Clad Degradation  FEPs Screening Arguments 
Prepared for: 
U.S. Department of Energy 
Office of Civilian Radioactive Waste Management 
Office of Repository Development 
1551 Hillshire Drive 
Las Vegas, Nevada 89134-6321 
Prepared by: 
Bechtel SAIC Company, LLC 
1180 Town Center Drive 
Las Vegas, Nevada 89144 
Under Contract Number 
DE-AC28-01RW12101 
QA: QA 
ANL-WIS-MD-000008 REV 01 
March 2004 DISCLAIMER 
This report was prepared as an account of work sponsored by an agency of the United States Government. Neither 
the United States Government nor any agency thereof, nor any of their employees, nor any of their contractors, 
subcontractors or their employees, makes any warranty, express or implied, or assumes any legal liability or 
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Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, 
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expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof. 
Clad Degradation  FEPs Screening Arguments 
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March 2004 iv ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
EXECUTIVE SUMMARY 
Twenty-four features, events, and processes (FEPs) have been identified as being related to the 
degradation of cladding (including, in some cases, cladding and fuel pellet degradation) on 
commercial spent nuclear fuel during the post-closure period, with an emphasis on zirconiumalloyed 
cladding. This report also addresses the effect of some FEPs on both the cladding and all 
waste forms where it was considered appropriate to address the effects on both materials together 
(FEP 2.1.09.09.0A, 2.1.09.11.0A, 2.1.11.05.0A, 2.1.12.02.0A, and 2.1.12.03.0A). Defense spent 
nuclear fuel, including fuel with aluminum cladding, is addressed in the general waste form FEPs 
screening report and not in this report. Four FEPs have been identified as having sufficient 
probability of occurrence and having sufficient effect on the resulting radionuclide exposures to 
the reasonably maximally exposed individual as to be included in the TSPA-LA. Summaries of 
their inclusion are presented in this document. Eighteen FEPs have been identified as being 
excluded from the TSPA-LA because their omission would have low consequence (i.e., they 
would not significantly change the magnitude and time of the resulting radiological exposures to 
the reasonably maximally exposed individual or radionuclide releases to the accessible 
environment). Two FEPs have been excluded based on low probability. Arguments for their 
exclusion are presented in this report. 
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vi March 2004 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
CONTENTS 
Page 
1. PURPOSE........................................................................................................................1 
1.1 SCOPE.......................................................................................................................... 5 
1.2 FEPS IDENTIFICATION AND ANALYSIS................................................................. 6 
2. QUALITY ASSURANCE.........................................................................................................6 
3. USE OF SOFTWARE ...............................................................................................................7 
4. INPUTS.........................................................................................................................7 
4.1 DIRECT INPUTS............................................................................................................ 7 
4.1.1 Data...................................................................................................................... 7 
4.1.2 Parameters and Parameter Uncertainty................................................................ 7 
4.1.3 Other Model / Analyses Inputs and Technical Information............................... 10 
4.2 CRITERIA..................................................................................................................... 11 
4.3 CODES AND STANDARDS........................................................................................ 11 
5. ASSUMPTIONS....................................................................................................................11 
6. SCIENTIFIC ANALYSIS DISCUSSION...............................................................................11 
6.1 DEGRADATION OF CLADDING FROM WATERLOGGED RODS....................... 17 
6.2 DEGRADATION OF CLADDING PRIOR TO DISPOSAL ....................................... 19 
6.3 GENERAL CORROSION OF CLADDING................................................................. 20 
6.4 MICROBIALLY INFLUENCED CORROSION (MIC) OF CLADDING .................. 26 
6.5 LOCALIZED (RADIOLYSIS ENHANCED) CORROSION OF CLADDING........... 28 
6.6 LOCALIZED (PITTING) CORROSION OF CLADDING.......................................... 31 
6.7 LOCALIZED (CREVICE) CORROSION OF CLADDING ........................................ 36 
6.8 ENHANCED CORROSION OF CLADDING FROM DISSOLVED SILICA............ 38 
6.9 CREEP RUPTURE OF CLADDING............................................................................ 39 
6.10 INTERNAL PRESSURIZATION OF CLADDING..................................................... 42 
6.11 STRESS CORROSION CRACKING (SCC) OF CLADDING.................................... 44 
6.12 HYDRIDE CRACKING OF CLADDING.................................................................... 47 
6.12.1 Hydride Embrittlement from Zirconium Corrosion (of Cladding) .................... 50 
6.12.2 Hydride Embrittlement from Galvanic Corrosion of Waste Package Contacting 
Cladding............................................................................................................. 50 
6.12.3 Delayed Hydride Cracking (of Cladding).......................................................... 52 
6.12.4 Hydride Reorientation (of Cladding) ................................................................. 54 
6.12.5 Hydride Axial Migration (of Cladding)............................................................. 59 
6.12.6 Hydride Embrittlement from Fuel Reaction (Causes Failure of Cladding)....... 60 
6.13 CLADDING UNZIPPING............................................................................................. 60 
6.14 MECHANICAL IMPACT ON CLADDING................................................................ 61 
6.15 NAVAL SNF CLADDING ........................................................................................... 63 
6.16 DIFFUSION-CONTROLLED CAVITY GROWTH (DCCG) IN CLADDING .......... 63 
6.17 LOCALIZED (FLUORIDE ENHANCED) CORROSION OF CLADDING............... 65 
March 2004 vii ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
CONTENTS 
Page 
6.18 ROCKFALL .................................................................................................................. 68 
6.19 VOLUME INCREASE OF CORROSION PRODUCTS IMPACTS CLADDING ..... 69 
6.20 ELECTROCHEMICAL EFFECTS IN EBS ................................................................. 69 
6.21 CHEMICAL EFFECTS OF WASTE-ROCK CONTACT............................................ 72 
6.22 THERMAL EXPANSION/STRESS OF IN-PACKAGE EBS COMPONENTS ......... 73 
6.23 GAS GENERATION (HE) FROM WASTE FORM DECAY...................................... 75 
6.24 GAS GENERATION (H2) FROM WASTE PACKAGE CORROSION...................... 76 
7. CONCLUSIONS....................................................................................................................79 
8. REFERENCES ........................................................................................................................81 
8.1 DOCUMENTS CITED.................................................................................................. 81 
8.2 CODES, STANDARDS, REGULATIONS, AND PROCEDURES............................. 90 
8.3 SOURCE DATA, LISTED BY DATA TRACKING NUMBER ................................. 90 
March 2004 viii ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
FIGURES 
Page 
6.3-1. The Iso-Corrosion Diagram for Zirconium in Sulfuric Acid (H2SO4) ...........................22 
6.3-2. A Comparison of Oxide Thicknesses for Different Zirconium Alloys...........................25 
6.5-1. pH Profiles Showing How the Radiolysis Inputs Affect the pH ....................................29 
6.6-1. Comparison of In-Package Chemistry with Chemical Conditions Necessary for 
Pitting........................................................................................................................33 
6.6-2. Chemistry Where Pitting of Zirconium Has Been Observed..........................................36 
6.9-1. Creep Failure Fraction as a Function of Peak Cladding Temperature............................41 
6.9-2. Comparison of Creep Damage in Dry Storage and Repository Emplacement...............41 
6.12-1. Fracture Toughness vs. Hydrogen Content of Zircaloy-4 ..............................................49 
6.12-2. Comparison of the YMP Reorientation Equation with Additional Experimental Data..56 
6.12-3. Dissolution and Precipitation Solubilities for Hydrogen in Irradiated Zirconium 
Metal ...............................................................................................................................59 
TABLES 
Page 
1-1. FEPs Related to the CSNF Clad Degradation Component...............................................3 
1-2. Changes to the Cladding FEPs from TSPA-SR to TSPA-LA ..........................................4 
4-1. Information Sources for FEPs Included in TSPA-LA ......................................................7 
4-2. Information Sources for FEPs Excluded from TSPA-LA ................................................8 
6-1. LA FEP List for Cladding...............................................................................................13 
6-2. Corroborating Models and Information Used to Develop This Report ..........................16 
6.3-1. Percent of Cladding Remaining in Waste Packages that Fail at Closure .......................24 
6.5-1. Chloride, Ferric Iron, and Hydrogen Peroxide Molality.................................................30 
6.5-2. Half-Life of Hydrogen Peroxide at Various Temperatures ............................................31 
6.6-1. Corrosion and Repassivation Potentials for In-Package Chemistry Condition ..............33 
6.10-1. Effect of Helium Production on Rod Pressure................................................................43 
6.11-1. Pressures, Stresses and Stress Intensity Factors for Fuel Rods During Repository 
ssd) for Hydrogen in Unirradiated Zirconium as a Function of 
Emplacement...................................................................................................................46 
6.12-1. Puls' Zircaloy-2 Strain Tests on Zirconium with Reoriented Hydrides..........................57 
6.12-2. Saturation Limits (T 
Temperature ....................................................................................................................58 
6.17-1. Compositions of J-13 Well Water and Its Concentrates.................................................67 
7-1. Summary of Cladding Degradation and Waste Form FEPs ...........................................80 
March 2004 ix ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
ACRONYMS AND ABBREVIATIONS 
Argonne National Laboratory 
American Society for Testing and Materials 
ANL 
ASTM 
boiling water reactor BWR 
Complementary Cumulative Distribution Function 
commercial spent nuclear fuel 
CCDF 
CSNF 
Diffusion Controlled Cavity Growth 
Delayed Hydride Cracking 
U.S. Department of Energy 
Defense High Level Waste 
Defense Spent Nuclear Fuel 
DCCG 
DHC 
DOE 
DHLW 
DSNF 
EBS Engineered Barrier System 
Features, Events, and Processes FEPs 
HLW High Level Waste 
International Atomic Energy Agency 
Idaho National Engineering and Environmental Laboratory 
IAEA 
INEEL 
kgU kilogram Uranium 
KTI Key Technical Issue 
LA License Application 
LWR Light Water Reactor 
MIC Microbially Influenced Corrosion 
MWd Megawatt days 
U.S. Nuclear Regulatory Commission NRC 
Office of Civilian Radioactive Waste Management OCRWM 
PWR pressurized water reactor 
reasonably maximally exposed individual 
RMEI 
SCC Stress Corrosion Cracking 
SCE Standard Calemel Electrode scale 
SNF Spent Nuclear Fuel 
SR Site Recommendation 
SRB 
STP 
Sulfate Reducing Bacteria 
Standard Temperature and Pressure 
Total System Performance Assessment 
Total System Performance Assessment and Integration 
Total System Performance Assessment  License Application 
Total System Performance Assessment  Viability Assessment 
TSPA 
TSPAI 
TSPA-LA 
TSPA-VA 
Yucca Mountain Project YMP 
March 2004 x ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
1. PURPOSE 
The purpose of this report is to document the screening of the cladding degradation features, 
events, and processes (FEPs) for commercial spent nuclear fuel (CSNF). This report also 
addresses the effect of some FEPs on both the cladding and the CSNF, DSNF, and HLW waste 
forms where it was considered appropriate to address the effects on both materials together (FEP 
2.1.09.09.0A, 2.1.09.11.0A, 2.1.11.05.0A, 2.1.12.02.0A, and 2.1.12.03.0A). This report 
summarizes the work of others to screen clad degradation FEPs in a manner consistent with, and 
used in, the Total System Performance Assessment  License Application (TSPA-LA). This 
document was prepared according to Technical Work Plan for Waste Form Degradation 
Modeling, Testing, and Analyses in Support of LA (BSC 2004a [DIRS 167796]). 
The development of a comprehensive list of FEPs potentially relevant to postclosure 
performance of a repository at Yucca Mountain is an ongoing, iterative process based on sitespecific 
information, design, and regulations. The approach for developing an initial list of FEPs 
in support of Total System Performance Assessment for the Site Recommendation (CRWMS 
M&O 2000a [DIRS 153246]) was documented in The Development of Information Catalogued 
in REV00 of the YMP FEP Database (BSC 2001c [DIRS 154365]). The initial FEP list 
contained 328 FEPs, of which 176 were included in the TSPA-SR model (CRWMS M&O 2000a 
[DIRS 153246], Tables B-9 through B-17). To support the TSPA-LA model, the FEP list was 
re-evaluated in accordance with The Enhanced Plan for Features, Events, and Processes (FEPs) 
at Yucca Mountain (BSC 2002 [DIRS 158966], Section 3.2). This was supplemented by KTI 
Letter Report Response to Additional Information Needs on TSPAI 2.05 and TSPAI 2.06 
(Freeze 2003 [DIRS 165394]). This re-evaluation resulted in an initial list of FEPs for TSPALA, 
documented in DTN: MO0307SEPFEPS4.000 [DIRS 164527]. 
Since the initial list was developed, some modifications have occurred. As noted above, FEPs 
2.1.09.09.0A, 2.1.09.11.0A, 2.1.11.05.0A, 2.1.12.02.0A, and 2.1.12.03.0A were re-assigned to 
this report. FEP 2.1.02.07.0A was re-assigned from this report to a different FEP analysis report. 
The description for FEP 2.1.02.11.0A was modified for clarity. The description for FEP 
2.1.02.15.0A was modified since H2O2 was incorrectly identified as ions in solution. The 
description of FEP 2.1.02.19.0A was modified to reflect the current temperature limit for creep 
of 400oC. The description of FEP 2.1.02.24.0A was modified to add impact from waste package 
internals and to remove rockfall damage from seismic events, which is now addressed in FEP 
1.2.03.02.0B. 
Table 1-1 provides the modified list of cladding FEPs, including their screening decisions 
(include or exclude). The primary purpose of this report is to identify and document the analysis, 
screening decision, and TSPA-LA disposition (for included FEPs) or screening argument (for 
excluded FEPs) for these FEPs related to clad degradation. In some cases, where a FEP covers 
multiple technical areas and is shared with other FEP reports, this cladding FEP report may 
provide only a partial technical basis for the screening of the FEP. The full technical bases for 
these shared FEPs are addressed collectively by all of the sharing FEP reports. The screening 
decisions and associated TSPA-LA dispositions or screening arguments from all of the FEP 
reports will be cataloged in a project-specific FEPs database. This report, along with the other 
FEP reports, and the database are being used to document information related to the FEPs 
screening to assist reviewers during the license review process. 
March 2004 1 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 1-2 summarizes the changes to the cladding FEPs from TSPA-SR (CRWMS M&O 2000g 
[DIRS 153947]) to the modified TSPA-LA list, including the changes in assignment and 
descriptions noted previously. Two major analysis changes have occurred. In TSPA-SR a creep 
model was included in the TSPA analysis although the temperatures of the fuel were sufficiently 
low that no cladding failures were predicted from this cause. The creep model included internal 
pressurization of cladding and thermally induced stress changes. Since no failures are predicted 
to occur because of creep, the FEPs dealing with creep (FEPs 2.1.02.19.0A, 2.1.02.20.0A, 
2.1.11.05.0A) are now excluded from TSPA-LA. The SR model contained a non-mechanistic 
corrosion model that was used to include failure from various corrosion mechanisms. Since SR, 
a pitting model and an in-package chemistry model have been developed. When these models 
were combined, no cladding failure from corrosion was predicted and these FEPs (FEPs 
2.1.02.14.0A, 2.1.02.15.0A, 2.1.02.16.0A, 2.1.02.27.0A) were excluded. 
The four FEPs included in the TSPA-LA are described in detail in Clad Degradation  Summary 
and Abstraction for LA (BSC 2003a [DIRS 162153]) and summarized in this report. Clad 
Degradation  Summary and Abstraction for LA (BSC 2003a [DIRS 162153]) also describes one 
FEP (2.1.02.24.0A) Mechanical Impact on Cladding, as included. At that time it was included 
because the cladding is modeled to be damaged by rock overpressure after the drip shield and 
waste package deteriorates. This damage occurs well after the 10,000-year regulatory period and 
is now characterized as excluded. For excluded FEPs, the technical bases for their exclusion are 
documented here. 
This analysis is limited to fuel exposed to normal operation and anticipated operational 
occurrences (i.e., events that are anticipated to occur within a reactor lifetime), and is not 
applicable to fuel that has been exposed to severe accidents. Fuel burnup projections have been 
limited to the current commercial reactor licensing environment with restrictions on fuel 
enrichment (5 percent for material shipment and fuel manufacturing, shipment, and storage), 
cladding oxide coating thickness, and rod plenum pressures. This is consistent with the 
projections for advanced reloads. The fuel considered has burnup up to 75 MWd/kgU and half 
of the fuel is above 44.7 MWd/kgU, todays typical PWR burnup range. Ranges and 
uncertainties have been defined in Initial Cladding Condition (CRWMS M&O 2000c [DIRS 
151659], Section 6.2). Information provided in this FEPs screening will be used to evaluate 
which cladding degradation mechanisms are included in the postclosure performance of the 
repository in relation to waste form degradation. 
The CSNF cladding performance is used to bound naval fuel cladding behavior. The Naval 
Nuclear Propulsion Program has demonstrated that the performance of naval spent nuclear fuel 
cladding is better than or equal to that of cladding on CSNF(BSC 2001a [DIRS 152059], p. 36 
and Figure 6.1-2). Cladding degradation for defense spent nuclear fuel (DSNF), including 
aluminum-clad fuel, is not addressed in this report, but is addressed in a general waste form 
screening report Miscellaneous Waste-Form FEPs (CRWMS M&O 2001c [DIRS 153938]). 
March 2004 2 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 1-1. FEPs Related to the CSNF Clad Degradation Component 
FEP Name Section FEP Number Shared 
With 
Screening 
Decision 
6.1 2.1.02.11.0A Degradation of cladding from waterlogged rods Excluded 
6.2 2.1.02.12.0A Degradation of cladding prior to disposal Included 
6.3 2.1.02.13.0A General corrosion of cladding Excluded 
6.4 2.1.02.14.0A Microbially influenced corrosion (MIC) of cladding Excluded 
6.5 2.1.02.15.0A Localized (radiolysis enhanced) corrosion of cladding Excluded 
6.6 2.1.02.16.0A Localized (pitting) corrosion of cladding Excluded 
6.7 2.1.02.17.0A Localized (crevice) corrosion of cladding Excluded 
6.8 2.1.02.18.0A Enhanced corrosion of cladding from dissolved silica Excluded 
6.9 2.1.02.19.0A Creep rupture of cladding Excluded 
6.10 2.1.02.20.0A Internal pressurization of cladding Excluded 
6.11 2.1.02.21.0A Stress corrosion cracking (SCC) of cladding Excluded 
6.12 2.1.02.22.0A Hydride cracking of cladding Excluded 
6.13 2.1.02.23.0A Cladding unzipping Included 
6.14 2.1.02.24.0A Mechanical impact on cladding Excluded 
6.15 2.1.02.25.0B Naval SNF Cladding Included 
6.16 2.1.02.26.0A Diffusion-controlled cavity growth (DCCG) in cladding Excluded 
6.17 2.1.02.27.0A Localized (fluoride enhanced) corrosion of cladding Excluded 
6.18 2.1.07.01.0A Rockfall Excluded WP, EBS 
6.19 2.1.09.03.0A Volume increase of corrosion products impacts cladding Included 
6.20 2.1.09.09.0A Electrochemical effects in EBS Excluded WP 
6.21 2.1.09.11.0A Chemical effects of waste-rock contact Excluded 
6.22 2.1.11.05.0A Thermal expansion/stress of in-package EBS Excluded 
components 
EBS 6.23 Excluded Gas generation (He) from waste form decay 2.1.12.02.0A 
WP, EBS 6.24 Excluded 2.1.12.03.0A Gas generation (H2) from waste package corrosion 
NOTE: WP = waste package, EBS = engineered barrier system 
March 2004 3 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 1-2. Changes to the Cladding FEPs from TSPA-SR to TSPA-LA 
SR-NameA SR-#A LA FEP Name LA # 
of Cs, I 
Significant Scope and Screening 
Changes from TSPA-SRA to TSPA-LA 
2.1.02.07.00 Gap and Grain Release Not a cladding FEP for LA. FEP now 
addressed by WF in BSC 2004c [DIRS 
167321] 
2.1.02.11.00 Waterlogged Rods No change, both exclude 2.1.02.11.0A Degradation of 
No change, both include 2.1.02.12.00 Cladding Degradation 
Before YMP Receives It 
cladding from 
waterlogged rods 
2.1.02.12.0A Degradation of 
cladding prior to 
disposal 
No change, both exclude 
cladding 
2.1.02.13.0A General corrosion of 2.1.02.13.00 General Corrosion of 
Cladding 
LA: Excluded with new pitting model, SR: 
Included as part of non-mechanistic 
corrosion model 
2.1.02.14.0A Microbially influenced 2.1.02.14.00 Microbiologically 
corrosion (MIC) of 
cladding 
Influenced Corrosion 
(MIC) of Cladding 
2.1.02.15.0A Localized (radiolysis 2.1.02.15.00 Acid Corrosion of 
Cladding From 
Radiolysis 
LA: Excluded with new pitting model and 
in-package chemistry, SR: Included as part 
of non-mechanistic corrosion model 
enhanced) corrosion 
of cladding 
2.1.02.16.0A Localized (pitting) 
corrosion of cladding 
2.1.02.16.00 Localized Corrosion 
(Pitting) of Cladding 
LA: Excluded with new pitting model, SR: 
Included as part of non-mechanistic 
corrosion model 
No change, both exclude 2.1.02.17.00 Localized Corrosion 2.1.02.17.0A Localized (crevice) 
corrosion of cladding (Crevice Corrosion) of 
Cladding 
No change, both exclude 2.1.02.18.0A Enhanced corrosion 2.1.02.18.00 High Dissolved Silica 
of cladding from 
dissolved silica 
Content of Waters 
Enhances Corrosion of 
Cladding 
2.1.02.19.00 Creep Rupture of 
Cladding 
2.1.02.19.0A Creep rupture of 
cladding 
LA: Exclude, SR: Included in creep model 
but no cladding failures when model was 
used 
2.1.02.20.0A Internal 2.1.02.20.00 Pressurization From He LA: Exclude, SR: Included in creep model 
but no cladding failures when model was 
used 
pressurization of 
cladding 
2.1.02.21.0A Stress corrosion LA: Exclude, SR: Included assuming high 
iodine and stresses present 
Production Causes 
Cladding Failure 
2.1.02.21.00 Stress Corrosion 
Cracking (SCC) of 
Cladding 
cracking (SCC) of 
cladding 
2.1.02.22.0A Hydride cracking of 2.1.02.22.00 Hydride Embrittlement No change, both exclude 
cladding of Cladding 
No change, both include 2.1.02.23.0A Cladding unzipping 2.1.02.23.00 Cladding Unzipping 
2.1.02.24.0A Mechanical impact 2.1.02.24.00 Mechanical Failure of 
Cladding on cladding 
LA Scope expanded to address impact 
from waste package internals, reduced by 
having seismic rockfall now addressed as 
part of shared DE and EBS FEP 
1.2.03.02.0B. 
LA: Exclude, SR: Included from rock 
overpressure after regulatory period 
New FEP since SR, LA: Include 
No change, both exclude 
2.1.02.25.0B Naval SNF Cladding 
2.1.02.26.0A Diffusion-controlled 2.1.02.26.00 Diffusion-Controlled 
cavity growth Cavity Growth 
(DCCG) in cladding 
March 2004 4 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 1-2. Changes to the Cladding FEPs from TSPA-SR to TSPA-LA (Continued) 
SR-NameA SR-#A LA FEP Name LA # Significant Scope and Screening 
Changes from TSPA-SRA to TSPA-LA 
LA: Excluded, low Fluoride concentrations, 2.1.02.27.0A Localized (fluoride 
enhanced) corrosion 
2.1.02.27.00 Localized Corrosion 
Perforation from pH>3.18 , SR: Included as part of nonmechanistic 
corrosion model Fluoride of cladding 
2.1.07.01.0A Rockfall 2.1.07.01.00 Rockfall (Large Block) LA Scope reduced by having seismic 
rockfall now addressed as part of shared 
DE and EBS FEP 1.2.03.02.0B , LA and 
SR both exclude 
2.1.09.03.0A Volume increase of 2.1.09.03.00 Volume Increase of 
Corrosion Products corrosion products 
impacts cladding 
LA Scope reduced to address impacts on 
cladding only. Impacts on other EBS 
components are now addressed in WP 
FEP 2.1.09.03.0B and EBS FEP 
2.1.09.03.0C, LA and SR both include 
2.1.09.09.00 * Electrochemical Effects No change, both exclude 2.1.09.09.0A Electrochemical 
effects in EBS (Electrophoresis, 
Galvanic Coupling) in 
Waste and EBS 
2.1.09.11.0A Chemical effects of 2.1.09.11.00 *Waste-Rock Contact No change, both exclude 
waste-rock contact 
2.1.11.05.0A Thermal LA Scope addresses all in-package issues 2.1.11.05.00 * Differing Thermal 
expansion/stress of 
in-package EBS 
components 
Exclude, SR included as part of creep 
analysis 
2.1.11.07.00 
Expansion of Repository that were in the two SR FEPs. LA: 
Components 
Thermally-Induced 
Stress Changes in 
Waste and EBS 
No change, both exclude 2.1.12.02.0A Gas generation (He) 2.1.12.02.00 * Gas generation (He) 
from fuel decay from waste form 
decay 
No change, both exclude 2.1.12.03.0A Gas generation (H2) 2.1.12.03.00 * Gas generation (H2) 
from metal corrosion from waste package 
corrosion 
* These were not addressed as Cladding FEPs for SR, but were addressed elsewhere (WF and/or WP and/or EBS). 
A Source: CRWMS M&O 2000g [DIRS 153947] 
1.1 SCOPE 
This report addresses clad degradation FEPs, which represent the key phenomena that result in 
degradation of cladding and, in some cases, the CSNF pellets themselves. Clad degradation 
FEPs addressed in this report are provided in Table 1-1. 
The repository design continues to evolve in preparation for the license application. In general, 
these design changes can directly influence the screening arguments for FEPs and their inclusion 
in or exclusion from the waste form degradation model. The elimination of backfill is an 
important design change that was made after the development of the current waste form 
degradation model. The primary effect of the elimination of backfill is the decrease of peak 
temperatures inside the waste package, which is beneficial to reducing uncertainty (CRWMS 
M&O 2001a [DIRS 151662], Section 6.2). However, the absence of backfill does not affect the 
cladding degradation FEPs included in the TSPA-LA analysis. The original screening (earlier 
version of this report) was performed with backfill temperatures and contained many of the same 
conclusions. 
March 2004 5 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
1.2 FEPS IDENTIFICATION AND ANALYSIS 
To support the TSPA-LA model, the FEP list was reevaluated in accordance with The Enhanced 
Plan for Features, Events, and Processes (FEPs) at Yucca Mountain (BSC 2002 [DIRS 158966], 
Section 3.2) and supplemented by KTI Letter Report Response to Additional Information Needs 
on TSPAI 2.05 and TSPAI 2.06 (Freeze 2003 [DIRS 165394]). The resulting 24 cladding FEPs 
are listed in Table 1-1. 
The U.S. Nuclear Regulatory Commission (NRC) requires the consideration and evaluation of 
FEPs as part of the performance assessment activities. More specifically, the NRC regulations 
allow the exclusion of FEPs from the TSPA if they can be shown to be of low probability or 
consequence. The specified criteria can be summarized as follows. 
1. The event has at least one chance in 10,000 of occurring over 10,000 years (see 
10 CFR 63.114(d) [DIRS 156605]). 
2. The magnitude and time of the resulting radiological exposure to the RMEI, or 
radionuclide release to the accessible environment, would be significantly 
changed by its omission (see 10 CFR 63.114 (e and f) [DIRS 156605]). 
Additionally, the Acceptance Criterion in Yucca Mountain Review Plan, Final Report (NRC 
2003 [DIRS 163274], Section 2.2.1.2.1.3) allows the exclusion of FEPs from TSPA if they are 
inconsistent with specifications within the NRC regulations. This criterion can be summarized in 
the form of a third FEP screening statement. 
3. The FEP is not excluded by regulation. 
Evaluating the FEPs against these screening statements may be done in any order. If there are 
affirmative conditions for all three screening criteria, a FEP is Included in the TSPA-LA 
model. If there is a negating condition in any of the three screening criteria, the FEP is 
Excluded from the TSPA-LA model. 
For the cladding FEPs, all three screening criteria were considered, but only two (Criteria 1 and 2 
above) were used. Consequence-based screening arguments can be established in a variety of 
ways, including TSPA-LA sensitivity analyses, modeling studies outside of the TSPA-LA, or 
reasoned arguments based on literature research. Probability-based screening arguments are 
based on a comparison of the FEP probability of occurrence with the regulatory probability 
criterion. 
2. QUALITY ASSURANCE 
The Quality Assurance program applies to the development of this document because this model 
will be part of the TSPA-LA safety analysis. This document was prepared in accordance with 
Technical Work Plan for Waste Form Degradation, Modeling, Testing, and Analyses in Support 
of LA (BSC 2004a [DIRS 167796]). This report does not directly impact structures, systems, or 
components classified in accordance with AP-2.22Q, Classification Analyses and Maintenance 
of the Q-List. The technical work plan contains the Process Control Evaluation used to evaluate 
the control of electronic management of data (BSC 2004a [DIRS 167796], Attachment I) during 
March 2004 6 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
modeling and documentation activities. This evaluation determined that the methods identified 
in the implementing procedures are adequate. There were no deviations from these methods. 
This report was prepared using AP-SIII.9Q, Scientific Analyses. 
3. USE OF SOFTWARE 
No software is used for analysis in this report. The analyses and arguments presented herein are 
based on regulatory requirements, results of analyses presented and documented in other reports, 
or technical literature. This report was documented using only commercially available software 
(Microsoft Word 97, SR2) for word processing, which is not required to be qualified or 
documented per AP-SI.1Q, Software Management. There were no additional applications 
(routines or macros) developed for documentation using this commercial software. SigmaPlot, 
Scientific Graphic Software, Version 2.0, Jandel Corporation is used to plot data. No 
calculations are performed with this software. This software is not required to be qualified or 
documented per AP-SI.1Q, Software Management, as specified in Section 2.1.2. 
4. INPUTS 
4.1 DIRECT INPUTS 
This section identifies input data, parameters, and other forms of analysis inputs used in this 
report. The data and technical information used in this report were obtained from controlled 
source documents and other appropriate sources. 
4.1.1 Data 
No data were used as direct input. 
4.1.2 Parameters and Parameter Uncertainty 
The analyses and arguments presented herein are based on guidance and regulatory 
requirements, results of analyses presented and documented in other reports or other technical 
literature. The sources for the detailed discussions of included FEPs are provided in Table 4-1. 
Sources for excluded FEPs are listed in Table 4-2. The justification for using the technical 
information listed in Table 4-2 is presented in Appendix C. Models developed in the supporting 
documents are cited for traceability and transparency purposes; however, they were not used 
directly in development of these analyses and arguments presented herein. 
Source 
BSC 2003a [DIRS 162153], Section 6.2.4 2.1.09.03.0A 
products impacts cladding 
Table 4-1. Information Sources for FEPs Included in TSPA-LA 
FEP Name 
Section 
Used 
BSC 2003a [DIRS 162153], Sections 6.2.1, 6.2.2 
BSC 2003a [DIRS 162153], Section 6.2.4 
BSC 2003a [DIRS 162153], Section 6 
7 
FEP Number 
6.2 2.1.02.12.0A Degradation of cladding prior to 
disposal 
6.13 2.1.02.23.0A Cladding unzipping 
6.15 2.1.02.25.0B Naval SNF Cladding 
6.19 Volume increase of corrosion 
ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation  FEPs Screening Arguments 
Table 4-2. Information Sources for FEPs Excluded from TSPA-LA 
FEP Number FEP Name 
All NA 
Section 
Used 
Through 
Source 
10 CFR 63.114(d, e, f) [DIRS 
156605]. out 
6.1 2.1.02.11.0A Degradation of 
cladding from 
waterlogged rods 
NRC 1997 [DIRS 101903], 
Section 8.V.1 
BSC 2003a [DIRS 162153], 
Section 6.2.1 
6.3 2.1.02.13.0A General corrosion 
of cladding 
Hillner et al. 1998 [DIRS 
100455], 
Pourbaix 1974 [DIRS 100817], 
p. 226 
6.4 2.1.02.14.0A Microbially Wolfram et al. 1996 [DIRS 
influenced corrosion 165268], pp. iii, iv 
(MIC) of cladding Hillner et al. 1998 [DIRS 
100455], p. 11 
6.5 2.1.02.15.0A Localized 
(radiolysis 
enhanced) 
corrosion of 
BSC 2003f [DIRS 161962], 
Attachment III 
BSC 2003a [DIRS 162153] 
BSC 2003g {DIRS 164667] 
cladding 
6.6 2.1.02.16.0A Localized (pitting) BSC 2003f [DIRS 161962], 
Attachment III 
BSC 2003g [DIRS 164667], 
corrosion of 
cladding 
Section 8.2 
CRWMS M&O 2000b [DIRS 
151561] 
6.7 2.1.02.17.0A Localized (crevice) 
corrosion of 
cladding 
Yau and Webster 1987 [DIRS 
100494], p. 717 
Greene et al. 2000 [DIRS 
145073], p. 7) 
Brossia et al. (2002 [DIRS 
161988]) 
6.8 2.1.02.18.0A Enhanced corrosion 
of cladding from 
dissolved silica 
Hansson 1984 [DIRS 101676] 
Yau and Webster 1987 [DIRS 
100494], Table 6 
6.9 2.1.02.19.0A Creep rupture of NRC 2002 [DIRS 164593], p. 2 Little creep damage below 400C 
cladding 
6.10 2.1.02.20.0A Internal Piron and Pelletier 2001 [DIRS 
165318], Section 5.3 pressurization of 
cladding 
6.11 2.1.02.21.0A Stress corrosion 
cracking (SCC) of 
cladding 
Beckman 2001 [DIRS 156122], Iodine-induced SCC unlikely 
p. 103 
BSC 2003a [DIRS 162153] Passive layer preserved 
BSC 2003g {DIRS 164667] 
Kreyns et al. 1996 [DIRS 
100462], Figure 5 
6.12.0 2.1.02.22.0A Hydride cracking of 
cladding (General) 
6.12.1 2.1.02.22.0A Hydride cracking of Lanning et al. 1997 [DIRS 
101704], Volume 1, p. 8.4, 
Figure 8.2 
cladding (zirconium 
alloys corrosion) 
6.12.2 2.1.02.22.0A Hydride cracking of 
cladding (galvanic 
corrosion) 
Yau and Webster 1987 [DIRS 
100494], p. 718, Table 15 
Hansson 1984 [DIRS 101676], 
p. 6 
8 ANL-WIS-MD-000008 REV 01 
Information 
FEP screening considerations 
Waterlogging not a problem 
Fraction of failed rods 
Corrosion rate 
Stability of ZrO2 
MIC and Zirconium alloys 
MIC and Zirconium alloys 
In-package chemistry 
Pitting potential not met 
In-package chemistry 
Pitting calculation for waste 
package 
MIC little effect on geochemistry 
Resistance to Crevice corrosion 
Resistance to Crevice corrosion 
Resistance to Crevice corrosion 
No accelerated corrosion 
No accelerated corrosion 
Helium pressurization 
Fracture toughness vs. hydrogen 
content 
Hydrogen absorption from 
general corrosion 
Galvanic corrosion 
Rapid passivation 
March 2004 Clad Degradation  FEPs Screening Arguments 
Table 4-2. Information Sources for FEPs Excluded from TSPA-LA (Continued) 
Section 
Used FEP Number FEP Name 
6.12.3 2.1.02.22.0A Hydride cracking of 
cladding (delayed 
hydride cracking) 
6.12.4 2.1.02.22.0A Hydride cracking of 
cladding (hydride 
reorientation) 
6.12.5 2.1.02.22.0A Hydride cracking of 
cladding (axial 
migration) 
6.12.6 2.1.02.22.0A Hydride cracking of 
cladding (fuel 
reaction) 
6.14 2.1.02.24.0A Mechanical impact 
on cladding 
6.16 2.1.02.26.0A Diffusion-controlled 
cavity growth 
(DCCG) in cladding 
6.17 2.1.02.27.0A Localized (fluoride 
enhanced) 
corrosion of 
cladding 
6.18 2.1.07.01.0A Rockfall 
6.20 2.1.09.09.0A Electrochemical 
effects in EBS 
6.21 2.1.09.11.0A Chemical effects of 
waste-rock contact 
6.22 2.1.11.05.0A Thermal 
expansion/stress of 
in-package EBS 
components 
6.23 2.1.12.02.0A Gas generation 
(He) from waste 
form decay 
6.24 2.1.12.03.0A Gas generation (H2) 
from waste package Tables 1 through 4 
corrosion 
NA NA App. A 
ANL-WIS-MD-000008 REV 01 
Source 
Peehs 1998 [DIRS 109219], 
pp. 5, 6 
Cragnolino et al. 1999 [DIRS 
152354], p.4-21 
EPRI 2002 [DIRS 161421] 
Cragnolino et al. 1999 [DIRS 
152354], p. 4-22 
Cunningham et al. 1987 [DIRS 
101591], Appendix C 
Cappelaere et al. 2002 [DIRS 
164195], p. 5 
Edsinger 2000 [DIRS 154433] Hydrating of cladding requires 
reducing environment in BWRs 
CRWMS M&O 2001b [DIRS 
153937]. 
BSC 2003c [DIRS 163935], 
Section 6.5.1.2.1.3.2 
NRC 2000 [DIRS 147797] 
Hayes et al. 1999 [DIRS 
164598], Figures 2, 5, 6, 8, 
and 11 
Pourbaix 1974 [DIRS 100817], 
p. 583 
BSC 2003f [DIRS 161962], 
Attachment III) 
CRWMS M&O 2001b [DIRS 
153937]. 
Yau and Webster 1987 [DIRS 
100494], p. 718, Table 15 
Hansson 1984 [DIRS 101676], 
p. 6 
Adler, Flitton, et al. 2002 [DIRS Cladding/rock contact test 
161991], p. 4 
BSC 2003a [DIRS 162153] 
BSC 2004c [DIRS 167321], 
Section 6.2.1 
BSC 2004d [DIRS 167619], 
Section 7.5.3 
Piron and Pelletier 2001 [DIRS 
165318], Section 5.3 
Clayton 1989 [DIRS 149208], 
IAEA 1998 [DIRS 150560], p. 
92 
Pourbaix 1974 [DIRS 100817], 
p. 226 
BSC 2003b [DIRS 166463], 
Section 6.2 
9 
Information 
DHC not expected 
DHC not expected 
No reorientation in dry storage 
No reorientation of Tmax < 290C 
Axial migration is minor 
Axial migration is minor 
Rockfalls will not damage the 
waste package 
Volumes and dimensions for 
oxides 
Little creep damage under 400C 
DCCG is unlikely if T < 330 - 
400C 
Hydrofluoric acid exists for pH < 
3.18 
pH > 3.5 
Rockfalls will not damage the 
waste package 
Zirconium is a noble metal 
Passive film forms in seconds 
Fuel pellet fragmentation 
Glass temperatures 
Helium pressurization 
H2 not absorbed through passive 
layer 
H2 not absorbed through passive 
layer 
Stability of Zr in water 
Temperature history of waste 
package surface 
March 2004 Clad Degradation  FEPs Screening Arguments 
4.1.3 Other Model / Analyses Inputs and Technical Information 
The list of cladding degradation FEPs was extracted from the LA FEP list documented in 
DTN: MO0307SEPFEPS4.000 [DIRS 164527] and is considered as input for this analysis. 
The low consequence requirements are stated in 10 CFR 63.114 (e and f) [DIRS 156605]: 
(e) Provide the technical basis for either inclusion or exclusion of specific 
features, events, and processes in the performance assessment. Specific features, 
events, and processes must be evaluated in detail if the magnitude and time of the 
resulting radiological exposures to the reasonably maximally exposed individual, 
or radionuclide releases to the accessible environment, would be significantly 
changed by their omission. 
(f) Provide the technical basis for either inclusion or exclusion of degradation, 
deterioration, or alteration processes of engineered barriers in the performance 
assessment, including those processes that would adversely affect the 
performance of natural barriers. Degradation, deterioration, or alteration 
processes of engineered barriers must be evaluated in detail if the magnitude and 
time of the resulting radiological exposures to the reasonably maximally exposed 
individual, or radionuclide releases to the accessible environment, would be 
significantly changed by their omission. 
and supported by 10 CFR 63.342 [DIRS 156605]: 
DOEs performance assessments need not evaluate the impacts resulting from any 
features, events, and processes or sequences of events and processes with a higher 
chance of occurrence if the results of the performance assessments would not be 
changed significantly. 
Although NRC regulations do not define the terms significantly changed and changed 
significantly, this report infers the absence of significant change to be equivalent to having no 
or negligible effect (i.e., inclusion of effect would have negligible effect on consequence). 
The low-probability criterion is stated in 10 CFR 63.114(d)[DIRS 156605]: 
Consider only events that have at least one chance in 10,000 of occurring over 
10,000 years. 
and supported by 10 CFR 63.342 [DIRS 156605]: 
DOEs performance assessments shall not include consideration of very unlikely 
features, events, or processes, i.e., those that are estimated to have less than one 
chance in 10,000 of occurring within 10,000 years of disposal. 
The low-probability criterion (i.e., very unlikely FEPs) is stated as less than one chance in 
10,000 of occurring in 10,000 years (10-4/104 yr), a 10-8 annual-exceedance probability. 
March 2004 10 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
4.2 CRITERIA 
The low consequence and low probability requirements stated in 10 CFR 63.114 (d, e and f) 
[DIRS 156605] and supported by 10 CFR 63.342 [DIRS 156605] can be considered criteria, as 
discussed in Section 4.1.3. 
4.3 CODES AND STANDARDS 
American Society for Testing and Materials (ASTM) Standard C1174-97, Standard Practice for 
Prediction of the Long-Term Behavior of Materials, Including Waste Forms, Used in Engineered 
Barrier Systems (EBS) for Geological Disposal of High-Level Radioactive Waste (ASTM 1998 
[DIRS 105725]), is used to support the degradation model development methodology, categorize 
the model developed with respect to its usage for long-term TSPA, and relate the 
information/data used to develop the model to the requirements of the standard. 
For the repository, cladding is not an engineered barrier. It is not designed and controlled by the 
project to reduce release of radionuclides. The behavior can be modeled but the design 
characteristics are not within the project controls. 
5. ASSUMPTIONS 
There are no assumptions used in this report. 
6. SCIENTIFIC ANALYSIS DISCUSSION 
This section discusses each of the 24 FEPs dealing with cladding degradation (Table 6-1). The 
purpose of this report is to document the screening of the cladding degradation FEPs. For 
included FEPs, a summary of their implementation in TSPA-LA is presented here. As shown in 
Table 4-1, detailed documentation of their inclusion is presented in BSC 2003a [DIRS 162153]. 
BSC 2003c [DIRS 163935] provides additional support for FEP 2.1.09.03.0A. For excluded 
FEPs, the arguments for their exclusion are documented here. The excluded FEPs are listed in 
Table 4-2 with the sources for the technical information for their exclusion. Table 6-2 contains 
additional sources for corroborating models and information used to develop this report. 
For cladding FEPs that are excluded from the TSPA, the screening arguments are based on low 
probability of occurrence or low consequence, in accordance with the criteria identified in 
Section 4.1.3. As appropriate, screening arguments cite work done outside this activity, such as 
in other scientific analyses. 
In a few cases, a low-consequence argument, but based on probability, has been invoked. This is 
acceptable if a FEP can be defined in terms of its potential to affect the behavior (or response) of 
the disposal system (consequence), rather than solely in terms of the probability of the 
independent phenomenon. To evaluate this aspect one defines a threshold value at which a FEP 
has the potential to affect repository performance, and then defines the probability of the 
threshold being violated. This use of probability to support a low-consequence argument is 
particularly germane to FEPs involving potential breaching of containers. An example of this 
approach is FEP 2.1.02.16.0A, Localized (pitting) corrosion of cladding. The FEP is defined in 
terms of the consequence to the cladding from localized corrosion. The probability of localized 
March 2004 11 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
corrosion was determined to be low but could not be quantified for the purpose of comparison 
with the low probability criterion. However, it could be determined that the associated 
consequence would not be significant. In other words, under the range of expected repository 
conditions, the threshold value for localized corrosion (i.e., the value above which significant 
damage and resulting effects on behavior would occur) is not achieved. Therefore, localized 
(pitting) corrosion of cladding can be excluded based on low consequence, because it is not 
expected to occur at a magnitude sufficient to cause a significant effect on behavior. 
For FEPs that are included in the TSPA, the TSPA Disposition section of each FEP discussion 
includes a short summary of how the FEP has been incorporated in the TSPA models. Details of 
the scientific analysis that describes the disposition are presented in Clad Degradation  
Summary and Abstraction for LA (BSC 2003a [DIRS 162153]). 
The output of this analysis (TSPA-LA dispositions for included FEPs and screening arguments 
for excluded FEPs) is intended to be used to support the project-specific FEP database and to 
promote traceability and transparency regarding FEP screening. 
March 2004 12 ANL-WIS-MD-000008 REV 01 ANL-WIS-MD-000008 REV 01 
Table 6-1. LA FEP List for Cladding 
Screening 
Decision FEP Name Section FEP Number Description 
6.1 2.1.02.11.0A Degradati of on Excluded Failed fuel rods (attributed to breaches caused by manufacturing defects and reactor 
operations) comprise a small fraction of the fuel rods that are currently being stored in 
commercial reactor spent fuel pools. Failed fuel contains water in the fuel rod void space 
that may promote degradation of the spent fuel cladding. Such fuel is referred to as 
waterlogged. The moisture remaining in a dried fuel rod is used to determine the extent 
of degradation of spent fuel cladding. 
cladding from 
waterlogged rods 
6.2 2.1.02.12.0A Degradati of on Included Certain aspects of cladding degradation occur before the spent fuel arrives at Yucca 
Mountain. Possible mechanisms include rod cladding degradation during reactor operation, 
degradation during wet spent fuel pool storage, degradation during dry storage, and rod 
degradation during shipping (from creep and from vibration and impact) and fuel handling. 
cladding prior to 
disposal 
Exclude 6.3 2.1.02.13 G .0A eneral corrosion of General corrosion of cladding could expose large areas of fuel and produce hydrides. 
cladding 
6.4 2.1.02.14.0A Microbially Excl infl uenced uded Microbially Influenced Corrosion (MIC) of cladding potentially may be a local cladding 
corrosion mechanism where microbes produce a local acidic environment that could 
produce multiple penetrations through the fuel cladding. 
corrosion (MIC) of 
cladding 
Excluded 6.5 2.1.02.15.0A Localized (radiolysis Radiolysis in a nitrogen/oxygen gas mixture with the presence of water film results in the 
formation of nitric acid (HNO3). Hydrogen peroxide (H2O2) is formed in the water from 
radiolysis. These chemicals can enhance corrosion of the fuel cladding. 
enhanced) corrosion 
of cladding 
13 of 90 
6.6 2.1.02.16.0A Localized (pitting) Exclude Localized corrosion in pits could produce penetrations of cladding. 
corrosion of cladding 
6.7 2.1.02.17.0A Localized (crevi ce) Excluded Localized corrosion in crevices could produce penetrations of cladding. 
corrosion of cladding 
6.8 2.1.02.18.0A Enhanced corr osion Excluded It must be determined if the high dissolved silica content of waters enhances corrosion of 
cladding. of cladding from 
dissolved silica 
Excluded 6.9 2.1.02.19 Cr .0A eep rupture of At high temperatures (>400C) for sufficiently long time intervals, creep rupture of Zircaloy 
cladding on spent fuel can occur and produce small perforations in the cladding to relieve 
stress. After the waste package fails, the fuel can react with water and radioisotopes can 
thereby escape over time from the fuel rod. 
cladding 
Excluded 6.10 2.1.02.20.0A Internal Increased pressure within the fuel rod due to the production of helium gas could contribute 
to cladding failure. pressurization of 
cladding 
6.11 2.1.02.21.0A Stress corrosion Excluded Stress corrosion cracking mechanisms can contribute to cladding failure. These 
mechanisms can operate both from the inside out from the action of fission products, or 
from the outside in from the actions of salts or other chemicals within the waste package. 
cracking (SCC) of 
cladding 
March 2004 
ANL-WIS-MD-000008 REV 01 
Table 6-1. LA FEP List for Cladding (Continued) 
Screening 
Decision FEP Name Section FEP Number Description 
Excluded Hydride cracking of 2.1.02.22.0A 6.12 Cladding contains hydrogen after reactor operation. The cladding might also pick up more 
hydrogen from cladding general corrosion (wet oxidation) after the waste package is 
breached. The hydrogen can exist both as zirconium hydride precipitates and as hydrogen 
in solid solution with zirconium. Hydrides might also form from UO2 oxidation (after waste 
package and cladding perforation). In addition, hydrides may dissolve in warmer areas of 
the cladding and migrate to cooler areas. Hydrogen can also move from places of low 
stress to places of high stresses, causing hydride reorientation or delayed hydride cracking 
(DHC). The buildup of hydrides can cause existing cracks to propagate by DHC or hydride 
embrittlement. 
cladding 
6.13 2.1.02.23.0A Claddi unzip ng ping Included In either dry or wet oxidizing conditions and with perforated fuel cladding, the UO2 fuel can 
oxidize. The volume increase of the fuel as it oxidizes can create stresses in the cladding 
that may cause gross rupture of the fuel cladding (unzipping). 
Exclude Mechanical impact on 2.1.02.24.0A 6.14 Mechanical failure of cladding may result from external stresses such as rockfall or impact 
from waste package internals. Seismic induced impacts are addressed in a separate FEP. cladding 
Included Naval SNF cladding 2.1.02.25.0B 6.15 DSNF to be disposed of in Yucca Mountain has a variety of fuel types that may not be 
similar to the CSNF to be disposed. Some of the fuel types may have initial claddingdegradation 
characteristics that are different from those for the CSNF. Therefore, the 
effectiveness of DSNF cladding as a barrier to radionuclide mobilization might be different 
from CSNF. This FEP addresses Naval SNF cladding only. 
14 of 90 
6.16 2.1.02.26.0A Diffusion-controlled Excluded Diffusion-Controlled Cavity Growth (DCCG) was once thought to be a possible creep 
rupture mechanism that could occur under the temperature and pressure conditions that 
prevailed during dry storage of spent fuel and might occur during disposal. 
cavity growth (DCCG) 
in cladding 
Excluded 6.17 2.1.02.27.0A Localized (fluoride Fluoride is present in Yucca Mountain groundwater, and zirconium has been observed to 
corrode in environments containing fluoride. Therefore, fluoride corrosion of cladding may 
occur in waste packages. 
enhanced) corrosion 
of cladding 
Excluded Rockfalls may occur with blocks that are large enough to mechanically tear or rupture drip 6.18 2.1.07.01.0A Rockfall 
shields and/or waste packages. Seismic induced rockfall is addressed in a separate FEP. 
Volume increase of 
corrosion products 
impacts cladding 
Included 2.1.09.03.0A 6.19 Corrosion products have a higher molar volume than the intact, uncorroded material. 
Increases in volume during waste form and cladding corrosion could change the stress 
state in the material being corroded and lead to cladding unzipping. 
6.20 2.1.09.09.0A Electrochemical Excluded Electrochemical effects may establish an electric potential within the drift or between 
materials in the drift and more distant metallic materials. Migration of ions within such an 
electric field could affect corrosion of metals in the EBS and waste, and could also have a 
direct effect on the transport of radionuclides as charged ions. 
effects in EBS 
March 2004 
ANL-WIS-MD-000008 REV 01 
Table 6-1. LA FEP List for Cladding (Continued) 
Screening 
Decision FEP Name Section FEP Number Description 
Chemical effects of 
waste-rock contact 
Excluded 2.1.09.11.0A 6.21 Waste (CSNF, DSNF, and HLW) and rock are placed in contact by mechanical failure of the 
drip shields and/or waste packages. Chemical effects on the waste (e.g., dissolution) of 
waste may be enhanced or altered in a system where waste, rock minerals, and water are 
all in physical contact, relative to a system where only waste and water are in physical 
contact. 
6.22 2.1.11.05.0A Thermal expansion/ Excluded Thermally induced stresses could alter the performance of the waste or EBS. For example, 
thermal stresses could cause the waste form to develop cracks and create pathways for 
preferential fluid flow and, thereby, accelerate degradation of the waste. 
stress of in-package 
EBS components 
Gas generation (He) 
from waste form 
2.1.12.02.0A 6.23 Excluded Helium (He) gas production may occur by alpha decay in the waste. Helium production 
might cause local pressure buildup in cracks in the fuel and in the void between fuel and 
cladding, leading to cladding and waste-package failure. decay 
Excluded Gas generation can affect the mechanical behavior of the host rock and engineered 
barriers, chemical conditions, and fluid flow, and, as a result, the transport of radionuclides. 
Gas generation due to oxic corrosion of waste containers, cladding, and/or structural 
materials will occur at early times following closure of the repository. Anoxic corrosion may 
follow the oxic phase if all oxygen is depleted. The formation of a gas phase around the 
waste package may exclude oxygen from the iron, thus inhibiting further corrosion. 
Gas generation (H2) 
from waste package 
2.1.12.03.0A 6.24 
corrosion 
March 2004 15 of 90 
Clad Degradation  FEPs Screening Arguments 
Appendix A of this report contains tables of mean and maximum temperatures expected for the 
cladding, which are needed to evaluate some of the FEPs. For FEPs included in the TSPA-LA, 
alternative conceptual models are discussed in the specific report for the included FEP (BSC 
2003a [DIRS 162153]). For excluded FEPs, the discussions of alternative conceptual models, 
where appropriate, are included in this report. The output of this analysis (TSPA-LA 
dispositions for included FEPs and screening arguments for excluded FEPs) is intended to be 
used to support the project-specific FEP database and to promote traceability and transparency 
regarding FEP screening. 
Table 6-2. Corroborating Models and Information Used to Develop This Report 
Source Information 
Solubility of oxygen in zirconium 
Zircaloy-clad fuel rods before and after in-water spent fuel pool storage 
Baker 1992 [DIRS 149104] 
Bradley, et al. 1981 [DIRS 101564] 
Brossia et al. 2002 [DIRS 161988], 
Hydride reorientation 
Experiments with Zircaloy-4 fasteners 
Hydride reorientation 
SCC and surface passivity 
Threshold Intensity Factors 
Fuel not damaged in transportation 
Cladding degradation discussion 
Fuel assembly description 
Oxidation of UO2 is slower in steam than in air 
Accelerated high temperature tests 
Fuel/Zirconium reaction, pressures 
hydride reorientation 
Fuel Temperatures 
Oxygen effects on corrosion of zirconium alloys and gas effects 
Threshold stress intensity properties 
Pitting experiments Figure 3 
Chan 1996 [DIRS 111876] 
Clayton 1984 [DIRS 131741] 
Coleman 1982 [DIRS 111999] 
Cox 1973 [DIRS 152920] 
Cox 1990 [DIRS 152778] 
Debes 1999 [DIRS 161193] 
Dieter 1961 [DIRS 147973] 
DOE 1996 [DIRS 100320] 
Einziger and Kohli 1984 [DIRS 101605] Creep rupture studies 
Einziger 1994 [DIRS 100442] 
Einziger et al. 1982 [DIRS 101604] 
Garde 1986 [DIRS 101651] 
Garde 1989 [DIRS 113614] 
Garde 1991 [DIRS 101652] 
Garzarolli et al. 1979 [DIRS 149256] 
Huang 1995 [DIRS 101683] 
Jangg et al. 1978 [DIRS 110544], Tables Pitting experiments 
1, 6 
Knoll and Gilbert 1987 [DIRS 123682] Cask-drying 
Kohli and Pasupathi 1986 [DIRS 
Delayed hydride cracking unlikely at temperatures above 260C 
Pin degradation during reactor operation 
Figure of hydride concentration vs. burnup 
Precipitation and dissolution solubilities of hydrogen 
Comparison of creep models for Zircaloy 
Water-logged spent fuel storage rods 
131519] 
Little and Wagner 1996 [DIRS 131533] Microbially influenced corrosion 
Maguire (1984, Table 6)[DIRS 101717] Pitting experiments 
Mahmood et al. 2000 [DIRS 152241] 
Manaktala 1993 [DIRS 101719] 
Mardon et al. 1997 [DIRS 109213] 
McMinn et al. 2000 [DIRS 112149] 
McNeil and Odom 1994 [DIRS 131537] Sulfate-reducing bacteria do not affect zirconium alloys 
Peehs and Fleisch 1986 [DIRS 102065] Water-logged spent fuel storage rods 
Pescatore and Cowgill 1994 [DIRS 
102066] 
March 2004 16 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 6-2. Corroborating Models and Information Used to Develop This Report (Continued) 
Source Information 
Survey of all cladding failure modes, SSCs, creep, DHC, and others 
Series of strain tests on Zircaloy-2 
Stress intensity at crack tip 
Pitting experiments 
General corrosion, helium pressure, and DHC 
Corrosion rates of alloys 
Threshold stress intensity factor 
Hydrides, hydrogen absorption and oxygen dissolution in zirconium 
Pescatore et al.1990 [DIRS 101230] 
Puls 1988 [DIRS 102067] 
Reed-Hill 1973 [DIRS 121838] 
Riley and Covino 1982 [DIRS 161993], 
p. 11 
Rothman 1984 [DIRS 100417] 
Seibold et al. 2001 [DIRS 164458] 
Shi and Puls 1994 [DIRS 102084] 
Simpson and Chow 1987 [DIRS 164483] Hydride reorientation 
Smith 1966 [DIRS 149107] 
alloys 
electrophoresis or electro-osmosis Soderman and Jonsson 1996 [DIRS 
149441] 
Half life of hydrogen peroxide tefanic and LaVerne 2002 [DIRS 
166303] 
Iodine-induced SSC was not a major failure mechanism 
General Corrosion, pitting 
General, crevice and SCC tests with zirconium alloys 
Fluorides and pitting 
Tasooji et al. 1984 [DIRS 102093] 
Wallace et al. 1989 [DIRS 112344] Hydride reorientation 
Wolery and Daveler 1992 [DIRS 100097] Description of EQ6 
Yau 1983 [DIRS 149233] 
Yau 1984 [DIRS 102050] 
Yau and Maguire, M. 1990 [DIRS 
110761] 
6.1 DEGRADATION OF CLADDING FROM WATERLOGGED RODS 
FEP Number: 
2.1.02.11.0A 
YMP FEP Description: 
Failed fuel rods (attributed to breaches caused by manufacturing defects and reactor 
operations) comprise a small fraction of the fuel rods that are currently being stored in 
commercial reactor spent fuel pools. Failed fuel contains water in the fuel rod void space 
that may promote degradation of the spent fuel cladding. Such fuel is referred to as 
waterlogged. The moisture remaining in a dried fuel rod is used to determine the 
extent of degradation of spent fuel cladding. 
Descriptor Phrases: 
Residual moisture from storage enhances degradation of cladding 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
March 2004 17 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
2.1.12.03.0A, Gas generation (H 
Related FEPs: 
2.1.02.12.0A, Degradation of cladding prior to disposal 
2) from waste package corrosion 
Screening Argument: Degradation of cladding and waste form from waterlogged rods is 
excluded from the TSPA-LA because of low consequence. 
Waterlogged rods are already failed, and further cladding failures of intact rods and waste form 
degradation are not expected because of the small amount of water present. Few rods are 
breached at the time they are received at the repository (FEP 2.1.02.12.0A; BSC 2003a [DIRS 
162153], Section 6.2.1). The supply of water that remains in the fuel after cask drying will not 
be sufficient to oxidize the fuel to an unacceptable level (NRC 1997 [DIRS 101903], 
Section 8.V.1; Knoll and Gilbert 1986 [DIRS 123682], p. iii). Moisture is effectively removed 
from defective rods during the cask drying operations. The residual moisture of the cask 
atmosphere can be estimated based on the drying conditions where the cask or waste package is 
vacuum dried at 5 mbars (maximum water vapor pressure, equivalent to 7.7 g or 0.43 moles of 
water). The consequence of a small amount of remaining moisture on the waste form, waste 
package, and cladding is negligible. 
The fraction of fuel rods with breached cladding is currently estimated to be 0.01 to 1 percent 
(FEP 2.1.02.12.0A; BSC 2003a [DIRS 162153], Section 6.2.1), or about 5.6 rods in an average 
pressurized water reactor waste package. 
Since cask drying operations will remove most of the water from the fuel rods, it is reasonable to 
expect that the supply of water that remains in the fuel after cask drying will not be sufficient to 
oxidize the fuel. Standard Review Plan for Dry Cask Storage Systems (NRC 1997 [DIRS 
101903], Section 8.V.1) describes the cask drying criteria with reference to Knoll and Gilbert 
1987 [DIRS 123682]. Less than 0.43 mole (7.7 g) of H2O is expected to be present in a 7-m3 
cask after drying. This amount of water produces an insignificant potential for corrosion of the 
cladding during dry storage or during disposal (CRWMS M&O 1995 [DIRS 102829], p. 16). 
A PWR waste package contains approximately 2,500 kg (3.4  105 moles) of zirconium alloys 
(CRWMS M&O 2000 [DIRS 151659], Table 2). Less than 0.43 mole (7.7 g) of H2O is 
expected to be present (Knoll and Gilbert 1987 [DIRS 123682]) and general corrosion of the 
zirconium alloys from the water would only degrade a small fraction (0.2 mole). Hydrogen 
uptake would be insignificant compared to existing hydrides on the cladding. 
Kohli and Pasupathi (1986 [DIRS 131519], p. iii) discuss removal of water from waterlogged 
fuel rods. Two reactor-breached fuel rods were tested, along with two fuel rods that were 
intentionally defected after irradiation. A predetermined amount of moisture was added to the 
intentionally defected rods to enable the extent of the moisture released during the drying to be 
determined. The rods were dried in flowing argon at atmospheric pressure while being heated in 
a furnace. The center 1.8 m of the furnace was heated to 400C; the remainder was heated to 
200C. The reactor operations breached rods were dried in the as-received condition, then a hole 
was drilled in the cladding, water was injected, and the experiment was repeated. In the reactor 
breached rods, the bulk of the uncombined water was removed in 1 to 1.3 hours and all 
March 2004 18 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
measurable releases ended after 3.7 hours. This set of experiments demonstrated that standard 
cask drying procedures would remove water from failed rods. 
Peehs and Fleisch (1986 [DIRS 102065], pp. 199 to 202) described the behavior of waterlogged 
PWR fuel rods on heating at 400C in a hot cell. The bulk of the water was released during the 
cask drying operation. Testing results showed that the moisture can be removed from defective 
rods during the cask drying operations and the residual moisture of the cask atmosphere can be 
minimized. 
An alternative conceptual model for waterlogged rods would be that the failed rods contain 
water, which is then released into the waste package to be consumed by the carbon steel rack 
inside the waste package. This alternative conceptual model is possible because of the rate at 
which carbon steel racks react with water vapor to form rust is faster than that for zirconium and 
UO2. The hydrogen that is produced tends not to be absorbed by the zirconium alloys 
(FEP 2.1.12.03.0A). It would lead to no significant damage because of the limited amount of 
water and large volume of steel in a waste package. 
In conclusion, cladding and waste form degradation from waterlogged fuel rods is excluded from 
TSPA-LA. There are few failed rods in any waste package that could be waterlogged and the 
volume of water inside a rod is quite limited. Because the drying procedure is effective, the 
quantity of water is limited could only affect a small amount of fuel. Cladding failure due to rod 
waterlogging has a low consequence, and is excluded from further consideration. The magnitude 
and time of the resulting radiological exposures to the reasonably maximally exposed individual, 
or radionuclide releases to the accessible environment, would not be significantly changed by the 
omission of this FEP (degradation from waterlogged rods) from the performance assessment 
(TSPA-LA) model. 
6.2 DEGRADATION OF CLADDING PRIOR TO DISPOSAL 
FEP Number: 
2.1.02.12.0A 
YMP FEP Description: 
Certain aspects of cladding degradation occur before the spent fuel arrives at Yucca 
Mountain. Possible mechanisms include rod cladding degradation during reactor 
operation, degradation during wet spent fuel pool storage, degradation during dry storage, 
and rod degradation during shipping (from creep and from vibration and impact) and fuel 
handling. 
Descriptor Phrases: 
Degradation of cladding during dry storage 
Degradation of cladding during reactor operation 
Degradation of cladding during shipment and handling 
Degradation of cladding during wet (pool) storage 
Screening Decision: 
Included 
March 2004 19 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
TSPA-LA Disposition: Degradation of cladding prior to disposal is included in TSPA-LA 
cladding degradation abstraction. The failure rate from prior degradation is based on historical 
data on reactor operation. It also includes failure from wet pool storage and transportation 
(negligible), dry storage, and handling (including spent pool events). In the TSPA, this 
percentage of rods is available for radionuclide release through fast release and axial splitting 
when the waste package fails. It is specified as a 0.01 to 1 percent log uniform distribution 
(0.1 percent median). The TSPA abstraction models that all stainless steel cladding has failed 
and places the stainless steel cladding into waste packages as it arrives at the repository. This 
results in 3.5 to 7 percent (uniform distribution) of the waste packages containing stainless-steelclad 
fuel rods. These waste packages contain 15 to 30 percent stainless-steel-clad fuel rods that 
are failed and available for fast release and unzipping upon waste package failure. In the TSPALA, 
waste packages containing stainless-steel-clad fuel rods are considered a different fuel type 
group with a high initial cladding failure percent. The scientific analysis that describes the 
disposition in greater detail is presented in Clad Degradation  Summary and Abstraction for LA 
(BSC 2003a [DIRS 162153]). 
Related FEPs: 
2.1.02.11.0A, Degradation of cladding from waterlogged rods 
2.1.02.19.0A, Creep rupture of cladding 
6.3 GENERAL CORROSION OF CLADDING 
FEP Number: 
2.1.02.13.0A 
YMP FEP Description: 
General corrosion of cladding could expose large areas of fuel and produce hydrides. 
Descriptor Phrases: 
Dry oxidation of cladding 
General (uniform) corrosion of cladding 
Thermal effects on corrosion of cladding 
Wet oxidation (aqueous corrosion) of cladding 
Zirconium oxidation of cladding 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.15.0A, Localized (radiolysis enhanced) corrosion of cladding 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
2.1.02.17.0A, Localized (crevice) corrosion of cladding 
2.1.02.22.0A, Hydride cracking of cladding 
March 2004 20 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Screening Argument: General corrosion of cladding has been excluded from the TSPA-LA on 
the basis of low consequence. 
2SO4 represents a theoretical pH of about 1.15, well below anything expected in the waste 
2SO4, a protective 
The in-package chemistry model predicts that in most cases the pH remains above 3.5. When 
uncertainties are added, minimum pHs in the range of 2.5 to 3 are possible. These low pHs are 
caused by sulfur in the carbon steel rack being released and forming sulfuric acid (H2SO4). This 
period of low pH lasts for the time period when the carbon steel is corroding (see Figure 6.5-1 
for pH profile). Yau and Webster (1987 [DIR 100494], pp. 709  710, Figs. 5, 7, and Table 6) 
review the corrosion potential for zirconium alloys in sulfuric acid. They note that zirconium 
alloys resist attack from H2SO4 at all concentrations up to 70 percent and at temperatures to 
boiling (their Figure 5, reproduced in this report as Figure 6.3-1). A concentration of 70 percent 
H
package. In the range that zirconium alloys show corrosion resistance in H 
film is formed on the zirconium that is predominantly cubic zirconium oxide (ZrO2) with only 
traces of monoclinic phases. At higher concentrations than 70 percent, zirconium corrodes 
because loose films form that are zirconium disulfate tetrahydrate and partially zirconium 
hydrides. In concentrations of less than 65 percent H2SO4, zirconium can tolerate some amounts 
of strong oxidizing agents such as 200 ppm Fe3+ and 200 ppm NO3
- (Yau and Webster (1987 
[DIR 100494], Fig. 7). Moreover, in 20 percent or less H2SO4 (pH > -0.61), zirconium can 
tolerate a great amount of strong oxidizing agents. Consequently, zirconium equipment is often 
used in steel pickling. Zirconium alloys are used in the chemical industry under low pH 
conditions. In the manufacturing of H2O2, zirconium alloys are used to contain up to 65 percent 
H2SO4 at up to 150C. In the manufacturing of HNO3 zirconium alloys are used to contain the 
acid up to 65 percent concentrations and temperatures to 204C. A pH of 1.5 (minimum in waste 
package with uncertainties) represents only 0.15 weight percent of H2SO4 and is not expected to 
cause accelerated corrosion. 
At low pHs (below 0.6), the ZrO2 film will start to slowly dissolve. This is shown in Figure 6-2 
of Pitting Model for Zirconium-Alloyed Cladding at YMP (BSC 2003g [DIRS 164667]), where 
there is discontinuity in the corrosion potential at about pH = -0.6; below this pH, the corrosion 
potential starts to rapidly decrease. This change in behavior is attributed to dissolution of the 
general oxide surface. Since the in-package chemistry model predicts pHs above 0.6, 
accelerated general corrosion is excluded from TSPA-LA performance models. The Pourbaix 
diagram for zirconium (Pourbaix 1974 [DIRS 100817], p. 226) shows that ZrO2 could start to 
dissolve at a pH of 4. In the text describing the diagram, the author (Pourbaix 1974 [DIRS 
100817], pp. 228, 229) recognizes that the solubility is much too high (1.2 x 10-3 moles/liter) and 
not consistent with experimental observations of solubility (10-7 moles/liter). He attributes these 
high solubilities to not having modeled the dominant form of oxide (such as cubic ZrO2), which 
controls the dissolution. 
General corrosion is synonymous with zirconium oxidation for this repository application. The 
outer surface of the cladding becomes oxidized with a ZrO2 film, which adheres to the surface 
and slows down further oxidation (IAEA 1998 [DIRS 150560], Section 4.2.4). The oxidation 
could be from O2 consumption (dry oxidation) or H2O consumption (wet oxidation). For the fuel 
in the repository, this corrosion does not occur until the waste package is penetrated. 
March 2004 21 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Source: Yau and Webster, 1987 [DIR 100494], Figure 5 
Figure 6.3-1. The Iso-Corrosion Diagram for Zirconium in Sulfuric Acid (H2SO4) 
There are three possible effects of surface oxidation: 
1. The oxidation could thin the cladding, contributing to cladding failure by creep 
rupture (see FEP 2.1.02.19.0A). Wet oxidation generates hydrogen, and some of 
the hydrogen is absorbed into the cladding to form hydrides. 
2. This hydrogen pick up could lead to delayed hydride cracking (DHC), or general 
hydride embrittlement (see FEP 2.1.02.22.0A). 
3. In the extreme, the oxidation could lead to cladding failure and expose the fuel 
pellets to the waste package environment. 
The in-package chemistry model predicts that in most cases the pH remains above 3.5. Under 
these nominal chemical conditions (pH is greater than 3.5) in the repository, general corrosion 
failures of the cladding are unlikely. Waterside Corrosion of Zirconium Alloys in Nuclear Power 
Plants (IAEA 1998 [DIRS 150560]) summarizes much of the research on zirconium corrosion. 
Hillner et al. (1998, p. 9) [DIRS 100455] studied corrosion of Zircaloy and published a Zircaloy 
corrosion correlation based on Bettis Atomic Power Laboratory experiments. Bettis developed 
Zircaloy for naval reactors in the early 1950s and has an extensive database on Zircaloy 
performance, including continuous autoclave corrosion tests on some samples for 30 years. 
22 of 90 ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation . FEPs Screening Arguments 
Some samples have developed oxide thickness as great as 110 m, greater than those expected 
during repository corrosion based on Hillner et al. calculations. The experiments are consistent 
with diffusion of oxygen ions through the corrosion film being the rate-limiting phenomenon. 
This corrosion film is generated in layers, with the physical characteristics of lower layer staying 
consistent (uniform). The consistency of the lower 2 m of oxide film leads to a steady 
corrosion rate after a transition period. The recommended post-transition rate equation is: 
9 ) .Th = 1.7210  exp(.11452 /T (Eq. 6.3-1) 
where 
.Th = oxide growth rate, m/yr 
T = temperature, K 
This corrosion rate equation is similar to the equations developed by others, but predicts a 
slightly higher corrosion rate. This rate for unirradiated metal, from Equation 7 of Hillner at al. 
(1998 [DIRS 100455]), is doubled to represent the effect of irradiation (Hillner et al. 1998 [DIRS 
100455], pp. 6, 9). The correlation is converted to micrometers per year (1m = 14.7 mg/dm2) 
(IAEA 1998 [DIRS 150560], p. 178). The pretransition rate is slower than the posttransition 
rate. The effect of irradiation conditioning before corrosion is to accelerate the corrosion rate for 
a few micrometers. To be conservative, Hillner et al. (1998 [DIRS 100455]) doubled the 
corrosion rate for the full thickness of the metal. The correlation shows a strong Arrhenius 
temperature relationship with the corrosion rate becoming small below 200C. The Arrhenius 
temperature relationship is consistent with the low temperature data from Hillner et al. (1998 
[DIRS 100455]). Corrosion tests at 270C for 8.2 years have produced approximately 4 m of 
oxide, while corrosion tests at 360C have produced films 88 m thick in 7.8 years (Hillner et al. 
1998 [DIRS 100455], p. 25). Using this equation, one predicts an oxide thickness of only 0.22 
m at a long term repository temperature of 40C for a million years. At 80C, in one million 
years, the loss would be 14 m, a small fraction of the cladding thickness (570 m for a 
Westinghouse 1717 design). Hillner et al. also address other corrosion mechanisms, such as the 
instability of second-phase particles (or intermetallics), and identify no corrosion mechanism that 
is expected to fail the cladding. 
If the waste package fails at repository closure, the cladding will be exposed to the high 
temperature profiles shown in Figures A-4, A-5, and A-6 in the presence of oxygen. At these 
elevated temperatures some oxidation of the cladding will occur. The amount of oxide film from 
general cladding corrosion has been calculated with failure at closure and extending the 
calculation out to the end of the regulatory period (10,000 years). The amount of cladding that is 
consumed is approximately 57 percent of the oxide thickness (Eq. 6.3-1), a portion due to the 
volume increase associated with Zircaloy oxidation. The cladding oxidation calculation was 
done for the design basis (hot) rod in the hottest waste package, average rod in average waste 
package and average rod in a cool waste package. Three rod locations were considered: the 
center rod (hottest), an interior rod at a radius of 60 percent of the outer fuel radius, and the outer 
rod. Table 6.3-1 provides the cladding thickness for the nine rods considered. The hottest rod 
loses less than 10 percent of its wall thickness and is not expected to fail. All other rods have 
very little thinning (approximately 2 percent or less). If the waste package fails after the thermal 
peak (after 100 years; see Figures A-4, A-5, and A-6), little or no corrosion occurs. 
March 2004 23 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 6.3-1. Percent of Cladding Remaining in Waste Packages that Fail at Closure 
Rod (%) 
ANL-WIS-MD-000008 REV 01 
Center Rod, 60% ROD Outer Rod 
% (%) 
99.14 97.98 90.41 
99.68 99.08 98.42 
99.93 99.78 99.59 
Hot Waste Package, Hottest Fuel Loading 
Mean Waste Package, Uniform Fuel Loading 
Cool Waste Package, Uniform Fuel Loading 
Hillner et al. (1998 [DIRS 100455], Figure 5) compare the weight gain of the samples in water, 
which correlates to the corrosion rate in water (the correlation used here) with that of steam. The 
steam corrosion rate is about 30 to 40 percent slower. A humid-air environment is expected to 
last for 470 to 3,500 years (Table A-4). Einziger (1994 [DIRS 100442], p. 556, Equation 14) 
states that dry oxidation of zirconium is slightly slower than the wet corrosion rate. 
There are experimental corroborating observations about the slow corrosion rates of zirconium 
alloys near ambient temperature (27C). Bradley et al. (1981 [DIRS 101564], p. 38) performed 
metallurgical examinations of Zircaloy-clad fuel rods from two assemblies (0551 and 0074) of 
the Shippingport PWR Core 1 blanket fuel after extended in-water spent fuel pool storage (21 
years for 0551, and 16 years for 0074). The oxide film thickness on the Shippingport fuel rods 
after reactor operation was reported to be an average cladding oxide film thickness of 1.8 m 
(0551) and 2.4 m (0074). After extended in-water spent fuel pool storage, the average cladding 
oxide film thickness was found to be 1.7 m (0551) and 2.3 m (0074) (Bradley et al. 1981 
[DIRS 101564], p. 38). The slight disagreement in these values is attributed to differences in 
measurement technique and experimental error. These results led to the conclusion that no 
significant change in oxide thickness occurred even after 16 to 21 years of pool storage. This 
conclusion is further supported by the observation that Zircaloy tube sheets (that had been cut to 
remove assembly 0551 fuel rods in 1960) stored in water for over 20 years were unblemished 
and showed no evidence of reaction with water. 
Rothman (1984 [DIRS 100417], pp. 6 to 13, Table 3) discusses cladding oxidation in repository 
conditions in great detail and compares the predicted cladding loss using six different oxidation 
correlations and predicts cladding thinning of 4 to 53 m after 10,000 years at 180C 
(a conservative temperature condition since the repository cools after a few hundred years) 
which would not lead to rod failure. 
The cladding corrosion analyses (Hillner et al. 1998 [DIRS 100455], Rothman 1984 [DIRS 
100417]) are based on the performance of Zircaloy-4 (Zirc-4), the cladding used in most PWRs 
before the year 2000. Today, most American PWRs use zirconium-niobium alloys for their 
reloads. Fuel supplied by Westinghouse (including some Combustion Engineering designs) is 
clad with ZIRLO (a 1 percent niobium alloy) and most fuel supplied by Framatome ANP uses 
M5 (trademark of Framatome ANP, also 1 percent niobium). Seibold et al. (2001 [DIRS 
164458], Figure 2, reproduced here as Figure 6.3-2) compare the oxidation rate of various 
cladding alloys as a function of burnup. In Figure 6.3-2, the HPA 4 alloys include ZIRLO and 
M5 is represented in the group Zr-1Nb. Current NRC guidance requires fuel oxide thickness 
to be less than 100 m at end of life. The figure shows that Zirc-4 is limited to fuel cycles less 
than about 54 MWd/kgU and that advanced alloys are required for higher burnups. ZIRLO has 
24 of 90 March 2004 Clad Degradation  FEPs Screening Arguments 
an oxidation rate of about half that of Zirc-4. In PWR environments, M5 has an oxidation rate 
of about one fifth that of Zirc-4. The hydride content of the cladding is directly proportional to 
the oxide thickness. At a given burnup, ZIRLO will have about half of the hydrides and M5 
will have one fifth of the hydrides as Zirc-4. Therefore, the new cladding is less likely to fail 
from general corrosion or hydride embrittlement (FEP 2.1.02.22.0A). The thinner oxide 
thickness produces lower fuel pellet temperatures and, thus, lowers fission gas release fractions. 
The remaining metal thickness is also greater. Therefore, cladding stresses are lower in 
advanced alloys and less creep is expected (FEP 2.1.02.19.0A). 
As alternative conceptual models for general corrosion, Hillner et al. (1998 [DIRS 100455], 
Table 4) provides the expected corrosion for cladding after 10,000 years at 180C using eight 
different corrosion equations developed by others. These alternative models show that Hillners 
equation is conservative and general corrosion is not expected to be significant in the repository. 
In conclusion, cladding degradation from general corrosion is excluded from TSPA-LA. The 
small amount of corrosion that will occur during the regulatory period will not penetrate the 
cladding and therefore will not affect the release of radionuclides. Cladding failure due to 
general corrosion has a low consequence, and is excluded from further consideration. The 
magnitude and time of the resulting radiological exposures to the reasonably maximally exposed 
individual, or radionuclide releases to the accessible environment, would not be significantly 
changed by the omission of this FEP (general corrosion of cladding) from the performance 
assessment (TSPA-LA) model. 
Source: Seibolt et al. 2001 [DIRS 164458], Figure 2 
Figure 6.3-2. A Comparison of Oxide Thicknesses for Different Zirconium Alloys 
ANL-WIS-MD-000008 REV 01 25 of 90 
Maximum Oxide Thickness (m) 
Average Fuel Rod Burnup 
March 2004 Clad Degradation  FEPs Screening Arguments 
6.4 MICROBIALLY INFLUENCED CORROSION (MIC) OF CLADDING 
FEP Number: 
2.1.02.14.0A 
YMP FEP Description: 
Microbially Influenced Corrosion (MIC) of cladding potentially may be a local cladding 
corrosion mechanism where microbes produce a local acidic environment that could 
produce multiple penetrations through the fuel cladding. 
Descriptor Phrases: 
Microbial induced pH change enhances localized corrosion of cladding 
Microbiologically influenced corrosion (MIC) of cladding 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.15.0A, Localized (radiolysis enhanced) corrosion of cladding 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
2.1.02.17.0A, Localized (crevice) corrosion of cladding 
Screening Argument: Microbially influenced corrosion (MIC) activity is excluded as a 
cladding failure mechanism because of low consequence. Microbes are expected to be present in 
the repository, but MIC of cladding is not expected to cause cladding failure and, therefore, not 
have a significant effect on radionuclide exposures to the RMEI. 
The term microbially influenced corrosion (MIC) is commonly used to designate corrosion 
caused by the presence and activities of microorganisms at the surfaces of metals. 
Two studies of MIC on spent nuclear fuel have been performed. Wolfram et al. (1996 [DIRS 
165268], pp. iii, iv) measured microbial activity in spent fuel pools. They concluded that all 
spent fuel pools tested contained microbial colonies. They also performed a literature search and 
concluded that There was no evidence in the literature that zirconium or its alloys are 
susceptible to MIC. 
Hillner et al. (1998 [DIRS 100455], p. 11) studied the corrosion of Zircaloy-clad fuels under 
repository conditions. They indicate that there are two major forms of MIC for materials being 
considered for waste packages. They are (1) sulfide attack through the action of sulfate reducing 
bacteria (SRB) and (2) corrosion induced by organic acids secreted from certain bacteria. With 
respect to attack by SRB, Hillner et al. (1998 [DIRS 100455]) reference the work of McNeil and 
Odom (1994 [DIRS 131537], p. 176), which indicates by thermodynamic calculations that SRB 
do not affect zirconium alloys. With respect to corrosion induced by organic acids, Hillner et al. 
(1998 [DIRS 100455], p. 11) noted that it is most unlikely because of zirconiums tolerance of a 
wide range of pHs and it is unlikely that production of weak organic acids will have an adverse 
March 2004 26 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
effect on the passivation of Zircaloy by a ZrO2 film. Yau and Webster (1987 [DIRS 100494], 
p. 717) also note that zirconium alloy resists a wide range of organic compounds, including 
acetic acid, acetic anhydride, formic acid, urea, ethylene dichloride, formaldehyde, citric acid, 
lactic acid, oxalic acid, tannic acid, and trichloroethylene. This supports the concept that organic 
solutions produced by MIC are unlikely to cause significant acceleration of the corrosion of 
zirconium alloys. In-Drift Microbial Communities (CRWMS M&O 2000b [DIRS 151561], p. 
154) evaluated in-drift microbial communities and concluded that the estimates of microbial 
masses growing in the potential repository system suggest that the effect to the in-drift 
geochemistry should be small. 
Little and Wagner (1996 [DIRS 131533], p. 367) corroborate this information in an overview of 
MIC of metals and alloys used in the storage of nuclear wastes. They indicate that MIC is a form 
of localized corrosion that results in pitting, selective leaching, crevice corrosion, under-deposit 
corrosion, and enhanced erosion and corrosion. Little and Wagner (1996 [DIRS 131533], pp. 
367 and 368) describe several mechanisms for MIC. In addition, various case studies are 
presented that document MIC of alloys of iron, nickel, and copper. However, it should be noted 
that there is no indication in the literature of MIC occurring on zirconium metal or alloys. Yau 
and Webster (1987 [DIRS 100494], p. 709) report no corrosion of zirconium alloy from marine 
organisms was found during sea water corrosion tests for 129 days. 
MIC is excluded as a component of the localized (pitting) corrosion model where MIC could 
cause a localized suppression of the water pH and permit other aggressive species to attack 
cladding. The zirconium alloy pitting model (BSC 2003g [DIRS 164667], Section 8) shows that 
pitting is dependent on the concentration of chlorides, ferric ions, and hydrogen peroxide (from 
radiolysis). 
One U.S. commercial nuclear plant spent fuel pool experienced a significant MIC event that 
lasted for about four years (Ralph et al. 2002 [DIRS 161992]). After an extended lay-up period, 
the spent fuel pool water was found to contain a significant amount of algae and bacteria. 
Biological agents were purged using controlled additions of chlorine and hydrogen peroxide 
before the pool was returned to normal operating chemistries. 
The assessment revealed that the steel rack corrosion products were up to 2.5 cm thick, and they 
had started to engulf the individual fuel rods or flow channels of the stored assemblies in the 
region where the rack and plates contacted the fuel assemblies. The corrosion product had 
adhered to the fuel. The iron oxide was composed of FeO, Fe2O3, and Fe3O4. Ralph et al. (2002 
[DIRS 161992], p. 6) states: 
One fuel assembly was removed from its storage location and its channel was 
removed. The oxide from contact with the carbon steel rack was removed with a 
water lance utilizing 350 to 700 kg/cm2 of water pressure. A camera with 
resolution of 0.025 mm was used to inspect the channel. The channel surface 
appeared uniform and smooth. No pitting, white discoloration or surface 
anomalies were observed. 
The paper concluded: The fuel cladding was not affected through any type of corrosion. 
Therefore, the corrosion did not change the classification of the fuel as intact or damaged. As 
March 2004 27 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
with the experiments by Yau (1983 [DIRS 149233]), the lack of pitting or stress corrosion 
cracking implies that the corrosion potential (Ecor) was not elevated to exceed Erp, even with 
Fe3O4 present and adhering to the zirconium oxide film. MIC colonies could also have locally 
suppressed the pH but, again, no localized corrosion was observed. 
Adler Flitton et al. (2002 [DIRS 161991], p. 4) buried various metal samples in an arid vadose 
zone environment for three years. They reported indications of pitting from MIC on some of the 
metals, but observed no pitting on the zirconium alloy samples. 
In summary, microbes are expected to be present in the repository, but MIC of cladding is not 
expected to cause cladding failure. The NRC requirements in 10 CFR 63.114 (e and f) allow the 
exclusion of MIC from the TSPA-LA because the omission would not significantly change the 
magnitude and time of the resulting radiological exposures to the reasonably maximally exposed 
individual (RMEI), or radionuclide releases to the accessible environment. 
6.5 LOCALIZED (RADIOLYSIS ENHANCED) CORROSION OF CLADDING 
FEP Number: 
2.1.02.15.0A 
YMP FEP Description: 
Radiolysis in a nitrogen/oxygen gas mixture with the presence of water film results in the 
3). Hydrogen peroxide (H2O2) is formed in the water from formation of nitric acid (HNO 
radiolysis. These chemicals can enhance corrosion of the fuel cladding. 
Descriptor Phrases: 
Localized corrosion of cladding enhanced by radiolysis induced pH change 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
2.1.02.17.0A, Localized (crevice) corrosion of cladding 
2.1.13.01.0A, Radiolysis 
Screening Argument: Radiolysis, as a cladding failure mechanism, is excluded from the 
TSPA-LA on the basis of low consequence. 
The in-package chemistry model (BSC 2003f [DIRS 161962], Attachment III) addressed the 
change of water chemistry with the inclusion of nitric acid and hydrogen peroxide production 
from radiolysis. In the analysis, all of the nitric acid that can be produced in a moist waste 
package was absorbed into the water film on the cladding surface. The radiation field was 
modeled as being constant at the dose at 500 years although it decreases with time. 
March 2004 28 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
The effect of radiolysis on the in-package chemistry was handled in a nonmechanistic manner 
(i.e., EQ6 does not have the facility to model the radiolysis process directly), therefore, only the 
products of radiolysis were included as inputs in EQ6 simulations. EQ6 is a reaction path code 
which models water, rock or other material interactions or fluid mixing in either a pure reaction 
progress mode or a time mode (Wolery and Daveler 1992 [DIRS 100097]). A series of runs was 
performed where nitric acid and hydrogen peroxide were included as inputs in EQ6 runs and the 
results of these simulations were compared to their non-radiolysis counterparts. Two base case 
runs (C12C25 and C22C25, BSC 2003f [DIRS 161962], Section 6.5.4) were used to test the 
effects of radiolysis. These runs were chosen because they represent the median fuel exposure 
value and the low and median water flux values. It would be expected that the effects of nitric 
acid and hydrogen peroxide additions would be greater at the low end of the flux range. Two 
simulations were performed for each file, the first using the base case nitric acid and hydrogen 
peroxide rates and the second multiplying the HNO 
The results of the simulations are displayed in Figure 6.5-1, where the pH profiles for the various 
runs are displayed versus time. These results show that neither the base case nor the 10 base 
case generation rates of HNO3 and H2O2 had an impact on the in-package pH. Therefore, it may 
be concluded that if radiolysis only affects the chemistry via HNO3 and H2O2 generation, then it 
will not be a significant process with regard to influencing the in-package chemistry. The 
radiolysis did not significantly affect the concentrations of Cl-, Fe3+, or H2O2 and therefore did 
not change the corrosion potential of the passive film on the zirconium alloy. 
Table 6.5-1 provides the ferric iron, chloride, and hydrogen peroxide maximum concentrations 
for the simulations. Again, runs with HNO3 and H2O2 input show little deviation compared to 
the runs without. 
9.0 
8.0 
7.0 
6.0 
5.0 
4.0 
3.0 
1.0E-02 1.0E+00 1.0E-01 
Source: BSC 2003f [DIRS 161962], Attachment III 
Figure 6.5-1. pH Profiles Showing How the Radiolysis Inputs Affect the pH 
ANL-WIS-MD-000008 REV 01 29 of 90 
pH 
3 and H2O2 rates by a factor of ten. 
C22C25 
C22C base case 
C22C base case 10x 
C12C25 
C12C base case 
C12C base case 10x 
1.0E+04 1.0E+03 1.0E+02 1.0E+01 
Time (yrs) 
1.0E+05 
March 2004 Clad Degradation  FEPs Screening Arguments 
Table 6.5-1. Chloride, Ferric Iron, and Hydrogen Peroxide Molality 
Maximum Molalities 
Cl- Fe+++ H2O2(aq) 
C22C25 6.6E-04 8.8E-11 3.7E-19 
C22CBC 6.6E-04 8.8E-11 3.7E-19 
C22BC10x 6.5E-04 8.8E-11 3.7E-19 
C12C25 9.7E-04 5.3E-11 3.7E-19 
C12CBC 9.6E-04 5.3E-11 3.7E-19 
C12BC10x 9.2E-04 5.4E-11 3.7E-19 
Source: BSC 2003f [DIRS 161962] 
The Center for Nuclear Waste Regulatory Analysis (CNWRA) performed a series of corrosion 
tests where hydrogen peroxide was added to the ongoing test while the corrosion potential was 
being measured. Greene et al. (2000 [DIRS 145073], Figure 8) shows two experiments where 
H2O2 was added and the corrosion potential was measured. In one test, the H2O2 was added two 
different times. In the three cases where H2O2 was added, the effect of the hydrogen peroxide 
rapidly died out. In another test (Greene et al. 2000 [DIRS 145073], Figure 11), a sample that 
was oxidized in air at 200C was exposed to a solution of 1M NaCl. When 5 mM H2O2 was 
added, the corrosion potential increased by 0.275 VSCE (Volts, Standard Calemel Electrode scale) 
and pitting was observed. In this experiment the corrosion potential normally is nominally 0.07 
VSCE, and the repassivation potential is 0.04 VSCE, so the increase in corrosion potential is 
significant. The concentrations of chloride and hydrogen peroxide in this experiment are many 
orders of magnitude higher than expected in the waste package (see concentrations in 
Table 6.5-1). 
Brossia et al. (2002 [DIRS 161988], Figure 3) report two experiments where H2O2 was added to 
ongoing corrosion potential tests. The metal samples had oxide coatings of 1.7 m and 3.4 m 
thick. The initial solution contained 0.1 M NaCl at 95C and 5 mM H2O2 was added. In both 
tests the corrosion potential initially increased, but later one test showed decreasing corrosion 
potentials. Pitting was not observed in either experiment. Again, these concentrations are higher 
than expected in the in-package chemistry. 
The experiments by Greene et al. (2000 [DIRS 145073], Figure 8) discuss the stability of H2O2. 
IAEA (1998 [DIRS 150560], p. 220) and suggest that H2O2 is known to decompose catalytically 
on the surfaces of various types of materials at ambient temperatures, and that zirconium oxides 
enhance decomposition. H2O2 decomposes in bulk solutions at elevated temperatures (tefanic 
and LaVerne 2002, [DIRS 166303], Abstract). Table 6.5-2 provides the half-life of H2O2 at 
various temperatures and shows that bulk decomposition during the regulatory period will be 
significant. 
The in-package model, coupled with the pitting experiments and pitting model (see FEP 
2.1.02.16.0A and Table 6.6-1), shows that radiolytic production of nitric acid and hydrogen 
peroxide is not sufficient to influence the corrosion potentials significantly at the steady state 
concentrations expected in the repository and produce pitting. Zirconium alloys have been 
shown to be relatively inert in both nitric acid and hydrogen peroxide as discussed by Yau and 
Webster (1987 [DIRS 100494]) and in Clad DegradationLocal Corrosion of Zirconium and Its 
30 of 90 
EQ6 Input File 
ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation  FEPs Screening Arguments 
Alloys Under Repository Conditions (CRWMS M&O 2000d [DIRS 136058], Sections 6.1.6 and 
B.4). For example, the chemical processing industry uses peroxide strengths of 90 percent with 
zirconium equipment. The service life has been increased by an order of magnitude compared to 
graphite components previously used, which were generally considered to be inert. In nitric 
acid, zirconium and its alloys are inert up to acid concentrations of 65 weight percent. Since 
radiolysis does not produce nitric acid and hydrogen peroxide at greater concentrations, there 
will be no impact (low consequence) on the uniform corrosion rate as a result of radiolysis. 
Table 6.5-2. Half-Life of Hydrogen Peroxide at Various Temperatures 
FEP Number: 
2.1.02.16.0A 
TSPA-LA Disposition: 
NA 
ANL-WIS-MD-000008 REV 01 
In conclusion, cladding degradation from radiolysis enhanced corrosion is excluded from TSPALA. 
Radiolytic production of nitric acid and hydrogen peroxide was included in the in-package 
chemistry model (BSC 2003f [DIRS 161962], Attachment III) and this analysis showed that 
radiolysis had a small effect on the chemistry. Experiments where hydrogen peroxide was added 
to tests show that in many cases the effect of the hydrogen peroxide quickly becomes negligible. 
Radiolysis by itself is not expected to damage the cladding (low consequence). Cladding failure 
due to radiolysis enhanced corrosion has a low consequence, and is excluded from further 
consideration. The magnitude and time of the resulting radiological exposures to the reasonably 
maximally exposed individual, or radionuclide releases to the accessible environment, would not 
be significantly changed by the omission of this FEP (radiolysis enhanced corrosion) from the 
performance assessment (TSPA-LA) model. 
6.6 LOCALIZED (PITTING) CORROSION OF CLADDING 
YMP FEP Description: 
Localized corrosion in pits could produce penetrations of cladding. 
Descriptor Phrases: 
Pitting corrosion of cladding 
Screening Decision: 
Excluded - Low Consequence 
Temperature (C) 
25 46 
50 4.9 
73 0.83 
95 0.19 
100 0.14 
Source: tefanic and LaVerne 2002 [DIRS 
166303], Abstract 
31 of 90 
Half-Life (Days) 
March 2004 Clad Degradation  FEPs Screening Arguments 
Related FEPs: 
2.1.02.15.0A, Localized (radiolysis enhanced) corrosion of cladding 
2.1.02.17.0A, Localized (crevice) corrosion of cladding 
2.1.02.13.0A, General corrosion of cladding 
Screening Argument: Localized corrosion of the cladding from pitting has been excluded on 
the basis of low consequence. A comparison of the expected in-package chemistry to the 
chemical composition where pitting is observed shows that pitting is not expected. Therefore, 
the inclusion of a pitting model in TSPA-LA would not have a significant effect on the 
magnitude and time of radiological exposures or radionuclide releases (low consequence). 
A zirconium-pitting model (BSC 2003g [DIRS 164667]) was developed to investigate the 
chemical conditions at which pitting occurs. Zirconium alloys are susceptible to pitting in a 
particularly aggressive combination of chloride (Cl-) ions, ferric ions (Fe+3), or hydrogen 
peroxide (H2O2). In order to predict cladding failure from chloride pitting, a review of the 
literature for pitting rates and electrochemical data for various zirconium alloys was conducted. 
Based on this review of the literature, failure criteria were constructed based on an 
electrochemical definition of pitting as the condition at which the corrosion potential for aspolished 
metal exceeds repassivation potential (i.e., Ecor > Erp ). Corrosion potential and 
repassivation potential values were obtained for as-polished zirconium alloys in various solution 
concentrations of Cl-, Fe+3, and H2O2 using measurements obtained from various experiments. 
The model to predict repassivation potential depends only on chloride concentration in the 
solution. The corrosion potential for as-polished metal (Ecor) was modeled by performing a 
regression analysis to fit experimental data with varying molar concentrations of Cl-, Fe+3, and 
H2O2. The model describes the conditions where pitting was observed in experiments. High 
concentrations of chlorides at extremely low pH (below -0.6) can lead to the general dissolution 
of the protective zirconium oxide film. 
This model was evaluated using in-package chemistry, including the production of nitric acid 
and hydrogen peroxide from radiolysis (see FEP 2.1.02.15.0A and Table 6.5-1). Table 6.6-1 
provides the corrosion potential for as-polished metals, repassivation potential, and potential 
differences for the cases described in Table 6.5-1. Table 6.6-1 shows that pitting is not expected 
because the repassivation potential exceeds the corrosion potential for as-polished metal. 
Figure 6.6-1 shows the chemical conditions where pitting would occur and the chemical regions 
predicted. No pitting was predicted to occur for any conditions that were associated with 
Zircaloy cladding in the repository. In a sensitivity study with acid production from radiolysis 
increased by a factor of ten, no pitting was predicted to occur. 
The in-package chemistry model predicts that in most cases the pH remains above 3.5. When 
additional uncertainties are added, minimum pHs in the range of 2.5 are possible. To address the 
potential for pitting at low pHs, an in-package chemistry sensitivity study was performed. The 
radiolytic production of nitric acid and hydrogen peroxide was 10 times nominal (case 
C12BC10x in Table 6.6-1) but the amount of sulfur in the carbon steel rack was increased by a 
factor of 20. Sulfur in the steel suppresses pH, as shown in Figure 6.5-1. Increasing the amount 
of sulfur by a factor of 20 reduces the pH to a minimum of 2.2. At this pH, the concentration of 
Fe3+ to 3.8 x 10-6 molality and chloride is 9.7  10-4 molality. Hydrogen peroxide is less than 
10-12 molality. For this chemical composition, the pitting model predicts a corrosion potential of 
March 2004 32 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
-0.2 VSCE and a repassivation potential of +0.29 VSCE. Since the repassivation potential exceeds 
the corrosion potential, no pitting is predicted. At the minimum pHs predicted in the in-package 
chemistry model, no pitting is predicted to occur. 
Table 6.6-1. Corrosion and Repassivation Potentials for In-Package Chemistry Condition 
Ecor 
VSCE 
10
1 
0.1 
No 
Pitting 
10 1 
0.01 
Chloride Ion Concentration [Cl-], M 
Case 
# 
C22C25 -0.20 0.30 
C22CBC -0.20 0.30 
C22BC10x -0.20 0.30 
C12C25 -0.20 0.29 
C12CBC -0.20 0.29 
C12BC10x -0.20 0.29 
Source: BSC 2003g [DIRS 164667] 
NOTE: a Negative difference implies pitting is possible. 
0.001 
0.0001 
Source: BSC 2003g [DIRS 164667] 
Y, In-package Chemistry 
Fe3+ < 1 ppm 
Ferric Ion Concentration [Fe3+], PPM 
Erp 
VSCE 
x = Model 
Y = In-package chemistry 
100 
x 
Differencea (Erp-Ecor) 
VSCE 
0.50 
0.50 
0.50 
0.49 
0.49 
0.49 
FeCl3-3g.spw 
Pitting 
x Uncertainty 
xx 
x 
1000 10000 
Figure 6.6-1. Comparison of In-Package Chemistry with Chemical Conditions Necessary for Pitting 
The pitting model was generated with data for as-polished samples, but CSNF is coated with 
thick oxides (54 m mean, 5 to 95 percent m range = 112 m to 5.3 m, CRWMS M&O 2000c 
[DIRS 151659], Section 6.4). These oxides will affect the measured corrosion potential (open 
circuit potential) because of the high electrical resistance of the coating, but will not increase 
susceptibility to pitting corrosion. This sensitivity to oxide thickness was demonstrated in a 
series of four tests performed on zirconium samples at the Center for Nuclear Waste Regulatory 
ANL-WIS-MD-000008 REV 01 March 2004 33 of 90 Clad Degradation  FEPs Screening Arguments 
Analyses. Two tests had oxide thickness of 1.7 m and 2 tests had 3.4 m coatings (Brossia et 
al. 2002 [DIRS 161988], Figure 3). These samples were exposed to 0.1M of NaCl. Prior to 
addition of oxidants to the solution, the highest corrosion potential was 0.67 VSCE, well above the 
repassivation potential of 0.12 VSCE. Two other samples also had potentials above the 
repassivation potential. No pitting was observed in the samples. Yau and Maguire (1990 [DIRS 
110761], Fig. 4) showed that zirconium annealed in air (air oxidized) also had a corrosion 
potential that was both above the as-polished value and the repassivation potential. Again, no 
pitting was observed. These observations show that the corrosion potential for oxide coated 
material can exceed the repassivation potential and pitting will not occur. Ecor > Erp is a 
necessary but not sufficient condition for pitting. For the case where the corrosion potential is 
raised above the repassivation potential only due to oxide formation, pitting will not occur, as 
discussed below. 
When pitting occurs on metal surfaces, the pit behaves as an anode and the surrounding metal 
surface behaves as a cathode. The metal in the pit gives up electrons, becomes oxidized, and 
goes into solution. To support pitting, the electrons must be conducted to the free surface of the 
surrounding metal and be accepted by an oxidizing agent (e.g., H2O, Fe3+, or O2). The potential 
for pitting starts when some event creates a site with less protection than the surrounding metal. 
For pitting to then occur, three conditions must be satisfied: 
1. The surface of the surrounding metal must be able to transport the electrons into 
the solution 
2. Oxidizing agents in the solution must be present to accept the electrons 
3. The pit site must support transport of the metal ions into solution. 
Zirconium oxide is a poor conductor and a thick oxide coating restricts the current of electrons 
from the metal into the solution (Condition 1 above). This insulation results in the measured 
corrosion potential increasing from the Zircaloy corrosion potential toward the oxide-surface 
redox-reaction potential as the oxide insulation increases. The redox potential is not relevant to 
pitting susceptibility but is a consequence of inerting of the Zircaloy metal. If the oxide is 
defected, the metal corrosion potential decreases to that of the bare metal, and pitting occurs only 
if oxidizing species in solution or an applied potential forces the bare-metal Zircaloy corrosion 
potential above the repassivation potential. If there are insufficient oxidizing agents in the 
solution to support a critical current (Condition 2 above), then a new protective oxide surface 
will form on the potential pit surface and prevent dissolution of the metal (Condition 3 above). 
In summary, although the corrosion potential (Ecor) is greater than the repassivation potential 
(Erp) for zirconium with thick oxide layers, both empirical evidence and understanding of the 
response to local breakdown of the protective oxide layer (e.g., due to oxide layer cracking) 
indicate that pitting does not occur. The empirical evidence discussed in the above paragraph 
(Brossia et al. 2002, [DIRS 161988] Figure 3, Yau and Maguire 1990 [DIRS 110761], Fig. 4) 
showed that zirconium with thick oxide layers did not pit. The oxide layer on the zirconium 
surrounding the breakdown region prevents sufficient cathodic current to support pitting. 
Therefore, both the anodic and cathodic reactions supporting corrosion occur in the breakdown 
region, a condition that can not occur for pitting. In addition, upon breakdown of the oxide, the 
March 2004 34 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
corrosion potential drops to about the corrosion potential of the as-polished zirconium. In effect, 
the breakdown region will not be influenced by the surrounding oxidized zirconium surface and 
will behave as as-polished zirconium. Therefore, results obtained using as-polished zirconium 
coupons apply in assessing the corrosion behavior in the breakdown region. For the experiments 
described by Brossia et al. (2002, [DIRS 161988], Figure 3), the corrosion potential for an aspolished 
surface (-0.11 VSCE) is significantly less than the repassivation potential (0.12 VSCE), 
and pitting was neither observed nor expected to occur. 
As further validation of the pitting model and results, a review of the literature for pitting 
observations for various zirconium alloys was conducted. Figure 6.6-2 provides the chemical 
environment where pitting is observed. Figure 6.6-2 contains experimental results by Yau 
(CRWMS M&O 2000d [DIRS 136058], Table 4), Maguire (1984 [DIRS 101717], Table 6), 
Riley and Covino (1982 [DIRS 161993], p. 11), Jangg et al. (1978 [DIRS 110544], Tables 1 and 
6), Greene et al. (2000 [DIRS 145073], Figure 8) and Brossia et al. (2002 [DIRS 161988], Figure 
3). Each experiment is shown with a different symbol and whether pitting is observed (symbols 
q, r, s, u). Yaus data (symbols a, q) show that pitting does not always occur under similar 
conditions. 
The in-package chemistry model (BSC 2003f [DIRS 161962], Attachment III) calculated the 
expected concentrations of chemicals in a thin film of water on the fuel. Table 6.5-1 provides 
the maximum concentrations of Cl-, Fe3+, and H2O2 expected with the inclusion of radiolysis. 
Figure 6.5-1 provides the evolution of the pH over time for the thin film of water. The pH stays 
above 3.5, which means few, if any, ferric ions (Fe3+) would exist. The concentrations of Cl- and 
Fe3+ (Table 6.5-1) are not shown in Figure 6.6-2 because they are well to the bottom and left of 
the scales shown. A comparison of the chemistry and the regions where pitting is observed 
shows that cladding failure from pitting is unlikely. 
Most of the tests shown in Figure 6.6-2 were performed on as-polished metal samples while 
CSNF cladding is coated with thick oxides (mean value = 54 m, CRWMS M&O 2000c [DIRS 
151659], p. 30). Two tests were performed on zirconium samples with oxide thickness of 1.7 
and 3.4 m (Brossia et al. 2002 [DIRS 161988], Figure 3). These samples were exposed to 0.1M 
of NaCl and either 5 mM FeCl3 or 5 mM H2O2. Pitting was not observed in any of these four 
tests (see symbol e in Figure 6.6-2 for the two ferric chloride tests). 
Application of the pitting model (BSC 2003g [DIRS 164667]) to predict the pitting/no-pitting 
boundary is also illustrated in Figure 6.6-2. The lines (with uncertainty) where the model 
predicts the onset of pitting are shown as straight lines in log/log scale. Each line represents the 
chemical composition (Cl- and Fe3+) where Ecor = Erp. A straight line is expected since both Erp 
and Ecor linearly depend on the log of the Cl- and Fe3+ concentrations. The application of the aspolished 
pitting model to CSNF (with thick oxides) is expected to be conservative based on 
increased difficulty to provide the cathodic currents needed to support pitting. 
In conclusion, cladding degradation from localized (pitting) corrosion of the cladding is excluded 
from TSPA-LA. A comparison of the expected in-package chemistry to the chemical 
composition where pitting is observed shows that pitting is not expected. Cladding failure due to 
pitting has a low consequence, and is excluded from further consideration. The magnitude and 
time of the resulting radiological exposures to the reasonably maximally exposed individual, or 
March 2004 35 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
radionuclide releases to the accessible environment, would not be significantly changed by the 
omission of this FEP (localized (pitting) corrosion of the cladding) from the performance 
d t t 
c 
100 
Ferric Ion Concentration [Fe3+], PPM 
FeCl3-3f.spw a ra 
aq Pitting 
s 
b r u 
cs 
c 
c 
c e c
s
s b r 
s 
x
x c
c
c 
x c
c 
10000 1000 
assessment (TSPA-LA) model. 
Source: This report. 
Figure 6.6-2. Chemistry Where Pitting of Zirconium Has Been Observed 
Chloride Ion Concentration [Cl-], M 
10 a a xa a 
a q q a q a r r q 
x
x
a a a a q q q q a a 
1 
No 
Pitting 
0.1 
0.01 
a, b, c, d, e = No Pitting 
q, r, s, t, u = Pitting 
a, q = Yau (CRWMS M&O 2000d [DIRS 136058]) 
b, r = Maquire (1984 [DIRS 101717]) 
c, s = Jangg et al. (1978 [DIRS 110544]) 
d, t = Riley and Covino (1982 [161993]) 
e = Brossia et al. (2002 [DIRS 161988]) 
U = Green et al. (2000 [145073]) 
x = as-polished model and range 0.001 
6.7 LOCALIZED (CREVICE) CORROSION OF CLADDING 
FEP Number: 
2.1.02.17.0A 
YMP FEP Description: 
Localized corrosion in crevices could produce penetrations of cladding. 
Descriptor Phrases: 
Crevice corrosion of cladding 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
ANL-WIS-MD-000008 REV 01 March 2004 36 of 90 Clad Degradation  FEPs Screening Arguments 
Related FEPs: 
2.1.02.13.0A, General Corrosion of Cladding 
2.1.02.14.0A, Microbially influenced corrosion (MIC) of cladding 
2.1.02.15.0A, Localized (radiolysis enhanced) corrosion of cladding 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
Screening Argument: Localized (crevice) corrosion of the cladding is excluded from the 
TSPA-LA on the basis of low consequence. 
Yau and Webster (1987 [DIRS 100494], p. 717) report: Of all the corrosion-resistant structural 
metals, zirconium and tantalum are the most resistant to crevice corrosion. In low-pH chloride 
solutions or chlorine gas, for example, zirconium is not subject to crevice attack. Greene et al. 
(2000 [DIRS 145073]) and Brossia et al. (2002 [DIRS 161988]) performed pitting and crevice 
corrosion tests on Zircaloy-4. They covered temperatures from 25 to 95C, chloride 
concentrations from 0.001 to 4.0 M, and pH from 2.1 to 10.7. The solutions also contained the 
predominant anions in the groundwater. Some of their tests had sufficiently aggressive solutions 
to cause pitting on exposed surfaces. Other tests had voltages applied to the sample to raise the 
corrosion potential above the repassivation potential and cause pitting on exposed surfaces. 
They report that no crevice corrosion is observed under the same environment and 
electrochemical conditions that promote pitting corrosion on exposed surfaces. In summary, 
crevice corrosion is not observed under severe conditions that promote pitting on the exposed 
surfaces. 
Additional corroborating information is also available. More detailed information is provided by 
Yau (1983 [DIRS 149233]) showing that zirconium and Zr-1.5 percent Sn were resistant to 
crevice corrosion after 14 days exposed to boiling (107C), saturated NaCl solution with the pH 
adjusted to 0 by the addition of HCl. Clad Degradation-Local Corrosion of Zirconium and its 
Alloys Under Repository Conditions (CRWMS M&O 2000d [DIRS 136058]) shows that 
zirconium is not susceptible to crevice corrosion. Section 4.1.3 of that report (CRWMS M&O 
2000d [DIRS 136058]) discusses the crevice corrosion resistance of zirconium in various 
chemical solutions, summarizes seven crevice corrosion tests, and reports that crevice corrosion 
was not observed. The U-bend tests discussed in Section 4.1.4 of Clad DegradationLocal 
Corrosion of Zirconium and Its Alloys Under Repository Conditions (CRWMS M&O 2000d 
[DIRS 136058]) are also designed to produce crevice corrosion under the U-bend test washers. 
In these tests, no crevice corrosion was reported. Section 6.1.10 of Clad DegradationLocal 
Corrosion of Zirconium and Its Alloys Under Repository Conditions (CRWMS M&O 2000d 
[DIRS 136058]) discusses the theoretical reasons why zirconium is immune to this type of 
corrosion. 
In conclusion, cladding degradation from localized (crevice) corrosion is excluded from TSPALA. 
Crevice corrosion of zirconium under repository in-package chemistry conditions is not 
expected. NRC requirements in 10 CFR 63.114 (e and f) allow the omission because crevice 
corrosion is not expected under repository conditions and therefore will not significantly change 
the magnitude and time of the resulting radiological exposures to the reasonably maximally 
exposed individual (RMEI), or radionuclide releases to the accessible environment. 
March 2004 37 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
6.8 ENHANCED CORROSION OF CLADDING FROM DISSOLVED SILICA 
FEP Number: 
2.1.02.18.0A 
YMP FEP Description: 
It must be determined if the high dissolved silica content of waters enhances corrosion of 
cladding. 
Descriptor Phrases: 
Enhanced corrosion of cladding from dissolved silica 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.14.0A, Microbially influenced corrosion (MIC) of cladding 
2.1.02.15.0A, Localized (radiolysis enhanced) corrosion of cladding 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
2.1.02.27.0A, Localized (Fluoride enhanced) corrosion of cladding 
Screening Argument: Enhanced corrosion of cladding due to high dissolved silica content in 
waters is excluded from the TSPA-LA on the basis of low consequence. 
Silicon dioxide reacts with hydrofluoric acid to form fluorosilicic acid, which is highly corrosive 
to zirconium in high concentrations (10 weight percent) (Yau and Webster 1987, p. 712). This 
reaction is not expected to occur because the pH is generally too high (pH>3.18) for hydrofluoric 
acid to exist and fluoride concentrations are too low (J13 well water contains only 2.2 ppm) to 
form significant concentrations of fluorosilicic acid. The fluoride corrosion itself is addressed in 
another FEP (2.1.02.27.0A) and is excluded. The potential for silica itself degrading the 
cladding is negligible. Hansson (1984 [DIRS 101676]) reports corrosion tests with concrete pore 
fluids which normally contain silica. Yau (1983 [DIRS 149233]) reports corrosion tests in sea 
water which also contains silica. Neither experimenter reports significant corrosion. Yau and 
Webster (1987 [DIRS 100494], Table 6) review the corrosion potentials for zirconium and report 
no corrosion with sodium silicate concentrations from 0 to 100 weight percent at ambient 
temperature to 100C. 
In conclusion, enhanced cladding degradation from dissolved silica is excluded from TSPA-LA. 
The NRC requirements in 10 CFR 63.114 (e and f) allow the omission because silica will not 
degrade cladding and therefore will not significantly change the magnitude and time of the 
resulting radiological exposures to the reasonably maximally exposed individual (RMEI), or 
radionuclide releases to the accessible environment. 
March 2004 38 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
6.9 CREEP RUPTURE OF CLADDING 
FEP Number: 
2.1.02.19.0A 
YMP FEP Description: 
At high cladding temperatures (>400C) for sufficiently long time intervals, creep rupture 
of Zircaloy cladding on spent fuel can occur and produce small perforations in the 
cladding to relieve stress. After the waste package fails, the fuel can react with water and 
radioisotopes can thereby escape over time from the fuel rod. 
Descriptor Phrases: 
Thermally induced cladding creep (rupture) 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.26.0A, Diffusion-controlled cavity growth (DCCG) in cladding 
Screening Argument: Creep rupture of cladding is excluded from the TSPA-LA on the basis of 
low consequence. 
The Spent Fuel Project Office of the NRC issued Interim Staff Guidance  11, Revision 2 (NRC 
2002 [DIRS 164593], p. 2), which set a maximum temperature limit for the cladding (400C) to 
prevent damage from creep or hydride reorientation during dry storage (similar temperature 
histories as repository closure). Since peak cladding temperatures in the repository are expected 
to be less than 268C (Appendix A), damage from creep is expected to be minimal. Figure 6.9-2 
compares cladding damage for dry storage and repository emplacement as a function of peak 
cladding temperatures, and shows that similar temperature limits apply to both environments and 
the analysis presented in Clad Degradation  Summary and Abstraction (CRWMS M&O 2001a 
[DIRS 151662], Section 6.2). 
Additional corroborating information is available. A statistical analysis of creep failure was 
performed (CRWMS M&O 2001a [DIRS 151662], Section 6.2), in which a burnup distribution 
(rod average = 44 MWd/kgU, range = 2 to 75 MWd/kgU) was used and a distribution of rod 
properties (including stress) was developed. The Murty creep correlation was selected after 
comparing six different correlations with results from five different experiments. It was then 
modified to better predict irradiated cladding creep data. The fuel rods were exposed to two 
consecutive temperature histories before being placed in the repository. They were exposed to 
24 hours of vacuum drying with a peak temperature of 430C, followed by 20 years dry storage 
with a peak temperature of 350C. 
March 2004 39 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
The sensitivity study considered creep strain for only 1,000 years of repository thermal history 
because after that time, the fuel is too cool for creep to occur. Uncertainties in the temperatures 
and strain rate were included. The radial temperature distribution across the waste package was 
also modeled. The waste package surface temperature was shifted upward or downward and the 
fraction of rods that fail from creep was calculated. Failure was predicted when the creep strain 
of a rod exceeds a creep failure criterion. The upper limit was a 1 percent strain failure criterion 
and lower limit was 6 percent strain. The best estimate creep failure criterion was a 
complementary cumulative distribution function (CCDF) based on 52 failure tests. Figure 6.9-1 
shows the results of this sensitivity study. Rod failures started to occur during repository closure 
at a peak cladding temperature of 400C. Cladding failure reached 1 percent at a peak 
temperature of about 430C (best estimate). In the current repository design, the cladding 
temperatures are below 268C (Table A-2) and rod failures from creep are not expected. 
The Dry Storage Characterization Project (EPRI 2002 [DIRS 161421], p. xii) studied the 
condition of fuel assemblies that were exposed to various thermal transients (peak temperatures 
to 415C) followed by 15 years of dry storage (peak temperature at 342C and slowly 
decreasing). These temperatures are higher than expected at the repository (Appendix A of this 
report). The Project concluded little or no cladding creep occurred during the thermal 
benchmark testing and dry storage. It is anticipated that the creep would not increase 
significantly during additional storage due to the low temperature after 15 years, the continual 
decrease in temperature from the reduction of decay heat, and the concurrent reduction in 
pressure and stress. 
Over time, helium production from alpha decay could generate pressures that exceed the pressure 
at the peak temperature (see FEP 2.1.02.20.0A). The creep rate has a strong temperature 
dependency (Figures 6.9-1 and 6.9-2) and the slightly higher stress that occurs at 10,000 years 
does not change the creep results because the temperatures are less than 73C and no observable 
creep occurs at these temperatures and strains. 
In conclusion, creep rupture of cladding is excluded from the TSPA-LA on the basis of low 
consequence. Repository temperatures are too low to have any significant creep. The NRC 
requirements in 10 CFR 63.114 (e and f) allow the omission because the small amount of creep 
that will occur during the regulatory period will not significantly change the magnitude and time 
of the resulting radiological exposures to the reasonably maximally exposed individual (RMEI), 
or radionuclide releases to the accessible environment. 
March 2004 40 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
0.10 
0.08 
0.06 
0.04 
0.02 
0.00 
300 250 
Source: CRWMS M&O 2001a [DIRS 151662] 
Figure 6.9-1. Creep Failure Fraction as a Function of Peak Cladding Temperature 
0.020 
CP-vdr.spw 
0.015 Repository 
(Peak VD=430C, 
Peak DS=350C) 
0.010 
0.005 
0.000 
300 
Peak Cladding Temperature at Each Phase 
Source: CRWMS M&O 2001a [DIRS 151662] 
Figure 6.9-2. Comparison of Creep Damage in Dry Storage and Repository Emplacement 
Fraction of Rods with Creep Failures Fraction of Rods with Creep Failure 
ANL-WIS-MD-000008 REV 01 
Upper Limit 
Best Estimate 
Lower Limit 
400 350 
Peak Cladding Temperature,oC 
Dry Storage 
(DS) 
(Peak VD 
= 430C) 
600 500 400
41 of 90 
creep-mc.spw 
500 450 
Vacuum 
Drying (VD) 
700 
March 2004 Clad Degradation  FEPs Screening Arguments 
6.10 INTERNAL PRESSURIZATION OF CLADDING 
FEP Number: 
2.1.02.20.0A 
YMP FEP Description: 
Increased pressure within the fuel rod due to the production of helium gas could 
contribute to cladding failure. 
Descriptor Phrases: 
Internal gas pressure from He (cladding damage) 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.19.0A, Creep Rupture of Cladding 
2.1.02.21.0A, Stress corrosion cracking (SCC) of cladding 
2.1.02.22.0A, Hydride cracking of cladding 
2.1.12.02.0A, Gas generation (He) from waste form decay 
Screening Argument: Internal pressurization of the cladding is excluded from the TSPA-LA on 
the basis of low consequence. 
Piron and Pelletier (2001 [DIRS 165318], Section 5.3) investigated the pressurization of the fuel 
rods from helium production (alpha decay). They concluded that fuel (47.5 MWd/kgU) would 
produce 1,171 cm3 (STP) of helium in a rod after 10,000 years. Piron and Pelletiers (2001 
[DIRS 165318], Section 5.3) values are based on having all of the helium released. Piron and 
Pelletiers values were adjusted for burnup (36 MWd/kgU) and temperature to be consistent with 
the earlier analysis reported in Initial Cladding Condition (CRWMS M&O 2000c [DIRS 
151659], Section 6.3). The resulting values are presented in Table 6.10-1. Their analysis 
produced a total rod pressure 30 to 50 percent higher than predicted in Initial Cladding 
Condition (CRWMS M&O 2000c [DIRS 151659], Section 6.3.4). Also given in this table are 
the pressures for the average burnup rod (44.1 MWd/kgU). This is shown in the last three lines 
of Table 6.10-1. The peak pressure (13.3 MPa) would have to be significantly higher (about 
33 MPa to produce the necessary stress intensity for crack propagation) for the cladding to fail 
from delayed hydride cracking (FEP 2.1.02.22.0A, see Section 6.12.3). Even when using values 
provided by Piron and Pelletier (2001 [DIRS 165318]), the change in pressure is not significant 
and no cladding failure from helium production is expected. 
March 2004 42 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 6.10-1. Effect of Helium Production on Rod Pressure 
Total Helium He % of total Fission Gas 
Time Temperature Pressure Pressure Pressure Pressure 
(years) (MPa)a (C) (MPa) (MPa) (%) 
0 135 0.03 5.4 5.5 0.46 
10 268 0.13 7.2 7.3 1.8 
90 212 0.52 6.5 7.0 7.4 
950 131 1.9 5.4 7.3 26.4 
3.7 4.6 8.3 44.3 10,000 73 
10,000 73 6.3 4.6 10.9 57.8 
7.8 5.5 13.3 58.5 10,000 73 
Source: CRWMS M&O 2000c [DIRS 151659], Table 7 
NOTES: a Initial fill and fission gas pressure was 4 MPa at 27C. 
b Helium pressure based on Piron and Pelletier 2001 [DIRS 165318] 
c Helium pressure based on Piron and Pelletier 2001 [DIRS 165318], average burnup 
of 44.1 MWd/kgU. 
d Mean value of fill and fission gas pressure from CRWMS M&O 2000c [DIRS 
151659], Table 12 
Additional corroborating information is available. Initial Cladding Condition (CRWMS M&O 
2000c [DIRS 151659], Section 6.3.4) provides the initial screening for the effect of helium 
production. This analysis could be considered an alternative conceptual model to Piron and 
Pelletier. The internal gas in a fuel rod consists of initial fill gas, fission product gases, and 
helium gas from alpha decay. The helium gas pressure will slowly increase over time by the 
production of helium as a result of alpha decay. Manaktala (1993 [DIRS 101719], Figure 3-4, 
p. 3-12) presents the helium pressure buildup for 100C as a function of time for a PWR fuel rod 
with 36 MWd/kgU burnup and a conservative analysis that 100 percent helium is released from 
the fuel into the fuel rod gap. This figure was used to develop an equation for helium buildup in 
a fuel rod. Table 6.10-1 provides the pressure change as a function of time for a typical rod (36 
MWd/kgU). The pressures have been adjusted to reflect the current maximum temperatures of 
268C (peak) and 73C (10,000 years) using the ideal gas law. Ambient (27C) pressure was 
increased by 1.80 (541 K/300 K) for the peak temperature and 1.15 (346 K/300 K) for 73C 
temperature. The table shows that there are two competing effects. Over time, the rod cools, 
decreasing the pressure and the helium production increases the pressure. After 10,000 years, 
the pressure is doubled from that at the time of peak temperature. This increase is not enough to 
change any failure mechanisms that were evaluated at the peak temperature. Cladding failure 
from pressure, stress, or stress intensity factors is not expected. Creep rupture (FEP 
2.1.02.19.0A) is also driven by stress and is not in expected to contribute to cladding failure. 
This is discussed further in FEPs 2.1.02.19.0A, 2.1.02.21.0A, and 2.1.02.22.0A where the 
individual failure modes are discussed. 
Other corroborating information is available. Rothman (1984 [DIRS 100417], Table 6) predicted 
that helium pressure buildup would be offset by the cooling of the rods and the pressure at 
10,000 years would be slightly less that at the peak temperature of 322C. He also considered 
that 100 percent of the gas was released. Both Rothman and Piron and Pelletier concluded that 
the pressurization from helium after 10,000 years is not significant to cause rod failure. 
43 of 90 ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation  FEPs Screening Arguments 
Although these results are based on complete release of the helium, this is actually unlikely to 
occur. Piron and Pelletier (2001 [DIRS 165318], p. 232) point out that UO2 can hold 0.66 cm3 
(STP) of helium per gram of fuel. A rod weighs approximately 2,200 gm (201 cm3 (CRWMS 
M&O 2000c [DIRS 151659], Table 2)  10.97 gm/cm3 (BSC 2003a [DIRS 162153], Table 8)). 
The fuel in a rod can contain 1,500 cm3 (STP) of helium, larger than the 1,171 cm3 (STP) 
expected to be formed. Therefore, even if microscopic bubbles of gas are formed, total gas 
release is not expected. M. Peehs (1998 [DIRS 109219], p. 5) shows that fission gas release 
during dry storage is not expected because the diffusion coefficients decrease by approximately 8 
orders of magnitude for dry storage temperatures (hotter than repository temperatures) compared 
to reactor operation. While helium diffusion is greater than fission gas diffusion, complete 
release is both conservative and unlikely. 
In conclusion, cladding degradation from internal pressurization of the cladding is excluded from 
TSPA-LA. The NRC requirements in 10 CFR 63.114 (e and f) allow the omission because 
internal pressurization of the fuel rods is too low to cause damage to the cladding and therefore 
will not significantly change the magnitude and time of the resulting radiological exposures to 
the reasonably maximally exposed individual (RMEI), or radionuclide releases to the accessible 
environment. 
6.11 STRESS CORROSION CRACKING (SCC) OF CLADDING 
FEP Number: 
2.1.02.21.0A 
YMP FEP Description: 
Stress corrosion cracking mechanisms can contribute to cladding failure. These 
mechanisms can operate both from the inside out from the action of fission products, or 
from the outside in from the actions of salts or other chemicals within the waste package. 
Descriptor Phrases: 
Cladding high stress locations (external stress corrosion crackings) 
Cladding radiolysis (internal stress corrosion cracking) 
Cladding threshold stress (stress corrosion cracking) 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
2.1.02.20.0A, Internal Pressurization of Cladding 
2.1.09.09.0A, Electrochemical effects in EBS 
Screening Argument: Stress corrosion cracking (SCC) of cladding is excluded from the TSPALA 
on the basis of low consequence. 
March 2004 44 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
SCC is the formation of brittle cracks on a metal surface through the simultaneous action of a 
tensile stress and a corrosive environment. SCC requires a certain chemical environment and 
sufficiently high stresses (stress intensity factors at the crack tip). Chloride-induced SCC could 
occur on the outside of the cladding. Chloride induced stress corrosion cracking requires that the 
passive layer of oxides on the zirconium surface be unstable (Cragnolino et al. 1999 [DIRS 
152354], p. 4-15; Cox 1973 [DIRS 152920], Abstract; Farina et al. 2002 [DIRS 163639], p. 5; 
Yau and Webster 1987 [DIRS 100494], p. 718). These are the same conditions under which 
pitting occurs. As demonstrated in FEP 2.1.02.16.0A, the chemical conditions for pitting and 
SCC do not exist in the waste package. Therefore, chloride-induced SCC is not expected. 
Even if the chemical environment existed, stresses and stress intensities are too low for SCC to 
occur. Table 6.11-1 provides the pressure, stresses and stress intensity factors for cladding at 
three different conditions. The first column contains ambient temperature conditions at closure, 
from Initial Cladding Condition (CRWMS M&O 2000c [DIRS 151659], Sections 6.3.6, 6.3.7, 
and 6.10.2). They represent a statistical distribution for PWR fuel with burnups ranging to 75 
MWd/ kgU. Only upper values of the distributions are provided because these are of interest in 
this type of failure. Column 2 contains the values at the time of the peak temperature (10 years 
after closure, 268C for rods located in the center of the waste package), and is calculated from 
column 1 using the ideal gas law. Column 3 represents the pressure, stress and stress intensity 
factor at the end of the 10,000-year regulatory period. At this time, the maximum center rod 
temperature has decreased to 73C (Table A-3). This has been offset by the buildup of helium 
pressure from alpha decay (FEP 2.1.02.20.0A). Table 6.10-1 (last row) concludes that if all of 
the helium were released from the fuel matrix, the helium would contribute 7.7 MPa of pressure 
to the initial pressure distribution for the initial fill gas and released fission gas. Stresses and 
stress intensity factors are then estimated. 
During the times of interest (peak temperature or peak helium pressure) the maximum stress 
intensity factors vary from 0.47 mean value to 2.73 MPa-m0.5. This is less than the threshold 
stress intensity factor for SCC, the value at which crack propagation will start. Table 10b of 
Clad Degradation  Summary and Abstraction (CRWMS M&O 2001a [DIRS 151662]) provides 
the threshold values for various chemical solutions. Table values are greater than or equal to 
4 MPa-m0.5. The threshold stress intensity factor (KISCC) of 28 MPa-m0.5 for moist chlorine is 
included in this table and pertinent to the case of external cracking (Cox 1990 [DIRS 152778], 
Figure 20, p. 15). This value is for 70oC. The threshold stress intensity decreases with 
increasing temperature (Cox 1990 [DIRS 152778], Figure 14). At a boiling point of 100oC, 
when water, possibly carrying chlorides, could enter the WP, the threshold stress intensity could 
be about 24 MPa-m0.5. This value is still well above values predicted in Table 6.11-1. Since the 
estimated stress intensities are less than the threshold values, chloride-induced cracking is not 
expected. 
March 2004 45 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table 6.11-1. Pressures, Stresses and Stress Intensity Factors for Fuel Rods During Repository 
Emplacement 
ANL-WIS-MD-000008 REV 01 
Ambient Properties at Properties at Peak 
Peak Temperature, Temperature 
Properties Temperature 
Time, years 0 10 
Temperature, C 
Pressure 
Mean, MPa 4.8 8.7 
95% MPa 7.3 13 
98%, MPa 8.9 16 
Max, MPa 18 32 
Stress 
Mean, MPa 38 69 
95%, MPa 62 111 
98%, MPa 76 136 
Max, MPa 146 263 
Stress Intensitya 
Mean, MPa-m0.5 0.26 0.47 
95%, MPa-m0.5 0.61 1.09 
98%, MPa-m0.5 0.77 1.38 
Max, MPa-m0.5 1.52 2.73 
Source: CRWMS M&O 2000c [DIRS 151659] 
NOTE: A For a sharp crack, (Reed-Hill 1973 [DIRS 121838], p. 800) 
Stress corrosion cracking on the interior surface of the cladding from iodine is not expected. In 
NRC Issue Resolution Status Report Key Technical Issue: Container Life and Source Term 
(Beckman 2001 [DIRS 156122], p. 103), the NRC concluded: 
The possibility of SCC induced by iodine as discussed in the WF PMR [Process 
Model Report] does not appear so important because it is limited essentially by 
the availability of iodine. The phenomenon as such has been postulated as the 
cause of pellet cladding interaction failure under reactor operating conditions 
following steep power ramps, but it does not seem plausible under repository 
conditions. 
Iodine concentrations are too low for SCC (Beckman 2001 [DIRS 156122], p. 103). Failure has 
not been observed in dry storage tests with similar conditions to early repository closure. Stress 
intensities for a sharp crack, (a limiting case, Reed-Hill 1973 [DIRS 121838], p. 800)(up to 2.73 
MPa-m0.5) are lower than the range of threshold stress intensity factors (4.0  15 MPa-m0.5) 
identified for iodine induced SCC (Tasooji et al. 1984 [DIRS 102093], Figure 12, p. 612). 
Cragnolino et al. (1999 [DIRS 152354], p. 4-27) suggested that lower threshold stress intensity 
factors could cause failure when considering small cracks. The Dry Storage Characterization 
Project (EPRI 2002 [DIRS 161421]) studied the condition of fuel assemblies that were exposed 
to various thermal transients (peak temperatures to 415C) followed by 15 years of dry storage 
(peak temperature at 342C and slowly decreasing). These temperatures are higher than 
10,000 
27 268 73 
13 
16 
18 
28 
107 
137 
154 
233 
0.73 
1.34 
1.56 
2.42 
46 of 90 
Regulatory Limit 
March 2004 Clad Degradation  FEPs Screening Arguments 
expected at the repository (Appendix A of this report). Rod failure during dry storage was not 
observed, suggesting that iodine induced SCC from small cracks did not fail rods. 
In conclusion, the magnitude and time of the resulting radiological exposures to the reasonably 
maximally exposed individual (RMEI), or radionuclide releases to the accessible environment, 
would not be significantly changed by the omission of this FEP (stress corrosion cracking (SCC) 
of cladding) from the performance assessment (TSPA-LA) model. 
6.12 HYDRIDE CRACKING OF CLADDING 
FEP Number: 
2.1.02.22.0A 
YMP FEP Description: 
2 
Cladding contains hydrogen after reactor operation. The cladding might also pick up 
more hydrogen from cladding general corrosion (wet oxidation) after the waste package 
is breached. The hydrogen can exist both as zirconium hydride precipitates and as 
hydrogen in solid solution with zirconium. Hydrides might also form from UO 
oxidation (after waste package and cladding perforation). In addition, hydrides may 
dissolve in warmer areas of the cladding and migrate to cooler areas. Hydrogen can also 
move from places of low stress to places of high stresses, causing hydride reorientation or 
delayed hydride cracking (DHC). The buildup of hydrides can cause existing cracks to 
propagate by DHC or hydride embrittlement. 
Descriptor Phrases: 
Delayed hydride cracking (DHC) (cladding) 
Galvanic coupling (cladding  carbon steel) 
Hydride cracking (cladding) 
Hydride embrittlement (cladding) 
Hydride reorientation in cladding 
Hydrogen induced cracking (HIC) cladding 
Zirconium oxidation of cladding 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.13.0A, General Corrosion of Cladding 
2.1.02.20.0A, Internal Pressurization of Cladding 
Screening Argument: Hydride cracking and embrittlement of the cladding is excluded from the 
TSPA-LA on the basis of low consequence. 
The stresses in the cladding are not sufficient to fail the cladding at the repository temperatures 
and experimental data indicate that the in-package environment and cladding stresses are not 
March 2004 47 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
conducive to hydride cracking and embrittlement. A more detailed discussion is provided 
elsewhere (Sections 6.12.1 through 6.12.6 of this report) 
As the waste package internals corrode, hydrogen is generated, although little is expected to be 
absorbed directly by the fuel cladding, because H2 molecules do not migrate through the highdensity 
ZrO2 fuel cladding layer (FEP 2.1.12.03.0A). Available data on zirconium hydriding 
indicate that corrosion of waste package internals will not result in hydriding of fuel cladding, as 
long as an oxidizing environment exists in the waste package (IAEA 1998 [DIRS 150560], 
p. 92). 
Cladding surface oxidation is minor at repository temperatures and hydrogen absorption will be 
negligible. Hydride embrittlement from galvanic corrosion of waste package contacting cladding 
has been excluded based on low consequence. Cladding has a thick, electrically insulating, 
oxide layer that is produced during reactor operation. This film prevents both direct absorption 
of hydrogen gas in the environment and galvanic coupling to dissimilar metals. If the passive 
film has been mechanically removed, the unprotected cladding oxidizes within seconds and 
forms a passive layer if exposed to water or humid air. Therefore, cladding would undergo little 
hydrogen charging because the oxide layer prevents hydrogen absorption in the metal (FEP 
2.1.12.03.0A). 
Cladding failure by DHC is unlikely and has not been included in the abstraction for the TSPALA. 
Stresses (and stress intensity factors) are too low for crack propagation (see Section 6.12.3 
of this report). 
Cladding failure by hydride reorientation is unlikely (see Section 6.12.4 of this report), because 
the maximum temperatures are too low to dissolve much hydrogen and most rods have stresses 
too low for reorientation. The cladding material will maintain sufficient strength, even if hydride 
reorientation occurs, such that failure would not be expected. 
Hydrogen axial migration will be limited at the temperatures expected during emplacement 
(268C maximum). Failure of the cladding by hydrogen embrittlement is unlikely. Hydrogen 
absorption in the cladding from UO2 fuel corrosion only occurs in fuel with already failed 
cladding. Such reaction, if it should occur, has little consequence (see Section 6.12.5 of this 
report). 
Hydrogen embrittlement results in a generally reduced resistance to fracture. In Zircaloy, 
hydrogen embrittlement is normally caused by precipitation of zirconium hydride. Since the 
hydride precipitates are quite brittle, a crack can propagate more readily by preferentially 
following the hydrides. Resistance to fracture (fracture toughness KIC) is a measure of resistance 
to crack propagation through the material. Fracture toughness is typically measured in terms of 
the critical stress intensity factor, that is, the stress intensity factor value that will cause growth of 
a crack. The stress intensity factor is proportional to the far-field stress times the square root of 
the crack length. Kreyns et al. (1996 [DIRS 100462], Figure 5, reproduced here as Figure 6.12- 
1) show that for both irradiated and unirradiated material, such hydrides could decrease the 
fracture toughness (KIC) from 42 MPam 0.5 to 8 MPam0.5 as the hydrogen content increases 
from zero to 4,000 ppm. As shown in Table 6.11-1 the maximum stress intensity (KI) for the 
statistical distribution of rods and crack sizes varies from 0.47 to 2.73 MPa m0.5 and therefore, 
March 2004 48 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
failure is not expected, even with hydride concentrations of 4,000 ppm. In the limit (100 percent 
hydride and no metal), the fracture toughness is about 1 MPam0.5. The outer surface of the 
cladding could be fairly brittle (hydrogen content greater than 800 ppm) but much of the 
cladding thickness has a reasonable toughness. 
In conclusion, hydride cracking and embrittlement of the cladding is excluded from the TSPALA 
on the basis of low consequence. The NRC requirements in 10 CFR 63.114 (e and f) allow 
this omission because the small amount of hydride cracking and embrittlement that will occur 
during the regulatory period will not significantly change the magnitude and time of the 
radiological exposures to the reasonably maximally exposed individual (RMEI), or radionuclide 
releases to the accessible environment. The stresses in the cladding are not sufficient to fail the 
cladding at the repository temperatures and experimental data indicate that the in-package 
environment and cladding stresses are not conducive to hydride cracking and embrittlement. 
Source: Kreyns et al. 1996 [DIRS 100462] 
Figure 6.12-1. Fracture Toughness vs. Hydrogen Content of Zircaloy-4 
49 of 90 ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation  FEPs Screening Arguments 
6.12.1 Hydride Embrittlement from Zirconium Corrosion (of Cladding) 
The hottest rods will pick up only small quantities of hydrogen from cladding surface general 
corrosion after the waste package fails. Waste packages are failed at various times, permitting 
water or moist air to enter. Wet oxidation (or steam oxidation if the local temperature is above 
boiling) occurs, and approximately 17 percent of the hydrogen released from the water is 
absorbed by the fuel cladding in the failed waste package (Lanning et al. 1997 [DIRS 101704], 
Volume 1, p. 8.4, Figure 8.2). However, the fuel cladding picks up little hydrogen because the 
corrosion rate is so slow (see FEP 2.1.02.13.0A). Even the hottest rods will pick up only small 
quantities of hydrogen from cladding surface general corrosion. 
6.12.2 Hydride Embrittlement from Galvanic Corrosion of Waste Package Contacting 
Cladding 
Hydride embrittlement from galvanic corrosion of waste package contacting cladding has been 
excluded based on low consequence. Cladding has a thick oxide layer that is produced during 
reactor operation. The oxide coating is electrically insulating and will prevent electric contact of 
dissimilar metals. It is kinetically noble (Yau and Webster 1987 [DIRS 100494], Table 15, p. 
718) and, for zirconium, the most noted effect is accelerated corrosion of the other material. If 
passive film has been mechanically removed the unprotected cladding oxidizes and forms a 
passive layer within seconds if exposed to water or humid air (Hansson 1984 [DIRS 101676], p. 
6). Inside the failed waste package, the cladding would behave as a cathode and the steel box 
would be an anode if they were electronically coupled (Yau and Webster 1987 [DIRS 100494], 
Table 15, p. 718). The cladding would undergo little hydrogen charging because the oxide layer 
prevents hydrogen absorption in the metal (IAEA 1998 [DIRS 150560], p. 92, see FEP 
2.1.12.03.0A). The consequence of hydrogen charging would be to hydride the zirconium. 
Direct absorption of hydrogen gas in the environment into zirconium has occurred with clean 
zirconium metals that have nickel coatings to prevent oxidation of the surface. It is unlikely to 
occur with irradiated cladding, which has heavy oxide coatings. 
Other corroborating information is available. The importance of galvanic hydrogen conditions 
can be inferred from results in Corrosion Properties of Zirconium in Chloride Solutions (Yau 
1983 [DIRS 149233], p. 26/10). Yau discusses a series of experiments in which zirconium alloy 
U-bend samples were exposed to boiling seawater for 365 days. Each sample was loaded by an 
uninsulated steel coupling. Three compositions (Zr 702, Zr 704 with up to 1.3 percent nickel, 
and Zr 704 without nickel) were considered and tested in both unwelded and welded conditions. 
Of the six samples, only the welded sample of Zr 704 with nickel showed hydrogen pickup. 
Clayton (1984 [DIRS 131741]) also observed the role of nickel in hydrogen absorption. The 
composition limit for these materials is the maximum hydrogen content of 50 ppm. Except for 
the welded sample of Zr 704 with nickel, the U-bend samples had hydrogen contents of 5 to 9 
ppm at the end of the test. It is apparent that contact with a steel surface alone does not lead to 
galvanic hydrogen charging. Other factors, such as relative surface areas of the coupled anode 
and cathode, also affect the degree of radiolysis. 
It is also important to compare the amount of nickel in various zirconium alloys. Zr 704 contains 
up to 1.3 percent nickel, Zr 702 contains up to 0.8 percent nickel, Zircaloy-2 contains up to 
0.08 percent nickel, and Zircaloy-4 contains no more than 0.007 percent nickel. It is clear that 
March 2004 50 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
the Zircaloys contain small amounts of nickel, so hydriding as a result of contact with the steel 
basket tubes is not expected. It might be argued that, because of its nickel content, Zircaloy-2 
cladding could be hydrided. However, Zircaloy-2 fuel cladding is used only for boiling water 
reactors, and most of the fuel cladding will be separated from the basket tubes by the fuel 
channels. Electrical coupling does not require physical contact, but the physical separation 
reduces the risk of galvanic corrosion. However, physical contact is required when insulating 
surface films (e.g., oxide coating) exist. Since the assemblies are loosely placed in the waste 
package and coated with thick oxides, they will not undergo hydrogen absorption. Therefore, 
this FEP is screened out as low consequence. 
Hydriding of zirconium alloys as a result of contact with dissimilar materials has been observed, 
so it is important to consider the conditions that promote hydriding. In this case, the dissimilar 
metal forms a hydrogen window, not galvanic coupling. Clayton (1984 [DIRS 131741], p. 578) 
lists a series of experiments in which Zircaloy-4 fasteners clamped onto a sample of nickel-base 
alloy were hydrided by exposure to hot water with dissolved hydrogen. Clayton (1984 [DIRS 
131741], p. 573) discusses the mechanism: 
. . . in mechanically attaching a Zircaloy fuel rod fastener to an Inconel support 
plate, relative motion occurred during assembly. The relative motion was 
sometimes sufficiently severe and the bearing stress sufficiently high to overcome 
the protective effects of the corrosion oxide film and graphite lubricant. Smearing 
and bonding of Inconel onto local regions of the mating Zircaloy contact surface 
occurred and provided the potential for accelerated hydriding. Inconel has a 
relatively high permeability for hydrogen and acts as a window for hydrogen 
entry into the Zircaloy. 
A similar but localized effect was noted for Zircaloy-4 smeared with Inconel alloy (75.2 percent 
nickel, 16.7 percent chromium, 3.5 percent copper, and 4.6 percent iron). 
Although hydriding was observed in the experiments discussed above, it does not follow that 
hydriding will occur under repository conditions. As is noted in Clayton (1984 [DIRS 131741], 
p. 587), Nickel alloy smearing and bonding to filmed Zircaloy, rather than just tight surface 
contact, is necessary for accelerated hydriding. Since fuel assemblies will be lowered carefully 
into waste package baskets, high contact pressures typical of fastener tightening are not 
expected. After emplacement, the fuel assembly will simply rest on the basket. Therefore, 
smearing and bonding should not occur. Few, if any, rods might undergo hydride embrittlement 
from galvanic coupling and, therefore, this FEP is being screened out as low consequence. 
However, if it is supposed that there is some bonding between the fuel assembly and the basket, 
the significance of the effect must be examined. It should be noted that accelerated hydriding is 
a transient, not a persistent, effect. Clayton (1984 [DIRS 131741], p. 572) notes that the initial 
high accelerated hydrogen ingress rate was effectively shut off during exposure at 271C in 
about 25 days. Because of the extremely small size of contact spots, corrosion of the carbon 
steel rack would quickly break the electrical contact between the two materials. Since the time 
of contact will be short, the amount of hydrogen absorbed as a result of contact will also be small 
(low consequence). This conclusion is supported by the discussion of U-bend hydriding above. 
March 2004 51 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Ralph et al. (2002 [DIRS 161992]) report an example of intimate contact between cladding and 
steel (due to corrosion of the fuel racks) in a spent fuel pool at a U.S. commercial nuclear plant. 
After an extended lay-up period, the spent fuel pool water was found to contain a significant 
amount of algae and bacteria. During the period when the water chemistry in the spent fuel pool 
was not maintained, this water still could be a poorer electrolyte than water seeping into the 
failed repository waste package. Biological agents were purged using controlled additions of 
chlorine and hydrogen peroxide before the pool was returned to normal operating chemistries. 
The assessment revealed that the steel rack corrosion products were up to 2.5 cm thick, and the 
corrosion products had started to engulf the individual fuel rods or flow channels of the stored 
assemblies in the region where the rack and plates contacted the fuel assemblies. The corrosion 
product had adhered to the fuel. The iron oxide was composed of FeO, Fe2O3, and Fe3O4. Ralph 
et al. (2002 [DIRS 161992], p. 6) states: 
One fuel assembly was removed from its storage location and its channel was 
removed. The oxide from contact with the carbon steel rack was removed with a 
water lance utilizing 350 to 700 kg/cm2 of water pressure. A camera with 
resolution of 0.025 mm was used to inspect the channel. The channel surface 
appeared uniform and smooth. No pitting, white discoloration or surface 
anomalies were observed. 
The paper concluded: The fuel cladding was not affected through any type of corrosion. 
Therefore, the corrosion did not change the classification of the fuel as intact or damaged. 
While it is not likely that galvanic corrosion occurred, if it had, the passivated zirconium alloy 
would have served as the cathode. Thus, a lack of zirconium corrosion does not demonstrate that 
galvanic corrosion did not occur. A lack of zirconium hydriding, however, would. 
6.12.3 Delayed Hydride Cracking (of Cladding) 
Failure of the cladding by DHC has been excluded from the TSPA-LA analysis because few, if 
any, fuel rods will fail from DHC (low consequence, would not have a significant effect on the 
resulting radionuclide exposures to the RMEI). The stresses (and stress intensity factors, KI) are 
lower than the threshold stresses and stress intensity factors (KIH) and therefore, few, if any, 
cracks will propagate and fuel rod failure is not expected. 
Peehs (1998 [DIRS 109219], pp. 5, 6) concluded that DHC would not occur in dry storage. The 
thermal history for dry storage is similar to the earlier period of repository closure, except the 
repository is cooler (i.e., a maximum temperature of 268C in [Appendix A of this report] vs. 
400C [NRC 2002 [DIRS 164593]]). Cragnolino et al. (1999 [DIRS 152354], p.4-21) reviewed 
the stress intensity factors and threshold values and also concluded that DHC is not relevant 
under repository conditions. 
Additional corroborating information follows. Rothman (1984 [DIRS 100417], pp. 33 to 39) 
reviewed DHC in Zircaloy cladding in a repository. He concludes that DHC is unlikely unless 
the fuel rods have large existing cracks (exceeding approximately 50 percent of wall thickness) 
and high stresses (exceeding approximately 137 MPa). Reviews of DHC are available in 
Section 6.3 of Hydride-Related Degradation of SNF Cladding Under Repository Conditions 
(CRWMS M&O 2000f [DIRS 148251]) and Initial Cladding Condition (CRWMS M&O 2000c 
March 2004 52 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation C FEPs Screening Arguments 
[DIRS 151659], Section 6.10.2). During delayed hydride cracking, hydrides slowly form at a 
crack until it propagates through the hydride region at the crack tip and stops. This sequence 
repeats itself as the crack propagates slowly through the metal. The hydrides preferentially 
collect at the crack tip. The hydrogen diffuses to a stressed crack tip because the locally high 
stress reduces the hydrogen chemical potential. The higher tensile stress also reduces the 
solubility in that region. The critical stress intensity factor (KIH) is the minimum stress intensity 
that will permit any DHC, regardless of velocity (velocity approaches 0). For this analysis 
(CRWMS M&O 2000d [DIRS 136058]), the stress intensities (KI) were calculated and compared 
to the threshold stress intensity factor KIH. If KI > KIH, then the crack will start to propagate and, 
because of long repository times, failure will occur. The threshold stress intensity for DHC is the 
minimum stress intensity where crack propagation is possible (at a slow velocity). DHC failure 
occurred in some zirconium coolant tubes in a CANDU reactor where high temperature gradients 
caused excess hydride buildup in a specific location (Cox 1990 [DIRS 152778], p. 2). 
The stress intensity factor, KI, is a measure of the increased stress at the tip of a crack. The stress 
intensity factor is proportional to the far-field stress times the square root of the crack depth. For 
a sharp crack, a limiting case, the stress intensity factor is (Reed-Hill 1973 [DIRS 121838], p. 
800): 
w K = (Eq. 6.12-1) I 
where 
KI = Stress intensity factor, MPa.m0.5 
St = Cladding stress, MPa 
w = Crack depth, m 
St   
Equation 6.12-1 is slightly modified from the form given by Reed-Hill; the crack depth w is used 
in place of c/2, where c is the crack length (Dieter 1961 [DIRS 147973], p. 194). 
DHC is analyzed in Section 6.10.2 of Initial Cladding Condition (CRWMS M&O 2000c [DIRS 
151659]). The calculated crack size distribution is given in Figure 17 of Initial Cladding 
Condition (CRWMS M&O 2000c [DIRS 151659], Section 6.6). The median (P = 50 percent) 
value is 13 m and the average crack is 18.6 m. The largest size crack of the 2000 samples is 
119 m. The calculated stress distribution (CRWMS M&O 2000c [DIRS 151659], Section 6.7) 
is given in Figures 18 and 26 of that report. DHC is unlikely at temperatures above 260C 
(Mahmood et al. 2000 [DIRS 152241], p. 139), because of the plasticity of the material. 
Rothman (1984 [DIRS 100417], p. 37) reports that DHC is unlikely above 250C because of the 
plasticity of the material. For this calculation, the temperature of 268C (maximum cladding 
temperature, see Table A-2) is used and the pressure is adjusted accordingly. The peak stress 
intensity factor, KI, was calculated for both peak temperature and temperature after 10,000 years 
of alpha decay and ranges in value from 0.47 to 2.73 MPa-m0.5 (Table 6.11-1). 
The work of Shi and Puls (1994 [DIRS 102084], p. 239, Fig. 7), shows experimental KIH in the 
range of 5 to 12 MPa-m0.5 for zirconium alloy containing 2.5 percent Nb. Rothman (1984 [DIRS 
100417], p. 37), reports a KIH of 6 MPa-m0.5 for Zircaloy-2. Pescatore et al. (1990 [DIRS 
101230], Table 6, p. 50) report values of 5 and 14. Huang (1995 [DIRS 101683], p. 195) shows 
KIH for irradiated Zircaloy-2 approached 6 MPa-m0.5. For this report, Huangs and Rothmans 
53 of 90 March 2004 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
value for irradiated cladding of 6 MPa-m0.5 was used. Because the KI values are well below the 
threshold values, exactly which KIH value is selected is not important. No KI values in this 
report's sampling (Table 6.11-1) of rods are near the threshold stress intensity value. The 
maximum KI was 2.73 MPa-m0.5 and the mean value was 0.73 MPa-m0.5. 
Cragnolino et al. (1999 [DIRS 152354], p. 4-27) suggested that, when considering small cracks, 
lower threshold stress intensities (KIH) could cause failure. The Dry Storage Characterization 
Project (EPRI 2002 [DIRS 161421]) studied the condition of fuel assemblies that were exposed 
to various thermal transients (peak temperatures to 415C) followed by 15 years of dry storage 
(peak temperature at 342C and slowly decreasing). These temperatures are higher than 
expected at the repository (Appendix A of this report). Rod failure during dry storage was not 
observed suggesting that delayed hydride cracking from small cracks did not fail rods, even in a 
15 year cooldown. 
6.12.4 Hydride Reorientation (of Cladding) 
Hydride reorientation has been excluded on the basis of low consequence since few if any rods 
will undergo reorientation and reoriented rods might not fail. Because of the low peak 
temperatures (less than 270C for center rod), few rods will undergo reorientation. Hydride 
reorientation does not necessarily indicate cladding failure. Puls (1988 [DIRS 102067]) 
measurements showed significant strength after reorientation. Dry storage tests (EPRI 2002 
[DIRS 161421]), which exposed irradiated fuel rods to numerous thermal cycles, showed little or 
no reorientation. 
Cragnolino et al. (1999 [DIRS 152354], p. 4-22) concluded that reorientation is unlikely if 
stresses are below 100 MPa and the maximum temperatures are below the hydrogen solvus 
temperature (290-300C). Table 6.11-1 provides the stress distribution value for the hottest rods 
(center location of the hottest waste package). Only about 8.4 percent of the center rods exceed 
100 MPa stress at peak temperature. At other locations in the waste package, a lower percent 
will exceed 100 MPa. No rods are expected to exceed the solvus temperature since the 
maximum temperature is 268C (Appendix A of this report). Using the screening arguments 
suggested by Cragnolino et al. (1999 [DIRS 152354]), few, if any, rods are expected to undergo 
reorientation. 
Additional corroborating information follows. Section 6.4 of Hydride-Related Degradation of 
SNF Cladding Under Repository Conditions (CRWMS M&O 2000f [DIRS 148251]) reviewed 
hydride reorientation. The total hydrogen content in the fuel cladding is not as important as the 
amount of hydrogen in hydrides aligned perpendicular to the largest principal tensile stress. In 
commercial reactor cladding during irradiation, hydrides form in a circumferential orientation 
(the normal to the platelet is in the radial direction). Such hydrides do not significantly weaken 
the cladding against hoop stress. Reorientation usually occurs under tensile stresses ranging 
from 69 to 208 MPa (with temperature dependency). Test rods at the low end of this range, but 
significantly higher than the stress expected for most rods at the repository temperatures, showed 
no hydride reorientation. This lack of reorientation indicates that there should be little, if any, 
cladding degradation due to hydrides under normal repository temperatures (Einziger et al. 1982 
[DIRS 101604], p. 65). In one dry storage test, reorientation to the radial direction was observed 
in one rod (Einziger and Kohli 1984 [DIRS 101605], p. 119) although rod failure did not occur. 
March 2004 54 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
This occurred in a fuel rod with high stresses (145 MPa at 323C). Reorientation was not 
observed in fuel rods with lower stresses (13 to 26 MPa). The hydride reoriented so that the 
normals in the circumferential direction could possibly weaken the cladding. Reorientation was 
also observed in CANDU reactor coolant pipes. Hydride reorientation under repository 
conditions was investigated as a potential cladding degradation mode. 
Using reorientation data collected by Pescatore et al. (1990 [DIRS 101230], pp. 52 to 55), 
Section 6.2.6 of Clad Degradation  Summary and Abstraction (CRWMS M&O 2001a [DIRS 
151662]) developed a reorientation equation that defines the temperatures and stresses where 
reorientation is observed. This line can be defined as: 
Sr = 335  0.6  TC (Eq. 6.12-2) 
where 
Sr = stress, MPa, required for reorientation 
T = Peak Temperature, C, required for reorientation. 
Additional data (Cappelaere et al. 2002 [DIRS 164195], p. 9, Chan 1996 [DIRS 111876), 
Coleman 1982 [DIRS 111999], Einziger et al. 1982 [DIRS 101604], Einziger and Kohli 1984 
[DIRS 101605], Simpson and Chow 1987 [DIRS 164483], Wallace et al. 1989 [DIRS 112344], 
p. 68) are compared with the equation in Figure 6.12-2. Compared to the repository cooldown 
period, the time periods for those tests were short, but inspection of rods exposed to dry storage 
cooldown (15 years cooldown) also showed no hydride reorientation. If Equation 6.12-2 is 
considered a model, then Figure 6.12-2 could be considered validation of this model since the 
sources cited above are independent of the source for the equation (Pescatore et al. 1990 [DIRS 
101230], pp. 52 to 55). 
Figure 6.12-2 and Equation 6.12-2 show that at the cladding maximum temperature of 268C, 
the stress would have to be greater than 174 MPa before reorientation would occur. This 
corresponds to about 94 MPa at ambient temperature. Based on Initial Cladding Condition 
(CRWMS M&O 2000c [DIRS 151659]), an estimated 0.45 percent of the rods could have 
stresses this high and might undergo some reorientation. This is a small amount of the total 
inventory and supports the screening argument of low consequence. Rothman (1984 [DIRS 
100417]) also studied cladding degradation in a repository and concluded that hydride 
reorientation would not occur. He did not consider fuel with burnups and stresses as high as 
considered in Initial Cladding Condition (CRWMS M&O 2000c [DIRS 151659]). 
Pescatore et al. (1990 [DIRS 101230], pp. 54, 69) state that even with hydride reorientation, 
stress levels will be insufficient to result in DHC and clad failure. Rothman (1984 [DIRS 
100417], pp. 33 to 39) concludes that hydride reorientation is also unlikely because of the lack of 
large temperature gradients in the repository and the cladding stresses are lower than needed for 
reorientation. Peehs (1998 [DIRS 109219], pp. 5, 6) concluded that hydride reorientation would 
not occur in dry storage. The temperature history for dry storage is similar to that of the early 
times (first 100 years) of repository closure and, therefore, similar hydride motion would be 
expected to be negligible. 
March 2004 55 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
nn n 
x 550 YMP Equation 
n 500 
n 
y y 
450 
y y n 400 
y 350 
n 
300 ny y y y
n y n n y y y 
250 
No 
Reorientation 
200 
100 0 
Source: CRWMS M&O 2001a [DIRS 151662] 
Figure 6.12-2. Comparison of the YMP Reorientation Equation with Additional Experimental Data 
Puls (1988 [DIRS 102067], p. 1507-1522) performed a series of strain tests on Zircaloy-2 with 
reoriented hydrides. His results are summarized in Table 6.12-1. He took samples of CANDU 
coolant tubing and performed strain tests in the circumferential direction. The tubing is made in 
a similar fashion as cladding and develops hydride platelets with their normals in the radial 
direction. All tests were performed at ambient temperature. He used samples with both 20 ppm 
and 90 ppm hydrogen content and used two reorientation techniques to form hydrides of various 
lengths. One technique cooled the samples from 250C with the stress near the yield point 
(designated y in Table 6.12-1). Other samples were cooled from 350C at a stress of 200 MPa. 
This stress is more than twice that expected for most cladding in the repository (CRWMS M&O 
2000c [DIRS 151659], Figure 26). Longer hydrides were produced by cooling samples in a 
furnace, while shorter ones were produced by bench cooling samples. Table 6.12-1 provides the 
range of hydride lengths. Two types of samples were used, smooth and notched (designated by 
n in the table). The table provides the stress (y) for which 0.2 percent strain was measured. 
Also provided is the ultimate stress at which necking and imminent failure was observed (or, for 
the arrested tests, was expected). For some experiments, the tests were stopped when the sample 
started to neck, but before failure (designated a for arrested in the table). In all of these tests, the 
reoriented hydrides did not significantly change the stress for 0.2 percent strain or the ultimate 
stress. Both stresses are much higher than those expected in repository cladding. The yield and 
ultimate strains reported are also higher than the strain failure criteria of 1 percent used in the 
cladding strain failure analysis developed by Peehs and Fleisch (1986 [DIRS 102065]). 
ANL-WIS-MD-000008 REV 01 March 2004 56 of 90 
Max. Temperature, oC 
y = Reorientation 
n = No reorientation 
x = YMP Equation 
Experiments, Temperature 
cycles < 10 
Reorientation 
Region 
y 
y nn y n 
y yn n 
n n 
y y 
y 
reor-b.spw 
x 
400 300 200 500 
Stress, MPa Clad Degradation  FEPs Screening Arguments 
Table 6.12-1. Puls' Zircaloy-2 Strain Tests on Zirconium with Reoriented Hydrides 
Tensile Stress 
(MPa) 
650 
Hydride Length 
(m) 
Initial Material 
Yield Stress (0.2%) 
(MPa) 
627 
57 of 90 
Type* 
a 
7-20 - 632 678 4.7/15.8 
7-20 y 627 675 4.7/15.8 
7-20 - 612 659 4.7/15.8 
7-20 n 783 885 - 
7-20 n, y 774 882 - 
7-20 n 933 1095 - 
7-20 n 766 858 - 
7-20 628 698 6 / 9 
a 
30-60 - 627 689 4.7/14.3 
30-60 y 605 661 4.7/14.3 
30-60 n 861 958 - 
30-60 n, y 776 921 - 
50-90 - 1079 1160 4.1/13.6 
50-90 y 689 741 4.1/13.6 
50-90 - 625 647 4.1/13.6 
50-90 n 721 803 - 
50-90 n, y 923 1032 - 
50-90 n 811 936 - 
50-90 a 633 701 - / 6 
50-90 643 730 - / 6 
*Type, y= hydride reoriented near yield stress, n - notched, a = arrested (test terminated before failure) 
Source: Puls 1988 [DIRS 102067], Tables 1,3,6 
Hydride reorientation might require all the hydride platelets to be dissolved before cool-down 
and reprecipitation starts. This could be necessary because the hydrides prefer to reprecipitate on 
existing platelets and these earlier, existing platelets are oriented in the circumferential direction. 
For all the hydrides to dissolve at a maximum cladding temperature of 268C, the initial 
concentration must be less than 44 ppm (see Table 6.12-2). Initial Cladding Condition 
(CRWMS M&O 2000c [DIRS 151659], Figure 15) shows that less than 5 percent of the fuel has 
an average concentration this low. Knowing that the hydride concentration is directly 
proportional to the oxide thickness, Initial Cladding Condition (CRWMS M&O 2000c [DIRS 
151659], Figure 12) shows the fuel that could reorient under this hypothesis is the fuel with the 
lower burnups and, therefore, lower stresses. Mardon et al. (1997 [DIRS 109213], Figure 3, p. 
408) show hydride content as a function of burnup. To have less than 21 ppm, the figure shows 
that Zircaloy 4 fuel with burnups less than 10 MWd/kgU has low enough hydride concentrations 
to dissolve all the hydrides at 350C. Again, these fuels would have the lowest stresses. Rods 
cladded with M5 or ZIRLO have thinner oxide coatings (see Figure 6.3-1) and therefore, 
lower hydride content. These rods would be susceptible to reorientation at higher burnups if 
complete dissolution is needed. The rod that Einziger observed to undergo reorientation was 
exposed to a 570C transient which would support a solubility of 729 ppm (Table 6.12-2). It is 
possible that all the hydrides were dissolved in that experiment. 
ANL-WIS-MD-000008 REV 01 
Uniform/Total 
Strain (%) 
- 
March 2004 Clad Degradation  FEPs Screening Arguments 
Table 6.12-2. Saturation Limits (Tssd) for Hydrogen in Unirradiated Zirconium as a Function of 
Temperature 
H Concentration (ppm) 
March 2004 
Temperature (C) 
40 0.1 300 66 
50 0.2 310 75 
60 0.3 320 85 
70 0.4 330 96 
80 0.6 340 108 
90 0.9 350 120 
100 1.2 360 134 
110 1.6 370 149 
120 2.1 380 165 
130 2.8 390 182 
140 3.6 400 201 
150 4.6 410 221 
160 5.8 420 242 
170 7.3 430 264 
180 9.0 440 288 
190 11 450 313 
200 13 460 339 
210 16 470 367 
220 19 480 396 
226 22 490 427 
230 23 500 459 
240 27 510 493 
250 32 520 529 
260 38 530 565 
270 44 540 604 
280 50 550 644 
290 58 560 686 
300 66 570 729 
Source: Pescatore et al. 1990 [DIRS 101230], Eq. 6, p. 44 
Further support for these conclusions is provided by Figure 8 of Effects of Irradiation and 
Hydriding on the Mechanical Properties of Zircaloy-4 at High Fluence (Garde 1989 [DIRS 
113614]), which shows a clear correlation between oxide thickness and hydrogen content. Of 
greater significance however is the conclusion that low ductility values were obtained on guide 
tube tensile samples (dog-bone and ring tensile samples) at 300C because of the high value of 
the hydride orientation factor and the fact that the hydrogen concentration was barely above the 
solubility limit for these samples (see comment column in Garde's Table 5). This observation 
implies that the hydrogen content of these samples was sufficiently low to go back into solution 
and then reorient on favorable precipitation sites. By contrast, cladding samples with higher 
hydrogen concentrations retained a number of hydride platelets and these sites were preferential 
locations for reprecipitation. 
ANL-WIS-MD-000008 REV 01 
H Concentration (ppm) Temperature (C) 
58 of 90 Clad Degradation  FEPs Screening Arguments 
Recently, fuel from a Dry Storage Characterization Project at the Idaho National Engineering 
and Environmental Laboratory was examined after about 15 years of dry storage. Rods were 
exposed to stresses up to 62 MPa at temperatures approaching 415C (EPRI 2002 [DIRS 
161421], p. II-4-6) and there were few, if any, radial hydrides in the Surry fuel rods. These rods 
were exposed to a slower cooldown period (15 years) than those used in the reorientation 
experiments plotted in Figures 6.12-2 and 6.12-3, yet reorientation was not observed. 
Comparison of these temperatures and pressures with the equation plotted in Figure 6.12-2 
suggests that reorientation would not be expected. Recent NRC Interim Staff Guidance (NRC 
2002 [DIRS 164593], Appendix, p. 2) for dry storage and transportation suggests that 
reorientation is not significant if temperatures are below 400C, which is well above the 
maximum temperature of 268C. 
In summary, few, if any, rods will undergo hydride reorientation and their failure is unlikely. If 
cladding failure from hydride reorientation had been included in the abstraction for TSPA-LA no 
changes would be seen. Stresses and temperatures are too low for hydride reorientation to occur 
in most of the fuel, and since the cladding material will maintain sufficient toughness even if 
hydride reorientation did occur, failure would not be expected. 
Source: McMinn et al. 2000 [DIRS 112149], Figure 15, p. 187 
Figure 6.12-3. Dissolution and Precipitation Solubilities for Hydrogen in Irradiated Zirconium Metal 
6.12.5 Hydride Axial Migration (of Cladding) 
Hydride axial migration has been excluded from further consideration based on low consequence 
of occurrence. It is unlikely that sufficient hydrogen can be moved because of a lack of large 
temperature gradients (limits the driving force) in the waste packages and the low peak 
temperatures (limiting the amount of hydrogen in solution). 
Hydrides can form in cooler parts of the rod (end sections) because the hydrogen can dissolve 
into the fuel cladding metal matrix at a warmer area, diffuse toward the cooler area, and 
March 2004 59 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
condense there. The effect was studied for dry storage (Cunningham et al. 1987 [DIRS 101591], 
Appendix C) and, for a 90-year period, was determined not to be a problem. Cappelaere et al. 
(2002 [DIRS 164195], p. 5) studied axial hydrogen redistribution under dry storage conditions 
and concluded that with a maximum cladding temperature of 470C, and an axial temperature 
gradient of 180C, after 100 years, the amount of hydrogen in the cold end would increase by 
110 ppm. They conclude that axial hydrogen migration towards the ends of the rods is small. 
With the lower peak temperatures (268C) and temperature differences from rod center to ends 
(19C) (CRWMS M&O 2000e [DIRS 147881], Table 6-4) in the repository, even less migration 
is expected. Replacing the helium fill gas with humid air would increase the temperature 
difference, but not to the degree reported by Cappelaere et al. 
As the repository cools, the driving forces for this redistribution (hydrogen solubility, 
temperature gradient, and diffusion rate) all decrease. The waste package internals will act to 
minimize the temperature variation along the length of the fuel assembly. Figure 15 by McMinn 
et al. (2000 [DIRS 112149]) (reproduced here as Figure 6.12-3) gives the solubility of hydrogen. 
The solubility varies whether one dissolves or precipitates the hydrogen to form hydrides. 
Zirconium hydrides have a larger volume than the zirconium metal that they displace and the 
difference in the dissolution solubility (TSSd) and the solubility for precipitation (TSSp) 
represents a type of super-saturation. This difference in temperatures is required to have axial 
migration along the rod. The figure shows that in order to move the hydrogen, a minimum 
temperature difference of 42C between the hot location and cold location is needed. At 
temperatures below 200C, temperature differences over 75C are needed to move the hydrogen. 
At a peak cladding center temperature of 268C, a difference of at least 50C is needed to move 
and precipitate hydrogen. After repository closure, temperature differences of 19C are 
predicted (CRWMS M&O 2000e [DIRS 147881], Table 6-4). Therefore, little or no axial 
migration is expected based on the difference between condensation and dissolution solubility. 
6.12.6 Hydride Embrittlement from Fuel Reaction (Causes Failure of Cladding) 
Cladding failure from hydride embrittlement due to fuel reaction has been excluded from further 
consideration based on low consequence. Clad degradation from fuel reaction is always 
associated with further degradation of failed fuel, and is mostly observed in failed BWR rods 
Edsinger 2000 [DIRS 154433]. Both high temperatures and steam environments are required for 
rod unzipping (propagation of the initial failure from a DHC type cracking). The conditions 
inside the fuel rod become reducing with the UO2 consuming any free oxygen. The protective 
oxide film on the cladding interior becomes disturbed and hydrogen is absorbed into the 
cladding. In the repository reducing conditions are not expected and therefore hydrogen 
obsorption from fuel/cladding interaction would be minor. In an intact rod, the pellet gap and 
plenum are inert, so the fuel does not react. Failure of intact rods is not expected. In summary, 
hydride cracking of the cladding has been excluded from TSPA-LA because of low consequence 
since the stresses are not sufficient to fail the cladding at repository temperatures. 
6.13 CLADDING UNZIPPING 
FEP Number: 
2.1.02.23.0A 
March 2004 60 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
YMP FEP Description: 
In either dry or wet oxidizing conditions and with perforated fuel cladding, the UO2 fuel 
can oxidize. The volume increase of the fuel as it oxidizes can create stresses in the 
cladding that may cause gross rupture of the fuel cladding (unzipping). 
Descriptor Phrases: 
Cladding unzipping after initial perforation 
Dry oxidation/unzipping of cladding 
Volume increase of corrosion products (cladding unzipping) 
Wet oxidation/unzipping of cladding 
Screening Decision: 
Included 
TSPA-LA Disposition: Cladding axial splitting or unzipping is included in the TSPA-LA as 
being instantaneous once the waste package fails and the cladding is perforated. The cladding 
degradation abstraction models all failed rods (after initial perforation) as bare fuel pellet 
fragments for the full length of the fuel rod. Failed fuel rod unzipping (cladding axially splits 
down its length) is caused by the volume increase of corrosion products, FEP 2.1.09.03.0A, (fuel 
or cladding). It is based on experimental observations at Argonne National Laboratory, where 2 
of 2 rods unzipped in less than 2 years. Unzipping leaves fuel pellets exposed to the waste 
package internal environment. Details of the scientific analysis that describes the disposition in 
greater detail is presented in Clad Degradation  Summary and Abstraction for LA (BSC 2003a 
[DIRS 162153]). 
Unzipping by dry oxidation (oxidation of UO2 to U3O8) of the fuel requires low humidity and 
high temperature conditions. Dry oxidation is expected to occur in the repository if the waste 
package fails at closure and the fuel is exposed to the temperature transients presented in 
Appendix A of this report. If it should occur, it also would cause rapid unzipping and is well 
modeled with the instant unzipping model used in TSPA-LA. Unzipping by wet oxidation was 
not observed in the ANL tests because unzipping due to the fuel-side cladding corrosion 
occurred first. Wet oxidation is still fast and can be modeled by the instant unzipping model 
where necessary. 
Related FEPs: 
2.1.02.12.0A, Degradation of cladding prior to disposal 
2.1.09.03.0A, Volume increase of corrosion products impacts cladding 
6.14 MECHANICAL IMPACT ON CLADDING 
FEP Number: 
2.1.02.24.0A 
YMP FEP Description: 
Mechanical failure of cladding may result from external stresses such as rockfall or 
impact from waste package internals. Seismic induced impacts are addressed in a 
separate FEP. 
March 2004 61 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Descriptor Phrases: 
Rockfall (cladding damage) 
Waste package internals  cladding contact (cladding damage) 
Screening Decision: 
Excluded  Low Probability 
TSPA-LA Disposition: 
NA 
Related FEPs: 
1.2.03.02.0A, Seismic ground motion damages EBS components 
2.1.07.01.0A, Rockfall 
2.1.09.03.0A, Volume increase of corrosion products impacts cladding 
Screening Argument: Mechanical failure of the cladding from external stresses originating 
outside the waste package has been excluded from the TSPA-LA on the basis of low probability. 
According to FEPs Screening of Processes and Issues in Drip Shield and Waste Package 
Degradation (CRWMS M&O 2001b [DIRS 153937], FEP 2.1.07.01.0A), nominal rockfall could 
damage the drip shield but will not deflect the drip shield sufficiently to contact the waste 
package. Waste package damage from rockfall was screened out based on low probability of 
occurrence. Therefore, during the 10,000-year regulatory period, no damage to the waste 
package is expected. The CSNF (waste form and cladding) is located inside the waste package 
and no damage is expected. 
Mechanical failure of the cladding from external stresses originating inside the waste package, 
such as from the degradation of basket components and other WP internals, has also been 
excluded from the TSPA-LA on the basis of low probability of occurrence. Volume increase 
from the waste package corrosion products is addressed in BSC 2003c [DIRS 163935], 
Section 6.5.1.2.1.3.2. Initially the WP void fraction is 64% (4.7 m3 free volume) but after the 
WP internals corrode, the void fraction is 26% (1.9 m3 free volume). Because of the high void 
fraction, even after corrosion, the waste package corrosion products do not mechanically load 
and fail the cladding and there is sufficient free volume for any helium that is released from the 
fuel (see FEP 2.1.12.02.0A). 
Another approach is to consider dimensional changes in the fuel channel that could be produced 
from hematite (Fe2O3) formation. The inside width of fuel channel is 22.64 cm and the steel 
component (fuel basket tube) is 0.5 cm thick (BSC 2003c [DIRS 163935], Table 34). The iron 
(density = 7.85 g/cm3) would change to hematite (density = 5.24 g/cm3 (BSC 2003c [DIRS 
163935], Table 11). Considering the corrosion products can only form in the fuel channel, the 
inside dimension would decrease by 0.5 cm [(2  0.5  7.85/5.24)  2  0.5 = 0.5 cm]. The 
standard Westinghouse 1717 assembly is 21.4 cm wide (DOE 1996 [DIRS 100320], p. 123). 
Although the free clearance would decrease from 1.2 cm to 0.7 cm after the formation of the 
hematite, fuel failure is not expected. In a fuel assembly, the rods do not contact each other and 
there is coolant spacing between the rods. Since no fuel failure is expected from the iron 
corrosion, this FEP is excluded in the TSPA-LA. 
March 2004 62 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Effects of longer term loading due to volume increase from waste form and cladding corrosion 
products (i.e., internal stresses) are addressed in FEP 2.1.09.03.0A, Volume Increase of 
Corrosion Products Impacts Cladding. Seismic-induced rockfall, drift degradation, and ground 
motion are also not treated within this FEP. A full discussion of seismic effects is contained in 
FEPs 1.2.03.02.0A, Seismic Ground Motion Damages EBS Components; 1.2.03.02.0B, Seismic 
Induced Rockfall Damages EBS Components; and 1.2.03.02.0C, Seismic Induced Drift Collapse 
Damages EBS Components. The effect of seismic events on fuel and cladding is addressed in 
Seismic Consequence Analysis (BSC 2004e) [DIRS 167780]. 
6.15 NAVAL SNF CLADDING 
FEP Number: 
2.1.02.25.0B 
YMP FEP Description: 
DSNF to be disposed of in Yucca Mountain has a variety of fuel types that may not be 
similar to the CSNF to be disposed. Some of the fuel types may have initial claddingdegradation 
characteristics that are different from those for the CSNF. Therefore, the 
effectiveness of DSNF cladding as a barrier to radionuclide mobilization might be 
different from CSNF. This FEP addresses Naval SNF cladding only. 
Descriptor Phrases: 
Degradation of Naval SNF cladding 
Screening Decision: 
Included 
TSPA-LA Disposition: Naval SNF cladding is modeled in the TSPA cladding degradation 
abstraction as CSNF cladding. The 300 waste packages containing naval SNF have been added 
to the CSNF inventory (approximately 7,500 waste packages) and both of these fuels are 
modeled as CSNF in the TSPA-LA. More information is available in Initial Radionuclide 
Inventories (BSC 2003d [DIRS 161961]). A previous comparison with an equivalent amount of 
Zircaloy-clad CSNF indicates that the effect on the resulting radionuclide exposures to the RMEI 
from the TSPA simulation using the commercial-fuel equivalent is significantly higher than the 
effect on the resulting radionuclide exposures to the RMEI from the source-term simulation for 
naval SNF (BSC 2001a [DIRS 152059], p. 36 and Figure 6.1-2). This analysis shows that 
modeling naval spent nuclear fuels as CSNF is conservative because of the robustness of naval 
SNF. Details of the scientific analysis that describes the disposition in greater detail is presented 
in Clad Degradation  Summary and Abstraction for LA (BSC 2003a [DIRS 162153]). 
Related FEPs: 
All CSNF FEPs addressed in this report 
6.16 DIFFUSION-CONTROLLED CAVITY GROWTH (DCCG) IN CLADDING 
FEP Number: 
2.1.02.26.0A 
March 2004 63 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
YMP FEP Description: 
Diffusion-Controlled Cavity Growth (DCCG) was once thought to be a possible creep 
rupture mechanism that could occur under the temperature and pressure conditions that 
prevailed during dry storage of spent fuel and might occur during disposal. 
Descriptor Phrases: 
Diffusion-controlled cavity growth (DCCG) cladding creep (rupture) 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.19.0A, Creep rupture of cladding 
Screening Argument: Diffusion-Controlled Cavity Growth (DCCG) in cladding is excluded 
from the TSPA-LA on the basis of low consequence. 
Glide and Coble creep rupture processes are included in the analysis and would account for creep 
failed fuel. Applicants for dry storage licenses for CSNF were once required by the U.S. Nuclear 
Regulatory Commission (NRC) to model diffusion controlled cavity growth (DCCG) to evaluate 
dry storage designs. That is, NUREG-1536 (NRC 1997 [DIRS 101903], Section 4.V.1) required 
the use of the DCCG method to calculate a maximum cladding temperature limit for a dry 
storage design. However, this design limit is overly restrictive and relatively inflexible. The 
literature does not support the use of this model for zirconium-based materials (Pescatore and 
Cowgill 1994 [DIRS 102066], p. 83 to 85) since it has not been validated, and voids and cavities 
are rarely seen in irradiated Zircaloy. Pescatore and Cowgill (1994 [DIRS 102066], p. 85) 
recommend a methodology of calculating the amount of creep and comparing it to a creep failure 
criterion. The earlier NRC Interim Staff Guidance (ISG) Number 11 (NRC 2000 [DIRS 
147797]) for transportation and storage recognizes the controversy with the DCCG conceptual 
model and permits license applicants to use other creep models in their license application. ISG 
11, Rev 2 (NRC 2002 [DIRS 164593]) does not address specific creep models but concludes that 
creep failures are unlikely if peak temperatures are below 400C. The temperature profiles for 
dry storage (time at temperature, see Figure 6 of Clad Degradation  Summary and Abstraction 
(CRWMS M&O 2001a [DIRS 151662])) are similar to those in the early periods of repository 
closure when DCCG might occur. With peak cladding temperatures below 268C (Appendix A 
of this report), creep failure (including DCCG failures) is not expected. 
Hayes et al (1999 [DIRS 164598], Figures 2, 5, 6, 8, and 11) concluded that failure from DCCG 
is unlikely if the peak temperature is less than 330 to 400C depending on the specific DCCG 
model (LLNL vs. PNNL), initial stress, and heat decay curve. None of their work suggests 
failure by DCCG at temperatures as low as 268C. In summary, omission of this FEP is justified 
on the basis that DCCG has a low consequence of occurrence because DCCG, as a mechanism to 
fail Zircaloy cladding, is not expected to occur at the temperatures (peak temperatures below 
268C) predicted at the repository. 
March 2004 64 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
The creep analysis described in FEP 2.1.02.19.0A is an alternative conceptual model for DCCG 
model. The creep analysis was selected because of the stronger experimental basis for this type 
of strain. 
In conclusion, DCCG in cladding is excluded from the TSPA-LA on the basis of low 
consequence. The NRC requirements in 10 CFR 63.114 (e and f) allow this omission because 
DCCG has not been observed experimentally and therefore will not significantly change the 
magnitude and time of the radiological exposures to the reasonably maximally exposed 
individual (RMEI), or radionuclide releases to the accessible environment. 
6.17 LOCALIZED (FLUORIDE ENHANCED) CORROSION OF CLADDING 
FEP Number: 
2.1.02.27.0A 
YMP FEP Description: 
Fluoride is present in Yucca Mountain groundwater, and zirconium has been observed to 
corrode in environments containing fluoride. Therefore, fluoride corrosion of cladding 
may occur in waste packages. 
Descriptor Phrases: 
Localized corrosion of cladding enhanced by fluoride 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.15.0A, Localized (radiolysis enhanced) corrosion of cladding 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
Screening Argument: Fluoride enhanced corrosion of cladding is excluded from the TSPA-LA 
on the basis of low consequence. 
Hydrofluoric acid can contribute to an accelerated general corrosion with fluoride concentrations 
greater than 5 ppm and pH less than 3.18 (Pourbaix 1974 [DIRS 100817], p. 583). The inpackage 
chemistry model (BSC 2003f [DIRS 161962], Attachment III) predicts pHs greater than 
3.5 and J13 well water contains only 2.2 ppm of fluoride. Since neither of the conditions is met, 
accelerated corrosion from fluoride is not expected. Even if pH dropped below 3, fluoride 
enhanced localized corrosion of the cladding would not occur because the fluoride concentration 
would still be less than 5 ppm. 
As corroborating evidence, corrosion of zirconium by fluorides is addressed in Clad 
DegradationLocal Corrosion of Zirconium and Its Alloys Under Repository Conditions 
(CWRMS M&O 2000d [DIRS 136058], Sections 4.1, 6.1.5, 6.2.2.3, and III.4). Zirconium 
resists attack by most halides, including halogen acids. The major exceptions are hydrofluoric 
March 2004 65 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation C FEPs Screening Arguments 
acid and ferric chloride (localized (pitting) corrosion of cladding, FEP 2.1.02.16.0A). As shown 
in Section 6.1.3 of Clad DegradationCLocal Corrosion of Zirconium and Its Alloys Under 
Repository Conditions (CWRMS M&O 2000d [DIRS 136058]), zirconium is corrosion resistant 
to certain fluorides when the pH is sufficiently high. Low fluoride ion concentrations (F- ions), 
on the order of a few ppm, in city or ground water have little effect on zirconium's excellent 
corrosion resistance. However, a few ppm of hydrofluoric acid (HF molecule in solution) will 
noticeably increase zirconium's corrosion rate. Hydrofluoric acid only exists in solution at pHs 
below 3.18 (Pourbaix 1974 [DIRS 100817], p. 583). 
For accelerated corrosion to occur, the fluoride must be present as free ions (i.e., not complexed 
as compounds), and the pH must be low. A high insoluble fluoride concentration (in essence a 
low fluoride ion concentration) would not be expected to have much impact on the standard 
zirconium corrosion rate. Section 4.1.1 of Clad DegradationCLocal Corrosion of Zirconium and 
Its Alloys Under Repository Conditions (CWRMS M&O 2000d [DIRS 136058], Test 12) shows 
that fluoride ion concentrations of less than 5 ppm, even at pH values as low as 1, produce 
similar corrosion rates to those with zero fluoride ion concentration. Thus, it is reasonable to 
conclude that low fluoride ion concentrations, as distinct from total fluoride content, will have 
limited impact on the uniform Zircaloy corrosion rate. 
Zirconium and its alloys generally exhibit low corrosion rates in fluoride solutions, including 
relatively high fluoride ion content solutions, if the temperature is sufficiently low and the pH is 
sufficiently high. This is illustrated with the results in Section 4.1.1 of Clad DegradationCLocal 
Corrosion of Zirconium and Its Alloys Under Repository Conditions (CWRMS M&O 2000d 
[DIRS 136058], Test 13). However, if the metal is in contact with solutions containing HF, the 
corrosion rate can increase rapidly. From the Pourbaix (1974 [DIRS 100817], p. 583) diagram, 
HF can exist when the pH is less than 3.18, although this does not necessarily mean that all 
fluoride ions are immediately converted to HF below this value. The data in Section 4.1.1 of 
Clad DegradationCLocal Corrosion of Zirconium and Its Alloys Under Repository Conditions 
(CWRMS M&O 2000d [DIRS 136058], Tests 10 through 13) can be divided into fluoride ioncontaining 
solutions and HF-containing solutions using a pH of 3.18 as the demarcation point. 
The calcium fluoride, due to its low solubility, does not contribute to the fluoride ion 
concentration (about 2 ppm at 25C and less than 3 ppm at 90C). 
If the fluoride were to react with the cladding, the amount of corrosion would be limited by the 
amount of fluoride entering the waste package because the fluoride is consumed in the reaction 
Zr + 4F  Zr F4. For example, if 1 liter of J-13 well water enters a waste package each year, the 
total contained fluorine content is 2.2 mg (2.2 wt. ppm). If all the available fluorine reacts with 
zirconium to produce ZrF4, the maximum quantity of zirconium that could be corroded away is 
2.6 mg per year (4.0 x 10-4 cc/year). This volume represents a general corrosion mechanism and 
would be distributed over the wet area. If it was concentrated at one spot of cladding surface 
area, 0.008 square centimeters of fuel would be exposed (using nominal wall thickness of 500 
microns). This quantity of fuel exposure is not significant and, therefore, fluoride corrosion 
would be a low consequence degradation mode. Much larger quantities of water containing 
fluorides than were used in the example must enter the waste package to fail significant 
quantities of fuel. 
March 2004 66 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Repository conditions (as represented by J-13 well water) would not be expected to produce any 
significantly different corrosion rates in zirconium and its alloys than in general corrosion (FEP 
2.1.02.13.0A). The well water analysis (Table 6.17-1) shows a neutral solution with impurity 
concentrations too low to be corrosive to zirconium and its alloys. However, it has been 
hypothesized that ground water entering the repository may be concentrated in impurities as a 
result of evaporation. Table 6.17-1 provides the predicted values. Of particular note is the fact 
that the halide content could become enriched due to the high solubility of most chlorides and 
fluorides. As a result, the corrosive potential of the water increases as the halide concentrations 
increase. However, the pH of the solution increases at the same time, and this is favorable 
because zirconium and its alloys are generally corrosion resistant at the higher pH values. That 
is, J-13 well water will not become oxidizing when the pH is so high. Conceivably, solution pH 
within crevices may become acidic. As discussed in Section 6.1.10 of Clad DegradationLocal 
Corrosion of Zirconium and Its Alloys Under Repository Conditions (CWRMS M&O 2000d 
[DIRS 136058]) the condition in a crevice would be too reducing to support the formation of 
oxidizing ions. 
Clad Degradation  Summary and Abstraction (CRWMS M&O 2001a [DIRS 151662]) 
contained an alternative conceptual model for fluoride corrosion. In that model the fluoride, at 
any concentration, was modeled to concentrate onto a 1-cm length of one rod and consume that 
rod until failure and then concentrate onto another rod. This sequential failure occurred well 
after the regulatory period and was not consistent with experimental observations cited above. 
This model has been dropped from TSPA-LA. 
Table 6.17-1. Compositions of J-13 Well Water and Its Concentrates 
Ion J-13 
Concentration (mg/liter) 
Beaker Evaporation 
(1000) 
15,700 
6,120 
6,730 
1,520 
31,471 
37,700 
3,720 
7 
0 
7,124 
9.9 
SO4 18.4 13,000 
Cl 7.14 7,200 
NO3 8.78 6,440 
F 2.18 1,580 
HCO3 128.9 47,326 
Na 45.8 42,500 
K 5.04 3,580 
Ca 13.0 3 
Mg 2.01 1 
SiO2 (aq) 61.0 109 
pH 7.4 10.1 
NOTE: Beaker Evaporation 1000X solution and Long Term Test 1000X solution agree reasonably well 
except the concentration of SiO2 (aq). This disagreement could be attributed to the attack of the 
In conclusion, the magnitude and time of the resulting radiological exposures to the reasonably 
maximally exposed individual (RMEI), or radionuclide releases to the accessible environment, 
would not be significantly changed by the omission of this FEP (fluoride enhanced corrosion) 
from the performance assessment (TSPA-LA) model. Few, if any, rods would experience 
67 of 90 
Long Term Test Solution 
(1000) 
glass beaker by fluoride ions. 
Source: BSC 2001b [DIRS 155640], Table 18 
ANL-WIS-MD-000008 REV 01 
90C / 
85% Relative Humidity 
29,500 
14,800 
14,200 
3,400 
11,370 
77,400 
9,700 
25 
0 
22,500 
10 
March 2004 Clad Degradation  FEPs Screening Arguments 
accelerated corrosion because the pH is too high (> 3.18) for the formation of hydrofluoric acid 
and the concentration of fluorine is too low (<5 ppm). 
6.18 ROCKFALL 
FEP Number: 
2.1.07.01.0A 
YMP FEP Description: 
Rockfalls may occur with blocks that are large enough to mechanically tear or rupture 
drip shields and/or waste packages. Seismic induced rockfall is addressed in a separate 
FEP. 
Descriptor Phrases: 
Rockfall (cladding damage) 
Screening Decision: 
Excluded  Low Probability 
TSPA-LA Disposition: 
NA 
Related FEPs: 
1.2.03.02.0A, Seismic ground motion damages EBS components 
1.2.03.02.0B, Seismic induced rockfall damages EBS components 
2.1.02.24.0A, Mechanical impact on cladding 
Screening Argument: Rockfall damage of the cladding has been excluded from the TSPA-LA 
on the basis of low probability of occurrence because the drip shield prevents rocks from striking 
the waste package and possibly damaging the cladding. According to FEPs Screening of 
Processes and Issues in Drip Shield and Waste Package Degradation (CRWMS M&O 2001b 
[DIRS 153937]), nominal rockfall could damage the drip shield but will not deflect the drip 
shield sufficiently to contact the waste package. Therefore, during the 10,000-year regulatory 
period, no damage to the waste package is expected. The CSNF (waste form and cladding) is 
located inside the waste package and no damage is expected. While the fuel and cladding might 
experience some minor vibration from rockfall, no damage is expected. Fuel and cladding are 
exposed to vibration during transportation and damage is not observed (Debes 1999 [DIRS 
161193]). 
Seismic induced rockfall and drift degradation are not considered within this FEP. A full 
discussion of seismic effects is contained in FEPs 1.2.03.02.0A, Seismic Ground Motion 
Damages EBS Components; 1.2.03.02.0B, Seismic Induced Rockfall Damages EBS 
Components; and 1.2.03.02.0C, Seismic Induced Drift Collapse Damages EBS Components. 
The effect of seismic events on fuel and cladding is addressed in Seismic Consequence Analysis 
(BSC 2004e [DIRS 167780]). 
March 2004 68 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
6.19 VOLUME INCREASE OF CORROSION PRODUCTS IMPACTS CLADDING 
FEP Number: 
2.1.09.03.0A 
YMP FEP Description: 
Corrosion products have a higher molar volume than the intact, uncorroded material. 
Increases in volume during waste form and cladding corrosion could change the stress 
state in the material being corroded and lead to cladding unzipping. 
Descriptor Phrases: 
Volume increase of corrosion products (cladding unzipping) 
Screening Decision: 
Included 
TSPA-LA Disposition: The volume increase of corrosion products causes cladding axial 
splitting, or unzipping, and is included in TSPA cladding degradation abstraction (BSC 2003a 
[DIRS 162153], Section 6.2.4). This FEP applies to failed cladding where water or moist air can 
interact with the fuel or cladding interior. The volume increase of corrosion products inside the 
cladding causes stress on the cladding and the cladding to tear open. This tearing is modeled to 
be instantaneous. All failed rods contain fuel pellet fragments for the full length of the fuel rod 
that are available for dissolution. Failed fuel rod unzipping (cladding axially splits down its 
length) is caused by the volume increase of corrosion products (fuel or cladding). It is based on 
experimental observations of two rods at ANL where both rods unzipped in less than two years. 
Unzipping leaves the fuel pellets exposed to the waste package internal environment. The 
scientific analysis that describes the disposition in greater detail is presented in Clad 
Degradation  Summary and Abstraction for LA, (BSC 2003a [DIRS 162153], Section 6.2.4). 
Unzipping by dry oxidation (oxidation of UO2 to U3O8) of the fuel requires low humidity and 
high temperature conditions. It is expected to occur in the repository if the waste package fails at 
closure and the fuel is exposed to the temperature transients given in Appendix A of this report. 
If dry oxidation should occur, it also would cause rapid unzipping and is well modeled with the 
instant unzipping model used in TSPA-LA. 
The effects of basket component degradation on external cladding integrity have been evaluated 
in the FEP Mechanical Impact on the Cladding, FEP 2.1.02.24.0A. 
Related FEPs: 
2.1.02.23.0A, Cladding unzipping 
2.1.02.12.0A, Mechanical Impact of Cladding 
6.20 ELECTROCHEMICAL EFFECTS IN EBS 
FEP Number: 
2.1.09.09.0A 
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YMP FEP Description: 
Electrochemical effects may establish an electric potential within the drift or between 
materials in the drift and more distant metallic materials. Migration of ions within such 
an electric field could affect corrosion of metals in the EBS and waste, and could also 
have a direct effect on the transport of radionuclides as charged ions. 
Descriptor Phrases: 
Electrophoresis/electro-osmosis 
Galvanic coupling (cladding) 
Galvanic coupling (waste form) 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.22.0A, Hydride cracking of cladding 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
Screening Argument: Electrochemical effects (electrophoresis and galvanic coupling) are 
excluded from the TSPA-LA on the basis of low consequence. 
Cladding, with its thick oxide layer produced in reactor operation, is kinetically noble (Yau and 
Webster 1987 [DIRS 100494], Table 15, p. 718). Any unprotected cladding oxidizes and forms 
a passive layer within seconds if exposed to water or humid air after being removed 
mechanically (Hansson 1984 [DIRS 101676], p. 6). Inside the waste package, the cladding 
would behave as a cathode and the steel box would be an anode if they are electronically 
coupled. The cladding can not undergo hydrogen charging because the oxide layer prevents 
hydrogen absorption in the metal (See FEP 2.1.12.03.0A). The consequence of such hydrogen 
charging would be to hydride the zirconium. Direct absorption of hydrogen gas in the 
environment into zirconium has occurred with clean zirconium metals with nickel coatings to 
prevent oxidation of the surface. It is unlikely to occur with irradiated cladding, which has heavy 
oxide coatings. 
The Center for Nuclear Waste Regulatory Analysis (Cragnolino 1999 [DIRS 152354], p. 4-13) 
surveyed various corrosion mechanisms for zirconium cladding under repository conditions. 
They concluded: 
Zr is not susceptible to galvanic corrosion because the protective ZrO2 passive 
film leads to Ecorr values in the galvanic series in flowing seawater close to those 
of noble metals and graphite but slightly lower than that of Ag (Yau and Webster, 
1987). However, local corrosion promoted by galvanic coupling to a more noble 
metal may occur if the film is mechanically disrupted. Nevertheless, the 
repassivation rate of Zr and its alloys is sufficiently fast in many aqueous 
solutions that unless fretting is continuously occurring no substantial corrosion 
can be expected. 
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Possible effects of galvanic coupling between cladding and the carbon steel rack are likely to 
have a low consequence. After reactor operation, fuel assemblies are coated with a thick oxide, 
which has a high electrical resistance and would minimize the galvanic effect. After the waste 
package is breached, the carbon steel rack would behave as an anode if coupled with the 
cladding and could corrode more quickly. An increase in corrosion of the carbon steel rack at 
assembly contact points would have a low consequence. Stainless steel and aluminum plates 
would maintain the geometry for some time. Galvanic corrosion of cladding or fuel boxes is 
excluded from the TSPA because of low consequence. 
Electrochemical effect on the waste form will also be minimal. The CSNF waste form is 
surrounded by the split zirconium cladding and would not be in direct contact with the waste 
package internals. If the UO2 did contact the steel fuel boxes and galvanic coupling did occur, 
the steel boxes would see the accelerated reaction and not the fuel pellets. As the UO2 corrodes, 
it coats itself with reaction products, which will minimize galvanic effects. Omission of 
electrochemical effects (electrophoresis and galvanic coupling) in the waste would not have a 
significant effect on the resulting radionuclide exposures to the RMEI because those effects are 
much smaller than the effects of modeling with a minimum flow rate through a failed container 
of 15 liters/yr used in performing the equilibrium-model calculations (CRWMS M&O 2000d 
[DIRS 136058]). This flow rate has a much greater effect than can be created by electrophoresis 
or electro-osmosis (Soderman and Jonsson 1996 [DIRS 149441]). 
Galvanic coupling could theoretically affect the rate of redox reactions involved in waste form 
matrix degradation when the waste form matrix is an electronic conductor. For example, 
galvanic coupling between the CSNF matrix (a semiconductor) and cladding could influence the 
corrosion potential, and hence, the rate of oxidative dissolution of the CSNF matrix. However, 
data from tests performed on CSNF rod segments (which include the cladding) and on CSNF 
fragments in Zircaloy holders indicate that such hypothetical galvanic coupling has a negligible 
effect on the rate of corrosion of the fuel matrix and the associated radionuclide release rate. The 
effects of galvanic coupling on the corrosion and associated radionuclide release rate from DSNF 
are also negligible because the base case model for DSNF corrosion causes it to corrode 
completely in one TSPA time step. Electrochemical effects on DHLW degradation are 
negligible because the principal reactions involved are glass network hydrolysis reactions that 
are not influenced by electrochemical effects. 
Electrochemical effects of various solutions that might cause pitting of the cladding are 
addressed in FEP 2.1.02.16.0A. The pitting model (BSC 2003g [DIRS 164667]) is an 
electrochemical model and the in-package chemistry model (BSC 2003f [DIRS 161962]) 
addresses the effect of EBS and waste package corrosion on chemistry. Figure 6.5-1 shows pH 
variation with time, including the effects of radiolysis. The depression of pH in this figure is 
caused by waste package steel corrosion. The result (Section 6.5 of this report (FEP 
2.1.02.15.0A) and Section 6.6 of this report (FEP 2.1.02.16.0A)) is that no failure is expected. 
In conclusion, electrochemical effects (electrophoresis and galvanic coupling) are excluded from 
the TSPA-LA on the basis of low consequence. The NRC requirements in 10 CFR 63.114 (e and 
f) allow this omission because electrochemical effects are expected to be small and will not 
significantly change the magnitude and time of the resulting radiological exposures to the 
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reasonably maximally exposed individual (RMEI), or radionuclide releases to the accessible 
environment. 
6.21 CHEMICAL EFFECTS OF WASTE-ROCK CONTACT 
FEP Number: 
2.1.09.11.0A 
YMP FEP Description: 
Waste (CSNF, DSNF, and HLW) and rock are placed in contact by mechanical failure of 
the drip shields and/or waste packages. Chemical effects on the waste (e.g., dissolution) 
of waste may be enhanced or altered in a system where waste, rock minerals, and water 
are all in physical contact, relative to a system where only waste and water are in physical 
contact. 
Descriptor Phrases: 
Chemical effects of waste-rock contact 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.13.0A, General corrosion of cladding 
2.1.02.14.0A, Microbially influenced corrosion (MIC) of cladding 
2.1.02.15.0A, Localized (radiolysis enhanced) corrosion of cladding 
2.1.02.16.0A, Localized (pitting) corrosion of cladding 
Screening Argument: The chemical effects of wasterock contact are excluded from the TSPALA 
on the basis of low consequence. 
Wasterock contact will have no effect on CSNF, DSNF, and HLW dissolution. Water that 
contacts SNF will have previously been in contact (i.e., equilibrium with the host rock) and 
therefore, placement of the rock in physical contact with the SNF will not affect the 
waste/rock/water system chemistry. Furthermore, In-Package Chemistry Abstraction (BSC 
2003f [DIRS 161962]) demonstrated that variations in the chemistry of the water contacting the 
waste package internal components, including the SNF, had an insignificant effect on the pH and 
total carbonate, the two key chemical parameters controlling the solubility of CSNF, DSNF, and 
HLW, of the in-package fluids. Waste form and cladding interaction with igneous intrusions are 
addressed in FEP 1.2.04.04.0A, Igneous intrusion interactions with EBS components. This FEP 
is addressed in Igneous Intrusion Impacts on Waste Package and Waste Forms (BSC 2003e 
[DIRS 165002]). 
The chemical effect of waste-rock contact has been excluded from TSPA-LA because of low 
consequence. Even if near-field chemistry contacts the cladding or waste form, accelerated 
corrosion is not expected. Tests involving the contact of zirconium with soils have not shown 
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corrosion. Adler Flitton et al. (2002 [DIRS 161991], p. 4) buried various metal samples 
including zirconium in an arid vadose zone environment for three years. These samples were in 
intimate contact with the vadose (unsaturated) zone environment. They reported indications of 
pitting on some of the other metals, but observed no pitting on the zirconium samples when in 
contact with rock. Other corroborating information is also available. The waste form is 
protected from the rocks by both the drip shield and waste package during the 10,000-year 
regulatory period. Zirconium is impervious to most chemical solutions (CRWMS M&O 2000d 
[DIRS 136058]) and only severe chemicals will pit cladding (BSC 2003g [DIRS 164667]). Yau 
(1984 [DIRS 102050]) performed corrosion tests with zirconium in geothermal fluids and reports 
no corrosion. Yau and Webster (1987 [DIRS 100494]) review the limited conditions where 
corrosion of zirconium is observed. This review suggests that contact with YMP tuff would not 
corrode the cladding. 
The CSNF and DSNF waste forms will have little contact with the rock because the stainless 
steel inner barrier and Alloy 22 outer barrier waste package will prevent contact. For the CSNF 
and some forms of DSNF, the zirconium alloy cladding surrounding the UO2 pellets will prevent 
contact. Contact of the UO2 pellets with the rock would have little effect. The uranium in the 
UO2 first oxidizes (U4+ to U6+), dissolves, and then precipitates immediately on the pellet surface 
as a U6+ mineral. The rock would have little contact with the UO2 pellets and little effect on the 
CSNF corrosion. DSNF and DHLW are also unlikely to contact the rock and direct effects of 
such contact, were it to occur, are also negligible. Indirect effects that could occur through the 
water chemistry (e.g., effects on dissolved silicon concentration) which would feed back into the 
rate of glass dissolution are small compared to the effects of the glass dissolution, as indicated by 
the results of tests conducted in EJ-13 water (i.e., J-13 well water preconditioned by reaction 
with tuff). 
In conclusion, the chemical effects of wasterock contact are excluded from the TSPA-LA on 
the basis of low consequence. NRC requirements in 10 CFR 63.114 (e and f) allow this 
omission because contact is not expected. Even if it did occur, the resulting chemical effects 
would not significantly change the magnitude and time of the resulting radiological exposures to 
the reasonably maximally exposed individual (RMEI), or radionuclide releases to the accessible 
environment. If contact were to occur during the regulatory period, no chemical effect is 
expected. 
6.22 THERMAL EXPANSION/STRESS OF IN-PACKAGE EBS COMPONENTS 
FEP Number: 
2.1.11.05.0A 
YMP FEP Description: 
Thermally induced stresses could alter the performance of the waste or EBS. For 
example, thermal stresses could cause the waste form to develop cracks and create 
pathways for preferential fluid flow and, thereby, accelerate degradation of the waste. 
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Descriptor Phrases: 
Thermal-mechanical effects on cladding 
Thermal-mechanical effects on CSNF waste form 
Thermal-mechanical effects on DSNF waste form 
Thermal-mechanical effects on HLW waste form 
Screening Decision: 
Excluded  Low Consequence 
TSPA-LA Disposition: 
NA 
Related FEPs: 
2.1.02.19.0A, Creep rupture of cladding 
Screening Argument: Thermal expansion and stresses of in-package EBS components, 
including the waste form, are excluded from the TSPA-LA on the basis of low consequence. 
The waste package and its internals are designed for the thermal expansion from the thermal 
cycle shown in Figure A-1. The CSNF and DSNF are designed for the thermal cycles expected 
in reactors, which are more severe than repository conditions. As discussed in Section 6.2.1 of 
CSNF Waste Form Degradation: Summary Abstraction (BSC 2004c [DIRS 167321]), the inreactor 
thermal cycles (principally that associated with the initial power escalation) result in 
extensive cracking of the fuel matrix. The effects of this cracking are included in the specific 
surface area parameter. Glass logs crack because of the cooldown during manufacturing. The 
cracking that results from this cooldown is included in the DHLW model surface area parameter. 
This cooldown (from molten glass, about 950C, BSC 2004d [DIRS 167619], Section 7.5.3) is 
more severe than repository conditions. 
Commercial nuclear fuel operates at higher temperatures than expected during the post-closure 
period at the repository. Under normal conditions, typical cladding operates at about 320C 
(Garde et al. 1991 [DIRS 101652], p. 582) with fuel centerline temperatures reaching 1,800C 
(Lanning et al. 1997 [DIRS 101704], V3, p. 3.2, Figure 3.1). Fuel is also designed to undergo 
anticipated operating occurrences (off normal transients that occur during the design life) without 
damage. These are more severe thermal cycles than normal reactor operation or repository 
closure. Every time a reactor shuts down and goes to cold shutdown, the fuel is cooled to below 
100C (coolant is less than boiling). These temperature transients are more severe than 
repository closure (Figure A-1). DSNF is also exposed to reactor transients more severe than the 
post-closure cooldown. Since the temperature transients for spent fuel from normal in-reactor 
operations and for DHLW from normal manufacturing cooldown are more severe than the 
transient associated with repository closure, no further degradation (cracking) is expected from 
thermal expansion or stress of in-package EBS components. 
In conclusion, thermal expansion and stress of in-package EBS components, including the waste 
form is excluded from the TSPA-LA on the basis of low consequence. The NRC requirements 
in 10 CFR 63.114 (e and f) allow this omission because thermal expansion will not significantly 
change the magnitude and time of the resulting radiological exposures to the reasonably 
maximally exposed individual (RMEI), or radionuclide releases to the accessible environment. 
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6.23 GAS GENERATION (He) FROM WASTE FORM DECAY 
FEP Number: 
2.1.12.02.0A 
YMP FEP Description: 
Helium (He) gas production may occur by alpha decay in the waste. Helium production 
might cause local pressure buildup in cracks in the fuel and in the void between fuel and 
cladding, leading to cladding and waste-package failure. 
Descriptor Phrases: 
Chemical effects from He generation 
Screening Decision: 
TSPA-LA Disposition: 
Internal gas pressure from He (cladding damage) 
Excluded  Low Consequence 
NA 
Related FEPs: 
2.1.02.20.0A, Internal pressurization of cladding 
2.1.09.03.0A, Volume increase of corrosion products impacts cladding 
Screening Argument: Effects of helium gas generation from alpha decay on the cladding in 
spent fuel are excluded from the TSPA-LA on the basis of low consequence. 
Piron and Pelletier (2001 [DIRS 165318], Section 5.3) investigated pressurization of the fuel 
rods from helium production (alpha decay) and concluded that fuel (47.5 MWd/kgU) would 
produce 1,171 cm3 (STP) of helium in a rod after 10,000 years, too low a quantity to damage the 
fuel. The peak pressure of 13.3 MPa (FEP 2.1.02.20.0A, Table 6.10-1) would have to be 
significantly higher (about 33 MPa) for the cladding to fail from delayed hydride cracking. 
2 
For cladding, delayed hydride cracking, stress corrosion cracking, and strain failures are driven 
by the cladding stress, which may be caused by the internal gas (including initial fill gas, fission 
product gases, and helium gas from alpha decay) pressure buildup. The gas pressure will slowly 
increase over time by the production of helium as a result of alpha decay. For failed rods, the 
helium would be released into the waste package, possibly increasing the pressure there. Using 
complete helium release from the fuel is conservative at these temperatures. Helium is an inert 
gas and will not chemically react with in-package components (internals, cladding or UO 
pellets). After waste package failure, it will also tend to diffuse out. The helium would have 
little effect on the corrosion of the UO2 pellets. UO2 corrosion is an oxidation/dissolution 
phenomenon and the presence of traces of an inert gas will not affect this. The effects of 
microcracking due to pressure buildup in gas bubbles within the fuel matrix on the rate of matrix 
corrosion are expected to be negligible. As described in Sections 6.2.1 and 6.2.2 of CSNF Waste 
Form Degradation: Summary Abstraction (BSC 2004c [DIRS 167321]), evidence from CSNF 
testing indicates that the corrosion process is a general corrosion process occurring 
predominantly at the periphery of the fuel fragments. The effective specific surface area of the 
corroding fuel is comparable to the geometric surface area and is not sensitive to the internal 
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grain boundary decohesion or microcracking that may result from helium buildup in the matrix. 
Likewise, because DHLW dissolution occurs at the periphery of glass fragments the internal 
microstructural features (including helium gas bubbles) do not affect the dissolution rate. 
After the waste package is placed in the repository and closure occurs, the temperature and 
pressure peak. The waste package then cools down, decreasing the pressure. This is countered 
by a steady production of helium by alpha decay of transuranic elements. After 10,000 years, 
total system pressure will increase to about 60 percent higher than at the thermal peak. Because 
this pressure is not high enough to cause any damage to the cladding (FEP 2.1.02.20.0A), there is 
a low consequence of this FEP. 
For an intact waste package, the free volume is 4.7 m3 (FEP 2.1.09.03.0A). If the waste package 
has failed, any helium that is released from the failed fuel can escape the waste package. It is 
expected that less than 1 percent of the fuel is failed (FEP 2.1.02.12.0A). If all of the rods were 
considered failed and all of the helium was released from the fuel matrix, then approximately 
1,500 cm3 (STP) of gas (FEP 2.1.02.20.0A) would be released per rod. With 5,544 rods per 
waste package (BSC 2003a [DIRS 162153], Table 8), there would be 8.3106 cm3 (STP) of 
helium available. When distributed into the free volume, 4.7 m3 (FEP 2.1.09.03.0A), of the 
waste package at 73C (peak temperature at 10,000 years, Table A-3), the pressure would 
increase by 0.22 MPa (2.2 atm). This pressure rise is too small to damage the waste package. 
With most of the fuel rods intact, and most of the helium tied up in the fuel matrix, the actual 
pressure rise would be small. An alternative conceptual model for helium generation is 
discussed in FEP 2.1.02.20.0A. 
In conclusion, effects of helium gas generation from alpha decay on the cladding are excluded 
from the TSPA-LA on the basis of low consequence. The NRC requirements in 10 CFR 63.114 
(e and f) allow this omission because the pressures generated by helium generation are 
insufficient to damage the cladding and, therefore, will not significantly change the magnitude 
and time of the resulting radiological exposures to the reasonably maximally exposed individual 
(RMEI), or radionuclide releases to the accessible environment. 
6.24 GAS GENERATION (H2) FROM WASTE PACKAGE CORROSION 
FEP Number: 
2.1.12.03.0A 
YMP FEP Description: 
Gas generation can affect the mechanical behavior of the host rock and engineered 
barriers, chemical conditions, and fluid flow, and, as a result, the transport of 
radionuclides. Gas generation due to oxic corrosion of waste containers, cladding, and/or 
structural materials will occur at early times following closure of the repository. Anoxic 
corrosion may follow the oxic phase if all oxygen is depleted. The formation of a gas 
phase around the waste package may exclude oxygen from the iron, thus inhibiting 
further corrosion. 
Descriptor Phrases: 
Hydride cracking (cladding) 
Internal gas pressure from H2 (waste package damage) 
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Screening Decision: 
TSPA-LA Disposition: 
Excluded  Low Consequence 
NA 
Related FEPs: 
2.1.03.01.0A, Waste package general corrosion 
2.1.02.22.0A, Hydride cracking of cladding 
Screening Argument: Hydrogen gas (H2) generation from waste package corrosion, as a 
degradation mechanism for cladding, is excluded from the TSPA-LA on the basis of low 
consequence. The cladding absorbs little or no hydrogen from the outside environment such as 
waste package corrosion and hydride embrittlement will not occur. 
Many investigators have considered the hydriding of zirconium alloys (for example, the papers 
cited in Clayton 1989 [DIRS 149208], Tables 1 through 4). Many of these investigations have 
been straightforward measurements of the rate of hydriding under various conditions. However, 
at least one set of experiments directly determined the origin of hydrogen in the metal. 
Waterside Corrosion of Zirconium Alloys in Nuclear Power Plants (IAEA 1998 [DIRS 150560], 
p. 92) discusses experiments in which zirconium-base alloys were oxidized in normal water 
(H2O) with dissolved tritium gas (T2). This experiment is extremely sensitive. If even one ppm of 
the hydrogen was from dissolved tritium gas, the radioactivity of T2 would result in thousands of 
decays per second for one square centimeter of surface. This level of activity would be readily 
detected. IAEA (1998 [DIRS 150560], p. 92) gives the following discussion of the experiment: 
Oxidation studies using T2/H2O mixtures ... have shown that, during normal 
oxidation, no T2 enters the metal ... until the thermally-induced exchange reaction 
has progressed to the point where a measurable fraction of HTO has been formed. 
Thus, the hydrogen isotopes which enter the metal do so as an integral part of the 
reaction of the zirconium with water molecules, and not by reaction with any 
dissolved hydrogen in the water. Studies have shown that this situation persists ... 
until hydrogen over-pressures in the system of tens of MPa are present. 
Note that, according to this quotation, the hydrogen pressures required to cause hydriding are 
quite large. By comparison, the highest hydrogen pressure in a breached waste package is pure 
hydrogen gas at atmospheric pressure, or about 0.1 MPa. 
Additional corroborating information follows. The oxide film on the surface of the metal is 
important in preventing the uptake of hydrogen from the surrounding environment. Under 
specific experimental conditions, this oxide layer can be removed. At 200C, the solubility of 
oxygen in zirconium is greater than 1 percent by weight, and the solubility of oxygen increases 
with increasing temperature (Baker 1992 [DIRS 149104], p. 2-326). As a result, the oxide film 
on the surface of a piece of zirconium is normally not thermodynamically stable. In his 
discussion of experimental procedure, Smith (1966 [DIRS 149107], p. 325) notes, Zirconium 
samples were first annealed at 700C under vacuum (~ 10-3 mm Hg) to remove any oxide film. 
The film dissolved into the samples, leaving them a bright metallic color. It is clear that the 
oxide film can be damaged or even destroyed by heat treatment in a suitable environment. 
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However, the film can be maintained if there is a supply of oxygen. Water will serve as a source 
of oxygen, because the electrochemical domain of stability for zirconium metal lies well below 
that of water (Pourbaix 1974 [DIRS 100817], p. 226). Spent nuclear fuel has a robust oxide 
layer from reactor operation. Under the oxidizing and humid environment of a failed waste 
package, it is expected that the oxide layer will be preserved. 
Garzarolli et al. (1979 [DIRS 149256], p. 64) studied the effect of the oxygen supply and stated: 
The effect of the composition of the gas atmosphere on the electrical properties of 
ZrO2 corrosion films was measured. The results revealed a large decrease in the 
electric resistance when the atmosphere changed from oxidizing to non-oxidizing, 
indicating a drastic change of the morphology (passivity of the oxide film)  the 
obvious implication of all available results is that massive hydriding can start 
when the availability of oxygen to continuously repair the protective oxide film 
falls below a critical value. 
Clayton (1989 [DIRS 149208], p. 270) quantified the conditions for hydriding in the equation: 
3 / 1 ( 2 . 0 p ) protective  ) p (Eq. 6.24-1) H2 ( H2O 
H 
H 
H2 2O 
where p 2O and pH2 are the pressures of H2O and H2, respectively. For Equation 6.24-1 to be 
applicable, both pressures must be given in torr (millimeters of Hg head). This fact is deduced as 
follows. First, Clayton (1989 [DIRS 149208], Tables 1 through 4) indicates that pressures are 
measured in mm. Second, pressures of 760 mm (of Hg) (= atmospheric pressure) occur 
many times in these tables. Third, the Critical p 2O/pH2 values given in the tables are consistent 
since both are expressed in the same units. For p = 101 kPa = 760 torr, it is found that pH 
(protective) = 0.24 kPa = 1.8 torr. Note that, by atmospheric standards, this pH2O corresponds to a 
dry gas. For comparison, the vapor pressure of water at 25C is about 3.2 kPa. For this small 
amount of humidity, corrosion of the waste package internals is not likely. If no corrosion 
occurs, no hydrogen is produced, so hydriding is impossible. 
It should be noted that hydriding of zirconium by absorption of gas has been observed in the 
laboratory. For example, Smith (1966 [DIRS 149107], Table 3) provides data on hydrogen 
absorption. However, the environment for these experiments was hot, extremely pure hydrogen. 
Smith (1966 [DIRS 149107], Table 3) states that the temperatures for the hydrogen absorption 
experiments were 210 C to 700 C. Smith (1966 [DIRS 149107], p. 325) notes that hydrogen 
was purified by passing it through a Deoxo unit, a bed of platinized asbestos (300C), a tube of 
P2O5 and a liquid nitrogen trap. The evident intention is to react any oxygen impurities, and 
absorb or condense any water vapor that is formed. After this treatment, little oxygen would have 
been available to maintain the oxide film. Such an environment is not relevant for a repository at 
Yucca Mountain because air, water vapor, or liquid water will be present and will maintain the 
protective oxide film. 
It is understood that corrosion of waste package internals will occur at temperatures below the 
range for which the Equation 6.24-1 was developed. Data on hydriding at lower temperatures 
were not available in the literature. 
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In some respects, it can be argued that it is conservative to apply Equation 6.24-1 to the 
repository. Equation 6.24-1 was developed for coupons of Zircaloy-2. This alloy is more 
susceptible to hydriding than is Zircaloy-4 (Clayton 1989 [DIRS 149208], Table 5), and coupons 
are more susceptible to hydriding than is tubing (Clayton 1989 [DIRS 149208], Table 5). 
Therefore, Equation 6.24-1 should give conservative predictions of the susceptibility of spent 
fuel cladding to hydriding. 
In summary, hydrogen will be generated in the waste package as the waste package internals 
corrode. This hydrogen is not directly absorbed, because the H2 molecules do not migrate 
through the high-density ZrO2 layer on the cladding. Hydrogen is introduced into the coolant in 
PWRs to reduce oxygen ions and reduce corrosion of components and yet it is not absorbed into 
the cladding. 
Hydrogen would have no effect on in-package chemistry because it would immediately diffuse 
out of the package before reacting with anything. Also, hydrogen is a reduced gas and is, 
therefore, unlikely to form in an oxidizing environment. The uranium in the UO2 first oxidizes 
(U4+ to U6+), dissolves, and then precipitates immediately on the pellet surface as a schoepitetype 
mineral. Although the kinetics of hydrogen gas reactions are sluggish under the conditions 
expected in a breached waste package, the presence of hydrogen would reduce the concentration 
of oxidizing agents and hence the corrosion rate of the UO2; neglecting this effect is therefore, 
conservative. The presence of some hydrogen gas would not affect the hydrolysis reactions that 
control the rate of DHLW dissolution. The hydrogen would be generated for the first 500 years 
while the waste package internals are corroding so this conservatism exists for a limited time. 
In conclusion, hydrogen gas (H2) generation from waste package corrosion, as a degradation 
mechanism for cladding, is excluded from the TSPA-LA on the basis of low consequence. The 
cladding absorbs little or no hydrogen from the outside environment such as waste package 
corrosion and hydride embrittlement will not occur. The NRC requirements in 10 CFR 63.114 (e 
and f) allow this omission because H2 will not damage the cladding and therefore will not 
significantly change the magnitude and time of the resulting radiological exposures to the 
reasonably maximally exposed individual (RMEI), or radionuclide releases to the accessible 
environment. 
7. CONCLUSIONS 
This report summarizes the screening of 24 waste form and cladding degradation FEPs. This 
screening is for the current LA design without backfill. The earlier screening of the FEPs 
(CRWMS M&O 2000g [DIRS 153947]) was based on backfill with higher temperatures and 
generally similar screening was developed. The degradation of cladding is not sensitive to 
temperature changes when the maximum temperatures are below 400C. Twenty-four (24) FEPs 
relevant to cladding degradation processes have been screened (see Sections 6.1 through 6.24) 
and are summarized in Table 7-1. The last five FEPs in this table also address waste form 
degradation. This table provides the FEP number, name, screening decision (included/excluded), 
and basis for the exclusion (i.e., low probability or low consequence). Low consequence means 
that omitting the FEP will not significantly change the magnitude and time of the resulting 
radiological exposures to the reasonably maximally exposed individual (RMEI), or radionuclide 
releases to the accessible environment (10 CFR 63.114 e and f). 
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For the first nineteen FEPs in Table 7-1, the FEP screening applies to CSNF and naval spent 
fuel. The last five FEPs in Table 7-1 address both cladding and all waste forms including 
defense spent nuclear fuel (DSNF) and high-level waste (HLW). Uncertainties are addressed in 
the included FEPs and are described in Clad Degradation  Summary and Abstraction for LA 
(BSC 2003a [DIRS 162153]). The list of Cladding Degradation FEPs is provided in Section 1.0 
and FEP screening analyses are provided in Sections 6.1 through 6.24 of this report. 
For the repository, cladding is not an engineered barrier. It is not designed and controlled by the 
project to reduce release of radionuclides. The behavior can be modeled but the design 
Screening Basis 
Low Consequence 
NA 
Excluded Low Consequence 
Excluded Low Consequence 
Low Consequence Excluded 
Low Consequence Excluded 
Low Consequence Excluded 
Low Consequence Excluded 
Included NA 
Low Probability Excluded 
NA Included 
Excluded Low Consequence 
Excluded Low Consequence 
Low Probability Excluded 
Low Consequence Excluded 
Low Consequence Excluded 
Excluded Low Consequence 
Low Consequence Excluded 
Excluded Low Consequence 
characteristics are not within the project controls. 
Table 7-1. Summary of Cladding Degradation and Waste Form FEPs 
General corrosion of cladding Excluded Low Consequence 
Sec FEP Number FEP Name Screening 
tion Decision 
6.1 2.1.02.11.0A Degradation of cladding from waterlogged rods Excluded 
6.2 2.1.02.12.0A Degradation of cladding prior to disposal Included 
6.3 2.1.02.13.0A 
6.4 2.1.02.14.0A Microbially influenced corrosion (MIC) of 
cladding 
6.5 2.1.02.15.0A Localized (radiolysis enhanced) corrosion of 
cladding 
6.6 2.1.02.16.0A Localized (pitting) corrosion of cladding Excluded Low Consequence 
6.7 2.1.02.17.0A Localized (crevice) corrosion of cladding Excluded Low Consequence 
6.8 2.1.02.18.0A Enhanced corrosion of cladding from dissolved Excluded Low Consequence 
silica 
6.9 2.1.02.19.0A Creep rupture of cladding 
6.10 2.1.02.20.0A Internal pressurization of cladding 
6.11 2.1.02.21.0A Stress corrosion cracking (SCC) of cladding 
6.12 2.1.02.22.0A Hydride cracking of cladding 
6.13 2.1.02.23.0A Cladding unzipping 
6.14 2.1.02.24.0A Mechanical impact on cladding 
6.15 2.1.02.25.0B Naval SNF cladding 
6.16 2.1.02.26.0A Diffusion-controlled cavity growth (DCCG) in 
cladding 
6.17 2.1.02.27.0A Localized (fluoride enhanced) corrosion of 
cladding 
Rockfall 6.18 2.1.07.01.0A 
6.19 2.1.09.03.0A Volume increase of corrosion products impacts Included NA 
cladding 
6.20 2.1.09.09.0A Electrochemical effects in EBS 
6.21 2.1.09.11.0A Chemical effects of waste-rock contact 
6.22 2.1.11.05.0A Thermal expansion/stress of in-package EBS 
components 
6.23 2.1.12.02.0A Gas generation (He) from waste form decay 
6.24 2.1.12.03.0A Gas generation (H2) from waste package 
corrosion 
March 2004 80 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
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Materials Behavior. 4A-3. 14. Erlangen, Germany: Framatome ANP GmbH. Copyright 
Requested. TBV-5665. 
Shi, S.Q. and Puls, M.P. 1994. "Criteria for Fracture Initiation at Hydrides in Zirconium Alloys I. 
Sharp Crack Tip." Journal of Nuclear Materials, 208, (3), 232-242. Amsterdam, The 
Netherlands: North-Holland Publishing Company. TIC: 237135. 
Simpson, L.A. and Chow, C.K. 1987. "Effect of Metallurgical Variables and Temperature on the 
Fracture Toughness of Zirconium Alloy Pressure Tubes." Zirconium in the Nuclear Industry, 
Seventh International Symposium, Strasbourg, France, 24-27 June 1985. Adamson, R.B. and 
Van Swam, L.F.P. eds. ASTM STP 939. Pages 579-596. [Philadelphia, Pennsylvania: American 
Society for Testing and Materials]. TIC: 254718. 
Smith, T. 1966. "Kinetics and Mechanism of Hydrogen Permeation of Oxide Films on 
Zirconium." Journal of Nuclear Materials, 18, 323-336. Amsterdam, The Netherlands: Elsevier. 
TIC: 247624. 
March 2004 88 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Soderman, O. and Jonsson, R. 1996. "Electro-Osmosis: Velocity Profiles in Different 
Geometries with Both Temporal and Spatial Resolution." Journal of Chemical Physics, 105, 
(23), 10300-10311. New York, New York: American Institute of Physics. TIC: 247669. 
tefanic I. and LaVerne, J.A. 2002. "Temperature Dependence of the Hydrogen Peroxide 
Production in the y-Radiolysis of Water." Journal of Physical Chemistry, 106, (2), 447-452. 
Washington, D.C.: American Chemical Society. TIC: 255323. 
Tasooji, A.; Einziger, R.E.; and Miller, A.K. 1984. "Modeling of Zircaloy Stress-Corrosion 
Cracking: Texture Effects and Dry Storage Spent Fuel Behavior." Zirconium in the Nuclear 
Industry, Sixth International Symposium, Vancouver, British Columbia, June 28- July 1, 1982. 
Franklin, D.G. and Adamson, R.B., eds. ASTM STP 824. Pages 595-626. Philadelphia, 
Pennsylvania: American Society for Testing and Materials. TIC: 241417. 
Wallace, A.C.; Shek, G.K.; and Lepik, O.E. 1989. "Effects of Hydride Morphology on Zr-2.5Nb 
Fracture Toughness." Zirconium in the Nuclear Industry: Eighth International Symposium, held 
19-23 June 1988 at San Diego, California. ASTM STP 1023. Pages 66-88. Philadelphia, 
Pennsylvania: American Society for Testing and Materials. TIC: 241414. 
Wolery, T.J. and Daveler, S.A. 1992. EQ6, A Computer Program for Reaction Path Modeling of 
Aqueous Geochemical Systems: Theoretical Manual, User's Guide, and Related Documentation 
(Version 7.0). UCRL-MA-110662 PT IV. Livermore, California: Lawrence Livermore National 
Laboratory. ACC: MOL.19980701.0459. 
Wolfram, J.H.; Mizia, R.E.; Jex, R.; Nelson, L.; Garcia, K.M. 1996. The Impact of Microbially 
Influenced Corrosion on Spent Nuclear Fuel and Storage Life. INEL-96/0335. Idaho Falls, 
Idaho: Idaho National Engineering Laboratory. ACC: MOL.20030925.0038. 
Yau, T-L. 1983. "Corrosion Properties of Zirconium in Chloride Solutions." Corrosion 83, 
International Corrosion Forum, Anaheim, California, April 18-22, 1983. Paper No. 26. Pages 
26/1 - 26/13. Houston, Texas: National Association of Corrosion Engineers. TIC: 243849. 
Yau, T. L. 1984. "Zirconium Versus Corrosive Species in Geothermal Fluids." Paper 140 of 
Corrosion 84. Houston, Texas: National Association of Corrosion Engineers. TIC: 238987. 
Yau, T.L. and Maguire, M. 1990. "Control of Localized Corrosion of Zirconium in Oxidizing 
Chloride Media." Advances in Localized Corrosion, Proceedings of the Second International 
Conference on Localized Corrosion, June 1-5, 1987, Orlando, Florida. Isaacs, H.S.; Bertocci, 
U.; Kruger, J. and Smialowska, S., eds. Pages 311-319. Houston, Texas: National Association of 
Corrosion Engineers. TIC: 247083. 
Yau, T.L. and Webster, R.T. 1987. "Corrosion of Zirconium and Hafnium." In Corrosion, 
Volume 13, Pages 707-721 of ASM Handbook. Formerly 9th Edition, Metals Handbook. 
Materials Park, Ohio: ASM International. TIC: 240704. 
March 2004 89 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
8.2 CODES, STANDARDS, REGULATIONS, AND PROCEDURES 
10 CFR 63. Disposal of High-Level Radioactive Wastes in a Geologic Repository at Yucca 
Mountain, NV. Readily available. 
10 CFR 60. Energy: Disposal of High-Level Radioactive Wastes in Geologic Repositories. 
Readily available 
AP-2.22Q, Rev. 1. Classification Analyses and Maintenance of the Q-List. Washington, D.C.: 
U.S. Department of Energy, Office of Civilian Radioactive Waste Management. ACC: 
DOC.20030807.0002. 
AP-SI.1Q Rev. 5. ICN 2 Software Management. Washington, DC: U.S. Department of Energy, 
Office of Civilian Radioactive Waste Management. ACC: DOC.20030902.0003. 
AP-SIII.2Q Rev. 1. ICN 2 Qualification of Unqualified Data. Washington, DC: U.S. Department 
of Energy, Office of Civilian Radioactive Waste Management. ACC: DOC.20040127.0008. 
AP-SIII.9Q, Rev. 1. ICN 3. Scientific Analyses. Washington, D.C.: U.S. Department of Energy, 
Office of Civilian Radioactive Waste Management. ACC: DOC.20040301.0002. 
ASTM C 1174-97. 1998. Standard Practice for Prediction of the Long-Term Behavior of 
Materials, Including Waste Forms, Used in Engineered Barrier Systems (EBS) for Geological 
Disposal of High-Level Radioactive Waste. West Conshohocken, Pennsylvania: American 
Society for Testing and Materials. TIC: 246015. 
8.3 SOURCE DATA, LISTED BY DATA TRACKING NUMBER 
MO0307SEPFEPS4.000. LA FEP List. Submittal date: 07/31/2003. 
March 2004 90 of 90 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
CLADDING TEMPERATURE ESTIMATES FOR POSTCLOSURE 
ANL-WIS-MD-000008 REV 01 
APPENDIX A 
A-1 March 2004 Clad Degradation  FEPs Screening Arguments 
INTENTIONALLY LEFT BLANK 
March 2004 A-2 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
APPENDIX A - CLADDING TEMPERATURE ESTIMATES FOR POSTCLOSURE 
Many of the features, events, and processes (FEPs) depend on temperature. Peak temperatures 
are important for some FEPs such as delayed hydride cracking. Other FEPs such as general 
corrosion or creep depend on both the temperature levels and how long the temperatures stay 
elevated. This appendix will develop both best estimate and peak temperatures for the cladding. 
The cladding temperatures are predicted by adding the interior temperature rise across the waste 
package to the waste package surface temperature. The waste package surface temperatures 
(BSC 2003b [DIRS 166463], Section 6.2) vary with waste package location in the repository, 
water infiltration flux, and time (from radioactive decay). The temperature rise across the waste 
package (BSC 2004b) [DIRS 166727] depends on the thermal loading (decay heat power) of the 
waste package and the loading pattern. The peak cladding temperature for a specific rod will 
also depend on what location the rod is loaded in the waste package. The maximum cladding 
temperatures tend to occur in the center location and the coolest locations are along the outer row 
of locations. 
This appendix will first address the expected waste package surface temperatures. It will then 
discuss the radial temperature distribution across the waste package for the fuel. Finally, these 
two pieces of information will be combined to give cladding temperatures. A comparison of the 
current temperature estimates with those used in the cladding degradation evaluation for SR 
(CRWMS M&O 2000a [DIRS 153246]) will also be made. 
Waste Package Surface Temperature 
The waste package surface temperatures (BSC 2003b [DIRS 166463], Section 6.2) depend on 
the location within the repository, the water infiltration flux, and the thermal loading. Table A-1 
(BSC 2003b [DIRS 166463], Section 6.2) provides the minimum, maximum and mean waste 
package surface temperatures for different infiltration fluxes and thermal conductivity of the 
rock. The overall temperatures (last row of Table A-1) have a minimum temperature of 74C, 
mean temperature of 138C, and the maximum temperature of 177C. The waste package 
temperature profiles for Bin 4, low infiltration, low thermal conductivity (bold row in Table A-1) 
will be used for estimating cladding temperatures. The selection of this group is conservative for 
estimating mean temperatures because it represents the hottest group. These temperatures are 
used in the first row of Table A-2 to estimate cladding temperatures. Figure A-1 provides the 
values of waste package surface temperature of the hottest waste package group (lowest flux, 
low thermal conductivity). This plot shows that after repository closure (after the 50 year 
ventilation period) the temperatures increase and reach a peak after about 20 years. They then 
decrease as the decay heat decreases exponentially (close to a straight line on the semi-log scale 
of Figure A-1). By 1000 years, the maximum waste package surface temperature is about 
114C. At 10,000 years, the maximum waste package has cooled to about 68C. 
March 2004 A-3 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Table A-1. The Minimum, Mean,and Maximum of the Peak CSNF Temperatures (C) for All Bins for the 
Three Infiltration Flux Cases for the TSPA-LA Base Case 
Thermal CSNF 
Conductivity Minimum 
Bin 
Number 
Infiltration 
Flux 
CSNF CSNF 
Mean Maximum 
(C) (C) (C) 
129 138 150 
123 133 142 
74 87 96 
144 155 170 
112 121 128 
122 138 152 
121 136 145 
121 134 143 
136 156 172 
112 123 130 
124 142 156 
121 139 145 
118 137 144 
138 160 176 
110 125 130 
123 144 156 
122 140 146 
114 137 144 
137 163 177 
108 125 130 
135 148 156 
133 142 145 
129 139 142 
151 166 176 
119 127 129 
1 Lower Mean 
1 Mean Mean 
1 Upper Mean 
1 Lower Low 
1 Upper High 
2 Lower Mean 
2 Mean Mean 
2 Upper Mean 
2 Lower Low 
2 Upper High 
3 Lower Mean 
3 Mean Mean 
3 Upper Mean 
3 Lower Low 
3 Upper High 
4 Lower Mean 
4 Mean Mean 
4 Upper Mean 
4 Lower Low 
4 Upper High 
5 Lower Mean 
5 Mean Mean 
5 Upper Mean 
5 Lower Low 
5 Upper High 
Overall 74 138 177 
Source: BSC 2003b [DIRS 166463], Section 6.2 
The predictions of the waste package surface temperatures have evolved over time as both the 
analytical tools and the design have evolved. The surface temperatures used for the cladding 
degradation evaluation for TSPA-SR were from an earlier evaluation (CRWMS M&O 2001a 
[DIRS 151662], Section 6.2.1) and represented waste packages with backfill. In that analysis, 
the mean value for waste package peak temperature reached 277C. An uncertainty of 13.5 
percent produced a maximum peak temperature of 315C. These temperatures are used in the 
first row of Table A-2 to estimate cladding temperatures. 
A-4 ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation  FEPs Screening Arguments 
180 
160 
140 
120 
100 
80 
60 
40 
Source: BSC 2003b [DIRS 166463], Section 6.2 
Figure A-1. WP Surface Temperature History for Hottest Group 
Interior Temperatures 
Sensitivity studies have been performed on the possible loading patterns for assemblies with high 
decay heat (hot assemblies) (BSC 2004b) [DIRS 166727]. The cladding temperature has a 
distribution that varies across the radius of the waste package, peaking in the center. The 
maximum thermal loading of the waste package is 11.8 kW. If the heat load is uniformly 
distributed to the 21 assemblies (562 W/assembly) then the temperature distribution across the 
waste package is given in Figure A-2. The center rod is the hottest, starting out 82C hotter than 
the waste package surface at closure and the temperature rise decreases with time (due to fission 
product decay). The outer rod (bottom curve) is the coolest (starting at 37C above waste 
package surface and cooling). The intermediate profile represents a rod located about 60 percent 
of the radial distance toward the outer edge. While only one assembly is placed in the center 
location, the remaining 20 assemblies are in cooler locations. Twelve assemblies are located in 
the coolest, outer rows of the waste package. Uniform loading represents the nominal or mean 
temperature rise. Hotter cladding temperatures are possible if high power assemblies are located 
in the center region and cooler temperatures are possible with hotter assemblies located in the 
outer region. This interior temperature rise will be added to the waste package surface 
temperatures to estimate cladding temperatures. 
ANL-WIS-MD-000008 REV 01 A-5 
Waste Package Surface Temperatures, oC 
Low infiltration, 
Low thermal conductivity 
Scenario 
Maximum 
Mean 
Minimum 
10 1 0.1 
Time from Closure, Years 
IntWP6.spw 
10000 1000 100 
March 2004 Clad Degradation  FEPs Screening Arguments 
80 
60 
40 
20
0 
0.1 1 
Source: BSC 2004b [DIRS 166727] 
Figure A-2. Average Temperature Rise Across the Waste Package 
Sensitivity studies of loading patterns were performed to estimate maximum cladding 
temperatures. Figure A-3 shows the temperature rise when a 2,100 Watt assembly is loaded in 
the center location and the remaining 20 assembly locations are loaded with 485 Watt 
assemblies. This is the hottest permitted loading pattern because the center rod is predicted to 
reach 350C with a simulated repository closure occurring after 25 years and using conservative 
design heat transfer codes (used for design conservatism). The current repository design and 
TSPA are based on 50 years of ventilation before closure. The center rod is the hottest, starting 
out 123C hotter than the waste package surface at closure and the temperature rise decreases 
with time (due to fission product decay). The outer rod (bottom curve) is the coolest (starting at 
38C above waste package surface and again decreasing with time). It is important to note that 
the temperature rise decreases significantly in the first 100 years. This interior temperature rise 
will be added to the waste package surface temperatures to estimate cladding temperatures. 
In the earlier TSPA-SR analysis, Thermal History of Cladding in a 21 PWR SNF WP Loaded 
with Average Fuel (CRWMS M&O 2000e [DIRS 147881], Table 6-2) was used to give the time 
dependent temperature distribution across the waste package for the initial average thermal 
loading of 9.1 kW. The earlier analysis of cladding degradation (CRWMS M&O 2001a [DIRS 
151662]) used a 31C temperature rise for the nominal SR case (thermal loading of 9.1 kW). 
This design also had a shrink fit between the inner and outer waste package so there was a much 
smaller temperature rise across the two barriers. The model only addressed uniform fuel loading. 
The peak temperature occurred three years after closure. The analysis used an uncertainty of 
13.5 percent to predict the maximum temperature rise of 35C. The 13.5 percent uncertainty was 
based on waste package loading errors and not on variation in waste package power. 
ANL-WIS-MD-000008 REV 01 A-6 
Temperature Rise from WP Outer Surface, oC 
IntWP1.spw 
Center Rod 
Rod 60% of Radius 
Outer Rod 
Uniform Fuel Loading 
10000 1000 100 10 
Time from Closure, Years 
March 2004 Clad Degradation  FEPs Screening Arguments 
Minimum temperature rise across the waste package would occur if the hotter assemblies were 
loaded on the outside row of the waste package or the waste package has a lower thermal loading 
than 11.8 kW per waste package (maximum design value). Such loading patterns were not 
studied and the mean thermal loading will be used to estimate the lower range of cladding 
temperatures. 
120 
100 
80 
60 
40 
20
0 
1 0.1 
Source: BSC 2004b [DIRS 166727] 
Figure A-3. Maximum Temperature Rise Across the Waste Package 
Cladding Temperatures 
The calculated center rod temperature is the sum of waste package surface temperature (see 
Figure A-1) and the temperature rise across the waste package (either Figure A-2 or A-3). The 
maximum cladding temperature is generated by placing the hottest possible fuel loading into the 
hottest waste package. Figure A-4 provides the temperature history of the hottest center rod, rod 
at 60 percent of radius and outer rod. The current repository design has the maximum cladding 
temperatures peak at 268C. The outer rod temperature peaks at 201C. Figure A-5 provides the 
temperature history of the average rod loaded into the average waste package. Here the center 
rod temperature peaks at 221C and the outer rod peaks at 186C. 
ANL-WIS-MD-000008 REV 01 A-7 
Temperature Rise from WP Outer Surface, oC 
IntWP2.spw 
Center Rod 
Rod 60% of Radius 
Outer Rod 
Hottest Fuel Loading 
10000 1000 100 10 
Time from Closure, Years 
March 2004 Clad Degradation  FEPs Screening Arguments 
280 
260 
Cladding Temperatures, oC Cladding Temperatures, oC 
240 
220 
200 
180 
160 
140 
120 
100 
80 
60 
0.1 
Source: This report 
Figure A-4. Temperature Profiles for Maximum Fuel Loading into Hottest Waste Package 
260 
240 
Mean WP Temperatures, 
Uniform Fuel Loading 
220 
200 
180 
160 
140 
120 
100 
80 
60 
40 
0.1 
Source: This report 
Figure A-5. Temperature Profiles for Mean Fuel Loading into Mean Waste Package 
ANL-WIS-MD-000008 REV 01 
Center Rod 
Rod 60% of Radius 
Outer Rod 
10 1 
Time from Closure, Years 
Center Rod 
Rod 60% of Radius 
Outer Rod 
10 1 
Time from Closure, Years 
A-8 
IntWP3.spw 
Hottest WP, 
Hottest Fuel Loading 
10000 1000 100 
IntWP4.spw 
10000 1000 100 
March 2004 Clad Degradation  FEPs Screening Arguments 
Figure A-6 provides the cladding temperature histories for an average loading of rods into the 
coolest waste package in the group being evaluated (Bin 4, lower infiltration flux, low thermal 
conductivity, see Table A-1). Here the center rod temperature peaks at 194C and the outer rod 
peaks at 160C. This group contains the hottest waste package. Colder waste packages are 
contained in other groups so Figure A-6 can not be considered minimum cladding temperatures 
for the repository. 
260 
240 
220 
200 
180 
160 
140 
120 
100 
80 
60 
40 
0.1 
Cladding Temperatures, oC 
Source: This report 
Figure A-6. Temperature Profiles for Mean Fuel Loading into Coolest Waste Package 
Table A-2 provides the cladding temperatures for the rods in the center of the waste package. 
Two cases are given. The case labeled Earlier Analysis shows the temperatures calculated in 
the cladding degradation for SR (CRWMS M&O 2001a [DIRS 151662]). In this case, maximum 
cladding temperatures peaked at 350C. These temperatures were higher than currently expected 
because they represented a backfill design evaluated in TSPA-VA. Even at these high 
temperatures, degradation from thermal effects was not expected. The case labeled Current 
Analysis shows the temperatures used for TSPA-LA. 
Table A-3 shows the center rod temperature histories for the average fuel loading into the 
average waste package and the maximum loading into the hottest WP. The temperatures are 
above 200C for less than 130 years. The peak temperature of 268C will be used to evaluate the 
FEPs in this report. Table A-3 also estimates the temperature of the cladding at the end of the 
regulatory period. The maximum cladding temperature is 73C at 10,000 years. This 
temperature will be used to evaluate FEPs addressing helium production from alpha decay of 
actinides. 
ANL-WIS-MD-000008 REV 01 A-9 
IntWP5.spw 
Center Rod 
Rod 60% of Radius 
Outer Rod 
Minimum WP Temperatures, 
Uniform Fuel Loading 
10000 1000 100 10 1 
Time from Closure, Years 
March 2004 Clad Degradation  FEPs Screening Arguments 
Table A-2. Peak Waste Package Surface Temperatures and Center Rod Cladding Temperatures 
Current Analysis 
Mean Maximum 
(C) 
Earlier Analysis 
Mean 
(C) 
Minimum 
(C) 
Maximum 
(C) 
Minimum 
(C) 
171 
97 
268 
(C) 
161 
60 
221 
315 
35 
350 
277 
31 
308 
245 
27 
272 
Waste package surface 
Temperature rise across waste package 
Center rod cladding peak temperature 
134 
60 
194 
Source: CRWMS M&O 2000e [DIRS 147881], BSC 2003b [DIRS 166463], Section 6.2, 
BSC 2004b [DIRS 166727] 
Table A-3. Time Dependency of Temperatures for Center Rod, Average and Maximum 
Temperature, 
Maximum Fuel 
Loading, Maximum 
Waste Package 
Temperature 
Temperature, 
Average Loading, 
Average Waste 
Package 
Temperature 
(C) 
Time 
(Years) (C) 
0 159 214 
5 208 260 
10 219 268 
266 15 221 
20 220 264 
30 214 254 
90 181 212 
130 169 197 
290 145 170 
950 111 131 
99 3000 82 
85 6000 71 
73 10,000 61 
Source: BSC 2003b [DIRS 166463], Section 6.2, 
BSC 2004b [DIRS 166727] 
Radionuclide transport depends on the water condensing or dripping. For local transportation of 
radionuclides to occur, the rods must cool to below the boiling point of the water (about 96C at 
the repository altitude). Table A-4 provides the approximate time to cool to boiling for rods in 
three radial locations of three different waste packages. 
Table A-4. Time from Closure to Cool Cladding to Boiling Point (96C) 
Outer Rod 
(Years) 
Rod 60% 
Radius (Years) 
Center Rod 
(Years) 
470 590 670 Cold waste package, Average Loading 
1460 1600 1680 Average waste package, Average Loading 
2600 2840 3520 Hottest waste package, Hottest Loading 
Source: BSC 2003b [DIRS 166463], Section 6.2 
BSC 2004b [DIRS 166727] 
March 2004 A-10 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
APPENDIX B 
DATA QUALIFICATION PLAN 
B-1 March 2004 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
INTENTIONALLY LEFT BLANK 
March 2004 B-2 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
APPENDIX B - DATA QUALIFICATION PLAN 
B-3 ANL-WIS-MD-000008 REV 01 
March 2004 Clad Degradation  FEPs Screening Arguments 
INTENTIONALLY LEFT BLANK 
March 2004 B-4 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
APPENDIX C 
QUALIFICATION OF TECHNICAL INFORMATION THAT SUPPORTS CLAD 
DEGRADATION - FEPS SCREENING ARGUMENTS (ANL-WIS-MD-000008) 
C-1 ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation  FEPs Screening Arguments 
INTENTIONALLY LEFT BLANK 
March 2004 C-2 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
APPENDIX C - QUALIFICATION OF TECHNICAL INFORMATION THAT 
SUPPORTS CLAD DEGRADATION - FEPS SCREENING ARGUMENTS 
(ANL-WIS-MD-000008) 
C.1 PURPOSE 
The Clad Degradation - FEPs (Features, Events and Processes) Screening Arguments 
(analysis) relies on technical information that may not have been collected under an approved 
quality assurance program that meets the requirements of 10 CFR Part 63, Subpart G [DIRS 
156605] or its predecessor, 10 CFR Part 60 [DIRS 103540]. This appendix qualifies the technical 
information that the cladding degradation analysis uses as direct input that is important to waste 
islolation. With the following exception, the appendix was prepared according to the Project's 
Procedure "Qualification of Unqualified Data" (AP-SIII.2Q, Revision 1, ICN 2). This appendix 
assumes "technical information" and "data" are synonymous; the procedure does not1. This 
qualification is being performed under the plan presented in Section C-5. 
C.2 METHODS AND CRITERIA 
The procedure offers five methods by which unqualified data may be qualified. The method of 
technical assessment was chosen because documentation or proof of proper data, or in this case 
information acquisition is unavailable for review. The assessment confirms that the information 
has been used in similar applications. According to AP-SIII.2Q (Attachment 3, Section 5.), this 
confirmation comprises: "A discussion and documentation that the data have been used in 
applications that are similar to those for which the data will be used." 
In addition to technical assessment, the qualification process considers at least one of the 
"Qualification Process Attributes" of AP-SIII.2Q Attachment 4. 
 Qualifications of personnel or organizations generating the information, 
 Prior peer or other professional reviews of the information and their results, and 
 Extent and quality of corroborating information. 
These criteria also served as the "data evaluation criteria" on which the Data Qualification Team 
provided judgements (AP-SIII.2Q, Section 5.1.2 (b)(4)). 
The team was comprised of two members: Eric Siegmann, a cladding expert and the analysis 
originator, chaired the team and Chris Pflum, who is technically competent in radioactive waste 
management, assisted. As the analysis originator, Eric Siegmann is not independent of the 
information to be qualified (AP-SIII.2Q, Section 5.1.2 (b)(3). However, the procedure requires 
that the originator chair the team " when the qualification is performed within the Analysis or 
1 This analysis and attachment were under review when the governing procedures: "Scientific Analyses" (APSIII.
9Q Rev. 1, ICN 3, 3/1/04) and "Managing Technical Product Inputs" (AP-3.15Q, Rev. 4, ICN 2, 2/25/04) 
were revised. The new procedures do not recognize "technical information". This attachment considers the 
factors that AP SIII.9Q (Section 5.2.1 (m)) lists for data that are suitable for intended use. AP- 3.15Q Rev. 4, ICN 
3 does not apply to this analysis because the analysis was finished within 30 days of the procedures effective 
date. 
March 2004 C-3 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Model Report" (AP-SIII.2Q Section 5.1.1). Chris Pflum did not participate in the acquisition or 
development of the information, and was an independent reviewer of the analysis. 
C.3 QUALIFICATION OF TECHNICAL INFORMATION 
The procedure does not apply to "Established Fact data or numerical data obtained from an 
established/authoritative data source". Because some of the technical information comes from 
authoritative sources (other than the U.S. Department of Energy, Office of Civilian Radioactive 
Waste Management), it does not need to be qualified. The remaining information is qualified 
only for its intended use i.e., to justify excluding from performance assessment twenty-one 
modes of cladding degradation. Some of the exclusionary arguments depend on qualitative 
arguments from the open literature that the analysis cites as technical information. 
Table C-1 lists sources of technical information and the criteria used to qualify the information. 
When the source is an authoritative source, the table provides only the sources' names. For the 
remaining sources, the Table designates the attributes or criteria that are used, in conjunction 
with the technical assessment, to qualify the information. 
Table C-1. Sources of Technical Information and Criteria Used to Qualify the Information 
Sources Criteria 
Beckman, D.A. 2001. Memorandum from D.A. Beckman (BSC), 
February 21, 2001, distributing: "NRC Issue Resolution Status 
Report on Container Life and Source Term, Revision 3," Division of 
Waste Management, Office of Nuclear Material Safety and 
Safeguards, U.S. Nuclear Regulatory Commission. ACC: 
MOL.20010418.0048. 
DIRS 156122 
Cragnolino, G.A.; Dunn, D.S.; Brossia, C.S.; Jain, V.; and Chan, 
K.S. 1999. Assessment of Performance Issues Related to Alternate 
Engineered Barrier System Materials and Design Options. CNWRA 
99-003. San Antonio, Texas: Center for Nuclear Waste Regulatory 
Analyses. TIC: 248875. 
DIRS 152354 
10 CFR 63. Energy: Disposal of High-Level Radioactive Wastes in a 
Geologic Repository at Yucca Mountain, Nevada. Readily available. 
DIRS 156605 
EPRI (Electric Power Research Institute) 2002. Dry Cask Storage 
Characterization Project. EPRI TR-1002882. Palo Alto, California: 
Electric Power Research Institute. TIC: 253737. 
DIRS 161421 
IAEA (International Atomic Energy Agency) 1998. Waterside 
Corrosion of Zirconium Alloys in Nuclear Power Plants. IAEATECDOC-
996. Vienna, Austria: International Atomic Energy Agency. 
TIC: 248234. 
DIRS 150560 
Lanning, D.D.; Beyer, C.E.; and Painter, C.L. 1997. FRAPCON-3: 
Modifications to Fuel Rod Material Properties and Performance 
Models for High-Burnup Application. NUREG/CR-6534. Volume 1. 
Richland, Washington: Pacific Northwest National Laboratory. TIC: 
238923. 
DIRS 101704 
Authoritative Source / Excluded from 
Qualification 
U.S. Nuclear Regulatory Commission 
Authoritative Source / Excluded from 
Qualification 
Center for Nuclear Waste Regulatory 
Analyses 
Authoritative Source / Excluded from 
Qualification 
U.S. Nuclear Regulatory Commission 
Authoritative Source / Excluded from 
Qualification 
Electric Power Research Institute 
Authoritative Source / Excluded from 
Qualification 
International Atomic Energy Agency 
Authoritative Source / Excluded from 
Qualification 
Contractor for the U.S. Nuclear Regulatory 
Commission 
C-4 ANL-WIS-MD-000008 REV 01 March 2004 Clad Degradation  FEPs Screening Arguments 
Table C-1. Sources of Technical Information and Criteria Used to Qualify the Information (Continued) 
Sources Criteria 
NRC (U.S. Nuclear Regulatory Commission) 1997. Standard 
Review Plan for Dry Cask Storage Systems. NUREG-1536. 
Washington, D.C.: U.S. Nuclear Regulatory Commission. ACC: 
MOL.20010724.0307. 
DIRS 101903 
NRC (U.S. Nuclear Regulatory Commission) 2000. "Interim Staff 
Guidance - 11. Storage of High Burnup Spent Fuel." ISG - 11. 
Washington, D.C.: U.S. Nuclear Regulatory Commission. Accessed 
March 21, 2000. TIC: 247227. 
http://www.nrc.gov/OPA/reports/isg11.htm 
DIRS 147797 
NRC (U.S. Nuclear Regulatory Commission) 2002. Interim Staff 
Guidance - 11, Revision 2. Cladding Considerations for the 
Transportation and Storage of Spent Fuel. ISG-11, Rev. 2. 
Washington, D.C.: U.S. Nuclear Regulatory Commission. TIC: 
254441. 
DIRS 164593 
Pourbaix, M. 1974. Atlas of Electrochemical Equilibria in Aqueous 
Solutions. Houston, Texas: National Association of Corrosion 
Engineers. TIC: 208955. 
DIRS 100817 
Yau, T.L. and Webster, R.T. 1987. "Corrosion of Zirconium and 
Hafnium." In Corrosion, Volume 13, Pages 707-721 of ASM 
Handbook. Formerly 9th Edition, Metals Handbook. [Materials Park, 
Ohio]: ASM International. TIC: 240704. 
DIRS 100494 
Adler Flitton, M.K.; Mizia, R.E.; and Bishop, C.W. 2002 
"Underground Corrosion of Activated Metals in an Arid Vadose Zone 
Environment." Corrosion/2002, [57th Annual Conference & 
Exposition, April 7-11, 2002, Denver, Colorado]. Paper No. 02531. 
Houston, Texas: NACE International. TIC: 253838. 
DIRS 161991 
Brossia, C.S.; Cragnolino, G.A.; and Dunn, D.S. 2002. "Effect of 
Oxide Thickness on the Localized Corrosion of Zircaloy." 
Corrosion/2002, [57th Annual Conference & Exposition, April 7-11, 
2002, Denver, Colorado]. Paper No. 02549. Houston, Texas: NACE 
International. TIC: 253839. 
DIRS 161988 
Cappelaere, C.; Limon, R.; Bredel, T.; Herter, P.; Gilbon, D.; 
Bouffioux, P.; and Mardon, J-P. 2002. "Long Term Behavior of the 
Spent Fuel Cladding in Dry Storage Conditions." [ICEM '01, 
Proceedings of the 8th International Conference on Radioactive 
Waste Management and Environmental Remediation, held in 
Burges, Belgium, September 30-October 4, 2001]. [Taboas, A.; 
Vanbrabant, R.; and Benda, G., eds]. [1], [New York, New York]: 
American Society of Mechanical Engineers. TIC: 254663. 
DIRS 164195 
Clayton, J.C. 1989. "Internal Hydriding in Irradiated Defected 
Zircaloy Fuel Rods." Zirconium in the Nuclear Industry: Eighth 
International Symposium, held June 19-23, 1988 San Diego, 
California. ASTM STP 1023. Pages 266-288. Philadelphia, 
Pennsylvania: American Society for Testing and Materials. TIC: 
241414. 
DIRS 149208 
Cunningham, M.E.; Simonen, E.P.; Allemann, R.T.; Levy, I.S.; 
C-5 ANL-WIS-MD-000008 REV 01 
Authoritative Source / Excluded from 
Qualification 
U.S. Nuclear Regulatory Commission 
Authoritative Source / Excluded from 
Qualification 
U.S. Nuclear Regulatory Commission 
Authoritative Source / Excluded from 
Qualification 
U.S. Nuclear Regulatory Commission 
Authoritative Source / Excluded from 
Qualification 
National Association Of Corrosion 
Engineers 
Authoritative Source / Excluded from 
Qualification 
American Society for Metals - Handbook 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Peer / professional review 
Personnel / organization qualifications 
March 2004 Clad Degradation  FEPs Screening Arguments 
Table C-1. Sources of Technical Information and Criteria Used to Qualify the Information (Continued) 
Sources Criteria 
Extent and quality of corroborating 
information. 
Hazelton, R.F.; and Gilbert, E.R. 1987. Control of Degradation of 
Spent LWR Fuel During Dry Storage in an Inert Atmosphere. PNL- 
6364. Richland, Washington: Pacific Northwest Laboratory. ACC: 
HQO.19941222.0016. 
DIRS 101591 
Edsinger, K. 2000. "A Review of Fuel Degradation in BWRs." 
Proceedings of the 2000 International Topical Meeting on Light 
Water Reactor Fuel Performance, Park City, Utah, April 10-13, 
2000. La Grange Park, Illinois: American Nuclear Society. TIC: 
248973. 
DIRS 154433 
Greene, C.A.; Brossia, C.S.; Dunn, D.S.; and Cragnolino, G.A. 
2000. "Environmental and Electrochemical Factors on the Localized 
Corrosion of Zircaloy-4." Corrosion/2000, [55th Annual Conference 
& Exposition, March 26-31, 2000, Orlando, Florida]. Paper No. 
00210. Houston, Texas: NACE International. TIC: 246988. 
DIRS 145073 
Hansson, C.M. 1984. The Corrosion of Zircaloy 2 in Anaerobic 
Synthetic Cement Pore Solution. SKB TR-84-13. Stockholm, 
Sweden: Svensk Krnbrnslefrsrjning A.B. TIC: 206293. 
DIRS 101676 Verified 
Hayes, T.A.; Rosen, R.S.; and Kassner, M.E. 1999. Critical Analysis 
of Dry Storage Temperature Limits for Zircaloy-Clad Spent Nuclear 
Fuel Based on Diffusion Controlled Cavity Growth. UCRL-ID- 
131098. Livermore, California: Lawrence Livermore National 
Laboratory. TIC: 254551. 
DIRS 164598 
Hillner, E.; Franklin, D.G.; and Smee, J.D. 1998. The Corrosion of 
Zircaloy-Clad Fuel Assemblies in a Geologic Repository 
Environment. WAPD-T-3173. West Mifflin, Pennsylvania: Bettis 
Atomic Power Laboratory. TIC: 237127. 
DIRS 100455 
Kreyns, P.H.; Bourgeous, W.F.; White, C.J.; Charpentier, P.L.; 
Kammenzind, B.F.; and Franklin, D.G. 1996. "Embrittlement of 
Reactor Core Materials." Zirconium in the Nuclear Industry, 
Eleventh International Symposium held in Garmisch-Partenkirchen, 
Germany, September 11-14, 1995. Bradley, E.R. and Sabol, G.P., 
eds. ASTM STP 1295. Pages 758-782. West Conshohocken, 
Pennsylvania: American Society for Testing and Materials. TIC: 
237256. 
DIRS 100462 
Peehs, M. 1998. Assessment of Dry Storage Performance of Spent 
LWR Fuel Assemblies with Increasing Burn-Up. Erlangen, Germany: 
Bereich Energieerzeugung. TIC: 245171. 
DIRS 109219 
Piron, J.P. and Pelletier, M. 2001. "State of the Art on the Helium 
Issues." Section 5.3 of Synthesis on the Long Term Behavior of the 
Spent Nuclear Fuel. Poinssot, C., ed. CEA-R-5958(E). Volume 1. 
[Paris], France: Commissariat  l'nergie Atomique. TIC: 253976. 
DIRS 165318 
Wolfram, J.H.; Mizia, R.E.; Jex, R.; Nelson, L.; and Garcia, K.M. 
1996. The Impact of Microbially Influenced Corrosion on Spent 
Nuclear Fuel and Storage Life. INEL-96/0335. Idaho Falls, Idaho: 
Idaho National Engineering Laboratory, Lockheed Martin Idaho 
Technologies Company. ACC: MOL.20030925.0039. 
DIRS 165268 
C-6 ANL-WIS-MD-000008 REV 01 
Personnel / organization qualifications 
Peer / professional review 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Peer / professional review 
Extent and quality of corroborating 
information. 
Personnel / organization qualifications 
Extent and quality of corroborating 
information 
March 2004 Clad Degradation  FEPs Screening Arguments 
Both the discussions above and in the following sections fulfill requirements of AP-SIII.2Q to 
discuss the data (in this case, technical information) for qualification, the method of 
qualification, and the evaluation criteria (See AP-SIII.2Q, Section 5.3.1, paragraphs (a) (1) 
through (3)). 
A discussion of the remaining requirements follows: 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(4), the team considers the information 
technically correct because it has undergone a technical assessment and it fullfills at 
least one of the Qualification Process Attributes. However, AP-SIII.2Q 
(Attachment 3, paragraph 1) does not require "an evaluation of the data quality and 
correctness" for data that would be qualified by the technical assessment method 
(referred to as Method 5). 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(5), no data are generated by this evaluation. 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(6), the team finds that the information is 
qualified for its use only within this document. 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(7), the team recommends that the information 
be given the status of qualified data should the applicable procedures no longer 
recognize technical information. Until then, the status should remain as technical 
information. 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(8), the team did not abandon any 
qualification method. 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(9), the technical information is qualified only 
to justify excluding from performance assessment twenty-one modes of cladding 
degradation. It is not qualified for other potential uses. 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(10), supporting or corroborating information 
is identified in Section C.3 when used in the qualification effort. 
 In regard to AP-SIII.2Q, Section 5.3.1 (a)(11), the qualification was performed under the 
data qualification plan presented in Section C.4 of this appendix. 
C.3.1 Qualification of Adler Flitton, M.K.; Mizia, R.E.; and Bishop, C.W. 2002 [DIRS 
161991] 
Technical AssessmentThe Cladding Degradation Analysis (i.e., this analysis, Section 6.21) 
applies the cladding/rock contact test in a manner that is consistent with the test results. Both the 
analysis and the test conclude that corrosion of zirconium was not observed after burial in soil. 
Personnel / Organization QualificationsThe tests were performed at a national laboratory 
(Idaho National Engineering and Environmental Laboratory). 
Peer / Professional ReviewThe source was peer-reviewed. 
March 2004 C-7 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Extent and Quality of Corroborating InformationAn authoritative source, Yau and Webster 
1987, corroborates this source's observation that zirconium resists corrosion in soil. 
C.3.2 Qualification of Brossia, C.S.; Cragnolino, G.A.; and Dunn, D.S. 2002. [DIRS 
161988] 
Technical AssessmentBoth this analysis (Section 6.7) and the above source apply consistent 
conclusions regarding localized (crevice) corrosion of the cladding. Both conclude that any 
crevice corrosion is not observed. 
Personnel / Organization QualificationsThe authors work for The Center for Nuclear Waste 
Regulatory Analyses, Southwest Research Institute. This work was sponsored by the NRC and 
performed under a QA program that is approved and audited by the NRC. 
Peer / Professional Review The source was peer-reviewed. 
Extent and Quality of Corroborating InformationTwo sources corroborate the above source. 
Yau and Webster (1987 [DIRS 100494], p. 717), an authoritative source, reported that zirconium 
resists crevice corrosion. In low-pH chloride solutions or chlorine gas, for example, zirconium is 
not subject to crevice attack. Greene et al. (2000 [DIRS 145073]) performed pitting and crevice 
corrosion tests on Zircaloy-4. They report that no crevice corrosion is observed under the same 
environment and electrochemical conditions that promote pitting corrosion on exposed surfaces. 
C.3.3 Qualification of Cappelaere, C.; Limon, R.; Bredel, T.; Herter, P.; Gilbon, D.; 
Bouffioux, P.; and Mardon, J-P. 2002. [DIRS 164195] 
Technical AssessmentBoth this analysis (Section 6.12.5) and the above source apply consistent 
conclusions regarding axial migration of hydrogen in the cladding. Both report minor migration. 
Personnel / Organization QualificationsThe authors work for CEA, which is a reputable 
source and an authoritative source for this type of information. 
Peer / Professional Review The source was peer-reviewed. 
Extent and Quality of Corroborating InformationCunningham et al. 1987 [DIRS 101591] 
corroborates this source's analysis of hydrogen migration. 
C.3.4 Qualification of Clayton, J.C. 1989 [DIRS 149208] 
Technical AssessmentBoth this analysis (Section 6.24) and the above source (Clayton, J.C. 
1989 [DIRS 149208], Tables 1 through 4, page 270) apply consistent conclusions regarding 
hydrogen absorption into zirconium. Both report that absorption only occurs under high 
hydrogen pressure. 
Personnel / Organization QualificationsThe experiments were performed at the Bettis Atomic 
Power Laboratory. 
Peer / Professional ReviewThe source was peer-reviewed, as stated in the publication. 
March 2004 C-8 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
C.3.5 Qualification of Cunningham et al. 1987 [DIRS 101591] 
Technical AssessmentBoth this analysis and the above source (page 5) apply consistent 
conclusions regarding axial migration of hydrogen in the cladding. Both report minor migration. 
Personnel / Organization QualificationsThe analysis was performed at Pacific Northwest 
Laboratory. 
Extent and Quality of Corroborating InformationCappelaere et al. 2002 [DIRS 164195] 
corroborates this source's analysis of hydrogen migration. 
C.3.6 Qualification of Edsinger 2000 [DIRS 154433] 
Technical AssessmentBoth this analysis and the above source apply consistent conclusions 
regarding hydriding cladding and the requirement for a reducing environment. 
Personnel / Organization QualificationsThe analysis was performed at Global Nuclear Fuel 
(previously known as General Electric Nuclear Fuel). 
Peer / Professional ReviewThe source was peer-reviewed. 
C.3.7 Qualification of Greene, C.A.; Brossia, C.S.; Dunn, D.S.; and Cragnolino, G.A. 
2000. [DIRS 145073] 
Technical AssessmentBoth this analysis (Section 6.7) and the above source apply consistent 
conclusions regarding localized (crevice) corrosion of the cladding. Both conclude that any 
crevice corrosion is not observed. 
Personnel / Organization QualificationsThe lead author works for the NRC. The other 
authors work for The Center for Nuclear Waste Regulatory Analyses, Southwest Research 
Institute. This work was sponsored by the NRC and performed under a QA program that is 
approved and audited by the NRC. 
Peer / Professional Review The source was peer-reviewed. 
Extent and Quality of Corroborating InformationTwo sources corroborate the above source. 
Yau and Webster (1987 [DIRS 100494], p. 717), an authoritative source, reported that zirconium 
resists crevice corrosion. In low-pH chloride solutions or chlorine gas, for example, zirconium is 
not subject to crevice attack. Brossia et al. (2002 [DIRS 161988]) performed pitting and crevice 
corrosion tests on Zircaloy-4. They report that no crevice corrosion is observed under the same 
environment and electrochemical conditions that promote pitting corrosion on exposed surfaces. 
C.3.8 Qualification of Hansson 1984 [DIRS 101676] 
Technical Assessment This analysis and the above source (page 6) assess the extent to which 
silica could affect cladding corrosion. The analysis considers high dissolved silica content in 
water, the source considers the silica content in concrete, and an authoritative source (Yau and 
Webster, 1987 [DIRS 100494]) considers the silica concentration in sea water. All three 
March 2004 C-9 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
anticipate little cladding degradation that can be attributed to silica concentrations that far exceed 
those that the cladding would encounter in the repository. 
Personnel / Organization QualificationsThis work was performed at the Danish Corrosion 
Center. 
Peer / Professional ReviewThis report was peer-reviewed within the SKB before publication. 
Extent and Quality of Corroborating Information An authoritative source, Yau and Webster, 
1987 [DIRS 100494] corroborates this source's assessment of cladding's resistance to silica 
corrosion. 
C.3.9 Qualification of Hayes et al. 1999 [DIRS 164598] 
Technical Assessment This analysis and the above source (Figures 2, 5, 6, 8, and 11) assess 
creep as a mechanism to fail Zircaloy cladding. The analysis does not expect creep at peak 
repository temperatures (< 268C), and the source expects little creep damage under 400oC. 
Personnel / Organization QualificationsThe analysis was performed at Lawrence Livermore 
National Laboratory. 
Extent and Quality of Corroborating InformationAn authoritative source, NRC 2002 [DIRS 
164593], corroborates this source's assessment of creep. Both state that cladding failure 
attributed to creep is unlikely if peak temperatures are below 400C. 
C.3.10 Qualification of Hillner et al. 1998 [DIRS 100455] 
Technical Assessment This analysis and the above source (page 11) do not expect microbes to 
cause cladding failure. Both consider the corrosion of Zircaloy-clad fuels under repository 
conditions and both support the concept that organic acids produced by microbes are unlikely to 
significantly accelerate zirconium corrosion. 
Personnel / Organization QualificationsThe experiments and analysis were performed at the 
Bettis Atomic Power Laboratory. 
Peer / Professional ReviewThe source was peer-reviewed within the Bettis organization before 
publication as a WAPD report. It was later accepted, peer-reviewed for publication by the 
Journal of Nuclear Material. 
Extent and Quality of Corroborating InformationThree independent sources corroborate 
Hillner et al. 1998. Wolfram et al. 1996 [DIRS 165268] found no evidence in the literature that 
zirconium or its alloys are susceptible to microbial-induced corrosion. Yau and Webster, 1987 
[DIRS 100494], an authoritative source, states that zirconium resists organic acids that microbes 
produce. And the Yucca Mountain Project (CRWMS M&O 2000b, [DIRS 151561]) evaluated 
in-drift microbial communities and concluded that they would have little effect on the in-drift 
geochemistry. 
March 2004 C-10 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
C.3.11 Qualification of Kreyns et al. 1996 [DIRS 100462] Figure 5 
Technical AssessmentHydride embrittlement could cause cladding failure. Test results from 
Kreyns et al. shows that the fracture toughness of the zirconium-alloyed cladding remains 
sufficiently high for hydrogen contents up to 4000 ppm that cladding failure is not expected. 
Personnel / Organization QualificationsThe experiments and analysis were performed at the 
Bettis Atomic Power Laboratory. 
Peer / Professional ReviewThe source was peer-reviewed within the Bettis organization before 
publication. 
Extent and Quality of Corroborating InformationThis reference is cited by the NRC in "NRC 
Issue Resolution Status Report on Container Life and Source Term, Revision 3" (Beckman 
2001 (DIRS 156122]). 
C.3.12 Qualification of Peehs 1998 [DIRS 109219] 
Technical AssessmentHelium produced in fuel rods could pressurize the cladding and cause it 
to fail. The Cladding Degradation Analysis and this source (pages 5 and 6) consider that the 
complete release of helium and the resulting pressurization are highly conservative. Piron and 
Pelletier 2001 [DIRS 165318], who are discussed in Section C.3.13 of this appendix, show that 
after 10,000 years at repository temperatures, the peak pressure would not cause the cladding to 
fail. These results assume complete release of the helium. This source shows that fission gas 
release during dry storage is not expected because the diffusion coefficients decrease by 
approximately 8 orders as the fuel cools from reactor temperatures. Because repository 
temperatures are well below dry storage temperatures, assuming complete release of helium is 
highly conservative. 
Personnel / Organization QualificationsThis work was performed at Siemens KWU-NBT, a 
nuclear vendor. 
Extent and Quality of Corroborating InformationManaktala 1993 [DIRS 101719] and 
Rothman 1984 [DIRS 100417] corroborate this source's assertion that complete release of helium 
is conservative. Both found that helium pressure buildup would be offset by the cooling of the 
rods. Consequently, the helium pressure that Cladding Degradation Analysis assumes at 10,000 
years is conservative. 
C.3.13 Qualification of Piron and Pelletier 2001 [DIRS 165318] 
Technical AssessmentThis analysis and the above source consider whether helium could 
pressurize the cladding and cause it to fail. This source indicates that after 10,000 years at 
repository temperatures, the peak pressure would reach 13.3 MPa, but pressure would have to 
reach about 33 MPa for the cladding to fail from delayed hydride cracking. Both conclude that 
helium production at peak temperatures would not elevate pressure in the fuel rod to a level that 
would cause cladding failure. 
March 2004 C-11 ANL-WIS-MD-000008 REV 01 Clad Degradation  FEPs Screening Arguments 
Personnel / Organization QualificationsThe French Atomic Energy Commission (CEA) is a 
reputable and an authoritative source for this type of information. 
Peer / Professional ReviewThe source was peer-reviewed. 
Extent and Quality of Corroborating InformationManaktala 1993 [DIRS 101719] and a 
Rothman 1984 [DIRS 100417] corroborate this source's assertion that helium production would 
not fail the cladding. Both found that helium pressure buildup would be offset by the cooling of 
the rods. Consequently, the pressure at 10,000 years would be less than what this source 
predicted, and the likelihood of cladding failure would be even more remote. 
C.3.14 Qualification of Wolfram et al. 1996 [DIRS 165268] 
Technical AssessmentThis analysis and the above source (pages iii and iv) do not expect 
microbes to cause cladding failure. Both found no evidence in the literature that zirconium or its 
alloys are susceptible to microbial-induced corrosion. 
Personnel / Organization QualificationsThis work was performed at the Idaho National 
Engineering Laboratory. 
Extent and Quality of Corroborating InformationThree independent sources corroborate 
Wolfram et al. 1996 [DIRS 165268]. Hillner et al. 1998 [DIRS 100455], supports the concept 
that organic acids produced by microbes are unlikely to significantly accelerate zirconium 
corrosion. Yau and Webster, 1987 [DIRS 100494], an authoritative source, states that zirconium 
resists organic acids that microbes produce. And the Yucca Mountain Project (CRWMS M&O 
2000b, [DIRS 151561]) evaluated in-drift microbial communities and concluded that they would 
have little effect on the in-drift geochemistry. 
C.4 CONCLUSION 
In conclusion, the data or technical information is qualified for its intended use. 
March 2004 C-12 ANL-WIS-MD-000008 REV 01