A Quantitative Impact 

Assessment of 

Hypothetical Spent Fuel 

Reconfiguration in Spent 

Fuel Storage Casks and 

Transportation Packages

Office of Nuclear Regulatory Research

NUREG/CR-7203

ORNL/TM-2013/92
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A Quantitative Impact 

Assessment of 

Hypothetical Spent Fuel 

Reconfiguration in Spent 

Fuel Storage Casks and 

Transportation Packages

Manuscript Completed: April 2015

Date Published: September 2015

Prepared by

J. M. Scaglione, G. Radulescu, W. J. Marshall, K. R. Robb

Oak Ridge National Laboratory

Managed by UT-Battelle, LLC

Oak Ridge, TN 37831-6170

Prepared for: 

Division of Systems Analysis 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001

M. F. Bales, NRC Project Manager

NRC Job Code N6789/N6686

Office of Nuclear Regulatory Research

NUREG/CR-7203

ORNL/TM-2013/92

iii

ABSTRACT

Based on the current knowledge, commercial spent nuclear fuel (SNF), including high burnup 

fuel (burnup > 45 GWd/MTU) in the United States can be stored and transported safely in 

accordance with the respective regulatory requirements. The NRC periodically conducts 

research activities to confirm the safety of operations and enhance the regulatory framework to 

address any changes in technology, science, and policies. This includes analyses of beyond 

design basis conditions to confirm that regulatory requirements continue to provide reasonable 

assurance for safe storage and transportation of spent nuclear fuel. The research documented 

in this report is an effort of such nature.

This report documents an evaluation of the impact of a wide range of extremely unlikely 

postulated fuel reconfigurations under non-mechanistic causes of fuel assembly geometry 

change with respect to four technical disciplines: criticality, shielding (dose rates), containment, 

and thermal. The term “fuel reconfiguration” refers to any change to the storage and 

transportation system nominal intact fuel assembly configuration used for the basis of cask 

certification. Many configurations were considered to be physically unlikely realizable scenarios. 

Three fuel reconfiguration categories were considered: configurations characterized by (1) 

cladding failure, (2) rod/assembly deformation without cladding failure or (3) changes to 

assembly axial alignment without cladding failure. The analyses considered representative SNF 

designs and storage cask/ transportation packages, and a range of fuel initial enrichments, 

discharge burnup values, and decay times. 

Overall, the safety impacts of fuel reconfiguration are system design, content type, and 

loading dependent. The areas and magnitude of the impact vary from cask/package design to 

cask/package design. It should also be noted that some of the scenarios are extreme and 

physically unlikely to occur; they represent bounding values. The spent fuel storage systems 

and transportation packages approved by the NRC to date provide reasonable assurance that 

they are safe under normal, off-normal, and hypothetical accident conditions as prescribed in 

10 CFR Part 71 and 72 regulations. 

v

FOREWORD

Part 72 and Part 71 of Title 10 of Code of Federal Regulations (10 CFR) govern the storage and 

transportation of spent nuclear fuel (SNF). Based on the current knowledge of material 

properties and mechanical performance of fuel cladding, the Nuclear Regulatory Commission 

(NRC) has reasonable assurance that spent nuclear fuel, including high burnup fuel (burnup > 

45 GWd/MTU), is safe for storage and transport under normal, off-normal, and hypothetical 

accident conditions as prescribed in 10 CFR Part 72 and Part 71 for all the storage systems and 

transportation packages approved to date. The NRC periodically conducts research activities to 

confirm the safety of operations and enhance the regulatory framework to address any changes 

in technology, science, and policies. This includes analyses of beyond design basis conditions 

to confirm that regulatory requirements continue to provide reasonable assurance for safe 

storage and transportation of spent nuclear fuel. The research documented in this report is an 

effort of such nature. 

This study performed a quantitative assessment of the impact of very unlikely beyond design 

basis hypothetical changes of fuel geometry on the safety of spent nuclear fuel storage casks 

and transportation packages. Specifically, this study examined the potential changes to 

criticality, shielding, confinement/containment, and thermal characteristics of the systems due to 

change of fuel geometry. The motivation of this study is, in part, to help understand the 

characteristics of the system’s responses to fuel geometry changes. 

vii

TABLE OF CONTENTS

Section Page

ABSTRACT................................................................................................................................iii

FOREWORD ..............................................................................................................................v

LIST OF FIGURES ....................................................................................................................ix

LIST OF TABLES.......................................................................................................................xi

EXECUTIVE SUMMARY ......................................................................................................... xiii

ACKNOWLEDGMENTS............................................................................................................xv

ACRONYMS........................................................................................................................... xvii

DEFINITIONS.......................................................................................................................... xix

1. INTRODUCTION ..................................................................................................................1

2. FAILED FUEL ASSEMBLY CONFIGURATION CATEGORIES.............................................3

2.1 CATEGORY 1: CLADDING FAILURE ......................................................................... 3

2.2 CATEGORY 2: ROD/ASSEMBLY DEFORMATION .................................................... 6

2.3 CATEGORY 3: CHANGES TO ASSEMBLY AXIAL ALIGNMENT...............................13

3. CONSEQUENCE ANALYSIS..............................................................................................19

3.1 CATEGORY 1: CLADDING FAILURE ........................................................................19

3.1.1 Criticality...........................................................................................................19

3.1.2 Shielding ..........................................................................................................23

3.1.3 Containment.....................................................................................................34

3.1.4 Thermal............................................................................................................38

3.2 CATEGORY 2: ROD/ASSEMBLY DEFORMATION ...................................................44

3.2.1 Criticality...........................................................................................................44

3.2.2 Shielding ..........................................................................................................46

3.2.3 Containment.....................................................................................................48

3.2.4 Thermal............................................................................................................48

3.3 CATEGORY 3: CHANGES TO INTACT ASSEMBLY AXIAL ALIGNMENT.................50

3.3.1 Criticality...........................................................................................................50

3.3.2 Shielding ..........................................................................................................53

3.3.3 Containment.....................................................................................................55

3.3.4 Thermal............................................................................................................55

4. SUMMARY AND CONCLUSIONS ......................................................................................57

5. REFERENCES ...................................................................................................................67

APPENDIX A. CRITICALITY EVALUATIONS FOR FUEL RECONFIGURATIONS................A-1

APPENDIX B. SHIELDING EVALUATIONS FOR FUEL RECONFIGURATIONS ..................B-1

APPENDIX C. CONTAINMENT EVALUATIONS FOR FUEL RECONFIGURATIONS ...........C-1

APPENDIX D. THERMAL EVALUATIONS FOR FUEL RECONFIGURATION 

CATEGORIES................................................................................................D-1

ix

LIST OF FIGURES

Page

Figure 1 Illustration of fuel rod failure mechanisms ........................................................... 4

Figure 2 Photographs of dummy fuel assemblies after 9-m end drops.............................. 7

Figure 3 Photographs illustrating assembly fuel rod buckling after a 9-m end drop test .... 8

Figure 4 Representative BWR and PWR fuel assemblies ................................................. 9

Figure 5 Schematic representation of PWR fuel assembly before and after horizontal 

drop event..........................................................................................................10

Figure 6 Schematic representation of BWR fuel assembly before and after horizontal

drop event..........................................................................................................10

Figure 7 Schematic representation of PWR and BWR fuel assemblies before and after 

a vertical drop event...........................................................................................11

Figure 8 Example configuration for non-uniform pitch expansion .....................................12

Figure 9 Schematic representation of fuel assembly within a typical basket cell ..............14

Figure 10 Example storage pack configuration for both vertical (a) and horizontal

(b) storage modules ...........................................................................................15

Figure 11 Example rail spent fuel transportation package..................................................15

Figure 12 Photo of BWR assembly after drop test .............................................................16

Figure 13 Change in keff in GBC-32 as a function of number of rods removed ...................20

Figure 14 Change in keff in GBC-68 as a function of number of rods removed ...................21

Figure 15 Total dose rate profiles along the PWR transport package external top 

surface (a), radial surface (b), and bottom surface (c) for the intact fuel 

configuration and the fuel reconfiguration with 25% fuel mixture 

collected into the assembly bottom region, NCT ................................................31

Figure 16 Total dose rate profiles at 2 m from the PWR transport package external 

top surface (a), radial surface (b), and bottom surface (c) for the intact fuel 

configuration and the fuel reconfiguration with 25% fuel mixture 

collected into the assembly top region, NCT ......................................................32

Figure 17 Allowable leakage rate variation as a function of spent fuel rod cladding 

breaches for the GBC-32 cask and 40 GWd/MTU burnup at different 

decay times........................................................................................................35

Figure 18 Allowable leakage rate variation as a function of spent fuel rod cladding 

breaches for the GBC-68 cask and 40 GWd/MTU burnup at different

decay times........................................................................................................35

Figure 19 Effects of the location from which the fuel fines may originate on allowable 

leakage rate for the GBC-32 cask at 65 GWd/MTU burnup................................37

Figure 20 Structural temperatures with 100% rod breach—vertical orientation ..................40

Figure 21 Structural temperatures with 100% rod breach—horizontal orientation. .............40
x

LIST OF FIGURES (continued)

Figure 22 Change in peak neutron absorber temperature with failure of one assembly—

vertical cask .......................................................................................................42

Figure 23 Change in peak neutron absorber temperature with failure of one assembly—

horizontal cask...................................................................................................42

Figure 24 Change in peak neutron absorber temperature with failure of all assemblies in 

vertical cask .......................................................................................................43

Figure 25 Change in peak neutron absorber temperature with failure of all assemblies in 

horizontal cask...................................................................................................44

Figure 26 GBC-32 criticality analysis results from pin pitch variation..................................45

Figure 27 GBC-68 criticality analysis results from pin pitch variation for channeled fuel.....45

Figure 28 GBC-68 criticality analysis results from pin pitch variation for unchanneled 

fuel.....................................................................................................................46

Figure 29 Change in peak cladding temperature vs. rod pitch to diameter ratio .................49

Figure 30 Recirculating mass flow rate vs. rod pitch to diameter ratio for vertical 

package.............................................................................................................50

Figure 31 GBC-32 criticality analysis results from axial misalignment................................51

Figure 32 GBC-68 criticality analysis results from axial misalignment................................51

Figure 33 Peak neutron absorber temperature variation from all fuel reconfigurations—

vertical cask .......................................................................................................66

Figure 34 Peak neutron absorber temperature variation from all fuel reconfigurations—

horizontal cask...................................................................................................67
xi

LIST OF TABLES

Page

Table 1 Case descriptions to evaluate the effects of breached spent fuel rods ................ 5

Table 2 Case descriptions to evaluate the effects of damaged fuel.................................. 6

Table 3 Parameters to evaluate the effects of rod/assembly deformation resulting 

from side/horizontal drop events in each technical discipline .............................11

Table 4 Parameters to evaluate the effects of rod/assembly deformation resulting 

from end/vertical drop events in each technical discipline ..................................12

Table 5 Parameters to evaluate the effects of fuel assembly components being 

axially displaced within the basket cavity of a transportation package in each 

technical discipline.............................................................................................17

Table 6 Criticality results for scenario S1(a)—breached spent fuel rods .........................21

Table 7 Combined rod removal and displaced fuel model results in GBC-32..................22

Table 8 Criticality results for scenario S1(b)—damaged spent fuel rods .........................22

Table 9 PWR package maximum dose rate change for NCT..........................................26

Table 10 BWR package maximum dose rate change for NCT..........................................28

Table 11 PWR package maximum dose rate change for HAC..........................................29

Table 12 BWR package maximum dose rate change for HAC..........................................30

Table 13 Maximum dose rate change at 1 m from a storage cask ....................................33

Table 14 Site boundary dose rate change ........................................................................34

Table 15 Thermal results for scenario S1(a)—breached spent fuel rods...........................38

Table 16 Single assembly failure results summary ...........................................................41

Table 17 Failure of all assemblies results summary..........................................................43

Table 18 PWR package maximum dose rate change for NCT..........................................47

Table 19 BWR package maximum dose rate change for NCT..........................................47

Table 20 Package maximum dose rate change for HAC...................................................48

Table 21 Rod/assembly deformation results summary......................................................49

Table 22 Criticality results for axial displacement of intact fuel..........................................52

Table 23 PWR package maximum dose rate change for NCT..........................................53

Table 24 BWR package maximum dose rate change for NCT..........................................54

Table 25 GBC-32 maximum dose rate change for HAC....................................................54

Table 26 Assembly axial alignment shift results summary ................................................56

Table 27 Fuel reconfiguration results summary pertaining to criticality .............................58

Table 28 Fuel reconfiguration results summary pertaining to shielding .............................60

Table 29 Summary of maximum thermal variations for vertical cask.................................64

Table 30 Summary of maximum thermal variations for horizontal cask.............................65

xiii

EXECUTIVE SUMMARY

This research performed a quantitative assessment of the safety impact of unlikely beyonddesign-basis hypothetical geometric changes of the fuel in spent fuel storage casks and 

transportation packages. The motivation of this study is, in part, to help understand the 

characteristics of the system’s responses to hypothetical fuel geometry changes. This study 

analyzed potential changes in system characteristics with respect to criticality, shielding, 

containment/confinement, and thermal safety parameters under a wide range of fuel 

reconfiguration scenarios. The following is a summary of the results of this study. 

For criticality, it was postulated that the cask/package was fully flooded under normal and 

accident conditions, although it is very unlikely. NUREG-2125 indicates that if an accident 

were to occur, there is about a one-in-a-billion chance that the accident would result in loss of 

containment, which is a prerequisite for flooding cask internals. The actual probability of 

flooding is even lower when one factors in the probability of such an accident in presence of a 

sufficient depth of water to enable full flooding of the cask internals, as was postulated in this 

study for criticality. The assessment results showed that the majority of the evaluated 

pressurized water reactor (PWR) and channeled boiling water reactor (BWR) fuel 

reconfigurations result in increases in system reactivity, keff, less than 4 %∆keff when the 

reconfigured fuel remains inside the neutron absorber panel envelope. For reconfigurations 

with fuel materials beyond the neutron absorber panel envelope and others involving 

unchanneled BWR fuel, the system reactivity varies widely, ranging from decreases to 

increases greater than 5 %∆keff, depending on the water-to-fuel ratio and length of fuel 

outside the absorber envelope. However, the results are dependent on the modeling 

assumptions and canister characteristics, particularly the available volume above and below 

the neutron absorber panel envelope for the fuel rods and failed fuel fragments to 

accumulate. This indicates that using axial spacers to limit axial movement of fuel is an 

effective measure to maintain criticality safety under such conditions. 

With respect to shielding design, the results showed that fuel redistribution within the central 

axial region of the fuel basket results in relatively small changes (up to ~10%) in the 

maximum dose rates at the transportation package external surfaces relative to the nominal 

intact fuel configuration. However, fuel redistribution toward the package internal cavity 

bottom and/or top regions would significantly increase (factor of 2.1 to 23.6) the dose rates at 

the transportation package external surfaces. For storage systems, fuel configuration 

changes can also cause a significant increase (factor of 1.2 to 9.6) in dose rates near the 

storage cask air vents. The results indicate that using axial spacers to limit axial movement of 

fuel is an effective measure to limit the impact on dose rates under such conditions. At 

locations away from storage cask air vents, the change in radiation dose rate is insignificant. 

The allowable leakage rates for a storage cask and transportation package were analyzed for 

postulated reconfigurations under various assumptions of numbers of breached fuel rods and 

various release fractions. The results showed that the allowable leakage rate exhibits greatest 

sensitivity to changes in the fractions of fuel released as fuel fines due to cladding breaches. 

The results also show that the fractional contribution to radioactive material release from the 

pellet region increases as fuel burnup increases. Fuel released as fuel fines from the highburnup rim structure (i.e., rim pellet region) yields a smaller allowable leakage rate than the 

fuel released as fuel fines from the non-rim pellet region. Because the allowable leakage rate 

is a function of decay-time, increases in fuel rod failure rates are offset by longer decay times. 

The results indicate that for the BWR fuel, the same allowable leakage rate can be obtained 
xiv

for failed fuel fractions of 0.03, 0.4, and 0.7 at 5-, 40-, and 100-year decay times, respectively. 

The results indicate that for PWR fuel, the same allowable leakage rate can be obtained for 

failed fuel fractions of 0.03, 0.10, and 0.15 at 5-, 40-, and 100-year decay times, respectively. 

With respect to thermal safety, the results show that postulated fuel reconfiguration resulting 

in relocation of the fuel pellets had the largest impact on the temperatures of the internal 

components of the cask. Cladding breach without fuel pellet relocation has a minor impact on 

component temperature for vertically oriented packages that rely on convective heat transfer 

because release of fission gases improves heat transfer. For bounding configurations in a 

vertical cask where the failed fuel was represented as a particle bed at the bottom of each 

basket cell, the maximum increase in component temperatures observed was 128 ̊C relative 

to the nominal intact fuel configuration. In horizontal casks that rely on conductive heat 

transfer, the release of the fission product gases from all assemblies would increase the 

internal component temperatures by about 42 ̊C, and for particle bed configurations, the 

maximum increase observed in component temperatures was 31 ̊C relative to the nominal 

intact fuel configuration. Depending on the timing, the thermal impacts of fuel failure may be 

offset by the decreased heat load of the fuel because the decay heat load is decreasing as a 

function of decay time. Results evaluated showed decreases in component temperatures by 

greater than 220 ̊C between 20 and 60 year decay times. Axial shifting of the assemblies and 

variations in the rod pitch has minimal impact on both internal component temperatures and 

the external surface temperature. 

Overall, the safety impacts of fuel reconfiguration are system design, content type, and 

loading dependent. The areas and magnitude of the impact vary from cask/package design to 

cask/package design. It should also be noted that some of the scenarios are extreme and 

physically unlikely to occur; they represent bounding values. The spent fuel storage systems 

and transportation packages approved by the NRC to date provide reasonable assurance that 

they are safe under normal, off-normal, and hypothetical accident conditions as prescribed in 

10 CFR Part 71 and 72 regulations.
xv

ACKNOWLEDGMENTS

The authors wish to thank Michelle Flanagan, Project Manager, and Mourad Aissa of the U.S. 

Nuclear Regulatory Commission’s Office of Regulatory Research; and Zhian Li and Meraj 

Rahimi of the NRC Office of Nuclear Materials Safety and Safeguards for their help in initiating, 

planning, and implementing this research effort. The authors also wish to thank John Wagner 

and Steve Bowman for their reviews, and Debbie Weaver, Katherine Ragle, and Wendy Hames

for assistance in formatting and preparing the final document. 

xvii

ACRONYMS

ANS American Nuclear Society

ANSI American National Standards Institute

BWR boiling water reactor

CFR Code of Federal Regulations

COBRA-SFS Coolant Boiling in Rod Arrays—Spent Fuel Storage

CSAS Criticality Safety Analysis Sequence

DBTT ductile-to-brittle transition temperature

EALF Energy of average lethargy of a neutron causing fission

ENDF/B Evaluated Nuclear Data Files, Part B

F/I failed to intact configuration dose rate ratio

GBC generic burnup credit

GE General Electric

HAC hypothetical accident conditions

IGSCC intergranular stress corrosion cracking

ISG Interim Staff Guidance

keff effective neutron multiplication factor

LWR light water reactor

MAVRIC Monaco with Automated Variance Reduction using Importance 

Calculations

MPC multipurpose canister

NA not applicable

NCT normal conditions of transport

NRC Nuclear Regulatory Commission

NUREG NRC technical report designation (Nuclear Regulatory Commission)

OFA optimized fuel assembly

ORIGEN-ARP Oak Ridge Isotope Generation – Automatic Rapid Processing

ORIGEN-S Oak Ridge Isotope Generation in Scale

ORNL Oak Ridge National Laboratory

PWR pressurized water reactor

SNF spent nuclear fuel

STARBUCS Standardized Analysis of Reactivity for Burnup Credit using Scale

TRITON Transport Rigor Implemented with Time-Dependent Operation for 

Neutronic depletion

W Westinghouse

wppm parts per million by weight

xix

DEFINITIONS

Assembly defect Any change in the physical as-built condition of the assembly with the 

exception of normal in-reactor changes such as elongation from irradiation 

growth or assembly bow. Examples of assembly defects: (a) missing rods; 

(b) broken or missing grids or grid straps (spacers); and (c) missing or 

broken grid springs, etc. An assembly with a defect is damaged only if it 

can't meet its fuel-specific and system-related functions required by the 

applicable regulations. [1]

Breached spent

fuel rod

Spent nuclear fuel (SNF) rod with cladding defects that permit the release 

of gas from the interior of the fuel rod. A breached spent fuel rod may also 

have cladding defects sufficient to permit the release of fuel particulate. A 

breach may be limited to a pinhole leak or hairline crack, or may be a gross 

breach. [1]

Damaged spent 

nuclear fuel

Any fuel rod or fuel assembly that cannot fulfill its fuel-specific or systemrelated functions. [1]

Grossly 

breached spent 

fuel rod

A subset of breached rods. A breach in spent fuel cladding that is larger 

than a pinhole leak or a hairline crack. An acceptable examination for a 

gross breach is a visual examination that has the capability to determine 

the fuel pellet surface may be seen through the breached portion of the 

cladding. [1]

Intact spent 

nuclear fuel

Any fuel that can fulfill all fuel-specific and system-related functions, and 

that is not breached. Note that all intact SNF is undamaged, but not all 

undamaged fuel is intact, since under most situations, breached spent fuel 

rods that are not grossly breached will be considered undamaged. [1]

Normal events 

and conditions

The maximum level of an event or condition expected to routinely occur. 

The cask system is expected to remain fully functional and to experience 

no temporary or permanent degradation from normal operations, events 

and conditions. [2]

Off-normal 

events

and conditions

The maximum level of an event or condition that although not occurring 

regularly can be expected to occur with moderate frequency and for which 

there is a corresponding maximum specified resistance, limit of response, 

or requirement for a given level of continuing capability. “Off-Normal” 

events and conditions are similar to “Design Event II” of ANSI/ANS 57.9. 

An independent spent fuel storage installation structure, system, or 

component is expected to experience off-normal events and conditions 

without permanent deformation or degradation of capability to perform its 

full function (although operations may be suspended or curtailed during offnormal conditions) over the full license period. [2]

Ready retrieval The ability to move a canister containing spent fuel to either a

transportation package or to a location where the spent fuel can be

removed. Ready retrieval also means maintaining the ability to handle

individual or canned spent fuel assemblies by the use of normal means. [3]

Undamaged 

spent nuclear 

fuel

SNF that can meet all fuel-specific and system-related functions.

Undamaged fuel may be breached. Fuel assembly classified as

undamaged SNF may have “assembly defects.” [1]

1

1. INTRODUCTION

In the United States, spent nuclear fuel (SNF) is expected to remain in storage for an extended 

period of time. The spent fuel will then be transported to a final repository or interim storage 

facilitie(s). One of the possible concerns associated with spent fuel storage and transportation 

safety is cladding material hydride reorientation because an elevated amount of hydride 

reorientation in the cladding material could impair the mechanical performance of fuel assembly. 

This potential phenomenon is of particular concern with high burnup fuel (HBU, >45 GWd/MTU) 

because the potential of hydride reorientation increases as fuel burnup increases. To obtain 

material property data and assess the mechanical performance of high burnup cladding, 

Argonne National Laboratory examined the material property of some samples from high burnup 

fuel rods with static tests [Ref. 1, 2] and Oak Ridge National Laboratory conducted bending 

vibration tests [Ref. 3, 4] on high burnup fuel rods. The initial results of static and vibration tests 

of HBU Zircaloy-4 cladding have confirmed that HBU fuel is structurally robust. Based on results 

of these tests, it is determined that there is reasonable assurance that spent nuclear fuel, 

including high burnup (burnup > 45 GWd/MTU) fuel, is safe for storage and transport under 

normal, off-normal, and hypothetical accident conditions as prescribed in 10 CFR Part 71 and 

Part 72. Note that these confirmatory tests are not inclusive for all cladding materials. Also, the 

fuel assemblies will cool off over time and the external environment may change as well. Both 

factors may affect the mechanical performance of the fuel cladding and the fuel assembly as a 

whole after an extended period of storage. 

This study performed a quantitative assessment of the impacts of fuel geometry changes on 

criticality, shielding, containment/confinement, and thermal safety of SNF storage casks and 

transportation packages under an unlikely hypothetical event of fuel reconfiguration. 

Specifically, this study examined the potential changes of these characteristics of the systems 

resulting from changes in fuel assembly geometry in comparison with the nominal intact fuel 

configuration values of these parameters. 

This research has two objectives. The first one is to gain a better understanding of the safety 

performance characteristics of the storage casks and transportation package with respect to 

change in fuel geometry. The second one is to obtain insights regarding areas of the cask and 

package where design enhancements would result in maximum safety benefits. 

Three fuel reconfiguration categories were considered and they were characterized by: (1) 

cladding failure, (2) rod/assembly deformation without cladding failure or (3) changes to 

assembly axial alignment without cladding failure. The analyses considered representative SNF 

designs and storage cask/ transportation packages, and a range of fuel initial enrichments, 

discharge burnup values, and decay times. The analyses encompass the impact from normal,

off-normal, and accident conditions of storage and transportation on the fuel and fuel assembly 

structures. The impact on the storage cask/ transportation package system mechanical and 

structural components was beyond the scope of this analysis. 

Under the above-mentioned three categories, a wide range of fuel reconfiguration scenarios 

were analyzed. The generic pressurized water reactor (PWR) and boiling water reactor (BWR) 

storage cask and transportation package models used in previous studies [4, 5] were adapted 

for use in the analyses. The PWR models contain thirty two (32) 17×17 PWR fuel assemblies 

representative of a Westinghouse (W) optimized fuel assembly (OFA) design and the BWR 

models contain sixty eight (68)10×10 BWR fuel assemblies representative of a General Electric14 (GE14) design. Both the PWR and BWR models are representative of high-capacity-type 

casks/packages, and are referred to as generic burnup credit (GBC)-32 and GBC-68, 
2

respectively. Different initial fuel enrichments (1.92 to 5.0 wt % 235U), burnups (0 to 70 

GWd/MTU), and decay times (5 to 300 years) were considered in the analyses. The impacts on 

system safety in each technical discipline with respect to fuel configuration changes were 

assessed relative to the corresponding nominal intact fuel configuration. However, there was no 

study or comparison of mechanical properties, finite element analysis stress distribution, or 

physical testing to evaluate the likelihood of each reconfiguration scenario. The likelihood of 

any particular fuel reconfiguration and the impact of that configuration on each technical 

discipline are dependent on many factors, and will need to be addressed on a specific basis for 

each storage and transportation system design. In certain instances, a reconfiguration scenario 

used in this study may be extreme (i.e. not physically likely). In those cases, the reconfiguration 

scenarios are used to represent a worst case scenario rather than a physical realization. 

The baseline intact configuration consists of 0% failed fuel (i.e., intact fuel conditions). The 

fuel assemblies are assumed to be positioned in the center of the fuel cells of the baskets for 

criticality and thermal evaluations. The fuel and assembly materials are assumed to be 

homogenously distributed inside each fuel cell for shielding and containment analyses. 

Specific cases evaluated for each category were tailored with respect to each technical 

discipline to reflect the extent of mechanical considerations governed by the requirements of 

the separate nuclear safety analyses. For example, in the criticality evaluations, the 

cask/package is modeled as fully flooded under normal and accident conditions, although it is 

very unlikely (i.e., conditional probability of accident that results in loss of containment breach 

that could allow flooding is estimated to be 1.08×10-9 [6] which is consistent with other studies 

[i.e., 7.8×10-9 ] [7]). The actual probability of flooding is even lower when one factors in the 

probability of such an accident in presence of a sufficient depth of water to enable full flooding 

of the cask internals.

Definitions of the various fuel reconfiguration categories selected and the basis for choosing 

them are provided in Sect.2. Pertinent results obtained from these analyses are provided in 

Sect.3; conclusions and recommendations are provided in Sect.4. Detailed descriptions of the 

cases developed to represent the different configurations for the technical disciplines criticality, 

shielding, containment, and thermal are provided in Appendixes A, B, C, and D, respectively.
3

2. FAILED FUEL ASSEMBLY CONFIGURATION CATEGORIES 

To evaluate the potential safety implications of fuel configuration changes that can result from 

normal, off-normal, and accident conditions of spent fuel storage, and normal and accident 

conditions of transportation, three fuel reconfiguration categories were considered:

(Category 1) cladding failure

(Category 2) rod/assembly deformation without cladding failure

(Category 3) changes to assembly axial alignment without cladding failure 

Within configuration categories 1 and 2, multiple scenarios were identified:

Cladding failure category

Scenario 1(a)—breached spent fuel rods

Scenario 1(b)—damaged spent fuel rods

Rod/assembly deformation category

Scenario 2(a)—configurations associated with side drop

Scenario 2(b)—configurations associated with end drop

The configuration category of assembly axial alignment changes was not broken into multiple 

scenarios; instead, Scenario 3 encompasses all changes to assembly axial alignment. 

Each technical discipline has specific analyses that are used to demonstrate cask system 

performance under normal, off-normal, and accident conditions of spent fuel storage, and 

normal and accident conditions of transportation. Therefore, parameters were developed and 

analyzed for each scenario that would have the most significant implications in each technical 

discipline. 

The following subsections will discuss each fuel reconfiguration category, including the safetysignificant parameters important for the four technical disciplines, example initiating events that 

could result in the subject configurations, and the assumptions and conservatisms used in the 

analysis for the subject configurations. 

2.1 CATEGORY 1: CLADDING FAILURE 

The structural integrity of the cladding for high-burnup fuel may become challenged under a 

variety of cladding degradation mechanisms [8, 9, 10]. Several of the identified mechanisms 

that could lead to degradation of cladding structural integrity are driven by mechanical property 

changes because of hydrogen-related phenomena, including hydride embrittlement, delayed 

hydride cracking, and hydride reorientation. Hydrogen uptake by clad during reactor operation is 

a known phenomenon that affects the mechanical properties of the cladding. The hydrogen 

typically remains in solution at elevated temperatures and precipitates out during cooling as 

platelets primarily oriented in the circumferential direction [11, 12]. During spent fuel caskdrying operations, cladding temperatures become elevated because of decreased heat removal 

because of the removal of water. At elevated temperatures some of the hydrogen goes into 

solution while the rod-internal-pressure-induced stress increases relative to in-reactor operation 

(i.e., reactor primary coolant is pressurized). As the cladding cools under stress, some of the 
4

dissolved hydrogen reprecipitates in the radial direction across the cladding wall if the hoop 

stress is high enough. This phenomenon is known as hydride reorientation. Radial hydrides can 

degrade the ductility of the cladding. After cooling to about 200°C, most of the dissolved 

hydrogen will have reprecipitated as hydrides, with further cooling during storage potentially 

resulting in radial-hydride-induced embrittlement. The temperature at which embrittlement 

occurs is referred to as the ductile-to-brittle transition temperature (DBTT) [11].

The cladding temperature of fuel that has been in storage for a long time may drop below DBTT. 

Once the cladding temperature goes below DBTT, the fuel rods are more susceptible to failure 

under load impacts. In addition, measured data are limited and typically are characterized by a 

relatively high level of uncertainty regarding mechanical performance of high-burnup fuel 

cladding. Given the uncertainty and variability of fuel rod material properties, different degrees 

of fuel rod cladding failure followed by release of material into the canister cavity during normal 

handling and transfer operations before and during transportation should be considered. 

Grossly breached spent fuel rods within the canister potentially may occur under the impacts of 

natural phenomena such as earthquakes or tornados, resulting in cask tip-over. Horizontal 

drops can induce pinch forces that can result in longitudinal tearing as described in Ref. [13]

and illustrated in Figure 1 resulting in fuel particle relocation. During transportation, if the 

cladding temperature is below DBTT, failure of fuel rods within the package may occur under 

normal conditions of transport (NCT) as prescribed in 10 CFR 71.71, as well as a result of 

impacts from accident loads specified in 10 CFR 71.73, particularly the 30-ft drop and puncture 

testing in the orientations that cause the most damage to the package containment vessel.

Figure 1 Illustration of fuel rod failure mechanisms [13]
5

Analyses of the cladding failure category was designed to represent the effects of two 

scenarios: S1(a), breached spent fuel rods where the cladding has failed to the extent to allow 

for a loss of gas and fuel particles from single or multiple locations with the rod segment and 

fuel fragments collecting at different locations within the cask, and S1(b), damaged spent fuel 

rods where the cladding has failed to the extent to allow free movement of fuel particles and 

pellets within a basket cell. Damaged SNF is a bounding condition that represents the maximum 

degree of fuel configuration change that may occur under cladding failure scenarios. 

Safety-significant parameters with respect to this category for each scenario in each technical 

discipline are provided in Table 1 and Table 2. 

Table 1 Case descriptions to evaluate the effects of breached spent fuel rods

Technical 

discipline

Configuration 

scenario

Parameter(s) 

evaluated Rationale for parameter selection

Criticality S1(a) Lattice positions 

from where fuel 

particulate could be 

displaced 

Commercial light water reactor (LWR) fuel assemblies 

are designed to be undermoderated. Fuel rods that 

suffer cladding failure can result in sections of fuel 

within the lattice being replaced with water, which 

increases moderation and potentially increases 

reactivity. The issue of concern is whether the 

reactivity increase exceeds the original licensed 

design basis subcritical limit.

Shielding S1(a) Fraction of spent 

fuel redistributed

and canister basket 

cavity regions 

where particulate 

accumulates

Storage and transportation overpacks are designed to 

maintain dose rates within acceptable limits assuming 

known source distributions based on intact 

assemblies. Clad failure can result in relocation of the 

fuel mixture and associated radiation source terms, 

thereby potentially affecting the design basis external 

dose rates.

Containment S1(a) Fraction of 

breached spent fuel 

rods; in addition for 

high-burnup fuel, 

release fractions for 

the contributors to 

the releasable 

activity and pellet 

region from which 

the radioactive 

material originates

Evaluate impacts on allowable leakage rates of 

different spent fuel rod breach fractions for NCT and 

releasable source term distribution in high-burnup fuel 

caused by rim effect. A sensitivity analysis is

performed for high-burnup fuel because the range of 

parameters important for the containment analysis 

(i.e., the release fractions and breached spent fuel rod

fraction) has not been established yet for NCT.

Thermal S1(a) Fraction of spent 

fuel rods 

experiencing 

cladding failure that 

release fission 

product and rod 

backfill gases 

(varied from 0

to100%)

The release of fuel rod inert gas and fission product 

gases affects the thermophysical properties of the 

canister gas space. The change in gas properties can 

affect the efficiency of heat removal, and therefore 

component temperatures, to varying degrees 

depending on the cask/package design and 

orientation. The sensitivity analysis will investigate the 

impact on component temperatures.
6

Table 2 Case descriptions to evaluate the effects of damaged fuel

Technical 

discipline

Configuration 

scenario

Parameter(s) 

evaluated Rationale for parameter selection

Criticality S1(b) Geometry changes 

and modeling 

homogenous 

versus 

heterogeneous 

representations of 

fuel debris mixture

Geometry and neutron energy spectrum are 

fundamental parameters used to define a design basis 

configuration. Any changes in these parameters will 

affect the system keff. These cases evaluate the 

effects of having an uncontrolled geometry, and the 

effects of different modeling simplifications to 

represent regions and packing fractions where 

particulate could relocate.

Shielding S1(b) Regions where fuel 

particles could 

redistribute 

Fuel particle relocation has a potential of increasing 

cask external dose rates at the top or bottom because 

of its proximity to those surfaces and as a result of 

intensifying the source term where the fuel has 

relocated. This can also affect the dose at the

controlled area boundary. In addition, neutron 

streaming above or below the radial neutron shield in 

a transportation package may significantly increase 

the neutron dose rate on the package external radial 

surface. 

Containment S1(b) For high-burnup 

fuel, varying 

release fractions for 

the contributors to 

the releasable 

activity and pellet 

region from which 

the radioactive 

material originates

Evaluate impacts on allowable leakage rates for 

hypothetical accident conditions (HAC) considering 

releasable source term distribution in high-burnup fuel 

caused by the rim-effect. A sensitivity analysis is

performed for high-burnup fuel because the range of 

parameters important for the containment analysis 

(i.e., the release fractions) has not been established 

yet for HAC.

Thermal S1(b) The number of 

assemblies (1 or 32 

[all]) and the 

packing fraction of 

the debris (0.612–

0.313) to 

investigate the 

impact of fuel

redistribution on

component 

temperatures.

Changes to the internal canister gas properties as well 

as location and geometry of the heat source within the 

canister can alter the convective flow paths within the 

canister. These phenomena can affect the efficiency 

of heat removal from the cask/package and therefore 

the temperatures of components important to safety. 

2.2 CATEGORY 2: ROD/ASSEMBLY DEFORMATION

Depending upon the fuel temperature and mechanical properties of the fuel assembly, a range 

of potential internal configurations are possible when a spent fuel package is involved in an 

impact event. The side, end, and corner drops may result in geometry changes depending on 

the orientation of the cask during the drop, the magnitude of the impact, and the mechanical 

properties of the fuel assembly components. 

Packages used for storage and transportation have been certified via analysis or testing and, in 

many instances, by using a combination of analysis and testing. A number of studies and tests 
7

have been performed over the years to investigate the impact of NCT and HAC on fuel 

assemblies that are contained within a transportation package [13, 14, 15, 16, 17, 18]. 

Additionally, in the early 2000s, TN International and International Nuclear Services started a 

joint program, the Fuel Integrity Project, to assess potential damage to fuel assemblies and 

confirm assumptions used in criticality-safety studies [15]. Overall, analyses and tests have 

indicated that during horizontal (side) drop, the fuel rods are primarily subjected to loads that 

can result in fuel rod bending or some degree of plastic deformation when the cladding is still 

ductile so that the lattice pitch of the fuel assemblies tends to reduce. In a vertical (end) drop 

orientation, the axial loading can lead to buckling of the fuel rods. Several pictures showing the 

resultant configurations of “dummy” fresh fuel assemblies after drop tests are presented in 

Figure 2 and Figure 3. The difference between actual fuel assemblies and the “dummy” fuel 

assemblies is that actual enriched UO2 fuel pellets were substituted with depleted UO2 pellets 

with similar mechanical properties [15]. Note that these drop tests are from fuel contained in a 

fresh fuel package, which experiences higher g-loads than a spent fuel package would receive, 

but they are illustrated here to show how fuel may deform after an impact event. 

(a) PWR fuel assembly (b) BWR fuel assembly

Figure 2 Photographs of dummy fuel assemblies after 9-m end drops [15]
8

Figure 3 Photographs illustrating assembly fuel rod buckling after a 9-m end drop test

[19]

Effects that can influence material strength and structural integrity of the cladding and fuel 

assembly include neutron fluence (e.g., grid spring relaxation, irradiation hardening, growth, 

cladding creep down), corrosion (e.g., thinning, oxidation, hydrogen uptake), operating 

conditions (e.g., temperature), and drying conditions (e.g., temperature, residual moisture). 

Analytical methods for calculating load responses and characteristics of fuel rods during and 

after impact events require assumptions that are difficult to determine, such as percentage of 

fuel mass that is bonded to or participates with the cladding during the buckling process as 

discussed in Ref. [20]. Additionally, BWR and PWR fuel assemblies are designed differently, 

and some of these differences result in different mechanical responses on the fuel rods under 
9

impact events. Representative BWR and PWR fuel assemblies are shown in Figure 4 to 

illustrate some of the design differences. Besides the BWR fuel assembly being channeled, the 

fuel rods are connected to upper and lower tie plates. In the PWR fuel assembly, the fuel rods 

are not directly connected to the upper and lower end fittings, leaving a small gap between the 

ends of the rods and the end fittings. Under horizontal drop events, this design difference does 

not result in noticeable differences on fuel rod response between a PWR and BWR fuel 

assembly; however, it can alter the response under vertical drop events [15]. This fuel 

reconfiguration category investigates the impact of pin pitch changes that could result as 

potential end states associated with side/horizontal and end/vertical drop events consistent with 

NCT as prescribed in 10 CFR 71.71, as well as a result of impacts from accident loads specified 

in 10 CFR 71.73.

Figure 4 Representative BWR and PWR fuel assemblies [21]

Analyses in this fuel reconfiguration category evaluate the impact of fuel rod and assembly 

deformation when the fuel cladding is able to absorb the loads of the initial impact event and 

remain intact (i.e., cladding does not fail). All configurations that involve cladding failure are 

evaluated in the cladding failure category (see Sect. 2.1). Two fuel reconfiguration scenarios are 

considered: S2(a), configurations associated with side/horizontal drop; and S2(b), configurations 

associated with end/vertical drop.

(a) BWR fuel assembly (b) PWR fuel assembly
10

Schematic illustrations of the resultant changes to the fuel assembly lattice geometry 

considered for a horizontal drop event are illustrated in Figure 5 and Figure 6 for a PWR and a 

BWR package, respectively. The parameters identified to represent the most significant 

implications of the horizontal drop event in each technical discipline are provided in Table 3.

Figure 5 Schematic representation of PWR fuel assembly before and after horizontal 

drop event (figure adapted from Ref. [15])

Figure 6 Schematic representation of BWR fuel assembly before and after horizontal 

drop event (figure adapted from Ref. [15])

Nominal PWR 

assembly

PWR assembly after 

horizontal drop

Nominal BWR 

assembly

BWR assembly after 

horizontal drop
11

Table 3 Parameters to evaluate the effects of rod/assembly deformation resulting from 

side/horizontal drop events in each technical discipline

Technical 

discipline

Configuration 

scenario

Parameter(s) 

evaluated Rationale for parameter selection

Criticality S2(a) Pin pitch 

contraction

Evaluate impacts of changes to design basis 

neutron energy spectrum

Shielding S2(a) Assembly lattice 

collapse

Localized source displacement toward cask 

radial surface can affect radial dose rate 

requirements as a result of reduced 

geometric attenuation

Containment S2(a)

Fraction of crud 

that spalls off 

cladding (varied 

from 0.05 to 1.0)

Evaluate sensitivity to fraction of crud 

removed. 

Thermal S2(a) Pin pitch 

contraction

The contraction of the fuel assembly lattice 

can affect the efficiency of heat removal from 

the fuel because of changes in flow area 

(convection), conduction lengths 

(conduction), and radiation view factors 

(thermal radiation). 

Schematic illustrations of the resultant changes to the fuel assembly lattice geometry 

considered for a vertical drop event are illustrated in Figure 7 and Figure 8. Note that 

compaction is shown for the BWR assembly (also referred to as bottlenecking), and expansion 

is shown for the PWR assembly (also referred to as birdcaging), but a combination of the two 

along the axial length of the fuel assembly is also considered. The parameters identified to 

represent the most significant implications of the vertical drop event in each technical discipline 

are provided in Table 4.

(a)PWR assembly before and after end drop (b) BWR assembly before and after end drop

Figure 7 Schematic representation of PWR and BWR fuel assemblies before and after a 

vertical drop event (figure adapted from Ref. [15])
12

(a) PWR assembly (b) BWR assembly

Figure 8 Example configuration for non-uniform pitch expansion

Table 4 Parameters to evaluate the effects of rod/assembly deformation resulting from 

end/vertical drop events in each technical discipline

Technical 

discipline

Configuration 

scenario

Parameter(s) 

evaluated Rationale for parameter selection

Criticality S2(b)

Uniform and nonuniform radial and 

axial pin pitch 

changes (birdcaging 

and bottlenecking)

Evaluate impacts of changes to design basis geometry 

and neutron energy spectrum

Shielding S2(b) Not applicable

Bounding configuration for this category is scenario S3 

(Table 5) where fuel assembly axial displacement 

toward the bottom lid is evaluated.

Containment S2(b)

Fraction of crud that 

spalls off cladding 

(varied from 0.05 to 

1.0)

Evaluate sensitivity to fraction of crud removed. 

Thermal S2(b) Assembly pin pitch 

expansion

The expansion of the fuel assembly lattice can affect

the efficiency of heat removal from the fuel because of

changes in flow area (convection), conduction lengths 

(conduction), and radiation view factors (thermal 

radiation).
13

2.3 CATEGORY 3: CHANGES TO ASSEMBLY AXIAL ALIGNMENT 

Fuel baskets used in typical storage and transportation packages contain neutron absorbers 

affixed to the basket cell walls. Most of the absorbers present in currently deployed spent fuel 

canisters are in plate form that does not extend the full length of the basket; they are held in 

place by a thin gauge stainless steel sheath. A schematic representation of a fuel assembly 

within a basket cell is illustrated in Figure 9. Some of the more recent basket designs made of 

metal matrix composites integrate the absorber into the basket material, in which case the 

absorber does extend the full length of the basket [22]. 

Different types of overpacks are used for storage and transportation. The storage system 

typically consists of a thick storage overpack made of steel, concrete, or a combination of the 

two that fully encompasses the spent fuel canister. Figure 10 illustrates a typical storage 

overpack configuration. For transportation, the overpack typically consists of a layered shell with 

several different materials to provide shielding for gamma rays and neutrons, as well as to 

provide a means for heat removal. The axial extent of the package radial neutron shielding does 

not always extend the full length of the containment vessel because the package needs an 

allowance for attaching the impact limiters on the ends. Hence, the cavity volume of the canister 

may not be fully covered with shielding material. Figure 11 illustrates a typical configuration for a 

rail spent fuel transportation package.
14

.

Figure 9 Schematic representation of fuel assembly within a typical basket cell

 
15

Figure 10 Example storage pack configuration for both vertical (a) and horizontal (b) 

storage modules [23]

Figure 11 Example rail spent fuel transportation package [24]

The safety features of the canister fuel basket and the transportation overpack (e.g., neutron 

absorber plates, gamma and neutron shield, and cooling fins) typically have been designed 

presuming that the fuel assembly remains in a fixed geometric location within the package

under normal and accident conditions of transportation. Fuel spacers are designed to restrain 

the fuel assembly within the basket cell to ensure axial alignment of the active fuel region within 

the neutron absorber envelope, as well as to provide a fixed source location within the extent of 

the transportation overpack. 

(a) Illustration of vertical aboveground 

concrete or steel structure.

(b) Illustration of horizontal aboveground 

concrete bunker.
16

While the degree of axial movement should be limited, transportation package testing has 

shown that post-buckling bending deformations, regions of lattice expansion, and interaction 

with deformed nozzles, such as illustrated in Figure 2(a) above, can result in some fuel rods 

sliding past the end fitting, resulting in a loss of axial geometry control from the as-designed 

configurations. Other drop test results such as that shown in Figure 12 illustrate how the end 

region of an assembly can become crushed. 

Figure 12 Photo of BWR assembly after drop test [25]

Additionally, the presence of some residual moisture in the canister after drying is expected, and 

can promote corrosion and/or stress corrosion cracking of fuel assembly hardware components 

while the SNF is in dry storage. Intergranular stress-corrosion cracking (IGSCC) is a known 

failure mechanism that can result in dislocation of the top nozzle end fitting from the remainder 

of the assembly [26], leaving space for fuel rod axial shifting.

Finally, sequences that can lead to axial movements include handling operations while 

preparing a package for transportation, under NCT as prescribed in 10 CFR 71.71, and HAC 

loads as specified in 10 CFR 71.73. These phenomena (i.e., end fitting deformation, IGSCC, 

lattice expansion), individually or in combination indicate that changes to the axial alignment of 

the fuel within the cask should be considered and are being evaluated with regards to the 

potential effects on the safety functions.

Analyses in this fuel reconfiguration category evaluate the impact of changes to assembly axial 

alignment assuming the fuel cladding is able to absorb the loads of the initial impact event and 

remain intact (i.e., cladding did not fail). All configurations that involve cladding failure are 

evaluated in the cladding failure category (see Sect. 2.1). Safety-significant parameters with 

respect to this category for each technical discipline are provided in Table 5.
17

Table 5 Parameters to evaluate the effects of fuel assembly components being axially 

displaced within the basket cavity of a transportation package in each 

technical discipline

Technical 

discipline

Configuration 

scenario Parameter(s) evaluated Rationale for parameter selection

Criticality S3 Fuel assembly axial position Evaluate effects of active fuel region 

being outside neutron absorber plate 

envelope in basket cell.

Shielding S3 Fuel assembly axial position This configuration has the potential to 

significantly increase dose rate at the 

cask top and bottom surfaces as well as 

at the radial surface above and below 

the neutron shield because the radiation 

sources are moved closer to these 

surfaces.

Containment S3 Fraction of crud that spalls off 

cladding (varied from 0.05 to 

1.0)

Evaluate sensitivity to fraction of crud 

removed. 

Thermal S3 Fuel assembly axial position Axial shifting of the assembly changes 

the heat source location within the 

canister. This can affect the heat 

removal via convection within vertical 

casks/packages. Shifting of assemblies 

is likely to have a minor impact on 

component temperatures. However, 

cases were investigated where all the 

assemblies are fully shifted axially 

within their respective basket cells to 

investigate the impact on component 

temperatures.

19

3. CONSEQUENCE ANALYSIS

This section addresses the consequences of fuel configuration changes with respect to the 

different technical disciplines. Nominal intact fuel configuration cases were developed for each 

technical discipline against which the impacts of the fuel reconfiguration could be measured. 

The nominal intact fuel configuration models and corresponding analysis approach are 

described in Appendixes A, B, C, and D for the criticality, shielding, containment, and thermal 

analyses, respectively. The consequences are reported as relative changes between the fuel 

reconfiguration and the nominal intact configuration. Criticality consequences are associated 

with a change in keff, shielding consequences are associated with changes to external dose 

rates, containment consequences are associated with release rates, and thermal consequences 

are concerned with component (e.g., seals, cladding, basket, neutron absorber, and surface) 

temperature changes. For all analyses, it was assumed that the containment barrier had not 

been breached for a sufficient length of time to make oxidation of structures, systems, or 

components a consideration. The basis for this assumption is that this allows the effects of 

geometry changes to be evaluated individually for impact on system performance. 

The following descriptions apply when referring to a canister, cask, or package: The spent fuel 

is emplaced within a canister (as shown in Figure 10) which is then either represented as being 

surrounded by a concrete storage cask (see Figure 10) or a multi-layered transportation cask

(see Figure 11). The combined canister and transportation cask is referred to as a 

transportation package. The terms GBC-32 and GBC-68 identify whether the canister contains 

PWR or BWR fuel, respectively. 

3.1 CATEGORY 1: CLADDING FAILURE

Cladding failure includes breached spent fuel rods, grossly breached spent fuel rods, and 

damaged fuel. This fuel reconfiguration category encompasses the greatest amount of 

variability for consideration when developing explicit models.

3.1.1 Criticality

The consequences of cladding failure with respect to criticality control require criticality 

calculations to investigate the changes associated with both grossly breached and damaged 

spent fuel geometries. To have an effect on system performance the SNF rods must be grossly 

breached such that the geometry and neutron energy spectrum present within the SNF canister 

are altered from the design basis configuration. The primary requirements for demonstrating 

subcriticality include 10 CFR 71.55(b), (d), and (e). 

Because of the effects of geometry changes, the cladding failure category can result in a large 

number of potential configurations with a wide range of reactivity effects. Two configuration 

scenarios were developed to evaluate the effects of cladding failure: S1(a), breached spent fuel 

rods; and S1(b), damaged spent fuel rods. 

Breached spent fuel rods. To evaluate the effects of breached spent fuel rods, including 

grossly breached spent fuel rods, cases were developed to model the effects of rod failure and 

changes in moderator-to-fuel ratios. Multiple rod removal patterns were evaluated by removing 

groups of rods in symmetric patterns until a peak keff value was identified. Figure 13 shows the 

results for the GBC-32 cases and Figure 14 shows the results for the GBC-68 cases. For these 

two sets of cases, the displaced fuel was omitted from the model, assuming that it would be 
20

sufficiently dispersed away from the lattice in a thin, undermoderated heap resulting in a less 

reactive system than if incorporated into the model. The maximum ∆keff value (i.e., a keff

increase of 1.87% ∆keff) for the GBC-32 cask representation occurs for the 5 wt % 235U initial 

enriched fuel with a burnup of 44.25 GWd/MTU and 300-year decay time case when 44 rods 

have been removed from the lattice configuration. The maximum ∆keff value (i.e., keff increase of 

2.40% ∆keff) for the GBC-68 cask occurs for the 5 wt % 235U initial enriched fuel with a burnup of 

35 GWd/MTU and 5-year decay time case when 18 rods have been removed from the lattice 

configuration with the channel present. The maximum ∆keff value changes for this configuration 

scenario are provided in Table 6.

Figure 13 Change in keff in GBC-32 as a function of number of rods removed

-0.5

0.0

0.5

1.0

1.5

2.0

0 5 10 15 20 25 30 35 40 45 50

Increase in 

keff (%

Δ

keff)

Number of rods removed

1.92 wt% U-235, Fresh 3.5 wt% U-235, 25.5 GWd/MTU

5 wt% U-235, 44.25 GWd/MTU 5 wt% U-235, 70 GWd/MTU
21

Figure 14 Change in keff in GBC-68 as a function of number of rods removed

Table 6 Criticality results for scenario S1(a)—breached spent fuel rods 

Parameter

PWR system

Burnup (GWd/MTU) 0 25.5 44.25 70

Initial enrichments (wt % 235U) Maximum change in keff (%∆keff)

1.92 0.04 NC NC NC

3.5 with 5-year decay time NC 1.07 NC NC

5.0 with 5-year decay time NC NC 1.86 1.69

5.0 with 80-year decay time NC NC 1.86 1.62

5.0 with 300-year decay time NC NC 1.87 1.62

BWR system

Burnup (GWd/MTU) 0 35 70

Initial enrichment (wt % 235U) Maximum change in keff (%∆keff) 

channeled/unchanneled

5.0 with 5-year decay time 2.24/2.11 2.40/2.30 2.30/2.20

5.0 with 80-year decay time NC 2.40/2.31 2.31/2.18

5.0 with 300-year decay time NC 2.42/2.29 2.32/2.20

Note: NC = not calculated.

An additional subset of cases was developed to investigate the combined effects of rod removal 

and spent fuel displacement to regions outside the neutron absorber envelope. Only the PWR 

fuel was represented with 5.0 wt% 235U initial enrichment at 44.25 GWd/MTU burnup and 5-year 

decay time. For these cases, the region above the active fuel zone and outside the neutron 

absorber plate envelope was modeled with fuel at different volume fractions simulating various 

sized particles with different bulk densities. The volume displacement of hardware components 

in this region was omitted from these models to maximize the reactivity effect. Note that this is a 

significant conservative assumption. Results provided in Table 7 show that combining the rod 

removal with displaced fuel distributed at ratios approaching optimum moderation conditions 

outside the absorber envelope yields higher changes in keff than the rod removal configurations 

alone. However, the magnitude of the effect is highly sensitive to the available void volume 

0.0

0.5

1.0

1.5

2.0

2.5

0 5 10 15 20

Increase in 

keff (%

Δ

keff)

Number of failed rods

5 wt% U-235, Fresh 5 wt% U-235, 35 GWd/MTU 5 wt% U-235, 70 GWd/MTU
22

outside the neutron absorber panel envelope, which is canister basket, fuel spacer, and 

assembly design dependent. 

Table 7 Combined rod removal and displaced fuel model results in GBC-32 

Number of 

failed rods

Volume 

fraction of 

displaced fuel

Length of region 

above absorber filled 

(cm)

Change in keff

(%∆keff)

Change in keff

(%∆keff) (Rod 

removal only)

40 0.524 30 3.89 1.82

28 0.341 30 4.91 1.52

8 0.146 20 3.55 0.57

Damaged fuel. To evaluate the consequences of extensive cladding failure, cases were 

developed to represent scenario S1(b), damaged spent fuel rods. Two subclasses of cases 

were used for this evaluation – one where the fuel is represented as a homogenous rubble 

mixture distributed throughout the basket cavity, and the other where the fuel is represented as 

a uniform pellet array to simulate a heterogeneous mixture. The fuel mixtures were allowed to 

be distributed throughout the entire basket cell cavity including outside the neutron absorber 

envelope. Fuel spacers were not included in the representations, and assembly hardware (i.e., 

cladding and guide tubes) were only included in the homogeneous rubble mixture 

representations. The results for this fuel reconfiguration scenario are provided in Table 8 and 

show that the modeling approximations based on homogenous representations are less 

conservative than the heterogeneous models. While models of 100% fuel rubblization are 

extreme and not physically likely under normal, off-normal, and accident conditions of storage, 

and under NCT as prescribed in 10 CFR 71.71, or as a result of impacts from accident loads 

specified in 10 CFR 71.73, they provide a simplified approach for estimating an upper bound on 

the potential reactivity increase. The consequences for this configuration scenario are very 

sensitive to the modeling approximations and resultant water to fuel ratio represented above or 

below the neutron absorber panel envelope. 

Table 8 Criticality results for scenario S1(b)—damaged spent fuel rods

Parameter

PWR system

Burnup (GWd/MTU) 0 25.5 44.25 70

Initial enrichments (wt % 235U) Maximum change in keff (%∆keff) limiting pellet 

array/homogeneous mixture

1.92 11.09/6.66 NC NC NC

3.5 with 5-year decay time NC 20.20/13.95 NC NC

5.0 with 5-year decay time NC NC 21.37/14.30 21.43/14.20

5.0 with 80-year decay time NC NC 22.21/15.29 21.63/14.77

5.0 with 300-year decay time NC NC 22.21/15.34 21.77/14.90

BWR system

Burnup (GWd/MTU) 0 35 70

Initial enrichment (wt % 235U) Maximum change in keff (%∆keff) limiting pellet array 

(homogeneous mixture channeled/unchanneled)

5.0 with 5-year decay time 28.12 

(21.68/22.90)

34.40 

(28.58/29.36)

35.22

(29.31/29.93)

5.0 with 80-year decay time NC 34.88

(29.12/29.87)

35.57

(29.74/30.33)

5.0 with 300-year decay time NC 34.87

(29.13/29.83)

35.63

(29.81/30.40)

Note: NC = not calculated.
23

Overall, the most limiting configurations to evaluate the consequences of cladding failure on 

criticality safety are represented by models where displaced fuel is relocated to regions above 

or below the neutron absorber panel envelope. Configurations that only consider rod removal 

may be under-predicting the impact of cladding failure by not accounting for the displaced fuel. 

Configurations that represent the fuel as damaged and distributed throughout the basket cavity 

are bounding but may not be physically likely under normal, off-normal, or accident conditions of 

storage, or normal and accident conditions of transportation. 

3.1.2 Shielding

Understanding the consequences of cladding failure on external radiation dose rates requires

shielding analyses to investigate the changes associated with both breached and damaged fuel 

geometries. The primary transportation requirements that shielding analyses address are 

identified in 10 CFR 71.47 for NCT and 10 CFR 71.51 for HAC. The requirements describe 

package external dose rate limits for specific distances from the package or transportation 

vehicle. For storage of SNF, dose limits at the controlled area boundary are specified in 10 CFR

72.104 for normal and off-normal conditions and 10 CFR 72.106 for accident conditions.

Regulatory guidance [2] indicates that a shielding analysis of a single cask and a generic array 

of casks at large distances may be used to demonstrate that the radiation shielding features of a 

proposed dry storage system are sufficient for it to meet the radiation dose requirements in 10 

CFR 72.104 and 10 CFR 72.106(b). The minimum distance required between spent fuel and the 

controlled area boundary specified in 10 CFR 72.106 is 100 m. For storage, the effects of fuel 

configuration changes on dose rate were determined at 1 m from a generic storage cask and at 

100 m from a 4×2 cask array (see Appendix B.3.2).

The consequences of fuel configuration changes were evaluated by comparing external dose 

rate values between the nominal intact and fuel reconfigurations. This approach helps identify 

the fuel reconfigurations that yield higher external dose rates than the nominal intact fuel 

configuration and thus have the potential to exceed the regulatory limits. Package/storage cask

external gamma and neutron dose rate values were evaluated separately because fuel 

configuration changes have different effects on gamma and neutron radiation, and the overall 

impact of fuel configuration changes on the external dose rate depends on the individual 

gamma and neutron dose rate contributions to the total dose rate. Generic PWR and BWR SNF 

transportation package/storage cask models based on GBC-32 and GBC-68, respectively, were 

developed for the shielding analysis (see Appendix B). The transportation package models have 

general shielding characteristics (e.g., thick inner gamma shield and outer neutron shield) 

similar to real high-capacity SNF transportation packages. The storage cask model has a 

vertical concrete overpack. 

The W 17×17 OFA and GE14 assemblies with 5 wt % 235U initial enrichment, 65 GWd/MTU 

burnup, and typical axial burnup profiles were used in this analysis (see Appendix B). The effect 

of fuel configuration changes on external radiation dose rates was determined as a function of 

decay time (i.e., time after fuel discharge from the reactor). The fuel decay times of 5 and 40 

years were analyzed because the contribution to cask external dose rates of 60Co (half-life 

[t1/2] = 5.271 years), which is an activation source in the assembly plenum and lower and upper 

hardware regions, is significant for a 5-year decay time and negligible for a 40-year decay time.

Neutron and gamma spectra variations are relatively small for decay times relevant to this 

analysis (e.g., <120 years) beyond 5 and 40 years after fuel discharge, respectively.
24

For the 5-year decay time, the gamma dose rate at the package bottom and top surfaces is high 

primarily because of the proximity of the 60Co activation source to those surfaces, the activation 

source strength, and energetic gamma rays of 1.25-MeV average energy. The 60Co activation 

source has negligible contributions to gamma dose rates for the 40-year decay time and 

beyond. Therefore, the contribution of the 60Co activation source to the gamma dose rates at the 

package bottom and top surfaces varies as a function of decay time. A large contribution of the 

60Co activation source to the gamma dose rate, as in the case of a 5-year decay time, equates 

to a relatively small increase of the gamma dose rate because of fuel relocation to the assembly 

hardware regions. A small or negligible contribution of the 60Co activation source to the gamma 

dose rate, as in the case of a 40-year decay time, equates to a relatively large increase of the 

gamma dose rate because of fuel relocation to assembly hardware regions. That is, the effects 

on the gamma dose rates at the package bottom and top surfaces from fuel relocation into 

assembly hardware regions are significantly smaller for the 5-year decay time than those for the 

40-year decay time.

Breached spent fuel rods. The effects of breached spent fuel rods (i.e., source term 

redistribution) on package external dose rates was analyzed by considering different 

percentages of failed fuel rods per assembly (i.e., 10% and 25% for the PWR fuel and 11% for 

the BWR fuel) and fuel redistribution to different assembly axial regions, including the active fuel 

region, and the lower and upper hardware regions. Important parameters used in the calculation 

models: 

(1) Relocation of fuel fragments and particulates from a failed assembly is confined within 

the space delimited by the fuel assembly basket plates and by the assembly top and 

bottom spacers.

(2) One hundred percent of the fuel in the failed fuel rods is displaced to a different 

assembly region.

(3) Fuel fragments and particulates from failed fuel rods form a closely packed mixture with 

a mass packing fraction of 0.58 based on powder mechanics for particles similar to sand

[27].

Because the packing fraction, or particle density, is dependent upon the particle size distribution 

a second higher mass packing fraction value of 0.67 was also used for sensitivity calculations. 

The specific fuel rod failure percentage values were selected arbitrarily because actual data on 

the expected fuel rod failure rate is not available. However, the whole range of fuel failure with 

respect to the amount of fuel mixture (0 – 100%) that may be displaced to canister cavity 

regions is analyzed in this report under the different fuel reconfiguration categories. The 25%

fuel rod failure value for the PWR SNF was analyzed to determine whether there is a direct 

proportionality between the amount of redistributed fuel mixture and the external dose rate (i.e., 

whether an increase in the displaced fuel mixture quantity from 10% to 25% equates to an 

increase in external dose rate by a factor of 2.5).

Damaged fuel. To evaluate the consequences of extensive cladding failure representative of 

scenario S1(b)—damaged spent fuel rods, two cases were analyzed: (1) fuel mixture assumed 

to form closely packed powder in the canister cavity bottom; and (2) fuel mixture 

homogeneously distributed within the canister cavity volume. These configurations are bounding 

for cladding failure scenarios resulting in fuel fragments and particulates being collected into the 

inner cavity regions below or above the fuel spacers or between the canister basket outer plates 

and radial canister wall. The latter modeling approach is often used to generate conservative 

dose rate estimates because the model is characterized by reduced gamma self-shielding (i.e., 

reduced gamma radiation absorption within the fuel mixture with lower particle density) as well 
25

as reduced gamma and neutron geometric attenuation (i.e., radiation source is closer to inner 

cavity walls) compared to the nominal intact fuel configuration. From all cases analyzed in this 

study, the damaged fuel configurations produced the greatest changes in the maximum dose 

rate on the cask external surfaces relative to the nominal intact fuel configuration.

The ratio of the maximum surface dose rate value for a fuel reconfiguration to the maximum 

surface dose rate for the nominal intact fuel configuration, which is identified as the failed to 

intact (F/I) configuration dose rate ratio, is summarized in Table 9 (cases 1 to 8 for the breached 

spent fuel rod cases, and cases 9 and 10 for the damaged fuel cases) for the evaluated PWR 

fuel reconfigurations and in Table 10 (cases 1 to 4 for the breached spent fuel rod cases, and 

cases 5 and 6 for the damaged fuel cases) for the evaluated BWR fuel reconfigurations, under 

NCT. For the analysis considering NCT, the F/I dose rate values are provided at the package 

external surfaces and at the 2 m locations. For the analysis considering HAC, the outer neutron 

shield considered in the package models for NCT was replaced with air and the dose rates were 

calculated at 1 m from the package external surfaces. The F/I dose rate ratio values for HAC

are provided in Table 11 and Table 12 for PWR and BWR package models, respectively.

For the breached spent fuel rod cases analyzed:

• the change in maximum gamma and neutron dose rates on the external surfaces of the 

casks was insignificant ( ~10% for both PWR and BWR fuel) for fuel collected into the 

middle portion of the active fuel (i.e., away from the upper and lower ends of the active 

fuel);

• a significant increase of the maximum neutron dose rate (e.g., ~6.2, 3.5, and 2.8 times 

as large as the reference maximum top, radial, and bottom dose rate values, 

respectively, for the analyzed PWR cases) was determined for fuel mixture collected into 

either assembly bottom or top regions; 

• the impact on dose rates from fuel redistribution to the assembly top and bottom regions 

was more pronounced for BWR fuel than for PWR fuel (e.g., ~22 versus 5 times as large

as the nominal intact configuration maximum top neutron dose rate values, respectively);

• for shorter decay times (e.g., 5-years), fuel configuration changes had a greater impact 

on the neutron dose rate than on the gamma dose rate at the external surfaces of the 

packages (e.g., maximum dose rate on the radial surface from all analyzed cases 

increased by ~30% for the gamma radiation and by a factor of 3.5 for the neutron 

radiation); 

• for longer decay times (e.g., 40-years), fuel configuration changes had a greater impact 

on the gamma dose rate than on the neutron dose rate at the external surfaces of the 

packages (e.g., maximum dose rate on the top surface from all analyzed cases 

increased by a factor of ~85 for the gamma radiation and by a factor of ~24 for the 

neutron radiation); and

• a larger percentage (i.e., 25% versus 10% analyzed for PWR fuel) of fuel mixture 

distributed to the assembly axial regions did not increase the dose rates at the package 

axial surfaces proportionally to the source strength increase (e.g., the F/I ratio values for 

the neutron dose rate at the top external surface were ~6.2 and 5 for the 25% and 10% 
26

quantities, respectively) because of an increased geometric attenuation associated with 

the larger percentage value. 

 

For the damaged fuel cases analyzed:

• the maximum top, radial, and bottom neutron dose rate values were ~6.7, 3.9, and 4.2 

times as large as the corresponding nominal intact configuration dose rate values for the 

PWR package model; the maximum top, radial, and bottom neutron dose rate values 

were ~23.5, 3.3, and 6 times as large as the corresponding nominal intact configuration

dose rate values for the BWR package model;

• for the 40-year decay time, the maximum top and bottom PWR gamma dose rates were 

~14 and 7 times, respectively, as large as the corresponding PWR nominal intact 

configuration dose rate values; the maximum top and bottom BWR gamma dose rates 

were ~84 and 27 times, respectively, as large as the corresponding BWR nominal intact 

configuration dose rate values; and

• the gamma dose rate increase for shorter decay times (e.g.5-years) is significantly 

smaller than that for longer decay times (e.g., 40-years).

Table 9 PWR package maximum dose rate change for NCT

Package external surfaces Top Side Bottom

Case 

# Scenario Fuel relocation 

regionb

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

1 10% fuel rod 

failure Assembly active fuel

5 1.09 1.01 1.07 1.01 1.02 1.01 1.04 1.00 1.02

40 1.09 1.00 1.09 1.01 1.01 1.00 1.03 1.00 1.03

2 10% fuel rod 

failure

Assembly lower end 

fitting

5 1.06 1.00 1.05 1.62 0.99 0.99 2.52 1.34 2.10

40 1.06 1.00 1.06 1.59 0.98 1.32 2.53 5.17 2.67

3 10% fuel rod 

failure Assembly plenum 

5 3.96 0.98 3.43 2.37 0.99 1.48 1.03 1.01 1.02

40 3.96 3.35 3.96 2.38 0.99 1.96 1.02 0.99 1.02

4 10% fuel rod 

failure

Assembly upper end 

fitting

5 4.98 0.93 4.27 2.75 0.99 1.66 1.05 1.01 1.01

40 4.99 11.55 5.00 2.77 0.99 2.34 1.02 1.00 1.02

5 25% fuel rod 

failure Assembly active fuel

5 1.12 1.00 1.10 1.06 1.12 1.09 1.00 1.00 1.00

40 1.13 1.00 1.13 1.06 1.10 1.07 1.00 0.96 0.99

6 25% fuel rod 

failure

Assembly lower end 

fitting

5 1.05 1.00 1.04 1.76 0.98 0.96 2.76 1.33 2.25

40 1.05 1.00 1.05 1.76 0.97 1.45 2.76 5.16 2.88

7 25% fuel rod 

failure Assembly plenum 

5 4.59 0.97 3.95 2.79 0.95 1.69 1.00 0.99 0.99

40 4.58 3.34 4.57 2.80 0.96 2.30 0.99 0.96 0.99

8 25% fuel rod 

failure

Assembly upper end 

fitting

5 6.16 0.93 5.23 3.50 0.95 2.03 1.00 1.00 1.00

40 6.13 11.59 6.14 3.51 0.96 2.94 0.99 0.95 0.99

9 Damaged Canister cavity 

bottom

5 0.32 0.001 0.26 2.57 1.19 1.47 4.18 1.72 3.30

40 0.32 0.02 0.32 2.59 1.21 2.18 4.13 8.37 4.34

10 Damaged Entire canister cavity

5 6.69 0.84 5.56 3.89 1.10 2.48 3.04 1.43 2.47

40 6.67 14.10 6.68 3.89 1.24 3.40 3.05 7.30 3.21
27

2 m from the package surfaces Top Side Bottom

Case 

# Scenario Fuel relocation 

regionb

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

1 10% fuel rod 

failure Assembly active fuel

5 1.07 1.01 1.05 1.02 1.00 1.01 1.02 1.00 1.01

40 1.08 1.00 1.08 1.02 1.00 1.01 1.02 0.99 1.02

2 10% fuel rod 

failure

Assembly lower end 

fitting

5 1.04 1.01 1.03 1.02 0.99 1.00 2.29 1.33 1.92

40 1.06 1.02 1.06 1.06 1.01 1.05 2.23 4.82 2.40

3 10% fuel rod 

failure Assembly plenum 

5 3.48 0.99 2.93 1.02 0.98 1.00 1.04 1.01 1.03

40 3.36 1.26 3.36 1.07 1.01 1.05 1.00 0.98 1.00

4 10% fuel rod 

failure

Assembly upper end 

fitting

5 4.27 0.94 3.54 1.01 0.99 1.00 1.04 1.01 1.00

40 4.19 2.85 4.21 1.08 1.01 1.07 1.00 0.98 1.00

5 25% fuel rod 

failure Assembly active fuel

5 1.11 1.00 1.09 1.04 1.04 1.04 0.99 0.99 0.99

40 1.11 0.99 1.11 1.02 1.00 1.02 0.98 0.96 0.98

6 25% fuel rod 

failure

Assembly lower end 

fitting

5 1.04 1.00 1.03 0.99 0.96 0.98 2.51 1.34 2.06

40 1.05 0.96 1.05 1.03 1.00 1.02 2.39 4.76 2.54

7 25% fuel rod 

failure Assembly plenum 

5 4.08 0.98 3.40 1.06 0.95 0.97 0.99 0.99 1.00

40 4.00 1.28 3.99 1.09 1.01 1.07 0.96 0.96 0.96

8 25% fuel rod 

failure

Assembly upper end 

fitting

5 5.27 0.95 4.32 1.25 0.95 0.96 0.99 0.99 0.99

40 5.13 2.88 5.14 1.15 1.01 1.10 0.97 0.95 0.97

9 Damaged Canister cavity 

bottom

5 0.30 0.03 0.23 1.02 1.07 1.02 4.01 2.17 3.30

40 0.29 0.26 0.29 0.94 0.94 0.92 3.95 10.56 4.38

10 Damaged Entire canister cavity

5 5.94 1.04 4.86 1.28 1.11 1.08 2.90 1.78 2.47

40 5.79 5.02 5.83 1.24 1.32 1.19 2.80 9.11 3.21

aRelative error (95% confidence level) less than 7% for all values except for the gamma F/I value for the case #9, top 

surface, the relative error is 15%. bFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions; axial 

location is described in Sect. B.5.1

Table 9 PWR package maximum dose rate change for NCT (continued)
28

Table 10 BWR package maximum dose rate change for NCT

Package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

regionb

Decay 

time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

1 11% fuel rod 

failure

Assembly active 

fuel

5 1.13 0.98 1.11 1.02 1.02 1.02 1.06 0.99 1.02

40 1.14 1.01 1.14 1.03 1.01 1.02 1.07 0.97 1.07

2 11% fuel rod 

failure

Assembly lower tie 

plate and 

nosepiece

5 1.14 1.02 1.12 1.43 0.97 0.99 4.37 1.21 2.88

40 1.14 1.00 1.13 1.42 0.97 1.22 4.44 24.39 4.58

3 11% fuel rod 

failure Assembly plenum 

5 12.74 1.40 11.05 1.78 0.96 1.11 1.05 0.99 1.02

40 12.98 13.67 12.98 1.75 0.98 1.48 1.08 0.97 1.08

4 11% fuel rod 

failure

Assembly upper tie 

plate and handle

5 21.78 3.22 19.02 2.42 0.96 1.65 1.06 1.01 1.04

40 21.77 84.60 21.83 2.41 0.99 2.12 1.07 0.98 1.07

5 Damaged Canister cavity 

bottom

5 0.60 0.003 0.51 1.84 1.35 1.63 5.99 1.17 3.72

40 0.61 0.15 0.61 1.81 1.40 2.12 6.04 32.19 6.22

6 Damaged Entire canister 

cavity

5 23.48 2.90 20.41 3.27 0.98 2.25 4.31 0.94 2.72

40 23.52 84.18 23.58 3.26 1.11 2.90 4.37 26.71 4.53

2 m from the package surfaces Top Side Bottom

Case # Scenario Fuel relocation 

regionb

Decay 

time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

1 11% fuel rod 

failure

Assembly active 

fuel

5 1.11 0.99 1.08 1.04 0.98 1.02 1.04 0.99 1.01

40 1.11 0.98 1.11 0.98 1.00 0.97 1.03 0.99 1.03

2 11% fuel rod 

failure

Assembly lower tie 

plate and 

nosepiece

5 1.12 1.00 1.09 1.04 0.96 1.01 3.78 1.19 2.43

40 1.09 1.01 1.09 1.05 1.00 1.04 3.79 19.79 3.97

3 11% fuel rod 

failure Assembly plenum 

5 10.62 1.44 8.70 1.03 0.96 1.11 1.04 0.99 1.01

40 10.15 1.37 10.15 1.07 0.99 1.06 1.05 1.00 1.05

4 11% fuel rod 

failure

Assembly upper tie 

plate and handle

5 17.12 3.21 14.21 1.02 0.97 1.00 1.02 1.00 1.01

40 16.24 4.94 16.33 1.04 0.99 1.03 1.05 1.01 1.05

5 Damaged Canister cavity 

bottom

5 0.53 0.12 0.42 2.19 1.29 1.76 5.62 1.34 3.38

40 0.52 0.46 0.52 2.01 1.24 1.90 5.57 32.09 5.86

6 Damaged Entire canister 

cavity

5 25.23 3.28 7.77 1.23 0.98 0.98 3.93 1.07 2.44

40 18.88 6.07 18.99 1.11 1.18 1.08 3.89 26.50 4.14

a Relative error (at the 95% confidence level) less than 7% for all values except for the gamma F/I value for case #5, 

top surface, the relative error is 20%. b

Farthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions; axial

location is described in Sect. B.5.3.
29

Table 11 PWR package maximum dose rate change for HAC

1 m from the package surfaces Top Side Bottom

Case # Scenario Fuel relocation 

regionb

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

1 10% fuel rod 

failure Assembly active fuel

5 1.07 1.00 1.06 1.01 1.00 1.01 1.02 1.00 1.01

40 1.07 0.98 1.08 1.02 0.97 1.02 1.02 0.96 1.02

2 10% fuel rod 

failure

Assembly lower end 

fitting

5 1.04 0.99 1.02 1.02 1.00 1.02 2.14 1.33 1.96

40 1.05 1.00 1.05 1.02 0.97 1.02 2.19 4.89 2.27

3 10% fuel rod 

failure Assembly plenum 

5 3.35 0.98 2.79 1.02 1.00 1.02 1.03 1.01 1.02

40 3.36 1.28 3.44 1.00 0.97 1.00 1.02 0.95 1.02

4 10% fuel rod 

failure

Assembly upper end 

fitting

5 4.12 0.93 3.37 0.99 0.97 1.01 1.13 1.01 1.10

40 4.17 3.31 4.19 1.01 0.96 1.01 1.09 0.95 1.01

5 25% fuel rod 

failure Assembly active fuel

5 1.10 0.99 1.07 1.01 1.02 1.01 0.99 0.98 0.99

40 1.10 1.04 1.11 1.01 0.99 1.01 1.16 0.94 1.16

6 25% fuel rod 

failure

Assembly lower end 

fitting

5 1.03 1.00 1.02 1.00 0.96 0.99 2.42 1.35 2.18

40 1.03 1.01 1.03 0.99 0.94 0.99 2.37 4.88 2.37

7 25% fuel rod 

failure Assembly plenum 

5 3.92 0.98 3.23 0.99 0.97 0.99 0.99 0.99 0.99

40 3.99 2.02 3.99 0.98 1.08 0.98 0.98 1.05 0.98

8 25% fuel rod 

failure

Assembly upper end 

fitting

5 5.05 0.93 4.08 0.99 0.96 0.99 1.00 0.99 1.00

40 5.12 4.74 5.21 0.98 1.09 0.98 0.98 1.05 0.98

9 Damaged Canister cavity 

bottom

5 0.29 0.02 0.22 0.78 0.99 0.81 3.79 2.23 3.44

40 0.29 0.26 0.21 0.80 1.05 0.81 3.82 10.19 4.01

10 Damaged Entire canister cavity

5 5.80 1.06 4.69 0.96 1.14 0.99 2.73 1.83 2.53

40 5.90 5.43 5.93 0.93 1.17 0.93 2.69 8.66 2.89

aRelative error (at the 95% confidence level) less than 5% for all values except for the gamma F/I value for case #9, 

top surface, the relative error is 20%. bFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions, as 

described in Sect. B.5.1.
30

Table 12 BWR package maximum dose rate change for HAC

1 m from the package surfaces Top Side Bottom

Case # Scenario Fuel relocation 

regionb

Decay 

time (year)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

1 11% fuel rod 

failure Assembly active fuel

5 1.04 1.01 1.08 1.04 0.98 1.03 1.08 0.99 1.05

40 0.99 0.99 0.99 1.02 0.98 1.02 1.06 0.96 1.06

2 11% fuel rod 

failure

Assembly lower end 

fitting

5 1.04 1.00 1.10 1.01 0.96 1.00 3.57 1.21 2.79

40 1.01 1.01 1.00 1.00 0.97 1.01 3.65 16.99 3.72

3 11% fuel rod 

failure Assembly plenum 

5 6.86 1.42 6.65 1.04 0.97 1.03 1.04 1.00 1.03

40 6.72 1.58 6.65 1.01 0.98 1.01 1.07 0.98 1.07

4 10% fuel rod 

failure

Assembly upper end 

fitting

5 11.36 3.18 11.15 1.00 0.97 1.00 1.06 1.00 1.04

40 11.06 6.49 11.00 1.00 0.97 1.00 1.05 0.94 1.05

5 Damaged Canister cavity 

bottom

5 0.43 0.06 0.45 0.77 1.02 0.80 5.26 1.37 3.98

40 0.44 0.33 0.43 0.77 1.03 0.78 5.33 26.24 5.45

6 Damaged Entire canister cavity

5 12.84 3.28 12.56 0.83 0.99 0.85 3.69 1.09 2.83

40 12.74 7.39 12.90 0.83 1.05 0.84 3.68 21.69 3.78

aRelative error (at the 95% confidence level) less than 5% for all values except for the gamma F/I value for case #5, 

top surface, the relative error is 20%.

bFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions, as 

described in Sect. B.5.3.

Graphs illustrating comparisons of the neutron and gamma dose rate profiles between the 

nominal intact fuel configuration and the fuel reconfigurations analyzed are provided in 

Appendix B for the 40-year decay time. The effects on the total dose rate profiles at the PWR 

package external surfaces and at the 2 m locations are illustrated in Figure 15 and Figure 16, 

respectively, for the fuel reconfiguration with 25% fuel mixture collected into the assembly 

bottom region. The relative errors (at the 95% confidence level) are 3% and 5% for the total 

dose rates at the package external surfaces and at the 2 m locations, respectively. Note the 

total dose rate profiles illustrated in the figures are specific to the package model used in this 

analysis, which is described in Sect. B.3.1, and are not applicable to different package designs.
31

(a)

(b)

(c)

Figure 15 Total dose rate profiles along the PWR transport package external top surface 

(a), radial surface (b), and bottom surface (c) for the intact fuel configuration 

and the fuel reconfiguration with 25% fuel mixture collected into the assembly 

bottom region, NCT
32

(a)

(b)

(c)

Figure 16 Total dose rate profiles at 2 m from the PWR transport package external top

surface (a), radial surface (b), and bottom surface (c) for the intact fuel 

configuration and the fuel reconfiguration with 25% fuel mixture collected into 

the assembly top region, NCT
33

The dose rates were calculated at 1 m from either a PWR or BWR storage cask (see Sect.

B.3.2 for a description of the models) for fuel configuration changes in casks with a vertical 

orientation. The changes (F/I dose rate ratio) relative to the nominal intact configuration of the 

maximum neutron, gamma, and total dose rates at 1 m from the storage cask are summarized 

in Table 13. Fuel configuration changes cause significant dose rate changes relative to the 

nominal intact configuration in the cask outer regions that face air vent locations, (i.e., receive 

radiation directly from streaming through the air vents). At locations away from air vents, the 

change in radiation dose rate is either small (e.g., ~30% for damaged fuel configurations) or 

negligible. 

Table 13 Maximum dose rate change at 1 m from a storage cask

1 m from the PWR storage cask surfaces Topa Sidea Sideb

Case # Scenario Fuel relocation regionc Decay time 

(years)

F/I dose rate ratio

n γ Total n γ Total n γ Total

2 10% fuel rod 

failure

Assembly lower end 

fitting

5 1.09 0.70 0.70 1.68 1.87 1.87 1.05 1.00 1.00

40 1.09 0.90 0.90 1.68 2.91 2.86 1.05 1.04 1.03

9 Damaged Canister cavity bottom

5 0.22 0.07 0.08 2.73 4.17 4.05 0.97 1.20 1.20

40 0.22 0.18 0.18 2.73 4.09 4.02 0.97 0.77 0.85

1 m from the BWR storage cask surfaces Topa Sidea Sideb

Case # Scenario Fuel relocation regionc Decay time 

(years)

F/I dose rate ratio

n γ Total n γ Total n γ Total

2 11% fuel rod 

failure

Assembly lower tie plate 

and nosepiece

5 1.02 0.56 0.56 2.16 1.78 1.79 1.05 0.97 0.97

40 1.02 0.82 0.79 2.16 2.73 2.70 1.05 1.00 1.01

5 Damaged Canister cavity bottom

5 0.22 0.15 0.15 3.23 5.58 5.51 0.92 1.09 1.08

40 0.22 0.40 0.39 3.23 9.58 9.23 0.92 1.31 1.24

aFacing air vent locations; relative error (at the 95% confidence level) is 10% for the radial surface and 20% for the 

top surface. bAway from air vent locations; relative error (at the 95% confidence level) is 6%. cFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions, as 

described in Sect. B.5.1.

Neutron and gamma dose rates at the controlled area boundary from a generic 4×2 storage 

cask array were calculated at 100 m from the array and are shown in Table 14. The results 

show that, relative to the nominal intact fuel configuration, the closely packed fuel mixture model 

reduces site boundary dose rates by ~70% for gamma radiation and ~30% for neutron radiation, 

whereas the models with the fuel mixture homogeneously distributed within the entire canister

cavity increase the site boundary dose rates by a factor of ~2.4 for gamma radiation and by a 

factor of ~2.7 for neutron radiation. The dose rate changes are caused by source geometry 

changes and gamma self-shielding effects associated with the different fuel configurations.
34

Table 14 Site boundary dose rate change 

Case description

F/I dose rate ratioa

n γ Total

PWR fuel mixture collected into canister cavity bottom 0.58 ± 0.07 0.22 ± 0.03 0.21 ± 0.03

BWR fuel mixture collected into canister cavity bottom 0.65 ± 0.01 0.30 ± 0.01 0.30 ± 0.01

PWR fuel mixture homogeneously distributed within canister 

cavity

2.20 ± 0.08 1.75 ± 0.12 1.76 ± 0.13

BWR fuel mixture homogeneously distributed within canister 

cavity

2.74 ± 0.04 2.41 ± 0.05 2.38 ± 0.03

aF/I dose rate ratio ± 2 sigma statistical error.

3.1.3 Containment

Understanding the consequences of cladding failure on containment performance requires 

analyses to investigate the changes associated with varying the fraction of rods that develop 

cladding breach under NCT, from 0.01 to 1 (all rods breached), and specifically for high-burnup 

fuel analyses, varying the release fractions for the contributors to the releasable source terms in 

addition to the fraction of breached spent fuel rods. This technical approach was used because 

the range of parameters for the containment analysis of high-burnup fuel (i.e., the release 

fractions and breached spent fuel rod fraction) has not been established yet for high-burnup 

fuel. The release fractions for gases, volatiles, and fuel fines provided in NUREG-1617, Table 4-

1 [28], were used in the containment analysis of low-burnup fuel (i.e., 40 GWd/MTU). However, 

the applicability of the release fractions described in NUREG-1617 to long-term storage should 

also be confirmed in future evaluations. For the high-burnup fuel analysis, the fraction of fuel 

rods that develop breaches was varied from 0.01 to 1, the fraction of crud that spalls off the 

cladding varied from 0.15 to 1, the fraction of gases varied from 0.1 to 0.4, the fraction of 

volatiles varied from 2 × 10-4 to 2 × 10-3

, and the mass fraction of fuel fines was varied from 3 ×

10-5 to 3 × 10-4

. Total releasable activity, effective A2 value, allowable radionuclide release rate, 

and allowable leakage rate were calculated for the generic PWR and BWR transportation 

packages referred to as GBC-32 and GBC-68, respectively (see Appendix B for model 

specifications). Representative fuel assemblies selected for analysis are the PWR W 17×17 

OFA and BWR GE14 10×10 lattice with 40- and 65 GWd/MTU burnup values. The impact on 

the releasable activity of the discharged radioactive material original location (i.e., non-rim 

region of the fuel pellet or the peripheral rim structure) was evaluated for the 65 GWd/MTU 

burnup value. The results of the containment analysis are provided in Appendixes C.4 and C.5

for the fuel assemblies with low-burnup (i.e., 40 GWd/MTU) and high-burnup (i.e., 65 

GWd/MTU), respectively.

Figure 17 and Figure 18 illustrate the allowable leakage rate variation as a function of the 

fraction of spent fuel rods with cladding breaches for the GBC-32 and GBC-68 containing fuel 

assemblies, respectively, with 40 GWd/MTU burnup. With a more restrictive allowable leakage 

rate, GBC-68 can accommodate greater fuel rod failure rates than GBC-32. For GBC-68 with 

low-burnup fuel, an allowable leakage rate of ~2.4 × 10-5 cm3

/s was obtained for the 0.03, 0.4, 

0.7, and 1.0 breached spent fuel rod fractions at the 5-, 40-, 100-, and 300-year decay times, 

respectively. For GBC-32 with low-burnup fuel, an allowable leakage rate of ~1 × 10-4 cm3

/s was 

obtained for the 0.03, 0.10, 0.15, and 0.25 breached spent fuel rod fractions at the 5-, 40-, 100-, 

and 300-year decay times, respectively.
35

Figure 17 Allowable leakage rate variation as a function of spent fuel rod cladding 

breaches for the GBC-32 cask and 40 GWd/MTU burnup at different decay 

times

Figure 18 Allowable leakage rate variation as a function of spent fuel rod cladding 

breaches for the GBC-68 cask and 40 GWd/MTU burnup at different decay 

times

The containment analysis for high-burnup fuel under NCT evaluated the impact on the allowable 

leakage rate of varying the fraction of spent fuel rods that develop breaches (0.01 to 1), fraction 

of crud (0.15 to 1) that spalls off cladding, fraction of gases (0.1 to 0.4), volatiles (2 × 10-4 to 2 ×

10-3

), fuel fines (3 × 10-5 to 3 × 10-4) released due to cladding breach, and the fuel pellet radial 

region from which the releasable activity originates. For each fuel type, decay time, and fraction 

of fuel rods that develop breaches, 15 different cases were analyzed, each case using a 
36

combination of the values previously described for the parameters important to containment 

analysis. The calculated allowable leakage rates are provided in the tables and graphs in 

Appendix C.5.4 with a summary of the results provided as follows.

Decay time. Decay time has a significant impact on the releasable activity and allowable 

leakage rate primarily because of the decay of 60Co (t1/2 = 5.271 years), which is the source of 

radioactivity in crud, and 85Kr (t1/2 = 10.76 years), which dominates the gaseous species. The 

allowable leakage rate increases with increasing decay time.

Fraction of fuel rods that develop cladding breaches. The allowable leakage rate decreases 

with increasing fraction of breached spent fuel rods. The decrease is relatively small for short 

decay times when crud has a large contribution to the total releasable activity (e.g., a 15% 

decrease because of an increase in spent fuel rod breach fraction from 0.01 to 0.03 for the 

BWR fuel with 5-year decay time) and is fairly large when the fuel fines dominate the total 

releasable activity (e.g. a decrease by a factor of 2.5 because of an increased failed fuel rod 

fraction from 0.01 to 0.03 for the BWR fuel with 40-year decay time). 

Fraction of crud that spalls off cladding. Crud is an important factor in the calculation of the 

allowable leakage rate for the time interval 5 to 40 years after fuel discharge from the reactor 

because of its relatively high contribution to the total releasable activity. Crud has a more 

pronounced effect on the BWR packages than on the PWR packages because of the larger 

60Co activity per fuel rod surface area associated with the BWR fuel. As a result, the BWR 

package allowable leakage rates are smaller (i.e., more restrictive) than the PWR package 

allowable leakage rates. For GBC-68 and the 5-year decay time, an increase in the fraction of 

crud that spalls off the cladding by a factor of two would cause a decrease in the allowable 

leakage rate by a factor of approximately two. 

Fraction of gases released due to cladding breach. For each decay time and fraction of 

breached spent fuel rods, an increase in the fraction of gaseous species from 0.1 to 0.4 would 

cause a relatively small decrease (up to ~30%) in the allowable release rate. 

Fraction of volatiles released due to cladding breach. The volatile source term is dominated 

by 137Cs (t1/2 = 30.07 years) and 90Sr (t1/2 = 28.78 years). An increase in the fraction of volatile 

source term by one order of magnitude would cause a maximum decrease in the allowable 

leakage rate by a factor of approximately two.

Fraction of fuel fines released due to cladding breach. Allowable leakage rate exhibits the 

greatest sensitivity to changes in the mass fraction of fuel released as fuel fines due to cladding 

breach. Depending on the crud contribution to the total releasable activity, an increase in the 

fraction of fuel fines by a factor of 10 would cause a decrease in the allowable leakage rate by a 

factor of ~1.5 to 10.

Fuel pellet radial region. Allowable radionuclide release rate and leakage rate for high-burnup 

fuel vary as a function of the pellet regions from which the radioactive material is released. 

Radioactive material released from the pellet peripheral region produced smaller allowable 

leakage rates than the radioactive material released from the non-rim region of the fuel pellet. 

The extent to which the pellet radial region from which the radioactive material is released

affects the allowable leakage rate depends on the contribution of crud to the total releasable 

activity. Small effects are observed for GBC-32 with 5-year decay time and a small fraction of 

breached spent fuel rods (e.g., 0.01), as well as for GBC-68 with 5-year decay time and any 

fraction of breached spent fuel rods, as crud has a significant contribution to the total releasable 
37

activity for these cases. For the cases where crud has a small or insignificant contribution to the 

total releasable activity (e.g., failed fuel fraction of 0.1 and decay time of 40 years), the 

allowable leakage rate based on radioactive material from the outer rim structure is 

approximately half the allowable leakage rate based on radioactive material from the non-rim 

region. The importance of the pellet region from which the radioactive material is released

increases with increasing decay time and breached spent fuel rod fraction for NCT. This is 

illustrated in Figure 19, which shows the effects of the pellet location from which the releasable 

activity may originate (i.e., either from the non-rim region of the fuel pellet or only from the rim 

structure) on the allowable leakage rate for the GBC-32 cask with decay times of 5 and 40 

years, and varying the fraction of fuel rods with cladding breaches. The release fractions for 

gases, volatiles, fuel fines, and crud that were used to calculate the allowable leakage rates 

illustrated in Figure 19 are those typically used for low-burnup fuel as provided in NUREG-1617

[29].

Figure 19 Effects of the location from which the fuel fines may originate on allowable 

leakage rate for the GBC-32 cask at 65 GWd/MTU burnup

The containment analysis for high-burnup fuel under HAC evaluated the impact on the allowable 

leakage rate of fraction of gases (0.3 and 0.4), volatiles (2 × 10-4 and 2 ×10-3

), and fuel fines 

(3 ×10-5 and 3 × 10-4

) released due to cladding breach, as well as the fuel pellet radial region 

from which the releasable activity originates. The fraction of fuel rods that have cladding 

breaches was 1, and the fraction of crud that spalls off the cladding was 1. For each fuel type

and decay time, five different cases were analyzed, with each case using a combination of the 

values previously described for the parameters important to containment analysis. The 

calculated allowable leakage rates are provided in the tables and graphs in Appendix C.5.4.

Similar to the NCT cases, the HAC allowable leakage rate has the largest sensitivity to the mass 

fraction of fuel released as fines. Depending on the decay time, an increase in the mass fraction 

of fuel released as fuel fines by a factor of 10 can cause a decrease in the allowable leakage 

rate by a factor of ~5 to 10 for the GBC-32 package and by a factor of ~3.5 to 10 for the GBC-68

package.
38

3.1.4 Thermal

The consequences of cladding failure require thermal calculations to investigate the changes 

associated with both breached and damaged fuel geometries. The thermal evaluations are 

limited to a representative GBC-32 transportation package with representative W 17×17 OFA

SNF as described in Appendix D. It is anticipated that the thermal response of a similarly 

designed BWR transportation package would follow similar trends as the analyzed PWR 

system. Both the vertical and horizontal configurations were considered. The primary 

parameters of interest are the component and cladding temperatures, and the peak surface 

temperature (with respect to 10 CFR 71.43). 

For the thermal analysis results presented, the case identifier (ID) nomenclature is as follows: 

The first letter indicates whether the package is in the vertical, “V,” or horizontal, “H,” orientation. 

The next number indicates the number of years of fuel decay time (i.e., “40” is 40-year decay 

time). The nominal intact configuration case is 40 years. The next set of numbers preceding the 

“pr” indicates the percentage of rods assumed to be ruptured (i.e., “3.125pr” indicates 3.125% of 

the rods have been breached). A number followed by “pf” indicates the fuel is assumed to be in 

a particle bed geometry, and the number indicates the packing fraction of the debris (i.e.,

“.313pf” is a packing fraction of 0.313). A number followed by “pdr” indicates the assembly fuel 

rod pitch has been altered, and the number indicates the pitch to diameter ratio (i.e., “116pdr” is 

a pitch to diameter ratio of 1.16). 

This fuel failure configuration category considered two sets of cases. The first set is 

representative of breached spent fuel rod configurations where a fraction of the fuel rod cladding 

is assumed to fail, releasing the gaseous contents, but otherwise remaining in their nominal 

geometry. The second case set represents damaged fuel and evaluates either one assembly or 

all assemblies damaged within their respective basket cells.

Breached spent fuel rods. Depending on the package orientation (i.e., horizontal or vertical) 

release of fission product gases into the canister gas space can cause either a significant 

increase or a significant decrease to component peak temperatures as illustrated in Table 15.

 

Table 15 Thermal results for scenario S1(a)—breached spent fuel rods

Case

ID

Fraction 

of rods 

failed

Change in temperature (∆°C)

Max.

clad

Min.

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSb

Max.

neutron 

shield

V.40a 0.0 0 0 0 0 0 0 0

V.40.3.125pr 0.3125 -14 +2 0 +1 -13 -13 0

V.40.10pr 0.10 -29 +8 -1 +2 -27 -27 0

V.40.50pr 0.50 -59 +19 -2 +5 -56 -56 -2

V.40.100pr 1.00 -74 +24 -2 +6 -71 -71 -2

H.40 0.0 0 0 0 0 0 0 0

H.40.3.125pr 0.3125 +4 0 0 0 +3 +3 0

H.40.10pr 0.10 +11 +1 0 0 +10 +10 0

H.40.50pr 0.50 +33 +2 0 0 +30 +31 -1

H.40.100pr 1.00 +45 +3 0 0 +42 +42 -1 aNominal intact configuration.

bSS = stainless steel.
39

For packages in a vertical orientation, the release of fission product gases leads to an overall 

decrease in peak component temperatures, while resulting in an increase in minimum 

temperatures. The recirculating mass flow rate for 100% rod break is 22.78 kg/min compared to 

0.57 kg/min for the 0% rod break nominal intact configuration case. This is primarily because of

the increase in gas density with fission product gas release but, it is also because of higher 

predicted recirculation velocities (e.g., 0.055 m3

/s for 100% rod break vs. 0.030 m3/s for 0% rod 

break at 204 ̊C). The specific heat capacity of the gas phase decreases with increasing fission 

product gases. However, the net effect of increased density and decreased heat capacity is a 

higher volumetric heat capacity of the flow. Thus, the convective flow is able to transport more 

heat with increasing Xe and Kr gas content. In other words, the convective flow can transport 

the same amount of heat at lower temperature differentials (e.g., the convective flow can 

transport approximately 175 W/K at 100% rod failure vs 50 W/K at 0% rod failure). This results 

in a flattening of the temperature profiles within the package (compare Figure D.17 to Figure 

20).

For the horizontal package, the release of fission gas products leads to an overall increase in 

peak and minimum component temperatures. In the horizontal package orientation, there is 

minimal convective heat transport within the canister. Therefore, the heat removal relies upon 

thermal radiation and conduction. The release of the fission products has a negligible impact on 

thermal radiation heat transfer; however, it reduces the thermal conductivity of the canister gas. 

The reduced thermal conductivity of the gas results in poorer heat transfer through the gas 

space and higher internal component temperatures. The change in gas properties has a 

negligible impact on the axial temperature profile of components within the package (compare 

Figure D.18 to Figure 21).
40

Figure 20 Structural temperatures with 100% rod breach—vertical orientation

Figure 21 Structural temperatures with 100% rod breach—horizontal orientation.
41

Damaged fuel. The fuel reconfiguration where a single, central, assembly was represented as

a debris pile that inside its basket cell had a minor impact on the component temperatures (see 

Table 16). The single assembly fuel reconfiguration had a larger impact on the component 

temperatures for the vertical orientation than for the horizontal orientation. Most of the impact is 

because of the release of the fission product gases into the canister space as opposed to the 

formation of the particle bed (Figure 22 and Figure 23). This can be seen by comparing the case 

where 3.125% of the rods are breached (equal to one assembly) but stay in their nominal 

geometry (cases V.40.3.125pr and H.40.3.125pr in Table 16) to the cases where the same 

amount of gas is released, but the assembly forms a debris bed (Table 16). The packing fraction 

of the debris bed had a minor impact on the predicted temperatures.

Table 16 Single assembly failure results summary

Case

ID

Packing 

fraction

Change in temperature (∆°C)

Max.

intact

clad

Min.

intact

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSb

Max.

neutron 

shield

V.40a NA 0 0 0 0 0 0 0

V.40.3.125pr NA -14 +2 0 +1 -13 -13 0

V.40.3.125pr.313pf 0.313 -11 +2 0 +1 -9 -8 0

V.40.3.125pr.320pf 0.320 -11 +2 0 +1 -8 -8 0

V.40.3.125pr.417pf 0.417 -12 +3 0 +1 -11 -11 0

V.40.3.125pr.574pf 0.574 -13 +3 0 +1 -13 -13 0

V.40.3.125pr.626pf 0.626 -14 +3 0 +1 -14 -14 0

H.40 NA 0 0 0 0 0 0 0

H.40.3.125pr NA +4 0 0 0 +3 +3 0

H.40.3.125pr.313pf 0.313 +3 0 0 0 +2 +2 0

H.40.3.125pr.417pf 0.417 +2 0 0 0 +1 +1 0

H.40.3.125pr.626pf 0.626 +1 +1 0 0 0 0 0

Note: NA = not applicable. aNominal intact fuel configuration.

bSS = stainless steel.
42

Figure 22 Change in peak neutron absorber temperature with failure of one assembly—

vertical cask

Figure 23 Change in peak neutron absorber temperature with failure of one assembly—

horizontal cask

For configurations where all of the assemblies are represented as debris piles, which remain 

inside their respective basket cells, large impacts on the predicted internal cask temperatures

were observed as shown in Table 17. Significant increases in package component temperatures 

could result in the package not meeting the thermal requirements for such systems. For the 

vertical orientation, the maximum temperature of the basket stainless steel walls and neutron 

absorber material increased by over 79°C compared to the nominal intact configuration case as 

shown in Figure 24. The assembly debris is assumed to block the basket cell channels causing 

a reduction in the convective heat transfer within the canister. In addition, the release of the 

fission product gases results in a decrease in gas thermal conductivity, resulting in lower heat 

conduction and heat transfer through the gas space. Both effects—lower convection and lower 

gas thermal conductivity—result in the large increase in internal temperatures. Increasing the 

packing fraction of the debris caused a greater increase in the maximum basket wall 

temperature. This is expected as the debris, generating the same amount of heat, has less 
43

debris-to-wall contact area for heat removal as the packing fraction is increased. Axial 

conduction reduces the thermal impact to structures toward the outer surfaces of the cask.

Table 17 Failure of all assemblies results summary

Case

ID

Debris 

packing 

fraction

Change in temperature (∆°C)

Max.

intact

clad

Min.

intact

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSb

Max.

neutron 

shield

V.40a NA 0 0 0 0 0 0 0

V.40.100pr NA -74 +24 -2 +6 -71 -71 -2

V.40.100pr.313pf 0.313 NA NA -5 +10 +79 +79 -6

V.40.100pr.417pf 0.417 NA NA -5 +8 +102 +102 -5

V.40.100pr.626pf 0.626 NA NA -3 +7 +127 +128 -2

H.40a NA 0 0 0 0 0 0 0

H.40.100pr NA +45 +3 0 0 +42 +42 -1

H.40.100pr.313pf 0.313 NA NA -4 -1 +30 +31 -5

H.40.100pr.417pf 0.417 NA NA -2 -2 +31 +31 -3

H.40.100pr.626pf 0.626 NA NA -1 -2 +21 +21 -2

Note: NA = not applicable. aNominal intact fuel configuration. 

bSS = stainless steel.

Figure 24 Change in peak neutron absorber temperature with failure of all assemblies in 

vertical cask

For the horizontal orientation, the fuel reconfiguration (i.e., all assemblies represented as debris 

piles within their respective basket cells) also increased the maximum temperature of the basket 

stainless steel walls and neutron absorber material. The internal heat transport for the horizontal 

case relies on heat conduction and thermal radiation as there is limited internal convection both 

in the nominal geometry and in the fuel reconfiguration geometry. The debris was represented 

as being evenly distributed along the length of the basket cell with a uniform decay heat profile. 

By spreading out the energy source term and putting the debris in direct contact with the basket 

cell walls, the maximum basket temperatures are increased by an amount less than the 
44

increase predicted for the case where 100% of the rods break, but the nominal geometry of the 

fuel latice is retained (case H.40.100pr). At a packing fraction of 0.313, the debris fully fills the 

basket cell and comes in contact with the top wall of the basket cell. The increased debris-wall 

contact area, for the 0.313 packing fraction case, causes the non-linearity in the trend in

Figure 25.

Figure 25 Change in peak neutron absorber temperature with failure of all assemblies in 

horizontal cask

3.2 CATEGORY 2: ROD/ASSEMBLY DEFORMATION

Fuel reconfigurations associated with rod/assembly deformation investigate the impact of pin 

pitch changes that could result as potential end states associated with spacer grid failure, and 

side and end drop events consistent with normal, off-normal, and accident conditions of storage 

as well as under normal and accident conditions of transportation. 

3.2.1 Criticality

Because of the effects of geometry changes on criticality calculations, changes in pin pitch can 

result in a wide range of reactivity effects. Both uniform and non-uniform pin pitch variations 

were considered to investigate the effects of axial and radial rod/assembly deformation on 

criticality calculations. To investigate the effects associated with side impact/drop events, cases 

were evaluated to represent configuration scenario S2(a)—configurations associated with side 

drop, as depicted in Figure 5 and Figure 6 for the PWR and BWR assembly, respectively.

Because LWR assemblies are under-moderated, configurations that result in pin pitch 

contraction will result in decreases in keff. 

Assembly deformation from end drop events is more significant for criticality safety implications. 

Axial effects to represent the phenomenon referred to as “birdcaging” and/or “bottlenecking” 

were evaluated by expanding pitches in one region and contracting pitches in another region. 

Results for configuration scenario S2(b)—configurations associated with end drop— are 

illustrated in Figure 26 for PWR assemblies, and in Figure 27 and Figure 28 for BWR 

assemblies with and without channels, respectively. Maximum changes to keff occurred when 
45

radial nonuniform pin pitch variations were modeled. A cross-sectional view of a PWR and 

BWR assembly with this configuration was illustrated in Figure 8. 

Figure 26 GBC-32 criticality analysis results from pin pitch variation

Figure 27 GBC-68 criticality analysis results from pin pitch variation for channeled fuel

0.00

0.50

1.00

1.50

2.00

2.50

3.00

3.50

4.00

4.50

1.25 1.30 1.35 1.40 1.45 1.50

Increase in keff (% 

Δ

keff)

Fuel Rod Pitch in Central Uniform Region (cm)

Uniform pitch 

expansion

Nonuniform 

pitch expansion

0.0

0.5

1.0

1.5

2.0

2.5

3.0

1.25 1.3 1.35 1.4 1.45 1.5 1.55 1.6

Increase in 

keff (%

Δ

keff)

Fuel rod pitch (cm)

Uniform pitch expansion

Nonuniform pitch expansion
46

Figure 28 GBC-68 criticality analysis results from pin pitch variation for unchanneled fuel

The effects of axial varying pitch were bounded by the nonuniform radial pin pitch variations. 

Investigation of the combined effect of at least one pin pitch contracted region with the other 

regions in the bounding radial nonuniform expansion cases showed essentially no change in 

results when compared with the bounding radial nonuniform expansion (i.e., 3.89 %∆keff versus 

3.90 %∆keff).

3.2.2 Shielding

Fuel rod lattice collapse against the fuel basket plates within a transportation package was 

analyzed for the rod/assembly deformation category (see Appendix B, Figure B.8 for model 

illustration). The assembly axial location for this fuel reconfiguration is identical to that of the 

nominal intact configuration. The calculation results indicate that fuel rod lattice collapse causes 

a small increase in the radial dose rates relative to intact fuel for both PWR and BWR SNF

assemblies. For the specific models used in this study, the maximum radial dose rate increase 

was ~20%. Collapsed fuel rod lattices also cause an increase in neutron dose rates at the 

package bottom and top surfaces because of neutron streaming. For the specific models used 

in this study, the maximum axial neutron dose rate increase was ~50%. The change in gamma 

dose rate at the package top and bottom surfaces was negligible (i.e., within statistical error).

The fuel reconfiguration analyzed is considered to be bounding for assembly deformation 

involving fuel rod lattice collapse with respect to package radial and axial dose rates because it 

provides the greatest neutron streaming and radiation source density near the package radial 

surface.

The ratio of the maximum surface dose rate value for a fuel reconfiguration to the maximum 

surface dose rate value for the nominal intact fuel configuration, identified as the F/I dose rate 

ratio, is summarized in Table 18 and Table 19 for the evaluated PWR and BWR fuel 

reconfigurations, respectively.

0

2

4

6

8

10

12

14

1.25 1.30 1.35 1.40 1.45 1.50 1.55 1.60 1.65 1.70 1.75

Increase in 

keff (%

Δ

keff)

Fuel rod pitch (cm)

Uniform Pitch Expansion Nonuniform Pitch Expansion
47

Table 18 PWR package maximum dose rate change for NCT

Package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

11

Assembly 

lattice 

collapse

See Figure B.8 (a) 

5 1.36 0.96 1.29 1.13 1.11 1.12 1.22 0.93 1.10

40 1.36 0.87 1.36 1.14 1.10 1.13 1.19 0.64 1.16

12

Assembly 

lattice 

collapse

See Figure B.8 (b)

5 1.27 0.99 1.22 1.03 0.93 0.98 1.17 0.96 1.10

40 1.16 0.89 1.16 1.04 0.96 1.03 1.16 0.80 1.14

Two meters from package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

11

Assembly 

lattice 

collapse

See Figure B.8 (a) 

5 1.31 0.98 1.24 1.15 1.10 1.12 1.19 0.93 1.09

40 1.32 1.04 1.31 1.17 1.09 1.15 1.17 0.73 1.14

12

Assembly 

lattice 

collapse

See Figure B.8 (b)

5 1.22 1.00 1.17 1.04 0.94 0.99 1.15 0.95 1.07

40 1.12 0.96 1.12 1.04 0.96 1.02 1.11 0.77 1.09

aRelative error (95% confidence level) less than 5%.

Table 19 BWR package maximum dose rate change for NCT

Package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

7 Assembly lattice 

collapse See Figure B.8 (a) 

5 1.48 1.01 1.41 1.07 1.08 1.07 1.26 1.00 1.14

40 1.49 1.01 1.49 1.06 1.07 1.06 1.27 0.89 1.27

Two meters from package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

7 Assembly lattice 

collapse See Figure B.8 (a) 

5 1.41 1.00 1.32 1.06 1.05 1.06 1.20 1.00 1.09

40 1.37 0.96 1.37 1.04 1.05 1.04 1.19 0.89 1.19

aRelative error (95% confidence level) less than 5%.
48

For the analysis under HAC, the dose rate at 1 m from the package external surfaces was

calculated with a model using air in place of the outer neutron shield considered in the package

models for NCT. The ratio of the maximum surface dose rate value for the fuel reconfiguration

analyzed to the maximum surface dose rate value for the nominal intact configuration (as the F/I 

dose rate ratio), is summarized in Table 20 for the PWR and BWR package models. The 

maximum radial neutron and gamma dose rate values increased by ~15%; the maximum 

neutron dose rates on the package top and bottom surfaces increased by ~30% and 20%, 

respectively.

Table 20 Package maximum dose rate change for HAC

1 m from the PWR package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

11 Assembly lattice 

collapse See Figure B.8 (a) 

5 1.31 0.99 1.24 1.12 1.09 1.12 1.19 0.93 1.13

40 1.30 1.04 1.30 1.13 1.06 1.13 1.17 0.66 1.15

12 Assembly lattice 

collapse See Figure B.8 (b)

5 1.13 1.00 1.10 1.04 0.94 1.03 1.13 0.95 1.09

40 1.13 0.91 1.13 0.95 0.90 0.94 1.14 0.74 1.13

1 m from the BWR package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

7 Assembly lattice 

collapse See Figure B.8 (a) 

5 1.07 1.00 1.04 1.03 1.06 1.04 1.20 1.00 1.13

40 1.03 1.00 1.03 1.03 1.04 1.03 1.21 1.02 1.21

aRelative error (95% confidence level) less than 5% for all values.

3.2.3 Containment

The rod/assembly deformation category does not include cladding failure. All cladding failure 

scenarios are evaluated under Category 1 (Sect. 3.1). Normal, off-normal, and accident 

conditions of spent fuel storage and normal and accident conditions of transportation that can 

cause rod/assembly deformation can affect the amount of crud that spalls off the cladding. 

Hence, the allowable leakage rate was calculated as a function of the fraction of crud that spalls 

off the cladding to evaluate the effect of varying crud releasable activity (see Appendix C.6). The 

allowable leakage rate is inversely proportional to the fraction of crud assumed to spall off the 

cladding. For the 5-year decay time, an increase in the fraction of crud from 0.05 to 1 would 

cause a decrease in the allowable leakage rate from 6.58 × 10-4 cm3

/s to 3.29 × 10-5 cm3/s for 

GBC-32 and from 7.54 × 10-5 cm3

/s to 3.77 × 10-6 cm3/s for GBC-68. For the 40-year decay 

time, an increase in the fraction of crud from 0.05 to 1 would cause a decrease in the allowable 

leakage rate from 6.56 × 10-2 cm3

/s to 3.28 × 10-3 cm3/s for GBC-32 and from 7.52 × 10-3 cm3/s 

to 3.76 × 10-4 cm3

/s for GBC-68.

3.2.4 Thermal

Depending on the initial assembly geometry and whether the package relies on convection for 

heat transfer, changes to the lattice pitch could cause either an increase or decrease in the 

package component temperatures as shown in Figure 29 and Table 21. For the thermal 

analysis results presented, the case ID naming convention is as described in Sect. 3.1.4.
49

Figure 29 Change in peak cladding temperature vs. rod pitch to diameter ratio

Table 21 Rod/assembly deformation results summary

Case

ID

Pitch to 

diameter 

ratio

Change in Temperature (∆°C)

Max.

clad

Min.

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSb

Max.

neutron 

shield

V.40.116pdr 1.16 -51 +11 -2 +3 -51 -51 -1

V.40.125pdr 1.25 -30 +5 -1 +2 -30 -30 0

V.40.135pdr 1.35 -1 0 0 0 0 0 0

V.40a 1.38 0 0 0 0 0 0 0

V.40.140pdr 1.40 -4 0 -1 +1 -3 -3 0

H.40.116pdr 1.16 +4 0 0 0 +4 +4 0

H.40.125pdr 1.25 +7 0 0 0 +6 +6 0

H.40.135pdr 1.35 +3 0 0 0 +3 +3 0

H.40a 1.38 0 0 0 0 0 0 0

H.40.140pdr 1.40 -5 0 0 0 -5 -5 0 aNominal intact fuel configuration. bSS = stainless steel.

For the vertical orientation, both increasing and decreasing the fuel lattice pitch caused a 

decrease in the maximum cladding, basket wall, and neutron absorber temperature as shown in 

Table 21. The heat transport within the canister relies on complex parallel and intersecting 

paths of conduction, convection and thermal radiation. Decreasing the lattice pitch resulted in a 

higher predicted recirculating mass flow rate within the canister as shown in Figure 30, thereby

increasing convective heat transport. However, the flow loss coefficients for spacer grids and

entrance/exit losses were not modified between cases. Depending on how the assembly was 

deformed, the spacer grids and/or the entrance/exit flow losses may be impacted. Increasing 

the lattice pitch increases the cladding-to-basket wall view factors, thereby increasing thermal 

radiation heat transport. Of the cases analyzed, the nominal intact configuration case resulted in 

the highest temperatures.
50

For the horizontal orientation, changes to the fuel lattice pitch had only a minor impact on the 

package temperatures as shown in Table 21. Of the cases analyzed, the case with a rod pitchto-diameter ratio of 1.25 resulted in the highest temperatures.

Figure 30 Recirculating mass flow rate vs. rod pitch to diameter ratio for vertical package

3.3 CATEGORY 3: CHANGES TO INTACT ASSEMBLY AXIAL 

ALIGNMENT 

Changes to the nominal intact axial alignment of the spent fuel within the canister basket can 

impact the design basis safety analyses that are based on a fixed geometry.

3.3.1 Criticality

The consequences of changes to intact assembly axial alignment require criticality calculations 

to investigate the changes in keff associated with active fuel exposure above and below the 

absorber plate envelope within a basket cell. The upward (toward the lid) misalignment is of 

higher importance for SNF criticality evaluations because the reactivity of the assembly is driven 

by the burnup gradient near the top of the assembly. Because of differences in canister design, 

fuel spacers, and assembly designs, the amount of space available for the active fuel to 

potentially be outside the absorber envelope will vary, which affects the extent of the range of 

possible movement. As such, representative ranges of active fuel exposure were evaluated. 

The PWR and BWR fuel axial displacement models were developed such that all the fuel 

assemblies within a canister are moved uniformly up or down relative to their original position. 

Exposed fuel ranges were varied to a maximum of 30 cm above the absorber. The reactivity 

increase is calculated at several points over the range to map the response; the maximum 

misalignments considered are intended to be larger than realistic misalignments to remove any 

design-specific constraints. The variation of the keff change as a function of axial position is

shown in Figure 31 for several PWR fuel initial enrichment and burnup combinations and in 

Figure 32 for the BWR fuel initial enrichment and burnup combinations. Maximum changes in 
51

keff for 30 cm versus 20 cm for PWR fuel, and for 31.78 cm versus 13.65 cm for BWR fuel are 

provided in Table 22 for comparative purposes.

Figure 31 GBC-32 criticality analysis results from axial misalignment

Figure 32 GBC-68 criticality analysis results from axial misalignment

-1

0

1

2

3

4

5

6

7

8

9

-15 -10 -5 0 5 10 15 20 25 30

Change in 

keff (%

Δ

keff)

Assembly axial displacement (cm)

1.92 wt% U-235, Fresh 3.5 wt% U-235, 25.5 GWd/MTU, 5 year

5 wt% U-235, 44.25 GWd/MTU, 300 year 5 wt% U-235, 70 GWd/MTU, 300 year

-2

0

2

4

6

8

10

12

14

16

18

20

22

-15 -10 -5 0 5 10 15 20 25 30 35

Increase in 

keff (%

Δ

keff)

Uncovered Fuel (cm)

Fresh - Channeled Fresh - Unchanneled

35 GWd/MTU, 80 year - Channeled 35 GWd/MTU, 80 year - Unchanneled

70 GWd/MTU, 80 year - Channeled 70 GWd/MTU, 80 year - Unchanneled

Uncovered fuel at top of assembly Uncovered fuel at bottom 

of assembly
52

Table 22 Criticality results for axial displacement of intact fuel 

Parameter

PWR system

Burnup (GWd/MTU) 0 25.5 44.25 70

Initial enrichments (wt % 235U) Maximum change in keff (%∆keff)

(30 cm/20 cm exposed fuel)

1.92 1.52/0.06 NC NC NC

3.5 with 5-year decay time NC 7.07/3.22 NC NC

5.0 with 5-year decay time NC NC 7.69/3.64 8.06/4.13

5.0 with 80-year decay time NC NC 8.28/4.18 8.28/4.51

5.0 with 300-year decay time NC NC 8.26/4.15 8.34/4.54

BWR system

Burnup (GWd/MTU) 0 35 70

Initial enrichment (wt % 235U) Maximum change in keff (%∆keff) channeled (31.78 cm 

exposed fuel/13.65 cm exposed fuel)

5.0 with 5-year decay time 12.52/0.33 19.40/6.29 20.47/8.03

5.0 with 80-year decay time NC 19.84/6.70 20.73/8.52

5.0 with 300-year decay time NC 19.83/6.66 20.76/8.49

Maximum change in keff (%∆keff) unchanneled (31.78

cm exposed fuel/13.65 cm exposed fuel)

5.0 with 5-year decay time 12.02/0.27 18.66/6.07 19.71/7.78

5.0 with 80-year decay time NC 19.10/6.49 19.96/8.24

5.0 with 300-year decay time NC 19.06/6.42 20.00/8.20

Note: NC = not calculated.

Overall, the results indicate that the effects of axial misalignment of intact fuel assemblies 

increase with increased active fuel exposure beyond the basket neutron absorber envelope, and 

that displacement towards the canister bottom is inconsequential. For most canister systems, 

the available displacement distance will be limited because of the presence of assembly end

fitting components and fuel spacers. The available displacement distance will be canister 

design, fuel assembly spacer design, and assembly design dependent.
53

3.3.2 Shielding

The consequences associated with changes to the intact assembly axial alignment were

analyzed assuming that the fuel assemblies could reach either the top or bottom surface of the 

inner package cavity. Fuel assembly displacement towards the canister cavity top causes an 

increase in dose rates at the external top surface and at the external radial surface above the 

neutron shield. Fuel assembly displacement toward the canister cavity bottom causes an 

increase in dose rates at the external bottom surface and at the external radial surface below 

the neutron shield. However, for the PWR package radial surface, maximum neutron and 

gamma dose rate values were only greater by ~15% and 30%, respectively, whereas the BWR 

package maximum radial neutron and gamma dose rate values did not change relative to the 

nominal intact fuel configuration. The PWR package maximum neutron and gamma dose rate 

values were ~40% and 80% greater at the external bottom surface and ~30% and 70% greater 

at the external top surface, respectively, relative to the nominal intact fuel configuration. The 

BWR package maximum neutron and gamma dose rate values were ~20% and 55% greater at 

the external bottom surface and ~15% and 40% greater at the external top surface, relative to 

the nominal intact fuel configuration. The maximum dose rate increase for the package axial 

surfaces depends on the free void volume available above or below the fuel assemblies, which 

varies as a function of fuel assembly length and package design. For the models analyzed, the 

free void volume was smaller for the BWR fuel than for the PWR fuel because the GE14 fuel 

assembly is longer than the W 17×17 OFA.

The ratio of the maximum surface dose rate value for the fuel reconfiguration to the maximum 

surface dose rate value for the nominal intact configuration (the F/I dose rate ratio) is 

summarized in Table 23 and Table 24 for the evaluated PWR and BWR fuel reconfigurations, 

respectively, under NCT.

Table 23 PWR package maximum dose rate change for NCT

Package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

13 Alignment 

changes See Figure B.9 (a)

5 0.80 0.60 0.76 1.15 1.00 0.99 1.35 1.66 1.46

40 0.80 0.58 0.80 1.14 1.00 1.00 1.34 1.77 1.36

14 Alignment 

changes See Figure B.9 (b)

5 1.26 1.59 1.32 1.10 1.29 1.20 0.78 0.59 0.71

40 1.26 1.68 1.27 1.10 1.01 0.99 0.77 0.57 0.76

2 m from package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

13 Alignment 

changes See Figure B.9 (a)

5 0.79 0.60 0.75 1.01 1.00 1.01 1.30 1.67 1.45

40 0.80 0.86 0.80 1.01 1.01 1.00 1.29 1.74 1.32

14 Alignment 

changes See Figure B.9 (b)

5 1.25 1.61 1.33 1.02 1.00 1.01 0.79 0.59 0.71

40 1.28 1.57 1.28 1.02 1.00 1.02 0.78 0.57 0.77

aRelative error (95% confidence level) less than 5%.
54

Table 24 BWR package maximum dose rate change for NCT

Package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

8 Alignment 

changes See Figure B.9 (a)

5 0.88 0.69 0.85 1.01 1.00 1.01 1.16 1.44 1.29

40 0.89 0.71 0.88 1.01 1.01 1.01 1.17 1.54 1.17

9 Alignment 

changes See Figure B.9 (b)

5 1.12 1.42 1.17 1.00 1.00 1.00 0.86 0.71 0.79

40 1.14 1.39 1.14 1.00 1.01 1.00 0.87 0.72 0.87

2 m from package external surfaces Top Side Bottom

Case # Scenario Fuel relocation 

region

Decay time 

(years)

F/I dose rate ratioa

n γ Total n γ Total n γ Total

8 Alignment 

changes See Figure B.9 (a)

5 0.88 0.70 0.85 1.01 0.99 1.01 1.13 1.42 1.28

40 0.88 0.92 0.88 1.01 1.01 1.01 1.18 1.53 1.19

9 Alignment 

changes See Figure B.9 (b)

5 1.15 1.41 1.20 1.01 1.00 1.00 0.85 0.72 0.78

40 1.14 1.08 1.14 0.97 1.02 0.97 0.87 0.75 0.87

aRelative error (95% confidence level) less than 5%.

For the analysis under HAC, dose rates at 1 m from the package were calculated with a model 

using air in place of the outer neutron shield considered in the package models for NCT. The 

ratio of the maximum surface dose rate value for each fuel reconfiguration analyzed to the 

maximum dose rate value for the nominal intact configuration (the F/I dose rate ratio) is 

summarized in Table 25 for the PWR and BWR package models. The maximum neutron and 

gamma dose rate values on the package top and bottom surfaces increased by ~30% and 70%, 

respectively.

Table 25 GBC-32 maximum dose rate change for HAC

1 m from PWR package external surfaces

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation region Decay 

time (y)

F/I neutron dose rate 

ratioa

F/I gamma dose rate 

ratioa

13 Alignment

changes See Figure B.9 (a)

5 0.80 0.99 1.30 0.60 1.00 1.66

40 0.81 0.99 1.29 0.76 0.97 1.73

14 Alignment 

changes See Figure B.9 (b)

5 1.25 1.01 0.79 1.57 1.00 0.60

40 1.26 1.01 0.78 1.30 0.96 0.55

1 m from BWR package external surfaces

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation region Decay 

time (y)

F/I neutron dose rate 

ratioa

F/I gamma dose rate 

ratioa

8 Alignment 

changes See Figure B.9 (a)

5 0.90 1.00 1.16 0.70 0.99 1.41

40 0.89 1.00 1.16 0.92 0.99 1.35

9 Alignment 

changes See Figure B.9 (b)

5 1.09 1.00 0.87 1.41 0.99 0.72

40 1.06 0.99 0.89 1.10 0.99 0.88

aRelative error (95% confidence level) less than 5%.
55

The effects of assembly alignment change [See Figure B.9 (a)] were also evaluated at 1 m from

an individual SNF storage cask in a vertical orientation. The maximum dose rate increase, at 

1 m from the cask side surface in a plane through the cask air vent locations, was a factor of 

~1.4 for the neutrons and a factor of ~2.7 for gamma rays. The maximum value of the total 

dose rate increased by a factor of 2.7.

3.3.3 Containment

The containment consequences for this category are the same as for Category 2 (Sect. 3.2.3) 

because cladding failure is not included in the changes with assembly axial alignment. All 

cladding failure scenarios are evaluated under Category 1 (Sect. 3.1). Normal, off-normal, and 

accident conditions of spent fuel storage and normal and accident conditions of transportation 

that can cause a change in SNF axial alignment can affect the amount of crud that spalls off the 

cladding. Hence, allowable leakage rate was calculated as a function of the fraction of crud that 

spalls off cladding to evaluate the effect of varying crud releasable activity (see Appendix C.6).

The allowable leakage rate is inversely proportional to the fraction of crud assumed to spall off

the cladding. For the 5-year decay time, an increase in the fraction of crud from 0.05 to 1 would 

cause a decrease in the allowable leakage rate from 6.58 × 10-4 cm3

/s to 3.29 × 10-5 cm3/s for 

GBC-32 and from 7.54 × 10-5 cm3

/s to 3.77 × 10-6 cm3/s for GBC-68. For the 40-year decay 

time, an increase in the fraction of crud from 0.05 to 1 would cause a decrease in the allowable 

leakage rate from 6.56 × 10-2 cm3

/s to 3.28 × 10-3 cm3/s for GBC-32 and from 7.52 × 10-3 cm3/s 

to 3.76 × 10-4 cm3

/s for GBC-68.

3.3.4 Thermal

Changes in assembly axial alignment within the basket cells had a minor impact on the 

component temperatures as provided in Table 26. For the thermal analysis results presented, 

the case ID naming convention is as described in Sect. 3.1.4. 

For the vertical orientation, shifting the assemblies upwards caused an increase in the 

maximum component temperatures while lowering the minimum component temperatures. 

Shifting the assemblies upward moves the heat source higher in the canister. This results in 

less of a “chimney” to drive the circulation within the canister. The recirculating mass flow is 

0.498 kg/min when the assemblies are shifted up versus 0.566 kg/min for the nominal intact 

configuration case. When the assemblies are shifted downward, there is a larger “chimney” to 

drive the internal recirculating flow, and the peak temperatures are decreased.

For the horizontal orientation, shifting the assemblies had a minor impact on the maximum 

component temperatures. The maximum component temperatures occur toward the center of 

the canister. Shifting the assemblies mainly shifts the maximum temperature locations. The 

minimum cladding temperature was impacted more than the other component temperatures. 

The minimum cladding temperature is impacted by the heat loss through the ends of the 

package. By shifting the assemblies such that they become in contact with the ends of the 

cask/package, the minimum cladding temperature is decreased.
56

Table 26 Assembly axial alignment shift results summary

Case

ID

Axial 

shift

(cm)

Change in temperature (∆°C)

Max.

clad

Min.

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSb

Max.

neutron 

shield

V.40.up +27.27 +8 -10 +2 -2 +8 +9 +3

V.40a 0 0 0 0 0 0 0 0

V.40.dn -8.89 -3 +4 -1 +1 -2 -2 -1

H.40.up +27.27 +2 -12 +1 -3 +2 +2 +1

V.40a 0 0 0 0 0 0 0 0

H.40.dn -8.89 -1 +5 0 -2 -1 -1 0 aNominal intact fuel configuration.

bSS = stainless steel.
57

4. SUMMARY AND CONCLUSIONS

This report investigates the consequences of fuel configuration changes with regards to SNF

storage and transportation systems in four technical disciplines: criticality, shielding (dose 

rates), containment, and thermal. Three fuel reconfiguration categories were considered:

configurations characterized by (1) cladding failure, (2) rod/assembly deformation without 

cladding failure or (3) changes to assembly axial alignment without cladding failure. The 

analyses used representative SNF designs and storage cask/ transportation packages, and a 

range of fuel initial enrichments, discharge burnup values, and decay times. Consequences 

with respect to the different technical disciplines for the fuel reconfiguration categories were 

provided relative to the nominal intact configuration in Sect. 3, and detailed descriptions of the 

analyses performed are provided in Appendices A, B, C, and D. The findings of this research 

effort are intended to inform the safety evaluation of SNF storage and transportation systems by 

identifying fuel reconfigurations that would have the most significant impact on safety-significant 

parameters with respect to the four technical disciplines. However, there was no study or 

comparison of mechanical properties, finite element analysis stress distribution studies, or 

physical testing to evaluate the likelihood of each configuration. 

In all technical disciplines evaluated, the postuated configurations involving cladding failure 

exhibited the largest impact on safety-significant parameters. Therefore, it is recommended that 

these configurations be evaluated further on a cask-specific and assembly-specific basis to 

identify the impact of changes in analysis assumptions and geometry conditions. Further, future 

work to evaluate the credibility of these fuel reconfigurations is suggested to fully understand the 

actual potential impacts on safety when mechanical properties for high-burnup fuel are better 

understood. A summary of impacts associated with fuel configuration changes for each of the 

technical disciplines is provided below.

Criticality

A summary of pertinent results for the fuel reconfiguration categories as they apply to criticality

safety considerations is provided in Table 27. The results reported are with respect to the 

following nominal intact configuration cases: (1) PWR fuel at 44.25 GWd/MTU burnup for 5 wt% 

235U initial enriched fuel and 5-year decay time for all three categories; and (2) channeled BWR 

fuel at 5 wt% 235U initial enrichment with 35 GWd/MTU burnup and 5-year decay time for 

Categories 1 and 3, and for 5 wt% 235U initial enriched fresh fuel for Category 2. 

The maximum increase caused by a single failed rod was 0.10% ∆keff for the GBC-32 model 

and 0.29% ∆keff for the GBC-68 model. The failure of multiple rods increases the reactivity 

impacts to 1.86% ∆keff for GBC-32 and 2.40% ∆keff for GBC-68. The damaged fuel

configurations resulted in significantly higher reactivity increases. Having complete rubblization 

of the fuel is not physically likely under NCT as prescribed in 10 CFR 71.71, or as a result of 

impacts from accident loads specified in 10 CFR 71.73. These representations were meant to 

provide an estimate of the maximum potential increase (upper bound) in keff available, and to 

assess the impact of modeling simplifications such as representing the fuel reconfiguration as a 

homogeneous distribution rather than a heterogeneous distribution. The results indicated that 

homogeneous representations were non-conservative primarily due to the resonance selfshielding effects of low enriched fuel.
58

No significant differences were observed in trends between configurations that evaluated fuel at 

44.25 GWd/MTU and 70 GWd/MTU.

Table 27 Fuel reconfiguration results summary pertaining to criticality

Configuration 

scenario Case description Parameter varied

Maximum keff increase 

(% ∆keff)

(GBC-32/GBC-68)

Category 1: Cladding Failure

S1(a)—breached spent 

fuel rods

Multiple rod removal Multiple missing rod 

combinations until an upper 

limit is identified

1.86/2.40

Combination of multiple rod 

removal and rubble 

extended beyond absorber 

envelope (displaced fuel 

volume fraction = 0.341)

Number of missing rods and 

distribution of displaced fuel 

outside neutron absorber 

envelope

4.91b/NC

S1(b)—damaged SNF Uniform pellet arraya

(includes SNF distributed 

outside neutron absorber 

panel envelope) 

The pellet spacing, and thus 

the debris bed size, was 

varied to find the largest Δkeff

21.37/34.40b

Category 2: Rod/Assembly Deformation

S2(a)—configurations 

associated with side 

drop

Pin pitch contraction Pin pitch reduced from 

nominal intact configuration

Not applicable. Nominal 

intact configuration 

bounding for this 

configuration 

S2(b)—configurations 

associated with end 

drop

Uniform and non-uniform 

radial and axial pin pitch 

changes (birdcaging and 

bottlenecking).

Pitch of all rods radially 

expanded over different axial 

lengths until outer pins are in 

contact with basket cell wall

2.65 to 3.90/ 2.09 to 

13.31

Category 3: Changes to Assembly Axial Alignment

S3—axial displacement 

of intact fuel

Assembly shift exposing 

active fuel outside neutron 

absorber envelope

Length of active fuel above 

or below the neutron 

absorber plate

3.64 at 20 cm / 6.29 at 

20 cm

Note: NC = not calculated. aCase is bounding but not considered physically likely. bMaximum value from cases evaluated but optimum missing fuel and volume fraction distribution were not 

determined.

The effect of rod/assembly deformation was modeled with a series of configurations of fuel pin 

pitch expansions that could result from side or end drop events. For both PWR and BWR fuel, 

increases in pin pitch cause keff to increase, and the peak keff value was not reached before the 

pitch expansion became restrained by the fuel basket walls. For representations with uniform 

pin pitch expansion, the maximum increase in keff for PWR fuel I in the GBC-32 was 2.65%. 

Configurations where a non-uniform radial pin pitch was created with larger pin pitches in the 

center of the assembly and smaller pitches along the edges resulted in larger keff increases

relative to the nominal intact configuration. The largest PWR package keff increase resulting 

from this non-uniform radial pin pitch expansion was 3.90% ∆keff. For BWR fuel in the GBC-68, 

the largest increase in keff for dechanneled fuel was 13.16% ∆keff and 13.31% ∆keff for the 

uniform pin-pitch expansion and non-uniform pin-pitch expansion cases, respectively. 

Channeled BWR fuel resulted in an increase of 2.09% ∆keff. The effects of axial non-uniform 

pin-pitch expansion were also examined for both BWR and PWR fuel, and the reactivity 

increases observed were consistent with the non-uniform radial pin-pitch expansion results.
59

The effects of axial alignment changes of intact fuel assemblies increase with increased active 

fuel exposure beyond the basket neutron absorber envelope. The intact assemblies in both 

GBC-32 and GBC-68 were displaced axially through a range of distances to determine the 

reactivity effect of axial alignment changes. Axial alignment shifts toward the canister lid were

observed to be the most limiting. Misalignments of 20 cm result in changes of 3.64% ∆keff for 

PWR fuel and 6.29% ∆keff for BWR fuel. The available displacement distance is canister design, 

fuel assembly spacer design, and assembly design dependent. 

Further investigations should be conducted regarding fuel relocation outside the neutron 

absorber panel envelope. No credit was taken for volume displacement of guide tubes, other 

assembly hardware components, or fuel assembly spacers in the models which could 

significantly reduce the amount of reactivity increase. In addition, sensitivity analyses should 

also be performed to associate fuel displacement amounts with fuel particle size, hydrogen-tofuel ratio, and resonance absorption effects for low enriched fuel to better understand the 

modeling approximations for spent fuel distributed outside the neutron absorber panel region.

 

Shielding

A summary of pertinent results for the fuel reconfiguration categories as they apply to 

shielding/dose rate considerations is provided in Table 28. The results reported in the table are 

with respect to a radiation source term for 5.0 wt% 235U initially enriched fuel at a burnup of 65 

GWd/MTU and 40-year decay time. Those results are specific to the generic PWR and BWR 

shielding analysis models used in this report. Descriptions of the analysis models and 

evaluations with differences in burnup and decay time are provided in Appendix B. The impact 

of fuel configuration changes on transportation package/storage cask external dose rates was 

evaluated by comparing the dose rate values between intact and fuel reconfigurations. 

External gamma and neutron dose rate values were evaluated separately because fuel 

configuration changes have different effects on gamma and neutron radiation dose rates. The 

overall impact on the external dose rates depends on the individual gamma and neutron dose 

rate contributions to the total dose rate, which vary with package/cask design, assembly type, 

burnup, and decay time. The neutron dose rate changes were very similar for the two decay

times analyzed (5-year and 40-year decay times) because the neutron spectrum exhibits small 

variations as a function of decay time within the time interval of 5 to 40 years. Relative to the 

nominal intact configuration, the impact of fuel configuration changes on the external gamma 

dose rate was small for short decay times (e.g., 5 years) compared to longer decay times (e.g., 

40 years) because of the varying contribution as a function of decay time of the 60Co (t1/2 = 

5.271 years) activation sources in the assembly top and bottom hardware regions to total dose 

rate.

Characteristics of the canister internal components such as the heights of assembly hardware 

regions, fuel axial burnup profile, the lengths of the fuel assembly top and bottom spacers, and 

the available free volume within the canister cavity, can affect the relative changes in dose rate 

at the package axial surfaces caused by fuel configuration changes. In this study, the BWR fuel 

reconfigurations had more pronounced effects on the external dose rates than the PWR fuel 

reconfigurations. For the intact fuel configuration, assuming the same assembly burnup and 

decay time, the dose rates at the top and bottom surfaces of the generic BWR package were 

lower than the dose rates at the corresponding surfaces of the generic PWR package primarily 

because the BWR assembly used in this study has longer hardware regions and a more pointed 

axial source profile than the PWR fuel assembly. Relative to the nominal intact configuration, 
60

fuel relocation to the canister bottom or top regions caused a larger change in the dose rate at 

the outer axial surfaces of the BWR package compared with the PWR package.

Table 28 Fuel reconfiguration results summary pertaining to shielding

Configuration 

scenario Case description Parameter varied

Relative change in 

maximum dose ratea

Category 1: Cladding Failure

S1(a) – breached spent 

fuel rods

Transportation package; 

combination of multiple rod 

failure and source relocation 

maintained within the active 

fuel region, distributed to the 

top end-fitting, or distributed 

to the bottom end-fitting 

Number of missing rods 

and distribution of 

displaced fuel particulates 

at middle of active fuel 

region

PWR

Top: 1.1

Radial: 1.1

Bottom: 1.0

BWR

Top: 1.1

Radial: 1.0

Bottom: 1.1

Number of missing rods 

and distribution of 

displaced fuel particulates 

to top or bottom of 

assembly

PWR 

Top: 6.1

Radial: 3.0

Bottom: 2.9

BWR 

Top: 21.8

Radial: 2.1

Bottom: 4.6

One meter from a storage 

cask; multiple rod failure 

and source relocation 

distributed to the bottom 

end-fitting

Source distribution PWRb

Radial: 2.9

BWRb

Radial: 2.7

S1(b) – damaged SNF Transportation package; 

homogeneous fuel mixture 

distribution settled at bottom 

or uniformly distributed 

throughout the package 

cavity

Source distribution PWR 

Top: 6.7

Radial: 3.4

Bottom: 4.3

BWR 

Top: 23.6

Radial: 2.9

Bottom: 6.2

1 m from a storage cask; 

homogeneous fuel mixture 

distribution settled at bottom

Source distribution PWRb

Radial: 4.1

BWRb

Radial: 9.2

4×2 storage array 

evaluation at controlled area 

boundary

Source distribution PWR

1.8

BWR

2.4

Category 2: Rod/Assembly Deformation

S2(a) – configurations 

associated with side 

drop

Pin pitch contraction with 

fuel rods collapsed against 

fuel basket plates

Source distribution within 

basket cell

PWR

Top: 1.4

Radial: 1.1

Bottom: 1.2

BWR

Top: 1.5

Radial: 1.1

Bottom: 1.3

S2(b) – configurations 

associated with end 

drop

None. This case is bounded 

by Category 3 

representations

None. NC

Category 3: Changes to Assembly Axial Alignment

S3 – axial 

displacement of intact 

fuel

Transportation package; 

assembly shift allowing fuel 

assemblies to reach top or 

bottom surface of the 

canister cavity

Source location PWR 

Top: 1.3

Radial: 1.0

Bottom: 1.4

BWR 

Top: 1.2

Radial: 1.0

Bottom: 1.2

1 m from a storage cask; 

assembly shift allowing fuel 

assemblies to reach bottom 

surface of the inner cavity

Source location PWRb

Radial: 2.7 

BWRb

Radial: 1.2

Note: NC = not calculated. aRelative change in the maximum value of the total dose rate at each package/storage external surface. bLocations that receive radiation streaming through cask air vents.
61

The conclusions reported for the transportation package under NCT are based on the 

calculation results for the package external surface. As the results of the calculations show, the 

effects of fuel configuration changes on dose rates at the external surface of a package are 

either similar or bounding for the dose rates at the other relevant locations evaluated (i.e., 2 m

from the package as required by 10 CFR 71.47 for NCT and at 1 m from the package external 

surfaces as specified in 10 CFR 71.51 for HAC).

The effect of multiple fuel rod failures (Scenario S1[a]) on cask external dose rates was 

analyzed for different percentages of failed fuel rods per assembly with source redistribution to 

different assembly axial regions, including the active fuel and lower and upper hardware 

regions. An important parameter associated with the mass density and source strength of the 

fuel mixture that may collect into the available free space is the particle packing fraction. The 

shielding analysis used a closely packed mixture with a packing fraction of 0.58 based on 

powder mechanics for normal settling; with sensitivity calculations performed using a 0.67 

packing fraction. The results based on the two packing fraction values were within the statistical 

error of the Monte Carlo calculations (i.e., 7% at the 95% confidence level). 

For the multiple fuel rod failure cases analyzed, the fuel reconfigurations with fuel mixture being 

relocated into either the assembly bottom or top region caused a significant increase in the dose 

rate at the respective package outer bottom or top region. For example, the maximum value of 

the total dose rate at the PWR package bottom surface increased by a factor of 2.7 for the case 

using 10% of the assembly fuel mixture relocated into the assembly bottom region. A larger 

percentage (i.e., 25% versus 10% analyzed for the PWR fuel) of fuel mixture distributed to the 

assembly bottom region increased the maximum value of the total dose rate at the package 

bottom surface by a factor of 2.9. Hence, the effect of a larger failed fuel percentage was not

proportional to the source strength increase because of increased geometric attenuation 

associated with the larger mass density.

Models with homogenously distributed radiation sources within the inner cavity volume/bottom, 

simulating damaged fuel conditions (Scenario S1[b]), are bounding for other fuel failure 

configurations. Note that this configuration produces maximum dose rate increases that are 

larger than or similar to the analyzed PWR and BWR fuel rod failure cases with 25% and 11%, 

respectively, of the fuel mixture relocated to different fuel assembly axial locations. However, 

this type of configuration is not physically likely under normal, off-normal, and accident 

conditions of spent fuel storage and normal and accident conditions of transportation. This fuel 

reconfiguration may be used to evaluate the impact of radiation source relocation to the canister 

wall adjacent regions (i.e., closer to the package external surfaces) due to fuel configuration 

changes.

Configurations with fuel rod lattice collapse against the fuel basket plates was analyzed for the 

rod/assembly deformation category. This fuel reconfiguration is considered to be bounding for 

assembly deformation involving assembly lattice pitch changes with respect to the package

radial and axial dose rates because it provides greater radiation source density near the 

package radial surface and greater neutron axial streaming. For the specific models used in this 

study, the maximum increase in the total dose rates at the radial surfaces of the PWR and BWR 

packages increase was ~10%. The collapsed fuel rods cause an increase in dose rates at the 

package bottom and top surfaces because of neutron streaming. For the specific models used 

in this study, the maximum axial dose rate increase was ~50%. 

The effects of axial alignment shift are dependent on the available volume above or below the 

fuel assembly and result in increases in the dose rate at the package axial surfaces. For the 
62

PWR package model, the maximum values of the total dose rate at the top and bottom surfaces 

increased by ~ 30% and 40%, respectively; for the BWR package model, the maximum values 

of the total dose rate at the top and bottom surfaces increased by ~ 20%. The maximum value 

of the total dose rate at the cask radial surface did not change relative to the nominal intact

configuration. 

The dose rate increases at typical lid, vent port, and drain port o-rings, which are used as seals 

in some transportation packages, were similar to those obtained for the external surfaces of 

packages under NCT.

Fuel configuration changes within a SNF storage cask can cause significant dose rate changes 

relative to the nominal intact fuel configuration in the cask outer regions that receive radiation 

streaming through cask air vents. For example, for the scenario considering homogeneous fuel 

mixture distribution in the canister cavity bottom, the dose rate increased by a factor of ~4 for 

the PWR fuel and by a factor of ~9 for the BWR fuel. Away from the air vent locations, the 

change in radiation dose rate is either small (e.g., 30% increase for damaged fuel 

configurations) or negligible. 

The dose rates at the controlled area boundary were calculated at 100 m from a generic 4×2 

storage cask array using two damaged fuel configurations. The results show that the closely 

packed fuel mixture model reduces the site boundary dose rate by ~70%, whereas the model 

with the fuel mixture homogeneously distributed within the entire canister cavity increases the 

site boundary dose rate contribution from the storage casks by a factor of ~2.4, relative to the 

nominal intact configuration. The dose rate changes are caused by source geometry changes 

and gamma self-shielding effects associated with the fuel reconfiguration.

Containment

The containment analysis of low-burnup fuel used varying fractions of rods that develop 

cladding breach (i.e., fuel reconfiguration Category 1) under NCT, from 0.01 to 1, and the 

release fractions provided in NUREG-1617 [29] for the contributors to the releasable source 

term. For the high-burnup fuel analysis the parameters important for containment analysis were 

varied across the following ranges: the fraction of fuel rods assumed to fail was varied from 0.01 

to 1; the fraction of crud that spalls off cladding was varied from 0.15 to 1; the fraction of gases 

was varied from 0.1 to 0.4, the fraction of volatiles was varied from 2 × 10-4 to 2 × 10-3

, and the 

mass fraction of fuel released as fuel fines was varied from 3 × 10-5 to 3 × 10-4

. For each fuel 

type, decay time, and fraction of fuel rods assumed to fail, multiple cases were analyzed using 

different combinations of the values for the containment analysis. Additionally, the impact on the 

releasable activity of the discharged fuel particulate original location (i.e., non-rim pellet region 

or the peripheral rim structure) was also evaluated for high-burnup fuel. Fuel reconfiguration

categories identified as either Category 2 or Category 3 did not include cladding failure, as all 

cladding failure configurations were evaluated under Category 1, so the containment analysis 

varied the fraction of crud that may spall off the cladding. The different parameter combinations 

from those listed in NUREG-1617 [29] were studied because insufficient data currently exist to 

establish values for the parameters important to the containment analysis of high-burnup fuel. 

Total releasable activity, effective A2 value, allowable radionuclide release rate, and allowable 

leakage rate were calculated for the representative PWR and BWR transportation packages 

referred to as GBC-32 and GBC-68, respectively. The W 17×17 OFA and GE14 assemblies 

with 40 and 65 GWd/MTU burnup values were used in the analysis. The decay times

considered were 5, 40, 100, and 300 years.
63

Relevant observations with respect to the containment analyses are as follows: 

• Overall, the results indicated that under NCT, increases in fuel rod failure rates can be 

off-set by longer decay times. For the GBC-68 cask and low-burnup fuel, an allowable 

leakage rate of ~2.4 × 10-5 cm3

/s was obtained for the 0.03, 0.4, 0.7, and 1.0 failed fuel 

fractions at the 5-, 40-, 100-, and 300-year decay times, respectively. For the GBC-32 

cask and low-burnup fuel, an allowable leakage rate of ~1 × 10-4 cm3

/s was obtained for 

the 0.03, 0.10, 0.15, and 0.25 failed fuel fractions at the 5-, 40-, 100-, and 300-year 

decay times, respectively.

• The allowable leakage rate decreases with increasing failed fuel rod fractions. The

decrease is relatively small (e.g., a 15% decrease because of an increase of failed fuel 

rod fraction from 0.01 to 0.03) if crud has a large contribution to the total releasable 

activity and is fairly large if the fuel fines dominate the total releasable activity (e.g., a 

decrease by a factor of 2.5 because of an increase of failed fuel rod fraction from 0.01 to 

0.03). 

• Crud is an important factor in the calculation of the allowable leakage rate for the time 

interval 5 to 40 years after fuel discharge from the reactor because of its relatively high 

contribution to the total releasable activity. Crud had a more pronounced effect on the 

BWR packages than on the PWR packages because of the larger 60Co activity per fuel 

rod surface area associated with BWR fuel. As a result, the BWR package allowable 

leakage rates are smaller (i.e., more restrictive) than the PWR package allowable 

leakage rates. For GBC-68 with 65 GWd/MTU burnup and 5-year decay time, an 

increase in the fraction of crud that spalls off the cladding by a factor of two causes a 

decrease in the allowable leakage rate by a factor of approximately two. However, the 

GBC-32 and GBC-68 total releasable activities as well as allowable leakage rates were 

similar for fuel decay times of ≥40 years because the crud contribution becomes

negligible at these decay times. 

• For each decay time and fraction of failed fuel rods, an increase in the fraction of 

gaseous species from 0.1 to 0.4 causes a relatively small decrease (up to ~30%) in the 

allowable release rate for high-burnup fuel. 

• An increase in the fraction of volatile source term by one order of magnitude causes a 

decrease in the allowable leakage rate for high-burnup fuel by a factor of approximately 

two.

• The allowable leakage rate showed the greatest sensitivity to changes in the mass 

fraction of fuel released as fuel fines because of cladding breach. Depending on the

decay time and crud contribution to the total releasable activity of high-burnup fuel, an 

increase in the fraction of fuel fines by a factor of 10 can cause a decrease in the 

allowable leakage rate by a factor of ~1.5 (BWR fuel with 5-year decay time) to 10 (PWR 

fuel with 40-year decay time).

• Similar to the NCT cases, the HAC allowable leakage rate has the largest sensitivity to 

the mass fraction of fuel released as fines. For the high-burnup fuel, an increase in the 

mass fraction of fuel released as fines by a factor of 10 can cause a decrease in the 
64

allowable leakage rate by a factor of ~5 to 10 for GBC-32 and by a factor of ~3.5 to 10 

for GBC-68 as the decay time increases from 5 to 300 years.

• Allowable radionuclide release rate and leakage rate for high-burnup fuel vary as a 

function of the pellet regions from which the radioactive material is released. Radioactive 

material released from the pellet peripheral region produced smaller allowable leakage 

rates than the radioactive material released from the non-rim region of the fuel pellet. 

The extent to which the pellet radial region, from which the releasable activity originates,

affects the allowable leakage rate depends on the contribution of crud to the total 

releasable activity. Small effects are observed for GBC-32 with 5-year decay time and 

small fraction of failed fuel rods (e.g., 0.01), as well as for GBC-68 with 5-year decay 

time and any fraction of failed fuel rods, as crud has a significant contribution to the total 

releasable activity for these cases. For the other cases, the allowable leakage rate 

based on releasable activity from the outer rim structure is approximately half the 

allowable leakage rate based on releasable activity from the non-rim region. The 

importance of the pellet region from which the radioactive material is released increases 

with increasing decay time and failed fuel fraction for NCT.

The currently recommended parameters in NUREG-1617 [29] for containment analysis should 

be evaluated in the future for their applicability to fuel in long-term dry storage and to highburnup fuel, and updated if data for these fuels become available.

Thermal

A summary of pertinent results for the fuel reconfiguration categories as they apply to thermal 

analysis considerations is provided in Table 29 and Table 30 for the vertical and horizontal 

orientations, respectively. The results reported are with respect to a thermal source term for 5.0 

wt % 235U initially enriched W 17×17 OFA fuel at a burnup of 65 GWd/MTU and a 40-year decay 

time unless identified otherwise. Of the fuel failure categories investigated, Category 1, where 

cladding failure occurs, has the largest impact on the component temperatures.

Table 29 Summary of maximum thermal variations for vertical cask

Category Description Cases compared

Peak cladding or

neutron absorber*

temp. variation

(∆°C)

- Decay time 20 to 60 years -221

- Decay time 40 to 60 years -45

- No insolation yes to no insolation -10

1

Failure of one assembly:

only gaseous release

0% to 100% failure -14

Failure of one assembly: 

gaseous release and particle bed

0% to 100% failure, 

0.626 packing fraction

-14

Failure of all assemblies:

only gaseous release

0% to 100% failure -71*

Failure of all assemblies: 

gaseous release and particle bed

0% to 100% failure, 

0.626 packing fraction

+127*

2 Rod pitch to diameter ratio 1.38 to 1.16 -51

3 Shifting all assemblies top to bottom -11
65

Table 30 Summary of maximum thermal variations for horizontal cask

Category Description Cases compared

Peak cladding or

neutron absorber*

temp. variation

(∆°C)

- Decay time 20 to 60 years -226

- Decay time 40 to 60 years -51

- No insolation yes to no -8

1

Failure of one assembly:

only gaseous release

0% to 100% failure +4

Failure of one assembly: 

gaseous release and particle bed

0% to 100% failure, 

0.313 packing fraction

+3

Failure of all assemblies:

only gaseous release

0% to 100% failure +42*

Failure of all assemblies: 

gaseous release and particle bed

0% to 100% failure, 

0.417 packing fraction

+31*

2 Rod pitch to diameter ratio 1.40 to 1.25 -12

3 Shifting all assemblies top to bottom +3

For Category 1, for the vertical package orientation where convective heat transport plays a 

role, the release of fission product gases from the rods into the canister reduced peak internal 

temperatures and caused an overall flattening of the axial temperature profiles. In contrast, for 

the horizontal orientations, where the package does not rely on convective heat transfer, the 

release of fission product gases caused a significant increase in internal temperatures. In all 

cases, releasing the fuel rod gaseous contents increases the canister internal pressure. Failure 

of all the assemblies and the formation of a debris bed resulted in a significant increase in the 

internal component peak temperatures.

For Category 2, packages in the vertical orientation that rely on internal convection for heat 

transport may be impacted by variations in assembly pitch more than horizontal or vertical 

package designs that do not rely on internal convection. In general, changes to the assembly 

pitch had a minor impact on the component temperatures.

For Category 3, where the axial alignment of the assembly was altered, the impact on 

component temperatures was only minor. Again, vertical packages relying on internal 

convection for heat transport may be impacted more than horizontal or vertical package designs 

that do not rely on internal convection.

From 20 to 100 years of decay time, the peak cladding temperature is predicted to decrease by 

272°C to 287°C depending on the package orientation. Other internal components also 

experience temperature reductions on this order. By accounting for insolation, the peak cladding 

temperature can change by 10°C to 8°C for the 40-year cooled cases. These variations provide 

a point of comparison as to the impact of fuel configuration changes. Depending on the package 

design, loading, and time frame for fuel failure, the thermal impacts of fuel failure may only be of 

secondary importance as compared to the decreased heat load of the fuel. If fuel failure occurs 

in a package that is already near its design basis (i.e., freshly loaded with decay heat loads near 

the design basis), the fuel failure could cause the package not to meet the thermal 

requirements. Conversely, fuel failure in packages loaded much below design basis heat loads

may have little consequence on meeting the thermal requirements.
66

An example of this is illustrated in Figure 33 for the vertical package and Figure 34 for the 

horizontal package where the peak neutron absorber temperature predicted by all fuel 

reconfigurations and cases is plotted. For the package loading analyzed, 40-year cooled fuel 

with 65 GWD/MTU burnup, no fuel reconfiguration resulted in temperatures outside the thermal 

limits for the neutron absorber. However, a similar package loaded with 25-year cooled fuel, 

may meet the normal operating condition thermal requirements but may not meet the 

requirements under fuel reconfiguration scenarios.

In all fuel configurations, the same amount of decay heat must be removed from the package. 

The fuel reconfigurations redistribute the energy source term within the canister and alter the 

heat transport paths from the fuel to the canister wall. The possibility for localized energy 

deposition, or localized isolation from effective heat transfer paths create the concern for 

localized hot spots and component failure. The current study highly discretized the canister solid 

material into sections less than 11.8 × 9.8 × 0.75 cm (basket cell wall). This discretization is 

complemented by modeling every assembly sub-channel. Thus, localized temperature 

variations of canister structures are resolved to approximately this scale or smaller in this study.

The analysis for Category 1 assumed a 30% release fraction of fission product gases 

(consistent with NUREG-1536 and NUREG-1617) [2, 29] regardless of whether only the 

cladding was breached or the entire assembly formed into a particle bed. The release fraction 

from rubblized fuel, or high-burnup fuel, may be different from that assumed.

Figure 33 Peak neutron absorber temperature variation from all fuel reconfigurations—

vertical cask
67

Figure 34 Peak neutron absorber temperature variation from all fuel reconfigurations—

horizontal cask

The fuel failure analysis was performed for only one set of nominal intact boundary conditions 

(i.e., 100°F still ambient air with insolation). The thermal impact under accident condition 

boundary conditions was not analyzed as part of this study. 

Finally, the change in the package temperature profile from fuel reconfigurations may affect the 

thermal stresses experienced by the various components. These would need to be resolved and 

analyzed for structural analysis of the package. 

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A-1

APPENDIX A. CRITICALITY EVALUATIONS FOR FUEL 

RECONFIGURATIONS

This appendix investigates and quantifies the impacts of changes to calculated keff values 

considering the fuel reconfiguration categories evaluated in this report. The term “fuel 

reconfiguration” refers to any change to the storage and transportation system nominal intact 

fuel assembly configuration used for the basis of cask certification. For all of the criticality 

evaluations, the basket structures and neutron poison materials were assumed to remain in the 

as-loaded condition. Additionally, all evaluations assume that the canister is flooded with water. 

A series of cases has been generated for pressurized water reactor (PWR) and boiling water 

reactor (BWR) storage cask/ transportation package models for several initial enrichment, 

burnup, and decay time combinations for intact fuel assemblies [30]. This set of cases is used in 

this report to form a basis of comparison for the various fuel reconfigurations considered. The 

representative PWR transportation package/cask system is referred to as the generic burnup 

credit-32 (GBC-32) package/cask, which has typically been used in the past as a GBC

benchmark model [4] loaded with Westinghouse (W) 17×17 optimized fuel assembly (OFA)

assemblies. The representative GBC-68 BWR cask system [5] was derived from the Holtec HISTORM 100 multipurpose canister (MPC) 68 cask loaded with representative 10×10 General 

Electric (GE) 14 BWR fuel assemblies. Note that use of a representative Holtec design in this 

work is not an endorsement of any design or vendor relative to any others; it was selected so 

that prior work performed for the Department of Energy Office of Nuclear Energy Fuel Cycle 

Technologies Used Fuel Disposition Campaign [30] could be leveraged where applicable. 

The Scale code system [31] was used to perform the keff and depletion calculations necessary 

for these analyses. All calculations used the 238-group neutron data library based on Evaluated 

Nuclear Data Files, Part B (ENDF/B)-VII.0, distributed with the Scale system. 

The KENO V.a and KENO-VI Monte Carlo codes were used for keff calculations within the 

appropriate Criticality Safety Analysis Sequence-5 (CSAS5) and CSAS6 sequences. Both 

codes use Monte Carlo transport to solve the keff eigenvalue problem. KENO-VI uses a 

generalized geometry process and is used for the fuel pellet array configuration and some 

increased fuel rod pitch configurations. KENO V.a has a more restrictive geometry package but 

is significantly faster because of the simpler geometry treatment. KENO V.a is used for the 

majority of configurations considered in this analysis. The KENO codes and CSAS sequences 

are further described and documented in Ref. [31]. 

For calculations involving irradiated fuel compositions, the isotopic compositions were 

generated with the Standardized Analysis of Reactivity for Burnup Credit using Scale 

(STARBUCS) sequence [31]. STARBUCS is a sequence to perform criticality calculations for 

spent fuel systems employing burnup credit. The STARBUCS sequence uses the Oak Ridge 

Isotope Generation and Depletion Code–Automatic Rapid Processing (ORIGEN-ARP)

methodology to generate depleted fuel compositions and uses the compositions in a KENO 

model to calculate keff. 

For the PWR cask model criticality evaluations, the initial enrichment and burnup combinations 

were selected from a representative loading curve such that the calculated keff was 0.94 after 5-

year decay time up to an initial enrichment of 5.0 wt % 235U. For the BWR cask model 

evaluations, only one initial enrichment was considered, 5.0 wt % 235U. For both cask models, to 

facilitate evaluations for high-burnup fuel, cases were also developed for a burnup of 
A-2

70 GWd/MTU. The range of parameters considered for the representative nominal intact 

configuration cases is provided in Table A.1. 

Table A.1 Range of parameters for criticality nominal intact configurations

Parameter

PWR system

Burnup (GWd/MTU) 0 25.5 44.25 70

Initial enrichments (wt % 235U) Nominal intact keff a/EALF

1.92 0.940/0.152 NA NA NA

3.5 NA 0.940/0.247 NA NA

5.0 NA NA 0.940/0.294 0.850/0.311

5.0 with 80-year decay time NA NA 0.900/0.297 0.789/0.323

5.0 with 300-year decay time NA NA 0.905/0.293 0.795/0.316

BWR system

Burnup (GWd/MTU) 0 35 70

Initial enrichment (wt % 235U) Nominal intact keff a/EALF channeled/EALF unchanneled

5.0 0.968/0.334/0.312 0.833/0.365/0.340 0.767/0.377/0.351

5.0 with 80-year decay time NA 0.824/0.366/0.341 0.753/0.382/0.355

5.0 with 300-year decay time NA 0.825/0.364/0.340 0.754/0.379/0.353

Notes: NA = not applicable; EALF = energy of average lethargy of a neutron causing fission (eV). aSigma value for keff calculations was 0.00010.

A.1 CRITICALITY ANALYSIS TO ASSESS IMPACTS OF

CATEGORY 1—CLADDING FAILURE

The design basis geometric orientation of the fuel, materials, and neutron energy spectrum are 

fundamental parameters that must be defined when developing the criticality safety licensing 

basis. Commercial light water reactor (LWR) fuel assemblies are designed to be undermoderated. The introduction of additional moderator into the assembly lattice can result in a net 

increase in reactivity depending on location. Fuel rods that suffer cladding failure (i.e., change in 

design basis geometry) can result in sections of fuel within the lattice being replaced with water, 

resulting in additional moderator being introduced to the lattice (i.e., change in design basis 

neutron energy spectrum) and the potential for reactivity increase. 

To have an effect on criticality safety performance, the spent nuclear fuel (SNF) rods must be 

grossly breached such that the geometry and neutron energy spectrum present within the 

canister are substantially altered from the design basis configuration. To investigate the effects 

of cladding failure on criticality control performance objectives, cases were developed for two 

configuration scenarios: (1) breached spent fuel rods where assemblies are partially intact; and 

(2) damaged fuel where the fuel assembly is represented as a rubble mixture allowing free 

movement of fuel particles and pellets within a basket cell. Case subclass descriptions are 

provided in Table A.2 and described in more detail below. 
A-3

Table A.2 Cases developed to address Category 1—Cladding Failure

Identifier

Configuration 

scenario Subclass description Parameter varied

S1(a)C1 Breached spent 

fuel rods

Single fuel rod failure Individual lattice positions 

modeled as missing rod

S1(a)C2 Breached spent 

fuel rods

Multiple fuel rod failure Multiple missing rod 

combinations until an upper limit 

is identified

S1(b)C3 Damaged fuel Gross assembly failure consisting 

of homogenized debris

The debris bed size, and thus 

internal moderation, was varied 

to find the largest ∆keff

S1(b)C4 Damaged fuel Gross assembly failure 

representing debris as discrete 

lumps

The pellet spacing, and thus the 

debris bed size, was varied to 

find the largest ∆keff

A.1.1 Breached Spent Fuel Rods

Two sets of cases were developed for analyses of breached spent fuel rods. The first is an 

evaluation of lattice position reactivity worth on an individual basis to be used in reducing the 

combinatorial phase space for multiple rod removal, and the second considers groups of 

multiple rod removal from the lattice. For these two sets of cases, the displaced fuel was omitted 

from the model, assuming that it would be sufficiently dispersed from the lattice in a thin, 

undermoderated heap, resulting in a less reactive system if incorporated into the model. An 

additional subset of cases was included with the multiple rod removal cases to evaluate the 

combined effects of rod removal and the distribution of the removed fuel at regions outside the 

neutron absorber envelope, varying the moderator-to-fuel ratio for the displaced fuel.

S1(a)C1: Single fuel rod failure—A set of cases was developed to evaluate the associated 

reactivity impact of missing fuel rod segments from each unique lattice position within an 

assembly. The results of this sensitivity analysis can be used in a combinatorial evaluation to 

identify a representative bounding lattice configuration with missing fuel rods. Each of the 39 

eighth-assembly symmetric rods in the W 17×17 PWR assembly was removed individually to 

determine its relative worth. Full rod removal is modeled to bound the effects of partial rod 

removal. The same process was used to evaluate the GE 10×10 BWR fuel assembly; however,

half-assembly symmetry was used and resulted in 51 unique fuel lattice locations for evaluation. 

Cross-sectional lattice arrays showing the symmetry positions for evaluation are presented in 

Figure A.1 and Figure A.2 for the PWR and BWR assemblies, respectively. 

Results showing the change in keff as a function of position for different initial enrichments, 

burnups, and decay times for the W 17×17 PWR assembly type in a representative GBC-32 

package are provided in Table A.3 through Table A.9. Similar results as a function of position for 

different initial enrichments, burnups, and decay times for the GE 10×10 BWR assembly type in 

the GBC-68 cask are provided in Table A.10 through Table A.16
A-4

A B C D E F G H I J K L M N O P Q

Figure A.1 Sketch of symmetry, row, and column labels for W 17×17 fuel assembly

Figure A.2 Sketch of symmetry, row, and column labels for GE 10×10 fuel assembly

Half-assembly symmetric portion of 

GE 10 × 10 fuel assembly

A B C D E F G H I J

1

2

3

4

5

6

7

8

9

10

1/8th symmetric portion of 

W 17 × 17 fuel assembly

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

Fuel rod

Guide tube

Water 

moderator

Water 

moderator

Fuel rod

Water rod
A-5

Table A.3 Change in keff (%∆keff) caused by individual removal of each fuel rod in

W 17×17 fuel in GBC-32, Fresh fuel, 1.92 w/o 235U (one-sigma uncertainty of 

all values is 0.014 %∆keff)

A B C D E F G H I

1 -0.14 -0.13 -0.12 -0.10 -0.10 -0.12 -0.14 -0.11 -0.14

2 -0.09 -0.08 -0.06 -0.07 -0.05 -0.07 -0.07 -0.08

3 -0.03 -0.03 -0.04 GT -0.04 -0.02 GT

4 GT -0.02 0.00 -0.01 0.00 -0.01

5 0.01 0.00 0.02 0.01 -0.01

6 GT 0.01 0.01 GT

7 0.02 0.02 0.01

8 0.04 0.02

9 IT

Notes: GT = guide tube location; IT = instrument tube location.

Table A.4 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

W 17×17 fuel in GBC-32, 44.25 GWd/MTU burnup, 5-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I

1 -0.15 -0.11 -0.11 -0.09 -0.12 -0.11 -0.08 -0.09 -0.09

2 -0.05 0.02 0.00 0.02 -0.00 0.02 -0.02 -0.00

3 0.04 0.03 0.04 GT 0.05 0.05 GT

4 GT 0.04 0.07 0.08 0.07 0.06

5 0.09 0.07 0.07 0.10 0.07

6 GT 0.08 0.08 GT

7 0.08 0.08 0.09

8 0.09 0.08

9 IT

Notes: GT = guide tube location; IT = instrument tube location.

Table A.5 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

W 17×17 fuel in GBC-32, 44.25 GWd/MTU burnup, 80-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I

1 -0.14 -0.09 -0.07 -0.08 -0.08 -0.09% -0.08 -0.06 -0.09

2 -0.01 0.00 0.03 0.03 -0.01% 0.01 0.01 0.00

3 0.06 0.04 0.05 GT 0.06 0.04 GT

4 GT 0.05 0.06 0.08 0.07 0.06

5 0.07 0.08 0.08 0.09 0.08

6 GT 0.07 0.08 GT

7 0.09 0.10 0.08

8 0.09 0.08

9 IT

Notes: GT = guide tube location; IT = instrument tube location.
A-6

Table A.6 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

W 17×17 fuel in GBC-32, 44.25 GWd/MTU burnup, 300-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I

1 -0.17 -0.12 -0.10 -0.09 -0.10 -0.09 -0.09 -0.08 -0.10

2 -0.01 0.00 0.00 0.02 0.01 0.01 0.01 -0.02

3 0.01 0.04 0.02 GT 0.04 0.05 GT

4 GT 0.06 0.05 0.06 0.06 0.07

5 0.05 0.06 0.09 0.09 0.09

6 GT 0.08 0.08 GT

7 0.10 0.08 0.07

8 0.08 0.08

9 IT

Notes: GT = guide tube location; IT = instrument tube location.

Table A.7 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

W 17×17 fuel in GBC-32, 70 GWd/MTU burnup, 5-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I

1 -0.14 -0.12 -0.11 -0.06 -0.11 -0.11 -0.09 -0.08 -0.10

2 -0.05 -0.01 0.02 -0.01 0.00 0.00 0.00 -0.02

3 0.00 0.03 0.03 GT 0.05 0.03 GT

4 GT 0.05 0.06 0.08 0.06 0.05

5 0.05 0.05 0.07 0.09 0.06

6 GT 0.07 0.07 GT

7 0.07 0.08 0.06

8 0.08 0.08

9 IT

Notes: GT = guide tube location; IT = instrument tube location.

Table A.8 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

W 17×17 fuel in GBC-32, 70 GWd/MTU burnup, 80-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I

1 -0.11 -0.10 -0.08 -0.05 -0.06 -0.07 -0.09 -0.06 -0.08

2 -0.01 0.01 0.02 0.02 0.01 -0.01 0.01 0.02

3 0.02 0.04 0.06 GT 0.05 0.06 GT

4 GT 0.05 0.06 0.07 0.07 0.08

5 0.05 0.08 0.08 0.08 0.08

6 GT 0.06 0.08 GT

7 0.10 0.09 0.09

8 0.08 0.09

9 IT

Notes: GT = guide tube location; IT = instrument tube location.
A-7

Table A.9 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

W 17×17 fuel in GBC-32, 70 GWd/MTU burnup, 300-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I

1 -0.12 -0.09 -0.08 -0.08 -0.08 -0.07 -0.08 -0.09 -0.07

2 -0.02 -0.01 0.01 0.03 0.02 0.02 0.02 0.01

3 0.05 0.04 0.02 GT 0.04 0.06 GT

4 GT 0.05 0.06 0.07 0.07 0.06

5 0.08 0.08 0.10 0.09 0.07

6 GT 0.07 0.09 GT

7 0.10 0.08 0.09

8 0.09 0.10

9 IT

Notes: GT = guide tube location; IT = instrument tube location.

Table A.10 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

GE 10×10 fuel in GBC-68, fresh 5 wt % 235U fuel (one-sigma uncertainty of 

all values is 0.014 %∆keff)

A B C D E F G H I J

1 -0.51 -0.21 -0.14 -0.11 -0.10 -0.11 -0.12 -0.15 -0.21 -0.49

2 0.06 0.15 0.18 0.20 0.18 0.17 0.15 0.05 -0.24

3 0.24 0.28 0.25 0.18 0.17 0.21 0.14 -0.16

4 0.26 0.20 WT WT 0.19 0.16 -0.12

5 0.11 WT WT -.19 0.19 -0.12

6 0.11 0.20 0.26 0.19 -0.10

7 0.27 0.29 0.18 -0.13

8 0.23 0.16 -0.14

9 0.08 -0.23

10 -0.50

Note: WT = one quarter of a water tube.

Table A.11 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

GE 10×10 fuel in GBC-68, 35 GWd/MTU burnup, 5-year decay time 

(uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I J

1 -0.43 -0.21 -0.12 -0.07 -0.09 -0.10 -0.11 -0.11 -0.20 -0.41

2 0.06 0.12 0.17 0.19 0.16 0.14 0.12 0.05 -0.20

3 0.23 0.24 0.23 0.18 0.17 0.19 0.12 -0.13

4 0.26 0.17 WT WT 0.16 0.14 -0.11

5 0.09 WT WT 0.18 0.16 -0.09

6 0.08 0.19 0.24 0.18 -0.08

7 0.26 0.25 0.15 -0.09

8 0.22 0.16 -0.12

9 0.06 -0.18

10 -0.44

Note: WT = one quarter of a water tube.
A-8

Table A.12 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

GE 10×10 fuel in GBC-68, 35 GWd/MTU burnup, 80-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I J

1 -0.39 -0.17 -0.11 -0.10 -0.07 -0.06 -0.09 -0.12 -0.19 -0.42

2 0.06 0.17 0.16 0.17 0.16 0.15 0.16 0.06 -0.19

3 0.24 0.26 0.24 0.19 0.15 0.21 0.14 -0.11

4 0.27 0.21 WT WT 0.19 0.15 -0.10

5 0.13 WT WT 0.17 0.17 -0.07

6 0.09 0.20 0.23 0.19 -0.08

7 0.27 0.26 0.17 -0.07

8 0.23 0.13 -0.11

9 0.07 -0.18

10 -0.40

Note: WT = one quarter of a water tube.

Table A.13 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

GE 10×10 fuel in GBC-68, 35 GWd/MTU burnup, 300-year decay time (onesigma uncertainty of all values is 0.014% ∆keff)

A B C D E F G H I J

1 -0.42 -0.17 -0.09 -0.08 -0.05 -0.07 -0.07 -0.10 -0.17 -0.41

2 0.08 0.15 0.15 0.20 0.17 0.16 0.15 0.07 -0.17

3 0.23 0.25 0.25 0.19 0.20 0.21 0.15 -0.13

4 0.28 0.19 WT WT 0.18 0.17 -0.09

5 0.09 WT WT 0.19 0.19 -0.06

6 0.13 0.20 0.26 0.18 -0.08

7 0.28 0.26 0.19 -0.09

8 0.25 0.15 -0.10

9 0.08 -0.17

10 -0.44

Note: WT = one quarter of a water tube.

Table A.14 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

GE 10×10 fuel in GBC-68, 70 GWd/MTU burnup, 5-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I J

1 -0.40 -0.17 -0.11 -0.10 -0.07 -0.07 -0.09 -0.11 -0.18 -0.40

2 0.02 0.12 0.15 0.16 0.13 0.14 0.10 0.04 -0.19

3 0.21 0.26 0.21 0.14 0.14 0.20 0.13 -0.13

4 0.25 0.17 WT WT 0.15 0.14 -0.10

5 0.09 WT WT 0.18 0.15 -0.09

6 0.08 0.19 0.22 0.17 -0.07

7 0.25 0.22 0.14 -0.09

8 0.22 0.12 -0.12

9 0.06 -0.15

10 -0.39

Note: WT = one quarter of a water tube.
A-9

Table A.15 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

GE 10×10 fuel in GBC-68, 70 GWd/MTU burnup, 80-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I J

1 -0.37 -0.18 -0.11 -0.09 -0.07 -0.06 -0.08 -0.10 -0.17 -0.37

2 0.04 0.12 0.17 0.16 0.17 0.13 0.12 0.05 -0.15

3 0.21 0.24 0.22 0.17 0.14 0.16 0.11 -0.11

4 0.24 0.19 WT WT 0.16 0.13 -0.09

5 0.10 WT WT 0.18 0.14 -0.08

6 0.12 0.18 0.22 0.16 -0.08

7 0.25 0.24 0.15 -0.08

8 0.21 0.12 -0.08

9 0.05 -0.16

10 -0.39

Note: WT = one quarter of a water tube.

Table A.16 Change in keff (%∆keff) caused by individual removal of each fuel rod in 

GE 10×10 fuel in GBC-68, 70 GWd/MTU burnup, 300-year decay time (onesigma uncertainty of all values is 0.014 %∆keff)

A B C D E F G H I J

1 -0.36 -0.16 -0.11 -0.07 -0.07 -0.06 -0.08 -0.09 -0.16 -0.38

2 0.07 0.11 0.18 0.16 0.14 0.15 0.12 0.04 -0.17

3 0.23 0.24 0.23 0.16 0.16 0.21 0.13 -0.11

4 0.25 0.18 WT WT 0.17 0.14 -0.08

5 0.11 WT WT 0.18 0.17 -0.08

6 0.10 0.20 0.23 0.17 -0.05

7 0.26 0.23 0.18 -0.07

8 0.22 0.13 -0.10

9 0.07 -0.16

10 -0.38

Note: WT = one quarter of a water tube.

The maximum keff increase observed for the PWR assembly was 0.10% ∆keff and is associated 

with rod H5 in the 5.0 wt % 235U initial enrichment, 44.25 GWd/MTU burnup with 5-year decay 

case. In general, the other cases at different decay times and burnup showed that the lattice 

position reactivity worths remained nearly constant for different decay times and burnups. It 

should be noted that several rods across many of the cases have a reactivity worth that is 

statistically equivalent to the limiting case. Many other rods have a reactivity increase that is 

less than one standard deviation from 0 and thus have no statistically significant impact on 

system keff.

For the BWR fuel assembly, the maximum keff increase was 0.29% ∆keff and is associated with 

rod H7 for fresh 5 wt % 235U channeled fuel. Cases have been evaluated investigating how 

dechanneled BWR fuel assemblies respond to lattice moderator changes and have shown that 

the keff increase for channeled fuel assemblies is greater than for unchanneled assemblies [30]. 

This is likely an artifact of the difference in neutron energy spectrum that occurs from the 

channel displacing moderator between adjacent fuel regions. The different burnup and decay 

time sensitivity studies for the BWR fuel also indicate no significant changes in lattice position 

worth as a function of burnup or decay time.
A-10

S1(a)C2: Multiple fuel rod failure—Cases representing multiple fuel rod failures were 

developed to identify an upper bound on the potential reactivity increase because of the effects 

of increased moderation resulting from fuel rod segment relocation. A combinatorial evaluation 

was performed for different groups of fuel rod failure represented. The high worth lattice 

locations identified in case set S1(a)C1 were used to reduce the sample space of potential 

combinations to consider from the original lattice orientation. This allowed rod combinations to 

be selected based on lattice position worth. Multiple rods were removed in symmetric patterns 

covering different combinations of rods removed until a reactivity peak was identified. For each 

group of rods removed, various combinations of removal patterns were evaluated. 

For the PWR assembly cases, groups of 2, 4, 8, 16, 24, 28, 32, 36, 40, 44, and 48 rods were 

considered for removal patterns. Results showing the maximum keff change that was observed 

from the patterns evaluated for the different enrichment and burnup combinations are provided 

in Table A.17. The maximum ∆keff value (i.e., a keff increase of 1.87 %∆keff) for the GBC-32 cask 

representation occurs for the 5 wt % 235U initial enriched fuel with a burnup of 44.25 GWd/MTU

and 300-year decay time when 44 rods have been removed from the lattice configuration. 

Results for a range of initial enrichment and burnup compositions with 5-year decay time as a 

function of rods removed are illustrated in Figure A.3, and a cross-sectional view of the 

corresponding limiting lattice configuration is illustrated in Figure A.4. The higher burnup and

longer decay time results indicate that there is no significant increase in reactivity change 

caused by the different nuclide compositions.

Table A.17 Multiple rod removal results for W 17×17 OFA in GBC-32

Enrichment

(wt % 235U)

Burnup

(GWd/MTU)

Decay time 

(years)

Number of 

rods removed

Maximum increase in keff

(%∆keff)

1.92 0 0 4 0.04

3.5 25.5 5 36 1.07

5 44.25 5 44 1.86

5 44.25 80 44 1.86

5 44.25 300 44 1.87

5 70 5 44 1.69

5 70 80 44 1.62

5 70 300 44 1.62
A-11

Figure A.3 Increase in keff in GBC-32 cask as a function of number of rods removed

Figure A.4 Limiting multiple rod removal lattice (44 rods removed)

An additional subset of cases was developed to investigate the combined effects of rod removal 

and spent fuel displacement in regions outside the neutron absorber envelope. Only the PWR 

fuel was represented with the 5.0 wt % 235U initial enrichment at 44.25 GWd/MTU burnup and 

5-year decay. For these cases, the available void region was modeled with fuel at different 

volume fractions simulating various sized particles with different bulk densities. The void 

-0.5

0.0

0.5

1.0

1.5

2.0

0 5 10 15 20 25 30 35 40 45 50

Increase in 

keff (%

Δ

keff)

Number of rods removed

1.92 wt% U-235, Fresh 3.5 wt% U-235, 25.5 GWd/MTU

5 wt% U-235, 44.25 GWd/MTU 5 wt% U-235, 70 GWd/MTU
A-12

displacement of hardware components was omitted from these models. Results are provided in

Table A.18 and show that combining the rod removal with displaced fuel distributed at ratios 

approaching optimum moderation conditions outside the neutron absorber plate envelope yields 

higher changes in keff than the displaced fuel configurations alone, which were 1.82, 1.52, and 

0.57 %∆keff for the 40-, 28-, and 8-rod removal cases, respectively.

Table A.18 Combination rod removal results with heterogeneous distribution of 

displaced fuel outside neutron absorber plate envelope for GBC-32

# Failed rods Volume fraction of 

displaced fuel

Length of void region 

filled (cm)

Change in keff

(%∆keff)

40 0.524 30 3.89

28 0.341 30 4.91

8 0.146 20 3.55

For the BWR assembly cases, groups of 2, 4, 8, 12, 14, 16, 18, and 20 rods were considered for 

removal patterns. Results in Table A.19 show the maximum keff change observed from the 

patterns evaluated for the different enrichment and burnup combinations presented as a 

function of rods removed. The maximum ∆keff value (i.e., a keff increase of 2.42 %∆keff) for the 

GBC-68 representation occurs for the 5 wt % 235U initial enriched fuel with a burnup of 35 

GWd/MTU and 300-year decay time when 18 rods have been removed from the lattice 

configuration with the channel present. Results for the 35 and 70 GWd/MTU burnup with 5-year

decay time compositions and fresh 5 w/o 235U fuel as a function of rods removed are illustrated 

in Figure A.5, and a cross-sectional view of the corresponding limiting lattice configuration is 

illustrated in Figure A.6. The results for higher burnups and longer decay times indicate no 

significant reactivity increases resulting from the use of different nuclide compositions.

Table A.19 Multiple rod removal results for GE 10×10 fuel in GBC-68

Burnup

(GWd/MTU)

Decay time

(years)

Channel 

present

Number of 

rods removed

Maximum increase in keff

(%∆keff)

0 0 Yes 16 2.24

35 5 Yes 18 2.40

35 80 Yes 18 2.40

35 300 Yes 18 2.42

70 5 Yes 18 2.30

70 80 Yes 20 2.31

70 300 Yes 18 2.32

0 0 No 16 2.11

35 5 No 18 2.30

35 80 No 18 2.31

35 300 No 18 2.29

70 5 No 18 2.20

70 80 No 20 2.18

70 300 No 18 2.20
A-13

Figure A.5 Increase in keff in GBC-68 as a function of number of rods removed

Figure A.6 Limiting multiple rod removal lattice (18 rods removed)

A.1.2 Damaged Fuel

Two configurations for the physical form of the failed fuel are considered in these analyses: the 

first is a homogeneous mixture of fuel, cladding materials, and water; and the second is a 

dodecahedral array of fuel pellets suspended in water. The selected configurations are 

expected to encompass the range of potential reactivity changes for different rubble 

configurations. Modeling an ordered array of pellets provides an upper bound of the reactivity of 

the fuel rubble because low enriched fuel is more reactive lumped when compared with a 

homogeneous mixture because of resonance self-shielding effects.

0.0

0.5

1.0

1.5

2.0

2.5

0 5 10 15 20

Increase in 

keff (%

Δ

keff)

Number of failed rods

5 wt% U-235, Fresh 5 wt% U-235, 35 GWd/MTU 5 wt% U-235, 70 GWd/MTU
A-14

S1(b)C3: Homogenized rubble—Extensive cladding failure could lead to a relatively large 

debris bed composed of fuel pellet fragments of varying sizes as well as assembly hardware 

pieces distributed throughout the basket cavity region. The height of the debris bed is varied, 

changing the volume of the debris and the moderation ratio. The modeled heights vary from the 

entire internal height of the canister to the volume of the fuel and cladding material in the 

assembly. Cases were developed to evaluate a range of debris concentrations and distributions 

within the basket cavity. The total fuel mass is conserved in each case. The fuel compositions 

are represented as stratified to maintain the burnup profile established in the assembly during 

power operation. A subset of cases was also included to evaluate fractional debris bed heights

constrained within the basket absorber panel region. Fuel debris may have limited axial mobility 

because of the combined effects of hardware (e.g., end fittings and spacers), fuel debris 

characteristics, and available basket clearances restricting large volumes of fuel from migrating 

beyond the original axial region of the basket. Additionally, this set of cases provides insight into 

the possible response of the newer metal matrix composite baskets that have neutron absorber 

extending the length of the basket. 

S1(b)C4: Loose fuel pellets—In the event of extensive cladding failure, representing the debris 

as discrete lumps may result in a higher keff value for low enriched fuel than a homogeneous 

debris because of resonance self-shielding in 238U. To evaluate this condition, a range of 

uniform, dodecahedral pellet arrays was considered. The pitch of the pellets was varied to 

change the debris bed heights and moderation ratios. The pellet arrays are represented as

stratified to maintain the burnup profile established in the assembly during power operation. 

The same axial compositions are used in the fuel as in the nominal intact configuration case. 

For PWR fuel, the pellet pitch varied from approximately 0.742 cm to 0.772 cm corresponding to 

the height of the nominal intact fuel assembly to the internal height of the canister cavity. For 

BWR fuel, the pellet pitch varied from approximately 0.771 cm to 0.879 cm, also ranging from 

the nominal intact assembly height to the internal height of the canister cavity. A subset of cases 

was also included to evaluate fractional debris bed heights constrained within the basket 

neutron absorber panel envelope. Fuel debris may have limited axial mobility because of the 

combined effects of hardware (e.g., end fittings and spacers), fuel debris characteristics, and 

available basket clearances restricting large volumes of fuel from migrating beyond the original 

axial region of the basket. Additionally, this set of cases provides insight into the possible 

response of the newer metal matrix composite baskets that have neutron absorber extending 

the length of the basket.

The two damaged fuel configurations described above were investigated for PWR and BWR

systems. Axial representations of a basket cell are shown in Figure A.7 and Figure A.8 for the 

homogeneous rubble and ordered pellet array cases, respectively. The results for both 

configurations with fresh fuel and 5-year decay time conditions are provided in

Table A.20 for the maximum increase cases, showing that the keff increase can be substantial 

(>20 %∆keff) depending on how the system is modeled. For the cases evaluated, the 

44.25 GWd/MTU burnup case with 5 wt % 235U initial enrichment had the largest change for the 

GBC-32 cask, and the 35 GWd/MTU with 5 wt % 235U initial enrichment had the largest change 

for the GBC-68 cask. In both casks, the limiting case occurs when the fissile material is 

distributed uniformly throughout the basket cell region, extending from the base plate to the lid 

with the ordered pellet array case being more limiting than the homogeneous rubble case. A 

floating pellet array of all the fuel is not physically likely, but represents an upper limit for the 

maximum reactivity increase for damaged fuel. This set of cases demonstrates that 

homogenous fuel rubble configurations may produce non-conservative results. 
A-15

The results for fuel compositions with higher burnups and longer decay times for PWR fuel and 

BWR fuel are provided in Table A.21 and Table A.22, respectively. The results indicate that the 

reactivity impact of gross assembly failure increases slightly at higher burnups and increases 

with longer decay time. The additional impacts are small, less than 1.5 %∆keff, and are less than 

the margin (i.e., reduction in keff) provided with the additional burnup and decay time. 

GBC-32 GBC-68

Figure A.7 Limiting homogeneous rubble configuration for GBC-32 and GBC-68

18 homogeneous 

fuel, cladding, and 

water mixtures

25 homogeneous 

fuel, cladding, and 

water mixtures
A-16

GBC-32 GBC-68

Figure A.8 Limiting ordered pellet array configuration for GBC-32 and GBC-68

Table A.20 Increase in keff because of damaged fuel (fissile material located outside 

neutron absorber panel envelope)

Cask Enrichment

(wt % 235U)

Burnup

(GWd/MTU)

Decay time

(years)

Maximum increase in keff

(%∆keff)

GBC-32

Limiting ordered pellet array

1.92 0 0 11.09

3.5 25.5 5 20.20

5.0 44.25 80 22.21

Homogeneous rubble

1.92 0 0 6.66

3.5 25.5 5 13.95

5.0 44.25 300 15.34

GBC-68

Limiting ordered pellet arraya

5.0 0 0 28.12

5.0 35 80 34.88

5.0 70 300 35.63

Homogeneous rubbleb

5.0 0 0 21.68/22.90

5.0 35 80 29.12/29.87

5.0 70 300 29.81/30.40 aThe BWR pellet array was only represented without the fuel assembly channel.

bBWR results are presented with channel / without channel.
A-17

Table A.21 Change in keff for damaged fuel in GBC-32

Burnup

(GWd/MTU)

Decay time

(years)

Change in keff

(%∆keff)

Ordered pellet array

44.25 5 21.37

44.25 80 22.21

44.25 300 22.21

70 5 21.43

70 80 21.63

70 300 21.77

Homogeneous rubble

44.25 5 14.30

44.25 80 15.29

44.25 300 15.34

70 5 14.20

70 80 14.77

70 300 14.90

Table A.22 Change in keff for damaged fuel in GBC-68

Burnup

(GWd/MTU)

Decay time

(years)

Change in keff

(%∆keff)

Ordered pellet array, channel removed

35 5 34.40

35 80 34.88

35 300 34.87

70 5 35.22

70 80 35.57

70 300 35.63

Homogeneous rubble, channel included

35 5 28.58

35 80 29.12

35 300 29.13

70 5 29.31

70 80 29.74

70 300 29.81

Homogeneous rubble, channel removed

35 5 29.36

35 80 29.87

35 300 29.83

70 5 29.93

70 80 30.33

70 300 30.40

The resulting change in keff is significantly reduced if the fuel debris is maintained within the 

neutron absorber panel envelope. The maximum reactivity impact for the ordered array of 

pellets is reduced from 22.21 %∆keff to 4.40 %∆keff for 5 wt % 235U initial enriched fuel with 

44.25 GWd/MTU burnup and 80-year decay time in the GBC-32 model. The reduction in the 

MPC-68 model is from 35.63% ∆keff to 10.81% ∆keff for 5 wt % 235U initial enriched fuel with 70 

GWd/MTU burnup and 300-year decay time.
A-18

Mixed size, irregular shaped particle packing can vary. To observe the effects of different bulk 

density changes on system reactivity, a range of homogeneous rubble cases within the neutron 

absorber panel envelope was evaluated and is provided in Table A.23. The cases resulting in 

the maximum reactivity increase for GBC-32 are 5.0 wt % 235U initial enrichment, 44.25 

GWd/MTU burnup with 5-year decay time isotopic compositions; and for GBC-68 are for 

5.0 wt % 235U enriched fresh fuel. Overall, these results show a decrease in reactivity with 

increased bulk density (i.e., tighter fuel particle packing).

Table A.23 Change in keff because of damaged fuel (homogeneous rubble debris within 

neutron absorber panel envelope)

Fraction of nominal intact

assembly height

Change in keff (%∆keff)

GBC-32

GBC-68

Channeled Unchanneled

1.0 -4.64 7.40 9.49

0.9 -7.05 6.65 9.12

0.8 -10.16 5.06 8.10

0.7 -14.36 2.30 6.16

0.6 -20.16 -2.57 2.66

0.5 -28.34 -11.07 -3.62

0.4 -39.10 -25.64 -15.10

0.36 (Fully compressed rubble) -45.50 -34.23 -31.44a

aUnchanneled fraction of nominal intact assembly height is 0.32

A.2 CRITICALITY ANALYSIS TO ASSESS IMPACTS OF

CATEGORY 2 – ROD/ASSEMBLY DEFORMATION

Fuel assembly distortion, twist, and rod bow are known phenomena present in current spent 

fuel. Each of these effects can alter the geometry and neutron energy spectrum and hence can 

affect keff. This configuration category investigates the impact of pin pitch changes that could 

result as potential end states associated with side and end impact events which can be 

exacerbated by grid failure. Side impact events are represented by configurations where the 

assembly lattice contracts and the fuel assembly is oriented against the basket side. Typical 

design basis criticality evaluations represent the fuel assembly as centered across the basket 

cell. The effects of end impact look at both uniform and non-uniform pin pitch expansion. The 

non-uniform pin-pitch variations expand on the uniform pin-pitch expansion cases to determine 

if a potential configuration from non-uniform pin-pitch expansion can result in a higher reactivity 

increase than the uniform pin-pitch expansion cases. Additionally, fuel assembly axial 

compositions vary with burnup and irradiation history. This set of cases also evaluates the 

impacts of multiple axial effects that could potentially be present because of lattice pin-pitch 

changes and varying nodal zone lengths. Case subclass descriptions are provided in Table 

A.24 and described in more detail below.
A-19

Table A.24 Criticality cases developed to address Category 2—rod/assembly 

deformation

Case set 

identifier

Configuration 

scenario Subclass description Parameter varied

S2(a)C1 Configurations 

associated with 

side drop

Pin pitch contraction PWR pin pitch contacted to 

95.7% of nominal, cross sectional 

position of contracted assembly 

within basket cell

BWR pitch contracted to 97.4% 

of nominal, cross sectional 

position of contracted assembly 

within basket cell, channel 

present

S2(b)C1 Configurations 

associated with 

end drop

Uniform pin pitch variation PWR pin pitch of all rods in an 

assembly expanded up to 

104.6% of nominal

BWR pin pitch of all rods in an 

assembly expanded up to 

102.3% and 121.8% of nominal 

for channeled and unchanneled 

models, respectively

S2(b)C2 Configurations 

associated with 

end drop

Radial non-uniform pin pitch 

variation

PWR pin pitch of different radial 

regions of the assembly 

expanded up to 116.9% of 

nominal

BWR pin pitch of different radial 

regions of the assembly 

expanded up to 108.9% and 

134% of nominal for channeled 

and unchanneled fuel, 

respectively

S2(b)C3 Configurations 

associated with 

end drop

Axial non-uniform pin pitch 

variation (birdcaging and 

bottlenecking)

PWR pin pitch of different axial 

regions of the assembly 

expanded up to 104.6% in 

expanded regions and contracted 

to 95.7% of nominal in contracted

regions

BWR pin pitch of different axial 

regions of the unchanneled 

assembly expanded up to 

121.8% in expanded regions and 

contracted to 97.4% of nominal in 

contracted regions

Configuration S2(a)C1: Pin pitch contraction—Side drop events are expected to include

configurations where the pin pitch is reduced. The contracted pitch was determined by the size 

of non-fuel guide, instrument, or water tubes in the various fuel assembly designs that are 

assumed to maintain nominal dimensions. This resulted in pitch contraction percentages of 

95.7% and 97.4% of nominal for the PWR and BWR assemblies, respectively. Because LWR 

assemblies are under-moderated, pin pitch contraction will result in decreases in keff. To 

illustrate the amount of decrease, two subcases for PWR and four for BWR fuel were evaluated. 

The subcases are based on the fuel assembly being centered within the basket cell or resting 

against the side of a basket cell, and being channeled versus unchanneled for BWR fuel. 
A-20

Illustrations of the models are show in Figure A.9 and Figure A.10 for PWR fuel and BWR fuel, 

respectively. Results for the selected cases are shown in Table A.25, where the PWR fuel 

corresponds to 5.0 wt % 235U initial enrichment with burnup of 44.25 GWd/MTU and 5-year 

decay time, and the BWR fuel corresponds to 5.0 wt % 235U enrichment fresh fuel. 

(a) (b)

Figure A.9 Pin pitch contraction cases in GBC-32
A-21

(a) (b)

(c) (d)

Figure A.10 Pin pitch contraction cases for GBC-68

Table A.25 Criticality results for pin pitch contraction cases

PWR fuel

Subcase ID Description Change in keff (%∆keff)

S2(a)C1 Pitch contraction (95.7% of nominal) with assembly 

centered within cell (Figure A.9 [a]) -3.63

S2(a)C2 Pitch contraction (95.7% of nominal) with assembly 

resting against cell (Figure A.9 [b]) -4.05

BWR fuel

S2(a)C4 Pitch contraction (97.4% of nominal) with assembly 

centered within cell (Figure A.10 [a]) -2.44

S2(a)C4 Pitch contraction (97.4% of nominal) with assembly 

resting against cell (Figure A.10 [b]) -3.48

S2(a)C4 Pitch contraction (97.4% of nominal) with unchanneled 

assembly centered within cell (Figure A.10 [c]) -2.48

S2(a)C4 Pitch contraction (97.4% of nominal) with unchanneled 

assembly resting against cell (Figure A.10 [d]) -4.42

S2(b)C1: Uniform pin pitch variation—A set of cases was developed to evaluate the 

associated reactivity impact of uniform pin pitch variations. For this set of cases, the change in 

pin pitch is applied uniformly, meaning that all pins at all axial heights experience the same 
A-22

changes at the same time. Pin pitches were expanded in several steps from nominal until the 

outer row of pins was in contact with the storage basket wall (for PWR fuel and dechanneled 

BWR fuel), representing the maximum amount of pin pitch expansion without a deformed basket 

structure.

For both GBC-32 and GBC-68 models, the rod/assembly deformation was first modeled with the 

pin pitch expanded uniformly across the assembly until the outer fuel rod unit cells were 

coincident with the inner surface of the storage cells. Once the fuel enrichment, burnup, and 

decay time, which results in the largest increase in reactivity, have been determined, this limiting 

condition is expanded until the fuel rods are modeled in contact with the storage cell walls. An 

illustration of the maximum uniform pin pitch configuration for the GBC-32 is shown in Figure 

A.11.

Two different fuel assembly representations are considered for the GBC-68 cask model—with 

and without the channel present. For the models where the channel is present, the pin pitch 

expansion is constrained by the contact of the fuel pins with the assembly channel. For the 

models with no channel present, the pin pitch is constrained when the outer fuel rods make 

contact with the neutron absorber wrappers and storage cell walls as shown in Figure A.12.

Figure A.11 Maximum uniform pitch expansion case in GBC-32

Basket cell wall
A-23

Figure A.12 Maximum uniform pitch expansion configuration in GBC-68

The results of the uniform pitch expansion cases with the outer unit cell boundary in contact with 

the storage cell wall are shown in Table A.26 and Table A.27 for GBC-32 and for unchanneled 

fuel in the GBC-68 models, respectively. For the cases evaluated the 44.25 GWd/MTU burnup 

case with 5 wt % 235U initial enrichment and 5-year decay time had the largest change for the 

GBC-32 system, and the fresh fuel with 5 wt % 235U initial enrichment had the largest change for 

the GBC-68 system. For both the PWR and BWR fuels, the results indicate that the reactivity 

impact of uniform pitch expansion is reduced at higher burnups and increased decay times.

Table A.26 Increase in keff caused by uniform fuel pin pitch expansion in GBC-32 (pin 

pitch at 102.7% of nominal, 1.2941 cm)

Burnup

(GWd/MTU)

Decay time

(years)

Increase in keff

(%∆keff)

1.92 0 0.78

3.5 25.5 1.48

44.25 5 1.69

44.25 80 1.67

44.25 300 1.66

70 5 1.53

70 80 1.44

70 300 1.42

Steel wrapper

Cell wall
A-24

Table A.27 Increase in keff caused by uniform fuel pin pitch expansion in GBC-68

Burnup

(GWd/MTU)

Decay time

(years)

Increase in keff

(%∆keff)

Channel intact, pin pitch 102.3% of nominal, 

1.3249 cm

0 0 2.06

35 5 1.76

35 80 1.76

35 300 1.72

70 5 1.64

70 80 1.55

70 300 1.60

Channel removed, pin pitch 117.5% of nominal, 

1.5222 cm

0 0 12.07

35 5 10.56

35 80 10.45

35 300 10.48

70 5 9.64

70 80 9.40

70 300 9.43

The maximum uniform expansion in the GBC-32 system results in a pitch of 1.3179 cm, which is 

104.6% of the nominal pitch. The resulting reactivity increase is 2.65 %∆keff. The maximum 

uniform increase for unchanneled BWR fuel is a pitch of 1.5773 cm, which is 121.8% of the 

nominal pitch. The resulting reactivity increase is 13.22 %∆keff. For channeled BWR fuel, the 

maximum pitch is 1.3249 cm, which is 102.3% of the nominal pitch. The results for all burnup 

and decay time combinations are shown in Table A.27 and show that the maximum reactivity 

increase is 2.06 %∆keff for fresh 5 wt % 235U fuel.

S2(b)C2: Radial non-uniform pin pitch variation—Results from the uniform pin pitch 

expansion cases indicate that the reactivity increase was continuing to rise as the pitch was 

expanded up to the basket cell wall. This indicates that additional pitch expansion of the inner 

region of the fuel assembly can result in even higher keff values. To evaluate this potential 

further, the limiting cases from the uniform pin pitch expansion cases were modeled with the pin 

pitch of the rods located in the inner region of the assembly allowed to expand until the outer 

rows were in contact with each other—at the basket boundary for the PWR fuel and the 

dechanneled BWR fuel, or at the channel inner boundary for channeled BWR fuel. Each of 

these configurations results in non-uniform radial pin pitch variations present across the 

assembly. An example model is shown in Figure A.13. For these models, the pitch in each of 

the outer rows is constant within the row and is equal to the pitch that caused that row to make 

contact with the outboard row or the basket wall. 
A-25

PWR Unchanneled BWR

Figure A.13 Example configuration for non-uniform radial pitch expansion

The results of the calculations with increasing pitch are shown in Figure A.14, Figure A.15, and

Figure A.16 for the GBC-32 model and the GBC-68 model with and without BWR channels, 

respectively. The results are presented as a function of the pitch of the inner, uniform portion of 

the assembly (i.e., maximum pin pitch modeled in assembly). The inner region pin pitch 

expansion was continued until the keff increase began to diverge. For the GBC-32 results, the 

first five points show the increase in keff associated with uniform pitch expansion, with 

nonuniform expansion beginning when the fuel rod pitch is in excess of 1.32 cm. The GBC-32 

maximum keff increase was 3.90 %∆keff vs. 2.65 %∆keff when using a uniform pin pitch 

expansion. For the GBC-68 results, nonuniform radial expansion causes a reactivity increase for 

channel constrained pin pitch expansion compared to uniform pin pitch expansion. The 

maximum increase for radially nonuniform pin pitch expansion is 2.80 %∆keff compared with

2.09 %∆keff for uniform radial pin pitch expansion constrained by a nominal size channel. The 

first point in Figure A.15 shows the impact of uniform pin pitch expansion, and the remaining 

four points are for non-uniform pin pitch expansion. The results for unchanneled fuel are shown 

in Figure A.16; the first six points show the increase in keff associated with uniform pin pitch 

expansion, and nonuniform pin pitch expansion begins when the fuel rod pitch is in excess of 

1.58 cm. The GBC-68 maximum keff increase is 13.31 %∆keff vs. 13.22 %∆keff when using a 

uniform pin pitch expansion. 
A-26

Figure A.14 Increase in keff in GBC-32 as a function of fuel rod pitch

(5.0 wt % 235U initial enrichment, 44.25 GWd/MTU burnup)

Figure A.15 Change in keff in GBC-68 as a function of fuel rod pitch for channeled fuel

(5.0 wt % 235U initial enrichment fresh fuel)

0.00

0.50

1.00

1.50

2.00

2.50

3.00

3.50

4.00

4.50

1.25 1.30 1.35 1.40 1.45 1.50

Increase in keff (% 

Δ

keff)

Fuel Rod Pitch in Central Uniform Region (cm)

Uniform pitch 

expansion

Nonuniform 

pitch expansion

0.0

0.5

1.0

1.5

2.0

2.5

3.0

1.25 1.3 1.35 1.4 1.45 1.5 1.55 1.6

Increase in 

keff (%

Δ

keff)

Fuel rod pitch (cm)

Uniform pitch expansion

Nonuniform pitch expansion
A-27

Figure A.16 Change in keff in GBC-68 as a function of fuel rod pitch for unchanneled fuel 

(5.0 wt % 235U initial enrichment fresh fuel)

S2(b)C3: Axial non-uniform pin pitch variation—Cases were developed to evaluate axially 

non-uniform pin pitch variations to represent the effects of “birdcaging” and/or “bottlenecking” of 

a fuel assembly. Models were constructed with two different pin pitches at different axial 

heights. The increased and decreased pitch variations were applied over discrete sections of 

the fuel rods, not as continuous changes as a function of height. The region of expanded pitch 

was selected based on the burnup of the assembly, near the top end for fuel irradiated to a 

burnup of more than 30 GWd/MTU or near the middle for fresh fuel. The length of the expanded 

zone was varied to evaluate changes in reactivity. The contracted pitch was determined by the 

size of non-fuel guide, instrument, or water tubes in the various fuel assembly designs which are 

assumed to maintain nominal dimensions. This resulted in contracted pitches of 95.7% and 

97.4% of nominal for the PWR and BWR assemblies, respectively. The contracted pitch zone 

was also varied in length to evaluate the impacts on reactivity. The expanded pin pitch zones 

included the pitch required for the fuel rod unit cells to touch the inside of the canister basket 

walls. Only unchanneled BWR fuel was considered.

The limiting pin pitch expansion case for the GBC-32 model was for the 5 wt % 235U initial 

enrichment fuel with a burnup of 44.25 GWd/MTU, so the most reactive axial section is near the 

top of the fuel assembly. The irradiated fuel was represented with 20.32 cm length segments to 

capture the axial burnup profile, and these segments are used to delineate the pin pitch 

variation. The size of the contracted pin pitch region was varied from one to four segments, and 

the expanded pitch section at the top of the assembly ranged from two to eight segments in 

length in an effort to identify the maximum change in keff attributable to birdcaging. An example 

with four segments in the contracted region and four segments in the upper expanded region is 

shown in Figure A.17. Slight reactivity increases were observed in the cases with four or more 

fuel segments in the expanded pin pitch zone. Results for the 5 wt % 235U initial enrichment fuel 

with a burnup of 44.25 GWd/MTU and 5-year decay time are shown in Table A.28. The 

maximum ∆keff change (i.e., 1.74%) is 0.05 %∆keff beyond the 1.69 %∆keff resulting from the 

0

2

4

6

8

10

12

14

1.25 1.30 1.35 1.40 1.45 1.50 1.55 1.60 1.65 1.70 1.75

Increase in 

keff (%

Δ

keff)

Fuel rod pitch (cm)

Uniform Pitch Expansion Nonuniform Pitch Expansion
A-28

uniform pin pitch expansion configuration. This additional increase in keff is considered 

negligible.

Notes: Fuel in expanded pin pitch segments is shown as black, regardless of isotopic composition.

Fuel in contracted pin pitch segments is shown in yellow, regardless of isotopic composition.

Large gaps between pairs of fuel rods indicate the presence of guide tubes.

Figure A.17 Assembly model with axially varying pin pitch in GBC-32

81.28 cm (contracted region)

203.2 cm (expanded region)

81.28 cm (expanded region)
A-29

Table A.28 Increase in keff in GBC-32 for axial nonuniform pin pitch expansion (5.0 wt 

% 235U initial enrichment, 44.25 GWd/MTU burnup, 5-year decay time)

Top expanded

zone length

(cm)

Contracted zone

length (cm)

Bottom expanded

zone length

(cm)

Increase in keff

(%∆keff)

40.64 60.96 264.16 1.00

60.96 20.32 284.48 1.58

60.96 40.64 264.16 1.58

60.96 60.96 243.84 1.57

60.96 81.28 223.52 1.58

81.28 20.32 264.16 1.72

81.28 40.64 243.84 1.70

81.28 60.96 223.52 1.70

81.28 81.28 203.2 1.71

101.6 40.64 223.52 1.71

101.6 60.96 203.2 1.71

101.6 81.28 182.88 1.72

121.92 60.96 182.88 1.72

121.92 81.28 162.56 1.73

142.24 60.96 162.56 1.74

142.24 81.28 142.24 1.73

162.56 60.96 142.24 1.72

162.56 81.28 121.92 1.72

The bounding axially non-uniform pitch expansion case was combined with the bounding 

radially non-uniform pin pitch expansion case to observe the net effect of birdcaging. The 

resulting keff increase compared to the nominal intact configuration is 3.89 %∆keff. This is 

statistically equivalent to the increase in keff for the axially uniform, radially nonuniform pin pitch 

expansion case (i.e., 3.90 %∆keff). The radial nonuniform pin pitch expansion bounds the 

potential effects of birdcaging.

 

The limiting pin pitch expansion case for the GBC-68 cask contains fresh fuel, so the most 

reactive axial portion of the assembly will be at the center. For that reason, the birdcaging 

analysis includes two contracted pitch sections, each 30.48 cm in length, symmetrically 

positioned above and below the mid-plane of the assembly. A range of center section lengths 

was considered, but no keff increase was observed in any case containing the contracted pitch 

sections. Results for the 5 wt % 235U enrichment fresh fuel cases are shown in Table A.29. The 

effects of birdcaging do not result in any additional keff increase beyond the 13.22 %∆keff

associated with the uniform pin pitch expansion configuration for fresh fuel in the GBC-68 cask.

Table A.29 Change in keff in GBC-68 for axial nonuniform pin pitch expansion (5.0 wt % 

235U enriched fresh fuel)

Central expanded

zone length

(cm)

Contracted

zone lengths

(cm)

End expanded

zone lengths

(cm)

Change in keff

(%∆keff)

137.16 60.96 60.96 12.24

198.12 60.96 30.48 12.77

228.60 60.96 15.24 12.91

259.08 60.96 0 13.02
A-30

A.3 CRITICALITY ANALYSIS TO ASSESS IMPACTS OF

CATEGORY 3 – CHANGES TO ASSEMBLY AXIAL ALIGNMENT

The neutron absorber panels in fuel storage and transportation packages are designed to 

extend beyond the length of the active fuel region within the fuel assembly, but they typically do 

not extend the full length of the basket. The primary function of the neutron absorber is to 

provide sufficient thermal neutron removal along the active fuel length between adjacent 

assemblies. Changes to the axial alignment of the assembly within the basket cavity can result 

in adjacent cells with fuel assemblies that do not have absorber panel material between them. 

Because spent fuel has axially varying fuel compositions, alignment changes above and below 

the absorber plate region are being evaluated. The upward (toward the lid) misalignment is 

likely to be of higher importance for SNF criticality evaluations because the reactivity of the 

assembly is driven by the burnup gradient near the top of the assembly. Case descriptions are 

provided in Table A.30 and described in more detail below

Table A.30 Cases developed to address Category 3—changes to assembly axial 

alignment

Identifier

Configuration 

scenario

Subclass

description

Parameter

varied

S3(a)C1 Axial 

displacement of 

intact fuel

Not 

applicable

Length of active fuel above or 

below absorber region

S3(a)C1: Axial displacement of intact fuel—Cases were developed to evaluate the impact of 

changes to intact fuel axial alignment. The parameter of interest for criticality safety is the length 

of fuel that is beyond the neutron absorber panel region of the basket. Because of differences in 

canister design, fuel spacers, and assembly designs, the amount of space available for the 

active fuel to potentially be outside the absorber envelope will vary by canister system and site 

and will affect the extent of the range of movement possible. 

The PWR fuel axial displacement models were developed by translating all the fuel assemblies 

within a canister uniformly up or down to achieve fixed uncovered fuel lengths for evaluation. 

The GBC-32 basket as described in Ref. [4] has the fuel and basket at the same length. To be 

more representative of typical storage and transportation package designs, the stainless steel 

portion of the basket was extended in these models to go from the canister baseplate to 5 cm 

below the lid. The length of fuel exposed above and below the absorber envelope was varied up 

to a maximum of 30 cm in 5 cm increments to map the reactivity response as a function of 

uncovered fuel length. The variation of the keff change as a function of axial position is shown in 

Figure A.18 for several fuel initial enrichment and burnup combinations. The assembly 

displacements in the GBC-32 model are equal to the length of exposed fuel because absorber 

panels in the GBC-32 model coincide with the active fuel length of the fuel assembly in the 

nominal alignment. A summary of GBC-32 results for displacements of 20 cm and 30 cm above 

the neutron absorber panel is shown in Table A.31. These results show that larger reactivity 

changes can occur for higher burnups and longer decay times. However, the larger reactivity 

increase is less than the reduction in nominal intact configuration case keff caused by the 

additional burnup and decay time. The nominal intact configuration keff values for each of the 

sensitivity cases are provided in Table A.1, resulting in margins of 3.5 to 15.1 %∆keff compared 
A-31

to the 44.25 GWd/MTU, 5-year decay time case. The largest ∆keff is associated with this 

moderate burnup, and short decay time point, even though a larger reactivity increase is 

associated with higher burnups and longer decay times.

Figure A.18 Change in keff in GBC-32 as a function of assembly axial displacement

Table A.31 Change in keff for assembly axial displacement in GBC-32 (displacement 

relative to the neutron absorber panel)

Enrichment

(wt % 235U)

Burnup

(GWd/MTU)

Decay time

(years)

Increase in keff (%∆keff)

30 cm 

displacement

20 cm 

displacement

1.92 0 0 1.52 0.06

3.5 25.5 5 7.07 3.22

5 44.25 5 7.69 3.64

5 44.25 80 8.28 4.18

5 44.25 300 8.26 4.15

5 70 5 8.06 4.13

5 70 80 8.28 4.51

5 70 300 8.34 4.54

The BWR fuel axial displacement models were developed translating all the fuel assemblies 

within a canister uniformly up or down to achieve fixed uncovered fuel lengths for evaluation. 

Exposed fuel ranges were varied up to a maximum of 31.78 cm above the absorber in 4 cm 

increments to map the reactivity response as a function of uncovered fuel length. The variation 

of the keff change as a function of axial position is shown in Figure A.19. The results for 

displacements of 13.65 cm and 31.78 cm above the neutron absorber panel are shown in 

Table A.32 and Table A.33, respectively. The results indicate that the effect of misalignment 

increases with burnup and is higher at longer decay times. As with the PWR cases, the larger 

-1

0

1

2

3

4

5

6

7

8

9

-15 -10 -5 0 5 10 15 20 25 30

Change in 

keff (%

Δ

keff)

Assembly axial displacement (cm)

1.92 wt% U-235, Fresh 3.5 wt% U-235, 25.5 GWd/MTU, 5 year

5 wt% U-235, 44.25 GWd/MTU, 300 year 5 wt% U-235, 70 GWd/MTU, 300 year
A-32

reactivity increases are less than the reduction in the nominal intact configuration case keff

caused by the additional burnup and decay time. The nominal intact configuration keff values for 

each of the sensitivity cases are provided in Table A.1, resulting in margins of 0.75 to 8.0 %∆keff

compared to the 35 GWd/MTU, 5-year decay time case, which is more than 11 %Δkeff less 

reactive than the 0.95 limit. The largest fuel reconfiguration keff is associated with this moderate 

burnup, and short decay time point, even though a larger reactivity increase is associated with 

higher burnups and longer decay times.

Figure A.19 Increase in keff in GBC-68 as a function of assembly axial displacement.

Misalignment toward the bottom of the canister causes a significantly smaller keff increase 

because the fuel at the bottom end of the assembly has lower reactivity than that at the top end. 

The misalignment toward the canister base plate differs for GBC-68 compared to GBC-32 

because the GBC-68 model has more distance below the fuel in the nominal intact 

representation, so larger misalignments are shown. 

-2

0

2

4

6

8

10

12

14

16

18

20

22

-15 -10 -5 0 5 10 15 20 25 30 35

Increase in 

keff (%

Δ

keff)

Uncovered Fuel (cm)

Fresh - Channeled Fresh - Unchanneled

35 GWd/MTU, 80 year - Channeled 35 GWd/MTU, 80 year - Unchanneled

70 GWd/MTU, 80 year - Channeled 70 GWd/MTU, 80 year - Unchanneled

Uncovered fuel at top of assembly Uncovered fuel at bottom 

of assembly
A-33

Table A.32 Increase in keff caused by loss of assembly position control in GBC-68, 

limited displacement of 13.65 cm above absorber

Burnup

(GWd/MTU)

Decay time

(years)

Increase in keff

(%∆keff)

Channel Intact

0 0 0.33

35 5 6.29

35 80 6.70

35 300 6.66

70 5 8.03

70 80 8.52

70 300 8.49

Channel Removed

0 0 0.27

35 5 6.07

35 80 6.49

35 300 6.42

70 5 7.78

70 80 8.24

70 300 8.20

Table A.33 Increase in keff for limited assembly axial displacement in GBC-68, 

displacement of 31.78 cm above absorber

Burnup

(GWd/MTU)

Decay time

(years)

Increase in keff

(%∆keff)

Channel Intact

0 0 12.52

35 5 19.40

35 80 19.84

35 300 19.83

70 5 20.47

70 80 20.73

70 300 20.76

Channel Removed

0 0 12.02

35 5 18.66

35 80 19.10

35 300 19.06

70 5 19.71

70 80 19.96

70 300 20.00

B-1

APPENDIX B. SHIELDING EVALUATIONS FOR FUEL 

RECONFIGURATIONS

B.1 OBJECTIVE

The objective of the shielding analysis is to identify cask external dose rate increases resulting 

from fuel configuration changes that may affect the ability of transportation packages/storage 

casks to comply with regulatory dose rate limits. This appendix describes a technical approach 

for shielding analyses of fuel reconfigurations in transportation packages and storage casks and 

provides reference values for such analyses. The term “fuel reconfiguration” refers to any 

change to the storage and transportation system nominal intact fuel assembly configuration 

used for the basis of cask certification. The analyzed fuel reconfiguration categories are:

(1) cladding failure where fuel fragments and particulates from multiple failed fuel rods can 

relocate near to and far from fuel cladding breach locations; 

(2) rod/assembly deformation; and 

(3) changes to assembly axial alignment. 

The specific fuel rod failure percentage values used in this study were selected arbitrarily 

because information about the expected fuel rod failure rate was not available for high-burnup 

fuel or for fuel in long-term storage.

This study primarily addresses fuel reconfigurations in transportation packages because specific 

regulatory dose rate limits have been developed for transportation packages under normal 

conditions of transport (NCT) and hypothetical accident conditions (HAC) (10 CFR 71.47 and 

71.51). For storage of spent nuclear fuel (SNF), annual dose limits (i.e., no specific dose rate 

limits) at the controlled area boundary are specified in 10 CFR 72.104 and 72.106 for normal,

off-normal, and accident conditions, respectively. Compliance with 10 CFR 72.104 can be 

demonstrated only on a site-specific basis because compliance depends on both cask and site 

characteristics. In this study, a non-site-specific shielding analysis (NUREG-1536) is performed 

for a generic 4×2 cask array to determine the impact on dose rate at the controlled area 

boundary for damaged SNF in accordance with 10 CFR 72.106. An analysis of the storage cask 

surrounding dose rates that contribute to occupational exposure (10 CFR 20.1201 

requirements) is also provided in the report. The effects of fuel configuration changes on the 

neutron and gamma dose rates were determined at 1 m from a generic storage cask.

The consequences of fuel configuration changes were evaluated by comparing external dose 

rate values between the nominal intact and fuel reconfigurations. This approach helps identify 

the fuel reconfigurations that yield higher external dose rates than the nominal intact fuel 

configuration and thus have the potential to exceed the regulatory limits. An advantage of this 

approach is that the conclusions of the analysis generally can be applied to similar package 

shielding analysis models because the relative effect of fuel configuration changes on package

external dose rates is less sensitive to the specific model parameters used in the analysis than 

the absolute dose rate values. Package/storage cask external gamma and neutron dose rate 

values were evaluated separately because fuel configuration changes have different effects on 

gamma and neutron radiation (e.g., internal high-Z materials have pronounced gamma

self-shielding effects), and the overall impact of fuel configuration changes on the external dose 

rate depends on the individual gamma and neutron dose rate contributions to the total dose 

rate. 
B-2

Generic package models for pressurized water reactor (PWR) and boiling water reactor (BWR)

SNF, identified as GBC-32 and GBC-68, respectively, were developed for this study and the

Westinghouse (W) 17×17 optimized fuel assembly (OFA) and General Electric (GE) 14 10×10 

fuel assemblies were selected as representative PWR and BWR fuel assembly types, 

respectively. The cask models have general shielding characteristics (e.g., thick inner gamma 

shield and outer neutron shield) similar to real high-capacity SNF transportation packages. The 

initial enrichment, burnup, and decay time values for the PWR and BWR fuel assemblies were 

5 wt % 235U, 65 GWd/MTU, and 5 and 40 years, respectively. The axial variation of the gamma 

and neutron radiation sources was based on typical PWR and BWR fuel axial burnup profiles.

Applicable codes, guidance, and regulatory criteria related to transportation/storage cask 

shielding performance were used as the basis for model development and analysis approach. 

The computer programs and calculation method employed in this analysis are provided in

Sect. B.2. The reference transportation and storage cask models as well as representative fuel 

assembly models for the shielding analysis are described in Sect. B.3. The fuel configurations 

and specific calculation parameters (e.g., failed fuel fraction) for transportation 

packages/storage casks evaluated in this analysis are described in Sect. B.4. Shielding analysis 

results are provided in Section B.5.

B.2 USE OF SOFTWARE

The radiation source terms for the shielding analysis were determined with the depletion and 

decay capabilities of the Scale 6.1.2 code system [31], including Transport Rigor Implemented 

with Time-Dependent Operation for Neutronic depletion (TRITON), Oak Ridge Isotope 

Generation and Depletion Code–Automatic Rapid Processing (ORIGEN-ARP), and Oak Ridge 

Isotope Generation in Scale (ORIGEN-S). The neutron and gamma radiation source terms were 

calculated in the group structure of the Scale 27N-19G Evaluated Nuclear Data Files, Part BVII.0 (ENDF/B-VII.0) shielding library.

The Scale 6.1.2 shielding analysis sequence Monaco with Automated Variance Reduction using 

Importance Calculations (MAVRIC) [31] and the Scale 27N-19G ENDF/B-VII.0 shielding library 

were used to perform Monte Carlo transport and dose rate calculations. MAVRIC uses Denovo, 

a discrete ordinates code [31], to determine particle importance as a function of position and 

energy and uses Monaco to perform Monte Carlo transport calculations. Radiation transport 

optimization is accomplished by: (1) sampling more often source particles that have an ability to 

produce a significant dose rate value outside the source regions, and (2) reducing the variance 

of particle scores in the spatial region of interest. The MeshView utility in the Scale code system 

enables visualization of detailed radiation dose maps produced by MAVRIC.

The American National Standards Institute/American Nuclear Society Standard 6.1.1-1977 [32]

flux-to-dose-rate conversion factors were used in all dose rate calculations, as recommended in

NUREG-1617 [29] and NUREG-1536 [2].

B.3 GENERIC TRANSPORTATION PACKAGE/STORAGE CASK AND 

REPRESENTATIVE FUEL ASSEMBLY MODELS

The generic models and representative PWR and BWR fuel assemblies described in this 

section were used in shielding calculations to determine reference dose rate values for failed 

fuel in transportation packages and storage casks. A transportation package model (see Sect.

B.3.1) and a storage cask model (see Sect. B.3.2) were developed because transportation and 
B-3

storage overpacks are typically different. Descriptions of the representative PWR and BWR 

intact fuel assembly models are provided in Sect. B.3.3 and B.3.4, respectively. The 

transportation package and storage cask models have shielding features similar to actual 

transportation packages and storage casks. Therefore, the trends identified in this analysis for 

dose rate change because of fuel configuration changes relative to the nominal intact 

configuration are applicable to actual transportation packages and storage casks. In addition to 

the overpack shielding design, the geometry and source term characteristics of the canister 

internal components— such as activation source strength, assembly hardware heights, fuel 

axial burnup profile, the length of the assembly top and bottom spacers, and the available free 

volume within the canister cavity—also affect the relative changes in neutron and gamma dose 

rates caused by fuel configuration changes. That is, variations in these characteristics will cause 

different changes in the external dose rates because of fuel configuration changes relative to 

the nominal intact configuration.

B.3.1 Generic Transportation Package Models

The GBC-32 and GBC-68 [4, 5] cask/package models originally were developed for criticality 

safety analyses of PWR and BWR SNF, respectively, and do not include appropriate shielding 

features. Simplified SNF transportation package and storage cask models, which are variations

of the GBC-32 and GBC-68 original models, were developed for this shielding analysis. The 

simplified transportation package models include a thick stainless steel (304) cylindrical shell, 

which serves as a gamma shield, and outer radial and bottom resin neutron shields with the 

elemental composition described in [33]. Cutaway views of the PWR and BWR package

shielding models showing the back half and bottom cask sections are illustrated in Figure B.1

(a) and (b) and Figure B.2 (a) and (b), respectively. As seen in the figures, the radial neutron 

shield does not extend the entire length of the package radial surface, which is a typical feature 

of transportation packages. The top, radial, and bottom thicknesses of the stainless steel 

cylindrical shell are 35, 24.5, and 24.5 cm, respectively; the radial and bottom thicknesses of the 

neutron shield are 7.5 and 3.5 cm, respectively. The canister inner radius is 87.5 cm for both 

models. The basket cell inside dimension and cell wall thickness are 22 cm and 0.75 cm, 

respectively, for the PWR package model and ~15 cm and 0.75 cm, respectively, for the BWR 

package model. The lower and upper fuel spacers that establish assembly axial position within 

the basket cell were not explicitly modeled; however, the nominal intact fuel configuration 

includes void regions in place of the spacers (20 and 10 cm long for the PWR and BWR 

package models, respectively) above and below the fuel assembly, as seen in the figures. Note 

that the lengths of the axial void regions are different for the PWR and BWR fuel assemblies 

because the BWR fuel assembly is longer than the PWR fuel assembly (see Table B.1 and

Table B.2). Radial steel fins, pocket trunnions, and small holes that typically cause radiation 

streaming were not included in the simplified models for the shielding analysis.

The fuel assembly model includes assembly active fuel, plenum, and upper and lower hardware 

regions, with contents homogenized within the volume delimited by the adjacent basket plates 

and the region axial height. This simple assembly representation is typically used in shielding 

analyses because it simplifies the model geometry description, produces slightly conservative 

results (primarily because of slightly reduced gamma self-shielding of the homogenized fuel 

contents), and reduces the computer time. The active fuel region was subdivided into axial 

zones to facilitate the description of the gamma and neutron source axial profiles. 
B-4

Stainless steel

cask body

Resin neutron 

shield

(a) (b)

Figure B.1 Cutaway view of the reference GBC-32 model showing (a) back half and (b) 

bottom half sections

Stainless steel

cask body

Resin neutron 

shield

(a) (b)

Figure B.2 Cutaway view of the reference GBC-68 model showing (a) back half (b) 

bottom half sections
B-5

Table B.1 PWR fuel assembly model description

Assembly axial 

region

Height 

(cm)

Stainless 

steel 304 

(kg)

Inconel 

718 (kg)

Zircaloy4 (kg)UO2

(kg) U (kg)

Smear 

density 

(g/cm3)

a

Free void 

volume 

(cm3)

Upper end fitting 13.44 6.89 0.96 8.58 - - 1.59 4211.29

Upper fuel plenum and 

upper end spacer grid 

14.66 0.11b 0.885 3.27 - - 0.74 6475.72

Active fuel 365.76 0.00 0.0 113.43c 480.191 426.00 3.35 102,136.74d

Lower fuel plenum and 

end fitting

11.95 5.90 0.0 2.72 - - 1.36 4627.33

aContents homogenized within basket cell volume corresponding to the assembly axial region (22 cm width). bIncludes SS304 grid sleeve and SS302 plenum spring. 

cIncludes fuel cladding, guide tubes, instrument tube, and in-core spacers. dFree void volume within basket cell and outside fuel rods, guide tubes, instrument tube, and in-core spacers; total 

volume for the region is 22×22×365.76 cm3 = 177,028 cm3.

Table B.2 BWR fuel assembly model description

Assembly axial 

region

Height 

(cm)

Stainless 

steel 304 

(kg)

Zircaloy-2 

(kg)

UO2

(kg) U (kg)a

Smear 

density 

(g/cm3)

b

Free void 

volume 

(cm3)

Upper tie and 

expansion spring 

22.29 2.0 1.57 ˗ ˗ 0.71 4552.74

Upper fuel plenum 28.55 2.32 6.78 ˗ ˗ 1.41 5136.20

Active fuel 368.91 ˗ 68.88 204.67 180.40 3.28 56,009.03c

Lower tie plate 18.76 4.77 1.86 ˗ ˗ 1.56 3360.93

aRef. [34]. bContents homogenized within basket cell volume corresponding to the assembly axial region (15.0435 cm width). cFree void volume within basket cell and outside fuel rods, water rods, and spacers.

B.3.2 Storage Cask Models

The SNF storage cask model for the analysis of dose rate surrounding a single storage cask is 

illustrated in Figure B.3. The storage cask model includes a 70 cm thick radial concrete shield 

and 40 cm thick top and bottom concrete shields, similar to some existing cask designs. The 

cask model also includes air vents located at cask top and bottom regions. PWR and BWR 

intact fuel configurations are identical to the GBC-32 and GBC-68 models, respectively. Nonsite-specific shielding analysis (NUREG-1536) was performed for a generic 4×2 storage cask 

array to determine the impact of damaged SNF on the dose rate at the controlled area 

boundary, which was located 100 m from the cask array. The generic cask array is shown in 

Figure B.4 along with a cutaway view of the storage cask model with intact PWR fuel. The 

skyshine calculation model includes the storage cask array, a 60-cm thick concrete storage pad, 

a 100 cm thick soil layer, and dry air within a 120×120×100 m3 volume surrounding the cask 

array. The model for the intact fuel configuration, which is simplified to facilitate skyshine 

calculations, has radially homogenized fuel contents. However, the cask models for intact PWR 

and BWR SNF assemblies specify appropriate fuel axial burnup profiles.
B-6

Figure B.3 Storage cask model with intact fuel assemblies

Figure B.4 Illustration of a 4×2 storage cask array model for site boundary dose rate 

calculations

B.3.3 Description of Representative PWR Fuel Assembly

The representative PWR fuel assembly for radiation source term and shielding calculations is 

the W 17×17 OFA with fuel assembly design parameters documented in Refs. [35] and [36].

The fuel assembly model consists of four axial regions: lower end fitting, active fuel, plenum, 
B-7

and upper end fitting regions. The height, fuel and non-fuel material weights, smear density, and 

free void volume are provided in Table B.1.

The weights for stainless steel and Inconel assembly parts were used to calculate activation 

source terms, primarily consisting of 60Co, for assembly hardware and fuel plenum regions. The 

volume occupied by each hardware component was calculated as the component weight 

divided by its material mass density. The transversal dimension of a GBC-32 assembly basket 

cell used in free void volume calculations is 22 cm. The free void volume within the rectangular 

parallelepiped axially delimited by the assembly axial region height and transversally delimited 

by the basket cell plates was used in subsequent source intensity spatial distribution and mass 

density calculations for the different fuel configuration models.

Radiation source terms were determined for the PWR fuel assembly with a 65 GWd/MTU 

burnup, a 5 wt % 235U initial enrichment, and 5- and 40-year decay times. The axial distributions 

of gamma and neutron radiation sources in the active fuel region were determined in 

accordance with the 18-zone assembly burnup profile documented in Ref. [30] for criticality 

calculations. The gamma radiation source axial profile was based on the direct proportionality 

between gamma radiation source strength and fuel burnup; the neutron radiation source axial 

profile was based on the variation of the neutron source strength as the burnup value raised by 

the power of 4.2 [33]. The profiles thus determined are illustrated in Figure B.5 (a) for the

gamma radiation source and Figure B.5 (b) for the neutron radiation source. The active fuel 

region of the PWR assembly was subdivided into 18 axial zones to facilitate the description of 

the axial gamma and neutron radiation source distributions.

(a) (b)

Figure B.5 Source intensity distribution as a function of PWR fuel axial zone for (a) 

gamma radiation source and (b) neutron radiation source

B.3.4 Description of the Representative BWR Fuel Assembly

The representative BWR fuel assembly for radiation source term and shielding calculations is 

the GE14 fuel assembly type with fuel assembly design parameters documented in Refs. [37]

and [31]. The fuel assembly model consists of four axial regions: lower tie plate, active fuel, 

plenum, upper fuel plenum, and upper tie and expansion spring regions. The height, fuel and 

non-fuel material weights, smear density, and free void volume values are provided in Table 

B.2. Fuel assembly hardware weight and material specifications for a Quad Cities 1 reactor 8×8 
B-8

reload fuel assembly [38] were used to calculate the activation sources for the representative 

GE14 assembly because this information was not available for the GE14 assembly. The volume 

occupied by each assembly hardware component was calculated as the component weight 

divided by its material density. The transversal dimension of a GBC-68 assembly basket cell 

used in free void volume calculations is ~15 cm. The calculated free void volume within the 

rectangular parallelepiped axially delimited by the height of each assembly axial region and 

transversally delimited by the basket cell plates was used in subsequent spatial source 

distribution and mass density calculations for the different fuel configuration models.

Radiation source terms were determined for the BWR fuel assembly corresponding to 65 

GWd/MTU burnup, 5 wt % 235U initial enrichment, and 5- and 40-year decay times. Gamma and 

neutron sources in the active fuel region have an axial distribution based on the 25-zone 

assembly burnup profile from LaSalle Unit 1 documented in Ref. [30] for criticality calculations. 

The gamma radiation source axial profile was developed based on the direct proportionality 

between the gamma source strength and the burnup value; the neutron radiation source axial 

profile was developed based on the variation of the neutron source strength as the burnup value

raised by the power of 4.2 [33]. The profiles thus determined are illustrated in Figure B.6 (a) for 

the gamma radiation source and Figure B.6 (b) for the neutron radiation source. In the cask 

model, the active fuel region of the BWR assembly was subdivided into 25 axial zones to 

facilitate description of the axial gamma and neutron radiation source distributions. Note that the 

gamma radiation source axial profile is relatively flat over a large fuel axial region, whereas the 

neutron radiation source axial profile is significantly pointed in the middle. The axial distribution 

shape of the BWR neutron radiation source also significantly differs from the axial distribution 

shape of the PWR neutron radiation source, which is flatter, as seen in Figure B.5.

(a) (b)

Figure B.6 Source intensity distribution as a function of BWR fuel axial zone for (a) 

gamma radiation source and (b) neutron radiation source

B.4 DESCRIPTION OF FUEL RECONFIGURATION CATEGORY

CALCULATION MODELS FOR SHIELDING

The calculation models developed for the fuel reconfiguration categories are described in this 

section. The configuration categories include cladding failure (see Sect. B.4.1), rod/assembly 

deformation without cladding failure (see Sect. B.4.2), and changes to assembly axial alignment 
B-9

without cladding failure (see Sect. B.4.3). A variety of fuel failure effects are addressed in this 

calculation, including assembly deformation or alignment changes and relocation within different 

canister cavity regions of fuel mixture from varying percentages of failed fuel rods. 

B.4.1 Category 1: Cladding Failure

To investigate the effects of cladding failure on shielding objectives, two configuration scenarios

were developed: (1) breached spent fuel rods where assemblies have partially failed and 

partially intact fuel rods; and (2) damaged fuel representing the maximum degree of fuel 

configuration change that may occur under cladding failure scenarios.

S1(a)—Breached spent fuel rods

The following configurations were evaluated for the transportation package models:

• Fuel mixture relocated to the middle section of the fuel assembly;

• Fuel mixture relocated to the bottom section of the fuel assembly;

• Fuel mixture relocated to the top section of the fuel assembly.

For the PWR fuel, the evaluated percentages of fuel mixture relocated to different assembly 

regions are 10% (27 rods/assembly) and 25% (66 rods/assembly). For the BWR fuel, the

evaluated percentage of fuel mixture relocated to different assembly regions is 11% 

(10 rods/assembly). Dose rate calculations were performed for the PWR and BWR package 

models under both NCT and HAC. The fuel rod failure rate values were arbitrarily selected 

because this information is not available for commercial high-burnup fuel or for spent fuel in 

long-term storage. However, the whole range of fuel failure with respect to the amount of fuel 

mixture (0–100%) that may be displaced to inner cavity regions is analyzed in this report under 

different consequence categories. The 25% value for PWR SNF was analyzed to determine 

whether there is a direct proportionality between the amount of relocated fuel mixture and 

external dose rate (i.e., whether an increase in the displaced fuel mixture quantity from 10% to 

25% of the assembly fuel mixture equates to an increase in external dose rate by a factor of 

2.5). The effects of a lower percentage of failed fuel, such as single fuel rod failure, were not 

considered because the change in cask external dose rates is estimated to be comparable to 

calculation statistical error (i.e., non-discernible effect). 

This configuration scenario assumes that the fuel pellets are fractured and that the fuel 

fragments and particulates can be released from fuel rods that have developed cladding 

breaches. Pellet fracturing is caused by fuel exposure to thermal gradients early in life. The 

number of fragments observed for typical fuel varies from 20 to 50 fragments per pellet and 

more fragmentation has been observed for higher burnup fuel [39]. For this configuration, 

assembly spacers are assumed intact so that fuel configuration changes cannot occur beyond 

the assembly location. The fuel material will likely collect beneath its parent rods close to the 

cladding breach locations (if the package is in the horizontal orientation) or into the assembly 

lower region (if the package is in the vertical orientation). However, fuel in the horizontal 

orientation, when subject to transportation vibrations, could relocate axially within the cask from 

the active fuel region to either bottom or top regions. Furthermore, these fragments and 

particulates may be spread across the available volume or be packed closely together within a 

small volume. It is assumed that the most limiting mass packing fraction of the released fuel 

fragments and particulates that results in the largest source intensity per volume unit is 0.58. 

This value is based on powder mechanics for particles similar to sand [27]. A mass packing 

fraction of 0.67 was also used in sensitivity calculations.
B-10

Axial locations beyond the normal active fuel location are considered in the analysis because 

fuel rubble collected into such locations has a greater potential of increasing external top or 

bottom dose rate because of its proximity to those surfaces. In addition, neutron streaming 

above or below the radial neutron shield may significantly increase the neutron dose rate on the 

cask external radial surface. Therefore, fuel relocation beyond the active fuel region is expected 

to produce significant dose rate increase relative to the nominal intact fuel configuration for the 

package external upper and lower regions.

Input parameter calculation approach:

Fuel reconfiguration models require input parameters that appropriately describe radiation 

source and fuel mixture spatial distributions within the package/cask inner cavity. For a realistic 

dose rate evaluation of fuel reconfigurations, the total source intensity and mass balance must 

be maintained within the cask. This means that if the source intensity is increased for the 

assembly region into which fuel material is collected, the source intensity for the active fuel 

region must be proportionally reduced so that the total source intensity does not change relative 

to the nominal intact fuel configuration.

 

Calculation of the radiation source terms and mass density values is described for the 10% fuel 

failure rate configuration in which fuel material released from 27 PWR failed fuel rods per 

assembly is collected into the lower assembly region assuming a 0.58 mass packing fraction. 

To simplify the radiation source calculation for the fuel reconfiguration, it is assumed that 10% of 

the total fuel mass per assembly is released into the lower assembly region, i.e., the source 

intensity of the relocated material is 10% of the total source intensity. The weight of the released 

fuel material is ~49 kg per assembly, and the minimum volume outside the fuel rods that may be 

occupied by the released fuel material is ~8227 cm3

, based on the 0.58 packing fraction. The 

free void volume corresponding to the assembly lower end fitting region in Table B.1 is 4627.3 

cm3

, which indicates that up to 56% (i.e., 4627.3/8227×100) of the released fuel mixture can be 

collected into the assembly lower end fitting region, and the remainder can be collected into the 

first (bottom) axial zone of the active fuel region. Because of fuel relocation, mass densities, 

elemental compositions, and radiation source terms for assembly lower end fitting and active 

fuel axial zones will change relative to the nominal intact assembly configuration. For example, 

the activation source of the assembly lower end fitting must be combined with 56% of the 

gamma and neutron source of the released fuel (i.e., 5.6% of the total gamma and neutron 

source intensities) to determine new source terms for the lower end fitting. The source intensity 

in the active fuel region will be reduced, relative to the nominal intact fuel intensity, by an 

amount proportional to the relocated fuel (i.e., 10%). Then 44% of the radiation source of the 

released fuel (i.e., 4.4% of the total gamma and neutron source intensities) is added to the 

radiation source of the first active fuel axial zone.

S1(b)—Damaged fuel

The evaluated configurations for the damaged fuel scenario are

• homogenized rubble packed closely together into the bottom of the inner cavity assuming 

a 0.58 mass packing fraction, as illustrated in Figure B.7 (a) for the transportation package 

model; and

• homogenized rubble assuming the rubble occupies the whole inner cavity, as illustrated in 

Figure B.7 (b) for the transportation package model. 
B-11

Configuration (a) is consistent with a package/storage cask in the vertical orientation. 

Configuration (b) is consistent with a package/storage cask in the horizontal orientation. A 

transportation package may have both vertical and horizontal orientations during normal 

operations. These configurations are bounding for damaged SNF resulting in fuel fragments and 

particulates being collected into inner cavity regions below or above the assembly spacers or 

between fuel basket outer plates and the canister radial wall.

Homogenized 

basket plate 

material

Homogenized 

assembly and 

basket plate 

material

(a) (b)

Figure B.7 Illustration of the transportation package models for damaged SNF

configurations with (a) fuel rubble collected into package cavity bottom and 

(b) fuel rubble homogenized within package cavity
B-12

B.4.2 Category 2: Rod/Assembly Deformation

The evaluated configurations consist of intact fuel rods collapsed against the basket plates of a 

transportation package, as illustrated in Figure B.8 (a) and (b), potentially resulting from a

horizontal drop or tip-over. Dose rate calculations were performed for the PWR package model

under both NCT and HAC, and for the BWR package model with fuel rods collapsed as shown 

in Figure B.8 (a) under NCT.

(a) (b)

Figure B.8 Illustration of the cask model for the collapsed fuel rod configuration

B.4.3 Category 3: Changes to Assembly Axial Alignment

Fuel spacers are typically used to ensure that the fuel assemblies are aligned within a fixed 

orientation within the storage cask or transportation package. This category evaluates the 

potential impacts of changes to the original fuel assembly axial alignment, which could result in 

higher dose rates at the package top and/or bottom external surfaces.

The configurations evaluated for this category are

• intact fuel assemblies axially displaced to the cavity bottom surface, as illustrated for 

the PWR fuel in Figure B.9 (a); and

• intact fuel assemblies axially displaced to the cavity top surface as illustrated for the 

PWR fuel in Figure B.9 (b).
B-13

(a) (b)

Figure B.9 Illustration of the transportation cask models for axial assembly 

displacement to (a) canister cavity bottom and (b) canister cavity top

Dose rate calculations were performed for the PWR and BWR package models (under both 

NCT and HAC) and for the PWR and BWR storage cask models. 

B.5 SHIELDING ANALYSIS RESULTS

The generic PWR and BWR packages were assumed to be transported by exclusive shipment. 

For normal conditions of transport, a package transported by exclusive shipment must not 

exceed the 10 CFR 71.47 dose rate limits for the package outer surfaces (1000 mrem/h), outer 

surfaces of the conveyance (200 mrem/h), 2 m from the conveyance outer lateral surfaces 

(10 mrem/h), and any normally occupied space (2 mrem/h). Gamma and neutron dose rates for 

a package under NCT were calculated for the package top, side, and bottom external surfaces 

and at 2 m from the conveyance, the lateral surfaces of which were assumed to be located at 

30 cm from the package external surfaces. Although the distance between the package outer 

surfaces and the 2 m locations may vary for the actual transportation packages, the dose rate 

effects on each surface of interest due to fuel configuration changes relative to the nominal 

intact fuel configuration are relatively similar at different locations surrounding the package. 

For HAC (design basis fire and 30 ft drop), the neutron shield was assumed to be completely 

lost. The dose rate values for a package under HAC were evaluated at 1 m from the package 

external surfaces. 

The ratio of the maximum surface dose rate for a fuel reconfiguration to the maximum surface 

dose rate for the intact fuel configuration, which is identified as the failed to intact (F/I)
B-14

configuration dose rate ratio, is provided in the tables included in this section. Note that the 

locations of the maximum dose rate at the package side surface for intact fuel and failed fuel 

may be different. 

The gamma dose rates at the bottom and top surfaces vary significantly as a function of decay 

time because of the decay of 60Co, which is an activation source in the assembly plenum and 

lower and upper hardware regions. For the 5-year decay time, the gamma dose rate at the 

package bottom and top surfaces is high primarily because of the proximity of the 60Co

activation source to those surfaces, activation source strength, and energetic gamma rays of 

1.25 MeV average energy. The 60Co activation source has negligible contributions to gamma 

dose rate for the 40-year decay time and beyond (t1/2=5.271 years). Therefore, the contribution 

of the 60Co activation source to the gamma dose rates at the package bottom and top surfaces 

varies as a function of decay time. A large contribution of the 60Co activation source to the 

gamma dose rate, as in the case of a 5-year decay time, equates to a relatively small increase 

of the gamma dose rate because of fuel relocation to assembly hardware regions. A small or 

negligible contribution of the 60Co activation source to the gamma dose rate, as in the case of a

40-year decay time, equates to a relatively large increase of the gamma dose rate because of 

fuel relocation to assembly hardware regions. That is, the effects on the gamma dose rates at 

the package bottom and top surfaces of fuel relocation into the assembly hardware regions are 

significantly smaller for the 5-year decay time than those for the 40-year decay time.

Beyond 5 years after fuel discharge from the reactor, the F/I neutron dose rate ratio is relatively 

insensitive to the decay time because neutrons are primarily produced from spontaneous fission 

and (alpha,n) reaction because of 244Cm (t1/2 = 18.1 years) and 238Pu (t1/2 = 87.7 years). The F/I 

gamma dose rate ratio is relatively insensitive to decay times larger than 40 years because the 

activation source in the assembly plenum and hardware regions is negligible and the fuel 

gamma dose rate is dominated by 137Cs (t1/2 = 30.07 years). 

B.5.1 PWR Package Model Assuming Normal Conditions of Transport

The changes (F/I dose rate ratio) relative to the nominal intact fuel configuration of the 

maximum neutron and gamma dose rates on the PWR package external surfaces and at 2 m

from the package are summarized in Table B.3 and Table B.4, respectively. The neutron and 

gamma dose rate profiles on the PWR package external surfaces and at 2 m from the package 

are shown in Figure B.10 through Figure B.35 for the 40-year decay time for each fuel 

reconfiguration case and the nominal intact configuration for comparison.
B-15

Table B.3 PWR package maximum dose rate change for NCT: package external 

surfaces

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation region a

Decay 

time 

(years)

F/I neutron dose rate 

ratio b

F/I gamma dose rate 

ratio b

1 10% fuel rod 

failure Assembly active fuel

5 1.09 1.01 1.04 1.01 1.02 1.00

40 1.09 1.01 1.03 1.00 1.01 1.00

2 10% fuel rod 

failure Assembly lower end fitting

5 1.06 1.62 2.52 1.00 0.99 1.34

40 1.06 1.59 2.53 1.00 0.98 5.17

3 10% fuel rod 

failure Assembly plenum 

5 3.96 2.37 1.03 0.98 0.99 1.01

40 3.96 2.38 1.02 3.35 0.99 0.99

4 10% fuel rod 

failure Assembly upper end fitting

5 4.98 2.75 1.05 0.93 0.99 1.01

40 4.99 2.77 1.02 11.55 0.99 1.00

5 25% fuel rod 

failure Assembly active fuel

5 1.12 1.06 1.00 1.00 1.12 1.00

40 1.13 1.06 1.00 1.00 1.10 0.96

6 25% fuel rod 

failure Assembly lower end fitting

5 1.05 1.76 2.76 1.00 0.98 1.33

40 1.05 1.76 2.76 1.00 0.97 5.16

7 25% fuel rod 

failure Assembly plenum 

5 4.59 2.79 1.00 0.97 0.95 0.99

40 4.58 2.80 0.99 3.34 0.96 0.96

8 25% fuel rod 

failure Assembly upper end fitting

5 6.16 3.50 1.00 0.93 0.95 1.00

40 6.13 3.51 0.99 11.59 0.96 0.95

9 Damaged Canister cavity bottom

5 0.32 2.57 4.18 0.001 1.19 1.72

40 0.32 2.59 4.13 0.02 1.21 8.37

10 Damaged Entire canister cavity

5 6.69 3.89 3.04 0.84 1.10 1.43

40 6.67 3.89 3.05 14.10 1.24 7.30

11 Assembly 

deformation See Figure B.8 (a) 

5 1.36 1.13 1.22 0.96 1.11 0.93

40 1.36 1.14 1.19 0.87 1.10 0.64

12 Assembly 

deformation See Figure B.8 (b)

5 1.27 1.03 1.17 0.99 0.93 0.96

40 1.16 1.04 1.16 0.89 0.96 0.80

13 Alignment 

changes See Figure B.9 (a)

5 0.80 1.15 1.35 0.60 1.00 1.66

40 0.80 1.14 1.34 0.58 1.00 1.77

14 Alignment 

changes See Figure B.9 (b)

5 1.26 1.10 0.78 1.59 1.29 0.59

40 1.26 1.10 0.77 1.68 1.01 0.57

aFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions, as 

described further in this section.

bRelative error (95% confidence level) less than 5% for all values except for the value for case #9, gamma dose rate 

on top surface, the relative error is 15%.
B-16

Table B.4 PWR package maximum dose rate change for NCT: 2 m from the package 

surfaces

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation region a

Decay 

time 

(years)

F/I neutron dose rate 

ratio b

F/I gamma dose rate 

ratio b

1 10% fuel rod 

failure Assembly active fuel

5 1.07 1.02 1.02 1.01 1.00 1.00

40 1.08 1.02 1.02 1.00 1.00 0.99

2 10% fuel rod 

failure Assembly lower end fitting

5 1.04 1.02 2.29 1.01 0.99 1.33

40 1.06 1.06 2.23 1.02 1.01 4.82

3 10% fuel rod 

failure Assembly plenum 

5 3.48 1.02 1.04 0.99 0.98 1.01

40 3.36 1.07 1.00 1.26 1.01 0.98

4 10% fuel rod 

failure Assembly upper end fitting

5 4.27 1.01 1.04 0.94 0.99 1.01

40 4.19 1.08 1.00 2.85 1.01 0.98

5 25% fuel rod 

failure Assembly active fuel

5 1.11 1.04 0.99 1.00 1.04 0.99

40 1.11 1.02 0.98 0.99 1.00 0.96

6 25% fuel rod 

failure Assembly lower end fitting

5 1.04 0.99 2.51 1.00 0.96 1.34

40 1.05 1.03 2.39 0.96 1.00 4.76

7 25% fuel rod 

failure Assembly plenum 

5 4.08 1.06 0.99 0.98 0.95 0.99

40 4.00 1.09 0.96 1.28 1.01 0.96

8 25% fuel rod 

failure Assembly upper end fitting

5 5.27 1.25 0.99 0.95 0.95 0.99

40 5.13 1.15 0.97 2.88 1.01 0.95

9 Damaged Canister cavity bottom

5 0.30 1.02 4.01 0.03 1.07 2.17

40 0.29 0.94 3.95 0.26 0.94 10.56

10 Damaged Entire canister cavity

5 5.94 1.28 2.90 1.04 1.11 1.78

40 5.79 1.24 2.80 5.02 1.32 9.11

11 Assembly 

deformation See Figure B.8 (a) 

5 1.31 1.15 1.19 0.98 1.10 0.93

40 1.32 1.17 1.17 1.04 1.09 0.73

12 Assembly 

deformation See Figure B.8 (b)

5 1.22 1.04 1.15 1.00 0.94 0.95

40 1.12 1.04 1.11 0.96 0.96 0.77

13 Alignment 

changes See Figure B.9 (a)

5 0.79 1.01 1.30 0.60 1.00 1.67

40 0.80 1.01 1.29 0.86 1.01 1.74

14 Alignment 

changes See Figure B.9 (b)

5 1.25 1.02 0.79 1.61 1.00 0.59

40 1.28 1.02 0.78 1.57 1.00 0.57

aFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions, as 

described further in this section.

bRelative error (95% confidence level) less than 7% for all values except for the value for case #9, gamma dose rate 

on top surface, the relative error is 15%.
B-17

Cases 1 to 4 evaluate the impact of 27 failed fuel rods per assembly (~10% failure rate) 

distributed to different axial locations. Discussions of the results are as follows:

Case 1. The redistributed fuel mixture is located at 285.5 to 326 cm relative to the package 

bottom surface. This location corresponds to assembly axial fuel zones 12 and 13. The fuel 

relocation caused ~9% increase in the maximum neutron dose rate on the top surface of the 

package. The gamma dose rate change relative to the nominal intact configuration was 

negligible for all surfaces and decay times evaluated. The neutron and gamma dose rate 

profiles along the PWR package outer surfaces and at the 2 m locations are shown in 

Figure B.10 and Figure B.11, respectively, for the 40-year decay time. 

Case 2. The redistributed fuel mixture is located at 50 to 82.3 cm relative to the package bottom 

surface. This location corresponds to the assembly lower end fitting and bottom axial fuel zone. 

The maximum neutron dose rate at the package bottom surface was ~2.5 times as large as that 

of the nominal intact configuration. For the 40-year decay time, the maximum gamma dose rate 

on the package bottom surface increased by a factor of ~5. Similar effects were obtained for the 

maximum neutron and gamma dose rates at 2 m from the package surfaces. The neutron and 

gamma dose rate profiles along the PWR package outer surfaces and at the 2 m locations are 

shown in Figure B.12 and Figure B.13, respectively, for the 40-year decay time. This case was 

also analyzed for a mass packing fraction of 0.67. The neutron and gamma dose rate values 

based on the 0.67 mass package fraction were within the statistical errors (i.e., 7% at the 95% 

confidence level) of the values based on the 0.58 mass package fraction.

Case 3. The redistributed fuel mixture is located at 407.4 to 442.4 cm relative to the package

bottom surface. This location corresponds to the assembly top axial fuel and plenum zones. The 

maximum neutron dose rates on the package top and side surfaces were ~4 and ~2.5 times, 

respectively, as large as the corresponding dose rates for the nominal intact configuration. At 2 

m from the package top surface, the maximum neutron dose rate increased by a factor of ~3.5. 

The gamma dose rate increase relative to the nominal intact configuration was negligible for the 

5-year decay time. For the 40-year decay time, the maximum gamma dose rate at the package 

top surface increased by a factor of ~3.5. The neutron and gamma dose rate profiles along the 

PWR package outer surfaces and the 2 m locations are shown in Figure B.14 and Figure B.15, 

respectively, for the 40-year decay time.

Case 4. The redistributed fuel mixture is located at 427.7 to 455.8 cm relative to the package

bottom surface. This location corresponds to the assembly top axial fuel and plenum zones. The 

maximum neutron dose rates on the package top and side surfaces were ~5 and ~3 times, 

respectively, as large as the corresponding dose rates for the nominal intact configuration. At 2 

m from the package top surface, the maximum neutron dose rate increased by a factor of ~4.5. 

The gamma dose rate increase relative to the nominal intact configuration was negligible for the 

5-year decay time. For the 40-year decay time, the maximum gamma dose rate at the package

top surface increased by a factor of ~12. At 2 m from the package top surface, the maximum 

gamma dose rate increased by a factor of ~3. The neutron and gamma dose rate profiles along 

the PWR package outer surfaces and the 2 m locations are shown in Figure B.16 and 

Figure B.17, respectively, for the 40-year decay time.
B-18

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time. 

Figure B.10 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into the free void volume of two axial fuel zones—PWR package out

(Case 1)
B-19

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.11 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into the free void volume of two axial fuel zones—2 m locations, PW

(Case 1)
B-20

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.12 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly lower end fitting region—PWR package outer surface
B-21

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.13 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly lower end fitting region—2 m locations, PWR packag
B-22

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.14 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into the top axial fuel and plenum regions—PWR package outer sur
B-23

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.15 Dose rate comparison between the intact configuration and the fuel reconfiguration assum

mixture collected into the top axial fuel and plenum regions—2 m locations, PWR package 
B-24

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.16 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly plenum and upper end fitting regions—PWR package 

NCT (Case 4)
B-25

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.17 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly plenum and upper end fitting regions—2 m locations

(Case 4)
B-26

Cases 5 to 8 evaluate the impact of 66 failed fuel rods per assembly (~25% failure rate) 

distributed to different axial locations. Discussions of the results are as follows:

Case 5. The redistributed fuel mixture is located at 244.8 to 326 cm relative to the package

bottom surface. This location corresponds to the assembly axial fuel zones 10 through 13. This 

fuel reconfiguration caused ~6% and ~12% increases in the maximum neutron dose rates on 

the side and top surfaces of the package, respectively, and ~12% increase in the maximum 

gamma dose rate on the side surface of the package, relative to the nominal intact 

configuration. Similar effects were obtained at 2 m from the package surfaces. The neutron and 

gamma dose rate profiles along the PWR package outer surfaces and the 2 m locations are 

shown in Figure B.18 and Figure B.19, respectively, for the 40-year decay time.

Case 6. The redistributed fuel mixture is located at 50 to 122.9 cm relative to the package

bottom surface. This location corresponds to the assembly lower end fitting and axial fuel zones 

1 through 3. The maximum neutron dose rate at the package bottom surface was ~3 times as 

large as that of the nominal intact configuration. The maximum gamma dose rate increase on 

the package bottom surface was ~30% for the 5-year decay time. For the 40-year decay time, 

the maximum gamma dose rate on the package bottom surface increased by a factor of ~5. 

Similar increases were obtained for the maximum neutron and gamma dose rates at 2 m from 

the package surfaces. The neutron and gamma dose rate profiles along the PWR package 

outer surfaces and the 2 m locations are shown in Figure B.20 and Figure B.21, respectively, for 

the 40-year decay time.

Case 7. The redistributed fuel mixture is located at 366.8 to 442.4 cm relative to the package

bottom surface. This location corresponds to the axial fuel zones 16 through 18 and to the 

plenum region. The maximum neutron dose rate values for the package top and side surfaces 

were ~4.6 and ~2.8 times, respectively, as large as those of the nominal intact configuration. At 

2 m from the package top surface, the maximum neutron dose rate increased by a factor of 

~4.1. Gamma dose rate increase relative to the nominal intact configuration was negligible for 

the 5-year decay time. For the 40-year decay time, the maximum gamma dose rate at the 

package top surface increased by a factor of ~3.5. The neutron and gamma dose rate profiles 

along the PWR package outer surfaces and the 2 m locations are shown in Figure B.22 and 

Figure B.23, respectively, for the 40-year decay time. 

Case 8. The redistributed fuel mixture is located at 387 to 455.8 cm relative to the package

bottom surface. This location corresponds to the assembly axial fuel zones 17 and 18, plenum 

region, and upper end fitting. The maximum neutron dose rate values for the package top and 

side surfaces were ~6 and ~3.5 times, respectively, as large as those of the nominal intact 

configuration. At 2 m from the package top surface, maximum neutron dose rate increased by a 

factor of ~5. Gamma dose rate increase relative to the nominal intact configuration was 

negligible for the 5-year decay time. For the 40-year decay time, the maximum gamma dose 

rate at the package top surface increased by a factor of ~12. At 2 m from the package top 

surface, the maximum gamma dose rate increased by a factor of ~3. The neutron and gamma 

dose rate profiles along the PWR package outer surfaces and the 2 m locations are shown in 

Figure B.24 and Figure B.25, respectively, for the 40-year decay time.
B-27

Topa Sidea Bot

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.18 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into the free void volume of four axial fuel zones—PWR package ou

(Case 5)
B-28

Topa Sidea Bot

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.19 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into the free void volume of four axial fuel zones—2 m locations, PW

(Case 5)
B-29

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.20 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly bottom region—PWR package outer surfaces, NCT (C
B-30

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.21 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly bottom region—2 m locations, PWR package NCT (Ca
B-31

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.22 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into top axial fuel and plenum regions—PWR package outer surface
B-32

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.23 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into top axial fuel and plenum regions—2 m locations, PWR packag

 
B-33

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.24 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into top axial fuel, plenum, and upper end fitting regions—PWR pac

NCT (Case 8)
B-34

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.25 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into top axial fuel, plenum, and upper end fitting regions—2 m locat

NCT (Case 8)
B-35

Cases 9 and 10 evaluate the impact of extensive cladding failure representative of damaged 

fuel, where the fuel mixture is represented as collecting into the internal regions adjacent to the 

canister wall such as the space between the outer basket plates and canister wall. Discussions 

of the results are as follows:

Case 9. For this case, the stainless steel basket plate materials and fuel rubble from the 

damaged fuel rods are homogenously distributed within the canister bottom region as shown in

Figure B.8 (a). The height and mass density of the homogenous mixture were based on a mass 

packing fraction of 0.58 for the fuel rubble. The maximum neutron dose rate values for the 

bottom outer region of the package were ~4 times as large as those of the nominal intact 

configuration, and the maximum neutron dose rate at the package side surface increased by a 

factor of ~3. For the 40-year decay time, the maximum gamma dose rate at the package bottom 

surface increased by a factor of ~8, and the maximum gamma dose rate at 2 m from the 

package bottom surface increased by a factor of ~11.The neutron and gamma dose rate profiles 

along the PWR package outer surfaces and the 2 m locations are shown in Figure B.26 and 

Figure B.27, respectively, for the 40-year decay time. This case was also analyzed for a mass 

packing fraction of 0.67. The neutron and gamma dose rate values based on a 0.67 mass 

packing fraction were within the statistical errors (i.e., 7% at the 95% confidence level) of the 

results based on the 0.58 mass packing fraction.

Case 10. This case represents fuel rubble from the damaged fuel rods as homogeneously 

distributed within the canister cavity as shown in Figure B.8 (b). This case is a bounding 

representation of extensive fuel damage that may cause the fuel mixture to collect into internal 

regions adjacent to the canister wall. The fuel mixture representation corresponds to a 0.22 

mass packing fraction. The neutron and gamma dose rates significantly increased outside the 

package. Compared to the nominal intact configuration, the maximum neutron dose rate values 

were ~7, 4, and 3 times as large for the package top, radial, and bottom surfaces, respectively 

and ~6, 1.5, and 3 times as large at 2 m from the package top, radial, and bottom surfaces, 

respectively. For the 40-year decay time, the maximum gamma dose rates at the package top 

surface, the package bottom surface, the 2 m top locations, and the 2 m bottom locations were 

~14, 7, 5, and 9 respectively, as large as those of the nominal intact configuration. The neutron 

and gamma dose rate profiles along the PWR package outer surfaces and the 2 m locations are 

shown in Figure B.28 and Figure B.29, respectively, for the 40-year decay time.
B-36

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PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.26 Comparison between dose rate values for the intact fuel configuration and the damaged S

with fuel rubble collected into canister cavity bottom—PWR package outer surfaces, NCT 
B-37

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.27 Comparison between dose rate values for the intact fuel configuration and the damaged S

with fuel rubble collected into canister cavity bottom—2 m locations, PWR package NCT (
B-38

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.28 Comparison between dose rate values for the intact fuel configuration and the damaged S

with fuel rubble homogeneously distributed within canister cavity—PWR package outer su

10)
B-39

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PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.29 Comparison between dose rate values for the intact fuel configuration and the damaged S

with fuel rubble homogeneously distributed within canister cavity—2 m locations, PWR pa

10)
B-40

Cases 11 and 12 evaluate the impact of fuel rods collapsed against the basket plates of a 

transportation package representative of rod/assembly deformation, and Cases 13 and 14 

evaluate the impacts of fuel axial displacement. Discussions of the results are as follows:

Case 11. All the PWR assembly fuel rods in the package are represented as collapsed against 

the basket plates as depicted in Figure B.9 (a). The maximum neutron dose rates on the top, 

radial, and bottom surfaces increased by ~40%, 15%, and 20%, respectively, relative to the 

nominal intact configuration. A relatively small increase (~10%) of the maximum gamma dose 

rate was obtained for the radial surfaces. The neutron and gamma dose rate profiles along the 

PWR package outer surfaces and the 2 m locations are shown in Figure B.30 and Figure B.31, 

respectively, for the 40-year decay time.

Case 12. All the PWR assembly fuel rods in the package are represented as collapsed against 

the basket plates as depicted in Figure B.9 (b). The maximum neutron dose rates on the top 

and bottom surfaces increased by ~30% and 20%, respectively, relative to the nominal intact 

configuration. The neutron and gamma dose rate profiles along the PWR package outer 

surfaces and the 2 m locations are shown in Figure B.32 and Figure B.33, respectively, for the 

40-year decay time. 

Case 13. All the PWR fuel assemblies in the package are displaced toward the canister bottom, 

as depicted in Figure B.7 (a). Relative to the nominal intact configuration, the maximum neutron 

dose rate value for the package radial surface increased by ~15%; the maximum neutron and 

gamma dose rate values for the bottom region increased by ~35% and 70%, respectively. The 

neutron and gamma dose rate profiles along the PWR package outer surfaces and the 2 m

locations are shown in Figure B.34 and Figure B.35, respectively, for the 40-year decay time. 

Case 14. All the PWR fuel assemblies in the package are displaced toward the canister cavity 

top, as depicted in Figure B.7 (b). Relative to the nominal intact configuration, the maximum 

neutron dose rate values for the top and radial surfaces increased by ~30% and 10%, 

respectively. The maximum gamma dose rate values for the top surfaces increased by ~70%. 

The maximum gamma dose rate on the package radial surface was ~30% higher than that of 

the nominal intact configuration for the 5-year decay time. The neutron and gamma dose rate 

profiles along the PWR package outer surfaces and the 2 m locations are shown in Figure B.36

and Figure B.37, respectively, for the 40-year decay time.
B-41

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PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.30 Comparison between dose rate values for the intact fuel configuration and the configurati

rods collapsed against the basket plates—PWR package outer surfaces, NCT (Case 11)
B-42

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.31 Comparison between dose rate values for the intact fuel configuration and the configurati

rods collapsed against the basket plates—2 m locations, PWR package NCT (Case 11)
B-43

Topa Sidea Bott

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PWR assemblies of 65 GWd/MTU burnup and 40-year decay time. 

Figure B.32 Comparison between dose rate values for the intact fuel configuration and the configurati

rods collapsed against the basket plates—PWR package outer surfaces, NCT (Case 12)
B-44

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time. 

Figure B.33 Comparison between dose rate values for the intact fuel configuration and the configurati

rods collapsed against the basket plates—2 m locations, PWR package NCT (Case 12)
B-45

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time. 

Figure B.34 Comparison between dose rate values for the intact fuel configuration and the configurati

displacement of the fuel assemblies to canister cavity bottom—PWR package outer surfac
B-46

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 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time. 

Figure B.35 Comparison between dose rate values for the intact fuel configuration and the configurati

displacement of the fuel assemblies to canister cavity bottom—2 m locations, PWR packa
B-47

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 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.36 Comparison between dose rate values for the intact fuel configuration and the configurati

displacement of the fuel assemblies to canister cavity top—PWR package outer surfaces, 
B-48

Topa Sidea Bott

 a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time. 

Figure B.37 Comparison between dose rate values for the intact fuel configuration and the configurati

displacement of the fuel assemblies to canister cavity top—2 m locations, PWR package N
B-49

B.5.2 PWR Package Assuming Hypothetical Accident Conditions

For the transportation package HAC calculations, the neutron shield considered in the PWR 

package model for NCT was replaced with air. The ratio of the maximum surface dose rate 

value for each fuel reconfiguration analyzed to the maximum dose rate value for the nominal 

intact configuration, which is identified as F/I dose rate ratio, is summarized in Table B.5. For 

illustrative purposes, the radial and axial dose rate values are shown in Figure B.38 for the fuel 

reconfiguration assuming 10% fuel failure rate with the fuel mixture redistributed into the

assembly bottom region. The values in the graphs correspond to a 40-year decay time.

The radial dose rate change relative to the nominal intact configuration was negligible for the 

majority of the fuel reconfigurations analyzed. The maximum increase in the radial neutron and 

gamma dose rates was approximately 20%. However, the neutron and gamma dose rates 

significantly increased at the top and bottom surfaces relative to the nominal intact configuration

for most fuel reconfigurations. Relative to the nominal intact configuration, fuel damage 

represented with fuel rubble homogeneously distributed within the canister cavity yielded the 

greatest increase in the neutron dose rate at the top surface, by a factor of ~6, and fuel damage 

represented with fuel rubble collected into the canister cavity bottom yielded the greatest 

increase in the neutron dose rate at the bottom surface, by a factor of ~4. For the 5-year decay 

time, the maximum F/I gamma dose rate ratio values on the top and bottom surfaces were ~1.6 

for assembly displacement to the canister cavity top and ~2.2 for damaged fuel represented 

with fuel rubble redistributed into the canister cavity bottom based on a mass packing fraction of 

0.58. For the 40-year decay time, the maximum F/I gamma dose rate ratio values on the top 

and bottom surfaces were ~6 and 10, for damaged fuel represented with fuel rubble 

homogeneously distributed within the canister cavity and damaged fuel represented with fuel 

rubble into the canister cavity bottom, respectively.
B-50

Table B.5 PWR package maximum dose rate change for hypothetical accident 

conditions

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation regiona

Decay 

time 

(years)

F/I neutron dose rate 

ratiob

F/I gamma dose rate 

ratiob

1 10% fuel rod 

failure Assembly active fuel

5 1.07 1.01 1.02 1.00 1.00 1.00

40 1.07 1.02 1.02 0.98 0.97 0.96

2 10% fuel rod 

failure Assembly lower end fitting

5 1.04 1.02 2.14 0.99 1.00 1.33

40 1.05 1.02 2.19 1.00 0.97 4.89

3 10% fuel rod 

failure Assembly plenum 

5 3.35 1.02 1.03 0.98 1.00 1.01

40 3.36 1.00 1.02 1.28 0.97 0.95

4 10% fuel rod 

failure Assembly upper end fitting

5 4.12 0.99 1.13 0.93 0.97 1.01

40 4.17 1.01 1.09 3.31 0.96 0.95

5 25% fuel rod 

failure Assembly active fuel

5 1.10 1.01 0.99 0.99 1.02 0.98

40 1.10 1.01 1.16 1.04 0.99 0.94

6 25% fuel rod 

failure Assembly lower end fitting

5 1.03 1.00 2.42 1.00 0.96 1.35

40 1.03 0.99 2.37 1.01 0.94 4.88

7 25% fuel rod 

failure Assembly plenum 

5 3.92 0.99 0.99 0.98 0.97 0.99

40 3.99 0.98 0.98 2.02 1.08 1.05

8 25% fuel rod 

failure Assembly upper end fitting

5 5.05 0.99 1.00 0.93 0.96 0.99

40 5.12 0.98 0.98 4.74 1.09 1.05

9 Damaged Canister cavity bottom

5 0.29 0.78 3.79 0.02 0.99 2.23

40 0.29 0.80 3.82 0.26 1.05 10.19

10 Damaged Entire canister cavity

5 5.80 0.96 2.73 1.06 1.14 1.83

40 5.90 0.93 2.69 5.43 1.17 8.66

11 Assembly 

deformation See Figure B.8 (a) 

5 1.31 1.12 1.19 0.99 1.09 0.93

40 1.30 1.13 1.17 1.04 1.06 0.66

12 Assembly 

deformation See Figure B.8 (b)

5 1.13 1.04 1.13 1.00 0.94 0.95

40 1.13 0.95 1.14 0.91 0.90 0.74

13 Alignment 

changes See Figure B.9 (a)

5 0.80 0.99 1.30 0.60 1.00 1.66

40 0.81 0.99 1.29 0.76 0.97 1.73

14 Alignment 

changes See Figure B.9 (b)

5 1.25 1.01 0.79 1.57 1.00 0.60

40 1.26 1.01 0.78 1.30 0.96 0.55

aSee Sect. B.5.1 for case description. bRelative error (at the 95% confidence level) less than 5% for all values except for the maximum gamma dose rate 

change for case #9, top surface, the relative error of which is 20%.
B-51

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.38 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly bottom region—1 m from package outer surfaces, HA
B-52

B.5.3 BWR Package Assuming Normal Conditions of Transport

The changes (F/I dose rate ratio) relative to the nominal intact fuel configuration of the 

maximum neutron and gamma dose rates on the BWR package external surfaces and at 2 m

from the package are summarized in Table B.6 and Table B.7, respectively. The neutron and 

gamma dose rate profiles on the BWR package external surfaces and at 2 m from the package 

are shown in Figure B.39 through Figure B.56 for the 40-year decay time for each fuel 

reconfiguration and the nominal intact configuration for comparison. 

Table B.6 BWR package maximum dose rate change for NCT: package external 

surfaces

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation regiona

Decay 

time 

(years)

F/I neutron dose rate 

ratiob

F/I gamma dose rate 

ratiob

1 11% fuel rod 

failure Assembly active fuel

5 1.13 1.02 1.06 0.98 1.02 0.99

40 1.14 1.03 1.07 1.01 1.01 0.97

2 11% fuel rod 

failure

Assembly lower tie plate and 

nosepiece

5 1.14 1.43 4.37 1.02 0.97 1.21

40 1.14 1.42 4.44 1.00 0.97 24.39

3 11% fuel rod 

failure Assembly plenum 

5 12.74 1.78 1.05 1.40 0.96 0.99

40 12.98 1.75 1.08 13.67 0.98 0.97

4 11% fuel rod 

failure

Assembly upper tie plate and 

handle

5 21.78 2.42 1.06 3.22 0.96 1.01

40 21.77 2.41 1.07 84.60 0.99 0.98

5 Damaged Canister cavity bottom

5 0.60 1.84 5.99 0.003 1.35 1.17

40 0.61 1.81 6.04 0.15 1.40 32.19

6 Damaged Entire canister cavity

5 23.48 3.27 4.31 2.90 0.98 0.94

40 23.52 3.26 4.37 84.18 1.11 26.71

7 Assembly 

deformation See Figure B.8 (a) 

5 1.48 1.07 1.26 1.01 1.08 1.00

40 1.49 1.06 1.27 1.01 1.07 0.89

8 Alignment 

changes See Figure B.9 (a)

5 0.88 1.01 1.16 0.69 1.00 1.44

40 0.89 1.01 1.17 0.71 1.01 1.54

9 Alignment 

changes See Figure B.9 (b)

5 1.12 1.00 0.86 1.42 1.00 0.71

40 1.14 1.00 0.87 1.39 1.01 0.72

aFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions, as 

described further in this section.

bRelative error (at the 95% confidence level) less than 5% for all values except for the maximum gamma dose rate 

change for case #5, top surface, the relative error is 20%.
B-53

Table B.7 BWR package maximum dose rate change for NCT: 2 m from package 

surfaces

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation regiona

Decay 

time 

(years)

F/I neutron dose rate 

ratiob

F/I gamma dose rate 

ratiob

1 11% fuel rod 

failure Assembly active fuel

5 1.11 1.04 1.04 0.99 0.98 0.99

40 1.11 0.98 1.03 0.98 1.00 0.99

2 11% fuel rod 

failure

Assembly lower tie plate and 

nosepiece

5 1.12 1.04 3.78 1.00 0.96 1.19

40 1.09 1.05 3.79 1.01 1.00 19.79

3 11% fuel rod 

failure Assembly plenum 

5 10.62 1.03 1.04 1.44 0.96 0.99

40 10.15 1.07 1.05 1.37 0.99 1.00

4 11% fuel rod 

failure

Assembly upper tie plate and 

handle

5 17.12 1.02 1.02 3.21 0.97 1.00

40 16.24 1.04 1.05 4.94 0.99 1.01

5 Damaged Canister cavity bottom

5 0.53 2.19 5.62 0.12 1.29 1.34

40 0.52 2.01 5.57 0.46 1.24 32.09

6 Damaged Entire canister cavity

5 25.23 1.23 3.93 3.28 0.98 1.07

40 18.88 1.11 3.89 6.07 1.18 26.50

7 Assembly 

deformation See Figure B.8 (a) 

5 1.41 1.06 1.20 1.00 1.05 1.00

40 1.37 1.04 1.19 0.96 1.05 0.89

8 Alignment 

changes See Figure B.9 (a)

5 0.88 1.01 1.13 0.70 0.99 1.42

40 0.88 1.01 1.18 0.92 1.01 1.53

9 Alignment 

changes See Figure B.9 (b)

5 1.15 1.01 0.85 1.41 1.00 0.72

40 1.14 0.97 0.87 1.08 1.02 0.75

aFarthest region for fuel relocation identified. Fuel mixture may extend into adjacent assembly axial regions, as 

described further in this section.

bRelative error (at the 95% confidence level) less than 7% for all values except for the maximum gamma dose rate 

change for case #5, top surface, the relative error is 20%.
B-54

Cases 1 to 4 evaluate the impact of 10 failed fuel rods per assembly (~11% failure rate) 

distributed to different axial locations. Discussions of the results are as follows:

Case 1. The redistributed fuel mixture is located from 221.1 to 260.6 cm relative to the package 

bottom surface. This location corresponds to the axial fuel zones 12 and 13. This fuel 

reconfiguration caused ~15% increase in the maximum neutron dose rate on the top surface of 

the package. The gamma dose rate change relative to the nominal intact configuration was 

negligible for all surfaces and decay times evaluated. The neutron and gamma dose rate 

profiles along the BWR package outer surfaces and the 2 m locations are shown in Figure B.39

and Figure B.40, respectively, for the 40-year decay time.

Case 2. The redistributed fuel mixture is located from 40 to 73.5 cm relative to the package 

bottom surface. This location corresponds to the assembly lower region. Relative to the nominal 

intact configuration, the maximum neutron dose rate at the package bottom surface increased 

by a factor of ~4.5. For the 40-year decay time, the maximum gamma dose rate at the package

bottom surface was ~24.5 times as large as that of the nominal intact configuration. Similar 

effects were obtained for the maximum neutron and gamma dose rate at 2 m from the package 

surfaces. The neutron and gamma dose rate profiles along the BWR package outer surfaces 

and the 2 m locations are shown in Figure B.41 and Figure B.42, respectively, for the 40-year 

decay time.

Case 3. The redistributed fuel mixture is located from 427.7 to 456.2 cm relative to the package

bottom surface. This location corresponds to the assembly plenum region. The maximum 

neutron dose rates on package top and radial surfaces were ~13 and ~2 times, respectively, as 

large as those of the nominal intact configuration. At 2 m from the package top surface, the 

maximum neutron dose rate increased by a factor of ~11. The gamma dose rate increase 

relative to the nominal intact configuration was small (e.g., 40% in the top outer region of the 

package) for the 5-year decay time. For the 40-year decay time, the maximum gamma dose 

rate at the package top surface increased by a factor of ~14. The neutron and gamma dose rate 

profiles along the BWR package outer surfaces and the 2 m locations are shown in Figure B.43

and Figure B.44, respectively, for the 40-year decay time.

Case 4. The redistributed fuel mixture is located from 456.2 to 478.5 cm relative to the package

bottom surface. This location corresponds to the assembly upper tie plate and handle region. 

The maximum neutron dose rates on the package top and radial surfaces were ~22 and ~3 

times, respectively, as large as those of the nominal intact configuration. At 2 m from the 

package top surface, the maximum neutron dose rate increased by a factor of ~17 relative to 

the nominal intact configuration. The maximum gamma dose rate increased relative to the 

nominal intact configuration by a factor of ~3 for the 5-year decay time. For the 40-year decay 

time, the maximum gamma dose rate at the package top surface increased by a factor of ~85. 

At 2 m from the package top surface, the maximum gamma dose rate increased by a factor of 

~5. The neutron and gamma dose rate profiles along the BWR package outer surfaces and the 

2 m locations are shown in Figure B.45 and Figure B.46, respectively, for the 40-year decay 

time.
B-55

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.39 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into the free void volume of two axial fuel zones—BWR package out

(Case 1)
B-56

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.40 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into the free void volume of two axial fuel zones—2 m locations, BW

(Case 1)
B-57

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.41 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly lower region—BWR package outer surfaces, NCT (Ca
B-58

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.42 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly lower region—2 m locations, BWR package NCT (Cas
B-59

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.43 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly plenum region—BWR package outer surfaces, NCT (C
B-60

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.44 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly plenum region—2 m locations, BWR package NCT (C
B-61

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.45 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly top region—BWR package outer surfaces, NCT (Case 
B-62

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.46 Dose rate comparison between the intact fuel configuration and the fuel reconfiguration a

fuel mixture collected into assembly top region—2 m locations, BWR package NCT (Case 
B-63

Cases 5 and 6 evaluate the impact of extensive cladding failure representative of damaged fuel, 

where the fuel mixture is represented as collecting into the internal regions adjacent to the 

canister wall such as the space between the outer basket plates and canister wall. Discussions 

of the results are as follows:

Case 5. This case represents fuel rubble from the damaged fuel rods being axially located from 

30 to 212 cm relative to the package bottom surface, as shown in Figure B.7 (a). The maximum 

neutron dose rate values for the bottom outer region of the package were ~6 times as large as 

those of the nominal intact configuration. The maximum neutron dose rate at the package radial 

surface increased by a factor of ~2. For the 40-year decay time, the maximum gamma dose rate 

values for the bottom outer region of the package increased by a factor of ~32. The maximum 

gamma dose rate values for the outer radial region of the package were higher by ~40% than 

those of the nominal intact configuration. The neutron and gamma dose rate profiles along the 

BWR package outer surfaces and the 2 m locations are shown in Figure B.47 and Figure B.48, 

respectively, for the 40-year decay time. 

Case 6. This case represents fuel rubble from the damaged fuel rods as homogeneously 

distributed within the canister cavity, as shown in Figure B.7 (b). This case is a bounding 

representation of extensive fuel damage that may cause the fuel mixture to collect into internal 

regions adjacent to the canister wall. The fuel mixture representation corresponds to a 0.23 

mass packing fraction. The neutron and gamma dose rates significantly increased outside the 

package. Compared to the nominal intact configuration, the maximum neutron dose rate values 

were ~24, 3.5, and 4.5 times as large for the package top, radial, and bottom surfaces, 

respectively, and ~25, 1.5, and 4 times as large at 2 m from the package top, radial, and bottom 

surfaces, respectively. For the 40-year decay time, the maximum gamma dose rate at the 

package top surface, package bottom surface, the 2 m top locations, and the 2 m bottom 

locations were ~84, 27, 6 and 27 respectively, as large as those of the nominal intact 

configuration. For the 5-year decay time, the maximum gamma dose rate values for the top 

outer region of the package increased by a factor of ~3. The neutron and gamma dose rate 

profiles along the BWR package outer surfaces and the 2 m locations are shown in Figure B.49

and Figure B.50, respectively, for the 40-year decay time.
B-64

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.47 Comparison between package dose rate values for the intact fuel configuration and the da

configuration with fuel rubble collected into canister cavity bottom—BWR package outer s

5)
B-65

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BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.48 Comparison between package dose rate values for the intact fuel configuration and the da

configuration with fuel rubble collected into canister cavity bottom—2 m locations, BWR p

5)
B-66

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 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.49 Comparison between package dose rate values for the intact fuel configuration and the da

configuration with fuel rubble homogeneously distributed within canister cavity—BWR pa

surfaces, NCT (Case 6)
B-67

Topa Sidea Bott

 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.50 Comparison between package dose rate values for the intact fuel configuration and the da

configuration with fuel rubble homogeneously distributed within canister cavity—2 m loca

package NCT (Case 6)
B-68

Case 7 evaluates the impact of fuel rods collapsed against the basket plates of a transportation 

package representative of rod/assembly deformation, and Cases 8 and 9 evaluate the impacts 

of fuel axial displacement. Discussions of the results are as follows:

Case 7. All the BWR assembly fuel rods in the package are represented as collapsed against 

the basket plates as depicted in Figure B.8 (a) for the PWR fuel. The maximum neutron dose 

rate increase of ~10%, 50%, and 30% relative to the nominal intact configuration was obtained 

for the radial, top, and bottom surfaces, respectively. A relatively small increase of 10% in the 

maximum gamma dose rate was obtained for the package outer radial region. The neutron and 

gamma dose rate profiles along the BWR package outer surfaces and the 2 m locations are 

shown in Figure B.51 and Figure B.52, respectively, for the 40-year decay time.

Case 8. All the BWR fuel assemblies in the package are displaced toward the canister bottom, 

as shown in Figure B.9 (a) for the PWR fuel. The maximum neutron and gamma dose rate 

values for the package bottom surface increased by ~20% and 50%, respectively, relative to the 

nominal intact configuration. Similar increase percentages were obtained for the 2 m bottom 

locations. The neutron and gamma dose rate profiles along the BWR package outer surfaces 

and the 2 m locations are shown in Figure B.53 and Figure B.54, respectively, for the 40-year 

decay time.

Case 9. All the BWR fuel assemblies in the package are displaced toward the canister top, as 

shown in Figure B.9 (b) for the PWR fuel. The maximum neutron dose rate values for the top 

outer region of the top increased by ~15% relative to the nominal intact configuration. The 

maximum gamma dose rate on the package top surface was ~40% higher than that of the 

nominal intact configuration for the 5- and 40-year decay times. Similar gamma dose rate 

increase percentage was obtained for the 2 m top locations and 5-year decay time. The neutron 

and gamma dose rate profiles along the BWR package outer surfaces and the 2 m locations are 

shown in Figure B.55 and Figure B.56, respectively, for the 40-year decay time. 
B-69

Topa Sidea Bott

 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.51 Comparison between package dose rate values for the intact fuel configuration and the co

assuming all fuel rods collapsed against the basket plates—BWR package outer surfaces, 
B-70

Topa Sidea Bott

 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.52 Comparison between package dose rate values for the intact fuel configuration and the co

assuming all fuel rods collapsed against the basket plates—2 m locations, BWR package 
B-71

Topa Sidea Bott

 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.53 Comparison between package dose rate values for the intact fuel configuration and the co

assuming displacement of the fuel assemblies to canister cavity bottom—BWR package o

(Case 8)
B-72

Topa Sidea Bott

 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.54 Comparison between package dose rate values for the intact fuel configuration and the co

assuming displacement of the fuel assemblies to canister cavity bottom—2 m locations, B

(Case 8)
B-73

Topa Sidea Bott

 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.55 Comparison between package dose rate values for the intact fuel configuration and the co

assuming displacement of the fuel assemblies to canister cavity top—BWR package outer 

(Case 9)
B-74

Topa Sidea Bott

 a

BWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

Figure B.56 Comparison between package dose rate values for the intact fuel configuration and the co

assuming displacement of the fuel assemblies to canister cavity top—2 m locations, BWR 

9)
B-75

B.5.4 BWR Package Assuming Hypothetical Accident Conditions

For the transportation package under HAC, the neutron shield considered in the BWR package 

model for NCT was replaced with air. The ratio of the maximum surface dose rate value for 

each fuel reconfiguration analyzed to the maximum surface dose rate value for the nominal 

intact configuration, which is identified as F/I dose rate ratio, is summarized in Table B.8.

Table B.8 BWR package maximum dose rate change for hypothetical accident 

conditions

Top Side Bottom Top Side Bottom

Case # Scenario Fuel relocation regiona

Decay 

time 

(years)

F/I neutron dose rate 

ratiob

F/I gamma dose rate 

ratiob

1 11% fuel rod 

failure Assembly active fuel

5 1.04 1.04 1.08 1.01 0.98 0.99

40 0.99 1.02 1.06 0.99 0.98 0.96

2 11% fuel rod 

failure

Assembly lower tie plate and 

nosepiece

5 1.04 1.01 3.57 1.00 0.96 1.21

40 1.01 1.00 3.65 1.01 0.97 16.99

3 11% fuel rod 

failure Assembly plenum 

5 6.86 1.04 1.04 1.42 0.97 1.00

40 6.72 1.01 1.07 1.58 0.98 0.98

4 11% fuel rod 

failure

Assembly upper tie plate and 

handle

5 11.36 1.00 1.06 3.18 0.97 1.00

40 11.06 1.00 1.05 6.49 0.97 0.94

5 Damaged Canister cavity bottom

5 0.43 0.77 5.26 0.06 1.02 1.37

40 0.44 0.77 5.33 0.33 1.03 26.24

6 Damaged Entire canister cavity

5 12.84 0.83 3.69 3.28 0.99 1.09

40 12.74 0.83 3.68 7.39 1.05 21.69

7 Assembly 

deformation See Figure B.8 (a) 

5 1.07 1.03 1.20 1.00 1.06 1.00

40 1.03 1.03 1.21 1.00 1.04 1.02

8 Alignment 

changes See Figure B.9 (a)

5 0.90 1.00 1.16 0.70 0.99 1.41

40 0.89 1.00 1.16 0.92 0.99 1.35

9 Alignment 

changes See Figure B.9 (b)

5 1.09 1.00 0.87 1.41 0.99 0.72

40 1.06 0.99 0.89 1.10 0.99 0.88

aSee Sect. B.5.3 for case description. bRelative error (at the 95% confidence level) less than 5% for all values except for the maximum gamma dose rate 

change for case #9, top surface, the relative error is 20%.

The radial dose rate change relative to the nominal intact configuration was within statistical 

error, i.e., negligible. However, the neutron and gamma dose rates significantly increased at the 

top and bottom surfaces relative to the nominal intact configuration for most fuel 

reconfigurations. Relative to the nominal intact configuration, fuel damage represented with fuel 

rubble homogeneously distributed within the canister cavity yielded the greatest increase in the 

neutron dose rate at the top surface, by a factor of ~13, and fuel damage represented with fuel 

rubble collected into the canister cavity bottom yielded the greatest increase in the neutron dose 

rate at the bottom surface, by a factor of ~5.5. For the 5-year decay time, the maximum 

increase in axial gamma dose rate relative to the nominal intact configuration was by a factor of 

~1.4 for assembly displacement to the package cavity top and bottom. For the 40-year decay 

time, the maximum F/I gamma dose rate ratio values on the top and bottom surfaces were ~7.5 
B-76

and 26, for damaged fuel represented with fuel rubble homogeneously distributed within the 

canister cavity and damaged fuel represented with fuel rubble into the canister cavity bottom, 

respectively.

B.5.5 Individual Storage Cask

Dose rates were calculated at 1 m from a concrete storage cask containing either PWR or BWR 

fuel assemblies of 65 GWd/MTU burnup and either 5- or 40-year decay time. The scenarios 

analyzed include: cladding breach (fuel mixture from 10% fuel assembly rods collected into the 

available free volume within assembly lower hardware region), fuel damage (fuel rubble 

collected into the canister cavity bottom), and changes to assembly axial alignment (fuel 

assembly displaced to the canister cavity bottom). The changes (F/I dose rate ratio) relative to 

the nominal intact fuel configuration of the maximum neutron and gamma dose rates at 1 m

from a concrete storage cask are summarized in Table B.9 and Table B.10 for PWR and BWR 

fuel assemblies, respectively. Figure B.57 shows the effects of 10% fuel rod failure with the fuel 

mixture redistributed into the assembly lower region on the PWR neutron and gamma dose rate 

profiles for the 40-year decay time. Fuel configuration changes cause significant dose rate 

increases relative to the nominal intact fuel configuration in the cask outer regions that face air 

vent locations. Away from air vent locations, the change in radiation dose rate is either small 

(e.g., 30% for damaged fuel configurations) or negligible. 

Table B.9 Maximum dose rate change at 1 m from a PWR storage cask

Topa Sidea Sideb Topa Sidea Sideb

Case # Scenario Fuel relocation region

Decay 

time 

(years)

F/I neutron dose rate 

ratioc

F/I gamma dose rate 

ratio

2 10% fuel rod 

failure Assembly lower end fitting

5

1.09 1.68 1.05

0.70 1.87 1.00

40 0.90 2.91 1.04

9 Damaged Canister cavity bottom

5

0.22 2.73 0.97

0.07 4.17 1.20

40 0.18 4.09 0.77

11 Alignment 

changes See Figure B.9 (a)

5

0.87 1.35 1.02

1.09 2.02 1.00

40 0.80 2.73 0.95

aFacing air vent locations; relative error (at the 95% confidence level) is 10% for the radial surface and 20% for the 

top surface. bAway from air vent locations; relative error (at the 95% confidence level) is 5%. cF/I neutron dose rate ratio values are very similar for the 5- and 40-year decay times.
B-77

Table B.10 Maximum dose rate change at 1 m from a BWR storage cask

Topa Sidea Sideb Topa Sidea Sideb

Case # Scenario Fuel relocation region

Decay 

time 

(years)

F/I neutron dose rate 

ratioc

F/I gamma dose rate 

ratio

2 11% fuel rod 

failure

Assembly lower tie plate and 

nosepiece

5

1.02 2.16 1.05

0.56 1.78 0.97

40 0.82 2.73 1.00

5 Damaged Canister cavity bottom

5

0.22 3.23 0.92

0.15 5.58 1.09

40 0.40 9.58 1.31

7 Alignment 

changes See Figure B.9 (a)

5

0.77 1.19 0.99

0.54 1.28 1.01

40 0.73 1.13 0.99

aFacing air vent locations; relative error (at the 95% confidence level) is 10% for the radial surface and 20% for the 

top surface. bAway from air vent locations; relative error (at the 95% confidence level) is 6%. cF/I neutron dose rate ratio values are very similar for the 5- and 40-year decay times.
B-78

Top surfacea,b Side surfacea,b Side su

a

PWR assemblies of 65 GWd/MTU burnup and 40-year decay time.

b

Dose rate profiles in a vertical plane through air vent locations.

c

Dose rate profiles in a vertical plane not through air vent locations.

Figure B.57 Neutron and gamma dose rate profiles at 1 m from a PWR storage cask assuming 10% fue

into assembly lower region
B-79

B.5.6 Generic Storage Cask Array

A site boundary dose rate comparison was performed based on a 4×2 array of storage casks 

using intact and damaged SNF conditions. The site boundary was assumed to be 100 m from 

the cask array (10 CFR 72.106[b]). The model for the cask array and cask internal configuration 

of the nominal intact configuration for the PWR fuel is illustrated in Figure B.4. A similar model 

was used for the BWR fuel, except the active fuel region was subdivided into 25 axial regions to 

facilitate the description of the BWR gamma and neutron radiation source axial profiles. The 

impact of changes in fuel configuration was evaluated for the damaged SNF case with fuel

rubble collected into the canister cavity bottom based on a 0.58 mass packing fraction (see 

Sect. B.4.1), and for the case with fuel rubble homogeneously distributed within the canister 

cavity. 

Table B.11 presents the ratio of site boundary dose rate for the cask array with failed fuel to the 

site boundary dose rate for the cask array with intact fuel for the PWR and BWR fuel, which is 

identified as F/I dose rate ratio. The results showed that closely packed fuel rubble reduces the 

site boundary dose rate by ~80% for gamma radiation and ~40 % for neutron radiation, whereas 

fuel rubble dispersed within the entire canister cavity increases the site boundary dose rate ~2.4 

times for gamma radiation and ~2.7 times for neutron radiation relative to the nominal intact fuel 

configuration. The dose rate changes are caused by source geometry changes and gamma 

self-shielding effects associated with the different configurations.

Table B.11 Site boundary dose rate change because of fuel configuration changes

Case description

F/I gamma 

dose rate 

ratioa

F/I neutron 

dose rate 

ratioa

PWR fuel rubble collected into canister cavity bottom 0.22 ± 0.03 0.58 ± 0.07

BWR fuel rubble collected into canister cavity bottom 0.30 ± 0.01 0.65 ± 0.01

PWR fuel rubble homogeneously distributed within canister cavity 1.75 ± 0.12 2.20 ± 0.08

BWR fuel rubble homogeneously distributed within canister cavity 2.41 ± 0.05 2.74 ± 0.04 aF/I dose rate ratio ± 2 sigma statistical error.

B.5.7 O-ring Locations

Lid, vent port, and drain port o-rings may be installed at cask top and bottom regions. O-rings 

serve as seals that provide primary containment. Fuel configuration changes may increase 

radiation dose rate and temperature at the o-ring locations, affecting the safety of the package. 

The F/I ratio values were calculated at typical vent and drain port o-ring locations within the 

package top lid and lower region, respectively, and are applicable to both NCT and HAC. The 

F/I ratio values for the PWR and BWR packages are provided in Table B.12 and Table B.13, 

respectively.
B-80

Table B.12 PWR package dose rate change at o-ring locations

Top Bottom Top Bottom

Case # Scenario Fuel relocation regiona

Decay

time 

(years)

F/I neutron dose rate 

ratiob,c

F/I gamma dose rate 

ratioc

1 10% fuel rod 

failure Assembly active fuel

5

1.07 1.02

1.02 1.01

40 0.98 1.00

2 10% fuel rod 

failure Assembly lower end fitting

5

1.05 1.97

1.02 1.23

40 1.00 4.33

3 10% fuel rod 

failure Assembly plenum 

5

3.29 1.02

1.02 1.00

40 3.30 0.99

4 10% fuel rod 

failure Assembly upper end fitting

5

3.96 0.99

0.97 1.00

40 11.23 1.00

5 25% fuel rod 

failure Assembly active fuel

5

1.10 0.99

1.01 0.98

40 0.98 0.98

6 25% fuel rod 

failure Assembly lower end fitting

5

1.03 1.90

1.01 1.24

40 0.99 4.45

7 25% fuel rod 

failure Assembly plenum 

5

3.85 0.98

1.00 0.98

40 3.31 0.98

8 25% fuel rod 

failure Assembly upper end fitting

5

4.92 0.98

0.94 1.00

40 11.29 0.97

9 Damaged Canister cavity bottom

5

0.29 2.83

0.0002 3.16

40 0.0001 15.06

10 Damaged Entire canister cavity

5

5.68 2.33

1.25 2.66

40 17.93 13.27

11 Assembly 

deformation See Figure B.8 (a) 

5

1.26 1.19

1.01 1.01

40 0.95 1.02

12 Assembly 

deformation See Figure B.8 (b)

5

1.20 1.13

1.00 0.93

40 0.90 0.91

13 Alignment 

changes See Figure B.9 (a)

5

0.80 1.40

0.61 1.86

40 1.68 0.31

14 Alignment 

changes See Figure B.9 (b)

5

1.26 0.77

1.68 0.37

40 0.60 4.10

aSee Sect. B.5.1 for case description. bF/I neutron dose rate ratio values are very similar for the 5- and 40-year decay times. cRelative error (95% confidence level) less than 5% for all values except for the value for case #9, gamma dose rate 

on top surface, the relative error is 15%.
B-81

Table B.13 BWR package dose rate change at o-ring locations

Top Bottom Top Bottom

Case # Scenario Fuel relocation regiona

Decay 

time 

(years)

F/I neutron dose rate 

ratiob,c

F/I gamma dose rate 

ratioc

1 10% fuel rod 

failure Assembly active fuel

5

1.12 1.05

1.01 1.00

40 0.99 0.99

2 10% fuel rod 

failure

Assembly lower tie plate and 

nosepiece

5

1.11 2.99

1.00 1.20

40 0.97 11.07

3 10% fuel rod 

failure Assembly plenum 

5

10.15 1.04

1.39 1.00

40 12.44 0.98

4 10% fuel rod 

failure

Assembly upper tie plate and 

handle

5

16.26 1.05

3.30 1.05

40 80.84 0.99

5 Damaged Canister cavity bottom

5

0.52 3.32

0.0004 2.22

40 0.02 30.71

6 Damaged Entire canister cavity

5

19.46 3.27

3.64 1.60

40 97.73 22.31

7 Assembly 

deformation See Figure B.8 (a)

5

1.34 1.19

1.00 1.01

40 0.96 1.01

8 Alignment 

changes See Figure B.9 (a)

5

0.88 1.20

0.69 1.55

40 0.73 2.05

9 Alignment 

changes See Figure B.9 (b)

5

1.13 0.85

1.44 0.57

40 1.40 0.57

aSee Sect. B.5.3 for case description. bF/I neutron dose rate ratio values are very similar for the 5- and 40-year decay times. cRelative error (95% confidence level) less than 5% for all values except for the value for case #9, gamma dose rate 

on top surface, the relative error is 15%.

C-1

APPENDIX C. CONTAINMENT EVALUATIONS FOR FUEL 

RECONFIGURATIONS

Containment is associated with the capability of a transportation package system to retain the 

radioactive material during transport within the limits established in 10 CFR 71.51 for normal 

conditions of transport (NCT) and hypothetical accident conditions (HAC). Current regulatory 

recommendations for the fractions of failed fuel rods and radioactive material release for use in 

the containment analysis of transportation systems have been based on data available for 

low-burnup fuel.

High-burnup fuel may have different characteristics than low-burnup fuel with respect to crud 

thickness, cladding oxide thickness and hydride content, radionuclide inventory and distribution, 

heat load, fuel grain size, fuel fragmentation, and fission gas release to the rod plenum. 

High-burnup fuel cladding properties such as reduced ductility because of hydriding and 

increased stiffness because of binding of the fuel to the cladding are of concern for the handling 

of high-burnup fuel after in-reactor irradiation. This study provides a brief review of the 

phenomena associated with high-burnup fuel that are relevant to the containment analysis. 

Evaluation of the applicability of the current regulatory recommendations to high-burnup fuel 

and long-term dry storage is necessary. However, determination of fuel failure rate and 

radionuclide release fractions for high-burnup fuel is outside the scope of this report.

Releasable activity, allowable radionuclide release rate, and allowable leakage rate for the 

generic pressurized water reactor (PWR) and boiling water reactor (BWR) transportation 

packages referred to as generic burnup credit-32 (GBC-32) and GBC-68, respectively, were 

calculated with the formulas presented in Sect. C.2. The fuel assemblies selected for analysis 

were representative of the Westinghouse (W) 17×17 optimized fuel assembly (OFA) and 

General Electric (GE) 14 assemblies with 40 and 65 GWd/MTU burnup values. For low-burnup 

(e.g., 40 GWd/MTU) fuel in long-term dry storage, releasable activities of gas, volatile, and fine 

species were calculated as a function of decay time and fraction of fuel rods that develop 

breach based on NUREG-1617, Table 4-1 [28] recommended release fractions for gases (0.3), 

volatiles (2×10-4

), and fuel fines (3×10-5) (see Table C.1). For high-burnup fuel (e.g., 

65 GWd/MTU), a sensitivity analysis is provided. Parameters important to the containment of 

failed fuel (i.e., the release fractions and failed fuel fraction) were varied, and their impact on the 

allowable leakage rate was evaluated. The impact on the allowable leakage rate of the 

discharged radioactive material original location (i.e., non-rim pellet region or the rim structure) 

was also evaluated. 

Table C.1 Release fractions for the contributors to the releasable PWR and BWR 

source terms

NCT HAC

Fraction of fuel rods that develop cladding breaches 0.03 1.0

Fraction of gases released due to a cladding breach, fGa 0.3 0.3

Fraction of volatiles released due to a cladding breach, fVa 2 × 10-4 2 × 10-4

Mass fraction of fuel released as fines due to cladding breach, 

fF

3 × 10-5 3 × 10-5

Fraction of crud that spalls off cladding, fCb 0.15 1.0 aIn accordance with NUREG/CR-6487 [40], gas species include 3H, 129I, 81Kr, 85Kr, and 127Xe; volatile species include 134Cs, 135Cs, 137Cs, 103Ru, 106Ru, 89Sr, and 90Sr. bThe source of radioactivity in crud is 60Co on fuel rods. At the time of discharge from the reactor, the crud surface

activity, SC, is estimated to be 140 μCi/cm2 for PWRs and 1254 μCi/cm2 for BWRs. Total 60Co activity is this estimate 

times the total surface area of all rods in the cask. Decay of 60Co to determine activity at the minimum time before 

loading is acceptable.
C-2

Currently applicable regulatory requirements and guidance used in the containment analysis are 

provided in Sect. C.1. The computer codes used to calculate the radiation source terms and 

mathematical formulas for calculating the quantities relevant to containment analyses are 

described in Sect. C.2. The parameters that were considered in this study are described in 

Sect. C.3. The calculation results for low-burnup and high-burnup fuel assemblies that develop 

rod breaches (Category 1) are provided in Sections C.4 and C.5, respectively. The calculation 

results for fuel assembly deformation (Categories 2 and 3) are provided in Section C.6. 

C.1 CURRENTLY APPLICABLE REGULATORY REQUIREMENTS 

AND GUIDANCE 

The regulatory activity limits for NCT and HAC are specified in terms of A2 quantities, as defined 

in 10 CFR Part 71.51 and Part 71–Appendix A. Following are the requirements for a spent fuel

package:

1. Loss or dispersal of radioactive contents should not exceed 10-6 A2 per hour under NCT.

2. Krypton-85 should not exceed 10 A2 in 1 week under HAC.

3. Other radioactive material should not exceed a total amount A2 in 1 week under HAC.

Current regulatory guidance for the fractions of failed fuel rods and radioactive material release 

to be used in the containment analysis of transportation systems are provided in NUREG-1617 

[29], and are summarized in Table C.1. A fraction of 0.03 is recommended for fuel rods that 

develop cladding breach under NCT. This fraction has been used in NUREG/CR-6487 [40] as a 

reasonable bounding value based on in-reactor and out-of-reactor cladding breach frequencies 

for fuel rods as of Dec. 31, 1986. The fraction of fuel rods assumed to fail as a result of HAC is 1

(i.e., 100% of the fuel rods). The release fractions for the gaseous, volatile, and fuel fine species 

have been justified in Ref. [13] based on the fuel test series described in NUREG/CR-0722 [41]

that had the objective to investigate fission products released in steam and dry air in the 

temperature range 500°C to 1200°C. The release fractions are typically used for low-burnup fuel 

because the maximum burnup of the fuel experimentally analyzed was ~30 GWd/MTU. The 

release fractions are also applicable to an inert atmosphere (i.e., not chemically reducing 

atmosphere as in the case of air ingress into the cask that would cause fuel oxidation and 

significant increases in the releases of fission product nuclides) [41]. The fuel tests 

demonstrated that burst release of the volatile species is significantly higher than release by 

diffusion. However, the diffusional release of cesium radionuclides is important for a rod break 

or other cladding failure that exposes a large amount of fuel directly to the cask environment 

[41]. 

Although the gap inventories of fission product nuclides appear to be identical, the release 

fractions of the radionuclides categorized as gaseous and volatile species are significantly 

different because of their different physico-chemical properties (e.g., boiling temperatures of 

elemental iodine and cesium are ~184°C and 678°C, respectively).

C.2 USE OF SOFTWARE AND APPLICABLE FORMULAS 

The radionuclide activities of the W 17×17 OFA and GE14 assemblies were calculated with the 

depletion and decay capabilities of the Scale 6.1.1 code system [31], including Transport Rigor 

Implemented with Time-Dependent Operation for Neutronic depletion (TRITON) and Oak Ridge 

Isotope Generation in Scale (ORIGEN-S). The Scale 238-group Evaluated Nuclear Data Files, 

Part B-VII.0 (ENDF/B-VII.0) nuclear data library was used in the TRITON depletion calculations.
C-3

Analysis results are provided for releasable source term, the effective A2 value for the total 

source term, allowable radionuclide release rate, and allowable leakage rate at operating 

conditions for the GBC-32 and GBC-68 casks described in the shielding analysis section. The 

calculation methodology described in NUREG/CR-6487 and the containment acceptance 

criteria provided in 10 CFR Part 71 were used in this study. Releasable activity for each 

individual contributor to the releasable source terms was calculated as:

RA f N A C C AC = ; (C.1)

RA f f N A G FFR G A G = ; (C.2)

RA f f N A V FFR V A V = ; (C.3)

RA f f N A F FFR F A F = ; (C.4)

where

RAC = releasable activity (Ci) as a result of crud spallation;

RAG = releasable activity (Ci) from gases as a result of cladding breach;

RAV = releasable activity (Ci) from volatile radionuclides as a result of cladding 

breach;

RAF = releasable activity (Ci) from fuel fines as a result of cladding breach;

NA = number of assemblies in the cask;

AC = crud activity (Ci/assembly);

AG = activity (Ci/assembly) of gaseous species;

AV = activity (Ci/assembly) of volatile species;

AF = activity of fuel fines (Ci/assembly);

fFFR = fraction of fuel rods that develop cladding breach;

 fG, fV, fF, and fC are the fractions for individual contributors to the releasable source term

(i.e., gases, volatiles, fuel fines, and crud, respectively, provided in Table C.1).

Crud primarily consists of 60Co with a half-life of ~5.3 years. Surface activity at the time of fuel 

discharge, which has been estimated to be 140 μCi/cm2 for PWRs and 1254 μCi/cm2 for BWRs 

(see Table C.1), was used to calculate crud releasable activity.

Total releasable activity was calculated as the sum of RAC, RAG, RAV, and RAF.

The effective A2 value for each individual contributor that consists of a mixture of radionuclides 

was calculated in accordance with 10 CFR 71, Appendix A:

=

=

∑

2

1 2

1

,mixture N

n

n ,n

A

f

A

,

(C.5)
C-4

where

A2,mixture = the A2 value (Ci) for each individual contributor to the releasable source term 

(i.e., gases, volatiles, and fuel fines);

fn = the fraction of activity of radionuclide n in the mixture for each individual 

contributor to the releasable source term;

A2,n = the A2 value (Ci) for radionuclide n, which is provided in 10 CFR71, 

Appendix A.

Allowable radionuclide release rates for NCT and HAC were determined using the containment 

requirements in 10 CFR 71.51(1) and 71.51(2), respectively. For NCT, “there would be no loss 

or dispersal of radioactive contents—as demonstrated to a sensitivity of 10-6 A2 per hour”

(10 CFR 71.51[1]). For HAC, “there would be no escape of krypton-85 exceeding 10 A2 in 1 

week, no escape of other radioactive material exceeding a total amount A2 in 1 week”

(10 CFR 71.51[2]). Hence, allowable radioactivity release rate for NCT, RRN (Ci/s), is 

RRN ≤ A2 × 2.78 × 10-10. Allowable radioactivity release rate for HAC, RRA (Ci/s), is 

RRA ≤ A2 × 1.65 × 10-6

; and the bounding value release rate for HAC is 4.46 × 10-3Ci/s (i.e., 10 

times the A2 value for 85Kr).

Allowable leakage rate, in cm3

/s, at operating conditions was calculated as: LRN = RRN/CN and 

LRA = RRA/CA, for NCT and for HAC, respectively, where CN and CA are the total source term 

activity concentration, in Ci/cm3

, for NCT and for HAC, respectively [40]. The free void volumes 

for the GBC-32 and GBC-68 cask models described in Appendix B were 7.06 × 106 cm3 and 

5.68 × 106 cm3

, respectively.

C.3 DESCRIPTION OF FUEL RECONFIGURATION CATEGORY

CALCULATION MODELS FOR CONTAINMENT

The containment analysis includes selected values for the fraction of fuel rods that may develop 

cladding breach. Currently, the applicability of the release fractions described in NUREG-1617 

to long-term storage and for high-burnup fuel is uncertain. In this respect, additional variations 

on the parameters important for containment analysis are included making this is a parametric 

analysis.

C.3.1 Category 1: Cladding Failure

Configurations evaluated:

a) Fuel assembly parameters: 

• 40 GWd/MTU burnup

o 3 wt % 235U initial enrichment 

o 5-, 40-, 100-, and 300-year decay times 

• 65 GWd/MTU burnup

o 5 wt % 235U initial enrichment 

o 5-, 40-, 100-, and 300-year decay times 

b) For the 40 GWd/MTU assembly burnup, the fraction of rods that develop cladding 

breach was varied from 0.01 to 1. The release fractions for gases, volatiles, and fuel 

fines from Table 4-1 in NUREG-1617, [28], are assumed to be applicable to longterm dry storage.

c) For the 65 GWd/MTU assembly burnup, the following parameters were varied: the 

fraction of rods that develop cladding breach (0.01, 0.03, 0.10, and 1); and the 
C-5

release fractions for crud (0.15, 0.3, 0.5, and 1), gases (0.1, 0.3, and 0.4), volatiles 

(2 × 10-4

, 1 × 10-3, and 2 × 10-3), and fuel fines (3 × 10-5, 1.5 × 10-4, and 3 × 10-4). In 

addition, the impact on the releasable activity of the radioactive material original 

location (i.e., inner pellet region or the rim structure) was evaluated. 

C.3.2 Category 2: Rod/Assembly Deformation

Rod/assembly deformation is assumed not to cause cladding breach. Crud is assumed to be 

released as a result of rod/assembly deformation. The allowable leakage rate was calculated as 

a function of the fraction of crud (0.05 to 1) that spalls off cladding for the 5- and 40-year decay 

times. The releasable activity for this category is independent of assembly burnup.

C.3.3 Category 3: Changes to Intact Assembly Axial Alignment 

Axial displacement of an intact assembly is assumed not to cause cladding breach. Crud is 

assumed to be released as a result of assembly axial displacement. The allowable leakage rate 

was calculated as a function of the fraction of crud (0.05 to 1) that spalls off cladding for the 5-

and 40-year decay times. The releasable activity for this category is independent of assembly 

burnup. 

C.4 LOW-BURNUP FUEL

A containment analysis for fuel rods that develop cladding failure for NCT was performed for the 

PWR GBC-32 and BWR GBC-68 transportation package models containing low-burnup (e.g., 

40 GWd/MTU) fuel assemblies. The total releasable activity, source term effective A2 values, 

allowable radionuclide release rates, and allowable leakage rates calculated as a function of 

fraction or rods that develop cladding breach and fuel decay time are provided in Table C.2 and 

Table C.3 for the GBC-32 and GBC-68 packages, respectively. The fraction of fuel rods that 

develop cladding breach under NCT was varied from 0.01 to 1. The value typically used for the 

fraction of fuel rods assumed to fail in NCT containment analyses, as recommended in NUREG1617 [28], is 0.03. Releasable activity for gas, volatile, and fine species was calculated as a 

function of decay time and fraction of fuel rods that develop breach assuming currently 

recommended release fractions for gases (0.3), volatiles (2×10-4

), and fines (3×10-5) (see

Table C.1). 

For the GBC-32 cask, the allowable leakage rate value based on the 0.03 fuel rod failure rate 

was ~1.0 × 10-4 cm3

/s for a 5-year decay time. An equivalent allowable leakage rate value was 

obtained for PWR fuel rod failure rates of ~0.10, 0.15, and 0.25, for 40-, 100-, and 300-year 

decay times, respectively. 

For the GBC-68 cask, the allowable leakage rate value based on the 0.03 fuel rod failure rate 

was ~2.4 × 10-5 cm3

/s for a 5-year decay time. An equivalent allowable leakage rate value was 

obtained for the BWR fuel rod failure rates of ~0.4, 0.7, and 1.0 for 40-, 100-, and 300-year 

decay times, respectively. With a more restrictive allowable leakage rate, the GBC-68 cask can 

accommodate greater fuel rod failure rates than the GBC-32. Crud contribution to GBC-68 

releasable activity is significantly higher than that of the GBC-32 cask for low decay times and a 

small fraction of failed fuel rods. As a result, the GBC-68 cask allowable leakage rates are 

smaller (i.e., more restrictive) than the GBC-32 cask allowable leakage rates for low decay 

times. However, the GBC-32 and GBC-68 total releasable activity and allowable leakage rates 

were similar for fuel decay times of 40, 100, and 300 years because the crud contribution effect 

becomes negligible at these decay times.
C-6

Table C.2 GBC-32 source term, allowable release rate, and allowable leakage rate: 

40 GWd/MTU

5-year decay time 40-year decay time

Failed 

fuel 

fraction

Releasable 

activity 

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

Releasable 

activity 

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

0.01 4.97E+02 3.99E+01 1.11E-08 1.57E-04 4.54E+01 2.49E+01 6.93E-09 1.08E-03

0.02 8.97E+02 5.62E+01 1.56E-08 1.23E-04 8.97E+01 2.53E+01 7.03E-09 5.53E-04

0.03 1.30E+03 6.66E+01 1.85E-08 1.01E-04 1.34E+02 2.54E+01 7.07E-09 3.72E-04

0.04 1.70E+03 7.38E+01 2.05E-08 8.52E-05 1.79E+02 2.55E+01 7.09E-09 2.80E-04

0.06 2.50E+03 8.32E+01 2.31E-08 6.53E-05 2.67E+02 2.55E+01 7.10E-09 1.87E-04

0.08 3.30E+03 8.90E+01 2.47E-08 5.29E-05 3.56E+02 2.56E+01 7.11E-09 1.41E-04

0.10 4.10E+03 9.30E+01 2.58E-08 4.45E-05 4.45E+02 2.56E+01 7.12E-09 1.13E-04

0.15 6.10E+03 9.89E+01 2.75E-08 3.18E-05 6.67E+02 2.56E+01 7.12E-09 7.53E-05

0.20 8.10E+03 1.02E+02 2.84E-08 2.47E-05 8.89E+02 2.56E+01 7.13E-09 5.65E-05

0.25 1.01E+04 1.04E+02 2.90E-08 2.02E-05 1.11E+03 2.56E+01 7.13E-09 4.53E-05

0.30 1.21E+04 1.06E+02 2.94E-08 1.71E-05 1.33E+03 2.57E+01 7.13E-09 3.77E-05

0.35 1.41E+04 1.07E+02 2.97E-08 1.49E-05 1.55E+03 2.57E+01 7.13E-09 3.23E-05

0.40 1.61E+04 1.08E+02 2.99E-08 1.31E-05 1.78E+03 2.57E+01 7.13E-09 2.83E-05

0.50 2.01E+04 1.09E+02 3.03E-08 1.06E-05 2.22E+03 2.57E+01 7.13E-09 2.27E-05

0.60 2.41E+04 1.10E+02 3.05E-08 8.92E-06 2.66E+03 2.57E+01 7.13E-09 1.89E-05

0.70 2.81E+04 1.10E+02 3.07E-08 7.69E-06 3.11E+03 2.57E+01 7.13E-09 1.62E-05

0.80 3.21E+04 1.11E+02 3.08E-08 6.76E-06 3.55E+03 2.57E+01 7.14E-09 1.42E-05

0.90 3.61E+04 1.11E+02 3.09E-08 6.03E-06 4.00E+03 2.57E+01 7.14E-09 1.26E-05

1.00 4.01E+04 1.11E+02 3.09E-08 5.44E-06 4.44E+03 2.57E+01 7.14E-09 1.13E-05

100-year decay time 300-year decay time

Failed 

fuel 

fraction

Releasable 

activity 

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

Releasable 

activity 

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

0.01 1.56E+00 1.31E+00 3.65E-10 1.65E-03 2.72E-02 3.72E-02 1.04E-11 2.68E-03

0.02 3.12E+00 1.31E+00 3.65E-10 8.27E-04 5.45E-02 3.72E-02 1.04E-11 1.34E-03

0.03 4.67E+00 1.31E+00 3.65E-10 5.51E-04 8.17E-02 3.72E-02 1.04E-11 8.93E-04

0.04 6.23E+00 1.31E+00 3.65E-10 4.13E-04 1.09E-01 3.72E-02 1.04E-11 6.70E-04

0.06 9.35E+00 1.31E+00 3.65E-10 2.76E-04 1.63E-01 3.72E-02 1.04E-11 4.47E-04

0.08 1.25E+01 1.31E+00 3.65E-10 2.07E-04 2.18E-01 3.72E-02 1.04E-11 3.35E-04

0.10 1.56E+01 1.31E+00 3.65E-10 1.65E-04 2.72E-01 3.72E-02 1.04E-11 2.68E-04

0.15 2.34E+01 1.31E+00 3.65E-10 1.10E-04 4.08E-01 3.72E-02 1.04E-11 1.79E-04

0.20 3.12E+01 1.31E+00 3.65E-10 8.27E-05 5.45E-01 3.72E-02 1.04E-11 1.34E-04

0.25 3.89E+01 1.31E+00 3.65E-10 6.62E-05 6.81E-01 3.72E-02 1.04E-11 1.07E-04

0.30 4.67E+01 1.31E+00 3.65E-10 5.51E-05 8.17E-01 3.72E-02 1.04E-11 8.93E-05

0.35 5.45E+01 1.31E+00 3.65E-10 4.73E-05 9.53E-01 3.72E-02 1.04E-11 7.66E-05

0.40 6.23E+01 1.31E+00 3.65E-10 4.13E-05 1.09E+00 3.72E-02 1.04E-11 6.70E-05

0.50 7.79E+01 1.31E+00 3.65E-10 3.31E-05 1.36E+00 3.72E-02 1.04E-11 5.36E-05

0.60 9.35E+01 1.31E+00 3.65E-10 2.76E-05 1.63E+00 3.72E-02 1.04E-11 4.47E-05

0.70 1.09E+02 1.31E+00 3.65E-10 2.36E-05 1.91E+00 3.72E-02 1.04E-11 3.83E-05

0.80 1.25E+02 1.31E+00 3.65E-10 2.07E-05 2.18E+00 3.72E-02 1.04E-11 3.35E-05

0.90 1.40E+02 1.31E+00 3.65E-10 1.84E-05 2.45E+00 3.72E-02 1.04E-11 2.98E-05

1.00 1.56E+02 1.31E+00 3.65E-10 1.65E-05 2.72E+00 3.72E-02 1.04E-11 2.68E-05 aTotal releasable activity of crud and fuel fission gas, volatile, and fine species. 
C-7

Table C.3 GBC-68 source term, allowable release rate, and allowable leakage rate: 

40 GWd/MTU

5-year decay time 40-year decay time

Failed 

fuel 

fraction

Releasable 

activity 

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

Releasable 

activity 

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

0.01 1.04E+03 1.58E+01 4.39E-09 2.67E-05 4.70E+01 2.17E+01 6.03E-09 8.14E-04

0.02 1.40E+03 2.03E+01 5.64E-09 2.55E-05 8.73E+01 2.35E+01 6.54E-09 4.76E-04

0.03 1.77E+03 2.44E+01 6.78E-09 2.44E-05 1.27E+02 2.43E+01 6.75E-09 3.36E-04

0.04 2.13E+03 2.82E+01 7.83E-09 2.34E-05 1.68E+02 2.47E+01 6.87E-09 2.60E-04

0.06 2.86E+03 3.48E+01 9.68E-09 2.15E-05 2.48E+02 2.51E+01 6.99E-09 1.79E-04

0.08 3.58E+03 4.05E+01 1.13E-08 2.00E-05 3.29E+02 2.54E+01 7.06E-09 1.36E-04

0.10 4.31E+03 4.55E+01 1.26E-08 1.86E-05 4.09E+02 2.55E+01 7.10E-09 1.10E-04

0.15 6.12E+03 5.53E+01 1.54E-08 1.59E-05 6.10E+02 2.57E+01 7.15E-09 7.44E-05

0.20 7.94E+03 6.26E+01 1.74E-08 1.39E-05 8.11E+02 2.58E+01 7.18E-09 5.62E-05

0.25 9.75E+03 6.83E+01 1.90E-08 1.24E-05 1.01E+03 2.59E+01 7.19E-09 4.51E-05

0.30 1.16E+04 7.29E+01 2.03E-08 1.11E-05 1.21E+03 2.59E+01 7.21E-09 3.77E-05

0.35 1.34E+04 7.66E+01 2.13E-08 1.01E-05 1.41E+03 2.59E+01 7.21E-09 3.24E-05

0.40 1.52E+04 7.97E+01 2.22E-08 9.26E-06 1.62E+03 2.60E+01 7.22E-09 2.84E-05

0.50 1.88E+04 8.46E+01 2.35E-08 7.93E-06 2.02E+03 2.60E+01 7.23E-09 2.27E-05

0.60 2.25E+04 8.82E+01 2.45E-08 6.94E-06 2.42E+03 2.60E+01 7.23E-09 1.90E-05

0.70 2.61E+04 9.11E+01 2.53E-08 6.16E-06 2.82E+03 2.60E+01 7.24E-09 1.63E-05

0.80 2.97E+04 9.33E+01 2.59E-08 5.55E-06 3.23E+03 2.60E+01 7.24E-09 1.43E-05

0.90 3.33E+04 9.52E+01 2.65E-08 5.04E-06 3.63E+03 2.61E+01 7.24E-09 1.27E-05

1.00 3.70E+04 9.67E+01 2.69E-08 4.62E-06 4.03E+03 2.61E+01 7.25E-09 1.14E-05

100-year decay time 300-year decay time

Failed 

fuel 

fraction

Releasable 

activity 

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

Releasable 

activity

(Ci)a

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

0.01 1.41E+00 1.36E+00 3.77E-10 1.70E-03 2.39E-02 3.76E-02 1.05E-11 2.78E-03

0.02 2.82E+00 1.36E+00 3.77E-10 8.49E-04 4.78E-02 3.76E-02 1.05E-11 1.39E-03

0.03 4.23E+00 1.36E+00 3.77E-10 5.66E-04 7.17E-02 3.76E-02 1.05E-11 9.26E-04

0.04 5.63E+00 1.36E+00 3.77E-10 4.25E-04 9.56E-02 3.76E-02 1.05E-11 6.95E-04

0.06 8.45E+00 1.36E+00 3.77E-10 2.83E-04 1.43E-01 3.76E-02 1.05E-11 4.63E-04

0.08 1.13E+01 1.36E+00 3.77E-10 2.12E-04 1.91E-01 3.76E-02 1.05E-11 3.47E-04

0.10 1.41E+01 1.36E+00 3.77E-10 1.70E-04 2.39E-01 3.76E-02 1.05E-11 2.78E-04

0.15 2.11E+01 1.36E+00 3.77E-10 1.13E-04 3.59E-01 3.76E-02 1.05E-11 1.85E-04

0.20 2.82E+01 1.36E+00 3.77E-10 8.49E-05 4.78E-01 3.76E-02 1.05E-11 1.39E-04

0.25 3.52E+01 1.36E+00 3.77E-10 6.79E-05 5.98E-01 3.76E-02 1.05E-11 1.11E-04

0.30 4.23E+01 1.36E+00 3.77E-10 5.66E-05 7.17E-01 3.76E-02 1.05E-11 9.26E-05

0.35 4.93E+01 1.36E+00 3.77E-10 4.85E-05 8.37E-01 3.76E-02 1.05E-11 7.94E-05

0.40 5.63E+01 1.36E+00 3.77E-10 4.25E-05 9.56E-01 3.76E-02 1.05E-11 6.95E-05

0.50 7.04E+01 1.36E+00 3.77E-10 3.40E-05 1.20E+00 3.76E-02 1.05E-11 5.56E-05

0.60 8.45E+01 1.36E+00 3.77E-10 2.83E-05 1.43E+00 3.76E-02 1.05E-11 4.63E-05

0.70 9.86E+01 1.36E+00 3.77E-10 2.43E-05 1.67E+00 3.76E-02 1.05E-11 3.97E-05

0.80 1.13E+02 1.36E+00 3.77E-10 2.12E-05 1.91E+00 3.76E-02 1.05E-11 3.47E-05

0.90 1.27E+02 1.36E+00 3.77E-10 1.89E-05 2.15E+00 3.76E-02 1.05E-11 3.09E-05

1.00 1.41E+02 1.36E+00 3.77E-10 1.70E-05 2.39E+00 3.76E-02 1.05E-11 2.78E-05 aTotal releasable activity of crud and fuel fission gas, volatile, and fine species.
C-8

C.5 HIGH-BURNUP FUEL

Physical phenomena relevant to the containment analysis of high-burnup fuel are described in 

Sect. C.5.1. A description of the high-burnup depletion calculation models for the W 17×17 OFA 

and GE14 assemblies is provided in Sect. C.5.2 and the Scale 6.1.1-calculated radionuclide 

activities are provided in Sect. C.5.3. The containment analysis for high-burnup fuel is provided 

in Sect. C.5.4. 

C.5.1 Physical Phenomena Relevant to Containment Analysis of High-Burnup Fuel

A limited review of publicly available documents addressing the properties of high-burnup fuel 

assemblies relevant to spent fuel management was performed for this study. Major differences 

between high- and low-burnup light water reactor fuels concern radiation source term, pressure 

of fission gas in the fuel cladding gap, cladding mechanical properties, and the fuel rim-effect 

associated only with high-burnup fuel. These characteristics affect the mechanisms by which the 

fuel can breach and the amount of fuel that can be released from failed fuel rods. Hence, the fuel 

rod failure rate for NCT as well as radionuclide release fractions based on low-burnup studies 

may need to be reevaluated for their applicability to high- burnup fuel, which is outside the scope 

of this report. Major high-burnup fuel physical characteristics relevant to containment analysis 

are summarized below. 

The mechanical properties of high-burnup fuel cladding can be adversely affected by embrittling 

factors such as in-reactor irradiation damage, oxidation, and circumferential hydriding (i.e. uptake 

of hydrogen from corrosion in the coolant). The cladding of discharged high-burnup PWR fuel 

contains an oxide coating up to 100 microns and between 50 and 700 parts per million by weight

(wppm) hydrogen, depending on the zirconium alloy composition; in comparison, the cladding of 

low-burnup PWR fuel has less than 100 wppm hydrogen. The cladding of high-burnup BWR fuel 

contains a thin uniform oxide layer with nodular corrosion of larger extent and lower hydrogen 

concentrations than the high-burnup PWR fuel [42]. Fission gas released from the fuel matrix to 

the fuel cladding gap is significantly larger in high-burnup fuel rods compared to the low-burnup 

fuel rods [43]. Excessive hoop stress because of higher internal pressurization and thinner 

cladding may cause cladding to be more susceptible to failure. In addition, the formation of radial 

hydrides can result in reduced ductility and tensile strength of the cladding [11].

The properties of high-burnup fuel pellets are more heterogeneous than for low-burnup fuel. 

Higher burnup (>55 GWd/MTU) and in-reactor fuel temperatures below ~1100 ̊C cause a 

restructuring of the UO2 matrix in the pellet outer region [39, 44]. The resulting microstructure is 

usually referred to as a high-burnup rim structure. The rim structure is characterized by submicron grains, which are depleted of fission gas, and much higher closed porosity (15 to 20%) 

than the bulk material (6%), which contains fission gas at high pressures [42]. The inside of the 

closed porosity tends to be lined with many submicron-sized fuel particulates in the 10–200 nm 

range. Hence, fission gas and fuel particles in the respirable range would be released when the 

fission gas bubbles in the rim structure burst.

High-burnup fuel pellets exhibit greater fragmentation than low-burnup fuel pellets. Pellet 

fracturing is caused by fuel exposure to thermal gradients early in life. The number of fragments 

observed for typical fuel varies from 20 to 50 fragments per pellet; more fragmentation has been

observed for higher burnup fuel [39]. Cracks observed in the rim structure are caused by strong 

pellet-cladding bonding—because of inter-penetration between the rim structure and the 

zirconia formed at the cladding inner surface after closure of the pellet-to-cladding gap—and by
C-9

the differential thermal expansion of the pellet and cladding during the cooling phases [45]. 

Microcracks in the rim structure also may be caused by overpressurization within the 

intergranular fission gas bubbles coupled with sudden relief of pellet cladding mechanical 

interaction restraints by cladding [46]. Greater fuel pellet fragmentation causes a greater amount 

of fuel particulates to be released from high-burnup failed fuel rods. 

Fuel fragmentation and fission gas release are correlated. The larger quantity of fission gas 

released to the fuel rod free-void volume increases fission gas pressure, which increases the 

risk of cladding breach. The fission gas bubbles formed in the central portion of a fuel pellet 

become interlinked with continued irradiation releasing the fission gas and volatile products to 

the fuel cladding gap. Fission gas release from the pellet to the fuel cladding gap increases with 

increasing fuel burnup (e.g., from 5 to 25% for burnup increases of 20 to 100 GWd/MTU 

burnup) [47]. Fission gas release from the rim structure has been evaluated to be generally low 

[45]. However, significantly more fission gas may be released under rapid heating conditions 

such as transient and off-normal irradiation conditions, as demonstrated by rapid heating tests, 

or because of the occurrence of microcracks upon the sudden relief of pellet cladding 

mechanical interaction restraints by cladding. Under certain heating conditions, such as those 

used in rapid or slow heating tests to very high temperatures (e.g., 1500-1800 ̊C), fission gas 

release fraction from high-burnup fuel (with burnup greater than 60 GWd/MTU) may exceed the 

0.3 value currently used in the containment analyses. Rapid heating tests that simulate transient 

and off-normal irradiation conditions have measured fission gas release fractions up to ~46% for 

a 74 GWd/MTU burnup fuel specimen [46]. 

The average amount of crud for high-burnup fuel assemblies has been estimated to be similar 

to that observed on lower burnup fuel assemblies [42].

C.5.2 Depletion Calculation Models

Neutron capture is more pronounced in the fuel pellet outer region. As a result, the power 

production at the center of a fuel rod is slightly less than that at the surface. As fuel burnup 

increases, the distributions of fissile material, fission rate, and fission products in irradiated fuel 

develop peaks at the pellet surface [48, 49]. A mathematical model, shown in Eq. (C.6), has 

been developed for the rim structure width as a function of pellet burnup based on best 

available data obtained by optical microscopy [42]. The rim width as a function of pellet burnup 

is illustrated in Figure C.1. Based on Eq. (C.6), the width of the rim structure corresponding to 

an assembly average burnup value of 65 GWd/MTU is approximately 150 microns. 

w = 1.439e-6 × burnup4.427, (C.6)

where

w = the rim width where the grains are fully transformed (microns), and

burnup = the pellet-average burnup in GWd/MTU. 
C-10

Figure C.1 Rim width as a function of pellet burnup

The TRITON 2-D depletion calculation models for the W 17×17 OFA and GE14 assembly with a 

5 wt % 235U initial enrichment and 65 GWd/MTU assembly average burnup are illustrated in 

Figure C.2 and Figure C.3, respectively. In these models, the UO2 fuel pellet consists of 14 

concentric rings, five of which are in the pellet rim structure. The simulated assembly operating 

conditions include: 

• four 12-month irradiation cycles with a 30 day down time period between the cycles,

• cycle-specific power values varying from 60 to 37 MW/MTU from cycle 1 to cycle 4, and

• fuel temperature variation as a function of pellet radial region and irradiation time.

Figure C.2 Illustration of the TRITON 2-D model for the PWR high-burnup assembly

0

200

400

600

800

1000

1200

0 20 40 60 80 100 120

Rim width (μm)

Pellet burnup (GWd/MTU)
C-11

Figure C.3 Illustration of the TRITON 2-D model for the BWR high-burnup assembly

C.5.3 Radionuclide Inventory

Within the rim structure, there is a pronounced pellet radial burnup increase near the pellet outer 

radius [50], as illustrated by the calculated 148Nd concentration variation as a function of pellet 

radial location for the W 17×17 OFA (Figure C.4). The isotope 148Nd is a burnup monitor used 

as a standard test method for atom fission in ASTM E 321 [51]. Pellet radial burnup variation 

results in significant variation in local nuclide concentrations primarily near the pellet periphery. 

Figure C.4 Variation of 148Nd concentration as a function of pellet radial location

6.0E+02

7.0E+02

8.0E+02

9.0E+02

1.0E+03

1.1E+03

1.2E+03

0 1000 2000 3000 4000

Nd-148 (g/MTU)

Pellet radius (microns)

Varying axial burnup

Average burnup
C-12

The calculated radionuclide activity values, in Ci per assembly, for W 17×17 OFA and GE14 fuel 

assemblies with a 65 GWd/MTU assembly average burnup value are tabulated in Table C.4 and

Table C.5, respectively, as a function of fuel decay time. The radionuclide activity values are 

separately shown for the crud, radioactive gas species, radioactive volatile species, and fuel 

fines. Crud was assumed to consist of 60Co and the total 60Co activity was calculated as the 

surface activity times the total surface area of all rods in the cask. The average amount of crud 

formation on the fuel pin surface for high-burnup fuel assemblies was estimated with currently 

recommended 60Co surface activities of 140 μCi/cm2 for PWRs and 1254 μCi/cm2 for BWRs [28]

at the time of discharge from the reactor. 

Radionuclide activity, in Ci per unit heavy metal mass (e.g., one metric ton uranium, or 1 MTU), 

is approximately constant in the pellet inner region (rings 1 through 8 of the pellet model) and 

then varies as a function of pellet radius because of varying pellet radial burnup. This variation 

is illustrated by the values provided in Table C.6 and Table C.7 for the W 17×17 OFA and the 

decay times of 5 and 300 years, respectively. The tables present the ratio of radionuclide 

activity in a pellet radial region (in Ci per MTU), to pellet radionuclide activity (in Ci per MTU). 

As can be seen in the table, radionuclide activity per unit heavy metal mass in the pellet 

periphery region is significantly greater than the pellet-averaged radionuclide activity for the 

majority of radionuclides (e.g., ~3.5 times as much for 252Cf). Similar values were obtained for 

the GE14 assembly with 65 GWd/MTU burnup.
D R A F T – 8/21/13

C-13

Table C.4 Radionuclide activities (Ci/assembly) for W 17×17 OFA with 65 GWd/MTU 

burnup

Nuclide/ decay 

time 5 years 40 years 100 years 300 years

Crud (60Co) activity (Ci/assembly)

2.01E+01 2.02E-01 7.55E-05 0.00E+00

Total fuel activity (Ci/assembly)

4.00E+05 1.30E+05 3.26E+04 3.25E+03

Gas species (Ci/assembly)

3H 2.95E+02 4.11E+01 1.41E+00 1.82E-05

129I 2.30E-02 2.30E-02 2.30E-02 2.30E-02

81Kr 4.38E-07 4.38E-07 4.38E-07 4.38E-07

85Kr 6.21E+03 6.51E+02 1.36E+01 3.44E-05

127Xe 1.02E-19 0.00E+00 0.00E+00 0.00E+00

Total 6.51E+03 6.92E+02 1.50E+01 2.30E-02

Volatile species (Ci/assembly)

134Cs 2.20E+04 1.74E-01 3.12E-10 0.00E+00

135Cs 3.42E-01 3.42E-01 3.42E-01 3.42E-01

137Cs 7.55E+04 3.37E+04 8.45E+03 8.40E+01

103Ru 1.18E-11 0.00E+00 0.00E+00 0.00E+00

106Ru 6.30E+03 3.14E-07 6.86E-25 0.00E+00

89Sr 2.74E-08 0.00E+00 0.00E+00 0.00E+00

90Sr 5.00E+04 2.15E+04 5.07E+03 4.11E+01

Total 1.54E+05 5.52E+04 1.35E+04 1.25E+02

Fuel finesa (Ci/assembly)

225Ac 3.82E-07 5.60E-07 1.24E-06 7.85E-06

227Ac 2.78E-06 1.23E-05 2.23E-05 4.91E-05

241Am 8.87E+02 2.63E+03 2.76E+03 2.02E+03

242mAm 4.75E+00 4.00E+00 2.98E+00 1.11E+00

242Am 4.72E+00 3.98E+00 2.96E+00 1.11E+00

243Am 3.09E+01 3.08E+01 3.07E+01 3.01E+01

137mBa 7.13E+04 3.18E+04 7.97E+03 7.93E+01

247Bk 1.49E-11 1.47E-11 1.42E-11 1.29E-11

144Ce 2.82E+03 8.76E-11 0.00E+00 0.00E+00

248Cf 3.95E-09 1.14E-20 0.00E+00 0.00E+00

249Cf 8.86E-05 8.34E-05 7.40E-05 4.99E-05

250Cf 7.67E-04 1.20E-04 4.99E-06 1.31E-10

251Cf 6.84E-06 6.65E-06 6.35E-06 5.44E-06

252Cf 3.97E-04 4.12E-08 6.10E-15 0.00E+00

242Cm 8.21E+00 3.29E+00 2.45E+00 9.16E-01

243Cm 1.89E+01 8.20E+00 1.97E+00 1.68E-02

244Cm 5.03E+03 1.32E+03 1.32E+02 6.23E-02

245Cm 1.13E+00 1.12E+00 1.12E+00 1.10E+00

246Cm 3.21E-01 3.20E-01 3.17E-01 3.08E-01

247Cm 1.57E-06 1.57E-06 1.57E-06 1.57E-06

248Cm 6.04E-06 6.04E-06 6.04E-06 6.04E-06

154Eu 3.85E+03 2.29E+02 1.81E+00 1.78E-07

155Eu 1.51E+03 9.15E+00 1.45E-03 3.11E-16

93mNb 4.17E-01 1.18E+00 1.38E+00 1.39E+00

236Np 2.98E-06 2.98E-06 2.98E-06 2.97E-06

237Np 2.95E-01 3.18E-01 3.72E-01 5.25E-01

239Np 3.09E+01 3.08E+01 3.07E+01 3.01E+01

231Pa 1.43E-05 1.89E-05 2.68E-05 5.33E-05
D R A F T – 8/21/13

C-14

Nuclide/ decay 

time 5 years 40 years 100 years 300 years

233Pa 2.95E-01 3.18E-01 3.72E-01 5.25E-01

147Pm 2.02E+04 1.94E+00 2.53E-07 2.82E-30

210Po 8.89E-08 8.34E-07 9.24E-06 1.63E-04

144Pr 2.82E+03 8.76E-11 0.00E+00 0.00E+00

236Pu 3.02E-01 6.25E-05 3.72E-07 3.72E-07

238Pu 3.65E+03 2.77E+03 1.72E+03 3.56E+02

239Pu 1.77E+02 1.76E+02 1.76E+02 1.75E+02

240Pu 3.43E+02 3.52E+02 3.53E+02 3.46E+02

241Pu 6.88E+04 1.26E+04 6.87E+02 1.14E+00

242Pu 2.05E+00 2.05E+00 2.05E+00 2.05E+00

244Pu 5.73E-07 5.73E-07 5.73E-07 5.73E-07

223Ra 2.78E-06 1.24E-05 2.23E-05 4.91E-05

224Ra 4.50E-02 5.17E-02 2.83E-02 3.78E-03

225Ra 3.82E-07 5.60E-07 1.24E-06 7.85E-06

226Ra 9.79E-08 2.49E-06 1.72E-05 2.06E-04

228Ra 2.05E-10 6.11E-10 1.19E-09 3.23E-09

106Rh 6.30E+03 3.14E-07 6.86E-25 0.00E+00

222Rn 9.79E-08 2.49E-06 1.72E-05 2.06E-04

125Sb 1.54E+03 2.33E-01 6.58E-08 9.81E-30

151Sm 1.85E+02 1.41E+02 8.88E+01 1.90E+01

121mSn 1.13E+01 7.24E+00 3.40E+00 2.73E-01

99Tc 1.07E+01 1.07E+01 1.06E+01 1.06E+01

125mTe 3.77E+02 5.72E-02 1.62E-08 0.00E+00

227Th 2.75E-06 1.22E-05 2.20E-05 4.84E-05

228Th 4.50E-02 5.17E-02 2.83E-02 3.78E-03

229Th 3.82E-07 5.60E-07 1.24E-06 7.85E-06

230Th 4.69E-05 2.87E-04 8.93E-04 3.79E-03

231Th 6.25E-03 6.26E-03 6.27E-03 6.30E-03

232U 5.84E-02 5.03E-02 2.75E-02 3.68E-03

233U 3.11E-05 7.75E-05 1.68E-04 5.61E-04

234U 5.82E-01 8.98E-01 1.27E+00 1.76E+00

236U 1.96E-01 1.96E-01 1.97E-01 1.99E-01

90Y 5.00E+04 2.15E+04 5.07E+03 4.11E+01

93Zr 1.43E+00 1.43E+00 1.43E+00 1.43E+00

Total 2.40E+05 7.36E+04 1.91E+04 3.12E+03 aNuclides with individual activity greater than 10-4 times total assembly activity and nuclides with 

A2 (Ci) values less than 1; 240Cm, 210mBi, 254Cf, 148Gd, and 230U are not shown because these 

nuclides have zero or negligible activities. 

Table C.4 (continued)
C-15

Table C.5 Radionuclide activities (Ci/assembly) for the GE14 assembly with 

65 GWd/MTU burnup

Nuclide/ decay 

time 5 years 40 years 100 years 300 years

Crud (60Co) activity (Ci/assembly)

6.65E+01 6.67E-01 2.50E-04 0.00

Total fuel activity (Ci/assembly)

1.70E+05 5.49E+04 1.38E+04 1.38E+03

Gas species (Ci/assembly)

3H 1.25E+02 1.74E+01 5.95E-01 7.71E-06

129I 9.74E-03 9.74E-03 9.74E-03 9.74E-03

81Kr 2.13E-07 2.13E-07 2.12E-07 2.12E-07

85Kr 2.63E+03 2.76E+02 5.77E+00 1.46E-05

127Xe 5.40E-20 0.00E+00 0.00E+00 0.00E+00

Total 2.76E+03 2.93E+02 6.38E+00 9.76E-03

Volatile species (Ci/assembly)

134Cs 9.58E+03 7.57E-02 1.36E-10 0.00E+00

135Cs 1.50E-01 1.50E-01 1.50E-01 1.50E-01

137Cs 3.20E+04 1.43E+04 3.58E+03 3.56E+01

103Ru 5.01E-12 0.00E+00 0.00E+00 0.00E+00

106Ru 2.65E+03 1.32E-07 2.89E-25 0.00E+00

89Sr 1.16E-08 0.00E+00 0.00E+00 0.00E+00

90Sr 2.12E+04 9.11E+03 2.15E+03 1.74E+01

Total 6.54E+04 2.34E+04 5.73E+03 5.31E+01

Fuel Finesa (Ci/assembly)

225Ac 1.85E-07 2.59E-07 5.43E-07 3.27E-06

227Ac 1.23E-06 5.41E-06 9.70E-06 2.11E-05

241Am 3.69E+02 1.09E+03 1.15E+03 8.40E+02

242mAm 2.05E+00 1.72E+00 1.28E+00 4.80E-01

242Am 2.04E+00 1.71E+00 1.28E+00 4.77E-01

243Am 1.26E+01 1.25E+01 1.24E+01 1.22E+01

137mBa 3.02E+04 1.35E+04 3.38E+03 3.36E+01

247Bk 1.17E-11 1.15E-11 1.12E-11 1.01E-11

144Ce 1.20E+03 3.71E-11 0.00E+00 0.00E+00

248Cf 3.20E-09 9.23E-21 0.00E+00 0.00E+00

249Cf 6.48E-05 6.10E-05 5.42E-05 3.65E-05

250Cf 5.30E-04 8.29E-05 3.45E-06 9.10E-11

251Cf 5.12E-06 4.98E-06 4.76E-06 4.08E-06

252Cf 2.89E-04 3.00E-08 4.44E-15 0.00E+00

242Cm 3.42E+00 1.42E+00 1.06E+00 3.95E-01

243Cm 8.51E+00 3.69E+00 8.85E-01 7.55E-03

244Cm 2.26E+03 5.90E+02 5.93E+01 2.79E-02

245Cm 5.83E-01 5.81E-01 5.78E-01 5.69E-01

246Cm 1.65E-01 1.64E-01 1.62E-01 1.58E-01

247Cm 9.34E-07 9.34E-07 9.34E-07 9.34E-07

248Cm 3.76E-06 3.76E-06 3.76E-06 3.76E-06

154Eu 1.74E+03 1.03E+02 8.15E-01 8.02E-08

155Eu 6.41E+02 3.89E+00 6.16E-04 1.32E-16

93mNb 1.76E-01 4.98E-01 5.83E-01 5.90E-01

236Np 1.38E-06 1.37E-06 1.37E-06 1.37E-06

237Np 1.21E-01 1.30E-01 1.53E-01 2.17E-01

239Np 1.26E+01 1.25E+01 1.24E+01 1.22E+01

231Pa 6.29E-06 8.26E-06 1.16E-05 2.29E-05
C-16

Nuclide/ decay 

time 5 years 40 years 100 years 300 years

233Pa 1.21E-01 1.30E-01 1.53E-01 2.17E-01

147Pm 8.71E+03 8.39E-01 1.09E-07 1.22E-30

210Po 4.24E-08 3.48E-07 3.90E-06 6.98E-05

144Pr 1.20E+03 3.71E-11 0.00E+00 0.00E+00

236Pu 1.35E-01 2.79E-05 1.72E-07 1.72E-07

238Pu 1.62E+03 1.23E+03 7.64E+02 1.58E+02

239Pu 7.85E+01 7.85E+01 7.84E+01 7.80E+01

240Pu 1.59E+02 1.63E+02 1.63E+02 1.60E+02

241Pu 2.86E+04 5.24E+03 2.86E+02 5.87E-01

242Pu 7.85E-01 7.85E-01 7.85E-01 7.85E-01

244Pu 2.62E-07 2.62E-07 2.62E-07 2.62E-07

223Ra 1.23E-06 5.42E-06 9.70E-06 2.11E-05

224Ra 2.01E-02 2.31E-02 1.26E-02 1.69E-03

225Ra 1.85E-07 2.59E-07 5.43E-07 3.27E-06

226Ra 3.94E-08 1.04E-06 7.27E-06 8.85E-05

228Ra 8.53E-11 2.60E-10 5.15E-10 1.41E-09

106Rh 2.65E+03 1.32E-07 2.89E-25 0.00E+00

222Rn 3.94E-08 1.04E-06 7.27E-06 8.85E-05

125Sb 6.58E+02 9.96E-02 2.82E-08 4.20E-30

151Sm 8.42E+01 6.43E+01 4.05E+01 8.68E+00

121mSn 4.84E+00 3.12E+00 1.46E+00 1.18E-01

99Tc 4.52E+00 4.52E+00 4.52E+00 4.52E+00

125mTe 1.62E+02 2.45E-02 6.93E-09 0.00E+00

227Th 1.21E-06 5.34E-06 9.57E-06 2.08E-05

228Th 2.01E-02 2.31E-02 1.26E-02 1.69E-03

229Th 1.85E-07 2.59E-07 5.43E-07 3.27E-06

230Th 1.91E-05 1.21E-04 3.81E-04 1.64E-03

231Th 2.66E-03 2.66E-03 2.66E-03 2.68E-03

232U 2.61E-02 2.25E-02 1.23E-02 1.64E-03

233U 1.34E-05 3.24E-05 6.93E-05 2.31E-04

234U 2.43E-01 3.83E-01 5.48E-01 7.65E-01

236U 8.57E-02 8.59E-02 8.62E-02 8.72E-02

90Y 2.12E+04 9.12E+03 2.15E+03 1.74E+01

93Zr 6.05E-01 6.05E-01 6.05E-01 6.05E-01

Total 1.02E+05 3.12E+04 8.11E+03 1.33E+03 aNuclides with individual activity greater than 10-4 times total assembly activity and nuclides with 

A2 (Ci) values less than 1; 240Cm, 210mBi, 254Cf, 148Gd, and 230U are not shown because these 

nuclides have zero or negligible activities.

Table C.5 (continued)
C-17

Table C.6 Pellet radial region radionuclide activity (Ci/MTU) relative to average pellet 

radionuclide activity (Ci/MTU) for a 5-year decay time

Pellet ring number 1-8 9 10 11 12 13 14

Volume fraction 0.8 0.0682 0.0264 0.0264 0.0264 0.0264 0.0264

Gas species

3H 0.95 1.02 1.07 1.12 1.20 1.36 1.72

129I 0.94 1.02 1.07 1.13 1.23 1.41 1.84

81Kr 0.94 1.05 1.10 1.16 1.26 1.43 1.85

85Kr 0.98 1.02 1.04 1.06 1.10 1.17 1.33

127Xe 0.94 1.04 1.09 1.15 1.25 1.43 1.85

Total 0.97 1.02 1.04 1.06 1.10 1.18 1.35

Volatile species

134Cs 0.95 1.06 1.10 1.14 1.21 1.33 1.62

135Cs 0.97 0.99 1.03 1.07 1.14 1.26 1.56

137Cs 0.96 1.02 1.06 1.10 1.18 1.31 1.62

103Ru 0.93 1.03 1.10 1.17 1.30 1.53 2.09

106Ru 0.91 1.04 1.11 1.20 1.35 1.62 2.27

89Sr 0.97 1.00 1.03 1.07 1.14 1.26 1.56

90Sr 0.98 1.02 1.03 1.06 1.09 1.15 1.30

Total 0.96 1.03 1.06 1.10 1.16 1.27 1.54

Fuel fines

225Ac 1.00 1.01 1.01 1.02 1.02 1.02 1.03

227Ac 1.01 0.97 0.97 0.97 0.97 0.96 0.96

241Am 0.90 1.04 1.12 1.23 1.40 1.71 2.45

242mAm 0.91 1.02 1.10 1.20 1.37 1.68 2.42

242Am 0.91 1.02 1.10 1.20 1.37 1.68 2.42

243Am 0.88 1.08 1.18 1.30 1.49 1.85 2.69

137mBa 0.96 1.02 1.06 1.10 1.18 1.31 1.62

247Bk 0.84 1.16 1.28 1.43 1.65 2.06 3.06

144Ce 0.96 1.01 1.05 1.10 1.18 1.32 1.67

249Cf 0.84 1.16 1.29 1.43 1.66 2.07 3.07

250Cf 0.83 1.18 1.31 1.46 1.69 2.12 3.14

251Cf 0.83 1.18 1.31 1.45 1.68 2.10 3.11

252Cf 0.81 1.21 1.35 1.51 1.76 2.20 3.27

254Cf 0.79 1.26 1.42 1.59 1.85 2.33 3.48

242Cm 0.90 1.03 1.12 1.23 1.41 1.73 2.50

243Cm 0.89 1.06 1.16 1.27 1.46 1.80 2.61

244Cm 0.87 1.10 1.20 1.33 1.53 1.90 2.78

245Cm 0.87 1.09 1.20 1.32 1.52 1.89 2.77

246Cm 0.85 1.13 1.25 1.38 1.60 1.99 2.93

247Cm 0.85 1.14 1.26 1.39 1.61 2.00 2.96

248Cm 0.84 1.17 1.29 1.44 1.67 2.08 3.08

154Eu 0.95 1.02 1.06 1.12 1.21 1.37 1.77

155Eu 0.94 1.03 1.08 1.13 1.23 1.41 1.82

93mNb 0.97 1.02 1.04 1.07 1.12 1.20 1.40

236Np 1.01 0.97 0.97 0.97 0.97 0.97 0.97

237Np 0.99 1.03 1.04 1.05 1.05 1.06 1.07

239Np 0.88 1.08 1.18 1.30 1.49 1.85 2.69

231Pa 1.01 0.98 0.97 0.97 0.97 0.97 0.96

233Pa 0.99 1.03 1.04 1.05 1.05 1.06 1.07

147Pm 0.96 0.99 1.02 1.07 1.15 1.29 1.65

210Po 1.00 0.99 0.99 0.98 0.98 0.99 0.99
C-18

Table C.6 (continued)

Pellet ring number 1-8 9 10 11 12 13 14

Volume fraction 0.8 0.0682 0.0264 0.0264 0.0264 0.0264 0.0264

144Pr 0.96 1.01 1.05 1.10 1.18 1.32 1.67

236Pu 0.99 1.01 1.02 1.02 1.03 1.03 1.04

238Pu 0.97 1.04 1.07 1.09 1.13 1.19 1.33

239Pu 0.92 1.00 1.07 1.16 1.30 1.57 2.20

240Pu 0.97 0.91 0.96 1.02 1.13 1.34 1.82

241Pu 0.90 1.04 1.13 1.23 1.40 1.71 2.46

242Pu 0.88 1.07 1.17 1.29 1.48 1.83 2.65

244Pu 0.87 1.10 1.20 1.33 1.53 1.90 2.79

223Ra 1.01 0.97 0.97 0.97 0.97 0.96 0.96

224Ra 1.00 1.01 1.01 1.02 1.02 1.03 1.04

225Ra 1.00 1.01 1.01 1.02 1.02 1.02 1.03

226Ra 1.01 0.98 0.97 0.97 0.97 0.97 0.97

228Ra 1.00 1.01 1.01 1.01 1.02 1.02 1.02

106Rh 0.91 1.04 1.11 1.20 1.35 1.62 2.27

222Rn 1.01 0.98 0.97 0.97 0.97 0.97 0.97

125Sb 0.94 1.02 1.07 1.13 1.23 1.42 1.86

151Sm 0.95 1.01 1.05 1.11 1.21 1.39 1.81

121mSn 0.95 1.02 1.06 1.12 1.21 1.39 1.80

99Tc 0.96 1.01 1.04 1.09 1.16 1.29 1.60

125mTe 0.94 1.02 1.07 1.13 1.23 1.42 1.86

227Th 1.01 0.97 0.97 0.97 0.97 0.96 0.96

228Th 1.00 1.01 1.01 1.02 1.02 1.03 1.04

229Th 1.00 1.01 1.01 1.02 1.02 1.02 1.03

230Th 1.01 0.97 0.97 0.97 0.97 0.97 0.97

231Th 1.03 0.91 0.90 0.89 0.89 0.88 0.87

232U 1.00 1.01 1.01 1.02 1.02 1.03 1.04

233U 1.01 0.97 0.96 0.96 0.96 0.96 0.96

234U 1.00 0.98 0.98 0.98 0.98 0.98 1.00

236U 1.00 1.00 1.00 1.00 1.00 1.00 1.00

90Y 0.98 1.02 1.03 1.06 1.09 1.15 1.30

93Zr 0.97 1.02 1.04 1.07 1.12 1.21 1.42

Total 0.94 1.02 1.07 1.13 1.23 1.41 1.84
C-19

Table C.7 Pellet radial region radionuclide activity (Ci/MTU) relative to average pellet 

radionuclide activity (Ci/MTU) for a 300-year decay time

Pellet ring number 1-8 9 10 11 12 13 14

Volume fraction 0.8 0.0682 0.0264 0.0264 0.0264 0.0264 0.0264

Gas species

3H 0.95 1.02 1.07 1.12 1.20 1.36 1.72

129I 0.94 1.02 1.07 1.13 1.23 1.41 1.84

81Kr 0.94 1.05 1.10 1.16 1.26 1.43 1.85

85Kr 0.98 1.02 1.04 1.06 1.10 1.17 1.33

Total 0.94 1.02 1.07 1.13 1.23 1.41 1.84

Volatile species

135Cs 0.97 0.99 1.03 1.07 1.14 1.26 1.56

137Cs 0.96 1.02 1.06 1.10 1.18 1.31 1.62

90Sr 0.98 1.02 1.03 1.06 1.09 1.15 1.30

Total 0.96 1.02 1.05 1.09 1.15 1.26 1.52

Fuel fines

225Ac 0.98 1.02 1.04 1.07 1.10 1.16 1.30

227Ac 1.02 0.93 0.92 0.92 0.91 0.91 0.90

241Am 0.90 1.04 1.12 1.23 1.40 1.71 2.45

242mAm 0.91 1.02 1.10 1.20 1.37 1.68 2.42

242Am 0.91 1.02 1.10 1.20 1.37 1.68 2.42

243Am 0.88 1.08 1.18 1.30 1.49 1.85 2.69

137mBa 0.96 1.02 1.06 1.10 1.18 1.31 1.62

247Bk 0.84 1.16 1.28 1.43 1.65 2.06 3.06

249Cf 0.84 1.16 1.29 1.43 1.66 2.07 3.07

250Cf 0.83 1.18 1.31 1.46 1.70 2.12 3.14

251Cf 0.83 1.18 1.31 1.45 1.68 2.10 3.11

242Cm 0.91 1.02 1.10 1.20 1.37 1.68 2.42

243Cm 0.89 1.06 1.16 1.27 1.46 1.80 2.61

244Cm 0.87 1.09 1.20 1.33 1.53 1.89 2.78

245Cm 0.87 1.09 1.20 1.32 1.52 1.89 2.77

246Cm 0.85 1.13 1.25 1.38 1.60 1.99 2.93

247Cm 0.85 1.14 1.26 1.39 1.61 2.00 2.96

248Cm 0.84 1.17 1.29 1.44 1.67 2.08 3.07

154Eu 0.95 1.02 1.06 1.12 1.21 1.38 1.77

155Eu 0.94 1.03 1.08 1.14 1.23 1.41 1.82

93mNb 0.97 1.02 1.04 1.07 1.12 1.21 1.42

236Np 1.01 0.97 0.97 0.97 0.97 0.97 0.97

237Np 0.95 1.04 1.08 1.13 1.21 1.35 1.68

239Np 0.88 1.08 1.18 1.30 1.49 1.85 2.69

231Pa 1.02 0.93 0.92 0.92 0.91 0.90 0.90

233Pa 0.95 1.04 1.08 1.13 1.21 1.35 1.68

147Pm 0.96 0.99 1.02 1.07 1.15 1.29 1.65

210Po 0.99 1.01 1.02 1.03 1.05 1.08 1.16

236Pu 1.01 0.97 0.97 0.97 0.97 0.97 0.97

238Pu 0.97 1.04 1.07 1.09 1.13 1.19 1.33

239Pu 0.92 1.00 1.07 1.16 1.30 1.57 2.20

240Pu 0.96 0.92 0.97 1.03 1.15 1.36 1.85

241Pu 0.87 1.09 1.20 1.32 1.52 1.88 2.76

242Pu 0.88 1.07 1.17 1.29 1.48 1.83 2.65

244Pu 0.87 1.10 1.20 1.33 1.53 1.90 2.79

223Ra 1.02 0.93 0.92 0.92 0.91 0.91 0.90
C-20

Table C.7 (continued)

C.5.4 Containment Analysis

A sensitivity analysis was performed in this study because the range of parameters important for 

the containment analysis of high-burnup fuel (i.e., the release fractions and breached spent fuel 

rod fraction) has not been established yet for normal and accident conditions of transport. 

Normal Conditions of Transport

The allowable leakage rate for NCT was calculated as a function of decay time (5, 40, 100, and 

300 years), pellet radial region from which the releasable radioactive material originates

(non-rim region and rim structure corresponding to ring No. 14 in the calculation model 

previously described in Sect.C.5.2), and fraction of fuel rods assumed to fail (0.01, 0.03, 0.10, 

and 1) using the following release fractions:

• crud: 0.15, 0.3, 0.5, and 1;

• gaseous species: 0.1, 0.3, and 0.4;

• volatile species: 2 × 10-4, 1 × 10-3, and 2 × 10-3; and

• fuel fines: 3 × 10-5, 1.5 × 10-4, and 3 × 10-4.

where, the underlined values are the release fractions typically used for low-burnup fuel (see 

Table C.1). 

Using these release fractions, 15 different cases were developed for each decay time, pellet 

radial region, and fraction of fuel rods assumed to develop cladding breaches. Case 1 is the 

Pellet ring number 1-8 9 10 11 12 13 14

Volume fraction 0.8 0.0682 0.0264 0.0264 0.0264 0.0264 0.0264

224Ra 0.99 1.01 1.01 1.02 1.02 1.03 1.04

225Ra 0.98 1.02 1.04 1.07 1.10 1.16 1.30

226Ra 0.99 1.01 1.02 1.03 1.05 1.09 1.16

228Ra 1.00 1.00 1.00 1.00 1.00 1.00 1.00

222Rn 0.99 1.01 1.02 1.03 1.05 1.09 1.16

125Sb 0.94 1.02 1.07 1.13 1.23 1.42 1.86

151Sm 0.95 1.01 1.05 1.11 1.21 1.39 1.81

121mSn 0.95 1.02 1.06 1.12 1.21 1.39 1.80

99Tc 0.96 1.01 1.04 1.09 1.16 1.29 1.60

227Th 1.02 0.93 0.92 0.92 0.91 0.91 0.90

228Th 0.99 1.01 1.01 1.02 1.02 1.03 1.04

229Th 0.98 1.02 1.04 1.07 1.10 1.16 1.30

230Th 0.98 1.01 1.03 1.04 1.06 1.10 1.19

231Th 1.03 0.92 0.90 0.90 0.89 0.89 0.88

232U 0.99 1.01 1.01 1.02 1.02 1.03 1.04

233U 0.97 1.03 1.06 1.09 1.14 1.23 1.44

234U 0.98 1.02 1.04 1.05 1.08 1.12 1.22

236U 1.00 1.00 1.00 1.00 1.00 1.00 1.01

90Y 0.98 1.02 1.03 1.06 1.09 1.15 1.30

93Zr 0.97 1.02 1.04 1.07 1.12 1.21 1.42

Total 0.92 1.02 1.09 1.18 1.32 1.59 2.20
C-21

nominal intact configuration that uses the release fractions typically used for low-burnup fuel. 

For cases 2 through 10, the varying parameters were changed one at a time. Cases 2 through 4 

use varying fraction of crud that spalls off cladding; cases 5 and 6 use a varying fraction of 

gases released due to a cladding breach; cases 7 and 8 use a varying fraction of volatiles 

released due to a cladding breach; and cases 9 and 10 use a varying mass fraction of fuel 

released as fuel fines due to a cladding breach. For cases 11 through 15, multiple varying 

parameters were changed simultaneously because fuel fragmentation and fission gas release 

are expected to be correlated. The NCT allowable release rates are summarized in Table C.8

and Table C.9 for the GBC-32 and GBC-68, respectively. The free void volumes of the GBC-32 

and GBC-68 used in the calculations were 7.05E+06 cm3 and 5.68E+06 cm3

, respectively. The 

graphs in Figure C.5 and Figure C.6 illustrate the effects of varying release fractions on the 

GBC-32 and GBC-68 allowable leakage rates, respectively. For each decay time, the effects of 

each parameter relevant to failed fuel containment are further analyzed.

Decay time

Decay time has a significant impact on the releasable activity and allowable leakage rate 

primarily because of the decay of 60Co (t1/2 = 5.271 years), which is the source of radioactivity in

crud, and 85Kr (t1/2 = 10.76 years), which dominates the gaseous species. The allowable leakage 

rate increases with increasing decay time.

Fuel pellet radial region

 

The extent to which the pellet radial region from which the releasable activity originates affects 

the allowable leakage rate depends on the contribution of the crud to the total releasable 

activity. Small effects are observed for GBC-32, the 5-year decay time, and small fraction of 

breached spent fuel rods (e.g., 0.01), as well as for GBC-68, the 5-year decay time, and any 

fraction of breached spent fuel rods, as crud has a significant contribution to the total releasable 

activity for these cases. For the other cases, the allowable leakage rate based on releasable 

activity from the outer rim structure is approximately half the allowable leakage rate based on 

releasable activity from the non-rim region. 

Fraction of fuel rods that develop cladding breaches

Allowable leakage rate decreases with increasing fraction of breached spent fuel rods. The 

decrease is relatively small (e.g., a 15% decrease because of an increase in spent fuel rod 

breach fraction from 0.01 to 0.03) if crud has a large contribution to the total releasable activity 

(e.g., cases 1 through 4) and is fairly large if the fuel fines dominate the total releasable activity 

(e.g. a decrease by a factor of 2.5 because of an increase of failed fuel rod fraction from 0.01 to 

0.03). 

Fraction of crud that spalls off cladding

Crud is an important factor in the calculation of the allowable leakage rate for the time interval 

5 to 40 years after fuel discharge from the reactor because of its relatively high contribution to 

the total releasable activity. Crud has a more pronounced effect on the BWR packages than on 

the PWR packages because of the larger 60Co activity per fuel rod surface area associated with 

the BWR fuel. As a result, the BWR package allowable leakage rates are smaller (i.e., more 

restrictive) than the PWR package allowable leakage rates. For GBC-68 and the 5-year decay 

time, an increase in the fraction of crud that spalls off the cladding by a factor of two would 

cause a decrease in the allowable leakage rate by a factor of approximately two. 
C-22

Table C.8 GBC-32 allowable leakage rate for NCT

5-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 1.3E-04 7.3E-05 2.9E-05 3.3E-06 1.0E-04 4.8E-05 1.7E-05 1.8E-06

2 0.3 2.0E-4 3.0E-5 0.3 8.2E-05 5.5E-05 2.5E-05 3.2E-06 6.8E-05 3.9E-05 1.6E-05 1.8E-06

3 0.3 2.0E-4 3.0E-5 0.5 5.5E-05 4.1E-05 2.2E-05 3.1E-06 4.8E-05 3.2E-05 1.4E-05 1.8E-06

4 0.3 2.0E-4 3.0E-5 1.0 3.0E-05 2.5E-05 1.6E-05 3.0E-06 2.8E-05 2.1E-05 1.2E-05 1.7E-06

5 0.1 2.0E-4 3.0E-5 0.15 1.5E-04 8.8E-05 3.6E-05 4.3E-06 1.1E-04 5.6E-05 2.0E-05 2.2E-06

6 0.4 2.0E-4 3.0E-5 0.15 1.3E-04 6.8E-05 2.6E-05 2.9E-06 9.5E-05 4.4E-05 1.6E-05 1.7E-06

7 0.3 1.0E-3 3.0E-5 0.15 1.1E-04 5.4E-05 1.9E-05 2.1E-06 7.9E-05 3.5E-05 1.2E-05 1.2E-06

8 0.3 2.0E-3 3.0E-5 0.15 8.8E-05 4.0E-05 1.4E-05 1.5E-06 6.3E-05 2.6E-05 8.4E-06 8.7E-07

9 0.3 2.0E-4 1.5E-4 0.15 7.3E-05 3.2E-05 1.1E-05 1.1E-06 4.3E-05 1.6E-05 5.2E-06 5.3E-07

10 0.3 2.0E-4 3.0E-4 0.15 4.7E-05 1.8E-05 5.9E-06 6.0E-07 2.5E-05 9.1E-06 2.8E-06 2.8E-07

11 0.3 2.0E-3 3.0E-4 0.15 4.0E-05 1.5E-05 4.8E-06 4.9E-07 2.2E-05 7.8E-06 2.4E-06 2.4E-07

12 0.1 2.0E-3 3.0E-4 0.15 4.1E-05 1.6E-05 5.0E-06 5.1E-07 2.2E-05 8.0E-06 2.5E-06 2.5E-07

13 0.4 2.0E-3 3.0E-4 0.15 4.0E-05 1.5E-05 4.7E-06 4.8E-07 2.2E-05 7.7E-06 2.4E-06 2.4E-07

14 0.4 2.0E-3 3.0E-4 0.5 2.8E-05 1.3E-05 4.5E-06 4.8E-07 1.8E-05 7.1E-06 2.3E-06 2.4E-07

15 0.4 2.0E-3 3.0E-4 1.0 2.0E-05 1.1E-05 4.2E-06 4.8E-07 1.4E-05 6.4E-06 2.2E-06 2.4E-07

40-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 6.8E-04 2.3E-04 7.0E-05 7.0E-06 3.4E-04 1.2E-04 3.5E-05 3.5E-06

2 0.3 2.0E-4 3.0E-5 0.3 6.6E-04 2.3E-04 6.9E-05 7.0E-06 3.4E-04 1.1E-04 3.5E-05 3.5E-06

3 0.3 2.0E-4 3.0E-5 0.5 6.3E-04 2.2E-04 6.9E-05 7.0E-06 3.3E-04 1.1E-04 3.5E-05 3.5E-06

4 0.3 2.0E-4 3.0E-5 1.0 5.8E-04 2.2E-04 6.8E-05 7.0E-06 3.1E-04 1.1E-04 3.4E-05 3.5E-06

5 0.1 2.0E-4 3.0E-5 0.15 7.1E-04 2.4E-04 7.4E-05 7.4E-06 3.6E-04 1.2E-04 3.6E-05 3.6E-06

6 0.4 2.0E-4 3.0E-5 0.15 6.6E-04 2.2E-04 6.8E-05 6.8E-06 3.4E-04 1.1E-04 3.4E-05 3.4E-06

7 0.3 1.0E-3 3.0E-5 0.15 4.8E-04 1.6E-04 4.9E-05 4.9E-06 2.6E-04 8.8E-05 2.7E-05 2.7E-06

8 0.3 2.0E-3 3.0E-5 0.15 3.5E-04 1.2E-04 3.6E-05 3.6E-06 2.0E-04 6.8E-05 2.0E-05 2.1E-06

9 0.3 2.0E-4 1.5E-4 0.15 1.6E-04 5.5E-05 1.6E-05 1.6E-06 7.8E-05 2.6E-05 7.8E-06 7.8E-07

10 0.3 2.0E-4 3.0E-4 0.15 8.4E-05 2.8E-05 8.4E-06 8.4E-07 4.0E-05 1.3E-05 4.0E-06 4.0E-07

11 0.3 2.0E-3 3.0E-4 0.15 7.5E-05 2.5E-05 7.5E-06 7.5E-07 3.7E-05 1.2E-05 3.7E-06 3.7E-07

12 0.1 2.0E-3 3.0E-4 0.15 7.6E-05 2.5E-05 7.6E-06 7.6E-07 3.7E-05 1.2E-05 3.7E-06 3.7E-07

13 0.4 2.0E-3 3.0E-4 0.15 7.5E-05 2.5E-05 7.5E-06 7.5E-07 3.7E-05 1.2E-05 3.7E-06 3.7E-07

14 0.4 2.0E-3 3.0E-4 0.5 7.4E-05 2.5E-05 7.5E-06 7.5E-07 3.6E-05 1.2E-05 3.7E-06 3.7E-07

15 0.4 2.0E-3 3.0E-4 1.0 7.3E-05 2.5E-05 7.5E-06 7.5E-07 3.6E-05 1.2E-05 3.7E-06 3.7E-07
C-23

100-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 1.1E-03 3.7E-04 1.1E-04 1.1E-05 5.1E-04 1.7E-04 5.1E-05 5.1E-06

2 0.3 2.0E-4 3.0E-5 0.3 1.1E-03 3.7E-04 1.1E-04 1.1E-05 5.1E-04 1.7E-04 5.1E-05 5.1E-06

3 0.3 2.0E-4 3.0E-5 0.5 1.1E-03 3.7E-04 1.1E-04 1.1E-05 5.1E-04 1.7E-04 5.1E-05 5.1E-06

4 0.3 2.0E-4 3.0E-5 1.0 1.1E-03 3.7E-04 1.1E-04 1.1E-05 5.1E-04 1.7E-04 5.1E-05 5.1E-06

5 0.1 2.0E-4 3.0E-5 0.15 1.1E-03 3.7E-04 1.1E-04 1.1E-05 5.1E-04 1.7E-04 5.1E-05 5.1E-06

6 0.4 2.0E-4 3.0E-5 0.15 1.1E-03 3.7E-04 1.1E-04 1.1E-05 5.1E-04 1.7E-04 5.1E-05 5.1E-06

7 0.3 1.0E-3 3.0E-5 0.15 9.5E-04 3.2E-04 9.5E-05 9.5E-06 4.6E-04 1.5E-04 4.6E-05 4.6E-06

8 0.3 2.0E-3 3.0E-5 0.15 8.1E-04 2.7E-04 8.1E-05 8.1E-06 4.1E-04 1.4E-04 4.1E-05 4.1E-06

9 0.3 2.0E-4 1.5E-4 0.15 2.3E-04 7.6E-05 2.3E-05 2.3E-06 1.1E-04 3.5E-05 1.1E-05 1.1E-06

10 0.3 2.0E-4 3.0E-4 0.15 1.1E-04 3.8E-05 1.1E-05 1.1E-06 5.3E-05 1.8E-05 5.3E-06 5.3E-07

11 0.3 2.0E-3 3.0E-4 0.15 1.1E-04 3.7E-05 1.1E-05 1.1E-06 5.1E-05 1.7E-05 5.1E-06 5.1E-07

12 0.1 2.0E-3 3.0E-4 0.15 1.1E-04 3.7E-05 1.1E-05 1.1E-06 5.1E-05 1.7E-05 5.1E-06 5.1E-07

13 0.4 2.0E-3 3.0E-4 0.15 1.1E-04 3.7E-05 1.1E-05 1.1E-06 5.1E-05 1.7E-05 5.1E-06 5.1E-07

14 0.4 2.0E-3 3.0E-4 0.5 1.1E-04 3.7E-05 1.1E-05 1.1E-06 5.1E-05 1.7E-05 5.1E-06 5.1E-07

15 0.4 2.0E-3 3.0E-4 1.0 1.1E-04 3.7E-05 1.1E-05 1.1E-06 5.1E-05 1.7E-05 5.1E-06 5.1E-07

300-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 2.1E-03 6.8E-04 2.1E-04 2.1E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

2 0.3 2.0E-4 3.0E-5 0.3 2.1E-03 6.8E-04 2.1E-04 2.1E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

3 0.3 2.0E-4 3.0E-5 0.5 2.1E-03 6.8E-04 2.1E-04 2.1E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

4 0.3 2.0E-4 3.0E-5 1.0 2.1E-03 6.8E-04 2.1E-04 2.1E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

5 0.1 2.0E-4 3.0E-5 0.15 2.1E-03 6.8E-04 2.1E-04 2.1E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

6 0.4 2.0E-4 3.0E-5 0.15 2.1E-03 6.8E-04 2.1E-04 2.1E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

7 0.3 1.0E-3 3.0E-5 0.15 2.0E-03 6.8E-04 2.0E-04 2.0E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

8 0.3 2.0E-3 3.0E-5 0.15 2.0E-03 6.8E-04 2.0E-04 2.0E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

9 0.3 2.0E-4 1.5E-4 0.15 4.1E-04 1.4E-04 4.1E-05 4.1E-06 1.7E-04 5.6E-05 1.7E-05 1.7E-06

10 0.3 2.0E-4 3.0E-4 0.15 2.1E-04 6.8E-05 2.1E-05 2.1E-06 8.4E-05 2.8E-05 8.4E-06 8.4E-07

11 0.3 2.0E-3 3.0E-4 0.15 2.1E-04 6.8E-05 2.1E-05 2.1E-06 8.4E-05 2.8E-05 8.4E-06 8.4E-07

12 0.1 2.0E-3 3.0E-4 0.15 2.1E-04 6.8E-05 2.1E-05 2.1E-06 8.4E-05 2.8E-05 8.4E-06 8.4E-07

13 0.4 2.0E-3 3.0E-4 0.15 2.1E-04 6.8E-05 2.1E-05 2.1E-06 8.4E-05 2.8E-05 8.4E-06 8.4E-07

14 0.4 2.0E-3 3.0E-4 0.5 2.1E-04 6.8E-05 2.1E-05 2.1E-06 8.4E-05 2.8E-05 8.4E-06 8.4E-07

15 0.4 2.0E-3 3.0E-4 1.0 2.1E-04 6.8E-05 2.1E-05 2.1E-06 8.4E-05 2.8E-05 8.4E-06 8.4E-07 aFraction of gases released due to a cladding breach

bFraction of volatiles released due to a cladding breach 

cMass fraction of fuel released as fines due to a cladding breach 

dFraction of crud that spalls off cladding.

Table C.8 (continued)
C-24

Table C.9 GBC-68 allowable leakage rate for NCT

5-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 2.3E-05 2.0E-05 1.3E-05 2.6E-06 2.2E-05 1.8E-05 1.0E-05 1.6E-06

2 0.3 2.0E-4 3.0E-5 0.3 1.2E-05 1.1E-05 8.7E-06 2.3E-06 1.2E-05 1.0E-05 7.3E-06 1.5E-06

3 0.3 2.0E-4 3.0E-5 0.5 7.3E-06 7.0E-06 6.0E-06 2.1E-06 7.2E-06 6.7E-06 5.3E-06 1.4E-06

4 0.3 2.0E-4 3.0E-5 1.0 3.7E-06 3.6E-06 3.3E-06 1.6E-06 3.7E-06 3.5E-06 3.1E-06 1.2E-06

5 0.1 2.0E-4 3.0E-5 0.15 2.4E-05 2.1E-05 1.5E-05 3.3E-06 2.2E-05 1.9E-05 1.2E-05 2.0E-06

6 0.4 2.0E-4 3.0E-5 0.15 2.3E-05 1.9E-05 1.3E-05 2.3E-06 2.2E-05 1.7E-05 9.7E-06 1.5E-06

7 0.3 1.0E-3 3.0E-5 0.15 2.2E-05 1.8E-05 1.1E-05 1.7E-06 2.1E-05 1.5E-05 8.0E-06 1.1E-06

8 0.3 2.0E-3 3.0E-5 0.15 2.1E-05 1.6E-05 8.6E-06 1.2E-06 1.9E-05 1.3E-05 6.3E-06 8.1E-07

9 0.3 2.0E-4 1.5E-4 0.15 2.0E-05 1.4E-05 6.9E-06 9.1E-07 1.7E-05 1.0E-05 4.3E-06 5.0E-07

10 0.3 2.0E-4 3.0E-4 0.15 1.7E-05 1.0E-05 4.3E-06 5.0E-07 1.3E-05 6.7E-06 2.5E-06 2.7E-07

11 0.3 2.0E-3 3.0E-4 0.15 1.6E-05 9.0E-06 3.6E-06 4.2E-07 1.2E-05 5.9E-06 2.1E-06 2.3E-07

12 0.1 2.0E-3 3.0E-4 0.15 1.6E-05 9.3E-06 3.7E-06 4.3E-07 1.2E-05 6.0E-06 2.2E-06 2.4E-07

13 0.4 2.0E-3 3.0E-4 0.15 1.6E-05 8.9E-06 3.6E-06 4.1E-07 1.2E-05 5.9E-06 2.1E-06 2.3E-07

14 0.4 2.0E-3 3.0E-4 0.5 6.4E-06 4.9E-06 2.7E-06 3.9E-07 5.7E-06 3.8E-06 1.8E-06 2.2E-07

15 0.4 2.0E-3 3.0E-4 1.0 3.5E-06 3.0E-06 2.0E-06 3.7E-07 3.2E-06 2.5E-06 1.4E-06 2.2E-07

40-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 4.9E-04 1.9E-04 5.9E-05 6.0E-06 3.0E-04 1.1E-04 3.3E-05 3.4E-06

2 0.3 2.0E-4 3.0E-5 0.3 4.1E-04 1.7E-04 5.8E-05 6.0E-06 2.7E-04 1.0E-04 3.3E-05 3.4E-06

3 0.3 2.0E-4 3.0E-5 0.5 3.4E-04 1.6E-04 5.6E-05 6.0E-06 2.3E-04 9.8E-05 3.2E-05 3.4E-06

4 0.3 2.0E-4 3.0E-5 1.0 2.3E-04 1.3E-04 5.2E-05 6.0E-06 1.8E-04 8.7E-05 3.1E-05 3.4E-06

5 0.1 2.0E-4 3.0E-5 0.15 5.1E-04 2.0E-04 6.2E-05 6.4E-06 3.1E-04 1.1E-04 3.5E-05 3.5E-06

6 0.4 2.0E-4 3.0E-5 0.15 4.8E-04 1.8E-04 5.8E-05 5.9E-06 2.9E-04 1.1E-04 3.3E-05 3.3E-06

7 0.3 1.0E-3 3.0E-5 0.15 3.7E-04 1.4E-04 4.2E-05 4.3E-06 2.3E-04 8.3E-05 2.5E-05 2.6E-06

8 0.3 2.0E-3 3.0E-5 0.15 2.8E-04 1.0E-04 3.1E-05 3.2E-06 1.8E-04 6.4E-05 2.0E-05 2.0E-06

9 0.3 2.0E-4 1.5E-4 0.15 1.3E-04 4.6E-05 1.4E-05 1.4E-06 7.4E-05 2.5E-05 7.6E-06 7.6E-07

10 0.3 2.0E-4 3.0E-4 0.15 7.0E-05 2.4E-05 7.2E-06 7.2E-07 3.8E-05 1.3E-05 3.9E-06 3.9E-07

11 0.3 2.0E-3 3.0E-4 0.15 6.3E-05 2.2E-05 6.5E-06 6.5E-07 3.5E-05 1.2E-05 3.6E-06 3.6E-07

12 0.1 2.0E-3 3.0E-4 0.15 6.4E-05 2.2E-05 6.5E-06 6.5E-07 3.5E-05 1.2E-05 3.6E-06 3.6E-07

13 0.4 2.0E-3 3.0E-4 0.15 6.3E-05 2.1E-05 6.5E-06 6.5E-07 3.5E-05 1.2E-05 3.6E-06 3.6E-07

14 0.4 2.0E-3 3.0E-4 0.5 6.0E-05 2.1E-05 6.4E-06 6.5E-07 3.4E-05 1.2E-05 3.5E-06 3.6E-07

15 0.4 2.0E-3 3.0E-4 1.0 5.5E-05 2.0E-05 6.4E-06 6.5E-07 3.3E-05 1.2E-05 3.5E-06 3.6E-07
C-25

100-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 9.6E-04 3.2E-04 9.6E-05 9.6E-06 5.1E-04 1.7E-04 5.1E-05 5.1E-06

2 0.3 2.0E-4 3.0E-5 0.3 9.6E-04 3.2E-04 9.6E-05 9.6E-06 5.1E-04 1.7E-04 5.1E-05 5.1E-06

3 0.3 2.0E-4 3.0E-5 0.5 9.6E-04 3.2E-04 9.6E-05 9.6E-06 5.1E-04 1.7E-04 5.1E-05 5.1E-06

4 0.3 2.0E-4 3.0E-5 1.0 9.6E-04 3.2E-04 9.6E-05 9.6E-06 5.1E-04 1.7E-04 5.1E-05 5.1E-06

5 0.1 2.0E-4 3.0E-5 0.15 9.6E-04 3.2E-04 9.6E-05 9.6E-06 5.1E-04 1.7E-04 5.1E-05 5.1E-06

6 0.4 2.0E-4 3.0E-5 0.15 9.6E-04 3.2E-04 9.6E-05 9.6E-06 5.1E-04 1.7E-04 5.1E-05 5.1E-06

7 0.3 1.0E-3 3.0E-5 0.15 8.3E-04 2.8E-04 8.3E-05 8.3E-06 4.5E-04 1.5E-04 4.5E-05 4.5E-06

8 0.3 2.0E-3 3.0E-5 0.15 7.1E-04 2.4E-04 7.1E-05 7.1E-06 4.0E-04 1.3E-04 4.0E-05 4.0E-06

9 0.3 2.0E-4 1.5E-4 0.15 2.0E-04 6.6E-05 2.0E-05 2.0E-06 1.0E-04 3.5E-05 1.0E-05 1.0E-06

10 0.3 2.0E-4 3.0E-4 0.15 1.0E-04 3.3E-05 1.0E-05 1.0E-06 5.2E-05 1.7E-05 5.2E-06 5.2E-07

11 0.3 2.0E-3 3.0E-4 0.15 9.6E-05 3.2E-05 9.6E-06 9.6E-07 5.1E-05 1.7E-05 5.1E-06 5.1E-07

12 0.1 2.0E-3 3.0E-4 0.15 9.6E-05 3.2E-05 9.6E-06 9.6E-07 5.1E-05 1.7E-05 5.1E-06 5.1E-07

13 0.4 2.0E-3 3.0E-4 0.15 9.6E-05 3.2E-05 9.6E-06 9.6E-07 5.1E-05 1.7E-05 5.1E-06 5.1E-07

14 0.4 2.0E-3 3.0E-4 0.5 9.6E-05 3.2E-05 9.6E-06 9.6E-07 5.1E-05 1.7E-05 5.1E-06 5.1E-07

15 0.4 2.0E-3 3.0E-4 1.0 9.6E-05 3.2E-05 9.6E-06 9.6E-07 5.1E-05 1.7E-05 5.1E-06 5.1E-07

300-year decay time

Non-rim pellet region Rim pellet region

Fraction of breached fuel rods Fraction of breached fuel rods

0.01 0.03 0.10 1.00 0.01 0.03 0.10 1.00

Case# fGa fVb fFc fCd Allowable leakage rate (cm3/s) Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 0.15 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.5E-04 2.8E-04 8.5E-05 8.5E-06

2 0.3 2.0E-4 3.0E-5 0.3 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.5E-04 2.8E-04 8.5E-05 8.5E-06

3 0.3 2.0E-4 3.0E-5 0.5 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.5E-04 2.8E-04 8.5E-05 8.5E-06

4 0.3 2.0E-4 3.0E-5 1.0 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.5E-04 2.8E-04 8.5E-05 8.5E-06

5 0.1 2.0E-4 3.0E-5 0.15 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.5E-04 2.8E-04 8.5E-05 8.5E-06

6 0.4 2.0E-4 3.0E-5 0.15 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.5E-04 2.8E-04 8.5E-05 8.5E-06

7 0.3 1.0E-3 3.0E-5 0.15 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

8 0.3 2.0E-3 3.0E-5 0.15 1.8E-03 5.9E-04 1.8E-04 1.8E-05 8.4E-04 2.8E-04 8.4E-05 8.4E-06

9 0.3 2.0E-4 1.5E-4 0.15 3.6E-04 1.2E-04 3.6E-05 3.6E-06 1.7E-04 5.6E-05 1.7E-05 1.7E-06

10 0.3 2.0E-4 3.0E-4 0.15 1.8E-04 5.9E-05 1.8E-05 1.8E-06 8.5E-05 2.8E-05 8.5E-06 8.5E-07

11 0.3 2.0E-3 3.0E-4 0.15 1.8E-04 5.9E-05 1.8E-05 1.8E-06 8.5E-05 2.8E-05 8.5E-06 8.5E-07

12 0.1 2.0E-3 3.0E-4 0.15 1.8E-04 5.9E-05 1.8E-05 1.8E-06 8.5E-05 2.8E-05 8.5E-06 8.5E-07

13 0.4 2.0E-3 3.0E-4 0.15 1.8E-04 5.9E-05 1.8E-05 1.8E-06 8.5E-05 2.8E-05 8.5E-06 8.5E-07

14 0.4 2.0E-3 3.0E-4 0.5 1.8E-04 5.9E-05 1.8E-05 1.8E-06 8.5E-05 2.8E-05 8.5E-06 8.5E-07

15 0.4 2.0E-3 3.0E-4 1.0 1.8E-04 5.9E-05 1.8E-05 1.8E-06 8.5E-05 2.8E-05 8.5E-06 8.5E-07 aFraction of gases released due to a cladding breach. bFraction of volatiles released due to a cladding breach. cMass fraction of fuel released as fines due to a cladding breach. dFraction of crud that spalls off cladding.

Table C.9 (continued)
C-26

(a)

(b)

(c)

Figure C.5 Illustration of the effects of varying release fractions on GBC-32 NCT 

allowable leakage rate at: (a) 5, (b) 40, and (c) 100 years after fuel discharge 

from the reactor
C-27

(a)

(b)

(c)

Figure C.6 Illustration of the effects of varying release fractions on GBC-68 NCT

allowable leakage rate at: (a) 5, (b) 40, and (c) 100 years after fuel discharge 

from the reactor
C-28

Fraction of gases released due to cladding breach

For each decay time and fraction of breached spent fuel rods, an increase in the fraction of 

gaseous species from 0.1 to 0.4 would cause a relatively small decrease (up to ~30%) in the 

allowable release rate. 

Fraction of volatiles released due to cladding breach

The volatile source term is dominated by 137Cs (T1/2 = 30.07 years) and 90Sr (T1/2 = 28.78 years). 

An increase in the fraction of volatile source term by one order of magnitude would cause a 

maximum decrease in the allowable leakage rate by a factor of approximately two (see 

Table C.9, 40-year decay time, cases 1 and 8).

Fraction of fuel fines released due to cladding breach

Allowable leakage rate exhibits the greatest sensitivity to changes in the mass fraction of fuel 

released as fuel fines due to cladding breach. Depending on the crud contribution to the total 

releasable activity, an increase in the fraction of fuel fines by a factor of 10 would cause a 

decrease in the allowable leakage rate by a factor of ~1.5 to 10.

Hypothetical Accident Conditions

The allowable leakage rate for HAC was calculated as a function of decay time (5, 40, 100, and 

300 years), and pellet radial region from which the fission products and fuel fines originate 

(non-rim region and rim structure corresponding to ring No. 14 in the calculation model 

previously described in Sect. C.5.2) assuming a breached fuel rod fraction of 1 and the following 

release fractions:

• crud: 1.0;

• gaseous species: 0.3 and 0.4;

• volatile species: 2 × 10-4, and 2 × 10-3; and

• fuel fines: 3 × 10-5 and 3 × 10-4.

where the underlined values are the release fractions typically used for HAC of low-burnup fuel.

The calculated allowable release rates for HAC are summarized in Table C.10 for both generic 

casks. Similar to the NCT results, the HAC allowable leakage rate has the largest sensitivity to 

the mass fraction of fuel released as fines. Depending on the decay time, an increase in the 

mass fraction of fuel released as fuel fines by a factor of 10 would cause a decrease in the 

allowable leakage rate by a factor of ~5 to 10 for the PWR fuel and by a factor of ~3.5 to 10 for 

the BWR fuel.
C-29

Table C.10 GBC-32 and GBC-68 allowable leakage rate for HAC

GBC-32

Decay time (years) 5 40 100 300

Pellet region Non-rim Rim Non-rim Rim Non-rim Rim Non-rim Rim

Case# fGa fVb fFc Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 1.78E-02 1.03E-02 4.14E-02 2.07E-02 6.56E-02 3.05E-02 1.22E-01 5.00E-02

2 0.4 2.0E-4 3.0E-5 1.60E-02 9.43E-03 4.03E-02 2.03E-02 6.56E-02 3.05E-02 1.22E-01 5.00E-02

3 0.3 2.0E-3 3.0E-5 8.42E-03 5.08E-03 2.13E-02 1.22E-02 4.82E-02 2.44E-02 1.21E-01 4.98E-02

4 0.3 2.0E-4 3.0E-4 3.52E-03 1.67E-03 4.98E-03 2.35E-03 6.82E-03 3.13E-03 1.22E-02 5.00E-03

5 0.4 2.0E-3 3.0E-4 2.83E-03 1.41E-03 4.46E-03 2.17E-03 6.58E-03 3.05E-03 1.22E-02 5.00E-03

GBC-68

Decay time (years) 5 40 100 300

Pellet region Non-rim Rim Non-rim Rim Non-rim Rim Non-rim Rim

Case# fGa fVb fFc Allowable leakage rate (cm3/s)

1 0.3 2.0E-4 3.0E-5 9.67E-03 7.07E-03 3.53E-02 1.99E-02 5.69E-02 3.00E-02 1.06E-01 5.02E-02

2 0.4 2.0E-4 3.0E-5 9.06E-03 6.65E-03 3.44E-02 1.95E-02 5.69E-02 3.00E-02 1.06E-01 5.02E-02

3 0.3 2.0E-3 3.0E-5 5.74E-03 4.09E-03 1.86E-02 1.17E-02 4.21E-02 2.39E-02 1.05E-01 5.00E-02

4 0.3 2.0E-4 3.0E-4 2.69E-03 1.51E-03 4.28E-03 2.29E-03 5.91E-03 3.09E-03 1.06E-02 5.02E-03

5 0.4 2.0E-3 3.0E-4 2.23E-03 1.29E-03 3.85E-03 2.11E-03 5.70E-03 3.01E-03 1.06E-02 5.02E-03 aFraction of gases released due to a cladding breach. bFraction of volatiles released due to a cladding breach. cMass fraction of fuel released as fines due to a cladding breach.

C.6 CONFIGURATION CATEGORIES 2 AND 3

Crud was assumed to be released as a consequence of normal, off-normal, and accident 

conditions of spent fuel storage, and normal and accident conditions of transportation that do 

not cause cladding breach (i.e., Category 2—rod/assembly deformation; and Category 3—

changes to assembly axial alignment). The GBC-32 and GBC-68 NCT allowable leakage rate 

as a function of the fraction of crud that spalls off cladding and decay time is provided in Table 

C.11. The allowable leakage rate is inversely proportional to the fraction of crud assumed to 

spall off the cladding. Hence, allowable leakage rate was calculated as a function of the fraction 

of crud that spalls off the cladding to evaluate the effect of varying the crud releasable activity

(see Appendix C.6). For the 5-year decay time, an increase in the fraction of crud from 0.05 to 1 

would cause a decrease in the allowable leakage rate from 6.58 × 10-4 cm3

/s to 3.29 × 10-5

cm3

/s for the GBC-32 package and from 7.54 × 10-5 cm3/s to 3.77 × 10-6 cm3/s for the GBC-68

package. For the 40-year decay time, an increase in the fraction of crud from 0.05 to 1 would 

cause a decrease in the allowable leakage rate from 6.56 × 10-2 cm3

/s to 3.28 × 10-3 cm3/s for 

the GBC-32 package and from 7.52 × 10-3 cm3

/s to 3.76 × 10-4 cm3/s for the GBC-68 package.
C-30

Table C.11 Allowable leakage rate for rod/assembly deformation and changes to axial 

alignment categories

GBC-32

5-year decay time 40-year decay time

Fraction 

of crud 

released

Releasable 

activity 

(Ci)

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

Releasable 

activity 

(Ci)

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

0.05 3.22E+01 1.08E+01 3.00E-09 6.58E-04 3.23E-01 1.08E+01 3.00E-09 6.56E-02

0.1 6.44E+01 1.08E+01 3.00E-09 3.29E-04 6.46E-01 1.08E+01 3.00E-09 3.28E-02

0.15 9.66E+01 1.08E+01 3.00E-09 2.19E-04 9.68E-01 1.08E+01 3.00E-09 2.19E-02

0.2 1.29E+02 1.08E+01 3.00E-09 1.64E-04 1.29E+00 1.08E+01 3.00E-09 1.64E-02

0.3 1.93E+02 1.08E+01 3.00E-09 1.10E-04 1.94E+00 1.08E+01 3.00E-09 1.09E-02

0.4 2.58E+02 1.08E+01 3.00E-09 8.22E-05 2.58E+00 1.08E+01 3.00E-09 8.20E-03

0.5 3.22E+02 1.08E+01 3.00E-09 6.58E-05 3.23E+00 1.08E+01 3.00E-09 6.56E-03

0.6 3.86E+02 1.08E+01 3.00E-09 5.48E-05 3.87E+00 1.08E+01 3.00E-09 5.47E-03

0.7 4.51E+02 1.08E+01 3.00E-09 4.70E-05 4.52E+00 1.08E+01 3.00E-09 4.68E-03

0.8 5.15E+02 1.08E+01 3.00E-09 4.11E-05 5.16E+00 1.08E+01 3.00E-09 4.10E-03

0.9 5.80E+02 1.08E+01 3.00E-09 3.65E-05 5.81E+00 1.08E+01 3.00E-09 3.64E-03

1 6.44E+02 1.08E+01 3.00E-09 3.29E-05 6.46E+00 1.08E+01 3.00E-09 3.28E-03

GBC-68

5-year decay time 40-year decay time

Fraction 

of crud 

released

Releasable 

activity 

(Ci)

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

Releasable 

activity 

(Ci)

Effective 

A2 value 

(Ci)

Allowable 

release 

rate (Ci/s)

Allowable 

leakage 

rate 

(cm3/s)

0.05 2.26E+02 1.08E+01 3.00E-09 7.54E-05 2.27E+00 1.08E+01 3.00E-09 7.52E-03

0.1 4.52E+02 1.08E+01 3.00E-09 3.77E-05 4.54E+00 1.08E+01 3.00E-09 3.76E-03

0.15 6.79E+02 1.08E+01 3.00E-09 2.51E-05 6.80E+00 1.08E+01 3.00E-09 2.51E-03

0.2 9.05E+02 1.08E+01 3.00E-09 1.88E-05 9.07E+00 1.08E+01 3.00E-09 1.88E-03

0.3 1.36E+03 1.08E+01 3.00E-09 1.26E-05 1.36E+01 1.08E+01 3.00E-09 1.25E-03

0.4 1.81E+03 1.08E+01 3.00E-09 9.42E-06 1.81E+01 1.08E+01 3.00E-09 9.40E-04

0.5 2.26E+03 1.08E+01 3.00E-09 7.54E-06 2.27E+01 1.08E+01 3.00E-09 7.52E-04

0.6 2.71E+03 1.08E+01 3.00E-09 6.28E-06 2.72E+01 1.08E+01 3.00E-09 6.27E-04

0.7 3.17E+03 1.08E+01 3.00E-09 5.39E-06 3.17E+01 1.08E+01 3.00E-09 5.37E-04

0.8 3.62E+03 1.08E+01 3.00E-09 4.71E-06 3.63E+01 1.08E+01 3.00E-09 4.70E-04

0.9 4.07E+03 1.08E+01 3.00E-09 4.19E-06 4.08E+01 1.08E+01 3.00E-09 4.18E-04

1 4.52E+03 1.08E+01 3.00E-09 3.77E-06 4.54E+01 1.08E+01 3.00E-09 3.76E-04
D-1

APPENDIX D. THERMAL EVALUATIONS FOR FUEL 

RECONFIGURATION CATEGORIES

D.1 THERMAL ANALYSES

The objective of the thermal analysis was to identify cask temperature variations resulting from 

fuel configuration changes that can affect the ability of transportation packages and storage 

casks to comply with regulatory thermal limits. This appendix describes a technical approach for 

thermal analyses of fuel reconfigurations in transportation packages and storage casks and 

provides reference values for such analyses. The term “fuel reconfiguration” refers to any 

change to the storage and transportation system nominal intact fuel assembly configuration 

used for the basis of cask certification. The analyzed fuel reconfiguration categories are: (1) 

cladding failure where fuel fragments and particulates from multiple spent nuclear fuel (SNF)

rods can relocate near to and far from fuel cladding breach locations; (2) rod/assembly 

deformation; and (3) changes to assembly axial alignment.

D.1.1 Applicable Codes, Guidance, Regulatory Criteria

During normal and off-normal conditions of storage and accident conditions, the cask must 

reject the decay heat to the environment without exceeding the operational temperature ranges 

of the components important to safety.

To maintain fuel rod integrity for normal conditions of storage the fuel must be maintained at a 

sufficiently low temperature to preclude fuel rod cladding deterioration. To ensure fuel rod 

integrity, the maximum allowable cladding temperature under normal conditions of storage and 

short-term loading operations is 400°C (752°F) for zirconium-based alloys and stainless steel 

cladding [52]. The maximum cladding temperature under off-normal and accident conditions 

must remain below 570°C (1,058°F) for zirconium-based alloys and stainless steel cladding [52]. 

In addition, all components important to safety (confinement, shielding, criticality, and heat 

removal) must remain within their individual material temperature limits.

It is important to note that the thermal characteristics and response of the cask are also integral 

to the cask structural requirements. The thermal characteristics can affect the failure limits of 

components (yield strength, creep, and fatigue) and the mechanical loading of components 

(thermally induced stresses and system pressure).

In transportation, the accessible surfaces of the cask must have temperatures at or below 85°C 

(185°F) for an exclusive use shipment (10 CFR 71.43 [g]).

D.2 USE OF SOFTWARE

The thermal analysis used the Coolant Boiling in Rod Arrays—Spent Fuel Storage (COBRASFS) cycle 3 and the RADGEN cycle 3 computer codes [53]. COBRA-SFS has been validated 

and used extensively to analyze SNF storage casks and transportation packages [54], [55]. The 

software was developed by Pacific Northwest National Laboratory (PNNL) and tailored 

specifically for transportation package/storage cask analysis. The code includes convection, 

conduction and thermal radiation heat transfer and can accommodate a range of environmental 

boundary conditions, fuel assemblies and transportation package/storage cask designs. View 

factor information for the assemblies, which is required by COBRA-SFS, is generated using the 
D-2

RADGEN code. The RADGEN code determines thermal radiation view factors analytically for 

each pin in an assembly as well as other cavity view factors [56].

Both the COBRA-SFS cycle 3 and RADGEN cycle 3 computer codes will be used “as-is”. Code 

modifications were not included within the work scope. Any code deficiencies or limitations 

found during the conduct of this work are reported and discussed.

D.3 GENERIC CASK AND REPRESENTATIVE FUEL ASSEMBLY 

DESCRIPTIONS

The gernic burnup credit-32 (GBC-32) package for SNF was previously developed as a 

computational benchmark problem for reactivity and burnup credit analysis [4]. A variant of the 

GBC-32 package suitable for thermal analysis, the GBC-32T, was developed and used in this 

study. Many aspects of the original GBC-32 package are retained; however, the GBC-32T 

package system model contains some modifications and additional details necessary for 

thermal analysis.

The GBC-32T package system is composed of the following:

• Spent fuel canister – contains up to 32 pressurized water reactor (PWR) spent fuel 

assemblies; placed inside the transportation overpack for off-site transportation or a 

storage cask for on-site dry storage

• Rail Transportation Package – package for rail transportation consisting of the spent fuel 

canister, the transportation overpack, and associated impact limiters

The GBC-32T system is intended for testing thermal analysis methodologies and for the general 

impact of system and modeling variations. While the GBC-32T system shares similarities to 

other transportation package/storage cask systems, it was not designed to meet manufacturing, 

vendor, or regulatory requirements for such physical systems. Because thermal performance is 

intimately coupled with structural design, analyses with this system should not be interpreted as 

defining performance limits for actual canisters at design-basis decay heat loadings.

The spent fuel canister is comprised of a steel cylinder with an internal basket that can contain 

up to 32 PWR assemblies. The spent fuel canister is placed inside the transportation overpack 

for rail transportation. Two orientation cases are considered; horizontal geometry and vertical. 

Horizontal orientation corresponds to the configuration for transportation by rail, and the rail 

transportation package includes impact limiters. Vertical orientation of the rail transportation 

package is assumed for the package on a concrete pad awaiting loading onto a rail car, before 

placement of impact limiters.

Illustrations of the GBC-32T system are given in Figure D.1 to Figure D.5. The material 

specifications and system dimensions are given in Table D.1 and Table D.2, respectively.

D.3.1 Canister

The GBC-32T canister holds a basket with space for up to 32 PWR assemblies. The basket is 

constructed with stainless steel walls with neutron absorber plates sandwiched in between, as 

shown in Figure D.5. The neutron absorber plates are a composite with an outer layer of 

aluminum that sandwiches a BORAL® panel. The neutron absorber plate was represented as

extending from the top to the bottom of the basket. Stainless steel support pieces connect the 

basket to the canister wall as shown in Figure D.4. These support pieces are welded to both 
D-3

structures. They extend from the top of the basket to 5.08 cm (2 in.) above the bottom of the 

basket. The bottom of the basket sits on the bottom of the canister. Rectangular openings 

through the bottom of the basket walls measure 5.08 cm (2 in.) high by 19.69 cm (7.752 in.) 

wide. The canister is evacuated and backfilled with helium at 2 atm abs at a reference 

temperature of 21°C (70°F). At an average gas temperature of 177°C (350°F), the canister 

would be at 3.06 atm abs.

D.3.2 Assemblies

The canister is fully loaded with 32 Westinghouse (W) 17×17 optimized fuel assembly (OFA)

SNF assemblies. The assembly contains 264 fueled rods with an active fuel length of 144 in. All 

assemblies are assumed to be radially centered within their respective basket cells. This is 

because of a limitation of the current version of the RADGEN code. The assemblies sit upon

bottom fuel spacers which are 8.89 cm (3.5 in.) tall and made of stainless steel. Top fuel 

spacers are present above the fuel assemblies to prevent axial shifting of the assemblies during 

transport. The length of the spacers varies based on the assembly length and canister design. 

For the W 17×17 OFA assembly, the top spacer length used is 27.27 cm (10.735 in.).

D.3.3 Rail Transportation Package

The canister resides inside a transportation package. The gas in between the canister and 

package is helium at 2 atm (at 21°C). At an average gas temperature of 93°C (200°F), the 

canister would be at 2.49 atm abs. The transportation package walls include a 25.4 cm (10 in) 

thick inner wall made of stainless steel. Outside of the stainless steel wall is a neutron shield 

consisting of resin epoxy placed inside 36 can-like structures of carbon steel that are uniformly 

spaced around the package, as shown in Figure D.3. The neutron shield starts 30.48 cm (12 in) 

above the bottom of the package and ends 30.48 cm (12 in.) below the top of the package, for a 

total 457.5 cm (180.125 in.) axial length. The top lid and package base (bottom) are made of 

stainless steel and are 7 and 11 in. thick, respectively. The impact limiters are made of foamed 

aluminum with a bulk thermal conductivity of 4.68 W m-1 K-1

. The impact limiters are 50.8 cm 

(20 in.) thick in both the axial and radial direction. The impact limiters cover the first 30.48 cm 

(12 in.) of the transportation package in the axial direction, as shown in Figure D.6.
D-4

Figure D.1 Package component identification
D-5

Figure D.2 Illustration of canister and basket

Figure D.3 Illustration of canister and basket within transportation package
D-6

Figure D.4 Illustration of one-quarter of basket with single assembly

 

Figure D.5 Illustration of basket cell with assembly
D-7

Figure D.6 Illustration of transportation package in horizontal orientation

Table D.1 Package material summary

Part Material

Transportation package outer wall Carbon steel (1010)

Transportation package neutron shield Epoxy resin

Transportation package inner wall Stainless steel 304

Transportation package lid and base Stainless steel 304

Transportation package impact limiters Al foam, closed cell

Canister Stainless steel 304

Basket walls Stainless steel 304

Basket supports Stainless steel 304

Basket neutron absorber BORAL® plates—Composite 

B4C & Al

Assembly spacers Stainless steel 304

Assembly rod cladding Zicaloy-4

Canister void space Helium—2 atm (at 21°C)

Transportation package void space Helium—2 atm (at 21°C)

Impact Limiters
D-8

Table D.2 GBC-32T package dimensions for thermal analysis

Parameter cm in.

Fuel Assembly (W 17x17 OFA)

Active fuel length 365.76 144.00

Bottom hardware length 15.00 5.9055

Top hardware length 25.04 9.8595

Total assembly length 405.80 159.765

Pin pitch 1.26 0.4960

Clad outside diameter 0.914 0.3600

Guide/thimble tube outside diameter 1.2040 0.4740

Fuel Spacer

Total height—top 27.27 10.735

Material thickness—top 1.27 0.50

Total height—bottom 8.89 3.50

Material thickness—bottom 1.27 0.50

Basket

Total basket length 436.88 172.00

Basket length between openings 431.80 170.00

Gap between basket and canister top 5.08 2.00

Length of opening at bottom of basket 5.08 2.00

Width of opening at bottom of basket 19.69 7.752

Neutron absorber plate length 431.80 170.00

Basket cell inside width 22.00 8.6614

Basket wall material thickness 0.750 0.2953

Neutron absorber plate total thickness 0.257 0.1010

Neutron absorber plate width 19.05 7.50

Basket to canister wall support thickness 1.76 0.6916

Basket to canister wall support length 431.80 170.00

Canister

Total cavity length 441.96 174.00

Total exterior length 472.44 186.00

Top lid thickness 25.4 10.00

Base thickness 5.08 2.00

Outer shell thickness 1.27 0.50

Outer shell inside diameter 175.00 68.8976

Outer shell outside diameter 177.54 69.8976

Transportation Package

Total cavity length 472.76 186.125

Total outside height 518.48 204.125

Top lid thickness 17.78 7.00

Base thickness 27.94 11.00

Total outside shell thickness 34.29 13.50

Stainless steel inner shell thickness 25.40 10.00

Neutron shield shell thickness 7.62 3.00

Carbon steel outer shell thickness 1.27 0.50

Cavity inside diameter 178.18 70.1476

Outer shell outside diameter 246.76 97.1476
D-9

D.4 NOMINAL THERMAL MODELING SETUP WITH THE COBRA-SFS 

CODE

Thermal models of the GBC-32T were developed for the COBRA-SFS cycle 3 code. Two 

orientations, vertical and horizontal, were modeled and are described in Sect. D.4.1. Using 

COBRA-SFS convention, the package is divided axially in three zones: top plenum, bottom 

plenum, and the assembly zone, as illustrated in Figure D.7. In the horizontal orientation, the 

impact limiters are included in the top and bottom plenum. The COBRA-SFS modeling of each 

zone is discussed in Sect. D.4.2 and D.4.3.

Figure D.7 Illustration of COBRA-SFS modeling zones
D-10

D.4.1 Package Orientation Cases

D.4.1 Case 1—Rail Transportation Package in Vertical Configuration on Pad

The vertical configuration represents the transportation package waiting to be placed onto the 

rail transportation car or after being removed from the rail transportation car and placed back on 

the ground. In general, long term storage casks are vented, allowing air—t he ultimate heat 

sink—to cool the internal canister. The transportation package, in contrast to long-term storage 

casks, does not allow the ultimate heat sink to contact the internal canister and contains 

additional layers of material between the canister and ultimate heat sink. Therefore, the 

transportation package in a vertical configuration would be hotter than a vertical storage cask

and represents the bounding case for vertical orientations. For this reason, the vertical 

orientation of the transportation package, as illustrated in Figure D.8, was selected for analysis.

Figure D.8 Illustration of transportation package in vertical orientation

The assemblies are centered radially within each basket cell. The canister is centered radially 

within the transportation package and rests on the bottom plate of the transportation package. 

The transportation package rests on a 25.4 cm (10 in.) thick concrete slab. Under the concrete 

slab is earth.
D-11

D.4.1.2 Case 2—Rail Transportation Package in Horizontal Configuration

The horizontal configuration represents the package during transportation. The assemblies are 

centered radially within each basket cell. The canister is centered horizontally and rests on the 

lower surface of the transportation package inner shell. Impact limiters are placed on the ends 

of the transportation package, as illustrated in Figure D.6.

D.4.2 Assembly Zone

D.4.2.1 Flow Channel and Solid Material Discretization

The GBC-32T was divided into 52 “assemblies” and 516 solid conduction nodes. The full 

package was modeled, and no symmetry was assumed. Illustrations of the discretization are 

provided in Figure D.9 to Figure D.11.

Of the 52 assemblies, 32 model fuel assemblies as detailed rod-and-subchannel arrays, and the 

remaining 20 model the open flow channel regions between the fuel basket and the canister. All 

289 rods and 324 flow subchannels between rods were modeled for each of the 32 fuel 

assemblies. Each assembly contains 25 guide/thimble/instrument tubes.

The basket walls consist of 304 solid conduction nodes of stainless steel 304, and 152 nodes 

are used to represent the neutron absorber plates between the basket walls. Each of the 20 

basket-to-canister supports is modeled with a node. The canister is divided into eight sections 

circumferentially and two sections radially, resulting in 16 nodes. Finally, the transportation 

package is divided into eight sections circumferentially and five sections radially (three across 

the stainless steel wall, one across the neutron shielding material, and one across the carbon 

steel outer wall), resulting in 40 nodes.

As illustrated in Figure D.7 for the vertical orientation, the assembly zone extends from the top 

of the cutout at the bottom of the basket (5.08 cm [2 in.] up from the canister base) to the top of 

the basket for a total of 4.318 m (170 in.). The assembly zone is divided into 44 uniformly 

spaced sections.

In the COBRA-SFS model, the fuel rods, guide/thimble tubes and instrument tube are modeled 

with the appropriate diameters and spacing. However, as discussed in Sect. D.4.2.5, the 

radiation view factors for the bundle assume all rods are of the same diameter. Future code 

enhancements of RADGEN could eliminate this modeling artifact.

Three modeling simplifications are employed in the assembly zone of the model. First, the fuel 

rods are modeled as extending from the bottom to the top of the assembly zone. This extends 

the axial length of the rods by approximately 48 cm (18.9 in.), relative to their actual overall 

length of 388.9 cm (153.1 in.). Second, the grid spacers are not explicitly modeled, and 

therefore their effect on heat transport is not modeled. This is a minor conservatism in the 

model, because the grid spacers do not substantially affect heat transfer within the assembly. 

The effect of the grid spacers on momentum transport in the helium gas, however, is not 

neglected. This is modeled with form loss coefficients appropriate to the hydrodynamic losses 

associated with the particular grid design. Finally, the small cruciform shaped channels formed 

between the neutron absorber panels of neighboring basket cells, as shown in Figure D.5, are 

not modeled as flow channels. However, the thermal radiation and conductive heat transfer from 

one basket wall, across the cruciform gap, to the neighboring basket wall are included in the 
D-12

model. The gas volume in the cruciform gaps is included in the canister free volume when 

determining gas mixture properties. 

Figure D.9 Nodalization illustration of package (not to scale)
D-13

 

Figure D.10 Nodalization illustration of one-quarter of package (not to scale)
D-14

Figure D.11 Nodalization illustration of a basket cell with assembly subchannels (not to 

scale)
D-15

D.4.2.2 Gap Thermal Resistances

Thermal resistances were included between certain solid nodes to account for gas gaps 

between surfaces. Table D.3 and 

Table D.4 summarize the assumed gap widths and the fill gas thermal conductivity for the 

vertical and horizontal package orientations, respectively. For the horizontal configuration, the 

gaps between the canister and package varied along the circumference. Figure D.12 illustrates

the positions of the gaps, A through D, specified in Table D.3 and 

Table D.4.

Table D.3 Assumed gap thermal resistances, vertical orientation

Location

Gap width Gas Thermal Conductivity

(cm) (in.) (W/m K) (Btu/h ft °F)

Basket SS wall—

Neutron Absorber Plate 0.0254 0.01

0.2155

(He, 3.06 atm. 204°C)

0.1245 

(He, 3.06 atm. 400°F)

Outer Canister Shell—

Inner Package Wall (A–D) 0.3175 0.125

0.1793

(He, 2.49 atm. 93.3°C)

0.1036

(He, 2.49 atm. 200°F)

Package Outer SS Shell—

Package Inner Neutron Shield 0.0254 0.01

0.0267

(He, 1.00 atm. 38.8°C)

0.0154 

(He, 1.00 atm. 100°F)

Note: SS = stainless steel.

Table D.4 Assumed gap thermal resistances horizontal orientation

Location

Gap width Gas thermal conductivity

(cm) (in.) (W/m K) (Btu/h ft °F)

Basket SS wall—

Neutron Absorber Plate 0.0254 0.01

0.2155

(He, 3.06 atm. 204°C)

0.1245 

(He, 3.06 atm. 400°F)

Outer Canister Shell—

Inner Package Wall

Top Solid Nodes (A) 0.5556 0.2188

0.1793

(He, 2.49 atm. 93.3°C)

0.1036

(He, 2.49 atm. 200°F)

Outer Canister Shell—

Inner Package Wall

Side Upper Solid Nodes (B) 0.3969 0.1563

0.1793

(He, 2.49 atm. 93.3°C)

0.1036

(He, 2.49 atm. 200°F)

Outer Canister Shell—

Inner Package Wall

Side Lower Solid Nodes (C) 0.2381 0.0938

0.1793

(He, 2.49 atm. 93.3°C)

0.1036

(He, 2.49 atm. 200°F)

Outer Canister Shell—

Inner Package Wall

Bottom Solid Nodes (D) 0.0794 0.0313

0.1793

(He, 2.49 atm. 93.3°C)

0.1036

(He, 2.49 atm. 200°F)

Package Outer SS Shell—

Package Inner Neutron Shield 0.0254 0.01

0.0267

(He, 1.00 atm. 38.8°C)

0.0154 

(He, 1.00 atm. 100°F)

Note: SS = stainless steel.
D-16

Figure D.12 Canister-package gap, outer wall, and failed assembly location 

identification

D.4.2.3 Assembly Decay Heat, Profile, and Nuclide Inventory

The package is uniformly loaded with 32 W 17×17 OFA assemblies assumed to have an 

average burnup of 65 GWd/MTU. There are 264 fueled rods, 24 locations that are guide/thimble

tubes and one instrument tube, as shown in Figure D.5. Each assembly has the same total 

decay heat, decay heat axial power profile, and radioisotope inventory.
D-17

The total decay heat per assembly is given in Table D.5 for various decay times and shown in 

Figure D.13. The assembly decay heat and radioisotope inventory were calculated using Oak 

Ridge Isotope Generation and Depletion Code–Automatic Rapid Processing (ORIGEN-ARP)

[57] assuming a PWR assembly with the following characteristics:

• 0.426 MTU,

• 65 GWd/MTU burnup, and

• 5 wt % 235U initial enrichment.

Table D.5 Assembly total decay heat as a function of decay time

Decay time

(years)

Decay heat

(W/assembly)

5 1800.

20 870.5

40 445.5

60 344.0

100 233.2

Figure D.13 Assembly total decay heat as a function of decay time

All assemblies have the axial decay heat profile given in Table D.6. The axial profile is derived 

from profiles representative of PWR fuel with >30 GWd/MTU burnup [58]. The profile power 

peaking factors are linearly interpolated within the COBRA-SFS code between points in the 

input table. The inventory of fission gas products in the assembly is given in Table D.7.
D-18

Table D.6 Assembly axial decay heat profile

Normalized height 

with respect to

active fuel length

Normalized 

height with 

respect to

assembly zone

Power 

profile

- 0.000 0.000

0.000 0.044 0.000

0.028 0.067 0.652

0.083 0.114 0.967

0.139 0.161 1.074

0.194 0.208 1.103

0.250 0.255 1.108

0.306 0.302 1.106

0.361 0.349 1.102

0.417 0.397 1.097

0.472 0.444 1.094

0.528 0.491 1.094

0.583 0.538 1.095

0.639 0.585 1.096

0.694 0.632 1.095

0.750 0.679 1.086

0.806 0.726 1.059

0.861 0.773 0.971

0.917 0.820 0.738

0.972 0.867 0.462

1.000 0.891 0.000

- 1.000 0.000

Note: The first column is the power profile, with respect to 

the active fuel length from DOE-RW-0472 [58]. In the 

second column the axial power profile was shifted and renormalized to correctly position the profile within the 

assembly channels.

Table D.7 Assembly inventory of fission product gases and volatile species

Decay

time

(years)

Fission product gases

(moles/assembly)

Volatile fission products

(moles/assembly)

Xe Kr He H Cs I Ru

5 33.54 3.27 1.05 0.013 15.18 1.32 19.03

20 33.54 3.16 1.33 0.008 13.20 1.32 19.01

40 33.54 3.11 1.64 0.005 11.54 1.32 19.01

60 33.54 3.09 1.90 0.004 10.49 1.32 19.01

100 33.54 3.09 2.35 0.003 9.41 1.32 19.01
D-19

Figure D.14 Power profile with respect to assembly zone height

D.4.2.4 Flow Pressure Loss Modeling

The pressure loss experienced by the flow is determined as the summation of the frictional 

losses and any form losses, as shown in Eq. (D.1). 

∆P = ∆Pfriction + ∆Pform (D.1)

The flow pressure losses because of obstructions or other local effects (form losses), across the

assembly zone, are prescribed to the model. These losses are specified as loss coefficients, 

Cd, as shown in Eq. (D.2).

∆Pform = Cd

1

2 ρV2 . (D.2)

where 

ρ = the fluid density, and 

V = the fluid velocity.

Table D.8 summarizes the position and value of the loss coefficients for the fueled assembly 

locations. The top and bottom of the channel have a loss coefficient of 0.5 to account for the 90 

degree turn in the flow. The bottom of the assembly has a loss coefficient of 0.4 representative 

of flow contraction from a large plenum into a small channel. The top of the assembly has a loss 

coefficient of 1.0 representative of flow expansion from a small channel to a large plenum. 

Finally, the losses across the eight spacer grids are modeled using a flow loss coefficient of 2.0, 
D-20

which is typical of vaned mixing grids. The empty flow channels around the periphery of the 

basket include only the top (0.5) and bottom (0.5) loss coefficients.

Table D.8 Modeled drag coefficients for assemblies

Position

from bottom

(cm) (in.) Location x/L Cd

431.8 170.00 Top Channel 1.00 0.50

409.6 161.27 Top Assembly. 0.95 1.00

364.5 143.51 Grid 0.84 2.00

319.4 125.76 Grid 0.74 2.00

274.3 108.01 Grid 0.64 2.00

229.3 90.26 Grid 0.53 2.00

184.2 72.51 Grid 0.43 2.00

139.1 54.76 Grid 0.32 2.00

94.0 37.00 Grid 0.22 2.00

48.9 19.25 Grid 0.11 2.00

3.8 1.50 Bottom Assembly 0.01 0.40

0.0 0.00 Bottom Channel 0.00 0.50

Notes: x/L = normalized length of assembly zone;

Cd = loss coefficient.

The pressure loss because of frictional losses across the assembly zone is calculated based on 

Eq. (D.3). The frictional pressure loss coefficient , f, is modeled as the maximum of either the 

predicted laminar or turbulent frictional pressure loss coefficients, as shown in Eq. (D.4), that is

calculated by the code.

∆Pfriction= ffL

D

1

2 ρV2 (D.3)

f=max�0.1∙Re-0.3 , 96∙Re-1.0� (D.4)

where

L = the axial segment length,

D = the characteristic channel diameter, and

f = the frictional pressure loss coefficient

Re = the local Reynolds number in the flow channels

D.4.2.5 Radiation View Factors

The RADGEN code was used to determine gray body radiation view factors within the fuel rod 

assembly and between the basket walls and the fuel assembly. The view factors generated by 

RADGEN are provided as input to COBRA-SFS, and are used in the energy equation solution to 

calculate rod-to-rod, rod-to-wall, and wall-to-wall radiation heat transfer based on local surface 

temperatures of the rods and walls at each axial level in the assembly zone.

Black body radiation view factors were determined for the 16 open channels around the 

periphery of the basket that did not contain fuel assemblies. The RADGEN code was used to 

determine the black body view factors through entering the geometry information and specifying 

the emissivity as 1.0. These black body view factors are included in the COBRA-SFS input, and 

the COBRA-SFS code internally converts the black body view factors into gray body view 
D-21

factors, based on the specified surface emissivity values defined for the nodes representing the 

surfaces enclosing the flow channel.

D.4.2.6 Internal Convective Heat Transfer

The heat transfer between the fluid flow channels and the rods or wall surface nodes is modeled 

with a user-specified heat transfer coefficient correlation, expressed as a Nusselt number, Nu. 

The transition between laminar and turbulent flow conditions is captured automatically in the 

code by the convention of selecting the local maximum of the user-specified laminar and 

turbulent flow correlations. For this geometry, the Dittus-Boelter correlation is specified for 

internal, fully developed, turbulent flow; and for laminar flow, a constant Nusselt number of 3.66 

is specified, as shown in Eq. (D.5). The value of 3.66 is based on validation of the COBRA-SFS

code against experimental data obtained from numerous full-scale tests of SNF casks loaded 

with spent fuel.

Nu=max�0.023∙Re0.8Pr

0.4 , 3.66� (D.5)

where

Pr = Prandtl number

D.4.2.7 Package Outer Surface Thermal Boundary Condition

The ambient still air temperature is assumed to be 38°C (100°F). 

For the vertical package orientation, the convective thermal boundary condition on the vertical 

sides of the package is based on natural convection along a vertical flat plate [59] and is shown 

in Eq. (D.6), where RaL is the Rayleigh number. The region of applicability for Eq. (D.6) is for 

RaL < 109

. The convective thermal boundary condition on the horizontal top of the package is 

based on natural convection from a horizontal, upward facing heated plate [59], as shown in Eq. 

(D.7). The region of applicability for Eq. (D.7) is for RaL between 107 and 1011.

Nu=0.10∙RaL

0.333 , (D.6)

Nu=0.15∙RaL

0.333 . (D.7)

For the horizontal cask orientation, the convective thermal boundary condition on the curved 

outer walls of the cask was approximated as follows. The top two solid nodes, labeled as “1” in

Figure D.12, were approximated as horizontal upward-facing flat plates. The boundary condition 

for these two solid nodes was modeled based on natural circulation using Eq. (D.7). The four 

solid nodes on the side of the package, labeled as “2” in Figure D.12, were approximated as 

vertical plates. The boundary condition for these four solid nodes was modeled based on natural 

circulation using Eq. (D.6). Finally, the bottom two slabs, labeled as “3” in Figure D.12, were 

approximated as horizontal downward-facing flat plates. The boundary condition for these two 

solid nodes was modeled based on natural circulation using Eq. (D.8), and is applicable for RaL

between 105 and 1010 [59]. The convective thermal boundary condition on the vertical ends of 

the impact limiters of the package is based on natural convection along a vertical flat plate and 

is shown in Eq. (D.6).
D-22

Nu=0.27∙RaL

0.25 . (D.8)

Insolation was applied to the external surfaces of the package. For the vertical orientation case, 

the imposed heat load is 387 W/m2 (123 Btu/h-ft2

) on the top lid of the transportation package 

and 97 W/m2 (30.7 Btu/h-ft2

) on the transportation package side walls. For the horizontal 

orientation case, the imposed heat load is 194 W/m2 (61.5 Btu/h-ft2

) on the curved horizontal 

surfaces of the package, and 97 W/m2 (30.7 Btu/h-ft2

) on the vertical outer faces of the ends of 

the impact limiters. The total insolation for the vertical package was 5.12 kW and 8.15 kW for 

the horizontal configuration. 

Heat rejection via thermal radiation was included for the outer surfaces of the package. For 

thermal radiation to ambient, the outer surface of the package is assumed to have an emissivity 

of 0.80.

D.4.3 Upper and Lower Plenum

The 32 fueled assemblies and 16 periphery flow channels are connected at the top and bottom 

via upper and lower plenums. The upper and lower plenums allow for fluid exchange between 

assemblies and also include modeling the heat loss paths through the top and bottom of the 

package. The fluid space in the upper and lower plenum is modeled as completely mixed with a 

single fluid temperature.

D.4.3.1 Vertical Orientation Plenum Modeling

For the package in the vertical orientation, the upper plenum (see Table D.9) was divided into 

nine axial sections to represent the layers of the canister and package lids. The first section of 

the upper plenum represents the helium gas region above the basket connected to all assembly 

zone channels. The second region accounts for axial heat conduction from the canister walls to 

canister lid. The 10 in. thick canister lid is represented with two nodes (Regions 3 and 4), and 

Region 5 represents the gap between the top of the canister lid and the inner surface of the 

package lid. Both conduction and thermal radiation are modeled across this gap. Three nodes 

(Regions 6 through 8) are used to represent conduction through the package lid. Region 9 

represents the surface of the package lid and is connected to the appropriate boundary type to 

represent free convection and thermal radiation to the environment from this horizontal surface.
D-23

Table D.9 Upper plenum axial sections, vertical orientation

Region Section

length

(cm)

length

(in.)

1 conduction through helium gas 5.08 2.00

2 conduction through canister walls to canister 

lid

9.982 3.93

3 conduction through canister lid (1/2) 12.7 5.00

4 conduction through canister lid (1/2) 12.7 5.00

5 conduction through Nitrogen gap 0.318 0.125

6 conduction through package lid (1/3) 5.918 2.33

7 conduction through package lid (1/3) 5.918 2.33

8 conduction through package lid (1/3) 5.918 2.33

9 convection off the top horizontal surface - -

Similarly, the lower plenum (see Table D.10) was divided into 10 sections. The first section 

represents the helium gas region beneath the basket that connects to all assembly zone 

channels. Region 2 represents paths for conduction from the basket and canister wall nodes to 

the canister base. The effective length for this heat transfer path is from the center of the first 

axial node of the basket slabs in the assembly zone. Regions 3 and 4 represent conduction 

through the 2 in. thick canister base, and Regions 5 through 7 represent conduction through the 

package base. Regions 8 and 9 represent conduction through the concrete pad, and Region 10 

models heat transfer through the underlying soil to the earth sink temperature. This is modeled 

with a typical representative value of 4.4°C (40°F). Because of the extreme weight of the entire 

package, contact conductance between adjacent components of the base is assumed to be 

nearly perfect, and thermal resistance because of contact gaps is not included in this model.

Table D.10 Lower plenum axial sections, vertical orientation

Region Section

length

(cm)

length

(in.)

1 conduction through helium gas 5.08 2.00

2 conduction through basket and canister 

walls to canister base

9.982 3.93

3 conduction through canister base (1/2) 2.54 1.00

4 conduction through canister base (1/2) 2.54 1.00

5 conduction through package base (1/3) 9.322 3.67

6 conduction through package base (1/3) 9.322 3.67

7 conduction through package base(1/3) 9.322 3.67

8 conduction through concrete pad (1/2) 12.7 5.00

9 conduction through concrete pad (1/2) 12.7 5.00

10 conduction through earth 182.88 72

D.4.3.2 Horizontal Orientation Plenum Modeling

As similar to the vertical orientation, the upper and lower plenum of the package in the 

horizontal orientation were each divided into 11 axial sections to represent the layers of the 

canister, package, and impact limiters (see Table D.11 and Table D.12). The first section 

represents the helium gas region between the basket and the canister ends. The second region 

accounts for axial heat conduction from the canister walls and the basket in the lower plenum to
D-24

the canister lid. The canister lid is represented with two nodes (Regions 3 and 4), and Region 5 

represents the gap between the canister lid or base and the inner surface of the package lid or 

base. Both conduction and thermal radiation are modeled across this gap. Three nodes 

(Regions 6 through 8) are used to represent conduction through the ends of the package. Two 

nodes (Regions 9 and 10) represent conduction across the impact limiters placed on the ends of 

the package. Finally, Region 11 represents the surface of the impact limiter and is connected to 

the appropriate boundary type to represent free convection and thermal radiation to the 

environment from this vertical surface.

Table D.11 Upper plenum axial sections, horizontal orientation

Region Section

length

(cm)

length

(in.)

1 conduction through helium gas 5.08 2.00

2 conduction through canister walls to canister lid 9.982 3.93

3 conduction through canister lid (1/2) 12.7 5.00

4 conduction through canister lid (1/2) 12.7 5.00

5 conduction through Nitrogen gap 0.159 0.0625

6 conduction through package lid (1/3) 5.918 2.33

7 conduction through package lid (1/3) 5.918 2.33

8 conduction through package lid (1/3) 5.918 2.33

9 conduction through impact limiter (1/2) 25.4 10.00

10 conduction through impact limiter (1/2) 25.4 10.00

11 convection off the vertical surface - -

Table D.12 Lower plenum axial sections, horizontal orientation

Region Section

length

(cm)

length

(in.)

1 conduction through helium gas 5.08 2.00

2 conduction through basket to canister base 9.982 3.93

3 conduction through canister base (1/2) 2.54 1.00

4 conduction through canister base (1/2) 2.54 1.00

5 conduction through Nitrogen gap 0.159 0.0625

6 conduction through package base (1/3) 9.322 3.67

7 conduction through package base (1/3) 9.322 3.67

8 conduction through package base(1/3) 9.322 3.67

9 conduction through impact limiter (1/2) 25.4 10.00

10 conduction through impact limiter (1/2) 25.4 10.00

11 convection off the vertical surface - -

D.4.4 Material Properties

The modeled solid material properties are provided in Table D.13. The zircaloy emissivity is for 

that of highly oxided cladding. The helium properties are provided in Table D.14. Additional 

material properties related to fuel configuration changes are discussed in their respective 

sections.

To determine the effective thermal conductivity of the neutron shield, the shield was divided into 

two radial sections. The outer section is a 0.5 in. thick carbon steel wall on the outside of the 
D-25

package. The inner section is 3 in. thick and consists of 36 periodic regions of carbon steel and 

resin epoxy, as shown in Figure D.3. A single, effective, thermal conductivity was used to model 

the inner section. The effective thermal conductivity of the region was determined, using Eq. 

(D.9), to be 2.05 Btu h-1 ft-1 °F-1

.

keff=

kAl∙nsections∙tAl∙lnro

r

i

2∙π∙Lpath

+

kre∙nsections∙tre∙lnro

r

i

2∙π∙Lpath

, (D.9)

where:

ro = Outer radius of material

ri = Inner radius of material

Lpath = Radial thickness of material

nsections = Number of repeating sections of carbon steel and resin epoxy

tAl = Width of carbon steel

tre = Width of resin epoxy 

kre = Thermal conductivity of resin epoxy

kAl = Thermal conductivity of carbon steel

keff = Effective thermal conductivity

Table D.13 Modeled solid material properties

Material

Thermal conductivity

(Btu/h ft °F) Emissivity

Stainless Steel 304 5.901468+0.004870*T(R) 0.36

BORAL® 61.033037-0.011262*T(R) -

Carbon Steel 1010 28.89 0.8

Neutron Shield 0.289 -

Effective Neutron Shield 2.05 -

Concrete 1.068 -

Earth 0.347 -

Foamed Al Impact Limiters 2.70 0.8

Zircaloy 8.83 0.8

Uranium Dioxide 3.47 -
D-26

Table D.14 Helium gas properties

Temp.

(°F)

Enthalpy

(Btu/lbm)

Thermal

conductivity

(Btu/h ft °F)

Specific

heat

(Btu/lbm °F)

Specific

volume

(ft3/lbm)

Viscosity

(lbm/ft h)

0 573.2 0.0807 1.24 27.45 0.0432

50 635.3 0.0867 1.24 30.43 0.0464

100 697.3 0.0925 1.24 33.41 0.0494

150 759.4 0.0981 1.24 36.39 0.0524

200 821.4 0.1036 1.24 39.37 0.0553

300 945.5 0.1143 1.24 45.34 0.0610

400 1069.7 0.1245 1.24 51.30 0.0665

500 1193.8 0.1344 1.24 57.26 0.0718

600 1317.9 0.1440 1.24 63.22 0.0769

700 1442.0 0.1533 1.24 69.18 0.0819

800 1566.1 0.1624 1.24 75.14 0.0868

900 1690.2 0.1713 1.24 81.10 0.0916

1000 1814.3 0.1800 1.24 87.07 0.0963

1500 2434.9 0.2212 1.24 116.87 0.1186

2000 3055.5 0.2593 1.24 146.68 0.1393

D.4.5 Computational Setup

The COBRA-SFS convergence criteria were set to 0.0003 for the total energy error and 0.0001 

for the fluid enthalpy error. All simulations were performed as steady state calculations.

D.5 DESCRIPTION OF THE FAILED FUEL CATEGORIES AND 

CALCULATION MODELS

D.5.1 Category 1: Cladding Failure

D.5.1.1 Breached Spent Fuel Rods

This cladding failure scenario investigates the impact of the release of backfill and fission 

product gases into the canister gas space from the fuel rods. The rods are assumed to breach

in such a manner that the cladding remains in its nominal geometry (i.e., the ruptured cladding 

does not block flow paths). 

The assumed amount of gases released by the ruptured rods is 100 percent of the initial fill gas 

and 30 percent of the fission product gases generated within the fuel rods during operation 

(consistent with NUREG-1536 [2]). The impact of He, Xe, and Kr is included in the 

determination of the package gas phase mixture thermo-physical properties. Volatile 

radionuclides, such as Cs, I, and Ru are not included in determination of the mixture 

thermophysical properties. The gas is assumed to be homogeneously mixed within the canister.

The source term from the ruptured rods, with respect to decay heat, is modeled as remaining 

inside the rods. Although Xe, Kr, and radionuclides such as Cs may be released into the 

canister gas space, the decay heat source term is modeled as remaining inside the rods. 

COBRA-SFS does not contain the capability of radionuclide tracking/deposition or the capability 

for the gas phase to have volumetric heat generation. Also, these nuclides contribute a small 
D-27

percentage of the overall decay heat. Modeling all of the decay heat as staying within the fuel 

rods is conservative with respect to the peak cladding temperature prediction.

In the vertical package orientation, cases with 0%, 3.125%, 10%, 50%, and 100% of the total 

fuel rods developing breaches were investigated. The case with 3.125% of the fuel rods 

developing breaches coincides with the number of fuel rods in one assembly. 

Gas Inventory

The free space of the canister is estimated to be 6.30 m3

. To backfill the canister to 2 atm at 

21°C requires 521 mol of helium.

The as-manufactured free space of each fuel rod is estimated to be 16.0 × 10-6 m3. The initial 

helium backfill pressure is assumed to be 350 psig standard temperature and pressure. It 

follows that the total initial helium backfill of the fuel rods is approximately 4.67 mol per 

assembly or 149 mol per 32 assemblies.

The amount of fission product gases within each assembly was previously provided in

Table D.7. Table D.15 summarizes the total canister gas inventory and the gas mixture in the 

canister free space for various fractions of rod breach.

Table D.15 Total inventory and canister gas mixture for various rod failure fractions

Species

Total 

inventory

(mol)

Canister gas mixture (mol)

for rod failure fraction]

0.0 0.03125 0.10 0.50 1.00

He 722.4 521 526.1 537.5 603.4 685.7

Xe 1073.2 0 10.1 321.2 161.0 322.0

Kr 99.4 0 0.9 3.0 14.9 29.8

Total 1895.1 521 537.1 572.7 779.3 1037.5

Gas Mixture Properties

The initial canister pressure at 177°C (350°F) was previously given as 3.06 atm. For a given rod 

failure fraction, the canister average pressure, at 177°C (350°F), was determined using the ideal 

gas law. At a constant volume and temperature, the change in pressure caused by a change in 

the number of moles of gas can be determined using the ideal gas law, as shown in Eq. (D.10), 

where P is the absolute pressure, n is the number of moles of gas, and the subscripts 1 and 2 

indicate the initial and final states, respectively.

P2=P1

n2

n1

. (D.10)

The free volume of the canister was determined by summing the initial canister free volume and 

the free volume added by the ruptured fuel rods. The average canister gas density was then 

determined by summing the masses of the individual species and dividing by the free volume. 

At the pressure, P2, determined by Eq. (D.10), this canister average gas density would occur at 

177°C (350°F).
D-28

The canister was then assumed to be at the pressure, P2, determined by Eq. (D.11). The 

density of the gas mixture at various temperatures at this pressure was then determined by 

using the ideal gas law shown in Eq. (D.11), where T1 is 450 K (177°C or 350°F), T2 is the 

temperature at which the density, ρ2, is to be determined, and ρ1 is the canister average density.

ρ2=ρ1

T1

T2

. (D.11)

The thermal conductivity of the gas mixture was determined using Eqs. (D.12) through (D.14)

[60].

kmix= ∑ � kixi

xi+ ∑ (1-δij)φijxj n

j=1

� nn

ii , (D.12)

where 

φij=θij �1+2.41 �Mi-Mj��Mi-0.142Mj�

�Mi+Mj�

2 � , (D.13)

and

θij=

�1+�

ki

kj

�

1/2�

Mi

Mj

�

1/4�

2

23/2�1+Mi

Mj

�

1/2 , (D.14)

where

δ = Kronecker Delta (equals 1 for i=j, 0 otherwise)

n = Number of components in mixture

M = Molecular weight

x = Mole fraction

k = Thermal conductivity

φ = Parameter defined by Eq. (D.13)

θ = Parameter defined by Eq. (D.14)

μ = Viscosity

The viscosity of the gas mixture was determined using Eqs. (D.15) and (D.16) [60].

μmix= ∑ � kiμj

∑ (1-δij)φijxj n

j=1

� n

i , (D.15)

and

θij=

�1+�

μi

μj

�

1/2

�

Mj

Mi

�

1/4�

2

23/2�1+Mi

Mj

�

1/2 . (D.16)

The specific heat and enthalpy are also required as input for a COBRA-SFS model. The specific 

heat was determined by weight averaging the individual species specific heats. The enthalpy, as 

a function of temperature, was then determined using the gas mixture specific heat and using 

0 K as the reference point.

The properties for the individual gas species were obtained from the National Institute of 

Standards and Technology database for a temperature range of 17.8°C to 426.7°C [61]. 
D-29

The density, viscosity, and thermal conductivity of the gas mixtures were evaluated as functions

of temperature from 17.8°C to 426.7°C (0°F to 800°F). Table D.16 summarizes the evaluated 

canister gas mixture properties. When implementing the gas mixture properties into the 

COBRA-SFS models, an additional point at 906.7°C (1600°F) was included and was derived by 

linear extrapolation from the 426.7°C data point.

Table D.16 Evaluated gas mixture properties for various fractions of fuel rods that 

develop cladding breaches

Failed rod fraction (-) 0.00 0.03125 0.10 0.50 1.00

Free volume [m3] 6.299 6.303 6.308 6.313 6.316

Mole fraction He (-) 1.000 0.980 0.939 0.774 0.661

Mole fraction Xe (-) 0.000 0.019 0.056 0.207 0.310

Mole fraction Kr (-) 0.000 0.002 0.005 0.019 0.029

Average canister

pressure

(at 350°F) (atm) 3.06 3.16 3.36 4.58 6.09

Specific heat 

(kJ/kg 

K) 5.28 3.18 1.80 0.67 0.46

Densitya

(kg m-3)

Temp. 

(°C)

-17.8 0.583 0.979 1.850 6.861 13.008

37.8 0.479 0.804 1.519 5.635 10.684

93.3 0.406 0.683 1.289 4.781 9.064

148.9 0.353 0.593 1.119 4.152 7.871

204.4 0.312 0.524 0.989 3.669 6.955

260.0 0.279 0.469 0.886 3.287 6.231

315.6 0.253 0.425 0.802 2.976 5.643

371.1 0.231 0.388 0.733 2.720 5.156

426.7 0.213 0.357 0.675 2.504 4.747

Thermal

conductivity

(W m-1 K-1)

Temp. 

(°C)

-17.8 0.140 0.128 0.109 0.063 0.045

37.8 0.160 0.147 0.125 0.073 0.053

93.3 0.179 0.165 0.141 0.083 0.060

148.9 0.198 0.182 0.156 0.092 0.067

204.4 0.215 0.199 0.171 0.102 0.074

260.0 0.233 0.215 0.185 0.110 0.081

315.6 0.249 0.230 0.199 0.119 0.087

371.1 0.265 0.245 0.212 0.127 0.093

426.7 0.281 0.260 0.225 0.135 0.099

Viscosity

(Pa s)

Temp. 

(°C)

-17.8 1.79E-05 1.94E-05 2.12E-05 2.27E-05 2.24E-05

37.8 2.04E-05 2.23E-05 2.46E-05 2.68E-05 2.65E-05

93.3 2.29E-05 2.51E-05 2.79E-05 3.07E-05 3.06E-05

148.9 2.52E-05 2.78E-05 3.11E-05 3.45E-05 3.45E-05

204.4 2.75E-05 3.04E-05 3.41E-05 3.82E-05 3.83E-05

260.0 2.97E-05 3.30E-05 3.71E-05 4.18E-05 4.19E-05

315.6 3.18E-05 3.54E-05 4.00E-05 4.52E-05 4.55E-05

371.1 3.39E-05 3.78E-05 4.27E-05 4.86E-05 4.89E-05

426.7 3.59E-05 4.01E-05 4.55E-05 5.18E-05 5.22E-05 a At average canister pressure.
D-30

D.5.1.2 Damaged SNF

This cladding failure scenario investigates the impact of redistribution of SNF within the canister.

For the vertical package orientation, two configurations were investigated. First, a single 

assembly near the center of the basket, and indicated by “F” in Figure D.12, is assumed to have 

cladding failure to the extent to allow free movement of fuel particles and pellets within a basket 

cell. The whole assembly is assumed to form a debris bed that slumps and stays within the 

basket cell. A gas space is formed within the basket cell above the debris. A few cases are 

simulated assuming a range of debris bed porosities (debris bed heights). The second 

configuration is based on the same assumptions as the first configuration except all the 

assemblies are assumed to have clad failure.

Two configurations were also investigated for the evaluation of gross assembly failure with the 

package in a horizontal orientation. The first configuration assumes a single assembly fails, 

forms a debris bed, and slumps within a basket cell. The debris bed slumps downward onto the 

lower face of the basket cell wall, leaving a gas space above the debris bed. The second 

configuration is similar to the first configuration except all the assemblies are assumed to fail.

To model the material of a damaged fuel assembly, the assembly is represented as a block of 

homogeneous material within the basket cell. The material comprising each assembly is 

assumed to stay within its respective basket cell within the assembly zone of the model (see 

Figure D.7) and does not spill out into the plenum regions. The free space within the basket cell

that is not occupied by the assembly material is also modeled as blocks of homogeneous 

material with the properties of the canister gas phase. The discretization of these blocks of 

material (solid conduction nodes) is illustrated in Figure D.15 for the vertical and Figure D.16 for 

the horizontal case. The entire assembly zone, including the basket cell and reconfigured 

material, is divided into 44 sections in the axial direction (i.e., 9.81 cm sections). In the radial 

cross section of the basket cell, the material is divided into a uniform 4×4 grid (i.e., 5.5×5.5 cm 

sections).

The entire assembly, including the fuel, cladding, and spacer grids, was assumed to fail. The 

mass of each component was based on Table B.1. As the shape of the particulate debris is 

unknown, the packing fraction was varied from 0.612 to 0.313 to simulate a range of possible 

bed porosities. A packing fraction of 0.612 is near the limit for randomly packed spheres of the 

same size and results in half of the basket channel being filled with debris. At a packing fraction 

of 0.313, the debris completely fills the basket channel.

Heat transfer via thermal conduction was modeled between adjacent solid nodes representing

the debris, gas space, and basket walls. No contact resistance was modeled between the debris 

and the basket cell walls. Thermal radiation across the gas space was not included in the 

model. Any possible convective heat transfer within the gas space above the debris was not 

modeled. In addition, no gas flow through the debris was modeled. Not modeling these heat 

transfer paths is conservative and results in higher debris bed temperature predictions.
D-31

 

Figure D.15 Illustration of failed fuel basket cell nodalization for vertical package

orientation
D-32

 

Figure D.16 Illustration of failed fuel basket cell nodalization for horizontal package

orientation

The thermal conductivity of the particle bed, kbed, was determined using the method developed 

by Zehner and Bauer as described by Tsotsas and Martin [62]. The method used assumes 

stagnant gas within the debris bed and does not account for thermal radiation heat transfer 

within the bed. The method uses a deformation parameter, β, which allows for the extension of 

the model to beds of non-spherical particles. The deformation parameter is only a function of the 

bed porosity, α, a measureable quantity. For the current simulations, the bed porosity, which is 

one minus the packing fraction, is specified. The method has been successfully applied to a 

wide range of dispersed systems and conditions [62]. The average thermal conductivity of the 

solid debris, kd, was determined by volume weighting the individual components from Table B.1

using the thermal conductivity values given in Table D.13. The gas thermal conductivity, kg, was 

taken as that of the canister gas space at 204°C (400°F).

kbed=kg �1 − √1 − αα + √1 − αα �

2

NN� �

ββ

NN2

kkdd − kkgg

kkdd

llll � kkdd

ββ ∙ kkgg

� − ββ + 1

2 − ββ − 1NN �� , (D.17)

where: β= �

1-α

α �

10/9

and N=1-

β∙kg

kd

.

The gas space above and within the debris bed is assumed to have the same composition and 

material properties as that computed in Sect. D.5.1.1 to account for fuel failure of either 3.125% 

(for the single-assembly failure configuration) or 100% of the rods (for failure of all assemblies). 

For the gas space above and within the debris bed, the material properties are not specified as 
D-33

a function of temperature and instead are provided for an assumed temperature of 204°C

(400°F).

The decay heat, from the originally intact assembly, is applied uniformly, both axially and 

radially, to the debris bed. This assumes the decay heat source term relocates with the bulk 

material and stays within the basket cell. In reality, some species may volatilize or become 

suspended, transport to another location within the canister and possibly deposit onto other 

surfaces. This also assumes the debris becomes homogeneous (i.e., well mixed) with respect to 

the decay heat. In reality, there is a possibility that the debris may become stratified with heavier 

or smaller debris sinking with larger and lighter debris on top.

D.5.2 Category 2: Rod/Assembly Deformation

This fuel reconfiguration category investigates the impact of changes to the assembly lattice 

pitch.

Five cases were analyzed including the expansion and contraction of the lattice pitch. The 

assembly pitch is assumed to change uniformly throughout the rod array cross-section and 

along the length of the assembly. In all cases, the entrance and exit pressure losses and the 

pressure losses because of spacer grids are assumed to remain unchanged. Also, the rods are 

assumed to remain intact without gaseous leakage into the canister space, as all cladding 

failure scenarios are evaluated under Category 1.

The nominal pitch of the W 17×17 OFA assembly is 1.26 cm (0.496 in.), with a pitch-to-diameter 

ratio of 1.38. In the case of the vertical orientation, assuming the assembly is centered within 

the GBC-32T basket cell; this leaves a 1.013 cm (0.399 in.) gap around the periphery between 

the outer fuel rods and the basket wall. Currently, RADGEN, the code used to generate the 

rod-to-rod and rod-to-wall thermal radiation view factors, is limited to pitch-to-diameter ratios of 

less than √2. Therefore, the largest pitch considered was 1.28 cm (0.50 in., a pitch-to-diameter 

ratio of 1.4). The smallest pitch considered was 1.06 cm (0.42 in.) where the fuel rods and 

guide/thimble tubes are nearly in contact. Finally, intermediate pitches of 1.143 cm (0.45 in.)

and 1.234 cm (0.486 in.) were also considered to provide additional resolution of the trends.

D.5.3 Category 3: Changes to Intact Assembly Axial Alignment

This fuel reconfiguration category investigates the impact of changes to axial position of the fuel 

assembly within the basket.

Two cases were analyzed in which the end fittings that support the assembly within the canister 

were assumed to fail or deform, allowing the assemblies to shift upward or downward within 

their respective basket cells. For the first case, the assembly is relocated 27.27 cm (10.735 in.) 

toward the canister lid. For the second case, the assembly is relocated 8.89 cm (3.5 in.) toward 

the canister base. In both cases, the rods are assumed to remain in the nominal intact geometry 

without gaseous leakage into the canister space, as all cladding failure scenarios are evaluated 

under Category 1. 

An assumption and a modeling simplification were employed to model these cases. First, no 

heat conduction was modeled from contact between the fuel rods and the canister base or lid. 

This is a conservative assumption limiting the axial heat removal from the fuel rods. Second, the 

flow loss coefficient for the assembly inlet (or outlet), which is shifted into the plenum region, 
D-34

was added to the basket entrance (or exit) flow loss coefficient for the assemblies. This 

simplification was made because of the lack of detailed geometric information as to how the 

spacers would fail or how the assemblies would relocate into the plenums. This simplification 

only affects the vertical cask orientation cases because the horizontal cask orientation cases 

have very limited internal flow.

D.6 THERMAL ANALYSIS RESULTS

D.6.1 Discussion of Comparison Methodology

Two simulations were selected to represent the nominal intact configuration—one for the 

vertical package orientation and one for the horizontal package orientation as described in Sect.

D.4. The decay time selected was 40 years. All the simulations were performed under the 

nominal boundary conditions described in Sect. D.4.2 (38 ̊C [100°F]) in still ambient air with 

insolation).

The thermal impacts of fuel configuration changes were evaluated by comparing predicted 

component temperatures between the nominal intact configuration and the fuel reconfiguration. 

This approach helps identify the fuel reconfiguration scenarios that yield the largest temperature 

changes in the package components. 

An alternative approach would be to compare predicted package thermal response to regulatory 

limits. This approach, evaluating the margin to the regulatory limits, would require many 

simulations for multiple boundary conditions/scenarios (i.e., hot storage, cold storage, 

transportation, accident conditions). Comparing the margins to the regulatory limits would also 

be strongly coupled to the decay heat load of the package (decay time) and the particular 

materials and package design. Ultimately such comparisons may be required for real-world 

systems and may be informed by the results of the current study.

Specific component responses were selected to quantify the impacts of fuel configuration

changes. The maximum cladding temperature is reported for its importance for fission product 

retention and maintaining assembly integrity. The minimum cladding temperature is also 

reported as it relates to phenomena that may occur at low temperature such as embrittlement. 

The maximum outer surface temperature is reported for its regulatory importance in 

transportation systems. The maximum neutron absorber, basket, and neutron shield 

temperatures are presented as they are important to the safety function of the package.

Finally, as noted previously, the thermal response of the package is integral to the structural 

performance. This report does not discuss phenomena coupled with the structural performance 

such as thermal gradients, which cause thermal stresses, and thermal expansion/contraction, 

which can cause mechanical stresses and geometry alterations.

D.6.2 Nominal Intact Configuration Results

The predicted package temperatures within the assembly zone (see Figure D.7) are 

summarized in Table D.17. As noted previously, this package design was not thoroughly 

designed to meet all regulatory requirements for such real-world systems. The cladding, neutron 

absorber, basket stainless steel, and neutron shield materials are within the temperature limits 

generally applicable for such materials. The maximum package outer surface temperatures are 

greater than 85 ̊C; however, the 85 ̊C regulation refers to shaded conditions, whereas the 
D-35

current simulations include insolation. A personnel barrier could be included in the design to 

meet this regulation for transportation packages.

Table D.17 Thermal nominal intact configuration results summary

Case

ID

Decay

time

(years)

Total

decay

heat

(kW)

Temperature ( ̊C)

Max.

clad

Min.

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SS

Max.

neutron

shield

V.40 40 14.26 267 104 90 73 262 262 95

H.40 40 14.26 326 97 103 78 322 322 108

Temperature profiles, at a number of positions within the package are illustrated in Figure D.17

for the vertical package orientation case and in Figure D.18 for the horizontal package 

orientation case.

The vertical package orientation case has considerable convection (approx. 0.57 kg/min 

recirculating flow) in contrast to the horizontal package orientation case. This produces top 

peaked temperature profiles for the vertical package, while the horizontal package is center 

peaked. The convective flow, the lower total insolation, and the absence of the large impact 

limiters result in the vertical package orientation case generally having lower predicted 

temperatures than the horizontal case.

Figure D.17 Structure temperatures—vertical orientation
D-36

Figure D.18 Structure temperatures— horizontal orientation

The nominal intact configuration models were used to investigate the package thermal response 

as a function of fuel decay times. Decay times less than 20 years were not simulated. The 

particular package design being investigated is not suitable for such high heat loads. A nonuniform loading pattern with cooler assemblies is one possible way to incorporate higher decay 

heat load assemblies. However, such mixed loading configurations were not investigated as 

part of this study. The results are summarized in Table D.18. As shown in Table D.17, the 

results are provided as the change in temperature from their respective nominal intact 

configuration case (i.e., +176 for the maximum cladding temperature for the V.20 case indicates 

the predicted maximum cladding temperature is 176 ̊C greater than that of the nominal intact 

configuration case, V.40). The change in component temperature is nearly linear with respect to 

the total package decay heat. An additional set of simulations was performed to investigate the 

impact of the solar insolation on the package. The results shown in Table D.19 indicate that 

removing the insolation heat load lowered the maximum component temperatures 

approximately 8 ̊C to 15 ̊C. These simulations with varying decay times as well as the cases of 

no insolation were conducted to provide perspective for the results from the fuel reconfiguration

investigation.
D-37

Table D.18 Thermal results vs. decay time

Case

ID

Decay

time

(years)

Total

decay

heat

(kW)

Change in temperature (∆°C)

Max.

clad

Min.

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSb

Max.

neutron

shield

V.20 20 27.86 +176 +44 +27 +19 +176 +176 +32

V.40a 40 14.26 0 0 0 0 0 0 0

V.60 60 11.01 -45 -12 -8 -5 -44 -44 -9

V.100 100 7.46 -96 -26 -18 -10 -94 -94 -20

H.20 20 27.86 +175 +35 +33 +15 +175 +175 +37

H.40a 40 14.26 0 0 0 0 0 0 0

H.60 60 11.01 -51 -9 -9 -5 -51 -51 -10

H.100 100 7.46 -112 -20 -19 -10 -112 -112 -21 a

Nominal intact configuration. b

SS = stainless steel.

Table D.19 Thermal results vs. insolation heat load

Case

ID

Insolation

(kW)

Change in temperature (∆°C)

Max.

clad

Min.

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSb

Max.

neutron 

shield

V.40a 5.12 0 0 0 0 0 0 0

V.40.no 0.0 -10 -7 -8 -7 -9 -9 -8

H.40a 8.15 0 0 0 0 0 0 0

H.40.no 0.0 -8 -10 -14 -13 -9 -9 -15 a

Nominal intact configuration. b

SS = stainless steel.

D.6.3 Particle Bed Modeling Sensitivities

As the damaged fuel assembly configuration may deviate from that assumed for the analyses, 

additional simulations were performed to help understand the impact of the modeled effective 

bed thermal conductivity and the decay heat axial power profile in the particle bed.

The effective thermal conductivity of the particle bed was calculated using the methodology 

discussed in Sect. D.5.1.2. This methodology assumed the contents of the bed were broken into 

pieces much smaller than the model nodalization size (2.17 cm [0.85 in.]) for the bed and 

homogeneously mixed. If pieces of the assembly remained largely intact, such as the spacer 

grids or fragments of the cladding, they would provide non-interrupted heat conduction paths 

augmenting the effective axial and/or radial bed thermal conductivity. Conversely, the fuel 

thermal conductivity, one component of the particle bed effective thermal conductivity, 

decreases with burnup because of void formation and fission products. The impact of voids on 

thermal conductivity is accounted for in the model through the packing fraction. If the thermal 

conductivity of the fuel is reduced to zero in the model, the predicted bed effective thermal 

conductivity is decreased by 17%, 13% and 10% for bed packing fractions of 63%, 42%, and 

31%, respectively. Finally, different assembly types may contain different volumes of fuel, 

cladding, spacer grids, and free space that can affect the effective bed thermal conductivity.
D-38

To investigate the impact of the particle bed thermal conductivity, four additional cases were 

simulated where the effective particle bed thermal conductivity was increased or decreased by 

25% or 50%. The cases assume all of the fuel assemblies are damaged and configured into 

particle beds with packing fractions of 62.6%, filling half of the height of the basket. The 

transportation package is assumed to be in the vertical orientation. The results are presented as 

the change in structure temperatures with respect to the nominal intact fuel configuration case.

Variations in the particle bed effective thermal conductivity did affect the peak basket stainless 

steel and neutron absorber temperatures (see Table D.20). 

Table D.20 Thermal results vs. particle bed effective thermal conductivity

Case

ID

Debris 

packing 

fraction

Change in temperature (∆°C)

Max.

intact

clad

Min.

intact

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSc

Max.

neutron 

shield

V.40a NA 0 0 0 0 0 0 0

V.40.100pr.626pf.+50% 0.626 NA NA -3 +7 +99 +99 -2

V.40.100pr.626pf.+25% 0.626 NA NA -3 +7 +112 +112 -2

V.40.100pr.626pfb 0.626 NA NA -3 +7 +127 +128 -2

V.40.100pr.626pf.-25% 0.626 NA NA -3 +7 +146 +147 -2

V.40.100pr.626pf.-50% 0.626 NA NA -3 +7 +169 +169 -2 a

Nominal intact configuration. b

100% damaged fuel, nominal bed thermal conductivity. c

SS = stainless steel.

A uniform axial decay heat profile was applied to the particle bed. Depending on how the 

fragments of a failed fuel assembly settle within the basket, the decay heat power profile in the 

debris may be top or bottom skewed. To investigate the impact of the particle bed decay heat 

profile, two additional cases were simulated where the decay heat in the bed was either top or 

bottom peaked. For the top peaked profile case, the relative power profile linearly varied from 

0.75 at the bottom of the bed to 1.25 at the top of the bed. For the bottom peaked profile case, 

the relative power profile linearly varied from 1.25 at the bottom of the bed to 0.75 at the top of 

the bed. The cases assume all of the fuel assemblies are damaged and reconfigured into 

particle beds with packing fractions of 62.6%, filling half of the height of the basket. The 

transportation package is assumed to be in the vertical orientation.

The decay heat axial power profile had a relatively minor impact on the peak basket 

temperatures. The impact of the axial power distribution did not propagate to the outer package 

structures (see Table D.21).
D-39

Table D.21 Thermal results vs. particle bed axial power profile

Case

ID

Debris 

packing 

fraction

Change in temperature (∆°C)

Max.

intact

clad

Min.

intact

clad

Max. 

outer 

surface

Min. 

outer 

surface

Max.

neutron

absorber

Max.

basket 

SSc

Max.

neutron 

shield

V.40a NA 0 0 0 0 0 0 0

V.40.100pr.626pf.TopPeak 0.626 NA NA -3 +7 +138 +139 -3

V.40.100pr.626pfb 0.626 NA NA -3 +7 +127 +128 -2

V.40.100pr.626pf.BotPeak 0.626 NA NA -2 +7 +126 +126 -2 a

Nominal intact configuration.

b

Uniform axial power profile. c

SS = stainless steel.

D.6.4 Seal Temperature

The integrity of the canister seal under various scenarios of fuel configuration changes is a 

combined structural–thermal–materials problem and is seal design dependent. However, the 

conditions at the seal location are important in understanding the containment performance of 

the system.

The modeled GBC-32 storage/transport system is generic and does not have a seal design. The 

canister seal would be located toward the top edge of the canister. The COBRA-SFS model of 

the GBC-32 modeled the upper and lower regions of the package in a 1-D fashion as described 

in Sect. D.4.3. Thus, the simulation results provide an average temperature for the lid and other 

upper and lower plenum structures. Table D.22 provides the 1-D temperature results for the 

inner layer of the canister lid and base for the various scenarios investigated. However, a 

detailed thermal–structural simulation may be warranted for various fuel failure scenarios to 

understand the seal integrity for a particular design.
D-40

Table D.22 Summary of canister lid and base temperature changes because of fuel 

configuration changes

Case

ID

Rod 

Pitch

to Diam.

Ratio

Rods

Ruptured

(%)

Rods

Damaged

(%)

Packing

Fraction

(%)

Assem.

Axial

Shifting

(cm)

Change 

in Lid 

Temp.

(∆°C)

Change 

in Base 

Temp.

(∆°C)

V.40a 1.38 0 0 NA 0 0 0

V.40.3.125pr 1.38 3.125 0 NA 0 +1 +3

V.40.10pr 1.38 10.0 0 NA 0 -3 +10

V.40.50pr 1.38 50.0 0 NA 0 -11 +21

V.40.100pr 1.38 100.0 0 NA 0 -15 +28

V.40.3.125pr.313pf 1.38 3.125 3.125 31.3 0 -1 +3

V.40.3.125pr.320pf 1.38 3.125 3.125 32.0 0 -1 +3

V.40.3.125pr.417pf 1.38 3.125 3.125 41.7 0 -1 +4

V.40.3.125pr.574pf 1.38 3.125 3.125 57.4 0 -1 +4

V.40.3.125pr.626pf 1.38 3.125 3.125 62.6 0 -1 +4

V.40.100pr.313pf 1.38 100.0 100.0 31.3 0 -18 +65

V.40.100pr.417pf 1.38 100.0 100.0 41.7 0 -19 +75

V.40.100pr.626pf 1.38 100.0 100.0 62.6 0 -19 +96

V.40.116pdr 1.16 0 0 NA 0 -2 +12

V.40.125pdr 1.25 0 0 NA 0 0 +6

V.40.135pdr 1.35 0 0 NA 0 0 0

V.40.140pdr 1.40 0 0 NA 0 0 0

V.40.up 1.38 0 0 NA +27.27 3 -10

V.40.dn 1.38 0 0 NA -8.89 -1 +4

H.40a 1.38 0 0 NA 0 0 0

H.40.3.125pr 1.38 3.125 0 NA 0 0 0

H.40.10pr 1.38 10.0 0 NA 0 0 +1

H.40.50pr 1.38 50.0 0 NA 0 0 +2

H.40.100pr 1.38 100.0 0 NA 0 0 +3

H.40.3.125pr.313pf 1.38 3.125 3.125 31.3 0 0 +1

H.40.3.125pr.417pf 1.38 3.125 3.125 41.7 0 0 +1

H.40.3.125pr.626pf 1.38 3.125 3.125 62.6 0 0 +1

H.40.100pr.313pf 1.38 100.0 100.0 31.3 0 +2 +13

H.40.100pr.417pf 1.38 100.0 100.0 41.7 0 +2 +13

H.40.100pr.626pf 1.38 100.0 100.0 62.6 0 +2 +13

H.40.116pdr 1.16 0 0 NA 0 0 0

H.40.125pdr 1.25 0 0 NA 0 0 0

H.40.135pdr 1.35 0 0 NA 0 0 0

H.40.140pdr 1.40 0 0 NA 0 0 0

H.40.up 1.38 0 0 NA +27.27 +1 -11

H.40.dn 1.38 0 0 NA -8.89 0 +5

Note: NA = not applicable. aNominal intact configuration.




NUREG/CR-7203 A Quantitative Impact Assessment of Hypothetical Spent Fuel Reconfiguration 

in Spent Fuel Storage Casks and Transportation Packages