Lafayette Dam Stability Report - East Bay Municipal Utility District

GeotechnicalEnvironmental andWater ResourcesEngineeringDynamic Stability Reviewof Lafayette Dam ( 1 of 2 )ReportSubmitted to:East Bay Municipal Utility DistrictOakland, CAPrepared By:Gilles Bureau, P.E., G.E.William A. Rettberg, P.E.Date 08/16/05Project 04035-0

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/051. SUMMARYThis report presents the results of our dynamic stability review of Lafayette Dam, ContraCosta County, CA. The dam is referred to as Dam No. 31-2 by the State of California,Department of Water Resources, Division of Safety of Dams (DSOD). The dam is ownedand operated by the East Bay Municipal Utility District (EBMUD, or the District).The project inspection and preparation of this report were done by GEI Consultants, Inc.(GEI), under the direction of Gilles Bureau, P.E., G.E., Project Manager, Bill Rettberg, P.E.,Project Director, and with assistance from Carol Buckles, P.E., G.E. We also acknowledgeour Subconsultants: William Cole, R.G., C.E.G. from Cotton, Shires and Associates, Inc.(CSA), who reviewed the site geology and tectonic environment; Mark McKee, P.E. ofRobert Y. Chew Geotechnical, Inc. (RYCG), who participated in the review andinterpretation of previous data; and Sangeeta Lewis, P.E. of Lewis Engineering (LE), whoperformed the slope stability analyses. We also acknowledge the contributions of Dr. I.M.Idriss, P.E., G.E. who independently reviewed the draft of this report and provided usefulcomments.In the preparation of this report, GEI and project team members reviewed previous reportsand other information updated since the last safety evaluations of Lafayette Dam by Shannonand Wilson (1966) and W.A. Wahler and Associates (1976); performed site and geologicinspections; defined current seismic requirements; reevaluated the liquefaction potential ofthe embankment and foundation soils; and performed limited new analyses to assess thestability of the embankment and estimate the potential for earthquake-induced deformations.We reviewed existing project reports describing construction and performance history,previous field exploration and laboratory testing programs, geology, engineering drawings,plans, specifications and other documents provided by the District and the DSOD. Theseprevious data, which cover and summarize 77 years of project history and performance sincethe beginning of construction of Lafayette Dam, provided substantial project information.The Phase I Inspection Report (National Dam Inspection Program) prepared by the DSOD(1980) was another useful source of information.We gratefully acknowledge the assistance of District personnel in providing feedback andcontinuous interaction and assembling and making the extensive project data available to us.We appreciated the cooperation of Xavier Irias, Manager of Engineering Services, AttaYiadom, Project Manager, Fred Starr, Senior Civil Engineer, and Hon Fung Chan, AssistantEngineer, who coordinated our access to the District’s project files. DSOD staff, includingTina Glorioso and Chuck Wong, facilitated our review of the State’s files.GEI Consultants - 1 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05No supplemental field exploration or laboratory testing was conducted as part of this review.To prepare this report, we exclusively relied on knowledge of the project acquired fromexisting reports, re-interpreted and updated as required, and on applicable recent literature,discussions with DSOD and EBMUD project engineering staff, and our field observations atthe site.1.1. Lafayette DamLafayette Dam was built from August 1927 to 1933. It is mostly founded on valley fillalluvium, underlain by soft sedimentary bedrock of the Tertiary Orinda Formation. A majorfailure of the downstream (D/S) slope occurred during construction in 1928. The failed slopewas rebuilt at a flatter angle by placement of additional fill and the dam completed at a lowercrest elevation than originally designed. At the time of failure, the constructed crest elevationwas at about El. 476, hence 9 feet higher than the final as-built crest elevation (El. 467).Since its completion in 1933, the dam has performed satisfactorily. The reservoir is operatedas a standby emergency storage and recreational facility.The dam is a rolled zoned earthfill embankment, 132 feet high, and with a crest 1,200 feetlong. The upstream dam face has a slope of 3H to 1V (horizontal to vertical), with two 15-foot wide berms originally designed at El. 450 and El. 400, respectively. The downstreamslope varies at 2.5H to 1V, 4.0H to 1V and 8.0H to 1V and has a 10-foot wide berm, locatedat El. 400. The dam crest is 210 feet wide.The dam has a central clayey core with upstream and downstream slopes, originally built at0.5H:1V. The constructed top width of the core is about 22 feet. A cutoff trench wasexcavated beneath the core into the alluvium during the early phase of the dam construction.A steel sheet pile curtain and a short concrete cutoff wall (pile cap) were installed in thecutoff trench, before placement of the core materials.The reinforced concrete outlet tower, approximately 120-foot high with an annular crosssection,provides reservoir drawdown capacity. The dam does not have a channel spillway.The top rectangular port in the tower (2.5’ x 3’) is ungated and provides limited spillwaydischarge capacity, with a spillway crest elevation at El. 449.2. Reservoir control is providedby four rectangular ports in the tower wall, of same size as the spillway port and equippedwith slide gates.1.2. Field and Geologic InspectionsOur engineers and geologist inspected the dam, reservoir slopes and surrounding area duringthree field visits. The dam embankment does not display significant cracks, recent horizontalor vertical displacements or shear failure, or any other visible signs of active deformations,GEI Consultants - 2 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05tectonic features. The Hayward Fault is located 8.8 km to the southwest, and the CalaverasFault is 9.8 km to the southeast. Although more distant (39 km), the San Andreas Fault isalso significant to the seismic setting of Lafayette Dam, because it is capable of locallygenerating moderate to strong shaking with long duration (60 seconds or more). These threemajor active faults and other smaller related faults form a complex tectonic environment.Additional information is provided below regarding the faults most critical to the dam.The Hayward Fault (about 87 km long) and its northern portion, which is the closest to thedam site, represent a relatively well-defined tectonic feature. It has an average slip rate of 9mm/year. Segmentation of the fault, which has been considered in some studies, no longerseems clearly defined, based on the results of current research. Based on the lack ofindisputable evidence for segmentation, we have assigned a Moment Magnitude of 7.25(Mw) to the Hayward Fault.The Calaveras Fault (nearly 120 km long) is a major active feature that has been subdividedinto three segments, based on geologic, geomorphic and seismic data. The northern segment,about 42 km long, is the closest to the Lafayette site, but has been less active (average sliprate 6 mm/year) than the central and southern segments (average slip rate 15 mm/year).Geologists generally agree that segmentation of that fault and the different rates of slip alongthe three segments justify consideration of shorter lengths of rupture for upper-boundmagnitude estimation purposes, although the possibility of rupture propagating from onesegment to the other cannot be completely ruled out. Based on assumed segmentation, wehave assigned a Moment Magnitude (Mw) of 7.0 to the Calaveras Fault in this study.The northern segment of the San Andreas Fault (474 km long) is the dominant regionaltectonic structure accommodating right-lateral, translational motion and represents theboundary between the North American and Pacific plates. It has been the most seismicallyactive of the faults present in the Bay Area and has the highest rate of slip, between 17 and24 mm/year for its segment the closest to Lafayette Dam. Lafayette Dam is approximately39 km away from the northern portion of the San Andreas Fault. Earthquake magnitudeestimates range from Mw 7.1 for individual segments up to M w 7.9 for the entire northernSan Andreas Fault. Despite its distance, it represents a significant earthquake hazard forLafayette as being potentially associated with large magnitudes and long durations ofshaking. We used M w 7.9 in this study.Other smaller faults, such as the Lafayette-Reliez Valley (LRV), at a distance of 3.0 km,Franklin (6.4 km), Miller Creek (9.5-11.3 km) and Concord (12.8 km) could generatesignificant motion at the site, although of shorter duration than upper-bound magnitudeevents along the Hayward or Calaveras faults. These faults are all less than 20 km long.Based on recent studies, the Lafayette-Reliez Valley and Franklin faults, as well as folds andthrusts in the vicinity of the Lafayette site, seem related to a transfer to the west, toward theHayward Fault, of tectonic stresses associated with the northern extremity of the CalaverasGEI Consultants - 4 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Fault. We assigned upper bounds of magnitude of 6.5 (Mw) to all four of these smallerfaults. Because of its short distance, the LRV Fault could potentially be associated withstronger shaking at the site than either the Hayward or Calaveras faults, but of considerablyshorter duration.USGS geologists (Graymer, et al., 1994) have suggested that a previously unrecognizedinferred fault might pass under the dam, parallel to the crest. Insufficient information isavailable to confirm or reject such assertion. If such inferred fault were present, it wouldlikely not be active, because the orientation of its strike is not consistent with the trend ofother well-recognized tectonic features. In the absence of more specific information, thepresence of such inferred fault appears questionable. Even if it were present, we concludedthat it would represent an insignificant hazard to the dam as a potential seismic source, or interms of secondary (sympathetic) movement potentially triggered by a major rupture of anyof the major faults identified in the greater site area, because of its short length and favorableorientation with respect to the dam axis.Lastly, several poorly understood lineaments, the Russell Peak lineament, 2 km north of thedam, and another lineament, 1.5 km west of the reservoir, have been recognized since 2002and could represent potentially active or capable tectonic features. These lineaments are notstrongly pronounced, but their origin is not clearly explained. Based on the data reviewed todate, we cannot eliminate the hypothesis that these lineaments could be fault-related. Suchlineaments should have negligible impact on Lafayette Dam.1.4. Updated Seismic CriteriaLafayette Dam is a “high risk” facility located near faults with a high (1 to 9 mm/yr) or “veryhigh slip rate” (greater than 9 mm/yr). The “high risk” classification assigned to this dam inthe National Inventory of Dams (NID) reflects its potential for extreme human and economicconsequences in case of failure, due to heavy downstream development. Based on suchconsiderations, we updated the applicable seismic requirements based on deterministicconcepts and 84th percentile criteria, as is required by the DSOD for the correspondingcombination of hazard and consequences. To quantify potential ground motion at the site,we used three sets of ground motion attenuation equations, previously accepted and used bythe DSOD in recent dam studies, and developed horizontal and vertical response spectrarepresenting the maximum level of shaking that could be potentially generated at the site byeach of the active faults near or in the project’s greater vicinity. The influence of the localsite conditions, the foundation alluvium and the soil-like characteristics of the local“bedrock” (Orinda Formation) were taken into consideration in developing the groundmotions.GEI Consultants - 5 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Because the site is located within less than 10 kilometers of the northwestern portions of theHayward and Calaveras faults, near-field and directivity effects could potentially affect localground motions characteristics. As a significant portion of the rupture along these faultscould propagate toward Lafayette Dam, under conceivable earthquake scenarios, this couldincrease spectral accelerations at periods greater than 0.6 sec and simultaneously reduce theoverall duration of shaking at that particular location. Such effects could also affect groundmotion emanating from either the San Andreas or the LRF faults. Directivity and near-fieldeffects have been included in the development of our updated seismic criteria.Our ground motion estimates were expressed as horizontal and vertical peak groundaccelerations (PGA) and sorted by decreasing “Earthquake Severity Index” (ESI). The ESI isrelated to both the horizontal PGA and, through the magnitude, to the expected duration ofshaking (Bureau, et al., 1985). It quantifies the potential for earthquake-induced damdeformations better than the PGA. The fault most critical to Lafayette Dam is the HaywardFault, with an estimated horizontal PGA of 0.60g and an ESI of 12.48. Next come the SanAndreas and Calaveras faults with ESIs of 10.61 and 8.11, respectively. Computed PGAs forthe MCE’s assigned to these faults are 0.52g (Calaveras) and 0.27g (San Andreas). Hence,despite its significantly lower PGA, the San Andreas Fault represents an appreciable level ofrisk for Lafayette Dam. The LRV Fault could generate the largest horizontal accelerations(0.76g) because of its proximity (3 km), but would be associated with significantly shorterdurations of shaking. The ESI of the LRV Fault is 6.11.It should be noted that three of these faults could generate significant vertical accelerations atthe site, because of their short distances. Hence, for possible future seismic evaluationpurposes, we have updated both the horizontal and vertical spectra applicable to the Lafayettesite.We compared our recommended response spectra with the response spectra of theacceleration time histories used in 1976 by W.A. Wahler & Associates to represent theHayward (“Earthquake A”) and San Andreas (“Earthquake B”) earthquake scenarios. Webelieve that the spectral accelerations of Earthquake A were insufficiently conservative, by afactor of between 2 and 3, at the periods of significance to the response of Lafayette Dam.Earthquake B was sufficiently conservative.1.5. Liquefaction PotentialThe liquefaction potential of the various dam zones and of the foundation was reviewed. No“loose” saturated silts or sands, generally acknowledged the most susceptible to liquefaction,have been encountered in the borings. The embankment materials are classified as clays,sandy clay, or silty clay. CL and CH are the dominant soil classifications in the dam andfoundation materials, with ML occasionally encountered. The average plasticity index (PI)GEI Consultants - 6 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05and liquid limits (LL) of most of the core, shell and foundation materials fall outside the“liquefiable” or “potentially liquefiable” zones, based on a recent interim soil typesclassification using Atterberg limits (Seed, R.B., et al., 2003). The few data points that fallwithin the “potentially liquefiable” zone fail a supplementary test that would indicateliquefaction susceptibility and, therefore, do not represent any particular concern.While the dam and foundation materials are classified as “non-liquefiable” based on theirclay content and Atterberg limits, they might be susceptible to straining due to the largecyclic stress ratios that would be induced under the most demanding earthquake scenarios.While not “liquefiable” in the “classic” term as clean loose sands or silts would be, the damand foundation materials could be sensitive to progressive loss of strength with remolding ormonotonic accumulation of shear deformations under the most severe earthquake loading.1.6. Previous AnalysesMost previous analyses found the Lafayette embankment to have adequate static and seismicstability, with factors of safety generally complying with evaluation procedures, formulationsand criteria applicable at the time when such analyses were performed. The embankmentwas concluded to meet stability guidelines for static normal operating, rapid drawdown andseismic loading conditions, except in early slope stability studies (Dukleth, 1956). In suchstudies, the dam was concluded to be unsafe for rapid drawdown condition, should thereservoir level be raised 8 feet from its current operating level (El. 448), and for earthquakecondition (pseudo-static, 0.10g) with a reservoir level at only El. 420. The 1956 studies werebased, however, on very conservative (i.e. low) strength parameters. Subsequent static orpseudo-static (0.10g) analyses by Shannon & Wilson (1966) demonstrated acceptableperformance. These early definitions of the seismic requirements are insufficientlyconservative by current standards.W.A. Wahler & Associates (Wahler) performed equivalent-linear (EQL) dynamic responsefinite element analyses in 1976. The San Andreas event was found the most critical of thethree “Maximum Probable” earthquake scenarios considered (Hayward, Calaveras or SanAndreas). We concluded that the methodology and dam and foundation modelingassumptions used in these analyses raise some questions regarding prediction of the damresponse, due to the potentially insufficient resolution of the finite element grid (mesh toocoarse), the use of now obsolete equivalent-linear properties, and because of recentlyrecognized limitations of the analysis and interpretation procedures then implemented (e.g.inability to rigorously address the problem of large, non-recoverable soil deformations).Furthermore, the Hayward response spectrum used in 1976 is significantly below thepresently recommended response spectra at the periods of interest to Lafayette Damresponse. While the 1976 analyses represented the state-of-the-art at the time whenperformed, it appears from our review that no sufficiently reliable and detailed dynamicresponse analyses of Lafayette Dam are available, especially for the Hayward event. ThisGEI Consultants - 7 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05needs to be considered in relationship with the eventful construction history of the dam andits classification as a “high risk” facility.The most recent slope stability and simplified deformation analyses by the CaliforniaDepartment of Water Resources, Division of Safety of Dams (DSOD, 2003), using thematerial properties established in 1966 by Shannon and Wilson and updated seismic criteria,concluded that the dam should have adequate seismic stability and reserve freeboard.However, because the reservoir is located within a heavily urbanized area and in proximity tothe Hayward Fault, the DSOD suggested that the District perform its own review of the dam,which is the subject of this report.1.7. Updated Simplified AnalysesFor our static and pseudo-static slope stability analysis of Lafayette Dam, we reviewed thematerial properties previously used, and updated the dam section geometry and analysisproperties. We estimated strength properties from considerations such as the location of thesamples tested and the conditions of confinement that prevail in the field. We used theconsolidated-undrained triaxial tests (TXCU) results as the primary basis to define strengthproperties. We then performed slope stability analyses and implemented simplifieddeformation analysis procedures to evaluate the performance of Lafayette Dam and compareit with the previous static and dynamic analyses.We performed static and pseudo-static stability analyses of the upstream and downstreamslopes of Lafayette Dam using the computer program XSTABL. We successively consideredtwo methods of analysis (Janbu and Spencer). Our analysis model is generally similar tothose previously used, except for the foundation alluvium, which we represented with twozones. The alluvium below the core and upstream shell seems stronger than the alluviumbelow the downstream shell, which was affected by the 1928 failure.For steady-state seepage static condition and a reservoir elevation at El. 449, the lowestfactor of safety we calculated is 2.3 for the downstream slope, and 2.5 for the upstream slope.These values confirm the satisfactory performance of the embankment to-date. For thepartial rapid drawdown condition (repeat of the maximum historic reservoir drawdown to El.431), we calculated a minimum factor of safety of 2.0. We also postulated a rapid completedrawdown to the elevation of the lowest outlet port, and obtained a minimum factor of safetyof 1.7.We performed pseudo-static analyses to calculate the yield accelerations under simulatedseismic loading condition. We found the yield acceleration of the downstream slope (0.14g)to be substantially lower than that of the upstream slope (0.29g). The downstream yieldacceleration is slightly lower than obtained in previous investigations. This is because weused total-stress strength parameters for the pseudo-static analysis and took into account theGEI Consultants - 8 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05results of static consolidated-undrained (CU) triaxial compression test data by W.A. Wahlerand Associates (1976). The potential critical failure surface identified would involve theentire downstream shell and extend deep within the foundation alluvium. The soil volumebounded by this failure surface presents similarities with the soil mass involved in the 1928failure.Next, we estimated potential permanent earthquake-induced deformations using sevensimplified or empirical methods. Such methods only make approximate prediction of thedam performance. They are useful as a screening technique, however, to assess whether adam is safe by an appreciable margin, needs more rigorous investigations, or is potentiallyunsafe. We obtained deformations estimates for the upstream or downstream slopes, or forthe dam as a whole, depending on what procedure was implemented. As expected, differentmethods provide different estimates, which must be pondered by recognizing the limitationsof each simplified procedure implemented. We obtained “best” estimates by averaging crestsettlements obtained with the upstream or downstream yield accelerations and the variousprocedures implemented.Our seismic settlement estimates for the dam crest range from 0.9 to 4.5 feet for the HaywardEarthquake, which is the most critical event considered. An average settlement of 2.7 feet isour “preferred” prediction. We have chosen the term “preferred” to indicate that we havesimultaneously considered several methods and analysis assumptions to computedeformations or settlements. Such methods involve simplified procedures, which all havelimitations on how they can be applied to specific seismic, embankment and foundationconditions. Important factors, such as the presence of the alluvium and how the seismicloading would be truly applied, are only approximately or not taken into account in some ofthese procedures. The use of a “preferred” settlement estimate reduces the potential marginof error that would be associated with only considering the lowest or the largest estimateddam movements. Computed crest settlements for the Calaveras, San Andreas or Lafayette-Reliez Valley earthquakes were found to be less than for the Hayward Earthquake.In all of the simplified analysis procedures implemented, directly or indirectly computedmaximum settlements are less than the available freeboard (17.8 ft). Our estimated averagemaximum crest settlement (2.7 feet) would leave over 15 feet of residual freeboard, which isa considerable margin of safety. Alternatively, if an upper-bound settlement of 4.5 feet wereto occur, this would still leave 13.3 feet of residual freeboard. Hence, based on theseprocedures, Lafayette Dam will likely safely impound the reservoir under earthquake loadssimilar to or less than postulated. However, because of the history of the dam, we believethat simplified analysis procedures, which are often conservative when a large margin ofsafety is available, may not be sufficient to demonstrate that large non-recoverabledeformations could not occur under the worst conceivable earthquake scenarios and,especially, considering that the embankment fill and foundation materials that failed duringconstruction were never removed. A 210-feet wide crest, as is present, improves the safetyGEI Consultants - 9 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05of Lafayette Dam. However, because of the uncertain extent of potentially weakerfoundation materials, the wide crest may not necessarily guarantee that slope movementswould be limited to the vicinity of the dam slopes. Earthquake-induced movements couldpotentially involve a large volume of soils. The 1928 failure caused large crest settlements(16 to 24 feet) over the entire width (about 160 feet) of the construction crest of theunfinished dam. The present crest (210-feet wide) is only about 30 percent wider than wasthe top of the unfinished embankment that experienced failure in 1928. Although the loadconditions would be different from the previous failure, it is conceivable, although unlikely,that the existing dam crest could settle and the embankment and foundation alluviumexperience large deformations, in a pattern similar to 1928, if the site were to experienceextreme seismic shaking.Finally, we compared estimated field cyclic stress ratios (CSR) for an “equivalent referencemagnitude” (M 7.5), based on H.B. Seed’s simplified method as updated by Idriss (1999),with laboratory cyclic stress ratios causing 10 percent axial strain in 15 uniform stress cycles.In the case of the Hayward Earthquake (M w 7.25), many equivalent field CSRs equal orexceed the laboratory CSRs causing 10 percent axial strain or greater, after correction forfield condition, for the applicable number of cycles. Such comparison suggests thatsimplified procedures may not be sufficient to fully assess the seismic performance ofLafayette Dam. Earthquake-induced deformations larger that those computed might occur,under some of the severe earthquake scenarios postulated, and may need to be furtherevaluated. This suggestion is based on our review of the 1928 failure, rather than on thecomputed deformations, which would likely be acceptable considering the large freeboardavailable, had Lafayette Dam not experienced an extensive historic slope failure.1.8. Operation, Maintenance and Monitoring DataOperation and maintenance of Lafayette Dam are considered adequate. Instrumentationincludes 24 crest monuments that were installed in grid pattern on the embankment, and aremonitored about every year, occasionally twice a year. In the last fifteen years, maximummeasured horizontal and vertical displacements have been less than 3.5 inches. Time versusdisplacements graphs for individual monuments show stable movements, with no significantincreases or adverse trends in recent years.The dam is equipped with 18 active open-standpipe piezometers, monitored about once amonth. Most piezometers were installed in 1965 or in 1973-1974 as part of safetyinvestigations of the embankment. They replaced original observation wells installed duringand shortly after the dam construction. The wells and some piezometers have frequentlyindicated erratic water levels, higher than expected based on reservoir surface elevations, orhave shown fluctuations indicative of questionable functioning. Especially, seven of theeight crest piezometers consistently indicated water levels higher than the reservoir level.Five were replaced in 1992, and another in 1996.GEI Consultants - 10 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05In recent years, nine piezometers still indicate seasonal water levels higher than expectedfrom the corresponding reservoir levels. The others indicate no unusual trends orfluctuations. It has been suspected that surface water runoff from the wide dam crest afterhigh rainfall and seepage or runoff from the abutments affect piezometric readings.Overall, for a dam built with an “impervious” core, the phreatic surface is unusually high inthe downstream shell of Lafayette Dam. This dam behaves as a homogeneous dam regardingthe position of the phreatic surface. This has not been a problem, because the embankment isvery wide and the seepage is collected in drains. The similar permeability characteristics ofthe core and shell materials or perhaps seepage from the abutments may contribute to suchobservation.Seepage through the dam is collected by tunnel and embankment subdrains. A 24-inchconduit, which was installed in a 60-inch diameter concrete conduit, runs along the leftabutment of the dam. Outlet tunnel leakage is collected in a sump box, located near the westend of the toe of the dam. Seepage through the dam is collected by a pipe subdrain system,which runs perpendicular to the dam axis and along the toe of the dam. Seepage collected inthe pipes is evacuated through a seepage collection box, located near the right abutment atthe toe of the dam. Seepage is regularly monitored by the District and by the DSOD atapproximately bi-yearly inspections. Our review of these inspection reports and correlationsestablished over a 10-year period indicate that tunnel and subdrains flows are typically verylow, ranging from near zero to about 10 gpm for the toe drain, and from less than 1 gpm to 5gpm for the tunnel flows.1.9. Conclusions and RecommendationsThe purpose of this investigation was to perform a detailed review and update of the seismicstability of Lafayette Dam. We carefully reviewed existing data made available to theproject team and updated the seismic requirements. Our conclusions and recommendationsrely solely on a reinterpretation of such information, on a review of applicable literature, onvisual inspection of the dam and site, and on simplified slope stability and dam deformationanalyses.The watershed area is very small (1.34 square miles). The dam crest and surrounding groundsare used as a regional park, and day use of the facilities is extensive. The reservoir ispresently used for emergency backup water supply and recreation, and is only subject tominor annual fluctuations. A substantial freeboard (17.8 feet) is maintained under normaloperating conditions.As previously mentioned, the dam experienced in 1928 a major downstream slope failureduring construction, and was built to a lower crest elevation than originally designed. TheGEI Consultants - 11 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05downstream slope was flattened by placement of additional fill, but the failed central sectionof the dam was left in place. The top of the failed embankment was kept near its incompleteheight and graded to form the present crest. The dam has satisfactorily performed sinceconstruction ended in 1933. A detailed seismic evaluation of this dam was performed in1976, using procedures then applicable, and concluded that it would be safe under thepostulated earthquake scenarios.We found no visual signs of deterioration, instability or inadequacy of the embankment.Seepage is low, consistent with the norms for the dam and the reservoir level, and is notdetrimental to the safety of the dam. The instrumentation is regularly monitored. Thephreatic surface within the embankment is high and indicates that Lafayette Dam performssimilar to a homogeneous embankment rather than as a zoned dam with an impervious core.This is probably because there are little differences between the physical characteristics ofthe core and shell materials, which all came from nearby borrow sources.The clayey nature of the embankment and foundation materials and other physical properties,such as liquid limit, plasticity index and water content, indicate that they are unlikely toliquefy and experience instantaneous loss of strength as a result of earthquake loading.Lafayette Dam is well maintained, in good visual condition, and the District should becommended for the obvious care it has given to this facility over the years. The very widecrest (210-feet) and the large freeboard significantly contribute to improving the safety ofthis dam.Considering the “high risk” rating of Lafayette Dam and its seismic exposure to the Haywardand other regional faults, we have found no condition that would require immediate action.However, based on our review and the simplified analyses performed, we recommend thatthe District consider implementing several action items to confirm the predicted seismicperformance of this dam.Lafayette Dam is sited within a complex geologic environment, and numerous old landslidessurround the reservoir. The dam was built on a thick layer of alluvium, up to 100 feet deep,and averaging 90 feet in its central portion. The dam itself was well constructed but,according to the post-failure investigation (Consulting Board,1929), the foundation alluviumwas the primary cause of failure during construction, in response to loads applied by perhapstoo rapid placement of the embankment materials. The 1928 failure was massive, andinvolved a considerable volume that included part of the core and downstream shell, as wellas the underlying foundation alluvium over most of its 90-feet thickness. The failedmaterials were not removed, were not strengthened through compaction, consolidationgrouting or other soil improvement techniques, and construction was completed by simplyleveling the dam crest, backfilling failure cracks and flattening the downstream slope byplacement of additional fill.GEI Consultants - 12 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05The upstream slope and impervious core, which experienced a substantial spread in crosssectionalshape (about 24 percent spread of its designed base width) during the failure, wereleft as originally built despite the significant movements and probable disturbances theyexperienced. The central portion of the upstream slope settled by about four feet in the yearsthat immediately followed the end of construction. The failed materials have stabilized, andundoubtedly consolidated and regained strength over time. Under static loading, possiblereactivation of movement along the 1928 failure surface, which was not clearly identified inthe borings drilled in the 1966 and 1976 investigations, is unlikely. The affected materialsshould have had ample time to stabilize in the 76 years since the failure occurred and most ofthe old failure surface has probably regained strength. Horizontal and vertical crest or slopemovements have been insignificant for many years, and are continuously monitored. Yet, it isnot clear whether the foundation alluvium has regained sufficient additional strength towithstand major earthquake loads. The foundation soils that failed in 1928 under rapidlyapplied construction loads experienced very large movements, and have neither beenremoved nor improved following the failure. This leaves open the question whether thepresence of the previous failure surface might affect potential deformations of the foundationsoils and embankment under rapidly applied, major seismic loads.Recent geologic literature mentions discontinuous lineaments and possible inferred faultingin the immediate project vicinity. Such features do not seem to represent any threat to thedam. The inferred faults are short, and the one that has been shown to potentially intersect thedam footprint is parallel, rather than perpendicular to the dam crest. Hence, if confirmed, itwould unlikely be critical in terms of direct or sympathetic relative movements, because ofits short length and favorable orientation. Additional field investigation of the inferred faultis not required because, should it be confirmed, it is not a seismogenic structure andsympathetic rupture would be very small and of no consequence to Lafayette Dam.Many old landslides along the reservoir rim are in a low slope position and presentlyinactive. One of these, however, is adjacent to the east margin of the dam and might impactthe lower portion of the embankment if it were to be reactivated as a result of strong groundshaking. Evaluation would be desirable if any signs of reactivation are observed in futureinspections. Such evaluation would include preparation of improved geologic maps andcross sections of this landslide to establish probable volume, dimensions and possibledirections of any movements.Existing field penetration data and laboratory test results in the alluvium below thedownstream slope of Lafayette Dam indicate that these materials are the weakest. Theinformation available is limited because of the wide spacing of the upper and lowerdownstream berms, which controlled the locations of the previous borings. As was the casein 1928, such materials may control the overall stability of Lafayette Dam. Further field andlaboratory testing in that area could be considered, including field exploration methods thatwould facilitate recognition of any thin and potentially weaker zones such as the old failureGEI Consultants - 13 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05surface. Cone penetration testing (CPT) and in-situ shear vane testing, combined withreference SPT testing, could be a rapidly implemented and cost-effective way to performsuch investigations.The response spectra developed for the Hayward and Lafayette-Reliez Valley faults, whichare potentially the most critical to this site, are significantly more demanding in the range ofperiods of interest to the response of Lafayette Dam and foundation than the responsespectrum of the acceleration time history used in 1976 to represent a “maximum probable”Hayward Earthquake. We suspect that the equivalent-linear dynamic analyses that were thenperformed did not fully capture the anticipated response of Lafayette Dam, due to thecoarseness of the finite element mesh used to represent the dam and its foundation.Our “best” estimates of potential earthquake-induced deformations, for the most criticalearthquake scenarios, suggest maximum average crest settlements on the order of 2 to 3 feet.Hence, Lafayette Dam is likely to maintain a large freeboard (greater than 15 feet).However, based on the yield acceleration (0.14g) computed for the most critical of thehypothetical failure surfaces considered, which would involve both the downstream shell ofthe dam and the underlying foundation alluvium, simplified methods of analysis predict awide range of deformations. Three of these methods lead to estimated upper bound crestsettlements (due to combined embankment slope and foundation deformations) that rangebetween 4 and 7 feet. The simplified Newmark and Seed-Makdisi procedures lead to upperbound combined slope deformations on the order of 8 to 13 ft for the downstream slope sideof the dam and underlying foundation. Furthermore, simplified methods of analysis are notalways conservative.More detailed exploration and testing are desirable, primarily in the portion of the foundationalluvium, below the downstream slope of the dam, which has not been explored in the courseof past safety evaluations. If strength properties higher than assumed were established forthat portion of the foundation alluvium, no further analysis would be required. However, ifstrength properties lower than those assumed were to be encountered, and because of thehistory of Lafayette Dam, we would suggest that EBMUD consider a detailed reanalysis,using updated material properties, modern computational techniques, updated accelerationtime histories consistent with the recommended response spectra, and constitutiverelationships that would simulate the behavior of the embankment and foundation materialsmore reliably than was possible in 1966 or 1976.In conclusion, Lafayette Dam is a well-maintained facility, and is reasonably safe for theMCE. However, because of the dam’s failure during construction, the District shouldsupplement the existing information regarding the downstream foundation alluvium.GEI Consultants - 14 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/052. DESCRIPTION OF PROJECTFEATURES2.1 GeneralThe following description of the project facilities is excerpted from the previous reports(Shannon & Wilson, 1966; W.A. Wahler & Associates, 1996; DSOD, 1980) with relevantupdates to reflect our observations from the current field inspection and data review.Lafayette Dam is owned by the East Bay Municipal Utility District (the District, or EBMUD)and is located about 1.6 km west of the center of the City of Lafayette, in Contra CostaCounty, CA. A map of the vicinity is presented on Figure 2-1. The dam was built acrossLafayette Creek, a small tributary of Las Trampas (Walnut) Creek that flows in a northerlydirection. The watershed has a very small drainage area, 1.34 square miles, and lies withinthe boundaries of the East Bay Regional Park System. The dam can be accessed at all times,and its crest is paved and actually used as a parking lot for visitors to the park.Unless otherwise stated, all elevations in this report are in feet and refer to United StatesGeological Survey (USGS) Mean Sea Level Datum (1929), which is referred to in someproject-related documents as National Geodetic Vertical Datum (NGVD-29).Based on DSOD and National Inventory of Dams (NID) classification guidelines, LafayetteDam is classified as a “large” (higher than 100 ft) and has a “high” hazard potential, becauseof its closeness to the City of Lafayette and densely developed areas. There have beennumerous recent commercial and residential developments downstream, which confirm thehazard classification of the dam.2.2 Embankment DamLafayette Dam was built from August 1927 to 1933 and was raised one foot in 1967 duringregrading of its crest. Plans and sections of the major project appurtenances are presented onFigures 2-2 and 2-3. The dam is a rolled zoned earthfill embankment, 132 feet high, andwith a crest 1,200 feet long. The crest of the dam is at El. 467. The upstream dam face has aslope of 3H to 1V (horizontal to vertical), with two 15-foot wide berms originally designed atEl. 450 and El. 400, respectively. The upstream face berms experienced up to about 1.5 feetof upward (lower berm) or downward (top berm) movement during the construction failurein 1928 (see Section 3.1.2). The upper berm settlement is visible on photographs (seeAppendix C), and appears to have increased since the end of construction. The downstreamGEI Consultants - 15 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05slope varies at 2.5H to 1V, 4.0H to 1V and 8.0H to 1V and has a 10-foot wide berm, locatedat El. 400. The dam crest is 210 feet wide. Most of the dam is founded on valley fillalluvium, underlain by soft sedimentary bedrock of the Tertiary Orinda Formation.The dam has a central “impervious” clayey core with upstream and downstream slopes,originally built at 0.5H:1V. The core slopes were also affected by the 1928 failure, seeSection 3.1.2. The constructed top width of the core is about 22 feet. A cutoff trench wasexcavated into the alluvium during the early phase of the dam construction. A steel sheetpile curtain and a short concrete cutoff wall (pile cap) were installed in the cutoff trench,before placement of the core materials.Principal project data regarding Lafayette Dam were obtained from the National Inventory ofDams (NID) and the Division of Safety of Dams (DSOD). The most important of these dataare summarized in Table 2-1.2.3 Spillway and Outlet WorksThe dam does not have a channel spillway, although adding one was considered in 1956(Dukleth, 1956). A reinforced concrete outlet tower, approximately 120-foot high, with anannular cross-section, provides reservoir drawdown capacity. The top platform of the toweris at El. 500, and the tower is free-standing from about El. 388. The interior space of theoutlet tower is separated into a spillway chamber portion and an outlet chamber portionthrough inner reinforced concrete partition walls. The top rectangular port in the tower (2.5’x 3’) is ungated and provides limited discharge capacity, with a spillway crest elevation at El.449.2. Spillway flow exits the tower through a 60-inch diameter conduit located at the baseof the structure. This conduit has a total length of 1,845 feet and terminates in a baffle box.Reservoir control is provided by four rectangular ports in the tower wall, of same size as thespillway port (see Table 1) and equipped with slide gates. The outlet portion of the tower isconnected to a 60-inch diameter conduit, which extends downstream of the dam toe. Fromthereon, outlet discharge is conveying through as 24-inch diameter steel pipe located a 60-inch diameter concrete conduit. The capacity of the spillway is very small and was showncapable of barely passing the Probable Maximum Flood in earlier studies. According to the“EBMUD Dam Guide” for Lafayette Dam, the District plans to modify the spillway toincrease its capacity in the future.International Civil Engineering Consultants, Inc. evaluated the structural performance of theLafayette outlet tower (ICEC, 1995). The structure was found to have insufficient seismiccapacity, and six conceptual upgrade alternatives were proposed. In a letter dated August2002, the District submitted to the DSOD documents and sketches describing its preferredupgrade alternative, which would consist of infilling the tower with mass concrete betweenGEI Consultants - 16 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05El. 379 and El. 432 (except for space used for three new 24-inch diameter spillway and outletconduits).2.4 Standard Operational ProceduresLafayette Reservoir is normally operated as a standby emergency storage and recreationreservoir and is maintained near its maximum operating level (El. 449). Personnelsupervising the operations of the East Bay Regional Park System (EBRPS) are continuouslypresent at the park headquarters during the day. District operation and maintenanceengineers and patrolmen frequently inspect the dam, with quarterly inspections by a DistrictSupervisor. Unscheduled inspections are performed on an as-needed basis, e.g., after any feltearthquakes in the vicinity. The DSOD has semi-annual scheduled inspections of the project,with inspections of the dam and spillway tower performed at regular intervals. Because thesurrounding watershed area is small, the potential for reservoir siltation and filling by debrisor sediments is low. There are no minimum downstream flow requirements for LafayetteReservoir, as it is filled with water imported from the Mokelumne Aqueduct.2.5 InstrumentationInstrumentation includes 24 crest monuments that were installed in grid pattern on theLafayette embankment, and are monitored about every year, occasionally twice a year. From1989 to present, maximum horizontal and vertical displacements recorded from all themonuments have been 3.36 inches and 3.12 inches, respectively. Review of the time versusdisplacements graphs for individual survey monuments show that slope horizontal andvertical movements are stable with no significant increase in recent years. Additional detailsare provided in Appendix A, Instrumentation Review.Originally installed observation wells and several older piezometers have been replaced byeighteen currently active standpipe piezometers, which are regularly monitored about onceevery month. Other instrumentation includes instruments to monitor flows from the outletconduits and seepage from the embankment toe drain. The District annually provides to theDSOD plots of horizontal and vertical movements, measured at the crest monuments, andpiezometric records. We reviewed such data as part of this investigation.Seepage is collected by tunnel and embankment subdrains. A 24-inch conduit, installed in a60-inch diameter concrete conduit, runs along the left abutment of the dam. Tunnel leakageis collected in a sump box, located near the west end of the toe of the dam. Seepage throughthe dam is collected by a subdrain system, composed of 6-inch and 8-inch pipes, which runperpendicular to the dam axis and along the toe of the dam. Seepage collected in the pipes isevacuated through a seepage collection box, located at the dam toe near the right abutment.GEI Consultants - 17 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/053. SUMMARY OF CONSTRUCTIONHISTORY AND OPERATIONThe following summary of the construction history and operation of the Lafayette Dam isexcerpted from the previous reports with relevant updates to reflect our observations fromthe current field inspection and data review.3.1 Design and Construction History3.1.1 Original Design and Initial ConstructionLafayette Dam was originally designed as a 140-foot high dam, with a 32 feet wide and1,400 feet long crest at El. 500. Dam zoning consisted of a central impervious “clay” core(referred as Zone 1 in this report), a “porous” downstream shell (Zone 2) and an upstreamshell to be built of “selected impervious” materials (Zone 3). The upstream slope wasdesigned at 3H:1V, with two 15-foot berms at El. 400 and El. 450, respectively. Thedownstream slope had one berm at El. 430 and a slope of 2.5H:1V above the berm, and3H:1V below. According to original construction drawings, the clay core was originallydesigned with upstream and downstream slopes of 0.5H:1V. Construction of the dam startedin August 1927.Dam construction started in August 1927 with the stripping of the creak banks (placed at thedam toe) and excavation of the cutoff trench. The trench was excavated a maximum of 20feet in the channel section. The materials to build the embankment were obtained from thebasin and side hills upstream of the dam and rolled in 12-inch lifts (less clayey materials) or8-inch lifts (more clayey materials). Although Lafayette Dam was designed as a zoned dam,all construction materials were “clayey in character” (Consulting Board, 1929). The Zone 1core materials were reported to contain “about 70 percent clay” and were obtained “fromborrow pits in the reservoir area”. The core “ is largely composed of clays from thefoundation alluvium”. The Zone 2 (D/S) materials were select fill from the OrindaFormation, described as “the most granular” but “with an average of about 20 percentclays”. The Zone 3 (U/S) materials “contain about 45 percent clay and 20 percent passingthe No. 14 sieve”, and are “practically watertight”. The dam was built over 50 to 90 feet ofalluvium, consisting mostly of clays, overlying the Orinda Formation. The upper 10 to 15feet of the alluvium were describes as “dry and firm”, while deeper alluvium was “moist andmore or less plastic”.GEI Consultants - 18 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Original design included a curtain of steel sheet piling, driven into the core trench beforeplacement of the core materials, and capped by a concrete cutoff wall. Based on a drawingprepared by U.S. steel, the pile curtain in the central portion of the dam (between Stations11+20 and 12+75) was mostly constructed of 45-foot long piles (M-105 or equivalent), witha few 60-foot long piles at about Station 11+80. The length of the piles was irregular andprogressively shorter toward the abutments, between Stations 9+50 and 11+20 at the leftabutment, and between Stations 12+75 and 14+30 at the right abutment. The concrete corewall is shown on existing drawings to be 12-foot high, 4-foot wide at the base, and 2-footwide at the top in the central portion of the dam (e.g. Station 12+00), but was 16-foot highnear the abutments, where no piles were driven (e.g. Station 9+00). Tile and gravelfoundation drainage was also provided at construction to a depth of about 5 feet in thedownstream part of the foundation. A concrete facing was also placed along the entireupstream face.3.1.2 The 1928 Construction FailureOn September 17, 1928, the crest had reached between El. 476 to El. 478, based on availabledata (or a dam height of between 116 feet and 118 feet at centerline). This was about 22 feetbelow the intended final height. The reservoir was at El. 392 on that day. Longitudinalcracking was first observed along the crest, accompanied by bulging of the ground surface atthe downstream toe. Additional cracks occurred in the following days, followed byprogressive failure of the downstream slope. Figure 3-1 shows an aerial view of the damtaken shortly after the failure. Figure 3-2 shows the outline of the failed zone and surfacemovements on a proposed reconstruction plan for the failed dam. The failed portion of theembankment reached a stable position on September 28, 1938. Hence, slope failuremovements lasted about eleven days. Based on cross-sections of the failed dam preparedafter the failure, see Figure 3-3, the top of the dam had dropped a maximum of 24 to 26 feetin its central portion (up to 22 percent of the constructed height), and across a width of about525 feet. The area affected by this movement was about 8 acres (340,000 ft 2 ). Maximumsettlement of the downstream edge of the constructed crest was about 20 feet.The top of the upstream shell, in-between the core material and the upstream edge of thecrest, settled about 10 feet and experienced severe cracking. The lower portion of thedownstream slope moved about 40 feet in the horizontal direction, while the upstream toemoved outward about 5 feet. A bulge or ridge of squeezed foundation soils, about 20-foothigh and up to 80-foot wide perpendicularly to the dam crest, formed along the edge of thedisplaced downstream toe. Although the upstream slope was considerably less affected thanthe downstream slope, it experienced curving of its berms at El. 450 and El. 400 withmaximum horizontal displacements of about 3.8 feet and 5.5 feet, respectively. No bulge wasobserved at the upstream toe. The upper berm was reported to have experienced a maximumof 1.5 feet of downward movement, while the lower berm experienced upward movement,with a maximum vertical displacement of 1.5 feet. Some of the concrete slabs wereGEI Consultants - 19 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05displaced, locally pulled apart, or thrust up to one foot on top of each other. A 1954 study ofsettlement records by the DSOD (E.V. Poe, Memorandum of 1/15/1954) suggested a markedincrease in the rate of observed settlement between 1949 and 1953 and that settlementrecords “did not show a smooth but humps and dips at different times”, reflecting influenceof changes in reservoir elevation and perhaps continuing consolidation of the failedmaterials. It should be noted that during our field inspection, we noticed that the upperberm of the upstream slope is curved downward by about 4 feet, suggesting further slopeadjustments in the years that followed the failure.A thorough investigation of the failure, including 23 borings (reported as 16 in some of thereports) along a line perpendicular to the dam axis (at Station 11+58) through the failedportion of the dam and foundation, was completed. This investigation focused on thealluvium material. It concluded that, for that material, “its most noticeable features as awhole were its plasticity when wet or moist and its lack of distinct and persistentstratification”. The investigation also concluded that the top of the foundation soils, belowthe failed portion of the dam, had settled up to a maximum of about 9 feet under thedownstream half of the core.The exploratory borings also indicated that the upstream corner of the base of the core hadmoved about 8 feet toward upstream, while the downstream corner had moved about 30 feettoward downstream, based on the contacts between core and shell materials interpreted fromthe borings. The two slopes of the core, after the failure, were found to be at approximately0.85H:1V (downstream slope), and 0.7H:1V (upstream slope), hence were substantiallyflatter than the designed (and presumably constructed) slopes present at the time of failure.Such finding, combined with the extensive measured crest settlements and measured slopemovements, suggest that the entire core of the dam and downstream shell experienced largenon-recoverable displacements and substantial remolding during the failure. The same intrue for the portion of the foundation soils located below the core and original displaceddownstream shell. Hence, the extent of the 1928 failure involved most of the core anddownstream shell materials, and supporting foundation soils.Perhaps ten longitudinal cracks and crevices, with depths of between 10 and 18 feet andwidths of one to four feet, are shown on a cross-section of the dam prepared after the failure(see EBMUD Drawing DH 1608-18, dated January 1929). Cracks ranging from hairline toup to 1.5 inch wide were also reported in the outlet conduits, but no change in grade oralignment of the conduits was reported.Water levels were measured in some of the borings drilled before the 1928 failure, andsuggested pore pressures increase averaging 1.35 times the height of the embankment at itspre-failure stage of completion. Hence, it was suspected that high pore pressures existedwithin the foundation alluvium at the time of failure. The post-failure borings did notGEI Consultants - 20 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05encounter any phreatic surface within the embankment materials, but a very small reservoir(El. 392) was already impounded.Additional details are provided in the report prepared by the Consulting Board hired in 1928by the District to investigate the cause of failure (Consulting Board, 1929). The Boardconcluded that the dam failed as a result of plastic flow within the foundation soils and in themost heavily loaded region. Some of these materials “were squeezed out toward the regionof least pressure”. “The dam would have shown no weakness had it rested upon a firmfoundation”. The exceptional conditions of the foundation alluvium that led to the damfailure were its “uncommon thickness and general plasticity”. It was also concluded, fromthe location of the displaced alluvium foundation surface and comparisons of failed materialvolumes, that significant consolidation had been experienced by the foundation soils duringthe failure. The Board report concluded that the dam was well built and that the failure wascaused by the sole foundation conditions. We did not find, in the materials reviewed, anysuggestion that the occurrence of the failure could have been related to the constructionsequence or the steeper D/S slope of the original design.Subsequent field investigations performed in 1976 by W.A. Wahler & Associates concludedthat a substantial downstream movement of the core wall had been associated with the 1928failure. The total movement of the core wall was then estimated to be 12 to 14 feet, and hadpreviously been undetected. Hence, the core wall and sheet piles may no longer function asan impervious barrier as originally intended, which may contribute to the high piezometricwater surface observed within the dam section.3.1.3 Revised Design and Final ConstructionIn 1929, the Board of Consultants who investigated the 1928 failure recommendedreconstruction of the dam with a crest width of 70 feet, at El. 460, and with flatter upstreamand downstream slopes of 5H:1V and 7H:1V, respectively, see Figure 3-2. However, thedam was actually redesigned and rebuilt using a less conservative (steeper) downstream slopethan recommended by the Review Board, but keeping the repaired and re-leveled wide top ofthe failed dam (220 feet wide) as a new crest. The Board accepted such redesign andconstruction was completed in 1933. The present downstream slope is flatter than that of theoriginal dam and, except for a very flat (8H:1V) bottom section below El. 380, consists of atop section at 3H:1V and an intermediate section at 4H:1V, separated by a 10-foot wide bermat El. 430. The original (but displaced) upstream slope was essentially left intact. Only theslope protection concrete slab facing was repaired (EBMUD, 1957).Hence, an essential aspect of the reconstruction process is that most of the failed materialswere neither removed nor recompacted. Cracks and scarps were filled at the dam crest. Thesurface of the crest of the failed embankment was regraded and the bulged foundation soil atthe downstream toe were removed, but the dam was redesigned simply by flattening theGEI Consultants - 21 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05downstream slope by placement of additional material and keeping a very wide crest at El.466. During additional grading in 1967, the dam crest was raised one foot to its presentelevation (El. 467). The present crest width is 210 feet.3.2 Dam OperationThe originally planned dam (crest El. 500) was originally intended to provide storage forabout 10,590 acre-feet for water pumped into the reservoir from the Mokelumne Aqueduct.Following reconstruction after the 1928 failure, the maximum storage capacity wasconsiderably reduced, due to the lower crest and spillway elevations. Furthermore, thereservoir water was prudently kept at reduced levels for an extended period of time.Over the years, the water was successively raised to El. 410, El.428 and, finally, to El. 448 in1937. Reservoir level has been restricted since that time to about that level and the presentwater level fluctuates between El. 441 and 449.2 (spillway crest). Current maximum storageis 4,250 acre-feet. The reservoir functions as a standby emergency storage facility to be usedin case of disruption in the Mokelumne Aqueduct Water Transmission System.It should be noted that Lafayette Reservoir, among those owned and operated by the District,is the only one that relies on the outlet tower for spillway functions.GEI Consultants - 22 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/054. GEOLOGIC AND SEISMICCONSIDERATIONSThe geologic and tectonic environment of the project area were reviewed using recentlypublished maps and reports on geology in the area, analysis of geomorphology profiles, airphoto interpretations, and geologic field mapping of the dam and reservoir area. Criticalfaults were identified, based on review of applicable literature and recent research on thesubject.4.1 Regional GeologyLafayette Dam is located in the East Bay Hills, a northwest-trending topographic uplandwithin the Coast Range geomorphic province, see Figure 4-1. The East Bay Hills consist ofTertiary-age sedimentary and volcanic rocks underlain by older plutonic and metamorphicrocks. The area is characterized by northwest-trending bedrock ridges and interveningvalleys filled with Quaternary sediments of varying thickness.The reservoir area is characterized by moderately steep to very steep (up to 35 degreesinclination) hillside terrain. Lafayette Dam is constructed across the neck of a small,northeast-trending valley eroded into Tertiary-age sedimentary rocks of the Contra CostaGroup (Orinda Formation). Fine-grained alluvial materials underlay the dam foundation.These clayey materials were likely deposited in a ponded depression setting, whichcontributed to their limited strength. The bedrock in the vicinity of the dam and reservoirconsists of interbedded conglomerate, sandstone, siltstone and claystone. These rockstypically are poorly cemented. However, hard and well-cemented units are locally present,and form ridges in the vicinity with dominant northwest-trending strike. Quaternary-agestream alluvium underlies the reservoir valley and central portion of the Lafayetteembankment. Only a thin mantle of soil and colluvium underlies ridges and most hillsides;however, these surficial materials are thicker in valleys and drainage swales. Landslides arepresent locally on many hillsides (see “Landslides” section).4.2 Geologic Structure and Tectonic SettingThe East Bay Hills are situated between the active Hayward and Calaveras faults, and havebeen strongly deformed by late Cenozoic folding and faulting associated with plate boundarytranspression, see Figure 4-1, where the regional faults are plotted. Traditionally, tectonicstructures within the East Bay Hills have been considered to be mostly compressional.However, re-evaluation of tectonic features in the East Bay Hills indicates that a significantGEI Consultants - 23 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05dextral (to the right) movement has also occurred. Unruh and Kelson (2002) propose a leftstepping,“restraining bend” step-over from the northern Calaveras Fault to a broad zone ofdistributed dextral slip within the East Bay Hills, either partly or wholly along previouslymapped faults and newly identified north-trending lineaments. Tectonic strain in thenorthern East Bay Hills may be accommodated by a complex interaction of dextral slip andreverse slip structures. The principal faults in the vicinity of Lafayette Dam are listed inTable 4-1.In the vicinity of Lafayette Dam, the Tertiary-age rocks are folded into parallel to subparallel,west- to northwest-trending anticlines and synclines with limbs characterized bymoderate to steep (30 to 65 degrees) northeast and southwest dips. Geologic maps byGraymer, et al. (1994) and Haydon (1995) depict a previously unidentified, possiblyquestionable fault extending below the dam (parallel to the crest). As mapped, the fault traceparallels the local fold axes to the east of the dam, but bends southward and offsets fold axesto the west of the dam. According to Graymer, (personal communication, May 2004), thisfault was inferred on the basis of the apparent truncation of two northwest-trending fold axesby Wagner (1978). Wagner conducted regional mapping of 12 quadrangle maps as part ofhis 1978 Ph.D. Research. His map displays the fault as a possible subsurface feature thatwould help explain an apparent change in structural trend across that location. As depicted inGraymer, et al., this inferred fault might connect to the northwest-trending Pinole fault,which is currently considered conditionally active.We conducted a reconnaissance of the aforementioned mapped fault trace and surroundingterrain. At the western termination of the mapped trace (near the inferred junction with thePinole fault), we observed an apparent faulted contact, with an east-west trend and steepnorthward dip (N75E, 50 degrees NW), between conglomerate and mudstone units. Theexposure appears to support the presence of an east-west trending fault at that location;however, it does not explain the approximately 1-mile-long northward bend to the dam, andwe did not observe obvious truncation of northwest-trending fold axes. We note that a mapof the same area (Dibblee, 1980) indicates geologic structure to be consistent across the faultmapped by Graymer and does not show that fault. A bedrock orientation (northeast strike,dipping 47 degrees southeast) shown on Wagner’s 1978 map, approximately 4,000 feet westof the dam, could be interpreted to represent a change in bedrock structure. However, itappears from our observations that this plotted bedrock attitude is located within a landslideand is likely not a reliable indicator of local structure. The lack of sufficient bedrockexposures along the remainder of the mapped trace does not allow us to confirm or reject thepresence and mapped location of that inferred fault and, if such presence was confirmed, toassess whether it is active or not. The inferred fault, would it be confirmed, is notseismogenic and could only rupture due to movement on a nearby active fault (sympatheticfaulting). The rupture would be small, less than a few inches, and favorably oriented withrespect to the dam, which has a size and configuration capable of readily accommodatingsuch unlikely movement. Hence, no additional investigation of this inferred fault is justified.GEI Consultants - 24 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/054.3 Recent Faulting and SeismicityLafayette Dam is located in a seismically active area, between the historically activeHayward Fault (8.8 km to the southwest) and the Calaveras Fault (9.8 km to the southeast).These distances were interpreted by our geologist and are shorter than the 9.6 km (Hayward)and 11.5 km (Calaveras) previously assumed for these faults by the DSOD (GeologicReview, J.L. Lessman, 2003). W.A. Wahler & Associates (1976) estimated the distancefrom the dam to the Hayward Fault as 10 km, but assumed the Calaveras Fault to be only 6.4km away, which cannot be substantiated.The northern Hayward Fault is a well-defined, active tectonic feature along the westernmargin of the East Bay Hills. In contrast, the Calaveras Fault, which is well definedgeomorphically to the south, becomes poorly defined and less active toward the north.Dextral slip north of the Calaveras Fault is likely transferred to multiple north- to northwesttrendingfaults located to the northwest or northeast. Traditionally, geologists have assumedthat dextral slip along the Calaveras Fault was being transferred to the Concord Fault, locatedto the northeast. However, there is a growing awareness that dextral slip may be transferredto the west via left-stepping, north-trending faults within the interior of the northern East BayHills (Taylor, 1992; Unruh and Kelson, 2002). Several northwest-trending linear valleysextend into the northern East Bay Hills from the town of Alamo, near the interpreted northernend of the active Calaveras Fault. These valleys, and related lineaments, appear to mergewith previously recognized faults and folds (e.g., by Saul, 1973; Dibblee, 1980; and Crane,1995), and newly identified lineaments and faults (Unruh and Kelson, 2002).The northern East Bay Hills are characterized by a moderate level of seismicity, compared tothe stronger level of seismicity associated with the dominant fault and fold zones in the SanFrancisco Bay Area. The 1977 sequence of moderate-sized earthquakes near BrionesReservoir (called the “Briones swarm”) may be typical of the style of earthquakes that occurin stepover zones between the endpoints of major faults (Oppenheimer and Macgregor-Scott,1992). Other “swarms”, including in 1970, 1976, 1990 and 2002, have occurred in theDanville, Alamo and San Ramon areas, near the northern termination of the Calaveras Fault.During the last twenty years, two significant earthquakes were experienced in the GreaterBay Area, but had no impact on Lafayette Dam. These were the April 24, 1984 Morgan Hill(M 6.2), and the October 17, 1989 Loma Prieta (M w 6.9). Other earlier historic events ofpotential significance to Lafayette Dam were the October 24, 1955 Concord/Walnut CreekEarthquake (M 5.4), a June 1, 1911 earthquake (estimated M 6.6), presumed centered alongthe Calaveras Fault, the 1906 San Francisco Earthquake (estimated M w 7.9), and two largeearthquakes centered along or near the Hayward Fault (1836 and 1868). The 1955earthquake was centered about 12 km away from the site. In the last century, there has beenGEI Consultants - 25 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05about 20 earthquakes of assigned magnitude greater than 4 within a 30 km radius from thesite.The faults of significance to the project site are described in the following paragraphs.Maximum earthquake magnitudes were obtained using empirical relationships betweenmoment magnitude (M w ) and fault rupture area developed by Wells and Coppersmith (1994).Hayward FaultThe Hayward Fault extends from Fremont northward to San Pablo Bay, for a length ofapproximately 87 km. The Hayward Fault is considered to be a part of the longer Hayward-Rodgers Creek fault system, which has been divided into three potential fault rupturesegments: Rodgers Creek (north of San Pablo Bay), northern Hayward, and southernHayward (WGCEP, 1999). This fault system is considered to be the primary dextral slipfault in the eastern San Francisco Bay area. Dextral movement is to the right, when lookingacross a strike-slip fault from the stable plate side.Portions of the Hayward Fault experience aseismic creep at about 4 to 5 mm per year, andlocally as high as about 9 mm per year in the Fremont area (Lienkaemper, 1992; Lettis,2001). A large (estimated M w 7.0) historic earthquake ruptured the southern, 50 to 55-kmsection of the fault in 1868, and paleoseismic investigations indicate that between four andseven large earthquakes have occurred on the northern Hayward Fault (considered to be 30 to35 km long) during the past 2,100 years. The long-term slip rate on the fault is estimated tobe 9 mm per year, which represents a “Very High Slip Rate”, per the fault slip ratingclassification and Consequence Hazard Matrix used by the DSOD (DSOD, personalcommunication).Previous segmentation models of the Hayward Fault were based on the extent of the 1868rupture and presumed location of another earlier, large earthquake (1836). The 1836earthquake, however, has recently been relocated (Toppozada and Borchardt, 1998) and isnow considered not associated with the Hayward Fault. The apparent lack of clear physical,geologic or seismic evidence indicating a segment boundary lead us to consider the fullHayward Fault length in our maximum earthquake magnitude estimate (M w 7.25).Calaveras FaultThe Calaveras Fault is a major northwest-striking tectonic structure. It is considered to be anactive tectonic feature for nearly 120 km, from south of the City of Hollister to Danville, andpossibly further north. The fault has been divided into three sections (northern, central andsouthern), based on geologic, geomorphic and seismic data. The Holocene slip rate decreasesfrom south to north along the length of the fault, from approximately 15 to 20 mm per yearalong the southern section to about 14 mm per year along the central section, and toGEI Consultants - 26 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05approximately 5 to 6 mm per year along the northern section (Sims, 1991; Simpson andothers, 1999). Seismic data indicate that dextral slip along the southern and central faultsections is partially transferred westward to the Hayward Fault in the vicinity of CalaverasReservoir. Active tectonic deformation on the northern Calaveras Fault appears to dissipateat its northern end, near Danville (Dibblee, 1980; Hart, 1981; Simpson and others, 1999).The northern portion of the Calaveras Fault, from Calaveras Reservoir to the Danville area, isapproximately 40 km in length, and has been associated with less seismic activity than thesouthern portion.Although there have been micro-earthquake swarms in the general vicinity of the northernCalaveras Fault, as discussed earlier, we found no well-constrained data on the timing of themost recent large earthquake. The northern segment can be considered seismicallydissimilar, compared with the central and southern sections of that fault, due to its lower sliprate and rate of seismic activity. Based on such differentiation, it seems excessivelyconservative to assign a moment magnitude using the entire 120-km fault length (M w 7.5) ofthe Calaveras Fault, but keeping in mind that the June 28, 1992 Landers Earthquake (M w 7.3)involved a rupture over 85 km long, with complex relative movements propagating alongthree different faults and a stepover segment (Lazarte, et al., 1994). Although unique, theLanders experience suggests that a major rupture along a well-defined fault segment couldpropagate to other segments of that same fault or even to nearby faults, increasing the overallrupture length and effective duration of associated ground motion. However, in consistencewith recent studies approved by the DSOD (Olivia Chen Consultants, Inc., 2003), weassumed segmentation of the Calaveras Fault to define our recommended upper boundestimate, M w 7.0, which is based on the 40 km long northern segment. This M w estimate islarger than M w 6.8 assigned to that same segment for near-source classification of faults forthe Uniform Building Code (Petersen, et al., 2000). We noted, however, that the DSODassigned an M w of 7.25 to the Calaveras Fault in its recent geologic review of Lafayette Dam(J.L. Lessman, 2003).San Andreas FaultThe San Andreas Fault is the dominant tectonic structure accommodating right-lateral,translational motion along the boundary between the North American and Pacific plates. Thetotal fault length, from Point Arena southward to the Gulf of California, is on the order of1,100 km. In northern California, the San Andreas Fault can be separated into two sections,the approximately 135-km-long creeping segment (from Cholame to San Juan Bautista), andthe approximately 450-km-long 1906 rupture segment (from San Juan Bautista to PointDelgada in Mendocino County). WGCEP (1999) divides the northern San Andreas Faultinto four discrete segments: North Coast North, North Coast South, Peninsula, and SantaCruz Mountains segments.GEI Consultants - 27 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Lafayette Dam is approximately 39 km from the northern San Andreas Fault; the closestlocation on the fault is near the interpreted break between the North Coast South andPeninsula segments (i.e., Golden Gate). Late Holocene slip rate estimates for these twoclosest segments to Lafayette Dam are 24 mm per year for the North Coast South segment,and 17 mm per year for the Peninsula segment. Paleoseismic investigations appear toindicate that the 1906 event was characteristic of large rupture events along the northern SanAndreas Fault (i.e., M w 7.9 with recurrence intervals on the order of several centuries).Current magnitude estimates for the San Andreas Fault range from M w 7.9 for the entirenorthern portion of the fault to M w 7.1 for individual segments (North Coast South andPeninsula). We have assigned M w 7.9 to the San Andreas Fault in this study.Concord FaultThe Concord Fault extends from Mount Diablo northwestward to Suisun Bay and isapproximately 14 to 24 km long and is about 12.8 km away from the dam. From Suisun Baynorthward, it is presumed that dextral slip along the Concord Fault is transferred to the east tothe Green Valley Fault, which extends to Wooden Valley in eastern Solano County. Thus,the Concord Fault is considered to be the southern segment of the Concord-Green Valleyfault system. We have assigned a moment magnitude M w 6.5 to the Concord Fault, henceappreciably below the M w 7.0 recently estimated by the DSOD (Geologic Review, J.L.Lessman, 2003). However, using the DSOD estimate, this fault would produce at the siteless severe shaking than could be produced by the Hayward or Calaveras faults and,therefore, such issue does not need to be investigated further.The Concord Fault experiences aseismic creep of about 3 to 4 mm per year, which appears toclosely match the geologic slip rate from sparse paleoseismic investigations (Borchardt andothers, 1999; Borchardt and Baldwin, 2001). WGCEP (1999) uses a slip rate of 4 +/- 2 mmper year for earthquake probability scenarios. The largest known historic earthquake (M 5.4)on the Concord Fault occurred on October 24, 1955, and no significant earthquakes (M > M6.0) have been reported along this fault in the past 225 years (Toppozada and others, 1986).Franklin FaultThe Franklin Fault is a southwest-dipping thrust fault that forms the southwestern boundaryof an apparent micro-structural block within the East Bay Hills (Crane, 1995). A structurallyrelated feature, the Southhampton Fault, forms the northeastern boundary of the block.Unruh and Kelson (2002) describe multiple, discontinuous geomorphic lineaments associatedwith the Franklin Fault that appear to suggest dextral motion. Consequently, the FranklinFault (and Southhampton Fault) may be a “pre-existing Tertiary fault that has beensubsequently deformed by dextral displacement” or has actually accommodated both thrustand strike-slip motion along a complex zone of faulting. The Franklin Fault is locatedapproximately 7.5 km from the dam. From the information reviewed, we assigned to theGEI Consultants - 28 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Franklin Fault a moment magnitude (M w ) of 6.75, based on probable rupture length vs.magnitude relationships (Wells and Coppersmith, 1994). The DSOD (2003) also assigned amoment magnitude of 6.75 to the Franklin Fault, but a shorter distance (6 km).Miller Creek FaultThe Miller Creek Fault is one of many recognized thrust or reverse faults within the interiorof the East Bay Hills. This fault is subparallel to the Hayward Fault, and has accommodateda large amount of late Cenozoic contraction. Most of the thrust faults in the area are notconsidered active tectonic features. However, paleoseismic investigation by Wakabayashiand Sawyer (1998) revealed that at least a portion of the Miller Creek Fault displays evidenceof lateral (strike-slip) movement during the late Quaternary. Their southern trench site,located on a north-trending section of the fault, revealed evidence of lateral offset of latePleistocene colluvium. However, a northern trench site, located north of a distinct westwardbend in the strike of the fault, demonstrated that the fault has not displaced late Pleistoceneunits at this second location. Although slip along the Miller Creek Fault might be transferrednorthward to other faults in the area, no breaks in geologic structure suggesting fault offsethave been observed to date.The southern extent and active portion of the fault is not clearly known. The fault may be ashort as 7 km in length, or could transfer slip to other faults further south for a possiblelength of approximately 20 km. We have assigned a moment magnitude of 6.5 to the MillerCreek Fault, based on probable rupture length vs. magnitude relationships. This estimate isconsistent with current interpretation by the DSOD (2003).Lafayette-Reliez Valley FaultsThe Lafayette and Reliez Valley faults are separate faults that merge at their northern end,but are subparallel to each other along most of their length (Graymer and others, 1994).Recent investigations of the area reveal that the faults are associated with stronglypronounced geomorphic features that may be indicative of Quaternary fault activity,including saddles, tonal lineaments, linear valleys, vegetation alignments, and closeddepressions (Unruh and Kelson, 2002). The Lafayette-Reliez Valley (LRV) fault system isapproximately 13 km in length, although possible connection with the Cull Canyon Fault(which has not been demonstrated to show dextral offset) could extend the length for another14 km. The LRV faults require further investigation to better constrain their possibleQuaternary activity. However, because of the geomorphic evidence indicating potentialQuaternary activity, we consider these faults active for dam safety evaluation purposes, andassigned them a moment magnitude (M w ) of 6.5, based on an assumed rupture length of 13km.GEI Consultants - 29 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Inferred FaultsAs discussed in Paragraph 4.2, Graymer and others (1994) have depicted an inferred, buried“fault” passing under the dam centerline. During our field reconnaissance, which waslimited to surficial observations and did not include any subsurface investigations, we did notobserve any geologic features that would substantiate such assertion. Our review of thegeologic map provided no strong evidence for the existence of such inferred fault. The samebedrock formation (Contra Costa Group) is present on both sides of this fault, and the localgeologic structure appears to be locally consistent with fold axes. The fold axes are parallelto the postulated alignment of the inferred fault, between the dam left (west) abutment to theeastern termination of the inferred fault (approximately 5,000 feet east) and are depicted asbeing truncated, about 2,000 feet west of the dam, by a northeast-southwest bend of thisfault. However, this apparent truncation in fold axes could also result from changes in thestrike of the folds, or might be associated with deformation along unrecognized north-southoriented faulting. Such north-south faulting has not been identified to date, but could explaina series of other discontinuous lineaments, as discussed under the next title of this section.In addition to the lack of strong geologic evidence supporting the presence of a fault underthe dam, we have not found in our review any data that would indicate that faulting wasobserved during excavation of the dam foundation. If such inferred fault were present, itwould likely not be active, because the orientation of its strike is not consistent with the trendof other well-recognized tectonic features (i.e., north-south to northwest-southeast “NorthernCalaveras” stepover features). In the absence of more specific information, the existence ofsuch inferred fault may be questionable. We conclude that it would represent a very lowhazard to the dam as a potential seismic source, or in terms of secondary (sympathetic)movement potentially triggered by a major rupture of any of the major faults identified in thegreater site area. Furthermore, its orientation, parallel to the dam crest, would reduce thepotential for any sympathetic movements affecting the dam adversely.Another unnamed small bedrock fault was mapped to the north of the dam, and shows somedextral displacement (Wagner, 1978; Graymer, et al., 1994; Haydon, 1995). This northtrendingfault could represent one of multiple, north-trending smaller faults that might beassociated with the transfer of slip from the northern Calaveras Fault.Other lineaments near Lafayette DamNorth of the dam, Unruh and Kelson (2002) have identified a 6.5-km-long, north-trendinglineament zone. This zone, called the Russell Peak lineament zone, appears to coincide withan unnamed bedrock fault, mapped to the north of the dam, that shows some dextraldisplacement (Wagner, 1978; Graymer and Other, 1994; Haydon, 1995). As mapped byUnruh and Kelson (2002), the Russell Peak lineament zone ends approximately 2 km northGEI Consultants - 30 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05of the dam. Southward projection of this lineament zone would pass through or near theeastern margin of Lafayette Reservoir. We evaluated available aerial photographs (from1928, 1939 1959 and 1986) to substantiate any indications of lineaments in the reservoir anddam area. Although several possible lineaments were observed near the eastern reservoirmargin, these are weak and discontinuous, and not necessarily along trend with the welldefinedzone located further north. A second possible zone of weak, discontinuouslineaments may be present approximately 1.5 km west of the reservoir.Based on our geologic inspection, we have not identified any obvious fault-relatedlineaments through the reservoir. A strong, northwest-trending strike ridge locatedapproximately 700 m south of the reservoir margin appears to extend unbroken across theprojection of the weak lineaments, see Figure 4-2.Overall, lineaments identified in the dam region are not strongly pronounced. However,when considered together with similar features mapped to the north and east, they couldrepresent conditionally active or capable faulting associated with the recent tectonic setting.At this time, we consider the lineaments to represent a low hazard to the dam as potentialseismic sources.To improve understanding of the local geology as it specifically relates to Lafayette Dam andto help resolve any issues regarding the inferred fault and the lineaments, more detailedgeologic mapping might be considered to clarify their geologic relationships and assesswhether they represent faulting and could have any potential for activity.4.4 LandslidesNumerous landslides are present along the margins of Lafayette Reservoir, includingprobable deep-seated landslides involving rock material and shallow failures involvingsurficial materials (soil and colluvium). The largest landslides identified during our review,based on aerial photographs and site reconnaissance, are also depicted on Figure 4-2. We alsoobserved abundant small soil slumps and shallow debris flows that were not mapped. Yet,most of the observed landslides appear to pose a low to moderate risk of significant impact tothe dam, because they are located high on the hillside and upslope from the reservoir, or areinterpreted to be relatively small in size. Several landslides may have a higher potential riskdue to their relatively large size, more definitive geomorphic form, and location relative tothe reservoir.An older landslide is located adjacent to the right (east) abutment. The landslide appears tobe old, based on a subdued landform, but still retains the distinctive remnants of a headscarpand landslide body. If the landslide actually underlies a corner of the embankment, andexperiences movement in the future, it could potentially damage the downstream toe of theembankment. This would not affect, however, the overall safety of the dam.GEI Consultants - 31 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05An active landslide is present above the parking lot for the Visitor Center, at the northwestcorner of the reservoir. This landslide may have been triggered, in part, by grading forreservoir facilities. We note that the landslide is not visible in 1928 photographs, but appearsto be fully developed by 1939.4.5 Foundation ConditionThe foundation for the dam could not be observed, except for the downstream abutmentcontacts. The abutments were observed to be in satisfactory condition. We did not observeany signs of foundation stability issues. We evaluated foundation conditions from theavailable subsurface exploration and laboratory testing data.Lafayette Dam is founded on alluvial sandy clays from the Orinda formation. The alluviumthickness extends from a few feet or less at the abutments, and averages about 90 feet in thecentral portion of the dam. Greatest depths of alluvium were encountered in Boring SS-2(100 feet), SS-22 (98 feet) and SS-1 (90 feet). Alluvium primarily consists of clay to sandyclay with occasional lenses of thin sands and gravel. The sand content of the alluvium variesfrom 5 to 40 percent (S&W, 1966), except for two soil samples retrieved from two borings(Fig. A.4 of the S&W report). W.A. Wahler & Associates (WA, 1976) reported that nosignificant lenses of sand were detected during their field exploration program. The mostrecent subsurface investigations, relying on soil data interpreted from their borings, did notidentify any specific layer or zoning within the foundation alluvium. The first geologicreport by Louderback (1927) described the upper 5 to 15 feet of the alluvium as a dark clay,behaving as a stiff adobe when dry, and as a plastic clay when wet. Below the upperalluvium, Louderback described a change in clay color, but not in lithology. He alsoidentified a “confined” water-bearing zone, about 50 feet below the original ground surfaceof the alluvium, with a water level 64 feet above the proposed reservoir level. Overall, noloose saturated silts or sands, generally acknowledged the most susceptible to liquefaction,have been encountered in these borings.The foundation alluvium is underlain by bedrock from the Orinda Formation. The OrindaFormation is of Pliocene origin. It is composed of partially consolidated claystone, sandstoneand conglomerate and was encountered below the alluvium under the dam footprint and atboth abutments. Based on boring logs and field penetration data, no firm rock wasencountered, and S&W (1966) stated that in some cases, the formation resembles a hardclay, rather that a rock and that dry densities in that formation range from 110 to 130 pcf.Following the 1928 failure, the failed foundation soils were not removed when the dam wasreconstructed. Although these soils have most probably consolidated with time, we judgethat the quality of the Lafayette dam foundation materials may not meet modern standards fordam foundation requirements within the area affected by such failure.GEI Consultants - 32 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/054.6 Seismic Criteria4.6.1 GeneralThe purpose of this section is to discuss the selection of criteria for the seismic stabilityreview of Lafayette Dam. As discussed in Section 4.3, various faults may affect LafayetteDam. Based on maximum magnitude and distance considerations and its high rate of slip, theHayward Fault is the most critical (controlling) feature. The San Andreas and Calaverasalso represent significant seismic hazard. Despite its shorter distance to the site and beingpotentially causative of large PGA’s, the Lafayette-Reliez Valley is associated with a lesserrisk than the aforementioned faults, based on its uncertain, lower rate of activity and muchshorter expectable durations of ground shaking in case of related earthquake occurrence. Theinferred fault (see Section 4.3) below the dam was not considered as a potential seismicsource.4.6.2 Basis for Seismic CriteriaThe seismic evaluation of dams whose failure would present a hazard to life (“high risk”dams) is presently based on the concept of Controlling Maximum Credible Earthquake(CMCE). The “high risk” classification assigned to Lafayette Dam in the National Inventoryof Dams (NID) reflects its potential for extreme human and economic consequences in caseof failure, due to heavy downstream development. The United States Committee on LargeDams (USCOLD, 1985), now the U.S. Society on Dams (USSD), defines the CMCE as themost severe of all Maximum Credible Earthquakes (MCE) capable of affecting a dam. TheMCE is the largest, reasonably conceivable earthquake that appears possible along either arecognized fault zone or within a geographically-defined tectonic province, under thepresently known or presumed tectonic framework. Little regard is given to its probability ofoccurrence. The MCE is assumed centered at the closest distance between the dam andcausative seismogenic source. More recent literature refers the MCE as the MaximumConsidered Earthquake.Because the faults that might generate ground motion potentially affecting Lafayette Damhave high rates of slip, it is appropriate to use 84 th percentile criteria to define the groundmotion that could be generated at the site in case of tectonic rupture along these faults. TheConsequence Hazard Matrix of the DSOD (first proposed October 4, 2002) describes suchrequirements. We developed seismic criteria for the four faults the most critical to LafayetteDam: the Hayward, San Andreas, Calaveras and Lafayette-Reliez Valley faults.GEI Consultants - 33 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/054.6.3 Influence of Local Site Conditions.A consideration to develop seismic criteria is the subsurface condition at the site. The centralportion of Lafayette Dam is founded an average of 90 feet of clayey alluvium. Furthermore,based on measured shear wave velocities and descriptions provided in the project files, thelocal bedrock, the Orinda Formation, is poorly cemented and characterized by low shearwave velocities (average Vs less than 1,300 ft/s). Therefore, the Orinda Formation cannotbe classified as a “hard rock”, or even as a “soft rock”. The 1929 Board reported that “whenthoroughly wet, the Orinda beds have probably little, if any, greater strength than wellcompactedlayers of gravel, sand and clay”. Both the valley alluvium and the OrindaFormation bedrock on which Lafayette Dam is founded are equivalent to a “firm soil”.Using shear wave velocities (1,180 ft/s

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05considered equations developed by Crouse and McGuire (1995), because they are applicableto the Western United States and differentiate between subsurface conditions classified perNEHRP site type, and by Campbell (1997, 2001), who is well recognized for his extensivework on the subject. We used Campbell’s and Abrahamson and Silva’s equations for thevertical component of motion, as the other authors did not include the vertical component ofmotion in their studies. Details of the numerical procedures followed to develop ourrecommended PGA's and complete references are presented in Appendix D.The attenuation equations used in our review apply to shallow crustal earthquakes in activetectonic regions such as Coastal California, which has provided the largest amount of data.We used parameters applicable to strike-slip predominant tectonic regime. These equationsdifferentiate between strike-slip or reverse tectonic environments, and between soil or rocksubsurface conditions. Depending on the equations considered, we have used soil or anaverage of soil and rock conditions for this site, as discussed in Appendix D. All equations,except Crouse and McGuire’s, which are based on the surface-wave magnitude (M s ), use themoment magnitude (M w ) to quantify the size of the earthquake. We obtained the peakground motion estimates for each fault by geometric (logarithmic) averaging of thepredictions of the three selected attenuation equations.Peak ground accelerations and corresponding Earthquake Severity Indexes (ESI) aresummarized in Table 4-2. The ESI is an indicator of the energy content of the earthquakemotion considered and is used to estimate earthquake-induced deformations (see AppendixD). The ESI is a convenient way to quickly rank the probable order in which the faults arethe most critical to the Lafayette site. The Hayward and San Andreas faults are the two mostcritical features. Faults other than those listed in Table 4-2 (e.g. Concord and Franklin faults)could generate significant ground motion at the site but, because of their associated distanceand magnitude, have less demanding ESIs and, therefore, do not control the MCEs forLafayette Dam.4.6.5 Response SpectraAttenuation equations for five percent damping pseudo-absolute spectral accelerations (PSA)at various periods are also provided in the same references as for the PGA. We used theseattenuation equations to develop response spectra that define the MCE ground motion atvarious frequencies. These spectra are representative of the “average” horizontal motion atthe site, since both the primary and secondary components of the records analyzed to developattenuation equation parameters were used. The process by which we obtained theseresponse spectra is also described in Appendix D. As for the PGA, mean + σ estimates wereaveraged to develop the recommended response spectra. We believe this averagingprocedure to be conservative.GEI Consultants - 35 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Because Lafayette Dam is located a short distance away from several faults of significantlength, the location where the fault rupture could originate along the causative fault and thedirection of propagation of such rupture, e.g., toward or away from the site, could inducenear-field and directivity effects and influence the response spectra. Such effects havetypically been observed as large velocity pulses (directivity) at the beginning of some strongmotion records and can result in large spectral amplitudes in a narrow band of periodslocated between 0.6 sec and 5 sec. In addition, in the near-field, the fault-normal (FN)component can be substantially larger than the fault-parallel (FP) component. The concept ofcorrecting response spectra for directivity and near-field effects is rather recent. As it issomewhat impossible to accurately predict at what periods of the spectra such effects wouldbe the most significant, we have corrected the 84 th percentile horizontal spectra to includenear-field and directivity effects using a recently recommended broad-band correctionprocess that depends on the type of faulting, magnitude and distances considered. Details onthe way these effects were implemented in our recommended horizontal spectra are providedin Appendix D. The 5 percent damping Hayward MCE response spectra, which would be themost demanding for Lafayette Dam because of the large associated magnitude, are presentedon Figure 4.3. The response spectra for the other faults are presented in Appendix D.Response spectra for damping values other than five percent are also needed, as significantdamping is expected in soil-like materials under demanding ground motions such asconsidered. For this purpose, we used a scaling procedure based on spectral amplificationratios published by Newmark and Hall (1982), and we multiplied the five percent dampingspectral amplitudes by appropriate scaling factors to obtain response spectra at 0.5, 2, 7, 10and 20 percent damping. The recommended scaling coefficients provide a smooth variationof the calculated spectral amplitudes as a function of period. Because of this scaling process,response spectra at damping values other than five percent are only approximate.The response spectra developed for the faults most critical to Lafayette Dam are tabulated inTables 4-3 through 4-10. These spectra conservatively define the frequency characteristicsof the ground motion representing the various MCEs.The recommended PGAs and spectral accelerations listed in Tables 4-3 through 4-10, whichare based on deterministic principles, can be compared with ground motion levels obtainedfor the Lafayette site (latitude: 37.885 o , longitude: 122.138 o ) through probabilistic seismichazard analysis by the USGS (Seismic Hazard Mapping Program), see Table 4-11. Twoconclusions can be drawn from comparing Table 4-11 with our recommended spectralvalues:• Although obtained deterministically, some of the PGAs and recommended spectralvalues have relatively high probabilities of occurrence, of about 10 percent in 50years or higher at some frequencies. This is not surprising, considering the high ratesof slip of the local faults.GEI Consultants - 36 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05• The recommended spectral accelerations do not represent upper bounds of groundmotion and could be exceeded during events of low probability of occurrence.Although they represent conservative requirements, 84th percentile response spectracould be exceeded at any period or range of periods of vibration.The above comparison with the USGS probabilistic estimates is only useful as a relativeindication of potential seismic hazard, and is not intended to provide an alternative way todefine seismic criteria for this project. However, as mentioned above, the 84 th percentiledeterministic spectra have a finite probability of being exceeded at any specific exposureperiod, and independently of how they compare with probabilistic estimates derived from thedata contained in the USGS database.We compared our recommended 5 percent damping response spectra with the responsespectra presented for the Hayward (Earthquake A) and San Andreas (Earthquake B) inputmotions previously considered in 1976 by W.A. Wahler & Associates, see Figure 4-4. The1976 response spectra were originally intended by their developers to represent bedrockmotion and, therefore, have a high energy content at periods between 0.2 sec and 0.5 sec, oflittle applicability to Lafayette Dam. Earthquake A was a modified “Lake Hughes” record,intended to represent a M 7.5 event, and Earthquake B was the Seed-Idriss 1969 syntheticearthquake, composed of accelerograms of shorter magnitudes and durations intended torepresent the propagation of the rupture for a M 8+ event. Periods of dynamic response ofsignificance to Lafayette Dam are comprised in the range of periods 0.6-2.0 sec. From thegraphical comparison shown on Figure 4-4, we conclude that Earthquake A is insufficientlyconservative by modern standards to represent the Hayward event, by a factor of between 2and 3 at some periods of potential significance to the dam response. Earthquake B isadequate and, indeed, quite conservative.GEI Consultants - 37 -

Table 4-1 – Faults in the vicinity of Lafayette DamFault orFaultSegmentSanAndreasDistanceFromLafayetteDam(km)BestEstimateRuptureLength(km)Down-DipWidth ofFaultRupture(km) 3ApproximateRupture Area(km 2 ) 4MagnitudeEstimate(based onrupturelength)MagnitudeEstimate(based onrupturearea) 639 474 2 12 5680 7.9 6 7.0Hayward 1 8.8 87 2 15 1300 7.25 5 6.75Northern,Central,andSouthernCalaveras 1 9.8 118 2 15 1770 7.25 6 7.25NorthernCalaveras 1 9.8 40 2 15 600 7.0 5 7.0Concord 1 12.8 14 2 15 210 6.5 6 6.5MillerCreek9.5/11.3 7– 20(notknown)__ 6.5 5 6.5Franklin 7.5 20(notknown)__ 6.75 5Lafayette–ReliezValley3.0 13(notknown)__ 6.5 51 Table A-1 (Working Group on Northern California Earthquake Probabilities, 1999)2 Best estimate of rupture length reported in Table A-1 (WGNCEP, 1999)3 Best estimate of down-dip width of fault rupture reported in Table A-1 (WGNCEP, 1999)4 Product of rupture length and down-dip width of fault rupture5 Magnitude vs rupture length estimated using Wells and Coppersmith (1994) Figure 96 Magnitude vs rupture area estimated using Wells and Coppersmith (1994) Figure 16Fault locations and distances based on:Hayward: (Lienkamper, 1992; Radbruch, 1969)Calaveras: (Hart, 1981; Crane, 1988, Simpson and others, 1992)Concord: (WGNCEP, 1999)Miller Creek: (Wakabayashi and Sawyer, 1995; Dibblee, 1980)Franklin: (Crane, 1988)Lafayette-Reliez Valley: (Unruh and Kelson, 2002)

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/055. INSTRUMENTATION5.1 Survey MonumentsThere are currently 24 survey monuments, arranged in a grid on the embankment ofLafayette Dam. The monuments are surveyed approximately once (and occasionally twice) ayear. EBMUD provided survey data measurements from June 1, 1989 to December 9, 2003for our review. These data were provided both as graphs and raw data.Tables and figures of maximum vertical and horizontal displacements recorded for themonuments from 1989 to the present have been prepared and are included in Appendix A.The maximum horizontal and vertical displacements for all the survey monuments duringthese 15 years of record are 3.36 inch and 3.12 inch, respectively. Review of the time versusdisplacement graphs for individual survey monuments shows that the horizontal and verticalmovements are stable, with no significant increase in recent years. Horizontal movement ofthe upstream slope is toward the upstream direction, while the downstream slope has movedslightly downstream.5.2 PiezometersThere are currently 18 active piezometers at Lafayette Dam. The District provided us timeversus-readinggraphs for the active piezometers from January 1989 to January 2004 for ourreview. Piezometers readings are taken approximately monthly. In addition, thecorresponding readings of reservoir level from 1989 to 2004, as well as rainfallmeasurements, were included in the graphs. Details regarding the installation history andreadings of the piezometers are provided in Appendix B.While several piezometers show only small variations consistent with the reservoir level andno specific trends over the last 10 years, others show fluctuations in excess of the reservoirlevel fluctuations and readings higher than the reservoir level.Following review of the piezometric data, we estimated the average phreatic level in themaximum cross-section of the embankment, see Figure 7-3. Comparison of the estimatedphreatic surface with the phreatic surface assumed in the Wahler (1976) stability analysesand by the DSOD (2003) indicates that our estimated phreatic surface is generally similar tothose used in previous studies, although somewhat lower in the upstream portion of the dam,and slightly higher in the downstream portion of the dam. Overall, such variation is probablynot significant to the stability and performance evaluation, and remains well within the levelof judgment needed to interpret the piezometric data.GEI Consultants - 38 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05The phreatic surface in the downstream portion of the embankment, as interpreted from theseopen-well piezometric data, is quite high for a dam with an “impervious core”. The core walland sheet piles probably no longer function as an impervious barrier as originally intended,which may contribute to the high water surface observed within the dam section. Somepiezometric readings could represent foundation water pressures, rather than indicate the truephreatic surface within the embankment.5.3 Seepage MonitoringSeepage is collected by tunnel and embankment subdrains. A 24-inch conduit, which wasinstalled in a 60-inch diameter concrete conduit, runs along the left abutment of the dam.Tunnel leakage is collected in a sump box, located near the west end of the toe of the dam.Seepage through the dam is collected by a subdrain system, composed of 6-inch and 8-inchpipes, which run perpendicular to the dam axis and along the toe of the dam. Seepagecollected in the pipes is evacuated through a seepage collection box, located near the rightabutment at the toe of the dam. Preliminary review of seepage data collected in the last tenyears indicate that tunnel leakage ranges from 1 gpm to less than 5 gpm, and seepage fromthe toe drain range from zero to less than 10 gpm. These quantities are well withinacceptable, and not indicative of any particular problem.5.4 Instrument EvaluationThe crest survey data are within the survey accuracy and indicates no significant vertical orhorizontal movement in the last 15 years. Due to the age of Lafayette Dam, we believe thatcrest monument surveys provide limited information, and could be conducted only after feltearthquakes. However, we understand that the District intends to continue surveying themonuments at regular intervals.Until 1992, the piezometers in Lafayette Dam have had both relatively consistent and erraticreadings with relatively high phreatic levels. The piezometers indicate a general downwardgradient in the downstream slope. The indicated phreatic levels remain in agreement with orare conservative with respect to the assumptions of the previous and present slope stabilityanalyses. Some piezometers were replaced, and still show fluctuations exceeding thereservoir water elevation, due to probable continued infiltration of surface water or“blinding”. Several of the older piezometers (installed 1956, 1965) may display signs ofplugging. Review of the installation details suggests that the piezometer screens may nothave been adequately designed to mitigate fines infiltration. Overall, the piezometersadequately monitor the continued good performance of the dam.Presently, leakage appears to generally be within the bounds of historic fluctuations. For adam of the size of Lafayette, average seepage of less than 10 gpm about is quite low.GEI Consultants - 39 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/056. FIELD INSPECTION6.1 GeneralThe dam and reservoir area were inspected on 2/24/2004 by Bill Cole (CSA) and MarkMcKee (RYCG), and on 4/1/2004 by Gilles Bureau and Carol Buckles (GEI). Bill Colereturned to the site on April 30, 2004 to specifically look for any evidence of an inferred faultmapped by Graymer, et al. (1994). The purpose of these visits was to observe and assess thelocal geologic conditions, compare the site layout with the available drawings, observe thecurrent condition of the facilities, and identify any conditions that would potentially impactthe seismic stability of the dam. The inspection was conducted by following standardinspection procedures (FERC dam inspection checklists). Selected inspection photographsare included as Appendix C.The project serves as an emergency standby storage facility. The reservoir is small (4, 250acre-feet) off-channel, and typically subject to insignificant daily fluctuations. Water surfaceelevation at the time of our second inspection was about one foot below the lowest pointalong the top berm (El. 445 ft, estimated, taking the downward curvature of the berm intoaccount). Lafayette Reservoir is open to the public for recreational purposes. The crest ofthe dam serves as parking for hikers and fishermen. However, access to the dam is limited tothe crest; the upstream and downstream slopes have restricted access. Permission wasgranted by EBMUD to walk in and inspect the restricted areas of the embankment. Thereservoir rim is heavily vegetated. No signs of reservoir rim instability were noted, butnumerous old landslides, discussed in Section 4.4, were observed.6.2 DamLafayette Dam is a zoned earth embankment. Its upstream slope is lined with concretepanels (slabs) intended for slope protection. The crest of the dam is paved with asphaltconcrete and is used for parking. The crest was observed to be in good condition with noobvious signs of settlement, misalignment, or cracks. Because of its wide area, surfacerunoff may accumulate locally, but a drainage collection system discharging into a pipe nearthe top of the upstream slope in the central portion of the dam appears to have been installed.The upstream slope has a very noticeable concave shape near the maximum section of thedam, with about 4 feet of vertical elevation differential at crest center, see Photos C-1 and C-2. Rainwater ponds near that location at the lowest point of the upper upstream berm. Basedon existing data and considering the vertical retaining wall along the upstream edge of thecrest, most of this settlement should have occurred shortly after the 1928 dam failure andGEI Consultants - 40 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05before the dam crest was brought to its present elevation. The 1929 report by the ConsultingBoard mentioned only 1.5 feet of subsidence for the upper upstream berm. Hence, anadditional 2.5 feet of vertical downward movement occurred at that location after the 1928post-failure surveys. Presently, the crest does not show any indication of settlement. Theupstream face concrete panels appear to be in good condition, except for minor cracking(Photo C-3) and some gaps, up to 4-inch wide, between adjacent concrete panels (Photo C-4). The joints have been filled with asphalt, but the filling is frequently deteriorated ormissing, and grass grows in open joint spaces, see Photo C-5.The downstream slope was covered with grass at the time of our inspection (Photo C-6), butwas observed to be free of erosion, excessive vegetation, and cracks or obvious settlement.No seepage was visible along that slope. The groin areas were observed to be dry. Various6-inch and 8-inch clay-tiled surface drains, dry (Photo C-7), or collecting minor surfacerunoff (Photo C-8), are located along the downstream face. Numerous rodent holes werenoticed. The dam is primarily founded on alluvium, but we noted no foundationdeficiencies along the sides or at the bottom of the embankment where it abuts the OrindaFormation. We observed the downstream toe of the dam and the collector outlet box for thesubsurface drains. Very small, clear outflow was observed at the downstream drainagecollection system (Photo C-9). Piezometer locations along the crest and two faces of thedam are well marked and protected with steel caps, see Photo C-10. No signs of adverseartesian pressures were observed downstream of the dam.Our review of the historical information on Lafayette Dam indicated that the 1928 failureoccurred after much of the upstream concrete slabs were in place. The Consulting Boardconvened after the failure (see Section 3.1.2) stated in its report that the upstream slope of thedam experienced much less movement than the crest and downstream slope, and that itsupper berm, originally constructed to be straight at El. 450, curved convexly (outwardhorizontal bulge) with a maximum horizontal displacement of 3.8 feet toward upstream as aresult of the failure. The central portion of the upper berm also showed a maximumsubsidence of 1.5 feet. The concrete slabs on the face were displaced but generally notbroken. Remedial work following the 1928 failure consisted of repairing or replacing theconcrete panels on the upstream face, but the post-failure slopes in-between the crest and theberms were not modified (Shannon & Wilson, 1966). Hence, the maximum verticalslumping reported immediately after the 1928 failure is less than the distortion of theupstream face presently observed (about 4 feet of downward movement). Shannon &Wilson (1996) reported that the dam settled 1.3 feet between 1933 and 1963, and predictedanother 0.4 ft in the following 30 years. Hence, some 0.8 ft of settlement may have occurredfrom 1928 to 1933. Review of the recent settlement data indicates only minor, insignificantmovement of the upstream slope. It is reasonable to conclude that most of the observeddistortion of the upstream face has occurred in the 35 years that followed the 1928 failure,and is not indicative of recent movement.GEI Consultants - 41 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/056.3 Outlet Tower and SpillwayThe dam does not have a separate spillway. A reinforced concrete outlet tower (Photo C-11),located in the reservoir approximately 320 ft from the crest, provides both spillway functionand reservoir drawdown capacity. The outlet tower was not accessible for inspection duringthis site visit. Spillway flows and releases from the reservoir exit the tower through twoadjacent 60-inch diameter underground conduits located below the left abutment portion ofthe embankment. The portion of the outlet conduit beyond the dam toe includes an internalsteel pipe. The outlet conduits terminate downstream of the dam toe in a baffle box acrossMount Diablo Blvd. The inspection team attempted to observe the baffle box and outletdischarge but was not able to gain access to the area due to heavy vegetation and fencing.Minor drainage was observed from the drainage collector at the right abutment side of thedam toe, as would be expected for this time of year. The seepage was clear and exiting at lowvelocity.The evaluation of the tower was not included in this review. The outlet tower appears to bein fair exterior condition, seen from the dam, but we are aware that its performance underearthquake loading is under review, based on recent studies by the District. No deficienciesother than insufficient seismic capacity have been reported in recent years. We understandthat seismic upgrade of that tower is presently contemplated.GEI Consultants - 42 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/057. EMBANKMENT STABILITYASSESSMENT7.1 Previous Field Exploration and Laboratory TestingProgramsVarious field exploration programs have been performed over the years at the site, includingprior to the dam construction. An inventory of the borehole information we collected andreviewed is shown in Table 7-1. In addition to the original field investigation and the boringsdrilled after the 1928 slope failure (see Figure 3-3), three exploration programs (1956, 1965and 1973) provided data useful to this review. Location of these borings is shown on Figure7-1, which is taken from the 1976 Wahler report. Additional boreholes were drilled in 1992and 1996 by EBMUD to install piezometers, but we did not locate any corresponding boringlogs. Summary information is provided below regarding these various field and laboratoryprograms. More complete information regarding laboratory testing data is provided inAppendix E.In general, we found many similarities between embankment and foundation materials. Thisis no surprise, since all construction materials were locally obtained from borrow areas in thealluvium and hillside upstream from the dam. Virtually all materials recovered either fromthe dam core and shells or from the underlying foundation alluvium are referred as clay orclay-like materials. Based on the data reviewed and our interpretations, we believe that thedam zones and foundation materials are essentially homogeneous, with occasional smallzones of more sandy materials, but with no continuous granular layers or significant lenses ineither the dam shells or foundation.7.1.1. Pre-Construction and 1929 Post-Failure Field Programs.Louderback (1927) described 12 test pits dug at various locations along the proposedcenterline of the dam, some to significant depths (11+05 pit was 32 feet deep). Borings wereadvanced from the bottom of three of these pits with a hand auger, and reached depths of 48to 79 feet. Louderback provided a description of the soils encountered in his report, but wefound little information of current engineering significance. We found no informationwhether groundwater was encountered in the borings drilled prior to the dam construction,but water was encountered at about 20 feet below ground surface in two wells, 74 feet and 94feet deep, respectively, drilled a short distance above the dam site (Board Report, 1929).GEI Consultants - 43 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05The 1929 investigation of the 1928 failure included 5 test pits along the upstream berm, 22borings through the failed portion of the dam along a line perpendicular to the dam axis (atStation 11+58.2), and 6 test holes along the downstream toe. The logs of these borings wereretrieved from the data files reviewed and are also plotted on a District drawing showing thefailed cross-section of the dam (DH 1608-18, dated January 1929). The corresponding logsonly describe the types of materials encountered, and their estimated condition of moistureand plasticity. No free water was encountered in these borings within the embankmentmaterials. Erratic water levels were encountered below the former ground surface (top ofalluvium). For example, based on the 1929 report, high water pressures were encounteredabout 50 feet below the original ground surface and the water rose in Hole 24, on theupstream side of the top the dam, to 64 feet above the reservoir water level (reported to be atEl. 392 on 9/17/1928). While this occurrence could have represented artesian pressure, itwas concluded that in the alluvium and, especially, near the downstream toe, the laterallycompressed failed materials probably significantly influenced pore pressures ratios.7.1.2. 1956 InvestigationIn 1956, the District drilled two borings, SS-1 and SS-2, by using its own rig, to obtainsamples from the foundation alluvium and embankment materials. Boring SS-1 was drilledfrom the dam crest to a depth of 203 feet, Boring SS-2 from the upper downstream berm to adepth of 182 feet. Samples were obtained with a 2-inch diameter modified California splitspoon sampler and Shelby tubes. We did not find records of any blow counts for these twoborings in the files made available to our project team. Soil samples from these two boringswere tested in the laboratory. The testing program included moisture and density tests (MD),Atterberg limits tests (AL), specific gravity (SG), consolidation (CONSOL) and confinedundrained(TXCU) and unconfined-undrained (TXUU) triaxial compression tests. Mostspecimens tested were 2-inch diameter by 4-inch long. TXCU tests were loaded to failure orup to 20% strain at a constant rate of loading of 0.05 inch/mn, with a confining pressuretargeted to be 0.6 times the overburden pressure at the depth where the sample wasrecovered. The strength testing included samples from the core (Zone 1), the upstream shell(Zone 3), and foundation materials “presumably failed during construction” (Zone 4?). TheTXUU tests were intended to measure the shearing resistance of the clayey soils.7.1.3. 1966 InvestigationTwelve borings designated as B-3 through B-14 were drilled in 1965 with a Failing 750 drillrig. The holes were augered to about 5 feet, then 5 to 8 feet of casing were driven from thetop of the hole, and the holes were completed to their final depth with rotary wash drillingwithout drilling mud. No further casing was placed. Two-inch O.D. standard SPTtesting/sampling was alternated with undisturbed sampling using a 3-inch diameter Pitcherbarrel. A few specimens were recovered by pushing the sampler into the soils, indicating thepresence of possibly softer zones. MD tests were performed on all samples recovered. ALGEI Consultants - 44 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05and TXUU tests were performed on selected samples, and torsion vane shear tests wereperformed “on all cohesive undisturbed samples”. One CONSOL test was performed on acore sample, and eight series of TXCU tests were performed on various materials.7.1.4 1976 InvestigationIn 1973 and 1974, W.A. Wahler & Associates (WA) directed a subsurface investigationconducted using truck-mounted (Failing 1500) and barge-mounted (B-40) drilling equipment.The drilling and sampling were actually performed by District’s engineering staff, who alsoprepared the boring logs. Seventeen borings were completed. The Wahler field program wasalso supplemented by a cross-hole geophysical investigation (Langenkamp and Nelson,1973). The DSOD questioned the 1973 geophysical data (shear velocities too high) and thespacing of the holes, and Woodward-Clyde Consultants (1975) conducted a second crossholegeophysical program. CalTrans also performed an independent downhole survey forresearch purposes, and the results were made available to the project team. The measuredvelocities were intermediate between those of the two other geophysical programs and wereused by Wahler as the primary basis to define the low-strain dynamic shear moduli of thematerials encountered. Wahler also performed an extensive laboratory testing program,including TXCU tests and stress-controlled and strain-controlled cyclic triaxial tests,isotropically (K c = 1) or anisotropically consolidated (K c = 1.5).7.2 Previous AnalysesThe stability of Lafayette Dam was previously investigated by the District (1956), Shannonand Wilson, Inc. (1966), and W. A. Wahler & Associates (1976). In 2003, the DSOD usedexisting data and performed additional slope stability and simplified deformation analyses(Newmark’s method). These previous studies are reviewed in the following paragraphs.7.2.1 Static and Pseudo-Static Analyses (1956)The District performed early stability studies in 1956. At that time, the reconstructed damimpounded a reservoir at El. 448 (since 1937), following two earlier reservoir raises at El.410 and at El. 428 in the years that followed reconstruction. The 1956 studies were toinvestigate the feasibility of raising the reservoir water level to El. 456 and constructing anew spillway, which was never built, through the left abutment. The 1956 studies were alsointended to compensate for the “poorly documented reconstruction of the dam”. As part ofthese studies, various options to upgrade the dam were proposed, including flattening of theslopes or constructing a drainage curtain.GEI Consultants - 45 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05The data from the 1956 laboratory testing were used to develop analysis parameters.Strength properties were obtained from triaxial compression test results, “reduced by aworking factor of 1.2” (EBMUD, 1956). However, strength of the foundation material wasback-calculated from the “at rest” post-failure position of the 1928 embankment slopes, and acohesion of 580 psf and a friction angle of 12 degrees appears to have been assigned to boththe foundation and the embankment fill in the final 1956 analyses (Dukleth, 1956).Stability analyses were performed using the method of slices. They included steady seepage,rapid drawdown, and pseudo-static analysis of the upstream (U/S) slope; and steady seepageand pseudo-static analysis of the D/S slope. The pseudo-static horizontal load coefficientwas taken as 0.10g. The studies concluded the upstream slope to be unsafe for rapiddrawdown, and the dam to be “unsafe under either proposed or present operation ifsubjected to a major earthquake of sufficient force to produce a horizontal acceleration of0.10g”. G. Dukleth (1956) presented factors of safety (FS) related to the 1956 studies in amemorandum. For rapid drawdown (El. 456 to El. 420), the computed FS was 0.88. The FSwas 1.12 for normal storage lowered to El. 420, and 0.63 for a reservoir at El. 420 andpseudo-static earthquake analysis (0.10g). The factors of safety computed by Dukleth areextremely low, but were based on post-failure strength estimates that should approach theresidual strength of the foundation and embankment materials. Undoubtedly, the failedembankment and foundation materials have gained strength with time, as indicated bysubsequent laboratory testing programs.7.2.2 Static and Pseudo-Static Analyses (1966)In 1966, and following a review of the 1956 studies by DSOD, Shannon and Wilson, Inc.(S&W, 1966) used the standard method of slices (no side forces) with postulated circularfailure surfaces. Based on the results of field exploration (12 additional borings) andlaboratory testing, they subdivided the dam section into three zones Zone 1 (core), Zone 2(U/S shell), Zone 3 (D/S shell) and the foundation alluvium into two zones, Zone 4 and Zone5. Zone 4 represented a “somewhat weaker section of the alluvium at shallow depth belowthe downstream section”. First, S&W performed a post-failure analysis of the 1928embankment, then static (steady seepage and rapid drawdown) and pseudo-static analyses ofthe maximum section of the reconstructed embankment. They also analyzed several wedgesto investigate the construction failure, assuming horizontal forces only. Pseudo-staticanalysis used a horizontal seismic coefficient equal to 0.10g.For the analyses “during construction”, S&W calculated a FS of about 1.0 for deep circlespassing through the foundation, and between 0.5 to 0.9 in the wedge-failure analysis. Theyconcluded that the dam in the 1928 final stages of construction was “marginally safe”, butstated that factors of safety obtained in their wedge failure analyses were unrealistically low.S&W used the slip circle analyses of the 1928 section as a basis to select shear strengths forstability analyses of the reconstructed dam.GEI Consultants - 46 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Based on their laboratory and post-failure analysis results, S&W selected effective stressstrength parameters to represent normal operating (“steady seepage”), and total stressstrength parameters for the construction failure and pseudo-static analyses of thereconstructed dam. Such strength parameters are summarized in Table 7-4. Steady seepageanalysis of the reconstructed dam section only considered the D/S slope. The lowest FS was1.9, and involved a circular failure surface (referred to as “trial circle 4A”) passing throughthe crest center and at a shallow depth within the foundation alluvium. Hence, such failurewould involve most of the D/S slope, but not the deeper part of the alluvium. Rapiddrawdownanalyses of the U/S slope involved critical circles passing through a berm at El.447 (normal pool) and through the foundation soils and dam toe, and corresponded to a FS of1.4. In pseudo-static analysis (0.10g), the same failure circle for the D/S slope as in thesteady seepage analysis (circle 4A) resulted in the FS, 1.2. A wedge analysis was alsoperformed for pseudo-static condition and had a FS of 1.7. The U/S slope was also notanalyzed for pseudo-static condition. Based on these results, S&W concluded the safety ofthe embankment to be adequate.7.2.3 Equivalent-Linear Dynamic Analyses (1976)In 1976, W. A. Wahler & Associates (WA) assessed the seismic stability of Lafayette Dam,assuming that general liquefaction of the dam materials would not occur. They conducted anew field exploration program to recover samples for cyclic triaxial testing. No“instantaneous or sudden” loss of strength occurred in any of the samples dynamicallytested. Wahler subcontracted downhole geophysical surveys (Langenkamp, 1973), but theresults were questioned by the State, presumably because the spacing of the holes wasexcessive. Woodward-Clyde Consultants (WCC, 1975) conducted new cross-hole anddown-hole seismic surveys at the site, but Wahler concluded that the new data probablyunderestimated V s in the alluvium. The WCC results and the results of another downholegeophysical survey by CalTrans (1975) in one single hole (SS-30) were used to define thelow-strain dynamic modulus of the soils encountered. The CalTrans variations of wavevelocities in the upper half of the alluvium appear quite large (500 fps to 2,500 fps), if noterratic. Overall, the shear wave velocities used by Wahler as a basis for estimating lowstraindynamic properties for dynamic analyses appear to be substantially higher thanmeasured by either WCC or CalTrans (see Fig. VI-6 of the Wahler report), but a review ofthe Wahler analyses performed in 1977 concluded that such selection of V s should notsignificantly affect the results of the analysis (Memorandum of Design Review, 3/4/1977).In its analyses, Wahler used zones 1 through 3 (U/S shell, D/S shell and core) to represent themaximum section of Lafayette Dam, but did not differentiate between Zone 4 and Zone 5 inthe foundation alluvium, as did Shannon & Wilson, and used the same engineering propertiesfor the entire alluvium layer below and beyond the dam. The response of Lafayette Damembankment to earthquake ground motion specified at bedrock level below the dam’sGEI Consultants - 47 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05foundation was estimated through equivalent-linear (EQL) dynamic finite element responseanalysis. Bedrock motions were estimated for three “Maximum Probable Earthquakes” asfollows: Hayward Fault (M 7.5) at 9.6 km distance, with a peak ground acceleration (PGA)of 0.60g; Calaveras Fault (M 7.3) at 6.4 km, with a PGA of 0.52g; and San Andreas Fault (M8.3) at 40 km, with a PGA of 0.40g. M designated the Richter Magnitude. Wahler selectedhorizontal acceleration histories with bracketed durations of shaking of 39 seconds(Hayward), 38 seconds (Calaveras), and 75 seconds (San Andreas). The bracketed durationis the interval of time between the first and last acceleration peaks of 0.05g or greater.Wahler calculated a fundamental period of 1.97 sec for the maximum section of the dam(Station 11+65). Such long period was obviously influenced by the thickness of alluvium atsuch section. It was concluded that the maximum section should be the most critical becausethe cyclic strength of the compacted shell material was between 15 and 50 percent higherthan that of the alluvium. As both the dam and foundation materials were “not liquefiable”,Wahler estimated earthquake-induced deformations based on the concept of “strainpotential”, using the laboratory-measured cyclic strengths and computed number ofequivalent cycles. Strain potentials calculated within each element of the numerical modelwere converted, using a simplified procedure, to “shear displacements” along selectedvertical columns through the embankment.In preliminary analyses, Wahler found the response to the San Andreas event to be the mostcritical, and the final dynamic analysis was only performed for that earthquake scenario. Wefound an early review memorandum in the files, dated March 18, 1974, that stated that “theseresults showed many elements in the foundation with factors of safety less than 1.0”. It is notclear whether that sentence pertained to preliminary or final results of the Wahler study.Thirty (30) equivalent uniform stress cycles were concluded to represent the averageresponse of the embankment to the San Andreas event, and average induced stresses werecompared with the cyclic strength of the dam and foundation materials at 5 percent and 10percent dynamic strain levels. The most critical area was found to be the upper half of thefoundation alluvium, below the upstream slope of the dam, with strain potentials at between5 and 10 percent (Figure VI-7). Maximum earthquake-induced displacements of 8 to 9 feetwere estimated for the section analyzed, and assumed to result in a maximum loss offreeboard of between 2 and 3 feet.7.2.4 Simplified Analyses by DSOD (2003)In a simplified reevaluation of the seismic stability of the Lafayette embankment, the DSODperformed additional slope stability and simplified deformation analyses. The DSODreviewers used the same dam zoning as S&W and Wahler and similar properties for Zones 1through 5, but introduced a new foundation Zone (Zone 4.5) below the lower portion of theD/S slope, where the slope is 8H:1V, “to account for some low blow counts possiblyindicative of residual soils”. The dam and foundation zoning considered by the DSOD isGEI Consultants - 48 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05shown on Figure 7-2. DSOD performed various stability analyses of the D/S slope, includingstatic, pseudo-static (0.15g coefficient) and post-earthquake static (using the residual strengthin Zone 4.5), and of the U/S slope, including pseudo-static (0.15g) and rapid drawdown. Allof these analyses resulted in computed factors of safety greater than 1.0. The earthquakeinduceddeformations were estimated at 1 to 3 feet using the Makdisi and Seed simplifiedprocedure (1977) with an 84 th percentile PGA of 0.69g, an estimated magnitude (M w ) of 6.5and a computed yield acceleration of 0.21g. Hence, based on these PGA and estimatedmagnitude, these deformations seem to correspond to a nearby event along the Lafayette-Reliez Valley Fault, rather than along the Hayward or Calaveras faults. Deformations werealso computed using the Newmark’s method (1965) by double integration of accelerationincrements of the Lucerne (1992 Landers Earthquake) above the yield acceleration. With thismethod, non-recoverable displacements less than one foot were estimated.7.2.5 Current Applicability of Previous AnalysesWe critically reviewed the results of the previous analyses discussed above. The standardmethod of slices (no side forces) was used in both 1956 and 1966, and analysis propertieswere estimated by back-calculations of the 1928 failure. Hence, the corresponding factors ofsafety were probably underestimated. The 1956 studies used excessively conservativeanalysis parameters, and should probably be disregarded. The 1966 analyses of thereconstructed dam led to acceptable factors of safety, but only considered the D/S slope. TheU/S slope could have been more critical, because of its higher phreatic surface and steeperslope, but its performance was likely acceptable, considering that a FS of 1.4 was obtainedfor rapid drawdown condition.The 1976 dynamic analyses were state-of-the-art procedures at the time when performed, butthe EQL method of dynamic analysis may not be sufficient for a high risk dam, potentiallysubjected to very demanding severe earthquake ground motion. For example, the frequencycontent of the input motion and the velocities of simulated earthquake waves affect thedynamic response analysis. Using modern finite element mesh sizing criteria, and dependingon the computer program to be used, the maximum element size should be smaller than 1/6 thto 1/10 th of the minimum shear wavelength to be transmitted. The low-strain shear-wavevelocities (V s ) of the embankment and foundation materials range from 500 to 1,500 ft/s.Assuming that frequencies of 5 Hz or below should be propagated without being overdamped,maximum element size (low-strain dynamic motion) should be between 10 feet and15 feet (for V s = 500 ft/s) and between 30 feet and 50 feet (for V s =1,500 ft/s). Hence, for themaximum dam section (132 feet high and 90 feet of alluvium) and low-strain condition, 7 to14 layers, and possibly more, might be required at low-strain levels. Under earthquakeloading, the dam and foundation material will lose stiffness and V s decrease, requiringadditional refinement of the analysis model and thinner layers. Wahler represented the damand foundation with 8 layers, a number probably insufficient to assure proper transmission ofthe seismic waves under MCE loading condition.GEI Consultants - 49 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05In his review, Dr. Idriss also indicated that the dynamic properties used in the 1976 analysesfor the clay shear modulus reduction and damping factors have been since judged inadequatefor such soils, and updated in recent work by Dr. K.H. Stokoe and his colleagues at theUniversity of Texas, Austin. However, we did not address this review comment further, asno dynamic analysis was contemplated in this study.The above modeling limitation may have contributed to lengthening the calculatedfundamental period of the dam and may explain why no amplification of the peak basemotion (0.40g) was computed at the crest (0.39g) in 1976. Most dams amplify groundmotion, based on observed performance of existing instrumented dams (USCOLD; 1992,2000). At such level of input motion, one would expect some amplification of the peakacceleration at the top of a dam such as Lafayette Dam, although its wide crest may reduceresponse at that level, compared with the base motion. This observation raises somequestions regarding the computed response. Quick checking that the dam response mighthave been underestimated as a result of insufficient resolution of the numerical model wasobtained by comparing the Wahler crest acceleration amplification ratio (0.97) with crestacceleration amplification ratios obtained in the Makdisi-Seed’s procedure (1977), seeSection 7.7 and Appendix F. Makdisi and Seed provided empirical formulas to estimate thespectral amplitudes of the first three periods of vibration of an embankment dam andcombine them into an approximate crest acceleration, based on mode superpositionprinciples. For a base motion with 0.60g PGA (Hayward event), the Seed-Makdisi's peakcrest acceleration ranges from about 1.1g to 1.3 g, depending on whether the alluvium isincluded in the analysis, hence is about twice the PGA of the input motion. For a SanAndreas event (0.29g PGA), crest accelerations estimated through the Makdisi-Seedprocedure range from 0.62g to 0.64g, hence provide similar amplification ratios. While thesevalues represent simplistic estimates and may be on the high side, they suggest that the crestacceleration computed in 1976 (0.39g for 0.40g PGA) was perhaps too low. Empiricalcorrelations developed from earthquake motions recorded on dams (Idriss, 2004) suggestthat, at about 0.40g peak base acceleration, dam crest accelerations could approach up to0.65g.The fundamental period of vibration of the dam reported in 1976, 1.97 sec, appears to bevery long, as compared with the periods calculated in the Makdisi-Seed procedure or thefundamental periods estimated from empirical formulas suggested by Dr. Idriss in his reviewof the GEI draft report. For the Hayward event (0.60g PGA), the Makdisi-Seed procedureleads to periods that range from 0.8 sec to 1.6 sec, whether or not the thickness of thealluvium is assumed to be included in the overall height of the dam. Periods calculatedthrough this procedure are shorter for the San Andreas event (0.29g PGA) than the Haywardevent, due to the lesser reduction in dynamic modulus stiffness, and range from 0.6 sec to1.25 sec. Based on Dr. Idriss’ formulas and assuming an average uniform low-strain V s of1,000 ft/sec (measured V s ranged from 500 ft/sec to 1,500 ft/sec), the low-strain fundamentalGEI Consultants - 50 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05period of a triangular wedge of same height as the dam (132 ft) would be 0.3 sec [ T 0 = 2.4H/V s ], and the period of a semi-infinite soil layer of thickness equal to the dam plus theunderlying alluvium (132 ft + 90 ft = 222 ft) would be about 0.9 sec [ T 0 = 4 H/V s ]. Hence,the low-strain period of the dam-foundation system, based on those formulas, should besomewhere between 0.3 and 0.9 sec. For average shaking conditions representing theHayward Earthquake, the Makdisi-Seed’s procedure indicate a modulus reduction of about80 percent, or an average V s equal to about 45 percent of the low-strain V s . This wouldcorrespond to a lengthened fundamental period of 0.7 sec for the triangular wedge and 2 secfor a semi-infinite layer. This range, 0.7-2.0 sec, is consistent with that obtained in theMakdisi-Seed procedure (0.8-1.6 sec) and suggests that the fundamental period computed byWahler was probably too long.As periods of significance to the Wahler response analysis were substantially longer than theperiod at which peak spectral amplitudes occur, this might have contributed tounderestimating the response of Lafayette Dam. The fundamental period of 1.97 seccalculated in the Wahler analysis for the San Andreas event may also result from too lowiterated dynamic moduli and too high damping values, combined with the probableinsufficient resolution (element size) of the numerical model.As Wahler found the simulated San Andreas event to be the most critical (Earthquake “B”),they did not include detailed results concerning their Hayward Earthquake (Earthquake “A”)in their report. We noted, however, that the response spectrum of Earthquake A wasinsufficiently conservative at the periods of significance to the dam response, a probablereason why the San Andreas event was found to be the most critical in 1976.As normally done for EQL analyses, different computer programs were used by Wahler forthe initial static (program LSTRN) and dynamic response analyses (program QUAD-4).Such programs are not fully compatible, which adds some uncertainty, although notsignificant, when their results are combined. LSTRN was described as a linear-elasticanalysis program, a possible limitation when used for soil materials. Static analysisprograms using hyperbolic constitutive relationships are now preferred for use in parallelwith EQL dynamic analysis. Also, Wahler used an early version of QUAD-4, and asubsequently recognized coding error in one of the subroutines may have affected thecomputed stresses. Such coding error was discussed in the review of the 1976 results by theDSOD (Memorandum of Design Review, dated 3/4/1977).Strain potentials were computed in a small number of elements using a decoupled procedure.An improved method for estimating earthquake-induced deformations from strain potentials(Serff, et al., 1976) was not yet commonly used at the time Wahler performed its EQLanalyses. Lastly, no post-earthquake static stability analyses that would take into accountany potential loss of strength experienced as a result of dynamic straining have beenperformed. As shear displacements of up to nine feet were predicted within the upstreamGEI Consultants - 51 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05foundation, the residual shear strength of the affected materials would be mobilized in thepost-earthquake condition. We also noted that the Wahler study found the upstreamfoundation to be the most critical under seismic loading, while the history of the dam andstrength data suggest that the downstream foundation is the weakest.While EQL analysis procedures are acceptable for non-liquefiable materials and when anappreciable margin of safety can be demonstrated, the early history of Lafayette Dam mustbe taken into consideration. Although the failed foundation and embankment materials musthave consolidated and gained substantial strength since 1928, one must keep in mind that thewide-crested incomplete dam experienced substantial flow failure displacements that led tocrest settlements of 24 to 26 feet in the central portion, under static loading conditions. Thisis more than the current freeboard (17.8 feet). Previously failed materials were neverremoved from the foundation and downstream slope of the dam, and could be sensitive tosustained dynamic loading. From sole consideration of the previous dynamic analysisresults, one cannot rule out that such materials might be prone to experiencing large nonrecoverabledisplacements, under exceptionally demanding earthquake loads and longdurations of shaking. Such consideration and the limitations of the 1976 dynamic analysesraise the question of whether the District should consider implementing modern evaluationprocedures to more reliably assess the seismic performance of Lafayette Dam.The simplified analyses by DSOD used the June 28, 1992 Lucerne Valley record (M w 7.3),which represents near-field (1.1 km distance) ground motion at a very shallow soil/rock site.We did not find information regarding which component of this record was used. TheLucerne acceleration histories contain near-field effects at periods well above 2 sec, but theirfrequency content at periods between 0.5 sec and 2 sec seem to fall substantially below therecommended 84 th percentile spectral requirements. Hence, if not modified in frequencycontent to match the requirements of the Lafayette site, deformations estimates obtained withthe Lucerne record(s) may be on the low side. Updated deformation estimates, usingsimplified evaluation procedures, are discussed in Section 7.8.7.3 Review of Material PropertiesThis section presents the results of our review of existing data describing the dam andfoundation material properties, and our selection of analysis parameters to perform simplifiedanalyses of Lafayette Dam. Table 7-2 presents the summary of our interpretation of averageunit weights and moisture contents, and Table 7-3 presents our recommended total stress andeffective strength parameters. Our estimated strength parameters are based on a criticalreview of the strength testing performed to date and, as needed, on our reinterpretation of thetriaxial test data and Mohr failure envelopes. Detailed review and our selection process forthe strength parameters are discussed in Appendix E. Dr. Idriss, in his review of the GEIdraft report, noted that assuming a cohesion intercept could be unconservative for the shallowGEI Consultants - 52 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05segments of any potential failure surface. However, since we found that the most criticalfailure surfaces involve the whole upstream slope, or the entire downstream slope plus thefoundation alluvium, we used the cohesion values shown in Table 7-3 for all zones andlocations within the analyzed cross-section.The principal soil parameters normally required for the detailed seismic evaluation of anembankment dam are unit weight, shear and bulk modulus, static (total or effective stress)strength, cyclic and residual shear strengths, and damping coefficient. In addition, and asneeded for the evaluation of dynamic pore pressures in sandy materials, gradationcharacteristics, relative density D R , porosity, degree of saturation, percent fines content andrate of build-up of excess pore pressures may be needed. Because of the clayey nature ofLafayette Dam and foundation materials, and because the previous studies only includedconventional slope stability or EQL dynamic analyses, some of the above parameters werenot required. Furthermore, dam studies from the 1960’s and 1970’s did not consider theconcept of residual strength. Our review and update of analysis parameters included thefollowing steps:• Review existing field and laboratory data;• Re-interpret some of the triaxial testing data, based on current practice;• Review previous static and dynamic analysis parameters; and• Select updated analysis parameters.We updated the material properties primarily from those tested by Shannon & Wilson (1966)and W.A. Wahler and Associates (1976). Based on our review, we distinguish the samethree zones within the dam, the core (Zone 1), the downstream shell (Zone 2) and theupstream shell (Zone 3), see Figure 7-2. However we believe that only two zones need to beconsidered in the foundation alluvium. Zone 4 comprises the foundation alluvium at about115 feet depth or less below the downstream slope surface, and at and beyond thedownstream toe. Zone 5 primarily includes the foundation soils below the core and upstreamshell, and is the same as shown on Figure 7-2. We also considered some of the deeperalluvium below the downstream shell as being part of Zone 5.Our Zone 4 includes weaker materials probably affected by the 1928 failure, as identifiedfrom available field penetration and laboratory testing data, and the Zone 4 and Zone 4.5shown on Figure 7-2. Based on the limited information available for the alluvium below thedownstream toe of the dam, we found no strong reason to differentiate between Zones 4 and4.5.Hence, our slope stability analyses assigned the same average material properties to Zone 4and 4.5 previously considered by the DSOD in its review. For our static and pseudo-staticslope stability analysis of Lafayette Dam, we assumed the material properties listed in Tables7-3 and 7-4.GEI Consultants - 53 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05For some of our estimations of earthquake-induced deformations, we used the genericmodulus reduction curve and lower bound damping curves originally proposed by Makdisi-Seed (1978). Considering the approximate nature of this simplified method to estimatedeformations, we consider the use of such generic curves sufficient. It is also appropriate, asMakdisi and Seed primarily used clayey dams in the development of their methodology andconsidered such curves to be suitable. In their EQL studies, Wahler used the Seed and Idriss(1970) shear modulus and damping curves, which are generally similar in shape and of acomparable vintage to the ones we used. We also specified an approximate averagemaximum shear wave velocity of 1,000 feet per second, and an average unit weight of 130pcf in our simplified analysis.7.4 Phreatic Surface AssumptionWe compared the phreatic surfaces assumed within the embankment in the S&W 1966 slopestability analyses and in the Wahler 1976 seismic stability evaluation with our interpretationof recorded piezometric measurements. Such comparison indicates slight differences betweenour and the earlier interpretations, or reflects improved readings from the piezometers thathave been replaced since 1976. While such readings have been consistent in recent years, allpiezometers are open-standpipe instruments and cannot reliably assess actual pore pressuresat specific locations within the embankment or foundation. Hence, there could be someuncertainty associated with past and current interpretations of the position of the phreaticsurface within the embankment, which does not seem to be influenced by the intended damzoning.The phreatic surface interpreted in earlier studies was higher in the upstream shell and core,but slightly lower in the D/S shell than presently assumed. S&W and Wahler took a nearhorizontalsurface to represent the water level in the U/S shell and in the U/S half of the core,and then assumed a sloping near-planar phreatic surface, extending from about the center ofthe core to slightly beyond the lower D/S berm, and then reaching the D/S toe at a flatterslope angle. From our interpretation, the average phreatic surface drops at a flatter slope,nearly constant from the U/S face of the dam to the downstream toe (see slope stabilityanalysis model), see Figure 7.3. Hence, our interpreted phreatic surface is lower than wasdefined by S&W and Wahler in the U/S shell, but slightly higher in the D/S shell. Suchdifferences should have little impact on the results of the slope stability analyses. Insummary, we draw two conclusions from our updated interpretation of the phreatic surface:• First, the core does not lower the phreatic surface within the embankment, as mightnormally be expectable from an “impervious” zone. This may be because the coreand shells have similar coefficients of permeability, due to their high fines content, orbecause of stratification. Hence, the zoned Lafayette Dam functions as aGEI Consultants - 54 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05homogeneous embankment regarding the location of the phreatic surface within thedam section.• Overall, the phreatic level within the D/S half of the embankment is rather high,considering the high clay content and expected low permeability of the embankmentmaterials. This could indicate that some seepage may travel through zones affectedby the 1928 failure, or that the embankment collects surface runoff from the crest andside hills.Overall, other than potentially making the local clays more plastic, the seepage through theembankment, as indicated from the seepage collection systems, is quite low, and isconsidered to represent satisfactory performance of the embankment.7.5 Updated Slope Stability Analysis7.5.1 Analysis PropertiesOur analyses for steady-state seepage (normal operating) condition are based on effectivestressstrength parameters (c’, Φ’), developed from updated interpretation of the availableisotropically-consolidated undrained (ICU) or anisotropically-consolidated undrained (ACU)triaxial (TX) tests. Because of the clayey nature of the embankment and foundationmaterials, we considered the use of the undrained shear strength (S u ) for the slope stabilityanalyses, but did not use such an approach because it is impractical to reliably defineincreases in S u as a function of depth in the slope stability model, where deep surfaces offailure need to be considered. Also, the variability of the measured S u s makes it difficult toselect truly representative values.Wahler reported that disturbance did occur during the sampling process, especially for deepsamples, and that the S u s obtained from unconfined compression (UC) or unconsolidatedundrained(UU) triaxial tests might significantly underestimate the actual strength availableat the depth where the samples were recovered. Yet, Wahler compared S u s in the core, shelland foundation materials, as obtained from UC, UU, or TXICU and TXACU laboratory tests(see Drawing V-2 of the 1976 report) and defined approximate linear increases in S u as afunction of the consolidation pressure P o . Because these S u -P o average relationships seemprimarily controlled by the CU data, we concluded the use of the effective stress strengthparameters derived from the CU tests would be more reliable than direct use of undrainedshear strengths.For slope stability analyses for rapid drawdown condition and to determine “yieldacceleration” coefficients (for conditions representing those that might occur when a slopeGEI Consultants - 55 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05under steady-state condition is subjected to undrained loading such as an earthquake), weperformed analyses based on total-stress shear strength envelopes (c, Φ) obtained from theCU triaxial tests.7.5.2 Analysis Model and ResultsWe performed slope stability analyses for both the upstream and downstream slopes ofLafayette Dam. We used the computer program XSTABL, version 5.2 (ISD, Inc., 1992-2003) for this purpose. The program calculates slope stability using a limit equilibriumanalysis based on the method of slices. The program calculates the forces that would causeand the forces that would resist movement of a soil mass bounded by a postulated failuresurface. The program automatically defines the most critical postulated failure surfacethrough a searching routine. In XSTABL, several methods of slope stability analysis can beactivated. These methods are derived from the standard method of slices and Bishop’smethod. The searching routine employs either the simplified Bishop or Janbu (1954)’smethods of analyses. The Janbu method is particularly useful to analyze the influence ofpartial submergence and drawdown conditions and, as needed, the effect of tension cracksand surcharge. More rigorous methods, such as Spencer’s Method (1967), the General LimitEquilibrium Method or Janbu’s Generalized Method of Slices may be subsequently used inXSTABL to further assess the safety evaluation of any single surface.We successively implemented the Janbu and Spencer’s methods of analysis. Searches forcritical surfaces were performed using Janbu’s method, and the critical failure surfaces werethen reevaluated using Spencer’s method. Spencer’s method is based on cylindrical failuresurfaces and satisfies two equations of equilibrium, the first with respect to forces, and thesecond with respect to moments. It assumes parallel inter-slice forces. For pseudo-staticseismic analysis, we represented inertial forces due to the earthquake loading with a constanthorizontal acceleration coefficient (a H ). As is often done in conventional slope stabilityanalyses, we did not use a vertical acceleration coefficient (a V ).Our slope stability analysis model is shown on Figure 7.3. As previously discussed, wecombined Zone 4 and Zone 4.5 of the DSOD model (Figure 7-2) as a single Zone 4. Thephreatic surface interpreted from the piezometric readings is shown on Figure 7.3. Asummary of the computed factors of safety is provided in Table 7-4. We obtainedcomparable factors of safety for either Janbu or Spencer’s methods, and Table 7-4 applies toeither of these two methods. The PMF load case, normally considered for dams, was outsidethe scope of our review.Slightly higher factors of safety were computed in Spencer’s Method. For steady-stateseepage static condition and a reservoir elevation at El. 449, the lowest factor of safety wecalculated is 2.3 for the downstream slope, and 2.5 for the upstream slope. These valuesGEI Consultants - 56 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05confirm the satisfactory performance of the embankment to-date. For partial rapiddrawdown condition (repeat of the maximum historic reservoir drawdown to El. 431), wecalculated a minimum factor of safety of 2.0. We also postulated a rapid completedrawdown to the elevation of the lowest outlet port, and obtained a minimum factor of safetyof 1.7.As this review was primarily intended to assess the seismic stability of Lafayette Dam, wehave presented detailed results only for the pseudo-static analysis cases and, especially, asused to verify K y , the “yield acceleration”. K y was obtained using the total-stress method ofanalysis by progressive increase of the pseudo-static acceleration, starting from the mostcritical surfaces obtained in a static analysis with the total stress strength parameters, until afactor of safety equal to or less than 1.0 was obtained. Hence, K y is the a H that correspondsto a factor of safety of exactly 1.0. Multiple trial failure surfaces were then re-analyzed withK y to assure that the most critical surface had been found. Cross sections of the dam showingthe locations of the most critical surfaces (a H = K y ) are shown on Figure 7-4 for the upstreamslope, and on Figure 7-5 for the downstream slope.A “yield acceleration” of 0.29g was computed for the upstream slope in both the Janbu andSpencer methods of analysis, using total stress strength parameters. The critical failuresurface is entirely located within the embankment materials. This is not surprising, as theassigned total strength parameters are higher for Zone 5 than for the core (Zone 1) orupstream shell (Zone 3) materials.For the downstream slope, we obtained a yield acceleration of 0.14g. The correspondingfailure surface passes through the foundation Zone 4, which has a lower assigned strength.Zone 4 of our model combines Zones 4 and 4.5 of the DSOD model. The existing field andlaboratory data may be insufficient to properly define the strength of the downstreamalluvium for the earthquake loading condition. For this reason, we have selected aconservative strength estimate, based on the available field penetration data and laboratorytesting results which suggest that weaker materials might still be present.For deformation analysis purposes, we took a K y of 0.29g for the upstream slope, and K y of0.14g for the downstream slope. Hence, the assumed weaker alluvium materials willinfluence our seismic analyses of the downstream slope.7.6 Evaluation of Liquefaction PotentialThe liquefaction potential of the various dam zones was reviewed. No “loose” saturated siltsor sands, generally acknowledged the most susceptible to liquefaction, have beenencountered in the borings. The embankment materials are classified as clays, sandy clays, orsilty clays. CL and CH are the dominant soil classifications in dam and foundation materials,with ML occasionally encountered. Neither continuous layers nor significant lenses of cleanGEI Consultants - 57 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05sands or silts have been reported in the previous field investigations. Shannon and Wilson(S&W, 1966) reported that “the soils encountered at the site are not sensitive toearthquakes…. as would be sensitive clays or loose sands” and W.A Wahler and Associates(WA, 1976) concluded that, in these materials, “no instantaneous or sudden loss of strengthoccurred during the dynamic testing” and that “significant liquefaction would not occur”.Using the “Modified Chinese Criteria” (Zhou, 1981) and a modified version of these criteria(Andrews and Martin, 2000), the DSOD concluded that “the embankment and foundationsoils are not expected to liquefy, due to high clay content and liquid limit”.In his review, Dr. Idriss indicated his belief that the Chinese criteria should not be used forcohesive soils. Yet, the clayey characteristics and plasticity indexes of the local materialscan be used in recently updated approaches to further evaluate whether or not these would besusceptible to liquefaction. Average plasticity index (PI) is 30 in the core materials, 15 in theD/S shell (Zone 2) and 22 in the U/S shell (Zone 3). The PI ranges from 18 to 21 in thefoundation materials, depending on which zone is considered. A new classification has beenproposed by R.B. Seed, at al. (2003) to replace the “Modified Chinese Criteria” and the“percent fines rule”, which are based on one single key parameter. This new interimclassification of “liquefiable” soil types is based on two “key” parameters, the LL and PI. AsS&W and Wahler performed numerous Atterberg limits (AL) tests, we recompiled these dataand plotted them on a reference chart that provides a way to quickly assess if the soils are“liquefiable”, see Figure 7-6 (S&W data) and Figure 7-7 (WA data). We also preparedsimilar plots, zone-by-zone (see Appendix E).Most AL tests on the dam and foundation soils fall outside of Zone A (“classic” cyclicliquefaction) or Zone B (potentially liquefiable). The core (Zone 1) materials clearly falloutside Zone A or Zone B. Three data points for the downstream shell (Zone 2) are withinZone A or Zone B. One data point for the upstream shell (Zone 3) is within Zone B andanother is on the boundary. Over half of the data points for the foundation (Zones 4, 4.5 and5) fall within Zone A or Zone B. However, only one of the data points within Zones A or Bfails the supplementary test [ w (%) > 0.8 LL ] regarding liquefaction susceptibility. Hence,based on AL tests and the high clay content of all the materials encountered in LafayetteDam and its foundation, we conclude that the embankment and foundation materials are notsusceptible of liquefying by sudden loss of strength.7.7 Computed Earthquake-Induced DeformationsNo detailed response analyses of Lafayette Dam were included in this review. However, weimplemented several simplified or empirical procedures to obtain estimates of potentialearthquake-induced deformations as a result of the MCE earthquake scenarios the mostcritical to Lafayette Dam. These methods include: Newmark (1965); Makdisi-Seed (1977);Bureau, et al. (1985, 1987); Jansen (1987); and Swaisgood (1995, 1998). We alsoimplemented a modified Makdisi-Seed’s procedure, based on recommendations made by Dr.GEI Consultants - 58 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Idriss in his review of our draft report (letter-report to GEI from I.M. Idriss, dated August 31,2004). Such approach, referred herein as the “Idriss’ procedure”, replaces the dam crest andfailed soil mass accelerations iteratively computed from the specified response spectrum inthe Makdisi-Seed’s method by estimates derived from empirical correlations between baseand crest peak horizontal accelerations obtained on several existing dams during historicearthquakes.Three of the methods implemented use the yield acceleration (K y ) as an input parameter,hence provide different displacement estimates for the upstream and downstream slopes.The two slopes were successively considered. The other methods depend on the energycontent of the specified input motion and calculate crest settlements, hence provide a singleestimate for a specified dam configuration and postulated earthquake scenario.Other assumptions significant to the implementation of the above procedures are brieflydiscussed below. A more detailed review of these procedures can be found in Appendix F.Dynamic Strength Reduction FactorAs is customary for earthquake-induced deformation estimates in clayey materials, thecomputed yield accelerations (K y ), which were based on undrained (total-stress) Mohr-Coulomb strength envelopes, were reduced by 20 percent in our slope deformation analysesto account for any loss of strength under cyclic loading and initiation of possible soilmovements. This procedure approximates the reduction in strength proposed by Makdisi andSeed (1978) for the K y calculations. Dr. Idriss suggested an initial 15 percent reduction ofthe static strength due to shaking. Such reduction was specifically applied in his procedure,in lieu of the 20 percent used in some of the other methods. Dr. Idriss further suggested that,“if considerable displacements were calculated (say one foot or so for a few cycles ofshaking), the strength would have to be reduced to a residual value, best established basedon in-situ shear vane tests”. As no information on the residual strength was available in thedata reviewed and no time-history analysis was performed, such additional correction wasnot implemented, but suggests that deformations computed in Dr. Idriss’ and other methodsbased on the concept of yield acceleration K y could be potentially underestimated. Instead,we used a conservative reduction of the calculated K y s, in recognition of such uncertainty.Influence of Foundation AlluviumExcept for Swaisgood’s method, the simplified analyses implemented do not consider thepresence of the potentially deformable foundation alluvium. Therefore, we first obtained“lower-bound” deformation estimates by assuming that Lafayette Dam is founded on a rigidfoundation, then “upper-bound” estimates by assuming the plastic foundation alluvium to bepart of the embankment (hence, we assumed an “equivalent” 222-foot high dam).GEI Consultants - 59 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Crest Settlement Versus DisplacementsAs some methods calculate maximum displacements along postulated failure surfaces, whileothers directly calculate maximum crest settlements, all results were approximatelyconverted to crest settlements by assuming that the maximum crest settlement would be halfof the corresponding computed maximum displacement. The factor 0.5 was selected afterobserving that maximum displacements reported after the 1928 failure were more than twicethe maximum crest settlement, and based on experience of GEI’s Project Manager withdetailed nonlinear dynamic response analyses of six embankment dams. For such timehistoryanalyses, performed from 1994 to 2001 by Mr. Bureau when employed by Dames &Moore, the average of the ratios of computed crest settlement and maximum slopedisplacement calculated from these studies was 0.46.ResultsWe obtained “best” estimates by averaging crest settlements obtained with the upstream ordownstream yield accelerations, and for all procedures implemented. The results of ouranalyses are summarized in Table 7-5 for each of the earthquake scenarios considered, andare presented in more detail in Appendix F, Tables F-1 to F-4. As expected, a wide range ofdeformations is predicted by the different methods, which all have limitations and apply to arange of conditions broader than considered herein.The results of the simplified Newmark, Seed-Makdisi and Idriss’ procedures are stronglyinfluenced by the computed yield acceleration of the downstream slope. Idriss’ procedureleads to considerably lower deformations than obtained in the Makdisi-Seed’s procedure, dueto a lower estimated crest and resulting failed soil mass accelerations (crest accelerations:Idriss, 0.76g; Makdisi-Seed, 1.08g or 1.33g, depending on whether the alluvium is includedor not).All computed or estimated crest settlements are significantly less than the availablefreeboard, but upper-bound estimates (dam + alluvium) range from about 2 feet to 7 feet, orabout five percent of the dam height (132 feet), for the most demanding earthquake scenario.Because the computed crest settlements are based on the combined deformations of theembankment plus alluvium, the corresponding cumulative strain levels in these materialsshould be less than about 3 percent. Without rigorous consideration of the residual strength(which might increase displacement estimates), the procedure recommended by Dr. Idrissleads to downstream slope deformation estimates ranging from 0.5 feet (U/S slope) to 2.6feet (D/S slope), for the most critical earthquake scenario.GEI Consultants - 60 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Hayward EarthquakeAfter considering all the procedures implemented as a whole, our “best” (weighted) estimatesof crest settlements for the Hayward Earthquake, which is the most critical earthquakescenario considered, vary from 0.9 to 4.5 feet. An average crest settlement of 2.7 feetrepresents our “preferred” mean prediction. We have chosen the term “preferred” to indicatethat we have simultaneously considered several methods to compute deformations orsettlements. Such methods involve simplified procedures, which all have limitations on howthey can be applied to specific seismic, embankment and foundation conditions. Importantfactors, such as the presence of the alluvium and how the seismic loading would be trulyapplied, are only approximately or not taken into account in these procedures. Our use of“preferred” average settlement estimates is intended to reduce the potential margin of errorassociated with the lowest or largest estimated dam movements.Other Earthquake ScenariosWe did not use the Idriss’ procedure for the other earthquake scenarios, as they all resulted insignificantly lower crest settlement estimates than for the Hayward Earthquake. Based on theresults obtained for the Hayward MCE, including the Idriss’ procedure in our averagingprocess would lower the “preferred” estimates described below by about 10 percent.For the Calaveras Earthquake, our mean estimates range from 0.8 to 2.4 feet, with a preferredprediction of 1.6 feet. For the San Andreas Earthquake, we estimated crest settlementsbetween 0.2 and 2.8 feet, with a preferred estimate of 1.5 feet. This result actually compareswell the 2.0 to 3.0 feet previously obtained in 1976 for the San Andreas event by Wahler,probably because our recommended San Andreas response spectrum and the responsespectrum of Earthquake “B” used by Wahler have reasonably compatible energy content inthe range of frequencies of interest to Lafayette Dam. Lastly, we calculated a range ofsettlements between 1.0 and 5.3 feet for the Lafayette-Reliez Valley Earthquake, and apreferred estimate of 2.2 feet.The average crest non-recoverable settlements obtained in our simplified analysis aresubstantially less than the available freeboard (17.8 ft). Hence, Lafayette Dam is likely tosafely retain the reservoir under the MCE. Our analysis procedures, under the mostconservative analysis conditions, result in estimates that range from several feet to about aquarter of the magnitude of the dam crest settlements and slope movements that occurred in1928 (which averaged about 20 and 40 feet, respectively). While such numbers suggest acondition less critical than previously encountered during construction, these computeddisplacements are large enough to partially or totally mobilize the residual shear strength,which was taken into consideration based on approximate strength reduction factors and noton measured values.GEI Consultants - 61 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Overall, Lafayette Dam is reasonably safe for the MCE, considering its large operatingfreeboard and wide crest. However, actual behavior of this dam may not be predicted withsufficient accuracy through simplified analysis procedures, considering the relatively highpotential earthquake loads to which it is exposed. Because the foundation soils were themain contributor to the dam failure in 1928, earthquake-induced seismic deformations wouldalso likely involve both the dam and underlying foundation alluvium, thereby increasing thecombined crest settlements. It is not clear whether the failure surface of the 1928 representsa potentially weaker, thin zone within the dam section. Such failure surface was not clearlyidentified in the previous field investigations, other than through recognition that severalzones are present within the alluvium. While the foundation alluvium has undoubtedlygained strength under the loading provided by the dam body, and excess pore water pressurescaused by the added weight of the newly constructed dam have now dissipated, we believethat, based on the history of this dam, further investigation would be desirable to confirm thein-situ strength of the downstream portion of the alluvium, and demonstrate thatunacceptable non-recoverable deformations cannot be induced by extreme earthquakescenarios.7.8 Review of 1976 Cyclic Triaxial Test DataW.A. Wahler & Associates (Wahler) performed in 1976 stress-controlled and straincontrolledcyclic triaxial tests. The Wahler dynamic testing laboratory was long consideredas being one of the best-equipped and best-experienced to perform such tests. The stresscontrolledcyclic tests were performed on isotropically consolidated (K c =1) or anisotropicallyconsolidated samples (K c =1.5), for various confining pressures (σ’ 3 ). Tests were performedon the core, shell and foundation materials, and were used as a basis to compute the strainpotential in some elements of the 1976 finite element model. None of the specimens testedexperienced instantaneous or sudden loss of strength, consistent with the clayey nature of theLafayette Dam and foundation materials. These test results suggest that a similar behavior(no “classic” liquefaction) would likely be observed in the field.Using the data presented in the Wahler report, it is possible to define cyclic strength curvesbased on the cumulative number of applied uniform cyclic deviator stress cycles when thesamples tested reached 5 percent, 10 percent axial strain, or other interpolated or extrapolatedaxial strain levels. For example, from the Wahler cyclic triaxial tests results and based on thecyclic deviator stress plots presented in the 1976 report, we calculated the laboratory cyclicstress ratios (CSR) causing 10 percent axial strain in the tested specimens at 10, 15, 20 and30 cycles, see Table 7-6. However, laboratory cyclic triaxial tests do not represent fieldconditions as well as simple cyclic shear tests would do. Yet, the CSR causing a specifiedlevel of axial strain in the laboratory tests (σ dc /2σ’ 3 ) can be approximately related to the fieldCSRs (τ h /σ’ v ) under multi-dimensional shaking conditions by the following expression:GEI Consultants - 62 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05τ h / σ’ v = C r σ dc / 2σ’ 3In such equation, τ h represents the earthquake-induced horizontal shear stress, σ’ v the verticaleffective overburden pressure, σ dc the applied deviatoric cyclic stress in the laboratorysample, and σ’ 3 is the lateral initial consolidation stress of such sample. The correctioncoefficient C r varies from about 0.6 for K o = 0.4, to about 0.9 to 1.0 for K o = 1, where K odesignates the coefficient of earth pressure at rest (Seed and Idriss, 1982).Overall, there seems to be little variability, relatively to each other, in the cyclic strengths ofthe foundation alluvium, shell or core materials. As Wahler did in 1976, we interpretedsome of the cyclic strength data based on trends as, at some confining stress levels, notenough samples were tested to fully define cyclic strength curves.The downstream foundation alluvium (Zone 4) shows lower laboratory CSRs than theupstream foundation (Zone 5). Based on only two series of cyclic triaxial tests performed forthe core materials, one series (6,000 psf) resulted in the lowest CSR at N equal 15 or greater,compared with the alluvium or the downstream shell materials, while the other (3,000 psf)led to higher CSRs than these other materials. Few samples were tested in the shell materials(Zone 2), but most of these tests yielded laboratory CSRs higher than for the other materials.Additional details are provided in Appendix E.Our slope stability analyses and the simplified methods we used to obtain earthquakeinduceddeformations did not require cyclic strength curves to make a preliminaryassessment of the performance of the dam. Cyclic shear strength data are normally used forinterpreting the results of more detailed numerical analyses, and we do not know K o and haveno reliable field CSRs. However, we computed the CSRs for an “equivalent referencemagnitude” of 7.5 (CSR eq ), using H.B. Seed’s simplified method as updated by Idriss (1999).While a detailed direct comparison of these CSR eq s with field CSRs converted from thelaboratory CSRs (C r σ dc / 2σ’ 3 ) shown in Table 7-6 is of limited value, considering theuncertainties of such simplified methods, we concluded that the CSR eq s, converted to thereference magnitude of 7.5 and estimated for the Hayward Earthquake (M w 7.25) by thesimplified Seed-Idriss procedure equal or exceed the laboratory CSRs causing 10 percentaxial strain or greater and corrected for field condition, for an equivalent number ofapplicable stress cycles (N=15).Hence, the above simplistic comparison suggests that simplified analysis procedures are notsufficient to fully assess the seismic performance of Lafayette Dam. Appreciable earthquakeinduceddeformations appear to be possible, under some of the severe earthquake scenariospostulated, and should be further evaluated.GEI Consultants - 63 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/058. ADEQUACY OF MAINTENANCEAND METHODS OF OPERATION8.1 Operation, Maintenance and Surveillance ProceduresOperational procedures that relate to project safety were briefly outlined in Section 2.4 of thisreport. Our inspection indicated no conditions that would require emergency action orchanges to the current operational procedures. Project facilities are visited regularly andmaintenance is scheduled as needed.During the last twenty years, maintenance of the dam has been regular and cosmetic, and hassimply involved installation of new piezometers. The maintenance performed on the damappears adequate with regard to maintaining this dam in excellent condition. Thesurveillance program appears to be adequate.8.2 EvaluationNo emergency maintenance measures are required for public safety under normal operatingcondition. We did not identify any inadequacy of methods of operation that wouldpotentially affect public safety. Based on the results of this review, we do not recommendany changes in maintenance procedures or methods of operation.GEI Consultants - 64 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/059. CONCLUSIONSThis section summarizes the findings of our inspection and seismic stability review ofLafayette Dam. This work was carried out in accordance with current, generally acceptedsimplified procedures for dam safety evaluation. The purpose of this evaluation was toassess the seismic safety of the structure for continuing operation in the interest of theDistrict’s water system requirements and maintenance of public safety, based on existingdata. No additional field exploration, laboratory testing, or detailed numerical analyses wereincluded in the scope of work.Our conclusions are based on observations and findings drawn from: (1) a detailed review ofexisting project files, made available to the project team by the District and the DSOD; (2)visual and geologic inspections of Lafayette Dam and adjacent areas, including the reservoirperimeter; (3) review of project maintenance, operation, and instrumentation monitoringrecords; (4) review of the seismic setting of the site and the development of response spectrafor the most critical MCE scenarios applicable to this site; and (5) performance of slopestability analyses and simplified evaluations of potential earthquake-induced deformations.Our safety review focused on the dam embankment and evaluating its ability to impound thereservoir in regard to its adequacy against possible catastrophic failure due to earthquakes.Other natural phenomena, such as flooding or excessive precipitation, were not investigatedbut are of lesser concern for this site because the drainage area is small and the reservoir ismostly filled with imported water, rather than with natural stream flow.The outlet tower, which also serves as spillway, was not included in our review, but has beenshown by others to have insufficient seismic capacity. No consideration has been given or isintended to those public safety aspects of the project features other than the dam stability.9.1 Construction HistoryLafayette Dam experienced in 1928 a major downstream slope failure during construction,and was built to a lower crest elevation than originally designed, keeping the final crestessentially at the same width and elevation as the incomplete embankment after failure. Thedam has satisfactorily performed since. The embankment was built on a thick layer of plasticclayey alluvium, up to an average of 90 feet deep below the central portion of the dam. Thefoundation alluvium was the primary cause of failure during construction, according to thereport that was prepared by a renowned Consulting Board who investigated the failure. Thefailed embankment and foundation materials were not removed. Construction was completedby leveling the crest of the failed dam, by backfilling failure cracks, and by flattening thedownstream slope by placement of additional fill. Such procedures would be unacceptableGEI Consultants - 65 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05by today’s standards in dam construction. The failed materials have stabilized shortly afterthe end of construction, and have undoubtedly consolidated and gained strength over time.9.2 Assessment of Dam and Reservoir PerimeterThe reservoir is used for emergency backup water supply and is only subject to minor levelfluctuations. A substantial freeboard (17.8 feet) is maintained. The watershed area is verysmall. All reservoir water is imported and no significantly stream inflow can occur, whichessentially eliminates risks of flood overtopping or uncontrolled spillway releases.Lafayette Dam is well maintained by the District staff and in good visual condition. No signsof instability or inadequacy of the embankment that would require emergency remedialaction were observed. Seepage levels are minimum and consistent with the norms for thedam and reservoir level. The limited seepage is not detrimental to the safety of the dam and iscontinuously monitored.Recent geologic literature mentions discontinuous lineaments and possible inferred faultingin the immediate project vicinity. As discussed in this report, such features do not seem torepresent a threat to the dam. The inferred faults are short, and one that has been shown topotentially intersect the dam footprint is parallel, rather than perpendicular to the dam crest.Hence, if further confirmed, it would unlikely be critical in terms of direct or sympatheticrelative movements, because of its short length and favorable orientation.9.3 Adequacy of Instrumentation, Monitoring ofInstrumentation and SurveillanceThe instrumentation and dam monitoring program are adequate. The existing seepagemeasurement systems properly monitor seepage flows. Several piezometers show highreadings and fluctuations during the wet season, possibly due to contributing surface runoff.Horizontal and vertical crest or slope movements have been insignificant for many years, andare continuously monitored. The settlement and movement surveys are inconsequential for adam over 76 years old. Methods of project operation relating to public safety are adequate.GEI Consultants - 66 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/059.4 Adequacy of Operation of Spillway and Outlet WorksThere is no separate spillway at Lafayette Dam, and infrequent releases are performedthrough the uncontrolled upper 2.5’ by 3’ open port in the outlet tower. The seismic stabilityof this tower has been questioned in recent studies initiated by the District. Emergencyrelease of the reservoir water, should it be required after an earthquake for upstream slopeinspection or safety purpose, might be impaired or impossible due to potential tower failure.Outflow of reservoir water through the outlet pipes and conduits, could be either uncontrolledor impaired by clogging of these elements by tower debris. We did not evaluate the seismicsafety of the outlet conduit.9.5 Updating of Seismic CriteriaThe response spectra developed for the Hayward and Lafayette-Reliez Valley faults, whichare potentially the most critical to this site, are significantly more demanding in the range ofperiods of interest to the response of Lafayette Dam and foundation than the responsespectrum of the acceleration time history used in 1976 to represent a “Maximum Probable”Hayward Earthquake; The input history used at that time to represent a “MaximumProbable” San Andreas Earthquake was sufficiently conservative.9.6 Assessment of Material PropertiesNo loose saturated silts or sandy silts, generally acknowledged to be the soil types the mostsusceptible to liquefaction, were encountered in the existing borings. We concluded that theembankment and foundation materials are not “liquefiable”. The available data may beinsufficient to properly define the strength of the downstream foundation alluvium (Zone 4),where existing field penetration data and laboratory testing results suggest that weakermaterials might be present. It is not clear whether the alluvium in that area has developedsufficient additional strength to withstand major earthquake loads. Possible reactivation ofmovement along the 1928 failure surface, which was not clearly identified in the boringsdrilled in the 1966 and 1976 investigations, is unlikely but remains a candidate potentialfailure mode under extreme seismic loading. As was concluded after the 1928 failure, it maystill be that “although the dam had come to rest, there is no evidence that the alluvium haslost its ability to move under sufficient load” (Consulting Board, 1929).GEI Consultants - 67 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/059.7 Assessment of Previous AnalysesWe concluded that the equivalent-linear dynamic analyses performed in 1976 might not fullycapture the possible response of Lafayette Dam, due to insufficient refinement of the finiteelement mesh used to represent the dam and its foundation.9.8 Simplified Stability Analysis and Adequacy of Factors ofSafetyWe have reanalyzed Lafayette Dam to assess stability under normal reservoir operating andrapid drawdown conditions and found it to have adequate stability and factors of safety incompliance with the requirements of modern dam safety evaluations. Pseudo-static loadingconditions and simplified deformation analysis suggest that the dam should have adequateseismic stability and reserve freeboard. Our “best” estimates of potential non-recoverableearthquake-induced deformations, for the most critical earthquake scenario, suggest crestsettlements of up to less than 3 feet. Hence, Lafayette Dam is likely to maintain sufficientfreeboard, as the dam is normally operated with a freeboard of 17.8 feet.The most critical hypothetical failure surfaces considered for the downstream slope of thedam would potentially include both that slope and the underlying foundation alluvium.Three simplified methods of analysis predicted upper-bound deformations that wereapproximately converted into crest settlements of 4 to 7 feet. Such settlements include thecombined contributions of both embankment and alluvium deformations. Estimateddisplacements are large enough that they could induce progressive degradation of theundrained strength of the affected zones. This was only considered in a simplistic fashion inseveral of the procedures (those based on K y ) implemented in this review by assuming thatthe degraded strength would be about 80 percent of the original strength.Conclusions similar to those described above were derived after comparing field cyclic stressratios (CSRs) for an “equivalent reference magnitude” (M 7.5) with laboratory cyclic stressratios causing 10 percent axial strain in 15 uniform stress cycles. In the case of the HaywardEarthquake, most equivalent field CSRs would equal or exceed the laboratory CSRs causing10 percent axial strain or greater, after correction for field condition, for the number of cyclesapplicable to the reference magnitude. Hence, the cyclic triaxial test results available forLafayette Dam suggest that simplified procedures may be insufficient to fully assess theseismic performance of the dam. Appreciable earthquake-induced deformations appearpossible, under the severe earthquake criteria postulated, and suggest that more detailedevaluation of the in-situ strength of the downstream foundation alluvium is desirable.GEI Consultants - 68 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/0510. RECOMMENDATIONSDespite the “high risk” rating of Lafayette Dam and its seismic setting, we have found nocondition that requires immediate action (Priority Level 1). However, based on our reviewand the limited analyses performed, we recommend that the District consider deferredimplementation of some action items to allow assessment of the seismic performance ofLafayette Dam with greater reliability. Such lesser priority action items have been assignedeither a moderate (Priority Level 2), or a low level of priority (Priority Level 3). Actionitems assigned a Priority Level 3 can be deferred until a related change of condition isobserved, suspected or concluded likely in future inspections or safety reviews.10.1 Recommendations for Optional Geologic InvestigationsLineaments: The series of lineaments observed in the vicinity of the dam could be bettercharacterized to determine their most likely cause of formation and, if found to be faultrelated,to help assess the associated rate of activity. However, because these lineamentsappear to have negligible potential impact on the dam if they were confirmed, such actionitem was not assigned any priority level. Should future changes in the understanding of thelocal geologic and tectonic setting occur that would justify a need to better identify suchfeatures, the discontinuous lineaments closest to the dam could be investigated by means ofadditional geologic fieldwork, including detailed surface mapping. If suitable trenching siteswere identified, subsurface exploration could be performed across discrete lineaments to helpdetermine the reason(s) of their existence.Landslides: Many of the large landslides along the reservoir rim are in a low slope positionand do not appear active. However, to a limited extent, there is always an unknown potentialfor reactivation during periods of elevated ground water or in response to strong groundshaking. The District should consider the long-term benefits of investigating the landslidesclosest to the dam (Priority Level 3), in order to assess any need to develop some mitigationalternatives to reduce any damage potential that could be associated with continued orrenewed movements. Of course, a higher priority level would be re-assigned to such actionitem, if any signs of landslide reactivation were observed during future inspections. Shouldsigns of reactivation become apparent, the old landslide adjacent to the east margin of thedam should be immediately evaluated through additional geologic fieldwork for potentialimpact on the lower portion of the embankment. A complete evaluation would includepreparation of geologic maps and cross sections to establish the probable volume anddimensions of this landslide, and assess whether it could impact a larger portion of the damin case of enhanced reactivation during earthquake loading.GEI Consultants - 69 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/0510.2 Recommendations for Optional Field and LaboratoryInvestigationsStrength Parameters for Foundation Alluvium: Existing field penetration data andlaboratory test results in the alluvium below the downstream slope of Lafayette Dam indicatethat these materials are weaker. As was the case in 1928, such alluvium materials maycontrol the overall seismic stability of Lafayette Dam. As the information in that area islimited, further field and laboratory testing in that area would be desirable (Priority Level 2).Field exploration methods that would facilitate recognition of any thin and potentially weakerzones, including the old failure surface, should be considered. Cone penetration testing(CPT) could be a rapidly implemented and cost-effective way to perform such investigations.CPT would be supplemented with SPTs for correlation purposes. Completion of in-situ vaneshear tests at various depths and a number of locations would also be essential. The primarypurpose of the vane shear tests would be to define the in-situ residual strength of the clayeymaterials, which cannot be established from correcting existing blowcount data because ofthe highly clayey nature of the materials encountered. The shear tests should be conducted ina manner consistent with measuring both the peak strength and ultimate (remolded orresidual) strength. Thus, it would be important to conduct these tests at depths that contain nogravel. As new holes would to be drilled to conduct the shear tests, it would be beneficial toinstall some multi-stage piezometers, which would provide more reliable information thanthe existing open-standpipe piezometers.10.3 Recommendations Regarding Stability AssessmentDetailed Seismic Analysis: Lafayette Dam is a well-maintained facility, and appears to beseismically safe based on simplified analyses. However, because of several factors including(1) the dam failure during construction, (2) insufficiently complete existing informationregarding the downstream foundation alluvium, and (3) the lack of a sufficiently rigorous andreliable seismic evaluation consistent with current standards, we recommend that EBMUDconsider, if found desirable after the recommended field and laboratory investigations, adetailed reanalysis, that would use confirmed alluvium properties, modern computationaltechniques, updated acceleration time histories compliant with the recommended responsespectra, and non-linear material properties and constitutive relationships that would simulatethe probable behavior of the embankment and foundation materials more closely than waspossible in 1966 or 1976 (Priority Level 3). The existing static dynamic laboratory testingdata already available could be used to define dynamic properties and improved constitutivemodels, and to characterize the strength degradation characteristics of the dam andfoundation materials under cyclic loading. The new CPT and vane-shear tests would be veryuseful to better define or confirm the parameters required for such analyses. Suchinformation would be used, as needed, as input to a more detailed, and more reliable methodof analysis than those used to date.GEI Consultants - 70 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/0511. REFERENCESAndrews, D. C.A.; Martin, G.R. (2000), “Criteria for Liquefaction of Silty Soils”,Proceedings, 0312, 12 th World Conference on Earthquake Engineering.Borchardt, G. and Baldwin, J. N. (2001), “Late Holocene Behavior and SeismogenicPotential of the Concord-Green Valley Fault System in Contra Costa and Solano Counties,California”, in Ferriz, H. and Anderson, R. (eds.), “Engineering Geology Practice inNorthern California”, Association of Engineering Geologists Special Publication 12 andCalifornia Division of Mines and Geology Bulletin 210, pp. 229-238.Bureau, G. (1996), “Numerical Analysis and Seismic Safety Evaluation of EmbankmentDams”, ASCE, BSCES Geotechnical Group, 1996 Lecture Series, “Dam Inspection,Analysis and Rehabilitation”, November 2, Bentley College, Waltham, MA, Proceedings, 28pp. plus Figures.Bureau, G.; Volpe, R.L.; Roth, W.R.; Udaka, T. (1985), "Seismic Analysis of Concrete FaceRockfill Dams", ASCE Int. Symp. on CFRD's, Detroit, Oct. 21, in "Concrete Face RockfillDams - Design, Construction and Performance", pp. 479-508, and Closure (1987), ASCEJourn. of the Geotechnical Eng. Div., Vol. 113, No. 10, October, pp. 1255-1264.Consulting Board (1929), “Report on Partial Failure During Construction of Lafayette Dam”,East Bay Municipal Utility District, January 12.Crane, R. (1988), Geologic Maps of the Las Trampas Ridge and Walnut Creek”, 7.5-minutequadrangles, 1:48,000 scale.Crane, R. (1995), “Geology of the Mt. Diablo Region and East Bay Hills”, in Sangines, E.M., Andersen, D. W., and Buising, A. V., (eds.), “Recent Geologic Studies in the SanFrancisco Bay Area”, Society of Economic Paleontologists and Mineralogists, PacificSection Vol. 76, pp. 87-114.Dibblee, T. W., Jr. (1980”, “Preliminary Geologic Maps of the Briones Valley, Hayward, LasTrampas Ridge, Niles, and Walnut Creek”, 7.5-minute quadrangles: United States GeologicalSurvey Open-File Reports, 1:24,000 scale.DSOD (2003), “Geologic Review of Seismic and Foundation Conditions – Lafayette Dam,No. 31-2, Contra Costa County”, Internal Memorandum by James L. Lessman, dated March14, 2003.GEI Consultants - 71 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05DSOD (1980), “Phase I Inspection Report for Lafayette Dam”, Report to Department of theArmy, Corps of Engineers, Sacramento District, July.Dukleth, G.W. (1956), “Lafayette Dam No. 31-2, Stability Study”, Internal Memorandum toMr. W.A. Brown, DSOD, March 8.East Bay Municipal Utility District (2002), “Lafayette Reservoir Tower Seismic Upgrade –Lafayette Reservoir No. 31-02”, letter to DSOD, with attachments, dated August 12.East Bay Municipal Utility District (1956-1957), “Lafayette Dam – Foundation Investigationand Stability Analysis (1956)”, Internal Report, Foundation Design Section, November 1957.East Bay Municipal Utility District (1929), “Revised Cross-Section Recommended byConsulting Board”, EBMUD Drawing DH 1609-18, January.Graymer, R. W., Jones, D. L., and Brabb, E. E. (1994), “Preliminary Geologic MapEmphasizing Bedrock Formations in Contra Costa County, California – A Digital Database”,United States Geological Survey Open-File Report 94-622, 1:75,000 scale.Hart, E. W. (1981), “Evidence for Recent Faulting, Calaveras and Pleasanton Faults, Diabloand Dublin Quadrangles, California”, California Division of Mines and Geology Open FileReport 81-09-SF.Haydon, W. D. (1995), “Landslide Hazards in the Martinez-Orinda-Walnut Creek Area.Contra Costa County, California”, California Division of Mines and Geology (CDMG),Open-File Report 95-12, Landslide Hazard Identification Map No. 32, 1:24,000 scale.Idriss, I.M. (2004), “Dynamic Stability Review of Lafayette Dam”, letter-report to GEI, Draft02, dated August 31, 2004, 9 pp.Interactive Software Designs, Inc. [ISD] (1992-2002), “XSTABL, Slope Stability AnalysisUsing the Method of Slices”, Sharma, S., Version 5-206, Moscow, ID 83843.International Civil Engineering Consultants (1995), “Seismic Response Analysis andPerformance Evaluation of the Inlet/Outlet Tower of Lafayette Reservoir”, Report preparedfor East Bay Municipal Utility District.Janbu, N. (1954), “Stability Analysis of Slopes with Dimensionless Parameters”, HarvardSoil Mechanics Series, No. 46, 81 pp.GEI Consultants - 72 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Jansen, R.B. (1987), "The Concrete Face Rockfill Dam. Performance of Cogoti Dam underSeismic Loading", discussion of a paper presented at ASCE's Symposium on Concrete FaceRockfill Dams, ASCE Journal of the Geotech. Engineering Div., Vol. 113, No. 10, October.Kelson, K. I. (2001), “Geologic Characterization of the Calaveras as a Potential SeismicSource in San Francisco Bay Area”, California, in Ferriz, H. and Anderson, R. (eds.),“Engineering Geology Practice in Northern California”, Association of EngineeringGeologists Special Publication 12 and California Division of Mines and Geology Bulletin210, pp. 179-192.Langenkamp, D.L.; Nelson, J.S. (1973), “Seismic Survey: P and S Wave Velocities,Lafayette Dam, Lafayette, CA”, report to W.A. Wahler & Associates, Job NO. 5949,002.01,December.Lazarte, C. A.; Bray, J.D.; Johnson, A.M. and Lemmer, R.E. (1994), “Surface Breakage ofthe 1992 Landers Earthquake and its Effects on Structures”, Bulletin of the SeismologicalSociety of America, Volume 84, No. 3, pp. 547-561, June.Lettis, W. R. (2001), “Late Holocene Behavior and Seismogenic Potential of the Hayward-Rodgers Creek Fault System in the San Francisco Bay Area, California”, in Ferriz, H. andAnderson, R. (eds.), “Engineering Geology Practice in Northern California”, Associationof Engineering Geologists Special Publication 12 and California Division of Mines andGeology Bulletin 210, pp. 167-178.Lienkaemper, J. J. (1992), “Map of Recently Active Traces of the Hayward Fault, Alamedaand Contra Costa Counties, California”, United States Geological Survey Map MF-2196,1:24,000 scale.Louderback, G.D. (1927), “Geological Report on the Site of the Proposed LafayetteReservoir”.Makdisi, F.; Seed, H.B. (1977), "A Simplified Procedure for Estimating Earthquake-InducedDeformations in Dams and Embankments" U. of California, Berkeley, EERC Report No.UCB/EERC-77/19, 33 pp, plus Appendices.Makdisi, F.I., and Seed, H.B., (1978). "Simplified Procedure for Estimating Dam andEmbankment Earthquake Induced Deformations". Journal of Geotechnical Engineering,ASCE, July.Makdisi, F.I., and Seed, H.B., (1979). "Simplified Procedure for Evaluating EarthquakeResponse", Journal of Geotechnical Engineering, ASCE, December.GEI Consultants - 73 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Newmark, N.M. (1965), "Effects of Earthquakes on Dams and Embankments", RankineLecture, Geotechnique 15, No. 2, pp. 139- 160.Olivia Chen Consultants, Inc. (2003), “Report on the Seismic Stability of Calaveras Dam”,Report to San Francisco Public Utilities Commission, Utilities Engineering Bureau, January.Oppenheimer, D. H., and Macgregor-Scott, N. (1992), “The Seismotectonics of the EasternSan Francisco Bay Region, in Borchardt, G., Hirschfeld, S. E., Lienkaemper, J. J.,McClellan, P., Williams, P. L., and Wong, I. G. (eds.), Proceedings of the 2 nd Conference onEarthquake Hazards in the Eastern San Francisco Bay Area, California Division of Minesand Geology Special Publication 113, pp. 11-16.Petersen, M.D.; Toppozada, T.R.; Cao, T.; and Cramer, C.H. (2000), “Active Fault Near-Source Zones Within and Bordering the State of California for the 1997 Uniform BuildingCode”, EERI, Earthquake Spectra, Volume 16, No. 1, February, pp. 69-83.Radbruch, D. H. (1969), “Areal and Engineering Geology of the Oakland East (7.5-minute)Quadrangle, California”, United States Geological Survey Miscellaneous GeologicQuadrangle Map GQ-769, 1:24,000 scale.Sarma, S.K. (1975), "Seismic Stability of Earth Dams and Embankments", Geotechnique 25,No. 4, pp. 743-761.Saul, R. B. (1973), “Geology and Slope Stability of the SW 1/4 Walnut Creek Quadrangle,Contra Costa County, California”, California Division of Mines and Geology Map Sheet 16,1:12,000 scale.Seed, H.B.; Idriss, I.M. (1970), "Soil Moduli and Damping Factors for Dynamic ResponseAnalysis", University of California, Berkeley, Report No. EERC/70-10, December, 15 pp.Seed, H.B., and Idriss, I.M. (1982) "Ground Motions and Soil Liquefaction DuringEarthquakes", Earthquake Engineering Research Institute, Monograph Series.Seed, R.B.; et al. (2003), “Recent Advances in Soil Liquefaction Engineering: A Unified andConsistent Framework”, 26 th ASCE Geotechnical Spring Seminar, Keynote Presentation,Long Beach, CA, pp. 1-71.Seed, R.B.; Harder, L.F. Jr. (1990), "SPT-based Analysis of Cyclic Pore Pressure Generationand Undrained Residual Strength", Presented at H. Bolton Seed Memorial Symposium,Proceedings, Volume 2, BiTech Publishers, Canada, May.GEI Consultants - 74 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Shannon & Wilson, Inc. (1966), “Review of Stability Lafayette Dam”, Report to East BayMunicipal Utility District, Oakland, CA, January, with amendment letters, February 15, 1966and February 21, 1966.Simpson, G. D., Baldwin, J. N., Kelson, K. I., and Lettis, W. R. (1999), “Late Holocene SlipRate and Earthquake History for the Northern Calaveras Fault at Welch Creek, Eastern SanFrancisco Bay Area, California”, Bulletin of the Seismological Society of America, vol. 89,no. 5, pp. 1250-1263.Sims, J. D. (1991), “Distribution and Rate of Slip Across the San Andreas TransformBoundary, Hollister Area, Central California”, Geological Society of America Abstracts withPrograms, Cordilleran section, vol. 23, no. 2, p. 98.Somerville, P.G.; Smith, N.F.; Graves, R.W.; Abrahamson, N.A. (1995), "Accounting forNear-Fault Rupture Directivity Effects in the Development of Design Ground Motions",Proceedings, ASME/SSME Conference, Hawaii, July.Spencer, E. (1967), “A Method of the Stability Analysis of Embankments Assuming ParallelInter-Slices Forces”, in Géotechnique, XII, No. 1, pp. 11-26.Swaisgood, J.R. (1995), "Estimating Deformation of Embankment Dams Caused byEarthquakes", ASDSO Western Regional Conference, Red Lodge, Montana, May 22-25.Swaisgood, J.R. (1998), “Seismically-Induced Deformation of Embankment Dams”, 6 th U.S.National Conference on Earthquake Engineering, Seattle, Washington, June 1998.Tokimatsu, K.; Yoshimi, Y. (1983), "Empirical Correlation of Soil Liquefaction Based onSPT N-Value and Fines Content", Soils and Foundations, Vol. 23, No. 4, December,Japanese Society of Soil Mechanics and Foundation Engineering, pp. 56-74.Toppozada, T. R., Real, C. R., and Parke, D. L. (1986), “Earthquake History of California”,California Geology, vol. 39, no. 2, pp. 27-33.Toppozada, T. R. and Borchardt, G. (1998), “Re-evaluation of the 1836 “Hayward Fault”Earthquake and the 1838 San Andreas Fault Earthquake”, Bulletin of the SeismologicalSociety of America, vol. 88, pp. 140-159.Unruh, J. R., and Kelson, K. I. (2002), “Critical Evaluation of the Northern Termination ofthe Calaveras Fault, Eastern San Francisco Bay Area, California”, Final Technical Report tothe United States Geological Survey, Award no. 1434-HQ-97-GR-03146, 72 p. with figures.GEI Consultants - 75 -

Dynamic Stability Review of Lafayette DamEast Bay Municipal Utility District08/16/05Vrymoed, J.L. (1996), "Seismic Safety Evaluation of Two Earth Dams", in "EarthquakeEngineering For Dams", Western Regional Technical Seminar, Ass. of State Dam SafetyOfficials, April 11- 12, Sacramento, pp. 215-234.W.A. Wahler & Associates (1976), “Seismic Stability Evaluation Lafayette Dam ContraCosta County, California”, Report to East Bay Municipal Utility District, Oakland, Ca, May.Wakabayashi, J. and Sawyer, T. L. (1998), “Paleoseismic Investigation of the Miller CreekFault, Eastern San Francisco Bay Area, California”, Final Technical Report to the UnitedStates Geological Survey, award no. 1434-HQ-97-GR-03141, 17 pp. with figures.Wagner, J. R. (1978), “Late Cenozoic History of the Coast Ranges East of San FranciscoBay”, Map 1:24,000 scale, University of California, Berkeley, Ph. D. Thesis, 161 pp.Wells, D. L. and Coppersmith, K. J. (1994), “New Empirical Relationships AmongMagnitude, Rupture Length, Rupture Width, Rupture Area and Surface Displacement”,Bulletin of the Seismological Society of America, vol. 84, pp. 974-1002.Woodward-Clyde Consultants (1975), “Lafayette Dam Seismic Evaluation – Cross-HoleShear Wave Velocity Survey”, prepared for East Bay Municipal Utility District, WCCProject G-13377, March.Working Group on California Earthquake Probabilities (1999), “Earthquake Probabilities inthe San Francisco Bay Region: 2000 to 2030 – A Summary of Findings”, United StatesGeological Survey Open-File Report 99-517.Zhou, S.G. (1981), “Influence of Fines on Evaluating Liquefaction of Sand by CPT”,Proceedings, International Conference on Recent Advances in Geotechnical EarthquakeEngineering and Soil Dynamics, Vol. 2, pp. 167-172GEI Consultants - 76 -

Dynamic Stability Reviewof Lafayette DamAPPENDICESAPPENDIX ASURVEY MONUMENTS REVIEWAPPENDIX BPIEZOMETRIC DATA REVIEWAPPENDIX CSELECTED PROJECT PHOTOGRAPHSAPPENDIX DDEVELOPMENT OF SITE-DEPENDENT SPECTRAAPPENDIX EREVIEW OF LABORATORY DATAAPPENDIX FCALCULATION OF EARTHQUAKE-INDUCED DEFORMATIONS- i -GEI Consultants Project # 04035-0

APPENDIX ASURVEY MONUMENTS REVIEWA-1 Survey Monuments LocationThere are currently 24 survey monuments, arranged in a grid on the embankment of LafayetteDam. The location of the monuments is shown on Figure A-1. Monuments BL-6 through BL-14are located on the centerline of the crest, which is considered the baseline. The numberrepresents the closest station, e.g., BL-6 is located at approximate Station 6+03, BL-8 is locatedat 8+03, etc. Monuments DIA-10 to DIA-14 are located on a diagonal line traversing theupstream face. Monuments 116-S-10 to 116-S-14 are located on the upstream berm, 116 ftupstream (south) of the baseline (e.g. 116-S-10 is located 116 ft south of the baseline at station10+03). Monuments 85-N-10 to 85-N-14 are located on the downstream berm, 85 ft downstream(north) of the baseline. Monuments 179-N-10 to 179-N-14 are located on the downstream berm,179 ft downstream (north) of the baseline. Monuments 396-N-10 to 396-N-14 are located on thedownstream berm, 396 ft downstream (north) of the baseline. Monuments 582-N-10 to 582-N-14are located on the downstream berm, 582 ft downstream (north) of the baseline.A-2 Recent ReadingsThe monuments are surveyed approximately once (and occasionally twice) a year. EBMUDprovided survey data measurements from June 1, 1989 to December 9, 2003 for our review.These data were provided both as graphs and raw data. The graphs are shown in the pages thatfollow Figure A-1.Table A-1 summarizes the maximum vertical and horizontal displacements recorded for themonuments from 1989 to the present. Figures A-2 and A-3 are graphs of the maximumhorizontal and vertical displacements at the survey monuments plotted along stations of the dam.It can be seen from Table A-1 that the maximum horizontal and vertical displacements for all thesurvey monuments are 3.36 inch and 3.12 inch, respectively. Review of the time versusdisplacement graphs for individual survey monuments shows that the horizontal and verticalmovements are stable, with no significant increase in the last 15 years.Horizontal movements of the upstream slope are toward the upstream direction, while thedownstream slope has moved slightly downstream. While these measured displacements arevery small, it is rather unusual that the slopes of a dam move in opposite directions, especiallyconsidering that the reservoir load is rather constant in the case of Lafayette Dam. This could beindicative of extremely slow creep taking place or continuing in the foundation alluvium.Similar dual directions of movement occurred during the 1928 failure. The dam displacementsare too small, however, to be of concern.The plots of survey monuments provided by the District also include precipitations. Damdisplacements should not be affected by precipitation, unless the monuments are too shallow.This does not seem to be the case, as precipitation seems to have no effects on the measured dammovements.A-1GEI Consultants Project # 04035-0

APPENDIX BPIEZOMETRIC DATA REVIEWB-1 Phreatic Level Monitoring BackgroundPrior to 1965, the water level within the Lafayette embankment was monitored with observationwells. The exact locations and date of installation of these wells are not known. Review ofrecords on Lafayette Dam at the DSOD indicates that the wells were installed in about 1932 tomonitor seepage of natural groundwater into the left abutment and under the downstream shell ofthe dam. The following excerpt is from a 1932 review by DSOD (Engle, 1932):“Seepage- Mr. Marliave has made a geological investigation as of January, 1932. I havediscussed with him the proposed repairs now before us for action. From a geologicalstandpoint he questions the formation of the left or west abutment, believing that naturalground water may find a passage from the abutment into and under the downstreamportion of the dam to create undesirable condition of saturation. He suggests that theowners put down test wells at the left end of dam to determine water condition of theabutment and suggest the advisability, if conditions warrant, of providing drainage forthe abutment- perhaps a drift run in from about the elevation of the downstream toe.”While no records of well installation were found in the DSOD or District’s files, it appears thatthese wells were actually installed. In 1944, another DSOD (Perkins, 1994) memorandumdiscussed the advisability of a foundation drainage program involving continuous pumping fromtwo of the wells that penetrate the foundation upstream from the cutoff.In 1956, a field investigation was conducted by EBMUD as part of a stability analysis of theupstream slope for a proposed spillway (never constructed) through the left abutment. Twoborings, SS-1A and SS-2, were drilled for this investigation. No record of piezometer installationhas been found, so it is assumed that piezometers were not installed at this time in these twoborings.In 1965, Shannon and Wilson and Burton Marliave performed field investigations and seismicstability analyses using a horizontal load coefficient. During that investigation, 12 borings weredrilled, SS-3 to SS-14. DSOD records indicate that 13 piezometers were installed in June andJuly 1965 in the core and alluvium foundation to monitor the dam. These piezometers are SS-1A(3 tip elevations), 3, 4, 5, 6 (2 tip elevations), 7, 8 (3 tip elevations), 9, 10, 11, 12, 13 and 14.After the piezometers were installed, the observation wells were abandoned.In 1973-1974, EBMUD conducted a new field investigation. The field investigation includeddrilling 17 rotary borings, SS-14 to SS-29. In 1975–1976, W.A. Wahler & Associates interpretedthe field results, performed laboratory tests and conducted another seismic evaluation usingdynamic equivalent-linear response analysis. Piezometers were not installed in the boringsdrilled for this investigation.Records obtained from DSOD files indicate that, in 1992, seven of the eight crest piezometerswere reading water surface elevations higher than the reservoir elevation. It was suggested thatB-1GEI Consultants Project # 04035-0

the malfunction could have been due to rain infiltration. This could also represent capillary rise.It was proposed to abandon the left abutment piezometer (SS-14) and replace the remaining sixpiezometers. Piezometer SS-14 was reported to be malfunctioning, but was not replaced becauseof its non-critical water level data. Piezometers SS-10 and SS-11 were also reported to bemalfunctioning, but were not replaced because of their non-critical water level data andenvironmental constraints. It was also noted that piezometers SS-5 to SS-9 in the downstreamslope area had shown fluctuations that could indicate that they were not working properly.Further evaluation of these piezometers was suggested.EBMUD records indicate that in 1992, piezometers SS-30A, SS-30B and SS-30C were installedto replace SS-1-AA, SS-1-AB, SS-1-AC. Drilling logs for these holes do not appear to beavailable. Whereas SS-1-AA, SS-1-AB and SS-1-AC had been installed in one hole, thereplacement piezometers were installed in three separate boreholes. Also in August 1992,piezometer SS-31A was installed to replace SS-3, and piezometer SS-32A was installed as areplacement for SS-13.In 1996, EBMUD records indicate that piezometer SS-33A was installed. It was not indicatedwhether this was a replacement piezometer, but due to its proximity to the location of SS-4 andthe fact that SS-4 is no longer monitored, we presume that it was installed as a replacement forSS-4. The drill log for this piezometer hole is not available.Table B-1 presents a summary of the 18 active piezometers at Lafayette Dam including thelocation and sensing depth of the instruments, as shown in EBMUD records. All piezometers areopen-standpipe piezometers.B-2 Time vs. Reading Graphs of DataFigure B-1 indicates the approximate location of the 18 active piezometers. This figure wasreprinted from the 1966 Shannon & Wilson report and includes the locations of severalpiezometers that are no longer active (SS-1A, SS-3, SS-4, SS-13 and SS-14). Time-versusreadinggraphs for the active piezometers from January 1989 to January 2004 are attached.Piezometers readings are taken approximately monthly. In addition, the corresponding reading ofreservoir level from 1989 to 2004 is included in the attached graphs as well as rainfallmeasurements.Table B-2 summaries the piezometric data from June 1998 (after several piezometers werecleaned) to January 2004. The maximum historic high and low water levels for this time periodfor each piezometer are indicated in the table along with the reservoir level reading for thecorresponding date. In addition, the most recent water level reading is also shown for eachpiezometer.B-3 EvaluationReview and evaluation of the piezometer data reveals that Piezometers SS-5A, SS-6A, SS-7A,SS-9A show seasonal fluctuation beyond what would be expected from response to the increaseB-2GEI Consultants Project # 04035-0

in reservoir level. This may be due to infiltration of rainwater into the piezometers, since theincrease in water level reading appears to correlate to periods of higher rainfall.Piezometers SS-8A, SS-8B and SS-8C show only small variation due to change in reservoirlevel. Within the last ten years, the piezometers show no upward trend and are within historicrange, with the exception of one reading in May 1998. Closer inspection of this data point seemsto indicate that the reading is in error. Subsequent readings have returned to expected levels.Piezometers SS-10A and SS-11A on the upstream face appear to be reading at or slightly belowthe reservoir level. Due to their location, these piezometers are submerged during high reservoirstage and are of limited value.Piezometer SS-12, located on the crest near the right abutment, continues to show water levelsabout 15 feet lower than the reservoir level and appears to fluctuate with changes in reservoirlevel. Within the last ten years, piezometer readings show no upward trend and are withinhistoric range, with the exception of one reading in February 1998. Closer inspection of this datapoint seems to indicate that this reading is in error, probably due to infiltration of surface watersince this occurred during a period of high rainfall. Subsequent readings have returned toexpected levels.Piezometers SS-30A, SS-30B and SS-30C are located on the crest towards the upstream slope.Since installation in 1992, Piezometer SS-30A (tip depth 34.0 ft) has shown water levels about15 feet above the reservoir level. This may be due to infiltration of surface water, influence ofground water or plugging of the piezometer. EBMUD records indicate that in May 1998 thispiezometer was cleaned, but water level readings continued to be high. Piezometer SS-30B (tipdepth 137.0 ft) currently indicates water levels about 15 ft below the reservoir level and SS-30C(tip depth 205.0 ft in the foundation bedrock) has water levels about 36 ft below the reservoirlevel. In 1998, SS-30B and SS-30C showed erratic readings for several months. EBMUD recordsindicate that in May 1998 these piezometers were cleaned and since that time they appear to beoperating properly. Readings in SS-30B and SS-30C show no upward trend in readings andappear to fluctuate with changes in reservoir level.Piezometers SS-31, SS-32 and SS-33 are located on the crest towards the downstream slope. SS-31 continues to have readings about 10 ft higher than the reservoir level. In May 1998, thispiezometer was cleaned and pumped out. Since that time, the water level has slowly risen andnow appears to be stable at its current reading. It can be concluded that this piezometer isplugged, and of little value. Piezometer SS-32 (tip at depth 102.0 in the foundation bedrock)shows water level readings approximately 34 ft below the reservoir level. Within the last tenyears, SS-32 readings show no upward trend and are within historic range with the exception ofone reading in 1997. EBMUD records indicate that, in May 1998, this piezometer was cleanedand appears to be operating properly since that time. The water level readings in piezometer SS-33, installed in January 1996, have risen slowly since installation to a level about 10 ft above thereservoir level. The water level in this piezometer appears to be stable at its current reading. Theunusual readings may be due to the influence of ground water from the right abutment, orperhaps this piezometer is plugged.B-3GEI Consultants Project # 04035-0

B-4 Phreatic Surface for Slope Stability AnalysesFollowing review of the piezometric data discussed above, we developed an estimate of thephreatic level through the maximum cross-section of the embankment. Figure B-2 shows theestimated phreatic level. The phreatic level is based on the high water level readings in thepiezometers. We used extrapolation and considerable judgment to estimate such phreatic level.The exact locations of the piezometers are not known, since the only available drawing (FigureB-1) showing current piezometer locations is only schematic. Also data from several piezometers(SS-30A, SS-31 and SS-33) are of questionable value, as discussed above.The piezometric readings show little influence of the dam “zoning” upon the position of thephreatic surface in the embankment. Two interpretations of phreatic levels are possible. Theupper level takes into consideration the readings in piezometer SS-6A, which are considerablyhigher than would be expected. While this piezometer may be subject to surface waterinfiltration, in which case the phreatic level at this location would be less than that indicated,there is no tangible evidence to discount this piezometer. The lower level, which does discountthe readings in SS-6A, is shown for illustrative purposes. We believe the upper level bestrepresents the actual maximum phreatic level through the embankment. It can also be seen fromFigure B-2 that the piezometers that monitor the pore pressures in the foundation show that theclayey foundation does not provide additional drainage, since the water level in thosepiezometers are consistent with those measured in the embankment.Comparison of the estimated phreatic surface with the phreatic surface assumed in the Wahler(1975) stability analyses and the DSOD (2003) review is shown in Figures B-3 and B-4,respectively. Our estimated phreatic surface is generally similar to those used in previous studies.Our interpreted phreatic surface is somewhat lower in the upstream portion of the dam, butslightly higher in the downstream portion of the dam. Overall, such variation is probably notsignificant, and well within the level of judgment needed to interpret the piezometric data.REFERENCESAPPENDIX BEngle, G.F. (1932), “Memorandum to Mr. Hawley – Lafayette Dam #31-2”. Comments onOwner’s Application for Repair of Dam, DSOD Internal Memorandum, Dated July 8, 1932.Perkins, W.A. (1944), “Memorandum to Mr. Holmes – Lafayette Dam #31-2”. Conference July20, 1944.B-4GEI Consultants Project # 04035-0

APPENDIX CSELECTED PROJECT PHOTOGRAPHSThe following photographs, taken during our field inspection of 4/01/04, describe the currentcondition of Lafayette Dam:Photo C-1: Upstream face. This is a view of the upstream face of Lafayette Dam, taken fromnear the park headquarters. Note the apparent settlement in the central portion of the concretelinedupstream slope, which occurred in the years that followed the 1928 failure.Photo C-2: Upstream face. This view is taken from the near the right abutment, and also showsthe upstream shell settlement and resulting ponding near the center of the top upstream berm.Photo C3: Typical cracking, upstream concrete slab panels. Minor random cracking wasobserved in some of the concrete slab facing of the upstream slope. Such cracking does notrepresent a safety concern, as settlements of the upstream shell have been negligible for severaldecades. Especially, the concrete slabs are used as slope protection, not as an impervious barrier.Photo C4: Typical gap between upstream concrete slab panels. Gaps, up to 4-inch wide, arepresent between several concrete panels, and probably result from the aforementioned settlementof the upstream shell. Although these open joints have been filled with asphalt, such filling isfrequently missing, deteriorated, or affected by the growth of weeds.Photo C-5: Upstream face. This view is taken from the top upstream berm, near the leftabutment. It shows the weeds growing in the joints between adjacent concrete slabs. This is anon-critical maintenance issue.Photo C-6: Downstream face. The downstream face is covered with grass, and appears to be ingood condition. No traces of seepage, slope movements, settlement, cracking or deteriorationare visible. Numerous small rodent holes were noticed, however.Photo C-7: Downstream face, typical surface drain. Clay-tiled surface drains collectrainstorm runoff along the downstream face. A piezometer cap is visible, left of the drainchannel. Note the excellent apparent condition of the downstream slope toward the leftabutment.C-1GEI Consultants Project # 04035-0

Photo C-8: Surface drain collector, bottom of downstream face. Note the good apparentcondition of the downstream slope toward the right abutment.Photo C-9: Exit box, downstream drainage collection system. Negligible (less than 1gpm)and clear seepage was observed on the day of our visit.Photo C-10: Typical piezometer cap. This picture shows a typical, well-maintained piezometerlocation (Piezometer SS-7).Photo C-11: Lafayette Dam outlet and spillway tower. The outside of the outlet tower, whichis inside the reservoir near its right bank, appears in good visual condition as seen from the dam.--o-oo-o--C-2GEI Consultants Project # 04035-0

APPENDIX DDEVELOPMENT OF SITE-DEPENDENT RESPONSE SPECTRAD-1 IntroductionThe characteristics of ground motion at a given distance from an earthquake of specifiedmagnitude is required for the seismic evaluation of existing dams. Attenuation equations(empirical relationships that relate the characteristics of ground motion to magnitude anddistance) are the most reliable at intermediate magnitudes (M w 5.5 to 7.0) and intermediatedistances (10 to 50 km), which represent the parameters for which the largest amount of strongmotion data is available. Hence, in the case of Lafayette Dam, estimates obtained fromapplicable attenuation equations are influenced by the large magnitudes (M w >= 7.0) and shortdistances (

seismological models of the fault rupture process. In California, the statistical method is themost frequently used, due to the existence of an appreciable set of strong motion records.Many attenuation equations started to be developed after the 1971 San Fernando earthquake, thefirst event to generate a large database of recorded ground motions. They have beencontinuously updated since, and every time that more records become available. RecentCalifornia events, such as the Loma Prieta (October 18, 1989), Landers (June 28, 1992) andNorthridge (January 17, 1994) earthquakes have significantly increased the existing database fornear-field records.Attenuation equations are based on different definitions of the applicable distance parameter,which is particularly critical in the near-field. The augmentation of the strong ground motiondatabase has improved the reliability of horizontal attenuation equations, allowing considerationof different types of faulting (strike- slip, reverse, and normal or extensional) and varioussubsurface conditions, typically simply differentiated as “hard rock”, “soft rock”, “firm soil”,“deep soil” or “soft soil”. Some site classification schemes rely on the average shear wavevelocity (V s ) in the upper 30 meters. As the Orinda Formation is intermediate between a softrock and a hard soil, the use of V s is more rigorous than a simple distinction between soil androck. Lastly, several sets of equations include vertical motion. Although simplified andequivalent-linear solutions typically ignore vertical motion, such component can be significantfor near-field sites and when more sophisticated analysis methods are implemented. It has beenincluded in our development of response spectra.Appreciable errors can be associated with strong motion predictions, when compared to past orfuture recordings, and attenuation equations typically consider mean (50 th percentile) or meanplus-one-standard-deviation(84 th percentile) estimates. Because Lafayette Dam is a “high-risk”dam, located in the vicinity of faults with a “high to very high slip rate”, we have used 84 thpercentile seismic criteria.For this study, we tested five sets of well-accepted attenuation equations for peak groundacceleration (PGA) and pseudo-absolute spectral accelerations (PSA) at 5 percent damping.Crouse and McGuire (1995) developed the first set. The others (Abrahamson & Silva; Boore, etal.; Sadigh, et al. and Campbell) were presented in the January/February 1997 Issue of the“Seismological Research Letters” of the Seismological Society of America (SSA). The Crouse-McGuire’s equations use the shear wave magnitude (M s ) to quantify the size of the earthquake,and the others use the moment magnitude (M w ). In our review of the Lafayette Project files, wefound out that the DSOD uses the Abrahamson & Silva, Boore, et al., and Sadigh, et al.’sequations for horizontal motion. Therefore, we also used these three sets of equations as theprimary basis for our horizontal ground motion estimates. We used the Crouse-McGuire’s andCampbell’s equations, however, for comparative purposes and to confirm the suitability of thepredictions obtained with the three other equations. All of these equations are applicable toCalifornia events (shallow crustal earthquakes), and are further discussed in the following pages.D-2GEI Consultants Project # 04035-0

• Crouse and McGuire (1995). Crouse and McGuire developed attenuation equationsapplicable to the Western U. S. to revise NEHRP Seismic Provisions. These equationsdepend on a V s -dependent classification of the site (Site Classes A through E). Based onmeasured shear wave velocities in the Orinda Formation, the Lafayette site falls in thelower-bound of the V s range for Soil Class C. Hence, for ground motion estimationpurposes, we averaged predictions obtained for soil Class C and Class D.• Abrahamson and Silva (1997). Abrahamson and Silva used a database of 655recordings from 58 earthquakes to develop empirical equations for peak and spectralresponse (horizontal and vertical). They distinguished between two types of sites, “rockand shallow soil”, and “deep soil”). Their equations include a correction factor todifferentiate between strike-slip (used herein) or reverse faulting, and are based on theclosest distance to the fault surface rupture (R up ). They also introduced a correctionfactor to account for differences in the ground motion on the “hanging wall” and “footwall” of dipping faults. After reviewing the site conditions present at the Lafayette site,we used the arithmetic average of Abrahamson and Silva’s “deep soil” and “rock andshallow soil” equations to develop MCE response spectra for Lafayette Dam.Abrahamson and Silva provide equations for both the horizontal and the verticalcomponents of ground motion.• Boore, Joyner, and Fumal (1997). In 1997, Boore, Joyner and Fumal upgraded Joynerand Boore’s 1982 equations. They added records from the 1989 Loma Prieta, 1992Petrolia and 1992 Landers records. They used as a distance parameter the closesthorizontal distance from the recording station to a point on the earth surface that liesdirectly above the fault rupture (R jb ). Their regression analysis uses shear wave velocity(V s ), averaged over the upper 30 meters of the site, as a way to differentiate betweenvarious subsurface conditions. Different equations apply to strike-slip or reverse-slipearthquakes, and we used the strike-slip equations. Boore, Joyner and Fumal did notprovide equations for vertical motion.• Sadigh, Chang, Egan, Makdisi and Youngs (1997). These equations, primarily basedon strong motion data from California earthquakes, have evolved over the years. The1997 equations include data from the 1994 Northridge Earthquake. Distance is definedas R up , the minimum distance to the fault rupture. Attenuation equations are presentedfor two general site categories, “rock” and “deep soil”. The authors also differentiatedbetween strike-slip and reverse faulting and concluded that ground motions from normaland strike-slip faulting do not significantly differ. For this study, we averaged thepredictions obtained for “rock” and “deep soil”, and used the equations for strike-slipfaulting.• Campbell (1997, 2001). In 1997, Campbell updated his 1994 equations for predictinghorizontal and vertical free-field PGA and PSAs at 5 percent damping. His 1997database included 645 horizontal recordings and 225 vertical recordings for PGA, and226 horizontal and 173 vertical recordings for PSA's. Source-to-site distance is definedas R seis , the shortest distance between the recording site and the interpreted zone ofseismogenic rupture on the fault. This implies that softer sediments and the upper 2 to 4D-3GEI Consultants Project # 04035-0

km of the fault zone are primarily non-seismogenic, at periods of engineering interest.Campbell also introduced a correction for long-period site response, using the depth tobasement rock (D b ) as a parameter. He differentiated between strike-slip and reversefaulting, and suggested that estimates for normal faulting be taken as the geometricaverage of those for strike-slip and reverse faulting. Different equations apply to“alluvium or firm soil”, “hard rock” and “soft rock”. For this study, we concluded thatthe arithmetic average of Campbell’s equations for “soft rock” and “alluvium” wouldbest represent the local ground motion. Campbell and Bozorgnia (2000) published anabbreviated and slightly different form of the 1997 equations and, in 2001, Campbellprovided an erratum and additional guidance on the use of his equations for sites simplyclassified as “soil” or “rock”. He has submitted for publication another new set ofequations (Campbell, 2003), which we did not consider. Like the Crouse-McGuire’sequations, we only used Campbell’s horizontal attenuation equations for comparativepurposes, but used his vertical ground motion estimates not to rely on a single set ofequations (Abrahamson and Silva).While the above relationships are well-accepted in common earthquake engineering practice,they have been recently updated or will be revised in the near future. Campbell and Bozorgnia(2003) published revised attenuation relationships to update Campbell’s 1997/2001 equations,using a database of records obtained throughout the world. We have tested the applicability ofthese new equations to California sites, such as the Lafayette site. For the MCEs and siteconditions considered, Campbell and Bozorgnia’s horizontal spectra are comparable to thoserecommended in this report, but the vertical spectra are lower. This is of no significance, asvertical spectra presented in this Appendix were included for informational purposes only. Theother attenuation relationships (Boore, et al., Abrahamson and Silva, and Sadigh, et al.) shouldbe updated around the end of 2005, when the Pacific Earthquake Engineering Research (PEER)Center publishes the findings of a recent workshop on the subject.D-3 MethodologyFirst, we calculated average response spectra for the distance parameters and upper bound ofmagnitude (see main part of the report, Section 4.3) that would result in the most severe groundmotion estimates at the site. The Hayward Fault is the controlling feature, in the case ofLafayette Dam, but we also developed ground motion estimates for the Calaveras, San Andreasand Lafayette-Reliez Valley faults. Such response spectra are designated as deterministic, anddescribe the motion from each critical fault, as opposed to probabilistic response spectra, e.g.USGS estimates. Probabilistic spectra include the contributions of all applicable faults andseismogenic zones that can produce ground motion with a specified probability of nonexceedanceat the site. In the case of “high risk” dams, the State Division of Safety of Dams(DSOD) requires deterministic criteria based on 84 th percentile estimates.To reduce the uncertainty associated with our estimates, we combined the horizontal responsespectra developed from the Abrahamson and Silva, Boore, et al. and Sadigh et al.’s attenuationequations, and the vertical spectra from the Abrahamson and Silva and Campbell’s equations,using a geometric averaging process. Geometric averaging is based on the natural logarithms ofthe calculated PGA's or PSA's) and is compatible with the regression analysis procedures usedD-4GEI Consultants Project # 04035-0

for the development of such equations. As these equations do not use the same periods ofvibration, we obtained intermediate estimates by interpolating from the natural logarithms of thePSA's calculated at the two closest periods values. All of the above relationships apply to theaverage of the primary and secondary components of horizontal ground motion.D-4 Near-Field And Directivity EffectsIt has been observed that increase or decrease of ground motion may occur as a result of thedirection of propagation of the rupture (a Doppler-like effect). Amplitudes in the forwarddirection of the rupture propagation will often be increased, while amplitudes in the backwarddirection reduced. Furthermore, systematic and significant differences have been observed inthe near-field between the fault-perpendicular (normal to the rupture plane) and the fault-parallelcomponents of long period ground motion. Although recognized, such effects have not beenused directly in the development of attenuation equations discussed in Section D-2.Sadigh, et al. (1993) first suggested that, within 10 km of the rupture surface, the fault normalcomponent be increased by 20 percent for spectral periods of 2.0 sec or greater, and the faultparallelcomponent decreased by 20 percent, compared to the geometric averages. Others(Ansary, et al., 1995) proposed that the largest component of motion be approximated from theaverage estimate using a 15 percent increase. Somerville, et al. (1997) concluded that near-fieldrupture directivity effects become significant around 0.6 sec period and increase in size withlonger periods. They proposed a simple geometric modification method based on empiricalanalysis of strong motion attenuation data. Abrahamson (2000) used a similar model to estimatedirectivity and radiation pattern effects on the fault-normal and fault-parallel components ofground motion.In 2000, Somerville indicated that his 1997 and the just developed Abrahamson models were toosimple, and that near-field and directivity effects rather manifest themselves as narrow-bandpulses at a period that appears to increase with the magnitude of the causative event. However,because it is impossible to reliably predict at what period such effects will occur, Campbell(2003) concluded that the 2000 Somerville directivity pulse model needed more developmentbefore it could be used for engineering purposes and that, until then, the 1997 Somerville and2000 Abrahamson models would remain the state-of-the-art.Because Lafayette Dam is a very wide-crested embankment dam, it is should respond at periodsof vibration of 0.6 sec or greater and near-field and directivity effects could influence itsresponse, especially in the case of the Lafayette-Reliez Valley, Calaveras and/or Hayward faults.The correction factor for directivity effects depends on the ratio s/L, where L represents the totalrupture length and s the distance from the epicenter to the normal projection of the site locationto the fault surface trace, and on the angle (θ) between the fault trace and a line joining the sitelocation to the epicenter. In the case of strike-slip faulting and rupture propagating toward thesite, maximum directivity effects result in a maximum increase of about 68 percent in thespectral acceleration at 5 sec period, compared with average predictions. Minimum directivityeffects result in a similar reduction, when the fault rupture propagates away from the site. Inaddition to directivity effects, near-field fault-normal (FN) and fault-parallel (FP) effects cancause up to about 107 percent increase (at the largest magnitudes) in fault-normal amplitudes,compared with average predictions, with a similar decrease for fault-parallel motion. AsD-5GEI Consultants Project # 04035-0

directivity and fault-normal/fault-parallel effects are cumulative, this could result for strike-slipfaulting in a cumulative directivity factor ranging from near-zero at 0.6 sec period to a maximumof about 3.5 at 5 sec period, for the most demanding combination of orientation, distance, faultrupture length and magnitude.For dam evaluation purposes, the DSOD recommends (record of conversation with Jeff Howard,4/30/2004) that directivity and fault-normal effects be considered for the largest component ofmotion (to be used perpendicularly to the dam crest), while the secondary component of motionis not reduced from the average motion based on attenuation equations. Directivity effects arenot necessarily broad-banded, based on DSOD’s own recent studies, which seems in consistencewith Somerville (2000). The DSOD indicated it will soon publish a position paper but, untilsuch paper is available, we have adopted Campbell’s recommendation and used the Somervilleand Abrahamson’s correction method to approximately account for directivity and near-fieldeffects in our development of recommended seismic criteria and to define the primary horizontalcomponent of motion (assumed to be oriented perpendicularly to the crest).Many combinations are possible between the parameters involved in estimating directivityeffects. As these depend on the unknown location of the initial rupture, a rigorous solution is notpossible. Furthermore, as we have recommended 84 th percentile criteria based attenuationequations developed for the average of the primary and secondary components of motion, someunknown portion of the directivity effects is already included in the error term (σ) of theattenuation equations. Yet, for this review, we have considered that maximum directivityeffects, which correspond to a value of the product (s/L) x cos(θ) greater than 0.4, would apply.For dam evaluation purposes, the most critical orientation of the dam is when its crest axis isparallel to the fault trace, as the largest shaking would occur perpendicularly to the crest, in theupstream to downstream direction. The crest of Lafayette Dam is not parallel to the Hayward,Calaveras or Lafayette-Reliez Valley faults, which should reduce fault-normal (FN) and faultparallel(FP) effects for the components of ground motion parallel or perpendicular to the damcrest (the most critical to the dam). The angle (α) between the axis of Lafayette Dam crest andthe faults considered are approximately 33 o , for the Hayward and San Andreas Fault, a minimumof about 38 o for the Calaveras and about 65 o for the Lafayette-Reliez Valley Fault. To take suchorientation into account, we have assumed the motion perpendicular to the crest of the dam to beequal to the FN motion, multiplied by cos(α). Hence, a slight reduction of the FN motion isaccounted for, based on the applicable crest orientation.The average horizontal and vertical 84 th percentile spectra predicted by the various attenuationequations are presented in Tables D-1 to D-8. They correspond to the applicable subsurfaceconditions and dominant strike-slip type of motion, for each of the critical faults potentiallyaffecting Lafayette Dam. The recommended near-field and directivity correction factors arelisted in Tables D-9 through D-12. The 50 th and 84 th percentile recommended average andcorrected spectra are plotted on Figures D-1 to D-4.The above response spectra correspond to 5 percent of critical damping, which is the dampingvalue commonly used for comparative studies of the characteristics of strong ground motion. Asother damping values are may be needed for analysis or simplified dynamic response evaluationpurposes, we also obtained response spectra for damping values other than 5 percent. In thisD-6GEI Consultants Project # 04035-0

purpose, we used a scaling procedure using spectral ratios derived from Newmark-Hall (1985).The corresponding spectral values (which also include directivity and near-field effects) arepresented in Section 4.6 of the main part of this report.D-7GEI Consultants Project # 04035-0

REFERENCESAPPENDIX DAbrahamson, N.A. (2000), “Effects of Rupture Directivity on Probabilistic Seismic HazardAnalysis”, CD-ROM Proceedings, 6 th International Conference on Seismic Zonation, November12-15, Palm Springs, CA, EERI, Oakland, CA.Abrahamson, N.A; Silva W.J. (1997), "Empirical Response Spectral Attenuation Relations forShallow Crustal Earthquakes", Seismological Research Letters, Seismological Society ofAmerica, Volume 68, No. 1, January/February, pp. 94-127.Ansary, M.A.; Yamazaki, F.; Katamaya, T. (1995), “Statistical Analysis of Peaks and Directivityof Earthquake Ground Motion”, in Earthquake Engineering and Structural Dynamics, Vol. 24,pp. 1527-1539Boore, D.B.; Joyner, W.B; Fumal, T.E. (1997), "Equations for Estimating Horizontal ResponseSpectra and Peak Acceleration from Western North American Earthquakes: A Summary ofRecent Work", Seismological Research Letters, Seismological Society of America, Volume 68,No. 1, January/February, pp. 128-153.Campbell, K.W. (2003), “Engineering Models of Strong Ground Motion”, Chapter 5, EarthquakeEngineering Handbook, W.F. Chen and C. Scawthorn, Editors, ICBO/SEA/CRC-Press, Editors,pp. 5.1-5.76Campbell, K.W. (2001), “Erratum: Empirical Near-Source Attenuation Equations Relationshipsfor Horizontal and Vertical Components of Peak Ground Acceleration, Peak Ground Velocity,and Pseudo-Absolute Acceleration Response Spectra”, Seismological Society of America,Seismological Research Letters, Vol. 72, p. 474.Campbell, K.W. (1997), "Empirical Near-Source Attenuation Equations Relationships forHorizontal and vertical Components of Peak Ground Acceleration, Peak Ground Velocity, andPseudo-Absolute Acceleration Response Spectra", Seismological Research Letters,Seismological Society of America, Volume 68, No. 1, January/February, pp. 154-179Campbell, K.W.; Bozorgnia, Y. (2003), “Updated Near-Source Ground-Motion (Attenuation)Relations for the Horizontal and Vertical Components of Peak Ground Acceleration andAcceleration Response Spectra”, Bulletin of the Seismological Society of America, Vol. 93,No.1, February, pp.314-331.Campbell, K.W.; Bozorgnia, Y. (2000), “New Empirical Models for Predicting Near-SourceHorizontal, Vertical and V/H Response Spectra: Implications for Design”, CD-ROMProceedings, 6 th International Conference on Seismic Zonation, November 12-15, Palm Springs,CA, EERI, Oakland, CA.Crouse, C.B.; McGuire, R., “Site Response Studies for Purpose of Revising NEHRP SeismicProvisions”, Proceedings, SMIP 94.D-8GEI Consultants Project # 04035-0

Sadigh, et al. (1993), "Specification of Long-Period Ground Motions: Updated AttenuationRelationships for Rock Site Conditions and Adjustment Factors for Near-Fault Effects", ATC-17-1 Seminar on Seismic Isolation, Passive Energy Dissipation and Active Control, March 11-12, San Francisco, CA, Proceedings, pp. 59-70.Sadigh, K.S.; Chang, C.Y.; Egan, J.A; Makdisi, F.; Youngs, R.R. (1997), "AttenuationRelationships for Shallow Crustal Earthquakes Based on California Strong Motion Data",Seismological Research Letters, Seismological Society of America, Volume 68, No. 1,January/February, pp. 180-189.Somerville, P.G. (2000), “New Developments in Seismic Hazard Estimation”, CD-ROMProceedings, 6 th International Conference on Seismic Zonation, November 12-15, Palm Springs,CA, EERI, Oakland, CA.Somerville, P.G; Smith, N.F.; Graves, R.W; Abrahamson, N.A. (1997), "Modification ofEmpirical Strong Ground Motion Attenuation Relations to Include the Amplitude and DurationEffects of Rupture Directivity", Seismological Research Letters, Seismological Society ofAmerica, Volume 68, No. 1, January/February, pp. 199-222.D-9GEI Consultants Project # 04035-0

APPENDIX EREVIEW OF LABORATORY DATAE.1 Static Analysis ParametersWe reviewed the extensive field and laboratory data developed over the years and, especially, theanalysis parameters selected for the S&W and WA studies. A summary of the availableinformation is presented in Tables 7-1, 7-2 and 7-3 of this report. Most of the static parametersrequired for our slopes stability analyses were obtained from such review.Table E-1 summarizes the field and laboratory unit weights and index properties of the variousdam and foundation zones identified by the DSOD (see Figure 7-2 of this report). Theseproperties include moisture content, Atterberg limits (AL), clay content (percent passing #200sieve) and gravel content (percent retained in #4 sieve). Average values, as reinterpreted fromour review, are also provided in Table E-1 for each of the dam and foundation zones.Table E-2 presents the results of the undrained shear strength (S u ) testing previously performedon the various materials, as summarized in the 1966 Shannon and Wilson (SW) and 1976 W.A.Wahler and Associates (WA) reports. SW provided a range of S u values for each type ofmaterials. In Table E-2, Zone 4 refers to either Zone 4 or Zone 4.5 previously used by theDSOD. As discussed subsequently, we considered a single zone, Zone 4, to represent these twozonesTable E-3 summarizes the results of isotropically consolidated undrained triaxial tests (TXICU)performed by SW. We reviewed the interpretation of strength envelopes provided in AppendixA of the SW report, and generally agree with such interpretations. Table E-4 summarizes theresults of all strength testing performed by WA, and includes unconfined compression tests(UC), unconfined undrained triaxial tests (TXUU), and isotropically consolidated undrainedtriaxial tests (TXICU) or anisotropically consolidated triaxial tests (TXACU or K o CU).Consolidated-undrained triaxial tests are also referred as CU tests in this report. As we foundsufficient details of the laboratory testing performed by WA, we independently re-estimated theunconfined shear strengths (S u ) and the total stress (c, Φ) and effective stress (c’, Φ’) strengthparameters from the raw laboratory test data included in one of the appendices the 1976 report.Our updated strength estimates are shown in Table E-4.We also assessed the liquefiability of the various dam and foundation zones, based on liquidlimits (LL), plasticity indexes (PI), and water contents (w, in % of dry unit weight). Section 7.6of the report discusses how these parameters were used to identify whether each AL data point,defined by LL and PI, is positioned within or outside of two critical zones, Zone A, “classiccyclic liquefaction” and Zone B “potentially liquefiable”. Potential liquefiability is furtherassessed through the results of a supplementary test, based on water content w and LL. Theresults of such classification and testing are presented in Table E-5 and on Figure E-1 through E-6, for each zone identified within the dam section and foundation alluvium. We found that asingle data point for a sample within Zone 5 of the alluvium was potentially liquefiable.Therefore and consistent with previous studies, we conclude based on the results of these testsE-1GEI Consultants Project # 04035-0

and because of the high percentage of clayey fines (materials passing the #200 sieve)encountered in all materials that they are not “liquefiable” (liquefiable being defined as potentialsudden loss of strength under applied cyclic loading).Based on the information contained in Tables 7-2 and 7-3, we plotted the effective stress andtotal stress Mohr envelopes for Zones 1, 2, 3, 4.5 and 5, as estimated by previous investigatorsand as re-estimated in this review. Zones 1, 2 and 3 represent the core, upstream anddownstream shells, respectively. Zone 4.5 and 5 are within the alluvium on which the dam isfounded, below the downstream shell and upstream shell, respectively. Based on the depth atwhich samples were recovered, we did not interpret any sample tested as belonging to Zone 4.Hence, based on their similar corrected penetration resistance and the lack of samples from Zone4, we combined Zone 4 and Zone 4.5 as a single Zone 4 in our slope stability calculations.In general, we have several comments regarding the way some of the CU triaxial tests wereperformed. These comments suggest that the testing performed 30 years ago or more may notmatch the quality of the testing that could be achieved today in a modern and well-equippedlaboratory:• The tests were performed very rapidly, especially considering the clayey nature of thesamples recovered. Typically, a maximum strain rate of 2 percent should be used, whichmeans that for the type of materials encountered at Lafayette Dam, such tests would nowbe performed over periods of time that would range from days to an entire week.• Results interpretation and, probably, data collection were done manually, which probablyexplains why the tests were performed so quickly. No computer-control was thenavailable to continue testing overnight.• Normally, change in sample volume should level out at the end of the initialconsolidation phase. Some tests plots indicate that the consolidation had not been fullycompleted before proceeding with the actual testing.• Some test were arbitrarily switched from stress-controlled to strain-controlled, sometimesseveral times during the testing.• The data presented in the appendices of the previous reports include no deviator stressesor pore-water pressures (u) when a state of failure has been reached.• Sample descriptions are sometimes incomplete or inconsistent with the field logs, e.g.sieve data for some sizes are not provided.We found one sample of the Wahler 1976 testing program (Zone 5, hole SS-23 at 105-10.7.5 ftdepth) to have a strength envelope somewhat lower than the other tests in that zone, due topossible disturbance of that sample upon recovery, or a perhaps simply indicating a local weakersample, not truly representative of the average strength of that zone. As the other tests in Zone 5materials were relatively consistent, we ignored that specific test in our recommended strengthestimate.E-2GEI Consultants Project # 04035-0

On Figures E-7 through E-11, effective stress strength envelopes developed from the S&W andWA testing programs, as well as those selected by the DSOD, are compared with thoserecommended after our review and interpretation of the available data. The results of the twoprevious testing programs are reasonably consistent, although some variability does exist. Totalstress strength envelopes are presented on Figures E-12 through E-16. Again, most of the testsprovide reasonably consistent results.As we have no direct detailed knowledge on how the tests were performed, we took a relativelyprudent, but not excessively conservative approach in our selection of strength envelopes foreach zone. Overall, specimens recovered from Zone 2 (downstream shell) display higherfrictional resistance and undrained strength than samples from the other zones, based on eithereffective stress or total stress strength considerations. Samples recovered from the core,upstream shell, or from the foundation do not exhibit strength characteristics that significantlydiffer from each other.E.2 Dynamic Analysis ParametersE.2.1 GeneralProperties significant to the dynamic response analysis of an embankment dam are the low-strainshear and bulk moduli, damping coefficients and the cyclic strength characteristics of thematerials encountered. Low-strain modulus and damping govern the dam response in the elasticrange. Undrained static and cyclic strengths define thresholds for possible occurrence of largestrains and plastic flow during dynamic response and, in the case of sandy materials, for anypotential development of excess pore pressure build up.Because no dynamic analyses were included in our review, we only performed a limitedevaluation of the 1976 dynamic testing results. Modern dynamic analysis methods give lessemphasis to laboratory cyclic triaxial testing than was given in the 1970’s. This is becausedynamic laboratory testing does not rigorously duplicate field conditions and because trulyundisturbed samples of a size sufficient for such tests to be meaningful are rarely recovered.While a few research laboratories still perform cyclic triaxial or cyclic direct shear tests, currentpractice favors the use of field penetrating testing (SPT or CPT) and shear wave velocitymeasurements to assess dynamic soil properties and lower-bound estimates of the cyclic strengthof the materials encountered.E.2.2 Dynamic Moduli and DampingWA used in 1976 the Seed and Idriss’ relationships (1970) for clay shear modulus and damping.The low-strain dynamic shear modulus of the dam and foundation materials was defined fromclassic relationships between G and S u . Field-measured shear wave velocities were also used inthat purpose. This methodology is reasonable, although the low-strain shear moduli may be onthe low side. This would be of limited significance in modern nonlinear analyses, except at verylow levels of dynamic shaking, of little interest to assessing the dam overall performance underearthquake loading. It should be noted that such relationships, in the case of clays, have beenupdated by Dr. K.H. Stokoe and his colleagues at the University of Texas, Austin.E-3GEI Consultants Project # 04035-0

E.2.3 Cyclic StrengthStress-controlled cyclic triaxial tests were performed in 1976 on selected specimens. There islittle basis to assess the quality of the dynamic testing program performed on the Lafayette Damand foundation materials, other than knowing that these tests were performed by W.A. Wahlerand Associates, a firm recognized at the time as a leader in the field. The WA laboratory waslong considered as being one of the best-equipped and best-experienced laboratories to performsuch tests.The WA 1976 laboratory testing provides a reasonable assessment of the cyclic strength of theLafayette materials. Stress-controlled cyclic tests were performed on isotropically consolidated(K c =1) or anisotropically consolidated samples (K c =1.5), for various confining pressures. Testswere performed on the core, shell and foundation materials. While the specimens tested did notexperience sudden loss of strength, cyclic strength curves were defined when the samples testedreach 5 percent or 10 percent axial strain under the applied cyclic deviator stress. The numbersof blows to reach such axial strain levels were recorded and define the “cyclic strength”. Thecyclic stress ratios (CSR) causing 10 percent axial strain in laboratory specimens are presented inTable 7-6 of the main part of this report.Cyclic triaxial tests do not represent field conditions as well as simple cyclic shear tests woulddo. However, the CSRs causing a specified level of axial strain (or liquefaction in the case ofsandy materials) in the triaxial tests (σ dc /2σ’ 3 ) can be related to the CSR under multi-dimensionalshaking conditions in the field (τ h /σ’ v ) by the following expression:τ h /σ’ v = C r σ dc /2σ’ 3 [ E-1 ]In the above equation, τ h represents the horizontal earthquake-induced shear stress, σ’ v thevertical effective overburden pressure, σ dc the applied cyclic deviatoric stress in the laboratorysample and σ’ 3 is the lateral initial consolidation stress of such sample. The correction coefficientC r varies from about 0.6 for K o = 0.4, to about 0.9 to 1.0 for K o = 1 (Seed and Idriss, 1982). Weinterpreted the WA cyclic tests results and computed the laboratory CSRs causing 10 percentaxial strain in the tested specimens at 10, 15, 20 and 30 cycles. Such laboratory CSRs can beconverted to field CSRs and were compared with estimated induced CSRs under the variousearthquake scenarios considered in order to quickly assess the “straining potential” of the damand foundation materials.The downstream foundation materials (Zone 4) appear to have lower laboratory CRSs that theupstream foundation (Zone 5). For example, see Table 7-6, for K c =1 and N=15 (whichcorresponds to the number of equivalent uniform stress cycles commonly used for the referencemagnitude M w 7.5), laboratory CSRs required to reach 10 percent axial strain, under 6,000 psfconfining pressure, are 0.35 in the downstream alluvium (Zone 4), but 0.40 and 0.55 in thecentral and upstream foundation (Zone 5). Similarly, for K c =1.5 and 6,000 psf, the CSR in Zone4 is 0.31, but 0.34 and 0.37 in Zone 5. For the two series of tests performed in the corematerials, one series (6,000 psf) resulted in the lowest CSR for N equal to 15 or greater,compared with the alluvium or the downstream shell materials, while the other series (3,000 psf)led to higher CSRs, compared with the other materials. Only a limited number of samples wereE-4GEI Consultants Project # 04035-0

tested in the shell materials (Zone 2), but these tests consistently yielding higher laboratory CSRsthan in the other materials.One could also consider using the corrected field penetration testing data to estimate lowerboundcyclic strengths for the various dam and foundation zones. But for the clayey materialsencountered, “cyclic failure” is progressive and would occur as a result of large straining, henceis more difficult to define than in the classic “liquefaction”. Hence, the applicability to clayeymaterials of standard relationships between cyclic stress ratios (CSR) at failure based on fieldobservations and SPT data corrected to cyclic simple shear condition is questionable and suchapproach was not considered.E.2.4 Post-Cyclic Failure Residual StrengthA parameter of potential significance to clayey dam analysis is the “post-cyclic” undrainedresidual strength, S u,r (Castro, et al., 1987). The concept of residual strength was unknown in1976, and has been controversial since first developed. It is now become better accepted in thegeotechnical engineering profession. While Castro's suggested laboratory testing proceduresremain difficult to implement and are sensitive to subjective interpretation, they wouldundoubtedly be helpful in the case of clayey materials such as present at Lafayette Dam and site,and could be supplemented by in-situ testing, as discussed subsequently. As for the cyclicstrength, simple estimates of the residual strength obtained through empirical relationships basedon the corrected standard penetration resistance (SPT) would not be reliable in the case of clayeymaterials such as those encountered. The relationships between the residual undrained shearstrength S ur and N 1 (60) cs based on various case histories and provided by Seed and Harder (1990)and other researchers (Seed, R.B., 1999) are most useful in the case of sandy materials, and theirapplicability to clayey materials would be highly questionable.In 1956, Dukleth back-calculated strength parameters for both the embankment and alluviumfrom the “at-rest” post-failure position of the 1928 failed embankment slope. He assigned acohesion of 580 psf and a friction angle of 12 degrees to these materials. Such calculationsprovide an indication of what the residual strength might be, and suggest a residual strengthsubstantially lower than the static strength parameters measured in 1966 and 1976 by the Districtand its Consultants.The residual strength of clayey materials remains difficult to measure in the laboratory, and isstill somewhat prone to controversy, when weak materials are present. This is because fieldconditions rarely yield “undisturbed” samples, and are difficult to reproduce in the laboratory.However, in-situ shear vane tests can be performed to obtain the peak and ultimate (remolded orresidual) strength of clayey materials. The use of carefully field-measured residual strengthswould be useful for any more detailed numerical analysis that the District might want to considerfor Lafayette Dam.E-5GEI Consultants Project # 04035-0

REFERENCESAPPENDIX ECastro, G. et al. (1987), "On the Behavior of Soils During Earthquakes -Liquefaction", in SoilDynamics and Liquefaction, A.S. Cakmak, Editor, Elsevier, Publisher, The Netherlands, pp. 169-204.Dukleth, G. W. (1956), “Lafayette Dam No. 31-2, Stability Study”, Internal Memorandum to Mr.A.W. Brown, DSOD, March 8, in East Bay Municipal Utility District’s project files.Seed, H.B.; Idriss, I.M. (1982), “Ground Motions and Soil Liquefaction During Earthquakes”,EERI Mongraph Series, 134 pp.Seed, H.B.; Idriss, I.M. (1970), "Soil Moduli and Damping Factors for Dynamic ResponseAnalysis", University of California, Berkeley, Report No. EERC/70-10, December, 15 pp.Seed, R.B. (1999), “Engineering Evaluation of Post-Liquefaction Residual Strengths”, presentedat TRB Workshop “New Approaches in Liquefaction Analyses”, Washington, D.C., 10 January,Proceedings, 10 pp.Seed, R.B.; Harder, L. F. H., Jr. (1990), "SPT-based Analysis of Cyclic Pore Pressure Generationand Undrained Residual Shear Strength", in Proceedings of H.B. Seed Memorial Symposium,Vol. II, May, pp. 351-376.E-6GEI Consultants Project # 04035-0

APPENDIX FCALCULATION OF EARTHQUAKE-INDUCED DEFORMATIONSIn this Appendix, we describe the various approaches taken to estimate the potential earthquakeinduceddeformations of Lafayette Dam.F-1 Simplified Newmark’s MethodIn 1965, Newmark estimated earthquake-induced displacements in embankments by assumingthat slope movements are initiated when inertia forces on a potentially sliding mass exceed theavailable yield resistance along the bounding surface of failure. Newmark’s method is normallyimplemented by double-integration of acceleration increments above the yield acceleration, ascalculated in a time history response analysis. When acceleration response histories have notbeen developed, as is the case herein, Newmark proposed a simpler approach where he relatedthe peak acceleration (PGA) and velocity (PGV) of the input motion to standard maximumdisplacement estimates depending on the sliding resistance and yield acceleration factor (N).The Newmark’s displacement estimates are, of course, significantly influenced by the computedyield acceleration (N or K y ). The PGV can be estimated from the applicable PGA and magnitudeof the earthquake considered.Newmark's method applies to dams on rigid foundations, hence may provide lower-bounddeformation estimates in the case of Lafayette Dam, which is mostly founded on deep alluvium.However, as discussed subsequently, we also took an upper-bound approach, which exaggeratesthe influence of the foundation alluvium.Newmark’s Method assumes rigid body movement of a failed portion of the embankment andformation of a distinct failure surface, two recognized potential limitations. Newmark's originalupper bounds of displacement applied to a "normalized" Maximum Probable Earthquake with aPGA of 0.50g and a PGV of 30 in/s. While the formulas can be used with other ground motionparameters, they imply a number of effective pulses for the standardized earthquake not greaterthan six. Newmark suggested that, for magnitudes larger than he considered in 1965, the numberof effective pulses be assumed to be proportional to the square root of the expected duration ofshaking. We made such an adjustment. Newmark also considered various standardizeddisplacement estimates for “symmetrical” or “unsymmetrical” resistance. In the case ofembankment dams, the DSOD (Vrymoed, 1996) extended Newmark’s procedure by applying itsuccessively to the upstream and downstream slopes. The crest settlement was obtained byvectorially combining the displacements of both slopes.In order to estimate crest settlement from the computed maximum slope displacements, wesimply took the crest settlement as half of the slope movements. As Lafayette Dam is notsymmetrical, we took Newmark’s standardized displacements for “unsymmetrical” resistance asthe basis for our estimated maximum displacements.F-1GEI Consultants Project # 04035-0

F-2 Makdisi-Seed’s Simplified ProcedureRecognizing that a dam responds as a flexible body, Sarma (1975) varied the accelerations as afunction of depth within the embankment. Makdisi and Seed (1977, 1978) further refined thatprocedure, using the examples of clayey dams, which actually included Lafayette Dam.Assuming a rigid foundation, they calculated the peak crest acceleration through a square-rootof-the-sum-of-the-squares(SRSS) combination of the peak accelerations of the first three modesof dam vibration. Standardized or project-dependent spectral shapes, such as those recommendedfor Lafayette Dam, see Figures D-1 to D-4 in Appendix D, can be used in that purpose. Then,using strain-dependent average dynamic properties (Seed and Idriss, 1970), they used theacceleration response of equivalent-linear (EQL) finite element models of several embankmentdams to correlate the average peak acceleration of the sliding mass with the depth of the failuresurface. They also related the normalized displacement of the soil mass with the ratio of yieldacceleration to average acceleration of soil mass (k y /k max ), for three different magnitudes, 6.5, 7.5and 8.0. The procedure can be refined using actual field or laboratory-tested dam properties.Makdisi and Seed used dams of moderate height (75, 132, and 150 feet) and their procedureapplies best to dams of similar height such as Lafayette Dam. In taller dams, peak accelerationsdo not necessarily increase regularly from base to crest, resulting in possible errors in estimatingk max . Although the Makdisi-Seed's procedure relates normalized displacement and dam heightthrough the calculated period, it is not clear whether the influence of the dam size is properlyaccounted for. For dams higher than 200 feet, a prudent approach may consist of increasingcalculated displacements proportionately to the dam height. Makdisi and Seed clearly stated thattheir procedure only applies to dams built of materials experiencing little or no strength lossduring earthquake shaking (such as densely-compacted sands or cohesive clays), a sometimesforgotten restriction. Such restriction does not apply in the case of Lafayette Dam. Due to theassumptions of no loss of strength during shaking and EQL properties, the procedure is,however, questionable for severe ground shaking (say 0.50g or greater) as plastic flow may beinduced even in competent embankment materials. Furthermore, it is very sensitive to crestacceleration estimates. As for the Newmark’s method, the Makdisi-Seed estimates are verysensitive to the calculated yield acceleration.Based on the location of the critical failure surfaces obtained in our static slope stability analyses,we assumed deep failure surfaces in our application of the Makdisi-Seed procedure to LafayetteDam. As we used peak strength parameters to obtain the yield accelerations, we lowered ourcalculated yield acceleration by 20 percent, which approximates the use a reduced strength of 80percent of the static undrained strength, as recommended by Makdisi and Seed in the applicationof their procedure.F-3 Bureau, et al.’s Empirical CorrelationsBureau, et al. (1985) used the observed performance of concrete face and earth core rockfilldams (ECRD's) to develop an empirical relationship between the local intensity of shaking,expressed as the Earthquake Severity Index (ESI), and the expected relative crest settlement forthis type of dam. Calculated settlements assume the dam to be founded on bedrock or hard soils,F-2GEI Consultants Project # 04035-0

although a few of the dams used in developing the correlation were constructed on alluvialfoundations. The ESI is defined as:ESI = PGA x (M - 4.5) 3where PGA represents the applicable peak ground acceleration in g's and M is the magnitude.[F-1]True nonlinear deformation analyses of a 325-foot high typical CFRD, using an elastic-plastic,Mohr-Coulomb model and stress-dependent friction angle, provided results consistent with theESI-settlement relationship (Bureau et al., 1985). As no volumetric changes were considered insuch verification analyses, the ESI-settlement relationship applies best to compacted rockfill, amaterial that does not develop significant loss of strength during shaking.In 1987, the authors tested the correlation with friction angles other than typically encountered inrockfill, using the results of physical model tests on dry sand embankments (Roth et al., 1986).The extended correlations are still primarily intended for rockfill dams, but can be used in thecase of densely compacted granular materials, using the applicable friction angle.In subsequent unpublished studies, Bureau tested the ESI relationships to estimate deformationsof embankments dams built of well-compacted soil, regardless of the materials they were builtof, and used comparison with the results of more detailed finite difference studies. He found thatthe ESI-relative crest settlement relationships led to reasonable estimates, in the case of materialsthat are not liquefiable and when computed deformations remain moderate. As the materials ofLafayette Dam have a high clay content, the application of this procedure, which relies primarilyon the effective friction angle, is conservative at low confining pressures, as the influence ofcohesion is essentially ignored. If failure surfaces are controlled by deep failure surfaces, thenthe frictional component of the strength envelope predominates.F-4 Jansen’s FormulaIn a discussion of the observed performance of Cogoti Dam, Chile, Jansen (1987) presented anempirical equation to estimate earthfill or rockfill dams settlement under earthquake loading. InJansen's equation, the estimated displacement U depends on the magnitude M, the yieldacceleration k y and the maximum crest acceleration k m . It is expressed as follows:U = 10 x (M/10) 8 x (k m - k y ) 1.5 / k y[F-2]Jansen did not describe how he developed his formula. A major shortcoming is that hiscalculated settlement is independent of the dam height. His verifications, using one hypotheticaland four existing dams, were based on "estimated", rather than calculated yield accelerations, butare influenced by the experience of this distinguished author. Jansen's equation implies that nodeformations will occur if the peak acceleration remains below the yield acceleration, andapplies to dams not vulnerable to liquefaction, such as Lafayette Dam. A correction factortaking into account the height of smaller dams was proposed by Bureau (1997).F-3GEI Consultants Project # 04035-0

F-5 Swaisgood’s ApproachSwaisgood (1995) estimated seismic crest settlements based on statistical treatment of empiricalinformation developed from a review of the seismic performance of 54 existing embankmentdams. He related the crest settlement (in percent of combined dam and alluvium thickness) to aSeismic Energy Factor (SEF) and three constants based on type of dam construction (K typ ), damheight (K dh ), and alluvial thickness (K at ) as follows:Relative Settlement (%) = SEF x K typ x K dh x K at[F-3]Like the ESI, the SEF depends on magnitude and PGA of the causative earthquake. Swaisgooddifferentiated between rockfill (ECRD's or CFRD's), earthfill (E) and hydraulic fill (HF) dams.Furthermore, his settlement estimate attempts to incorporate the effect of alluvial foundations.However, dam and foundation are assumed to contribute to the settlement in proportion of theirrespective heights, a potential limitation for weak dams on strong foundations, or the reverse.Material strength is not used other than through the consideration of three types of dams. Thelimited applicable data justify extreme caution when applying Swaisgood's formula to HF's orloose embankment dams that may experience significant excess pore pressures. This is not thecase for Lafayette Dam.In 1998, Swaisgood has submitted an updated version of his approach. He added 60 casehistories of embankment dams severely shaken by historic earthquakes and, following aregression analysis on the available data, expressed the computed dam crest settlement (CS) as afunction of the SEF and a resonance factor (RF), specified for three types of dams (earthfill,hydraulic fill and rockfill) as a function of the distance in km between the seismic energy sourceand the dam. Although Swaisgood’s equations are empirical and are based on the observedperformance of dams that may differ from Lafayette in size, slope geometry and zoning or typesof materials encountered, it is helpful to place deformation estimates in perspective with historicobservations. Swaisgood’s 1995 correlations represent the only simplified procedure thatdifferentiates between dams founded on bedrock or alluvium and, for that reason, were alsoconsidered in our deformation estimates for Lafayette Dam. We used a weighting factor of 0.5to each of his 1995 and 1998 equations.F-6 Idriss’s ApproachIn his review of the draft of our report, Dr. Idriss recommended a procedure similar to theMakdisi-Seed's procedure, but with the following modifications:F-4GEI Consultants Project # 04035-0

a) Use a 15 percent reduction in static undrained strength to compute the yieldacceleration (we have approximated a 20 percent reduction, as suggested in Makdisi-Seed's original paper).b) Estimate the transverse crest acceleration from empirical correlations based onobserved records, see Figure F-1, instead of the procedure recommended by Makdisi-Seed (which is based on the specified response spectrum and the estimated periodsof the first three modes of vibration of the dam).Other than the two above changes, the Makdisi-Seed’s and Idriss’ procedures are identical.They use the same empirical correlations between the depth y of the sliding mass and the ratiok max /a max of the effective acceleration of the sliding mass (k max ) and crest acceleration (a max ),and between the computed displacement and the ratio k y /k max , as established for magnitudes of6.5, 7.5 and 8 ¼.Modification a) slightly increases the effective yield acceleration (K y ) to be used, hence slightlyreduces computed deformation estimates, compared with Makdisi-Seed’s procedure.Modification b) has the most significant impact, since for the Hayward Earthquake (PGA =0.60g), the upper range crest acceleration computed from Fig. F-1 is 0.76g, hence quite lowerthan computed in the Makdisi-Seed's procedure, see Paragraph F-7. This appreciably reducesthe average acceleration k max of the sliding soil mass. Hence, both modifications a) and,especially, modification b) reduce estimated deformations, compared with the original Makdisi-Seed's method.It is not surprising that the range of crest accelerations obtained from Figure F-1 is lowerthan estimated in the Makdisi-Seed's procedure from the specified 84 th percentile responsespectrum. This is because Figure F-1 shows "upper-range" and "lower-range" relationshipsbetween base and peak accelerations, empirically developed from acceleration records obtainedat the crest of existing dams. Records from most "natural" earthquakes display a normalizedfrequency content significantly less complete (demanding) than the specified 84 th percentileresponse spectrum, except occasionally in a limited range of periods. Furthermore, the datashown on Figure F-1 include no base accelerations higher than 0.48g, and most likelycorrespond to magnitudes lower than that specified for the Hayward Earthquake (M w =7.25).Hence, while extrapolating the data shown on Figure F-1 a specified PGA of 0.60 g appearsreasonable, the procedure has not been verified for large magnitudes and short distancesbetween the site and potential earthquake sources, for which essentially no dam base/crestrecords have been obtained to date.F-7 Computed DeformationsComputed deformation estimates in the case of Lafayette Dam are presented on Tables F-1through F-4, for the four earthquake scenarios considered in this study, which correspond toMaximum Considered Earthquakes (MCE) centered along the Hayward, Calaveras, SanAndreas, or Lafayette-Reliez Valley faults at their closest approach to Lafayette Dam. For thoseof these methods that depend on the computed yield acceleration, separate estimates were madeF-5GEI Consultants Project # 04035-0

for the upstream and downstream slopes, as respective yield accelerations of 0.29g and 0.14gwere obtained, based on the total-stress method of analysis. The Idriss’ procedure was onlyapplied to the Hayward MCE, which is the most critical and controls the evaluation of theseismic safety of Lafayette Dam.Several of the analysis methods discussed in the previous paragraphs calculate estimatedmaximum deformations. To facilitate comparisons between the various procedures, weassumed that the ratio between maximum crest settlement and maximum displacement(deformation) would be 0.50. Such ratio was based on the two following lines of reasoning:a) A review of the maximum displacements and crest settlements reported in 1928 after thestatic slope failure of Lafayette Dam,b) Calculated average ratios between maximum dynamic crest settlements and maximumdisplacement vectors computed in detailed nonlinear time-history response analyses ofsix other embankment dams.In 1928, the incomplete top of the dam dropped a maximum of 24 to 26 feet in its centralportion. The lower portion of the downstream slope moved about 40 feet horizontally. Themaximum vector displacement of the downstream edge of the crest can be estimated at between50 and 60 feet, see Figure 3-3 of the main portion of this report. Hence, a ratio of 0.50 betweenmaximum settlement and displacement appears to be appropriate, in the case of Lafayette Dam,based on observations made in 1928.The choice of a value of 0.50 for such ratio is further substantiated in the results of detailednonlinear dynamic response analyses performed for six other dams and a pit slope by Mr. Bureauand his former colleagues at Dames & Moore. These studies were completed for PleasantValley, Magalia, Lake Madigan, Hidden and Los Angeles dams, in California, Los Leones Dam,Chile, and the South Slope of Pit O, a quarry facility operated by the Alameda County WaterDistrict. For these analysis cases, the specified seismic criteria were intended to representearthquakes between magnitudes 7 to 8+. Computed ratios between maximum non-recoverablecrest settlement and slope displacement ranged from 0.27 to 0.85, with a mean value of 0.46.Hence, in the absence of detailed studies, a ratio of 0.50 is a reasonable choice in the case of anembankment dam such as Lafayette Dam. Additional information on the aforementioneddetailed studies can be found in several related technical publications (Bureau, 1996, 1997, 1999;Bureau, et al., 1994, 1996; and Roth, et al., 1991)All methods, except Swaisgood, ignore the foundation soils in their calculations of deformationestimates. As the foundation alluvium probably significantly contributes to the expectable overalldeformations, as being potentially the weakest materials, we took two successive approaches:1) A “lower-bound” approach simply assumed that the dam is founded on a non-deformablefoundation, and displacements or relative crest settlements with respect to the rigid baseare calculated based on the dam height, or 132 feet.2) In an “upper-bound” approach, we assumed that the foundation alluvium, of maximumthickness 90 feet in the central part of the dam, would be part of the dam itself, whichF-6GEI Consultants Project # 04035-0

The San Andreas response spectrum specified in this study and the WA’s response spectrum forthat event are somewhat similar at the periods of interest to the dam response (0.5-2.0 sec), butthe WA spectrum is more conservative at high frequencies (period less than 0.5 sec), see Figure4.4 in the main text of this report. GEI’s Hayward and LRV response spectra are moredemanding than the WA Hayward spectrum, in that same range of periods.F-8GEI Consultants Project # 04035-0

REFERENCESAPPENDIX FBureau, G. (1999), “Seismic Analysis and Safety Evaluation of Embankment Dams”,Proceedings, Second US-Japan Workshop on Advanced Research on Dam EarthquakeEngineering, UJNR/JSDE, Tokyo, Japan, May 8-11, 16 pp.Bureau, G. (1997), “Evaluation Methods and Acceptability of Seismic Deformations inEmbankment Dams”, XIX th ICOLD Congress, Florence, Italy, Proceedings, pp. 175-200.Bureau, G. (1996), “Numerical Analysis and Seismic Safety Evaluation of Embankment Dams",Boston Society of Civil Engineers Section of ASCE (BSCES), Seminar on "Dam Inspection,Analysis and Rehabilitation", November 2, Bentley College, Waltham, MA, Proceedings, 54 pp.Bureau. G.; Edwards, R.; Blumel, A.S. (1994), “Seismic Design of Stage IV Raising, Los LeonesDam, Chile", 1994 Annual Conference, The Association of State Dam Safety Officials, Sep. 11-14, Boston, Massachusetts, in Proceedings Supplement, pp. 77-86, and ASDSO Newsletter, Vol.9, No.5, September, pp. 15-22.Bureau, G.; Inel, S.; Davis, C.A.; and Roth, W. H. (1996) “Seismic Response of Los AngelesDam, CA During the 1994 Northridge Earthquake", (co-authors: W.H. Roth, Sinan Inel &George Brodt), USCOLD Annual Meeting and Lecture, July 22-26, Proceedings., pp. 281-295.Bureau, G.; Volpe, R.L.; Roth, W.R.; Udaka, T. (1985), "Seismic Analysis of Concrete FaceRockfill Dams", ASCE Int. Symp. on CFRD's, Detroit, Oct. 21, in "Concrete Face RockfillDams - Design, Construction and Performance", pp. 479-508, and Closure (1987), ASCEJournal of the Geotechnical Eng. Div., Vol. 113, No. 10, October, pp. 1255-1264.Idriss, I.M. “Dynamic Stability Review of Lafayette Dam”, Letter-Report to GEI Consultants,Inc., August 31, 2004, Draft 2, 9 pp.Jansen, R.B. (1987), "The Concrete Face Rockfill Dam. Performance of Cogoti Dam underSeismic Loading", discussion of a paper presented at ASCE's Symposium on Concrete FaceRockfill Dams, ASCE Journal of the Geotech. Engineering Div., Vol. 113, No. 10, October.Makdisi, F.; Seed, H.B. (1977), "A simplified Procedure for Estimating Earthquake-InducedDeformations in Dams and Embankments" U. of California, Berkeley, EERC Report No.UCB/EERC-77/19, 33 pp. plus Appendices.Makdisi, F.I., and Seed, H.B., (1978). "Simplified Procedure for Estimating Dam andEmbankment Earthquake Induced Deformations". Journal of Geotechnical Engineering, ASCE,July.Makdisi, F.I., and Seed, H.B., (1979). "Simplified Procedure for Evaluating EarthquakeResponse". Journal of Geotechnical Engineering, ASCE, December.F-9GEI Consultants Project # 04035-0

Newmark, N.M. (1965), "Effects of Earthquakes on Dams and Embankments", Rankine Lecture,Geotechnique 15, No. 2, pp. 139- 160.Roth, W.H.; Bureau, G.; Brodt, G. (1991), “Pleasant Valley Dam: An Approach to Quantifyingthe Effects of Foundation Liquefaction”, 17 th ICOLD, Vienna, Austria, June, Proceedings, pp.1199-1223.Sarma, S.K. (1975), "Seismic Stability of Earth Dams and Embankments", Geotechnique 25, No.4, pp. 743-761.Seed, H.B.; Idriss, I.M. (1970), "Soil Moduli and Damping Factors for Dynamic ResponseAnalysis", University of California, Berkeley, Report No. EERC/70-10, December, 15 pp.Seed, H.B., and Idriss, I.M. (1982) "Ground Motions and Soil Liquefaction DuringEarthquakes", Earthquake Engineering Research Institute, Monograph Series.Swaisgood, J.R. (1995), "Estimating Deformation of Embankment Dams Caused byEarthquakes", ASDSO Western Regional Conference, Red Lodge, Montana, May 22-25.Swaisgood, J.R. (1998), “Seismically-Induced Deformation of Embankment Dams”, 6 th U.S.National Conference on Earthquake Engineering, Seattle, Washington, June 1998.Vrymoed, J.L. (1996), "Seismic Safety Evaluation of Two Earth Dams", in "EarthquakeEngineering For Dams", Western Regional Technical Seminar, Association of State Dam SafetyOfficials, April 11- 12, Sacramento, pp. 215-234.F-10GEI Consultants Project # 04035-0