Clad Degradation  FEPs Screening Arguments Rev 00, ICN 00

ANL-WIS-MD-000008

April 2000

1. PURPOSE 
The purpose of this report is to summarize the work of others into the Features, Events, and 
Processes (FEPs) for cladding degradation screening that is consistent with and used in the Total 
System Performance Assessment - Site Recommendation (TSPA-SR). This review is to describe 
what FEPs are to be included in the TSPA-SR and also document the reason for excluding 
others. In accordance with AP-2.13Q, Technical Product Development Planning, a work plan 
(CRWMS M&O 1999a) was developed, issued, and utilized in the preparation of this document. 
Under the provisions of the U.S. Department of Energys (DOEs) Revised Interim Guidance 
Pending Issuance of New U. S. Nuclear Regulatory Commission (NRC) Regulations (Revision 
01, July 22, 1999), for Yucca Mountain Nevada (Dyer 1999; and herein referred to as DOEs 
interim guidance), the DOE must provide a reasonable assurance that the regulatory-specified 
performance objectives for the Yucca Mountain Project can be achieved for a 10,000-year postclosure 
period. This assurance must be demonstrated in the form of a performance assessment 
that: (1) identifies the features, events, and processes (FEPs) that potentially affect the 
performance of the geologic repository; (2) examines the effects of such FEPs on the 
performance of the geologic repository; (3) estimates the expected annual dose to a specified 
nearby population group. The performance assessment must also provide the technical basis for 
inclusion or exclusion of specific FEPs. 
Although the NRC has not defined nor used the term "scenario" in the pertinent regulations, the 
Yucca Mountain Total System Performance Assessment - Site Recommendation(TSPA-SR) has 
chosen to satisfy the above-stated performance assessment requirements by adopting a scenario 
development process. This decision was made based on the Yucca Mountain TSPA-SR adopting 
a definition of "scenario" as not being limited to a single, deterministic feature of the system, but 
rather as a set of similar features that share common FEPs. The DOE has chosen to adopt a 
scenario development process based on the methodology developed by Cranwell et al. (1990) for 
the NRC. The first step of this process is the identification of FEPs potentially relevant to the 
performance of the Yucca Mountain repository; the second step includes the screening of each 
FEP. 
The primary purpose of this Analysis/Model Report (AMR) is to identify and document the 
analysis, screening decision, and Total System Performance Assessment (TSPA-SR) disposition 
or screening argument for FEPs related to clad degradation. Both primary and secondary FEPs 
are considered. The screening decisions, and associated TSPA-SR disposition or screening 
argument, for the subject FEPs will be cataloged separately in a project-specific FEPs database 
(see Section 1.4). This AMR and the database are being used to document information related to 
the FEPs screening decisions and associated screening argument and to assist reviewers during 
the license review process. 
Most commercial nuclear fuel is encased in Zircaloy cladding. This AMR does not address 
potential damage to assemblies that might occur at the YMP surface facilities. There are 
constraints, caveats and limitations to this report. This FEPs screening is based on commercial 
Pressurized Water Reactor (PWR) fuel with Zircaloy cladding but is applicable to Boiling Water 
Reactor (BWR) fuel. This analysis referenced in the other AMRs is also limited to fuel exposed 
to normal operation and anticipated operational occurrences (i.e. events which 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, oxide coating thickness, and rod plenum 
pressures. Ranges and uncertainties have been defined in the other AMRs. The information 
provided in this FEPs screening will be used in evaluating what cladding degradation 
mechanisms are included in the post-closure performance of the Monitored Geologic Repository 
(MGR) in relation to waste form degradation. 
1.1 SCOPE 
This AMR has been prepared to satisfy the FEP screening documentation requirements in the 
Work Scope/Objectives/Tasks section of the development plan entitled Cladding FEPs 
Screening Arguments (CRWMS M&O 1999a). 
The current FEPs list consists of 1786 entries (as described in Section 1.2). The FEPs have been 
classified as primary and secondary FEPs (as described in Section 1.2) and have been assigned to 
various Process Modeling Reports (PMRs). The assignments were based on the nature of the 
FEPs so that the analysis and resolution for screening decisions reside with the subject-matter 
experts in the relevant disciplines. The resolution of other than EBS FEPs is documented in 
AMRs prepared by the responsible PMR groups. Several relevant FEPs do not fit into the 
existing PMR structure. Criticality is the largest example, and is treated in FEP assignments as if 
were a separate PMR. Some FEPs were best assigned to the TSPA-SR itself (i.e., system-level 
FEPs), rather than to its component models. 
This AMR addresses the clad degradation FEPs. These FEPs represent the key features that 
result in degradation of the cladding. The clad degradation primary and secondary EBS FEPs 
addressed in this AMR are provided in Table 1.

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Table 1. FEPs Related to the CSNF Clad Degradation Component 
FEP Number FEP Title 
2.1.02.07.00 Gap and Grain Release of Cs, I 
2.1.02.11.00 Waterlogged Rods 
2.1.02.12.00 Cladding Degradation Before YMP Receives It 
2.1.02.12.01 Pin Degradation During Reactor Operation 
2.1.02.12.02 Pin Degradation During Spent Fuel Pool Storage 
2.1.02.12.03 Pin Degradation During Dry Storage 
2.1.02.12.04 Pin Degradation During Fuel Shipment and Handling 
2.1.02.13.00 General Corrosion of Cladding 
2.1.02.13.01 Cladding Degradation Mechanisms At YMP, Pre-Pin Failure 
2.1.02.13.02 Corrosion (of Cladding) 
2.1.02.14.00 Microbiologically Influenced Corrosion (MIC) of Cladding 
2.1.02.15.00 Acid Corrosion of Cladding From Radiolysis 
2.1.02.16.00 Localized (Pitting) Corrosion of Cladding 
2.1.02.17.00 Localized Corrosion (Crevice Corrosion) of Cladding 
2.1.02.18.00 High Dissolved Silica Content of Waters Enhances Corrosion of Cladding 
2.1.02.19.00 Creep Rupture of Cladding 
2.1.02.20.00 Pressurization From He Production Causes Cladding Failure 
2.1.02.21.00 Stress Corrosion Cracking (SCC) of Cladding 
2.1.02.21.01 Inside Out From Fission Products (Iodine) (Failure of Cladding) 
2.1.02.21.02 Outside In From Salts Or WP Chemicals (Failure of Cladding) 
2.1.02.22.00 Hydride Embrittlement of Cladding 
2.1.02.22.01 Hydride Embrittlement From Zirconium Corrosion (of Cladding) 
2.1.02.22.02 Hydride Embrittlement From WP Corrosion & H2 Absorption (of Cladding) 
2.1.02.22.03 Hydride Embrittlement From Galvanic Corrosion of WP Contacting Cladding 
2.1.02.22.04 Delayed Hydride Cracking (of Cladding) 
2.1.02.22.05 Hydride Reorientation (of Cladding) 
2.1.02.22.06 Hydrogen Axial Migration (of Cladding) 
2.1.02.22.07 Hydride Embrittlement From Fuel Reaction (Causes Failure of Cladding) 
2.1.02.23.00 Cladding Unzipping 
2.1.02.23.02 Dry Oxidation of Fuel (Causes Failure of Cladding) 
2.1.02.23.03 Wet Oxidation of Fuel (Causes Failure of Cladding) 
2.1.02.24.00 Mechanical Failure of Cladding 
2.1.02.27.00 Localized Corrosion Perforation from Fluoride 
2.1.02.28.00 Diffusion-Controlled Cavity Growth 
2.1.7.01.00 Rockfall (Large Block) 
2.1.09.03.00 Volume Increase of Corrosion Products 
2.1.11.07.00 Thermally-Induced Stress Changes in Waste and EBS 
1.2 FEPs IDENTIFICATION AND ANALYSIS 
For the YMP TSPA-SR, a scenario is defined as a subset of the set of all possible features of the 
fuel cladding that contains the features resulting from a specific combination of FEPs. The first 
step of the scenario development process is the identification of FEPs potentially relevant to the 
performance of the Yucca Mountain repository. 
The initial set of FEPs was created for the Yucca Mountain TSPA-SR by combining lists of 
FEPs previously identified as relevant to the Yucca Mountain Project (YMP) (e.g. Total System 
Performance Assessment-1995: An Evaluation of the Potential Yucca Mountain Repository, 
CRWMS M&O 1995a) with a draft FEP list compiled by the Nuclear Energy Agency (NEA) of

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the Organization for Economic Cooperation and Development (OECD). The NEA list is 
maintained as an electronic FEP database and is the most comprehensive list available 
internationally. The list currently contains 1261 FEPs from Canadian, Swiss, and Swedish spentfuel 
programs, intermediate and low-level waste programs of the U.K., and the U.S. Waste 
Isolation Pilot Plant (WIPP) program. An additional 292 FEPs have been identified from YMP 
literature and site studies, and 82 FEPs have been identified during YMP project staff 
workshops. These FEPs are organized under 151 categories, based on NEA category headings, 
resulting in a total of 1786 entries. Consistent with the diverse backgrounds of the programs 
contributing FEPs lists, FEPs have been identified by a variety of methods, including expert 
judgment, informal elicitation, event tree analysis, stakeholder review, and regulatory stipulation. 
All potentially relevant FEPs have been included, regardless of origin. This approach has led to 
considerable redundancy in the FEP list, because the same FEPs are frequently identified by 
multiple sources, but it also ensures that a comprehensive review of narrowly defined FEPs will 
be performed. The FEPs list is considered open and will continue to grow as additional FEPs are 
identified. 
There is no uniquely correct level of detail at which to define scenarios or FEPs. Decisions 
regarding the appropriate level of resolution for the analysis are made based on consideration of 
the importance of the scenario in its effect on overall performance and the resolution desired in 
the results. The number and breadth of scenarios depend on the resolution at which the FEPs 
have been defined: coarsely defined FEPs result in fewer, broad scenarios, whereas narrowly 
defined FEPs result in many narrow scenarios. For efficiency, both FEPs and scenarios should 
be aggregated at the coarsest level at which a technically sound argument can be made that is 
adequate for the purposes of the analysis. 
Consequently, each FEP has been identified as either a primary or secondary FEP. Primary 
FEPs are those FEPs for which the project proposes to develop detailed screening arguments. 
The classification and description of primary FEPs strive to capture the essence of all the 
secondary FEPs that map to the primary. For example, the primary FEP "Cladding Degradation 
Before YMP Receives It" can be used appropriately to resolve multiple and redundant secondary 
FEPs that address degradation during reactor operation, during spent fuel pool storage, etc. By 
working to the primary FEP description, the subject matter experts assigned to the primary FEP 
address all relevant secondary FEPs, and arguments for secondary FEPs can be rolled into the 
primary FEP analysis. Secondary FEPs are FEPs either that are completely redundant or that can 
be aggregated into a single primary FEP. 
To perform the screening and analysis, the FEPs have been assigned so that the analysis, 
screening decision, and TSPA-SR disposition reside with the subject matter experts in the 
relevant disciplines. The TSPA-SR recognizes that FEPs have the potential to affect multiple 
facets of the project, may be relevant to more than one PMR, or may not fit neatly within the 
PMR structure. For example, many FEPs affect waste form (WF), waste package (WP), and the 
EBS. Rather than create multiple separate FEPs, the FEPs have been assigned, as applicable, to 
one or more process modeling groups, which are responsible for the AMRs. 
At least two approaches have been used to resolve overlap and interface problems of multiple 
numbers for the same FEP. FEP owners from different process modeling groups may decide that 
only one PMR will address all aspects of the FEP, including those relevant to other PMRs.

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Alternatively, FEP owners may each address only those aspects of the FEP relevant to their area. 
In either case, the FEP AMR produced by each process modeling group lists the FEP and 
summarizes the screening result, citing the appropriate work in related AMRs as needed. 
1.3 FEPs SCREENING AND ANALYSIS PROCESS 
As described in Section 1.2, the first step in the scenario development process was the 
identification and analysis of FEPs. The second step in the scenario development process 
includes the screening of each FEP. Each FEP is screened for inclusion or exclusion in the 
TSPA-SR against three criteria, which are stated as regulatory requirements in NRCs proposed 
rule 10 CFR Part 63 (64 FR 8640), and in the U.S. Environmental Protection Agency's (EPA) 
proposed rule 40 CFR Part 197 (64 FR 46976). The screening criteria are discussed in more 
detail in Section 4.2 and are summarized here. FEPs are excluded from the TSPA-SR only if: 
 They are specifically ruled out by regulation, are contrary to the stated regulatory 
assumptions, or are in conflict with statements made in background information 
regarding intent or directions of the regulations. 
 They can be shown to have a probability of occurrence of less than 10-4 in 104 years. 
 Their occurrence can be shown to have no significant effect on the overall TSPA. 
The regulatory screening criteria contained in DOEs interim guidance (Dyer 1999) and in the 
proposed 40 CFR Part 197 (64 FR 46976) are relevant to many of the FEPs. FEPs that are 
contrary to DOEs interim guidance or to specific proposed regulations, regulatory assumptions, 
or regulatory intent are excluded from further consideration. Examples include: the explicit 
exclusion of all but a specified scenario to address treatment of human intrusion (10 CFR 
63.113(d)), assumptions about the critical group to be considered in the dose assessment (10 
CFR 63.115), and the intent that the consideration of "the human intruders" be excluded from 
the human intrusion assessment (64 FR 8640, Section XI. Human Intrusion). 
Probability estimates used in the FEPs screening process may be based on technical analysis of 
the past frequency of similar events (such as igneous and seismic events) or, in some cases, on 
expert elicitation. Probability arguments, in general, require including some information about 
the magnitude of the event in its definition. Probability arguments are also sensitive to the 
spatial and temporal scales at which FEPs are defined. For example, the definition of the 
probability of a seismic event depends on the magnitude of the event. Probability arguments are 
therefore made at reasonably coarse scales. 
Consequence-based screening arguments can be established in a variety of ways. Various 
methods include TSPA-SR sensitivity analyses, modeling studies outside of the TSPA-SR, or 
reasoned arguments based on literature research. For example, consequences of many 
geomorphic processes such as erosion and sedimentation can be evaluated by considering 
bounding rates reported in geologic literature. More complicated processes, such as igneous 
activity, require detailed analyses conducted specifically for the Yucca Mountain Project. Lowconsequence 
arguments are often made by demonstrating that a particular FEP has no effect on 
the distribution of an intermediate performance measure in the TSPA-SR. For example, by

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demonstrating that including a particular waste form has no effect on the concentrations of 
radionuclides transported from the repository in the aqueous phase, it is also demonstrated that 
including this waste form in the inventory would not compromise compliance with the 
performance objectives. Explicit modeling of the characteristics of this waste form could 
therefore be excluded from the TSPA-SR. 
Using the type of arguments discussed above, each FEP identified as relevant to the clad 
degradation was reviewed against the three exclusion criteria. Those that were determined to 
meet one of the three criteria were designated as excluded from further consideration within 
the TSPA-SR. Those that did not meet any of these criteria must, by definition, be included. 
1.4 ORGANIZATION OF FEP DATABASE 
Under a separate scope, the TSPA-SR team is constructing an electronic FEP database to assist 
project reviewers during the license review process. Each FEP has been entered as a separate 
record in the database. Fields within each record provide a unique identification number, a 
description of the FEP, the origin of the FEP, identification as a primary or secondary FEP for 
the purposes of the TSPA-SR, and mapping to related FEPs and to the assigned PMRs. Fields 
also provide summaries of the screening arguments with references to supporting documentation 
and AMRs, and, for all included or retained FEPs, statements of the disposition of the FEP 
within the TSPA-SR modeling system. The AMRs, however, contain the detailed arguments and 
description of the disposition of the subject FEPs. 
Alphanumeric identifiers (called the NEA category) previously used have been retained in the 
database for traceability purposes. Each FEP has also been assigned a unique YMP FEP 
database number, based on the NEA categories. The database number is the primary method for 
identifying FEPs, and consists of an eight-digit number of the form x.y.zz.pp.qq. The general 
structure of the database is reflected in the first two digits (x.y) as shown below: 
0.0. Assessment Basis 
1.0. External Factors 
1.1 Repository Issues 
1.2 Geological Processes and Effects 
1.3 Climatic Processes and Effects 
1.4 Feature Human Actions (Active) 
1.5 Other 
2.0. Disposal System - Environmental Factors 
2.1 Wastes and Engineered Features 
2.2 Geologic Environment 
2.3 Surface Environment 
2.4 Human Behavior 
3.0. Disposal System - Radionuclide/Contaminant Factors 
3.1 Contaminant Characteristics 
3.2 Contaminant Release/Migration Factors 
3.3 Exposure Factors

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The next six digits (zz.pp.qq) define a grouping structure for the FEPs, with zz designating the 
category, and pp designating the heading. The exact details of this grouping structure are not 
important to the evaluation, since each FEP will be evaluated regardless of the database 
organization. Finally, the last two digits (qq) signify whether the FEP is primary (00) or 
secondary (other than 00). Each heading has a primary FEP associated with it, and may or may 
not have any secondary FEPs. In those cases where secondary FEPs do exist, the primary FEP 
encompasses all the issues associated with the secondary FEPs. The secondary FEPs either 
provide additional detail concerning the primary, or are a restatement of the primary based on 
redundant input from a different source. 
2. QUALITY ASSURANCE 
The activities documented in this AMR were evaluated in accordance with QAP-2-0, Conduct of 
Activities and were determined to be quality affecting and subject to the requirements of the U.S. 
DOE Office of Civilian Radioactive Waste Management (OCRWM) Quality Assurance 
Requirements and Description (QARD) (DOE 2000). This evaluation is documented in Conduct 
of Performance Assessment (CRWMS M&O 1999c). Accordingly, analysis activities 
documented in this AMR have been conducted in accordance with the Civilian Radioactive 
Waste Management System Management and Operating Contractor (CRWMS M&O) quality 
assurance program, using approved procedures identified in Cladding FEPs Screening 
Arguments (CRWMS M&O 1999a). 
This AMR has been developed in accordance with procedure AP-3.10Q, Analyses and Models. 
Preparation of this analysis did not require the classification of items in accordance with QAP-2- 
3, Classification of Permanent Items. This activity is not a field activity. Therefore, an 
evaluation in accordance with NLP-2-0, Determination of Importance Evaluations was not 
required. 
3. COMPUTER SOFTWARE AND MODEL USAGE 
This AMR uses no computational software nor model. The AMR was documented using only 
commercially available software (Microsoft Word 97-SR2) for word processing, which is 
exempt from qualification requirements in accordance with AP-SI.1Q, Software Management. 
There were no additional applications (routines or macros) developed using this commercial 
software. The analyses and arguments presented herein are based on regulatory requirements, 
results of analyses presented and documented in other AMRs, or technical literature. 
4. INPUTS 
There are no input data sources used in this analysis. 
4.1 Data and Parameters 
This section is not applicable to this analysis.

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4.2 Criteria 
Programmatic requirements for this document are listed in the development plan (CRWMS 
M&O 1999a), which covers eight tasks needed to provide the submodels for the performance 
assessment Physical and Chemical Environment Abstraction Model. The development plan 
specifies that this document and all analyses described herein must adhere to the requirements of 
AP-3.10Q, Analyses and Models, and must address applicable NRC issue resolution status report 
(IRSR) criteria for the TSPA (NRC 1998). 
The U.S. Nuclear Regulatory Commissions Total System Performance Assessment and 
Integration (TSPAI) Issue Resolution Status Report (NRC 1998) establishes generic technical 
acceptance criteria considered by the NRC staff to be essential to a defensible, transparent, and 
comprehensive assessment methodology for the repository system. The NRCs IRSR Key 
Technical Issue: Container Life and Source Term (NRC 1999) establishes generic technical 
acceptance criteria considered by the NRC staff for the waste form, with the cladding 
degradation analysis being part of this KTI. Attachment IV of Initial Cladding Condition 
(CRWMS M&O 2000b) describes how IRSR issues and criteria are addressed. 
Technical screening criteria are provided in DOEs interim guidance (Dyer 1999) and have also 
been identified by the NRC in the proposed 10 CFR Part 63 (64 FR 8640) and by the EPA in the 
proposed 40 CFR Part 197 (64 FR 46976). Both proposed regulations specifically allow the 
exclusion of FEPs from the TSPA-SR if they are of low probability (less than one chance in 
10,000 of occurring in 10,000 years) or if occurrence of the FEP can be shown to have no 
significant effect on expected annual dose. There is no quantified definition of significant 
effect in the guidance nor proposed regulations. 
4.2.1 Low Probability 
The probability criterion is explicitly stated by the NRC in the proposed 10 CFR 63.114 (d): 
Consider only events that have at least one chance in 10,000 of occurring over 10,000 
years. 
The EPA provides essentially the same criterion in 40 CFR 197.40: 
The DOEs performance assessments should not include consideration of processes nor 
events that are estimated to have less than one chance in 10,000 of occurring within 
10,000 years of disposal. 
4.2.2 Low Consequence 
Criteria for low consequence screening arguments are provided in DOEs interim guidance (Dyer 
1999, Section 114(e) and (f)), which indicates that performance assessment shall: 
(e) Provide the technical basis for either inclusion or exclusion of specific features, 
events, and processes of the geologic setting in the performance assessment. 
Specific features, events, and processes of the geologic setting must be evaluated in

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detail if the magnitude and time of the resulting expected annual dose 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 
expected annual dose would be significantly changed by their omission. 
The EPA provides essentially the same criteria in 40 CFR 197.40: 
with the NRCs approval, the DOEs performance assessment need not evaluate, in 
detail, the impacts resulting from any processes and events or sequences of processes and 
events with a higher chance of occurrence if the results of the performance assessment 
would not be changed significantly. 
The terms [not] significantly changed and not .. changed significantly are undefined terms in 
DOEs interim guidance and in the EPAs proposed regulations. These terms are inferred for 
FEPs screening purposes to be equivalent to having no or negligible effect. Because the relevant 
performance measures differ for different FEPs (e.g., effects on performance can be measured in 
terms of changes in concentrations, flow rates, travel times, and other measures, as well as 
overall expected annual dose), there is no single quantitative test of significance. 
4.2.3 Reference Biosphere 
Both DOEs interim guidance (Dyer 1999) and EPAs proposed regulations specify assumptions 
(which in effect serve as criteria) pertinent to screening many of the EBS FEPs. Particularly 
germane are explicit assumptions regarding the reference biosphere (10 CFR 63.115), and less 
so are assumptions regarding the location and use of groundwater by the critical group used for 
calculation of exposure doses. 
The assumptions pertaining to the characteristics of the reference biosphere are presented in 
DOEs interim guidance (Dyer 1999, Section 115(a)1). The specified characteristics pertinent to 
the EBS FEPs are that: 
 Features, events, and processes shall be consistent with present knowledge of the 
conditions in the region surrounding the Yucca Mountain site. 
The EPA has specified a similar assumption in proposed 40 CFR 197.15. This assumption is 
stated as: 
 DOE must vary factors related to the geology, hydrology, and climate based on 
environmentally protective but reasonable scientific predictions of the changes that could 
affect the Yucca Mountain disposal system over the next 10,000 years.

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4.3 Codes and Standards 
ASTM C1174-97, Standard Practice for the Long-Term Behavior of Materials, Including Waste 
Forms, Used in Engineered Barrier Systems (EBS) for Geologic Disposal of High-Level 
Radioactive Waste is used to support the degradation analysis development methodology, 
categorize the analyses developed with respect to their usage for long-term TSPA-SR, and relate 
the information/data used to develop the analysis to the requirements of the standard. 
Material standards are also cited. ASTM B 551 and ASTM B811 are referenced in Section 
6.2.12. 
5. ASSUMPTIONS 
This AMR makes no assumptions beyond those reported and documented by the references. 
6. ANALYSIS/MODEL - FEPs RELATED TO CLADDING DEGRADATION 
To demonstrate that the regulatory-specified performance objectives of proposed 10 CFR Part 63 
and proposed 40 CFR Part 197 can be achieved for a 10,000-year post-closure period, the Yucca 
Mountain Project is implementing a stochastic scenario development methodology based on the 
work of Cranwell et al. (1990). The methodology provides a systematic approach for 
considering, as completely as practicable, the possible feature states of a repository system. It 
seeks to span the set of all possible feature states using a finite set of scenarios. Each scenario 
represents the ensemble of possible feature states corresponding to parameter and model 
uncertainties present in the group of FEPs composing the scenario. The methodology begins 
with a comprehensive FEP identification step followed by a rigorous FEP screening step. With 
its focus on FEPs related to cladding degradation, this Analysis/Modeling Report (AMR) 
considers these first two steps (i.e., FEP identification and FEP screening) of scenario 
development. 
6.1 Identification of FEPs 
The first step of the scenario development methodology is the identification of FEPs potentially 
relevant to the performance of the Yucca Mountain repository. This process is summarized in 
Section 1.2. 
Each FEP has been identified as either a primary or a secondary FEP. Primary FEPs are 
those FEPs for which the project proposes to develop detailed screening arguments. The 
classification and description of primary FEPs strive to capture the essence of each secondary 
FEP that maps to the primary FEP. For example, the primary FEP Meteorite Impact can be used 
appropriately to resolve multiple and redundant secondary FEPs that address size and effects of 
meteorite impacts. By working to the primary FEP description, the subject matter experts 
assigned to the primary FEP will address each relevant secondary FEP, and arguments for 
secondary FEPs will be rolled into the primary FEP analysis. 
Table 2 identifies both the primary and the secondary FEPs related to cladding degradation.

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6.2 Screening of FEPs 
The second step in the scenario development methodology includes the screening of each FEP. 
Each FEP is screened for inclusion in or exclusion from the TSPA-SR against three criteria, 
which are stated as regulatory requirements in the NRC proposed rule 10 CFR Part 63 and in 
EPA proposed rule 40 CFR Part 197. FEPs are excluded from the TSPA-SR for one of the 
following reasons: 
(1) They are specifically ruled out by regulation, are contrary to the stated regulatory 
assumptions, or are in conflict with statements made in background information regarding 
intent or directions of the regulations. 
(2) They can be shown to have a probability of occurrence of less than 10-4 in 104 years. 
(3) Their occurrence can be shown to have no significant effect on the overall TSPA. 
Examples of reason (1) are the explicit exclusion by the National Academy of Science from 
consideration of all but a specific scenario to address treatment of human intrusion (10 CFR 
63.113(d)) and assumptions about the critical group to be considered in the dose assessment (10 
CFR 63.115). 
Probability estimates used in the FEP screening process in relation to reason (2) may be based on 
technical analysis of the past frequency of similar events (such as seismic events), or, in some 
cases, on expert elicitation. Probability arguments, in general, require including appropriate 
information about the magnitude of the event in its definition. Probability arguments are also 
sensitive to the spatial and temporal scales at which FEPs are defined. For example, definition of 
the probability of meteorite impact depends on the size of the meteorite of interest and must 
consider that meteorite impacts are less likely in shorter time intervals and at smaller locations. 
Probability arguments are therefore made at reasonably coarse scales based on the judgment of 
the originator. 
Consequence based screening arguments in relation to reason (3) can be established in a variety 
of ways. Various methods include TSPA-SR sensitivity analyses, modeling studies outside of 
the TSPA-SR, or reasoned arguments based on literature research. For example, consequences of 
many geomorphic processes such as erosion and sedimentation can be evaluated by considering 
bounding rates reported in geologic literature. More complicated processes, such as criticality, 
require detailed analyses conducted specifically for the Yucca Mountain Project. Low 
consequence arguments are often made by demonstrating that a particular FEP has no effect on 
the distribution of an intermediate performance measure of the TSPA-SR. For example, to 
demonstrate that including a particular waste form does not compromise compliance with 
performance objectives, one may show that aqueous phase concentrations of radionuclides 
transported from the repository would be unaffected by the exclusion of the specific waste form. 
Explicit modeling of the characteristics of this waste form could then be excluded from the 
TSPA-SR. 
Table 2 presents the screening decision as either include or exclude for FEPs related to 
cladding degradation. However, for those cases in which there are secondary FEPs, one should 
be aware of the rule for assigning a composite decision to the primary FEP. Specifically, a 
primary FEP is designated as include if at least one of its secondary FEPs is designated as

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include. Conversely, a primary FEP is designated as exclude only if all of its secondary FEPs are 
also designated as exclude. Section 6.2 provides detailed documentation of both the screening 
argument and the TSPA-SR disposition for each of the primary FEPs. Table 2 lists the specific 
subsection where each FEP is discussed. If the primary FEP has no secondary FEPs, then the 
decision to include or exclude the primary FEP is made for that FEP itself. 
Also provided in each FEP section is a cross reference to key technical issues identified by the 
NRC as being important for the Yucca Mountain repository. These are identified as Issue 
Resolution Status Report (IRSR) issues. The key technical issues and subissues are listed below. 
The relevance of these subissues to the EBS FEPs is identified in Sections 6.2.1 through 6.2.15. 
Whenever the key technical issue (CLST, for example) is identified rather than a specific 
subissue, all subissues apply. All the FEPs discussed in this AMP fall into the container life and 
source term or the Total System Performance Assessment and Integration (TSPAI) subissues 
which are listed below: 
Container Life and Source Term (CLST) 
CLST1 The effects of corrosion processes on the lifetime of the containers 
CLST2 The effects of phase stability of materials and initial defects on the mechanical 
failure and lifetime of the containers 
CLST3 The rate at which radionuclides in spent nuclear fuel (SNF) are released from the 
EBS through the oxidation and dissolution of spent fuel 
CLST4 The rate at which radionuclides in high-level waste (HLW) glass are leached and 
released from the EBS 
CLST5 The effect of in-package criticality on WP and EBS performance 
CLST6 The effect of alternate EBS design features on container lifetime and radionuclide 
release from the EBS 
Total System Performance Assessment and Integration (TSPAI) 
TSPAI1 Demonstration of the overall performance objective 
TSPAI2 Demonstration of multiple barriers 
TSPAI3 Model abstraction 
TSPAI4 Scenario analysis 
TSPAI5 Transparency and traceability of the analysis

ANL-WIS-MD-000008 REV 00 19 April 2000 
Table 2. Screening Decisions for FEPS Related to Cladding Degradation 
YMP FEP 
Database 
Number 
NEA 
Category 
FEP Name Primary (P) 
or 
Secondary (S) 
FEP 
Screening 
Decision 
Section 
2.1.02.07.00 N/A Gap and Grain Release of Cs, I P Include 6.2.13.4 
2.1.02.11.00 WF-3 Waterlogged Rods P Exclude 6.2.1 
2.1.02.12.00 2.1.03ca Cladding Degradation Before YMP Receives 
It 
P Include 6.2.2 
2.1.02.12.01 2.1.03cb Pin Degradation During Reactor Operation S Include 6.2.2.1 
2.1.02.12.02 2.1.03cc Pin Degradation During Spent Fuel Pool 
Storage 
S Exclude 6.2.2.2 
2.1.02.12.03 2.1.03cd Pin Degradation During Dry Storage S Include 6.2.2.3 
2.1.02.12.04 2.1.03ce Pin Degradation During Fuel Shipment and 
Handling 
S Include 6.2.2.4 
2.1.02.13.00 2.1.03ch General Corrosion of Cladding P Exclude 6.2.3 
2.1.02.14.00 2.1.03ci Microbiologically Influenced Corrosion (MIC) 
of Cladding 
P Exclude 6.2.4 
2.1.02.15.00 2.1.03cj Acid Corrosion of Cladding From Radiolysis P Exclude 6.2.5 
2.1.02.16.00 2.1.03ck Localized Corrosion (Pitting) of Cladding P Exclude 6.2.6 
2.1.02.17.00 2.1.03cl Localized Corrosion (Crevice Corrosion) of 
Cladding 
P Exclude 6.2.7 
2.1.02.18.00 2.1.09f High Dissolved Silica Content of Waters 
Enhances Corrosion of Cladding 
P Exclude 6.2.8 
2.1.02.19.00 2.1.03cm Creep Rupture of Cladding P Include 6.2.9 
2.1.02.20.00 2.1.03cn Pressurization From He Production Causes 
Cladding Failure 
P Include 6.2.10 
2.1.02.21.00 2.1.03cp Stress Corrosion Cracking (SCC) of 
Cladding 
P Include 6.2.11 
2.1.02.21.01 2.1.03cq Inside Out From Fission Products (Iodine) 
(Failure of Cladding) 
S Include 6.2.11 
2.1.02.21.02 2.1.03cr Outside In From Salts Or WP Chemicals 
(Failure of Cladding) 
S Include 6.2.11 
2.1.02.22.00 2.1.03cs Hydride Embrittlement of Cladding P Exclude 6.2.12 
2.1.02.22.01 2.1.03ct Hydride Embrittlement From Zirconium 
Corrosion (of Cladding) 
S Exclude 6.2.12.1 
2.1.02.22.02 2.1.03cu Hydride Embrittlement From WP Corrosion 
& H2 Absorption (of Cladding) 
S Exclude 6.2.12.2

ANL-WIS-MD-000008 REV 00 20 April 2000 
YMP FEP 
Database 
Number 
NEA 
Category 
FEP Name Primary (P) 
or 
Secondary (S) 
FEP 
Screening 
Decision 
Section 
2.1.02.22.03 2.1.03cv Hydride Embrittlement From Galvanic 
Corrosion of WP Contacting Cladding 
S Exclude 6.2.12.3 
2.1.02.22.04 2.1.03cw Delayed Hydride Cracking (of Cladding) S Exclude 6.2.12.4 
2.1.02.22.05 2.1.03cx Hydride Reorientation (of Cladding) S Exclude 6.2.12.5 
2.1.02.22.06 2.1.03cy Hydrogen Axial Migration (of Cladding) S Exclude 6.2.12.6 
2.1.02.22.07 2.1.03dd Hydride Embrittlement From Fuel Reaction 
(Causes Failure of Cladding) 
S Exclude 6.2.12.7 
2.1.02.23.00 2.1.03da Cladding Unzipping P Include 6.2.13 
2.1.02.23.02 2.1.03db Dry Oxidation of Fuel (Causes Failure of 
Cladding) 
S Exclude 6.2.13.1 
2.1.02.23.03 2.1.03dc Wet Oxidation of Fuel (Causes Failure of 
Cladding) 
S Include 6.2.13.2 
2.1.02.24.00 2.1.03co Mechanical Failure of Cladding P Include 6.2.14 
2.1.02.27.00 N/A Localized Corrosion Perforation from 
Fluoride 
P Include 6.2.15 
2.1.02.28.00 N/A Diffusion-Controlled Cavity Growth P Exclude 6.2.16 
2.1.07.01.00 N/A Rockfall (Large Block) P Exclude 6.2.14 
2.1.09.03.00 N/A Volume Increase of Corrosion Products P Include 6.2.13.3 
2.1.11.07.00 N/A Thermally-Induced Stress Changes in Waste 
and EBS 
P Include 6.2.9

ANL-WIS-MD-000008 REV 00 21 April 2000 
6.2.1 Waterlogged Rods - YMP No. 2.1.02.11.00 
FEP Description Failed fuel rods that occur in up to 0.1% of the fuel rods are 
currently being stored in commercial reactor spent fuel pools. This is attributed to breaches 
caused by manufacturing defects and reactor operations. Failed fuel contains water in the fuel 
rod void space. 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. 
Screening Decision: Exclude, low consequence. 
Screening Argument: The fraction of fuel rods with breached cladding is small (up to 
0.1%). Moisture can be removed from defective rods during the cask drying operations. The 
residual moisture of the cask atmosphere can be estimated based on the drying conditions to 5 
mbars (= 500 Pa). The impact on the waste form, waste package, and cladding of a small amount 
of remaining moisture is negligible. 
TSPA Disposition Few rods are breached at the time they are received at the 
repository. Breached rods are dried effectively using standard cask drying techniques. The 
remaining moisture after drying has little or no effect on the fuel. Therefore this FEP is excluded 
from further evaluation based on the analyses described below. 
IRSR Issues: CLST1, CLST3, TSPAI4 
References: CRWMS M&O 1995b, EPRI 1997, Kohli, R. and Pasupathi, V. 
1986, Peehs, M. and Fleisch, J. 1986, NRC 1997, CRWMS M&O 2000b, Knoll and Gilbert 
1987. 
Basis for Screening Decision 
The fraction of fuel rods with breached cladding is currently estimated to be 0.164% (CRWMS 
M&O 2000b, Table 13), or about 9.1 rods in an "average" waste package. EPRI 1997 (p. 4-1 
and 4-2) gives the BWR and PWR fuel rod reliability as a function of calendar year based and 
their estimates are lower than that estimated in CRWMS M&O 2000b. 
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 to U3O8. NRC (1997, Section 8.V.1) describes the cask drying criteria with 
reference to Knoll and Gilbert 1987. Less than 0.43 mol of H2O are expected to be present in a 
7-m3 cask after drying. This volume of water produces an insignificant potential for corrosion of 
the cladding during dry storage or during disposal (CRWMS M&O 1995b, p. 16). 
Kohli and Pasupathi (1986, p. iii) discusses 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. Since the initial amount of water in the reactor operations breached rods is 
unknown before drying, 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.

ANL-WIS-MD-000008 REV 00 22 April 2000 
The rods were dried in flowing argon at atmospheric pressure while being heated in a furnace. 
The center 1.8m of the furnace was heated to 673 K; the remainder was heated to 473 K. 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 3600 to 4800 sec and all measurable 
releases ended after 13200 sec (3.7 hours). This set of experiments demonstrated that standard 
cask drying procedures would remove the water from the failed rods. 
Peehs and Fleisch (1986, pp. 199-202) described the behavior of waterlogged PWR fuel rods on 
heating at 400oC in a hot cell. The bulk of the water was released during the cask drying 
operation. Results of the testing were 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. 
If there were problems with the drying process and rods were filled with water, the extent of the 
potential corrosion can be estimated. CRWMS M&O (2000b Section 6.3.5) discusses the free 
volume in a rod and concludes that the average irradiated rod has 17.7 cm 3of free volume 
although the rod started out with a free volume of 23.3 cm 3. Using the as-manufactured void 
volume, it is found that there can be no more than 23.3 g of water in one fuel rod. By 
comparison, 88.2 g of water is required to oxidize the fuel in one fuel rod from UO2 to U 3O8. 
That is as much water as could be supplied by about 3.8 fully waterlogged fuel rods. 
In conclusion, waste form degradation from waterlogged fuel rods is excluded from TSPA-SR. 
There are few failed rods in any WP that could be waterlogged. Experiments have demonstrated 
that cask drying procedures are effective in removing water. Since the volume of water inside a 
rod is quite limited, the water could only affect a small amount of fuel. 
6.2.2 Cladding Degradation Before YMP Receives It - YMP No. 2.1.02.12.00 
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. 
Screening Decision: Include. 
Screening Argument: The performance of commercial Zircaloy clad fuel has improved 
with time. The overall rod failure rate is 0.01  0.05 percent. For the TSPA-SR model, CRWMS 
M&O (2000b, Section 7) gives the CCDF for the fraction of failed rods expected to be received 
at YMP. In terms of exposing UO2 fuel pellets to the environment after the waste package (WP) 
fails, most of the fuel that has failed during reactor operation will still have some protection 
provided by the remaining cladding. 
After discharge from the reactor, the fuel assemblies are stored in open spent fuel storage pools. 
An International Atomic Energy Agency (IAEA) survey (IAEA 1988, Table XXVI) reported no 
evidence of fuel degradation in spent fuel pools and no evidence of significant additional 
degradation of fuel damaged during reactor operation during residence in the spent fuel pool. 
The oldest fuel in the survey was Shippingport PWR fuel that has been in wet storage since

ANL-WIS-MD-000008 REV 00 23 April 2000 
1959. Other fuel reported to have no significant additional degradation during residence in the 
spent fuel pool has been in wet storage since 1962, 1966, 1968, and the 1970s. The importance 
of the spent fuel pool storage experience is that fuel failure or significant further degradation is 
not expected during pool storage, and the fuel failure rates observed from reactor operation are 
appropriate for the cladding degradation analysis. 
The DOE has sponsored a Spent Nuclear Fuel Integrity During Dry Storage-Performance Tests 
Program at the INEEL since 1984. Approximately 26,500 rods have been studied in various 
commercial dry storage casks. This program also demonstrates what can be expected to happen 
to fuel in repository WPs during the early thermal period when temperatures are elevated. The 
cover gas has been monitored and very few rod failures have been observed during dry storage. 
The importance of the spent fuel dry storage experience (domestic and foreign) is that fuel 
failure or degradation is expected to be very small during dry storage, and the fuel failure rates 
are approximated in the cladding degradation analysis by CRWMS M&O (2000b and 2000c). 
During normal shipping of fuel, no failures have been reported in the literature. Failure 
probabilities have been calculated for a specific drop test (9-m drop) of a cask with impact 
limiters, a 0.3-m drop, and normal transport (i.e., vibration). Results show that no yielding of the 
rods should occur below 63g acceleration. Normal transport would result in accelerations much 
below these values. IAEA surveyed shipping worldwide and reported in 1988 that there had been 
no major incidents during 30 years experience connected with irradiated fuel transport. The 
importance of the fuel transportation and handling experience (domestic and foreign) is in 
demonstrating that fuel failure or degradation during transportation from vibration or impact is 
expected to be very small, and that the fuel failure rates are appropriately estimated for the 
cladding degradation analysis. 
TSPA Disposition For the TSPA-SR analysis, initial cladding failures from reactor 
operation, dry storage and transportation has a median value of 0.0948 percent with a range of 
0.0155 percent to 1.28 percent for any group (bin) of WPs. In addition, cladding creep failures 
during dry storage and transportation are integrated into the creep failure analysis for the YMP 
post-closure clad degradation abstraction. Fuel degradation during spent fuel storage is excluded 
from TSPA-SR. 
IRSR Issues: CLST1, CLST3, TSPAI4 
References: Einziger, R.E.; Atkin, S.D.; Stellrecht, D.E.; and Pasupathi, V. 
1982, Einziger, R.E. and Kohli, R. 1984, EPRI 1997, IAEA 1988, Johnson, A.B., Jr. 1977, 
Johnson, A.B., Jr.; Bailey, W.J.; Schreiber, R.E.; and Kustas, F.M. 1980, Johnson, A.B., Jr.; 
Dobbins, J.C.; Zaloudek, F.R.; Gilbert, E.R.; and Levy, I.R. 1987, Manaktala, H.K. 1993, 
McDonald, S.G. and Kaiser, R.S. 1985, McKinnon, M.A. and Doherty, A.L. 1997, Ravier, G.; 
Masuy, G.; and Willse, J.T. 1997, Sanders, T.L.; Seager, K.D.; Rashid, Y.R.; Barrett, P.R.; 
Malinauskas, A.P.; Einziger, R.E.; Jordan, H.; Duffey, T.A.; Sutherland, S.H.; and Reardon, P.C. 
1992, Sasaki, S. and Kuwabara, S. 1997, Schneider, K.J. and Mitchell, S.J. 1992, Witte, M.C.; 
Chun, R.C.; and Schwartz, M.W. 1989, Yang, R.L. 1997, CRWMS M&O 2000b, CRWMS 
M&O 2000c.

ANL-WIS-MD-000008 REV 00 24 April 2000 
Basis for Screening Decision 
6.2.2.1 Pin Degradation During Reactor Operation - YMP No. 2.1.02.12.01 
The performance of commercial Zircaloy-clad fuel has improved with time. Early cores (up to 
1985) had a rod failure rate averaging 0.02 percent to 0.07 percent (EPRI 1997, p. 4-1). After 
1985, the rod failure rate decreased to 0.006 percent to 0.05 percent. The overall rod failure rate 
for both time periods is 0.01 percent to 0.05 percent. Manaktala (1993, p. 3-4) shows fuel 
reliability as a function of calendar year for both PWR and BWR fuel from 1969 through 1976. 
He shows early PWR cores had failure rates over 1 percent but the rates dropped below 0.1 
percent by 1973. The BWR failure rates fell below 0.1 percent after 1975. Yang (1997, Table 2, 
p. 10), summarized the frequency and type of assembly failures from 1989 through 1995. 
Yangs failure data are for assemblies, and can be converted to rod failure rates using 2.2 failed 
rods per failed assembly (EPRI 1997, p. 4-1), and an average of 221 rods per assembly. 
Occasionally a specific core will have a higher failure rate. These data are included in the EPRI 
fuel failure rate, and affect the averages slightly. After a steam generator replacement in 1981, 
one reactor operator reported 0.26 percent of the rods (32 percent of the assemblies) were 
damaged by fretting from foreign particles, such as debris from the steam generator replacement 
(McDonald and Kaiser 1985, pp. 2-4, 2-5). This case of severe core damage shows that a 
relatively small fraction of the rods was actually damaged. The French ( Ravier et al. 1997, 
Figure 4, p. 34) report annual rod failure rates from 5.510-4 to 0.0 from 1986 through 1996. 
Their reactors are PWRs based on the Westinghouse design, and have similar fuel designs as the 
United States reactors. The Japanese (Sasaki and Kuwabara 1997, pp. 13-14) report failure rates 
of 0.01 percent for BWRs and 0.002 percent for PWRs. Their BWR design is based on the 
General Electric (GE) design, and their fuel designs are similar to the United States reactors. 
Both rates are similar to the current United States observed rates. 
Rod damage characteristics in reactor operation (EPRI 1997, pp. 4-2, 4-3) are: 
Pinhole and through wall hairline cracks 80 to 90 percent 
Intermediate condition 10 to 20 percent 
Severe damage 0.04 to 0.9 percent 
In terms of exposing UO2 fuel pellets to the environment after the WP fails, most of the fuel that 
has failed during reactor operation will still have some protection provided by the remaining 
cladding. 
Reactor operation failed rods are included in TSPA-SR. The details of the analysis are presented 
in CRWMS M&O 2000b, which also develops the stress distribution used in the creep analysis. 
6.2.2.2 Pin Degradation During Spent Fuel Pool Storage -YMP No. 2.1.02.12.02 
Rod failure during spent fuel pool storage is excluded from TSPA-SR. The importance of the 
spent fuel pool storage experience is that fuel failure or further degradation of rods already failed 
during reactor operation is not expected during pool storage, and the fuel failure rates observed 
from reactor operation, dry storage, and shipping are appropriate for the cladding degradation 
analysis.

ANL-WIS-MD-000008 REV 00 25 April 2000 
After discharge from the reactor, the fuel assemblies are typically stored in spent fuel storage 
pools. An International Atomic Energy Agency (IAEA) survey (IAEA 1988, Table XXVI) 
reported no evidence of fuel degradation in spent fuel pools and no evidence of significant 
additional degradation of fuel damaged during reactor operation. The oldest fuel in the survey 
was Shippingport PWR fuel that has been in wet storage since 1959. Other fuel reported to have 
no further degradation has been in wet storage since 1962, 1966, 1968, and the 1970s. 
Under the DOE Spent Fuel and Fuel Pool Component Inventory Program, the effect of storing 
both fuel with intact cladding and fuel with failed cladding has been studied. An international 
survey of in-water storage (Johnson et al. 1980, p. iii) reports no cases of fuel cladding 
degradation during pool storage. 
Johnson (1977, p. 20), reports: Operators at several reactors have discharged, stored, and/or 
shipped relatively large numbers of Zircaloy-clad fuel which developed defects during reactor 
exposures, e.g., Ginna, Oyster Creek, Nine Mile Point, and Dresden units I and II. Several 
hundred Zircaloy-clad assemblies which developed one or more defects in-reactor are stored in 
the GE-Morris pool without need for isolation in special cases. Detailed analysis of the 
radioactivity in the pool water indicates that the defects are not continuing to release significant 
quantities of radioactivity. 
6.2.2.3 Pin Degradation During Dry Storage -YMP No. 2.1.02.12.03 
Creep strain damage and pin damage (failure) during dry storage are included in the creep 
degradation analysis described in CRWMS M&O (2000c, Section 6.2). It is assumed that all 
fuel undergoes 20 years of dry storage starting at 350oC. The amount of creep strain during dry 
storage is calculated and added to the creep from shipping and the YMP thermal transient 
(CRWMS M&O 2000c, Figure 5). In addition, 0.045 percent of the rods are assumed to have 
failed in dry storage from handling and other causes (CRWMS M&O 2000b, Table 13). 
The DOE has sponsored a Spent Nuclear Fuel Integrity During Dry Storage-Performance Tests 
Program at the INEEL since 1984 (McKinnon and Doherty 1997, pp. 2.1, 5.16). Approximately 
26,500 rods have been studied in various commercial dry storage casks. This program also 
demonstrates what can be expected to happen to fuel in repository WPs during early storage 
times when temperatures are elevated. The dry storage cask fuel cover gas has been monitored 
and very few fuel rod failures have been observed during dry storage. During fuel consolidation, 
approximately 10 rods are believed to have developed small leaks. Some rods leaked, but the 
release was over a period of approximately two months, a very slow gas release. Only 
0.5 percent of an individual rods 85Kr content was released from the leaking rods. After 
shipment to the INEEL, the observed rod failure was similar to that after reactor operation. For 
rods that were not consolidated, 2 rods leaked out of 16,700 rods producing a failure probability 
of 1.2  10-4 per rod, below the 0.045 percent failure rate used in TSPA-SR for failure in dry 
storage. 
Some utilities experimented with fuel consolidation before the introduction of dry storage 
facilities. After consolidation at INEEL (McKinnon and Doherty 1997, pp. 2.1, 5.16), 12 rods 
failed out of 9800 (failure probability =1.22  10-3), an increase by 1 order of magnitude from 
the unconsolidated failure rate. Although no utilities are currently using consolidation, the rod

failure rate on 0.045% used in TSPA-SR for dry storage contains a component of the failure rate 
from consolidation. 
Dry storage tests were performed at the Nevada Test Site (Johnson et al. 1987, p. iv) with 17 
spent fuel assemblies, each in an individual test. These tests contained 3468 rods, and cladding 
temperatures varied from 168 to 27XC. One of the fuel rods failed during these tests. This 
assembly was exposed to air at 275'C and had nine thermal cycles. The estimated hole size was 
1 pm. No further degradation was observed in this one failed rod after the initial failure. No 
visible damage was observed in the other tests. 
Accelerated high temperature tests were performed on 15 rods (Einziger et al. 1982, pp. 65, 69). 
Post irradiation studies of failure mechanisms of well-characterized pressurized water reactor 
rods were conducted for up to a year at 482'C, 510C, and 571C in limited air and inert gas 
atmospheres. No cladding breaches occurred. The cladding had crept away from the pellets and 
showed a smoother profile. Strains from 1.7 percent to 7 percent were measured (their Table 
IV). One rod had a local creep strain as high as 12 percent (their Figure 3, p.69). The extended 
lifetime is attributed to significant creep strain of the Zircaloy cladding, which decreases the 
internal rod pressure. The cladding creep also contributes to radial cracks through the external 
oxide layer and internal fuel-cladding chemical interaction layers (layer of a few micrometers 
where some zirconium/U02 interaction occurs), which propagated into and arrested in an oxygen 
stabilized alpha-Zircaloy layer. Since cracks extended only for a few micrometers and were 
arrested, their significance was small. There were no signs of either additional cladding 
hydriding, stress corrosion cracking (SCC), or fuel pellet degradation. 
A second series of tests (Einziger and Kohli 1984, pp. 107 and 114, Table 111) was performed on 
five PWR spent fuel rods. They were pressurized to a hoop stress in the range of approximately 
145 to 155 MPa, for times up to 2101 hours at 323'C. The conditions were chosen for limited 
annealing of in-reactor irradiation hardening. With the stresses in the range of 145 MPa, creep of 
0.004% to 0.16% was observed. No cladding breaches occurred, although significant hydride 
agglomeration and reorientation took place in one rod that cooled under stress. Einziger and 
Kohli (1984, pp. 107 and 114, Table 111) state that these high-temperature tests based on creep 
rupture as the limiting mechanism indicate that storage at temperatures between 400 and 440C 
may be feasible for annealed rods. 
Schneider and Mitchell (1992, p. 2.7) summarized experience in the foreign dry storage 
programs. At that time, seven countries had some fuel in dry storage. They conclude that Light 
Water Reactor (LWR) fuel can be stored for up to 100 years at temperatures of 320 to 400C in 
an inert atmosphere and, if exposed to air, will last comparable times if the temperature is limited 
to 135 to 160C. The Canadians have large quantities of irradiated fuel with Zircaloy cladding 
that has been exposed to air with favorable results. For over eight years, they tested the effects 
of exposing fuel with defective cladding to moist and dry air with favorable results (no observed 
strain from U02 oxidation nor cladding failure propagation). 
The importance of the spent fuel dry storage experience (domestic and foreign) is that fuel 
failure or degradation is expected to be very small during dry storage for the current fuels. It is 
also important to note that under repository conditions, most of the cladding creep or DHC 
failures would be expected to occur in the first 100 years after closure, this is the time period 
ANI-WIS-MD-000008 REV 00 26 Apr112000 

when fuel temperatures are highest and conditions are closest to dry storage conditions. Dry 
storage is considered to be a good representation of this period. 
6.2.2.4 Pin Degradation During Fuel Shipment and Handling - YMP No. 2.1.02.12.04 
During normal shipping of fuel, no failures have been reported in the literature. Sanders et al. 
(1992, Table 111-10, p. 111-137), analyzed transportation accidents. Failure probabilities were 
calculated for a specific drop test of nine meters (9-m), a 0.3-m drop, and normal transport (i.e., 
normal vibration). The 9-m drop was by far the most severe, with failure probabilities on the 
order of 2 x The normal transport failure is 2 x per rod. This reference generated 
acceleration versus frequency curves for trucking and rail shipping. It then looked at structural 
damage from a sampling of the hazard curves and the structural analysis and concluded that no 
additional damage is done in shipping. This value is conservatively bounded by a failure fraction 
of 1 x 10" to account for other transport accident conditions and is included in the TSPA-SR. In 
addition, it was assumed that all rods undergo three (3) weeks of transportation at a temperature 
of 35OoC when considering the creep strain damage. This inclusion is described in CRWMS 
M&O (2000b, Section 6.10.1) and (2000c, Section 6.2.1). 
Witte et al. (1989) performed an analytical evaluation of the potential impacts of all transport 
movements. They concluded (Witte et al. 1989, p. 194, Table 3) that no yielding of the rods 
would occur below 63g acceleration. Normal transport would result in acceleration far below 
these values. 
IAEA (1988, p. 114) surveyed shipping worldwide and reported that "to date, there have been no 
major incidents during 30 years experience connected with irradiated fuel transport." 
The importance of the fuel transportation handling experience (domestic and foreign) is in 
demonstrating that very little fuel failure or degradation is expected during transportation from 
vibration or impact, and that the fuel failure rates observed from reactor operation are 
appropriate for the cladding degradation analysis. 
6.2.3 General Corrosion of Cladding - YMP No. 2.1.02.13.00 
FEP Description General corrosion of cladding could expose large areas of fuel and 
produce hydrides. 
Screening Decision: Exclude, low consequence 
Screening Argument: General corrosion is synonymous with zirconium oxidation. The 
outer surface of the cladding becomes oxidized with ZrOa providing a protective film which 
adheres to the surface and slows down further oxidation. The oxidation could be from 0 2 
consumption (dry oxidation) or Hz0 consumption (wet oxidation) and the dry and wet oxidation 
rates are very similar. For the fuel in the repository, oxidation does not occur until the WP is 
penetrated. Various analyses have concluded that cladding oxidation under repository conditions 
would not lead to rod failure. The amount of oxide film resulting from general corrosion as a 
function of time since WP failure has been calculated but is insignificant if the WP remains 
sealed for the first 100 years. There was no measurable degradation of the Zircaloy cladding 
after 21 years of in-water spent fuel pool storage times. 
ANI-WIS-MD-000008 REV 00 27 Apr112000 

TSPA Disposition Exclude, low consequence 
IRSR Issues: CLST1, CLST3, TSPAI3, TSPAI4 
References: Bradley, E.R.; Bailey, W.J.; Johnson, A.B., Jr.; and Lowry, L.M. 
1981, Einziger, R.E. 1994, Hillner, E.; Franklin, D.G.; and Smee, J.D. 1998, IAEA 1998, 
Rothman, A. J. 1984 
Basis for Screening Decision 
General corrosion is synonymous with zirconium oxidation. The outer surface of the cladding 
becomes oxidized with a Zr02 film, which adheres to the surface and slows down further 
oxidation. The oxidation could be from 0 2 consumption (dry oxidation) or Hz0 consumption 
(wet oxidation). For the fuel in the repository, this does not occur until the WP is penetrated. 
The effect of surface oxidation is threefold. The oxidation could thin the cladding, contributing 
to cladding failure by creep rupture. Wet oxidation generates hydrogen, and some of the 
hydrogen is absorbed into the cladding to form hydrides. This could lead to delayed hydride 
cracking (DHC), or general hydride embrittlement. In the extreme, the oxidation could lead to 
cladding disintegration and expose the fuel pellets to the WP environment. 
Wet oxidation has been studied for over forty years because of its importance in reactor 
operation. Rothman (1984, pp. 6 - 13) discusses cladding oxidation in repository conditions in 
great detail and his Table 3 compares the predicted cladding loss using six different oxidation 
correlations and predicts cladding thinning of 4 to 53 pm after 10,000 years at 180C (a 
conservative temperature assumption) which would not lead to rod failure. 
Recently, Hillner et al. (1998, p. 9) published a recommended Zircaloy corrosion correlation 
based on Bettis Atomic Power Laboratory experiments. Bettis developed Zircaloy for Navy 
reactors in the early 1950s and has an extensive data base on Zircaloy performance. They have 
continuous autoclave corrosion tests on some samples for 30 years. Some samples have 
developed oxide thickness as great as 110 pm, greater than those expected during repository 
corrosion. The experiments show that the corrosion rate is determined by diffusion of oxygen 
ions through the corrosion film. This corrosion film is generated in layers, with the lower layer 
staying very consistent. The consistency of the lower 2 pm of oxide film leads to a steady 
corrosion rate after a transition period. The recommended post-transition rate equation is: 
ATh =1.69 x l o 9 x exp(-111452lT) (Eq. 1) 
where: 
ATh = oxide growth rate, pm/y 
T = temperature, K 
This correlation rate equation is similar to the equations developed by others, but predicts a 
slightly higher corrosion rate. It is taken from Equation 7 of Hillner at al. (1998), doubled for a 
conservative correction for irradiation (Hillner et al. 1998, pp. 6, 9) and converted to 
ANI-WIS-MD-000008 REV 00 28 Apr112000 

micrometers per year (lpm = 14.7 mg/dn?) (IAEA 1998, p. 178). The pre-transition rate is 
slower than the post-transition rate. The effect of irradiation conditioning before beginning the 
corrosion is to accelerate the corrosion rate for a few micrometers. To be conservative, Hillner et 
al. doubled the corrosion rate for all time. The correlation shows a strong Arrhenius temperature 
relationship with the corrosion rate becoming very small below 200C. This is consistent with 
the data from Hillner et al. Corrosion tests at 270C for 8.2 years have produced approximately 
4 pm of oxide, while corrosion tests at 360C have produced films 88 pm thick in 7.8 years 
(Hillner et al. 1998, p. 25). 
For general cladding corrosion, the WP must be breached. The amount of oxide film from 
general cladding corrosion as a function of time since WP failure has been calculated. These 
calculations were carried out from the WP failure time to one million years. The amount of 
cladding that is consumed is approximately 57 percent of the oxide thickness because of the 
volume increase associated with Zircaloy oxidation. The cladding oxidation calculation was 
done for the design bases (hot) rod in the hottest repository region. If the WP is not breached 
after 100 years from emplacement, little general corrosion is expected. Hillner et al. (1998, 
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 steam environment is expected to last for 1000 to 5000 years in the near 
field. Einziger (1994, p. 556, Equation 14) states that dry oxidation of zirconium is slightly 
slower than the wet corrosion rate. The conclusion from CRWMS M&O (1998, Section 
6.3.1.1.4.1), is that general cladding corrosion is not a problem for fuel cladding in the average 
WP but could be a problem for fuel cladding in the design WP if the design WP fails before 100 
years. 
Bradley et al. (1981, p. 38) performed metallurgical examinations of Zircaloy-clad fuel rods from 
two bundles (0551 and 0074) of the Shippingport PWR Core 1 "blanket" fuel after extended inwater 
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 pm (0551) and 2.4 pm (0074). After extended in-water 
spent fuel pool storage, the average cladding oxide film thickness was found to be 1.7 pm (0551) 
and 2.3 pm (0074)(Bradley et al. 1981, 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 bundle 0551 fuel rods in 1960) stored in water for over 20 years 
were unblemished and showed no evidence of reaction with water. 
In summary, the conclusion in this AMR is that general cladding corrosion is not a problem for 
fuel cladding in the average WP but could be a problem for fuel cladding in the design WP if the 
WP fails before 100 years. 
6.2.4 Microbiologically Influenced Corrosion (MIC) of Cladding - YMP No. 2.1.02.14.00 
FEP Description Microbiologically 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 
ANI-WIS-MD-000008 REV 00 29 Apr112000 

Screening Decision: Exclude, low probability. 
Screening Argument: There is no indication that MIC occurs on zirconium metal or 
alloys. The two major forms of MIC for materials being considered for waste packages 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 these forms of MIC: (1) SRB do not 
affect zirconium, and (2) corrosion induced by organic acids is unlikely because of zirconium's 
tolerance to organic acids and a wide range of pHs. 
TSPA Disposition Exclude. 
IRSR Issues: CLST1, CLST3, TSPAI3, TSPAI4 
References: Little, B. and Wagner P. 1996, Yau, T.L. and Webster, R.T. 1987, 
Hillner, E.; Franklin, D.G.; and Smee, J.D. 1998, McNeil, M., and Odom, A. 1994, CRWMS 
M&O 2000a 
Basis for Screening Decision 
The term microbiologically influenced corrosion (MIC) is commonly used to designate corrosion 
caused by the presence and activities of microorganisms at the surfaces of metals. Little and 
Wagner (1996, p. 367) published an overview of microbiologically influenced corrosion 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, underdeposit 
corrosion, and enhanced erosion/corrosion. Little and Wagner (1996, pp. 367-368) describe 
several mechanisms for microbiologically influenced corrosion. In addition, various case studies 
are presented that document microbiologically influenced corrosion of alloys of iron, nickel, and 
copper. However, it should be noted that there is no indication in the literature that MIC occurs 
on zirconium metal or alloys. Yau and Webster (1987, p. 709) report that no corrosion of 
zirconium metal from marine organisms was found during sea water corrosion tests for 129 days. 
Hillner et al. (1998, p. 11) indicate that there are two major forms of MIC for materials being 
considered for WPs. They are (1) sulfide attack through the action of sulfate reducing bacteria 
and (2) corrosion induced by organic acids secreted from certain bacteria. With respect to attack 
by SRB, Hillner et al. 1998 reference the work of McNeil and Odom (1994, p. 176). McNeil and 
Odom (1994, p. 176) indicate by thermodynamic calculations that SRB do not affect zirconium. 
With respect to corrosion induced by organic acids, Hillner et al. (1998, p. 11) noted that it is 
most unlikely because of zirconium's tolerance of a wide range of pHs and that it is unlikely that 
production of weak organic acids will have an adverse effect on the passivation of Zircaloy by a 
Zr02 film. Yau and Webster (1987, p. 717) also note that zirconium 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. 
CRWMS M&O 2000a summarizes the corrosion potential of zirconium to many chemicals. 
ANI-WIS-MD-000008 REV 00 

6.2.5 Acid Corrosion of Cladding From Radiolysis - YMP No. 2.1.02.15.00 
FEP Description Radiolysis in a nitrogenloxygen gas mixture with the presence of 
water film results in the formation of nitric acid (HNO3) and hydrogen peroxide (H202) ions that 
can enhance corrosion of the fuel cladding. 
Screening Decision: Exclude, low probability. 
Screening Argument: Radiolysis has the potential to produce chemical species that may 
not otherwise be anticipated in the repository. One example is the production of nitric acid as a 
result of the radiolytic reaction between nitrogen and oxygen with the production of nitrogen 
oxides and subsequent formation of nitric acid. Another example is the production of hydrogen 
peroxide. An analysis of the gamma and neutron dose at 1000 years at the WP surface, thus well 
before the first anticipated breach of the waste packages, provides a total dose of 0.466 remlhour 
at the waste package surface (CRWMS M&O 1999d, p. 17). The dose may be too low for the 
production of sufficient peroxide or nitric acid to cause a corrosion concern. 
Zirconium is used for fuel cladding because of its high resistance to corrosion in high radiation 
fields. Yau and Webster 1987 (p. 710) note that zirconium has excellent corrosion resistance to 
nitric acids. They also report that zirconium does not corrode in concentrated hydrogen 
peroxide. 
TSPA Disposition Exclude. 
IRSR Issues: CLST1, CLST3, TSPAI3, TSPAI4 
References: CRWMS M&O 2000a, CRWMS M&O 1999d, Hansson, C.M. 
1984, Van Konynenburg, R.A.; Curtis, P.G.; and Summers, T.S.E. 1998, Yau, T.L. and Webster, 
R.T. 1987, IAEA 1993, Hillner et al. 1998. 
Basis for Screening Decision 
It is recognized that radiolysis may create various chemical species that may not otherwise be 
expected in the repository. Examples are the production of nitric acid as a result of the radiolytic 
reaction between nitrogen and oxygen and the formation of hydrogen peroxide. In either case, 
accelerated corrosion of Zircaloy would not be predicted in the absence of irradiation since 
zirconium alloys have been shown to be relatively inert in both media as discussed in Yau and 
Webster 1987 and CRWMS M&O (2000a, Sections 6.1.6 and 11.4). As one example, zirconium 
equipment is used in the chemical processing industry where peroxide strengths of 90% are used. 
The service life has been increased by an order of magnitude compared to graphite components 
previously used, which were generally considered to be very inert. In nitric acid, zirconium and 
its alloys are inert up to acid concentrations of 65%. Provided radiolysis does not produce nitric 
acid at greater concentrations, there will be no impact on the uniform corrosion rate as a result of 
nitric acid production from radiolysis. 
Radiolysis was reviewed in detail in the IAEA 1993 report (pp. 82-92). Specifically the topic of 
radiolysis of "thick oxide film effects" was discussed. The report noted (IAEA 1993, p. 91) that 
"the radiation enhancement of Zircaloy corrosion in the aqueous phase arises from the 
ANI-WIS-MD-000008 REV 00 31 Apr112000 

synergistic interaction of radiation and water chemistry. It has been recognized that in the early 
stages of Zircaloy corrosion the acceleration by reactor radiation, which is usually observed in 
oxygenated water, is suppressed in the presence of excess hydrogen in the aqueous phase, but 
that beyond the threshold oxide thickness Zircaloy specimens exposed to low-oxygen water 
corrode at accelerated rates under irradiation as if they were immersed in oxygenated water." 
This implies that under the worst conditions of high radiation levels and excess oxygen ions, the 
corrosion rate is only accelerated by a factor of three. However, an analysis of the gamma and 
neutron dose in the repository at 1000 years, thus well before the first anticipated breach of the 
waste package, predicts a total dose of 0.466 remhour at the waste package surface (CRWMS 
M&O 1999d, p. 17). This dose is significantly lower than that in an operating reactor and may in 
fact be too low to produce sufficient peroxide to produce an accelerated corrosion. In the 
unlikely event that premature failure of a WP should occur, the waste package temperature 
would preclude the presence of liquid water, and thus the presence of peroxide, for at least 1000 
years. At this time the surface radiation dose will be very low level as noted above, so that the 
practical likelihood of producing peroxide is extremely low under any conditions. 
Further support for this position can be obtained from Hillner et al. 1998. On page 12, it is stated 
that "The corrosion rate of Zircaloy in water is limited by the slowest of the basic corrosion 
steps: dissociation of water into oxygen and hydrogen ions, diffusion of oxygen ions through the 
oxide film, oxidation of the Zr metal, diffusion of electrons through the oxide, and hydrogen-ion 
reduction by the electrons at the water-to-oxide interface. At temperatures for current testing, 
that is, above about 45OoF, diffusion of oxygen ions through the oxide film is rate limiting. In 
extrapolating to lower temperatures, one of the other basic steps could become rate limiting, 
which would decrease the corrosion rate to less than predicted by the extrapolation". In 
summary, if corrosion is controlled by the oxygen ion diffusion rate, the presence of peroxide 
would have no effect on Zircaloy corrosion and the corrosion rate would be expected to be that 
of uniform corrosion. 
Yau and Webster (1987, pp. 707-721) review the corrosion of zirconium under various chemical 
environments for commercial applications (corrosion rates and time scales of interest for 
industrial applications, not for repository time scales). Zirconium is resistant to corrosion from 
HC1 to temperatures well above boiling (Yau and Webster 1987, p. 710). In basic solutions, 
Hansson (1984, summary page) measured corrosion of Zircaloy 2 in anaerobic cement pore 
solutions of DHS of 12.0 to 13.8. She concludes: "Thus. it mav be concluded that active 
corrosion of Zircaloy 2 in anaerobic concrete will not occur and by comparison with 
measurements on steel, it is likely that the passive corrosion rates will be even lower in concrete 
than those measured in the synthetic pore solution." Yau and Webster 1997 report no corrosion 
in seawater, brackish water, and polluted water. Zirconium is resistant to corrosion from sulfuric 
acids in concentrations less than 20 percent and corrodes slowly in terms of commercial 
applications (not repository time scales) in sulfuric acid solutions below 65 percent. 
Van Konynenburg et al. (1998, p. 7-17) performed container material scoping tests using 
Zircadyne 702 (a Zirconium hafnium alloy) in 0.01 mol/L each of sodium formate (NaCOOH), 
nitric acid (HNO?), NaCI, H202, and 0.02 mol/L sodium oxalate (NazCzOd), with an initial pH of 
approximately 4, temperature of 9Of 5 "C, and typical test duration of 96 hours. The test solution 
was designed to represent highly concentrated forms of acids formed by radiolysis. Formic acid, 
ANL-WIS-MD-000008 REV 00 32 April 2000 

a potential pH buffer, was not present in the Zircadyne 702 test. They report a final pH of 4.26 
and corrosion rate of 50 pmlyr for zirconium in the acid solution. 
The duration of highly alkaline conditions in an emplacement drift depends on the assumed use 
of concrete. If concrete (emplacement drift liner material) is not used, the duration of highly 
alkaline conditions in the emplacement drift is less than 5000 years, with the long-term pH at 
about 8. If concrete material is used in the liner, the expectation is that the pH in the 
emplacement drift will be quite basic for the first 10,000 years. 
After emplacement of the waste packages, the radioactive decay of the waste will heat the drifts 
and disturb the normal percolation of water through the mountain. As the drifts cool, some of the 
ground water percolating through the mountain may drip into the drifts and some of the waste 
packages. Through time, the components of the engineered barrier are expected to degrade. The 
metallic materials of the WPs, pedestals, and drip shield are expected to undergo humid air and 
aqueous corrosion, and the concrete materials, if used, are expected to undergo hydrothermal 
degradation. As the materials deteriorate, the drift liner will fall onto and around the WP and the 
pedestals will collapse lowering the WPs toward the floor. The WPs and drip shield are expected 
to undergo humid corrosion, and if wetted by dripping or high relative humidity, to undergo 
aqueous corrosion. It is assumed that corrosion degradation of Zircaloy fuel cladding does not 
begin until the WP is breached (typically 10,000 years or more after emplacement). A steam 
environment is expected to last for 1000 to 5000 years in the near field. 
Once a WP is breached, water may enter the WP as water vapor or as drips. The WP and fuel 
will have cooled down by that time to an extent that will allow a water film to form on the WP 
surface and on the fuel cladding that is thick enough to support corrosion reactions. The radiation 
dose would also be a lowered by that time to a level that is considered insignificant to long-term 
engineered barrier performance. 
In summary, the WPs are expected to be intact for thousands of years after emplacement. During 
the first 1000 to 5000 years, the environment around the WPs is expected to be steam. This 
environment will prevent nitric acid or hydrogen peroxide from pooling on or coating the WP (or 
fuel cladding) and would thus control acid corrosion. 
6.2.6 Localized Corrosion (Pitting) of Cladding - YMP No. 2.1.02.16.00 
FEP Description Localized corrosion in pits could produce penetrations of cladding. 
Screening Decision: Exclude, low probability. 
Screening Argument: Corrosion tests are currently being performed for repository type 
chemicals, but results are not yet available. The probability of pitting, crevice corrosion, or some 
other form of localized corrosion depends on the local water chemistry and for predicted WP 
water chemistry is unlikely. Local corrosion of zirconium-based clad by pitting is observed to 
occur when: 
1. Chloride levels exceed 1 mmol/L; 
2. Inhibitor ions, such as nitrate, sulfate, and bicarbonate are absent; 
3. Ferric ion levels are high and the pH low; 
ANI-WIS-MD-000008 REV 00 33 

4. The passivating ZrOz film is thin, absent, or mechanically disrupted. 
Note that the passivating film is somewhat soluble at moderately low pH (pH < 3.5), though its 
dissolution, and that of the underlying metal, is primarily observed at very low pH. Since none 
of these conditions is predicted to occur inside the waste package, the process is unlikely. 
TSPA Disposition Exclude 
ZRSR Issues: CLST1, CLST3, TSPAI3, TSPAI4 
References: Yau, T.L. and Webster, R.T. 1987, Hansson, C.M. 1984, Rothman, 
A.J. 1984, Hillner, E.; Franklin, U.G.; and Smee, J.U. 1998, Yau, 1'. L., 1984, Van 
Konynenburg, R.A.; Curtis, P.G.; and Summers, T.S.E. 1998, Cragnolino, G. and Galvele, J.R. 
1978, Maguire, M. 1984, Krauskopf, D.K. and Bird, D.K. 1995, Uziemblo, N.H. and Smith, 
H.D. 1989, Greene et al. 2000, CRWMS M&O 2000f. 
Basis for Screening Decision 
Chloride-accelerated corrosion is neglected, in part, because the water in the vicinity of the 
repository (represented by water from well 5-13 and referred to as J13 water) contains only 0.2 
mmol1L chloride. This is roughly 115 of the amount required to cause pitting in solutions 
otherwise free of corrosion inhibitors, though evaporative concentration might in some cases 
raise 513 chloride levels. 
Yau and Webster (1987, pp. 707-721) review the corrosion of zirconium under various chemical 
environments for commercial applications (their corrosion rates and time scales are of interest 
for industrial applications, not for repository time scales). CRWMS M&O 2000a also review the 
potential for corrosion of Zircaloy cladding. Zirconium is resistant to corrosion from HC1 to 
temperatures well above boiling (Yau and Webster 1987, p. 710), basic solutions, sea water, 
brackish water, and polluted water. It is resistant to corrosion from sulfuric acids in 
concentrations less than 20 percent and corrodes slowly in terms of commercial applications (not 
repository time scales) in solutions below 65 percent. Zirconium is susceptible to corrosion from 
fluoride ions (> 100 ppm and pH dependent), hydrofluoric acid, and aqua regia (mixture of 
hydrochloric and nitric acid). Yau and Webster (1987, p. 717) report that zirconium is among 
the structural metals that are most resistant to crevice corrosion. If the zirconium oxide layer is 
intact, it is not susceptible to galvanic corrosion if coupled to a more noble metal (platinum, 
gold, graphite, titanium, and silver). It has a corrosion rate of 18 pdyr for unwelded specimens 
in 3.5 percent boiling NaCl at a pH of 5 but no observed corrosion at a pH of 6. Hansson (1984, 
summary page) reports no corrosion rate increase in anaerobic cement pore solutions of pHs of 
12.0 to 13.8. It is susceptible to pitting from FeC13 ions (CRWMS M&O 2000a, Section 6.2.2.1). 
FeC13 pitting requires femc ions, which are the dominant species at pH less than 3. Current 
analysis of the water chemistry inside the WP predicts pHs greater than 3.6 (CRWMS M&O 
2000f, Table 3). 
Rothman (1984, pp. 4-5) qualitatively lists the chemical components of J-13 water and the ability 
of these chemicals to attack zirconium. He concludes that Zircaloy corrosion from these 
chemicals is unlikely. Hillner et al. (1998, p. 11) also qualitatively compared Zircaloy corrosion 
ANL-WIS-MD-000008 REV 00 34 April 2000 

potential with J-13 components and concluded that corrosion the J-13 water, in the measured 
concentrations and pH levels, would not have a detrimental effect on the zirconium corrosion 
behavior. 
The possibility of chloride-accelerated corrosion is made more unlikely by the fact that J-13 
water contains anions that are known to inhibit pitting. Nitrate at 0.16 mmolIL, sulfate at 0.25 
mmollL, and bicarbonate at 2 mmollL are known inhibitors at these levels unless chloride levels 
are raised to - 1 mol1L (Greene et al. 2000). 
Ferric and cupric ions can cause pitting of Zircaloy. Although large amounts of ferric iron are 
expected to accumulate through degradation of breached waste packages, very little of it will 
remain in solution due to the low solubility of Fe(II1)-containing solids. Ferric ion levels are not 
predicted to exceed 4 x mmol1L. The predicted Fe levels are many orders of magnitude less 
than those that are observed to accelerate pitting (Greene et al. 2000). Moreover, ferric ion 
pitting typically requires pH below 2.5, well below the lowest value predicted to prevail in the 
WP. 
Passivating surface layers on spent nuclear fuel rods tend to be thick and well-developed; 5 to 
100 pm for PWRs and 5 to 20 pm for BWRs, malung them more resistant to pitting than solids 
used in many experiments (CRWMS M&O 2000a). 
Published results (Yau 1984, pp. 14011-14018) show general, crevice and stress corrosion 
cracking (SCC) tests with zirconium in concentrated acids to simulate concentrated geothermal 
salts. In terms of chemical composition, these tests are expected to be more severe than expected 
at YMP, although the test time periods are very short. Yau (1984) also looked at crevices and 
SCC with mixed salts. The solution contained 25% NaCl + 0.5% CH3COOH + 0.1% S t 
saturated H2S, with or without saturated C02. Both welded and non-welded samples were 
exposed in U-bend SCC tests, and general and crevice corrosion tests. Tests were conducted 
both at room temperature and under high temperaturelpressure autoclave conditions. The pH 
was between 2.3 and 3 (measured at room temperature), and the test duration was for 30 or 60 
days. The U-bend samples showed no cracking after 60 days. Steel coupled U-bend samples 
showed no hydrogen pickup, although the steel nuts and bolts corroded badly during the testing. 
These test solutions became contaminated with iron, but the iron contamination resulted in no 
degradation of zirconium's corrosion resistance. Crevice tests showed no crevice corrosion. 
Autoclave tests of general corrosion showed negligible corrosion, with "slight" weight gain of 
the samples (the rate was not given in the paper). The author also noted that fluoride pitting is 
not expected in geothermal fluids, because the presence of silica tends to trap the fluoride and 
form non-corrosive complex ions. 
Van Konynenburg et al. (1998, p. 7-17) performed certain material scoping tests using Zircadyne 
702 (a Zirconium hafnium alloy) in 0.01 mol1L each of sodium formate (NaCOOH), nitric acid 
(HNO3), NaC1, H202, and 0.02 mol1L sodium oxalate (NazCzOd), with an initial pH of 
approximately 4, temperature of 9Of5 "C, and typical test duration of 96 hours. The test solution 
was designed to represent highly concentrated forms of acids formed by radiolysis. Formic acid, 
a potential pH buffer, was not present in the Zircadyne 702 test. They report a final pH of 4.26 
and corrosion rate of 50 pmlyr for zirconium in the acid solution. 
ANI-WIS-MD-000008 REV 00 35 

The data presented in the papers by Cragnolino and Galvele (1978, p. 1056) and Maguire (1984, 
p. 177) show that Zircaloy could pit in solutions with chloride ions (1 mollL or higher and at 
room temperature). Cragnolino and Galvele studied zirconium and Zircaloy-4, but Maguire only 
studied zirconium and data on zirconium may not be directly applicable to Zircaloy-2 and -4. 
These experiments were performed in air but without any additional oxidants present. Maguire 
reports that the critical pitting potential values are not affected by pH from 0 to 12.8 or by 
temperature. He also argues that presence of oxidizing impurities such as Fe3+ and cu2+ would 
make Zirconium more susceptible to localized corrosion. Maguire (1984, p. 188) reports: "Our 
experience has been that if the pH is maintained above three no pitting occurs." A pH less than 3 
is required for the preferential formation of the Fe3+ ions (Krauskopf and Bird 1995, p.228, 
Figure 9-2) and the pH inside the WP is predicted to stay above 3.6 (CRWMS M&O 2000a. 
Table 3). 
Zircaloy is susceptible to pitting with F ions. Uziemblo and Smith (1989, pp. 10-19) measured 
corrosion rates at various pH and F ion concentrations. With 100 ppm F ions, they found that 
Zircaloy-4 corroded a negligible amount at pH > 8, 32 pmlyr at pH = 6 and 1 mmlyr at pH = 5. 
Yau (1984, p. 140/3), states that silica tends to trap the fluoride and to form non-corrosive 
complex ions. 
The experiments by Uziemblo and Smith (1989) show the importance of pH on the corrosion 
rate. It also shows the importance of estimating the composition of the water and the degree of 
concentration for the water that contacts the fuel, since the fluoride content of 5-13 is about 2.4 
ppm. Experiments with 500 ppm cu2+ in 3.5 percent boiling NaCl also show a strong pH 
dependency with a corrosion rate of 18 pmlyr at pH of 5 and nil at a pH of 6 and above (Yau and 
Webster 1987, pp. 716-717, Table 13). 
CRWMS M&O 2000f calculated the chemistry in the WP after WP failure. The analysis 
predicted that the water would be basic for most of the time but could turn acidic for a short 
period because of the sulfur within the carbon steel. The pH for all the cases studied was always 
above a pH of 3.6 (CRWMS M&O 2000f, Tables 2 to 5). Therefore water chemistries that 
would permit pitting are not expected. 
6.2.7 Localized Corrosion (Crevice Corrosion) of Cladding - YMP No. 2.1.02.17.00 
FEP Description Localized corrosion in crevices could produce penetrations of 
cladding. 
Screening Decision: Exclude. 
Screening Argument: Zirconium is very resistant to crevice corrosion. CRWMS M&O 
2000a discusses the crevice corrosion resistance of zirconium in various chemical solutions and 
in Section 4.1.3 summarizes seven (7) crevice corrosion tests and reports that crevice corrosion 
was not observed. The U-bend tests discussed in Section 4.1.4 are also designed to produce 
crevice corrosion testing under the U-bend test washers. In these tests, no crevice corrosion is 
reported. Section 6.1.10 discusses the theoretical reasons why zirconium is immune to this type 
of corrosion. Yau and Webster 1987 ( p. 717) report that zirconium is among the structural 
metals that are most resistant to crevice corrosion. Greene et al. 2000 ( p. 7) also reported no 
ANI-WIS-MD-000008 REV 00 36 Apr112000 

crevice corrosion. In summary crevice corrosion has not been observed in zirconium and is not 
expected at YMP conditions. 
TSPA Disposition This FEP is addressed in CRWMS M&O 2000a 
IRSR Issues: CLSTI, CLST3, TSPAI3, TSPAI4 
References: Yau and Webster 1987 ,CRWMS M&O 2000a ,Greene et al. 2000, 
CRWMS M&O 2000a 
Basis for Screening Decision: Excluded based on screening argument presented above. 
6.2.8 High Dissolved Silica Content of Waters Enhances Corrosion of Cladding - YMP 
No. 2.1.02.18.00 
FEP Description It must be determined if the high dissolved silica content of waters 
enhances corrosion of cladding. 
Screening Decision: Exclude, low consequence 
Screening Argument: Silica is very stable. It is practically insoluble in water or acids, 
except hydrofluoric acid. It is not corrosive to most materials. However, there may be some 
fluoride contamination in silica, in which case it becomes corrosive to some materials, including 
zirconium, because fluorides in silica are soluble in water. For information on fluoride see YMP 
No. 2.1.02.16.00, Section 6.2.6. 
TSPA Disposition Silica content is not an issue, but the fluoride content that is in the 
silica is an issue. The fluoride issue is discussed in YMP No. 2.1.02.27.00. Therefore this FEP 
is excluded from further evaluation based on the arguments described below. 
IRSR Issues: CLST1, CLST3, TSPAI3, TSPAI4 
References: CRWMS M&O 2000a, Yau and Webster 1987 
Basis for Screening Decision 
Silica content is not an issue, but the fluoride content that is in the silica is an issue. Although 
frequently used as a corrosion resistant coating on materials, silicon dioxide reacts with 
hydrofluoric acid to form fluosilicic acid. Since fluosilicic acid is highly corrosive to zirconium 
even at room temperature, silicon dioxide can not be used as a technique to prevent attack of 
zirconium by acidic fluorides. The fluoride corrosion itself is addressed in another FEP ( YMP 
No. 2.1.02.27.00) and is included. The potential for silica itself degrading the cladding is 
negligible. Hansson (1984) reports corrosion tests with concrete pore fluids which normally 
contain silica. Yau (1983) reports corrosion tests in sea water which also contains silicon. 
Neither experimenters report significant corrosion. Both CRWMS M&O 2000a and Yau and 
Webster (1987, pp. Table 6) review the corrosion potentials for zirconium and show that it is 
resistant to most chemicals. 
ANI-WIS-MD-000008 REV 00 

6.2.9 Creep Rupture of Cladding - YMP No. 2.1.02.19.00 
FEP Description At high temperatures, creep rupture of Zircaloy cladding on spent 
fuel can occur and produce small perforations in the cladding to relieve stress. After the WP 
fails, the fuel can react with water and radioisotopes can thereby escape over time from the fuel 
rod. 
Screening Decision: Include. 
Screening Argument: Only moderately high temperatures (<350C) are anticipated on the 
cladding surface and thus, creep rupture would not be a failure mechanism under storage 
conditions. Failure of Zircaloy cladding by creep rupture is included in the CSNF Cladding 
Degradation Component of the Waste Form Degradation Model as a perforation mechanism. 
Inclusion of the cladding creep process permits the TSPA-SR to evaluate, statistically, the 
potential for a small number of cladding ruptures because of a wide distribution of temperatures 
that may occur in the repository and a distribution of initial stresses that may be present in the 
spent fuel rods. Inclusion also permits an objective assessment of the cladding creep failures 
during storage and transportation prior to receipt at the repositoty. 
TSPA Disposition The manner in which creep rupture is discussed in detail in 
CRWMS M&O (2000~). In summary, to estimate the percent of the rods that will perforate from 
creep rupture, a creep failure criterion based on data from Chung et al. (1987, pp. 780-781) was 
used. Chung et al. (1987) conducted a series of 20 pressure burst tests at slow strain rates and 
mandrel tests with irradiated cladding. Four tests were on BWR Zircaloy 2 cladding with a 
burnup of 22 MWdIkgU and the remaining 16 tests were on PWR Zircaloy 4 cladding with a 
burnup of 28 MWdIkgU. For 18 of the 20 tests, strains when perforation occurred averaged 
3.3% with a range of 0.4% to 11.7%. They inspected the perforations with a scanning electron 
microscope and found evidence that, in 11 of the 20 tests, the failures occurred at cracks formed 
in the cladding. Hence, the use of the Chung et al. (1987) results to estimate a failure criterion 
indirectly addresses the potential for lower failure strains from pre-existing internal cracks and 
the situation that perforation might not be caused by pure material creep. 
While the strain rates used in the experiments are faster than expected under repository 
conditions, they are not rapid burst tests and therefore are more applicable to repository 
conditions. The strain rates in Chung's gas pressurization experiments were such that the tests 
lasted between 1 and 312 hours with an average of 152 hours. The duration of the mandrel tests 
varied between 0.3 to 691.8 hours, with an average of 236 hours. Use of the Chung et al. (1987) 
empirical data to estimate a failure criterion implicitly assumes that the strain failure criterion 
conservatively bounds a failure criterion at slower strain rates. This assumption is valid since the 
creep tests summarized in CRWMS M&O (2000d, pp. 11-1 to 11-14) extend up to 10,000 hours 
(-1.14 year) and show creep strains, without failure, in excess of the failure criterion being used. 
In the CSNF Cladding Degradation Component, the potential for creep rupture is expressed 
statistically (CRWMS M&O 2000~) For a sampling of twelve thousand fuel rods and an 
average WP peak surface temperature of 277OC, the center rod peak temperature was 30SC and 
the rod failure distribution was: 
ANI-WIS-MD-000008 REV 00 

Lower limit = 1.06 % 
Mean - - 3.39 % 
Upper limit = 24.8 % 
Most of the predicted creep failures occurred in the rods with the highest initial stresses and 
during dry storage and transportation to the repository when temperatures above 350C were 
possible. Only when WP peak surface temperatures reached 300C did rod perforations from 
creep rupture start to be influential while in the repository. 
The FEP Thermally-Induced Stress Changes in Waste and EBS (YMP No. 2.1.11.07.00) is 
included in the TSPA-SR in that the stress in the fuel rods is adjusted for the temperature at each 
time and rod location when analyzing both creep damage and the potential for SSC. 
IRSR Issues: CLST1, CLST3, TSPAI4 
References: CRWMS M&O 2000c, CRWMS M&O 2000d, Chung et al., 1987 
Basis for Screening Decision 
No further elaboration in addition to the Screening Argument is necessq. 
6.2.10 Pressurization From Helium Production Causes Cladding Failure - YMP No. 
2.1.02.20.00 
FEP Description Increased pressure within the fuel rod due to the production of 
helium gas could contribute to cladding failure. 
Screening Decision: Include. 
Screening Argument: Both delayed hydride cracking (DHC) 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. 
TSPA Disposition This FEP is addressed in CRWMS M&O (2000b, Section 6.3.4). 
Both delayed hydride cracking (DHC) 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 Manaktala (1993, Figure 3-4, p. 3-12) presents 
the helium pressure buildup for 100 "C as a function of time for a PWR fuel rod with 36 
MWdIkgU burnup and an assumed 100% helium release from the fuel into the fuel rod gap. This 
figure was used to develop an equation for helium buildup in a fuel rod. In TSPA-SR, the creep 
analysis was performed including the helium buildup at 100 years, a time well after the peak 
thermal period. 
IRSR Issues: CLST1, CLST3, TSPAI4 
ANI-WIS-MD-000008 REV 00 39 

References: Manaktala, H.K. 1993, CRWMS M&O 2000b 
Basis for Screening Decision 
No further elaboration in addition to the Screening Argument is necessary 
6.2.11 Stress Corrosion Cracking (SCC) of Cladding - YMP No. 2.1.02.21.00 
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. 
Screening Decision: Include. 
Screening Argument: SCC requires a susceptible material, an aggressive chemical 
environment, and high stress levels. Iodine-induced SCC requires an iodine concentration in the 
fuel-cladding gap greater than 5 x g/cm2 (Cunningham et al. 1987, pp. A.4, A.5). Below this 
threshold of free iodine concentration, Zircaloy cracking due to SCC has not been observed. In 
actual fuel rods, free iodine concentrations are expected to be negligible. However, over 
sufficiently long times at high stress and elevated temperatures the iodine may be present in 
sufficient quantities for SCC, such that, once cracking starts, there is sufficient time to propagate 
through the cladding (without considering crack velocities). As reported by Tasooji et al. (1984, 
p.600, their Figure 3), a rod with a stress above 180 MPa fails by SCC. Although few rods have 
such high stresses and the same rods that fail from SCC are also prone to fail from creep, the 
inclusion of SCC increases the mean failure from creep alone from 2.0% to 2.4% for dry storage 
and transportation (CRWMS M&O 2000~). 
The SCC model compares the existing stress with a SCC threshold stress and considers neither 
the crack size, chemical environment, nor the crack location. Therefore this model is applicable 
for Inside-Out SCC (FEP YMP No. 2.1.02.21.01) or Outside-In SCC (FEP YMP No. 
2.1.02.21.02). 
TSPA Disposition Over sufficiently long times at high stress and elevated 
temperatures, and with a sufficient concentration of free iodine present, SCC may occur in a few 
fuel rods. Once cracking starts, there is sufficient time to propagate through the cladding. 
IRSR Issues: CLST1, CLST3, TSPAI4 
References: Cunningham, M.E. ; Simonen, E.P.; Allemann, R.T.; Levy, I.S.; 
and Hazelton, R.F. 1987, Tasooji, A,; Einziger, R.E.; and Miller, A.K. 1984, CRWMS M&O 
2000c 
Basis for Screening Decision 
No further elaboration in addition to the Screening Argument is necessary. 
ANI-WIS-MD-000008 REV 00 

6.2.12 Hydride Embrittlement of Cladding - YMP No. 2.1.02.22.00 
FEP Description Cladding contains hydrides after reactor operation. The cladding 
might also pick up more hydrogen from cladding general corrosion (wet oxidation) after the WP 
is breached. Hydrides might also form from U02 oxidation (after WP 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. 
Screening Decision: Exclude, low probability. 
Screening Argument: Hydrogen will be generated in the WP as the WP internals corrode. 
This hydrogen is not expected to be absorbed directly by the fuel cladding, because the H2 
molecules are not expected to migrate through the high-density Zr02 layer on the fuel cladding. 
Available data on zirconium hydriding indicate that corrosion of waste package internals will not 
result in hydriding of fuel cladding. 
Although hydriding of Zircaloy cladding as a result of galvanic corrosion has been observed, 
Zircaloy cladding alloys resist galvanic corrosion when contacting carbon steel, and, if such a 
galvanic corrosion effect should occur, corrosion of the carbon steel would quickly break the 
electrical contact between the two materials. 
Failure of the cladding by hydride reorientation is unlikely. Stresses and temperatures are too 
low for hydride reorientation to occur and the cladding material will maintain sufficient strength 
even if hydride reorientation did occur such that failure would not be expected. 
Failure of the cladding by DHC is unlikely and has not been included in the abstraction for the 
TSPA-SR. Stresses (and stress intensity factors) are too low for crack propagation. 
Failure of the cladding by hydrogen embrittlement is unlikely. Hydrogen absorption from fuel 
cladding surface oxidation and WP corrosion will be negligible. Hydrogen migration will be 
limited at the temperatures expected during disposal. 
TSPA Disposition Exclude 
IRSR Issues: CLST1, CLST3, TSPAI4 
References: Baker 1992, ASTM B811, ASTM B855, Clayton, J. C. 1984, 
CRWMS M&O 1997, Cunningham, M.E.; Simonen, E.P.; Allemann, R.T.; Levy, I.S.; and 
Hazelton, R.F. 1987, Dieter, G.E. 1961, Einziger, R.E. et. al., 1982, Einziger, R.E. and Kohli, R. 
1984, Garzarolli et al. 1979, IAEA 1998, Kreyns, P.H.; Bourgeois, W.F.; White, C.J.; 
Charpentier, P.L.; Kammenzind, B.F.; and Franklin, D.G. 1996, Laming, D.D.; Beyer, C.E.; and 
Painter, C.L. 1997, Mahmood, S.T.; Farkas, D.M.; Adamson, R.B.; and Etoh, Y. 1998, 
McEachern, R.J. and Taylor, P. 1997, McMinn, A,; Darby, E.D.; and Schofield, J.S. 2000, Peehs, 
M. and Fleisch, J. 1986, Pescatore, C.; Cowgill, M.G.; and Sullivan, T.M. 1990, Puls, M.P. 1988, 
Reed-Hill, R.E. 1973, Rothman, A.J. 1984, Shi, S.Q. and Puls, M.P. 1994, Wasywich, K.M. and 
Frost, C.R. 1992, Yau, T. L., 1984, Yau, T.L. 1983, YMP 1998, Clayton 1989, CRWMS M&O 
ANI-WIS-MD-000008 REV 00 41 Apr112000 

2000b, CRWMS M&O 2000g, Smith 1966, Huang 1995, Peehs 1998, Paurbaix 1974, Mardon et 
al. 1997, ASM 1992. 
Basis for Screening Decision 
Hydrogen embrittlement results in a generally reduced resistance to fracture. In Zircaloy, 
hydrogen embrittlement is normally caused by formation of microscopic flakes of zirconium 
hydride. Since the hydride flakes 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 value of the stress intensity factor 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, p. 767, Figure 5) show that such 
hydrides could decrease the fracture toughness (KIc) from 42 MPa.m O 5 to 8 ~ ~ a . m ' ~ as the 
hydrogen content increases from zero to 4000 ppm. As shown in CRWMS M&O 2000b 
(Section 6.10.2) the maximum stress intensity (KI) for the statistical distribution of rods and 
crack sizes at 260C is 2.7 MPa mn and therefore even with hydride concentrations of 4000 
ppm, failure is not expected. In the limit, 100 percent hydride and no metal at all, the fracture 
toughness is about 1 ~ ~ a . m ' ~ . 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. 
6.2.12.1 Hydride Embrittlement from Zirconium Corrosion - YMP No. 2.1.02.22.01 
Additional hydrides could be formed by general corrosion of the cladding. WPs are assumed to 
fail at various times, permitting water or steam to enter the failed WP. Wet oxidation (or steam 
oxidation if the local temperature is above boiling) occurs, and approximately 17 percent 
(Lanning et al. 1997, Vol. 1, p. 8.4, Figure 8.2) of the hydrogen released from the water is 
absorbed by the fuel cladding in the failed WP. However, the fuel cladding picks up very little 
hydrogen because the corrosion rate is so slow. Even the hottest rods will pick up only small 
quantities of hydrogen from cladding surface general corrosion if the WP remains sealed for the 
first 100 years. 
6.2.12.2 Hydride Embrittlement from WP Corrosion & H2 Absorption - YMP No. 
2.1.02.22.02 
Many investigators have considered the hydriding of zirconium alloys (for example, the papers 
cited in Clayton 1989, 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. IAEA 1998 
(p. 92) discusses experiments in which zirconium-base alloys were oxidized in normal water 
(H20) with dissolved tritium gas (T2). This experiment is extremely sensitive. If even one part in 
lo6 of the hydrogen was from dissolved 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 (p. 92) gives the following discussion of the experiment: 
Oxidation studies using T2/H20 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 
ANL-WIS-MD-000008 REV 00 42 April 2000 

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 the quotation given above, the hydrogen pressures required to cause 
hydriding are quite large. By comparison, the highest credible hydrogen pressure in a breached 
waste package is pure hydrogen gas at atmospheric pressure, or about 0.1 MPa. 
It is instructive to examine why high hydrogen pressures can cause hydriding. Even at 200 OC, 
the solubility of oxygen in zirconium is greater than 1% by weight, and the solubility of oxygen 
increases with increasing temperature (Baker 1992, p. 2-326). As a result, the oxide film on the 
surface of a piece of zirconium is normally not thermodynamically stable. In its discussion of 
experimental procedure, Smith (1966 p. 325) notes "Zirconium samples were first annealed at 
700" C under vacuum (- mm Hg) to remove any oxide film. The film dissolved into the 
sample, 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. 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, p. 226). 
The effect of the oxygen supply has been studied. Garzarolli et al. (1979, p. 64) state that: 
the effect of the composition of the gas atmosphere on the electrical properties of ZrO, 
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. 
The conditions for hydriding have been quantified. Clayton (1989, p. 270) presents the equation 
where p ~ , o and p ~ , are the pressures of Hz0 and HZ, respectively. For Equation 2 to be 
applicable, both pressures must be given in torr (millimeters of Hg head). (This fact is deduced 
as follows. First, Clayton (1989, 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 values of "Critical P ~ , ~ / ~ H , " given in the tables are consistent with both . . 
values being in the same units.) ForpH, = 101 kPa = 760 tom, it is found that mzo (protective) = 
0.24 kPa = 1.8 tom. Note that, by atmospheric standards, this p ~ , o corresponds to a dry gas. For 
comparison, the vapor pressure of water at 25 "C is about 3.2 kPa. For this small amount of 
humidity, corrosion of the waste package internals is not credible. However, 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 (Table 3) gives data on hydrogen absorption. However, the 
environment for these experiments was hot, extremely pure hydrogen. Smith 1966 (Table 3) 
states that the temperatures for the hydrogen absorption experiments were 210 OC to 700 OC. 
Smith 1966 (p. 325) notes that "hydrogen was purified by passing it through a Deoxo unit, a bed 
of platinized asbestos (300" C), a tube of P2O5 and a liquid nitrogen trap." The evident intention 
ANI-WIS-MD-000008 REV 00 43 Apr112000 

is to react any oxygen impurities, and absorb or condense any water vapor that is formed. Mer 
this treatment, vety 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 2 was developed. Data on hydriding at lower temperatures were not 
available in the literature. 
In some respects, it can be argued that it is conservative to apply Equation 2 to a repositoty. 
Equation 2 was developed for coupons of Zircaloy-2. This alloy is more susceptible to hydriding 
than is Zircaloy-4 (Clayton 1989, Table 5), and coupons are more susceptible to hydriding than 
is tubing (Clayton 1989, Table 5). Therefore, Equation 2 should give conservative predictions of 
the susceptibility of spent fuel cladding to hydriding. 
In summary, hydrogen will be generated in the WP as the WP internals corrode. This hydrogen 
is not directly absorbed, because the H2 molecules do not migrate through the high-density Zr02 
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. This 
degradation mode is excluded from TSPA-SR. 
6.2.12.3 Hydride Embrittlement from Galvanic Corrosion of WP Contacting 
Cladding - YMP No. 2.1.02.22.03 
In current waste package designs, the fuel assemblies will be in contact with carbon steel tubes 
(CRWMS M&O 1997). Since the fuel cladding and WP basket tubes are of different materials, 
there may be galvanic corrosion with hydrogen charging of the fuel assembly. This would occur 
if there were electrical contact between the assembly and the fuel basket tube, oxidation of iron 
from the carbon steel fuel basket tubes, and reduction of hydrogen ions on the fuel assembly 
cladding. In addition, it would be necessary that at least some of the reduced hydrogen be 
absorbed by the metal of the fuel assembly cladding. 
It should be noted that there is only a limited period during which galvanic corrosion could occur 
in a specific WP. The fuel basket tubes will be completely corroded away within tens to a few 
hundreds of years after individual WP failures (YMP 1998, p. c-17). By the time the fuel basket 
tubes are completely corroded away in a specific failed WP, a thick layer of corrosion products 
will electrically insulate the fuel assembly cladding. 
It is important to understand the conditions for galvanic hydrogen charging. These can be 
inferred from results given in Yau (1983, 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 were considered: Zr 
702, Zr 704 with nickel, and Zr 704 without nickel. Each composition was tested in both 
unwelded and welded conditions. Of the six samples, only the welded sample of Zr 704 with 
nickel showed hydrogen pickup. ("Zr 702" and "Zr 704" are taken to be UNS R607024 and 
R6070, respectively. ASTM B55UB551M-97 (Table 1) gives composition limits for these 
materials; the maximum hydrogen content is 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. 
ANI-WIS-MD-000008 REV 00 44 Apr112000 

Although Yau (1983, p. 2612) does not give the compositions of the samples of Zr 704 with 
nickel and Zr 704 without nickel, it is instructive to compare the amount of nickel in various 
zirconium alloys. According to the minimum concentrations for other elements, the specification 
for UNS R60704 admits up to 1.3% nickel, and that for UNS R60702 admits up to 0.8% nickel 
(ASTM B551, Table 1). UNS R60802 (Zircaloy-2) contains 0.03% to 0.08% nickel, and UNS 
R60804 (Zircaloy-4) contains no more than 0.007% nickel (ASTM B811-90, Table 2). It is clear 
that the Zircaloys contain very 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 the fuel cladding will be separated from the basket tubes by the fuel 
channels. 
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. Clayton (1984, 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, 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 very localized effect was noted for Zircaloy-4 smeared with Inconel alloy (4.6% 
nickel - 16.7% chromium - 3.5% copper - 75.2% 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, p. 587), "Nickel 
alloy smearing and bonding to filmed Zircaloy, rather than just tight surface contact, is necessary 
for accelerated hydriding." Fuel assemblies will be lowered carefully into waste package baskets. 
The 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. 
If it is nevertheless 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, p. 572) notes that the initial high 
accelerated hydrogen ingress rate was effectively shut off during exposure at 271 "C in about 25 
days. Because of the extremely small size of contact spots, corrosion of the carbon steel would 
quickly break the electrical contact between the two materials. Since the time of contact will be 
small, the amount of hydrogen absorbed as a result of contact will also be small. This conclusion 
is supported by the discussion of U-bend hydriding above. 
ANI-WIS-MD-000008 REV 00 45 

The data above on zirconium hydriding indicate that corrosion of waste package internals will 
not result in hydriding of fuel cladding. Although hydriding as a result of galvanic corrosion has 
been observed, cladding alloys resist such corrosion on contact with carbon steel, and, if such an 
effect should occur, corrosion of the carbon steel would quickly break the electrical contact 
between the materials. 
6.2.12.4 Delayed Hydride Cracking (DHC) of Cladding - YMP No. 2.1.02.22.04 
CRWMS M&O (2000g, Section 6.3) reviews DHC and CRWMS M&O (2000b, Section 6.10.2) 
analyzes DHC with the predicted crack size distribution and rod stress distribution expected at 
YMP. The general DHC process begins with precipitation of a hydride at the crack tip. The 
stress gradient at the crack tip provides a driving force for preferential hydride precipitation in 
the tensile stress field in front of the crack tip. If the crack tip stress is greater than the fracture 
stress of the brittle hydride, the hydride fractures and crack growth occurs. 
There are three basic requirements for DHC: (1) incipient defects or cracks; (2) presence of 
hydride at or near the crack tip; and (3) sufficient stress to propagate the crack. Incipient cracks 
and defects are present in the cladding as a result of the manufacturing process and from 
irradiation. Hydrogen is present as a result of Zircaloy alloy impurities and in-reactor corrosion. 
Cooling of the fuel cladding after reactor operations (irradiation) promotes precipitation of 
hydrides. The high stresses required to crack a hydride exist only in front of a deep crack. The 
critical stress intensity for DHC is much greater than cladding stress intensities expected during 
dry storage. Additionally, at the high stress intensities required for DHC, SCC is the dominant 
crack growth mechanism. 
During delayed hydride cracking, hydrides slowly form at a crack until the crack propagates 
through the hydride region at the crack tip and stops. This sequence repeats itself and the crack 
propagates slowly through the metal. The hydrides preferentially collect at the crack tip because 
tensile stress reduces 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 2000a), the stress intensities (KI) will be calculated and 
compared to the KIH. If KI > KIH, then the crack will start to propagate and, because of long 
repositoty times, it is assumed that failure will occur. DHC failure occurred in some zirconium 
coolant tubes in a Candu reactor where high temperature gradients caused excess hydride 
buildup in a specific location. 
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 length. 
For a sharp crack a limiting case, the stress intensity factor is (Reed-Hill 1973, p. 800) 
where 
KI = Stress intensity factor, ~ ~ a . r n ' ~ 
o = Cladding stress, MPa 
w = Crack depth, m 
ANI-WIS-MD-000008 REV 00 

Equation 3 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, p. 194). 
DHC is analyzed in CRWMS M&O (2000b, Section 6.10.2). The calculated crack size 
distribution is given in CRWMS M&O (2000b, Section 6.6 Figure 17). The median (P = 
50 percent) value is 13 pm and the average crack is 18.6 pm. The largest size crack of the 2000 
samples is 119 pm. The calculated stress distribution (CRWMS M&O 2000b, Section 6.7) is 
given in Figures 18 and 26 of that AMR. DHC is unlikely at temperatures above 260C 
(Mahmood et al. 1998, p.20), because of the plasticity of the material. Rothman (1984, p. 37) 
reports that DHC is unlikely above 250C because of the plasticity of the material. For this 
calculation, the temperature of 260C is used and the pressure is adjusted accordingly. The crack 
size distribution and stress distribution can be combined to give the distribution of stress 
intensity factors, KT, in Figure 27 of that AMR. Some of the properties of this distribution of the 
stress intensity factor are: 
Mean: 0.47 M~a-m'.~ 
Median: 0.40 MPa-mo5 
maximum: 2.7 MPa-mo5 
minimum: 1.61E-03 MPa-mo5 
The work of Shi and Puls (1994, p. 239, Fig. 7), shows experimental KIH in the range of 5 to 12 
MPa-mo for zirconium alloy containing 2.5% Nb. Rothman (1984, p. 37), reports a KIH of 6 
MPa-mo for Zircaloy-2. Pescatore et al. (1990, Table 6, p.50) report values of 5 and 14. Huang 
(1995, p. 195) shows KIH for irradiated Zircaloy-2 approaching 6 MPa-mo 5. For this AMR 
Huang's and Rothman's value for irradiated cladding of 6 MPa-mo was used. Because the 
observed values of K I are well below these values, exactly which value of K IH is selected is not 
important. No KI values in this AMR's sampling of rods are near the threshold stress intensity 
value. The maximum observed KI was 2.7 MPa-mo and the mean value was 0.47 MPa-mo 5. 
Rothman (1984, pp. 33 - 39) reviewed DHC in Zircaloy cladding in a repository. Rothman 
concludes that DHC is unlikely unless the fuel rods have large existing cracks (exceeding 
approximately 50 percent of wall thickness) and very high stresses (exceeding approximately 
137 MPa). He also 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, pp. 5, 6) concluded that neither DHC nor hydride reorientation 
would occur in dry storage. 
In conclusion, failure of the cladding by DHC in dry storage is unlikely and has not been 
included in the abstraction for the TSPA-SR analysis. Stresses (and stress intensity factors) are 
too low for crack propagation. Since the threshold for embrittlement failure is always greater 
than the threshold for DHC ( K I ~ > KIH), embrittlement failures are also not expected. 
ANI-WIS-MD-000008 REV 00 47 Apr112000 

6.2.12.5 Hydride Reorientation - YMP No. 2.1.02.22.05 
CRWMS M&O (2000g, Section 6.4) reviews 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. Normally, the hydrides are oriented in the 
circumferential direction and do not weaken the cladding against a hoop stress. Reorientation 
usually occurs under tensile stresses ranging from 69 to 208 MPa. Test rods at the low end of this 
range, but significantly higher than the stress expected at the repository temperatures, showed no 
reorientation of hydrides. 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, 
p. 65). 
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. In one dry storage test, reorientation to the radial direction was 
observed in one rod (Einziger and Kohli 1984, p. 119) although rod failure did not occur. This 
occurred in a fuel rod that had very high stresses (145 MPa at 323'C). 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. 
Figure 1 summarizes the data that Pescatore et al. (1990, pp. 52-55) collected showing 
reorientation. The points from Marshall show different fractions of reorientation. Hardie's point 
was reported at minimum stress. The straight solid line suggests a region in temperature and 
stress where reorientation occurs. This figure shows that at the cladding maximum temperature 
of 35OoC, the stress would have to be greater than 120 MPa before reorientation would start to 
occur. This corresponds to about 60 MPa at room temperature. From CRWMS M&O 2000b 
(Figure 26), it is estimated that 5% of the rods could have stresses this high and might undergo 
some reorientation. Rothman (1984) also studied cladding degradation in a repository and 
concluded that hydride reorientation would not occur. He was not considering fuel with burnups 
and stresses as high as considered in CRWMS M&O 2000b. The predicted cladding 
temperatures will be less than 350C and therefore this comparison should be considered an 
upper bounding case. 
Pescatore et al. (1990, pp. 54, 69) state that even with hydride reorientation, stress levels will be 
insufficient to result in DHC and clad failure. For general DHC, they also said that crack 
propagation would tend to stop when the crack propagates into a hydride platelet. Puls (1988, p. 
1507-1522) performed a series of strain tests on Zircaloy-2 with reoriented hydrides. His results 
are summarized in Table 3. He took samples of Candu coolant tubing and performed strain tests 
in the circumferential direction. The initial 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 
room 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 3). Other samples 
were cooled from 350C at a stress of 200 MPa. This stress is more than twice than expected for 
ANI-WIS-MD-000008 REV 00 

380 
360 
3 340 " 320 
300 + $ 280 
260 
240 
Hardie minimum o 
220 
200 
120 140 160 180 200 220 
Stress (M Pa) 
Figure 1. Temperature and Stress at which Hydride Reorientation was Observed for Cold Worked 
Zircaloy-2, 4 (Pescatore et al. 1990) 
most cladding in the repository (CRWMS M&O 2000b, Figure 26). Longer hydrides were 
produced by cooling the samples in a furnace, while shorter ones were produced by bench 
cooling the samples. Table 3 gives the range of hydride lengths. Two types of samples were 
used, smooth ones and notched ones (designated by n in the table). The table gives the stress 
(oy) for which 0.2 percent strain was measured. Also given 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 0.2percent 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). 
AM-WIS-MDOOOOO8 REV 00 April 2000 

Table 3. Puls' Zircaloy-2 Strain Tests on Zirconium with Reoriented Hydrides 
A It - t - 
-lype, y= nyariae reorienrea near yiela srress, n - norcnea, a = arresred (resr rermmarea oerore failure) 
i 
Source: M.P. Puls (1988, Tables 1,3,6) 
JU-JU I IOU '1. 11 13.0 
Hydride reorientation might require all the hydride platelets to be dissolved before cool-down 
50-90 
50-90 
50-90 
50.90 
50-90 
and reprecipitation starts. This could be necessary because the hydrides prefer to condense at 
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 35OoC, the initial 
concentration must be less than 120 ppm (see Table 4). CRWMS (M&O 2000b, Figure 15) 
shows that only about 13 percent of the fuel has average concentrations this low. Knowing that 
the hydride concentration is directly proportional to the oxide thickness, CRWMS M&O (2000b, 
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 (their Figure 3, p. 408) shows hydride 
content as a function of burnup. To have less than 120 ppm, the figure shows that Zircaloy 4 
fuel with burnups less than 22 MWdIkgU has low enough hydride concentrations to dissolve all 
the hydrides at 350C. Again, these fuels would have the lowest stresses. The rod that Einziger 
observed reorientation was exposed to a 570C transient which would support a solubility of 729 
ppm (Table 4). It is possible that all the hydrides were dissolved in that experiment. 
In summary, failure of the cladding by hydride reorientation is unlikely and has not been 
included in the abstraction for TSPA-SR. Stresses and temperatures are too low for hydride 
reorientation to occur in most of the fuel and the cladding material will maintain sufficient 
toughness even if hydride reorientation did occur such that failure would not be expected. 
- 1 
Y 
n 
n. Y 
n 
AM-WIS-MDOOOOO8 REV 00 50 April 2000 
689 
625 
721 
923 
81 1 
741 
647 
803 
1032 
936 
4.1113.6 
4.1113.6 

Table4. Saturation Limits for Hydrogen in Zirconium as a Function of Temperature 
Temperature ("C) I H Concentration (ppm) I 
Arl I n I I 
Temperature ("C) 
300 
310 
320 
330 - 
I I 
Source: Pescatore et al. (1990. eq. 6) 
H Concentration (ppm) 
66 
75 
85 
96 - - 
I 44u I LOO 
6.2.12.6 Hydride Axial Migration - YMP No. 2.1.02.22.06 
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 
condense there. The effect was studied for dry storage (Cunningham et al. 1987, Appendix C) 
and for a 90-year period was determined not to be a problem. As the repository cools, the 
driving force for this redistribution (hydrogen solubility, temperature gradient, and diffusion 
rate) all decrease. The WP internals will act to minimize the temperature variation along the 
length of the fuel assembly. McMinn et al. (2000, Figure 15) show that a minimum of 42C 
temperature difference between the hot location and cold location in the cladding is needed to 
move the hydrogen. At temperatures below 20OoC, temperature differences over 75C are 
needed to move the hydrogen. It is unlikely that sufficient hydrogen can be moved because of a 
lack of large temperature gradients in the WPs. 
E E 
450 
6.2.12.7 Hydride Embrittlement from Fuel Reaction - YMP No. 2.1.02.22.07 
Hydrides can be formed in the cladding from the oxidation of UOz fuel (Wasywich and Frost 
1992, p. 1171). The fact that the fuel is oxidizing means that the cladding has already been 
perforated. This type of cladding degradation is only observed in BWRs and a high temperature 
steam environment is required for failure propagation. Such conditions are unlikely at YMP 
31 3 
ANI-WIS-MD-000008 REV 00 51 Apr112000 
I 
460 
470 
480 
490 
500 
510 
520 - 
339 
367 
396 
427 
459 
493 
529 - - 

after WP failure. If the embrittlement becomes severe and the cladding is mechanically loaded 
in excess of its fracture toughness, the cladding could fail exposing additional fuel. This type of 
failure has a small effect on the fuel oxidation rate since the fuel oxidizes after the first 
penetration. It does affect the amounts and rate of fission product migrating out of the secondary 
phases of the fuel dissolution. This type of secondary cladding failure has not been modeled. 
6.2.13 Cladding Unzipping - YMP No. 2.1.02.23.00 
FEP Description In either dry or wet oxidizing conditions and with perforated fuel 
cladding, the U02 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). 
Screening Decision: Dry Oxidation - Exclude, low probability. 
Wet Oxidation Include. 
Screening Argument: In the TSPA-SR model, wet unzipping is included as the key 
element in exposing and dissolving fuel from damaged cladding. Dry unzipping has been 
excluded because it is not expected to occur if the WP remains sealed for a few hundred years 
after repository closure and, if it were to occur, only a very small fraction of the fuel would 
undergo dry unzipping. 
TSPA Disposition Dry Oxidation - Exclude, low probability. 
Wet Oxidation - included as described in CRWMS M&O 2000e, 
and abstracted into TSPA-SR in CRWMS M&O 2000c. 
IRSR Issuance: CLST1, CLST3, TSPAI4 
References: Doubt, G. 1984, Einziger, R.E.; Atkin, S.D.; Stellrecht, D.E.; and 
Pasupathi, V. 1982, Gray, W.J. and Wilson, C.N. 1995, Hanson, B.D. 1998, IAEA 1988, 
Johnson, A.B., Jr. 1977, Johnson, L.H. and Taylor, P.; 1998, McEachern, R.J. and Taylor, P. 
1997, CRWMS M&O 1998, CRWMS M&O 1999c, CRWMS M&O 2000h, CRWMS M&O 
2000e 
Basis for Screening Decision 
6.2.13.1 Dry Oxidation of Fuel - YMP No. 2.1.02.23.02 
The cladding that has failed during reactor operation, from creep rupture, or from DHC generally 
has small cladding perforations and very small areas of fuel exposed to the near-field 
environment. When the WP fails, the inert atmosphere inside the WP is replaced with the air and 
steam from the near-field environment. After WP failure, the near-field environment is almost 
completely steam and dry oxidation is not expected. If the dry oxidizing conditions existed after 
WP failure and with perforated cladding, the UOZ phase in the spent fuel can oxidize in two 
stages, first to U409 and then to U3O8, causing an increase of the spent fuel matrix volume. 
These are nominal formulas only and do not necessarily describe the actual stoichiometry of 
these oxides. The stress from the volume increase can unzip the clad, causing a gross rupture of 
ANI-WIS-MD-000008 REV 00 52 Apr112000 

the clad and exposing the fuel inside. Oxidation also produces an increase in surface area and 
dissolution rate, the combined factor being 150 (Gray and Wilson 1995, p. vii). Dry fuel 
oxidation has been experimentally studied by Einziger et al. (1982, p. 65) and Hanson (1998, p. 
2.3). McEachern and Taylor (1997, p. i) summarize the work of others. An improved model for 
dry unzipping is presented in (CRWMS M&O 2000h). 
For defective WPs (assumed to have failed at emplacement), it may be possible to unzip the 
cladding of the fuel rods that are perforated. This will probably not actually occur because the 
environment is mostly steam and the 0 2 partial pressure is low at this early time. Therefore, dry 
oxidation is unlikely. However if it is assumed that dry oxidation does occur at this early time, 
the degree of unzipping for defective WPs can be estimated. The combination of defective WPs, 
which is taken as 0.02 percent (Doubt 1984, p. 30) with premature rod failure, which is 
1.25 percent including the stainless steel cladding, gives the fraction of rods expected to be 
converted to U3O8 to be 2.5 x (= 2 x l o 4 x 1.25 x lo-'). Overall, when considering the 
presence of steam, timing of WP failures and decreasing temperatures, it is expected that only 
very small fractions of fuel will be converted to U308, and these small fractions will not affect 
the PA. Earlier analysis presented in the TSPA-VA (CRWMS M&O 1998, Section 6.3.1.1.9) 
showed that the cladding could unzip if both the cladding is perforated and the WP also fails 
within the first 200 years. Such early failures of the WP are not expected. 
6.2.13.2 Wet Oxidation of Fuel - YMP No. 2.1.02.23.03 
Cladding could unzip in a wet environment. As the U02 is converted to secondary phases such 
as metaschoepite and Na-boltwoodite, the fuel volume increases. These phases will only form 
where there is available liquid volume, or, alternatively, they might form at solid surfaces and 
tear the cladding. Cladding unzipping has not been observed in spent fuel pools where fuel has 
been stored for tens of years as noted by the IAEA (1988, Table XXVI) and Johnson (1977, p. 
20). The Canadians tested a defective rod in a steam autoclave for ten years and did not measure 
any cladding strain (Johnson and Taylor 1998, p. 10). While wet unzipping has not been 
observed, the time periods are very short compared to the repository scale and a form of wet 
oxidation and secondary phase formation might tear the cladding. 
The current wet unzipping model is developed in CRWMS M&O 2000e and the abstraction is 
described in CRWMS M&O 2000c (Section 6.6). All failed Zircaloy clad rods, all stainless steel 
clad rods (assumed to be failed) are available to unzip. The unzipping velocity is dependent on 
the fuel intrinsic dissolution rate. This rate is evaluated at each TSPA-SR time step because it 
depends on the temperature and chemistry inside the WP. 
6.2.13.3 Volume Increase of Corrosion Products - YMP No. 2.1.09.03.00 
The volume increase of corrosion products is included in the TSPA-SR. It is this volume 
increase from the corrosion of U02 to other minerals that drives the cladding wet unzipping 
described above. 
6.2.13.4 Gap and Grain release of Cs, I - YMP No. 2.1.02.07.00 
The release of fission products from the gap and grain boundaries is included in the TSPA-SR 
and described in CRWMS M&O 2000c (Section 6.5) 
ANI-WIS-MD-000008 REV 00 

6.2.14 Mechanical Failure (of Cladding) - YMP No. 2.1.02.24.00 
FEP Description Mechanical failure of cladding may result from external stresses, 
such as ground motion and rockfall during earthquakes. 
Screening Decision: Include. 
Screening Argument: The analysis of seismic events is included in the TSPA-SR. 
Seismic events with a frequency of 1.1~10.~ eventslyear are assumed to break all of the fuel 
cladding. Such events are sampled, and, when such an event occurs, all cladding is failed and is 
considered to be available for unzipping. 
The effect of a rubble bed consisting of rocks from a drift collapse on bare fuel rods (no WP or 
possible drip shield protection) is not included in TSPA-SR because this would not occur until 
well after the 10,000 years considered for the TSPA-SR. Since high accelerations are needed to 
fail the cladding, rock drops onto an intact WP will not cause rod failure and therefore was not 
included in the TSPA-SR. 
TSPA Disposition Mechanical failure will be represented in the performance 
assessment for the TSPA-SR. 
IRSR Issues: CLST1, CLST3, TSPAI4 
References: CRWMS M&O 1999b, Witte et al. 1989 
Basis for Screening Decision 
Seismic analysis (CRWMS M&O 1999b, Section 6.1) shows that most of the rods in the WPs 
would fail from a very severe earthquake (about once per million years event) but no rods would 
fail for less severe and moderate frequency seismic events. This is consistent with studies (Witte 
et al. 1989, p. 194) of rod damage during transportation accidents that concluded that 63 g 
accelerations are needed to fail the rods in the shipping container (or WP). Therefore, the 
seismic failures have been included in the TSPA-SR as a disruptive event. 
The analysis of seismic events is included in the TSPA-SR. Based on the analysis (CRWMS 
M&O 1999b, Section 6.1), seismic events with a frequency of 1 . 1 ~ 1 0 - ~ eventslyear would break 
most of the fuel. Such events are sampled, and, when such an event occurs, all cladding is failed 
and is considered to be available for unzipping. 
6.2.14.1 Rockfall - YMP No. 2.1.07.01.00 
CRWMS M&O 1999b (Section 6.2) also considered the effect of a rubble bed consisting of 
rocks from a drift collapse on bare fuel rods (no WP or possible drip shield protection). The 
analysis showed that the bare fuel assemblies would fail under the static loading of the rocks. 
This effect was not included in the cladding degradation abstraction because it does not occur 
until after the WP no longer affords any protection. The first patches (about 100 cm by 100 cm) 
penetrate the WP in 50,000 to 60,000 years, and a significant number of patches (about 100 
patches) are open in about 350,000 years. With the potential for rubble bed damage not occuring 
until well after the 10,000 years considered for the TSPA-SR, rubble bed damage to the cladding 
ANI-WIS-MD-000008 REV 00 54 Apr112000 

was neglected. Witte et al. 1989 (p. 194, Table 3) analyzed rod damage during transportation 
accidents that concluded that 63 g accelerations are needed to fail the rods in the shipping 
container (or WP). This high acceleration indicates that rock drops onto an intact WP will not 
cause rod failure and therefore this FEP was not included in the TSPA-SR. 
6.2.15 Localized Corrosion Perforation from Fluoride - YMP No. 2.1.02.27.00 
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 WPs. 
Screening Decision: Include. 
Screening Argument: Numerous processes have been identified that might cause 
localized corrosion and eventually perforation of Zircaloy cladding on CSNF. Yet as 
summarized in FEP 2.1.02.16.00 and discussed in more detail in the AMR, Clad Degvadation- 
Local Covvosion of Zivconium and Its Alloys Undev Reposito~) Conditions (ANL-EBS-MD- 
000012) (CRWMS M&O 2000a), the conditions necessary for these processes generally do not 
exist in the repository. Even pitting corrosion in the presence of fluorides does not appear likely 
since the concentration of fluoride in pore water at Yucca Mountain is low, this low initial 
amount of fluoride is consumed during the corrosion process, and the pH of the in-package 
chemistry does not drop low enough to promote corrosion. Localized corrosion in the present of 
fluorides was included in the TSPA-SR to account for modeling uncertainty of the in-package 
chemistry since conditions for fluoride corrosion were considered more likely to occur relative to 
the other processes examined (CRWMS M&O 2000a). In the TSPA-SR, one important 
assumption was made to make the process feasible under the modeled conditions of the 
repository: the fluoride was assumed to remain at only one small segment of the cladding until 
the rod perforated, and not diffuse or migrate to other regions of the waste package (CRWMS 
M&O 2000~). 
TSPA Disposition Corrosion of zirconium has been observed in fluoride-containing 
environments. Since fluoride is present in Yucca Mountain groundwaters, localized corrosion 
from fluoride is considered as a perforation mechanism in the CSNF Cladding Degradation 
Component. Two scenarios for fluoride corrosion have been considered. In the first (water-filled 
WP scenario), the WP is full of water, and fluoride ions are transported to the cladding by 
aqueous diffusion. This scenario is less severe than the second one and is not addressed further. 
In the second (flow-through scenario), water enters the WP through a breach on the top and drips 
out through a breach on the bottom. The fluoride is assumed to be transported rapidly through 
the WP by advection, and fluoride attack is localized on a relatively small area of cladding 
(assumed to be 10 mm of rod length). It is also assumed that all the fluoride that enters the WP is 
consumed on the surface of the first rod it encounters until that rod fails. After one rod fails, the 
fluoride attacks the next rod. After WP breach, the fraction of cladding that fails is proportional 
to the volume of water that enters the package, reaching one (100%) when 2,424 m3 of water 
enters the WP. The uncertainty range placed on this estimate is plus or minus a factor of ten. In 
this approach, the fraction of cladding perforated is linearly dependent on the water inflow 
(percent failing equals 0.0413xm3 water in WP). The water flow into the WP increases with time 
as additional patches on the WP open. Cladding perforation rate also depends on the location of 
ANL-WIS-MD-000008 REV 00 55 April 2000 

the WP group in the repository because of different drip rates in different repository regions 
(CRWMS M&O 2000~). 
IRSR Issues: CLST1, CLST3, TSPAI4 
References: CRWMS M&O 2000a, CRWMS M&O 2000c 
Basis for Screening Decision 
No further elaboration in addition to the Screening Argument is necessary. 
6.2.16 Diffusion-Controlled Cavity Growth (DCCG) - YMP No. 2.1.02.28.00 
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. 
Screening Decision: Exclude, low probability (general creep rupture process, however, 
is included-see FEP 2.1.02.19.00). 
Screening Argument: Applicants for dry storage licenses for CSNF were once required 
by the U.S. Nuclear Regulatory Commission (NRC) to assume diffusion controlled cavity 
growth (DCCG) to evaluate dry storage designs. That is, the NUREG-1536 (NRC 1997, Section 
4.V.4.b) endorsement of the DCCG method which was used to calculate a maximum cladding 
temperature limit for a dry storage design. However, this design limit is overly restrictive and 
relatively inflexible. Recent literature does not support the use of this model for zirconiumbased 
materials (Pescatore and Cowgill, 1994, p. 83-85) since it has not been validated, and 
voids and cavities are rarely seen in irradiated Zircaloy. Pescatore and Cowgill (1994, p. 85) 
recommend a methodology similar to the approach used of calculating the amount of creep and 
comparing it to a creep failure criteria. The current NRC Interim Staff Guidance (ISG) Number 
11 (NRC 2000) recognizes the controversy with the DCCG conceptual model and permits 
license applicants to use other creep models in their license application. The use of creep model 
presented in CRWMS M&O (2000c, Section 6.2) is consistent with this ISG. The German dry 
storage creep analysis approach (Peehs 1998) is also consistent with the methodology used in 
CRWMS M&O (2000c, Section 6.2) and cited in Section 6.2.9 of this AMR. 
DCCG is excluded as a creep rupture mechanism. Creep rupture, as a cladding perforation 
process, however, is included in the CSNF Cladding Degradation Component of the Waste Form 
Degradation Model as described in FEP 2.1.02.19.00. The CSNF Cladding Degradation 
Component uses a creep strain method based on phenomena expected to be encountered under 
dry storage conditions as explained in the AMR Clad Degradation-Summary and Abstraction 
(CRMWS M&O 2000c) 
TSPA Disposition Exclude. 
IRSR Issues: CLST1, CLST3, TSPAI4 
ANI-WIS-MD-000008 REV 00 

References: Pescatore, C. and Cowgill, M. 1994, NRC 1997, NRC 2000, 
CRWMS M&O 2000c 
Basis for Screening Decision 
No further elaboration in addition to the Screening Argument is necessary 
6.3 NRC ISSUE RESOLUTION 
Both the expectations of and the strategic planning activities by the NRC emphasize the early 
identification and resolution of licensing issues, prior to the NRC's receipt of the license 
application to constmct a geologic repository. The NRC's objective is to reduce the number of 
issues and to better define in advance the issues that may be in dispute between licensee and 
regulator. NRC regulations and a 1993 agreement between NRC and DOE expand on this 
initiative by allowing NRC staff-level issue resolution to be achieved during the pre-licensing 
consultation period. Such resolution, however, would not preclude the issue's being raised and 
reconsidered during licensing proceedings. 
6.3.1 Staff-Level Issue Resolution 
To structure staff-level interactions, NRC has focused on the topics the NRC considers most 
critical to post-closure performance of the proposed geologic repository. At present, NRC staff 
has developed 10 Key Technical Issues, nine of which relate to post-closure performance 
assessment. The nine KTIs are identified in Table 5. 
Table 5. Key Technical Issues Related to Post-Closure Performance Assessment 
Each KT1 is configured with sub-issues, sub-issue components, and the NRC's acceptance 
criteria, thus facilitating the NRC's issuance of Issue Resolution Status Reports. As an important 
part of the staff-level interaction process, the IRSR provides the primary mechanism that NRC 
staff uses to provide feedback to DOE. 
Number 
1 
2 
3 
4 
Full resolution at the staff level concerning a particular issue is achieved during pre-licensing 
whenever the NRC staff have no further questions or comments regarding how the DOE program 
is addressing the issue, and the NRC has documented this situation. Furthermore, there may be 
some cases in which staff-level resolution during pre-licensing may be limited to documenting a 
common understanding regarding differences in NRC and DOE technical positions. 
Issue 
Total System Performance Assessment and Integration 
Container Lifetime and Source Term 
Evolution of the Near Field 
Radionuclide Transport 
Unsaturated and Saturated Flow Under Isothermal Conditions 
ANI-WIS-MD-000008 REV 00 57 Apr112000 
o 1 Structural Defornrawrl aru ~ a ~ s r r l ~ t i ~ r y 
9 I Igneous Activity 
1 

6.3.2 Cross-Reference Between FEPS Related to Cladding Degradation and NRC Key 
Technical Issues 
The following Tables 6 and 7 contain cross-references between NRC and Container Lifetime and 
Source Terms (CLST) and Total System Performance Assessment and Integration (TSPAI) Key 
Technical Issues and the cladding degradation FEPS. 
Table 6. Cross-Reference Between FEPS Related to Cladding Degradation and Container Life and Source 
Term Key Technical Issues 
rao on-c oes n ShF lconsoereo ne, processes for r i I ui I I uu lvvarenoggeo roos 
are reeaseo from me IShF oegraoal on an0 lne r e e a s l l 
1 
CLST 
Sub-Issue 
EBS through the 
oxidation and 
dissolution of sDent fl 
Acceptance Criteria FEPs 
3. The rate at which 14. DOE has identifled and I 
reep rupture of cladding 
-essurization from He Droduction causes claddina 
of radionuclidesfrom the EBS, as 
follws: dissolution of the I. lradlated U02 matrix. with the 
consequent formation of 
secondary minerals and colloids: 
prompt release Of radionuclides: 
degradation in the dry air 
environment: degradation and 
failure of fuel cladding: 
preferential dissolution of 
intermetallics in DOE SNF: and 
release of radionuclides from the 
WP emplacement drifts. 
iffusion-Controlled Cavity Growth 
ockfall (Large Block) 
2.1.02.13.00 
2.1.02.14.00 
2.1.02.15.00 
2.1 O2I6OO 
2.1 O2I7OO 
2.1.02.18.00 
2.1.02.19.00 
21022000 
ANI-WIS-MD-000008 REV 00 58 Apr112000 
" " 
General corrosion of cladding 
Microbiologically Influenced Corrosion (MIC) of cladding 
Acid corosion of cladding from radiolysis 
Localized corrosion (pitting) of cladding 
Localized corrosion (crevice corosion) of cladding 
High dissolved silica content of waters enhances 
corosion of cladding - 
CI - 
PI 

Table 7. Cross-Reference Between FEPS Related to Cladding Degradation and Total System Performance 
Assessment and Integration Issues 
coiosion of cladding 
2.1.02.19.00 Creep rupture of cladding 
2.1 0220.00 Pressurization from He production causes cladding 
failure - 
2.1.02.21.00 Si - 
2.1.02.22.00 H' - 
2.1.02.23.00 Cl - 
2.1 0224.00 Mecnanlcal mure or claaalnq 
tress cotros on cracn ng SCC of c aoo ng 
)onoe emonn emenl of caoo ng 
2 1 0- 01 00 I~UL1\1611 (LdlYC DlULR, 
2 1 09 03 00 1 Jo -me ncrease of Corros on P r ~ a ~ c r s 
2 1 11 --00 l ~ n e r m a .- no-cea Srress Cnanaes n Nasre ana 
7. CONCLUSIONS 
The following conclusions may be drawn: 
b This AMR contributes to the Yucca Mountain scenario development methodology by 
screening the FEPs related to cladding degradation. 
b This AMR develops screening arguments and TSPA-SR disposition statements for these 
FEPs. This AMR provides information for the YMP FEP Database and guidance to TSPASR 
analyses, which are appropriate for both site recommendation and license application 
documents. 
0 Screening decisions reached in this AMR are documented in Table 2. 
0 By providing references to appropriate other AMRs, this AMR provides a valuable link 
between issues related to cladding degradation and the research directed at their resolution. 
0 Finally, this AMR correlates FEPs related to cladding degradation with the NRC Container 
Life and Source Term and Total System Performance Assessment and Integration Key 
Technical Issues. 
This AMR uses no models and performs no analysis but only summarizes the models or analysis 
performed in the references. The references address cladding degradation before receiving it at 
YMP and after emplacement. Inherent is the assumption that the cladding is not damaged at the 
YMP surface facilities. This requires that the YMP will have both procedures and design 
ANI-WIS-MD-000008 REV 00 59 Apr112000 

features for the surface facilities to ensure that cladding degradation does not occur at the surface 
facilities. This AMR was developed using the methodology recommended in ASTM C 1174. 
This document may be affected by technical product input information that requires 
confirmation. Any changes to the document that may occur as a result of completing the 
confirmation activities will be reflected in subsequent revisions. The status of the input 
information quality may be confirmed by review of the Document Input Reference System 
database. 
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8.2 Codes, Standards, Regulations and Procedures 
64 FR 46976. 40 CFR 197: Environmental Radiation Protection Standards for Yucca Mountain, 
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64 FR 8640. Disposal ofHigh-Level Radioactive Wastes in a Proposed Geologic Repository at 
YuccaMountain, Nevada. Proposed rule 10 CFR 63. Readily Available 
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