80.0system returnHead spray(a) 206 19.88 40.0 23.22 80.0Vent(a) 204 19.65 40.0 20.36 80.0Drain(a) 22 9.85 40.0 10.00 80.02 in. instru- 12 T = 3.624 T Allowablementation avg 12(a) Nozzles qualified in fatigue under Article NB-3222,.4(d), Section III, ASME Code.TABLE X A-25 REACTOR PRESSURE VESSEL ELASTIC ANALYSISMaximumComputedCodeStress AllowableStress IntensityCategoryCriticalLoadingIntensity(ksi)Value(ksi)Primary Membrane SSE + Dead Load 15.1 Sm = 26.7SSE + Jet Load + 28.5 1.5 Sm = 40.1Dead LoadPrimary Plus 111.0Secondary StressIntensity Range3 Sm = 80.0

TABLE X A-26NATURAL PERIODS AND SPECTRALACCELERATIONSMode123456789101112131415161718192021Period (sec)0.33550.3039 (a)0.296710.28320.26530.22430.2117 (a)0.21070.1889 (a)0.18100.16850.14540.1400 (a)0.13780.12880.12710.11650.10470.09960.09300.0898Spectral Acceleration(in./sec2112.05115.69116.47117.92119.84142.40177.83180.53241.65270.87321.15358.98358.98356.71268.35252.36169.4884.7482.1378.7377.08(a) Note: Closely spaced periods.TABLE X A-27 CRITICAL STRESSES (psi)(a) Seismic Total Allowable MarginLoads Considered Stress Stress Stress Safety LocationP + W + DE 5,910 14,409 1 7 , 3 1 2 (b) 0.20 #44-12 in. ElbowP + W + MCE 13,180 21,000 2 5 , 9 6 9 (c) 0.24 Near ReactorP + W + DE 2,341 7,803 17,312 1.22 f #48-Branch aboveP + W + MCE 4,682 10,144 25,969 1.56 Gate Valve(a)(b)(c)P, pressure; W, weight; DE, design1.2 Sh.1.8 Sh.earthquake; and W, weightTable X A-24 - Table X A-27X-119/120

TABLE X A-28PERIODS OF VIBRATIONAND SPECTRAL ACCELERATIONSFOR MAIN STREAM LINEMode Period SpectralNo. (sec) Acceleration (g)1 0.338 0.472 0.187 1.813 0.158 1.834 0.106 0.35 0.101 0.36 0.099 0.37 0.058 0.658 0.054 0.489 0.053 0.42TABLE X A-29 CRITICAL STRESSES (psi) FOR MAIN STEAM LINESeismic Total Allowable MarginLoads Considered Stress Stress Stress Safety LocationP + W + DE 7,492 15,938 21,000 0.32 HPCI Tee (pt #118)P + W + DE + T (a) 4,643 45,285 52,550 0.16 @ 26 in. Tee (pt #79P + W + MCE 14,984 23,147 31,500 0.36 HPCI Tee (pt #118)(a) T, thermal expansion.TABLE X A-30 SUMMARY OF LOADS ON PIPE WHIP CONSTRAINTSCase I Case II Case IIIT (kip) S (kip) M (in.-kip) T (kip) S (kip) N (in.-kip)Act. Allow. Act. Allow. Act. Allow. Act. Allow. Act. Alloyw. Act. Allow.Load Load Load Load Load Load Load Load Load Load Load Load12 in. • 127 16922 in. 4 394 52528 in. 637 860I127394637155 1650 2020 90 122 90 122 1170 1580437 670 7420 279 324 279 324 4740 5500662 10800 14600 450 510 450 510 9900 11200I

TABLE X A-31 CRITICAL STRESSES (psi)Earthquake Total Seismic Allowable MarginDE 14,538 6,232 17,375 0.2MCE 23,263 14,957 27,800 0.2TABLE X A-32RESULTS OF DYNAMIC ANALYSIS OF BATTERY RACKSAcceleration (g)FundamentalDesign Earthquake(from FloorMaximum CredibleEarthquakeType of Frequency Response Spectrum) (Taken 2.4 TimesRack (cps) (2% Damping) Design Earthquake)HorizontalExcitation12- and 9.2 (hand17-ft racks computation)8.0 (computercheck) 0.18 0.436-ft racks 16.5 0.12 0.288VerticalExcitat-iron12- and17-ft racks 8.0 0.16 0.386-ft racks 13.9 0.17 0.41TABLE X A-33 HANGER ATTACHMENT TYPESHanger LengthTypes of(ft) No. of Trays Hanger Attachment63 to 16 1 swivel3 to 9 2 to 4 swivel16 and up 1 fixed9 and up 2 to 4 fixedTable X A-28 - Table X A-33X-121/122

Plant PWROBW RSAFETY STUDY DESIGN ADEQUACYREVIEW DATA SHEETFIRL Reviewer1Component Description:Drawing References:35Installed Elevation: 4 - Location:Design Criteria:6Data Basis Description:79- Estimated Weight: 8 - Fundamental Frequency:- Mathematical Model: .......10- Method of Analysis:11- Seismic Load Definition:12- Maximum Stresses: (LocationTotal Stress Seismic Stress Allowable MarginOBEDBE1314- Maximum Deflection: Location:- Comments:FIGURE X A-iDesign Adequacy Review Data Sheet

PlantWSAFETY STUDY DESIGN ADEQUACYSUPPLEMENTARY REVIEW DATA SHEETF I RL Reviewer1 - Component Description:Reviewed ItemSupplementary Data6FIGURE X A-2Supplementary Review Data SheetFig. X A-1 - Fig. X A-2X-123/124

ZCyx - -- + ±z:-C- "YOKE WEB DIMENSIONS AT WEAKEST CROSS SECTION+ -z b,= o.SON.-Y1, =- 7.64 IN0 (MOMENT OF INERTIA)A = 6.50 IN? (AREA)C= 2.25 IN.DIMENSION TOOUTER FIBER)h=4.501IN.b2= 2.00 IN.h2= 0.50 IN.FIGURE X A-3 Example of a Multi-Element Structure Connected toRigid Structures at Both Endswse MM40aS09•m 89coaM mus as 6 4 a D *A assi0A CLi wiOi M a" amS amSPECTRA ARE FOR A MAXIMUM HORIZONTAL GROUNDACCELERATION OF 7 PERCENT OF GRAVITYUPI K.17; T~- I4__3

Mg*O "I itCOWO$

. .. .,9, A;,., -. .. ... .. ." , ,,- . 7-• ..'-I"-., ". - - ',-., *-."• 4 / 4/.Frequency, cp•s.FIGURE X A-7 Response Spectra (MCE) - BWRFig. X A-3 - Fig. X A-7X-125/126

4ab

I Ar-, Rt cea.twls~(1,t.4v EL.47-4*XAA42L.T-Vta,~~-J* CL.1L~~aim2*'FIGURE X A-8Cross Section Through Reactor Building

a,(a) CONTAINMENT a INTERNAL STRUCTURE (b) MATHEMATICAL MODELMASS PROPERTIESm 1 = 2400 kip-sec 2 /ft* 2 = 475m 3 = 3103 = 20 ) 1O6 kip-sec 2 -ft0STIFFNESS PROPERTIESkI = 5 GR = 106 kip/ftk4 = 3 GR 3 = 2.7 x I09 ft-kip/radk3 = 0.3 x 106 kip/ftk2 = 0.31 x 106 kip/ftFIGURE X A-9Soil-Structure Interaction Model,Z4U8,K14/74LCUT OUT TO 14'ABOVE BASE MAT% 0 e?FIGURE X A-10 Containment Internal Structure -Cross Section of Lower Portion

STEAM GENERATORSTEAM GENERATORUPPER SUPPORT-STEAMLOWERV'00ZREACTORCOOLANTPUMP9g29FIGURE X A-11 Reactor Coolant System - Seismic Analysis ModelUsed by Suppliere-U.1.68Z00.50/. EQUIPMENT DAMPINGuJ-j1.0C-)40.350.140.21 _0.0I 0.070 0.125 0.90PERIOD (SEC)2.0FIGURE X A-12Horizontal OBE Floor Response Spectrum Used forReactor BuildingFig. X A-8 - Fig. X A-12X-127/128

STEAM GEN. CENTER LINERC PUMP CENTER LINEHOT LINEPUMP SUCTION LINECOLD LINEMAIN STEAM LINEFORCING FUNCTIONS1 2 3 4 570 37 17 20 19LEGEND1-2-3-26-4-5-6-78-35-9-1597-89-85-79-70-62-3738- 36-61-71-83-88-92-95-91-3782-76-66-45-1539-63-78-81-86-90-94-96-991-34-46-68-77-93-986 7 8 9 10 11 12 13 14 15 1618 38 61 71 95 91 66 45 39 90 98LOWER STEAM GENSNUBBERSSTEAM GEN. SWIVELEND COUPLINGSSTEAM GEN.SUPPORTSSTEAM GEN.SNUBBERSHANGINGUPPER30-5532-5629-5330-5416-5831-5733-5921-4724-50STEAM GEN. SNUBBER 22-48BOLTS 23-4925-8127-52RC PUMP 13-41SNUBBERS 14-43RC PUMP FEET 11,12,13,14REACTOR NOZZLES 97,99FIGURE X A-13A/E Mathematical Model for Dynamic Analysis ofPrimary Coolant Lop

THRUST BEARING.OIL LIFT PUMP + MOTORa4FIGURE X A-14Typical Reactor Coolant Pump

BAR STOCK AISI-1141MINIMUM YIELD STRENGTH=55,OOO PSI0TORFIGURE X A-15 Pipe-Whip Restraint - Main Steam tine (not to scale)INSIDE PRESSURE =-700 PSIDESIGN TEMPERATURE - 300' FFIGURE X A-16Accumulator Tank NozzleFig. X A-13 - Fig. X A-16X-129/130

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0 0.15FIGURE X A-17075 1.00 2.00PERIOD (SECHorizontal DBE Floor Response Spectrum Used forDiesel Generator Building300.5%4 EQUIPMENT DAMPING0- 2.01.O 1-.800C6O0.30 0.300 015 045FIGURE X A-181.00PERIOD(SEC)Vertical DBE Floor ResponseSpectrum Used for DieselGenerator Building200

FIGURE X A-19Day TankBACKFRAMEAIR BOTTLECLAMP STRAP(TYPICAL)FIGURE X A-20Air Bottle Support (Not to Scale)

OJC CMAS&I~h M~j1, TRUSS t.29" 0,WThL S191%MML~ rA~k 000~L23~ - AL'CO.I L21' 2Lth c2tmAes FA OOT -RAIL EL.1953 VIll.bIL 0I___ Fl IfC '/111F.4 e 2J2 '0sa~TINFfP1etPAS 4ySAC I"* WOUTaw0P'PFIGURE X A-21Elevation View Through BWR PlantFig. X A-17 - Fig. X A-21X-131/132

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EL. 291' - 3"Legend:0JointEL. 279' - 3"EL. 259' -6"00(4 IllReactor Bldg.0MemberSpringLumped MassEL. 234' -0"0R. VesselEL. 214' -0"EL. 195' -0"EL. 180' - 0"504 aS. Shield13017 EL. 213' - 0"16EL. 177'-0"EL. 165' - 0"EL 165' - 0"20014EL. 161' -0"EL. 135'-0" 4EL. 145' - 0"E• 135'-0" 0L.19,200/'- E L. 120'-0"EL. 92' - 6"FIGURE X A-22Reactor Building Mathematical Model(.

z0w-JwC-)00.1 0.2 0.4 0.6 0.8 1 2 4 6 8 10 20 40 60 80 100FREQUENCY (CPS)FIGURE X A-23Response Spectra5% g, 2% DampingComparison, Design Earthquake,FIGURE X A-24 Vacuum Breaker Valve

LEGEND 'Shell Stretchout - Inside View of Reinforc-3 ';4THICK PLATE - I'2"THICK PLATE ing Plates. Scale 1/16" - 0 .-2'/2 THICK PLATEI'"4"THICK PLATEFIGURE X A-25Missile Barriers Inside DrywellFIGURE X A-26 Safety ValveFig. X A-22 - Fig. X A-26X-133/134

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28" PIPERECIRCULATIONPUMPEL. 122*-4"FIGURE X A-27Recirculation LineFIGURE X A-28Main Steam IsolationValve

FIGURE X A-29Pressure ReliefValveREACTORNYxz14 OCN -22. GUARD PIPEa-14DCN- 12"FIGURE X A-30A/E Mathematical Model for Dynamic Analysis ofCore Spray Piping

FIGURE X A-31 Core Spray Pump andDrive - Lower PortionShowing Pump MountingFIGURE X A-32Core Spray Pump andDrive - Upper PortionShowing Motor andCoupling to Pump

4UhFIGURE X A-33Typical Valve MotorOperator Installation04FIGURE X A-34HPCIS Pump and InterstagePipingFig. X A-27 - Fig. X A-34X-135/136

~jnIFIGURE X A-35HPCIS Turbine DriveI,/FIGURE X A-36Watertight Entrywayto RHR Pump Room

FIGURE X A-37RHR Pump and DriveWM = 14,000 lb8 - 1" ) BoltsEWindow Aream- Shatt Couplin WS - 2,200 lb24 - 1 Y/.'" BoltsSuctionDischargeFIGURE X A-38RHR Pump and Motor

PeakIs-HE A -A J0 0.1 0.2 0.3APe IM I-0.4TimeI I I I IJ. ~ T0.5 0.6 0.7 0.8I ICEU'U0.0-T0 0.1 0.2 0.3 0.4 0.5 0.6Time0.001 " Typ.FIGURE X A-39Selected Displacement Records from RHR PumpVibration Tests Showing Cyclic DisplacementPeaks"IFIGURE X A-40Diesel GeneratorBuilding - SideFacing River

HJFIGURE X A-41Diesel GeneratorBuilding - OppositeSide161'151' DIESELGENERATORFLOOR127'I7 105'FIGURE X A-42Lumped Mass Model for DieselGenerator Building SeismicAnalysisFig. X A-35 - Fig. X A-42X-137/138

FIGURE X A-43Diesel GeneratorUnit - View fromGenerator EndFIGURE X A-44View of Diesel GeneratorUnit ShowingTurbo-Charger

POINTSOF ATTACHMENT BETWEENSUBBASE 8 SKIDFIGURE X A-45WF CROSS BEAMSDiesel Generator Dynamic ModelFIGURE X A-46Batteries and BatteryRacks - 125/250-VSystem4

FIGURE X A-47Batteries and BatteryRacks - 24-V System750 LB75C LBFIGURE X A-48K -16 1Lumped-Mass Model for Seismic Analysisof 58-Cell FTC-21 Battery Rack. (Typicalof Model Scheme for All BatteryRacks.)

WIFIGURE X A-49Electric Power andControl CablesInstalled in CableTraysFIGURE X A-504-kV Switchgear UnitFig. X A-43 - Fig. X A-50X-139ý/140

FIGURE X A-514-kV Switchgear.cksFIGURE X A-52 480-V Load Cerer

FIGURE X A-53480-V Motor ControlCentersFIGURE X A-54D-C DistributionPanels

pp,FIGURE X A-55RPS InstrumentationRacks (Front View)&FIGURE X A-56RPS InstrumentationRacks (Rear View)

I ;ýj -I ý 4 411-'4FIGURE X A-57 Reactor ProtectionSystem Logic CabinetFIGURE X A-58Control Room Panelsfor Reactor SafetySystemsAFig. X A-51 - Fig. X A-58X-141/142

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