Canister Design 2012 (pdf) (7.2 MB) - Posiva

POSIVA 2012-13
Canister Design 2012
Heikki Raiko
VTT
April 2013
POSIVA OY
Olkiluoto
FI-27160 EURAJOKI, FINLAND
Phone (02) 8372 31 (nat.), (+358-2-) 8372 31 (int.)
Fax (02) 8372 3809 (nat.), (+358-2-) 8372 3809 (int.)

ISBN 978-951-652-194-0
ISSN 1239-3096

Posiva-raportti – Posiva Report
Posiva Oy
Olkiluoto
FI-27160 EURAJOKI, FINLAND
Puh. 02-8372 (31) – Int. Tel. +358 2 8372 (31)
Raportin tunnus – Report code
POSIVA 2012-13
Julkaisuaika – Date
December 2012
Tekijä(t) – Author(s)
Heikki Raiko, VTT
Toimeksiantaja(t) – Commissioned by
Posiva Oy
Nimeke – Title
CANISTER DESIGN 2012
Tiivistelmä – Abstract
This report summarises the design of the Posiva disposal canister for spent nuclear fuel. The report
presents the design bases, the reference design, and summarises the performance analyses carried
out for the design. The report mainly addresses the long-term properties of the canister, but
operational safety aspects are also included because they may affect long-term safety.
Report presents the detailed design of the reference canister (for BWR fuel) and in more general
terms the variant designs for the other nuclear fuel types (VVER and EPR/PWR). The design
report also describes the main features of the canister manufacture, encapsulation and handling.
This report addresses aspects concerning the manufacture, quality control, mechanical strength,
chemical resistance, nuclear engineering analyses, thermal dimensioning, canister handling and
material ageing phenomena. Interaction of canister and other engineered barriers are included in
the study. The evolution of repository and its main drivers are considered if they have an impact on
the canister performance.
This report fulfils the requirements of the preliminary safety analysis-report (PSAR) concerning
the description of the disposal canister and of its performance qualification in the expected
repository conditions during its long-term evolution.
Avainsanat - Keywords
Final disposal canister, canister design, performance assessment, spent fuel.
ISBN
ISBN 978-951-652-194-0
Sivumäärä – Number of pages
156
ISSN
Kieli – Language
ISSN 1239-3096
English

Posiva-raportti – Posiva Report
Posiva Oy
Olkiluoto
FI-27160 EURAJOKI, FINLAND
Puh. 02-8372 (31) – Int. Tel. +358 2 8372 (31)
Raportin tunnus – Report code
POSIVA 2012-13
Julkaisuaika – Date
Joulukuu 2012
Tekijä(t) – Author(s)
Heikki Raiko, VTT
Toimeksiantaja(t) – Commissioned by
Posiva Oy
Nimeke – Title
KAPSELISUUNNITELMA 2012
Tiivistelmä – Abstract
Tämä raportti kokoaa yhteen Posivan käytetyn ydinpolttoaineen loppusijoituskapselin suunnittelun.
Raportti esittää suunnitteluperusteet ja referenssiratkaisun sekä vetää yhteen tulokset
suunnitelman toimintakykyanalyysien tuloksista. Raportti käsittelee pääasiassa kapselin pitkäaikaisturvallisuuteen
liittyviä kysymyksiä, mutta myös käytönaikaiseen turvallisuuteen liittyvät
kysymykset on käsitelty, koska osalla niistä voi olla vuorovaikutusta pitkäaikaisturvallisuuteen.
Raportti esittelee yksityiskohtaisesti referenssikapselin suunnitelman sekä yleisemmin sen
variaationa olevat muiden ydinpolttoainetyyppien kapselit. Varsinaiset toimintakykyanalyysit on
raportoitu pääreferensseinä esitellyissä raporteissa. Suunnitelmassa on käsitelty myös kapselin
valmistuksen, kapseloinnin ja kapselin käsittelyn pääpiirteet.
Tarkastelujen aihepiirit kattavat kapselin valmistuksen, laadunvalvonnan, mekaanisen lujuuden,
kemiallisen kestävyyden, ydinteknisen suunnittelun, lämpöteknisen mitoituksen, kapselin
käsittelyn sekä materiaalien ikääntymisilmiöt. Kapselin ja muiden päästöesteiden vuorovaikutus
on myös tarkasteluissa mukana. Loppusijoitustilan evoluutiota ja siellä vaikuttavia prosesseja
tarkastellaan niiltä osin kuin ne vaikuttavat kapselin toimintakykyyn.
Raportti sisältää alustavan turvallisuusselosteen (PSAR) tasoisen kuvauksen loppusijoituskapselista
ja sen vaatimuksenmukaisesta toimintakyvystä loppusijoitusolosuhteissa hyvin pitkällä
aikavälillä.
Avainsanat - Keywords
Loppusijoituskapseli, kapselin suunnitelma, polttoaine, toimintakykyarvio.
ISBN
ISBN 978-951-652-194-0
Sivumäärä – Number of pages
156
ISSN
Kieli – Language
ISSN 1239-3096
Englanti

1
TABLE OF CONTENTS
ABSTRACT
TIIVISTELMÄ
DEFINITIONS AND ABBREVIATIONS .......................................................................... 5
FOREWORD .................................................................................................................. 7
1 INTRODUCTION ................................................................................................... 9
2 CANISTER’S FUNCTION AS A PART OF ENGINEERED BARRIER SYSTEM
OF THE KBS-3 CONCEPT ................................................................................. 13
2.1 Safety functions .............................................................................................. 13
2.2 Performance targets ....................................................................................... 14
2.3 Safety functions, performance targets and safety-related guidance to design ...
................................................................................................................. 15
3 DESIGN BASES FOR CANISTER ...................................................................... 19
3.1 Sub-system requirements – Canister ............................................................. 20
3.2 Design requirements – Canister ..................................................................... 21
3.2.1 Canister performance .............................................................................. 21
3.2.2 Copper overpack requirements ............................................................... 22
3.2.3 Cast iron insert requirements .................................................................. 22
4 DESIGN LOADS ................................................................................................. 25
4.1 Handling loads ................................................................................................ 25
4.2 Incidents and accidents during encapsulation, transfer and disposal ............ 25
4.3 Internal loads .................................................................................................. 26
4.4 External mechanical loads ............................................................................. 28
4.5 External chemical loads ................................................................................. 31
4.6 Mechanical load combination ......................................................................... 31
5 NUCLEAR SAFETY CLASSIFICATION OF THE CANISTER ............................ 35
6 CANISTER SHAPE, DIMENSIONS AND SURFACE QUALITY ......................... 37
7 MECHANICAL AND PHYSICAL PROPERTIES OF THE
CANISTERCOMPONENT MATERIALS.............................................................. 49
7.1 Material qualities involved .............................................................................. 49
7.2 Mechanical properties .................................................................................... 50
7.3 Models for stress strain curves ....................................................................... 55
7.3.1 Models for stress strain curves for iron and steels .................................. 55
7.3.2 Model for stress strain curves of copper ................................................. 55
7.3.3 Copper creep model ................................................................................ 56
7.3.4 The copper creep model for multiaxial stress states ............................... 58
7.3.5 Comparison to copper creep tests for notched specimens ..................... 59

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7.3.6 Posiva’s copper creep testing and canister evaluation ........................... 59
7.4 Bentonite material model ................................................................................ 63
7.5 Physical properties of canister materials ........................................................ 65
8 VERIFICATION OF CANISTER DIMENSIONING .............................................. 67
8.1 Mechanical failure processes ......................................................................... 67
8.1.1 Copper overpack ..................................................................................... 67
8.1.2 Insert ....................................................................................................... 67
8.2 Mechanical failure criteria ............................................................................... 68
8.2.1 Copper overpack ..................................................................................... 68
8.2.2 Insert ....................................................................................................... 69
8.2.3 Summary of mechanical failure criteria relevance .................................. 74
8.3 Strength and damage tolerance ..................................................................... 74
8.3.1 Plastic collapse criteria ............................................................................ 75
8.3.2 Stress / strain criteria .............................................................................. 76
8.3.3 Fracture resistance criteria / allowable defect sizes ................................ 81
8.3.4 Essential design parameters ................................................................... 81
8.3.5 Strength of variant canister designs ........................................................ 83
8.4 Thermal behaviour ......................................................................................... 85
8.4.1 Temperature inside canister .................................................................... 85
8.4.2 Thermal expansion of canister components ........................................... 86
8.4.3 Thermal evolution of the canister surface ............................................... 89
8.4.4 Canister during permafrost ...................................................................... 91
8.4.5 Thermal behaviour of the variant canister designs ................................. 92
8.5 Cooling of the canister in all expected conditions .......................................... 92
8.5.1 Canister in encapsulation plant ............................................................... 92
8.5.2 Canister in transfer vehicle ...................................................................... 95
8.5.3 Fire .......................................................................................................... 96
8.5.4 Canister in repository .............................................................................. 97
8.5.5 Cooling capacity of variant canister designs ........................................... 99
8.6 Corrosion resistance ...................................................................................... 99
8.6.1 Atmospheric corrosion in the encapsulation plant ................................... 99
8.6.2 Corrosion during repository operation ................................................... 100
8.6.3 Corrosion in the repository under unsaturated, oxic conditions ............ 100
8.6.4 Corrosion in the repository under saturated as well as anoxic and
reducing conditions ............................................................................... 102
8.6.5 Stress corrosion cracking ...................................................................... 104
8.6.6 Corrosion inside the canister ................................................................. 105
8.6.7 Corrosion in the weld and grain boundaries .......................................... 106

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8.6.8 Conclusions about corrosion resistance ............................................... 107
8.7 Radiation dose rates .................................................................................... 108
8.8 Materials ageing due to radiation dose ........................................................ 111
8.8.1 Conclusion about radiation induced ageing .......................................... 112
8.9 Criticality safety ............................................................................................ 114
9 RATIONALE FOR THE SELECTION OF MANUFACTURING MATERIALS AND
THEIR SPECIFIED PROPERTIES ................................................................... 115
9.1 Insert materials ............................................................................................. 115
9.2 Overpack material ........................................................................................ 115
10 MANUFACTURING OF THE CANISTER COMPONENTS ............................... 117
10.1 Insert manufacturing ..................................................................................... 117
10.2 Copper overpack manufacturing .................................................................. 117
10.3 Component inspections ................................................................................ 118
10.3.1 Canister insert inspections .................................................................... 118
10.3.2 Copper overpack inspections ................................................................ 118
11 ENCAPSULATION PROCESS.......................................................................... 121
11.1 Fuel preparation ........................................................................................... 121
11.2 Fuel handling and packaging into canister ................................................... 121
11.3 Canister preparation before sealing ............................................................. 122
11.4 Sealing weld of the copper overpack ........................................................... 122
11.5 Final machining of welded surfaces ............................................................. 123
11.6 Weld controls ................................................................................................ 123
11.7 Final control .................................................................................................. 126
12 CANISTER TRANSFER AND DEPOSITION .................................................... 127
12.1 Canister handling and transfer ..................................................................... 127
12.2 Control of handling and transfer damages ................................................... 128
12.3 Canister deposition ....................................................................................... 128
13 CANISTER INITIAL STATE .............................................................................. 129
13.1 Fuel types ..................................................................................................... 129
13.2 Average fuel burnup, number of fuel assemblies and activity inventory ...... 131
13.3 Fillers and residual contents of canisters ..................................................... 132
13.4 Decay heat ................................................................................................... 132
13.5 Canister size, shape and material integrity .................................................. 133
13.6 Adverse effects of manufacturing process on the materials ......................... 133
13.6.1 Residual stresses in seal weld .............................................................. 133
13.6.2 Residual stresses in cast iron insert ...................................................... 134
13.6.3 Temper embrittlement in cast iron insert ............................................... 136

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14 ASSESSMENT OF CANISTER COMPLIANCE WITH DESIGN
REQUIREMENTS ............................................................................................. 137
14.1 Design analysis evidence against design requirements ............................... 137
14.2 Continuing research work on performance assessment .............................. 139
14.3 Uncertainty of analyses and assessments ................................................... 140
15 SUMMARY ........................................................................................................ 143
REFERENCES ........................................................................................................... 147

5
DEFINITIONS AND ABBREVIATIONS
Buffer
Compacted bentonite blocks and pellets surrounding the copper canister in the deposition
hole.
BWR
Boiling water reactor.
Canister
Metal container used for spent nuclear fuel disposal in bedrock.
Canister overpack
Corrosion protection and leak-tight shell around canister insert. Made of oxygen-free
copper.
Canister insert
Cast iron body of the disposal canister. Stands for mechanical loads, inhibits criticality,
part of radiation protection shield, part of decay heat transfer chain.
Deposition hole
The vertical hole where the canister and the surrounding buffer are emplaced in KBS-
3V concept.
Deposition tunnel
The tunnel, where the deposition holes are located in KBS-3V concept.
EPR
European pressurised water reactor.
EBS
Engineered Barrier System refers to the barrier system, which consists of host rock as a
natural barrier, canister, clay buffer and backfill of the deposition tunnels. It includes
also the backfill and plugs in other openings and the plugs of the deposition tunnels.
Initial state
The initial state is the state a given component has after it has been emplaced according
to its design when the direct control over that specific part of the system ceases and only
limited information can be made available on the subsequent development of conditions
in that part of the system or its near-field.
KBS
(Kärnbränslesäkerhet). The method for implementing the spent nuclear fuel disposal
concept based on multiple barriers.
KBS-3V
(Kärnbränslesäkerhet 3-Vertikal). The reference design alternative of the KBS-3
method, in which the spent nuclear fuel canisters are emplaced in individual vertical
deposition holes.

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LO1, LO2
Loviisa nuclear reactor units 1 and 2. Type VVER 440.
MCNP5, MCNP
A General Monte Carlo N-Particle Transport Code, Version 5, Los Alamos National
Laboratory.
MX-80 bentonite
High grade sodium bentonite, known by the commercial name MX-80, produced by
American Colloid Company in Wyoming, USA and distributed by Askania. MX-80 is a
blend of several natural sodium-dominated bentonite horizons, dried and milled to millimetre-sized
grains. The reference buffer material for Posiva Oy.
NDT
Non-destructive testing.
OL1, OL2, OL3, OL4
Olkiluoto nuclear reactor units 1 - 4. OL1 and OL2 are BWR-reactors in operation, OL3
is EPR-type (in construction) and OL4 is so far only a decision-in-principle.
PSAR
Preliminary Safety Analysis Report.
PWR reactor
The pressurised water reactor.
RSC
Rock Suitability Criteria.
SKB
Svensk Kärnbränslehantering Ab (Swedish Nuclear Fuel and Waste Management Company).
STUK
Radiation and Nuclear Safety Authority, Finland.
VAHA
Posiva’s requirement management system.
VTT
Technical Research Centre of Finland
VVER 440
The Russian pressurised water reactor type used in Loviisa.
Äspö HRL
Hard Rock Laboratory in Äspö, Sweden.

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FOREWORD
This report, “Canister design 2012”, is a summary of the design bases, expected loads
and design analyses of the disposal canister within the KBS-3 method for geologic disposal
of spent nuclear fuel. The report summarises the requirements concerning the
operational loads, mechanical strength, creep and fracture resistance, radiation shielding,
sub-criticality and corrosion resistance that affect the design and assesses the compliance
of the proposed design with these design requirements. Most of the analyses
have been performed and reported earlier during the long-standing evolution of the canister
design. The so-called reference canister (for BWR-type spent fuel) is analysed
thoroughly and its design variants (for VVER-440 fuel and for EPR/PWR reactor fuel)
are discussed only in the context of the features deviating from the reference design.
The writer and principal compiler of this report is Heikki Raiko (VTT). Many specific
areas are either written by subject-matter experts or their original reports are quoted.
Mechanical material properties are based mainly on material samples from SKB tests on
the serial production of the canisters. Mechanical strength, creep and fracture resistance
analyses are based mainly on activities carried out within the SKB-Posiva cooperation
or within SKB, which are in (Raiko et al. 2010). Some updated data of copper creeping
and cast iron fracture resistance is produced and published lately by Posiva.
The constitutive material modelling for copper behaviour in elastic, plastic and creep
condition is based on the work of Professor Rolf Sandström (KTH Royal Institute of
Technology, Sweden) and his colleagues. Modelling is originally reported in (Raiko et
al. 2010) and references therein. Additional copper creeping data, model for creeping
and finite element simulation of creeping of the canister overpack are produced by VTT
and published in 2012 by Posiva, too. Bentonite material modelling is based on work of
Dr Lennart Börgesson (Clay Technology) and his colleagues. The rock shear analyses
are made by Jan Hernelind (5T Engineering) and both the fracture resistance and the
canister buckling analyses are made by Peter Dillström (Inspecta Technology) and his
colleagues. These analyses and models are summarised in (Raiko et al. 2010) with
references to the original reports.
The thermal dimensioning of the repository and the cooling process inside the canisters
are analysed by Dr Kari Ikonen (VTT) and discussed in referenced reports (Ikonen &
Raiko 2012; Ikonen 2006). The basic mechanical analyses for the external pressure
loads for reference (BWR) canister and for its variant designs (VVER-440 and
EPR/PWR) are analysed by Kari Ikonen and reported in (Ikonen 2005).
The radiation-related design analyses (decay heat, radiation shielding, criticality safety,
neutron and gamma radiation dose on the insert material) are based on reports of
Markku Anttila and Anssu Ranta-aho (VTT).
The corrosion resistance of the canister is based mainly on Dr Fraser King’s work
published in (King et al. 2011b) and summarised in this report by Dr Barbara Pastina
(Saanio & Riekkola Oy) and Marjut Vähänen (Posiva Oy).
All experts are gratefully acknowledged.

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1 INTRODUCTION
The aim of this design report of the disposal canister for spent nuclear fuel is to present
the reference canister design (and its variants) and to assess its compliance with the design
requirements.
In a KBS-3 type disposal system, the canister is the principal engineered barrier between
the spent fuel and the environment. All the other barriers are, on one hand, supporting
and shielding systems for the containment function of the canister and, on the
other hand, a redundant barrier system that becomes in active use in case canister barrier
is unexpectedly damaged. This means that the neighbouring barriers both interact with
each other and act also independently as containment or a delay element for the releases
from the fuel. This is how bentonite buffer shields chemically, mechanically and biologically
and supports mechanically the canister and the surrounding bedrock shields
and supports the buffer and both of them act as reserve barrier to releases from canister
in case canister is damaged.
According to the regulations, the safety approach for disposal of spent fuel shall be that
the safety functions provided by the engineered barriers will limit effectively the release
of radioactive substances into bedrock for at least 10000 years (YVL D.5, Section 408).
Therefore, the canister safety function should be fulfilled with high reliability for at
least the time period referred to in the regulations.
Additional principles in the development of the canister structure are that the long life
expectation is based on natural materials that have verified the long life through natural
analogies and that the manufacturing methods of the canister components should be in
wide practical utilisation today and no further development of the manufacturing processes
can be postponed to future. However, the general development of the industrial
manufacturing processes during the utilisation era of the disposal canisters, some 100
years, can be utilised in due course, if seen beneficial.
The canister structure consists of a massive cast iron insert covered by a 49 mm-thick
copper overpack. Copper has been chosen as the shell material because of its wellknown
properties, its good thermal and mechanical properties and for its resistance to
corrosion in reducing environments. Cast iron has been chosen for the insert to provide
mechanical strength, radiation shielding and to maintain the fuel assemblies in the required
configuration. The copper canister lid is welded to the rest of the canister using
the electron beam welding (EBW) method. Friction stir welding (FSW) is the alternative
welding method considered. A detailed description of the canister design is given in
Chapter 6. The overall shape of the Posiva’s canister variants are shown in Figure 1.
The canister design and manufacture processes have been under development in Sweden
and Finland already for more than two decades. The reference design has changed
in small steps year after year. The current reference design is a product of a long chain
of scientific investigation, design analyses, process development, trial manufacturing,
testing and examination.

10
Figure 1. Copper-iron canisters for the spent fuel from the Loviisa 1-2 (VVER-440),
Olkiluoto 1-2 (BWR) and Olkiluoto-3 (EPR/PWR) reactors from left to right. All variants
of the canister have the same outer diameter of 1.05 m. The height of the canister
ranges from 3.55 to 5.22 m. Illustration by Posiva Oy.
The main documentation of the canister design in Posiva will consist of the following
principal documents that will be the base for preliminary safety analysis reporting
(PSAR) and the base for safety case of the long term safety to the extent that concerns
that engineered barrier system (EBS) canister. See Figure 2.

11
Canister production line 2012
POSIVA 2012-16
Canister design 2012
POSIVA 2012-13
Canister manufacture
POSIVA 2009-03
Canister sealing
POSIVA 2010-05
Canister strength
SKB TR-10-28
Component inspection
POSIVA 2012-35
Weld inspection
POSIVA 2010-04
Corrosion resistance
POSIVA 2011-01
Background information on:
• Material properties (SKB TR-10-28), (SKB TR-09-32), (SKBDoc 1207576),
(SKB R-09-14), (SKB R-10-64)
• Material models (SKB TR-10-28), (SKB TR-10-31), (SKB TR-09-32), (SKB R-09-14)
• Mechanical strength (SKBDoc 1206894), (SKB R-10-11), (WR 2005-12),
(SKB TR-05-18), (SKBDoc 1177857), (SKBDoc 1207429), (SKB TR-09-32)
• Fracture resistance (SKBdoc 1203550), (SKBDoc 1187725), (SKBDoc 1089758),
(SKB TR-10-29), (SKB R-10-11), SKBDoc 1206868
• Sub-criticality (WR 2005-13)
• Radiation dose rates (WR 2008-63)
• Decay heat (WR 2005-71)
• Thermal analyses (WR 2012-56), (WR 2006-19)
• Welding demonstrations (WR 2009-126)
Figure 2. The principal documents that specify and qualify the canister as a component
of the engineered barrier system. The main references of the principal reports are listed
in the lower part of the graph and they contain the main technical information of the
data and analyses.
This report is also properly connected to Design Basis and Performance Assessment
reports.

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2 CANISTER’S FUNCTION AS A PART OF ENGINEERED BARRIER
SYSTEM OF THE KBS-3 CONCEPT
The disposal system consists of the spent nuclear fuel, the canisters, the buffer, the tunnel
backfill and tunnel plugs, the auxiliary components, the geosphere and the biosphere
in the vicinity of the repository. According to the reference repository design, the spent
fuel canisters are disposed of vertically in deposition holes in a one-storey underground
facility with deposition tunnels at a depth of 400-450 m below ground (Saanio et al.
2012).
The safety concept for a KBS-3 repository according to Design Bases report is based on
the long-term isolation and containment of the spent fuel assemblies in the canisters and
on the retardation features of the disposal features in case of a release.
2.1 Safety functions
In accordance with the safety concept for a KBS-3 repository, safety functions are assigned
to the EBS and the host rock.
In Posiva's repository concept the safety functions of the EBS components according to
Design Bases report are described as follows:
1) The safety function of the canister is to:
Ensure a prolonged period of containment of the spent fuel. This safety function
rests first and foremost on the mechanical strength of the canister’s cast
iron insert and the corrosion resistance of the copper surrounding it.
2) The safety functions of the buffer are to:
contribute to mechanical, geochemical and hydrogeological conditions that
are predictable and favourable to the canister, and to protect canisters from
external processes that could compromise the safety function of containment
of the spent fuel and associated radionuclides, and
limit and retard radionuclide releases in the event of canister failure.
3) The safety functions of the deposition tunnel backfill and plug are to:
contribute to favourable and predictable mechanical, geochemical and hydrogeological
conditions for the buffer and canisters,
limit and retard radionuclide releases in the possible event of canister failure,
and
contribute to the mechanical stability of the rock adjacent to the deposition
tunnels.
4) The safety functions of the closure systems (backfill of underground openings
other than the deposition tunnels; various plugs and seals of shafts and tunnels)
are to:

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prevent the underground openings from compromising the long-term isolation
of the repository from the surface environment and normal habitats for
humans and other biota,
contribute to favourable and predictable geochemical and hydrogeological
conditions for the other engineered barriers by preventing the formation of
significant water conductive flow paths through the openings, and
limit and retard inflow to and release of harmful substances from the repository.
According to YVL D.5 the natural barriers and their safety functions may consist of:
"stable and intact rock with low groundwater flow rate around disposal canisters
rock around waste emplacement rooms where low groundwater flow, reducing
and also otherwise favourable groundwater chemistry and retardation of
dissolved substances in rock limit the mobility of radionuclides
protection provided by the host rock against natural phenomena and human
actions."
The surface environment is not given any safety functions; instead it is considered as
the object of the protection provided by the repository system.
2.2 Performance targets
Performance targets are specified for the safety functions of the EBS and target properties
are defined for the safety functions of the host rock. These targets indicate the extent
to which a safety function is fulfilled at various times during the repository evolution.
The actual performance targets for the disposal system are defined on the basis of such
properties that can be derived from measurable or otherwise observable properties -for
instance, with the aid of modelling.
In the case of the host rock, the term "target properties" is used in place of "performance
targets", since the properties of the rock are set by the natural features of the selected
site and the deposition hole location. On the other hand, the repository has to be adapted
to the local site conditions.
The performance targets for the canister are given in Table 1, along with a summary of
their rationales (for a more extensive discussion, see Performance Assessment report).

15
Table 1. Performance targets for the canister, main rationale and the related design
requirements according to Performance Assessment report, Table 2-1.
Performance target
Canister shall initially be intact
when leaving the encapsulation
plant for disposal except for
incidental deviations.
In the expected repository conditions
the canister shall remain
intact for hundreds of thousands
of years except for incidental
deviations.
Canister shall withstand corrosion
in the expected repository
conditions.
Canister shall withstand the
expected mechanical loads in
repository.
Canister shall not impair the
safety functions of other barriers.
Canister shall be subcritical in
all postulated operational and
repository conditions including
intrusion of water through
damaged canister wall.
The canisters shall be stored,
transferred and emplaced in a
way that the copper shell is not
damaged.
Design of the canister shall
facilitate the retrievability of
spent fuel assemblies from the
repository.
Main rationale
The main safety function of the canister is containment. The
performance of the welding method and NDT should be such
that containment is ensured at the initial state and for as long
as possible. After a period of a few hundreds of thousands of
years the radiological hazard of spent fuel will be similar to
the one posed by the uranium it was originally made of.
Corrosion is a potential mode of canister failure.
Mechanical loading is a potential cause of canister failure.
Spent fuel emits heat and radiation that are potentially detrimental
to the other barriers if the canister does not provide
adequate shielding. Furthermore, the canister itself will inevitably
undergo some chemical changes over time, the products
of which should not be detrimental to the other barriers.
If criticality is reached a large amount of energy is generated
causing damage to the canister and other barriers and widespread
release of radionuclides.
If the canister is damaged the containment is endangered.
As required in Decision in Principle 2000 and 2010.
(M 3/2010 vp, 6.5.2010)
2.3 Safety functions, performance targets and safety-related guidance to
design
The assessment of long-term safety is based on an understanding of the events and
processes that have the potential to impair the safety functions of the disposal system.
The starting point for the assessment of long-term safety is the initial state, which is
defined for each canister as the conditions prevailing in the near-field when the buffer
and backfill are installed (in other words, at the point when the direct influence and

16
monitoring possibilities of the disposed canister and buffer has come to an end). The
system is designed so that it evolves from the initial state through an early, transient
phase to a long-term evolution phase, during which changes are much slower. However,
during the entire regulatory compliance period, the slow changes and rare events with
related uncertainties need to be taken into account.
The safety assessment describes the performance and evolution of the repository over
time, taking into account all known phenomena and uncertainties that affect safety. The
way in which the system evolves, however, depends on its initial state, including uncertainties
and potential deviations, which in turn depend on decisions taken on repository
design and the implementation of such design. These decisions will be constrained by
design requirements, based in part on guidance from previous performance and safety
assessments. The development of the disposal system can, thereby, be considered as
continuous iteration between performance assessments and design bases, as shown in
Figure 3.
The long-term safety assessors provide feedback and guidance to the system design
concerning:
indications of the need for improved engineered robustness; these may be used
to increase confidence in the safety assessments; and
specifications of the uncertainties and deviations that can be tolerated such that a
target state is still achieved.
The iteration between the design and performance assessors is to ensure, as far as possible:
mutual compatibility of the engineered barriers with each other and with the
bed-rock, taking into account their respective safety functions;
resistance of the engineered barriers to the main thermal, hydraulic, mechanical
and chemical loads to which they will be subjected during evolution of the system;
and
robustness with respect to slow processes and unlikely events that may occur
over the regulatory compliance period.
The design bases presented in Chapter 3 are based on these principles.

17
Safety functions
Identify the external
loads/stresses
Design bases
Design base cases
Assumption of the
state of other barriers
Design criteria
Reference design
Assumption of the
design base cases of
other barriers
Quantitative
analysis of how the
loads affect on the
design
Sufficiency of
chosen set of
properties
Production line
reports
Figure 3. The iteration process between the mechanical design development of a component
and the safety assessment of the system.

19
3 DESIGN BASES FOR CANISTER
According to Design Bases report the system design premises comprise the objectives
set for the whole system, the limitations set by the environment, technology and knowledge
and the existing operating environment (regulations, responsibilities, organisations,
resources). These form the starting point for the definition of the design basis of
disposal operations.
The design basis refers to the current and future environmentally induced loads and
interactions that are taken into account in the design of the disposal system, and, ultimately,
to the requirements that the planned disposal system must fulfil in order to
achieve the objectives set for safety (i.e. the design premises).
In defining the design basis, Posiva shall, by regulation, on the one hand, assess the
likelihood of different scenarios and, on the other hand identify those deemed reasonable,
and assess those that may be possible but are considered highly unlikely. Although
only scenarios deemed reasonable are used as design basis scenarios, scenarios that are
deemed unlikely also need to be assessed in the safety case.
According to regulations (YVL D.5, paragraph 407), targets shall be specified for the
performance of each safety function. Safety functions are the main roles for each barrier,
from which performance targets for the engineered barriers and target properties for
the host rock are defined considering their respective safety functions. Individual performance
targets or target properties must be defined for each main component of the
repository system (canister, buffer, filling material, repository host rock).
The actual design requirements and design specifications are ultimately defined so as to
enable the achievement of the performance targets in the expected scenarios. The performance
targets have been set so that individual deviations or deficiencies will not endanger
the long-term safety of the whole disposal system. The performance targets form
the basis for the definition and implementation of the design requirements. The initial
state of the disposal system can be affected through the design requirements and system
implementation methods (up to the closing and sealing of each deposition hole or tunnel);
the degree to which the performance targets are met and the capability of the system
as whole to effectively isolate the radionuclides from the living environment is
evaluated through the assessment of the design basis scenarios of the evolution of the
disposal system.
Together with a number of issues determined by operational safety, environmental factors,
operational efficiency and quality assurance, it is the safety functions and performance
targets that govern the design bases of the canister. From these, the design and
implementation requirements and specifications have then been deducted on the basis of
development work and research data. The canister shall be designed in such a manner
that it meets the specified design requirements and manufactured in such a manner that
it fulfils the respective design specifications. The design base requirements are compiled
from the statements given in (YVL D.5; YVL D.3) and the retrievability is stated in the
Finnish Government Decision 478/1999.

20
These design bases lead furthermore to some design requirements of the canister design,
dimensions, material selection, metallurgical properties and chemical contents. The
bases also lead to some design requirements of the internal atmosphere, chemical
content and temperature of the canister void and internals. These are discussed later in
chapter 14 where the compliance with the requirements is assessed.
The quantitative canister design specifications derived from design requirements are the
principal target and output of this design report.
3.1 Sub-system requirements – Canister
Definition and objectives
Canister is a container with a water and gas tight shell and a mechanical loadbearing
insert in which the spent nuclear fuel is placed for final disposal in the repository. The
canister shall contain the spent fuel and prevent, and in the case of a leak, limit the
spreading of radioactive substances into the environment.
Containment
Canister shall initially be intact when leaving the encapsulation plant for disposal except
for incidental deviations.
In the expected repository conditions the canister shall remain intact for hundreds of
thousands of years except for incidental deviations.
Chemically resistant
Canister shall withstand corrosion in the expected repository conditions.
Mechanically resistant
The canister shall withstand the expected mechanical loads in repository.
Compatibility with the EBS and host-rock performance
The canister shall not impair the safety functions of other barriers.
Sub-criticality
The canister shall be sub-critical in all postulated operational and repository conditions
including intrusion of water through damaged canister wall.
Handling before disposal
The canisters shall be stored, transferred and emplaced in a way that the copper shell is
not damaged.

21
Retrievability
Design of the canister shall facilitate the retrievability of spent fuel assemblies from the
repository.
Safeguards
Encapsulation and disposal of the spent fuel shall be organised in a way that makes the
safeguards control of the nuclear material possible according to requirements of (YVL
D.5, Section 5.4).
3.2 Design requirements – Canister
Definition
The canister is composed of a leak-tight copper shell and of a load-bearing nodular cast
iron insert.
3.2.1 Canister performance
Chemically resistant
The copper overpack shall provide the corrosion resistance required in the postulated
repository.
Mechanical strength
The iron insert shall provide the mechanical strength required.
Sub-criticality
To ensure sub-criticality, the properties (e.g., enrichment, burnup) of the fuel inside the
canisters, as well as the internal geometry of the insert, shall be known precisely enough
to reach a high degree of confidence in criticality safety.
Limitation of radiation level
The shielding provided by the canister shall limit the dose rate to minimise radiolysis of
water outside the canister.
The fuel elements for encapsulation shall be selected in a pre-planned, controlled and
documented way to limit the radiation dose on the canister surface.
Limitation of heat generation
The heat generation inside the canister shall be limited in a way that the performance of
the other barriers is not impaired.

22
The fuel elements for encapsulation shall be selected in a pre-planned, controlled and
documented way to meet the decay heat limit set for each canister type.
Thermal conductivity
The canister materials shall have sufficient thermal conductivity so that the heat from
the spent nuclear fuel is effectively dissipated.
Canister geometry
The copper overpack and insert shall be dimensioned to that the insert can be installed
into the copper overpack.
3.2.2 Copper overpack requirements
Copper overpack is composed of a copper lid and a bottom welded into a copper tube or
of a copper lid welded into a copper tube with an integrated bottom.
Properties of the weld shall fulfil the same performance requirements as the rest of the
copper shell.
Corrosion resistance
The design, manufacturing and any further processing and handling of the canister shall
aim at limiting the risk for stress corrosion cracking in repository conditions.
Lifting and transfer
The copper overpack shall be designed to bear the load from canister handling and
transfer.
Dent marks and scratches on the copper surface shall be minimised during canister handling
and transport.
Copper overpack ductility
The canister copper overpack shall be designed to withstand the plastic deformation and
creep caused by any postulated mechanical or thermal load.
3.2.3 Cast iron insert requirements
Sub-criticality
The insert geometry and acceptance criteria for soundness shall be set so that subcriticality
is guaranteed.
Mechanical strength
The canister insert shall be designed to bear the hydrostatic pressure from groundwater
and from swelling of bentonite.

23
The canister insert shall be designed to bear the hydrostatic load caused by glaciation.
The canister insert shall be designed to bear unevenly distributed swelling loads.
The canister insert shall be designed to bear the loads from the postulated rock shear
displacements in the deposition hole.

25
4 DESIGN LOADS
Design loads for the canister structure are mechanical loads (pressure, local forces or
forced displacements), thermal loads (varying temperature in time or position), chemical
loads (chemical around the canister environment, including bacteria-induced chemical
loads) and radiation load (radiation embrittlement).
4.1 Handling loads
The lifting equipment and the shoulder in the copper lid collar and the whole of copper
overpack shall be dimensioned for the gravity load of the loaded canister weight multiplied
by the dynamic factor of lifting loads and required safety factor. During the encapsulation
process the canister is supported from the bottom lid until the point, where canister is transferred
from the encapsulation line trolley cradle onto the automatic guided crawler. At that
time and when the canister is loaded into the canister installation vehicle and installed into
the deposition hole, the canister is gripped and lifted from the lifting collar in the copper
lid. In these three cases all the canister weight is hinged on the copper lid collar only.
Description of canister handling in encapsulation plant is given in (Kukkola 2012, chapter
2). Basic dimensioning of the lifting shoulder is analysed in Section 8.3.2. The masses of
all variant canisters are given in Table 6, in Section 6 of this report. Canister lid lifting
shoulder strength including the effect of postulated cracks is assessed in the strength analysis
reports (Raiko et al. 2010) and originally in (Bolinder 2009).
Canister lifting from the lid collar is also necessary if the retrieval of the canister out of the
deposition hole becomes necessary. In such a case, the saturated and swollen bentonite is
first dissolved with high saline water and the canister is lifted up with normal type gripper
from the top lid collar. In this case the lifting load is not higher than during encapsulation,
vice versa, the buoyancy caused by the dissolving water makes the canister about 4 tons
lighter than as it is submerged. The retrieval of a disposed canister has been planned in
principle several years ago (Saanio & Raiko 1999). The important international test for
Posiva’s demonstration on this topic has been the canister retrieval test made by SKB at
Äspö in 2006. As a main conclusion the freeing trial showed that the tested method
works. The retrieval operations in various phases of disposal are described in (Saanio et
al. 2012, Section 5.4).
4.2 Incidents and accidents during encapsulation, transfer and disposal
The leading principle in encapsulation and disposal of spent fuel is that in case of incidents
or accidents that faces the canister the disposal process is halted for assessment of
possible damages caused on the canister. In case effective damages are detected or assessed,
the canister will be reversed in the process chain back to encapsulation station,
the outer lid will be machined loose and decoupled, and the canister will be docked to
the fuel handling chamber. After that the inner steel lid can be opened and removed with
the normal handling equipment in the chamber. After opening the canister, the fuel assemblies
are moved one by one from the canister into a temporary store in a drying vessel
inside the fuel handling chamber. Then the empty canister is loosened from the
docking station, transferred backwards below the receiving hall, wrapped in a contami-

26
nation protection foil and lifted up with a canister transport cradle into the receiving hall
for further repair or scrapping process. For details, see (Kukkola 2012, Sections 2.5.3
and 2.5.4).
The encapsulation and canister transfer process is planned in a way that canister cannot
fall or be dropped from a remarkable height as a result of single failure of an active
member or device of the canister lifting or transferring systems. Canister lifts are single
failure secured (reduplicated) for active components and the static components are accurately
designed and dimensioned against all postulated loads, thus the probability of
canister falling accident is very low.
The canister behaviour during a transport vehicle fire is analysed in (Lautkaski et al.
2003). In the case of vehicle fire, the temperature is not high enough to damage the canister
or fuel inside the canister, when the canister is inside the radiation shield cylinder
of the canister transfer vehicle. The analysis was made assuming the vehicle with heavy
rubber tires, but the vehicle fire load can be decreased using a crawler-type vehicle
equipped with tracks instead of rubber wheels. Thus the actual fire safety can be even
better if no rubber wheels are used.
Overall operational safety in encapsulation plant and in repository is analysed and
shown to fulfil the national nuclear safety regulation in (Rossi & Suolanen 2012).
In the worst accident cases the canister is assumed to break and all the fuel inside the
canister is assumed to be damaged. Even in the worst case, the radioactive releases outside
the plant do not exceed the safety limits of the regulation. Thus the operational
safety does not set any special or definite conditions for canister strength.
The leading design principle in canister handling accidents is as follows. The disposal
canister is not designed to maintain its long term properties after a major handling or
transport incident or accident in operation phase. If such accident happens, the canister
will be returned to the encapsulation plant, opened and unloaded, and the fuel will be reencapsulated
into an intact canister.
4.3 Internal loads
The spent fuel itself produces He-gas as a result of radioactive decay. The pressure generated
from this process in long-term is, however, lower than the external hydrostatic
pressure according to (Miller & Marcos 2007, Section 3.2.9), thus this process can be
ignored as a mechanical load.
Corrosion of the insert can generate H 2 -gas inside the canister, but this is possible only
after the canister has been breached and the water is filling the void inside canister according
to (Miller & Marcos 2007). Gas generation inside a leaking canister is not causing
any additional pressure load, because, due to leak, all pressure difference loads are
vanished.
The residual water trapped inside the canister during encapsulation might form nitrogen
acids with help of possible N 2 -gas and the radiation. These acids might cause anaerobic
corrosion inside a closed canister according to (Miller & Marcos 2007). To avoid this

27
risk, the canister inside atmosphere is changed from air to inert argon-gas (Ar) during
encapsulation and the amount of existing residual water in the fuel assemblies is minimised
through drying process during encapsulation. Thus even this type of corrosion
and gas production in such amount that could cause remarkable internal pressure load
can be ignored. The encapsulation process is generally described in (Kukkola 2012).
The insert with flat lids will be pressurised up to an internal overpressure of 0.1 MPa. This
is the load case during the encapsulation process when the normal atmospheric pressure of
inert gas is prevailing inside the insert and the canister is in a vacuum chamber for the
EBW of the copper lid. This load case is not applicable in case of the alternative welding
method FSW. FSW is made in normal atmospheric pressure. The flat lids are primarily
dimensioned for external pressure of 45 MPa. The only detail that needs to be examined
for this internal pressure load case during EBW welding is the top lid central screw and the
gaskets. During welding, all the pressure of the inside inert gas is loading the central screw
of M30 mm. M30 stands for a metric ISO standard screw thread and it is designated by
the letter M followed by the value of the nominal diameter in millimetres. The screw
calculation is given in Section 8.3.2.
The decay heat generation in canisters will depend on the amount, burnup, operational
history and cooling time of the fuel. The initial decay heat load in a canister is limited in
safety case assumptions to 1700 W (SKB 2009) in reference case for BWR canister. Accordingly,
the power limits are set for VVER-440 and EPR/PWR canisters as 1370 W and
1830 W, respectively. These numbers are derived from the reference design by weighting
with the respective canister’s cooling surface area; see Table 6 for surface areas. The thermal
conductivity of the copper body of the canister is two orders of magnitude higher than
the conductivity of the surrounding bentonite and rock in the repository. Therefore the
metallic canister surface will be practically at a uniform temperature and most of the thermal
gradient will exist in the bentonite and rock around the canister and in possible air
gaps between the solid material interfaces.
High neutron and gamma-radiation dose may cause embrittlement in the insert material.
Secondly, a too high gamma radiation dose rate outside canister may cause radiolysis in
the surrounding water, which may lead to increasing corrosion on the outer surface of the
canister or make changes in bentonite buffer. That is why the allowable dose rate on canister
surface is limited to 1 Gy/h to avoid excessive radiolysis of water outside canister.
These phenomena will be discussed more in Section 8.8 of this report.
The maximum fuel rod temperature inside the canister is estimated with numerical analyses
in (Ikonen 2006) to be about +230 °C. This is calculated assuming radiation heat transfer
and conduction through the Ar-gas between the fuel rods and the canister insert, when
the canister surface temperature is conservatively +100 °C. The 1.5 mm gap between the
insert and the copper overpack is also assumed to act as an isolator over which gap the
thermal flow is transferred by radiation only, because the gap may be in vacuum after the
EB-weld process for a while. The thermal expansion of various canister components will
be discussed later in Section 8.4.3 in this report.
If FSW sealing method is used instead of EBW, then the gap between insert and shell is air
filled with normal atmospheric pressure and the conductivity over the gap is initially re-

28
markably better and leads to lower maximum temperature in the insert and fuel. Thermal
conduction phenomena inside a canister are analysed and discussed in (Ikonen 2006). It is
worth noting that this particular analysis is made to ensure the fuel integrity after encapsulation
and it may be overly conservative in respect of insert temperature. In Section 8.4.2
there is made a more realistic comparison calculation of insert temperature between FSW
and EBW sealed canisters.
Excessive heat flow might lead to too high temperature in the canister components, which
may lead to unwanted chemical processes, excessive thermal deformations, strains and
stresses and change of material properties. The thermal properties of all the cooling chain
from fuel rods inside the insert up to ground surface above the repository govern the
maximum allowable decay power that can be accepted without of risk for local overheating.
The cooling chain is described in details for ex. in (Miller & Marcos 2007). The cooling
analyses of disposed canisters including the definition of minimum distances between
the canisters in repository are made and reported in (Ikonen & Raiko 2012).
4.4 External mechanical loads
Mechanical external loads come from the natural environment and from the behaviour
of the surrounding bentonite buffer. The depth of the Olkiluoto repository is defined to
be 400 - 450 m and the nominal depth is 420 m. Thus the maximum groundwater hydrostatic
pressure is 4.1 MPa. The maximum postulated 2.5 km ice layer during glaciation
at Olkiluoto area according to (Pimenoff et al. 2011). This 2.5 km ice sheet may
create an additional pressure of about 25 MPa to the groundwater pressure, if the effect
of ice layer is conservatively added to the hydrostatic pressure of the groundwater. The
bentonite buffer swells, when the bentonite is getting water-saturated. The wetting of
the bentonite buffer is expected to take place gradually after the deposition tunnel backfilling
and closing within some months or years depending on water inflow rate into the
tunnel and deposition holes. The bentonite swelling pressure is strongly dependent on
the final density of the buffer. The salinity of the absorbed water also affects swelling.
The specified final density for the buffer is 1950-2050 kg/m 3 (Design Bases, Section
5.5). This leads to a swelling pressure of 2-10 MPa, respectively, according to (Börgesson
et al. 2009). The maximum swelling pressure of bentonite depends on the density
and the chemical contents of the bentonite. In long term, the chemical contents of the
Na-bentonite may change. The Na-ion may be changed to Ca-ion. Ca-bentonite has a
remarkably higher swelling pressure than the Na-bentonite, up to 15 MPa in the high
density region, e.g. at a montmorillonite content of 0.83 and a saturated density of 2050
kg/m 3 (Karnland 2010, page 27). As a summary, the enveloping maximum sum of
isostatic pressure load for a canister at Olkiluoto site is about 44 MPa.
In Forsmark, Sweden, the design pressure for the canister has been taken as 45 MPa by
SKB. It is slightly higher than the calculated design pressure load at Olkiluoto area,
thus, when referring to the respective SKB strength analyses, Olkiluoto has some additional
safety margin in the assessment of the loads. All the referenced SKB strength
analyses are made using 45 MPa as the design pressure.
The bentonite swelling pressure can be somewhat unevenly developed and distributed,
especially during the water uptake in the early evolution but also in saturated condition,

29
if the dimensional tolerances of the deposition hole and the density of the bentonite are
remarkably variable. In earlier load specifications, (Raiko 2005) and (SKB 2006a), the
unevenly distributed swelling loads have been overly conservative assumptions. The
absolutely rigid supports and restraints assumed in the load specifications are mechanically
unfeasible in reality. It is essential to note that the swelling pressure of the bentonite
induces the strength of bentonite, but the strength is less than the pressure, see
material model for bentonite in Chapter 7.4.
The bentonite buffer applies a load on the canister and supports it inside the deposition
hole. The bentonite buffer has the same swelling properties on the locations of the postulated
load and on the location of the postulated support or reaction force; thus, the
maximum pressure acting on the canister surface inside bentonite buffer is limited to the
sum of hydrostatic pressure and the swelling pressure and, furthermore, the vectorial
sum of loads and supporting reaction forces have to be statically in balance. These
specifications lead to the following type of unevenly distributed load schemes for the
canister, which causes the maximum bending load on the canister, see Figure 4. The
derivation of the determining load cases in bentonite buffer wetting phase and during
saturated period are given in (Börgesson et al. 2009). The stressing effects due to the
maximum load cases are also preliminarily assessed in the same report.
Area 1 p
Area 3
p
L/4
L/8
Area 2
L·(½-⅛)
L/4
Figure 4. The most unfavourable swelling pressure distribution causing bending (Börgesson
et al. 2009).
The same report (Börgesson et al. 2009) also gives a heaviest load case expected from
the unevenly distributed swelling pressure for the copper overpack. This is the case,
where swelling pressure is different on top and bottom end of the canister and the shear
force from the radial swelling pressure will balance the canister loading. This load type
is shown in Figure 5.

30
MPa
2 = 573 kPa
m
1 =2550 kPa
MPa
Figure 5. The most unfavourable buffer swelling pressure distribution that causes shear
(Börgesson et al. 2009).
Figure 6. Rock shear load case variations for the canister according to the analysed
cases of (Hernelind 2010). Rock shear is a rare upset loading condition that may ask
for the most demanding deformation capability.

31
The integrity of the canister can be damaged due to shear-type rock movements if the
shear plane accidentally intersects the deposition hole and the shear amplitude is large
enough (see Figure 6 for load case variations). If the bentonite buffer (assumed to have
been transformed to Ca-bentonite) around the canister is 350 mm thick it is required to
withstand a rock shear of 5 cm with a velocity of 1 m/s according to (SKB 2009, Section
3.1.2). The bentonite material properties and swelling pressure is described in
(Börgesson et al. 2010). The risk caused by even larger rock movements, which may
occur during the melting phase of a major continental glacier, is minimised to allowable
level by locating the disposal gallery within justified respect distances outside major
fracture zones in the bedrock and by locating the deposition holes in such a way that is
not intersected by fractures with potential to undergo damaging shear movements (Hellä
et al. 2009). When the actual canister deposition hole positions are selected outside the
existing major fracture zones, the canisters will be exposed only to possible secondary
(new) fractures generated during future shear movement.
The possible presence of permafrost at repository depth is considered in the formulation
of scenarios (Marcos et al. 2011). For canister design assessment purposes, permafrost is
assumed to extend down to the repository level. The lowest temperature is assumed to be
-5 °C, at lowest. The low temperature of bentonite and water in it may lead to changes in
swelling pressure. This exceptional load phenomenon is assessed in Section 8.4.4.
4.5 External chemical loads
Chemical loads (corrosion) and effect of bacteria may cause some degradation of the leaktight
copper overpack structure in the long period perspective of time. The degradation
processes are assessed in postulated environmental evolution conditions. The chemical
corrosion resistance of the copper canister is discussed in this report in Section 8.9.
4.6 Mechanical load combination
The mechanical loads described above can be grouped as follows. The typical time of
occurrence is noted at the end of each type of load category.
1 Isostatic or asymmetric swelling pressure loads due to incomplete even or uneven
water saturation in Na-bentonite (effective local swelling pressure ~7.8
MPa), possible time of occurrence (0-100 y)
2 Asymmetric loads in saturated Ca-transformed-bentonite due to manufacturing
tolerances of deposition hole and of buffer density (swelling pressure difference
~7.8 MPa) (100 y - glaciation)
3 Groundwater pressure at the depth of repository (4.1 MPa) (100 y - glaciation)
4 Glacial pressure from 2.5 km thick glacier (25 MPa) (glaciation period)
5 Shear load due to rock displacement (5 cm, v=1 m/s) (permafrost and glaciation
period or later)
6 Combination of 2 + 3 (100 y - glaciation)
7 Combination of 2 + 4 (glaciation period)

32
8 Combination of 3 + 5 (before or after glaciation period)
9 Combination of 4 + 5 (glaciation period)
10 Lifting loads in operation phase (canister weight + dynamic extras)
11 Buffer swelling at lowest temperature -5 °C (glaciation period).
The temperature of the canister copper overpack will stay above room temperature
about 10000 years after deposition and then the temperature will slowly go down to
natural temperature of the Olkiluoto rock (10-11°C) within a few thousand years.
Around glaciation, before or after, the canister temperature may be lowered to close to 0
°C, if the cold period is long enough and there is not protecting glaciation or snow on
the ground. As a design assessment exercise, the repository level temperature is assumed
to go down to -5 °C. More details of the temperature evolution of the canister are
given in (Pastina & Hellä 2006, chapter 6.1). A detailed graph of calculated canister
surface temperature is given also in Figure 28 in Section 8.4.3 of this report. It clearly
shows the effect of bentonite buffer saturation rate in early years of the evolution in the
repository. This graph does not include the possible permafrost effects but assumes the
ground surface conditions to be unchanged. In safety case, the onset of the first cold
period is expected at about 50000 years with temperature and precipitation changes
leading to first permafrost development and later on to ice-sheet growth and advance
(Pimenoff et al. 2011).
Load types of 10 and 11 (above) are handled separately without any combinations. The
exceptionally low temperature of the bentonite buffer is lowering the swelling pressure,
thus there is no reason to recalculate any load combinations, for assessment see Section
8.4.4. The appearance and combination of various basic load cases and respective temperature
range are described in Table 2.
The saturation time of the bentonite buffer depends locally on wetting conditions. The
time for full saturation may vary considerably. The expected reference scenario for climate
evolution is given in detail in Chapter 4 of Performance Assessment report.
The maximum insert temperature of the canister may vary a bit depending on the canister
sealing method. If EBW is used, the gap between insert and the copper overpack
may stay in vacuum for some time after welding. The vacuum causes lower heat conductivity
over the gap. If FSW is used, the gap stays is atmospheric gas pressure all
time. In long perspective the conditions will be equal in both canister seal variations due
to gas diffusion through the insert lid gasket. The copper overpack temperature is not
depending on the conditions inside canister, because the very high thermal conductivity
of copper makes the temperature field even in thick copper shell.

33
Table 2. The canister loads extracted from postulated repository evolution. Loads that may act simultaneously shall be combined. Coloured
boxes correspond to the possible periods for appearance of the load case (1…5) in question.
Repository evolution phase
The reference scenario for climate evolution is referenced
in the text in this Section
Water saturation
0 a. – 100 a.
Canister temperature (°C) (EBW-sealed) < 140/100
Load case number Deformation rate (Fe/Cu parts)
1) Asymmetric loads due to uneven
water saturation and imperfections
in deposition hole geometry. No
simultaneous hydrostatic pressure.
Uneven water saturation effects
will decay later and be replaced by
permanent loads 2) and 3) acting in
saturated condition.
2) Permanent asymmetric loads due
to uneven bentonite density and
imperfections in deposition hole
geometry.
3) Groundwater hydrostatic pressure
+ even isostatic swelling pressure
of bentonite.
4) Glacial pressure (additional isostatic
pressure, only during glacial
period).
5) A shear displacement exceeding
5 cm at shear velocity higher than 1
m/s in fractures intersecting a canister.
Insert
Static
Copper overpack
Creep or static
Insert
Static
Copper overpack
Creep or static
Insert
Static
Copper overpack
Creep or static
Insert
Short-time forced
displacement
Copper overpack
Short-time forced
displacement
Water saturation effects
are assumed to reach
their maximum.
Load 1) can create
bending loads
Load 1) can create
compressive loads
Temperate
100 a. – 50 ka.
Permafrost Glacial Subsequent
permafrost and
glacial periods
100/70 > T > 15 15 > T > 0* 15 > T > 0 15 > T > 0
*Disturbance, -5 o
Bending loads from
load 2) and compressive
isostatic loads
from load 3)
Uneven pressure loads
from load 2) and
isostatic loads from
load 3)
Loads 2) and 3)
are expected to
act throughout the
analysis period
Loads 2) and 3)
are expected to
act throughout the
analysis period
Load 4) will
cause additional
isostatic pressure
on insert
Load 4) will
cause additional
isostatic pressure
on shell
Load 5) is of
most concern in
end-glacial
periods
Permafrost and
glacial conditions
are to reappear

35
5 NUCLEAR SAFETY CLASSIFICATION OF THE CANISTER
Systems, structures and components important to safety shall be designed, manufactured,
installed and operated in such a way that their quality level and the inspections and tests
needed to verify their quality level are commensurate with the importance to safety of each
item.
According to the YVL-guide B.2 components shall be assigned to safety class 2, if their
failure to operate would cause a considerable risk of uncontrolled criticality. On the other
hand, components shall be assigned to safety class 3, if their fault or failure would prevent
decay heat removal from spent fuel, or cause the dispersal of radioactive material.
The appendix of the YVL-guide B.2 gives an example that the storage racks for fresh and
spent fuel are assigned to safety class 2 and the storage of spent fuel and liquid wastes,
including pools and tanks, are assigned to safety class 3.
Basing on the rules and examples given in the YVL-guides B.2 and D.3 it can be concluded
that the canister as a whole system shall be classified as safety class 2. The canister
insert as a component shall be classified in safety class 2, because it has a prominent role in
criticality safety. The canister overpack could be classified according the rules in safety
class 3 as a barrier against dispersal of radioactive material. However, the especially long
postulated lifetime of the copper overpack is responded according to requirements of YVL
D.3 by increasing the safety classification up to class 2. The overpack consists of copper
overpack components and the sealing weld.
The steel lid of the insert is also classified in the safety class 2 like the insert body. The lids
are active mechanical load bearing members of the system. However, the fixing screw of
the insert lid is not safety related component and the gaskets of the lid have no requirements
in long term tightness or sub-criticality. These auxiliary components are classified as
safety class 3.
In practice, the safety classification governs the scope of design analyses and the quality
controls of the canister component manufacture.

37
6 CANISTER SHAPE, DIMENSIONS AND SURFACE QUALITY
The canisters size and shape have been derived on one hand from the space needed for
the actual spent fuel assemblies, and on the other hand from the mechanical strength,
radiation shielding and cooling capability. Economic optimisation has led to maximise
the number of the positions for assemblies in the canister and to minimise the size (or
weight) of the canister. Striving for these goals, the canister design has evolved over
several stages to result in the current reference canister design. In the following, the
detailed dimensions of the BWR canister are given in Figures 7 to 9 and Tables 3 to 5.
The dimensioning of the reference design fulfils the design bases given in chapter 3 of
this report. The design of the insert steel lid is still under consideration. The originally
planned gasket construction, a rubber O-ring, see Figure 8, is not acceptable as an organic
material inside canister. The new insert lid proposal is given in Figure 16, but the
non-organic gasket material has not yet been decided.
In Finland, there are three variants of the canister, one for each spent fuel type: BWR,
VVER-440 and EPR/PWR (Figure 1). The spent fuel is sealed in the canisters as whole
fuel assemblies including the fuel channel outside the VVER-440 and BWR fuel elements.
The copper overpack is identical for all three variants with the exception of
length. The cast iron insert for different variants has, accordingly, different length depending
on the actual length of the fuel elements in question. In addition, the fuel elements
have various size and shapes in cross section, too. Thus the various inserts have
different shapes and size openings for the fuel elements. The dimensions of various fuel
element types are given in Section 13.1. The canister variants are shown in Figures 10
to 13 and the main dimensions and masses are given in Table 6.
The sealing method of the copper canister has not yet been decided. The reference solution
is the EB welding and the alternative method is the FS welding. Figure 15 shows
the dimensions of the EB weld and Table 5 gives geometric dimensions for both alternatives.
Also, there are two possible alternatives for the copper overpack manufacture, namely
cylinder with an integral (flat) bottom, or a tube with a welded bottom end. The latter is
made by welding the bottom end likely with FSW method. In this case the bottom lid
will have additional extension on edge area like the top end lid. This makes the weldedbottom-lid
canister variant 75 mm longer than the integrated flat bottom variant. The
main dimensions for both variants are given in Tables 5 and 6.
The Posiva reference design for the canister is the BWR-type canister. The reference
design has an integrated bottom in the insert and in the copper overpack. However, there
are alternative options for manufacture, such as welded bottom for the copper overpack
and three optional manufacturing methods for the hot deformation of the copper tube.
The alternative to the reference sealing method for the copper overpack, the EBW
method, is the FSW method. Furthermore, the reference disposal method, KBS3-V, has
an alternative option of the horizontal deposition, KBS3-H, but this possible option has
no practical effects on canister design or load processes. KBS3-H is generally described
in (Posiva 2008, pages 6-7).

38
The identification of sealed canisters for bookkeeping, quality control and safeguards
purposes will be implemented by visible identification label (serial number) that will be
engraved in clear characters on the frontal surface of the lift shoulder. The size of the
characters will be about 10 mm. The location of the identification label is defined so
that it is easy to detect with cameras and, on the other hand, the engraving does not degrade
the corrosion resistance of the canister. The identification string location is
marked in Figure 15.
Figure 7. Canister insert dimensions. SKB illustration from (Raiko et al. 2010)

39
Table 3. Functional dimensions of BWR, VVER 440 and EPR/PWR canister insert with
tolerances. The dimensions refer as for BWR insert to Figure 7.
Dimension Description Nominal value Tolerance Reason
A
total length with
steel lid
4565 mm
3365 mm
5035 mm
B bottom thickness 60 mm*
70 mm*
85 mm*
C inside length 4450 mm
3245 mm
4900 mm
+0/-0.5 mm
+0/-0.5 mm
+0/-0.5 mm
+10.1/-5.6 mm
+10.1/-5.6 mm
+10.1/-5.6 mm
+5/-10 mm
+5/-10 mm
+5/-10 mm
gap for thermal
expansion
strength
fuel geometry
D outer diameter 949 mm +0.5/-0 mm strength, thermal
expansion
H neck thickness 33.3 mm
45.6 mm
50 mm
I outer radius 20 mm
(D193.7 mm)
25 mm
J
calculated distance
210 mm
210 mm
360 mm
K thickness 30 mm
16.2 mm
100 mm
L inner width**** 160 x 160 mm
D173.7 mm
235 x 235 mm
M profile thickness 10 mm
10 mm
12.5 mm
+10/-10 mm**
+10/-10 mm**
+10/-10 mm**
+5/-5 mm
-
+5/-5 mm
+1/-4 mm
+1/-4 mm
+3.6/-3.6 mm
+2.7/-4.6 mm
+2.7/-4.6 mm
+5/-5 mm
+3.8/-3.8 mm
gauge D168 mm
+5.1/-5.1 mm
+1/-1 mm
+1/-1 mm
+1.25/-1.25 mm
strength
strength
sub-criticality
strength, subcriticality
fuel geometry
stiffness
N lifting eye hole M48*** - 2 lifting eyes
*) The total bottom thickness is the sum of cast iron thickness and the steel cassette bottom
plate thickness.
**) Local tolerance. The eccentricity (variation between opposite neck average thicknesses) of
the structure may be +5/-5 mm.
***) Threaded lifting eye hole (the size is metric thread with diameter 48 mm).
****) The width and straightness of the openings are gauged, respectively, with 152 mm,
Ø168 mm and 224 mm full length gauge.

40
Figure 8. Principal scheme of the steel lid of the insert and the dimensions. This is a
SKB illustration from (Raiko et al. 2010). The lid and gasket system is under reconsideration.
See Figure 16 for a Posiva design draft.
Table 4. Functional dimensions of canister insert lid with tolerances. The insert lid gasket
system is still under re-consideration.
Dimension Description Nominal value Tolerance Reason
E diameter 910 mm +0/-0.09 mm assembly
F thickness 50 mm +0.1/-0.1 mm strength
G bevel angle 5° - assembly

41
Figure 9. The copper overpack of the canister. SKB illustration from (Raiko et al.
2010). Manufacturing and functional dimensions of the FSW variant of copper overpack.
The dimensions of EBW version are the same, but the weld orientation differs.
Table 5. Functional dimensions of the canister overpack with tolerances. The dimensions
refer to Figure 9. Dimensions are given for BWR, VVER 440 and EPR/PWR variants,
respectively, if they are not identical. Dimensions for FSW variation are given,
too.
Dimension Description Nominal value Tolerance Reason
A total length 4752 mm
3552 mm
5223 mm
+3.25/-2.75 mm
+3.25/-2.75 mm
+3/-3 mm
welded bottom
variant is 75 mm
longer
t wall thickness 49 mm +0.85/-0.85 mm corrosion
B outer diameter 1050 mm +1.2/-1.2 mm strength
C FSW inner diameter 850 mm +0.8/-0.8 mm FSW welded
bottom only
E inner diameter 952 mm +0.5/-0.5 mm strength
F inner diameter 821 mm +0/-0.5 mm strength
G inner diameter 850 mm +0.8/-0.8 mm strength
H EBW (lid/pipe) diameter 960 mm +0.3/0, +0.4/0.1 clearance fit
H FSW (lid/pipe) diameter 953 mm d8/H8 clearance fit
I corner radius 10 mm - stress
K dimension 35 mm +0.5/-0.5 mm strength
L dimension 50 mm +0.2/-0.2 mm gripper space

43
Table 5. Continued from preceding page.
Dimension Description Nominal value Tolerance Reason
M thickness 50 mm +0.6/-0.6 mm corrosion
N FSW FSW position 60 mm - FSW variant
N EBW EBW shoulder 50 mm - EBW variant
calculated inner free length 4567 mm
3367 mm
4900 mm
+0.6/-0.1 mm
+0.6/-0.1 mm
+0.6/-0.1 mm
fuel geometry
calculated
calculated
axial gap between
lids
radial gap between
cylinders
2 mm
2 mm
2.5 mm
1.5 mm
1.5 mm
1.5 mm
+1.1/-0.1mm
+1.1/-0.1mm
+1.1/-0.1mm
+0.25/-0.5 mm
+0.25/-0.5 mm
+0.25/-0.5 mm
thermal expansion
allowance
installation,
thermal expansion
allowance
Table 6. Main dimensions and masses of canisters for different types of spent fuel.
Loviisa 1-2
(VVER-440)
Olkiluoto 1-2
(BWR)
Olkiluoto 3
(EPR/PWR)
Outer diameter (m) 1.05 1.05 1.05
Height with flat bottom end* (m) 3.552 4.752 5.223
Thickness of copper cylinder, nominal, (mm) 49 49 49
Thickness of copper lid and bottom, nominal, (mm) 50 50 50
Thickness of iron insert bottom**, nominal, (mm) 70 60 85
Total volume of canister* (m 3 ) 3.03 4.07 4.47
Total area of canister outside surface *) (m 2 ) 13.67 17.63 19.18
Void space with fuel assemblies (m 3 ) 0.61 0.95 0.67
Number of fuel assemblies 12 12 4
Amount of spent fuel (tU) 1.4 2.2 2.1
Mass of fuel assemblies (ton) 2.6 3.6 3.2
Mass of iron (ton) 8.6 10.6 15.8
Mass of steel (ton) 2.0 3.0 2.1
Mass of copper* (ton) 5.6 7.3 8.0
Total canister mass *) , gross, (ton) 18.8 24.5 29.0***
*) If the welded bottom lid alternative is used for the copper overpack, then the total length
increases +75 mm, the total canister volume +0.024 m 3 , the total surface area +0.45 m 2 , and
the copper mass and the total canister mass +0.21 ton.
**) The total bottom thickness is the sum of cast iron thickness and the steel cassette bottom
plate thickness.
***) The possible effect of control rod absorbers (about 55 kg per element) is not included.

44
The surface roughness requirement for the weld preparation of EBW between the copper
lid and the cylinder is Ra = 3.2 m. All other machined outside surfaces of the copper
overpack have the surface requirement Ra = 6.3 m and inside surfaces 12.5 m. The
surface requirement for the alternative sealing method FSW surfaces is the same, Ra =
3.2 µm.
The machined surfaces of the cast iron insert are generally the same Ra = 6.3 m, with
the exception that the gasket surfaces of the steel lid gasket at the top part of the insert
and at the edge of the steel lid shall be of better surface quality. The required higher
surface quality of the cast iron insert in the gasket surface may be achieved by local
surface plating. The gasket design is still going on.
The insert is made of nodular graphite cast iron in one piece. The positions for fuel assemblies
are holes, which are dimensioned and formed either for BWR, VVER-440, or
EPR/PWR fuel assemblies. The inserts have an integral flat bottom of 60 mm cast iron for
BWR and VVER-440 type and 85 mm for the EPR/PWR type. Bottom thickness is higher
in EPR/PWR insert because of the need for increased strength for wider openings. The
total bottom thickness is the sum of cast iron thickness and the steel rack bottom plate
thickness.
All insert types have loose flat lids made of 50 mm steel plate on top ends. The top lids are
fixed centrally with 1 screw (size M30). Gaskets tighten the gaps between the lid edge and
the insert body and around the central screw. The gasket is proposed to be made of some
soft metallic material, such as indium. The purpose for the gasket is to keep the inert gas
inside the insert during the electron beam welding of the copper lid. The EB-weld is made
in vacuum. A manufactured demonstration canister is shown in Figure 10. The sections of
the insert types for various fuel assemblies are shown in Figures 11, 12 and 13. The canister
assembly, EB welded lid and the insert lid arrangement are shown in Figures 14 to 16.
Figure 10. A full-scale demonstration of a BWR canister (Posiva Oy/Jussi Partanen).

45
~ 33
160
160
210
949
A
210
50
A
Profiles: square tube 180x180x10
BWR-type
Figure 11. The section of the insert for BWR fuel assemblies.
~ 46
D174
210
949
A
210
36
A
Profiles: round tube 193.7x10
VVER 440-type
Figure 12. The section of the insert for VVER-440 fuel assemblies.

46
~ 50
949
A
235
360
125
A
235
360
Profiles: square tube 260x260x12.5
EPR-type
Figure 13. The section of the insert for EPR/PWR fuel assemblies. The sectional dimensions
are not identical with the SKB PWR insert section.
Figure 14. An exploded view of the BWR-canister main components from left; copper
overpack, cast iron insert, steel lid and copper lid. Illustration by Afore Oy.

47
D830
EBW
Chamfering 20 deg
D1050
Engraved label
D821
85
50
R3
D850
R10
Figure 15. Canister variant with EB welded copper lid. The dimensions reflect the geometry
before welding deformations. Also the location of the engraved identification
label is shown in the figure.
Figure 16. Detail of canister with EB weld sealed lid and metal gasket in the insert lid.
This is a draft drawing of new proposed gasket design by Optimik Oy. Fitting dimension
between copper lid and cylinder is the installation-time dimension; the shrinkage of
EBW will close the gap between cylinder and lid during welding. The axial gap
dimension is 2.4-3.6 mm for EPR/PWR type canister.

49
7 MECHANICAL AND PHYSICAL PROPERTIES OF
THE CANISTERCOMPONENT MATERIALS
7.1 Material qualities involved
Disposal canisters consist of two main components: the outer shell, made of oxygenfree
copper and the insert, mainly made of nodular graphite cast iron. The insert also
contains some steel parts like the cassette tubes, the lid and the lid fixing screw.
The canister materials and the manufacturing processes are being developed to produce
a canister that can fulfil the design bases given in chapter 3. The material of the copper
overpack is oxygen-free high conductivity copper (Cu-OF) with an addition of 30-100
ppm of phosphorus. The micro-alloying improves the creep strain properties of Cu-OF
especially in high temperature (200 to 300 C). Concurrently the alloying lowers the
risk of cracking during hot-deformation process.
The copper specification is given in specification KTS001 (Nolvi 2009) and the main
features are as follows: the material for copper canisters shall fulfil the specification in
EN 1976:1998 for the grades Cu-OFE or Cu-OF1 with the following additional requirements:
O

50
(with longitudinal weld) or hot-formed steel (seamless tubes). The material for the hotformed
rectangular hollow sections (RHS) fulfils the requirements in EN 10210-1 grade
S355J2H concerning chemical composition and mechanical properties (R eL , R m , A 5 ).
The material for the alternative cold-formed RHS sections fulfils the requirements in
EN 10219-1 grade S355J2H concerning chemical composition and mechanical properties
(R eL , R m , A 5 ). The material for steel plates and flat bars of the cassette assembly
shall fulfil the requirements in EN 10025 grade S235J2 or similar. The steel qualities
and properties in cassette are provided formally in the specification KTS022. The specifications
mentioned above are described in the manufacturing report of (Nolvi 2009).
The specifications are updated continuously to better reflect the requirements. The applicability
of the material standards referenced is limited to the current version. If the
standards are changed in future, the applicability shall be re-assessed.
7.2 Mechanical properties
The mechanical properties of the structural materials to be used in the design analyses
are based on a large amount of data from demonstration-manufacture tests. Thus the set
values reflect the actual material properties more than standard values. To handle the
scatter in data, basic statistical measures have been used. The lower 90 % confidence
value is used as a reference value, when appropriate. This ensures that material properties
are realistic and can be attained during production. However, for the stress-strain
relationship of cast iron, the average values taken from the test data are used for consistency.
In different type of loading analyses, the conservative stress-strain relationship
may be either the higher or the lower curve. A better strategy is to use the average value
and add the required safety margin in the final results, as opposed to adding it to all of
the input data in the calculations.
It should be pointed out that values used in the calculations may differ from what is
stated in the materials specifications. For instance, the critical fracture parameter for the
postulated defect is not specified at all and the elongation values are higher. The values
used in the analysis are not the lowest values defined in the specifications but the statistical
values achieved in the series of test manufacturing. The full background of the
used data is given in (Raiko et al. 2010).
The design basis mechanical load cases, glacial isostatic pressure and/or rock slide shear
deformation take place at temperatures that may be between 0 °C and +20 °C. Operational
handling or transfer loads and asymmetric bentonite swelling loads take place at
ambient temperature range, from +20 °C to +100 °C that will be typically also the copper
overpack temperature. During the early years after disposal the canister insert may
be, at maximum, at some 45 °C higher temperature than the copper overpack, in other
words, at +140 °C, for estimation, see Section 8.4.2. If FSW sealing is used instead of
EBW, the maximum insert temperature will be lower, some 100 °C according to calculation
given in (Ikonen 2006, figures 15 and 23). After water saturation of the bentonite
buffer, the canister temperature will decrease markedly (about 15 °C) from the calculated
maximum temperature due to the improved heat transfer between the canister surface
and the rock.

51
The mechanical properties of the structural materials of the insert are given in Tables 7
to 10 for cast iron, steel in the lid, as well as cold formed and hot formed steel for the
cassette hollow sections. Standard values for yield strength and ultimate strength are
taken to mean the minimum required material test values based on standard-type uniaxial
tension tests (EN 10002-1) with round samples and low strain rate tension tests.
From the mechanical point of view, the most demanding load cases are the isostatic
pressure under glacial period and the rock shear deformation. The strength values referenced
in the tables below are mainly based on either tension or compression tests depending
on load case type. The ultimate strength or the given stress/strain relationship
is converted from uniaxial test result to a true-strain/true stress relation. The test result
details are published in (Claesson 2009; Öberg & Öberg 2009b; Minnebo & Mendes
2004; Öberg & Öberg 2009a) and they provide the bases of the given design values of
the material properties.
Cast iron was widely investigated in conjunction with a probabilistic pressure test programme
(Nilsson et al. 2005; Martin et al. 2009; Ikonen 2005). A representative stressstrain
curve was then selected based on compressive testing. The curve and its origin are
given in (Raiko et al. 2010). For consistency, in the analyses the same compression data
set is used in all reports using compression data for the cast iron.
However, as concluded in (Raiko et al. 2010) this is a pessimistic approach since more
recent data from manufactured inserts have shown better values both in strength and
ductility. The selected curve is used anyway to model the strength of the insert material.
Tension data used in the analyses, mainly in rock shear analyses, are represented by the
stress-strain data from series of stress-strain tests made both in +21 °C and in 0 °C using
standard and elevated strain rate. The used data are taken from the strain-rate-dependent
testing of cast iron. The rock shear analysis used a strain-rate dependent material model,
so the material stress-strain curve was presented for static and for strain rate 0.5 s -1 case.
The base for selection of the strain rate 0.5 s -1 comes from the fact that the rock shear of
5 cm takes 0.05 s at 1 m/s shear velocity and the maximum strain appears to be some
2 % in cast iron. Thus the average maximum strain rate is 0.02/0.05 s = 0.4 s -1 . In rock
shear analysis, the instantaneous strain rate is then used for linear interpolation between
the static and dynamic strength.
Comparison to the data from the serial manufacturing tests shows good agreement between
the representation and actual measured values. This comparison is reported in
(Raiko et al. 2010), Section 4.1.1, and in its references.
Based on results from standard type testing of material from top sections of SKB inserts
I53-I57 the elongation at fracture for BWR-inserts can be evaluated and the 90 % confidence
interval calculated according to the methodology in summary report of materials
testing data for insert iron, (Dillström & Bolinder 2010b). This gives elongation at rupture
12.6 % < A5 < 14.8 % with 90 % confidence. The engineering stresses are converted
to true stresses thus they can be used directly as a material model for largedeformation
and large-strains analyses made by FE-methods. All stress-strain model
curves used for cast iron are given in Table 7.

52
The fracture toughness data measured for cast iron at 0 °C are 88.1 kN/m ≤ J 2 mm ≤ 93.5
kN/m, when expressed as a J-integral value. The J-integral is used to calculate the fracture
parameter of a postulated defect and then the parameter value is compared to the
measured fracture resistance. The J 2 mm means the J integral value corresponding crack
growth of 2 mm in the fracture resistance curve of the material. The number used for the
damage tolerance analyses (88 kN/m) for the rock shear case is taken from the measured
data with 90 % confidence. See reference (Dillström & Bolinder 2010b) for further information
about the evaluation of fracture toughness data measured for cast iron. The
measured fracture toughness data are in more detail presented in fracture mechanics
testing report (Öberg & Öberg 2009b). Recent fracture resistance measurements for cast
iron at 0 °C temperature made at VTT gave considerably higher results for J 2 mm . Now
the test samples were taken from the critical location, the insert cylindrical surface, the
test sample was larger, a 25 mm thick CT-specimen, and the sample orientation was
longitudinal and the initial crack was radially oriented from the cylindrical surface of
the insert. Typical J 2 mm -values measured were between 130 to 150 kN/m (Planman
2012). Using these thicker test samples, the actual fracture resistance could actually be
measured up to more than 2 mm crack growth.
In the case of the external pressure load, when the load controls the stresses and causes
primary stresses, the damage tolerance analysis is made using K Ic data that are based on
crack initiation, not for limited crack growth like J 2 mm . The fracture toughness K Ic data
measured for cast iron at 0 °C are 78.0 MPa√m with 90 % confidence, when declared as
stress intensity factor K Ic value. K Ic = 78 MPa√m is equivalent to J c = 33 kN/m.
Iron is being investigated by SKB under long-term loading conditions up to +125 °C
and preliminary tests only show a creep behaviour that is logarithmic in nature. The
preliminary tests show that the creep strain after long times, even at stress levels close
to the yield of the material, is likely to be small or negligible at all tested temperatures.
This is why the creep phenomenon in cast iron in repository condition is omitted in the
following mechanical analyses. Reporting of creep testing for cast iron has been published
(Martinsson et al. 2010). It is essential to notice that the higher temperature in
insert in the beginning of evolution is possible only in dry conditions of the buffer,
which means that the mechanical loads are not loading the canister at the time.
According to the specification given in SKB design premises report (SKB 2009), Section
3.1.5, the P content should be in the interval 30 to 100 ppm to ensure sufficient
creep ductility. The typical values for mechanical properties that are given are 69 MPa
for the yield strength and 220 MPa for the tensile strength in soft condition, for the
product form tube 25×1.7 mm and a grain size of 50 m. Since the canister material
will have a coarser grain size, its mechanical properties can be expected to have somewhat
lower values. In the canister production tests, yield strength values from 40 to 75
MPa have been obtained at normal strain rates for tensile testing (Sandström et al.
2009).
In the FEM-computations, model values for stress/strain curves have been used. The
model is described in Section 7.3. The elastic modulus used in these investigations is
120 GPa, which is an average value for pure copper, whose elastic modulus can vary
between 115 to 128 GPa (Metals Handbook 1990). The Poisson's ratio used was 0.308
according to (Metals Handbook 1990).

53
Table 7. Mechanical properties of the cast iron EN-GJS-400-15U (Raiko et al. 2010).
Property
Yield strength in tension
[MPa]
Yield strength in compression
[MPa]
Standard value
(EN 1563:2010)
Values in static
analyses
≥240 True stress [MPa] /
true strain [%]
0/0
267/0.1608
330/1.998
366/4.000
392/6.000
427/9.998
456/15.005
480/49.990
≥240 True stress [MPa] /
true strain [%]
0/0
270/0.1627
333/2
394/4
429/6
482/10
534/20
550/50
550/100
Values in dynamic shear
case at 0 °C and at strain
rate 0/s*
True stress [MPa] / plastic
strain [%]
293/0
324/1
349/2
370/3
389/4
404/5
418/6
428/7
438/8
-
447/9
456/10
465/11
472/12
478/13
484/14
488/15
491/16
Ultimate strength
[MPa] (in tension /
compression)
Elongation at failure
[%]
Fracture toughness, J-
integral J c [kN/m]
Fracture toughness, J-
integral J 2 mm [kN/m]
≥370 456/534 456/534
11 % (from samples
on casting)
12.6

54
Table 8. Mechanical properties for the lid steel EN 10025 S355J2 (t= 40-63 mm).
Property
Standard value
(EN 10025:2004)
Engineering stress*
[MPa] / strain [%]
Yield strength [MPa] ≥335 0/0
335/0.1595
470/15
470/20
Ultimate strength
[MPa]
Elongation at failure 19
[%]
Young’s modulus E
[GPa]
Poisson’s ratio ν
[-]
490-630 470 564
210 210 210
0.3 0.3 0.3
True stress* [MPa] /
logarithmic strain
[%]
0/0
335/0.1593
540/13.98
564/18.2
*) Engineering stress assumes that the area a force is acting upon remains constant, true stress
takes into account the variation in the cross sectional area as a result of the stress induced deformation
(strain) of a material. Stress-strain results from 1-D measurements are usually reported
as engineering stress whereas the true stress relation is used for the 3-D finite element
calculations.
Table 9. Mechanical properties for the cold-formed tube steel EN 10219-1 S355J2H
(thickness t=

55
Table 10. Mechanical properties for the hot-formed * tube steel EN 10210-1 S355J2H
(thickness t=

56
Figure 17. Slow rate tensile tests for cold worked Cu-OFP at 75ºC, strain rate 0.0001
s -1 , and at 20ºC, 0.001 s -1 . Note different scales (Yao & Sandström 2000).
The numerical modelling of copper stress-strain behaviour has been presented thoroughly
in (Sandström et al. 2009) and summarised in (Raiko et al. 2010). The temperature
dependency and the deformation rate dependency are included in the modelling.
7.3.3 Copper creep model
In the computations of creep deformation in the canister, models for the stationary and
non-stationary creep rate have been used (Sandström & Andersson 2008; Sandström &
Andersson 2007). These models are described in the present Section. At high temperatures
above half of its melting point Tm, climb is believed to control the deformation in
many types of metals including copper. For copper the half melting temperature (T m /2)
is about 400 ºC.
It is believed that the deformation is glide controlled at lower temperatures. A difficulty
with expressions for glide controlled deformation is that the values of the constants are
not known. However, there are similarities between the equations for climb and glide
control and such equations were combined into a unified model
3
3
b
Q
RT
1
(
2bc
D 0b
L s L
k
max
e BT
e
i
OFP
/ fP
h(
)
(1)
m k T Gb
B
2
)
The interpretation and values of the parameters in eq. (1) are given in Table 11.

57
Table 11. Values of constants used in the model in equation (1).
Parameter description Parameter Value
Burgers vector b 2.56·10 -10 m
Taylor factor m 3.06
Boltzmann’s constant k B 1.381·10 -23 J/grad
Shear modulus G G 4.75 10
4 17 T MPa, T in K
Dislocation line tension L 7.94·10 -16 MN at RT
Coefficient for self-diffusion D s0 1.31·10 -5 m 2 /s
Activation energy for self-diffusion Q 198000 J/mol
Strain hardening constant c L
57
Constant 0.19
Max back stress imax 257 MPa
Influence of phosphorus f P
3000 for T < 125ºC
Time at the start of primary creep t init 1 h
Time at minimum creep rate t min t R
/3, where t R
is the rupture time
Parameter in g rate 2
13.26 0.022T
, T in K
Omega c 0.45
Ratio between initial and stationary creep
rate
grate 2
/(12
)
( t min / tinit
)
Equation (1) is compared to experimental data in Figure 18.
In general an acceptable agreement is obtained between the model and the observation.
The difficulty is in the transition between 175 and 215 ºC, where the experiments show
a sharp transition in slope whereas the model transition is more gradual.
The numerical modelling of copper creep behaviour has been presented thoroughly in
(Andersson-Östling & Sandström 2009) and summarised in (Raiko et al. 2010).

58
Figure 18. Comparison of eq. (1) to creep data and slow strain rate (SSRTT) data Cu-
OFP (Andersson-Östling & Sandström 2009).
7.3.4 The copper creep model for multiaxial stress states
The traditional way to transform a uniaxial creep model to multiaxial stress state is described
in (Sandström & Andersson 2008). When this way was used in FEMcomputations
some difficulties appeared. As a consequence three new approaches were
developed and used (Jin & Sandström 2009) and (Sandström & Jin 2009). A summary
is given in (Andersson-Östling & Sandström 2009) and in (Raiko et al. 2010). The three
approaches give essentially the same results. According to one of the approaches the
starting equation is
p
deff
dt
h((
e
i)
e
p
eff
) grate
h(
ee
p
where eff
is the effective strain and e the effective stress. With the help of Odqvist's
equation, the individual components of the creep rate are obtained.
p
dε
dt
3
2
p
eff
d
dt
σ'
e
' is the deviatoric part of the stress tensor . The back stress is now a scalar. It can be
derived from the following equation, identifying the analogue with the uniaxial case
i
p
p
eff
c
(1 eff
i max e )
(4)
)
(2)
(3)

59
7.3.5 Comparison to copper creep tests for notched specimens
The creep lifetime under multiaxial stresses for notched round bars has proven to be
much longer than that for uniaxial specimens (Wu et al. 2009). This demonstrates notch
strengthening for the Cu-OFP material. If the rupture curves are extrapolated, the notch
strengthening factor in time is greater than 100. Metallographic examination has shown
that only limited number of pores and cavities are observed in ruptured specimens. This
demonstrates that the local creep ductility is high.
Comparison to finite element modelling is illustrated in Figure 19. The -model in eq.
(2) and the basic model for primary creep are used.
Using the basic model, the observed strain is somewhat overestimated. In particular the
initial strain is overestimated. With the model, eq. (2), on the other hand the strains
are underestimated by a factor of three. Considering that the time difference to the uniaxial
test results is more than a factor of 100, the comparison between the experiments
and the simulation must be considered as satisfactory. These results give some validation
of the multiaxial model formulation.
Later additional results for copper creep cracking and multiaxial phenomenon have been
reported in (Wu et al. 2011). This discusses about the big difference in copper creep
behaviour between the canister operational temperature (20 to 75 °C) and the elevated
creep test temperature (175 to 225 °C).
Figure 19. Comparison between experimental data and FEM results for a notched creep specimen
under a net section stress a) 215 MPa and b) 200 MPa. For the model marked SSR creep,
the basic model was used for primary and secondary creep. For the curves marked -model,
eq. (2) was used. The initial strain on loading is included in the experimental data (Andersson-
Östling & Sandström 2009).
7.3.6 Posiva’s copper creep testing and canister evaluation
Posiva has had copper creep investigations and creep tests at VTT for both base material
and for EB-welds in copper. These studies are mainly made to complete the test pro-

60
gramme executed in Sweden by SKB. The supplement testing has focused on EB-weld
material, as SKB has, instead, concentrated on testing FS-weld.
EB-weld material differs from hot-deformed copper base material as for grain size,
yield strength and creep properties. EB-weld material is in cast condition, the grains are
of millimetre size, but the chemical material contents are the same as in base material.
In EB-welding, no filler or any other additive material is used and the melting during
welding is made in inert condition, high vacuum.
Creep testing of the copper EB-welds has continued for years. Some preliminary results
are presented in (Holmström et al. 2012a). These creep tests have been accelerated by
elevated temperature. The test temperature has been usually 125 or 175 °C. So far all
the results (loading stress and resulting strain) has been presented only in engineering
measures, as A 5 strains and nominal stress.
If the results are converted to true stress and true (and local) strain, then the numbers
and models are changed drastically better. The reason of the misleading results of engineering
(standard) results is that for ex. specimens containing an EB-weld that is 5 to 8
mm wide is measured typically from points that are 50 mm apart and thus we get an
average strain for 50 mm length even if the most deformation takes place in the weld
that is only roughly one tenth of the measuring length. Another reason for misleading
result is the high ductility of copper. The standard type measurement of A 5 elongation in
several tensile tests has given 26 to 30 % elongation for the weld (see table 4 in Holmström
et al. 2012a), but when the local true strain is calculated from respective registered
reduction of area Z, 80 to 85 %, the true strain at rupture is true = ln (1/(1-Z)) =
160 to 190 %, respectively. True stresses are calculated from nominal stress by dividing
with relative necking area σ true = σ nominal /(1-Z). Thus, in this typical case the measured
engineering measures differ from the theoretically more exact measures by factor 6.
This is why the creep test result presentation and creep modelling should be made in
true-stress/true-strain space.
The creep test results that have yielded to rupture so far according to table 5 of (Holmström
et al. 2012a) are as follows in Table 12. The true stress and strain values at rupture
are calculated from reduction of area and engineering measures and added into the
table. Additional tests are going on with lower temperature and stress levels.
The canister structure is usually analysed for creep with finite element method using
large-deformation and large-strain modelling, which procedure leads to true strain and
true stress results. Thus it is essential that the respective material testing results are presented
and material modelling is made in true-stress and true-strain system to allow reasonable
comparison between them.

61
Table 12. Preliminary creep test results from the EB-weld in copper. True stresses and
strains are calculated at rupture from area reduction Z and engineering measures.
Specimen
ident
Nominal
stress σ
Time
Temperature
T
Engineering
strain f
Area
reduction
Z
True
stress
σ true
True
strain
e true
(°C) (MPa) (h) (%) (%) (MPa) (%)
8J 175 125 194 23 67 379 111
9J 175 120 471 25 59 305 89
9N 175 115 844 20 55 256 80
8N 175 100 4405 15 52 208 73
8I 175 100 4656 16 44 179 58
8C 225 95 210 10 34 144 42
If the external pressure load and the ambient temperature are high enough, a plastic deformation
will then take place in the copper overpack and the gap between the canister
and the insert will gradually be closed. In the beginning, creeping starts in the locations
where the existing stresses are higher due to geometric concentrations in structural discontinuities
or due to residual stresses. Both of these are typical secondary stresses and
the peak stresses are relaxed first. When the gap between the shell and the insert is
closed due to plastic or primary creep deformation, the deformation and strains stop to
grow and the remaining stresses continue to relax until they reach equilibrium without
causing additional strain. The creeping analyses are usually made with the most conservative
assumption on both temperature and load, in other words T=75 °C and p=15 MPa
are used. If one or the other of them is lower, then the creeping is consequently much
slower than the modelled condition.
Several creep analyses for the canister construction have been recently published. The
creeping time until the overpack contact against canister insert may vary many orders of
magnitude depending on assumed temperature, load level and creep model. However,
the maximum creep strain is always close to a constant number that only depends on
shape of geometric concentrations and size of the structural gaps. One of the latest
analyses of the structural creeping of the canister is discussed in (Sandström & Jin
2009). The analysis assumed 75 °C temperature and 15 MPa external pressure load. The
result was that the gap is closed in 10 years and that the highest creep strain (including
primary creep) was 10.6 % in a geometric concentration of a rounding in the copper lid
fillet, see Figure 20. The creep strain was

62
Figure 20. A sketch of the deformed shape of the canister overpack lid area during the
external pressure induced deformation and the location of the lid fillet (R10 mm), where
the highest peak stresses and strains exist. Another hot spot area is the EB-weld root
that is notch-like.
Updated creep analysis with updated geometric constraints has been made lately, and
continuation of creep analyses is included in Posiva’s further examinations’ programme.
This analysis, using new copper creep data and VTT’s creep model for copper is made for
canister overpack creep deformation and strain simulation and it is reported in (Holmström
et al. 2012b). The essential result is that the primary creep deformation takes place immediately
after the pressure load is applied and the essential gaps between copper overpack
and iron insert become into contact. The primary creep strains are generally

63
means that the effect of initial eccentricity of insert is negligible in comparison to the
forces causing plastic deformation in the copper overpack. In practise, the symmetric
plastic deformation (hourglass shape) has been observed in the retrieval test in Äspö,
where a real size canister was disposed in simulated repository conditions for several
years before retrieval.
The external pressure load may later in the canister evolution become even higher, but
because of the contact between copper overpack and the solid cast iron insert, the creep
deformation of the copper overpack cannot proceed farther.
Creeping of copper overpack of the disposal canister is limited to certain maximum by
the geometric constraints of the canister design. The increased rounding radius lowers
the peak stress/strain concentration in the corner of the lid and the controlled gap dimensions
between the insert and overpack lower the global deformation of the overpack
in a way that the maximum primary plus secondary creep strain in the overpack will be
limited to a few per cent. This amount of strain is assessed to be acceptable with respect
to the measured creep strength in respective conditions.
The FEA simulation (Holmström et al. 2012b) on canister overpack creeping includes
also cases, where residual stresses were modelled in the EB-weld area. The residual
stresses had no practical effect on the results. Additional analyses and research on copper
creep are still going on. The dependence of the gap dimension on the actual temperature
will be included into the FEA model, and the bimetallic behaviour of the ironcopper
structure, as well, when the plastically deformed structure is cooling down to
environmental temperature in long run.
In case of rock shear, an analysis is made on how high the creep can be in the copper
overpack after a rock shear due to residual load of the buffer. The analysis showed that
the additional creep strain after the instantaneous plastic deformation could be only
about 2 % and thus acceptable. See details of the analysis result in Hernelind (2010) in
Figure 9-18 on page 56.
7.4 Bentonite material model
As all external mechanical loads are transferred through the bentonite buffer to the canister,
the material properties of bentonite define important conditions for the design
analysis of the canister. Table 13 gives an overview of which are the dominating bentonite
properties in different load cases. For more information about bentonite data used
in the analyses see (Börgesson et al. 2010).
For FEM-analyses with the code ABAQUS the bentonite buffer is modelled with an
elastic-plastic material model. The swelling pressure and the yield strength of the saturated
bentonite strongly depend on the density of the bentonite. There are also different
strength and swelling pressure estimates for Na and Ca bentonites.
The bentonite material model is based on laboratory testing and it is essentially different
from material models for metals, because the stiffness and strength of bentonite depends
strongly on the swelling pressure, which in turn depends on the density. Bentonite has

64
two roles; it is a swelling pressure load generating media and, on the other hand, it is a
supporting and flexible material that mitigates the effect of rock shear on canister.
The resulting stress-strain relations are shown in Figure 21. The design basis material
model that is used in the calculations in (Hernelind 2010) is the model of calcium converted
MX-80 at water saturation and at the density ρ m =2050 kg/m 3 . The shear strength
of bentonite is also rate dependent.
The FE-model for the canister rock-shear case uses the elastic-plastic bentonite material
model only on compression state. This is secured by using contact elements between the
bentonite and canister interface. The contact at the material interface is opened, when
the stress component perpendicular to the interface becomes in tension.
Figure 21. Design basis strain-rate dependent stress-strain relation for the calcium
converted MX-80 buffer material with maximum density (Hernelind 2010).

65
Table 13. Overview of dominating bentonite properties for different load cases.
Loads
1. Asymmetric loads due to uneven water saturation and imperfections
in deposition hole geometry. No simultaneous
hydrostatic pressure. Uneven water saturation effects will
decay later and be replaced by permanent loads 2) and 3)
acting in saturated condition.
2. Permanent asymmetric loads due to uneven bentonite density
and imperfections in deposition hole geometry.
4. Glacial pressure (additional isostatic pressure, only during
glacial period).
5. Shear load due to rock displacement. Amplitude is 5 cm,
shear velocity 1 m/s.
Dominating property of
bentonite
Dry density, water absorption
rate, degree of water
saturation, swelling pressure.
Dry density, swelling pressure,
pore water pressure.
Dry density, swelling pressure,
pore water pressure.
Young’s modulus, strainrate
dependent material
model, von Mises stress at
failure.
7.5 Physical properties of canister materials
Physical properties for canister materials are used in thermal conduction analyses (stationary
or transient temperature), thermal-mechanical analyses (thermal deformation),
mechanical stress and thermal stress analyses. Data in Table 14 are typical for ductile
iron, structural steel and high-conductivity oxygen-free copper in room temperature.
They are from a steel product guide (Rautaruukki 1996, page 258), and from Copper
Development Association and Ductile Iron Society data sheets that are available on
internet. However, Young’s moduli in Table 14 are based on material testing results of
actual SKB canister materials.
Material properties are, at least in principle, functions of temperature. In canister design
calculations, however, the operating temperature range is narrow (roughly from 0 to 100
°C) and in long term the temperature is close to room temperature, so no temperature
dependency is modelled. The material properties are selected, however, conservatively,
as for temperature. In thermal expansion and creep analyses the actual temperature is, of
course, modelled as in reality.

66
Table 14. Physical properties of canister materials.
Material
Young’s
modulus
[GPa]
Cast iron 166*
(162-
170)
Structural
steel
210*
(206)
Copper 114*
(117)
Poisson’s
ratio
[-]
0.32*
(0.275)
Density
[kg/m 3 ]
7200****
(7100)
Thermal
conductivity
[W/mK]
Specific
heat
[J/kgK]
Thermal
expansion
[10 -6 K -1 ]
Reference
36 461 11.5 Rio
Tinto**
0.3 7850 52…63 500 12 (Rautaruukki
1996)
0.308* 8940 391 394 16.9 CDA***
(0.35)
*) Values used and rationalized in (Claesson 2009) and (Sandström & Andersson 2009).
**) Ductile iron data for design engineers. 1990. Rio Tinto Iron & Titanium Inc. Montreal,
Canada.
***) Copper data according to oxygen-free copper quality C10100 properties in Copper Development
Association data.
****) Information from abated SFS 3345 standard.

67
8 VERIFICATION OF CANISTER DIMENSIONING
8.1 Mechanical failure processes
8.1.1 Copper overpack
The following mechanical failure modes of the copper and should be considered under
the conditions given in Table 2:
• Fracture due to excessive plastic deformation.
• Rupture due to creep deformation.
The following potential failure processes are excluded:
• Brittle failure.
Cu-OFP is so ductile that unstable crack growth is not relevant at repository temperatures.
The fracture mechanics tests made on oxygen-free copper showed that the cracks
in the test specimen are blunted but not growing (Wells 2008). Unstable crack growth is
consequently not to be considered in the design.
• Plastic instability (buckling) due to excessive plastic deformation requires special
load cases that do not exist for the copper overpack. As far as the insert is supporting
the copper overpack, the shell cannot collapse inwards.
• Creep crack growth.
Creep tests on notched specimens at 20 and 75ºC show that initially sharp notches are
blunted due to the high ductility of Cu-OFP and creep crack growth cannot take place
(Andersson-Östling & Sandström 2009).
8.1.2 Insert
The following mechanical failure modes of the cast iron insert under the conditions
given in Table 2:
• Plastic collapse (buckling).
In compressive stress conditions, loss of stability may be involved. Effects that may
contribute to buckling tendency are low yield strength, geometric inaccuracy, nonsymmetry
of the structure or load.
• Crack initiation or stable crack growth.
• Ultimate tensile strength is exceeded.
The following potential failure mode is excluded:
• Brittle fracture.
Brittle fracture is possible only for brittle materials at low temperature. The tendency for
brittle fracture depends on material quality, the amount of some foreign elements in the
material and the ambient temperature. The fracture mechanics testing of the insert mate-

68
rial at 0 °C temperature showed that all the test samples had a ductile behaviour during
the testing conditions (Claesson 2009). In addition, a series of high loading rate tests
were conducted and the results showed that the static fracture resistance curves are representative
even for dynamic loads, and the higher loading rate does not lower the fracture
resistance of this insert material at this temperature according to measurements reported
in (Öberg & Öberg 2009b). Therefore, brittle fracture of the insert is not considered
to be an issue in repository conditions.
8.2 Mechanical failure criteria
8.2.1 Copper overpack
Relevance and criteria for potential failure mechanisms in the copper overpack:
• Fracture due to excessive plastic deformation.
This failure mode can best be represented by the reduction in area, which is 80 to 90 %
for Cu-OFP and FSW welds in the material (Andersson-Östling & Sandström 2009).
Very large deformations of the order of the reduction in area are needed to initiate this
type of failure. The design criterion is that the effective strain should not exceed 80 %.
This is derived from the fact that the reduction of area 80 to 90 % in uniaxial test
specimen corresponds to 160 to 230 % of true strain. And half of that is taken as allowable
strain in design. This allowable strain (40 %) is given as a design parameter in Table
16. Figure 22 shows the copper stress/strain models for static and dynamic load
conditions.
Figure 22. Copper stress/strain models for static and dynamic load conditions. Remark
the difference between engineering stress and true stress.

69
• Rupture due to creep deformation.
Creep tests defect free material of parent copper metal and friction stir welds of Cu-OFP
have given a creep elongation of 30 % or more in the temperature interval 75 to 175 ºC.
Multiaxial creep tests that have been performed at 75 ºC demonstrate that Cu-OFP is not
notch sensitive and that the creep rupture time can be estimated to be 100 times longer
than in uniaxial tests for the same net section stress. Local strains of 30 % can appear
without crack initiation (Andersson-Östling & Sandström 2009). The creep rate is essentially
controlled by the effective stress since the creep exponent is a high as 65. The
factor with the deviatoric stress plays only a secondary role since it enters the creep rate
only to the first order. If creep rupture would occur during the conditions in the repository
it would show a ductile behaviour.
To initiate creep rupture a spatially constant stationary effective stress must have been
established across a section of the canister. The design criterion is that such a stationary
stress should not exceed the uniaxial rupture stress. The safety factor is chosen to be 1.2
after considering the flatness of the rupture curve. This rule is converted to design parameter
in Table 16 as follows. As the creep elongation in temperature interval between
75 and 175 °C has been 30 % or more, and half of that has been taken as an allowable
creep strain in design. Thus the allowable creep strain as design parameter in Table 16 is
set as 15 %.
8.2.2 Insert
The failure criteria for the cast iron insert are classified as follows:
• Plastic collapse (buckling).
This criterion is used for external pressure loading cases of the insert.
Plastic collapse is the first and most common failure mode for an externally pressureloaded
thick wall shell supported by bulkheads. This phenomenon can be accounted for
in the analyses by using large deformation theory in the numerical models, when the
external pressure load cases are analysed. It can be analysed and assessed according to
the plastic collapse method described in ASME Code, Section III, Divisions 1 and 2.
(ASME III 2008). The Code requires that the operational load shall be less than 2/3 of
the limit load, which means, in other words, that the required safety factor against
(global) collapse load is 1.5. This criterion is used for load controlled cases, in other
words, for external pressure load cases.
The interpretation of this criterion is given so that, for basic dimensioning, the plastic
collapse load is determined for the load case through isostatic pressure, say p L , and then
the maximum allowable isostatic pressure in the design is taken to be 2p L /3. In other
words, a safety factor 1.5 is used for collapse load analyses for design pressure load.
In analysis of components used in the nuclear industry, acceptance criteria are usually
adopted from the code (ASME III 2008). In the design analysis report the criteria for
plastic analysis described in (ASME III, Div. 1, NB-3228.3) is used.

70
The purpose with the method described in NB-3228.3 is to show that the applied load
does not exceed 2/3 of the calculated plastic analysis collapse load and if this can be
shown then the limits of General Membrane Stress Intensity (NB-3221.1), Local Membrane
Stress Intensity (NB-3221.2), and Primary Membrane Plus Primary Bending
Stress Intensity (NB-3221.3) need not be satisfied at a specific location (ASME III
2008).
In (ASME III 2008, Div. 1, NB-3213.25), the definition of a plastic analysis collapse
load can be found. The following criterion for determination of the collapse load shall
be used. A load–deflection or load–strain curve is plotted with load as the ordinate and
deflection or strain as the abscissa. The angle that the linear part of the load–deflection
or load–strain curve makes with the ordinate is called . A second straight line, hereafter
called the collapse limit line, is drawn through the origin so that it makes an angle =
tan −1 (2 tan ) with the ordinate. The collapse load is the load at the intersection of the
load–deflection or load–strain curve and the collapse limit line. This is also, more
clearly, shown in (ASME VIII 2004, Div. 2, 6-153) using Figure 6-153, which shows
how to determine this collapse load (please, note that this is given in ASME version
2004).
• Crack initiation or stable crack growth.
This criterion is used for all types of loading of the insert.
In the case of the external pressure load case, when the load controls the stresses and
causes primary stresses, the damage tolerance analysis is made using K Ic -data that are
based on crack initiation, not for limited stable crack growth like J 2 mm . The safety factor
used for K I -parameter is √10 = 3.16, which is the ASME Code requirement for normal
operational loads. This means that crack initiation is not allowed during pressure type of
loading.
When doing a damage tolerance analysis of components with cracks, different approaches
may be used concerning method of analysis and decision of safety factors in
the assessment. In Sweden, the Swedish Radiation Safety Authority (SSM) published a
handbook (Dillström et al. 2008) which describes a procedure that can be used both for
assessment of detected cracks or crack-like defects and for defect tolerance analysis.
The method utilized in this procedure is based on the R6-method. This is also the
method chosen for the damage tolerance analysis of the insert in the case of an external
pressure load case (R6, option 1 failure assessment curve).
In the case of a displacement controlled load, i.e. a rock shear load, the damage tolerance
analysis is based on a J-integral analysis. The integrity assessments are partly
made from the stress and strain results using global models and partly from fracture
resistance analyses using the sub-modelling technique. The sub-model analyses utilize
the deformations from the global analyses as constraints on the sub-model boundaries
and more detailed finite-element meshes are defined with defects included in the models
together with elastic-plastic material models. The J-integral is used as the fracture parameter
for the postulated defects. The allowable defect sizes are determined using the

71
measured fracture resistance curves of the insert iron as a reference with respective
safety factors according to the ASME Pressure Vessel Code requirements.
Within the SSM reported procedure, a deterministic safety evaluation system is defined
(which is not present in the original version of the R6-method). When choosing safety
factors for nuclear applications, the objective has been to retain the safety margins expressed
in (ASME III 2008), and (ASME XI 2008). For ferritic steel components SF K =
3.16 (normal/upset load event) and SF K = 1.41 (emergency/faulted load event) as defined
in the SSM-handbook (when using a J-integral analysis, SF J should be used,
where SF J = (SF K ) 2 ).
These safety factors are taken from (ASME XI 2008, Div. 1, IWB-3612); acceptance
criteria based on applied stress intensity factor.
Doing a damage tolerance analysis, using these safety factors, does not imply that one
needs to fulfil other code requirements within the ASME code (regarding inspection,
fabrication etc.). The only purpose is to use established safety factors for nuclear applications
when doing a damage tolerance analysis.
The aspect ratio chosen for postulated (initial) defects is mainly related to the assumed
damage mechanism. When no damage mechanism is known, an aspect ratio
(length/depth) of 6 could be used for surface defects. In the damage tolerance analysis
for the insert, different assumptions regarding the aspect ratio has been used (both for
surface and subsurface defects). The purpose has been to show that it is possible to introduce
reasonable sized defects without jeopardising the integrity of the reference canister.
Regarding annual frequency of occurrence of the loading conditions and implicit conditional
probability of failure in the event of the service loading, it is believed that this is
fulfilled and also conservative.
Initiation of crack growth can be allowed for special load cases, but reasonable safety
margin shall be applied for stable crack growth. This means for ex. that the calculated
fracture parameter J may be higher than the J c that corresponds to the initiation of crack
growth but the design basis value could be J 2 mm that corresponds to the stable crack
growth of 2 mm, which is very moderate in a massive iron structure of typical dimension
of 1m. Reasonable small crack growth can be allowed, because limited local crack
growth does not lead to global rupture.
In the case of a displacement controlled load, like the rock shear, the influenced stresses
are secondary in character. The stable crack growth criterion is then taken as
J(a)

72
A safety factor of 2 for J-integral is equivalent to a safety factor √2=1.41 for K I -
parameter, which is the parameter that is primarily used in ASME Code. This discrepancy
in safety factors comes from the relation between J and K I as follows: (K I ) 2 =
J*E/(1-ν 2 ), where E is Young’s modulus and ν is Poisson’s ratio. This relation means
that J ~ K I 2 . The justification for classification of the shear load case as a low probability
case is based on reasoning in Section 2.3 of (SKB 2009), where it is calculated that 4
canisters out of 6000 may be subjected to shearing of magnitude of 5 cm or more. This
gives a probability of

73
Thus we can set an engineering type stress criterion for displacement controlled secondary
stresses that the effective stress may be, at maximum, the stress corresponding half
of the strain at the lower limit of the ultimate elongation (A 5 ) in uniaxial tensile testing
with 90 % reliability (12.6 %, see Section 4.1.1). Later (in Section 8.3.2) we will see
that much less (about 2 %) would be much enough. Thus the actual safety factor against
elongation at rupture is about 12.6/2 = 6.
The stress in the actual stress-strain curve of the insert iron corresponding a half of the
12.6 % elongation is 395 MPa. Figure 23 shows the relation of the allowable effective
stress in comparison to ultimate tensile stress (UTS). The stress-strain curve is the
multi-linear elastic-plastic relation used in the FEM-analyses. We can see that the strain
energy utilised at allowable stress level is less than a half of that at failure. This means,
in other words, that we set a requirement of a safety factor of 2 against the ultimate
strain for the effective stress. The rock shear case in ASME Service Conditions (load
classification) is Level C or D condition, for which the Code does not set any direct
requirements for the secondary stresses, according to (ASME III 2008, Figure NB-
3224-1). The selected engineering type stress criteria are clearly stricter in this case than
the ASME practise.
500
Uniaxial true stress/strain relation for insert iron
400
Stress (MPa)
300
200
100
0
0 2 4 6 8 10 12 14
Strain (%)
Figure 23. The ultimate strength is 456 MPa (red lines) and the corresponding strain is
12.6 % with the reliability of 90 % according the material testing. The maximum allowable
effective (von Mises) stress (green lines) is set in a way that the respective strain is
half of the strain (A 5 ) corresponding the ultimate tensile strength (UTS=456 MPa). The
green lines are at 6.3 % and 395 MPa. The basic curve is the average tension test results
(true stress) in 20 °C with standard test (no strain-rated). The maximum actual
plastic strain in rock shear case is about 2 %, thus the safety factor in strain is about 6.

74
Uniaxial uniform elongations for the insert material are 12.6 % < A5 < 14.8 % using
90 % confidence in test data. Elongation measured in the uniaxial tensile tests could be
used as a measure for maximum allowable strain, but multi-dimensional strain condition
should be taken into account. However, such a criterion for strains with a general acceptance
is not available. Thus we use the effective stress criteria (von Mises), as given
above, instead of equivalent plastic strain (PEEQ in ABAQUS terms).
8.2.3 Summary of mechanical failure criteria relevance
The failure criteria relevance for iron and copper are summarised in Table 15. Brittle
fracture in actual operating temperatures can be neglected because of the adequate fracture
resistance properties of the both materials. Iron and copper are very dissimilar metals,
thus different type of failure criteria are needed.
Table 15. Summary of the relevance of failure criteria.
Failure criteria
Relevance to cast iron insert
Plastic collapse (buckling) Yes (for primary stresses) -
Crack initiation or stable crack growth Yes (for all load types) -
Ultimate tensile strength is exceeded Yes (for secondary stresses) -
Fracture due to excessive plastic deformation - Yes
Creep - Yes
Brittle fracture (at relevant temperature) - -
Relevance to
copper overpack
8.3 Strength and damage tolerance
The basic design verification of the canister as a load carrying component has been conducted
according to mechanical design codes where applicable. As an example, the
ASME Boiler and Pressure Vessel Code, (ASME Section III 2008; ASME Section VIII
2004; ASME Section XI 2008), gives methods and application rules for design verification
of reactor pressure vessels. The code gives practical guidance to make the integrity
assessment. The only problematic area in the design verification and strength verification
is that the engineering materials of the canister are non-typical for customary vessels
and shells. Thus the examination of the properties of the construction material (cast
iron and oxygen-free copper) has been emphasized. The material testing and analysis of
test data that have been carried out are very comprehensive and extensive, and the experimentally
determined properties used in the design verification analyses are considered
very reliable.
As for acceptability of the calculated results, the failure criteria defined in Section 8.2
are used as a reference. All the set criteria are assessed separately in the following subsections.

75
8.3.1 Plastic collapse criteria
The basic design verification of the canister against the external pressure of 45 MPa has
been conducted according to the ASME Code guidance utilising the limit load method.
The canister is modelled using both 2 and 3-dimensional finite-elements, where all the
components, gaps, tolerances and materials are modelled as realistically as possible and
the limit load is estimated. The stress acceptance criterion is that the design load shall
not exceed 2/3 of the limit load. Separate analyses were made for insert cylinder, insert
bottom and the steel lid on the top end of the insert (Dillström et al. 2010a; Alverlind
2009a; Alverlind 2009b). All these analyses fulfilled the design criteria. Some other
analyses were made earlier on pressure resistance of the canister with similar results
(Ikonen 2005; Martin et al 2009).
In addition to basic design verification analyses, the strength was also investigated with
modelled deviations in nominal geometry, tolerances, lack of material or inclusions in
the cast, eccentric installation of steel cassette and rounding radii of the corners of the
square tubes. The canister strength was shown to be insensitive to these types of imperfections
(Dillström et al. 2010a).
The analyses of the cast iron insert and the steel lid show that the design is very rigid to
an external isostatic load of 45 MPa. Analysis of the damage tolerance of the cylindrical
part of the insert shows that large defects can be tolerated without jeopardising the canister
integrity. A defect size of a maximum of 20 mm can be accepted for both hole-type
and crack-like defects and a 10 mm off-set of the steel tube cassette can be accepted. An
off-set of 10 mm means that the edge distance, measure H in Figure 7, is accepted to be
reduced by 10 mm. The presented results are based on material data from inserts manufactured
some years ago. As stated in (Dillström et al. 2010a) data from more recently
manufactured inserts show that the performed analyses are pessimistic.
The analyses for the insert bottom and steel lid show that even when including the least
favourable combination of geometrical tolerances the margin to the collapse limit load
is high.
The values of the important design parameters for the insert with regard to an isostatic
load shall be:
• compression yield strength ≥ 270 MPa, the insert stands for the load elastically,
• fracture toughness K Ic > 78 MPa(m) ½ (90 % lower confidence) at 0°C to
withstand brittle fracture, and
• tensile yield strength of the steel lid material ≥ 335 MPa.
The pressure load capacity of the canister was demonstrated earlier by two model tests,
when 700 mm long sections of actual canisters were pressure tested up to the limit load.
The pressure tests showed that the collapse pressure was in both cases between 130 and
140 MPa (Nilsson et al. 2005); that is roughly 3 times the design pressure. The pressure
tests have been used also for calculation method’s validation. The aim of the programme
was to verify that the probability of a canister breakage in deep repository condition
is acceptably low (

76
robust and clear results in that the risk for global collapse is vanishingly small (10 -50 )
according to the probabilistic assessment of (Dillström 2009).
In general, the observed material properties are better than the minimum specified properties.
Thus the actual collapse pressure is expected to be higher than the lowest predicted
limit value from the numerical strength analyses, even if the test pieces contained
material faults and geometric imperfections. The damage mode of the canister collapsed
under external pressure is shown in Figure 24, for more details see (Nilsson et al. 2005).
Figure 24. Result of destructive pressure testing of a BWR type canister. Collapse pressure
load was 138 MPa. Photograph shows the second pressure test result (Nilsson et
al. 2005). The collapse mode of the insert is buckling of the wall between open positions
for fuel elements and then the shell rupture takes place through shearing of the copper
overpack along the insert lid.
8.3.2 Stress / strain criteria
The operational loads for the copper overpack have been analysed. The strength is
adequate in all the analysed cases.
An important fact when assessing the allowable stresses and strains is that in case of the
canister, the loads are not variable or cyclic in nature, but very stable and unique in
character, and thus there is no need for fatigue or cyclic crack-growth studies.

77
Cast iron insert
During encapsulation the insert lid and central screw are loaded by the 1 bar overpressure
of the inside inert gas. The loaded surface area of the lid is 0.949 2 *π/4 = 0.707 m 2
and the net section of the M30 mm central screw is 561E-6 m 2 . The stress caused by the
1 bar pressure load on the lid causes a tension stress of 0.707/561E-6*0.1 MPa = 126
MPa in the screw net section. The yield strength of an ISO 898 grade 8.8 screw is 80 %
of the ultimate strength 800 MPa that is 640 MPa. This means that the safety factor between
existing stress and yield stress is 5, which is much more than required.
For cases involving an external pressure load and design conditions the strains and
stresses are low. The bending effect of the postulated uneven bentonite pressure load
also does not lead to any risk of excessive stress or strain values in the insert. For the
strength analyses of the insert, the plastic collapse load method, with a safety factor 1.5,
was applied according to (ASME VIII 2004).
The only load case that may locally lead to significant yielding and plasticity of the insert
is the rock shear case. Rock shear is, however, a “displacement-controlled load” that
causes secondary stresses only according ASME nomenclature. If the load is secondary,
the possible local yielding or cracking leads to decreasing stiffness and increasing deformation
in the structure and, consequently, the load would decrease. That is why additional
safety factors are not needed in displacement-controlled load cases. However, the analysis
results for rock shear case show, that in case of 5 cm shear the plastic deformations and
strains in the canister insert are low (1-2 %), see (Raiko et al. 2010, Table 6-3) and material
rupture is not expected to take place. For the square steel tubes the stresses are higher than
for the cast iron part, but the steel has higher ductility and higher ultimate strength, so both
of insert materials fulfil the criteria.
The measured ductility (Claesson 2009; Öberg & Öberg 2009b) from manufactured inserts
is clearly sufficient for the insert in the postulated 5 cm rock shear case, as far as the
allowable size of existing cracks is concerned, (Dillström & Bolinder 2010b). The applied
metric was allowable equivalent stress according to von Mises. What most evidently justifies
the analysis as a whole is that if the effective plastic strain exceeds the ductility
limit, a crack can be expected to initiate. Both linear and non-linear fracture mechanics
analyses do however show that the integrity of the insert structure is not at jeopardy in
postulated load cases, even if very large cracks are postulated in critical locations and
orientations.
Copper overpack
The behaviour of the copper overpack during design bases load cases is described in
(Raiko et al. 2010, Section 6.2). High strains are observed only at geometric discontinuities.
These do not threaten the global integrity or leak tightness of the copper overpack.
Extensive copper creep is limited either by the supporting effect of the insert in the pressure
load cases or by the relaxation of the applied displacement boundary condition in the
bentonite buffer in the rock shear case. The results are shown to be acceptable for all postulated
load conditions and combinations.

78
The basic operational load condition for the copper overpack is the lifting of the loaded
canister during handling from the copper lid lift shoulder. The canister is handled either by
supporting through the bottom end or by hanging from the top end shoulder in the copper
lid. In the following the strength of the lid shoulder is verified. The shoulder is calculated
as a cantilever beam, whose length is 14.5 mm and height is 35 mm, the load bearing
width (the total width of the gripper jaws) is assumed to be 75 % of the total circumference.
The total circumference of the shoulder is 2.67 m. Thus the grip is assumed to load
the length 75 %2.67 = 2.00 m of the circumference. The shape of the shoulder is shown in
Figure 15.
The maximum shear mode and bending mode stresses are calculated at the root of the cantilever.
The shear stress is calculated according to the formula (6)
F
f
A
F
f
b
h , (6)
where F is the maximum weight of the canister, EPR type (285 kN),
f is the additional dynamic load factor (1.3)
A is the sectional area of the loaded part of the shoulder,
b is the load bearing width of the shoulder circumference (2.00 m), and
h is the section height at the butt of the shoulder (0.035 m).
The resulting shearing stress in the lifting shoulder is = 5.3 MPa.
The maximum bending stress is calculated assuming that the entire dead weight load is
concentrated onto the inner edge of the shoulder, see Figure 15. The bending stress component
is calculated according to the formula (7) for cantilever beam:
b
M
W
F
f l
2
, (7)
b
h
6
where M is the bending moment in the section,
W is the section modulus,
F is the maximum weight of the canister (285 kN),
f is the additional dynamic load factor (1.3),
l is the distance of the acting force from the section (0.0145 m),
b is the load bearing width of the shoulder circumference (2.00 m), and
h is the height of the shoulder section (0.035 m).
The resulting maximum bending stress in the lifting shoulder is b = 13.2 MPa.
The typical minimum yield stress of annealed copper is 50 MPa in room temperature and
45 MPa in the design temperature of +100 °C. The reduced stress is calculated combining
the bending and shearing stress components as follows (8):

79
1 1 2 2
red
2 2 4
(8)
We get the reduced stress of 15.1 MPa. Thus the actual safety factor against yielding in
design temperature is 3, which is generously acceptable.
No standards give minimum yield strength for hot deformed oxygen free copper, only
typical strength values are given. Now, as the copper overpack is used as a structural
member, when lifting the canister, the minimum yield strength shall be defined. Referring
to the strength calculation above, the yield strength of 40 MPa in room temperature for the
base copper material of the canister lid is adequate. This leads to safety factor 2.6 in canister
lift condition.
When the whole canister is lifted from the top lid corner the gravity load causes an average
axial membrane stress of 1.8 MPa to the copper cylinder of thickness 49 mm and an average
shear stress of 1.9 MPa to the weld between the lid and the cylinder. These stresses are
insignificantly low when compared to minimum copper yield strength specified above as
40 MPa.
In case of analysed pressure case (45 MPa); plastic and creep deformation levels in the
copper overpack are generally very low, below 1 %. In some parts, at the lid and base,
areas can be found where the creep strain approaches 12 %, see (Raiko et al. 2010, Section
6.2.1). In very local areas at the slits between the copper cylinder and the lid or base
plastic deformations of 30 % can be found (in FSW geometry), see (Raiko et al. 2010,
Section 6.2.3). However, creep rupture cannot be initiated because of this since a section
through the canister would then have to creep simultaneously, and this does not
take place because of the very local distributions. All together the results of the performed
analyses show that an isostatic load case of 45 MPa external pressure load will
not cause any rupture of the copper overpack within design lifetime if the design parameters
are fulfilled. The more detailed discussion of the item is given in (Raiko et al.
2010, Section 6.2).
The effects of uneven swelling pressure on copper overpack are analysed in (Andersson-Östling
& Sandström 2009), in Section 11.4. The plastic strain in copper in there
load cases is very low and the possibility of creep is very limited.
In rock shear case, the strain and stress levels in the copper overpack have been evaluated
using the global model (Hernelind 2010). Evaluations have been done including
short-term analyses to assess the plastic strains that the copper is exposed to during the
shear movement. Complementary analyses including assessing the creep in the copper
overpack after the shear movement have also been done to determine the levels of creep
strains that can be expected after a rock shear (long-term shear analyses). The creep
analyses have been done by incorporating the copper creep model described in (Raiko et
al. 2010; Andersson-Östling & Sandström 2009) into the global model described in
(Hernelind 2010).
The rock shear analysis is currently (2012-2013) updated. The parameters are varied in
a large number of analyses to get an idea of the probabilistic nature of the case. More

80
rock slice locations are calculated. Also the effect of the possible weak joint between
the steel cassette tubes and the cast iron of the insert are examined.
The short-term analyses showed that, for the case of rock shear perpendicular to the axis
of the canister, the maximum plastic strain in the copper overpack, in cylinder part outside
geometric disturbances, is generally less than 2 %. This is the same order of magnitude
as for creep strain in the long-term analyses. This means that most of the copper
strain is caused by the immediate plasticity during the rapid rock shear load case and the
creep after the shear case will only relax the stresses in the bentonite-copper-iron construction.
However, the highest local strains in copper overpack are localised to the radii in the lid
or welded base, and the maximum is 9.3 % in the fillet and 21-23 % in the singularity of
the slit between the cylinder and the lid, see (Raiko et al. 2010, Section 6.2.5 and table
6-4). Figure 25 shows the local character of the high strain locations in case of 5 cm
rock shear case.
The copper ductility is much enough to tolerate even the peak strain values in the geometric
notches, because the specified ultimate elongation of copper is ≥ 40 % in uniaxial
tests, see Table 16. The copper overpack is sufficiently ductile to tolerate the strain
values generated by the 5 cm rock shear case. Additional safety exists in plasticity and
creeping analysis result assessment from the fact that all material test data is generally
given in engineering units (stress and strain), whereas the FEA-simulation results are
always in true-stress and true-strain units.
Figure 25. Plastic strain of copper in canister lid area. The load case is 5 cm rock
shear. Maximum strain is 21 % but it is very local. Only the very limited grey coloured
volume has strain more than 2 %.

81
8.3.3 Fracture resistance criteria / allowable defect sizes
For the copper overpack, no kind of postulated crack, defect or cavity of postulated size
has proven to be critical. The operational loads for the copper overpack have been analysed
with postulated large circumferential cracks or with a lack of material. The copper
overpack withstands the design loads with a good margin even with these large postulated
defects. The material testing has shown that the copper cracks blunt under a tension
load and no kind of crack growth is detected at applicable temperatures.
The insert was analysed for postulated cracks and other types of material deficiencies.
With design pressure load the allowable ASME-type reference defect became a 32 mm
deep semi-elliptic crack in the BWR-insert and 31 mm in the PWR-insert, respectively.
The safety factor used was 3.16 according to ASME-requirements for normal operation
condition loads as expressed in stress intensity factor value K Ic , according to practice
described in (ASME XI 2008). The crack sizes are however limited to a maximum 80 %
of the material thickness and deeper cracks are thought to be possible without exceeding
the allowed K Ic .
The design basis load case with respect to allowable defect size proved to be the rock
shear case. This load is a rare case that will occur only for very few canisters or none at
all. The load will be very short lived and there is extremely low probability that the
shear movement will occur more than once on the same single canister. The temperature
is assumed conservatively be at 0 °C and a safety factor of 2 is used when defining the
allowable J-parameter value from the fracture test results corresponding the stable crack
growth of 2 mm at 0 °C. The rock shear is classified as a level D load case (emergency
condition) according to ASME Code (ASME III 2008) practice and the safety factor is
determined respectively.
The maximum allowable surface defect size on the cylinder surface is a 4.5 mm deep
and 27 mm long reference defect laying in a circumferential orientation. This damage
tolerance analysis is the design basis load case for the canister insert for close-to-surface
volumes. For more calculated results, see (Raiko et al. 2010, table 6-6). The resent results
from cast iron fracture resistance testing of I68 (Planman 2012) have given remarkably
higher results. If the better trend in fracture resistance is later statistically verified,
the allowable flaw size in the insert may be increased remarkably for the rock
shear case. The rock shear case is currently (in 2012) also re-analysed with probabilistic
assessment of existing combination of properties and circumstances.
The reference canister withstands the specified loads with an applicable safety margin
even if the material has the allowable size defects mentioned above.
8.3.4 Essential design parameters
Table 16 summarizes the essential design parameters that have an influence on the canister
integrity and have either an effect on the static strength or damage tolerance of the
canister. Table 16 also contains the failure modes which are affected by the parameter,
an estimate of the qualitative sensitivity and a reference to possible manufacturing
specification values.

82
Table 16. Essential mechanical design parameters.
Parameter Effects on Sensitivity Value derived from
the design analysis
Yield strength of cast
iron in compression
Yield strength of cast
iron in tension
Ultimate elongation
of cast iron
Fracture toughness of
cast iron
Minimum ligament
thickness of the insert
wall around
square openings (dimension
H in Figure
7)
Yield strength of
copper (design
strength)
Ultimate elongation
of copper
Creep ductility of
copper
Cold work of copper
(strain hardening,
hardness)
Gap dimensions between
insert and
copper overpack
Wall thickness of
copper overpack
Plastic deformation
and strain
Plastic deformation
and strain
Rupture
Ductile fracture
Strength and stability
Lifting safety of the
canister
Rupture
Important but adequate
in practice
Less important for
any load case as far
as the lower limit is
satisfied
Important but adequate
in practice
Important in lowtemperature-highload
condition
Collapse load is directly
ruled by the
weakest load carrying
member of the
insert
Important but adequate
in practice
Important but adequate
in practice
≥ 270 MPa (statistical
requirement)
≥ 267 MPa, upper
limit to be determined
(shear load
case)
≥ 12.6 (As defined in
Section 7.2, statistical
requirement)
K 1c > 78 MPa (m) 1/2
J 2 mm > 88.1 kN/m
statistical requirement)
< 10 mm deviation
from nominal value
≥ 40 MPa
≥ 40 % in uniaxial
tests
Creep rupture Important > 15 % in uniaxial
tests
Reduce creep ductility
Limit the plastic or
creep deformation of
the copper overpack
Corrosion resistance
Important
Sensitive and important,
but strictly set
tolerances keep the
effect within acceptable
limits
Non-dimensioning in
the mechanical design
analysis
Further investigations
are made
Axial gap 1.9-3.1
mm, radial 1.25-2.0
mm
Axial gap for EPR
canister is 2.4-3.5
mm
Nominal 49 mm of
which 35 mm without
flaws

83
8.3.5 Strength of variant canister designs
The variant canister designs are those for the VVER-440 and the EPR/PWR fuel types.
The collapse load case leads to limited plasticity and large deformation in the insert structure.
In addition, in case of square tube openings, the steel tubes tend to be separated from
the cast iron body due to weak interface strength between the steel tubes and the cast iron.
The material behaviour is modelled in this ultimate case including the post-yield condition.
The critical measures in this kind of analysis are the maximum strain and the maximum
deformation.
The limit load is the pressure load that induces plastic collapse of the insert structure. The
limit load analysis for the canister structure (insert + copper overpack) was made with finite
element method using non-linear (elastic-plastic) material modelling and large deformations
(Ikonen 2005). In the dimensioning analyses the effect of copper overpack was
conservatively omitted.
The yielding and strain hardening material behaviour was modelled with bit-by-bit linear
stress-strain relation based on the standard requirement for the yield (240 MPa at +20 °C)
and ultimate strength (370 MPa) and respective typical measured behaviour of the cast
material. The external pressure load acting on the model surface was incrementally increased.
The non-linear analysis was continued in load increments as far as to the point
that the structure became unstable due to exceeding large deformations caused by the external
load. The stepwise balance iteration was used to put the system converge with all
increments. The ratio between load and maximum displacement was very stable until
about 90 MPa pressure and after that the displacements started to increase more rapidly.
As long as the iteration converges, the strain state is stable and the load-carrying capacity
is not exceeded.
The results using standard strength values for cast iron showed that the pressure-load
carrying capacity of the cast insert is at least about 90 MPa in case of BWR-type insert
and far more with other type inserts. The mechanical dimensioning calculations of the
canister design with three variants (for BWR, VVER-440, and EPR/PWR fuel) are documented
in the report (Ikonen 2005). The mechanical strength of all variant design fulfils
the requirements for isostatic pressure load with high margins of safety. The minimum
collapse load with lowest standard material properties and with most unfavourable manufacturing
tolerances due to external pressure is 90 – 150 MPa depending on the canister
insert type. These analyses show that the VVER-440 and EPR/PWR variant designs of
canister insert are more robust and resistant against pressure load than the reference design
(BWR-type canister). The SKB analyses for their PWR type insert in (Raiko et al. 2010)
support the assessment given in (Ikonen 2005) of the strength of variant canisters against
pressure load.
As for bending type loads, the VVER-440 and EPR/PWR variant inserts have higher flexural
rigidity and, accordingly, higher sectional modulus than the BWR insert that has been
analysed for all load cases more thoroughly. The values are shown in Table 17 that are
scaled to the nominal geometry. Section modulus of VVER-440 insert is 4 % higher than
that of BWR and EPR/PWR insert has a modulus of 19 % higher than that of BWR. This
leads to the conclusion that the bending load resistance of the variant inserts is at least

84
equal to that of the reference BWR insert, even if the bending load is weighted according
to the square of the total length of the canister. The bending moment, and accordingly
bending stresses of the unevenly distributed swelling pressure, is proportional to the square
of the length according to (Börgesson et al. 2009, Section 2.2). The relative length between
canister variants are (BWR):(VVER-440):(EPR/PWR) = (4.752):(3.552):(5.223)
and the square of the ratios are after normalization 1 : 0.56 : 1.21. This simple comparison
shows that the bending type load resistance of the VVER-440 canisters is higher than that
of the reference canister and that the BWR and EPR/PWR variants are about the same.
This comparison is valid for stresses, but in case of fracture resistance, equal fracture resistance
is expected for all types of inserts. So far, the manufacturing demonstrations have
showed acceptable fracture resistance only for BWR type of insert. Further development of
casting process is needed for EPR/PWR type of insert, because of remarkably thicker wall
sections that do not show satisfactory ductility.
Table 17. Sectional properties of the cast insert for various types of insert. The numbers
are based on nominal dimensions of the section geometry.
Sectional area
A (m 2 )
Section modulus
W (m 3 )
Flexural rigidity
I (m 4 )
BWR -type insert 0.4001 0.05635 0.02674
VVER-440 -type insert 0.4230 0.05855 0.02778
EPR/PWR -type insert 0.4895 0.06689 0.03174
As for copper overpack under various load cases and in case of variant canister types it
can be concluded that the size and shape of the copper overpack is identical for all canister
variants with the exception of total length that varies -25 % or +10 % when VVER-
440 and EPR/PWR types, respectively, are compared to the reference canister type of
BWR. It is evident that the shear type loads for identical copper canister lids are the
same independently of the length of the canister and the bend type loads for shorter canister
are lower than for longer one, when the diameter is the same. The design basis
shear load is such that it is cutting the lid transversely off from the top of canister. Thus
for VVER-440 canister shell there are no doubts of its lower loads when compared to
the reference canister. For EPR/PWR canister the relative length is only 10 % more than
that of reference canister, thus the load estimates differ very little from the reference
canister, as shown in bending moment discussion for the swelling pressure load above.
The design basis load case for the canister shell is the transverse shear at middle plane
according to (Hernelind 2010). The maximum strains are located in the copper lid area;
see (Raiko et al. 2010, figures 6-16 and 6-17).
On the basis of these comparisons of the collapse load, section modulus and canister
length it can be concluded that the analyses made for BWR type reference canister
cover with adequate accuracy also the respective analysis needs for the canister variant
designs (VVER-440 and EPR/PWR).

85
8.4 Thermal behaviour
8.4.1 Temperature inside canister
Heat transfer between the cast iron insert and the copper overpack takes place by direct
conduction through the metal surfaces that are in direct contact, e.g. the bottom lids.
Moreover, heat is transferred by radiation over the gap between insert and shell. Conduction
in the gap is possible if the gap is gas filled. If the electron beam welding technology
is used to seal the lid, the gap will be in a vacuum but if the friction-stir welding
technology is used, the gap will stay gas filled with air at atmospheric pressure.
Temperature and heat fluxes in the fuel and canister cavities have been calculated by
(Ikonen 2006). Uncertainties in the heat transfer between the different components of
the canister exist. However, heat transfer among fuel rods and to the cast iron insert
takes place mainly by radiation. Heat transfer by radiation is proportional to T e 4 where
T e is the absolute temperature of the emissive surfaces. The maximum temperature in
the fuel is, according to the analysis, about 230 °C, if the canister outer surface temperature
is at typical 90 °C (Ikonen 2006). If FSW method is used to seal the canister lid
instead of EBW method, the maximum temperature in the fuel rods will be considerably
lower, about 150 °C (Ikonen 2006), Figure 15. This large difference in the maximum
fuel temperatures is due the difference in the thermal conduction of the gap between the
copper overpack and the insert. In EBW case, the gap is assumed to stay in absolute
vacuum and only radiation heat transfer is possible. In the FSW case, the gap is airfilled
at normal atmospheric pressure. The vacuum assumption in the gap between insert
and the copper overpack is very conservative, especially in long-term safety assessment.
Heat generation and increased temperature dominate other canister processes only for a
relatively short time period, in the order of 1000 years.
The thermal analyses of the canister inside temperature were partly repeated by using
initial data that correspond better to the current understanding of the system evolution.
The main goal of the analyses is to determine the temperature in the insert by using the
following assumptions. The decay power of all the fuel inside a reference canister is
1700 W. The canister outside surface is at 95 °C temperature. The conductivity of cast
iron is 36 W/m 2 /K. The gap between insert and copper overpack is 1.5 mm wide. The
emissivity of the gap surfaces are 0.22 and 0.6 for the copper and iron surfaces, respectively.
These are typical tabulated values (see Table 18) for matte metal surfaces, i.e.
unpolished surfaces. The insert is filled with argon at normal atmospheric pressure,
whose conductivity is 0.018 W/m 2 /K. The gap between insert and copper overpack is in
vacuum after the EB welding of the copper lid. Later, due to leakage through the insert
lid gasket, the gap may be filled with argon. If FSW is used instead of EBW for canister
sealing, then the gap remains filled with air in normal pressure. The conductivity of air
is 0.030 W/m 2 /K.
For completeness, the insert temperature is analysed in two cases; the vacuum gap between
insert and overpack, and air in the gap. The results from the analyses are such that
the maximum temperature in the insert (determined in the centre) is 139 °C, and 103 °C,
when the gap between copper and insert is in vacuum, or contains air, respectively. The
assumptions mentioned above give for the maximum temperature in the fuel 193 °C, or

86
166 °C (intermediate gas between fuel rods is argon in all cases). We can conclude that
the fuel temperature is well below the maximum allowable temperature of about 300
°C. The allowable temperature for spent fuel in canister is estimated from reactor conditions
for which the fuel is designed for. The reactor coolant temperature is about 300 °C
and now in canisters we have the fuel rods in inert gas and the thermal power of the fuel
is close to zero (only low decay is left). As a background, elevated temperature may
cause immediate damage to fuel rod typically at temperature above 600-800 °C.
From the results above we can conclude that the insert temperature is about 45 °C (139-
95=44 °C) higher than the copper overpack at maximum. If the gap is air-filled (FSW
case), the insert temperature is about 8 °C higher than the copper overpack. The analysis
is conservative, because it is made in cylindrical symmetric condition and the thermal
flux through the bottom and top end of the canister are ignored.
8.4.2 Thermal expansion of canister components
The canister insert is made of cast iron whose linear thermal expansion coefficient is
11.5*10 -6 K -1 . In comparison, the copper overpack has an expansion coefficient of
16.9*10 -6 K -1 . The expansion coefficients are given in Table 14 of this report. The
nominal size of the gaps between the insert and the copper overpack are according to
the dimensions in Table 5 of this report. Nominal dimensions refer to room temperature
(20 °C).
The internal gap (clearance) increases if the canister components are in constant elevated
temperature, because the expansion coefficient of copper is higher than that of
iron. However, when the insert is at higher temperature than the copper overpack, then
the gap decreases. The following scoping calculation was made: the smallest gap dimensions
(1.9 mm in axial and 1.0 mm in radial direction) were assumed and the temperature
is then varied from the assemblage temperature (20 °C). If the copper overpack
is assumed to be at highest temperature (100 °C), then the cast iron insert may be heated
up roughly to 170 °C before the axial gap becomes in contact and roughly to 300 °C
before the radial gap becomes in contact. See the red arrows in Figures 26 and 27. The
horizontal red arrows show the temperature increase that correspond the contact in respective
orientation.
The actual calculated temperature difference between insert and copper overpack is, at
maximum, 45 °C; see the analysis results in Section 8.4.1. This shows that even in this
case we have a margin for contact even with the smallest axial gap dimension 1.9 mm.
As discussed in Section 8.4.1, the calculated temperature difference 45 °C is kept conservative.

87
Figure 26. The thermal expansion of the BWR canister components in axial direction.
Red arrow shows the allowable temperature difference (~70 °C) between insert and
overpack before the insert and the shell will have a contact. The gap is assumed to be
the minimum (1.9 mm) at assemblage at room temperature.
Figure 27. The gap and the thermal expansion in radial direction. Red arrow shows the
allowable temperature difference before contact between the insert and the shell. This is
valid for all canister variations. The gap is assumed to be the minimum (1.0 mm) at
assemblage at room temperature.

88
The canisters reach their maximum temperature in the repository within 10 to 15 years
after disposal. A typical maximum temperature evolution is shown in Figure 28.
The case of simultaneous maximum temperature of the copper overpack and maximum
thermal deformation is also considered. In this case, the shell is also deformed due to plastic
or creep deformation under the external pressure load. As a result of these deformations
the radial and axial gaps between the overpack and insert are closed. The copper overpack
will be deformed until full contact is reached on all the surfaces between the shell and
the cast iron insert due to the slowly increasing external pressure load. When contact is
reached, the copper overpack material is in compressive stress state and at the yield point.
This kind of analysis will be simulated in conjunction of further canister creep analyses,
see (Holmström et al. 2012b).
In the following, a simplified manual calculation of the stresses induced in the copper
overpack during the long cooling period after the maximum creeping is given. Excessive
tensile stresses may increase the risk for stress-corrosion cracking in the copper material.
The temperature of this bimetallic structure is cooled within a few thousands of years from
the conservative maximum of +80 °C to assembly temperature of +20 °C and later in very
long term down to +10 °C that is roughly the natural ambient temperature in the rock at
that depth. It is pessimistically assumed that the insert is extremely stiff when compared to
the copper overpack and all the thermal deformation is concentrated into the copper overpack
only. The temperature decrease will now decrease the compression stress state into a
tensile stress in the copper overpack due to the fact that the copper is shrinking more than
iron when the temperature is decreasing. This is due to different thermal expansion properties
of copper and iron. The shrinking can be estimated conservatively by the formula (9)
T , (9)
where is the change of the strain due to temperature decrease,
T is the temperature change of the system (10 - 80 = -70 °C), and
is the difference of linear thermal expansion coefficients of the materials
in the shell and canister insert (16.910 -6 - 11.510 -6 = 5.410 -6
1/°C). Thermal expansion coefficients are given in Table 14 in Section
7.5.
From Equation (9) we get = 0.000378. If the residual stress level both in axial and
circumferential direction after creeping is assumed to be as low as -50 MPa (compression)
equalling yield stress of annealed copper, the compressive strain at the end of
creeping at higher temperature is originally (using general Hooke´s law in two dimensional
case) yield = = 1/114 GPa * [-50 MPa - 0.308 * (-50 MPa)] =
-0.00030. After cool-down the strain of the copper overpack against the assumed rigid
insert surface is increased from yield with leading to yield += -0.00030+0.000378 =
+0.000078 = +0.0078 %. This tensile strain, by using the same Hooke’s law, corresponds
to tension stress of 13 MPa that is only some 25 % of the yield stress of annealed
copper. This is valid both for axial and circumferential stress component (the
material physical properties are according to Table 14). The calculation shows that the
afterward cooling of the canister releases the post-creeping stress state close to zero

89
stresses. This level (13 MPa) of residual tension stress in copper (induced by bi-metallic
effect in lowering temperature) is of no importance if assessed against the risk of stress
corrosion cracking. Stress corrosion cracking is discussed more in Section 8.6.
8.4.3 Thermal evolution of the canister surface
The main process during the thermal evolution of the canister is heat transport to the
surrounding buffer and rock. Heat is generated by the radioactive decay of some of the
radionuclides in the spent fuel at a time-dependent rate depending on the characteristics
of spent fuel, as shown in Section 13.3. Heat is generated in the fuel pellets and is transported
in the fuel and cavity by conduction and radiation to the canister insert and then
through the insert material to the canister shell, bentonite buffer and to the near and far
field of the rock.
The heat transport is governed by the thermal properties of the thermal transfer pathways.
In solid materials, the heat is transferred by conduction and in gas-filled gaps
(such as those between the canister and the buffer) by radiation, conduction and, in case
of wider gaps, by convection.
The maximum design temperature allowed in the bentonite buffer is +100 °C, but in
design analysis there is used a 5-10 °C safety margin due to natural variations of thermal
properties of rock and the uncertainty in decay heat estimate. Thus the calculated
maximum operational temperature on the canister-bentonite-buffer interface in repository
condition is nominally set to 95 °C. This design limit is set in order to ensure the
chemical stability of the bentonite in the deposition hole. The thermal conductivity of
copper is about 391 W/m/K and that of cast iron is about 36 W/m/K. The thermal conductivity
of the copper body of the canister is two orders of magnitude higher than the
conductivity of the surrounding bentonite (1 W/m/K) and rock (2.82 W/m/K) in the repository
(for metal properties, see Table 14 in Section 7.5). Therefore, the copper canister
surface will be practically at a uniform temperature and the entire thermal gradient
will be transferred to the bentonite and rock around the canister and to possible air gaps
between the material interfaces (Ikonen & Raiko 2012).
For dimensioning purposes, the bentonite buffer is assumed to be in initial condition
and a 10 mm air-filled gap is assumed between the canister outer surface and the buffer
inside surface. Figure 28 shows the effect between the initial condition buffer and the
saturated buffer on the maximum temperature of the canister-buffer interface.
If the gap between the canister and buffer is open and the buffer in initial condition, the
temperature of the canister surface can be, at maximum temperature, 15-20 °C higher
than in case of solid contact and saturated buffer. The rock temperature at the edge of
the deposition hole does not depend on the heat transfer conditions between the canister
and the buffer. The maximum temperature of the canister will be reached after 10 to 15
years and that of the rock at the edge of the hole will be reached after about 60 years.
After about 600 years, the canister temperature goes below 50 °C and the effect of
buffer condition (dry or saturated) on canister temperature is only 3 °C.

90
In summary, the thermal evolution in the repository can be estimated from the two separate
analyses assuming different buffer conditions (see Figure 28): First, the temperature
follows the red line as far as the groundwater starts to reach the buffer. When time goes
forward, the saturation grade and conductivity of the bentonite buffer and the pellet slot
will increase and the temperature starts to tend to the blue line. Simultaneously the
ground water pressure starts to develop and reach finally the maximum 4.0 to 4.2 MPa.
Parallel to bentonite saturation, the bentonite swelling pressure is developed and
reached at full saturation the maximum, 2 to 10 MPa, depending on the final density of
bentonite. Thereafter, the temperature of canister will follow the blue line, the rock edge
of the deposition hole will follow the black line and the bentonite buffer temperature
will be between them. The curves in Figure 28 are valid until the climate on the ground
surface does not remarkably change. The onset of the first cold period is expected at
about 50000 years with temperature and precipitation changes leading to first permafrost
development and later on to ice-sheet growth and advance (Pimenoff et al. 2011).
At that time, the residual temperature effects of the decay heat are less than 5 °C at the
repository level.
TEMPERATURE (C)
100
90
80
70
60
50
40
30
20
10
0
CANISTER/BUFFER INTERFACE TEMPERATURE
Buffer in initial condition
Saturated buffer
Buffer/rock interface
1 10 100 1000 10000 100000
TIME (years)
Figure 28. Canister surface temperature estimates in repository (central area) using
the two extreme saturation degrees for the bentonite buffer. EPR canister, average burnup
50 MWd/kgU, canister distances in repository 10.5 m / 25 m, buffer conductivity is
1.0 W/m/K in initial condition and 1.3 W/m/K in saturated condition. In initial condition,
there is a 10 mm air gap between the canister and the buffer, in saturated condition
the gap is closed. The outer 50 mm gap between buffer and rock is assumed to be
filled with bentonite pellets that have conductivity of 0.2 W/m/K in initial condition and
0.6 W/m/K when saturated. The figure is according to results of (Ikonen & Raiko 2012).

91
8.4.4 Canister during permafrost
As stated in Section 4.4, for canister design assessment purposes, permafrost is assumed
to extend down to the repository level. The temperature is assumed to be -5 °C, at lowest.
The freezing of bentonite and the water in it may lead to changes in swelling pressure load
on the canister. The maximum compression capacity of freezing of water only is first considered
according to the pressure-temperature (p-T) diagram of water. This assessment is
made with very conservative assumptions:
The rock is extremely rigid, strong and leak-tight
The solid material proportion of the buffer is ignored
The increase of iron yield strength from room temperature to freezing point is ignored
(maybe 5 %).
The p-T-diagram of pure water shows that the freezing point of water decreases as the
pressure increases up to a pressure level of 200 MPa. At temperature -5 °C the respective
freezing/melting pressure is 60 MPa. This is very pessimistic estimate for the
maximum possible swelling pressure due to freezing at -5 °C, because it ignores the
major effect of solid bentonite fraction in the buffer.
Figure 29. The pressure along the melting and sublimation curve of ordinary water
substance. The pressure at T = -5 °C is 60 MPa (Revised Release on the Pressure along
the Melting and Sublimation Curves of Ordinary Water Substance 2008).
There are laboratory tests made on bentonite freezing and the results show, however,
that the actual swelling pressure from bentonite at -5 °C is remarkably lower than that of
freezing of pure water. Laboratory tests results and theoretical considerations are given
in (Schatz & Martikainen 2010; Birgersson et al. 2010). From a safety assessment point

92
of view, the findings indicate that possible lowering of repository temperature down to
-5 °C repository will not impose a problem. For a typical buffer swelling pressure of 7
MPa, the present results show that the critical temperature at where the swelling pressure
is lost is below -5 °C. Thus, actual freezing will not occur above -5 °C and no high
pressures are expected. Swelling pressure will only be lowered when the surrounding
rock is already frozen and advective transport mechanisms are deactivated. The swelling
pressure will also be regained before the surrounding rock is thawed.
The results of testing (Schatz & Martikainen 2010; Birgersson et al. 2010) show that the
swelling pressure of bentonite loading the canister in repository is not increasing but
decreasing, if the temperature goes down to -5 °C. In even lower temperature than -5 °C
the ice formation and pressure increase is detected.
8.4.5 Thermal behaviour of the variant canister designs
The variant canister designs (VVER-440 and EPR/PWR) are analysed separately as
individual cases with their actual geometry and decay heat for internal temperatures in
(Ikonen 2006). The design value for the allowable decay heat is calibrated so that the
allowable decay heat is linearly proportional to the relative cooling surface area of the
variant canister, if the decay heat of 1700 W is defined for the reference canister in
(SKB 2009). This analysis (Ikonen 2006) is primarily made to give a conservative assessment
of the fuel temperature inside the canister in repository conditions. The result
is that the fuel temperature is about 200 to 250 °C, at maximum, during the early decades
of disposal. In longer perspective, the fuel temperature goes down as the decay
power is decreased.
8.5 Cooling of the canister in all expected conditions
8.5.1 Canister in encapsulation plant
The canister components are typically at room temperature during the spent fuel encapsulation.
The thermal capacity of the canister is high, about 10 MJ/K (consisting of 13.6
ton iron or steel + 7.3 ton copper + 3.6 ton of fuel elements). This means that the 1700
W decay heat needs (10 MJ/K)/(1700 W) = 5880 s/K (=1.6 h/K) to warm up all the canister
mass of 1°C if ignoring the heat losses of cooling. This means that the thermal response
of the canister is slow in case of loss of cooling. In other words, the canister
temperature increases only 15 °C per day and night even if the cooling is lost totally
(i.e., in case of adiabatic insulation). The stationary cooling condition of a canister in
natural convection condition in an air conditioned room is reached only after several
days.
In stationary cooling condition, in an air-conditioned environment, the cooling of a canister
takes place through combined radiation and conduction of the natural circulation of
air on the cylindrical surface of the canister and the horizontal plane of the top end. The
bottom end is usually not in free contact with the air but covered by support structures
and it is ignored as cooling surface in the following estimates.

93
The cooling surface area of a reference canister is 15.7 m 2 from the cylinder and 1.1 m 2
from the lid, totalling 16.8 m 2 . The average heat transfer coefficient (h) of an up-right
cylinder surface due to natural convection of air in room temperature can be estimated
according to (Incropera & DeWitt 1996, example 9.10 on page 519) as
h(convection = ~1.4*{T/L} 0.25 (10)
where L is the height of the cylinder (4.7 m) and T is the temperature difference between
the surface and the surroundings. The area of bottom lid is ignored because the
canister may stand on a structure that may decrease the cooling of the bottom. The additional
part of the total heat transfer comes from radiation that can be estimated according
to Stefan-Bolzmann law with small temperature differences from the formula (11)
adapted from (Tekniikan käsikirja 1970), Vol. 5, formula (30) on page 396, and converting
into SI-units,
h(radiation) = ~ 4.6*10 -8 *T 3 (11)
where T is the average temperature of the surface and the surroundings in Kelvin degrees.
The emissivity of the copper surface is assumed to be 0.2 as reasoned in Section
8.4.2 before. Typical surface emissivities are listed in Table 18. From these formulas we
get the effective heat transfer coefficients as a sum and we can calculate easily the increase
of surface temperature of the canister, when the stationary decay power is the
maximum, 1700 W. The results are presented in Table 19. At stationary cooling condition
the surface temperature will be 322 K (49 °C) and the thermal heat transfer coefficient
is 2.20 W/m 2 K from natural convection and 1.34 W/m 2 K from radiation, totalling
3.5 W/m 2 K.
Table 18. Typical surface emissivities (Cole-Palmer Technical Library 2010).
Surface Quality Emissivity
Cast iron Oxidized 0.64
Copper Matte 0.22
Copper Black, oxidized 0.78
Paints All colours 0.92 – 0.96
Plastics Average 0.95
Concrete Rough 0.94
Rock Granite/Mica 0.45/0.75
Sand - 0.76

94
Table 19. The heat transfer coefficients of an up-right canister in natural convection
added by radiation condition at room temperature environment. The stationary maximum
decay heat power P = 1700 W is reached when interpolated from results to surface
temperature increase of 29 °C that is equal to 49 °C, if the environment is typically
in 20 °C. The emissivity of copper overpack surface is estimated to be 0.2.
T
(K)
T
(°C)
T
(K)
h(conv)
(W/m 2 K)
h(rad)
(W/m 2 K)
h(total)
(W/m 2 K)
P(total)
(W)
303 30 10 1.69 1.22 2.91 489
308 35 15 1.87 1.25 3.12 787
313 40 20 2.01 1.28 3.29 1107
318 45 25 2.13 1.32 3.44 1446
323 50 30 2.23 1.35 3.57 1801
328 55 35 2.31 1.38 3.69 2172
333 60 40 2.39 1.42 3.81 2558
338 65 45 2.46 1.45 3.91 2957
343 70 50 2.53 1.48 4.01 3370
This means that the surface temperature of a single canister will be typically about 50
°C or less in all phases in encapsulation plant operations or in canister buffer storage.
The canister buffer storage has a special air-conditioning (forced circulation and cooling)
system that keeps the room temperature constant independently of the amount of
decay heat of the variable number of stored canisters. However, in buffer storage a canister
may be surrounded by other similar canisters thus the radiation cooling may be lost
almost totally. In such a case, only the natural convection to cool the canister can be
used. From the numbers in Table 19, above, the difference in temperature is T =
42 °C, if all radiation cooling is ignored and thus thermal heat transfer coefficient is
2.42 W/m 2 K. This means that the maximum surface temperature of a canister may be up
to 42 + 20 = 62 °C in the buffer storage in case the canister is surrounded from all sides
with equal canisters. This is a conservative assumption, because the top end of each
canister can radiate in all cases and the surrounding canisters do not shield the innermost
canisters completely.
The canister surface temperature in encapsulation is therefore allowable up to 100 °C
(as in repository) and there are good margins before the fuel temperature inside canister
will become harmfully high. In reality, the canister buffer storage in encapsulation plant
is air-conditioned with forced air circulation, thus the cooling is remarkably better than
the calculated natural circulation case. Canister behaviour during loss of air conditioning
in the canister storage will be analysed in the future air conditioning report for the
plant.

95
8.5.2 Canister in transfer vehicle
When the canister is transferred in the repository level, the canister is on a vehicle and
inside a heavy radiation shield cylinder made of iron or steel. The cylinder has wall
thickness of 150 mm and the cylinder is lined on inside with a 50 mm polyethylene
layer containing boron as neutron absorber. Between the canister and the inside of the
plastic absorber there is an air gap of 10 mm in average varying between contact and 20
mm. The thermal conductivities of the materials are selected conservatively according
to Table 14 in Section 7.5. The emissivities of various surfaces are estimated according
to values given by Cole-Palmer Technical Library, as quoted in Table 18. The cooling
chain is described in Table 20.
Table 20. Thermal conductivities and surface emissivities of the cooling chain in case of
the canister inside the radiation shield of the canister transfer vehicle.
Sequence of
materials and
interfaces
Thermal conductivity
(W/m 2 K)
Surface heat
transfer coefficient
h (W/m 2 K)
Emissivity
(-)
Thickness
t (m)
Copper
overpack
Air gap
(average)
Polyethene
neutron
absorber
Cast iron
shield
cylinder
Air (conditioned)
386 0.030 0.3-0.4 36 0.030
- - - Due to
convection
only
- -
Rock or
concrete
walls
0.2 0.95 Painted
0.45-0.94
>0.9
0.05 0.010 0.05 0.15 >1 -
The effective thermal resistance R, the inverse of conductance, (R = {t/) for the
cooling chain from the canister outer surface to the outer surface of the radiation shield
is (using simple 1-dimensional calculation that is conservative in axisymmetric condition)
as follows
R = (t/) = (0.010/0.03 + 0.05/0.3 + 0.15/36) = 0.50417 mK/W (12)
where means summation over consequent material layers;t and are the respective
thickness and thermal conductivity of the material layer in question. The T from thermal
flow balance in stationary condition (taking only conduction into account in the air
gap estimation and omitting radiation and convection) can be calculated according to
(Tekniikan käsikirja 1970, page 398, formula (38)), as follows
P = A * T/R (13)

96
where A is the gap area (cylinder + both ends) (17.6 m 2 ), R the thermal resistance
(0.50417 mK/W) and P is the canister decay power (1700 W). Solving T from formula
(13) above we get T = 1700/17.6*0.50417 = 49 K. This is the conservative temperature
decrease from the canister outer surface to the outer surface of the radiation shield
in stationary cooling condition.
On radiation shield cylinder surface in the vehicle we can estimate that the heat transfer
coefficient is at least 5 W/m 2 K according to the single canister calculations in Table 19
above. In this case, the radiation component of the heat transfer is roughly 3-4 times
higher, because of the painted iron surface of the vehicle has the emissivity of >0.9 and
the surfaces of the surroundings (sand, rock or concrete) have also high emissivity of
0.45-0.94, which leads to a total emissivity (Tekniikan käsikirja 1970, formula (26),
page 396), of 0.43 to 0.86, as for the machined and slightly oxidised copper surface of
the canister has the emissivity of 0.2 as mentioned in 8.5.1. The convective heat transfer
of a horizontal cylinder with these dimensions is also a little better than that of a vertical.
The outside surface area of the radiation shield (L=5 m, D=1.45 m) in the vehicle is
about 26 m 2 . Then, using very conservative assumptions it can be calculated that that
the temperature increase on the radiation shield cylinder surface is T = 1700 W / (5
W/m 2 K * 26 m 2 ) = 13 K at stationary cooling condition, when the ultimate heat sink is
the air cooled atmosphere and the rock or concrete walls of the repository, which are
assumed to be in constant +20 °C temperature.
Summing all the temperature increments from environmental (room) temperature in the
repository up to the surface of the canister inside the radiation shield, we get 20 + 13 +
49 = 82 °C. Even if this seems to be quite high number, the radiation heat transfer is
increasing comparable to the T 4 , T being the absolute temperature, so the balance will
be reached with very moderate differences, if there are some inaccuracies in the estimated
properties.
The allowable canister surface temperature (the calculated maximum is up to 82 °C in
stationary condition) inside the radiation shield of the transfer vehicle is up to 100 °C, at
least, as in repository. There are no such uncertainties that this limit will be met and,
even then, there are good margins before the fuel temperature inside canister will become
harmful. Now we have calculated all temperatures as a stationary condition, but in
practise, the operational time during transfer and other encapsulation operation phases is
so short that stationary thermal cooling condition is hardly ever achieved with the exception
of canister buffer storage.
8.5.3 Fire
The canisters are transferred inside a heavy radiation shield cylinder in a rubber
wheeled transfer vehicle in the technical area and tunnels of the repository. The fire
resistance of the canister was analysed in (Lautkaski et al. 2003). According to this
analysis, the spent fuel in transportation shield will endure a fire with a flame temperature
of 1 000100 °C for 2.5 to 3 hours without any resulting damage of the fuel or canister
itself. The fire resistance of the canister inside the radiation shield is remarkably
more than the total available energy of a burning transfer vehicle can produce. A typical
simulated fire temperature field around the vehicle that is burning in a repository tunnel
is shown in Figure 30.

97
Figure 30. Gas temperature around transfer vehicle at t=500 seconds after the ignition
of the fire (Lautkaski et al. 2003).
8.5.4 Canister in repository
The thermal dimensioning of the repository is made in (Ikonen & Raiko 2012). The
report contains the temperature dimensioning of the KBS-3V type nuclear fuel repository
in Olkiluoto for the BWR, VVER and EPR/PWR fuel canisters, which are disposed
at vertical position in the horizontal tunnels in a rectangular geometry according to the
preliminary Posiva plan. This report concerns only the temperature dimensioning of the
repository and does not take into account the possible restrictions caused by the stresses
induced in the rock.
The far field rock acts as a temporary heat sink absorbing and storing the thermal energy
for a few thousands of years. The ground surface acts as the ultimate heat sink,
dispersing the excess heat by convection or by radiation to the atmosphere for a few
thousand years. Figure 28 shows the temperature evolution of the hottest canister in the
middle of a repository block assuming different buffer conditions. The decay heat in the
canisters is halved in the beginning about every 40 to 50 years and, depending on the
heat removal rate, the maximum temperature in fuel, canister and the near field is
reached in 15-20 years after disposal. The maximum rock temperature in near-field
around a canister is reached in about 50 to 60 years, see Figure 28. The thermal decay
from the canisters will no longer have a remarkable impact on the repository system in
approximately 10000 years after emplacement when the decay is some 13 W/tU. Some
of the heat from the canister is stored in the rock even longer due to the relatively low
thermal conductivity of the Olkiluoto rock.
The maximum temperature on the canister-bentonite interface is limited to 100 °C in
Design Basis report. However, due to uncertainties in some thermal analysis parameters

98
(like scattering in rock conductivity or in predicted decay power) the nominal calculated
maximum canister-bentonite-buffer interface temperature is set to 95 °C in (Ikonen &
Raiko 2012) giving a safety margin of 5 °C.
The temperature on the canisters and in the repository is controlled by selecting the fuel
elements for encapsulation in a way to control the heat generation inside the canister
(decay heat) and adjusting the space between adjacent canisters, adjacent tunnels and
the pre-cooling time affecting on power of the canisters. The temperature of canister
surfaces can be determined by superposing analytic line heat source models much more
efficiently than by numerical analysis, if the analytic model is first calibrated by numerical
analysis (by control volume method). This was done by comparing the surface
temperatures of a single canister calculated numerically and analytically.
For the Olkiluoto repository, one dimensioning panel having 900 canisters of BWR,
VVER or EPR/PWR spent fuel was analysed. The analyses were performed with an
initial canister decay power of 1700 W, 1370 W and 1830 W, respectively. These decay
heats are obtained when the pre-cooling times of the fuels are 32.9, 29.6 and 50.3 years
(corresponding the burnup values 40, 40 and 50 MWd/kgU, respectively). The analyses
gave as a result the canister spacing (6.0-10.5 m), when the tunnel spacing was 25 m, 30
m or 40 m. The canister maximum temperature inside a repository panel is not
depending on the size of the panel, if the panel is at least some hundreds of canisters.
The minimum canister distances are presented for reference canister (BWR type) in
Figure 31.
At the farthest edges of the panel with constant canister spacing the temperatures of the
canisters are somewhat lower than in the middle area of the repository. Thus it is possible
to pack the canisters denser on the edge areas of the panel.
Figure 31. Maximum BWR canister surface temperature as a function of canister spacing,
when tunnel spacing is 25 m, 30 m or 40 m and burnup of the spent fuel is 40
MWd/kgU. Initial decay power is 1700 W. Data are from (Ikonen & Raiko 2012).

99
8.5.5 Cooling capacity of variant canister designs
The variant canister constructions are analysed separately as individual cases with their
actual geometry and decay heat for repository conditions with their respective distances
in deposition tunnels in (Ikonen & Raiko 2012). The respective results are also summarised
in Section 8.5.4.
As for the cooling of variant canisters during encapsulation, buffer storage and transfer
in the repository, the analyses given in 8.5.1 and 8.5.2 are valid as such, because the
decay heat in each variant is directly proportional to the canister surface area, and respectively,
the cooling is directly proportional to the canister surface area, too. Thus the
thermal analyses are valid for all canister variants.
As for the fire case reported in 8.5.3, the margins in the allowable fire duration are so
large that the small variations in the canister thermal capacity per surface area –ratio
(that is proportional to the heat response) do not affect the conclusions concerning the
design acceptability. The ratio of canister’s cross mass/outside surface area is 4.6, 4.6
and 5.1 t/m 2 for BWR, VVER-440 and EPR/PWR canisters, respectively. This means
that the fire resistance time (estimated as the ratio between thermal capacities per surface
area) of BWR and VVER-440 canisters is equal and that of EPR/PWR a little
higher.
8.6 Corrosion resistance
The corrosion of the canister is a key process that needs to be understood for the canister
safety function of containment. Much is known about the general corrosion behaviour
of copper under repository conditions. Detailed thermodynamic analyses of possible
corrosion reactions have been performed, particularly in the Swedish and Finnish
programmes. In Canada, more emphasis has been placed on kinetic studies under wellcontrolled
mass-transport conditions. Combined, these complementary approaches provide
a detailed understanding of the general corrosion behaviour of copper canisters
under the evolving conditions expected in a repository. The results of laboratory studies
have been confirmed by the observations from long-term in situ tests under relevant
conditions in underground research laboratories.
Copper corrosion conditions vary during the various phases of the canister lifetime.
These are described below and summarised in Section 8.6.8. Corrosion processes are
described in detail in Features, Events and Processes report.
8.6.1 Atmospheric corrosion in the encapsulation plant
The atmospheric corrosion of the copper shell during the storage time before emplacement,
estimated to be a couple of months at most, is negligible in spite of the elevated temperature
of about 60-70 °C in the storage facility. A layer of copper oxide with a thickness of
a few tens to a few hundreds of nanometres will form on the canister surface. Even if
the storage time would extend up to 2 years, the total corrosion attack would be less that
1 m (King et al. 2011b, Section 4.1.2).

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8.6.2 Corrosion during repository operation
Damage of the canister surface caused by handling during emplacement is unlikely to
significantly affect the corrosion behaviour. Scratches and other defects in the surface
oxide caused by handling would rapidly oxidize when exposed to the repository environment
until the protective oxide layer has been reformed. Neither would handling
introduce stress raising defects of sufficient size to cause cracking in the absence of a
suitable environment for stress corrosion cracking. Plastically deformed material is
known to corrode more rapidly than unstrained material. However, this localized corrosion
would stop once the deformed material had been consumed. Copper is not susceptible
to galvanic corrosion due to embedded iron particles resulting from the use of steel
handling equipment. In fact, iron particles would temporarily galvanically protect the
canister surface (Gubner & Andersson 2007).
8.6.3 Corrosion in the repository under unsaturated, oxic conditions
Once the canister has been emplaced in the deposition hole there will be trapped atmospheric
oxygen in the gaps and in the bentonite surrounding it. In the early evolution,
trapped atmospheric O 2 is consumed by (i) corrosion of the canister, (ii) reaction with
oxidisable mineral impurities and sulphide in the clay, and (iii) microbial activity.
Therefore, the initial oxic conditions in the deposition hole will become progressively
anoxic and reducing. As the initially trapped O 2 is consumed, the rate of corrosion will
become limited by the diffusion of O 2 (or of Cu(II) formed by the homogeneous oxidation
of Cu(I) by O 2 ) to the canister surface.
The maximum possible corrosion attack from residual O 2 can be estimated from mass
balance considerations. The total volume of buffer and backfill in the deposition tunnel
and the deposition hole is assumed to be 150 m 3 per canister. The porosity in the bentonite
and the backfill material can be conservatively estimated to be 40 %. If all of this
porosity consisted of air, the amount of O 2 per canister would be 12.5 m 3 , or approximately
560 moles. Assuming that Cu 2 O is formed as the corrosion product, 2240 mol
of copper or 140 kg could be oxidised. This corresponds to a maximum depth of corrosion
of 840 m evenly distributed over the canister surface (SKB 2006b, Section 3.5.4).
In reality the corrosion will be considerably smaller since residual oxygen will also be
consumed through reaction with accessory minerals in the buffer and backfill and
through microbial activity (SKB 2010a; SKB 2010b).
The uneven swelling of the bentonite buffer around the canister could lead to localised
corrosion. Unless the bentonite is pre-wetted with an artificial supply of water before
backfilling the tunnel, the natural wetting process will most probably result in uneven
swelling of the buffer. As a consequence, the gap between the canister and the buffer
may close in some areas while it remains open in others. Those sites where the bentonite
first contacts the copper canisters are possible locations for the spatial separation
of anodic and cathodic processes. However, since the gaps close gradually as the bentonite
swells, the location of these sites will not be constant. Once the bentonite has
reached full saturation, the whole canister surface will have been exposed to these conditions.
The fact that some sites have been exposed to conditions that enable electrochemical
corrosion longer than others may cause slightly uneven corrosion. Apart from

101
that, the gradual closing of the gaps is not likely to result in any significant localised
effects.
Various experimental and modelling approaches have been developed to study the localized
corrosion of copper. Although the extensive database on the pitting of copper
water pipes provides some useful mechanistic information, the results of corrosion experiments
under simulated repository conditions suggests that canisters will not undergo
classical pitting, but rather a form of surface roughening, in which there is no permanent
separation of anodic and cathodic sites. The current approach for the copper corrosion
prediction model is to assume a certain degree of roughness (+/- 50 micrometres) representing
localized corrosion which is then added to the predicted depth of general corrosion
(SKB 2006b, Section 3.5.4).
The mechanistic copper pitting studies indicate that an oxidant (either O 2 or Cu(II)) is a
pre-requisite for pit propagation. Since the near-field environment in the repository will
evolve from initially oxidising to ultimately anoxic and reducing, this implies that pitting
will only be possible (if at all) in the early stages of the repository life. Pit propagation
may also occur in waters with high sulphate concentrations, Sulphate is aggressive
towards copper in a corrosion pit because it forms a complex with divalent copper. However,
since classical pitting is not expected to occur even during the aerobic phase sulphate,
sulphate will not propagate pitting and it is almost inert with respect to the general corrosion
(King et al. 2011b, Section 5.3.3). Thus, the environment within the repository is
evolving to one in which only general corrosion will occur. In addition, the difficult
problem of predicting localised corrosion is made easier by the fact that predictions only
have to be made for the early oxidising period.
As the saline or brackish groundwater (the salinity at initial state in the repository is
expected to be about 10-12 g/l and during the first thousands of years expected to decrease,
e.g. Löfman et al. 2009 saturates the buffer, the pore-water Cl - concentration will
gradually increase. Chloride ions are known to support general corrosion of copper. It is
reasonable to assume that the copper chlorides/hydroxy chlorides may also form as an
initial corrosion product in saline ground waters during the water saturation phase in
compacted pure bentonite. Although localised corrosion may initiate, it is unlikely to
propagate very deeply, mainly because the amount of O 2 trapped in the buffer material
is limited. Furthermore, as the relative humidity increases the surface will become more
uniformly wetted as less-deliquescent contaminants absorb moisture from the atmosphere.
Eventually the entire surface will be wetted and the differential O 2 concentration
cell that acted initially as the driving force for localised corrosion will disappear. At this
stage the surface will take on a generally roughened appearance (King et al. 2011b),
Section 5.
King et al. (2011b) considered the effect on the localised corrosion of copper of an increase
in pore-water pH due to an alkaline plume from cementitious material in the repository.
If the pore-water pH increases prior to the establishment of anoxic conditions,
the canister surface will passivate as the pore-water pH exceeds a value of ~pH 9. Passivation
will result from the formation of a duplex Cu 2 O/Cu(OH) 2 film. The corrosion
potential will be determined by the equilibrium potential for the Cu 2 O/Cu(OH) 2 couple
under oxic conditions, or by the Cu/Cu 2 O redox couple under anoxic conditions (in the

102
absence of sulphide). Pitting corrosion is only likely to occur early in the evolution of
the repository environment, whilst the canister is still relatively cool (100 Gy/h) and temperatures. At dose rates in the range of 10-100
Gy/h, the experimental data seem to indicate a lower corrosion rate in the presence of
radiation. The maximum surface dose rate for the canister is set by design at 1 Gy/h
(Chapter 3).
8.6.4 Corrosion in the repository under saturated as well as anoxic and reducing
conditions
Eventually, conditions will become anoxic and reducing for copper. The available experimental
evidence suggests that only general corrosion is expected under anoxic conditions
(King et al. 2011b). The most important parameters controlling the rate of general
corrosion are: the rates of mass transport of species to and from the canister surface,
the influx of Cl - ions from the groundwater, and the supply of sulphide ions to the canister.
In a sealed deposition hole, the extent of general corrosion is limited by the general lack
of oxidants. As mentioned earlier, trapped atmospheric O 2 is consumed during the saturation
phase. Under reducing conditions, corrosion will be supported by the slow supply
of sulphide to the canister surface.

103
The corrosion depth due to sulphide corrosion during the long-term evolution of the
canister depends on the assumptions on the origin of sulphide ions and the type of modelling
approach used (mass-balance/thermodynamic or kinetic approach). Sulphide is
present naturally in the groundwater at repository depth at Olkiluoto generally in concentrations
of less than 1 mg/L, the highest measured sulphide content is about 12 mg/L
at a depth of 367 m in borehole KR13. Other sources of sulphide may be the pyrite in
the buffer and backfill. These will contribute to corrosion of the copper but the corrosion
rate will be limited by diffusion through the bentonite barrier to the canister surface
(Pedersen 2010). A higher rate of diffusion of sulphide to the canister would, however,
be expected in the case of a defective buffer layer around the canister but this case is not
a design basis case although it is considered in the long-term safety assessment.
Copper corrosion depths in repository conditions can be predicted using thermodynamic
and kinetic approaches (King et al. 2011b, Section 5.2). A 1-D reactive transport model
has been developed to predict the evolution of the general corrosion behaviour of the
copper canisters taking into account the initially trapped O 2 in the proximity of the canister
as well as the contribution of chloride and sulphide ions to corrosion (King et al
2011c). Various sources of sulphide are considered in this model, including the microbial
reduction of sulphate in the buffer and backfill, the dissolution of pyrite impurities
and the groundwater sulphide itself. Different assumptions are made concerning the
value of different input parameters: e.g., groundwater sulphide or chloride concentrations,
microbial activity levels, rates of repository resaturation, the presence or absence
of backfill. The assumptions used for the Olkiluoto corrosion predictions were: [HS - ]
=12 mg/L (0.4 mmol/L), [Cl - ]=16 g/L (0.45 mol/L), transport through buffer, backfill,
rock, EDZ and a period of 200 years to resaturate the entire repository. The predicted
corrosion depth is up to 1 mm in 10 6 years in all but one scenarios (King et al. 2011c).
The vast majority of this corrosion occurs under anaerobic conditions (due to the formation
of Cu 2 S), with minimal corrosion predicted under aerobic conditions. The results of
the simulations suggest that the microbial reduction of sulphate in the backfill material
is the most important source of sulphide. In comparison, SRB activity in the bentonite,
diffusion of sulphide from the ground water and the dissolution of pyrite contribute relatively
little to the corrosion of the canister.
The main uncertainty is the availability of sulphide on the canister surface, depending
on the groundwater concentrations and on microbial activity. Microbially-induced sulphide
production in the buffer and in the backfill remains a subject for further studies.
However, microbial activity in the buffer is expected to be limited as long as the buffer
remains within the allowed density limits.
The conditions are expected to remain reducing indefinitely. Oxygen intrusion at repository
depth in conjunction of the melting of an ice sheet in the vicinity of the repository
is considered in the long-term safety assessment but it is not part of the design basis due
to the application of rock suitability criteria (RSC) that would lead to the exclusion of
canister positions from transmissive fractures that could have a potential for oxygen
transport. Even in the event of such O 2 intrusion, the supply of O 2 reaching the canister
would be limited, as during the early evolution of the canister.

104
Szakálos et al. (2007) raised the possibility of copper corrosion in pure water. The authors
claim to have observed a transition from O 2 -consuming to H 2 -evolving corrosion
of OFHC copper in deionised water. This corrosion process has been addressed in King
& Lilja (2011a). The results of Szakálos et al. (2007) are judged unreliable for the time
being since nobody has been able to reproduce the results and experimental conditions
were not thoroughly described. The possibility of copper corrosion by pure water is being
investigated in the framework of a Posiva-SKB cooperation project and in the independent
KYT programme. The aim is to repeat the test conducted by Szakálos et al.
(2007) under very low oxygen conditions and to verify whether the same phenomenon
can be observed.
8.6.5 Stress corrosion cracking
There is extensive experience with, and knowledge of, the SCC of copper alloys. Considerable
effort has gone into studying the mechanism of SCC of copper canisters.
These are described King and Newman (2010) and summarised in King et al. 2011b
(Chapter 6).
The three pre-requisites for SCC are a susceptible material, a tensile stress, and a suitably
aggressive environment. The proposed canister material cannot be claimed to be
immune to SCC since pure copper has been shown to be susceptible, especially phosphorous-containing
alloys. Tensile stresses on the canister surface are possible during
various stages in the evolution of the repository environment, either due to external
loads or from residual manufacturing stresses. Finally, it is not possible to exclude the
possibility that known SCC agents, i.e., ammonia, nitrite, sulphide or acetate, may be
present in the repository. Therefore, the possibility of SCC of copper canisters must be
considered.
As with other forms of corrosion, the available evidence indicates that the probability of
SCC of a copper canister will diminish with time as the repository environment evolves
from oxic to anoxic conditions. The period of highest SCC susceptibility is not known
with certainty, but is likely to be of the order of less than ten years.
Oxidants will be available in the form of trapped atmospheric O 2 and/or Cu(II) produced
by corrosion of the canister. Ammonia, and possibly acetate ions, will be present
in the groundwater and, possibly as a result of human activity during construction
(Saario et al. 1999), although it is highly unlikely that sufficient nitrite will be present to
cause SCC. Microbial production of ammonia, acetate and nitrite ions has also been
taken into account. Furthermore, the inhibiting effects of Cl - ions will not be fully felt
until the bentonite pore water has equilibrated with the groundwater. During this early
period, the outer copper shell may also be subject to considerable strain as the hydrostatic
load develops and the copper shell is deformed onto the inner cast iron insert. As
the available oxidant is consumed, as the pore water becomes more saline, and as the
buffer material saturates and restricts the transport of SCC agents from the groundwater
to the canister surface, the probability of SCC will diminish accordingly. Only the decrease
in repository temperature with time will tend to render the canister more susceptible
to SCC.

105
Although there is a higher probability of SCC during the initial warm, aerobic period,
there is some evidence to suggest that cracking can be sustained during the long-term
anaerobic phase in the absence of oxidants due to the presence of sulphide at sufficiently
high concentration (>5 mmol/L) (Taniguchi & Kawasaki 2008).
The work by Taniguchi & Kawasaki (2008) requires further investigation because, if the
claim is correct, cracking may then be possible during the long-term anaerobic period.
King and Newman (2010) argued that sulphide would not lead to cracking of the copper
overpack in repository conditions by any of the known SCC mechanisms. The rate of
general corrosion during the anaerobic phase will be determined by the rate of supply of
sulphide to the canister surface, partly because of the presence of highly compacted bentonite.
Therefore processes that have been shown to be possible in the bulk of sulphide
solutions are not necessarily possible at the copper surface because of the gradient in
sulphide concentration throughout the buffer.
8.6.6 Corrosion inside the canister
Corrosion inside the canister can occur only if residual water is present. Any residual
aerated water from the spent fuel pool water introduced with the fuel at the time of encapsulation
will be decomposed by radiolysis to generate small quantities of nitrogen
oxide species and even smaller quantities of hydrogen gas (H 2 ), oxygen gas (O 2 ), and
hydrogen peroxide (H 2 O 2 ). The products of radiolysis of aerated water or vapour will
then be converted to corrosive species such as nitric or nitrous acid.
Further reaction of these acids with the cast iron canister insert and the copper internal
surfaces could lead to accelerated corrosion.
To estimate the maximum amount of nitric acid that could be formed a sufficiently pessimistic
assumption on the residual amount of water in the canister is needed. To this
end, the residual water volume is assumed to be 600 g, which corresponds to the void
volume of a fuel rod (50 cm 3 ) assuming 12 leaking rods per canister. This is an extreme
assumption based on the number of leaking fuel rods expected and on the fuel encapsulation
process. With this amount of residual water, the radiolytic acid production yield
has been estimated to be between 1 and 450 g per canister depending on the conceptual
model, data and assumptions applied (Henshaw et al. 1990; Marsh 1990; Henshaw
1994; SKB 2010, Section 2.5.2). These nitric acid amounts are negligible compared to
the thickness of the copper overpack and the availability of other metallic structures in
the canister (King et al. 2011, Section 7.2).
Nevertheless, to reduce the risk of nitric acid production, all fuel elements are dried in a
drying unit using a combination of elevated temperature and vacuum. The air in the
insert will also be replaced by argon gas (Section 11.1).
Galvanic interaction between the copper shell and the insert has been considered given
that the copper shell is expected to be pressed against the insert by the swelling bentonite.
At the points of contact between the copper and the iron there could be galvanic
interaction because there is an ionic contact with an electrolyte (i.e. groundwater). The
external pressure applied by the swelling bentonite would increase the number of con-

106
tact points between the inner and outer parts. The possible effect of galvanic corrosion
due to anaerobic corrosion of the steel was investigated experimentally in Smart et al.
(2005) by measuring the hydrogen production rate. It was found that there was no significant
difference in the gas generation rate between the coupled and uncoupled specimens
and hence no galvanically enhanced crevice corrosion was observed when copper
and cast iron were coupled in deoxygenated conditions.
Stress corrosion cracking of the insert may occur due to static tensile stress on the cast
iron insert in the presence of corrosive chemical species (e.g. radiolytically generated
nitric acid, see above). Under the expected conditions in the repository, the canister will
be under uniform external pressure due to bentonite swelling and tensile stresses will
occur only on small, localised areas, and is not thought to be a significant contributor to
corrosion. Furthermore, the amount of radiolytically produced oxidants is negligible, as
discussed above.
General corrosion of the cast iron insert cannot occur until the copper shell has failed
and water can penetrate into the canister. In the long-term safety assessment, the insert
is expected to corrode and this is taken into account in the expected canister evolution.
Anaerobic corrosion of iron will generate magnetite (Fe 3 O 4 ) as the most likely corrosion
products and hydrogen gas, along with small concentrations of dissolved Fe(II). Because
general insert corrosion can occur only after the copper shell has already failed,
this process is not considered in the design bases for the canister.
8.6.7 Corrosion in the weld and grain boundaries
In any welded structure, the regions around the weld, including the weld material itself
as well as the heat-affected zone, can be locations of enhanced corrosion susceptibility.
Proper attention must be paid to the design of the weld and of the welding procedure in
order to minimise such effects. The growth of grains during welding can concentrate
impurities at the grain boundary due to a decrease in the relative volumes of the grain
body and the grain boundaries. Gubner and co-workers (Gubner & Andersson 2007,
Gubner et al. 2006) have studied the corrosion behaviour of welded OFP copper, produced
by both electron-beam welding (EBW) and friction-stir welding (FSW). The EBweld
used in the comparison was made using the SKB-developed low-vacuum EBWmethod,
which differs in some process details from the method developed by Posiva.
Overall, there were no indications for preferential corrosion of the weld. Galvanic currents
between the weld and base material for FSW were low and no significant potential
difference could be detected between the materials. The FSW tool is cathodic to the
weld material, so that any particles from the tool that become incorporated into the weld
material would be cathodically protected and would not corrode to create locally aggressive
conditions (as can occur when Fe particles from C-steel steels tools become
embedded in fabricated structures).
The conclusion from these studies is that there is no evidence to indicate that the weld
region should suffer higher corrosion rates than the rest of the canister shell. Both
welding procedures appear to produce welds of acceptable corrosion performance. Of
the two techniques, FSW provides better corrosion resistance than EB welding, possibly

107
because of the lower residual stresses, the minimal grain growth and the absence of any
resultant concentration of impurities at the grain boundaries for FSW.
8.6.8 Conclusions about corrosion resistance
Given the evolution of environmental conditions, the following statements can be made
regarding the expected general and localised corrosion behaviour of the canister. Initially,
general corrosion will be supported by the reduction of the atmospheric O 2
trapped in the deposition hole. Redox conditions will be relatively oxidising and the
corrosion potential of the canister surface will be relatively positive.
During the early part of the unsaturated transient, there is the possibility for localised
corrosion because of non-uniform wetting of the surface. Localised corrosion is possible
during this period due to the non-permanent separation of anodic and cathodic processes,
leading to a general roughening of the surface. The conditions during the unsaturated
phase are the most aggressive for the copper canister because of the presence of
residual oxygen and the potential for localised corrosion and stress corrosion cracking
However, uniform corrosion conditions are expected to be established by the time the
bentonite becomes fully saturated and the residual oxygen has been consumed. The rate
of long-term general corrosion will be limited by the rate of supply of sulphide to the
canister surface, and will fall to very low levels indefinitely.
Stress corrosion cracking is considered as a possible corrosion mode during the initial
part of the aerobic phase but it is believed to be unlikely in the long-term evolution of
the canister for several reasons: the maximum concentration of SCC agents (oxygen,
acetates, nitrates, ammonium and sulphides) and the corrosion potential lie below the
respective threshold values for SCC, and because the creep rate of copper will likely
exceed the crack growth rate.
General corrosion due to sulphide is the main long-term corrosion mode. The expected
canister corrosion depths have been calculated using different assumptions on the transport
of sulphide to the canister surface. Modelling results show that the general corrosion
depth due to sulphide general corrosion is 1 mm in 10 6 years.
Corrosion inside the canister does not lead to significant corrosion of the copper inner
surface. The most aggressive conditions are assumed to occur in the presence of residual
aerated water in the spent fuel cavities. This could lead to the formation of nitric acid
and other oxidants but the amount of residual air and water are so limited that the
amount of corrosive agents can be neglected. Insert corrosion is not part of the design
bases for the canister.
The design requirement for the corrosion barrier (i.e. copper shell) is assessed to be
about 30 mm and the nominal wall thickness of the copper overpack is set to 49 mm,
which gives reasonable conservatism and allowance for possible material defects or
other degradation. Table 21 shows that the maximum expected corrosion depth for the
canister over 10 6 years is no more than 2 mm. This is consistent with several corrosion
modelling prediction results obtained by different groups in different countries, as
shown in King et al. (2011b, Table 8.1). The design requirement that the copper canister

108
shall be intact at least 10 5 years is thus fulfilled in the conditions expected in the repository
during its evolution. Even though it is not possible to directly sum up all the corrosion
depths from the different corrosion mechanisms, it is possible to see that the copper
nominal thickness is not expected to be exceeded due to corrosion. More detailed corrosion
calculations are presented in Performance Assessment report.
Table 21. Estimated corrosion depths during the evolution of the canister in a repository
in Finnish/Swedish conditions (based on King et al. 2011b, Chapter 8 and Table
8.1).
Time period Corrosion depth Comments
Encapsulation
and pre-disposal

109
Table 22 gives the gamma and neutron dose rates on the outer copper overpack surface
of the three final disposal canisters as calculated with the detailed and homogeneous
models. The results calculated with the different models were very similar and differed
mostly by 5 %. Furthermore, the homogeneous models gave conservative results in
most cases.
The gamma dose rates on the surface of the copper overpack of the canisters were over
an order of magnitude higher than the neutron dose rates. The gamma and neutron dose
rate on the radial surface of the BWR canister was clearly the highest. On the other hand
the gamma and neutron dose rate on the top surface was the highest for the VVER-440
and EPR/PWR canisters. The limiting canister will therefore depend on the problem in
question. For example for radial penetration (walls) the BWR canister will be the limiting
case while for axial penetration (lids) the VVER-440 or EPR/PWR canisters may
give the highest dose rates.
When compared to the earlier studies (Anttila 2005a) the dose rates given in Table 22
are somewhat higher. This is because in the current design there is about 10 mm less
iron between the assemblies and the canister surface. The dose rate show a similar angular
dependency of the dose rate near the canister surface as in earlier 2D analyses reported
in (Anttila 2005a).
Figures 32 and 33 show the gamma and neutron dose rates at the outer surface of the
axial mid plane of the three disposal canisters as a function of distance from the canister
surface. The results are shown for the cases with a copper lid.
The results summarised in Table 22 show that the dose rates outside the canisters are
remarkably (at least 3 times) lower than the highest allowable dose rate (1Gy/h) given in
Chapter 3. The maximum sum of gamma and neutron dose rates in Sv/h can be directly
compared to the maximum allowable (absorbed) dose rate given in Gy/h with adequate
accuracy, because the gamma radiation is dominating.
The calculated dose rates are well below the design limit 1 Gy/h.
Table 22. The gamma and neutron dose rates (mSv/h) on the outer surface of the final
disposal canisters calculated with detailed and homogeneous MCNP5 models using
ICRP74 flux-to-dose conversion coefficients (Ranta-aho 2008).
Assembly Model Surface Top Bottom
Gamma Neutron Gamma Neutron Gamma Neutron
VVER-440 Detailed 165 9.9 25.8 2.0 6.2 0.9
VVER-440 Homogen. 172 9.9 27.6 2.1 8.2 1.0
BWR Detailed 199 15 16.5 2.1 20.5 3.1
BWR Homogen. 206 15 17.0 2.1 26.1 3.1
EPR/PWR Detailed 47 9.6 24.2 1.9 14.3 2.5
EPR/PWR Homogen. 45 9.5 28.2 1.9 17.0 2.5

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Figure 32. Average gamma dose rates (mSv/h) as a function of distance from the canister
outer surfaces. Results were calculated with the homogeneous models and the
ICRP74 flux-to-dose conversion coefficients. The keywords TOP and BOT refer to top
and bottom lid surfaces of the canister, respectively (Ranta-aho 2008).
Figure 33. Average neutron dose rates (mSv/h) as a function of distance from the canister
outer surfaces. Results were calculated with the homogeneous models and the
ICRP74 flux-to-dose conversion coefficients. The keywords TOP and BOT refer to top
and bottom lid surfaces of the canister, respectively (Ranta-aho 2008).

111
8.8 Materials ageing due to radiation dose
Radiation-induced damage mechanisms in the metal structures of the nuclear reactors
have been studied extensively during the last several decades. Neutron and gamma radiation
are very intense in an operating reactor. In the storage facilities of the spent nuclear
fuel neutron and gamma radiation are many orders of magnitude smaller than in
reactor cores. Therefore, the initial assumption was that in the final disposal canisters
radiation damages are also negligible (Guinan 2001). In some recent studies it has been
claimed that even if dose rates are low, they can have adverse effects on material properties
in some cases (Brissonneau, et al. 2004; Sandberg & Korzhavyi 2009). It has been
argued that long timescales and, for instance, low temperature may increase the effects
of neutron and gamma radiation in disposal canisters.
For further damage assessments the high-energy (E >1 MeV) neutron fluence and the
absorbed gamma dose at the inner parts of a BWR disposal canister were estimated in a
very straight-forward way. First, the neutron flux and the photon dose rates were calculated
for two or three volumes of the canister at a cooling time of 20 years assuming that
the canister is filled with the elements having a burnup of 60 MWd/kgU. Then simple
time integration was carried out. The results can be used to estimate long-term radiation
damage to the canister, especially to its cast iron insert.
Figure 34. MCNP geometry of the BWR insert according to Section 6.1 in (Ranta-aho
2008).

112
The neutron fluence was calculated and reported earlier in (Ranta-aho 2008) and the
MCNP input files were used in gamma dose calculations. The photon source spectrum
and photon-flux-to-dose conversion coefficients are taken from (Anttila 2005c).
Neutron fluence was calculated for the iron cross between the four innermost fuel element
holes (volume 8201 in Figure 34) and for the iron wall around those elements
(volume 8202). The neutron fluence up to the cooling time of 100000 years was ca.
4.2·10 15 neutrons/cm 2 for the volume 8201 and ca. 3.7·10 15 neutrons/cm 2 for the volume
8202 (Ranta-aho 2008).
The photon fluxes and absorbed gamma doses were calculated for two volumes (volumes
8301 and 8302), which were the one-meter long sections at the axial centre of the
volumes 8201 and 8202. An extra volume (volume 8308) was defined in the way as the
regions 8301 and 8302 for the innermost region of the copper wall of the canister (47.7
cm ≤ R ≤ 48.9 cm). The photon dose rates of the volumes were as given in Table 23.
Table 23. Absorbed gamma dose rate at the cooling time of 20 years.
Volume Dose rate (Gy/h)
8301 70
8302 63
8308 25
From the above mentioned calculations of the activity inventories (Anttila 2005c), the
total photon source and the total photon energy at 51 cooling time moments can be extracted.
Assuming the photon spectrum remains the same and the strength of the source
varies like the photon source (A in Table 24) or the total photon energy (B), the absorbed
dose can be integrated with a log-log integration method. The resulting doses of
the three volumes are shown in Table 24.
8.8.1 Conclusion about radiation induced ageing
The effects of radiation on metals depend on many variables. Therefore, the results of a
study may not be applicable in other cases. In (Brissonneau, et al. 2004) a long-term
interim storage low-carbon steel container was studied. The main radiation-induced
damage was due to copper in steel. In the referenced study, the copper concentration
was assumed to be from 0.05 to 0.25 %.
In the present Swedish and Finnish canister designs the copper content in cast iron inserts
is specified to be less than 0.05 %. This minimises the effect of possible radiation
embrittlement. In reactor materials the interesting neutron fluence that starts to cause
detectable embrittlement effects in steel is order of 10 18 neutrons/cm 2 . The calculated
neutron dose absorbed in canister internal parts during 100000 years is 4*10 15 neutrons/cm
2 . This means that the ratio in neutron fluence is almost 3 orders of magnitude
to the advantage of the canister. The calculated neutron and gamma doses absorbed in

113
disposal canister insert material in the canister lifetime perspective are so low that no
detectable material damage or ageing effects due to radiation are expected.
The need of possible limitation of phosphorus content in cast iron is under consideration.
Table 24. The absorbed dose at three volumes of a BWR disposal canister at seven
cooling times (the integration started at the cooling time of 20 years)
Cooling time (Years) Total absorbed dose (Gy)
A
B
A) Volume 8301
4.0E+01 9.6E+06 9.5E+06
1.0E+02 2.2E+07 2.1E+07
2.0E+02 2.6E+07 2.5E+07
1.0E+03 2.9E+07 2.6E+07
1.0E+04 3.3E+07 2.7E+07
1.0E+05 3.9E+07 3.1E+07
1.0E+06 6.4E+07 5.3E+07
B) Volume 8302
4.0E+01 8.7E+06 8.6E+06
1.0E+02 2.0E+07 1.9E+07
2.0E+02 2.3E+07 2.3E+07
1.0E+03 2.6E+07 2.3E+07
1.0E+04 3.0E+07 2.5E+07
1.0E+05 3.6E+07 2.8E+07
1.0E+06 5.8E+07 4.8E+07
C) Volume 8308
4.0E+01 3.5E+06 3.4E+06
1.0E+02 7.8E+06 7.7E+06
2.0E+02 9.3E+06 9.0E+06
1.0E+03 1.0E+07 9.3E+06
1.0E+04 1.2E+07 9.8E+06
1.0E+05 1.4E+07 1.1E+07
1.0E+06 2.3E+07 1.9E+07
A) The source spectrum assumed to be constant and the source strength assumed to
vary as the total number of photons.
B) As A) above, but the source strength assumed to vary as the total gamma energy
production.

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8.9 Criticality safety
Preliminary criticality safety analyses of the Finnish disposal canisters have been reported
in (Anttila 1999; Anttila 2005b). According to the international standards and regulation
guides a canister used for final disposal of nuclear fuel must be sub-critical also under very
unfavourable conditions, i.e. for instance, when:
the fuel in the canister is in the most reactive credible configuration,
the moderation by water is at its optimum, and
the neutron reflection on all sides is as effective as credibly possible.
The criticality safety calculations have been performed with the MCNP4C Monte Carlo
code using its standard data library.
In an earlier study (Anttila 1999) it was proved that a version of the VVER canister
loaded with twelve similar fresh VVER-440 assemblies with the initial enrichment of
4.2 % fulfils the criticality safety criteria. An earlier design of the BWR canister loaded
with twelve fresh BWR assemblies of so-called ATRIUM 10x10-9Q type with the initial
enrichment of 3.8 % and without burnable absorbers has also been proved to meet
the safety criteria. However, in these calculations the impact of various uncertainties
were not assessed thoroughly enough.
The main emphasis in the most recent study (Anttila 2005b) was on the EPR/PWR canister.
It was shown that this canister type fulfils the criticality safety criteria only if the
so called burnup credit principle is applied. The fuel elements to be loaded in an
EPR/PWR canister should have been irradiated at least to a burnup of 20 MWd/kgU, if
the initial enrichment is about 4 %.
Complementary studies concerning criticality safety will be carried out. Then, all the
relevant issues will be analysed in a systematic way according to the Finnish regulatory
guides and international standards. Some validation of the calculation system will be
performed and the impact of various uncertainties will be assessed. The possible longterm
phenomena will be also analysed.

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9 RATIONALE FOR THE SELECTION OF MANUFACTURING MATERIALS
AND THEIR SPECIFIED PROPERTIES
9.1 Insert materials
The insert material shall be resistant to long term mechanical load, static or dynamic, not
prone to radiation embrittlement, possible to manufacture with proven quality and integrity,
the raw material easily available at reasonable cost. The material shall have adequate
thermal conductivity for transferring the decay heat out of insert without excessive thermal
gradients or deformations. The radiation shielding and properties for reactivity shall also
be considered is selection of insert material.
The insert is planned to be made of nodular graphite cast iron. The cast iron quality is selected
in such a way that adequate strength is achieved and the ductility of the material is
adequate. The proposed quality is EN-GJS-400-15U according to the European Standard
(SFS-EN 1563). The manufacturing material, nodular cast iron, is selected because of its
superior manufacturing aspects compared with cast steel or any welded steel structure.
The insert will be exposed to internal corrosion if the void in the insert is filled with air.
The total void of the canister is about 0.61, 0.95 and 0.67 m 3 in VVER-440, BWR and
EPR/PWR type canister inserts with fuel assemblies in each position, respectively, see
Table 6. In addition, the residual water in the fuel cavities and the air inside the canister
insert may cause radiolytic formation of nitric acid and other oxidizing species (Section
8.6). These products can increase the risk of stress corrosion cracking (SCC) during the
canister evolution. To avoid the risk, the insert atmosphere is planned to be replaced by an
inert gas (argon or helium) during encapsulation. In addition, to minimize the corrosion
risk, the fuel elements are dried in a fuel drying system before they are installed into canister
in the encapsulation plant.
The insert material is not prone to creep in disposal temperatures (Martinsson et al. 2010)
and, thus, the shape and size of the insert will remain stable even in presence of the external
loads expected during the system evolution.
The insert is designed to be leak tight only during the welding of the copper overpack, if
EBW is used. In Posiva’s design, seal welding of the copper lid is planned to be made by
EBW in vacuum, however, the FSW is an alternative method. In the long-term safety assessment,
the insert tightness is neglected, that is, the insert does not provide any hindrance
for the contact of spent fuel with water or the release of radio-nuclides.
The insert also contains some steel parts like the cassette tubes, the lid and the lid fixing
screw. Steel is selected because of its high strength and ductility, good availability as
formed products and good compatibility with cast iron.
9.2 Overpack material
The canister overpack is designed for long-term corrosion resistance, and to bear the loads
during transportation and handling operations. The canister design aims at providing with a
high probability a corrosion resistance for hundreds of thousands of years except for in-

116
cidental deviations in the repository environment. The regulation requires a corrosion
resistance of 10000 years, at least (YVL-D.5, Section 408).
The lifetime of the shell may be limited by corrosion or mechanical failure. The stability of
oxygen-free copper in repository environment is widely studied in (King et al. 2011b).
Table 21 in Section 8.6 shows that the maximum expected corrosion depth for the canister
over 10 6 years is less than 30 mm. The design requirement that the copper canister
shall be intact for hundreds of thousands of years except for incidental deviations is thus
fulfilled in the conditions expected in the repository during its evolution.
The material selected for the overpack is oxygen-free high conductivity copper (Cu-OF
according to the Finnish standard (SFS 2905)) with an addition of 30 to 100 ppm phosphorus.
We call this micro alloyed quality in this report ‘Cu-OFP´. The micro-alloying is made
to improve the creep strain properties of Cu-OF copper in higher temperature (+200 to
+300 °C). This minimises also the risk of cracking during the hot-working process. Copper
is selected as a corrosion protection layer due to its corrosion resistance. It is a noble metal,
it is available at reasonable cost, and it does not require an oxide film on the surface to
withstand the corrosion as many other passive metals like titanium or aluminium do.
For conservative mechanical, manufacturing and corrosion resistance reasons the thickness
of the intact oxygen-free copper overpack is conservatively selected to be at least 35 mm.
For the safety of mechanical handling of canister the copper overpack shall be dimensioned
conservatively, because the canister lift is made from the copper lid shoulder and
thus all the weight of the canister and the spent fuel will be supported by the copper overpack.
The design nominal wall thickness of the copper overpack is 49 mm. The excess
dimension is reserved for manufacturing allowances (tolerances) and for postulated material
imperfections in the base material or welds and to give additional margins for handling
operation loads.
Copper is, in addition to good resistance against chemical processes, a very good conductor
for the decay heat generated in the fuel. The thick copper overpack makes the
temperature evenly distributed on the copper surface and copper overpack also conducts
the heat very effectively out of the canister. Its thermal conductivity is about 10 times
higher than that of cast iron and about 400 times higher than that of bentonite buffer.
Also copper overpack is very ductile; it can bear plastic deformation and creep due to
any postulated mechanical and/or thermal load.

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10 MANUFACTURING OF THE CANISTER COMPONENTS
The canister components are planned to be pre-fabricated in various workshops of subcontractors.
The canister components are manufactured, tested and quality controlled and then
transported into canister assembly workshop that may be common for Posiva and SKB.
The final machining, controls and assembly will take place in this workshop. The assembled
canister with one end open and the lids, screws and gaskets are transported to the encapsulation
site, where the loading of canister with fuel assemblies, final closing and final
controls will take place. A special cradle is used for handling and transportation of a prefabricated
canister.
10.1 Insert manufacturing
The insert is cast in a foundry, and no separate thermal treatment is done. The openings for
fuel assemblies are formed to the cast body by setting an equivalent steel structure into the
mould. The prefabricated structure is welded together from steel tubes with strict dimensional
tolerances. After casting the block is cleaned up by blasting, cut into the right length
and the outer surfaces are machined. Then a volumetric ND-examination is carried out. In
addition, the dimensions and machined surfaces are examined.
The lids are manufactured by machining from rolled structural steel plate. The current versions
of inserts have an integral bottom, so no lid, gasket or screws are needed for the bottom.
Manufacturing description and results of demonstration tests are described in (Nolvi 2009).
10.2 Copper overpack manufacturing
The cylinder of the shell can be manufactured by extrusion or pierce-and-draw process
from a single cast and pre-heated billet. The cylinder is then machined. Posiva has selected
the pierce-and-draw process as a reference method (Nolvi 2009).
The copper lids are manufactured by casting a cylindrical billet, hot pressing and machining.
When the pierce-and-draw process is used, a cylinder with an integral bottom lid is
produced. If the extrusion method is used, an open-ended cylinder is produced. In the latter
case, the bottom lid is welded onto the cylinder. In this case, the likely welding method
used will be FSW based on the methods available at the manufacturing facility. The prefabricated
shell is examined for dimensions, surface roughness, grain size and unacceptable
flaws. In case the bottom lid is welded, the weld will be inspected by sensitive NDT
method. The bottom lid weld is made and inspected in normal workshop conditions without
any disturbance caused by the radiation. Thus the reliability of the result is higher than
that of the top lid weld, which is made and inspected with remote controlled equipment
due to disturbing radiation.

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10.3 Component inspections
10.3.1 Canister insert inspections
The manufacture of the insert is carried out in steps. The manufacturing consists of the
following phases: manufacture of steel cassette for the mould of the insert, casting,
cleaning, testing samples, and machining. The quality controls of each step are described
below.
The quality controls of the manufacturing of the steel cassette are based on:
• Material certificates for steel products
• Ultrasonic testing (UT) of welds
• Dimensional controls
• The straightness and space of the channels are tested with a special gauge
• Visual inspection.
The quality controls of the casting are based on:
• The process control documentation
• The temperature control log
• The chemical analyses of melt.
Mechanical, metallurgical and chemical properties from the cast body are made from
special cast-on samples and from a sectional slice that is taken from the top part of the
casting. The geometry of the openings is examined with a gauge after casting. After
machining, the dimensions of the insert are controlled and the integrity of the cast material
is controlled with surface and volumetric inspection equipment. A special UT control
is made to register the thickness variations of the weakest section between the openings
and the outer cylindrical surface. The whole control programme for insert manufacture
is described in (Pitkänen 2012).
10.3.2 Copper overpack inspections
The inspection methods and equipment for canister component testing are described in
details in (Pitkänen 2012). The acceptance procedure and the allowable material imperfections
are given in the same report (Pitkänen 2012). The copper overpack weld controls
are discussed in Section 11.6.
The material manufacturing inspections include, in addition to NDT controls of the
component, also the chemical, metallurgical and mechanical inspections that are described
in the manufacturing specifications for the material. The detailed manufacturing
inspection activities are reported in (Nolvi 2009, Chapter 7).
The manufacturing inspections are carried out after each manufacturing step: raw material
inspections, hot deformed material inspections, pre-machined material inspections,
final machined product inspections. Visual inspections are also carried out after transport
for transport damages and cleanliness.

119
Tables 25 and 26 give the preliminary quality control plans for the canister shell and
canister lid manufacture.
The acceptance criteria for the copper base material faults in the canister shell are the
same as for seal weld and they are given in Section 11.6.
Table 25. Preliminary control plan for copper overpack manufacture.
QUALITY CONTROL PLAN
Manufacturing phase
Acceptance test of the billet
Acceptance test of the hot-deformed product
Acceptance test of pre-machined product
Acceptance test of final-machined product
Welding operation
Weld examination
CANISTER OVERPACK
Examination
Chemical composition
Surface flaw detection (piqued penetrant)
Visual inspection
Dimensional examination, weight
Material identification
Material identification
Grain size test
Tensile test
Material identification
Cleanliness test
Surface roughness test
Volumetric examination
Dimensional examination
Material identification
Cleanliness test
Surface roughness test
Surface flaw detection (eddy current)
Dimensional examination
Welding procedure qualification
Welding operator qualification
Welding equipment qualification
Welding process control (log)
Visual examination, weld surface
Surface flaw detection (eddy current)
Radiographic examination, welds
Ultrasonic examination, welds

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Table 26. Preliminary control plan for copper lid manufacture.
QUALITY CONTROL PLAN
COPPER LID (OR BOTTOM)
Manufacturing phase
Acceptance test of the billet
Examination
Chemical composition
Surface flaw detection (piqued penetrant)
Visual inspection
Dimensional examination, weight
Material identification
Acceptance test of the hot-deformed product
Material identification
Grain size test
Tensile test
Acceptance test of pre-machined product
Material identification
Cleanliness test
Surface roughness test
Volumetric examination
Dimensional examination
Acceptance test of final-machined product
Material identification
Cleanliness test
Surface roughness test
Surface flaw detection (eddy current)
Dimensional examination

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11 ENCAPSULATION PROCESS
11.1 Fuel preparation
Spent fuel elements are stored in pools at their respective nuclear power plant sites for a
minimum of 20 years and an average of 30 to 50 years before encapsulation. Some fuel
rods may be filled with pool water (in case the fuel rod is leaking). Some assemblies
may have water on the fuel surface after the transport in water-filled casks from the interim
storage locations to the encapsulation plant. To remove the maximum amount of
water from the fuel elements during the encapsulation process, all fuel elements are
dried in a drying unit using a combination of decay heat and vacuum using drying system
according to (Suikki et al. 2007). After drying, the fuel rod cavities in 12 fuel assemblies
in each canister are assumed to contain in total a maximum of 600 grams of
residual water according to (Miller & Marcos 2007, page 37). Due to the absence of
sufficient amounts of moderator (e.g. liquid water) inside the intact canister, no induced
fission takes place within the intact canister for the entire range of Finnish spent fuel
burnup. The transportation of fuel elements from interim storages outside encapsulation
site and the risks involved in transportation are described in (Suolanen et al. 2004). The
fuel transportation risk assessment report is under renovation in 2012.
To minimize the radiolysis of residual air and water inside the canister, during the encapsulation
process the void space of the cast iron insert is filled with argon at normal
temperature and atmospheric pressure. Argon is an inert gas and its role is to prevent the
radiolysis of residual air and moisture in the canister. Radiolysis of air and water mixtures
forms nitric acid and other oxidizing species which could promote stress corrosion
cracking of the canister components.
11.2 Fuel handling and packaging into canister
The canister is docked (Suikki 2006) with the handling cell of the encapsulation plant and
the fuel assemblies are moved one by one from the transport cask first into drying system
for drying (Suikki et al 2007) and then into the positions in the disposal canister insert by
the fuel handling machine. Finally the steel lid with gasket is installed, the gas atmosphere
in the canister cavity is then, according to proposed procedure, changed to some inert gas
and the lid fastening nut is tightened with a manipulator. The inert filling gas proposed is
argon and the filling gas pressure is equal to the atmospheric pressure. The schematic presentation
of the encapsulation process is shown in Figure 35. Encapsulation plant description
is given in more details in (Kukkola 2012).

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Figure 35. Encapsulation plant cross-section. The fuel handling cell is to the right of
the centre of figure (Kukkola 2012).
11.3 Canister preparation before sealing
The canister is transferred into the handling cell of the encapsulation plant and the fuel
assemblies are moved one by one from the fuel drying system rack into the positions in the
canister insert by the fuel handling machine. The gas atmosphere in the canister cavity is
then changed. Finally the steel lid with gasket is installed and the nut is tightened with a
manipulator. After loading of the fuel assemblies and closing of the insert lid the canister is
lowered from fuel handling cell to the transfer corridor, the copper lid is travelling on a
holder of the canister transfer trolley and the canister is transferred to welding position for
the sealing weld of the copper lid. The canister is transferred in the corridor on a canister
transfer trolley, which has a hoisting device, for details see (Kukkola 2012).
11.4 Sealing weld of the copper overpack
After loading the fuel assemblies into the canister and closing of the insert, the canister is
moved from the handling cell in vertical position to the electron beam welding position in
a vacuum chamber (Suikki & Wendelin 2008). After evacuating the chamber and the canister,
the copper lid is installed, fixed with a tool and welded. The welding is made in at
least two phases, first a tag weld all around the lid to ensure the lid positioning during the
full penetration weld.
The positioning of the electron beam must be targeted to the lid seam with an adequately
accurate tolerance during all the welding process. The positioning and the compensation of

123
thermal deformations during the welding process can be made with a programmed steering
automation or mechanical steering device, which follows the pre-fabricated trail close to
the seam to be welded. This kind of a local steering device allows the possible use of a
front bar on the seam to give additional material to the weld and to minimise the surface
craters caused by leaking of liquid weld metal.
Several programs for the development of EBW for 50 mm thick copper were carried out in
Finland since 1994. The experience from development work and welding tests are summarised
in (Meuronen & Salonen 2010).
EBW is the reference sealing method for the Posiva canisters but an alternative method,
FSW, is also investigated. The selection between them will be made not later than in 2013.
11.5 Final machining of welded surfaces
The welding surface is machined with help of a face-milling cutter. The cutter head is
moved in horizontally and the canister is rotated around its vertical axis. Milling is performed
remote-controlled. Milling is used instead of grinding, because milling chips are
easier to clean than grinding dust. During the machining, the chips are removed from
the canister lid groove by vacuuming. The required surface roughness of the machined
surface is Ra 6.3, which equals with the general outside surface roughness requirement
of the canister cylinder. Surface roughness is more discussed in Chapter 6.
11.6 Weld controls
The canister weld is inspected using three inspection methods based on: ultrasounds,
eddy current and X-ray. Visual examination is also used to get better overview of the
weld surface and the surface quality after machining. These inspection phases will take
place at the inspection station, (Suikki & Wendelin 2009), see Figure 36. The canister is
lifted with help of the canister transfer trolley into the inspection station. During the
inspection, the canister is rotated around its vertical axis.
The X-ray inspection is made with a high-energy accelerator that sends the X-rays
through the upper corner of the canister lid. A digital receiver unit is behind the object
and it collects the shaded picture of the trans-illuminated object.
The ultrasonic inspection is performed as follows. A water-filled box structure is attached
on upper part of the canister. The canister is rotated, and the ultrasonic PAprobes
(Phased Array - method) in the sleeve scan the canister weld from the side. As
an acoustic coupling, a water path between PA-probe and copper cylinder is used. After
the inspection, the box is drained and removed.
Eddy current inspection is performed as described above for ultrasonic, except water
layer is not used.

124
Figure 36. Inspection station model of the encapsulation plant. In front the ultrasonic
and eddy current inspection devices behind the shield wall, on backyard the X-ray inspection
device (Suikki & Wendelin 2009).
The control plan for the canister seal weld is given in Table 27. This is a preliminary
control plan that was originally made for the demonstration test welds in (Raiko et al.
2009).
Table 27. Control plan for the canister seal weld.
QUALITY CONTROL PLAN
Welding operation
Weld surface machining
Weld examination
CANISTER LID SEAL WELD
Welding procedure qualification
Welding operator qualification
Welding equipment qualification
Welding process control (log)
Visual examination (special camera)
Visual examination, weld surface
Surface flaw detection (eddy current)
Radiographic examination, welds
Ultrasonic examination, welds

125
In case of alternative sealing method, the FSW weld, the X-ray inspection is preliminarily
omitted from the control plan, because of the fact that FSW weld defects are not efficiently
revealed by X-ray but phased-array ultrasonic.
The inspection methods and equipment are described in details in (Pitkänen 2010). The
acceptance procedure and the allowable material imperfections are given in the same
report (Pitkänen 2010).
Acceptance criteria for welds in this welding test are presented in Table 28. Any two
adjacent defects separated by a distance smaller than the major dimension of the smaller
defect shall be considered as a single defect. The criterion covering the intact wall
thickness requirement of 35 mm in 100 % and 40 mm in 99 % of canisters is the master
requirement for acceptance, especially for combining defects. Penetration of the EBW
shall be limited between 42–70 mm in axial direction of the canister.
Table 28 gives the first screening criteria for the acceptance of the weld. In case the
indication in examination is below the values given in Table 28 it can be accepted without
further analysis. If the weld does not fulfil the first acceptance criteria then an additional
evaluation is made using more advanced detection and sizing techniques by two
separate persons qualified for detection and sizing. The main evaluation principle and
final accepting criterion is that the total intact material thickness in the copper overpack
is 35 mm at minimum. If even this criterion is not met, a further expert analysis is made.
The expert panel will consist of inspection, material and corrosion specialists, who then
will make the decision of acceptance or rejection. This kind of acceptance procedure is
typical for ASME-inspections for nuclear components according to (ASME XI 2008).
This is described in more detail in (Pitkänen 2010). The same acceptance criteria valid
for canister base material are also valid for the welds. The screening criteria given in
Table 28 are based on welding standards and practical expert reasoning.
Copper overpack scratches and indentations are preliminarily analysed in (Unosson
2009). Deeper indentations may cause remarkable plasticity and cold-hardening in the
copper material. In this aspect additional research is going on. Surface scratches are not
harmful for corrosion in long term, because the corrosion resistance of copper is not
based on oxide surface.

126
Table 28. Acceptance criteria for the canister sealing welds in the lid test series. The
acceptance criteria for different defect types are modified according to remaining wall
thickness criterion and completed with thick copper weld defects types (EB- and FSweld)
(Pitkänen 2010).
Defect No. Type of defect Maximum allowable size
100 Cracks l < 10 mm, h < 3 mm
2011, 200 Gas pore, porosity l < 25 mm, h < 6 mm, w ≤ 8 mm
2013 Clustered porosity l < 25 mm, h < 6 mm, w ≤ 8 mm
2014 Linear porosity l < 25 mm, h < 6 mm, w ≤ 8 mm
2015 Elongated cavity l < 25 mm, h < 6 mm, w ≤ 8 mm
5011, 5012
External undercut, defect on the
side of the weld originating machining
and welding, possible
repair by machining
402 Incomplete penetration
Cold lap
Symbols in Table above are as follows:
l length of defect, w width of defect, h height of defect.
l < 20 mm, h < 5 mm
l continuous, h < 8 mm,
Intact 42 mm
l < 50 mm, h < 10 mm
Joint like hooking
l continuous, h < 8 mm,
Intact 42 mm
401 Lack of fusion l < 50 mm, h < 10 mm
2016 Wormhole / Crater l ≤ 5 mm , w ≤ 3 mm, h ≤ 10 mm
300 Solid inclusions l ≤ 10 mm, w ≤ 3 mm, h < 10 mm
511 Incompletely filled groove l < 10 mm, w < 8 mm, h < 5 mm
Scratches
Indentation
Permitted locally
1 mm depth, large diameter indentation
(d > 10 mm), small and sharp
indentations (scratch-like) are allowed
11.7 Final control
After the final control in the encapsulation plant, the canister is accepted for interim
surface storage in the buffer storage before being transferred to the repository. The final
control in this phase consists of checking the canister visually (using cameras) that the
surface is clean and that there are not excessive surface scratches or punch marks. The
control documentation of the preceding controls is also checked in this phase to ensure
that all the proposed examinations are made; the results are documented, assessed and
accepted. A further surface control will be carried out after the canister is transferred to
the repository just before lowering the canister into the deposition hole.

127
12 CANISTER TRANSFER AND DEPOSITION
After the canisters have passed all the inspections, the accepted canisters are transferred
to the canister buffer storage in the encapsulation plant waiting to be transferred into the
repository. The automated guided fork lift transfers the canisters from the transfer corridor
to the buffer storage and further to the canister lift cage through the labyrinths that
are made for radiation protection purpose.
After encapsulation, the canisters are stored in the encapsulation plant storage area or in
the canister buffer storage at the repository level for days or, at maximum, for a couple
of months before emplacement. The canister surface temperature may increase within a
few days up to stationary temperature of about +50 °C, depending on the ventilation
condition in the encapsulation plant and in the canister storage areas, see Section 8.5.1.
12.1 Canister handling and transfer
The lift in the canister shaft is ad joint to the encapsulation plant. The canister lift is
used for transferring the fuel canisters from the encapsulation plant into the repository.
The transfer and installation vehicles are used for transferring and installing fuel canisters
and bentonite blocks in the repository facilities. The canister transfer and installation
vehicle is foreseen as rubber-wheeled variant of transfer vehicle, see Figure 37. The
vehicle will transfer the fuel canister from the loading station at the canister buffer store
to the deposition hole and install the canister into it. The vehicle is equipped with a radiation
shield shell, inside which the canister is transferred. The shield makes it possible
for personnel to work close to the vehicle.
Figure 37. Canister transfer and installation vehicle model. This is a trailer type vehicle
that needs a pull-tractor to be able to move (Cadring OY).

128
12.2 Control of handling and transfer damages
A final surface control will be done after canister transfer to the repository just before
lowering the canister into the deposition hole. This is done with cameras or special
eddy-current device that controls the canister surface condition, when the canister is
lowered from the vehicle into the deposition hole.
12.3 Canister deposition
The canister is placed into the bentonite-lined deposition hole using the canister transfer
and emplacement vehicle. The vehicle is parked in the designated locations above the
deposition hole. After this, the vehicle is lifted to rest on its struts and levelled. When
the vehicle is in place, the radiation shield can be rotated to a vertical position. At the
same time, the rear end of the radiation shield opens and the supporting wheels of the
radiation shield rest on the floor of the deposition tunnel. The installation vehicle is
shown in Figure 38 in installation position.
When the radiation shield is in a totally vertical position, the lowering of the canister
can begin. The lowering can be monitored using several cameras in order to ensure that
the process is successful and that the canister does not damage the bentonite buffer rings
already placed in the hole. Once the canister has been lowered onto the bottom of the
hole, the gripping mechanism can be removed from the canister cover lug and lifted
back inside the radiation shield. After this, the radiation shield can be rotated back to a
horizontal position and the vehicle can be driven out of the deposition tunnel (Wendelin
& Suikki 2008). The canister surface control in this phase is described in Section 12.2.
Figure 38. Placement of the canister into the deposition hole (Cadring OY).

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13 CANISTER INITIAL STATE
13.1 Fuel types
The spent fuel characteristics for Asea Atom BWR (nowadays Westinghouse Electric
Sweden), VVER-440, and EPR/PWR reactor are given in Table 29. The length of fuel rods
and assemblies may grow in length during the reactor irradiation some 10 to 25 mm depending
on the burnup. The values are from the earlier design report (Raiko 2005). The
typical enrichment and burnup is updated and the second clarifying note is added.
Table 29. Representative fuel characteristics for BWR, VVER-440 and EPR/PWR fuel.
FUEL TYPE
Asea-Atom BWR VVER-440 PWR EPR/PWR
Assembly sectional configuration square hexagonal square
Length of assembly (m) 4.127 3.217 4.865**
Sectional dimension (mm) 139* 144 215
Number of rods per assembly 63 - 96 126 265
Mass of uranium (kg) 172 - 184 120 - 126 530 - 533
Total assembly mass (kg) 292 - 331 210-214 785
Fuel channel dimension (mm) 139 (square) 144 (hexagonal) No channel
Total length with fuel channel (m) 4.398-4.421 3.217*** 4.865**
Anticipated maximum average burnup 55 57 55
of a fuel element (MWd/kgU)
Estimated average burnup of all the 39 - 40 40 - 41 45-46
fuel to be disposed (MWd/kgU)
Typical enrichment U-235 (%) 3.3 – 4.4 3.6 - 4.4 1.9 - 4.9
Minimum cooling time of a single fuel 20 20 20
element (years)
Minimum average cooling time for
33 30 42
encapsulation with average burnup
(years)
Allowable average decay heat at disposal
(full canisters) (W/tU)
806 950 862
*) The top end handle of the BWR fuel assembly has some more extensive details, whose
maximum sectional dimension is 151 mm.
**) The fuel element has leaf springs on the top tie-plate that extends the total length with
some tens of millimetres. The leaf springs are planned to be removed before disposal.
Geometric details of the fuel element are to be confirmed later. Control rod crown is
not included in the EPR/PWR fuel elements that are to be disposed. The weight of a set
of 24 absorber rods is roughly 55 kg.
***) Length of the control rod follower assembly is 3.200 m.

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The BWR fuel assemblies will be inserted with fuel channels into the canister. The canister
is dimensioned for 8x8-, 9x9- and 10x10-type fuel assemblies with the standard fuel channel.
The VVER-440 fuel assembly is an integral structure with the hexagonal fuel channel. The
channel makes the assembly much longer than the fuel rods.
The EPR/PWR fuel assembly is of 17x17-24 -type, which has no fuel channel at all. The
top end plate has some axial leaf springs that lengthen the maximum dimension (height) of
the assembly. These springs are planned to be removed before the assemblies are encapsulated.
The possible control rod crowns are also planned to be cut off before the fuel elements
are encapsulated with the possible absorber rods. This is done to minimise the
length of the EPR/PWR fuel assemblies to be disposed. Roughly one third of EPR/PWR
assemblies contain control rods when in reactor core. The control rod assemblies can be
recycled from one fuel element to another. It is estimated that 15 % of fuel elements to be
disposed will contain control rods in them. Representative pictures of fuel elements are
given in Figure 39.
Figure 39. Representative illustrations (from left) of BWR, EPR/PWR and VVER-440 type
fuel assembly. BWR and VVER-440 fuel elements are partly cut open. Pictures are not in
scale. Illustrations are from TVO/Olkiluoto and Fortum/Loviisa nuclear power plant brochures.

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13.2 Average fuel burnup, number of fuel assemblies and activity inventory
The basic dimensioning criteria for the encapsulation plant are that the fuel may have burnup
up to 60 MWd/kgU, the minimum cooling time of a single fuel element that can be
handled in the plant is 20 years and that up to 100 canisters can be encapsulated annually.
This minimum value has been used in radiation dose and shielding calculations. The actual
cooling time due to decay heat limitations is typically 30-50 years. Estimated burnup and
amount of spent fuel produced in future is given in Figure 40 and Table 30. They are based
on the operation plans of the relevant power companies.
Data for the planned OL4 unit are included, even if the decision of reactor type is not
yet decided. They will be estimated after the plant-type decision. The maximum total
amount of spent fuel from OL1-4 and LO1-2 units is estimated up to 9000 tU and number
of canisters up to 4500. These numbers in Table 30 include the effect of possible
power increases and plant life extensions in the future.
55
AVERAGE BURNUP (MWd/kgU)
50
45
40
35
30
25
OL2
OL3
20
1980 1990 2000 2010 2020 2030 2040 2050 2060 2070
YEAR
Figure 40. The development of the average discharge lot-specific burnups in different
nuclear power plant units. Projected figures are shown for 2012 onward.
LO1
LO2
OL1

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Table 30. Details of forecast fuel accumulations at the OL and LO power plant units.
OL1–2 OL3-4 LO1–2 Total
Planned operating life (a) 60 60 50 -
Average discharge burnup of all fuel 39.5 45.1 40.6 (42.7)
elements (MWd/kgU)
Numbers of canisters (pcs) 1400 2350 750 4500
Corresponding tonnage (tU) 2950 5000 1050 9000
The main characteristics of spent fuel relevant to the repository evolution are the decay
heat, reactivity, radionuclide inventory and its decay with time. The properties of various
fuel types are examined and reported in (Anttila 2005b; Anttila 2005c).
13.3 Fillers and residual contents of canisters
The canister void is filled with argon. There may be some residuals of air atmosphere,
but the volumetric amount of argon is 90 %, at minimum, see Design Basis report.
The residual water from outer surface of the fuel assemblies is dried in a vacuum dryer
during encapsulation. The amount of residual water on surfaces is minimal. However,
some residual water may be trapped inside leaking fuel rods. This is estimated very pessimistically
for the long-term safety assessment as 600 grams/canister. This number
comes from the very conservative estimate that 1 rod in every assembly (totally 12 in a
BWR canister) is leaking and the rod contains water as much as there is void inside a
rod, 50 grams. This leads to 600 grams per canister, see (Miller & Marcos 2007, page
85).
The canister contains intentionally no organic materials. This is to avoid the risk for any
bacterial activity inside the canister in the long time.
13.4 Decay heat
One of the main criteria guiding the selection of the fuel for encapsulation in a given
canister is the resulting decay heat power for the canister. The encapsulation process is
designed and planned so that the specified decay heat at the moment of disposal is fulfilled
at about ± 2 % accuracy. This is done by predicting the decay heat of each fuel
assembly using calibrated computer programs and selecting the assemblies with an optimization
procedure that leads to the specified decay power in each canister with sufficiently
high accuracy. The specified maximum decay powers for canisters are 1700,
1830 and 1370 W for BWR, EPR/PWR and VVER-440 canisters, respectively.
The design value of the decay heat of the reference canister is set to 1700 W in (SKB
2009). This is applied for Posiva canister, too, because of identical fuel and canister design.
The permissible decay power for other types of canisters have been derived from the
power of the reference canister (BWR) proportional to the cooling surface area of the canister
(the surface area of the copper exterior).

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13.5 Canister size, shape and material integrity
The documented quality control system will examine the material properties (chemical
contents, metallurgical and mechanical properties), size, shape, surface quality and the
material integrity in a way that the specified requirements are met with a high reliability.
The dimensions, especially those creating the gap between insert and the shell are essential
to be within the specification. The assembly of the canister components and the
thermal expansion shall be possible without violent contact and, on the other hand, the
gap shall be as small as possible to minimise the room for creep deformation of the
copper overpack when the external pressure loads have arisen. This is the reason for the
strict tolerances and controls given to the design specifications
The most important quality-related feature for long-term safety for the safety case is the
probability of a penetrating flaw through the canister wall. The probability analysis by
Holmberg & Kuusela (2011) suggests that a canister with an initial penetrating defect
may be present in the final repository. The Bayesian approach has been utilized to express
large uncertainties in the estimates. Varying the assumptions used as input data
result that the expected probability of at least one defective canister in the final repository
of the population of 4500 canisters is 1-10 %. Due to lack of sufficient statistical
data in particular from the canister seal welding and possible human errors in the
NDT process, expert judgements from a related welding method (FSW) and NDT praxis
have been used.
The minimum intact wall thickness of the copper overpack is guaranteed by NDT controls
to be 35 mm, out of the nominally 49-50 mm thick. All the screening criteria of the
various NDT processes aim to fulfil this target with 100 % reliability. The acceptance
criteria for copper weld and overpack are given in Section 11.6.
The insert as a load bearing component has also strict requirements for material
integrity. The isostatic pressure load is not critical even for moderate flaws or lack of
material of the insert, see (Dillström et al. 2010a), Section 7. However, the 5 cm rock
shear case is very sensitive for transverse surface flaws in the insert. The maximum
allowable surface crack size is 4.5 mm deep and 27 mm long semi-elliptic reference
crack the maximum internal crack size is 10x60 mm, see (Dillström & Bolinder 2010b),
Section 4.
The quality control procedures, technology and the acceptance criteria for canisters and
canister components are discussed in (Pitkänen 2010; Pitkänen 2012).
13.6 Adverse effects of manufacturing process on the materials
13.6.1 Residual stresses in seal weld
The interest of residual stresses in copper overpack is primarily related to the possibility
that stress corrosion might appear for unforeseen reasons or that disadvantageous de-

134
formation might take place. This item is more discussed in (Raiko et al. 2010, Section
6.2.10).
After welding and other high temperature processes, residual stresses appear in the material.
There are many factors that influence the distribution and magnitude of these
stresses: physical and mechanical properties of the material, structural dimension, restraint
conditions, and welding parameters such as heat input. It must also be taken into
account that the material properties vary across the welds and that the properties are
temperature dependent. A number of these parameters are simply not fully known. As a
consequence, fully reproducible measurements and accurate computations of residual
stresses cannot be achieved. This is even more relevant for copper because of the low
yield strength than for steels where most work on residual stresses has been performed.
Posiva has also investigated residual stresses and deformations in EB welded samples.
EBW seems to cause some more distortion than the FSW. The reason for that is that
EBW weld material is momentarily liquid and afterwards all the shrinking from solidification
down to room temperature causes remarkable shrinking. The transverse width
of the copper EBW is substantial, some 8 mm. In addition, if a second surface melting
pass is done, then the weld’s transverse sectional area becomes remarkably higher and
the distortion becomes, accordingly, higher. Distortion causes residual tensile stresses
both in transverse and longitudinal orientations of the weld pass.
After EB-welding, the tube will shrink by approximately 2 mm in diameter. The centre
of the lid will buckle 1-2 mm inward. These can be taken into account in design and
also in machining the outer surface of the canister so that after welding there is not barrelling
in the canister shape. The top surface of the lid will be machined for NDT. There
is no need for machining the outer surface of the canister after welding, unlike in the
case of FSW. Deformations are related to residual stresses. It has been reported previously
that high residual stresses may occur in the weld (Gripenberg & Hänninen 2006).
The welding process has been changed and now residual stresses are lower than previously
reported. A residual stress level of 0-30 MPa has been measured using the optical
Prism Hole Drilling (Prism HD) method (Laakkonen 2011). Residual stresses of the
same plate tests as used in Prism HD have also been measured by the contour method
(Romppainen & Immonen 2011). The contour measurement method showed that longitudinal
residual stresses of the plate weld are 40-55 MPa, at maximum. The maximum
longitudinal residual stress is 45.7-49.5 MPa (Romppainen & Immonen 2011) when
using typical welding parameters for lid welding. Modelling of the EBW plate welds
showed that longitudinal residual stresses are 47-48 MPa with the same welding power
and welding speed as in welding tests. Further measurements of the lid welds and plate
welds will be reported in near future.
13.6.2 Residual stresses in cast iron insert
Residual stresses induced in the material during manufacturing processes like casting,
hot-deformation, welding or machining, are secondary stresses. They do not have any
external driving force that would protract them after yielding or after thermal stress relief
treatment of the material.

135
The question of possible residual stresses in cast iron inserts was arisen for many years
ago. Some surface based examination methods for detecting residual stresses was first
used, but the results were very sensitive for measurement scattering and no information
beyond surface area was possible to achieve. In 2011, the best available technique was
used to get the general picture of the residual stresses in the whole volume of the cast
insert. The deep-hole drilling procedure (DHD) was used in England for detection of
residual strains in a 1 m long section of a full-scale insert that was earlier manufactured
in a series of demonstration manufacture. The general picture of the strains was measured
by 5 deep holes that were drilled radially inwards from the outer surface of the
insert. The deepest of the drillings extended to the centre of the insert section. Some of
the drillings penetrated also the steel square tubes that are used to form the square openings
for the spent fuel elements. All this DHD testing is reported in (Bowman 2011).
Under this work scope only the in-plane distortions of the reference holes were measured
following stress relief. The distortions were then converted into residual stresses using a
Young’s modulus, E, of 166 GPa for the cast iron and 210 GPa for the steel tubes. The
analysis used to convert the measured distortions into stresses assumes isotropic, plane
stress conditions and as such the Poisson’s ratio was not required.
The measured stress profiles from the 5 prescribed locations showed similar results where
the locations are comparable. The axial and hoop residual stresses were shown to be very
similar at all locations within the cast iron, but differed in the steel tube sections. In all
measurements made from the outer cast iron surface, except for measurement 1, the results
showed compressive peaks in the surface regions. The maximum compressive residual
stresses in both the axial and hoop directions were -113 MPa and -81 MPa respectively,
both found at the surface of measurement 3. Generally the compressive peaks were followed
by sharp rises into tension by both the axial and hoop stresses. Once in tension the
results showed tensile peaks occurring in most measurements before 20 mm deep. The
maximum tensile stress levels in the cast iron for both the axial and hoop directions were 58
MPa and 57 MPa respectively, both found at approximately 2 mm from the surface of
measurement 1. The exception to this was measurement 3 in which peaks, particularly in
the hoop direction, occurred at the specimen centre and at approximately 170 mm deep,
where the measurement was not adjacent to a steel tube wall. This was confirmed by measurement
5, which was made across the region between the inner and outer steel tubes, and
showed higher hoop stresses.
In general, the residual stresses in the cast iron insert are low and on the outer surface of the
insert the residual stress is in compression but in the small neck area around the outermost
square openings where the residual stress is on the surface in tension but close to zero.
Throughout the circumference the stresses turn to tension when going inwards from the
outer surface and then the stress converges towards zero typically by the depth of 20 mm.
However, the amount of tension is low, less than 60 MPa at maximum, thus the harmful
effect of the residual stresses in cast iron insert is low.
An overall nominal accuracy of approximately ±30 MPa is valid for the DHD residual
stress measurements at all depths, except at those within 1 mm of the specimen surface.
Based on this error bound, many of the “features” in the residual stress profiles were considered
to be measurement fluctuations and were not necessarily due to a changing stress
field, but more likely a result of the errors and inaccuracies of the measurement technique.

136
The origin of residual stress is the casting process, where the cylindrical surface is solidified
first and later the shrinking of the melt iron inside the thicker sections cause
tension, which causes compression in the surface areas as a balancing reaction.
The compressive residual stresses on component surfaces are beneficial for small surface
cracks due to a crack closing tendency. The residual stresses have no practical influence
on limit load or other higher loads that cause yielding, because the manufacturing-based
residual stresses are expected to vanish when the material yields.
13.6.3 Temper embrittlement in cast iron insert
Since the cast iron insert will operate the first hundreds of years in a temperature above
+70 °C, the possibility for temper embrittlement has been considered. Experience from
equal cast iron material has not shown any problem in this respect for example in marine
diesel engines that typically operate tens of years in equal temperature range.
The industry reference (Ductile iron data for design engineers 1990, page 3-53) reports
about temper embrittlement in ductile cast irons the following: “Temper embrittlement,
as found in certain quenched and tempered steels, may also occur in similarly treated
ductile irons with susceptible compositions. This form of embrittlement, which does not
affect normal tensile properties but causes significant reductions in fracture toughness,
can occur in ductile iron containing high levels of silicon and phosphorus which have
been tempered in the range 350-600 °C and cooled slowly after tempering. Although
normally associated with tempered martensitic matrices, temper embrittlement can also
occur if the matrix is tempered to the fully ferritic condition. Temper embrittlement can
be prevented by keeping silicon and phosphorus levels low, adding up to 0.15 % molybdenum
and avoiding the embrittling heat treating conditions.”
Posiva’s canister inserts are not tempered at all according to current manufacturing procedures.
The silicon content of the cast iron is typically about 2 % and the phosphorus
0.02 %. Under these circumstances no temper embrittlement of the insert material can
be expected.

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14 ASSESSMENT OF CANISTER COMPLIANCE WITH DESIGN
REQUIREMENTS
14.1 Design analysis evidence against design requirements
All the requirements given in chapter 3 of this report are quoted here and a short summary
of the qualification for the requirements is given with references.
Definition and objectives
Canister is a container with a water and gas tight shell and a mechanical loadbearing
insert in which the spent nuclear fuel is placed for final disposal in the repository. The
canister shall contain the spent fuel and prevent, and in the case of leak, limit the
spreading of radioactive substances into the environment.
The designed canister variants are shown to be able to accept all the Finnish spent fuel
types that are in use today or which have a construction license. The mechanical and
chemical resistance against postulated loads is shown to be adequate for the specified
lifetime (Sections 8.3-8.9).
Containment
Canister shall initially be intact except for incidental deviations when leaving the encapsulation
plant for disposal. In the expected repository conditions the canister shall remain
intact for hundreds of thousands of years except for incidental deviations.
The manufacturing, encapsulation, transfer and disposal processes are planned so that
the canister is intact after the process. The quality assurance programme steers and
documents the quality of the processes and the quality control programme that concentrates
on all activities of the encapsulation process verifies the quality of each canister.
Quality control systems are described in (Pitkänen 2010; Pitkänen 2012) and Sections
10, 11 and 12.
Chemically resistant
Canister shall withstand corrosion in the expected repository conditions.
The 5-cm thick oxygen-free copper overpack with sealing welds is able to resist the corrosion
processes in postulated repository environment for more than 100000 years according
to Section 8.6 and (King et al. 2011b).
Mechanically resistant
Canister shall withstand the expected mechanical loads in repository.
The design of the reference canister is shown to fulfil the mechanical strength and ductility
against the postulated loads (Raiko et al. 2010). The VVER-440 and EPR/PWR
variants of the canister are shown to be at least equally resistant against mechanical

138
loads (Section 8.3) and (Ikonen 2005). Also the resistance against the expected thermal
loads is good (Sections 8.4-8.5).
Compatibility with the EBS and host-rock performance
Canister shall not impair the safety functions of other barriers. Radiation dose shall be
limited so that EBS materials or host rock are not damaged.
The canister is in contact with the bentonite buffer and is in the vicinity of the near-field
rock. The canister has no mechanical or chemical effects on bentonite or the rock. The
radiation effect on bentonite or the rock is reduced by limiting the radiation dose rate to
a safe level of 1Gy/h. The canister variants fulfil this criterion with a margin of several
hundred per cents (Section 8.7) and (Ranta-aho 2008). The thermal effects of the canister
decay heat on bentonite are reduced by limiting the maximum temperature in the
canister/buffer interface to +100 °C by design (Section 8.5.4) and (Ikonen & Raiko
2012). The thermal effects of canister decay heat on near-field rock are moderately increased
compressive stresses in the rock and this is taken care by the buffer design that
shall have a sufficient supporting pressure against the rock surface in the deposition
holes.
Sub-criticality
Canister shall be sub-critical in all postulated operational and repository conditions including
intrusion of water through damaged canister wall.
The criticality safety of the spent fuel in canisters is fulfilled as long as the canister is
not filled with water. If the canisters start to leak for some reason, the reactivity of the
spent fuel configuration inside canister increases due to the moderating effect of the
water. The effective multiplication factor shall be less than 0.95 also when the canister
is in the most reactive credible configuration (optimum moderation and close reflection).
The BWR and VVER-440 canister variants have been shown to fulfil this criterion
even if fresh fuel elements are disposed (Anttila 2005b). The larger fuel element, the
EPR/PWR type, is more reactive. Thus when verifying its criticality safety, the so-called
burnup credit must be utilised. This means that the criticality safety calculations take
into account the fission and actinides that are produced in the reactor which lower the
reactivity of spent fuel. However, this entails some limitations on the selection of the
fuel elements to be disposed or the amount of fuel to dispose in a single canister. Further
analyses are going on (Section 14.2).
Handling before disposal
The canisters shall be stored, transferred and emplaced in a way that the copper overpack
is not damaged.
The canister handling, transfer and deposition has been planned to be made in a controlled
way to minimise the possible damages, disturbances or accidents in the process,
refs. (Kukkola 2012; Saanio et al. 2012; Rossi & Suolanen 2012) and Sections 12.1-
12.3. Excessively damaged canisters are not disposed, but returned to the fuel handling

139
cell, unload the fuel elements and re-encapsulate the fuel in intact canister. The canister
condition is assessed at several stages of the process including the surface condition
monitoring immediately before lowering the canister into the deposition hole (Section
12.2).
Retrievability
Design of the canister shall facilitate the retrievability of spent fuel assemblies from the
repository.
The retrieval of a disposed canister has been planned in principle in (Saanio & Raiko
1999). The important international test for Posiva’s demonstration on this topic has
been the canister retrieval test made by SKB at Äspö in 2006. As a main conclusion the
freeing trial showed that the tested method works. The retrieval operations in various
phases of disposal are described in (Saanio et al. 2012, Section 5.4).
Safeguards
Encapsulation and disposal of the spent fuel shall be organised in a way that makes the
safeguards control of the nuclear material possible according to requirements of (YVL-
D.5, Section 5.4).
The requirements to provide the safeguards controls in the spent fuel handling chain
are taken into account in the planning of the constructions and operations of the encapsulation
plant, transportation and repository (Saanio et al. 2012, Section 4.7).
This report summarises the fulfilment of all the design bases of the disposal canisters.
Even if some supplementary analyses and tests are still under way, no severe deficiencies
in the performance are detected or expected. For continuing research or development
activities, see Section 14.2.
14.2 Continuing research work on performance assessment
Research and testing is on-going in canister-related areas because of the very long testing
times or because of new information became available. The most important continuing
research and development areas are:
• Examination of residual (weld) stresses in copper
• Creep strength of copper (also weld) during very low-strain-rate creeping
• Copper corrosion in water
• Criticality safety especially of EPR/PWR canister
• Gasket design of the insert lid
• Comparison of alternative sealing methods, selection in 2013
• QA manual for canister manufacture
• Manufacturing demonstrations of especially VVER-440 and EPR/PWR type
canister versions will be continued
• Demonstration and documentation of bentonite buffer design and properties

140
• Demonstration and documentation of rock suitability criteria (RSC) for repository
and acceptable canister locations.
The canister design documentation will be updated accordingly for operation license
application, if new knowledge or analysis results are published by that stage.
Detailed design documentation of canister variants will be presented as a part of preexamination
documentation before the beginning of canister manufacture.
14.3 Uncertainty of analyses and assessments
Mechanical loads and strength analyses
The uncertainty of mechanical loads and strength analyses are discussed in chapter 7 of
the mechanical design report (Raiko et al. 2010). Generally, the ASME Code based
safety factors and analysis methods for nuclear power plant components are applied for
mechanical strength and fracture resistance analyses.
In mechanical analyses of canister, the material properties of the essential components
are based on demonstration manufacture and destructive testing of actual materials
taken from actual demonstration manufacture. The used material models (stress-strain
curve and fracture resistance curve) are based on lower bound values of tested properties
with 90 % reliability. Material tests are made in relevant temperature and, in case of
rock shear; even the strain rate (deformation speed) has been taken into account as far as
possible. Less important materials are modelled with lower bound values according to
relevant material standards. When we take into account the effect of wide quality control
programme applied on the safety class 2 components, the information of material
properties and the integrity of the materials of the canister components is highly reliable.
The load assumptions of the mechanical analyses are based on basic physics (hydrostatic
pressure, own weight), partly on laboratory experiments (bentonite swelling pressure),
climate evolution history (thickness of glaciation) or experience of geological
science (rock shear). Actual loads can vary, but the load assumptions are made conservatively.
Bentonite density is assumed to be at the maximum specified (corresponding
to density 2050 kg/m 3 ) all around the canister. In long term phenomenon, rock shear,
the bentonite is assumed to be chemically changed, from Na to Ca bentonite that induces
even higher swelling pressure and stiffness against shear.
The finite-element analyses for mechanical loads are made with well-known, widely
used and validated tools (Abaqus, Ansys and R6 computer programs) and using comprehensive,
usually 3D element models and experienced experts. Some of the strength
analyses are also repeated independently by various organisations (like JRC-Petten or
VTT) and the canister collapse pressure load is verified by testing.
Safety factors obeyed in the analysis results and shown existing additional margins for
mechanical strength are so robust that the smaller inaccuracies in the load postulations
or material modelling cannot change the acceptability of the assessment result.

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Thermal properties and thermal analyses
The thermal analyses are generally made with simple and theoretically well-founded
methods and the thermal properties are selected to reflect the real and expected values
of the properties and conditions or the system’s components. Conservatism has been
added to the allowable temperatures that are always set with a considerable margin
within the critical values.
Many of the thermal analyses concerning operational phase can be verified by simple
measurement during the pre-operation test phase of the encapsulation plant and repository.
As for large-scale rock thermal properties, additional data are collected during the
construction phase of the repository. Thermal dimensioning of the repository can be
updated later, if new data becomes available and the local canister distances can be
adapted accordingly even during plant operation.
The canister cooling condition in repository is analysed both with assumption of dry
conditions in the deposition hole and with water saturated condition. The maximum
temperature of the canister surface and the buffer takes place in about 15 years after the
disposal of the canister. The dimensioning temperature of the bentonite buffer is
+100 °C. In dry condition the nominal temperature is set to +95 °C. The 5 °C margin is
showed to be enough for natural variation of bedrock thermal properties in (Ikonen &
Raiko 2012). In normal water saturated conditions the cooling capacity is much higher
and the nominal maximum temperature will lay at about +75 °C. As for temperatures, it
is essential to remember that the canister is not loaded by major mechanical (hydrostatic
pressure or swelling pressure) loads when it is in dry conditions. In saturated condition,
instead, the pressure loads are possible. This means that for creeping analyses the highest
actual temperature with mechanical loads is about +75 °C.
In the dimensioning case, dry condition, the most important thermal resistance for canister
cooling chain is the 10 mm air gap between the canister and the buffer. The thermal
conduction over the gap is the sum of thermal radiation and conduction in the air.
Before the maximum temperature is reached after some 15 years, the elevated temperature
oxidizes the canister surface from the oxygen in the trapped residual air in the
deposition hole and tunnel. The increasing oxidation makes the canister surface emissivity
better and thus decreases the thermal resistance of the gap and lowers the canister
temperature. When the buffer gets water, the thermal resistance is remarkably lowered
and the temperature is lowered, too. The canister surface is matte after turning and
somewhat oxidized during storage after encapsulation in canister storage of encapsulation
plant or repository in a ventilated room for typically a couple of weeks. Thus the
emissivity coefficient used for dry condition analyses, 0.3, is conservative at least for
longer term perspective in repository condition, in other words, when the maximum
temperature is expected after a few years.
Environmental circumstances and corrosion
The very long-term corrosion behaviour is difficult to predict exactly because the environmental
circumstances may vary. Thus a remarkable conservatism been used for definition
of the corrosion resistant copper overpack wall thickness.

142
The corrosion analysis uses robust environmental and evolution postulates and concludes
that 30 mm copper is much enough. The actual thickness of the designed canister
construction is 49-50 mm, which gives allowance for material defects up to 14-15 mm.
Nuclear engineering analyses
For nuclear engineering analyses, the best estimate parameters have been selected and a
reasonable safety factor for the allowable values has been adopted.
More detailed analyses of criticality safety issues including some further validation of
nuclear engineering computer codes and an assessment of uncertainties will be performed
and reported in the near future.
The activity inventory and the decay heat propagation has been examined and reported
and the reliability of this data is high. The prediction of the decay heat of the computer
code that is used for analyses is verified with a wide BWR- and PWR-fuel calorimetric
testing in CLAB interim spent fuel storage. The prediction accuracy of the decay heat is
expected to be about ±2 %.
The radiation shielding calculations are made several times, first in 2D and later in 3D,
and the results have been compatible. And furthermore, the radiation dose rates outside
canister are well below specification, thus no problems are expected in this respect.

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15 SUMMARY
Design bases and canister dimensioning
This canister design report summarises all the design aspects that are set for the successful
performance of a disposal canister used in a KBS-3 type repository at Olkiluoto,
Finland. The canister shape, dimensions and materials are shown to fulfil all the set
practical and theoretical requirements.
Nuclear engineering design
The nuclear engineering design analyses describe how the activity inventory, decay
heat, radiation dose absorbed in the construction material, radiation dose rate outside the
canister, and the potential for criticality is assessed in the short and long term.
Canister manufacture
The overall manufacturing methods for canister components are described. Some alternative
manufacturing methods are discussed. The possible adverse effects of manufacture
are also discussed.
Encapsulation, canister handling and disposal
The fuel preparation, encapsulation and canister handling in the encapsulation plant are
described to give a general view of the disposal process. The canister transfer, buffer
storage and installation into the deposition hole are described. Possible disturbances and
handling errors are also discussed.
QA and QC
The obeyed quality assurance (QA) procedure for canister manufacture, encapsulation
and canister handling until disposal is described in canister manufacture manual. The
requirements are given as design requirements and manufacturing specifications. Then,
the quality control (QC) programme to verify the qualification of the components and
operations is described. The acceptance criteria are based with analyses and adequate
safety factors are used to increase the reliability of the final product.
Corrosion resistance
The maximum expected corrosion depth for the canister over 10 6 years is less than 30
mm. The design requirement that the copper canister shall be intact at least 10 5 years is
thus fulfilled in the conditions expected in the repository during its evolution (King et
al. 2011b).
Thermal analyses
Thermal behaviour of canister in repository and in a fire during transportation has been
thoroughly analysed and reported. The thermal behaviour of the fuel rods inside a can-

144
ister have also been analysed earlier. Now, in this design summary report, the thermal
analyses are expanded to cover the canister in encapsulation, buffer storage, and transportation.
The effect of welding method on the canister insert temperature is also analysed.
In addition, the long term behaviour of the canister in the repository is analysed
using alternative assumptions for the bentonite buffer condition around the canister. The
natural cooling properties of the canister in all postulated conditions from encapsulation
to final disposal seem to be sufficient with large safety margins. For operational safety
in case of loss-of-cooling the canister’s thermal capacitive response is analysed. Canister
temperature is increasing only 15 °C/day in case all cooling is accidentally lost for
some reason. This gives reasonable time to make corrective actions in case of failures in
ventilation operation.
Strength analyses and material properties
The mechanical strength of the canister has been studied. The loading processes are
adopted from Design Basis report and some of them, especially the uneven bentonite
swelling cases, are further developed in (Börgesson et al. 2009). The canister geometry
is described in detail including the manufacturing tolerances of the dimensions. The
canister material properties are summarised and the wide material testing programmes
and model developments are referenced. In addition to reference canister design, the
canister variants and some alternative manufacturing routes are also assessed.
The combination of various load cases are analysed and the conservative combinations
are defined. The probabilities of incidence of various load cases and combinations are
also assessed for setting reasonable safety margins. The safety margins are used according
to ASME Code principles for safety class 1 components.
The design bases load cases are analysed with 2D- or global 3D-finite-element models
including large-deformation and non-linear material modelling and, in some cases, also
creep. The integrity assessments are partly made from the stress and strain results using
global models and partly from fracture resistance analyses using the sub-modelling
technique. The sub-model analyses utilize the deformations from the global analyses as
constraints on the sub-model boundaries and more detailed finite-element meshes are
defined with defects included in the models together with elastic-plastic material models.
The J-integral is used as the fracture parameter for the postulated defects. The allowable
defect sizes are determined using the measured fracture resistance curves of the
insert iron as a reference with respective safety factors according to the ASME Pressure
Vessel Code requirements.
The asymmetric loads that may exist due to the uneven wetting process during the first
decades and due to density or geometry variations of the bentonite buffer later in the
saturated condition were shown not to be a design basis load case (Andersson-Östling &
Sandström 2009; Börgesson et al. 2009).
Based on the BWR canister analyses, the following conclusions can be drawn. The 45
MPa isostatic pressure load case shows very robust and clear results in that the risk for
global collapse is vanishingly small (10 -50 ) according to (Dillström 2009). Further, the
copper overpack will remain intact after such expected events despite that a number of

145
worst case events are taken into account. In addition to postulated loads, also a disturbance
scenario of freezing of bentonite buffer down to -5 °C during permafrost has been
analysed and shown not to be critical for mechanical strength of a canister.
For the shear load case the stresses and strains in the canister are high, depending on the
shear amplitude, shear angle and the intersection point. The design basis case for the
insert is perpendicular to the canister’s main axis at about ¾ of its length while the design
basis for the copper overpack is 22.5º to the main axis.
Mechanical modelling uncertainties
The main uncertainty in the calculations concerns the anticipated changes in the bentonite
properties when high strain rate data becomes available. The copper overpack is
made of soft (hot-deformed) copper and thus its ability to tolerate deformation is especially
high. The design case of the 5 cm rock shear leads to equivalent plastic strains
typically between 5 and 23 %, predominantly in locations of geometrical discontinuities
(or even at geometric singularities). This observation applies directly to the short-term
analysis and roughly the same results also apply to the creep analysis. This means that
creep has no important role in the rock shear case and that the plastic and creep elongation
in copper are so high that the copper overpack will manage the deformation. The
insert also may experience a slight plastic deformation due to the shear load, but the
effective stress remains below the ultimate tensile stress even in and around geometric
discontinuities; thus no damage on the insert is expected.
Damage tolerance
The damage tolerance analysis for the different load cases leads to a number of requirements
on inspection of the insert where the most rigorous requirements are derived
from the shear load case. The inspection requirements from the 45 MPa case are however
more modest. The rock shear case is the design basis case for the insert. The allowable
fault sizes are determined according to code practises for nuclear power components.
The resent results from cast iron fracture resistance testing (Planman 2012) have
given remarkably higher results. If the better trend is later statistically verified, the allowable
flaw size in the insert may be increased for the rock shear case.
For the copper overpack, it is important to avoid even smaller impact damage or other
cold work in the regions of the bottom and lid in order not to jeopardize the creep ductility,
this is planned to be confirmed by appropriate inspections before canister is disposed.
The lifting safety puts very modest requirements on the material integrity inspection
of the lifting flange.
Creep
The creeping of copper and especially the copper EB-weld has been examined for canister
reliability and lifetime. The copper overpack is plasticising or creeping due to external
loading in elevated temperature until the copper overpack goes into contact with the
cast iron insert that is essentially the load bearing member of the canister structure. The
maximum strain in copper due to this pre-planned deformation is, however, reasonably

146
low, only some per cent, at maximum (Holmström et al. 2012b). Thus the copper and
even the copper EB-weld creeping properties are ductile enough to bear this kind of
straining.
The canister has also been shown to have a good tolerance against material defects
(Raiko et al. 2010, Section 8.3.3).
The creep of iron at relevant temperature is ignored according to reasons given in (Martinsson
et al. 2010). In addition, higher temperature and remarkable loads on canister
insert are never present simultaneously.
Canister initial state
The canister’s initial state, concerning the fuel elements, inner materials, void atmosphere,
dimensions, and material integrity, is described and summarised in this report.
The initial state is the starting point for the long-term safety assessment of the engineered
barrier system.
Compliance with design requirements
All the requirements given in the design bases are presented in the report and it is
shown that the canister design fulfils the performance requirements of the long term
containment function. Uncertainties and safety margins are discussed and possible ongoing
research activities are referred to, where applicable.

147
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LIST OF REPORTS
POSIVA-REPORTS 2012
_______________________________________________________________________________________
POSIVA 2012-01
POSIVA 2012-02
Monitoring at Olkiluoto – a Programme for the Period Before
Repository Operation
Posiva Oy
ISBN 978-951-652-182-7
Microstructure, Porosity and Mineralogy Around Fractures in Olkiluoto
Bedrock
Jukka Kuva (ed.), Markko Myllys, Jussi Timonen,
University of Jyväskylä
Maarit Kelokaski, Marja Siitari-Kauppi, Jussi Ikonen,
University of Helsinki
Antero Lindberg, Geological Survey of Finland
Ismo Aaltonen, Posiva Oy
ISBN 978-951-652-183-4
POSIVA 2012-03 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Design Basis 2012
ISBN 978-951-652-184-1
POSIVA 2012-04 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Performance Assessment 2012
ISBN 978-951-652-185-8
POSIVA 2012-05 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Description of the Disposal System 2012
ISBN 978-951-652-186-5
POSIVA 2012-06 Olkiluoto Biosphere Description 2012
ISBN 978-951-652-187-2
POSIVA 2012-07 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Features, Events and Processes 2012
ISBN 978-951-652-188-9
POSIVA 2012-08 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Formulation of Radionuclide Release Scenarios 2012
ISBN 978-951-652-189-6
POSIVA 2012-09 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Assessment of Radionuclide Release Scenarios for the Repository
System 2012
ISBN 978-951-652-190-2

POSIVA 2012-10
Safety case for the Spent Nuclear Fuel Disposal at Olkiluoto - Biosphere
Assessment BSA-2012
ISBN 978-951-652-191-9
POSIVA 2012-11 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Complementary Considerations 2012
ISBN 978-951-652-192-6
POSIVA 2012-12 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto -
Synthesis 2012
ISBN 978-951-652-193-3
POSIVA 2012-13 Canister Design 2012
Heikki Raiko, VTT
ISBN 978-951-652-194-0