Materials Aging Issues and

Aging Management for Extended

Storage and Transportation of

Spent Nuclear Fuel

Office of Nuclear Material Safety and Safeguards

NUREG/CR-7116

SRNL-STI-2011-00005
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Materials Aging Issues and

Aging Management for Extended 

Storage and Transportation of 

Spent Nuclear Fuel

Manuscript Completed: September 2011

Date Published: November 2011

Prepared by

R. L. Sindelar, A.J. Duncan, M.E. Dupont, P.-S. Lam,

M.R. Louthan, Jr., T.E. Skidmore

Savannah River National Laboratory

Aiken, SC 29808

R.E. Einziger, NRC Technical Project Manager

NRC Job Code J5657

Office of Nuclear Material Safety and Safeguards

NUREG/CR-7116

SRNL-STI-2011-00005

iii

ABSTRACT

An evaluation of potential degradation of materials in the Systems, Structures, and 

Components (SSCs) of dry cask storage systems (DCSS) for Independent Spent Fuel 

Storage Installations (ISFSIs) has been performed to identify issues in the technical 

bases for regulating storage well beyond the present allowable storage time of 120 years 

that includes 60 years pool storage and 60 years dry storage. The evaluation included 

low (≤ 45 GWd/MTU) and high burnup (HBU) (> 45 GWd/MTU up to approximately 

62.5 GWd/MTU) fuel, and was performed to determine if: 1) degraded conditions could 

be anticipated to develop in the materials of the SSCs over an extended storage and 

transportation (EST) period (> 120 years); and 2) the anticipated degradation would 

challenge functions important to safety during storage and during post-storage 

transportation including fuel retrieval for final disposition.

The approach followed a general Aging Management Review process by: 1) identifying 

the materials of construction of the SSCs; 2) evaluating their susceptibility to degradation 

from environmental conditions of exposure; and 3) evaluating the potential for 

degradation to impact the safety functions required by the present Code of Federal 

Regulations for dry storage systems and spent fuel transportation in 10 CFR Part 72 and 

10 CFR Part 71, respectively. The evaluation was performed using extrapolation of 

available data and materials degradation models with consideration of service histories 

of similar materials systems, and with expert judgment.

With the present materials systems and controls for a DCSS, no materials degradation 

phenomena were identified that would be expected, with certainty, to be operative and 

cause an impact sufficient to violate the safety functions for storage and transportation. 

However, several data/information/technology gaps were identified that would require 

resolution to provide a sufficient technical basis to demonstrate the continuing integrity of 

the SSCs and their ability to maintain functions important to safety throughout EST.

Recommendations for activities to resolve the potential materials aging issues are 

provided. Aging management activities for preventive maintenance and periodic 

condition assessments through inspections of SSCs are outlined. Additional 

recommended activities include materials testing and a full cask demonstration project 

under conditions that challenge the SSCs of the DCSS to verify the expected materials’ 

behavior and provide confidence that the SNF, as is presently conditioned and drystored, can continue to be safely stored for greater than 120 years, and handled and 

transported following storage.

v

TABLE OF CONTENTS

ABSTRACT iii

TABLE OF CONTENTS v

FIGURES ix

EXECUTIVE SUMMARY xiii

FOREWORD xxv

ACRONYMS AND ABBREVIATIONS xxvii

1. INTRODUCTION 1

1.1 Objectives of the Aging Evaluation for EST 1

1.2 Approach 2

1.3 Assumptions 5

1.3.1 Management Strategy and Infrastructure ................................................................. 6

1.3.2 Functions Important to Nuclear Safety ..................................................................... 7

1.3.3 Potential Storage Environments ............................................................................... 8

2. AGING OF DRY STORAGE SSCS UNDER EST 11

2.1 Fuel Description 11

2.1.1 Early Domestic Cladding Alloy Types..................................................................... 14

2.1.2 Current Domestic Cladding Alloy Types................................................................. 16

2.1.3 Potential New Advanced Metal Alloy Cladding....................................................... 18

2.1.4 Additional Fuel Design - Fort St. Vrain SNF........................................................... 19

2.1.5 Additional Fuel Design - Mixed Oxide Fuel (MOX)................................................. 20

2.1.6 Effects of Burnup, Pool Storage, and Drying on Fuel Condition ............................ 23

2.1.7 Reference Fuel Assembly....................................................................................... 40

2.2 Fuel Degradation Phenomena 42

2.2.1 Creep ...................................................................................................................... 43

2.2.2 Hydrogen-Induced Degradation Processes - Overview ......................................... 47

2.2.3 Delayed Hydride Cracking (DHC) - Potential for DHC during Extended Storage .. 52
vi

2.2.4 Flaw Stability Assessment in Zircaloy Tube ........................................................... 58

2.2.5 Hydride Embrittlement – Reduction in Tensile Ductility.......................................... 64

2.2.6 Hydrogen Embrittlement - Hydrogen Induced Shifts in the DBTT.......................... 67

2.2.7 SCC of Fuel ............................................................................................................ 70

2.2.8 Fission Product and Helium Gas Pressurization .................................................... 72

2.2.9 Creep and Diffusion Controlled Cavity Growth (DCCG)......................................... 74

2.2.10 Air Oxidation of the Cladding – Off-Normal Condition............................................ 76

2.2.11 Air Oxidation of the Fuel Pellet Material – Off-Normal Condition ........................... 77

2.3 Baskets and Neutron Absorbers Description 79

2.4 Baskets and Neutron Absorbers Degradation Phenomena 80

2.5 Canisters and Weld Description 83

2.6 Canister Degradation Phenomena 86

2.6.1 Corrosion of Canister Materials .............................................................................. 90

2.6.2 Stress Corrosion Cracking and Pitting of Austenitic Stainless Steel Canisters...... 91

2.7 Bolted Closures Description (Bolts and gaskets) 99

2.8 Bolted Closures Degradation Phenomena 100

2.8.1 Metallic Seal Degradation..................................................................................... 100

2.8.2 Polymeric Seal Degradation ................................................................................. 104

2.8.3 Bolt Degradation ................................................................................................... 109

2.9 Casks Description 111

2.10 Casks Degradation Phenomena 113

2.10.1 Carbon Steel Material Corrosion .......................................................................... 113

2.10.2 Neutron Shielding Degradation Phenomena........................................................ 113

2.11 Concrete Pad Description 114

2.12 Concrete Pad Degradation Phenomena 115

2.12.1 Carbonation of Concrete....................................................................................... 117

2.12.2 Chloride Ingress.................................................................................................... 122
vii

2.12.3 Shrinkage of Concrete .......................................................................................... 124

2.12.4 Freeze-Thaw of Concrete ..................................................................................... 125

2.12.5 Calcium Leaching of Concrete.............................................................................. 128

2.12.6 Sulfate Ingress...................................................................................................... 130

2.12.7 Temperature Effects and Radiation Effects on Concrete ..................................... 131

2.12.8 Alkali-Silicate Reactions ....................................................................................... 133

3 RECOMMENDATIONS FOR ACTIVITIES TO ESTABLISH THE TECHNICAL 

BASES TO ENABLE EST 137

3.1 Maintenance Program 138

3.2 Monitoring & In-Service Inspection Program 138

3.3 Testing and Analysis 139

3.4 Cask Demonstration Considerations 143

4. OFF-NORMAL AND ACCIDENT CONSIDERATIONS 145

ix

FIGURES

Figure 1 Steps to Identify Materials Aging Issues for EST and Provide Resolution 

Through Aging Management for SSCs in DCSS. ...........................................4

Figure 2 PWR Fuel Assembly....................................................................................12

Figure 3 BWR Fuel Assembly [http://www.nucleartourist.com/images/bwrfuel1.jpg]. .12

Figure 4 Reactivity of MOX and UO2 fuel depending on burnup 

[reproduced from ref. 1]................................................................................21

Figure 5 Decay heat of UO2 and MOX FAs depending on burnup and cooling time 

[reproduced from ref 1].................................................................................21

Figure 6 Neutron activity of UO2 and MOX FAs dependent on burnup and cooling 

time [reproduced from ref. 1]........................................................................22

Figure 7 Fuel Cladding Changes with Burnup/Duty Cycle..........................................24

Figure 8 Maximum Wall Thickness Average (MWTA) Hydrogen Content in Low-tin 

Zircaloy-4 Cladding [Figure reproduced from ref. 10]. A “transition knee” in 

hydrogen pick-up with burnup at approximately 45 GWd/MTU is suggested

...................................................................................................28

Figure 9 A light micrograph showing a hydride layer and associated hydrides in the 

substrate beneath the layer in irradiated Zircaloy-4 cladding tube (average 

fuel burnup of 67 GWd/t and fast fluence of 1.3 x 1022 n/cm2

) 

(Courtesy of R. Daum, Argonne National Laboratory). 

[Figure reproduced from ref. 12]...................................................................29

Figure 10 Example of a hydride rim introduced in an unirradiated clad sample 

[Figure is reproduced from ref. 13, Fig. 5 - Wall-through crack in sample 

with hydride rim failed in room temperature burst test with pressurization 

rate of 1.9 GPa/s].........................................................................................30

Figure 11 Example of a hydride blister in an unirradiated sheet of zircaloy 

[Figure reproduced from ref. 14]...................................................................31

Figure 12 Example of a hydride blister in a fuel cladding 

[Figure reproduced from ref. 15]...................................................................31

Figure 13 Example of a hydride blister in a fuel cladding 

[Figure reproduced from ref. 16]...................................................................32

Figure 14 Example of ID crack in fuel cladding [Figure reproduced from ref. 17]. .......33

Figure 15 Example of ID crack in fuel cladding [Figure reproduced from ref. 18]. .......33

Figure 16 Hydrogen (H) solubility for zirconium alloys: () = Zr-2; () = Zr-4; 

and (+) = Zr2.5Nb [Figure reproduced from ref. 21]......................................36

Figure 17 TSS for Pure Zirconium [Figure reproduced from ref. 21]. ...........................37

Figure 18 Creep Mechanisms by atom or defect complex migration (left) and 

dislocation formation/moverment mechanisms (right). .................................44

Figure 19 Zr-H Phase Diagram [Figure reproduced from ref. 3]...................................50

Figure 20 Hydride Reorientation in the Near Fracture Region of a Zircaloy-2 Tensile 

Sample Tested at Room Temperature. A) Before Test, B) After Test 

(Archival data from M.R. Louthan, Jr.)..........................................................51

Figure 21 Postulated Flaws in the Cladding.................................................................59

Figure 22 An axial semi-elliptic crack with aspect ratio a/c in a fuel rod with inner 

radius Ri and cladding thickness t................................................................60

Figure 23 Stress intensity factors at 100 MPa uniform stress loading for semi-elliptic 

cracks with length 2c= 1.7t in a typical fuel rod as functions of penetration 

depth (a/c)....................................................................................................61
x

FIGURES (continued)

Figure 24 Comparison of stress intensity factors at crack tips of internal and 

external thumbnail cracks at 100 MPa uniform stress loading......................61

Figure 25 An axial throughwall crack with length 2c loaded by a uniform stress 

across the thickness t in a tube with mean radius R.....................................62

Figure 26 Stress intensity factors for axial flaws with 100 MPa uniform stress 

loading at various initial crack lengths..........................................................63

Figure 27 Impact of Hydrogen Content on the Tensile Properties of Zircaloy-4 

at Room Temperature [Figure reproduced from ref. 2]. ................................66

Figure 28 Ductile-to-brittle transitions in hydrided zirconium 

[Figure reproduced from ref. 2].....................................................................68

Figure 29 Exposure Time Dependence of Iodine-Vapor-Induced SCC 

[Figure reproduced from ref. 2].....................................................................71

Figure 30 Fission gas release from UO2 fuel with burnup 

[Figure reproduced from ref. 1].....................................................................73

Figure 31 Helium Build-up in Fuel Pin (CANDU Fuel) 

[Figure reproduced from ref. 2].....................................................................73

Figure 32 Typical Dry Cask Storage Systems. The cask may be an integrated cask 

with confinement and shielding in a integrated structure. The dry storage 

cask may include a separate canister (dual-purpose canister) over which 

an overpack is placed to provide shielding for storage or a cask is placed 

for transportation [http://www.nrc.gov/waste/spent-fuel-storage/diagramtypical-dry-cask-system.html].......................................................................84

Figure 33 Generic Canister Sketch Showing Basket Internals. ....................................86

Figure 34 Residual Fracture Toughness After Low Temperature (Tirr ~150°C) 

Neutron Irradiation on Type 304/304L Stainless Steel and Type 308 

Stainless Steel Weld Metals [Figure reproduced from ref. 3]........................87

Figure 35 Jdeformation-T Curves for Base (BLC, BCL), Weld (WLC, WCL), and HAZ 

(HLC, HCL) Components of Low-Temperature-Irradiated (Tirr ~150°C) 

Type 304 Stainless Steel. The CL crack orientation is parallel to the rolling 

direction of the plates used to fabricate the stainless steel tank 

[Figure reproduced from ref. 3].....................................................................88

Figure 36 Example of Canister Surface Temperature vs. Time 

[Figure reproduced from ref. 4].....................................................................89

Figure 37 SEM Micrographs of Stress Corrosion Crack in the Heat-Affected Zone 

in a Type 304L Stainless Steel Coupon with an Autogenous Weld. The 

specimen was exposed, at room temperature, to salt mixtures that 

contained about 0.6 wt% moisture [SRNL Micrograph]. ...............................91

Figure 38 Through Wall Stress Corrosion Cracks in a Welded Stainless Steel 

Container not Subjected to Applied Stresses [SRNL Micrograph]. ...............92

Figure 39 Pit and Stress Corrosion Crack due to Fly Ask Contact with Austenitic 

Stainless Steel [SRNL photograph]..............................................................93

Figure 40 Temperature Profile on MPC [Figure reproduced from ref. 5]. .....................95

Figure 41 SCC Growth Rate on Weldment Components of MPC 

[Figure reproduced from ref. 5].....................................................................95

Figure 42 General structure of ‘Helicoflex’ type metal seals 

[Figure reproduced from ref. 2]...................................................................100

Figure 43 Construction principles of HELICOFLEX® double jacket seals 

[Figure reproduced from ref. 2]...................................................................101
xi

FIGURES (continued)

Figure 44 Extrapolation of seal force decrease over 40 years under constant 

conditions [Figure reproduced from ref. 3]..................................................102

Figure 45 Time-Temperature Superposition of Compression Stress Relaxation 

Data for Viton GLT O-rings (Parker V0835-75) at 80 °C 

[Figure reproduced from ref. 7]...................................................................107

Figure 46 Time-Temperature Superposition of CSR data for EPDM seals 

[Figure reproduced from ref. 2]...................................................................107

Figure 47 Loss of Sealing Force Can Occur with Multiple Adjacent Bolt Breakage....110

Figure 48 Generic Dry Storage Cask Sketch for an Overpack on a Canister. ............112

Figure 49 Relative volume of possible rebar corrosion products 

[Figure reproduced from ref. 3]...................................................................115

Figure 50 Spallation of Concrete Induced by Rebar Corrosion [SRNL Photograph]..116

Figure 51 Carbonation depth with time for normal atmospheric CO2 exposure 

indoors of P-reactor building at the Savannah River Site plotted with 

experimental data by Nagataki et al. [Figure reproduced from ref. 3]. .......118

Figure 52 Cross-section of a 3.7 inch (9.4cm) diameter concrete core taken from the 

inside wall of the SRS P-reactor building. A phenolphthalein solution has 

been applied to indicate the extent of concrete carbonation. Lack of color 

indicates a pH below 10.2 showing that carbonation has progressed 

~ 1.4 inches (3.6cm) in from the surface of the inner wall. 

[Figure reproduced from ref. 10].................................................................119

Figure 53 Tight crack (see arrows) from surface (bottom) to interior of concrete 

providing a pathway for enhanced diffusion. Evidence of enhanced 

diffusion of CO2 adjacent to crack is observed. 

[Figure reproduced from ref. 10].................................................................120

Figure 54 Chloride ingress (left to right) limited to enhanced near-surface 

concentration in a concrete specimen from the SRS P-Reactor 

Building with 60 year’s service [Figure reproduced from ref. 10]. ...............123

Figure 55 Example of surface cracking on concrete pad at Test Area North in 

Idaho [Figure reproduced from ref. 4].........................................................126
xii

TABLES

Table 1 Nominal Compositions of Alloys Used as Nuclear Fuel Cladding .................14

Table 2 Influences of Heat Treatment on Creep of Zr-2.5Nb Alloy 

at 545°C (818K) [1] ......................................................................................45

Table 3 Predicted Hoop Strains of Nb-Modified Zirconium Cladding Alloy Under 

Extended Dry Storage Conditions [Table reproduced from ref. 2] ................45

Table 4 Sealing Force after Temperature Exposure for 1 Year [3] ..........................102
xiii

EXECUTIVE SUMMARY

Dry storage of spent nuclear fuel (SNF) has been implemented as an alternative to pool 

storage of fuel that has been removed from a reactor and is awaiting a long-term 

disposition decision. At this time, there are thousands of spent fuel elements safely 

stored at Independent Spent Fuel Storage Installations (ISFSI) around the world. Much 

is known about dry storage – the technical bases to demonstrate safe storage are 

sound, and the technologies for implementation of dry storage systems are mature. The 

Nuclear Regulatory Commission (NRC) has responsibility for regulating the dry storage 

of power reactor SNF in the United States (U.S.).

This document contains an evaluation of age-related degradation of the Systems, 

Structures, and Components (SSCs) of dry cask storage systems (DCSS) for ISFSIs 

and identifies gaps in the technical bases needed to demonstrate safe extended dry 

storage and transportation (EST) of low burnup (≤ 45 GWd/MTU) and high burnup 

(HBU) (>45 GWd/MTU up to approximately 62.5 GWd/MTU) spent nuclear fuel for its 

storage beyond the present allowable storage period of 120 years following reactor 

discharge, up to 300 years, followed by its transportation. The evaluation considers the 

variety of fuel designs and discharge conditions of SNF including MOX fuels.

There is a large body of information describing degradation mechanisms which can 

occur during dry storage of low burnup SNF. This information addresses fuels in dry 

storage for up to about 60 years and provides a firm technical basis for the licensability 

of dry storage facilities. Several critical investigations have considered the dry storage 

of this fuel in present storage facilities to 100 years. There is a trend to higher fuel 

burnup, new fuel compositions and new cladding alloys, and the possibility that the fuel 

will remain in dry storage for an even longer periods pending final disposition decisions.

This report serves to identify information needed to support this trend and extend the 

understanding of materials degradation to include the expected behavior of SNF and the 

materials of the DCSS for EST.

An evaluation of the potential age-related degradation of the fuel and materials in the 

SSCs of DCSS for ISFSIs during EST was made. This evaluation of the information 

needed to develop the technical basis to demonstrate that safety functions are met

during EST, showed four major data/information/technology gaps:

1. A wide variety of alloys are, and have been, used for fuel cladding in the U.S.

commercial fleet, and the response of these alloys to EST depends on chemistry, 

thermal mechanical process history, irradiation history and other 

fabrication/service variables. The information required to address the response 

of a specific cladding to a specific age-related degradation process is limited and 

in many cases simply not available.

2. The conditions required for corrosion-induced degradation processes such as 

stress corrosion cracking are, and will continue to be, present on the external 

surfaces of dry storage canisters, especially those fabricated from austenitic 

stainless steels. The information necessary to assess the time-temperature 

induced changes in cracking susceptibility is lacking and the technology 

necessary to inspect (remotely) for such cracking in the canisters has not been 

developed.
xiv

3. Delayed hydride cracking in the Zircaloy-clad fuel may occur and hydrogen 

embrittlement is anticipated as the fuel cladding temperatures fall during EST. 

The occurrence of these two phenomena would create cracks in a fuel cladding 

that has minimal fracture toughness. Limited data is available; however a 

validated fracture assessment technology is not available to establish the impact 

of these two degradation processes on assembly retrievability, handling, or 

transport.

4. Bolted closures with metallic seal materials will experience sealing force

reduction and may lose sealing capabilities during EST. The information needed 

to establish replacement schedules is lacking.

In addition to the four major gaps described above there are several other significant 

gaps in the available information. 

A) The heat output, attendant radiation, and reactivity of spent MOX fuels 

significantly exceed those of UO2 fuels; cask thermal and radiation design must 

address these conditions to avoid excessive temperatures for long times that 

would impact fuel and canister degradation.

B) There is a potential for delayed hydride cracking of fuel cladding and stress 

corrosion cracking of the canister occurring in the same dry storage cask. The 

impact of this occurrence on dry storage confinement, retrievability, handling, and 

post-storage transport has not been well developed.

C) The concrete pads, and also casks, will deteriorate and a path forward to 

address the impact of such deterioration has not been defined even though there 

are some ISFSI’s that are built on the only practical site at the reactor.

D) The lifetime of the ISFSIs, in some cases, will exceed the lifetimes of the 

corporation using the facility. The information and regulations required for 

successful transfer of data from a dying or defunct corporation to the new tenant 

is minimal at best. 

E) The aluminum based baskets and neutron absorbers could degrade during EST 

because of creep and/or corrosion processes. An evaluation is needed for 

canister internals with materials subject to creep deformation at anticipated 

storage temperatures (e.g., aluminum). Data are needed to demonstrate that 

such material degradation does not compromise the retrievability of the fuel 

assemblies or change the configuration of the absorber and limit its effectiveness 

to maintain subcriticality during re-submersion, if required. 

F) The technical basis for any assumptions in both the number of failed fuel rods, 

and radiological source terms from high burnup (HBU) fuel for normal and 

accident conditions needs to be developed for EST.

The evaluation that led to the points listed above included HBU (62.5 GWd/MTU) fuel 

and was performed to determine if degraded conditions could be anticipated to develop 

over a very long term (up to a nominal 300 years) storage period, and if the anticipated 

conditions could challenge functions important to safety during storage and/or 

transportation. It was clear that anticipated materials degradation processes could 
xv

cause failures of fuel and SSCs during the extended storage period and/or during poststorage transportation to a facility for final disposition. However, no materials 

degradation phenomena were identified that would, with certainty, occur to cause an 

impact sufficient to violate the safety functions for storage and transportation.

The approach to the degradation evaluation considered the types of materials and 

generic designs for the DCSS, and followed, in general, the Aging Management Review 

process listed in the Standard Review Plan for licensing and license renewal of an 

Independent Spent Fuel Storage Installation for 40 additional years of storage (NUREG1927 at http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1927/), with the 

exception that storage periods up to a nominal period of 300 years followed by handling

and transport, was considered. 

Additional environmental conditions (e.g., an increase in atmospheric carbon dioxide) 

due to man-made or natural climatic changes over EST were considered. The impact of 

off-normal/accident design-basis events such as floods, tornados, projectiles, 

temperature extremes, cask drop, cask tipover, and other unusual scenarios during EST 

on the materials of the DCSS were not considered in this report with the exception of 

uncontrolled air/water ingress into the canister due to failure of the canister.

The materials aging evaluation included a review of published material degradation 

processes and focused on well-referenced and highly-cited documents. The approach 

included extrapolation of available constitutive models and expert judgment applied to 

the materials used in the DCSSs with no specific time limit to the aging phenomenon, 

except for the general target of 300 years. 

There is a wide variety of fuel cladding alloys at various reactor service conditions, and 

several variations of DCSS design including materials of construction for the SSCs. A 

reference material and condition was constructed in some cases because of the 

multitude of materials options. The materials and fabrication of DCSS SSCs were 

assumed in the as-design condition; and in the case of fuel, the post-discharge condition 

of the fuel and the effects of drying were considered.

The table ES-1 summarizes results of the assessment of specific degradation processes 

(types of degradation) and their expected impact to safety functions of the SSCs of the 

DCSS. An important consideration is the degradation of the fuel cladding. While in the 

DCSS, a fuel cladding failure causing a release of radioisotopes to the confinement 

would increase the source term for confinement provided by the canister, but would not 

compromise confinement. It would impede the retrievability of the fuel by assembly, not 

canister. Additionally, cladding failure of the higher burnup fuel would be expected to 

have increased potential source term due to fission gases, volatiles, and fines produced 

from the rim that formed on the fuel pellets.

Aging management activities were recommended to address information gaps and 

ensure that the performance of the DCSS meets its safety basis during EST. A periodic 

monitoring/in-service inspection program is the principal strategy recommended to 

ensure safety function of the DCSS throughout its desired service. This program would 

include periodic inspection of the materials and the condition of materials and 

components of the SSCs of the DCSS. Acceptance criteria for disposition of indications 

attributed to aging-induced processes and other flaws must be established. Provisions 
xvi

for repair or replacement (recovery) would need to be available for SSCs that would fall 

outside the acceptance criteria.

The aging management activities also include preventive maintenance activities such as 

periodic replacement of gaskets and closure bolts. The gasket materials are expected to 

relax and lose sealing capacity during long term storage and the bolts may experience 

stress corrosion cracking during extended service.

General recommendations to obtain additional data/information to support resolution of 

the information gaps are outlined. The recommendations include those for laboratory 

tests and for a fuel storage demonstration project that would include storage and testing 

under challenging (e.g., HBU, partially-degraded fuel) conditions. It is recognized that 

programs to address information needs, including international programs, are in various 

stages of completion. The conclusions in this report would be subject to change as 

additional data and service experience with fuel in dry storage are obtained.
xvii

Table ES-1 Summary of the evaluation of fuel and SSCs of an ISFSI/DCSS for their susceptibility to materials a

under extended storage, and the potential impact of such degradation on the safety functions for EST. All credi

mechanisms for EST were considered. Generalized activities are proposed to resolve the issues. The set of de

mechanisms considered for the fuel are itemized in the table. The primary degradation mechanisms anticipated 

component are listed in the table. 

Fuel

(Zircaloy cladding 

failure mechanisms)

Degradation Under EST Impact on Safety Resolution (Recomm

Creep Creep leading to clad failure is not 

expected due to low temperature in 

EST, and the self-relief of the driving 

force (internal pressure) with creep 

deformation. If clad failure did 

occur, it would be a minor breach, 

not a large rupture. 

Clad failure could cause release of gases, 

volatiles, and some fines to the canister and 

impact the source term and fuel retrievability 

from dry storage.

Testing to verify existing 

model(s) applicability to t

behavior of HBU fuel cla

and to the new alloys, as 

to determine the univers

applicability of creep failu

strain of ~2%. Also, eval

phenomenon as part of c

demonstration project.

Delayed Hydride 

Cracking (DHC)

DHC, including the development of a 

through-wall (minor) crack, is 

possible, if not probable. However, if 

clad failure did occur, it would 

produce a minor breach, not a large 

rupture. Hydrogen diffusion and 

hydride reorientation are expected to 

be involved in the DHC processes.

Clad failure could cause release of gases, 

volatiles, and some fines to the canister. This 

would impact the source term and

retrievability from dry storage. DHC could also 

create seed flaws that are subject to instability

when the storage and/or transport 

temperatures are below DBTT. Growth of 

such flaws could cause gross failure during

handling and transportation and further

impact the retrievability and confinement.

Evaluate phenomenon d

fuel destructive evaluatio

part of cask demonstratio

project. Evaluate impact 

treatment on hydride 

microstructure to establis

fuel condition and asses

potential for DHC and HE 

EST.

Hydrogen 

Embrittlement 

(HE)

HE will reduce both the tensile

ductility and the fracture toughness

and shift in DBTT to higher 

temperatures. Higher burnup fuel 

has the potential for lower 

toughness and higher DBTT shifts 

than low burnup fuel due to

additional hydrogen accumulation 

and hydride precipitation. Hydride

orientation with respect to flaw

orientation affects the fracture 

toughness of cladding – flaws with 

radial hydrides causing low 

toughness for axially-oriented flaws.

Hydrogen embrittlement, in conjunction with 

cladding flaws could result in gross failure 

under high mechanical loads (off normal/ 

accident conditions and/or post storage 

handling) that may impact retrievability and 

cause a large cladding rupture. The cladding 

rupture would increase the source terms for 

transportation. Further, HBU fuel would have 

potentially higher source terms including fines 

due to rim effects.

Develop models with 

benchmarked data for te

fracture properties. Deve

consensus approach to f

control plan(s) for fuel ha

and transportation. Revis

consensus standard on d

provide explicit considera

hydride microstructure 

reconfiguration during dr

if practical, modify the st

reduce potential for hydr

reorientation that promot

and HE degradation.

xvii
xviii

Fuel, CONT’D

(Zircaloy cladding 

failure mechanisms)

Degradation Under EST Impact on Safety Resolution (Recomm

Stress 

Corrosion 

Cracking

SCC of cladding is not expected 

because most modern cladding 

alloys are lined with zirconium to 

protect against such degradation. If 

clad failure did occur, it would be a 

minor breach, not a large rupture.

Clad failure could cause release of gases, 

volatiles, and some fines to the containment 

of canister and impact post storage 

retrievability.

Evaluate, by destructive 

examination, the occurre

stress corrosion cracking 

cladding after prolonged 

Such evaluations should 

of cask demonstration pr

and not conducted simpl

assess SCC.

Helium 

Pressurization

Additional cladding internal 

pressurization by helium is not 

anticipated because the temperature 

will be decreasing and the increased 

helium concentration due to alpha 

decay is insignificant.

No degradation anticipated. None needed. 

Diffusion 

Controlled 

Cavity Growth 

(DCCG)

DCCG is a micromechanism of 

creep and is not anticipated.

Same as for creep. Same as for creep. 

Off-Normal Air 

Oxidation of 

Fuel Cladding

Oxidation of cladding with air inleakage that could result in metal 

wastage leading to a large rupture of 

fuel with time is not anticipated due 

to the relatively low cladding 

temperature.

No degradation anticipated. None needed for breach

Off-Normal Air 

Oxidation of 

Fuel Pellet

Oxidation of fuel pellets with air inleakage because of existing or 

developed though clad defects could 

result in pellet swelling and a large 

rupture of the fuel cladding with 

time. Air in-leakage could also affect 

cover gas conductivity, cladding 

oxidation and emissivity and alter 

heat transfer conditions. The 

potential impact of cladding 

oxidation is minimal compared to 

fuel pellet oxidation.

Gross failure during storage could cause 

release of gases, volatiles, and fines to the 

canister and impact the source term and 

retrievability. This would be especially 

significant to HBU fuel because of their higher 

source terms including fines.

Establish controls or mon

procedures to ensure ine

atmosphere is maintaine

xviii
xix

Fuel Canister 

Internals Degradation Under EST Impact on Safety Resolution (Recomm

Fuel Basket Loss of geometry of aluminium 

basket with temperature and time 

due to creep may occur. 

Loss of ready retrievability of fuel. Evaluate creep of alumin

basket materials for exte

storage thermal histories 

excessive, develop fuel r

strategy.

Neutron 

Absorbers

Design of some absorbers (the 

Boral absorber, B4C-Al in SS clad) 

could have led to inadvertent water 

ingress and with incomplete 

drying, and eventual corrosion (of 

aluminum) and H2 blistering with 

absorber crumbling. This is not 

expected in the newer design 

absorbers but may be present in 

the old design absorbers.

Loss of necessary criticality control for wet 

immersion only (not dry-to-dry transfer).

To assure criticality contr

during fuel retrieval, even 

neutron absorbers have 

degraded during storage 

plans to: 1) mitigate degr

by modifying drying stand

that assure water remova

developing dry-to-dry tra

procedures for fuel retrie

and/or 3) if wet transfer is 

necessary for shielding, r

the degraded neutron ab

in cask during fuel retriev

flooding with fluid such a

borated water.

xix
xx

Fuel Canister Degradation Under EST Impact on Safety Resolution (Recomm

Canister Body

(stainless steel)

The canister would be expected to 

experience pitting and cracking that 

could cause through-wall failure of 

canister under aggressive 

conditions. Increased condensation 

of moisture with time due to cooling 

exacerbates corrosion condition 

especially in DCSS located near 

the sea coast or highly industrial 

environments that are a source of 

aggressive species.

Canister/cask is credited as confinement 

boundary of the DCSS. Only a limited leak 

path is expected with a pitting/cracking breach 

of the canister. Release of gases, volatiles, 

and some fines could occur during storage. 

Cracks could open under high (e.g., from 

accident) loadings.

Establish in-service inspe

protocol with acceptance 

for pitting and cracking.

Closure 

Weldment

The weldment is susceptible to 

stress corrosion cracking that could 

cause through-wall failure. 

Aggressive conditions are the 

same as for the canister body. 

Canister is credited as confinement boundary. 

Only a limited leak path is expected if a 

cracking breach should develop.

Establish in-service inspe

protocol with acceptance 

for cracking.

Metallic Seals Loss of sealing force and seal 

resiliency is expected and that loss 

may cause a loss of sealing. Seals 

are also subject to corrosion 

including crevice corrosion.

Seal is necessary for confinement boundary of 

cask in storage.

Perform testing to evalua

sealing capability with los

sealing force with time. 

Replacement of metallic 

under EST is expected a

therefore facility (pool or 

system) may be needed.

Bolts Corrosion and stress corrosion 

cracking leading to breakage may 

occur.

Multiple adjacent bolt failures would cause loss 

of sealing force leading to loss of confinement 

of cask in storage.

Perform periodic bolt 

replacement as a mainte

item.

xx
xxi

Cask and Pad Degradation Under EST Impact on Safety Resolution (Recomm

Steel Cask

(carbon steel)

General and pitting corrosion if 

bare steel. No significant 

degradation expected contingent 

on periodic maintenance of coating.

Steel overpack provides shielding and 

structural stability functions and pitting not 

typically significant to these functions.

Establish plan for periodi

maintenance of steel (e.g

coat) as necessary.

Polymeric 

Shield Material

Medium potential for oxidationinduced mechanical property 

changes leading to slump within 

the overpack.

Polymeric materials provide n-shielding 

function. Shielding function of hydrogenous 

media would not diminish with structural 

change unless material re-configures in 

dimensions and slumps.

Establish plan for periodi

monitoring of shielding to 

for radiation leakage to a

present monitoring for DC

Concrete 

Overpack

Corrosion of steel in reinforced 

concrete leading to concrete 

spallation and loss of section is 

expected. Minimal impact to metalencased unreinforced concrete.

Concrete materials provide gamma and nshielding function and cooling geometry for 

maintaining thermal safety function.

Establish in-service inspe

protocol with acceptance 

Concrete Pad Corrosion of steel in reinforced 

concrete leading to concrete 

spallation and loss of structural 

function for cask and retrieval 

equipment foundation is expected.

Loss of structural stability of the pad under 

normal conditions would lead to cask tip-over 

potential with direct impact to retrievability and 

indirect impact (unquantified) to other safety 

functions of DCSS including shielding and 

thermal functions.

Establish in-service inspe

protocol with acceptance 

xxi

xxiii

ACKNOWLEDGMENTS

This report describes work done by the Savannah River National Laboratory for the U.S.

Nuclear Regulatory Commission (NRC). The activities reported here were carried out on 

behalf of the U.S. NRC, Office of Nuclear Material Safety and Safeguards. The authors 

greatly appreciate the contributions made by the materials experts in the NRC Division 

of Storage and Transportation in providing helpful discussions and critical reviews. 

This report is an independent product of the Savannah River National Laboratory and 

does not necessarily reflect the views or regulatory position of the U.S. NRC.

xxv

FOREWORD

This report is intended for use by the U.S. Nuclear Regulatory Commission (NRC) staff 

in its consideration of technical information needed to enable regulation of extended 

storage (up to 300 years) of spent fuel in a dry storage system at a fuel storage site and 

the subsequent safe transportation of the fuel. Specifically, the report provides 

information to support a qualitative prediction of material degradation with time for those 

Systems, Structures, and Components (SSCs) of a Dry Cask Storage System (DCSS) 

that are important to safety as required in part 72.24 of Title 10 of the Code of Federal 

Regulations, Part 72, “Licensing Requirements for the Independent Storage of Spent 

Nuclear Fuel, High-Level Radioactive Waste, and Reactor-Related Greater than Class C 

Waste.” It further provides information to support a qualitative prediction of the fuel 

condition for post-storage retrieval and transportation.

Additionally, because of the necessity to extrapolate engineering data for far beyond the 

time frames used for data gathering, some prediction of materials behavior must be 

based on engineering judgment. For example, several age-related degradation 

processes (e.g., stress corrosion cracking and delayed hydride cracking) are nucleation 

and growth processes and the data to demonstrate the lack of nucleation over a 

100+ year time frame does not exist. This report provides the evidence and background 

necessary for realistic, well-informed engineering judgments of SSC materials’ aging but 

does not attempt unrealistic extrapolations of short-term data.

This report also identifies the gaps in the technical bases for prediction of materials 

performance for extended storage, assesses the impact of the degradation to the safety 

functions, and provides general recommendations for activities to close the gaps and 

enable safe extended storage and post-storage transportation. The intent of the report 

is not to prescribe a detailed approach for issue resolution.

Furthermore, although the report adopts the general approach of the present Standard 

Review Plan (NUREG-1927 at http://www.nrc.gov/reading-rm/doc-collections/

nuregs/staff/sr1927/) to consider materials aging in a dry storage system that had an 

initial license period of 20 years with a 40-year extension, specific guidance for a 

licensing review is outside the scope of this report.

xxvii

ACRONYMS AND ABBREVIATIONS

ALARA

AMA

As Low As Reasonably Achievable

Aging Management Activity

AMP Aging Management Program

AMR Aging Management Review

ANL Argonne National Laboratory

ASME American Society of Mechanical Engineers

ASTM American Society for Testing and Materials

BRC Blue Ribbon Commission on America’s Nuclear Future

BWR Boiling Water Reactor

°C

CC

Degrees Celsius

Criticality Control

CFR Code of Federal Regulations

CoC Certificate of Compliance

DCSS Dry Cask Storage System

DE Destructive Examination

DHC Delayed Hydride Cracking

DSC Dry Storage Canister

DOE Department of Energy

DOT Department of Transportation

DSA Documented Safety Analysis

DWPF Defense Waste Processing Facility

EIS Environmental Impact Statement

EM Environmental Management (DOE)

EPRI Electric Power Research Institute

ER Environmental Report
xxviii

EST

ACRONYMS AND ABBREVIATIONS (continued)

Extended Storage and Transportation

°F

FSAR

Degrees Fahrenheit

Final Safety Analysis Report

GTCC Greater-Than-Class-C

GWd Giga-watt days

HBU High Burnup (> 45 GWd/MTU)

HE Hydrogen Embrittlement

HEPA High Efficiency Particulate Air

HLW High-Level Waste

HPGe High Performance Germanium detector

HRS Hydraulic Ram System (NUHOMS)

HSM Horizontal Storage Module (NUHOMS)

HT Heat Transfer

INL Idaho National Laboratory

INTEC Idaho Nuclear Technology and Engineering Center

ISFSI Independent Spent Fuel Storage Installation

ISG Interim Staff Guidance

ITS Important to Safety

KSI Kips per Square Inch (1000 lbs/in2)

LWR Light Water Reactor

MOX Mixed Oxide (typically plutonium and uranium oxides)

MPC Multipurpose Canister

MTHM Metric Ton Heavy Metal

MTU Metric Tonne of Uranium

MVDS Modular Vault Dry Storage
xxix

MW

ACRONYMS AND ABBREVIATIONS (continued)

Megawatt

N/A Not Applicable

NAC Nuclear Assurance Corporation

NDE Nondestructive Examination

NRC U.S. Nuclear Regulatory Commission

PB Pressure Boundary

PCI Pellet Cladding Interaction

PIE Post-Irradiation Examination

PM NRC Project Manager

PNNL Pacific Northwest National Laboratory

psig pounds per square inch, gauge

PVC Polyvinyl Chloride

PWR Pressurized Water Reactor

RS Radiation Shielding

SAR Safety Analysis Report

SCC Stress Corrosion Cracking

SER Safety Evaluation Report

SRP Standard Review Plan

SSC System, Structure, and Component

SME Subject Matter Expert

SNF Spent Nuclear Fuel

SRS Savannah River Site

SRNL Savannah River National Laboratory

TLAA Time-Limited Aging Analysis

TSS Terminal Solid Solubility
xxx

VCC

VLTS

VSC

ACRONYMS AND ABBREVIATIONS (continued)

Ventilated Concrete Cask

Very Long-Term Storage (alternate term used to describe 

consideration of dry storage beyond present licensing basis)

 

Ventilated Storage Cask System
1

1. INTRODUCTION

The United States strategy for the nuclear fuel cycle for its nuclear fleet, including the 

interim options for fuel management and waste form disposition, is under review by the 

Blue Ribbon Commission on America’s Nuclear Future (BRC).1

The U.S. Nuclear Regulatory Commission (NRC) has sponsored an independent 

evaluation of the bases for predicting materials behavior in dry cask storage systems 

(DCSS) for extended storage (> 120 years) and transportation (EST) because the spent 

nuclear fuels may need de facto storage in a DCSS for multiple decades. A time frame

up to 300 hundred years is considered a reasonable bound to this need and one

purpose of the evaluation is to summarize information, relative to licensing strategies,

needed to address such an extended storage schedule. This report describes the 

results of the material aging evaluation and the aging management activities considered 

necessary for extended storage and transportation.

 The final BRC report to 

the U.S. Department of Energy including a review of alternatives for storage, processing, 

and disposal of civilian used nuclear fuel and other materials is expected in January 

2012; however, no timetable has been established for decisions by the U.S. Department 

of Energy on alternatives selection. In addition, the technology development needed for 

full implementation of any selected alternative may require several additional decades of 

study and work. Therefore, fuel presently in either wet or dry storage systems may need 

to be stored well beyond the initial and re-license periods, and the transfer of used fuel 

from commercial utilities to the U.S. DOE may not occur in the near future.

1.1 Objectives of the Aging Evaluation for EST

Independent Spent Fuel Storage Installations (ISFSI) must provide for functions 

important to nuclear safety including thermal performance, radiological protection, 

confinement, sub-criticality control, and retrievability. These functions are required by 

the code of federal regulations governing irradiated fuel storage in 10 CFR Part 72 [2], 

and are provided by the systems, structures, and components (SSCs) of the dry storage 

systems as directed through the Standard Review Plan process [3, 4]. The original 

licenses granted by the NRC for DCSS’s were for 20 years.

The fact that extended dry storage of spent nuclear fuel in DCSSs may be required was 

recognized and a technical basis for such storage was developed and summarized in a 

major report in 2002 [5]. The focus of that work was on age related degradation of the 

fuel and cask materials. A potential storage period of up to 100 years was considered, 

and the work concluded that “the review conducted in this report finds no technical 

reason why extended storage (beyond 20 years) would lead to loss of any of the 

required safety functions.” Recently, formal consideration of the license period (and relicense period) was extended to 40 years through revision of 10 CFR Part 72 and the 

template of the Standard Review Plan process [6]. 

 

1 The Blue Ribbon Commission on America’s Nuclear Future (the Commission) was established in 

accordance with the provisions of the Federal Advisory Committee Act (FACA), as amended, 5 USC. App. 2, 

and as directed by the President’s Memorandum for the Secretary of Energy dated January 29, 2010: Blue 

Ribbon Commission on America’s Nuclear Future. This charter establishes the Commission under the 

authority of the US Department of Energy (DOE). [1]
2

The storage period for consideration in the current analysis was arbitrarily chosen as 

300 years, and includes high burn-up fuels that were not considered in the earlier 

evaluations. Therefore, one objective of the evaluations included in this present report is 

to expand the earlier considerations by extending the potential time of storage, including 

high burn-up fuels, and analyzing previous and emergent data for potential gaps in the 

data base. For example, the mechanical properties of zirconium based alloys are 

sensitive to crystallographic texture and texture is sensitive to the thermal-mechanical 

history of the alloy. These sensitivities raise a question about the relevance of 

mechanical property data developed on one alloy to another new alloy fabricated by 

different processing technologies.

The Standard Review Plan [6] for renewal of ISFSI licenses and DCSS CoC’s states that 

“the renewal is primarily a materials engineering effort” and provides a flow chart (Figure 

A of the SRP [6]) for the license renewal process. An Aging Management Review, the 

development of Aging Management Programs and the assurance of post-storage fuel 

retrievability are key elements to the flow chart. Previous evaluations have listed 

potentially active age related degradation mechanisms, in fuels for example, and 

evaluated the likelihood that the mechanisms will be active during the initial license 

period or during extended storage (Table 3-2 from reference [5]). 

The primary objective of this document is to identify and discuss materials degradation 

issues that may impact the safety functions required for the DCSSs to maintain low and 

high burn-up fuel in EST. The types of materials used in the existing DCSSs [7] were 

considered. The secondary objective of this document is to provide recommendations 

for activities (e.g., experiments, lead cask tests, monitoring programs, analyses, or other 

aging management activities) needed to address the aging issues, reduce uncertainties, 

and develop a sufficient technical basis for EST.

The 10 CFR Part 72 requirement for fuel performance in a DCSS is for “no gross rupture” 

of the fuel cladding for design basis conditions. The canister or cask is credited to 

provide the confinement function in a DCSS, and a defined leak rate of a shipping cask 

or overpack provides the confinement control for transportation; nevertheless, the 

viewpoint of the present evaluation is to avoid gross failure of the fuel under an extended 

dry storage period which is followed by transportation for final disposition, so that 

retrievability, even following normal conditions of transport, is maintained throughout 

EST. Transportation of irradiated fuel is governed by 10 CFR Part 71 [8] and the 

Standard Review Plan (SRP) for Transportation Packages for Spent Nuclear Fuel [9] 

provides guidance for the review and approval of applications for packages used to 

transport spent nuclear fuel under 10 CFR Part 71.

1.2 Approach

The degradation evaluation considered a set of materials and components currently 

used in DCSS and included selected emergent materials that may be used in future 

systems. The information for selecting the materials for evaluation was drawn from the 

general literature on dry storage systems. The degradation types and mechanisms 

potentially active for fuel and non-fuel materials in the SSCs under long-term dry storage 

are drawn from previous degradation analyses contained in key topical reports, and are 

supplemented with recent literature information and by considerations of the authors of 

this report. Nuclear power plant aging and aging management documents also provide 

input to this evaluation. Additionally, emergent information on HBU fuel and extended
3

storage systems and their impact on materials, components, and degradation 

evaluations are included.

Except for the fuel, the initial conditions of all other SSCs of the dry storage system are 

assumed to comply with the applicable national consensus codes and standards for 

design and construction, and have appropriate controls on workmanship and quality. 

Therefore, the as-designed, as-built conditions of the SSCs are not susceptible to failure 

as a result of defects in their initial material’s condition or because of poor fabrication 

practices.

The flowsheet in Figure 1 depicts the approach taken to identify materials issues or gaps 

in the technical bases necessary to demonstrate the capability of the SSCs to perform 

their intended safety functions under extended storage and during transportation after 

storage. Uncertainties in the information base regarding the ability of the SSCs to 

perform the safety functions are qualitatively discussed and engineering judgment is 

applied to the evaluation.

ASTM C-1562-10 (Standard Guide for Evaluation of Materials used in Extended Service 

of Interim Spent Nuclear Fuel Dry Storage Systems) [10] is a national consensus guide 

aimed at aging and degradation of materials of the SCCs of DCSS for a target time of up 

to 100 years total service. This guide was reviewed and provided many of the inputs for 

SSCs, their materials of construction, and potential degradation mechanisms that were 

considered in this report. 
4

Figure 1 Steps to Identify Materials Aging Issues for EST and Provide Resolution 

Through Aging Management for SSCs in DCSS.

Section 2 of this report evaluates the potential degradation phenomena and the impact 

of the degradation processes on the functions important to safety. The authors of this 

report and other subject matter experts evaluated potential degradation mechanisms 

and the potential of an operative mechanism to cause effects detrimental to the safety 

functions. The findings determine whether sufficient bases exist for extended dry 

storage and, when such bases do not exist, the report makes recommendations for 

Determine the Variety of Materials of Construction for 

Dry Cask Storage Systems for Present and Emergent 

Dry Storage Systems and the Varieties in Fuel to be 

Stored

Assess Materials and Storage Conditions to Establish 

Reference Situations for Evaluation of Component 

Response to Extended Storage Environments

Review/Assemble Available Information and Extrapolate 

Constitutive Models; Use Service Histories and Expert 

Judgement to Assist in Extrapolation of Degradation 

Processes. Project DCSS Material Component 

Conditions for Up to 300 Years of Storage

Evaluate the Available Information Base on Materials 

Degradation and the Impact of Potential Degradation on 

Required Safety Functions during Storage and PostStorage Transportation 

Identify Approaches to Resolve Information Gaps and 

Potential Materials Degradation Issues Significant to 

Safety Deficiencies Through Additional Tests, 

Analyses, or Monitoring. Outline Aging Management 

Activities Necessary to Resolve Potential Safety 

Issues. If no aging management option is available to 

avoid loss of safety function, assess 

repair/replacement strategy.
5

additional tests, analyses, periodic monitoring, and other activities to support the 

technical bases for EST.

Emphasis is placed on post-storage retrievability to assure that the 10 CFR Part 72 

requirement for retrievability from the dry storage system and the 10 CFR Part 71 

requirements for transportation and removal from the storage canister are met. 

Specifically, 10 CFR 72.122(h)(1) requires that spent fuel cladding must be protected 

from degradation that leads to gross rupture, or the fuel must otherwise be confined so 

that degradation of the cladding will not impose operational safety problems. Further, 

10 CFR 72.122(h)(1) requires that the storage system must be designed to allow ready 

retrieval of the spent fuel from the storage system for further processing or disposal.

The odd subsections 2.1, 2.3, etc. in Section 2 of this report describe the component and 

materials of the DCSS; the even subsections 2.2, 2.4, etc. describe the corresponding 

potential degradation types and their impact on the safety functions provided by the 

component. These even subsections include identification of the data needs and 

recommendations to address the needs. The GALL report [11] identifies Aging 

Management Programs (AMP) that were previously determined to be acceptable to 

manage aging effects of systems, structures, and components (SSC) in the scope of 

license renewal, as required by 10 CFR Part 54, “Requirements for Renewal of 

Operating Licenses for Nuclear Power Plants,” [12]. The template provided by the GALL 

report provides the foundation for the set of actions for the aging management program 

that this report recommends for the DCSS. 

Section 3 summaries the activities recommended in section 2 that, along with the body 

of work referenced in this report and supporting information, would establish the 

technical bases for enabling EST, and categorizes the activities into Maintenance; 

Monitoring/In-Service Inspection; and Testing & Analysis bins. Maintenance activities 

are periodic actions that are pre-emptive to loss of function. Monitoring/In-Service 

Inspection activities provide for a periodic condition assessment of the component and 

would disposition conditions found to be below the condition at installation. Testing & 

Analysis activities are primarily R&D activities to provide materials degradation 

information and strengthen the technical bases to predict and demonstrate component 

integrity throughout targeted service life. Section 3 also outlines the type of information 

to be gained from a test cask demonstration that would, for example, challenge (e.g., 

partially initially degraded, HBU) the fuel in a DCSS to provide validation of predictions 

and verification of materials’ conditions.

A risk-based evaluation to demonstrate the long-term acceptability of the DCSS is 

outside the scope of this report.

1.3 Assumptions

The first assumption in the development of this document was that the original 

placement of spent fuel in the DCSS conformed to the licensing requirements at the time 

of placement. Most of these requirements were based on an assumption that the time of 

storage in a dry cask system would be, at most, several decades prior to transport and 

repository disposal. The recent U.S. Department of Energy decision to review its 

policies on managing the back end of the nuclear fuel cycle, including alternatives for 

processing and disposal, has significantly lengthened the potential time requirement for 

storage. This lengthening of potential storage time prompted additional evaluations to 
6

explore extension of the licensing period for storage of nuclear fuel. The evaluations 

have focused on fuel discharged from the U.S. commercial fleet and subject to licensing 

by the U.S. Nuclear Regulatory Commission 

To provide a reasonable upper bound on the period for storage, a time of 300 years is 

considered. Therefore, the evaluations and recommendations are designed to extend 

the bases for safe storage from the present technical and licensing bases2

1) the personnel and infrastructure necessary to support the storage site will be 

maintained and available throughout the storage period;

to a time well 

beyond anticipated needs. Three basic assumptions to provide for safety and security of 

DCSS are made:

2) the SSCs of the DCSS important to safety must be capable of 

providing/performing their design safety functions throughout the storage period. 

This includes the assumption that the fuel cladding must be protected from 

degradation that leads to gross rupture throughout the storage period and also 

during post-storage, normal conditions of transportation and up to its acceptance 

for reprocessing or its placement in a repository for ultimate disposition; and

3) the potential storage environments and off-normal events outlined in NUREG1927 [6] represent the range of storage conditions encountered by the DCSS 

during the storage period with additional consideration of recent reports on 

general climate change conditions expected in the U.S..

These assumptions establish the foundation that provides the basis for this report to: 

1) define the potential for various age-related degradation processes, 2) establish the 

probable impact of potentially operative processes on functions important to safety, and 

3) recommend strategies to mitigate and/or repair the negative impacts of materials and 

system aging. Each of these assumptions is briefly discussed in Sections 1.3.1 through 

1.3.3

1.3.1 Management Strategy and Infrastructure

The long-term safety and security of the DCSS requires continual corporate control of 

the storage site. This control may be maintained using a variety of technologies, all of 

which will require facilities, trained staff and the resources to support both the personnel 

and infrastructure. A primary assumption of this report is that staff and facilities are 

available to provide the support necessary to conduct any operation required to 

safeguard the dry storage facility and assure nuclear safety throughout the storage 

period. The staff and/or facilities may be maintained at, or near, the storage site or may 

be readily mustered to the site to conduct the required operations. Staff functions 

throughout the storage period must include, but are not limited to:

1) implementation of security and controls to preclude unintentional and/or 

intentional access to the facility by unauthorized personnel and to preclude 

human theft or sabotage,

 

2 The present technical bases address materials and SCCs fuel storage to enable a regulatory basis for 

licensing for up to 60 years wet storage at a NPP followed by up to 60 years dry storage at an ISFSI.
7

2) implementation of a maintenance program and a monitoring program for the 

DCSS; and

3) capability to recover and re-pack a canister, repair degrading infrastructure, 

and/or relocate components and systems to nearby or remote facilities that have, 

or will have, an impact on the fuel or DCSS components important to safety.

1.3.2 Functions Important to Nuclear Safety

To support of 10 CFR Part 72, NRC’s NUREG-1536 [3] and NUREG-1567 [4] the 

functions important to safety are identified and a template for evaluation of the SSCs to 

meet the requirement of Part 72 is provided. Nuclear safety functions required by 10

CFR Part 72 and listed in NUREG-1536 include thermal performance, radiological 

confinement, subcriticality control, and retrievability. These safety functions are the 

same regardless of the differences among the storage or transport conditions and each 

of these functions relies on the continual integrity of the materials in the DCSS. 

Components of the fuel and container are especially important because of the potential 

for degradation processes to lead to fuel fragmentation, loss of container integrity, and 

other structural alterations that could directly impact confinement, sub-criticality control 

and/or retrievability. Each of these safety functions relies on the continual integrity of the 

materials in the DCSS. 

Additionally, deterioration of associated structures could indirectly impact safety by 

adversely impacting the monitoring program. Periodic monitoring is essential to assuring 

the long term nuclear safety because potential degradation mechanisms, including 

stress corrosion cracking and several of the hydrogen–induced degradation phenomena, 

are nucleation and growth processes, thus satisfactory behavior cannot be reliably 

extrapolated beyond the realm of experience. Reliable determinations of the behavior of 

materials and components during the time frames required for long term storage cannot 

be assured in the absence of a periodic monitoring program and such a program 

requires the continuing ability to access, and relocate if necessary, the dry cask, the fuel 

container and/or the fuel assembly. Therefore, maintenance of nuclear safety functions 

for dry cask storage system as required by 10 CFR Part 72 would mandate the following 

under extended storage:

1) determine materials degradation processes that could lead to corrosion, 

cracking, hydrogen induced degradation and other potentially operative age 

related processes;

2) provide verified, time-dependent models for any degradation process of potential

concern and continually provide the data necessary to assess the models used. 

If no suitable models are available, use expert engineering judgment to assess 

material behavior;

3) conduct the monitoring and analysis necessary to verify near-term extrapolations 

of the models, refine the models for longer term extrapolation and determine 

materials conditions; and 
8

4) establish provisions for satisfactory transfer, if necessary, of corporate 

responsibilities to other qualified entities even when the time frame of storage 

exceeds the corporate life of the responsible party.

1.3.3 Potential Storage Environments

The present environmental conditions to which the materials of the DCSS are 

anticipated to be exposed are essentially the local climate of storage [3]. Additional 

environments of storage may occur due to climate change over the extended storage 

timeframe of 300 years. The current predictions of global and local climate changes 

include only a low (several degrees Fahrenheit) increase in average global temperature, 

and local changes in total rainfall, in frequency of heavy precipitation; and/or average 

temperatures [13]. These conditions do not identify any factors which would exceed the 

current ISFSI design parameters.

The currently predicted climate changes [13] have minimal impact on the general 

degradation processes because of the range of climate conditions already considered in 

the current DCSSs. Nevertheless specific features of a global climate change (e.g., 

increased atmospheric CO2 concentration) may impact specific materials (e.g., 

reinforced concrete), and local climate considerations (e.g., local pollutants such as 

flyash) may impact specific materials (e.g., stainless steel canisters). Prolonged 

exposure to ambient conditions of a DCSS such as a marine environment may also lead 

to conditions (e.g., salt deposits on stainless steel canisters) detrimental to the materials 

of the DCSS. These specific conditions and their impact have been identified in the 

report.

Section 1 References

[1] Blue Ribbon Commission on America’s Nuclear Future, web home page. 

http://www.brc.gov/.

[2] NRC Regulations Title 10, Code of Federal Regulations, Part 72 - LICENSING 

REQUIREMENTS FOR THE INDEPENDENT STORAGE OF SPENT NUCLEAR 

FUEL, HIGH-LEVEL RADIOACTIVE WASTE, AND REACTOR-RELATED 

GREATER THAN CLASS C WASTE. http://www.nrc.gov/reading-rm/doccollections/cfr/part072/.

[3] NUREG-1536, Standard Review Plan for Dry Cask Storage Systems. 

http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1536/.

[4] NUREG-1567, Standard Review Plan for Spent Fuel Dry Storage Facilities. 

http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1567/.

[5] Technical Bases for Extended Dry Storage of Spent Nuclear Fuel, Electric Power 

Research Institute report number 1003416 Final Report, December 2002.

[6] NUREG-1927, Standard Review Plan for Renewal of Independent Spent Fuel 

Storage Installation Licenses and Dry Cask Storage System Certificates of 

Compliance. http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1927/.
9

[7] Electric Power Research Institute, report 1021048, “Industry Spent Fuel Storage 

Handbook,” July 2010.

[8] NRC Regulations Title 10, Code of Federal Regulations, Part 71 - PACKAGING 

AND TRANSPORTATION OF RADIOACTIVE MATERIAL. 

http://www.nrc.gov/reading-rm/doc-collections/cfr/part071/.

[9] NUREG-1617, Standard Review Plan for Transportation Packages for Spent 

Nuclear Fuel. http://www.nrc.gov/reading-rm/doccollections/nuregs/staff/sr1617/1617/sr1617.pdf.

[10] ASTM C-1562-10, Standard Guide for Evaluation of Materials used in Extended 

Service of Interim Spent Nuclear Fuel Dry Storage Systems, American Society for 

Testing and Materials, reapproved 2010.

[11] Generic Aging Lessons Learned (GALL) Report — Final Report (NUREG-1801, 

Revision 2, December 2010). http://www.nrc.gov/reading-rm/doccollections/nuregs/staff/sr1801/r2/index.html.

[12] NRC Regulations Title 10, Code of Federal Regulations, Part 54 -

REQUIREMENTS FOR RENEWAL OF OPERATING LICENSES FOR NUCLEAR 

POWER PLANTS. http://www.nrc.gov/reading-rm/doc-collections/cfr/part054/.

[13] Global Climate Change Impacts in the United States, Thomas R. Karl, Jerry M. 

Melillo, and Thomas C. Peterson, (eds.). Cambridge University Press, 2009.

11

2. AGING OF DRY STORAGE SSCS UNDER EST

Section 2 of this report evaluates the potential degradation phenomena and the impact 

of the degradation phenomena on the functions important to safety. Subsections 2.1, 

2.3, etc. in section 2 describe the component and materials of the DCSS, and

subsections 2.2, 2.4, etc. describe the corresponding potential degradation types and 

their impact on the safety functions provided by the component. Subsections 2.2, 2.4, 

etc. also include identification of the data needs and recommendations to address the 

materials aging issues.

The SSCs considered in the EST degradation evaluation include the past and present 

fuel designs, emergent fuels (e.g., MOX), and the non-fuel SSCs of DCSS that have 

been listed in references [1] and [2]. The importance to safety of the components of the 

storage systems and transportation packages are described in reference [3].

The initial or starting condition directly affects the type and extent of the degradation that 

might occur during long term dry storage. The materials and fabrication assumed in this 

report are listed as “reference conditions” for the respective SSC and are assumed to 

comply with applicable codes and standards for design and construction as directed in 

the SRPs for storage [4] and transportation [5]. Only the fuel is assumed to be in an 

initially “degraded” condition by virtue of its burn-up in the reactor.

Section 2 References

[1] ASTM C-1562-10, Standard Guide for Evaluation of Materials used in Extended 

Service of Interim Spent Nuclear Fuel Dry Storage Systems, American Society for 

Testing and Materials, reapproved 2010.

[2] Electric Power Research Institute, report 1021048, “Industry Spent Fuel Storage 

Handbook,” July 2010.

[3] NUREG/CR-6407, INEL-95/0551, Classification of Transportation Packaging and 

Dry Spent Fuel Storage System Components According to Importance to Safety, 

J.W. McConnell, Jr., A.L. Ayers, Jr., M.J. Tyacke, February 1996.

[4] NUREG-1536, Standard Review Plan for Dry Cask Storage Systems. 

http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1536/.

[5] NUREG-1617, Standard Review Plan for Transportation Packages for Spent 

Nuclear Fuel. http://www.nrc.gov/reading-rm/doccollections/nuregs/staff/sr1617/1617/sr1617.pdf.

2.1 Fuel Description

The fuel is considered as the entire assembly and includes the fuel pellets, fuel cladding, 

grid spacers, end pieces, and associated hardware. This “fuel” component is handled as 

an entire assembly in fuel movement and is the system that must be retrievable after 

EST. Figure 2 and Figure 3, respectively, show versions of PWR and BWR fuel 

assemblies. While some PWR assemblies have both top and bottom plenums/springs, 

the majority of the PWR assemblies in storage have only a top plenum and fuel spring.
12

Figure 2 PWR Fuel Assembly.

Figure 3 BWR Fuel Assembly [http://www.nucleartourist.com/images/bwrfuel1.jpg].
13

The non-cladding materials for components of the PWR and BWR fuel assemblies are 

summarized below. It is important to note that the non-cladding materials used will 

depend on the material used for the fuel cladding.

PWR Fuel Non-Cladding Materials: (Grid Spacers, Thimble Guide Tubes)

Fuel with Stainless Steel Cladding: Inconel grids and SS thimbles.

Fuel with Standard Zircaloy-4 Cladding: Inconel and Zircaloy-4 grids and Zircaloy-4 

thimbles.

Fuel with Improved Zircaloy-4 Cladding: Inconel and Zircaloy-4 grids and Zircaloy-4 

thimbles.

Fuel with ZIRLO Cladding: Some Zricaloy-4 grids and thimble, Mostly ZIRLO grids and 

thimbles and some Inconel grids.

Fuel with M5 Cladding: Some M5 thimbles and grids and some Zircaloy-4 thimbles and 

grids.

The assembly end pieces/nozzles are stainless steel and Inconel. 

BWR Fuel Non-Cladding Materials: (Grid Spacers, Channels)

Fuel with Standard Zircaloy-2 Cladding: Zircaloy channels.

Fuel with Improved Zircaloy-2 Cladding: Zircaloy channels.

The fuel pellets are currently uranium dioxide. However, mixed uranium-plutonium oxide 

fuel is planned for use in the existing fleet of LWRs.

The fuel material component that is most susceptible to degradation that could lead to a 

loss of retrievability is the fuel cladding. Although the other components of the fuel 

experience radiation damage and corrosion, and stress corrosion cracking of guide tube 

welds has created conditions that required repair to insure retrievability, the fuel cladding 

exhibits the greatest susceptibility to age related damage in terms of corrosion and 

hydride embrittlement. This higher susceptibility is due to the hydrogen uptake 

experienced during reactor service and the higher temperatures encountered during 

storage.

The wide variety of fuel cladding alloys used by the U.S. commercial fleet is significant 

because virtually every form of age-related degradation is sensitive to the composition of 

the alloy. Alterations in alloy chemistry designed to improve the in-reactor corrosion 

behavior, for example, may also alter the creep resistance, the susceptibility to hydrogen 

induced degradation and the mechanical properties of the alloy. Thermal/mechanical 

processing conditions, irradiation level, fuel drying conditions and other fabrication and 

service variables will also impact the behavior of the cladding during storage. Therefore 

the response of any fuel during EST will depend on the chemistry, thermal/mechanical 

and irradiation history and other material variables. Data and models to address, 

comprehensively across the variety of cladding alloys, the overall impact of these

variables on cladding behavior do not exist. Therefore, the age related degradation 
14

behaviors evaluated in this report do not describe the response of any specific cladding 

but use models and expert judgment to evaluate the generalized behavior of, for 

example, zirconium-based cladding alloys. Because of the very small inventory of 

stainless steel clad fuels, and their lack of susceptibility to aging mechanisms in dry

storage relative to zircaloy clad fuels, these fuels are not explicitly considered in this 

report.

Table 1 shows the composition of cladding alloys used at present in the commercial fleet 

in the U.S. Sections 2.1.1 and 2.1.2 briefly describe these alloys and selected early 

cladding alloys and section 2.1.3 addresses potential alloys for future applications.

Table 1 Nominal Compositions of Alloys Used as Nuclear Fuel Cladding

Table 1

Alloy Sn (wt%) Fe (wt%) Cr (wt%) Nb (wt%) Ni (wt%) O (wt%) C (wt%) Si (wt%) Zr (wt%)

Zr- 2 (BWR) 1.5 0.12 0.1 0.05 0.13 Balance

Zr- 2 (BWR) Improved 1.3 0.17 0.1 0.06 0.13 Balance

Zr -4 (PWR) High Tin 1.55 0.22 0.12 0.12 0.015 0.01 Balance

Zr -4 (PWR) Improved 1.3 0.22 0.12 0.12 0.012 0.01 Balance

M5 ® 0.04 1 0.14 Balance

ZIRLOTM 1 0.1 1 0.12 Balance

Optimized ZIRLOTM 0.7 0.1 1 0.12 Balance

Zircaloy 2 Specification 

ASTM B811 1.2 - 1.7 0.07 - 0.20 0.05 - 0.15 0.03 - 0.08 0.09 - 0.16

0.027 

max

0.012 

max Balance

Ziracloy 4 Specification 

ASTM B811 1.2 - 1.7 0.18 - 0.24 0.07 - 0.13 0.09 - 0.16

0.027 

max

0.012 

max Balance

Nominal Compositions of Alloys Used as Nuclear Fuel Cladding

2.1.1 Early Domestic Cladding Alloy Types

Stainless Steel

Initially and until the late-1970s, commercial fuel rods used austenitic stainless 

steel (primarily Type 304) as cladding. Some Type 348H stainless steel was also 

used [1], and one reactor system, Haddam Neck, used SS cladding through the 

1980s [1]. Issues with stress corrosion cracking in the stainless steel cladding 

resulted in a shift to Zircaloy-2 for BWR applications [2]. In PWR reactors the SS 

cladding performed well regarding fabricability and corrosion resistance during 

operation, but the relatively high thermal neutron absorption cross section 

reduced the neutron economy and stainless steel cladding was phased out and 

replaced by Zircaloy-4. Therefore, only a small fraction of the commercial spent 

fuel assemblies are SS clad.

The stainless steel clad fuel may have a higher margin against degradation 

during long term storage due to lower burnup and fuel duties, and the absence of 

hydrogen-induced degradation or hydride-related issues. Furthermore, in a dry 

storage environment with inert gas cover and negligible moisture, stress 

corrosion cracking of the SS clad fuel in EST is not likely.
15

Zircaloy 2 

Zircaloy 2 (Zr2) was chosen to replace SS in BWR cladding applications. The Zr 

2 alloy has about 98% zirconium with alloy additions of:

o Iron, chrome, and nickel to cause second phase precipitates for 

mechanical strength and improved corrosion resistance.

o Tin for solid solution strengthening and improved corrosion resistance, 

particularly to offset the adverse corrosion effects related to nitrogen 

contamination during production and abnormal coolant chemistry during 

operation.

o Oxygen is also typically added at about a 1200 ppm level for mechanical 

strength. During reactor operation corrosion of the exterior of the 

cladding by reaction with the coolant and the interior of the cladding 

reacts with the fuel to increase the oxygen level at the clad surfaces. At 

the operating temperatures the diffusion of the oxygen into the clad is 

relatively slow and most of the oxygen absorbed by the cladding is 

confined to surface oxide films. Oxygen diffusion is even slower at 

temperatures below the reactor operation temperature and little or no 

oxygen relocation is anticipated during long term dry storage.

Zircaloy 2 is generically categorized as grade R60802 in ASTM B811 and B353 

and is used extensively for BWR cladding. Prior to 1991 the applicable ASTM 

specification was B353. This specification was replaced by ASTM B811. 

Because of the operating conditions the typical uniform oxide layer on BWR fuel 

is less than that on PWR fuel; however, nodular corrosion of the Zircaloy 2 in a 

BWR environment could cause local thin spots that can impact cladding behavior 

during EST.

Zircaloy 4 (high tin)

Zircaloy 4 is defined by ASTM B811 grade R60804, although some vendors 

produce Zircaloy 4 cladding per their own specifications which provide for more 

specific controls within the variations allowed by the ASTM specification. Prior to 

1991, the applicable ASTM specification for Zircaloy 4 was B353.

Zircaloy 4 is similar to Zircaloy 2 except Zircaloy 4 has the nickel removed to 

reduce hydrogen absorption from corrosion reactions that occur in the PWR 

environment. To offset the strength reduction by nickel removal more iron was 

added to the Zircaloy 4 [2].

The Zircaloy 4 specification has a tin range of 1.25 to 1.7%. The initial versions 

of Zircaloy 4 were produced with tin levels at 1.55% near the upper tin limit. 

Carbon is an impurity in zirconium alloys which has positive effects on 

microstructure control but at high levels has negative effects on corrosion 

resistance and irradiation-induced growth. Because of the negative effects the 

maximum carbon level is limited to 270 ppm. However, in the late 1980s, 

improved versions of Zircaloy 4 were introduced that provided tighter controls on 

tin and carbon.
16

The high tin Zircaloy 4 cladding has a lower corrosion resistance than the current 

low tin version; thus, even though the burn up levels and duties for the early fuels 

are below the current levels. The lower corrosion resistance could lead to a high 

hydrogen concentration in the cladding. The clad oxide thicknesses range up to 

and in some cases exceed the guidelines of 100 microns of oxide. Therefore, 

because of the higher oxidation levels and the associated hydrogen absorption 

by the cladding, early fuels may be clad with alloys having higher than 

anticipated hydrogen contents.

Some of the Zircaloy 4 clad fuel has experienced significant oxide spallation. 

Although the spallation does not typically result in a clad breach, the spalling 

oxide presents a particulate source term and can promote localized hydriding 

that may impact cladding behavior during storage and transportation.

2.1.2 Current Domestic Cladding Alloy Types

There are primarily four cladding alloy types currently used in large quantities and in 

multiple fuel regions of commercial reactors. These alloys are lined Zircaloy 2, improved 

Zircaloy 4, M5®, and ZIRLOTM. Optimized ZIRLOTM is being introduced for very high duty 

applications and is present as multiple regions in a few reactors. 

The typical alloy element levels for these alloys are shown in Table 1, and the alloys are 

discussed in the following paragraphs. The cladding alloy of the fuel assembly is 

important to long-term storage considerations because the tendency for zirconium based 

alloys to experience age-related degradation depends on alloys content, fabrication 

history, crystallographic texture, thermal history, irradiation level and other 

materials/environmental variables. The impact of the cladding alloy and the mechanical 

history of the alloy on material behavior during long-term storage must be considered

before the anticipated long term behavior of any specific fuel can be estimated. The 

degradation processes such as creep, hydrogen-induced degradation, stress corrosion 

cracking and corrosion resistance are all dependent on the alloy type and processing 

history but are discussed in a general way in Section 2.2 of this report.

Zircaloy 2 – Lined

The bulk of the BWR fuels are now clad with lined Zircaloy 2. The Zircaloy 2 clad 

is similar to the earlier versions but has iron and nickel adjusted towards the 

upper limits and tin adjusted towards the lower limit of the ASTM range [3]. 

The ID liner which is “softer” than the base clad is pure zirconium or zirconium 

with small amounts of alloy additions. The liner is designed to accommodate the 

strains resulting from pellet expansions that occur when the reactor core power 

level changes. The lining also aides in mitigating stress corrosion assisted 

cracking from the clad ID surface; a condition that can lead to PCI (pellet clad 

interaction) failures.
17

Improved Zircaloy 4 – Low Tin

The increase in fuel duty and burn up required that the cladding have more 

margins to the corrosion and hydrogen design limits. For PWRs applications 

Zircaloy 4 was improved with a reduction in the tin level and improved controls 

over impurity/alloy elements such as carbon and nitrogen optimized through an 

annealing process to produce a consistent and desirable microstructure, referred 

to as A-Time control. The improved versions of Zircaloy 4 were introduced by 

fuel vendors in the late 1980’s. The improved corrosion resistance of this 

material supported operations to higher burn-up. The higher duties have caused 

some of the fuel cladding to have oxide and hydrogen levels near the design 

limits [4, 5, 6] thus clearly impacting the tendencies for hydrogen-induced 

degradation in this alloy.

M5® Alloy

AREVA introduced the alloy M5® for fuel applications in the 1990’s and 16 U.S.

commercial reactors have used or are using M5®clad fuel [7]. This alloy is now 

considered the AREVA PWR reference alloy. M5® is an alloy from the Zr – 

1% Nb family and has no additions of tin and chrome like the Zircaloys and only 

minor amounts of iron (400 ppm) and oxygen (1400 ppm). Sulfur at about the 

25 ppm level is also added to improve the creep strength. Performance data 

indicates that in reactor operation the clad oxidation is significantly reduced 

compared to Zircaloy 4 for the same burn up conditions. The typical M5® fuel 

would be expected to have less than 40 microns peak oxide at a discharge burn 

up of 62 GWD / MTU. At this same discharge burn up the hydrogen levels are 

also relatively low, and expected to be less than 200 ppm. M5® is also reported 

to have higher residual irradiated ductility compared to Zircaloy 4.

ZIRLOTM

ZIRLO is a Zr -1% Nb - 1% Sn - 0.1% Fe alloy with the typical addition of 

0.12% O [8]. Westinghouse introduced ZIRLO in early 1990’s as a replacement 

for Improved Zircaloy 4 for higher duty operations and it has become the 

reference alloy for Westinghouse fuel cladding with most if not all of the 

Westinghouse fueled domestic reactors using ZIRLO cladding. ZIRLO has 

enhanced corrosion resistance and lower hydrogen levels compared to Zircaloy 4 

for equivalent burn ups and fuel duties. With the tin and iron additions ZIRLO is 

a robust alloy that is resistant to coolant chemistry variation effects. ZIRLO is 

also resistant to oxide spalling which is observed in some high burn up Zircaloy 4 

fuel.

Optimized ZIRLOTM

Optimized ZIRLO was developed to provide added performance margin for high 

duty fuel applications [9]. Optimized ZIRLO has the basic characteristics of 

standard ZIRLO but has a tin level reduced to a range of 0.6% to 0.8% while 

maintaining the niobium at 1% and iron at 0.1% levels. The reduced tin improves 

the corrosion resistance compared to standard ZIRLO and the peak oxide 

thicknesses are less than 40 microns at discharge burn ups of 62 GWD/MTU. In 
18

2005 the NRC approved optimized ZIRLO for use in reactor fuel and as of 2010 

the alloy is in use as regions in a few commercial reactors.

2.1.3 Potential New Advanced Metal Alloy Cladding

The fuel vendors are evaluating other advanced alloys for future applications. For 

example an alloy called GNF-ZIRONTM which is a high iron modification to the improved 

Zircaloy 2 [10] is being tested and a new family of alloys in the AXIOM program [11] is 

being evaluated. AREVA has reported test results on a set of alloys which are a 

modified M5 with increased levels of tin and iron [12].

The diversity in fuel cladding alloy composition and the anticipated emergence of new 

cladding alloys couples with the alloy-to-alloy difference in cladding behavior to make the

estimation of a specific fuels response to EST difficult and make the data requirements 

for such an estimation extensive, if not prohibitive. Material behavior also varies with

numerous process and exposure variables and the data required to predict behavior of 

new and emergent alloys simply do not exist at this time. 

Section 2.1.1 – 2.1.3 References

[1] Reich, W.J., Moore, R.S., and Notz, K.J. Distribution of Characteristics of LWR 

Spent Fuel, ORNL/TM – 11670, January 1991.

[2] Schemel. J.H., Zirconium Alloy Fuel Clad Tubing Engineering Guide, Sandvik 

Special Metals Corporation, December 1989.

[3] Graham, R.A. and Eucken, C.M., Controlled Composition Zircaloy-2 Uniform 

Corrosion Resistance, ASTM STP 1132, 9th International Symposium on Zirconium

in the Nuclear Industry; 1991 pp. 279-303.

[4] Garde, A.M. et.al., Corrosion Behavior of Zircaloy-4 Cladding with Varying Tin 

Content in High-Temperature Pressurized Water Reactors; ASTM STP 1245, 10th

International Symposium on Zirconium in the Nuclear Industry; 1994 pp. 760-778.

[5] Mardon, J.P. et.al., Optimization of PWR Behavior of Stress-relieved Zircaloy-4 

Cladding Tubes by Improving the Manufacturing and Inspection Process; ASTM 

STP 1245, 10th International Symposium on Zirconium in the Nuclear Industry; 

1994 pp. 328-348.

[6] Sabol, G. et.al., In-Reactor Corrosion Performance of ZIRLOtm and Zircaloy 4; 

ASTM STP 1254 , 10th International Symposium on Zirconium in the Nuclear 

Industry, 1994 pp. 724-744.

[7] Mardon, J.P., Garner, G.L., and Hoffmann, P.B. M5® , a Breakthrough in Zr Alloy 

2010 LWR Fuel Performance Meeting Top Fuel; September 2010 Orlando 

Florida. 

[8] Sabol, G. et.al., Development of a Cladding Alloy for High Burnup ; ASTM STP 

1023 8th International Symposium on Zirconium in the Nuclear Industry, 1989 pp.

227-244.
19

[9] Wikmark, G. et.al., Cladding to Sustain Corrosion, Creep and Growth at High Burnups; Nuclear Engineering and technology, Vol. 41 No. 2 March 2009-Special Issue 

on the Water Reactor Fuel Performance Meeting 2008.

[10] Ishimoto, S. et.al., Improved Zr Alloys for High Burnup BWR Fuel, 2006 LWR Fuel 

Performance TopFuel; October 2006, Salamanca, Spain.

[11] Pan, G. et.al., Advanced Material for PWR Application: AXIOMtm Cladding 2010 

LWR Fuel Performance M Top Fuel; September 2010 Orlando Florida. 

[12] Chabretou, V. et.al., Ultra Low Tin Quaternary Alloys In-Pile Performance Impact of 

Tin Content on Corrosion and Mechanical Resistance; ASTM 16th International 

Symposium on Zirconium in the Nuclear Industry; May 2010 Chengdu, China.

2.1.4 Additional Fuel Design - Fort St. Vrain SNF

The following information describing Fort St Vrain was summarized from reference 1.

The Fort St Vrain reactor, a high-temperature, gas-cooled reactor (HTGR) was the only 

commercial thorium-fueled nuclear plant in the U.S. This reactor, developed from the 

AVR in Germany, was operated from 1976 until 1989. It was a high-temperature 

(700°C), graphite-moderated, helium-cooled reactor with a Th/HEU fuel designed to 

operate at 842 MWth (330 MWe). The fuel is in microspheres of thorium carbide and 

Th/U-235 carbide coated with silicon carbide and pyrolytic carbon to retain fission 

products. Fort St. Vrain spent fuel elements are graphite elements in hexagonal arrays 

with a 36.0 cm maximum cross section and 79.3 cm in length. The fuel achieved 

170,000 MWD/MT burn-up.

A total of 1464 HTGR spent fuel assemblies (15.4 MTHM) have been stored in the 

Modular Vault Dry Storage (MVDS) since 1991. The fuel is a highly enriched 

uranium/thorium carbide fuel that operated at a maximum heat output of 150 watts per 

assembly and an average heat output of 85 watts per assembly. The Fuel Storage 

Canister is a cylindrical carbon steel tube body with an exterior aluminum spray coating, 

double metal O ring seals, and a bolted lid. Nitrogen is used to backfill the Fuel Storage 

Canister (FSC).

The SiC coating of TRISO fuel (Fort St. Vrain) is the primary confinement of fission 

products and is expected to be stable for >1000 years in a repository environment [2]. 

This reference, coupled with the design of the fuel supports the conclusion that no fuelgenerated problems will result if the fuel is left in dry storage of 300+ years. Therefore, 

the behavior of this fuel is not discussed any further in this report. Monitoring of the 

carbon steel storage canisters and the double metal o-ring seals should be performed to 

detect the onset of any corrosion or leakage and to correct, as necessary.
20

Section 2.1.4 References

[1] IAEA TECDOC 1100, Survey of Wet and Dry spent Fuel Storage, July 1999.

[2] Nabielek, H. et al, Ceramic Coated Particles for Safe Operation in HTRS and in 

Long-Term Storage, Ceramics in Nuclear Applications, Volume 30, Issue 10, 

p. 196, Wiley, 2009. 

2.1.5 Additional Fuel Design - Mixed Oxide Fuel (MOX)

The following information describing Mixed Oxide Fuel was taken primarily from 

references [1, 2, and 3].

MOX fuel, consisting of about 7% plutonium mixed with depleted uranium, is equivalent 

to a uranium oxide fuel enriched to about 4.5% U-235, assuming that the plutonium has 

about two-thirds fissile isotopes. If weapons plutonium is used (>90% Pu-239), only 

about 5% plutonium is needed in the mix. Under the Plutonium Management and 

Disposition Agreement, Russia and the U.S. agreed in 2000 to dispose of 34 tonnes of 

weapons-grade plutonium deemed surplus to military requirements. A Mixed Oxide Fuel 

Fabrication Facility (MFFF) is being constructed at the Savannah River Site and will 

convert the U.S. plutonium to MOX fuel. Expected to begin operations in 2016, the 

MFFF is designed to turn 3.5 t/yr of weapons-grade plutonium into MOX fuel assemblies. 

This use of MOX fuels in U.S. reactors coupled with the trial irradiation of four MOX fuel 

lead test assemblies at the Catawba plant will place MOX fuel in the U.S. spent fuel 

inventory.

The post irradiation behavior of MOX differs from that of UO2 fuels including a difference 

in the reactivity changes with time.

The MOX reactivity curve is flatter than that for Uranium (Figure 4). The point of 

intersection depends on the relation between U-235 enrichment and Pu fissile content. 

A high preponderance of the Pu fissile content shifts the point of intersection to the left 

and leads to a considerably higher thermal load of the MOX-rods at burnups beyond 

20 to 30 GWd/MTHM. For dry storage the evolution of the decay heat and the neutron 

activity with burnup is important to assess cask temperatures and shielding capacities. 

A burnup increase of 25% from 55 to 69 GWd/MTHM increases the decay heat of both 

UO2 and MOX fuel assemblies (FAs) by roughly the same amount (Figure 5). However, 

the decay heat of a UO2-FA with around 2 kW at the beginning of storage after 5 years 

cooling time is less than a half of a MOX FA of comparable range of burnup. With 

increasing cooling time the decay heat decreases, however more and more slowly. 

Whereas for a UO2-FA it takes about 10 to 20 years to drop below 1 kW, it takes roughly 

100 to 200 years for a MOX FA. 

Additional information on weapons grade MOX (WG-MOX that contains Pu-239 only as 

initial fissile specie) characteristics for extended dry storage compared to UO2 fuel 

shows similar results to those below for the U-235/Pu-239 MOX [4].
21

Figure 4 Reactivity of MOX and UO2 fuel depending on burnup 

[reproduced from ref. 1].

Figure 5 Decay heat of UO2 and MOX FAs depending on burnup and cooling time 

[reproduced from ref. 1].
22

Figure 6 Neutron activity of UO2 and MOX FAs dependent on burnup and cooling time 

[reproduced from ref. 1].

The neutron activity exhibits somewhat different features than the decay heat (Figure 6). 

Between 10 and 100 years cooling time it drops by one order of magnitude and above 

100 years the reactivity remains relatively constant. The difference between MOX and 

UO2 FAs amounts to roughly a factor of ten. With regard to a storage extension, the 

graphs show that it takes about 100 years cooling time for both decay heat and neutron 

activity of MOX fuel to reach the UO2 level at the beginning of storage.

The large difference in the time-heat load profiles of MOX fuel compared to UO2 fuel 

mandates thermal management of the MOX fuels if it is to be loaded in a DCSS to avoid 

high temperatures that exceed design storage limits.

Section 2.1.5 References

[1] IAEA-TECDOC-1343, Spent Fuel Performance Assessment And Research, Final 

Report of a Coordinated Research Project on Spent Fuel Performance 

Assessment and Research (SPAR) 1997–2001, March 2003.

[2] Lovasic, Z. and Einziger, R.E., “International Atomic Energy Agency (IAEA) Activity 

on Technical Influence of High Burnup UOX and MOX Water Reactor Fuel on 

Spent Fuel Management,” paper 9065 in Waste Management 2009, proceedings of 

the International High-Level Radioactive Waste Management conference, Phoenix, 

AZ, March 1-5, 2009.

[3] IAEA Nuclear Energy Series, No. NF-T-3.8, Impact of High Burnup Uranium Oxide 

and Mixed Uranium-Plutonium Oxide Water Reactor Fuel on Spent Fuel 

Management, IAEA, 2011.

[4] Mendez-Torres, A., Vinson, D.W., and Branney, S., “A Comparison of the Isotopic 

Characteristics of WG-MOX and UOx fuel in Extended Dry Storage,” presented 
23

and to be published in the proceedings of the International High-Level Radioactive 

Waste Management conference, Albuquerque, NM, April 10-14, 2011.

2.1.6 Effects of Burnup, Pool Storage, and Drying on Fuel Condition

Low burn-up fuel has been defined for dry storage as being fuels at burn-up levels of 

45 GWd/MTU and lower. High burn-up fuel is above 45 GWd/MTU and nominally up to 

licensed limit of about 62.5 GWd/MTU. Changes occur in fuel with burnup or fuel duty 

as is shown pictorially in Figure 7. Both the fuel pellet and the cladding experience 

significant changes that can impact the performance of the fuel in DCSS and poststorage retrievability and transportation.

Some generic changes to the fuel rod during reactor operation include:

a) the fuel pellet develops fines at the rim of the pellet and undergoes cracking to 

enable fission gas release (FGR) to the cladding; and

b) the cladding undergoes radiation damage and embrittlement; corrosion and oxide 

scale formation/spallation; hydrogen ingress, hydriding, embrittlement; and 

creep.

In addition, fuel selected for placement in dry storage may have been in spent fuel 

storage pools for very long times and the physical characteristics of the fuel and fuel 

assemblies may have been altered by the storage and/or drying processes, and for this 

reason, the drying protocol used for the fuel is important. Pin holes, hairline cracks and 

other defects that would have released gaseous species but do not present a potential 

pathway for gross escape of fuel particles are considered minor, and fuel with such 

damage may be stored directly in the dry storage cask. Fuel with clad penetrations, 

however, is likely to contain water from the storage pool and needs to be carefully 

considered because of the potential for oxidation and swelling.

Fuels showing significant damage may also be stored in a cask after the damaged 

assembly and fuel debris are placed in a metallic fuel can that meets the fuel and system 

specific functions required for non-damaged fuels. Fuels placed in a dry storage cask, 

the non-damaged fuels, the fuels with insignificant damage and the degraded and 

canned fuels must all satisfy the same functions important to safety including, criticality, 

shielding, thermal and retrievability.

The changes to the pellet and cladding with burnup in the reactor, and the effects of pool 

storage and drying are summarized is this subsection. Information on the effects of 

burnup and fuel type on fuel characteristics is drawn from a recent IAEA publication [1] 

and other listed references in this subsection.
24

Figure 7 Fuel Cladding Changes with Burnup/Duty Cycle.

2.1.6.1 Fuel Duty Metric for In-Core Changes to Fuel 

Fuel duty is a general term that is used to describe the relative aggressiveness of 

operating conditions that tend to damage the fuel over time [2]. These operating 

conditions include heat flux, irradiation exposure, time-at-temperature, and coolant 

chemistry. While burn-up is one measure of the fuel condition it does not provide a 

distinctly defined set of characterizations. For example in determining the clad 

characteristics, two fuel rods may have the same burn-up but very different levels of 

surface oxide and clad hydrides depending on heat flux, coolant chemistry and other 

aspects of the fuel duty.

2.1.6.2 Uranium Dioxide Fuel Pellet Changes with Burnup

Changes to the typical UO2 fuel pellet occur during operation in a power reactor. For the 

fuel (pellets) there is both a microstructure change and a fission product build up with 

increased burn-up. These changes lead to pellet cracking that also occurs with 

increasing burn-up. The buildup of fission products is one of the major differences 

between low and high burn-up fuel. The high burn-up fuel will have more fission 

products, more residual radioactivity and a higher potential for energy release during a 

dry out anneal and a high temperature during dry storage.

Hydride Rim/Blister

Development

Radiation damage to

cladding → degradation of

The mechanical properties

H nuggets in cladding 

at the pellet-pellet 

interface

Thermal diffusion of

Hydride to colder surface

including clad/grid contact point

Large scale pellet

Cracking from

Normal operation

FGR in the gap → Rod internal pressure

Fission gas in the fuel matrix :

Inter- and intra-granular bubbles

(amount of gas, bubble size

distribution, bubble pressure)

Fuel-cladding chemical reactions; inner 

clad oxide film

Waterside

Corrosion and

Oxide Film

Buildup

H content in

the cladding

Development of 

rim region

Oxide

Spallation
25

The fuel pellet fractures because of the thermal and internal stresses from the fission 

product generation. In addition to cracking, the pellet microstructure changes during the 

high burn up to produce a rim at the outer surface. The rim has high porosity / low 

thermal conductivity thus leading to the development of higher thermal gradients. As the 

fuel burn up proceeds more fission products are generated and swelling occurs in the 

pellet matrix in response to the internal gas bubble formation and coalescence. The 

continued accumulation of fission gasses in the bubble leads to into high pressures and 

additional cracking in the matrix. 

The pellet matrix changes with burnup have been studied extensively [1, 3, 4, 5, and 6]. 

The following is a summary of the major characteristics of the changes in the fuel pellet 

that occur with burnup.

The conclusion of this present evaluation is that pellet aging does not represent a major 

concern for the fuel clad integrity (which is the primary fuel confinement) but that there is 

a need to have accurate data to support:

a) the estimation of the fission gas release and pellet swelling and their impacts on clad 

stress during storage and 

b) the internal heat generation and its impact on fuel and clad temperatures. 

Additionally, the potential release of fines from the rim region in a breached-clad fuel is a 

consideration that should be evaluated.

Pellet Rim Size – Source Term and Thermal Conductivity - The pellet rim size increases 

with burn up and is dependent on reactor operating conditions. Any increase in rim 

thickness increases the potential for the generation of fines and hence the source term 

for fuels with cladding penetrations. Additionally, the lower thermal conductivity of the rim 

will result in higher pellet centerline temperatures during storage but since the rim will 

also restrict heat transfer the clad temperature will not increase.

Pellet Rim Heat Conductivity - The pellet rim size is dependent on the burn up with 

higher rim thicknesses in the higher burn up fuel. The lower thermal conductivity of the 

rim will result in higher pellet centerline temperatures during storage but since it will 

restrict heat transfer to the clad it will not result in a higher clad temperature.

Pellet Rim Grain Size - The grain size in the rim is smaller than the remaining pellet. This 

smaller grain size needs to be included in storage evaluations of clad breach and fuel 

dispersion.

Fission gas release - During the fuel operation some of the fission products (FP) 

generated are elements that are normally in the gas phase. Most of these FPs are 

captured or contained within the fuel matrix and produce the pellet stress/swelling that is 

observed with burnup. A small percentage, nominally less than 10%, of the FP gas 

escapes the pellet and adds to the internal rod pressure. The internal rod pressure, 

acting in concert with welding and other residual stresses, produce the cladding stresses 

that are the drivers for age related degradation mechanisms such as creep, DHC and 

HE. Therefore, the contribution of fission gas release to cladding stresses needs to be 

well understood throughout the storage period. Measurements exist for fission gas 
26

release during normal operation, but there is insufficient data to assess fission gas 

release during long term storage.

Pellet Oxidation - During normal storage conditions the pellet is encapsulated in the 

cladding and is not exposed to oxygen in the environment. In some off-normal conditions 

such as failed fuel cladding prior to storage or fuel clad breach during storage that 

protective barrier is lost. When fuel is exposed to air/oxygen and heat there is a 

thermodynamic driving force for oxidation to occur and convert the UO2 to U3O8

(including some intermediate oxidation states). With the conversion to U3O8 the fuel will 

tend to become a powder and be more of a risk for dispersal from the fuel rod. It is 

reported [7, 8] that higher burn up fuel is more resistant to the oxidation so high burn up 

fuel should not be a greater concern or risk for pellet oxidation than the low burn up fuel. 

In addition with a He atmosphere provided in the container during planned storage 

conditions, even with a cladding breach, pellet oxidation would be precluded. If air or 

significant levels of humidity are in the container than the risk for pellet oxidation is 

increased.

Fission Product Generation - The chemical composition of the fuel changes with uranium 

burn up because a number of elements or fission products are formed from fission 

and/or neutron absorption. Some of these fission products are radioactive and their 

presence increases the activity of the fuel. After removal from the reactor, fission 

product decay causes the activity (and temperature) of the fuel to decrease with time. 

However, the heat produced by fission product decay necessitates fuel cooling during 

storage and handling. The heat produced is generally sufficient to necessitate pool 

storage and water cooling immediately after removal from the reactor. Dry storage 

systems can only be used after the spent fuel has cooled for a period of time which 

depends on the fuel and fuel service conditions. The higher burn up fuels will have 

higher quantities of fission products and thus will have higher heat loads than low burn 

up fuels and will generally have higher temperature profiles throughout the storage 

period.

Source Terms for Release - The source terms of HBU fuel are expected to be 

significantly different than the source terms for low burn up fuels, especially for failed 

rods. Reference [9] provides a comprehensive description of the features of the pellet 

rim, fuel grain size, fission gas release (FGR) to the rod plenum, CRUD thickness, 

cladding oxide thickness, and other fuel parameters on the source terms for HBU fuel for 

cask evaluation.

2.1.6.3 The Differences of Low and High Burn-up Cladding 

Low burn-up fuel has been defined for dry storage as being at burn-up levels of 

45 GWd/MTU and lower. High burn-up fuel is above 45 GWd/MTU and nominally up to 

licensed limit of about 62.5 GWd/MTU.

The following is a discussion comparing major characteristics of the cladding for low and 

high burn up fuel and the general conclusion is that there is not a clear property 

characteristic differentiation between low and high burn up fuel for long term storage 

evaluation. This is not the intuitive conclusion but it is supported by the fact that over 

time as the burn ups have increased, so have the fuel rod designs and the margins from 

material improvements increased. From a minor difference comparison, the higher burn 
27

up fuel will nominally have higher internal pressures and more hydride/oxide than the 

equivalent fuel design with a low burn up.

Oxide Thickness: In general it is expected that low burn up fuel will have slightly lower 

oxide thickness than the high burn up fuel but there is not a clear distinction. Even in a 

single-clad alloy system such as for low-tin Zircaloy-4 as shown in Figure 2-2 of 

reference 10, there is a high variability (tolerance interval from 20 to 100 μm at 

45 GWd/MT) in oxide thickness at a given fuel burn up.

The continued improvements in clad performance have resulted in the newer or 

improved clad versions that are more resistant to corrosion such that the newer versions 

may have less oxide and less hydride at 62 GWd/MTU burn-up than the older fuel rods 

had a 45 GWd/MTU. There is a driving force to extract as much energy from the fuel as 

can safely be accomplished. A general limit for both old and more recent fuel is a 

maximum OD oxide thickness of around 100 μm. The early cladding would reach that 

limit with burn ups near 45 GWd/MTU while the improved cladding can exceed 

62 GWd/MTU before that oxide thickness is reached.

Hydrogen levels: Hydrogen pick up by the cladding during irradiation occurs because 

the cladding alloy interacts with the water to form zirconium oxide and hydrogen. Some 

of the hydrogen formed is absorbed by the cladding and the amount absorbed varies 

with clad alloy. The amount of hydrogen absorbed in any given alloy generally increases 

with increasing operating temperatures and with the amount of oxide formed. Although 

the higher burn-up fuels typically have experienced a high duty and trend to see higher 

hydrogen than low burn up fuels, claddings with similar oxide thicknesses generally have 

similar levels of hydrogen, regardless of the fuel burn up. Some of the advanced alloys 

have been developed to minimize hydrogen absorption and thus, even at high burn ups, 

have a lower hydrogen pickup ratio (hydrogen absorbed into clad versus total hydrogen 

produced by clad oxidation) and lower hydrogen levels than the older low burn-up fuel 

[11].

Figure 8 below shows the hydrogen pickup corresponding to the oxide thickness for the 

low-tin Zircaloy-4 cladding shown in Figure 2-2 of reference 10. The oxide film growth 

and corresponding hydrogen pick up are not linear with burnup for this cladding system, 

but rather appear to show a trend that strongly suggests a change in mechanism for the 

ingress of hydrogen. 

A “transition knee” in hydrogen pick-up with burnup at approximately 45 GWd/MTU is 

suggested using an empirical bi-linear model for hydrogen pickup. It is assumed that a 

microstructure change or set of changes significant to enhance clad-water interaction

and hydrogen pickup may occur with burnup, although no formal model or detailed 

mechanistic explanation of this effect has been reported to the authors’ knowledge.
28

Figure 8 Maximum Wall Thickness Average (MWTA) Hydrogen Content in Low-tin

Zircaloy-4 Cladding [Figure reproduced from ref. 10]. A “transition knee” in 

hydrogen pick-up with burnup at approximately 45 GWd/MTU is suggested.

Hydride Microstructure in Cladding – The hydrogen in the cladding is partitioned as 

hydrogen in solid solution in the cladding, and, when local concentrations exceed the 

solubility limit, hydrides form. The formation of hydrides leads to a “hydride 

microstructure” in the cladding. 

A brief description of the cladding microstructure in terms of hydride formation and 

cracking that is representative of typical hydride structures generated during reactor 

operation and cool-down of zircaloy cladding follows. However, these microstructures 

are not expected to reflect the post-dry out microstructures that develop after the 

hydrogen has been re-dissolved and re-precipitated.

The, as irradiated, hydride microstructures that develop in the fuel clad have been 

grouped into three major types:

• hydride platelets;

• hydride rim; and

• hydride blisters.

Hydride platelets: Hydride platelets are the ZrHx phase that precipitates as individual or 

linked assemblages across the clad thickness. These platelets have preferred 

orientations depending on the stresses, thermal gradients, grain structure, 

crystallographic texture and the cooling rates. During reactor operation the bulk of the 

clad is at a high temperature and it is assumed (at this point in our 

understanding/evaluation) that in most cases platelets do not exist and that the hydrogen 
29

is in solution for most of the clad wall. However, at the clad OD where the temperature 

is cooler and the hydrogen solubility lower, a hydride rim may form. Upon cool down,

the temperature of the clad decreases and the platelets precipitate across the clad wall.

Figure 9 is an optical micrograph showing the Zircaloy-4 cladding microstructure in the 

clad of a fuel that experienced high burn up.

Figure 9 A light micrograph showing a hydride layer and associated hydrides in the 

substrate beneath the layer in irradiated Zircaloy-4 cladding tube (average 

fuel burnup of 67 GWd/t and fast fluence of 1.3 x 1022 n/cm2

) (Courtesy of 

R. Daum, Argonne National Laboratory). [Figure reproduced from ref. 12].

Hydride rim: A hydride rim is a semi-continuous hydride structure formed at the clad OD 

surface. It forms on the clad OD due to the thermal gradient across the clad during 

reactor operation which lead to an increasing hydrogen content as the OD of the clad is 

approached. The rim forms when this local hydrogen concentration exceeds the 

solubility limit for hydrogen in zirconium. Ultimately the local hydrogen content 

approaches that of solid ZrH2 and is approximately 10,000 wppm. The rim is of relatively 

uniform thickness axially and circumferentially proportional to the local oxide thickness. 

It is thicker at some local areas such as at the pellet to pellet interface because the clad 

is cooler at that location. 

Figure 9 and Figure 10 are examples of a hydride rim condition. The sample in Figure 

10 was hydrided ex-reactor to simulate the in-reactor rim formation.
30

Figure 10 Example of a hydride rim introduced in an unirradiated clad sample [Figure 

is reproduced from ref. 13, Fig. 5 - Wall-through crack in sample with 

hydride rim failed in room temperature burst test with pressurization rate of 

1.9 GPa/s].

Hydride blister: A hydride blister is a local area of high hydride concentration. It differs 

from the rim in that it is a local area of significant radial depth. It can be on the OD 

surface if there is some local corrosion acceleration. It also is associated with spalled 

oxide which changes the thermal gradient across the fuel and causes a local area of the 

clad to have a cooler temperature. 

ID blisters are also observed and are associated with internal hydriding either from 

hydrogenous materials present in the rod or from coolant that has entered the rod due to 

a cladding breach. Hydride blisters introduce stresses within the matrix due to the 

significant volume mismatch with the Zircaloy cladding. These stresses interact with the 

other cladding stresses and promote additional hydride formation along the resulting 

stress fields. Hydride alignment along the stress fields is manifested as hydride fingers 

emitting from blisters as shown in Figure 11, Figure 12, and Figure 13. Note that the 

radial alignment of the hydrides associated with a blister on the OD of the cladding 

(Figure 11) and the circumferential alignment of the hydrides associated with a blister on 

the clad ID (Figure 13). The importance of stress during hydride precipitation can be 

seen in Figure 13 where the hydrides near the clad OD tend to be radial while those 

near the clad ID tend to be circumferential.

Hydride 

rim
31

Figure 11 Example of a hydride blister in an unirradiated sheet of zircaloy 

[Figure reproduced from ref. 14].

Figure 12 Example of a hydride blister in a fuel cladding 

[Figure reproduced from ref. 15].
32

Figure 13 Example of a hydride blister in a fuel cladding 

[Figure reproduced from ref. 16].

Clad Cracks Due to Pellet Chips -

Figure 14 and Figure 15 show radial cracks in an irradiated fuel cladding. These cracks 

are associated with the as-fabricated-rod missing pellet surface (pellet chip). The 

cracking is in the general category of Pellet Cladding Interaction (PCI) cracking and can 

be associated with non-typical loads and cyclic stressing in the cladding due to the lack 

of contact with the underlying fuel pellet. The lack of contact results from a missing chip 

in the fuel pellet and is the result of fabrication errors. It is suggested that the 

operational stresses (coolant pressure, thermal, etc.) produce cyclic stresses in the clad 

in the region where the clad is not supported by the pellet. Local hydriding and/or 

iodine-induced stress corrosion cracking may also be contributors to the cracking 

process. Missing pellet chips have been observed in multiple rods from different 

vendors. Figure 14 is an example of an ID-initiated clad wall crack that is associated 

with a missing pellet surface chip. Many early vintage rods will have a potential for the 

presence of such chips and an associated partial wall crack. Such cracks will 

concentrate the stresses in the clad wall and may serve as sites for hydrogen induced 

clad degradation during EST. 
33

Figure 14 Example of ID crack in fuel cladding 

[Figure reproduced from ref. 17].

Figure 15 Example of ID crack in fuel cladding 

[Figure reproduced from ref. 18].
34

Irradiation Hardening and Microstructure Changes

The neutron fluence involved in fuel burn up cause displacement of atoms in the clad 

matrix and produce damage or a material hardening. This irradiation damage increases

with as the fuel burn up increases thus the strength of the clad increases and the 

ductility decreases with burn up. However, this effect is not linear with the neutron 

fluence and a near saturation of the irradiation hardening effects is reached typically 

during the first year of fuel operation [19]. Thus, there is essentially no significant 

difference in irradiation related clad strength or ductility between low and high burn up 

fuel.

There are, however, changes in clad and fuel micro structure that occur during 

irradiation and become more apparent as the fluence increases. Precipitate structures 

may develop in the cladding and the dislocation substructure is altered. Additionally, the 

clad density will change with fluence and the magnitude of the change is related to burnup. The drying operations required to prepare the fuel for dry storage will then alter the 

irradiation induced changes in microstructure and cladding properties.

Internal rod pressure, pellet swelling and clad stress

The clad stress states and levels are from a combination of loadings. The stresses are 

generated by the combination of pellet/cladding interactions, internal gas pressures, the 

residual stresses resulting from fuel fabrication and handling and the hydrostatic 

pressure exerted by the cooling water. Storage temperature and fission gas release 

influence the rod internal pressure and reactor operating conditions influence the free 

volume inside the fuel rod. Typically, the higher burn up rods will have higher internal 

pressures due to the combined effects of clad creep down, pellet swelling and fission 

gas release. Internal gas release from burnable absorbers will occur early in the life of 

the fuel and should not be significantly different for high burn up versus low burn up fuel. 

Taken together, the above factors combine to maximize the cladding stresses when the 

rod temperatures are the highest and before any post irradiation creep has occurred. 

Although typical cladding stresses decease as the storage time increases, stresses are 

anticipated to remain the highest in the high burn up fuels thus making these fuels more 

susceptible to most operative age-related degradation processes. However, the ultimate 

susceptibility to degradation is the product of the combined effects of available hydrogen 

content and hydride characteristics, applied and residual stresses, cladding 

characteristics, temperature and temperature gradients, time and irradiation. 

2.1.6.4 Fuel Drying and Impact on Fuel Initial Condition

The placement of spent fuel in a dry storage requires removal from pool storage and 

drying before placement in the DCSS. Drying is necessary fuel to prevent corrosion 

induced degradation of the fuel and/or canister and radiolytic reactions with moisture that 

can cause additional canister pressurization. However, the heating of the fuel to the 

anticipated drying temperature may alter the cladding microstructure by: 1) partially 

annealing of the microstructural changes induced by irradiation and 2) causing hydride 

re-solution during heating and re-precipitation on cooling. 

A national consensus standard for drying spent nuclear fuel [20] summarizes the 

practices for removal of water from fuel systems but does not provide sufficient 

justification for dryness of water logged rods or the resultant reconfiguration of the 
35

hydride microstructures. It is important to consider the potential for hydrogen 

redistribution and hydride reorientation in the cladding during the drying process 

because of the potential for hydrogen embrittlement processes during extended storage. 

Zirconium is an "exothermic occluder" of hydrogen, which means that hydrogen is more 

stable in solution in the zirconium matrix than as hydrogen gas and therefore is not 

removed from the zirconium alloys in a drying process.

Two different solubilities for hydrogen in zirconium are often referred to in the literature; 

namely the maximum solubility, and the terminal solid solubility (TSS). The maximum 

solubility is the total or maximum amount of hydrogen that can be absorbed under the 

exposure conditions (temperature and hydrogen pressure) and may involve the 

formation of zirconium hydride. The TSS is the amount of hydrogen that can be in 

solution in the zirconium matrix without the precipitation of a hydride phase. Reference 

[21] provides a treatise on the hydriding of zirconium alloys. 

Figure 16 and Figure 17 below, reproduced from reference [21], show that hydrogen 

maximum solubility and TSS, respectively for zirconium cladding alloys. Reference [21] 

summarizes additional literature references showing dependency of TSS on alloy 

additions, fabrication, and irradiation. 

Using the data and data trend from Figure 16, it is seen that the hydrogen solubility is 

about 100 ppm at 300°C and 200 ppm at 400°C. A high burn up fuel with an assumed 

clad content of 500 ppm H, including that in the rim, could dissolve up to 200 ppm during 

a dry out anneal at 400°C. The hydrogen would then precipitate as a hydride as the fuel

cools. A time delay for the precipitation process would be expected due to reaction 

kinetics and the “final” hydride morphology may not develop until the fuel experiences 

EST. 
36

Figure 16 Hydrogen (H) solubility for zirconium alloys: () = Zr-2; () = Zr-4; and 

(+) = Zr2.5Nb [Figure reproduced from ref. 21].
37

Figure 17 TSS for Pure Zirconium [Figure reproduced from ref. 21].
38

2.1.6.5 Fuel Initial Condition - Fuel Drying Data Needs and Recommendation

The ASTM standard guide [20] for drying spent fuel describes drying considerations, and 

lists the standard protocol for drying commercial spent nuclear fuel. The standard 

provides sufficient details to evaluate remaining free water in systems that do not have 

major impediments to drying, but it does not provide a quantitative approach to achieve 

dry-out of water-logged rods. The water from incomplete drying could lead to additional 

fuel corrosion and/or radiolysis of the water and hydrogen buildup inside the canister. 

The radiolytically produced hydrogen could be gettered by the cladding and add to its

hydrogen content. The contribution of hydrogen from water radiolysis has be discounted 

[22]; however a parametric evaluation of residual water in a canister free volume and 

potential hydride microstructure produced during cool-down over EST should be 

considered.

Most importantly, the present standard guide does not address a thermal drying 

treatment to minimize the potential for hydride reorientation from the predominantly 

circumferentially-oriented hydrides at reactor discharge to radially-oriented hydrides after 

drying. Changes to the fuel initial condition during dry-out can occur and are dependent 

on the thermal cycles used for drying. These changes may include some recovery of 

radiation damage and creep, and may alter the hydride microstructure. 

1. Testing & Analysis Activity: The following topics should be addressed in a 

revision to the existing ASTM drying standard or in a new drying standard guide. 

These topics should include explicit methods or controls (e.g., maximum cladding 

temperatures and appropriate cool-down rates to avoid radial hydrides), to 

assure water removal and prevent significant hydrogen induced degradation and

should be applicable to low and high burn up fuel:

• Dryout of water-logged rods

• Hydride microstructure reconfiguration

• Radiolysis of water with equilibrium back reactions and including gettering of 

H2 by fuel components

• Annealing of radiation damage

• Creep during dryout

Section 2.1.6 References

[1] IAEA Nuclear Energy Series, No. NF-T-3.8, Impact of High Burnup Uranium Oxide 

and Mixed Uranium-Plutonium Oxide Water Reactor Fuel on Spent Fuel 

Management, IAEA, 2011.

[2] Kaiser, R.S., The Fuel Duty Index (FDI)–A New Performance Measure of Fuel Rod 

Cladding Performance, International Topical Meeting on Light Water Reactor Fuel 

Performance, Park City, UT. American Nuclear Society, April 2000.

[3] Rondinella, V.V.and Wiss,T., The high burn-up structure in nuclear fuel, Materials 

Today, December 2010, Vol. 13 number 12.
39

[4] Adamson, R. et. al., High Burnup Fuel Issues, ZIRAT-8 Special Topics report, 

December 2003, ANT International.

[5] Matzke, H.J.; On the rim effect in high burnup UO2 LWR fuels; Journal of Nuclear 

Materials, Vol. 189 (1992) pp. 141-148.

[6] Ha, Yeong-Keong, et. al., Local Burnup Characteristics of PWR Spent Nuclear 

Fuels Discharged from Yeonggwang-2 Nuclear Plant, Nuclear Engineering and 

Technology, Vol 42, No 1 February 2010 pp. 79-88.

[7] Rouyer, J. et. al., Modeling of the spent fuel oxidation; Toward the operational 

model. Journal of Nuclear Materials, Vol. 395 (2009) pp. 89–98.

[8] Hanson, B. D.; The Burnup Dependence of Light Water Reactor Spent Fuel 

Oxidation; PNNL report 11929, 1998.

[9] Einziger, R.E. and Beyer, C., Characteristics and Behavior of High-Burnup Fuel 

That May Affect the Source Terms for Cask Accidents, Nuclear Technology, Vol. 

159, August 2007, pp. 134-146.

[10] Electric Power Research Institute, EPRI Synthetics Report 1015048, “Spent Fuel 

Transportation Applications – Assessment of Cladding Performance”, December 

2007.

[11] Mardon, J, Garner, G., and Hoffmann P., M5® a Breakthrough in Zr Alloy; 2010 

LWR Fuel Performance Meeting/Top Fuel/WRFPM; Orlando, Florida, September 

2010.

[12] Pierron, O.N. et al., The influence of hydride blisters on the fracture of Zircaloy-4, 

Journal of Nuclear Materials, Vol. 322 (2003) pp. 21–35.

[13] Nagase, F. and Fuketa, T., Investigation of Hydride Rim Effect on Failure of 

Zircaloy-4 Cladding with Tube Burst Test Japan Atomic Energy Research Institute, 

Tokai-mura, Naka-gun, Ibaraki 319-1195, Journal of Nuclear Science and 

Technology, Vol. 42, No. 1, p. 58–65 (January 2005).

[14] Raynaud, P. A., Koss, D. A., Motta, A. T., and Chan, K. S., Crack Growth in the 

Through-Thickness Direction of Hydrided Thin-Wall Zircaloy Sheet ASTM 16th 

Zirconium Symposium 2010 - Chang Du China.

[15] Garzarolli F., Von Jan R. and, Stehle H., “The main causes of fuel element failure 

in water cooled power reactors”, Atomic Energy Review, Vol. 17, pp. 31-128, 1979.

[16] Edsinger, K., “A Review of Fuel Degradation in BWR”, Proc. Intern. Topical 

Meeting, “LWR Fuel Performance,” American Nuclear Society, Park City Utah, 

2000, pp. 193-202.

[17] Nuclear Fuel Performance Presentation Department of Nuclear Regulation Atomic 

Energy Council, R.O.C AEC/NRC Bilateral Technical Meeting May 31 to June 2, 

2005.
40

[18] Olander, D., Consequences of pellet-cladding interaction (PCI) in Nuclear fuels –

Present and future, Journal of Nuclear Materials 389 (2009) pp. 1–22.

[19] Pettersson K., Vesterlund, G., and Andersson, T, Effect of Irradiation on the 

Strength , Ductility , and Defect Sensitivity of Fully Recrystallized Zircaloy Tube; 

ASTM STP 681, Zirconium in the Nuclear Industry, 1979.

[20] ASTM C1553-08, Standard Guide for Drying Behavior of Spent Nuclear Fuel, 

American Society for Testing and Materials, approved 2008.

[21] Cox, B. and Rudling, P., “Hydriding Mechanisms and Impact on Fuel 

Performance,” published by Advanced Nuclear Technology, Sweden AB, 2000.

[22] Einziger, R.E., McKinnon M.A., and Machiels, A.J., “Extending Dry Storage of 

Spent LWR Fuel for Up to 100 Years,” ANL/CMT/CP-96494, Argonne National 

Laboratory, presented at the International Symposium on Storage of Spent Fuel 

from Power Reactors, 9-13 November 1998.

2.1.7 Reference Fuel Assembly

The fuel assemblies placed in dry storage will be quite variable in terms of their physical 

characteristics. Typical variables include the fuel cladding alloy and fabrication 

technique, the fuel duty cycle, time since removal from the reactor, the presence or 

absence of pinholes or hairline cracks in the cladding, drying times and temperatures 

and other processing and service variables. These and other fuel assembly variables 

will alter the response of the fuel and cladding to any given set of dry storage conditions. 

The anticipated behavior of any given fuel may be very difficult to accurately establish 

because of the lack of directly applicable fuel/cladding response data. For example, the 

magnitude of the fuel cladding changes illustrated in Figure 7 will depend on fuel duty 

cycle and cladding alloy as well as other fabrication and service variables. Fuel cladding 

alloys progressively evolved to address specific economic and/or service problems (see 

Sections 2.1.1 and 2.1.2). Therefore, a single fuel cladding alloy was selected, by 

engineering judgment, to represent the variety of alloys that may be placed in dry 

storage. However, the evaluation also recognized that the lack of data on the alloy-toalloy variation in behavior represents a significant gap in the data necessary for accurate 

estimations of the response of the cladding to the very long term dry storage 

environment.

The attributes of the single fuel cladding selected as a representative of a reference fuel 

assembly for consideration of degradation during EST are:

Cladding Material: Standard Zircaloy-4 or Standard Zircaloy-2 (random grain texture, 

susceptible to hydride reorientation during thermal cycles and during long term

exposure).

Oxide Thickness (ZrO2): 100 μm including some spalled oxide in local areas.

Hydrogen Content: 600 ppm incorporated in a locally hydrided microstructure that 

contains predominantly circumferential hydrides.
41

Cladding Stress (from internal gas at RT): 75 MPa (includes impact of wall loss due to 

corrosion wastage)

CRUD: 25 μm (low burnup); 100 μm (high burnup). Note that the impact of CRUD due 

to the predominant radioisotope, Co-60, (T1/2 of 5.5 years) would be expected to be 

greatly lessened for EST.

Fuel Condition: Fuel in fragments including fines in rim at the cladding/fuel interface. 

It is also noted that there are some incipient conditions from both the fabrication (small 

laps in the clad, pellet chips, weld contamination, etc.) and reactor operation (vibration 

wear, non-breaching clad cracks, hydride blisters, etc.) that may be present on the fuel.

The detailed descriptions of the physical and material conditions at the beginning of dry 

storage illustrate two important features:

1) plots of fuel characteristics, including rod internal pressure, oxide thickness and 

hydrogen content as a function of rod average burn-up show considerable data 

scatter (for example hydrogen content of 400 ppm is found in rods with 35 to 

60 GWd/MTU burn-up), and

2) internal pressure, oxide thickness and hydrogen content generally increase with 

increasing burn-up, however, there is no discontinuity in the increase that 

categorically separates high burn-up from low burn-up conditions.

The data scatter shown in the reference [1] coupled with the lack of a discontinuous, 

burn-up induced transition in material condition is illustrated by the representative fuel 

cladding which, according to the specified oxide thickness, could have experienced a 

burn-up between about 40 and 60 GWd/MTU. The anticipated average hydrogen 

content in a cladding at these burn-up levels could vary from a low of about 100 ppm to 

a high of about 600+ ppm while the rod internal pressure could vary from 2.5 to 7.5 MPa. 

These variations in initial conditions could lead to large variations in the response of the 

cladding to the dry storage environment.

A discussion of the impact on fuel radiological source terms with burn-up that impact 

confinement is provided in reference 2.

Section 2.1.7 References

[1] Electric Power Research Institute, EPRI Synthetics Report 1015048, “Spent Fuel 

Transportation Applications – Assessment of Cladding Performance”, December 

2007.

[2] Einziger, R.E. and Beyer, C., Characteristics and Behavior of High-Burnup Fuel 

that may Affect the Source Terms for Cask Accidents, Nuclear Technology, Vol. 

159, August 2007, pp. 134-146.
42

2.2 Fuel Degradation Phenomena

There are three basic type processes that control age-related fuel degradation of fuel 

materials and components:

1) processes that involve the diffusion of point defects and defect complexes 

including dislocation movement in the material or component,

2) processes that involve the diffusion of specie atoms, especially hydrogen, 

and,

3) processes that involve chemical interaction among atoms within the 

materials and/or within the storage environment.

These processes are time-dependent, thermally-activated phenomena that are markedly 

influenced by alloy chemistry, microstructure, thermal-mechanical history, 

crystallographic texture, irradiation and other variables that depend on component 

fabrication and use as well as service and storage conditions. For example, much of

sections 2.2.x discuss the effect of zirconium alloy chemistry on material strength, 

corrosion resistance, in-reactor behavior and other performance measures. Such alloyto-alloy difference in performance translates directly into differences in the response to 

age-related degradation processes as shown in those same sections.

In spite of such well-established differences in alloy performance, there is a tendency to 

use long-term data developed for Alloy “A” to represent the anticipated behavior of Alloy 

B. For example, reference [1] states that “in creep tests at 250-400°C of Zircaloy 

cladding irradiated up to burn-up of 64 GWd/MTU, no failures have been observed 

below 2% strain – therefore a conservative strain limit of one percent has been used...” 

Similar statements that lump zirconium alloys as a single material system are frequently 

made, and while true to a very limited extent, alloy-to-alloy variations and burn-up 

related differences in susceptibility to age-related degradation processes are expected.

The following subsections under section 2.2 discuss the fuel degradation processes for 

zirconium cladding alloys from published information.

General Fuel Data Needs for Fuel in EST; General Recommendation for 

Resolution: A wide variety of alloys are, and have been, used for fuel cladding in the 

U.S. commercial fleet and the response of these alloys to EST will depend on chemistry, 

thermal mechanical process history, irradiation history and other fabrication/service 

variables. The information required to address the response of a specific cladding to a 

specific age-related degradation process is limited and in many cases simply not 

available. It is recommended that information generated (e.g., fracture toughness of 

newer cladding alloys with a hydride microstructure) should be explicitly compared to 

available data and improvements in the basis for constitutive model(s) for alloy 

performance should be made. Models should then be developed that strive to 

encompass the spent fuel inventory and these models should be assessed to validate 

their extrapolation to EST conditions.
43

Fuel Condition Data Needs for Storage Retrievability and for Source Term for 

Transportation; General Recommendation for Resolution: An important 

consideration for the stored fuel is the potential release of particulates, especially in 

evaluation of minor breaches that may occur in storage and impact retrievability during 

EST. Additionally, failures that may occur during EST should be assessed and the 

impact of such failures on confinement in transportation should be determined. As

discussed, reference [2] provides a comprehensive description of the features of the 

pellet rim, fuel grain size, fission gas release (FGR) to the rod plenum, CRUD thickness, 

cladding oxide thickness, and other fuel parameters on the gas, volatile, fines, and 

CRUD source terms of HBU fuel for cask evaluation. Many data needs related to these 

source terms were identified and are listed in Table II of reference 2. This present 

evaluation concurs with these needs. These needs for EST would best be evaluated in 

a destructive examination of HBU fuel after an extended exposure to dry-storage 

conditions. This examination could be performed in conjunction with the cask 

demonstration project described in section 3 of this report.

Section 2.2 References

[1] Electric Power Research Institute, report 1003416, “Technical Bases for Extended 

Dry Storage of Spent Nuclear Fuel” December 2002.

[2] Einziger, R.E. and Beyer, C., Characteristics and Behavior of High-Burnup Fuel 

That May Affect the Source Terms for Cask Accidents, Nuclear Technology, Vol. 

159, August 2007, pp. 134-146.

2.2.1 Creep

Creep is time dependent plastic deformation that is generally attributed to stressdirected, diffusional processes. The activation energy for diffusion is dependent on the 

path for vacancy and atom migration. The activation energy for grain boundary diffusion 

is typically less than the activation energy for lattice diffusion and greater than that for 

diffusion along an external surface. The variations in the activation energies for diffusion 

and the relative volumes of the diffusion pathways that exist on a polycrystalline alloy 

such as the zirconium based cladding alloys frequently causes creep to be divided into 

several distinguishable but related processes.

Several of the potential creep mechanisms and the associated diffusion pathways are 

illustrated in Figure 18. Coble creep, for example, involves grain boundary diffusion 

while lattice diffusion controls the power law creep process. The strain that is associated 

with any creep process is the sum of the strains caused by the operative creep 

mechanisms. Models to estimate creep strains or strain rates typically involve the 

diffusivity of base metal atoms in the lattice, D, temperature, T, stress, σ, grain size, d, 

material constants such as the elastic modulus, E, the Burger’s vector of dislocations, b, 

and constants that are dependent on the operating creep mechanism. For example, the 

creep strain rate may be expressed as:

= A(DEb/kT)(σ/E)n

(b/d)p [1]

where A, n and p are constants. The constant n is termed the stress exponent and p is 

the inverse grain size exponent. More elaborate creep models are available but these 
44

models also suffer from the weaknesses described and should not be extrapolated into 

time frames that extend far beyond the data used to develop the model. That is, the use 

of such an equation to extrapolate short term data into times relevant to long term 

storage of spent fuels is tenuous at best especially when alloy-to-alloy variations in 

material behavior are considered. Small changes in alloy composition can cause 

significant changes in creep rates. Microstructure and heat treatment also have a 

significant impact on the creep rates of any given alloy as shown in Table 2.

Figure 18 Creep Mechanisms by atom or defect complex migration (left) and 

dislocation formation/moverment mechanisms (right).

σ

σ

MATTER MIGRATES DISLOCATIONS MOVE

1

2

3

4

N-H CREEP

LATTICE 

DIFFUSION

COBBLE CREEP

GB 

DIFFUSION

H-D CREEP

Dislocation 

density 

Independent of σ

But Peierls stress

dependent

POWER-LAW

CREEP

Dislocation 

density a function 

of σ

FLOW OF

MATTER

in BULK

FLOW OF

MATTER AT 

GBs

T

T

T

T

1, 2 and 3: Creep mechanisms at low stresses

4: Creep mechanism at high stresses
45

Table 2 Influences of Heat Treatment on Creep of Zr-2.5Nb Alloy 

at 545°C (818K) [1]

Specimen Heat Treatment Test 

Temperature

Stress (MPa) Creep Rate (s-1)

A 650°C/3hr, FC 545°C 1.5 8.5x10-9

B 650°C/3hr, FC 

+ Aged at 

550°C/10 days

545°C 1.5 2.0x10-7

C 745°C/75hr, FC 545°C 1.5 2.0x10-9

Table 3 lists predicted creep strains, from a creep model benchmarked to short-term 

pressurized tube testing, for time and temperature conditions that span EST. The creep 

strain results clearly indicate that most of the strain has occurred within a 50-year period 

corresponding to the high temperature conditions of fuel storage.

Table 3 Predicted Hoop Strains of Nb-Modified Zirconium Cladding Alloy Under

Extended Dry Storage Conditions [Table reproduced from ref. 2].

Table 3

Predicted hoop strains of Nb-Zry clad under dry-storage conditions

Time (year) T (

o

C) Coble (%) Dislocation (%) Total strain (%)

Time-hardening Strain-hardening

0.02 295.6 1.66 x 10-4 5.97 x 10-6 1.72 x 10-4 1.41 x 10-4

1.19 330.4 3.12 x 10-2 2.72 x 10-3 3.39 x 10-2 3.40 x 10-2

5.43 330.1 0.160 7.41 x 10-3 0.167 0.180

10.94 321.7 0.294 1.00 x 10-2 0.304 0.338

25.18 298.0 0.483 1.27 x 10-2 0.496 0.559

51.31 265.2` 0.598 1.39 x 10-2 0.612 0.696

101.80 222.0 0.638 1.41 x 10-2 0.652 0.750

200.50 193.5 0.650 1.41 x 10-2 0.664 0.767

250.90 186.8 0.652 1.41 x 10-2 0.666 0.770

540.80 168.7 0.658 1.41 x 10-2 0.672 0.777

769.30 159.6 0.660 1.41 x 10-2 0.674 0.779

993.70 153.5 0.661 1.41 x 10-2 0.675 0.780

The pressure loading of the pin-design fuel is the primary driving load for cladding creep. 

Creep can be significant at high temperatures and/or over long time periods at moderate 

temperatures [3]. 

The initial cladding stresses are dependent on burnup. The fission gas inventory in a 

fuel rod increases with burn-up, while the fuel thermal conductivity decreases and 

changes of fuel microstructure reduce the ability of the fuel to retain the fission gas. 

This, together with the rod void volume reduction caused by cladding creep-down (in 

reactor) and fuel swelling, increases the potential for initial high rod internal pressures 

with increasing burn-up [3].
46

Assessment of Creep Degradation under EST - Creep-induced degradation of the 

spent fuel under EST conditions is not anticipated because of the lower cladding 

stresses due to reduced internal pressure with time, and the limited ability of 

substitutional atom diffusion at the temperatures anticipated (< 200°C) during extended 

dry storage.

Furthermore, creep of the cladding in a dry storage system would be self-limiting as the 

driving force decreases with dimensional changes. For example assuming typical spent 

fuel rod conditions (ID – 0.83mm, void volume -15 cc) a 1% diametrical creep will 

produce about a 15% increase in rod internal volume and an equivalent 15% reduction 

in stress; thus, reducing the creep rate. The internal volume increase with creep and the 

internal pressure reduction with cooling work to reduce clad stress and hence creep rate.

Should creep rupture occur, the cladding breach is expected to be limited in extent, and 

would result in gas and volatile specie release, and some small fraction of fines from the 

rod due to the much reduced internal pressure driving force with time. Axial splitting 

along the rod length is a beyond unlikely occurrence.

Creep Degradation Data Needs for EST and General Recommendations for 

Resolution - The present temperature limit, applicable to all Zircaloy fuel cladding types 

for dry storage, is 400°C, and is based on limiting the creep strain total deformation to 

1%. This limit is selected to avoid creep rupture that could occur at 2% strain [4, 5]. 

Primarily fuel at low burnup was used for this determination [citations in 5]. HBU fuel 

would have higher internal gas pressure and higher cladding stresses at this 

temperature by virtue of additional gas release to rod with burnup [6]. 

Other cladding phenomena, primarily hydride reorientation (temperatures affect hydride 

solubilities and strains affect reorientations) should be considered in creep evaluation 

and modeling to enable a more accurate prediction of spent fuel performance in very 

long storage conditions. Specifically, 

1. Testing & Analysis Activity: Verification that the creep models, and failure strain 

limits presently assumed for all zirconium cladding alloys are applicable for new 

and alloys and for HBU conditions.

2. Testing & Analysis Activity: Consideration of hydrogen effects on creep rate 

models. Claddings in HBU fuel may contain more hydrogen than LBU fuel 

claddings.

3. Testing & Analysis Activity: Consideration of higher pressure loading from HBU 

fuel on creep throughout storage period.

4. Cask Demonstration Activity: As part of a cask demonstration, creep data should 

be collected on HBU fuel and on new cladding alloys.

These recommended activities are to improve predictive accuracy especially for new 

alloys and at HBU conditions. Temperature limits for dry storage under short-term 

(< 60 years at present) are sufficient to protect against creep, contingent upon DCSS 

thermal management in dry storage to enable credit of low (e.g.,< 400°C and falling) 

temperatures in the EST period.
47

The probable significance of creep to EST is low considering: 1) that creep rates at the 

start of EST after 120 years is expected to be low due to very low fuel temperatures 

(< 200°C); 2) creep is self-limiting in a rod under internal pressure; and 3) any fuel rod 

breach that would occur due to creep would be a minor breach with a small release 

within the confinement provided by the canister of the DCSS.

Section 2.2.1 References

[1] Rao, P. R., paper TRA-80 in METALLO – 2007, International Conference on 

Metals and Alloys: Past, Present & Future, IIT Kanpur, December 2007. 

http://www.iitk.ac.in/infocell/announce/metallo/talk/pramarao.pdf.

[2] Zhou, Y., Devarajan, B., Murty, K.L., Short-term rupture studies of Zircaloy-4 and 

Nb-modified Zircaloy-4 tubing using closed-end internal pressurization, Nuclear 

Engineering and Design 228 (2004) pp. 3–13.

[3] Very High Burn-up in Light Water Reactors, ISBN 92-64-02303-8, Nuclear Energy 

Agency, NEA 6224, 2006.

[4] Einziger, R.E., Brown, C.L., Hornseth, G.P., and Interrante, C.G., Data Needs for 

Storage and Transportation of High-Burnup Fuel, Radwaste Solutions March/April 

2005.

[5] Interim Staff Guidance - 11, Revision 3, U.S. NRC, Spent Fuel Project Office, 

November 17, 2003; http://www.nrc.gov/reading-rm/doc-collections/isg/isg-11R3.pdf.

[6] Manzel, R. and Walker, C.T., “High Burnup Fuel Microstructure and Its Effect on 

Fuel Rod Performance,” Proceeding of the ANS International Topical Meeting. 

Light Water Fuel Performance, Park City, Utah, April 2000, American Nuclear 

Society (2000).

2.2.2 Hydrogen-Induced Degradation Processes - Overview

Hydrogen induced degradation is a generic term used to designate a number of 

processes or mechanisms by which the load carrying capabilities of metals and alloys 

are reduced by the presence of hydrogen [1]. Often the hydrogen acts in combination 

with applied and/or residual stresses on the material to produce the degradation 

process. The hydrogen may be present in the external environment, absorbed in the 

metal, precipitated as a second phase or adsorbed on the metal surface. Hydrogen 

found in the cladding of spent fuels is generated by in-reactor corrosion of the fuel 

cladding and its ingress increases the hydrogen content in the cladding alloy. The 

hydrogen induced degradation mechanisms generally considered applicable to the 

zirconium based alloys in spent nuclear fuel assemblies involve the precipitation of 

hydride phases and include [2, 3]: 

1) Delayed Hydride Cracking (DHC),

2) Hydrogen Embrittlement (HE) as pertains to reduces tensile ductility, and

3) Hydrogen Embrittlement (HE) as pertains to hydrogen/hydride induced shifts in 

the ductile-to-brittle transformation temperature. 
48

Each of these degradation mechanisms may involve hydride reorientation and hydrogen 

migration which have occasionally been considered degradation mechanisms [4] but, in 

reality, are enablers of one of the three mechanisms listed above. 

This report demonstrates that each of the three degradation mechanisms listed may 

cause degradation of zirconium based fuel cladding and assembly components during 

extended storage and transportation and that such degradation has the potential to 

impact the safety functions of the DCSS. Because of this potential, a brief background 

on hydrogen-in-zirconium alloys is presented.

Background for Hydrogen Effects on Zirconium Alloys - The adverse effects of 

hydrogen on the mechanical behavior of zirconium based nuclear fuel cladding alloys 

are well established and are generally attributed to the transgranular and/or intergranular 

precipitation of hydrides [5]. 

Hydrogen solubility is temperature dependent and at temperatures below approximately 

100°C, hydrogen is virtually insoluble as illustrated in Figure 19. This lack of hydrogen 

solubility at 100°C or below contrasts the approximately 200 ppm hydrogen that is 

soluble at PWR operating temperatures and at the 400°C maximum fuel drying 

temperature (see Figure 16). The decrease in hydrogen solubility as the cladding cools 

during dry storage, coupled with the fact that hydrogen was absorbed by the fuel 

cladding during reactor service provides the opportunity for hydride precipitation during 

cooling. 

Hydrogen remains mobile in many metals, including zirconium alloys, at room 

temperature, thus hydride precipitation in zirconium alloys may occur throughout the 

extended storage period. The temperature dependence of hydrogen diffusivity in alpha 

phase zirconium is [5]

D = 2.17 x 10-3 exp(-8380/RT) cm2 sec-1 [2]

Evaluation of the above equation at T = 300 K (approximately room temperature) gives a 

value for diffusivity of 1.7 x 10-9 cm2 sec-1

. The distance that atoms can diffuse, x, during 

a time, t, can be broadly approximated [6] by 

x = (2Dt)1/2 cm [3]

Thus, even at room temperature, hydrogen in zirconium alloys may migrate, by random 

walk diffusion, a distance of over 0.05 mm per day. The driving forces for hydrogen 

migration include concentration, thermal and stress gradients, all of which will exist in 

spent fuel cladding. Hydrogen will diffuse down a concentration or temperature gradient 

but up a stress gradient [5]. Therefore, there is tendency for hydrogen to migrate up a 

stress gradient and concentrate in regions of lattice dilation on the cold side of the fuel 

cladding [7]. These tendencies can increase the hydrogen flux into regions of lattice 

dilation and toward colder areas. The influence of the various gradients on hydrogen 

flux, J, is given by [5]

J = -D(ΔC) + D(Q*

C/RT2) ΔT + exp(C2 Δs) [4]
49

where D is the diffusivity, Q* is the heat of transport (approximately 6 kcal mol-1), C is the 

local hydrogen concentration, ΔC is the concentration gradient, ΔT is the thermal 

gradient, Δis the stress gradient and C1 and C2 are constants. 

Hydrogen, in solid solution, is a screened proton, occupies interstitial sites in the metal 

lattice and has a positive partial molar volume. The lattice strains caused by hydrogen 

absorption interact with other lattice strain such as those produced by imperfections or 

by stresses acting on the metal [7]. For example, hydrogen will concentrate on the 

tensile side of an edge dislocation and in the hydrostatic stress field ahead of a sharp 

crack tip.

There is a tendency for the hydride platelets to develop a preferred orientation during the 

nucleation and/or growth processes [8, 9]. This preferred orientation depends on the 

fabrication history of the cladding alloy, post-fabrication thermal-mechanical processing, 

existing temperature differentials and stress fields, and other material/environmental 

variables. The typical preferred orientation results from a combination of texture effects, 

temperature differentials and the tendency of platelets to align parallel to a compressive 

stress and perpendicular to a tensile stress. 

In spent fuel that is discharged from the reactor, the preferred orientation is 

predominantly dictated by the temperature differential that was present across the clad 

wall during operation and favors a circumferential alignment of hydride platelets in many 

cladding alloys. However, as shown above, the mobility of hydrogen in zirconium alloys 

is sufficient to allow hydride reorientation during spent fuel storage. In fact, the few 

hours required to conduct a room temperature tensile test and prepare the sample for 

metallographic examination are sufficient for hydride reorientation as shown in Figure

20. Examination of the hydrides in Figure 20 (magnification 100X) suggests that 

hydrogen must migrate approximately 0.005 cm during the time required for the tensile 

test and metallographic preparation. Although this time was not measured for the test 

specimen shown, typically, at least 24 hours would lapse between the start of the tensile 

test and the completion of the metallographic examination. The time required for 

hydrogen to diffuse a distance x, by random walk processes, is x2

/2D or approximately 

20 hours thus reorientation could occur during the time required to test and prepare the 

sample. Additionally, hydrogen mobility during the tensile test could be enhanced by 

dislocation transport and hydrogen diffusion necessary to cause hydride reorientation is 

not a random walk process but is directed motion toward regions of lattice dilation. 

In any event, hydrogen mobility at room temperature is relatively high and this high 

mobility, coupled with the positive partial molar volume of hydrogen in solution, provide a 

tendency for hydrogen to accumulate in the cold side of a thermal gradient, and the low 

density of zirconium hydrides, provides the thermodynamic driving force to cause:

1) The development of a hydride rim on the cooler surface of the cladding;

2) Reorientation of existing hydrides perpendicular to a tensile stress and the 

precipitation of stress oriented hydrides as the cladding cools;

3) Hydrogen accumulation and hydride precipitation in regions of tensile stress 

such as those that exist at a stress concentration or at the tip of existing 

hydride precipitates, and
50

4) Localized hydride concentrations in regions of the cladding adjacent to pelletto-cladding interaction interfaces.

Figure 15a in reference 8 illustrates how applied stresses during cooling can cause 

hydride platelets to form aligned clusters and the room temperature stress strain curve 

for Zircaloy-2 with radially-clustered hydrides shown in reference 11 demonstrates how 

severe the reduction in load carrying capacity can be when hydride clusters align 

perpendicular to the stress axis. Data presented in Reference [9] are consistent with the 

observation that aligned hydrides create low strength, brittle paths within the cladding 

and therefore may lead to any or all of the three hydrogen-induced degradation 

mechanisms. Additionally, the present observations are clearly consistent with the 

conclusion that “spent fuel rods would fail by DHC in dry storage as stress risers are 

present inside the cladding on cooling to below 180°C” [10] and provide a technical basis 

for the conclusion of this report that there is a clear potential for hydrogen-induced 

degradation and failure of spent fuel cladding during extended storage and 

transportation. Further discussion of the potential cladding degradation mechanisms is 

given in the subsections below.

Figure 19 Zr-H Phase Diagram [Figure reproduced from ref. 3]
51

Figure 20 Hydride Reorientation in the Near Fracture Region of a Zircaloy-2 Tensile 

Sample Tested at Room Temperature. A) Before Test, B) After Test 

(Archival data from M.R. Louthan, Jr.).

Section 2.2.2 References

[1] Kim, C.D., “Hydrogen Damage Failures” Metals Handbook, Ninth Edition, Vol. 11, 

Failure Analysis and Prevention, pp. 245-251, 1986.

[2] ASTM C-1562-10, Standard Guide for Evaluation of Materials used in Extended 

Service of Interim Spent Nuclear Fuel Dry Storage Systems, American Society for 

Testing and Materials, reapproved 2010.

[3] Figure 55, Environmentally Induced Corrosion, p. 175, ASM Metals Handbook, Vol. 

13, Corrosion: Fundamentals, Testing and Protection, ASM International, Materials 

Park, OH (2003).

[4] Kessler, J., “Technical Bases for Extended Dry Storage of Spent Nuclear Fuel”, 

EPRI Report 1003416, December 2002.

[5] Cox, B. and Rudling, P., “Hydriding Mechanisms and Impact on fuel Performance”, 

Advanced Nuclear Technology AB, ANT, Sweden, 2000.

[6] www.unc.edu/courses/2006spring/envr/210/001/chapt9old/chapt9old.doc.

[7] Oriani, R.A., “A Brief Survey of Useful Information about Hydrogen in Metals”, 

International Symposium on Cold Fusion and Advanced Energy Sources, 

Belarusian State University, Minsk, Belarus, 1994.

[8] Louthan, M.R., Jr. and Marshall, R.P., “Control of Hydride Orientation in Zircaloy”, 

Journal of Nuclear Materials, Vol. 9, pp. 170-184, 1963.

[9] Hsu, H.H. and Tsay, L.W., “Effect of Hydride Orientation on Fracture Toughness of 

Zircaloy-4 Cladding”, Journal of Nuclear Materials, Vol. 408, pp. 67-72, 2010.

[10] Kim, Y.S., Hydride Reorientation and Delayed Hydride Cracking of Spent Fuel 

Rods in Dry Storage, Metallurgical and Materials Transaction A, Vol. 40 pp. 2867-

2875, 2009.
52

[11] Marshall, R.P. and Louthan, M.R., Jr. “Tensile Properties of Zircaloy with Oriented 

Hydrides”, Trans. ASM, Vol. 56, pp. 693-698, 1963.

2.2.3 Delayed Hydride Cracking (DHC) - Potential for DHC during Extended 

Storage

Failures of zirconium alloy cladding tubes at end cap welds by delayed hydride cracking 

(DHC) during long-term storage at room temperature were first reported by Simpson and 

Ells in 1974 [1]. Subsequently leakage from Zr-2.5Nb pressure tubes in CANDU plants 

and Russian RMBK reactors has been attributed to DHC [2] and DHC has been 

observed in other cladding alloys and implicated to be the cause of long splits in some 

BWR fuel cladding, but the splits may also be attributed to oxidation of the fuel pellets.

Laboratory testing has demonstrated that the temperature dependence of DHC 

velocities in Zircaloy-4 is similar to those for Zr-2.5Nb (figure 8 in Reference 3). These 

and other similarities lead to the conclusion that the DHC behavior of Zircaloy-4 fuel 

cladding will be similar to that of the pressure tubes [3]. In addition to temperature, 

several material/environmental variables influence DHC velocities including hydrogen 

content, material chemistry and strength, microstructure, crack or flaw tip stress 

intensity, thermal history and crystallographic texture [2, 3, and 4]. Although validated 

crack growth models incorporating these influences are not available, the influence of 

some of the material/environmental variables has been established.

The temperature dependence of DHC velocities at temperatures below 275°C 

demonstrates that DHC is a thermally activated process. This temperature dependence 

does not mimic the activation energy for diffusion because hydride precipitation also 

depends on the hydrogen concentration reaching/exceeding the local hydrogen solubility 

limit. Hydrogen solubility is temperature dependent and, at a notch tip, dependent on 

the applied stress, the sharpness of the notch and the strength of the material [2, 5]. 

Other considerations that may impact DHC velocities include radiation effects and local 

variations in alloy chemistry.

Analysis of the impact of material and environmental variables on DHC growth rates in 

Zr-2.5Nb alloys [6] has shown that DHC growth rates increase with decreasing 

irradiation temperature, increasing operating flux and increasing operating time. These 

observations were made on surveillance samples and the experimentally observed axial 

DHC growth rates showed an Arrhenius-type temperature dependence (Figure 1 from 

Reference 6) with several orders of magnitude variation in crack velocities. Assessment 

of DHC in the surveillance samples in CANDU reactors is very prescriptive [7] and 

designed to provide a leak before break assessment of the pressure tubes [8]. 

A phenomenological description of DHC [9] concludes that the process proceeds as 

follows: “Hydrides nucleate and grow slowly in the high stress region of a stress riser 

such as a crack tip. When they reach a critical condition, probably related to size, they 

fracture, the crack extends and the process is repeated” [9]. The cracking process thus 

provides for a stable, time dependent, two-step crack growth mechanism. The first step 

is the diffusion of hydrogen to the crack tip region coupled with the nucleation and 

growth of a hydride in a region of high stress concentration, and this step is followed by 

hydride (or hydride metal interface) fracture and crack extension. Reference 9 

concludes that there are two technologically-relevant quantities for DHC: the critical 

stress intensity for crack initiation and the rate of crack propagation.
53

Measurements of the threshold (or critical) stress intensity for the initiation of DHC in 

Zr-2.5Nb alloys [4] suggest values in the range of 9 to 11 MPa m1⁄2 for samples tested 

between 205 and 283°C (from Figure 5 of reference 4). These threshold values are 

about one third of the critical fracture toughness values for Zircaloy-4 cladding that 

contains oriented hydrides and is tested at 25°C and much lower than the fracture 

toughness of the same alloy at 300°C (Figure 7 of reference 10). 

The difference between the threshold stress intensity values for DHC and the critical 

fracture toughness values demonstrates that, if DHC should initiate in the cladding of a 

spent fuel rod, crack growth would initially be stable. A fracture mechanics evaluation of 

flaw stability (see subsection 2.2.4 below) demonstrates that the crack would penetrate 

the cladding before becoming unstable. Clad penetration will relieve the internal

pressure in the fuel rod and minimize the driving force for crack growth. This conclusion 

is consistent with the assessments of DHC in Zr-2.5Nb pressure tubes which have been 

shown to leak as a result of stable DHC.

Evaluation of DHC from simulated debris fretting flaws showed that the threshold stress 

intensity for crack initiation was dependent on notch root radius and metallurgical 

condition as shown in Figure 5 of reference [11]. Additionally, the hydride morphology at 

the crack tips was markedly dependent on root radius (from Figures 10-12 of reference 

11) as was the probability of failure by DHC (from Figure 7 of reference 11) [11]. The 

difference in failure probability for samples stressed to equivalent applied stress 

intensities is due to the differences in crack tip shielding caused by hydride precipitation 

at the notch roots. 

Even though most of the DHC assessments have been for Zr-2.5Nb alloys, tests with six 

versions of Zircaloy-4 fuel cladding representing materials used in PWR, BWR, CANDU 

and Atucha reactors were basically consistent with results from Zr-2.5Nb alloys [9]. 

Material strength was a controlling factor for DHC crack velocities. The fastest crack 

rates were found in the strongest material. One basic difference in DHC in the Zircaloy4 relative to DHC in Zr-2.5Nb alloys was the lack of striations on the fracture surfaces. 

Reasons for this difference were not developed. 

The observations described above clearly demonstrate the possibility of DHC in fuel 

cladding during extended storage and transport and also suggest that the DHC behavior 

of zirconium based alloys is highly variable. The assembled evidence demonstrates that 

spent fuel rods may fail by DHC during storage upon cooling to temperatures where 

hydride precipitation may occur. However, such failures would produce small, through 

wall cracks rather than catastrophic tears in the cladding. Clearly the time/temperature 

profile necessary for the hydride precipitation required for DHC will depend on the 

hydrogen content in the cladding alloy which, as an upper bound in the reference fuel, 

was assumed to be 600 ppm. Reference to Figure 16 in Section 2.2.2 illustrates that 

hydride precipitation in an alloy containing 600 ppm hydrogen at high temperature would 

occur continuously as the fuel cools. Typical time/temperature profiles of fuel in dry 

storage suggest that a 300°C fuel cladding temperature is reached during the first few 

years of storage and that hydride precipitation should occur in the reference fuel. 

However, the onset of DHC also requires that several other conditions be satisfied.

The DHC phenomenon is generally subject to an incubation period and the impact of 

alloy chemistry, temperature, hydrogen content, cooling rate, time, stress intensity and 

other variables on this incubation period has not been established. For example, in 
54

fatigue pre-cracked samples loaded to an initial stress intensity of approximately 

17.3 MPa m1⁄2 showed large scatter in the incubation time for crack initiation and no 

evidence of any temperature dependence (from figure 3.13 of reference 9). 

Determinations of threshold stress intensities for the initiation of DHC used decreasing 

load techniques, thus the values determined physically approximated the stress intensity 

at cessation or arrest of measurable crack growth rather than the stress intensity 

required for DHC initiation. However, measurements with a variety of metals and alloys 

show that arrest and initiation Kth values are approximately the same, when 

measurements are made using reasonable experimental care. Therefore, the arrest 

threshold stress intensities shown in Figure 5 of reference 4 are assumed to represent 

the stress intensities required for DHC initiation. Clearly, metallurgical and 

compositional differences among the cladding alloys may alter the threshold stress 

intensity. In fact, in-service and laboratory experience with a variety of metals and alloys 

has shown that hydrogen induced cracking and stress corrosion cracking will occur on 

samples and components that have such minute defects that estimation of an applied 

stress intensity is extremely difficult. However, this difficulty may be overcome by 

evaluations using samples with smooth surfaces.

The behavior of statically loaded, smooth surface samples shows that environmentally 

induced cracking sequences can be divided into three stages: 1) an incubation stage 

that represents the time required for crack nucleation, 2) a crack propagation stage that 

represents the time required for a nucleated crack to grow the a critical size, and 3) final 

fracture. The effect of applied stress on this cracking sequence is illustrated in Figure 1 

of reference 12. This illustration assumes the existence of σth or stress level below which 

the degradation process will not initiate. In-situ studies [13] show that a stress of 75 to 

80 MPa is required for hydride re-orientation in Zircaloy. Earlier studies [14] showed that 

the threshold stresses was very dependent on fabrication history with values ranging 

from 40 to 120 MPa. Stress induced orientation of hydrides is assumed to be a critical 

component of the DHC process, thus σth for DHC should be approximately 80 MPa but 

could be as low as 40 MPa depending on the microstructure, texture, fabrication 

processes and other material/environmental variables associated with the particular 

zirconium based cladding.

The fact that smooth specimens, stressed above σth may be expected to fail by DHC can 

be conceptually combined with the threshold stress intensity for DHC initiation to 

estimate a “safe zone” where DHC should not develop. This concept is shown 

schematically in Figure 61 of reference 12 and demonstrates that a zone of conditions 

where DHC will not occur depends on both the stresses acting on the fuel cladding and 

the size of a flaw or crack in the cladding. Technologies for flaw characterization by nondestructive evaluation techniques are well established and have been applied to the 

identification of DHC induced flaws in zirconium alloys [15]. 

Interpretation that DHC will not occur in the absence of applied stresses (pressure for 

example) acting on the cladding must be viewed with caution. Residual stresses from 

fabrication, handling or in-service situations may provide local stresses that are sufficient 

for the incubation of DHC. Once initiated, DHC may become a self-sustaining process 

because of the stresses induced by the volume expansion associated with hydride 

precipitation. The potential for self-sustaining processes is evident in the appearance of 

the aligned hydride clusters.
55

Summary of Current Observations of DHC during Storage - There is no evidence 

that DHC is currently a significant problem in dry storage environments. However, the 

phenomenon has not been thoroughly studied by evaluating the cladding on either low 

or high burn-up fuels in dry storage. The evidence presented suggests that DHC is 

certainly possible and, given the time frames considered, the development of stable, 

throughwall cracks by DHC could occur. Cracks in the fuel cladding will provide 

pathways for fission product release and, if oxygen is present in the canister, will allow 

oxidation of the fuel. 

Additionally, cladding, irradiation, and handling variables will impact the potential for 

DHC and viable models to estimate the time to cracking as a function of storage 

conditions do not exist. Furthermore, even though there are numerous data for a variety 

of alloys that suggest, once DHC has initiated, the temperature dependence of DHC 

crack velocity shows an Arrhenius type behavior, the absolute value of crack velocity is 

sensitive to a variety of thermal, mechanical, and metallurgical variables. Sufficient 

information to assess the impact of material and environmental variables on DHC 

initiation and growth does not exist. 

A recent summary report [9] demonstrated that: 

a) DHC has been responsible for several failures in reactor components, 

b) Zircaloy fuel cladding is not immune from DHC, 

c) DHC cracking rates are independent of crack length and stress intensity, 

d) DHC crack velocities are highest in the highest strength materials and the added 

strength generated by irradiation increased crack velocities by a factor between ten 

and fifty, 

e) Threshold stress intensities for crack initiation increase with increasing 

temperature, and 

f) Failure of fuel cladding by DHC should not occur until the cladding has cooled to a 

temperature where DHC is possible. 

The evaluations discussed above show that if cladding cracks have been produced by 

DHC these cracks should penetrate the wall and allow release of the contained fission 

product gasses into the spent fuel canister and ingress of the canister environment into 

the fuel rod. Although the information on the impact of time, temperature, irradiation, 

hydrogen content etc., on DHC is not sufficient to predict the behavior of any specific 

fuel assembly, the occurrence of DHC will be stochastic. Only a fraction of the 

susceptible cladding will experience cracking, much like only a fraction of the stress 

corrosion cracking susceptible welds in reactors have cracked. As an example, stress 

corrosion cracking in the near weld regions of austenitic stainless steel primary coolant 

water piping in the Savannah River Site production reactors had been attributed to 

sensitization of the heat affected zone of the welds, residual stresses in the near weld 

region, and the presence of process water at approximately 100°C [16]. There were, 

literally, thousands of such welds and stress corrosion cracking was observed in less 

than fifty welds. Additionally, in the welds where cracking was observed, only a fraction 

of the weld length cracked and most of the weld was not degraded by the exposure. 
56

The tendency for DHC in the cladding of spent fuel rods is anticipated to follow a trend 

similar to that found for stress corrosion cracking, thus the absence of DHC in a few 

elements selected for evaluation does little to establish a firm technical basis to predict 

the behavior of the remaining fuel. An evaluation scheme that includes PIE of both high 

and low burn-up fuels is needed to maximize the technical gains that result from the 

limited number of evaluations that will be made during the surveillance of spent fuel 

during dry storage.

Assessment of DHC in Fuel Subject to Extended Storage and Transportation - The 

fact that DHC has been observed in zirconium based nuclear components coupled with 

the observation that the conditions that produce DHC may develop as the temperature of 

the fuel cladding decreases, demonstrate a significant potential for the failure, by DHC, 

of a number of fuel rods during long term storage. Such failure will only occur after the 

cladding has cooled down from the initial dryout anneal, during which about 

200-300 ppm of hydrogen will go into solution, depending on the temperature reached. 

As the clad cools the dissolved hydrogen will re-precipitate as hydrides in the clad 

structure. 

There is not sufficient information available to provide models to account for 

metallurgical, mechanical and chemistry effects on DHC and allow realistic predictions of 

cladding behavior during EST. 

Although additional data may improve modeling capabilities, a more realistic approach to 

assuring safe storage is to develop the handling and transport technologies necessary 

for fuels with minor failed cladding and to establish monitoring technologies to detect 

throughwall DHC. Cladding cracks should be anticipated and, as shown, the ductility 

and toughness of intact cladding may be significantly reduced by EST, especially at the 

lower temperatures encountered after decades of storage. The stresses necessary for 

the initiation of DHC in non-flawed regions may be as low as 40 MPa [14] and can 

readily be developed by the precipitation of a hydride nugget or rim or exist as residual 

stresses in the fuel rod.

Cracks in the fuel cladding are expected to be minor and not lead to gross rupture, but it 

would cause an impact to readily retrieve the fuel. Specifically, failure from DHC would

1) increase the source term in the canister/cask,

2) lessen the ready retrievability of the fuel during handling, post-storage transport or 

disposition, and

3) create a path to expose fuel pellets and fragments to oxygen if present in the 

internal canister environment because of bolting failures, seal leakage and/or 

stress corrosion cracking of the canister.

There would be no impact, by DHC, to cause a loss of confinement provided by the 

canister/cask of the DCSS. 

DHC Data Needs for EST and General Recommendations for Resolution - The data 

and models currently available to predict DHC during long-term dry storage are very 

limited and are of minimal use to the assessment of DHC in irradiated fuel cladding. 

Additional data are necessary and as data are accumulated, predictive models for DHC 
57

need to be developed, verified, and modified as appropriate. High burn-up fuels are 

anticipated to contain more hydrogen than low burn-up fuels, thus the temperature range 

where hydride precipitation will begin will differ for these two fuel classes. The 

susceptibility of zirconium alloys to stress orientation is influenced by the hydrogen 

content and thermal cycling, thus the data/model development schemes must include

fuels having a range of burnup and a range of hydrogen content in the fuel cladding. 

Additionally, concurrent with the development of DHC modeling, it is recommended that 

improved assembly temperature profile modeling be included. The temperature profiles 

both within and between stored assemblies from dry out annealing to the very long term 

storage will be critical parameters in evaluating fuel performance, including the potential 

for DHC. Specifically, 

1. Testing & Analysis Activity: Predictive model(s) for DHC including HBU fuel 

with high hydrogen contents with respect to thermal equilibrium at low 

(< 100°C) temperatures should be established for EST.

2. Cask Demonstration Activity: As part of a cask demonstration, DE should be 

performed with appropriate sampling to check for DHC. Fuel that has 

undergone a high temperature dryout cycle and has subsequently been at 

temperatures below approximately 300°C in dry storage and that has

experienced in-core failure (to provide “seed” cracks) is judged to be most likely 

to exhibit DHC. 

The impact of DHC to EST is anticipated to be troublesome but manageable 

considering: 1) that the DHC mechanism has the potential to form flaws in the cladding; 

2) any through-clad failure by DHC during dry storage would not be a gross cladding 

failure but may lead to an increase in source term within the confinement; and 3) flaws 

formed by DHC (and other existing cladding flaws) may be subject to extension with 

loading from handling and transportation.

Section 2.2.3 References

[1] Simpson, C.J. and Ells, C.E., “Delayed Hydrogen Embrittlement in Zr-2.5 wt% Nb”, 

Journal of Nuclear Materials, Vol. 52, pp. 289-295, 1974.

[2] Roth, M., Choubey, R., Coleman, C., and Ritchie, I., “Measurement of DHC 

Velocity in CANDU Pressure Tubes”, 17th International Conference on Structural 

Mechanics in Reactor Technology, Prague, Czech Republic, 2003.

[3] Coleman, C., et al “Delayed Hydride Cracking in Zircaloy Fuel Cladding- An IAEA 

Coordinated Research Program” Nuclear Engineering and Technology, Vol. 41, 

pp. 171-178, 2008.

[4] Singh, R.N., Stahle, P., Chakravarrtty, J.K., and Shmakov, A.A., “Threshold Stress 

Intensity Factor for Delayed Hydride Cracking in Zr-2.5% Nb Pressure Tube Alloy” 

Materials Science and Engineering A 523, pp. 112-117, 2009.

[5] Cox, B. and Rudling, P., “Hydriding Mechanisms and Impact on fuel Performance”, 

Advanced Nuclear Technology AB, ANT, Sweden, 2000.
58

[6] Gutkin, L. and Scarth, D.A., “Multi-Variable Engineering Model for Axial Delayed 

Hydride Cracking Growth Rate in Pre-Irradiated Zr-2.5%Nb” Proceedings of the 

ASME Pressure Vessels and Piping Division Conference, PRP20009, Prague, 

Czech Republic, 2009.

[7] Technical Requirements for In-Service Evaluation of Zirconium Alloy Pressure 

Tubes in CANDU Reactors, CSA N285. pp. 8-10.

[8] Pandey, M.D. and Sahoo, A.K., “A Risk-Informed Approach to Leak-Before-Break 

Assessment of Pressure Tubes in CANDU Reactors, Journal of Pressure Vessel 

Technology, Vol. 132, 021403-1 to 021403-8 (2010).

[9] “Delayed Hydride Cracking of Zirconium Alloy Fuel Cladding”, IAEA-TECDOC1649, International Atomic Energy Agency, Austria, 2010.

[10] Hsu, H.H. and Tsay, L.W., “Effect of Hydride Orientation on Fracture Toughness of 

Zircaloy-4 Cladding”, Journal of Nuclear Materials, Vol. 498, pp. 67-72 (2011).

[11] Cui, J., Shek, G.K., Scarth, D.A., and Wang, Z., “Delayed Hydride Cracking 

Initiation at Simulated Debris Fretting Flaws in Zr-2.5Nb Alloys”, Proceedings of 

PVP2006-ICPVT-11, Vancouver, BC (2006).

[12] ASM Metals Handbook, Vol. 13A, Corrosion: Fundamentals, Testing and 

Protection, ASM International, Materials Park, OH (2003).

[13] Colas, K.B., et al, “In-Situ Study of Hydride Precipitation Kinetics and Reorientation in Zircaloy using Synchrotron Radiation”, Acta Met, Vol. 58, pp. 6575-

6583 (2010).

[14] Louthan, M.R., Jr. and Marshall, R.P., “Control of Hydride Orientation in Zircaloy”, 

Journal of Nuclear Materials, Vol. 9, pp. 170-184 (1963).

[15] Nanekar, P.P., Mangsulikar, M.D., Cleveland, J., and Shah, B.K., “Flaw 

Characterization in PHWR Pressure Tubes by Ultrasonics: India’s Experience 

During IAEA CRP,” 12th A-PCNDT, Asia-Pacific Conference on NDE, Auckland, 

NZ (2006).

[16] WSRC-RP-90-97, “Reactor Materials Program – Process Water Piping Large 

Break Frequency,” W.L. Daugherty, Westinghouse Savannah River Co., Savannah 

River Site, 1990.

2.2.4 Flaw Stability Assessment in Zircaloy Tube

A simple assessment of axial flaws was performed to evaluate crack stability in a lowtoughness (embrittled) cladding. The evaluation was performed to investigate the 

potential for a gross cladding axial split to release the fuel materials and fission products 

from a hydride-embrittled cladding tube under internal pressure while in dry storage. 

The most limiting configuration under internal pressure loading is cracks aligned to grow 

axially along the tube while being driven by hoop stresses.
59

Figure 21 Postulated Flaws in the Cladding

Cracking emanating from either the ID or OD of the cladding is presumed as 

schematically shown in Figure 21. An axial thumbnail crack on the inner cladding 

surface is first considered. The thumbnail crack is simplified as a semi-elliptic crack with 

length 2c and depth a into the cladding (aspect ratio a/c) as shown in Figure 22 for a fuel 

rod with inner radius Ri and cladding thickness t. The crack is assumed to be loaded 

with a uniform circumferential stress resulting from fission gas pressure and other 

mechanical loadings. Linear elastic response is considered through this analysis so the 

Mode I stress intensity factors (SIF), denoted by K, can be calculated from handbook 

solutions [1]. The SIFs were calculated at the deepest penetration point and on the 

inner fuel rod surface, which provide insight on its potential to penetrate the cladding (by 

increasing a) or the potential to increase its crack length (2c). The membrane stress is 

assumed to be 100 MPa, which allows convenient scaling of the calculated SIFs that are 

proportional to the stress loading based on the principle of linear elasticity. A coldworked stress-relieved Zircaloy-4 clad fuel rod was chosen for the fracture analysis with 

low fracture toughness due to hydriding. This rod has an outside diameter of 10.71mm 

and a wall thickness of 0.62mm [2].

Outer Wall

of Cladding

Inner Wall

of Cladding

Oxide

(ZrO2)

Hydride Rim Uniform Corrosion

Zirconium Alloy Matrix

Crack

Circumferertial

Radial Hydrides Hydrides

σh

σh

Crack

Outer Wall

of Cladding

Inner Wall

of Cladding

Oxide

(ZrO2)

Hydride Rim Uniform Corrosion

Zirconium Alloy Matrix

Crack

Circumferertial

Radial Hydrides Hydrides

σh

σh

Outer Wall

of Cladding

Inner Wall

of Cladding

Oxide

(ZrO2)

Hydride Rim Uniform Corrosion

Zirconium Alloy Matrix

Crack

Circumferertial

Radial Hydrides Hydrides

σh

σh

Crack
60

Figure 22 An axial semi-elliptic crack with aspect ratio a/c in a fuel rod with inner 

radius Ri and cladding thickness t.

The calculated SIFs for the internal flaw are shown in Figure 23. Figure 24 shows the 

SIF results at the tip of the external flaw in comparison to the internal flaw as a function 

of flaw penetration depth. The crack length 2c is 1.054 mm, which is the maximum size 

limited by the validity of the solution [1]. This crack length is at 1.7 times the cladding 

thickness. Based on the result, the SIF for this crack configuration is far below the 

threshold value for DHC to initiate (in the range of 9 to 11 MPa m1⁄2 based on Zr-2.5Nb 

alloys which have been widely studies particularly for CANDU reactor systems). It is 

expected that the DHC threshold for Zircaloy-4 would be in the similar range. In 

addition, it was also reported in Table 3 of reference [3] that the fracture toughness is 

about 25 MPa m1⁄2 at room temperature for Zircaloy-4 sheet with random and mixed 

hydride platelet. These material property data suggest that it would require about 

250 MPa to initiate DHC at room temperature for this semi-circular crack (maximum 

SIF), and about 600 MPa to propagate this crack in a mechanically unstable manner.
61

Figure 23 Stress intensity factors at 100 MPa uniform stress loading for semi-elliptic 

cracks with length 2c= 1.7t in a typical fuel rod as functions of penetration 

depth (a/c).

Figure 24 Comparison of stress intensity factors at crack tips of internal and external 

thumbnail cracks at 100 MPa uniform stress loading.

In the case of an existing flaw which has already penetrated the cladding, the stability of 

the crack can be analyzed with an axial throughwall crack shown in Figure 25, where the 

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

Aspect Ratio a/c

SIF (MPa•

Crack Length 2C=1.054 mm (2c/t=1.7)

K at Deepest Point

K at Inner Surface

Axial Part-throughwall Crack

OD=10.71 mm, t=0.62 mm

Ri/t= 7.637 Stress=100 MPa
62

crack is under a uniform stress σm through the cladding thickness. The SIFs at a 

loading of 100 MPa were calculated up to the maximum crack length (17.7 mm or 

28.5 times the cladding thickness), allowed by the valid solution [4]. 

According to the previously quoted data for the fracture toughness of hydrided Zircaloy-4 

sheet at room temperature (about 25MPa m1⁄2 [3]), a throughwall crack must be 6 mm in 

length (or 10 times cladding thickness) to be split by the 100 MPa loading as indicated 

by Figure 26. Note that in this calculation, the cladding stress was assumed not to be 

relieved by the pre-existing crack. 

These results for the subcritical flaw growth and stability assessment suggest several 

conclusions:

1) internal pressure on the cladding is not sufficient to drive subcritical cracking; 

however, residual stresses and stresses generated by hydride precipitation 

may be sufficient to cause DHC.

2) should a subcritical flaw form during the dry storage period, the internal 

pressure on the cladding would not cause a gross rupture (e.g., axial split) of 

the cladding.

These conclusions are based on the cited published data for threshold stress intensities 

for DHC and room temperature fracture toughness value for a heavily hydrided 

microstructure. These results are consistent with the conclusion in reference [5] for no 

axial splitting failure of fuel rods. However, the analysis does not consider the 

development of a clustered, axially oriented hydride distribution. Current data suggest 

that such distributions may develop during long term storage and lead to stable, 

throughwall crack growth by DHC (see Section 2.2.3).

Figure 25 An axial throughwall crack with length 2c loaded by a uniform stress across 

the thickness t in a tube with mean radius R.
63

Figure 26 Stress intensity factors for axial flaws with 100 MPa uniform stress 

loading at various initial crack lengths.

Flaw Stability Evaluation Data Needs for EST and General Recommendations for 

Resolution - The methodology that is used for evaluating stability in Zircaloy rods 

involves comparison of the strain energy density against the material property of Critical 

Strain Energy Density (e.g., reference [6]). This methodology has been used in 

evaluation of fuel resistance to fracture in transportation loading events, but is not 

currently endorsed by the NRC. This methodology is not readily amenable to flawed 

materials due to discontinuities imposed by the flaws [5], and alternative methodologies 

are needed to provide improved accuracy and characterize flaw stability in a hydrided

rod with flaws. 

It is likely that fuel from EST will be subjected to loads during retrieval, handling, and 

transportation while the cladding temperature is below its DBTT. If this fuel contains

through-clad and/or non-through-clad flaws the fuel may be susceptible to brittle fracture. 

It is recommended that a consensus standard fracture approach be developed for fuel

flaw evaluation. This approach should, consider cladding type, temperature, hydride 

microstructure, and potential flaws as well as the loads developed during storage, 

handling, and transportation. Specifically, 

1. Testing & Analysis Activity: Develop fracture methodology, with national 

consensus, for flaw stability in Zircaloy fuel tubes with explicit consideration of 

subcritical growth leading to mechanical instability, multiple flaws, hydride 

(second phase) effects, flaw orientation, loading rates, and DBTT, etc. Existing 

fracture data for Zircaloy would be applied to this methodology and future 

materials properties and testing methods would be suggested for improvement in 

accuracy of flaw stability estimation.

0

10

20

30

40

50

60

70

0 5 10 15 20 25 30

2c/t

K (MPa•

Axial Throughwall Crack

OD=10.71 mm, t=0.62 mm

Ri/t= 7.637 Stress=100 MPa
64

The increased susceptibility of fuel to brittle failure due to both HBU (potential higher 

hydrogen content) and low temperature condition for EST make the fuel vulnerable 

during handling (with potential off-normal drop) and transportation loading conditions 

including accident. A comparison of the robustness and accuracy of fracture mechanics 

methods with the strain energy density method would establish an important technical 

basis for EST. 

Section 2.2.4 References

[1] Zahoor, A., Ductile Fracture handbook, NP-6301-D, V3, Electric Power Research 

Institute, Palo Alto, CA, January 1991, Vol. 3, Chapter 8, pp. 8.1-23.

[2] Einziger, R.E., et al, Examination Of Spent Pressurized Water Reactor Fuel Rods 

After 15 Years In Dry Storage, Nuclear Technology, Vol. 144, November 2003.

[3] Raynaud, P.A., Koss, D.A., Motta, A.T., and Chan, K.S., “Fracture Toughness of 

Hydrided Zircaloy-4 Sheet under Through-Thickness Crack Growth Conditions,” 

Journal of ASTM International, 2008, Vol. 5, No. 1, Paper ID JAI101183, ASTM 

International, West Conshohocken, PA 19428-2959.

[4] Zahoor, A., Ductile Fracture handbook, NP-6301-D, V2, Electric Power Research 

Institute, Palo Alto, CA, October 1990, Vol. 2, Chapter 6, pp. 6.1-5.

[5] Dry Storage of High-Burnup Spent Fuel Responses to Nuclear Regulatory 

Commission Requests for Additional Information and Clarification, EPRI 1009276, 

Final Report, November 2003.

[6] Spent Fuel Transportation Applications – Assessment of Cladding Performance, 

EPRI report number 1015048, December 2007.

2.2.5 Hydride Embrittlement – Reduction in Tensile Ductility

Hydride embrittlement will occur in zirconium alloy clad spent fuels placed in extended 

storage because reactor service has increased the hydrogen content in the cladding and 

the storage temperature will decrease to the point that hydrides precipitate in the 

cladding. The fact that hydride precipitation lowers the ductility of zirconium based 

alloys was well established by the early 1960s [1] and has been confirmed in numerous 

investigations. For example, Figure 27 shows the changes in the strength and ductility 

of Zircaloy-4 tensile samples as a function of hydrogen content. 

The literature clearly demonstrates that hydrogen uptake and hydride precipitation 

decrease the ductility of zirconium based alloys. The extent of hydride embrittlement is 

a function of the hydrogen content and the hydride orientation. Hydride orientation is so 

important to the embrittlement process that models to estimate the ductility (and 

strength) of Zircaloy as a function of fluence, burnup, oxide thickness, hydrogen 

concentration, temperature, strain rate and cold work contain the statement “the models 

described here apply only to cladding with circumferential hydrides and do not apply to 

cladding with radial hydrides or significant hydride blisters or spalling” [4]. The 

precipitation of radial hydrides can reduce the ductility of Zircaloy cladding alloys to 
65

essentially zero and cause normally ductile samples to fail while in the elastic portion of 

the stress-strain curve [1].

Irradiation also reduces the ductility of zirconium based alloys and the combined effects 

of irradiation show that when irradiated to a fluence of 1.6 to 9.5 x 1025 n/m2

, Zircaloy-2 

specimens loose essentially all their ductility when the hydrogen content reaches 600 

ppm (Figure 4 of Reference 4). The hydrogen concentration in irradiated zirconium 

alloys generally increases with increasing fluence (Figure 2 of Reference 4) although 

there is considerable scatter in the hydrogen concentration in fuel claddings after 

irradiation to high fluence. Therefore, although most HBU fuels have a range of fluence

between 9 and 11 x 1025 n/m2 the hydrogen content in the cladding alloys on those fuels 

can be quite variable [4]. For example, the corrosion and hydrogen levels at a burnup of 

60 GWd/MTU for low tin Zircaloy-4 can vary between 50 and 100 μm of oxide and 400 to 

800 ppm hydrogen [4]. Similar estimations for Zircaloy-2 from BWR fuel at the same 

burnup suggest that the oxide thickness can vary between 20 and 45 μm and the 

hydrogen content between 100 and 220 ppm. 

There are numerous literature examples demonstrating that the combination of 

irradiation and hydride precipitation can significantly reduce the ductility of zirconium 

base alloys. For example a 1994 summary paper [5] shows irradiation increasing the 

strength and decreasing the ductility of Zr-2 alloys. Figure 27 illustrates the effects of 

hydrogen on mechanical properties and reaffirms the hydrogen induced reductions in 

ductility. Another investigation [6, 7, and 8] of the reduction loss of ductility with 

hydrogen up to 4500 ppm shows the importance of the hydride microstructure on the 

ductility in zircaloy claddings.

These reductions are most apparent at low temperatures and therefore will become 

more serious as the temperature of the fuel cladding decreases during storage. There is 

not sufficient information available to assess the structural integrity of spent fuel during 

storage and transport when the cladding has experienced significant hydride 

embrittlement and fails under elastic stresses.
66

Figure 27 Impact of Hydrogen Content on the Tensile Properties of 

Zircaloy-4 at Room Temperature [Figure reproduced from ref. 2].

HE Ductility Reduction Data Needs for EST and General Recommendations for 

Resolution - Additional data should be developed and models, such as those published 

in Reference 4, must be modified to include cladding alloys with radial hydrides and also 

include the emerging cladding alloys that may show behaviors that differ significantly 

from those of Zircaloy. Emphasis should be placed on the HBU fuels because their 

hydrogen content is expected to exceed that for low burnup fuels. Specifically, 

1. Testing & Analysis Activity: Synthesize tensile data to establish model(s) for 

properties based on alloy composition, hydrogen content, hydride orientation, 

static and dynamic loading, bi-axial (mixed mode) loading, radiation damage and 

hydride microstructures, etc. to enable accurate prediction across fuel varieties.

The increased susceptibility of fuel to brittle failure due to both HBU (potential higher 

hydrogen content) and low temperature condition for EST make the fuel vulnerable 

during handling (with potential off-normal drop) and transportation loading conditions 

including accident. 

Section 2.2.5 References

[1] Louthan, M.R., Jr. and Marshall, R.P., “Control of Hydride Orientation in Zircaloy”, 

Journal of Nuclear Materials, Vol. 9, pp. 170-184 (1963).

[2] Cox, B. and Rudling, P., “Hydriding Mechanisms and Impact on Fuel 

Performance”, Advanced Nuclear Technology AB, ANT, Sweden, 2000.
67

[3] Huang, J.-H. and Huang, S.-P., “Effect of Hydrogen Contents on the Mechanical 

Properties of Zircaloy-4”, J. Nucl. Mater. Vol. 208, pp. 166-180, (1994).

[4] Geelhood, K.J., Luscher, W.G., and Beyer, C.E., “PNNL Stress/Strain Correlation 

for Zircaloy”, PNNL-17700, July 2008.

[5] Zirconium Alloys in Nuclear Applications, Chapter 7 in Materials Science and 

Technology, Vol 10B, Nuclear Materials, Part II, VCH New York.

[6] Arsene, S., Bai, J.B., and Bompard, P., “Hydride Embrittlement and Irradiation 

Effects on the Hoop Mechanical Properties of Pressurized Water Reactor (PWR) 

and Boiling-Water Reactor (BWR) ZIRCALOY Cladding Tubes: Part I. Hydride 

Embrittlement in Stress-Relieved, Annealed, and Recrystallized ZIRCALOYs at 

20°C and 300°C,” Met. Mat. Trans. A, Vol. 34A, Number 3, p. 553 (March 2003).

[7] Veleva, M., Arsene, S., Record, M., Bechade, J.L., and Bai, J.B., “Hydride 

Embrittlement and Irradiation Effects on the Hoop Mechanical Properties of 

Pressurized Water Reactor (PWR) and Boiling-Water Reactor (BWR) ZIRCALOY 

Cladding Tubes: Part II. Morphology of Hydrides Investigated at Different 

Magnifications and their Interacton with the Processes of Plastic Deformation,” 

Met. Mat. Trans. A, Vol. 34A, Number 3, p. 567 (March 2003).

[8] Arsene, S., Bai, J.B., and Bompard, P., “Hydride Embrittlement and Irradiation 

Effects on the Hoop Mechanical Properties of Pressurized Water Reactor (PWR) 

and Boiling-Water Reactor (BWR) ZIRCALOY Cladding Tubes: Part III. Mechanical 

Behavior of Hydride in Stress-Relieved Annealed and Recrystallized ZIRCALOYs 

at 20°C and 300°C,” Met. Mat. Trans. A, Vol. 34, Number 3, p. 579 (March 2003).

2.2.6 Hydrogen Embrittlement - Hydrogen Induced Shifts in the DBTT

Numerous references, including those highlighted in Section 2.2.5, show that the 

fracture resistance of zirconium based alloys decreases with increasing hydrogen 

content. The decrease includes a concomitant shift in the DBTT, as illustrated in Figure 

28 and the extent of the shift will vary with alloy composition, hydride orientation, 

temperature and irradiation level as well as hydrogen content. Although the hydrogen 

induced decrease in fracture resistance is well established, the database to analyze the 

influence of hydride orientation, hydride rims, alloy type and irradiation on fracture 

resistance is not comprehensive and virtually non-existent for newer cladding alloys. 

High burn-up fuel may be especially susceptible to hydrogen embrittlement and 

hydrogen induced shifts in the DBTT because of the higher hydrogen contents that result 

from the longer in-reactor time frames [1].
68

Figure 28 Ductile-to-brittle transitions in hydrided zirconium 

[Figure reproduced from ref. 2].

The shift in DBTT may become critical to the behavior of fuel cladding during post 

storage handling and transport. Significant reductions are possible if the combination of 

alloy type, thermomechanical history, hydrogen content and storage conditions lead to 

the formation of clustered, radial hydrides. Such hydride formations can make the 

cladding brittle in a room temperature tensile test, even when the hydrogen content is 

only 100 ppm and the alloy has not been irradiated [3].

The elevation of the DBTT is generally associated with the presence of hydrides in the 

microstructure and should therefore become increasing important as the storage 

temperature decreases. Therefore the potential for such a shift should increase with 

increasing storage time (decreasing storage temperature) and reach its maximum at the 

end of the fuel storage period. This potential for a hydrogen induced increase in the 

DBTT will generally increase as the fuel burnup increases because the hydrogen content 

of the cladding alloys tends to increase with increasing fuel burnup. However, the 

database necessary to analyze the time/temperature/burnup dependence of the DBTT 

for cladding alloys during very long term storage is insufficient and the data necessary to 

couple the DBTT with other cladding variables are minimal, at best.

These evaluations illustrate a significant potential for hydrogen induced reductions in the 

fracture resistance of the cladding on stored fuel assemblies and also illustrate that there 

are no models to accurately assess the magnitude of that potential. Additionally, even if 

the potential for hydrogen induced increases in the DBTT could be established, the 

models to assess the impact of DBTT increases on the handling and retrieval of spent 
69

nuclear fuels after very long term storage are not currently available, as discussed in 

Section 2.2.4. 

Data Needs for Hydrogen Induced Increases in the DBTT - The data and models 

currently available are not sufficient to assess the impact of long term storage on the 

DBTT of cladding alloys. Data are needed to establish the combined effects of alloy 

content, hydrogen content, irradiation, temperature and other cladding variables on the 

DBTT (and other measures of fracture resistance) of fuel cladding. Emphasis should be 

placed on HBU fuel [1] and efforts should be made to establish the singular and 

combined effects of irradiation and hydrogen content on fracture resistance. 

Specifically, 

1. Testing & Analysis Activity: Synthesize existing fracture data and develop the

additional fracture toughness data necessary to establish model(s) for properties 

based on temperature, alloy composition, hydrogen content, hydride orientation 

(with respect to crack orientation), static and dynamic loading, bi-axial (mixed 

mode) loading, combined radiation damage and hydride microstructures, etc. 

The models developed should enable accurate prediction across fuel varieties

and be validated using a variety of fuel claddings.

The increased susceptibility of fuel to brittle failure due to both HBU (potential higher 

hydrogen content) and low temperature condition for EST make the fuel vulnerable 

during handling (with potential off-normal drop) and transportation loading conditions 

including accidents. 

Measurements of the fracture resistance of irradiated cladding requires significant 

resources and, before such resources are allocated, attempts should be made to model 

the behavior of flawed fuel assemblies during postulated post storage handling and 

transport conditions when the fracture resistance of the cladding is minimal (represented 

by the lower shelf fracture toughness at room temperature). A satisfactory outcome for a 

validated model using lower shelf fracture toughness properties could mitigate the need 

for data on irradiated, high burn-up fuel cladding. However, if a satisfactory outcome 

does not develop, data gathering for HBU spent fuel cladding should become a data 

need focus.

Section 2.2.6 References

[1] Very High Burn-ups in Light Water Reactors, ISBN 92-64-02303-8, Nuclear Energy 

Agency, NEA 6224, 2006.

[2] Fig. 55, Environmentally Induced Cracking, p 175, ASM Metals Handbook, Vol. 13, 

Corrosion: Fundamentals, Testing and Protection, ASM International, Materials 

Park, OH (2003).

[3] Marshall, R.P., and Louthan, M.R., Jr. “Tensile Properties of Zircaloy with Oriented 

Hydrides”, Trans. ASM, Vol. 56, pp. 693-698, 1963.
70

2.2.7 SCC of Fuel

Stress corrosion cracking of Zircaloy clad SNF due to the presence of fission product 

iodine was observed during the 1960’s and continued through the early 1980’s. This 

cracking followed reactor power transients, particularly in BWRs. The stress corrosion 

process was identified as Pellet Cladding Interaction (PCI) and the cracks developed 

from the clad interior. To address this problem, a layer of pure zirconium was added to 

the inner surface of the cladding, effectively eliminating the susceptibility to SCC while 

the fuel was in the reactor [1]. 

However, many of the fuel rods without the pure zirconium layer have been placed into 

dry storage and the potential for recurrence of SCC as a result of radiolysis of CsI has 

not been thoroughly investigated. Additional studies of the potential for SCC in 

zirconium based fuel cladding alloys placed in long term storage are needed because 

even though the mechanism of cracking is well established, the failure propensities still 

remain unresolved for spent fuel rods. Although a number of works have been 

conducted on iodine SCC of Zircaloy in relation to the pellet clad interaction induced 

failures of LWR fuel rods, most of the experimental works in which Zircaloy was exposed 

to iodine and other chemical elements are within the range of operating temperatures 

below about 320°C and exposure times shorter than 259 ks [2]. There is little data 

related to longer time exposures or other temperature conditions.

Most of the fission-product iodine is tied up in compounds with zirconium and with 

fission-product cesium and exists in the fuel as ZrI4 or CsI. Experimental evidence 

indicates radiolysis can create an iodine partial pressure approximately 9 orders of 

magnitude greater than would be otherwise present in a spent fuel rod at 400°C [3]. 

The presence of this free iodine may lead to SCC in fuel with appropriate cladding 

stress/strain conditions. Additionally, as seen in Figure 29, the incubation time for 

cracking is highly dependent on the iodine vapor concentration.

In addition to the consideration of the classic PCI–SCC related cracking, the fission 

products present on the clad ID surface may synergistically contribute to crack 

propagation by interacting with the DHC mechanism. Most if not all of the DHC tests 

were done with de-fueled clad samples. These tests did not have the added parameter 

of the SCC factor from the I or Cs. Thus, future DHC testing should also consider the 

potential contribution of the elements/compounds that are present at the clad ID surface.

Top Nozzle – Stainless Steel Connecting Tubes Stress Corrosion Cracking - NRC 

issued Information Notice 2002-09 to illustrate the fact fuel handling events had resulted 

in the separation of the top nozzle from a fuel assembly. This separation caused the 

subsequent drop of the assembly during fuel movement [4]. The fuel assemblies were 

identified as Westinghouse fuel assemblies supplied in the early to mid-1980s that 

contained materials that were susceptible to intergranular stress-corrosion cracking 

(IGSCC). Several methods to allow this fuel to be loaded into dry storage have been 

used [5]. Fuel placed in dry storage should be evaluated for such cracking [4].
71

Figure 29 Exposure Time Dependence of Iodine-Vapor-Induced SCC 

[Figure reproduced from ref. 2].

Fuel SCC Data Needs for EST and General Recommendations for Resolution: 

Failures due to PCI-SCC have occurred with earlier generation Zircaloy cladding but 

have been generally mitigated in later generation fuel. Nevertheless, fuel susceptible to 

SCC is in dry storage, and radiolysis and exposure time may be sufficient for incubation 

of this degradation phenomenon. Fuel should be examined for SCC to provide 

additional information. Specifically, 

1. Cask Demonstration Activity: As part of a cask demonstration, DE of the cladding 

should be performed with appropriate sampling to check for SCC. 

Considerations include zirconium-lined and non-lined fuel. HBU fuel that may be 

susceptible to DHC should be evaluated to check for potential interactive effects.

The significance of cladding SCC to EST is considered minimal considering: 1) the issue 

for present fuel has been resolved with the addition of zirconium to the ID surface; 2) any 

fuel rod breach that would occur due to SCC would be a minor breach with a small 

release to the confinement provided by the DCSS.

Section 2.2.7 References

[1] Armijo, J.S., Coffin, L.F., Rosenbaum, H.S., Development of zirconium barrier fuel 

cladding, Zirconium in the Nuclear Industry – 10th International Symposium, ASTM 

STP 1245, Eds, A.M. Garde and E.R. Bradley, ASTM, West Conshohocken, PA, 

(1994), pp. 3–18.
72

[2] Mayuzumi, M., Onchi, T., and Matsuo, Y., “Failure Propensities Of Pressurized 

Zircaloy Tube Containing Iodine And Other Chemical Species,” Journal Of Nuclear 

Science And Technology, 30 [12], pp. 1288-1298 (December 1993).

[3] Konashi, K., Kamimura, K., and Yokouchi, Y., “Estimation of Irradiation Induced 

Iodine Pressure in an LWR Fuel Rod,” J. Nucl. Mater. Vol. 125, pp. 244-247, 

(1984).

[4] Information Notice 2002-09, “Potential for Top Nozzle Separation and Dropping of 

a Certain Type of Westinghouse Fuel Assembly”, on February 13, 2002.

[5] Industry Spent Fuel Storage Handbook EPRI 1021048 Final Report, July 2010.

2.2.8 Fission Product and Helium Gas Pressurization 

The stress acting on the cladding is an important factor in many of the age related 

degradation processes that have the potential to impact the performance of the spent 

fuel during long-term storage. For example, stress corrosion cracking, creep and 

hydrogen induced degradation require the presence of a tensile stress, and the potential 

for significant degradation increases as the stress increases. The dominant, nonlocalized stress on the fuel cladding results from the internal pressure in the fuel pin. 

This pressure will increase with fuel burnup due to increased fission gas release as 

suggested in Figure 30. 

The fuel pin internal pressure decreases with decreasing temperature and therefore will 

generally decrease with time during storage. However, the release of fission product 

gases, including helium and iodine, and the production of helium by alpha decay may 

occur during dry storage and could increase the internal pressure. There are, however, 

two factors that demonstrate that cladding stresses due to internal pressurization will not 

increase significantly during the storage period:

1. the production of fission product gasses will have essentially ceased when the 

fuel was removed from the reactor and the most of the release of these gasses 

from the fuel into the fuel pin interior occurred before the onset of dry storage, 

and 

2. the buildup of helium by alpha decay will not be significant until after thousands 

of years of storage.

These two factors, coupled with the decreasing temperatures within the dry storage cask 

system, lead directly to the conclusion that changes in helium pressurization of spent 

fuel pin in dry storage will not be significant during a 300-year storage period. This 

conclusion is consistent with the information presented in Figure 31 for CANDU fuel 

which illustrates that no significant increase in gas volume occurs until approximately 

800-years of storage.
73

Figure 30 Fission gas release from UO2 fuel with burnup 

[Figure reproduced from ref. 1].

Figure 31 Helium Build-up in Fuel Pin (CANDU Fuel) 

[Figure reproduced from ref. 2].
74

Section 2.2.8 References

[1] IAEA Nuclear Energy Series, No. NF-T-3.8, Impact of High Burnup Uranium Oxide 

and Mixed Uranium-Plutonium Oxide Water Reactor Fuel on Spent Fuel 

Management, IAEA, 2011.

[2] UNENE, McMaster University, Hamilton, Ontario, CANADA.

2.2.9 Creep and Diffusion Controlled Cavity Growth (DCCG) 

Diffusion-Controlled Cavity Growth is a micromechanism for creep degradation of the 

fuel cladding. Coble creep and other creep mechanisms that rely on atom migration by 

low activation energy diffusional processes can also cause measurable creep at 

temperatures below Tm/2. Studies with zirconium based alloys suggest that the lower 

temperature limit for significant creep in zirconium based alloys at realistic stress levels 

is approximately 350°C. Therefore, because the temperature decreases continually 

during dry storage, creep is not anticipated to produce significant strain after the initial 

storage period. However, there are significant alloy-to-alloy differences in creep rates 

(for example low tin SRA Zircaloy 4 creeps twice as fast as regular RXA Zircaloy 2) and 

estimates of creep strains during the initial storage period should be based on data and 

analysis directly applicable to the cladding alloy of concern and data gaps certainly exist 

in this area.

Creep processes will provide characteristic microstructural evidence that cladding 

damage is occurring prior to the onset of failure. However, collection of such evidence 

requires destructive evaluation of the material of interest. SEM examinations of the 

fracture surfaces and cross sections of the near fracture surfaces of the zirconium alloys 

that were undergoing creep revealed extensive cavitation damage. These results, 

combined with an examination of the initial material condition and the condition of 

samples during steady-state creep, suggest that failure of Zircaloy under conditions 

relevant to the dry storage of spent nuclear fuel occurs by a cavity nucleation, growth, 

and interlinkage mechanism [1].

Chin and Gilbert [2] constructed a fracture mechanism map based on Ashby from which 

they concluded that creep under the initial dry storage conditions would occur by 

diffusion controlled cavity growth (DCCG). They used the Raj and Ashby [3] model for 

DCCG fracture mechanism to predict the fracture lifetimes for the anticipated 

temperature–time variations during the initial period of dry storage. However, the 

acceptance of this model is not universal. 

Zircaloys are generally considered to be resistant to cavity formation and there is limited 

evidence of creep-induced grain boundary cavity formation in unirradiated or irradiated 

Zircaloy. Keusseyan et al.’s [4] work does show evidence for cavity formation and a 

potential for creep induced damage.

The vast majority of the literature, however, clearly shows that cavity formation is 

unlikely in Zircaloy clad spent fuels. Even high neutron radiation exposures do not 

provide evidence of voids or cavities in Zircaloys. This behavior is unlike that of 

stainless steels and other metals which show extensive evidence of creep-induced voids 

and cavities. Many investigators attempted to find cavities and reported black-dot 
75

microstructures comprising dislocation loops [5]. Yoo [6] presented a model based on 

the anisotropy of Zircaloys to rationalize the absence of voids and cavities. 

Based on the proposal by Chin and Gilbert [2], many investigators made detailed 

calculations assuming the predominance of DCCG. However, Pescatore and Cowgill [7] 

point out the probable lack of applicability of this model to creep in dry storage based 

again on the lack of sufficient experimental evidence for cavities in deformed Zircaloy [8]. 

Assessment of DCCG Degradation in Fuel Subject to Very Long-Term Storage -

The pressure loading on the cladding of the pin-design fuel is the driving load for clad 

creep. Creep can be significant at high temperatures and/or over long time periods at 

moderate temperatures with moderate to high stresses [9].

The fission gas inventory in a fuel rod increases proportionally with burn-up, while the 

decreasing fuel thermal conductivity and changes of fuel microstructure reduce the 

ability of the fuel pellet to retain the fission gas. This, together with the rod void volume 

reduction caused by cladding creep-down during reactor operations and fuel pellet 

swelling due to irradiation, increases the potential for high rod internal pressures with 

increasing burn-up [9]. Therefore, although creep damage is not anticipated, the 

potential for creep, especially when high burn-up fuels are stored and during high 

temperature processing such as dry out anneals is considered in the present technical 

bases for DCSS. However, potential creep concerns are basically related to the initial 

storage period and are not anticipated to be significant during long term dry storage. 

DCCG Degradation Data Needs for EST and General Recommendations for 

Resolution - The degradation mechanism of DCCG, if operative would cause local 

rupture of the cladding on fuel rods. This failure phenomenon is preceded by creep 

deformation and the driving force for deformation in the cladding is reduced as creep 

deformation takes place, as discussed in section 2.2.1. Under EST, creep damage is 

not considered likely due the low temperatures after 50-60 years of storage and the 

resulting low driving forces for creep. The recommendations for creep evaluations in 

section 2.2.1 address DCCG with the additional note to verify the presence or absence 

of triple boundary cavitation in materials that is thought to have undergone creep. 

Section 2.2.9 References

[1] Hayes, T. A. and Kassner, M.E., “Creep of Zirconium and Zirconium Alloys,” 

Metallurgical and Materials Transactions A, Volume 37A, August 2006.

[2] Chin,B.A.,Gilbert,E.R.,1989. Prediction of Maximum Allowable Temperature for Dry 

Storage of Zircaloy-Clad Spent Fuel in Inert Atmosphere. Nucl. Tech. 85, pp. 57–

65.

[3] Raj, R., Ashby, M.F., 1975, Intergranular Fracture at Elevated Temperature. Acta 

Met. 23, pp. 653–667.

[4] Keusseyan, R.L., Hu, C.P., Li, C.Y., 1979. Creep Damage In Zircaloy-4 At LWR 

Temperatures. J. Nucl. Mater. Vol. 80, pp. 390–392, (1979).

[5] Wolfenden, A., Farrell, K., 1972, On The Question Of Void Formation In Neutron 

Irradiated Zirconium, Scripta Met. 6, pp. 127–129.
76

[6] Yoo, M.H., 1994, Suppression of Void Formation in Zirconium. In: Zirconium in 

Nuclear Applications, ASTM STP 551, pp. 292–307.

[7] Pescatore,C.,Cowgill,M.G.,1994. Temperature Limit Determination for the Inert Dry 

Storage of Spent Nuclear Fuel, EPRI Report TR-103949.

[8] Short-term rupture studies of Zircaloy-4 and Nb-modified Zircaloy-4 tubing using 

closed-end internal pressurization, Y.Zhou, B.Devarajan, K.L.Murty, Nuclear 

Engineering and Design 228 (2004).

[9] Light Water Reactors, ISBN 92-64-02303-8, Nuclear Energy Agency, NEA 6224, 

2006.

2.2.10 Air Oxidation of the Cladding – Off-Normal Condition

An off-normal condition that may occur during EST is air-ingress due to failure of 

canister seals or weldments, or failure of the canister body itself. There is a potential for 

stress corrosion cracking of the canister (see section 2.6.2) and throughwall cracks will 

provide a pathway for air ingress as would seal or weldment failure. The ingress of air 

into the canister will provide an oxygen rich environment in the canister interior. This 

oxidizing environment may interact with the fuel cladding and, if the cladding is 

breached, with the fuel pellet. Both fuel pellet and cladding oxidation are strongly 

dependent on temperature, available oxygen supply and time at temperature. 

Excessive oxidation of the Zircaloy cladding combined with an internal stress can 

potentially breach the cladding. Oxidation of the Zircaloy [1] is a thermally induced 

process. Dry storage under inert gas conditions leads to no further increase in the oxide 

layer that existed on the cladding at final discharge from the reactor. Pool storage 

conditions (water temperature) mitigate the continued growth of the oxide and the 

absence of oxidizing species in the as-designed DCSS does not enable continued 

oxidation. However, if the canister is breached while the internal temperature is above 

300°C oxidation of the cladding may occur. This oxidation will reduce the effective wall 

thickness of the cladding and can lead to rupture if the combination of oxygen induced 

wastage and cladding stress is sufficient. Constitutive relationships for cladding oxidation 

are contained in reference [2].

Oxidation of the pellet will cause swelling and may cause the cladding to split locally and 

the fuel pin to fragment (see section 2.2.11). However, cladding oxidation at the low 

temperature (< 200°C) of EST is not considered significant and any air oxidation impact 

will be due to oxygen interactions with the fuel pellet.

Assessment of Air Oxidation of Cladding - Impact to Safety Function - Extensive 

cladding failure could result with cladding wall loss through continued oxidation of the 

cladding with long-term exposure to air at high cladding temperature (>> 300°C); 

however the occurrence of such oxidation is extremely unlikely during EST because of 

the low temperatures (< 300°C) of exposure.

Cladding Air Oxidation Degradation Data Needs for EST and General 

Recommendations for Resolution - Oxidation of the fuel cladding in low temperature 

considerations of EST is not a significant degradation mechanism. An evaluation should

be made for air oxidation of fuel in a demonstration program if the fuel is stored at the 
77

off-normal condition of air ingress to provide verification of lack of significant changes to 

the cladding as a result the event.

Section 2.2.10 References

[1] Rothman, A. J., “Potential Corrosion and Degradation Mechanisms of Zircaloy 

Cladding on Spent Nuclear Fuel in a Tuff Repository,” Report Attachment 10 to 

MRB-0418, JUCID-20172, Contract W-7405-Eng-48, Lawrence Livermore National 

Laboratory, Livermore, CA, 1984.

[2] Einziger, R.E., Baldwin, D.L., and Pitman, S.G., Data Needs for Long-Term Dry 

Storage of LWR Fuel, EPRI Report TR-108757, April 1998.

2.2.11 Air Oxidation of the Fuel Pellet Material – Off-Normal Condition

Oxidation of the fuel, UO2, is possible only if the fuel is exposed directly to air in the dry 

storage cask system. Such exposure is possible as illustrated in section 2.2.10 and a 

comprehensive review of the mechanisms and kinetics of fuel oxidation has been 

performed [1]. Typical spent fuels oxidize by first forming either U3O7 or U4O9, phases 

that are denser than the original UO2. This first transition is accompanied by a net 

contraction of the fuel, relieving mechanical stress on the cladding. Oxidation typically 

occurs first along the fuel grain boundaries. Thus, the release of fission gases, such as 

85Kr, to the DCSS relatively early during this transition is possible. Upon further 

exposure to air, the fuel is oxidized to U3O8, a phase which is approximately 36% less 

dense than the original fuel, thus oxidation ultimately results in swelling of the pellet. 

The swelling of the fuel as U3O8 forms has been shown to supply sufficient mechanical 

stress to split the cladding [2]. U3O8 formed by oxidation of the fuel is a fine powder that 

spalls from the fuel surface. The release of fines and/or fuel relocation from split cladding 

must be evaluated if U3O8 formation is suspected. The extent of oxidation of irradiated 

UO2 is a time and temperature dependent phenomenon. In addition, the oxidation of 

spent fuel to U3O8 has been shown [3] to be burnup-dependent. As burnup increases, 

the soluble fission products and actinides stabilize the matrix and inhibit the formation of 

U3O8. For example, at 229°C, the reported incubation time for crack propagation due to 

fuel oxidation was approximately 1000 h for a rod with an assembly average burnup of 

about 12 MWd/kg U, whereas a rod with an assembly average burnup of about 

30 MWd/kg showed no sign of splitting or dilation after 5962 h [2]. At the lower 

temperatures expected during extended storage, oxidation of even low-burnup fuels will 

take sufficiently long, at least months or years, that detection and mitigation of the offnormal condition should prevent fuel oxidation from being a concern. At higher burnup 

levels, under the conditions of extended storage, sufficient oxidation to propagate 

cladding defects should not occur and no mitigative measures may be needed to protect 

against pellet oxidation and swelling which might lead to cladding rupture. 

The potential for air oxidation of MOX fuels is likely to be higher than that for UO2 fuels 

because the flatter temperature profile during cooling will increase the time that the fuels 

and cask are at higher temperatures. This increase may promote stress corrosion of the 

cask and should stress corrosion occur promote oxidation of the fuel. Additionally, the 

handling difficulties associated with MOX fuel fragments must be evaluated.
78

Assessment of Air Oxidation of Fuel - Impact to Safety Function - There is a 

potential to load some fuel into the canister that initially has a pin hole type clad breach. 

Also, during long term storage some degree of clad breach may occur from operating 

mechanisms like DHC. Coupled with a failure of the canister and air ingress (loss of 

confinement), there is a potential to have gross clad failures from fuel oxidation. The 

oxidation process would expand the fuel and stress the cladding and could lead to gross 

rupture of the cladding. Additionally, the dimensional changes in the fuel oxide may 

cause a loss of fuel retrievability. The potential air oxidation rates will be coupled with 

the air ingress rates and will depend on the temperature of the fuel at the time of 

container compromise, the specific cladding alloy in the compromised container, the fuel 

type (uranium dioxide or MOX) and the irradiation history. Qualitatively, a loss of 

confinement provided by the canister and an increase in source term within the canister 

will result. The data necessary to evaluate the impact of these potential issues are not 

available; however, this assessment concluded that the impact to safety, if any, will 

involve an increase in the source term and a loss of retrievability.

Data Needs for Air Oxidation and Recommendations - Air oxidation of the fuel and 

fuel cladding will only occur if the cover gas in the containment system is compromised. 

Air ingress could occur if corrosion or stress corrosion cracking occurred in the canister 

or, for bolted closures, if the bolts failed or the seals relaxed. The rate of oxidation will 

depend on the fuel and cladding temperatures and physical conditions at the time of air 

ingress. The development of the data necessary to assess these quantities will require 

extensive resources and the return on investment will be minimal. Therefore, periodic 

monitoring for cover gas and containment efficacy is recommended throughout EST to 

assure that the inertness of the cover gas environment and avoid air oxidation.

1. Monitoring/In-Service Inspection Activity: Periodic monitoring for cover gas and 

confinement efficacy in necessary throughout EST and acceptance criteria 

should be developed as should remediation plans for action should confinement 

be lost. 

2. Testing & Analysis Activity: Evaluate the potential for MOX fuel pellet oxidation 

(as was done for UO2 fuel with burn up in PNNL 11929).

A very small fraction of the cladding on fuels in EST will contain minor breaches and only 

a small fraction of the canisters are anticipated to allow air ingress. Thus, air oxidation of 

fuel in EST is not likely but is clearly possible. However, since such oxidation could lead 

to cladding loading and large ruptures, if left unmitigated for a long time, an inspection 

protocol and mitigation path forward should be established.

Section 2.2.11 References

[1] McEachern, R.J., and Taylor, P., “A Review of the Oxidation of Uranium Dioxide at 

Temperatures below 400°C,” JNM 254, 1998, pp. 87-121.

[2] Einziger, R.E. and Cook, J.A., “Behavior of Breached Light Water Reactor Spent 

Fuel Rods in Air and Inert Atmospheres at 229°C,” Nuclear Technology, Volume 

69, 1985, pp. 55-71.

[3] Hanson, B.D., “The Burnup Dependence of Light Water Reactor Spent Fuel 

Oxidation;” PNNL 11929, 1998.
79

2.3 Baskets and Neutron Absorbers Description

Fuel transport and storage canisters incorporate baskets of aluminum alloys, carbon 

steel, or stainless steel that provide geometric configuration control of assemblies within 

the canister [1, 2]. The primary function of the basket is to maintain spacing during wet 

loading and unloading operations and ensure fuel retrievability. The fuel basket also 

provides structural support for the canister and conduction pathways for decay heat 

dissipation.

Neutron poisons are added to the fuel canister to provide criticality control. These 

neutron poisons are absorber plates that are held in place by the fuel basket or are 

integrated into the metallic basket matrix material. Criticality control within fuel canisters 

is required only during canister loading and unloading operations conducted in reactor 

spent fuel pools [3] and during transportation in the event of an accident that could lead 

to flooding. During dry storage, the inert atmosphere of the canister provides insufficient 

moderation to support criticality and subcriticality is maintained by moderator exclusion.

There are several neutron poisons that are in use in licensed dry storage casks. These 

include the following:

Aluminum Boron Carbide Cermets: BORAL®

BORAL® is widely used as a neutron absorber in fuel pool storage racks and in some dry 

storage casks. “Boral® is made by mixing boron carbide granules and aluminum powder 

inside an aluminum box, heating the box and its content to form an “ingot”, and then hotrolling the “ingot” to form a plate consisting of a coarse core of B4C-Al composite 

material bonded between two thin sheets of aluminum cladding” [4]. The application of 

BORAL in basket designs is the predominant component for criticality control during fuel 

retrieval in basins. The BORAL® remains with the fuel in dry storage and is expected to 

be present for fuel retrieval, as necessary. BORAL® has excellent boron loading 

capabilities and is a very effective neutron absorber, but has very poor mechanical and 

thermal performance.

Metal matrix composites: Metamic®, Bortec®

Metamic® is a fully-dense, discontinuously-reinforced, metal matrix composite material. 

It consists of high-purity Type 6061 aluminum (Al 6061) alloy matrix reinforced with Type 

1 ASTM C-750 isotopically-graded boron carbide (B4C). The boron carbide loading can 

be varied and includes concentrations up to 33% by weight. Metamic® is used as a 

neutron absorber in wet and dry storage of used nuclear fuel. Metamic is weldable and 

has favorable mechanical properties up to 400°F (204°C), making it suitable as a 

structural material in fuel storage racks and in low heat emitting fuel baskets in dry 

storage casks [5]. The material is produced at essentially maximum theoretical density 

with no internal interconnected porosity.

BORTEC® MMC is produced in rolled or extruded shapes and in B4C contents up to 32% 

by weight using various 5000, 6000, and 1000 series aluminum alloys as matrix 

materials. The resulting composites have outstanding properties for structural or nonstructural design applications, and are lightweight and adaptable for neutron capture in 

both wet and dry environments. BORTEC® MMC is produced with no interconnected 

porosity and is used in high areal density neutron absorber applications.
80

Borated Aluminum Alloys: BorAluminum®

BorAluminum® alloys are true alloys developed by incorporating boron as a second 

phase in standard aluminum compositions using alloys 1100 and 6351. BorAluminum®

alloys are used in several spent fuel storage and transportation applications. Nonstructural applications typically use 1100 alloy + boron, while applications requiring 

structural alloys most frequently use 6351 + boron that maintains the strength and 

thermal conductivity of a standard alloy.

Borated Stainless Steels: NeutroSorb® Plus, Neutronit® 

NeutroSorb Plus is a sintered borated stainless steel material, typically ASTM A887

Types 304B3 through 304B6 [6] while Neutronit® is an austenitic stainless steel 

containing boron. Neutronit can be characterized as a 316 stainless steel with a boron 

addition. Both borated stainless steel alloys have good mechanical properties and 

fabrication capabilities, but are challenged by limited boron loading capability [7].

Section 2.3 References

[1] Electric Power Research Institute, report 1021048, “Industry Spent Fuel Storage 

Handbook,” July 2010.

[2] ASTM C-1562-10, Standard Guide for Evaluation of Materials used in Extended 

Service of Interim Spent Nuclear Fuel Dry Storage Systems, American Society for 

Testing and Materials, reapproved 2010.

[3] NUREG-75/110, "Safety Evaluation Report for Preliminary Design Approval for 

GESSAR-238, "Nuclear Island Standard Design," U.S. Nuclear Regulatory 

Commission, December 1975, pp. 4-14.

[4] Resolution of Generic Safety Issues: Issue 196: Boral Degradation (NUREG-0933, 

Main Report with Supplements 1–33).

[5] Personal communication, J. Russell, Sales and Marketing Manager, Holtec 

International, to M. E. Dupont, September, 2010.

[6] Pasupathi, V., Waste Package Neutron Absorber, Thermal Shunt, and Fill Gas 

Selection Report, B00000000-01717-2200-00227 REV 00, January 2000.

[7] White Paper on the Proposed Design, Development, and Implementation of a 

Monitored Retrievable Storage Module and the Siting Criteria for Spent Nuclear 

Fuel, B. Villarreal and D. Knobeloch, Las Alamos National Laboratory, LA-13001, 

UC-721 January 1996.

2.4 Baskets and Neutron Absorbers Degradation Phenomena

Degradation of the fuel baskets could impact the ability of the fuel baskets to maintain 

the geometry of fuel assemblies, neutron poison plates, and integral neutron absorbers. 

Deformation of the baskets could cause fuel assemblies to physically bind within the 

basket, thereby impacting retrievability of the spent fuel assemblies. Creep is one 
81

mechanism potentially leading to basket deformation. Additionally, should the storage 

canister become breached (see section 2.2.10 for example) air ingress could lead to 

oxidation of basket material. 

For metallic materials, creep generally only becomes important for homologous 

temperatures greater than 0.4Tm where Tm is the absolute melting temperature [1]. 

Though carbon steel and stainless steel would not be subject to creep at expected 

service temperatures and conditions, the temperature within the canisters may be above 

the temperature at which aluminum baskets experience creep.

Fuel baskets may experience limited corrosion associated with residual moisture within 

the canister as a result of incomplete drying or waterlogged rods. This residual water 

would be consumed very early in the storage period through corrosion processes, after 

which time, the corrosion would cease. Corrosion of the fuel baskets by corrosion 

induced by residual water is expected to be higher for the aluminum and carbon steel 

baskets than for stainless steel baskets. Because this is a phenomenon that is limited to 

the early storage period, there is no additional impact to basket structural integrity for 

very long-term storage by corrosion of fuel baskets. However, air ingress into a 

breached canister could lead to continued oxidation of the basket material. This may be 

particularly significant to the aluminum based basket material and, after prolonged 

exposure to elevated temperature air the aluminum in the assembly may be converted to 

aluminum oxide. Such conversion would degrade the mechanical properties of the 

aluminum based materials to the point that structural integrity could not be assured and 

the basket/poison plate material may lose dimensional stability. This would impact the 

retrievability of the fuel and could impact re-immersion of the DCSS, if needed.

Degradation of the fuel basket would affect its thermal characteristics. Thinning of the 

baskets walls by corrosion will decrease the cross section of the conductor and therefore 

decrease the heat conducted by the basket. Deformation of the basket material will 

increase the defect concentration in the material and decrease its thermal conductivity. 

Oxidation of the basket surface will affect the emissivity of the basket surface, thereby 

impacting the radiative heat transfer characteristics within the canister.

Neutron poisons that are fixed within the fuel basket, in conjunction with soluble neutron 

absorbers, provide criticality control during flooding associated with loading and potential 

unloading operations in a spent fuel pool. During the storage period, fixed neutron 

poisons are not necessary to maintain criticality control, except for their contribution to 

moderator exclusion. The degradation of fixed neutron poisons within the canister may 

specifically impact retrievability of spent fuel assemblies by physically impeding the 

removal of spent fuel assemblies from the canister or by redistribution of neutron 

absorbers within the canister leading to a decreased effectiveness as a neutron 

absorber during flooding. 

BORAL® is used extensively in dry spent fuel storage applications. However, there are 

indications that BORAL® could form blisters between the boron carbide matrix and the 

aluminum sheet cladding as a result of infiltration of spent fuel pool water into the porous 

boron carbide matrix. The blisters may likewise develop during vacuum drying of the 

casks by vaporizing water which had infiltrated into the plates or from corrosion during 

long-term storage at temperatures on the order of 500°F (260°C). Under these 

conditions, it is possible the material may crumble or otherwise relocate in storage, or 

may be physically damaged by reflooding associated with retrieval operations [2]. 
82

Similar degradation may result if a canister with BORAL became breached during EST. 

Data are needed to demonstrate that aluminum clad/boron carbide absorbers with core 

porosities between one and three-percent will not develop blistering that impacts the 

effective multiplication factor during cask reflooding associated with cask retrieval (or 

loading) operations [3] and during transportation in the event of an accident that could 

lead to flooding. Similar data are needed to assess the performance of BORAL in a 

breached dry storage system. The effect of the degradation on criticality control will 

depend on the physical condition and location of the neutron poison material, thus, 

calculations of the effects require significant post storage information. To eliminate the 

expenditure of the resources needed for the collection of such information, it is 

recommended that, fuel retrieval be achieved dry-to-dry or by flooding the canister with a 

fluid such as borated water.

The metal matrix composites, Metamic® and Bortec®, are not subject to blister formation 

like the aluminum clad/boron carbide absorbers, because these composites are 

produced with a homogenous material at effectively the maximum theoretical density, 

and with “no interconnected porosity in which corrosion can occur causing hydrogen 

generation, internal pressure buildup and deformation of the matrix” [4]. Likewise, the 

borated aluminum and stainless steel poison materials are not subject to the 

mechanisms of blister formation that have been observed in BORAL®. Because of the 

general lack of ductility characteristic of clad absorbers, similar to BORAL®, there is a 

need to evaluate the potential for embrittlement and cracking, and associated change in 

reactivity, of clad neutron absorber plates due to thermal cycling of the canister during 

the drying, storage, and retrievability period. However, Bortec should be evaluated for it 

performance in a breached canister.

Exposure of neutron absorber materials to neutron radiation generated within a spent 

fuel storage cask during the service life of the cask is expected to produce negligible 

effects on the efficacy of the neutron absorber. The flux of neutrons within the canister 

is decreasing with storage time and the total neutron fluence, even after extended 

storage, is insignificant compared to the number of absorber atoms in the neutron 

absorber material. The effect of extending the storage times will, therefore, not affect 

the susceptibility of the absorber materials to absorber isotope depletion or radiation 

induced embrittlement or cracking.

Assessment of Degradation, Impact to Safety Function – Excessive creep 

deformation of the aluminum basket and absorber casing materials would potentially 

impact the retrievability of the fuel from the canister as would excessive oxidation in a 

breached canister. Even though the temperature of the fuel basket would decrease with 

storage time, creep is viable for aluminum baskets and neutron absorbers due to 

relatively high temperatures expected in the fuel basket relative to the homologous 

temperature of 0.4Tm for creep in aluminum, or approximately 100°C. Additionally, the 

long term (100+ years) performance of aluminum alloys at anticipated EST temperatures 

is not known. Deformation and slump analyses would be needed on a structure-specific 

basis to evaluate and demonstrate that creep does not impair load-bearing components 

of aluminum fuel baskets to perform their safety functions during very long-term storage. 

Similar analyses are needed to assess performance in breached canisters.
83

Basket/Absorber Data Needs and Recommendations: The neutron absorbers should 

be evaluated for retrievability of fuel and efficacy of absorber material after water/dryout 

cycling and long-term exposure to air at anticipated storage temperatures. Deformation 

and slump analyses are routinely done in the initial SAR and need to be extended, on a 

structure-specific basis to demonstrate that creep and oxidation, time dependent 

processes, do not impair load-bearing components of aluminum fuel baskets to perform 

their safety functions during very long-term storage.

1. Testing & Analysis Activity: Because of the general lack of ductility characteristic 

of clad absorbers, similar to BORAL®, there is a need to evaluate the potential for 

embrittlement and cracking, and associated change in reactivity, of clad neutron 

absorber plates due to thermal cycling of the canister during the drying, storage, 

and retrievability period. The impact of long-term exposure to air should also be 

evaluated.

2. Testing & Analysis Activity: Evaluate basket-specific designs of aluminum 

baskets for slump over EST.

Basket degradation during EST and the impact of the water/dryout degradation for 

BORAL has been addressed, and the creep damage of aluminum basket materials is not 

anticipated to have a high negative impact on retrievability since engineering methods 

could be developed and implemented to free a slumped basket from a canister or 

remove fuel that may be stuck in a basket. It is not anticipated that reconfiguration of the 

basket would be severe to violate criticality safety during re-immersion in a pool, if 

necessary. However, the impact of long-term exposure to air at anticipated storage 

temperatures should be established.

Section 2.4 References

[1] Cadek, J., Creep in Metallic Materials, Elsevier Science Publishing Company, Inc., 

1988.

[2] Handbook on Neutron Absorber Materials for Spent Nuclear Fuel Applications, 

EPRI 1011818, 2005.

[3] Interim Staff Guidance 23 (ISG-23), “Application of ASTM Standard Practice 

C1671-07 When Performing Technical Reviews Of Spent Fuel Storage And 

Transportation Packaging Licensing Actions,” Division of Spent Fuel Storage and 

Transportation, U.S. Nuclear Regulatory Commission, D.C., June 2009.

[4] Qualification of METAMIC® For Spent-Fuel Storage Application, EPRI report 

number 1003137, Palo Alto, CA: 2001.

2.5 Canisters and Weld Description

A variety of DCSSs have been approved by the NRC for storage only, and, if approved, 

for both storage and transport, the DCSS is termed a dual-purpose dry storage system 

[1, 2]. The two basic configurations for storage canisters in a DCSS are shown in Figure 

32.
84

 

Figure 32 Typical Dry Cask Storage Systems. The cask may be an integrated 

cask with confinement and shielding in a integrated structure. The dry 

storage cask may include a separate canister (dual-purpose canister) 

over which an overpack is placed to provide shielding for storage or a 

cask is placed for transportation [http://www.nrc.gov/waste/spent-fuelstorage/diagram-typical-dry-cask-system.html].

The weld-sealed canister is the component of the DCSS that constitutes the confinement 

boundary and is essential for fuel handling and retrievability in the DCSS design with a 

canister and overpack that includes an air gap for passive cooling between the canister 

and overpack. The canister is sometimes termed a “multi-assembly sealed basket” or 

“multi-purpose canister” (MPC) in DCSS designs [2]. A cask with a metallic gaskets and 

bolted lid closure constitutes the confinement boundary in a DCSS design of an integral 

basket and cask.

For transportation, the shipping cask or overpack is credited with a defined leakage rate 

and is the primary confinement boundary. In some cases, the MPC is desired to provide 

moderator exclusion during transportation, and therefore it must continue to remain as a 

leak-tight component in the transportation package.

The canister (MPC) is typically weld-sealed as illustrated in Figure 33. Also, the canister 

is constructed from rolled plate, with a typical plate thickness of 0.5 inches (1.3 cm). The 

plate is rolled into a cylinder and seam welded. Two or more cylinders are stacked and 

joined to form the canister, and a bottom plate is welded at the canister base. 

The cask, as an integrated container with confinement and shielding, is typically closed 

with a bolted, gasket seal.

Several materials are used for construction of the canister and dry storage casks 

including [2]:

• Austenitic stainless steel

• Carbon steel (coated for corrosion protection)
85

• Ductile cast iron (coated for corrosion protection)

The shell and bottom plates are joined by welding and the joined plates may be one to 

several inches thick [2]. The carbon steel and ductile cast iron canisters use a bolted 

closure with metallic gasket seals. The austenitic stainless steel canisters are closed 

with either a double seal weld or bolted closure with two concentric gaskets.

Most of the dry storage canisters that are placed in concrete overpacks are fabricated 

from austenitic stainless steels [2]. In general, the canisters will be marginally protected 

from the environment but will be exposed to the storage site atmosphere. As discussed 

in Section 2.6, pitting and stress corrosion cracking are the only expected type of 

degradation anticipated in canisters during EST. 

Section 2.5 References

[1] Source: U.S. Nuclear Regulatory Commission, http://www.nrc.gov/waste/spentfuel-storage/designs.html.

[2] Industry Spent Fuel Storage Handbook, EPRI Report 1021048, July 2010.
86

Figure 33 Generic Canister Sketch Showing Basket Internals.

2.6 Canister Degradation Phenomena

The storage environment (external) for the canister in EST will include mildly elevated 

temperatures, ambient air, low irradiation doses, and the ‘industrialized” atmosphere of 

the storage site. A gamma field of ~105 R/hr and a neutron flux (En > 1 MeV) of 104

-106

n/cm2

-s [1] are typically assumed [2] for the irradiation field of a DCSS for extended 

storage, but the gamma field would be expected to decrease with age of the fuel and the 

increase with burnup. The radiation fields for the MOX fuel on a per fuel assembly (FA) 

basis, as discussed in section 2.1.5, would be above that for UO2 fuel. Irradiation from 

Closure Weld Sealed

Fuel Basket

∪ ∪ ∪ ∪ ∪ ∪ ∪

∪ ∪ ∪ ∪ ∪ ∪ ∪ ∪ ∪

Lift

Upper 

Fuel 

Spacer

Base Plate

Fuel in Basket with n-Absorbers

Fabrication Weld

∪

Lower Fuel Spacer
87

spontaneous fission neutrons and/or (α, n) neutrons will not be significant to impact the 

properties of the canister. The very low dose, even after several hundred years of 

exposure, is insignificant compared to dose levels needed to manifest changes to the 

mechanical properties at the low temperature of irradiation. Even then, high residual 

fracture toughness levels are observed in the canister materials [3]. 

The stainless steel canisters would not suffer significant radiation damage over EST. 

Figure 34 shows the dose dependence of the normalized (unirradiated/irradiated) 

fracture toughness for Base (B), Heat-Affected-Zone (H), and Weld (W) materials as 

obtained from Compact Tension (C(T)) specimens oriented with the notch/crack parallel 

(CL) or perpendicular (LC) to the axis of the large diameter pipe (rolling direction of the 

initial plate) source material. Figure 35 shows the irradiated fracture toughness of the 

materials in the form of J-T curves. The lowest toughness was in the heat-affected zone 

region when a crack ran parallel to the rolling direction of the original plate used to

fabricate the large diameter piping; this toughness remained high even up to dose levels 

much higher than anticipated in the canisters.

Thermally-induced changes in the canister material are also not plausible due to the low 

temperatures anticipated for EST.

Figure 34 Residual Fracture Toughness After Low Temperature (Tirr ~150°C) Neutron 

Irradiation on Type 304/304L Stainless Steel and Type 308 Stainless Steel 

Weld Metals [Figure reproduced from ref. 3].

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4 5 6 7

dpa

J1mm Irr/J1mm Non-Irr

BLC

BCL

HLC

HCL

WLC

WCL
88

Figure 35 Jdeformation-T Curves for Base (BLC, BCL), Weld (WLC, WCL), and HAZ (HLC, 

HCL) Components of Low-Temperature-Irradiated (Tirr ~150°C) Type 304 

Stainless Steel. The CL crack orientation is parallel to the rolling direction of 

the plates used to fabricate the stainless steel tank [Figure reproduced from 

ref. 3].

The fuel canister is generally exposed to the external environment at slightly elevated 

temperatures due to the heat generated by the stored fuel assemblies. The surface 

temperatures of the canister depend on the temperature of the surrounding environment, 

the canister heat load and the design of the overpack system providing for canister 

cooling.

The temperature-time profile illustrated in Figure 36 provides an example of possible 

conditions at the surface of a spent fuel storage canister and demonstrates that the 

surface temperatures are in the regime for corrosion induced degradation of the canister 

but are too low to cause significant metallurgical changes. However, canister surface 

temperature vs. time data for specific applications should be determined using actual 

canister heat loads and atmospheric conditions because, as discussed below, time –

temperature profiles will impact the tendency for corrosion. 

0

100

200

300

400

500

600

700

800

900

1000

0 100 200 300 400 500 600

T (Tearing Modulus)

J-Integral (kJ/m2

)

BLC 1BB8

BLC 1BB16

BLC 1BB9

BLC 1BB11

BCL 4BB10

BCL 5BA7

BCL 4BB11

BCL 4BB12

HLC 3HA5

HLC 3HB4

HCL 7HA5

HCL 7HA7

HCL 7HA6

WLC 2W2

WCL 7W4
89

Figure 36 Example of Canister Surface Temperature vs. Time 

[Figure reproduced from ref. 4].

Austenitic stainless steels are susceptible to pitting, crevice, and stress corrosion 

cracking under conditions anticipated during EST. The tendency for cracking when 

exposed to marine environments is dependent on salt chemistry and content, 

temperature, relative humidity and time of exposure [5, 6, and 7]. Additionally, chloride 

leaching from paints, adhesives, tapes and labels is known to cause cracking of such 

steels at room temperature [8].

Section 2.6 References

[1] Johnson, A.B. and Gilbert, E.R., “Technical Basis for Storage of Zircaloy-Clad 

Spent Fuel in Inert Gases,” PNL-4835, Pacific Northwest National Laboratory, 

1983.

[2] Einziger, R.E., Baldwin, D.L., and Pitman, S.G., “Data Needs for Long-Term Dry 

Storage of LWR Fuel,” Electric Power Research Institute report TR-108757, April 

1998.

[3] Sindelar, R.L., Lam, P.S., Duncan, A.J., Wiersma, B.J., Subramanian, K.H., and 

Elder, J.B., “Development and Application of Materials Properties for Flaw Stability 

Analysis in Extreme Environment Service,” invited paper in Engineering Safety, 

Applied Mechanics & NDE – A Symposium in honor of Dr. Spencer H. Bush, 

Proceedings of ASME Pressure Vessels and Piping Conference, San Antonio, TX, 

Paper No. PVP2007-26660, 2007.
90

[4] Climatic Corrosion Considerations for Independent Spent Fuel Storage 

Installations in Marine Environments, EPRI, Palo Alto, CA: 2006. 1013524.

[5] Caseres, L. and Mintz, T.S., “Atmospheric Stress Corrosion Cracking Susceptibility 

of Welded and Unwelded 304, 304L, and 316L Austenitic Stainless Steels 

Commonly Used for Dry Cask Storage Containers Exposed to Marine 

Environments,” NUREG/CR-7030, October 2010.

[6] Effects of Marine Environments on Stress Corrosion Cracking of Austenitic 

Stainless Steels, EPRI, Palo Alto, CA: 2005. 1011820.

[7] Carroll, S., Craig, L., Wolery, T., “Deliquescence of NaCl–NaNO3, KNO3–NaNO3, 

and NaCl–KNO3 salt mixtures from 90 to 120°C,” Geochem. Trans. 2004; 6(2):19.

[8] Miller, J.C. and Dupont, M.E., Labels, Gaskets, and Chloride Cracking of Austenitic 

Stainless Steels, J Fail. Anal. and Preven. (2011) 11:26–28.

2.6.1 Corrosion of Canister Materials

If the storage conditions were truly dry, corrosion, which is an electrochemical process, 

would neither degrade the materials nor compromise the intended safety function of a 

system or component. Corrosion can only occur in the presence of an electrolyte and in 

a dry storage system the electrolyte is anticipated to be moisture (water) that has 

condensed or fallen (rain) on the external surfaces of DCSS components. Initially, the 

materials selection processes were based on a license period of twenty years and the 

system design life, although not formally stated, was approximately fifty years. The 

radiation levels and temperatures of material exposure decrease with time and the initial 

exposure conditions were generally considered to be the most aggressive. Therefore, if 

no significant degradation had occurred during the initial storage period, extension of the 

“design life” and license period was thought to be relatively straightforward. This thought 

is basically reflected in ASTM C1562-10 [1] which states that “--- the demands on 

materials properties for an additional 20 to 80 years of storage may be reduced ---”. The 

review also concluded that “corrosion of metal components exposed to air and humid 

environments will occur.” The outer surfaces of the canister, with its welds, seals and 

attachments, therefore represent the area most likely to experience corrosion induced 

degradation during extended dry storage.

Section 2.6.1 References

[1] ASTM C-1562-10, Standard Guide for Evaluation of Materials used in Extended 

Service of Interim Spent Nuclear Fuel Dry Storage Systems, American Society for 

Testing and Materials, reapproved 2010.

[2] Technical Bases for Extended Dry Storage of Spent Nuclear Fuel, EPRI report 

number 1003416, Final Report, December 2002.
91

2.6.2 Stress Corrosion Cracking and Pitting of Austenitic Stainless Steel 

Canisters

The austenitic stainless steel canisters are anticipated to experience corrosion-induced 

degradation during long-term dry storage. Stress corrosion cracking at welds is 

probable, especially in canisters exposed to marine environments and both stress 

corrosion cracking and pitting are anticipated at non-weld regions. Additionally, if 

crevices that intersect the canister surface are present, crevice corrosion may occur.

Canisters constructed from austenitic stainless steels (Types 304 and 304L) will be 

susceptible to stress corrosion cracking at ambient temperatures in EST. Legend 

formally held that stress corrosion cracking of austenitic stainless steel did not occur at 

temperatures below 40°C (the legend was so strong that 1970’s versions of ASM 

Handbooks published this as a lower temperature limit for SCC) but experience has now 

demonstrated that cracking occurs at room temperature. Welded tanks, forged rings for 

climbing equipment, bolts and piping made of austenitic steels have cracked during 

room temperature service. 

Recently, a welded, Type 304L stainless steel coupon test as part of a qualification 

program for plutonium storage containers developed stress corrosion cracks when 

exposed, at room temperature, to salt mixtures that contained only about 0.6 wt% 

moisture [1]. The stress corrosion cracks initiated in the heat affected region of the weld 

and propagated along the HAZ and into the base metal (Figure 37). Such observations 

are clearly relevant to the welded canisters placed in dry storage.

 

Figure 37 SEM Micrographs of Stress Corrosion Crack in the Heat-Affected Zone in a 

Type 304L Stainless Steel Coupon with an Autogenous Weld. The specimen 

was exposed, at room temperature, to salt mixtures that contained about 

0.6 wt% moisture [SRNL Micrograph].

Stress corrosion cracking involves the synergistic action of three variables: a tensile 

stress, a susceptible material, and a semi-specific environment (chloride ions on 

austenitic stainless steel for example). Remove one of the three variables and cracking 

will not occur. 
92

The tensile stress necessary to cause stress corrosion cracking may be either residual 

or applied. The residual stresses developed by operations such as forming, welding, 

and grinding are more than sufficient to support cracking in austenitic stainless steel. 

This is illustrated in Figure 38, which shows multiple throughwall cracks in the bottom of 

a welded, stainless steel container. There were no applied stresses acting on the 

container during the cracking process. 

Figure 38 ThroughWall Stress Corrosion Cracks in a Welded Stainless Steel

Container not Subjected to Applied Stresses [SRNL Micrograph].

The chloride ions and the humidity necessary to cause cracking in a Type 304 stainless 

steel container in a DCSS may come from a variety of sources. The most apparent 

source of chlorides is a sea coast environment. The importance of humidity to the 

cracking process is apparent from studies [1] that demonstrated the necessity of 

deliquescence to the cracking process. These studies were conducted in sealed 

containers and showed that no cracking would occur in the absence of deliquescence. 

This observation is consistent with more recent results from in salt fog tests on U-bend 

specimens which showed that all the 304 and 304L stainless steels maintained at 43°C 

experienced cracking while all specimens tested at 85 and 120°C showed no SCC [2]. 

These observations, coupled with time-temperature profiles shown in Figure 36, suggest 

that the austenitic stainless steel dry storage canisters may be more susceptible to SCC 

during long term storage than during the initial storage period. The high surface 

temperatures during the initial storage period could prevent deliquescence of the salt 

and therefore prevent cracking of the canister.

The elements necessary to cause stress corrosion cracking to initiate on the external 

surfaces and propagate through an austenitic stainless steel component will be present 

during long-term dry storage and the tendency for cracking may increase with time. 

Furthermore, stress corrosion cracking is a nucleation and growth process, thus the 

absence of cracking at time t, provides no assurance that cracking will not occur at time 

t + one day. The potential for SCC to produce a throughwall penetration in a dry storage 

canister provides a potential for oxygen exchange with the helium environment and 

could adversely impact the performance of the fuel cladding. If the cladding has been 

breached, the fuel pellets would be exposed to oxygen and could be oxidized and swell.
93

Pitting and/or crevice corrosion could also occur on the canister if an electrolyte 

develops on the canister surface. Both pitting and crevice corrosion could lead to a 

throughwall breach in the canister.

Although the preceding discussion mentioned a sea coast environment as a potential 

source of chlorides, other stress corrosion cracking causing species may contact the 

external surface of the canister. Figure 39 shows a pit and an associated stress 

corrosion crack that resulted from fly ash and the associated chlorides depositing on the 

surface of a weld-closed stainless steel drum. Industrial pollutants including flyash are 

well recognized to have the potential to cause SCC. Although the applied stresses on 

dry storage canister walls are minimal, residual stress (fixed displacement stresses) are 

clearly present and likely to be driving force for the cracking process. This was the case 

for the cracks shown in Figure 39.

Figure 39 Pit and Stress Corrosion Crack due to Fly Ask Contact with 

Austenitic Stainless Steel [SRNL photograph].

The cracks shown in Figure 39 are typical of cracks driven by residual or fixed 

displacement stresses. Such cracks tend to be tight and when the crack propagated 

from the outside in, the tightness of the crack increases as the inner surface of the 

container is approached.

Conditions for Pitting/Cracking on Stainless Steel Canisters: Aggressive ionic 

species and water are required for cracking of the austenitic stainless steel canisters. It 

is assumed that surface tensile stresses are present due to weld residual stresses or 

fabrication-induced stresses and the material is known to be susceptible to SCC. 
94

Recently, key work has been conducted on conditions for stress corrosion cracking and 

pitting susceptibility of the canisters in a dry storage system [2]. Welded and unwelded, 

single and double U-bend specimens of Types 304, 304L, and 316L stainless steel were 

subjected to two types of environmental tests, salt spray and salt fog, using simulated 

sea salt. Complementary tests with NaCl and MgCl2 salt droplets were also conducted. 

All specimens exhibited pitting and cracking within a 2-month exposure in the salt spray 

tests that included heated specimens at temperatures of 25°, 93°, and 176°C. In the 

case of salt fog test, all specimens showed cracking at the specimen test temperature of 

43°C whereas specimens tested at 85 and 120°C showed no SCC, consistent with the 

inability of the salt to deliquesce.

Experimental and modeling work (high temperature Pitzer ion interaction thermodynamic 

data base model) were performed to evaluate salt deliquescence for repository dust 

assemblages [6]. Specific to the system, the modeling results predict that mixtures of 

NaCl, KNO3, and/or NaNO3 can show deliquescence when the relative humidity is as low 

as 20% at 160°C. Reversed deliquescence experiments were conducted in 

NaCl-NaNO3-H2O and KNO3-NaNO3-H2O systems at 90oC to 120°C to determine 

relative humidity and solution composition. The modeling results were in moderate 

disparity at high temperatures (120°C) with the experimental results – the model 

underestimated relative humidity by as much as 8% and solution composition by as 

much as 50% in the KNO3- NaNO3- H2O system.

Two important observations are made: 1) salt assemblages can have a lower 

deliquescence relative humidity than single-component systems, and 2) assemblages 

can form brines above 100°C and at standard atmospheric pressure of 1.01325 bar. 

Although the deliquescence relative humidity of most pure salt minerals is known over a 

range of temperature, the behavior of salt mixtures at elevated temperatures is largely 

unknown.

With an assumed presence of marine salts or industrial pollutants such as flyash on a 

canister surface during EST, it is concluded that the stainless steel canister is 

susceptible to pitting and cracking when canister temperatures are below 100°C and that 

cracking can occur at more elevated temperatures if exposure conditions allow 

deliquescence of the salts.

SCC Growth Rates – Salt Deposits: Work done by CRIEPI on SCC growth in marine 

environments is included in reference [5]. This work, coupled with thermal analyses of 

temperature profiles anticipated in canisters (Figure 40) indicate that the canister could 

be at temperatures that readily support SCC (at 100°C) within several decades of 

storage. Test results for 3-point bend specimens such SCC crack growth rates for salt 

deposit conditions (Figure 41) are high, approximately up to 10 cm/year. These data 

suggest that once cracking is initiated, the time for a crack to grow through the wall of a 

0.5 inch thick canister will be within a one year timeframe.
95

Figure 40 Temperature Profile on MPC [Figure reproduced from ref. 5].

Figure 41 SCC Growth Rate on Weldment Components of MPC 

[Figure reproduced from ref. 5].

Impact of Cracking on Confinement - Fixed displacement stress corrosion cracks can 

be so tight that in some instances liquids do not leak from the container, even when the 

crack is throughwall [5]. This observation suggests that radioactive or contaminated 

particulates would be unlikely to pass through a stress corrosion crack in a canister wall 

unless the crack were opened by mechanical forces such as might be imposed during 

handling and transportation. Additionally, such transport will generally require a fluid or 

gas stream be available for particulate transport. Therefore, the loss of confinement 

would be limited to gases and volatiles that would be in the volume space of the 
96

canister. However, as previously mentioned, oxygen ingress through a SCC could 

significantly alter the condition of the stored fuel.

Residual Stress Distribution on Large Diameter Canisters: The residual stress field 

of multipass metal inert gas welds used to fabricate canisters is expected to exhibit both 

throughwall and azimuthal variation. Typically, tensile stresses will be found on both the 

inside diameter and outside diameter of 0.5 inch plate that is joined with multipass 

welds. Sufficient stresses to initiate and propagate a stress corrosion crack are therefore 

assumed to be present on canister external surfaces. This assumption is consistent with 

industrial experience that includes SCC failures in systems with no (very small) applied 

stresses on the failed component.

Effects of Low-Temperature Neutron Irradiation on SCC: Previous studies on SCC 

in fully aqueous conditions were performed to support the low temperature production 

reactors at the Savannah River Site. Corrosion testing of sensitized Type 304SS, 

including Type 304SS samples that had been neutron-irradiated at low temperature 

were performed [7]: Constant extension rate tests (CERT) were conducted to study of 

environmental effects on intergranular stress corrosion cracking of Type 304 stainless 

steel. Specimens of both furnace sensitized and annealed material were irradiated to 

fluence levels of 1 to 2x1021 n/cm2 (E> 0.1 MeV) at 150°C in a beryllium reflector position 

of the High Flux Isotope Reactor. The CERT test results showed the same pattern of 

dependence on test variables (high importance of chloride ion and temperature) as the 

non-irradiated specimens in an associated study and the annealed and irradiated 

specimens showed no evidence of irradiation assisted stress corrosion cracking. These 

results demonstrate that the low fluence that the canisters will receive will not 

significantly change the SCC behavior and suggest that industrial experience with 

austenitic stainless steel systems is relevant to long-term canister behavior.

Experimental work to characterize stress corrosion cracking in water systems (reactor 

grade pure water from 40 to 100°C, Cl < 40 ppb) in the Savannah River Site production 

reactors piping system [8] showed stress corrosion crack growth rates that varied from 

10-4 to 10-7 in/hr (2.2 to 2.2 x 10-3 cm/yr). These results show that cracking can occur at 

very low chloride concentrations and suggest that stress corrosion cracking of the 

canisters is possible even in high quality water if water is pooled on the canister in 

regions where tensile stress are present at the pool-canister interface.

Assessment of Corrosion Degradation of Canister under EST: It has been 

demonstrated in laboratory and in field studies plus observed in field experience with 

actual and simulated components that the stress corrosion cracking (SCC) of austenitic 

stainless steel can occur at low temperatures (down to room temperature) in marine-type 

environments. As discussed in detail in reference [5], the SCC propensity of austenitic 

stainless steels in marine environments is affected by many variables including the 

amount of moisture and time of wetness, airborne contaminants (e.g., chlorides, sulfur 

dioxide, carbon dioxide, etc.), distance, elevation, orientation and shelter, temperature, 

sunlight and wind [9].

Austenitic stainless steels are very tough, fracture resistant materials. The toughness of 

the steel should be sufficient to minimize the tendency for mechanically-induced crack 

growth even during post storage handling and transport operations. Tabulated fracture 

toughness data are available for Type 304, 304L, 316 and 316L stainless steel and, if 

stress corrosion crack lengths in a canister were established by non-destructive 
97

inspections techniques, available fracture data could be used to establish safe handling 

and transport protocols.

ASME code cases N-513 and N-705 provide consensus methods to disposition planar 

and non-planar flaws (cracks) as would occur from SCC and pitting. These code cases 

also provide consensus crack growth rates for high (> 200°F (93.3°C)) and low (< 200°F 

(93.3°C)) temperature water to support flaw disposition in high purity water conditions. 

The growth rates observed in samples exposed to salt water condition appear to exceed 

by an order of magnitude those for similarly stressed samples (common crack tip stress 

intensities) exposed to high purity water at a common stress intensity loading. Additional 

data are needed to establish consensus crack growth rates for canister materials 

exposed to conditions anticipated during EST.

Assessment of SCC and Pitting in Canister in EST, Impact to Safety Function: The 

potential for stress corrosion crack and/or pit development presents several challenges 

to the safety functions of the DCSS: 

1. a loss in confinement function of the storage canister, 

2. the loss of retrievability of the fuel assemblies due to degradation of the fuel by 

oxygen ingress and the potential for structural weakness in a canister containing 

cracks,

3. a loss of confinement for transport of a cracked or pitted canister. 

The toughness of the steel, combined with the tightness of potential stress corrosion 

cracks suggest that non-destructive inspection techniques could safely be used to 

identify crack sites and that actions could be taken to mitigate further risks and safely 

recover the canister and contents. However, it must be recognized that the potential for 

stress corrosion cracking in canister walls clearly exists. The time to initiate SCC in the 

stainless steel is highly variable. Low temperatures of the canister would allow moisture 

condensation and would be increase the potential for SCC. Therefore, SCC is 

anticipated to be significant to canister integrity in EST. 

Reference 6 reports that deliquescence of a salt mixture of NaCl–NaNO3–KNO3 can 

occur at a 20% relative humidity at 160°C. This data is used to suggest that if occluded 

areas are established, sufficient chemistry (aggressive species and water) could form, 

even at temperatures well above 100°C, on the canister surface to lead to pitting and 

cracking. This condition could adversely impact the canister integrity and confinement 

function.

Canister Corrosion Degradation Data Needs for EST and General 

Recommendations for Resolution - Available data and analyses demonstrate a very 

real potential for stress corrosion cracking of the austenitic stainless steel canisters. 

Additional information concerning this potential is unlikely to provide practical 

technologies to mitigate or even minimize the tendency for cracking. Therefore any 

research to address stress corrosion cracking susceptibility should focus on developing 

1) technologies to determine (remotely if practical) the location and size of any cracks 

that may occur during dry storage, and 2) analytical methodologies to assure that 

identified and characterized cracks do not compromise the handling and transport of the 

canister system. Specifically,
98

1. Monitoring/In-Service Inspection Activity: Prescribe in-service inspection for 

canister and develop inspection methods, as practical for canister to monitoring 

for pitting and cracking. Develop acceptance criteria for disposition of detected 

flaws.

2. Testing & Analysis Activity: Develop structural and fracture methodologies with 

consensus for disposition of canister flaws in terms of acceptance leakage, 

acceptable flaw size, structural resistant to handling and DBE loadings and other 

anticipated activities. 

3. Testing & Analysis Activity: Develop methods to effectively flush surface 

contaminants to minimize potential for deleterious salt and debris buildup on 

canister surface.

4. Cask Demonstration Activity: As part of a cask demonstration, canister should be 

monitored for: 

a. buildup of surface debris including salts and particles (e.g., a “swipe” of 

the surface taken to evaluated particulate deposits);

b. canister surface temperature and relative humidity should be monitored;

c. NDE methods should adapted/developed to identify and characterize 

pitting and cracking, remotely through vents in the overpack, as 

practicable.

5. Maintenance Activity: Development of a wet or dry transfer system with protocols 

for canisters that develop flaws that need remediation

The distinct probability that SCC will occur in a canister during EST illustrates the 

importance of the above activities because 1) the canister provides confinement of the 

DCSS; 2) the canister is susceptible to pitting and cracking with raising of ambient 

relative humidity and capability for salts to deliquesce on the canister surface as 

temperature decreases; and 3) the necessity for remediation of a flawed canister will be 

adversely impact cost and safety in the management of DCSS in EST.

Section 2.6.2 References

[1] Zapp, P.E., et al “Relative Humidity and the Susceptibility of Austenitic Stainless 

Steel to Stress Corrosion Cracking in an Impure Plutonium Oxide Environment”, 

p.4 Vol. XXXVIII, Number 3, Journal of Nuclear Materials Management, 2010.

[2] Caseres, L. and Mintz, T.S., “Atmospheric Stress Corrosion Cracking Susceptibility 

of Welded and Unwelded 304, 304L, and 316L Austenitic Stainless Steels 

Commonly Used for Dry Cask Storage Containers Exposed to Marine 

Environments,” NUREG/CR-7030, October 2010.

[3] Mickalonis, J. I. and Dunn, K.A., “Residual Stresses in 3013 Containers” p. 31, Vol. 

XXXVIII, Number 2, Journal of Nuclear Materials Management, 2010.
99

[4] Understanding and Managing Aging of Material in Spent Fuel Storage Facilities, 

IAEA TRS443, 2006.

[5] Effects of Marine Environments on Stress Corrosion Cracking of Austenitic 

Stainless Steels, EPRI 1011820, 2005.

[6] Carroll, S., Craig, L., Wolery, T., “Deliquescence of NaCl–NaNO3, KNO3–NaNO3, 

and NaCl–KNO3 salt mixtures from 90 to 120°C,” Geochem. Trans. 2004; 6(2):19.

[7] Caskey, G.R., Ondrejcin, R.S., Aldred, P., Davis, R.B., and Wilson, S.A., “Effects of 

Irradiation on Intergranular Stress Corrosion Cracking of Type 304 Stainless Steel,” 

in NACE Corrosion ’90, April 1990.

[8] WSRC-RP-90-95, “Reactor Materials Program – A Derivation of a Nominal IGSCC 

Growth Rate for SRS Reactors (U),” Westinghouse Savannah River Company 

Report, March 1990.

[9] Climatic Corrosion Considerations for Independent Spent Fuel Storage 

Installations in Marine Environments EPRI 1013524, June 2006.

[10] ASME Code Case N-513, Evaluation Criteria for Temporary Acceptance of Flaws 

in Moderate Energy Class 2 or 3 Piping.

[11] ASME/BPVC CASE N-705: N-705 Evaluation Criteria for Temporary Acceptance of

Degradation in Moderate Energy Class 2 or 3 Vessels and Tanks Section XI, 

Division 1 Supp 4.

2.7 Bolted Closures Description (Bolts and gaskets)

A component of the confinement boundary for some canister designs or for casks that 

are an integrated unit includes bolted closures using metallic seals [1, 2, and 3]. A 

variety of metallic gasket materials, including Inconel X730, Al-jacketed Nimonic 90, and 

Al-jacketed Inconel X750, are used. Bolt materials include 304L SS, 193 Grade B7, 

320 Grade L43, SB637 Grade NO7718, SA564 Type 630, and H1150. Bolts can fail if 

given an improper heat treatment (temper embrittlement) however that event is not 

aging-related.

Loss of confinement could occur if either the sealing capability of the gaskets was 

compromised due to viscous flow or creep, or the bolt closure stresses were reduced by 

processes such as stress corrosion cracking. The potential for these degradation 

processes is discussed in the following section of the report.

Polymeric seals are used in DCSS to provide secondary confinement and leak detection 

capability, and are used in transportation packages to provide confinement. A variety of 

polymeric compounds are used as listed in reference [3].
100

Section 2.7 References

[1] Source: U.S. Nuclear Regulatory Commission, http://www.nrc.gov/waste/spentfuel-storage/designs.html.

[2] Industry Spent Fuel Storage Handbook, EPRI Report 1021048, July 2010.

[3] ASTM C-1562-10, Standard Guide for Evaluation of Materials used in Extended 

Service of Interim Spent Nuclear Fuel Dry Storage Systems, American Society for 

Testing and Materials, reapproved 2010.

2.8 Bolted Closures Degradation Phenomena

Metallic seals and bolts are anticipated to be subject to degradation in EST. Given that 

these components are SSC and part of the confinement boundary, it is important to 

assess the sealing capability of a system under EST.

2.8.1 Metallic Seal Degradation 

Metallic gaskets are an essential component in the confinement of a bolted closure, and 

are subject to aging [1]. The degradation of a closure seal could occur if the gasket 

relaxed to the point that no sealing forces were present on the closed joint. The ability to 

forecast century-long performance from 25,000-hour-long experiments is a difficult 

process thus accurate assessments of gasket behavior during EST are not available 

although several experiments are in progress.

As an example of an approach to demonstrate metallic seal efficacy over long-term 

storage, an experimental program to quantify the relaxation of four types of metallic 

Helicoflex® seals for their use in spent nuclear fuel storage casks is being conducted [2].

The Helicoflex® seals (see Figure 42 and Figure 43) are constructed of a helical spring, 

surrounded by an inner jacket of stainless steel and an outer jacket of a malleable 

material, such as aluminum or silver. These seals are capable of sealing to leakage 

rates of less than 10-7 atm cc s-1

.

Figure 42 General structure of ‘Helicoflex’ type metal seals 

[Figure reproduced from ref. 2].
101

Figure 43 Construction principles of HELICOFLEX® double jacket seals 

[Figure reproduced from ref. 2].

Two types of linings were studied: aluminum and silver [3]. Tests longer than 10,000 hr 

were implemented only for silver. For each type of lining, two different section diameters 

were investigated. A primary goal of the work [3] was to determine the residual linear 

load that can be guaranteed for a seal after a particular time of relaxation. This 

relaxation depends on the seal temperature and time. To test this process, seals were 

tightened between two flanges and exposed to constant temperatures, including 100 and 

200°C. Residual load and ‘useful’ recovery was measured after the exposure. Results 

were interpreted according to two methods: a time extrapolation, and a time–

temperature equivalence parameter. Both methods are based on linear relationships 

and were assessed through a statistical analysis (calculation of scatter) which is also 

used to determine a minimum guaranteed residual closure load. Finite element 

simulations of the relaxation of a seal were also performed in order to show that the time 

extrapolation method is appropriately conservative. 

For silver lining seals, the authors of the referenced report conclude that use of a time–

temperature equivalence parameter equal to T (11 + log10 (t)) is justified. This 

relationship allows one to assess the maximum temperature at which seals can be 

‘safely’ used. However, the use of this relationship assumes that a residual linear load 

of at least 100 N/mm will maintain a seal after closing the cask. This time/temperature

equivalence may not be a universal relationship for all materials and designs. Additional 

work is needed to validate this assumption or to determine bounding (minimal) residual 

loads required to maintain leaktightness.

Studies completed and on-going at BAM indicate that relaxation of the sealing force of 

metallic seals is expected to occur over time because of plastic deformation of the outer 

jacket of the seal. Greater deformation is expected at higher temperatures. The extent 

of the relaxation is determined by the construction of the seal and the temperature of the 

sealing surfaces. Table 4 shows the extent of relaxation which occurs in Helicoflex® 

seals with outer jackets of aluminum or silver at two different temperatures after 1 year. 

In addition to the potential problems with stress relaxation, the crevices formed by 

gasket joint are subject to crevice corrosion conditions including potential galvanic 

effects. Moisture is assumed to be able to condense and build up in the crevice of the 

joint. The kinetics of crevice corrosion are difficult to predict reliably but crevice 
102

corrosion is supported by the same environmental conditions that support SCC and as 

indicated in section 2.2.6 those conditions will exist on canisters during EST.

Table 4 Sealing Force after Temperature Exposure for 1 Year [3]

Reduction of initial sealing force after 1 year

(% of initial force remaining)

Seal jacket material 150 °C 20 °C (room temperature)

Aluminum, Al 39% 78%

Silver, Ag 51% 81%

Figure 44 shows an extrapolation of the 150 °C data indicates the decrease in the 

sealing force due to stress relaxation would be expected to continue leaving but even 

after 40 years approximately 22% (aluminum) or 33% (silver) of the initial sealing force 

will remain. Further extrapolation of these data indicate that at least 10,000 weeks 

(192 years) is required to reach the 100 N/mm threshold value assumed above. Linear 

extrapolation goes well beyond the 10,000 week point, and is closer to 

20,000-30,000 weeks (384-576 years). Linear extrapolation over this range is tenuous 

but indicates significant lifetimes are possible, perhaps up to and beyond 300 years if no 

significant changes in the relaxation mechanism are observed and crevice corrosion 

does not occur.

Figure 44 Extrapolation of seal force decrease over 40 years under constant 

conditions [Figure reproduced from ref. 3].

Elastic rebound or resiliency is another parameter that is suggested [3] to be linked to 

sealability. This parameter is also subject to Arrhenius aging behavior or property 

reduction thus these parameters also depend on temperature and time. 
103

Assessment of Metallic Seal Degradation – Impact to Safety: As a consequence of 

the continued reduction in sealing force during EST, replacement of the metallic seals 

may be required at some point to insure the integrity of the primary confinement 

boundary of the cask. Otherwise, undetected loss of the seal could lead to loss of the 

inert cover gas in the cask and to infiltration of outside atmosphere. Such infiltration 

could potentially lead to degradation of the fuel and release of radioactive material.

Metallic Gasket Degradation Data Needs for EST and General Recommendations 

for Resolution: Metallic seals should not be assumed to remain leak tight throughout 

the entire 300 years target for EST. Available data are simply not sufficient to support 

that assumption. However, the gaskets may continue to provide a sufficient barrier to 

leakage for multiple decades even after seal forces have been greatly reduced (e.g., due 

to viscous flow of the seal). It is recommended that a plan for replacement of the 

metallic seals for EST be developed and included as part of periodic maintenance. This 

could involve a re-wet or a dry system to effect seal replacement. Specifically,

1. Monitoring/In-Service Inspection Activity: Periodic surveillance for leakage at the 

seal joint would be needed under EST with acceptance criteria for leakage 

detection.

2. Testing & Analysis Activity: The metallic leakage tests in progress (BAM) should 

be continued to evaluate the sealing force at which leakage occurs and also 

evaluate the tendency for crevice corrosion degradation. The effect of 

temperature cycling, service environment (air, moisture, chlorides, etc.) and other 

possible relevant parameters should be accounted for in the testing. The 

response of aged/relaxed metallic seals under transient excursion conditions (if 

deemed possible, such as an extended abnormally cold period) should be 

considered. Depending on cask/seal design and actual seal temperatures, such 

excursions may or may not pose a concern.

3. Maintenance Activity: Development of a wet or dry system with protocols for 

replacement of seals that develop leakage or experience crevice corrosion and 

need to be replaced.

Metallic seal degradation during EST is likely to be significant due to the confinement 

function provided by metallic seals, and the continuing loss of sealing force with time. 

The potential for crevice corrosion also contributes to the significance. 

Section 2.8.1 References

[1] “Ageing of metallic gaskets for spent fuel casks: Century-long life forecast from 

25,000-h-long experiments”, H. Sassoulas, L. Morice, P. Caplain, C. Rouaud, 

L. Mirabel, F. Beal, Nuclear Engineering and Design 236 (2006) pp. 2411–2417.

[2] General structure of ‘Helicoflex’ type metal seals [ref. D. Wolff, et al, “Safety 

Aspects of Long-Term Dry Interim Storage of Type-B Spent Fuel and HLW 

Transport Casks” RAMTRANS Vol. 15, Nos.3–4, pp. 207–213 (2004).]

[3] Völzke, H. and Wolff, D., “Aspects on Long Term Storage of Used Nuclear Fuel 

and High Active Waste in Germany”, BAM Federal Institute for Materials Research 

and Testing, Presentation at PATRAM 2010, October 07, 2010.
104

2.8.2 Polymeric Seal Degradation

Elastomeric seals are not credited for confinement in storage, though they may be part 

of the transport/storage cask design for leak testing purposes. Therefore, degradation of 

elastomeric seals is considered a relatively low priority item from a DCSS confinement 

perspective for EST. Elastomeric seals are used in post-storage transportation, and are 

important to safety if the elastomeric seals have to provide a confinement function. In 

that case, it may be prudent or necessary to evaluate elastomeric seals used in the 

DCSS design to predict when such seals may no longer provide the regulatory function.

The primary degradation mechanism for elastomeric seals exposed to an oxygenbearing environment is oxidation and the oxidation processes are enhanced by elevated 

temperature and radiation. The general literature strongly suggests that elastomers 

used as seals have lifetimes of no more than several decades [1-8], and therefore would 

require periodic replacement for continuing integrity during EST, and as a minimum, 

replacement must be made prior to post-storage transportation where integrity must be 

assured.

The lifetime of elastomeric seals is dependent on many factors, including the seal 

material, storage conditions, the acceptance criterion used to define lifetime, design 

parameters and sensitivity to oxidation. Seals that only have to maintain minimal sealing

integrity at well-controlled ambient conditions may function for many years. For 

example, some studies indicate that it may be possible to use of EPDM seals from 

150 years to several hundred years at well-controlled conditions (25°C) [3]. Seals that

are exposed to higher temperatures or radiation doses will likely fail sooner. 

Accelerated-aging studies based solely on leak rate measurements or monitoring of a 

mechanical parameter such as compression set or compression stress-relaxation are 

useful for prediction purposes, but may overlook slower degradation mechanisms such 

as diffusion-limited oxidation. Oxygen consumption analysis and other techniques are 

therefore proving useful in providing more reliable lifetimes for elastomeric seals [5, 6].

Elastomeric seals are not credited for confinement directly during storage; however, 

there may be a need to credit such seals for post-storage transportation. Therefore, the 

aging of elastomeric seals may still be an issue to consider for the very long-term dry 

storage period. The length of time the seal needs to function and what acceptance 

criterion should be established (helium leaktight per ANSI N14.5 or other) are not welldefined. 

The seals may be credited with being leak tight defined per ANSI N14.5 as 1 x 10-7 ref 

cc/sec or 2 x 10-7 cc/sec He. This is the typical requirement for elastomeric seals in a 

Type B shipping package such as the Model 9975 [7]. A common approach in such 

designs is to use dual O-rings, with one seal credited for confinement and the other one 

for leak testing of the volume in between the two seals. For transportation, such 

packages are generally demonstrated to be leak tight prior to loading, then the packages 

are loaded and given a post-assembly verification test, often using a pressure drop 

criteria with 1 x 10-4 cc/sec sensitivity. For short-term transport purposes, it is not 

necessary or required to verify helium leaktightness during the shipping period as the 

seals are replaced during maintenance. However, if the shipping package is then used 

for interim storage, it may be difficult if not impossible to verify that the seals are still 

leaktight after assembly unless the system is repackaged. 
105

The basis for elastomer seal qualification in shipping packages or casks is typically to 

prove that a new seal can be leaktight at room temperature and perhaps down to low 

temperatures (such as -40 °C) that might be encountered during transport, depending 

upon requirements specific to the package. Leak tests after elevated temperature 

exposure might also be performed as part of the package qualification, but usually the 

aging time involved is quite limited. It is important to note that the technical basis for 

“continuous” service temperature limits often quoted by elastomer seal manufacturers is 

often very limited and is typically based on test results from relatively short aging periods 

(e.g.1000 hrs. or so). Validation of long-term seal performance, regardless of 

temperature, is usually left to the end-user to perform.

Many polymers, including most elastomers, also exhibit a significant difference in 

radiation and temperature resistance between oxidizing and non-oxidizing or vacuum 

environments, particularly if the radiation dose rates are high. There is a strong 

synergistic relationship between radiation, temperature, and oxygen that will impact the 

performance of any polymer. While this relationship varies in significance among 

polymer types, the general observation is that materials exposed to radiation and/or 

temperature in limited oxygen environments or vacuum exhibit significantly higher 

resistance to degradation, as defined by measurement of certain mechanical properties 

such as tensile strength and elongation than materials exposed to high oxygen 

concentrations. Such properties are usually measured for convenience and relative 

comparison, but are not necessarily the most relevant for all applications such as 

O-rings or gaskets. In addition, the relative change in some properties may be higher or 

lower than the changes observed in other properties, making complete evaluation of 

environmental resistance more complex. Therefore, critical properties should be 

evaluated based on the nature of the application when possible. Nevertheless, a review 

of the common elastomeric seals in a transportable storage cask concludes that “All 

elastomer seals are unacceptable for exposures of 108 rads” [8]. 

The general aging methodology is to assume Arrhenius-type behavior, which assumes a 

single activation energy for the degradation mechanism over the entire 

temperature/dose rate range anticipated in service. Though useful for evaluation 

purposes and general prediction of service life, many studies over the years have shown 

evidence of non-Arrhenius aging behavior, diffusion-limited oxidation effects and other 

factors that complicate the life prediction of elastomeric materials [1-9].

Using the general Arrhenius-type relationship for reaction rate as a function of 

temperature, lifetime to failure (of varying definition) can be extrapolated for varying 

temperatures. Figure 45 and Figure 46 show examples of the trend of loss of sealing 

force using time-temperature superposition principles. All experimental data at various 

aging temperatures are superposed into a “master” curve using shift factors that can be 

translated to other service temperatures of interest.

Among the limitations of these techniques are: 1) the use of a constant activation energy 

over the temperature range anticipated may not be valid, and 2) extreme extrapolation of 

higher temperature results to lower temperatures does not always provide an accurate 

result due to competing degradation mechanisms. This is particularly true for materials 

experimentally aged at temperatures much higher than anticipated in service.
106

Due to the synergy of all competing mechanisms, variation in compound formulation, 

and specific service conditions, accelerated aging can never completely duplicate true 

long-term aging data in actual service. A useful approach is to periodically leak test 

sealed mock or surrogate vessels that are aged at relevant temperatures for the storage 

environment [7, 9-11].

The correlation between leak tightness and mechanical parameters such as elongation, 

compression set or even compression stress-relaxation (CSR) is not well-defined and 

likely varies with each design. Even with the same seal compound, some designs may 

be more or less sensitive to seal material property changes than other designs.
107

Figure 45 Time-Temperature Superposition of Compression Stress Relaxation 

Data for Viton GLT O-rings (Parker V0835-75) at 80 °C 

[Figure reproduced from ref. 7].

Figure 46 Time-Temperature Superposition of CSR data for EPDM seals 

[Figure reproduced from ref. 2].

0

0.2

0.4

0.6

0.8

1

1.2

1 10 100 1000 10000 100000 1000000

Time (Hours)

Retained Sealing Force 

175

235

250

300

350
108

Assessment of Polymer Seal Degradation – Impact to Safety: The function of 

sealing by elastomeric (polymeric) seals is typically not credited for long-term 

confinement in DCSS designs due to known active degradation. They are used as 

secondary seals to provide a leak check of the primary seal, and thus are not credited 

directly with a safety function.

Most polymeric materials are subject to degradation of their physical properties by 

oxidation, radiation, heat and/or hysteresis effects. However, the metrics of 

compression set and compression stress relaxation (CSR) are often used for predicting 

the service life and sealing capability of the elastomers. Oxygen consumption analysis 

and dynamic mechanical analysis can also be useful, particularly to identify possible 

non-Arrhenius behavior. Such metrics do not provide direct time to leakage failure, as 

leak rate depends on many design parameters, but do provide insight into long-term 

material performance.

No requisite data needs for EST are established. It is assumed that polymeric seals will 

not meet leak tightness for desired service after, at most several decades in DCSS 

service. Additional research may be warranted to predict seal lifetime in order to 

maximize seal longevity and minimize seal replacement efforts, during storage and for 

post-storage transportation. When seals are replaced (either due to degradation or at a 

prescribed maintenance period), it is recommended that the removed seals be retained 

for further analysis and characterization. Such data may be useful for other nuclear 

materials packaging applications and non-nuclear applications where the same 

elastomer compounds may be used.

Section 2.8.2 References

[1] “Validation of Improved Methods for Predicting Long-Term Elastomeric Seal 

Lifetimes from Compression Stress-Relaxation and Oxygen Consumption 

Techniques”, Polymer Degradation and Stability 82 (2003), K.T. Gillen, M. Celina, 

R. Bernstein.

[2] SAND98-1942, New Methods for Predicting Lifetimes in Weapons, Part I: 

Ultrasensitive Oxygen Consumption Measurements to Predict the Lifetime of 

EPDM O-Rings, K.T. Gillen, R.L. Clough, G.M. Malone, M. R. Keenan, J. Wise, 

Sandia National Laboratory, 1998.

[3] SAND-97-2181C, “Evidence that Arrhenius High-Temperature Aging Behavior for 

an EPDM O-Ring Does Not Extrapolate to Lower Temperatures”, K. T. Gillen, 

J. Wise, M. Celina, R.L. Clough, 1997.

[4] “Predicting and Confirming the Lifetime of O-Rings”, Polymer Degradation and 

Stability 87 (2005), K.T. Gillen, R. Bernstein, M.H. Wilson.

[5] “An Ultrasensitive Technique for Testing the Arrhenius Extrapolation Assumption 

for Thermally Aged Elastomers”, Polymer Degradation and Stability 49 (1995), 

J. Wise, K. T. Gillen & R. L. Clough.

[6] “Predicting and confirming the lifetime of O-rings”, K.T. Gillen, R. Bernstein, M.H. 

Wilson, Polymer Degradation and Stability 87 (2005) pp. 257-270.
109

[7] “LONG-TERM LEAK TIGHTNESS OF O-RING SEALS IN THE 9975 SHIPPING 

PACKAGE”, E.N. Hoffman, T. E. Skidmore, W.L. Daugherty, Proceedings of the 

ASME 2010 Pressure Vessels & Piping Conference, July 18-22, 2010, Bellevue, 

Washington, PVP 2010-25106.

[8] “Considerations Applicable to the Transportability of a Transportable Storage Cask 

at the End of the Storage Period,” T.L. Sanders, C.A. Ottlinger, J.L. Brimhall, J.M. 

Creer, E.R. Gilbert, R.H. Jones, P.E. McConnel, SAND88-2481, TTC-0841, UC820, November 1991.

[9] AGING PERFORMANCE OF VITON® GLT O-RINGS IN RADIOACTIVE 

MATERIAL PACKAGES, T.E. Skidmore, K.M. Counts, E.B. Fox, E.N. Hoffman, 

K.A. Dunn, Proceedings of PVP2007, July 22-26, 2007, San Antonio, Texas.

[10] “Seal life of EPDM O-rings at high temperature determined by unique method”, 

R. Marlier, R. Andre, P. Malesys and H. Issard, Packaging, Transport, Storage & 

Security of Radioactive Materials, 2006, VOL. 17 NO 1.

[11] “First tests results for determination of seal life of EPDM O-rings at high 

temperature (determined by unique method)”, Packaging, Transport, Storage & 

Security of Radioactive Material, 2010, VOL. 21 NO 1.

2.8.3 Bolt Degradation

Carbon and low alloy steel bolts are subject to general corrosion and creep in EST. With 

time, bolt failure can occur, and if sequential (Figure 47), a loss of gasket compression 

and leak tightness would be expected. A failure of one or several bolts may not cause a 

sealing (confinement) failure unless the failed bolts are adjacent to each other in the bolt 

pattern. 

Bolting is considered a maintenance item that would be replaced. Corrosion and stresscracking are common degradation phenomena in bolts, and a schedule for bolt 

replacement is suggested for aging management. Bolts would not be expected to 

remain viable for EST.
110

Figure 47 Loss of Sealing Force Can Occur with Multiple Adjacent Bolt Breakage.

Bolt Degradation Data Needs for EST and General Recommendations for 

Resolution: Bolts are not expected to remain capable of providing a sealing force 

throughout the entire 300 years target for EST. Single bolt failures are not expected to 

cause a significant loss of sealing capacity of the gasket/bolt joint. However, multiple 

sequential bolt failures would be expected to enable a leakage path and loss of 

confinement. It is recommended that a plan for periodic monitoring of the bolts (e.g., 

check torque) and/or bolt replacement be performed for EST as part of periodic 

maintenance. Specifically,

1. Monitoring/In-Service Inspection Activity: Periodic surveillance for leakage at the 

seal joint would be needed under EST with acceptance criteria for leakage 

detection.

2. Testing & Analysis Activity: Analysis of cask to determine the impact of sequence 

of bolt failure on leak tightness of gasket/bolt/cask system.

3. Maintenance Activity: Development plan for periodic bolt replacement.

The need for bolt replacement during EST is high due to the confinement function 

provided by metallic seal/bolt system, and the potential for bolt degradation and 

breakage to go undetected unless actively checked.
111

2.9 Casks Description

The casks of a licensed DCSS may be of several different designs [1, 2, and 3]. Figure 

48 shows a generic overpack on a canister that comprises the cask. Other cask designs 

are an integral container with shielding rather than the overpack. Cask materials are 

listed in the table in Annex 3 of reference [4]. 

Casks bodies are made of ferritic or ferritic-martensitic steel; stainless steel; cast iron 

and graphite. Some casks contain lead that, as a component of the system, provides 

shielding, confinement, and thermal performance functions.

Casks can be made of concrete, with or without reinforcement and with or without steel 

shells and lead. These casks provide a shielding and thermal performance functions 

and confinement is provided by the stainless steel canisters within the cask

Casks may also contain components that are layers of neutron absorber and/or gamma 

shields as part of the cask structure. Such layers provide an important safety function. 

Radiation shielding is incorporated into dry storage systems as a means of providing 

integral personnel protection. For example, polyethylene or other low Z material is often 

placed in a cask as a neutron shield. Similarly, gamma shielding is provided in dry 

storage systems, either as an add-on material (e.g., lead), or by appropriate selection of 

the dry storage system construction materials. For example, heavy steel or concrete 

casks, by their very nature, provide a measure of radiation shielding.

A variety of designs are used for neutron shielding [1]. Neutron shielding in the 

Transnuclear Inc. TN-40 cask is provided by a resin compound surrounding the body. 

The resin compound is enclosed in long, slender aluminum containers. The array of 

resin filled containers is enclosed within a smooth outer steel shell. A disk of 

polypropylene is attached to the cask lid to provide neutron shielding during storage. 

Polyethylene shielding rods are used as a neutron shield in CASTOR-V/21 PWR spent 

fuel storage cask. For the NAC-STC, gamma shielding is provided by a lead wall and 

neutron shielding by a layer of a solid borated synthetic polymer, which surrounds the 

outer shell along the cavity region.

The Westinghouse MC-10 cask body is made of low-alloy steel and has forged steel 

walls and bottom that provide radiation (gamma) shielding and structural integrity. 

Neutron shielding is provided by a layer of BISCo NS-3 cured in the cavity between the 

cask wall and outer protective skin. In concrete dry storage systems (casks and vaults), 

neutron and gamma radiation shielding are provided, in part, by the concrete structure. 
112

Figure 48 Generic Dry Storage Cask Sketch for an Overpack on a Canister.

Shield

Shield

Outlet

Ai

Sealed

Canister

Shield (concrete, steel, and/or polymers) 

Shield

Cooling 

Air Inlet

Inner

Shell

Wall

Outer 

Shell

Wall
113

Section 2.9 References

[1] Source: U.S. Nuclear Regulatory Commission, http://www.nrc.gov/waste/spentfuel-storage/designs.html.

[2] Industry Spent Fuel Storage Handbook, EPRI Report 1021048, July 2010.

[3] Data Needs for Long-Term Dry Storage of LWR Fuel, EPRI report number 

TR-108757, Interim Report, April 1998.

[4] ASTM C-1562-10, Standard Guide for Evaluation of Materials used in Extended 

Service of Interim Spent Nuclear Fuel Dry Storage Systems, American Society for 

Testing and Materials, reapproved 2010.

[5] IAEA TECDOC 1100, Survey of Wet and Dry Spent Fuel Storage, July 1999. 

2.10 Casks Degradation Phenomena

Excessive degradation of the materials of the casks could lead to failure to provide a 

shielding function, structural function, loss of environmental control and/or loss of 

configuration control for the fuel storage.

2.10.1 Carbon Steel Material Corrosion

Non-stainless metallic systems may be subject to corrosion over EST. The metal shell 

for thick-walled monolithic metal canister systems provides gamma radiation shielding 

which is provided by lead in multi-shelled metal casks and by the concrete in thin-walled 

metal casks. Carbon steel will require periodic maintenance to avoid general and pitting 

corrosion. This is a low priority data need due to the ease of detection and 

maintenance.

2.10.2 Neutron Shielding Degradation Phenomena

Polymeric materials and resins are used in transport and storage casks as neutron 

shields due to their high hydrogen contents. These materials are subject to degradation 

and possible relocation within the casks as a result of thermal and radiation exposures, 

as well as the passage of time. Chain scission can result in a reduction in molecular 

weight and an increased tendency of the polymer to creep. Crosslinking can increase 

the molecular weight of the polymer, increasing its hardness and resistance to creep. 

These radiation effects are sensitive to the applied dose, the dose rate and the presence 

of oxidants and other reactive species that make polymer performance application 

specific.

Polymeric materials and resins used for neutron shielding are subject to degradation and 

possible relocation as a result of thermal and radiation exposures, as well as the 

passage of time. Loss of shielding will be detectable during required periodic radiation 

monitoring and must be addressed as part of the life management of the dry storage 

facility. Materials for neutron shielding include polyethylene, polypropylene, and borated

polymer. During very long-term storage, these shielding materials are subject to 

thermal, radiation, and corrosion degradation phenomena that may lead to physical 
114

changes in the materials that may impair the ability of the neutrons shield to perform its 

intended safety function.

Periodic survey of radiation fields from the cask should be performed and administrative 

controls could be readily implemented if a loss of shielding was detected. This is a low 

priority item, especially since the radiation fields in a cask will reduce with time during 

EST.

2.11 Concrete Pad Description

A concrete pad is used as the foundation to stage the casks in the DCSS. The concrete 

pad is not generally within the scope of the materials degradation analysis for present 

licensing consideration unless the pad provides a safety function during a seismic event 

[1]. Several mechanisms for concrete degradation described below in section 2.12 of 

this report are anticipated to be significant to cause loss of compressive strength and/or 

reinforcement bar corrosion leading to spallation and further loss of material section over 

a very long term exposure period. These degraded conditions of the pad may impact 

the stability of the foundation for the casks under normal conditions. Additionally, 

concrete is also used in overpacks and is susceptible to the degradation mechanisms 

addressed.

Reference Concrete Pad Design: The concrete pad design and construction is 

assumed to be in accordance with the specifications and codes of the American 

Concrete Institute, ACI 318 [2].

Placement of concrete is important in its susceptibility to degradation. It is important to 

prepare concrete with 1) the correct positioning of the steel reinforcement, 2) an 

appropriate water/cement ratio and the water chemistry being low in detrimental 

constituents such as chloride species, 3) the correct use of aggregates and 4) other 

considerations in weathering regions. It is assumed that the as-fabricated condition of 

the pad provides for high-quality concrete in its initial condition. Guidance to assure a 

dense and well-cured concrete is provided in references 2-4.

The concrete is considered to have a compressive strength of 3000 psi. The 

reinforcement bar is considered to be #3 (0.375 inch (0.953cm) diameter) with a yield 

strength of 60 ksi. A cover of concrete of 1 inch is assumed over the reinforcement bar 

for the purpose of this assessment of degradation.

Section 2.11 References

[1] Section 2.4.3 of NUREG-1927, Standard Review Plan for Renewal of Independent 

Spent Fuel Storage Installation Licenses and Dry Cask Storage System 

Certificates of Compliance. http://www.nrc.gov/reading-rm/doccollections/nuregs/staff/sr1927/.

[2] ACI-318-08, Building Code Requirements for Structural Concrete and 

Commentary, American Concrete Institute, 2008.

[3] ACI-301-10, Specifications for Structural Concrete, American Concrete Institute, 

2010.
115

[4] ACI 201.2R-08, Guide to Durable Concrete, American Concrete Institute, 2008.

2.12 Concrete Pad Degradation Phenomena

The physical properties and/or physical condition of concrete can degrade and impact 

the ability of concrete structures to perform their desired function. The modeling of 

concrete degradation that occurs during the EST exposure requires a coupled chemicalphysical model to accurately describe the evolution of the degrading concrete. For 

example, the degrading cement phase in concrete may involve concomitant features 

including micropore and microcrack evolution, and transport/equilibrium of multiple 

species including species ingress (e.g.,O2, CO2, Cl-

) and leaching (e.g., Ca2+) may even 

include further hydration of unreacted cement constituents.

Several of the phenomena involve corrosion of the reinforcement bar. Typically, the 

highly alkaline (pH > 13) nature of the cement provides an environment around 

embedded steel and steel reinforcement, which protects them from corrosion [1, 2]. If 

the pH is lowered (for example, pH < 10) due to carbonation, intrusion of acidic materials 

through microcracks, and/or leaching of alkaline products through cracks, corrosion of 

the reinforcement bar may occur causing high internal stresses due to volume mismatch 

between the steel rebar removed (corroded) and the corrosion products formed as 

shown in Figure 49. The high stresses can lead to extensive cracking of the concrete 

and degradation such as that shown in Figure 50.

Figure 49 Relative volume of possible rebar corrosion products 

[Figure reproduced from ref. 3].
116

Figure 50 Spallation of Concrete Induced by Rebar Corrosion 

[SRNL Photograph].

Loss of strength may occur with leaching of species (e.g., Ca) that increase porosity, or 

by ingress of other species (e.g., sulfates) that cause alteration of calcium-silicatehydrate (CSH) and calcium hydroxide (CH) paste and replacement with a weak paste 

[2].

The degradation mechanisms relevant to concrete pads and concrete casks in a DCSS 

are described, and degradation models used for estimation of the condition of the 

concrete with service exposure and time are listed. An assessment of the significance of 

degradation in terms of affecting the physical properties or physical condition of the 

concrete over EST is made with regard to a single operative mechanism (e.g., CO2

diffusion) unless noted otherwise. An assessment of the potential impact to the safety 

functions of the DCSS concrete pad or concrete cask from the mechanism is provided. 

Ideally, the assessment would involve the use of a coupled-phenomena model but a 

model applicable to the times involved in extended storage is not currently available.

Resolution of concrete degradation issues will primarily be through monitoring/in-service 

inspection.

Section 2.12 References

[1] Troxell, G. E., Davis, H. E., and Kelly, J. W., Composition and Properties of 

Concrete, Second Edition (J. W. Kelly), McGraw-Hill, 1968.

[2] Walton, J. C. et al. “Models for Estimation of Service Life of Concrete Barriers in 

Low-Level Radioactive Waste Disposal,” NUREG/CR-5542, July, 1990.

[3] Figure from page 159 of Pierve R. Roberge, Handbook of Corrosion Engineering, 

McGraw Hill Professional, 2000.
117

2.12.1 Carbonation of Concrete

Carbonation or the ingress of CO2 into the concrete to alter its chemistry can lead to 

concrete degradation by reducing the high alkalinity of the pore water in the material in 

which reinforcement bars are embedded. Subsequent oxygen ingress into the pore 

water feeds the corrosion of the steel and the local expansion due to formation of 

corrosion products can cause large splitting (cracking > ~ 1 mm opening) exposing 

additional rebar to the ambient external air/water environment with eventual spallation of 

the concrete cover. In as-prepared concrete, the remaining water in the pores is at a 

high pH (above 13) from residual solutions saturated in Ca, K, and/or Na hydroxides. 

Reaction of CO2 from the atmosphere with cement can occur to form calcium carbonate 

and lower pH from basic to neutral so that reinforcing steel becomes susceptible to 

corrosion [1]. A summary description of carbonation processes in concrete is given in 

Reference 1 and is consistent with the experiences at the Savannah River Site. 

Mechanistically-based models show that both time and CO2 atmospheric concentration 

increase the depth of carbonation in concrete. The depth of carbonation is proportional 

to the square root of time. 

X A t C = [5]

Xc is the distance from the external surface of the concrete to the carbonation front, t is 

the time and A is a material constant. Papadakis et al. [2] developed an expression for 

A from theoretical basis.

[ ]

[ ( ) ( )] [ ] 0 0

2

0

2

3

2 2

Ca OH s CSH

CO D

A eCO

+ =

[6]

DeCO2 is the effective diffusivity of CO2 in the concrete, while [CO2]

0

, [Ca(OH)2(s)]0 , 

where Ca(OH)2 is the solid calcium hydroxide, and [CSH]0

, where CSH is the gelatinous 

calcium silicate hydrate, are the molar concentrations in the concrete. 

The effective diffusivity of CO2 is dependent on pore size. Pore size can be affected by 

several factors including the carbonation reaction itself, as well as moisture, and it must 

be accounted for in the relationship.

With the assumptions stated above, the carbonation depth can be predicted based on 

the known age and environmental conditions in service. Figure 51 shows one example 

of this for concrete exposed in indoor service conditions with data for two cases (w/c 

ratios of 0.28 and 0.52) from reference 2, and from Savannah River Site (SRS) data 

from the P-reactor building. For 10 years of service, the carbonation depths vary 

between 0.5 and 1 cm for the 2 cases shown. 
118

0.00

0.20

0.40

0.60

0.80

1.00

1.20

1.40

1.60

1.80

0 2 4 6 8

square root time (sqrt years)

Carbonation Depth (inches)

R w/c =0.52

P reactor

R w/c =0.48

Linear (R w/c =0.52)

10 yr 25 yr 50 yr

Figure 51 Carbonation depth with time for normal atmospheric CO2 exposure 

indoors of P-reactor building at the Savannah River Site plotted with 

experimental data by Nagataki et al. [Figure reproduced from ref. 3].

The depth of carbonation can be measured with a simple technique using a concrete 

core removed from the structure. When a 1% phenolphthalein solution is applied to 

normal concrete it will turn bright pink. If the concrete has undergone carbonation, no 

color change will be observed. An example of this is shown in Figure 52. This concrete 

core was removed from the P-reactor building, an above-ground facility at SRS after 

approximately 60 years of service. Extrapolation of the linear fit of the literature data [3] 

to 60 years of service would predict a lower carbonation depth than was actually 

measured in the sample from the inside wall from the SRS P-reactor as shown in Figure 

51. The difference could be attributed to different starting materials or environmental 

conditions. 
119

Figure 52 Cross-section of a 3.7 inch (9.4cm) diameter concrete core taken from the 

inside wall of the SRS P-reactor building. A phenolphthalein solution has 

been applied to indicate the extent of concrete carbonation. Lack of color 

indicates a pH below 10.2 showing that carbonation has progressed 

~ 1.4 inches (3.6cm) in from the surface of the inner wall. 

[Figure reproduced from ref. 10].

Cracking in concrete can lead to a short circuit in the depth of carbonation. Figure 53

shows carbonation from the faces of a microcrack in concrete that is deeper than the 

average carbonation depth from the surface exposed to atmosphere. 

Rebar
120

Figure 53 Tight crack (see arrows) from surface (bottom) to interior of concrete 

providing a pathway for enhanced diffusion. Evidence of enhanced 

diffusion of CO2 adjacent to crack is observed. 

[Figure reproduced from ref. 10].

Assessment of Degradation of Concrete Due to Carbonation under Extended 

Storage: Carbonation is expected to occur in concrete and penetrate to depths of the 

reinforcement steel well within the exposure time of 300 years. For example, the data in 

Figure 51 show carbonation to a depth of one inch within 50 years. The carbonation of 

concrete can adversely impact concrete integrity by creating a low pH environment at 

the rebar. In conjunction with oxygen and pore water, this condition promotes 

reinforcement bar corrosion [1, 3, and 4]. Oxygen diffusivity is greater than carbon 

dioxide diffusivity in cement [5], and the onset of general corrosion of the reinforcement 

bar would be controlled by carbonation. However, the rate of corrosion would be 

controlled by oxygen diffusion to the reinforcement bar. Both CO2 and oxygen diffusion 

can be short-circuited by microcracks in the concrete. 

With the increase in volume of the oxide product from that of the metal removed, tensile 

stresses, cracking and the opening of larger (~ 0.3 to 1 mm) cracks [6-8] expose 

additional rebar to the ambient external air/water environment leading to additional 

corrosion and cracking. Eventually, spallation of the concrete cover would be expected.

As stated, the depth of carbonation increases parabolically with time and CO2 level. 

Atmospheric concentration of CO2 is projected to rise from 380 ppm to several thousand 

ppm within a hundred years [9] and that increase would increase the rate of carbonation. 

Although the rate of carbonation would be concrete-condition-specific and environmentspecific (e.g., low RH conditions tend to increase carbonation rate), the time for 

carbonation to depths of the reinforcement bar should be well-within a 300-year period.
121

Assessment of Concrete DCSS Structures with Carbonation - Impact to Safety 

Function: There is a potential for extensive cracking of a concrete pad from 

carbonation and rebar corrosion. The ultimate result of the corrosion will be spallation of 

the concrete cover and could result in an unstable foundation for a cask. This would 

create a cask tip-over hazard with a loss of its thermal cooling configuration and its 

shielding function. Additionally retrieval of the casks for transportation would also be 

impacted with an unstable pad, even if tip over did not occur.

There is a potential that extensive cracking of a concrete cask containing reinforcement 

bar but without an encasement shell could result in concrete spallation with a loss of net 

section and shielding function. Concrete casks without reinforcement bars would not be 

subjected to carbonation-induced spallation damage.

Concrete Degradation by Carbonation, Data Needs for EST and General 

Recommendations for Resolution: The concrete in the pad and cask is subject to 

degradation by carbonation throughout EST. There is presently no program for 

monitoring carbonation and technologies to support a demonstration that the pH 

condition at the embedded reinforcement bars in the pad and shell is benign to corrosion 

is needed. Specifically, 

1. Monitoring/In-Service Inspection Activity: Develop monitoring/in-service 

inspection protocol for carbonation with acceptance criteria.

2. Testing & Analysis Activity: Investigate new technologies for NDE for carbonation 

in concrete. 

3. Cask Demonstration Activity: As part of a cask demonstration, establish baseline 

and periodically monitor for carbonation and other concrete degradation 

mechanisms. 

Carbonation degradation of the concrete is important because: 1) carbonation is 

expected throughout EST; 2) aging management for carbonation is not established; and 

3) remediation of carbonation degradation would be costly.

Section 2.12.1 References

[1] Walton, J.C. et al. “Models for Estimation of Service Life of Concrete Barriers in 

Low-Level Radioactive Waste Disposal,” NUREG/CR-5542, July, 1990.

[2] Papadakis, V.G.,“A Reaction Engineering Approach to the Problem of Concrete 

Carbonation, AIChE Journal Vol. 35, No. 10 p. 1639, (1989).

[3] Nagataki, S. et al., “Effect of Curing Conditions on the Carbonation of Concrete 

with Fly Ash and the Corrosion of Reinforcement in Long-Term Tests,” Proc. Int. 

Conf. on the Use of Fly Ash, Silica Fume, Slag and Natural Pozzolans in Concrete, 

ACISP 91-24, p. 521 (1986).

[4] Chi, J.M., Huang, R. and Yang, C.C., “Effects of Carbonation on Mechanical 

Properties and Durability of concrete Using Accelerated Testing Method,” Journal 

of Marine Science and Technology, Vol. 10, No. 1, pp. 14-20 (2002). [National 

Taiwan Ocean University].
122

[5] Houst, Y. and Wittmann, F.H., “The Diffusion of Carbon Dioxide and Oxygen in 

Aerated Concrete,” Swiss Federal Institute of Technology Lausanne, Switzerland, 

Cement and Concrete Research, Vol. 24, Issue 6, pp. 1165-1176, 1994.

[6] “Computation of Corrosion Distribution Of Reinforcing Steel in Cracked Concrete", 

S.C. Kranc and A.A. Sagues, in Proc. International Conference on Corrosion and 

Rehabilitation of Reinforced Concrete Structures, Orlando, FL, Dec. 7-11, 1998, 

CD ROM Publication No. FHWA-SA-99-014, Federal Highway Administration, 

1998.

[7] Andrade, C., Alonso, C., and Molina, F.J., “Cover cracking as a function of bar 

corrosion: Part I – Experimental test,” Materials and Structures, Vol. 26, pp. 453-

464, 1993.

[8] Molina, F.J., Alonso, C., and Andrade, C., “Cover cracking as a function of bar 

corrosion: Part 2 – Numerical model,” Materials and Structures, Vol. 26, pp. 532-

568, 1993. 

[9] Global Climate Change Impacts in the United States, Thomas R. Karl, Jerry M. 

Melillo, and Thomas C. Peterson, (Eds.). Cambridge University Press, 2009.

[10] Duncan, A. J. and Riegel, M. M., "Evaluation of the Durability of the Structural 

Concrete of Reactor Buildings at SRS” Report #SRNL-STI-2010-00729, Savannah 

River National Laboratory, Savannah River Nuclear Solutions, Aiken, SC 29808. 

February 2011.

2.12.2 Chloride Ingress

Chloride ingress can cause localized de-stabilization of the passive iron oxide film that is 

normally present on reinforcement bars embedded in the concrete. Such destabilization 

will lead to pitting corrosion. If carbonation, as described in section 2.12.1, would also 

reach the depth of the reinforcement bars, the oxyhydroxide film on steel would tend to 

dissolve, and gross steel corrosion would be expected leading to eventual spallation. 

That is, chloride ingress alone would cause localized (i.e., pitting) attack, but with 

carbonation, gross steel corrosion would occur and would be accelerated by the 

presence of chlorides.

The following discussion is drawn primarily from reference 1. Chloride attack has been 

modeled as a 2-step mechanism [1]. The first step is the time to initiation of corrosion 

due to chloride diffusion to the rebar. This initiation time can be given by a simplistic 

empirical model [2]:

( )[ ]

0.42

1.22 129

− = WCR Cl

X

t C

c

[7]

tc = time to onset (yr)

Xc = thickness of concrete cover (in.)

WCR = Water-to-cement ratio (mass)

Cl- = Cl- ion concentration at surface of concrete (ppm)
123

With an assumed concrete cover of 1 inch (2.54cm), a WCR of 0.45, and a chloride ion 

concentration of 10 ppm, a time of 109 years would be expected for chloride penetration 

to the rebar. With microcracks in the cover and the concomitant short-circuit transport 

compared to bulk diffusion, this time would be lessened.

Figure 54 shows the results of a chloride ingress characterization taken from the 

P-reactor building at SRS after approximately 60 years of service. With the assumptions 

above for surface chloride ion concentration and WCR, the predicted result of 

approximately 0.4-inch (1.0cm) ingress with the empirical model is congruent with the 

measurement of high chloride at the depth of 0.5 inches (1.3cm).

The second step is the corrosion rate of rebar is controlled by oxygen diffusion. 

Reference 1 describes reaction kinetics for iron corrosion and loss of reinforcement bar 

net section in terms of oxygen diffusion.

Figure 54 Chloride ingress (left to right) limited to enhanced near-surface 

concentration in a concrete specimen from the SRS P-Reactor Building with 

60 year’s service [Figure reproduced from ref. 10].

Assessment of Degradation of Concrete Due to Chloride Ingress under Extended 

Storage: With time, the diffusion of chloride and its ingress through the cover of 

concrete over the reinforcement bar can lead to pitting corrosion of the bar. Additionally, 

general corrosion of the rebar and potential spallation of the concrete through corrosion 

product expansion could occur because of the carbonation and oxygen diffusion that will 

also be taking place. A lowering of pH at the steel/concrete interface can be caused by 

diffusion of CO2 through the concrete cover. An oxygen feed and water are necessary 

conditions for the active degradation of the reinforcement bar by general corrosion. 

Such corrosion can lead to loss of tensile strength and concrete spallation. An 

oxygen/water supply can develop through microcracks in the concrete which can result 

from mechanical stresses and/or the combined effects of CO2, Cl-

, and O2 species. The 

0

50

100

150

200

250

0

200

400

600

800

1000

1200

1400

1600

1800

0 1 2 3 4 5 6

Chloride (mg/kg)

Sulfate mg/kg

Drill depth (inches)

Anions for B2S INT

SO4 (mg/kg)

Cl (mg/kg)
124

penetration of any species depends on the surface concentration of that species, thus, 

the driving force for chlorination in marine environments will be higher than in a nonmarine environment due the chloride concentration at the surface of the concrete. 

Clearly, the nature of the exposure environment will have a significant impact on the rate 

of concrete degradation.

Although the rate of chloride ingress would be concrete-condition-specific, the time to 

effect carbonation to depths of the reinforcement bar in a concrete pad should be wellwithin a 300-year period as shown in Figure 42. 

Assessment of Concrete DCSS Structures with Chloride Ingress - Impact to Safety 

Function: The potential for extensive cracking of the concrete from chloride ingress 

with rebar corrosion and spallation of the concrete cover is significant and could result in 

an unstable pad foundation for a cask. The degraded concrete would create a cask tipover hazard with the associated loss of thermal cooling configuration and shielding 

function. Additionally retrieval of the casks for transportation would also be impacted by 

an unstable pad. 

The potential for extensive cracking of a concrete cask containing reinforcement bar, but 

without an encasement shell, could result in concrete spallation with a loss of net section 

and a reduction in the shielding function. Concrete casks without reinforcement bars 

would not be subjected to chloride ingress induced spallation damage but would be 

subjected to freeze-thaw damage and would lack the strength of reinforced concrete 

structures.

Recommendations for Resolution: Age related degradation should be anticipated in 

any concrete structure used in a dry storage system. Models to accurately predict such 

degradation as a function of concrete and rebar specifications and exposure 

environment are not available. The resources required to develop and validate such 

models are extensive and investment of such resources is not likely to be cost effective. 

Additionally, concrete degradation generally initiates at the surface and can be noted by 

visual examination. Therefore, it is recommended that visual inspection and evaluation 

protocols be developed and applied as necessary. 

Section 2.12.2 References

[1] Walton, J. C. et al. “Models for Estimation of Service Life of Concrete Barriers in 

Low-Level Radioactive Waste Disposal,” NUREG/CR-5542, July, 1990.

[2] Clear, K. C., Time to Corrosion of Reinforcing Steel in Concrete Slabs, Vol. 3. 

Performance After 830 Daily Salt Applications, Federal Highway Administration 

Report No. FHWA-RD-76-70, NTIS PB-258 446, 1976.

2.12.3 Shrinkage of Concrete

Excess water is typically added to the concrete mix to improve its workability during 

forming of a concrete structure. Shrinkage of concrete occurs initially during curing as 

the result of the excess water leaving the concrete. This curing period typically 

continues several months after placement. As excess water evaporates into the 

surrounding environment, tensile stresses are induced in the concrete due to internal 
125

pressure from the capillary action of water movement, and cracks develop. Subsequent 

drying and shrinkage occurs in concrete for up to about 30 years. Over 90% of the 

shrinkage occurs during the first year and about 98% in the first five years [1]. 

Therefore, shrinkage cracking is not a significant aging degradation mechanism for 

extended storage.

Assessment of Degradation of Concrete Due to Shrinkage Cracking under 

Extended Storage: Microcracking due to shrinkage cracks start as surface cracks in 

concrete that can grow in depth with time [2]. It is assumed that the DCSS concrete 

structures were designed and fabricated with a high degree of workmanship, and that 

shrinkage cracking would be limited to minor superficial cracking.

However, degradation due to carbonation, chloride ingress, and chemical leaching can 

be promoted by surface and internal cracking in concrete.

Assessment of Concrete DCSS Structures with Shrinkage Cracking - Impact to 

Safety Function: Shrinkage cracks themselves are not significant to concrete structural 

performance. No impact to the safety function for either a concrete pad or a concrete 

cask is expected under EST.

Data Needs and Recommendation for Resolution: Shrinkage can cause surface 

cracking. Periodic monitoring of surface cracks to evaluate and trend distribution and 

opening width is recommended as part of a monitoring & in-service inspection program.

Section 2.12.3 References

[1] ACI 209R-92, Prediction of creep, Shrinkage, and Temperature Effects in Concrete 

Structures (Reapproved 2008), American Concrete Institute, 2008.

[2] ACI 224R-01, Control of Cracking in Concrete Structures, American Concrete 

Institute, 2001. 

2.12.4 Freeze-Thaw of Concrete

Repeated cycles of freezing and thawing may change the mechanical properties and 

physical form of concrete [1]. The durability of concrete to a freeze-thaw damage 

mechanism depends primarily on the porosity characteristics of aggregates, the 

presence of moisture to saturate the fines pores in aggregates, and the permeability of 

the hardened cement mortar matrix to the passage of water. A mathematical model for 

the basic processes involved in freezing and thawing of concrete is available [2], 

however, there is no rational model for crack growth in a freeze-thaw cycle, and any 

such model would be dependent on these many factors including concrete and rebar 

specifications, environmental conditions. Because of the large number of variables 

involved, such a model would be of little value to the assessment of dry storage facilities.

Freeze-thaw damage typically occurs on relatively flat surfaces such as pavement, 

where water remains in contact with the concrete. Concrete structures or components 

located in areas of the country that experience numerous freeze-thaw cycles with 

significant winter rainfall are more likely to exhibit damage than areas in milder climates. 

Specification C 33 [2] groups the areas of the U.S. into severe, moderate, and negligible 
126

weathering regions on the basis of the weathering index. This index is the product of the 

average annual number of freezing cycle days and the average annual winter rainfall in 

inches. However, climate change over many decades is expected to result in changes 

including local annual precipitation [3].

Reference 1 claims that for concrete structures or components located in a geographic 

region subject to negligible weathering conditions, that is, a weathering index of less 

than 100 day-inches per year, the freeze-thaw is not a significant aging degradation 

mechanism and requires no further evaluation. This claim would not be expected to be 

applicable to EST at any specific site because of the anticipated climate change during 

the storage period.

Figure 55 shows a sketch of the face map indicating cracks on the pad for the Idaho 

cask demonstration program at Test Area North. These cracks are assumed to be a 

combination of shrinkage and freeze-thaw damage.

Figure 55 Example of surface cracking on concrete pad at Test 

Area North in Idaho [Figure reproduced from ref. 4].

Assessment of Degradation of Concrete Due to Freeze-Thaw under Extended 

Storage: Concrete structures designed in accordance with the ACI 318-63, or a later 

version of ACI 318 [5], or ACI 349-85 [6], and constructed in accordance with 

ACI 301-66 or later [7] and using materials conforming to ACI and ASTM standards were 

suggested [1] to not be subject to the freeze-thaw aging degradation mechanism. 

Virtually all concrete contains pore water and surface/rain water can, with time, ingress 

into minor discontinuities which can be enlarged during freeze-thaw cycles. Therefore, it 

is expected that, with time and continued exposure to freeze-thaw cycles over the many 

decades of EST, freeze-thaw damage would initiate. Freeze-thaw damage originates at 

the surface of a concrete structure and propagates as surface defects are opened. The 

extent and depth progression of freeze-thaw cracking in a concrete pad or even a bare 
127

concrete cask shell would be expected to increase with repeated exposure cycles and 

therefore it is not self-limiting. Climate change can cause regions that have been initially 

classified as negligible in weathering to become severe, so freeze-thaw cannot be 

dismissed categorically for DCSS in EST.

Minor freeze-thaw cracking is localized on a structure and will typically not affect the 

strength of the concrete. Continued freeze-thaw cracks may cause “pop-outs” or local 

areas of spalled concrete and extensive cracking of the concrete would be expected with 

time under repeated freeze-thaw cycles.

In addition, degradation due to carbonation, chloride ingress, and chemical leaching can 

be accelerated by short-circuiting the diffusion process through surface cracking in 

concrete caused by freeze-thaw induced cracking.

Assessment of Concrete DCSS Structures with Freeze-Thaw Cracking - Impact to 

Safety Function: Minor freeze-thaw cracks themselves are not significant to concrete 

structural performance unless they continue to progress and cause “pop-outs” and 

extensive cracking. “Pop-outs” or local areas of spalled concrete can impact a 

component’s safety function as previously discussed. 

There is a potential for extensive cracking of a concrete pad from freeze-thaw damage 

and such cracking could result in an unstable foundation for a cask creating a cask tipover hazard with a loss of its thermal cooling configuration and its shielding function. 

Additionally retrieval of the casks for transportation would also be impacted with an 

unstable pad.

There is a potential for extensive cracking of a concrete cask without an encasement 

shell which could result in concrete spallation with a loss of net section and shielding 

function.

Data Needs and Recommendation for Resolution: Freeze-thaw can cause surface 

cracking and more extensive damage with time. Additionally, freeze-thaw processes 

may interact with other concrete degradation mechanisms to accelerate age related 

damage. Periodic monitoring of the concrete pad and DCSS components for cracking 

and results trending is recommended as part of a monitoring & in-service inspection 

program. Acceptance criteria would need to be established for the concrete inspection. 

The condition at which the concrete would need to be repaired or a new concrete pad 

would need to be constructed would be identified.

Concrete Degradation by Freeze-Thaw, Data Needs for EST and General 

Recommendations for Resolution: The concrete in the pad and cask is subject to 

degradation by freeze-thaw throughout EST, especially if climate change cause normal 

storage environments to produce many additional freeze-thaw cycles than typical for the 

present climate. Freeze-thaw degradation can be readily monitored in a periodic inservice inspection program. Specifically, 

1. Monitoring/In-Service Inspection Activity: Develop monitoring/in-service 

inspection protocol for freeze-thaw damage with acceptance criteria.
128

2. Testing & Analysis Activity: Investigate new technologies for NDE for 

characterizing freeze-thaw cracking in concrete to better evaluate crack depth 

rather than present methods of surface opening and mapping.

3. Cask Demonstration Activity: As part of a cask demonstration, establish baseline 

and periodically monitor for freeze-thaw cracking.

4. Maintenance: Develop/implement a repair technology to seal surface 

penetrations from continued freeze-thaw damage. 

The significance to EST for freeze-thaw degradation of the concrete structures is low. 

However, it is noted that: 1) freeze-thaw damage may occur due to climate changes in 

EST and may accelerate other degradation mechanisms, 2) aging management for 

freeze-thaw damage is not established; and 3) remediation of freeze-thaw degradation 

should be simple with surface sealing.

Section 2.12.4 References

[1] Class I Structures License Renewal Industry Report; Revision 1, EPRI TR-103842, 

July 1994.

[2] “Mathematical Model for Freeze-Thaw Durability of Concrete,” Z.P. Bazant, J.-C. 

Chern, A.M. Rosenberg, and J.M. Gaidis, J. Am. Ceramic Soc., 71 [9] (1988) 

pp. 777-783.

[2] Specification C-33: ASTM C 33, Figure 1, “Location of Weathering Regions.”

[3] Global Climate Change Impacts in the United States, Thomas R. Karl, Jerry M. 

Melillo, and Thomas C. Peterson, (eds.). Cambridge University Press, 2009.

[4] Dry Cask Storage Project – Phase 1: CASTOR V/21 Cask Opening and 

Examination, NUREG/CR-6745, INEEL/EXT-01-00183.

[5] ACI-318-08, Building Code Requirements for Structural Concrete and 

Commentary, American Concrete Institute, 2008.

[6] ACI-349-01/349R-01, Code Requirements for Nuclear Safety-Related Concrete 

Structures and Commentary.

[7] ACI-301-10, Specifications for Structural Concrete, American Concrete Institute, 

2010.

2.12.5 Calcium Leaching of Concrete

Cement components can be leached from concrete only when water passes over the 

surface or through cracks. Specifically, calcium compounds in the concrete can be 

dissolved in unsaturated water and washed away. Since the most readily soluble 

calcium compound in concrete is calcium hydroxide (lime), water can leach lime from 

concrete. When calcium hydroxide in the lime in concrete is leached away, other 
129

cementitious constituents become exposed to chemical decomposition, eventually 

leaving behind silica and alumina gels with little or no strength [1]. 

The extent of the leaching during any given exposure to water depends on the salt 

content and temperature of the water [2]. Both of these variables are associated with 

the atmospheric conditions at the storage site which are likely to change significantly 

during the storage period. Therefore, exposure criteria cannot be developed and 

modeling concrete behavior is tenuous, at best. Because such difficulties dominate the 

concrete degradation arena, visual inspection of the storage sites is the most applicable 

current technology for site evaluation.

However, in spite of the lack of an ability to quantify concrete behavior, leaching over 

long periods of time can increase the porosity and permeability of concrete, making it 

more susceptible to other forms of aggressive attack and reducing strength. Leaching 

can also lower the pH of the concrete and affect the integrity of the protective oxide film 

of reinforcement steel [2], similar in effect to carbonation.

The dissolving and leaching actions of the percolating water are related to permeability 

of concrete [3]. Resistance to leaching action can be enhanced by using a concrete with

low permeability [2]. A dense concrete with a suitable cement content that has been well 

cured will be less susceptible to leaching of calcium hydroxide because of its low 

permeability and low absorption.

Models for the several coupled processes involved in calcium leaching from cement 

hydrates have been developed [3, 4]. A recent model [4], benchmarked by immersion 

tests, includes a multi-component hydration model, a moisture transport/equilibrium 

model, and a micro-pore structure development model. The effects of w/c ratio, flyash 

content, and temperature on calcium leaching were assessed. Concrete with a low w/c 

ratio (0.25) and fly ash content of 20% were highly resistant to long-term calcium 

leaching showing a calcium-loss region of 5 cm after 10,000 years exposure.

Assessment of Degradation of Concrete Due to Calcium Leaching under Extended 

Storage: The requisite conditions of refreshed water and the overall kinetics of the 

coupled processes involved in calcium leaching renders it an insignificant degradation 

mechanism for the concrete in the pad or cask of the DCSS for EST. Calcium leaching 

is a slow process compared to the carbonation mechanism in the environment for the 

DCSS structures and is not rate limiting because it is not required for carbonation and 

other degradation processes to be operative.

Assessment of Concrete DCSS Structures with Calcium Leaching - Impact to 

Safety Function: Calcium leaching would reduce the strength and would increase 

micropore size in concrete; however, it is not considered to be a significant degradation 

mechanism so it would not be significant to impact the safety functions of the DCSS 

structures.

Calcium Leaching Data Needs and Recommendation for Resolution: Calcium 

leaching is not considered a significant degradation processes and therefore does not 

present any issues requiring resolution. Furthermore, the periodic monitoring of the 

concrete pad and DCSS components recommended as part of a monitoring & in-service 

inspection program to resolve other degradation issues are applicable to calcium 

leaching.
130

Section 2.12.5 References

[1] Troxell, G. E., Davis, H. E., and Kelly, J. W., Composition and Properties of 

Concrete, Second Edition (J. W. Kelly), McGraw-Hill, 1968.

[2] Class I Structures License Renewal Industry Report; Revision 1, EPRI TR-103842, 

July 1994.

[3] Bentz, D.P. and Garboczi, E.J., “Modeling the Leaching of Calcium Hydroxide from 

Cement Paste: Effects on Pore Space Percolation and Diffusivity,” Materials and 

Structures 25, pp. 523-533 (1992).

[5] Nakarai, K., Ishida, T., and Maekawa, K., “Modeling of Calcium Leaching from 

Cement Hydrates Coupled with Micro-Pore Formation,” Journal of Advanced 

Concrete Technology, Vol. 4, No., 3, pp. 395-407, October 2006. 

2.12.6 Sulfate Ingress

The ingress of sulfate ions from the environment can cause degradation of concrete 

structures and is termed sulfate attack. Damage due to sulfate interactions can cause 

softening and cracking of the concrete leading to spallation. The cracking is the result of 

expansive attack of calcium hydroxides and calcium-aluminates in the concrete. The 

sulfate ions are present in water on the concrete surface because of dissolved sodium, 

potassium or magnesium sulfates from the environment. 

Modeling of the rate of degradation due to sulfate ingress shows that the rate is directly 

dependent on the diffusivity of the sulfate ion in concrete and increases as the sulfate 

concentration in the exposure environment increases [1]. Other models for sulphate and 

freeze-thaw damage coupled model calculate the penetration depth based on sorptivity 

coefficient at current concrete surface temperature, current “duration” of time-ofwetness, and concrete porosity [2]. These models, using sorption as the dominant 

transport mechanism, show sulphate attack is due to reaction with cement to form 

expansive ettringite that leads to spallation of concrete, as described in reference [3].

Because the sulfate concentration in any water on the concrete pad will be low and the

time of exposure generally short, this degradation mechanism is not considered 

significant to long term storage of spent fuel. Clearly, acid rain and industrial 

environments will increase the sulfate concentration and promote sulfate ingress. 

However, even in heavy industrial environments, sulfate ingress is slow relative to 

carbonation and therefore it is not considered to be the primary degradation mechanism 

to concrete degradation during EST.

Assessment of Degradation of Concrete Due to Sulfate Ingress under Extended 

Storage: The requisite conditions of water transport of the sulphates dissolved in water 

renders sulphate ingress an insignificant degradation mechanism for the concrete in the 

pad or cask of the DCSS for EST. Sulfate ingress is a slow process compared to the 

carbonation mechanism in the environment for the DCSS structures.

Assessment of Concrete DCSS Structures with Sulfate Ingress - Impact to Safety 

Function: Sulfate ingress would reduce the strength and in the absence of carbonation 

could cause degradation of the concrete leading to spallation. However, it is not 
131

considered to be a significant relative to carbonation and will not impact the safety 

functions of the DCSS structures.

Sulfate Ingress Data Needs and Recommendation for Resolution: Sulfate ingress is 

not considered significant to cause concrete degradation to impact a safety function. 

However, tests to evaluate carbonation should include evaluation of sulfate ingress at 

the time of the test as a verification of limited sulphate ingress.

Section 2.12.6 References

[1] Walton, J. C. et al. “Models for Estimation of Service Life of Concrete Barriers in 

Low-Level Radioactive Waste Disposal,” NUREG/CR-5542, July, 1990.

[2] CONCLIFE: Service Life Predictions for Concrete Pavements and Bridge Decks 

Exposed to Sulfate Attack and Freeze-Thaw Deterioration, Software prepared for 

the FHWA by Building and Fire Research Laboratory 2001, National Institute of 

Standards and Technology.

[3] Atkinson, A., and Hearne, J.A., “Mechanistic Model for the Durability of Concrete 

Barriers Exposed to Sulphate-Bearing Groundwaters,” MRS Symp. Proc., Vol. 176, 

pp. 149-156, 1990.

2.12.7 Temperature Effects and Radiation Effects on Concrete

Concrete has been successfully used as a structural material in nuclear facilities for over 

50 years. These facilities include storage pools for spent fuel, reactor buildings, fuel 

reprocessing canyons and other areas that may be exposed to elevated temperatures, 

and that have received doses of α, β, γ, and n radiation. The structures are typically in 

very good condition.

A recent publication [1] that contains an extensive literature survey of the effects of 

temperature on the short-term mechanical and physical properties of concrete for the 

nuclear industry shows that elevated temperature properties are dependent on 

temperature (level), heating rate, loading, external sealing of the specimen (affects 

moisture loss during the test), and concrete composition including its moisture state, 

chemical structure, and physical structure. The publication [1] also compiled the 

available literature on laboratory testing and specimens harvested from serviceexperienced structures, albeit limited, on long-term aging. A rigorous test program 

conducted by Central Research Institute of Electric Power Industry in Japan show 

results [2] that indicate, in the worst case, there is only a small decrease in compressive 

strength (up to 25%) and a decrease in elastic modulus (up to 50%) in concrete exposed 

for 8 years for temperatures up to 110°C.

An earlier publication [3] evaluated the results of the “long-term” exposure (5000 days) 

exposure of concrete to 65°C temperatures. The results showed an increase in both the 

compressive strength and elastic modulus. Similar increases were seen in concrete 

samples exposed for shorter times (< 5000 days) at temperatures up to 121°C. 

However, exposure of concrete to temperatures above 149°C decreased the 

compressive strength and elastic modulus [3].
132

Reference 1 provides sufficient data to demonstrate that the present limit for long-term, 

steady-state state temperature exposure of concrete should be greater than the 

American Society of Mechanical Engineers Boiler and Pressure Vessel Code limit of 

65°C [1]. The primary mechanism of long-term, low elevated temperature (up to 

approximately 100°C) exposure that can affect the properties of concrete is the loss of 

pore water [1]. The data presented in references 1-3 suggest a threshold for thermal 

degradation during EST of approximately 90°C which exceeds the anticipated 

temperature of the concrete in DCSS service.

The only plausible degradation is water loss from the near surface regions. Shrinkage 

due to water loss may lead to some surface cracking that would be readily detectable. 

Surface cracking can promote other degradation mechanisms such as freeze-thaw 

cracking and the ingress of aggressive ions.

Generally, concrete is considered stable until 1020 neutrons/cm2 or 1010 rads of gamma. 

A literature review [4] confirms the stability under compressive loads, but suggests that 

the tensile properties begin to degrade after about 1019 neutrons/cm2

. However, it is 

worth noting that the review also stated that “all effects on concrete due to radiation, per 

se, were too slight to reliably measure because of the gross effects of the increased 

temperature during exposure.” Additionally, the total exposure anticipated for concrete 

shielding structures containing high burnup MOX fuel is less than 1020 neutrons/cm2

.

The literature, as summarized in references 1-4, suggests that exposure of concrete to 

low level irradiation of a DCSS and low, < 90°C, temperatures, even under long times, 

will not significantly degrade the properties of concrete.

Additional temperature limits for concrete have been recommended in the literature [5] to 

protect concrete against water loss and retention of shielding capability. For neutron 

shields, the maximum internal concrete temperature is suggested to be 88 ̊C, and for 

gamma shields, the maximum internal temperature is suggested to be 177  ̊C. These 

temperature limits are consistent with the 90°C suggested for limiting mechanical 

property impact during EST.

Assessment of Concrete DCSS Structures Exposed to Temperatures and 

Radiation Levels under EST: The impact of the temperature below 90°C and low 

radiation fields of both gamma and neutron expected during EST is expected to be 

insignificant. Surface cracks may form on the exposed surfaces due to loss of pore

water, and these cracks could serve to enhance other forms of age related degradation 

which require the ingress of moisture or aggressive species. However, the exposure 

temperatures and radiation levels per se should not significantly impact the bulk 

properties of concrete associated with a DCSS.

Temperature and Radiation Effects on Concrete - Data Needs and 

Recommendation for Resolution: Long-term exposure data do not exist, and because 

temperature and irradiation effects have been observed during (albeit very high dose) 

short-term exposures, relevant data should be collected during the aging management 

program activity of periodic monitoring of the physical condition of the concrete. Periodic 

radiation monitoring would provide a measure of loss of the shielding function.

The significance to EST for high temperature and radiation degradation to concrete is 

low. There is no significant degradation due to temperature or radiation effects expected 
133

to concrete structures under EST. Periodic monitoring of radiation fields during EST 

would identify loss of shielding function; however administrative controls could be readily 

established if shielding loss became significant.

Section 2.12.7 References

[1] Naus, D.J., “A Compilation of Elevated Temperature Concrete Material Property 

Data and Information for Use in Assessments of Nuclear Power Plant Reinforced 

Concrete Structures”, NUREG/CR-7031; ORNL/TM-2009/175, published 

December 2010.

[2] Kanazu, T., Matsumura, T., and Nishiuchi, T., “Changes in the Mechanical 

Properties of Concrete Subjected to Long-Term Exposure to High Temperatures,” 

Report No. U95037, Abiko Research Laboratory, Central Research Institute of 

Electric Power Industry, Japan, March 1996.

[3] Kassir, M.K., Bandyopadhyay, K.K., and Reich, M., “Thermal Degradation of 

Concrete in the Temperature Range from Ambient to 315°C”, BNL-52384, 1996.

[4] Fillmore, D.L., “Literature Review of the Effects of Irradiation and Temperature on 

the Aging of Concrete”, INEL/EXT-04-02319, 2004.

[5] Hungerford, H.E., Hönig, A., Desov, A.E., Du Bois, F., and Davis, H.S., “Concretes, 

Cements, Mortars, and Grouts,” Engineering Compendium on Radiation Shielding,” 

Vol. II, p. 117, R.G. Jaeger (Editor), Springer-Verlag, Berlin, Germany, 1975.

2.12.8 Alkali-Silicate Reactions

Alkali-silica reactions, although a major cause of concrete degradation, are not 

anticipated in concretes relevant to long term dry storage, because these concretes are 

assumed to have been constructed with aggregates that are compatible with the 

hydroxyl ions in the concrete mixture. There is a simple, non-destructive field test to 

demonstrate the lack of alkali-silica reactions in concrete structures [1] and it is 

recommended that, if necessary, tests to confirm this assumption be conducted.

Chemical reactions may develop between certain mineral constituents of aggregates 

and alkalis that compose the Portland cement [2]. These alkalis are largely introduced in 

the concrete by cement, but may also be present from improper admixtures and saltcontaminated aggregates [3]. Seawater and solutions of deicing salt can also inject 

alkalis into concrete by action of penetration. Three types of chemical reactions may 

occur depending upon the composition of the aggregates. They are alkali-aggregate 

reaction, cement-aggregate reaction, and expansive alkalis-carbonate reaction [3].

Alkali-aggregate reaction, also known as alkali-silica reaction, occurs when aggregates 

that contain silica get in contact with alkaline solutions. All silica minerals have the 

potential to react with alkaline solutions, but the degree of reaction and ultimate damage 

incurred can vary significantly [3]. Alkali-aggregate reactions can cause expansion and 

severe cracking of concrete structures [4]. A map and data showing geographic areas 

known to yield natural aggregates suspected of, or known to be capable of, alkali-silica 

reaction is included in ACI 201.2. The reactivity of such aggregates might not be 
134

recognized until the structures are over 20 years old, even if used in combination with 

high alkali cement [3]. It is recommended that if an NPP is close to one of the areas 

shown on the ACI 201.2 map, that the associated concretes be reviewed and, if 

necessary, evaluated to assure the absence alkali-silicate reactions.

Cement-aggregate reaction, which is a second type of reaction similar to alkaliaggregate reaction, occurs between the alkalis in cement and siliceous constituents of 

the aggregates, and is complicated by environmental conditions that produce high 

concrete shrinkage and alkali concentrations on the surface due to drying (aging) of the 

concrete [3].

Expansive alkali-carbonate reaction is a third type of reaction between certain carbonate 

aggregates and alkalis, which in some instances produces expansion and cracking of 

concrete. Certain limestone aggregates have been reported as reactive [3]. Aggregates 

that react with alkalis can induce expansion stress in the concrete to a varying degree of 

severity, and in some cases, the expansion stresses may be high enough to produce 

cracking of the concrete structures. This cracking is irregular and has been referred to 

as “map cracking” [3].

The operating histories of nuclear power plant concrete structures do not indicate that 

structural integrity of these concrete structures is significantly affected by alkaliaggregate reactions. Aggregates used in the concrete can be tested and 

petrographically examined in accordance with ASTM Practice C 295 and Test Method C 

227 to determine the potential for reactivity with alkalis. When aggregate reactivity is 

considered a possibility, accepted construction practices place limitations on cement 

alkalis used during construction and/or require an additive such as an effective pozzolan 

to be used in combination with the cement. Reaction with aggregates is not a significant 

aging degradation mechanism for concrete structures designed and constructed in 

accordance with provisions given in the ACI 201.2. However, in the unlikely event that 

concretes used in construction did not meet accepted practices, the inspection and 

monitoring protocols be modified to increase the frequency of evaluation.

Assessment of Degradation of Concrete Due to Alkali-Silicate Reactions under 

Extended Storage: Alkali-silicate reactions are caused by incompatibilities between the 

aggregate and cement in the concrete mixture. The assumption made in this evaluation 

is that the concrete was comprised of and fabricated with aggregates, cement, and 

admixture constituents that are compatible. No fundamental aging mechanism, 

including environmental effects, is involved in this degradation mechanism, and it is 

judged as insignificant to EST. 

Assessment of Concrete DCSS Structures with Alkali-Silicate Reactions- Impact to 

Safety Function: Alkali-silicate reactions can be insidious to impact strength by 

increasing micropore/microcrack size in concrete; however, it is not considered to be a 

significant degradation mechanism, so it would not be significant to impact the safety 

functions of the DCSS structures.

There is a simple, non-destructive field test to demonstrate the lack of alkali-silica 

reactions in concrete structures [1] and it is recommended that, if necessary, such a test 

be conducted to confirm this assumption. 
135

Section 2.12.8 References

[1] Colored Dye Field Test to Detect Alkali Silicate Reaction, ASR Detect*, U.S. Patent 

5,739,035 Los Alamos National Laboratory. 

[2] Suoulikidis, T., Tsakopoulos, A., and Moropoulos, T., “Accelerated Rebar 

Corrosion When Connected to Lightning Conductors and Protection of Rebars with 

Needle Diodes Using Atmosphere Electricity,” Contained in Publication ASTM STP 

906, “Corrosion Effects of Stray Currents and the Techniques for Evaluation 

Corrosion of Rebars in Concrete.”

[3] Class I Structures License Renewal Industry Report; Revision 1, EPRI TR-103842, 

July 1994.

[4] Stanlon, T. E., “Expansion of Concrete Through Reaction Between Cement and 

Aggregate,” American Society of Civil Engineers, Proceedings, Vol. 66, December 

1940.

137

3 RECOMMENDATIONS FOR ACTIVITIES TO ESTABLISH THE 

TECHNICAL BASES TO ENABLE EST

Recommendations are provided for activities to resolve materials aging issues in the 

technical bases needed to demonstrate the continuing integrity of SSCs and their ability 

to maintain functions important to safety during EST. Aging management activities, such 

as monitoring and preventive maintenance, are essential to provide condition 

assessments through the storage period. Other activities are suggested to strengthen 

the set of information to evaluate materials reliability in the technical bases for EST. 

The recommended activities are grouped in the following four areas:

1. Monitoring Program/In-Service Inspection – Periodic actions to provide a 

condition assessment of the SSCs through examinations and non-destructive 

testing are recommended. Acceptance criteria for service-induced 

degradation of the components should also be established. The decision to 

keep in service or repair/replace components and/or recover the fuel for 

repackaging would follow identification and characterization of service-induced 

flaws or degraded materials conditions.

2. Maintenance Activities – Periodic actions to provide preventive maintenance 

required to maintain the DCSS throughout the EST are required. These 

actions should address items identified in the inspection programs.

3. Testing and Analysis – Laboratory full-scale separate-effects testing and/or 

additional modeling and analysis of degradation phenomena to provide 

verification of the presence or absence of age-related degradation should be 

conducted. If degradation is expected, data to support constitutive model(s) 

for the exposure-time dependence of materials properties should be obtained. 

4. DCSS Demonstration – Integrated testing of SSCs of DCSS in storage 

setting. The demonstration would:

a. include a fuel safe storage demonstration, including those at challenging 

conditions (high burn-up fuels, for example)

b. confirm degradation and aging models

c. demonstrate NDE techniques needed for an ISI Program

d. provide a test bed for new remote sensors

e. provide information from fuel DE removed from the demonstration cask

Additionally, activities for EST storage will require plan development and potential new 

dry-transfer system design for fuel recovery, including for degraded fuel and SCCs, and 

especially for envisioned orphan sites.
138

3.1 Maintenance Program

Polymeric and metallic seals installed on casks with bolted closures are expected to 

deteriorate over time. The integrity of these gasketed seals will need to be periodically 

confirmed by leak test and provisions must be made to replace degraded seals. This will 

not be a trivial exercise as it may require removing the cask lid, replacing the seal(s), 

reinstalling the lid, and reestablishing the cask gas fill, all while maintaining personnel 

radiation exposure ALARA.

The surfaces of metal casks may need periodic re-coating. Bolts installed on casks with 

bolted closures may need periodic replacement. 

The selection of items for preventive maintenance and the preventive maintenance 

schedule should be integrated with the monitoring program.

3.2 Monitoring & In-Service Inspection Program

Monitoring and in-service inspection provides a condition assessment of the DCSS 

SSCs, and enables trending of material condition with time. A summary of domestic and 

international experience in DCSS surveillance (monitoring) is provided in chapter 4 of 

reference [1]. For EST, a standing program is recommended to be developed and 

implemented.

The inspection program recommended for EST should be consistent with the 

requirements of ASME Section XI, with scope components consistent with the Safety 

Analysis Reports, Certificates of Compliance, and Technical Specifications of the 

individual licensees. With storage times expected to extend to greater than 100 years, it 

is clear these inspection programs will have to include mechanisms for maintenance and 

repair of the ISFSI components which would be expected to degrade over time. 

Additionally, the inspection protocols will change as new data are obtained and new 

inspection technologies emerge.

The element of an ISI program for DCSS would follow those in an ASME SCXI-type 

program and include the scope, extent, NDE methods, frequency, acceptance criteria, 

and records. Through acceptance criteria of an SCC that is identified and characterized 

with service-induced degradation, decisions for repair/replacement and also for fuel 

recovery for repackaging would be determined.

Several SSCs are suggested to be subject to in-service inspection. Employment of 

inspection programs has shown to be an effective tool to ensure that confinement 

structures in the commercial nuclear industry meet design life requirements [2].

The canister and especially its closure weld are also subject to failure from stress 

corrosion cracking as a result of the interaction of environmental conditions with residual 

stresses associated with the welds. In general, the gas fill of welded containers is not 

monitored, so detection of an SCC failure (wall penetration) may be very difficult. 

Insipient cracking of containers could be detected by penetrant or ultrasonic 

interrogation of the canister welds, but would require access to the welds which is not 

currently available. Methods of remotely performing periodic inspections of container 

welds should be developed. Obtaining the access to perform necessary inspections in 

situ may require modification or redesign of the ISFSI.
139

As discussed, the concrete in casks, vaults, and pads will be expected to degrade as a 

result of environmental exposure, chemical attack, etc. This could lead to cracking, 

spalling, and/or cement/aggregate reactions, potentially leading to corrosion of any 

reinforcing steel members. Inspection of the concrete structures will primarily be 

performed by visual inspection, but may also include inspection by sonic or ultrasonic 

techniques, impact tests, radiography, radar, and thermal techniques. 

Periodic visual inspection of exposed concrete cover to monitor for signs of degradation 

(i.e., cracking, spalling, and alkali aggregate reactions) will be the primary foundation of 

this program element. ACI 201.1R (Guide to conducting a visual inspection of concrete 

in service) will serve as the basis by which these inspections are performed. A 

frequency will be selected to ensure the changes in condition are noted prior to 

degradation proceeding beyond the point of mitigation. Reports will detail the condition 

of the concrete in a manner such that further inspections will be able to discern if any 

change has occurred since previous inspections. If necessary, non-destructive (e.g., 

ultrasonic techniques for crack measurement) and destructive (e.g., core sampling) 

evaluation will be performed as part of the inspection to determine if degradation is in an 

active or dormant state.

3.3 Testing and Analysis

Several gaps in the technical bases for materials aging that may be addressed with 

testing and analysis include the following: 

DHC testing and modeling of fuel cladding alloys: There are many factors that affect 

delayed hydride cracking in the zirconium cladding alloys. These factors, combined with 

the very long time of storage, make it very challenging to realistically test and predict 

DHC performance. Hydride reorientation is a large factor in DHC. The hydride 

orientation is influenced by the alloy composition, temperature differentials, texture, grain 

boundary and microstructure, presence of hydrides (preference to continue growth of 

existing versus nucleate new hydrides), plastic strains and stress states, to list the major 

contributors. To accurately characterize DHC requires that all of these and other factors 

such as stress raisers from partial throughwall cracks and flaws need to be considered. 

Thus, it is suggested that initially a multiple approach be taken to develop basic data on 

DHC that includes a) worst case analysis of large quantity of clad breaches during long 

term storage, b) characterization of the morphology of a DHC throughwall crack and the 

risk for fuel relocation, and c) general characterization of the relative time frames when 

DHC has the most risk of occurring. Expansion of these three data needs suggests that:

a) In considering the worst case condition one approach could be to assume a short 

hairline crack exists in all or most of the fuel rods and determine what impacts the 

internal gas release would have on the safety aspects of the storage system. 

Additionally, one could determine how the release would affect retrieval and transport 

and what mitigating actions could be taken.

b) To characterize the DHC, crack samples of the various cladding types could be tested 

using existing pressurized tube samples to represent the nominal crack growth 

geometry. Relatively high stresses and temperatures near, but below, the DBTT could 

be used to enhance the probability of generating a DHC. Care should be taken to limit 

the volume or path of the pressurization gas so that there is not excessive stored energy 

in the system. Testing of different alloys in different conditions could be done and the 
140

resultant OD surface crack morphology then documented and evaluated for risk of fuel 

relocation.

c) At temperatures above the DBTT, the zirconium microstructure has some ductility 

even with the hydrides present. The ductility is strongly related to the irradiation 

damage, also, and the DBTT is affected by the hydrogen content as discussed in a 

previous section. Hydride reorientation also is a very important factor for consideration 

and is very alloy, stress, strain, microstructure and temperature dependent, and other 

factors also play a part as listed above. The DHC testing plan should consider using 

both internally pressurized clad samples and some fracture toughness-type samples 

with controlled cracks present. The samples should be representative of the 

microstructure condition present after a typical dry out anneal, including residual 

radiation damage. This is important, because if microstructure changes and some 

annealing of irradiation damage occur, these changes will affect the hydride orientations 

and basic material matrix properties. Testing temperatures should be chosen to map 

particularly the conditions around the DBTT. Data from these tests can be used to 1) 

estimate the probability for long term DHC failure rates and frequencies, 2) determine if 

there are temperatures above which the probability for DHC is very small and that 

handling and transport has an acceptably low risk, and 3) determine the effects of 

various parameters which can enhance future testing by allowing a focus on variables of 

consequence and limiting the testing scope by excluding variables that are not 

significant.

Hydrogen Embrittlement testing and modeling of fuel cladding alloys: Hydrogen 

embrittlement is expected to an extent in all zircaloy cladding alloys in EST. With 

temperature decrease, the cladding will transition from ductile tearing behavior to brittle 

fracture as the failure mode. Deformation mode will also be expected to change with 

temperature. There is a need to establish fracture behavior for a set of alloys, establish 

fracture methodology, and to establish analysis protocol for transport of fuel in a 

transport cask.

The hydrogen embrittlement concern is partially related to the DHC issue, but while DHC 

is a longer term crack velocity-type of mechanism, hydrogen embrittlement is associated 

with the general reduction of ductility of the clad metal matrix and the risk for mechanical 

failure due to excessive stresses/strains associated predominately with handling and 

accident conditions. After operation in the reactor, the fuel assembly cladding has both 

extensive radiation damage and a buildup of precipitated hydrides in the microstructure. 

Like other cladding material characteristics, the degree of hydride embrittlement will be 

affected by a number of variables, including alloy chemistry, microstructure, level of 

hydrogen in the clad, hydride orientations, test temperatures, irradiation history, and 

inherent cladding flaws, both from manufacture and from operation.

Embrittlement testing should focus on testing in the circumferential (hoop) direction 

using ring sample testing, for example, from which a lot of past data has been 

generated. The testing scope should also include axial testing to generate data for 

conditions where axial loading may be predominating. Some ring crush testing and 

expansion testing should be performed using expanding mandrels from which there is 

also some data generated in the past. The basic tests (ring and maybe axial) should 

encompass the range of expected temperature exposures from room temperature to 

400°C and address the major variables of alloy, hydrogen level, hydride orientation, 
141

microstructure (including crack / flaws) and strain rate. Data on tensile strengths, 

ductility, elastic constants, and hardness can be generated.

Data from these tests will support the analysis of fuel performance during both normal 

and off-normal storage, handling and transport conditions. Tests of samples from the 

same rod, but at different axial elevations such that there are significant differences in 

hydrogen levels, will also provide data on the relative effects of radiation and hydrogen 

on clad ductility.

Creep Testing: There are two major regions where gaps exist in creep data for longterm fuel storage; during high temperature dry out anneal where the temperatures and 

pressures are relatively high and during the long term storage where temperature and 

pressures are decreasing and relatively low for significant creep to occur. Creep data 

generated in these two regions can be used to characterize and predict the creep 

condition of the fuel cladding during long-term storage.

The typical creep testing procedures used for testing of unirradiated and as fabricated 

cladding are applicable to the spent fuel testing also. The techniques are based on an 

accurate measurement of the cladding OD to a 2 micron or better accuracy, typically 

using laser micrometers. There are additional problems with spent fuel samples since 

they will have an oxide coating that will have varying dimensions and also may 

experience minor spalling during the creep test. Consideration should be given to 

removal of small sections of the oxide to allow accurate dimensional data on exposed 

metal surfaces. The creep samples would be defueled fuel rod sections approximately 

25 cm long. For accurate creep measurements the gage length should be at least 10 

times the diameter to remove end fixturing effects. The clad sample would have end 

fittings that seal one end and allow internal pressurization through the other end. The 

creep samples would be placed in a vacuum or inert gas furnace and pressurized via a 

manifold that can serve multiple samples and maintain a constant pressure.

Testing parameters of time, pressure (hoop stress) and temperature would be chosen to 

encompass the predicted exposure conditions. For example if the dry out temperature 

can reach 400°C then testing should be done, for example, at temperature internals of 

350°C, 400°C and 450°C to cover the potential exposure ranges. Since the dry out 

time at these temperatures is relatively short the test times should be from a few hours 

to a few days. The samples can be measured after each time interval and then returned 

to the furnace for additional exposure. The pressures would be determined based on 

the range of representative fuel rod internal pressures at the test temperatures.

For the longer term lower temperature creep representing conditions after the dry out 

anneal similar samples and test procedures could be used except that the test 

conditions would be different. The test times may need to extend to 3 or 5 years. It is 

important in projecting the creep rate to the long storage times to have confidence in 

obtaining a quasi-steady state creep rate for the condition, or if still in primary creep to 

have an understanding of that (upper bounding) creep rate. 

Creep rates will be strongly alloy and microstructure dependent. Other factors that will 

affect the creep rates will be the hydride content and morphology, the degree of anneal 

during dry out, variations in oxygen levels and, maybe, the irradiation damage, although 

that will approach saturation early in the reactor operation.
142

DCCG: The relative influence of diffusion controlled cavity growth (DCCG) on cladding 

performance during long term dry storage is unknown, but there are opportunities to 

develop information without expending a lot of additional effort by using samples from 

other testing programs such as the creep, DHC, and hydrogen embrittlement tests. For 

example, after the creep tests are completed a section of the creep sample can be 

examined under SEM to determine the presence or lack thereof of cavities in the 

microstructure. Likewise, the fracture surface of DHC samples and of some hydrogen 

embrittlement samples can be examined to assess the presence of cavities. The 

examination scope should include different alloys, microstructures and temperatures.

SCC Testing and Model of Canister: The stainless steel canister is a critical 

component of the DCSS that provides the confinement barrier. Stainless steel is subject 

to stress corrosion cracking, and may be subject to pitting and crevice corrosion. For 

containers that are exposed to outside air, humid oxidation occurs on the outside of the 

container; especially in marine environments. With longer storage times, temperatures

drop to lower values and condensation on the canister would be more likely. 

Condensation allows the wet corrosion mechanisms to occur on the outside of 

containers exposed to outside air. 

A fabricated prototypic canister(s) is recommended for the testing. Tests and analysis 

would consider the following materials and environments: The critical regions for SCC 

consideration are the welded joints of the canister—these welded joint regions provide 

sites for residual stresses combined with either humid air or condensate to potentially 

lead to incidents of SCC in both the weld and weld HAZ regions. SCC at the weld 

regions of the DCSS could potentially impact both DCSS confinement integrity and fuel 

retrievability. 

SCC testing of prototypic weld region materials should be conducted under the following 

conditions: 1) environment—a) humid air, to include chlorides for marine environment 

simulation and additional germane atmospheric contaminants and, b) condensate, also 

to include chlorides for marine environment simulation and additional germane 

atmospheric contaminants, 2) temperature—isothermal temperature testing over a 

range of temperature to bound the anticipated canister wall temperatures, 3) coupled 

radiation—some testing should be conducted with coupled gamma irradiation in order to 

understand any impact of gamma radiation on the SCC corrosion processes and to 

assess the impact of radiolysis decomposition products on the SCC processes, and 4) 

test configuration—SCC testing should also include U-bend weld samples configured to 

provide varying levels of stress and to be tested in a controlled environment capsule test 

for the humid environment testing in radiation fields to capture radiolysis effects in 

air/moisture system. Testing for condensate effects can also be performed using weld 

u-bend samples immersed in shallow condensate vessels—the goal of these u-bend 

tests is to provide a go-no-go measure of SCC response of the weld region materials 

and also to provide an indication of the incubation time to crack initiation. Finally, testing 

to determine accurate growth rates for materials/stress levels that indicate a cracking 

response through u-bend testing can be conducted using fully instrumented bolt-loaded 

compact tension specimens. 

For long-term predictive capability, data from the aforementioned testing must be 

combined to produce a model to represent the effect of the various test variables on 

SCC response of the DCSS weld region materials.
143

3.4 Cask Demonstration Considerations

An integrated, full cask demonstration project could provide important degradation 

information and/or confirm degradation models to support establishment of the technical 

bases for EST. A cask demonstration could involve three types of monitoring and data 

collection: 1) in situ monitoring of cask internals with sensors; 2) periodic removal of fuel 

and other internals for destructive examination; and 3) external NDE of cask, canister, 

and pad components. The specification of functions and performance requirements for 

a cask demonstration project are outside the scope of this report. The following 

discussion lists considerations for a cask demonstration.

Cask external and internal parameters could be monitored. Environmental parameters 

important to degradation are: temperature, environment (including gas atmosphere), 

radiation, and chemical effects on the fuel in storage and all storage system 

components.

Monitoring the conditions inside the storage canister or cask would include, 

thermocouple probes to determine the temperature gradient along and across the fuel 

bundles, and periodic or continuous gas monitoring for pressure and gas composition 

that would correspond to fuel cladding breaches and fission gas/volatiles release.

It is not generally straightforward to obtain cladding material properties in situ during 

storage. An exception may be cladding emissivity that would change with oxide 

thickness and structure on the cladding. With development of remote technologies, 

monitoring the SNF in situ could provide information about changes in cladding ductility 

and the potential for cladding creep; the formation, orientation, and distribution of 

zirconium hydrides in the cladding; and possibly information about the condition of the 

fuel pellets. Periodic removal of the fuel for destructive examination would validate the 

remote technologies.

As an example, an instrumented dry storage cask could include sensors to monitor for:

o temperature (of fuel and cask internals), 

o radiation fields (gamma and spontaneous neutron fluxes),

o “air-borne” contamination (monitor for gaseous and volatile fission products),

o gas composition (monitor for initial cover gas and air contamination), 

o gas conductivity,

o gas pressure,

o gas moisture, and 

o camera(s) (for fuel cladding and cask internal conditions including corrosion, 

swollen fuel, other evidence of materials degradation).

Redundant thermocouples, passive sampling tubes, and fiber optics provide a robust, 

long-term platform for information-gathering. Information from the instrumented cask 

would serve to provide a real-time surveillance of the materials and also data to 

compare against predicted materials performance models.

External cask examinations could also be performed. For example, a bolt could be 

retrieved for destructive testing and a new bolt installed. Destructive examinations 

could be ultimately performed on a cask removed from service. 
144

The cask demonstration would also provide a test bed for new NDE methods and 

remote sensor technologies. New NDE methods, new sensor technology and remote 

telemetry of data, and new DE methods could be demonstrated and deployed in the 

cask demonstration program. Adaptation of these new technologies would be expected 

to follow national consensus of their efficacy and consensus for their application in 

DCSSs.

Section 3 References

[1] Technical Bases for Extended Dry Storage of Spent Nuclear Fuel, EPRI report 

number 1003416, Final Report, December 2002.

[2] NUREG/CR 4652 “Concrete-Component Aging and Its Significance Relative to Life 

Extension of Nuclear Power Plants,” D. J. Naus, September 1986.
145

4. OFF-NORMAL AND ACCIDENT CONSIDERATIONS

Normal conditions of operation include those environments specified in the 

Environmental Report, Site Characteristics, and/or the Operating Procedures for the 

local site.

Off-normal conditions/events are considered to include those events that may 

reasonably be expected to occur during the life of the cask system and that exceed 

normal conditions. Types of off-normal conditions include loads associated with 

equipment malfunction or inadvertent loads during equipment transfer and 

environmental conditions such as low or high extreme ambient temperatures. Events 

due to climate change, such as local heavy/low rainfalls and temperature 

increase/decreases discussed in [1] would be expected in EST for a DCSS. The extent 

of these changes will be dependent on geographic location and may be categorized 

under off-normal conditions. Even under climate change conditions for the next 

300 years, no conditions are expected to fall outside of the design space for normal 

operation environmental conditions (e.g., only several degrees change in average 

temperature would be expected) with the exception of increases in atmospheric CO2

concentration and industrial pollution.

Accident analyses are ISFSI-specific. The types of accidents considered are seismic, 

tornado (wind and missile), cask tip-over, fire, explosion, and ice/snow overload. These 

events should be at the same magnitude and frequency for both present license periods 

and EST period.

Explicit consideration of the impact of off-normal and accident events on dry storage 

system performance is given in references 2 and 3. The impact of off-normal or accident 

events on material condition of the SSCs of the dry storage systems is outside the scope 

of this report.

Materials must be considered in their degraded (if any) condition as required by 

reference 4 to withstand design-basis normal, off-normal, and accident conditions. The 

potential for materials to degrade and impact safety functions during normal conditions 

of EST were discussed in Section 2 of this report. Degraded conditions, that could 

impact the performance of the SSCs of the DCSS for normal conditions, would also 

impact off-normal and accident conditions.

The following list highlights the potential impact of low burnup and HBU fuel in EST with 

respect to off-normal and accident conditions. The specific response of an SSC in a 

degraded condition is dependent on the specific off-normal or accident event, and is 

outside the scope of this report.

• HBU fuel contains a fine-grain size rim region on the fuel pellet that would be 

expected to cause a larger fraction of a fines source term for accident conditions 

in transportation.

• Both low burnup and HBU fuel as susceptible to brittle fracture and may fail at 

low temperature, high loading off-normal or accident conditions.

• Degraded (cracked) canisters with tight hairline cracks may exhibit large crack 

openings if sufficiently loaded under off-normal or accident conditions.
146

• Degraded reinforced concrete pads and overpacks may not have requisite 

capacity in off-normal or accident events, as would undegraded concrete.

The aging management activities described in Section 3 of this report could provide 

resolution to these materials issues. For example, a crack identified and characterized 

in an inspection of a canister before its transportation could be dispositioned with 

acceptance criteria that include off-normal and design basis loading events.

Section 4 References 

[1] Global Climate Change Impacts in the United States, Thomas R. Karl, Jerry M. 

Melillo, and Thomas C. Peterson, (eds.). Cambridge University Press, 2009.

[2] NUREG-1536, Standard Review Plan for Dry Cask Storage Systems. 

http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1536/.

[3] NUREG-1567, Standard Review Plan for Spent Fuel Dry Storage Facilities. 

http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1567/.

[4] NUREG-1927, Standard Review Plan for Renewal of Independent Spent Fuel 

Storage Installation Licenses and Dry Cask Storage System Certificates of 

Compliance. http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr1927/.
BIBLIOGRAPHIC DATA SHEET

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NRC FORM 335

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NRC FORM 335 (12-2010)

NUREG/CR-7116 

Materials Aging Issues and Aging Management for Extended Storage and Transportation of 

Spent Nuclear Fuel 11 2011

R.L. Sindelar, A.J. Duncan, M.E. Dupont, P.-S. Lam, 

M.R. Louthan, Jr., T.E. Skidmore Technical

Savannah River National Laboratory, Building 773-41A, Aiken, SC 29808

Division of Storage and Transportation 

Office of Nuclear Safety and Safeguards 

U.S. Nuclear Regulatory Commission 

Washington, D.C. 20555-0001

An evaluation of potential degradation of materials in the Systems, Structures, and Components (SSCs) of dry cask storage systems 

(DCSS) for Independent Spent Fuel Storage Installations (ISFSIs) has been performed to identify issues in the technical bases for 

regulating storage well beyond the present allowable storage time of 120 years that includes 60 years pool storage and 60 years dry 

storage. The evaluation included low (45 GWd/MTU) and high burnup (HBU) (> 45 GWd/MTU up to approximately 62.5 GWd/

MTU) fuel, and was performed to determine if: 1) degraded conditions could be anticipated to develop in the materials of the SSCs 

over an extended storage and transportation (EST) period (> 120 years); and 2) the anticipated degradation would challenge functions 

important to safety during storage and during post-storage transportation including fuel retrieval for final disposition. 

The approach followed a general Aging Management Review process by: 1) identifying the materials of construction of the SSCs; 2) 

evaluating their susceptibility to degradation from environmental conditions of exposure; and 3) evaluating the potential for 

degradation to impact the safety functions required by the present Code of Federal Regulations for dry storage systems and spent fuel 

transportation in 10 CFR Part 72 and 10 CFR Part 71, respectively. The evaluation was performed using extrapolation of available 

data and materials degradation models with consideration of service histories of similar materials systems, and with expert judgment. 

Spent Nuclear Fuel Storage & Transportation 

Extended Storage & Transportation 

Degradation 

Material Aging


UNITED STATES 

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

--------------------

OFFICIAL BUSINESS
NUREG/CR-7116 Materials Aging Issues and Aging Management for Extended Storage and 

Transportation of Spent Nuclear Fuel