ML19204A285

Dry Storage and Transportation of High Burnup Spent Nuclear Fuel NUREG 2224 Report for ACRS Review Predecisional

ML19204A285
Person / Time
Issue date:	08/16/2019
From:	Wendy Reed Division of Spent Fuel Management
To:	Advisory Committee on Reactor Safeguards
Reed W
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ML19204A282	List: ML19204A283 ML19204A285
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NUREG-2224
Download: ML19204A285 (131)

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NUREG-2224 Dry Storage and Transportation of High Burnup Spent Nuclear Fuel Final Report Office of Nuclear Materials Safety and Safeguards

NUREG-2224 Dry Storage and Transportation of High Burnup Spent Nuclear Fuel Final Report Date Published: TBD Office of Nuclear Materials Safety and Safeguards

For ACRS Review Purposes Only 1 ABSTRACT 2 The purpose of this report is to expand the technical basis in support of the U.S. Nuclear 3 Regulatory Commissions (NRCs) guidance on adequate fuel conditions as it pertains to hydride 4 reorientation in high burnup (HBU) spent nuclear fuel (SNF) cladding. This guidance defines 5 adequate fuel conditions, including peak cladding temperatures during short-term loading 6 operations to prevent or mitigate degradation of the cladding. Time-dependent changes on the 7 cladding properties of HBU SNF are primarily driven by the fuels temperature, rod internal 8 pressure (and corresponding pressure-induced cladding hoop stresses), and the environment 9 during dry storage or transport operations. Historically, the potential for these changes to 10 compromise the analyzed fuel configuration in dry storage systems and transportation packages 11 has been addressed through safety review guidance.

12 Hydride reorientation is a process in which the orientation of hydrides precipitated in HBU SNF 13 cladding during reactor operation changes from the circumferential-axial to the radial-axial 14 direction. Research results over the last decade have shown that hydride reorientation can still 15 occur at temperatures and stresses lower than those assumed in the current staff review 16 guidance. Therefore, the NRC has since sponsored additional research to better understand 17 whether hydride reorientation could affect the mechanical behavior of HBU SNF cladding and 18 compromise the fuel configuration analyzed in dry storage systems and transportation 19 packages.

20 This report provides an engineering assessment of the results of research on the mechanical 21 performance of HBU SNF following hydride reorientation. Based on the conclusions of that 22 assessment, the report then presents example approaches for licensing and certification of HBU 23 SNF for dry storage (under Title 10 of the Code of Federal Regulations (10 CFR) Part 72, 24 Licensing Requirements for the Independent Storage of Spent Nuclear Fuel and High-Level 25 Radioactive Waste, and Reactor-Related Greater Than Class C Waste) and transportation 26 (under 10 CFR Part 71, Packaging and Transportation of Radioactive Material).

27 The NRC expects these example licensing and certification approaches, when followed by 28 applicants, to minimize or eliminate the need for requests for additional information during the 29 staffs safety review of applications for dry storage and transportation of HBU SNF. Further, the 30 NRC expects that future revisions of the Standard Review Plans for dry storage systems and 31 transportation packages will reference the licensing and certification approaches delineated in 32 NUREG-2224.

33 The information in this report is not intended for use in applications for wet storage facilities or 34 monitored retrievable storage installations licensed under 10 CFR Part 72.

35 Nothing contained in this report is to be construed as having the force or effect of regulations.

36 Comments regarding errors or omissions, as well as suggestions for improvement of this 37 NUREG should be sent to the Director, Division of Spent Fuel Management, U.S. Nuclear 38 Regulatory Commission, Washington, D.C., 20555-0001.

39 Paperwork Reduction Act 40 41 This NUREG provides guidance for implementing the mandatory information collections in 10 42 CFR Parts 71 and 72 that are subject to the Paperwork Reduction Act of 1995 (44 U.S.C. 3501 43 et. seq.). These information collections were approved by the Office of Management and Budget 44 (OMB) under control numbers 3150-0008 and 3150-0132. Send comments regarding this iii

For ACRS Review Purposes Only 1 information collection to the Information Services Branch, U.S. Nuclear Regulatory Commission, 2 Washington, DC 20555-0001, or by e-mail to Infocollects.Resource@nrc.gov, and to the Desk 3 Officer, Office of Information and Regulatory Affairs, NEOB-10202, (3150-0008, 3150 -0132) 4 Office of Management and Budget, Washington, DC 20503.

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For ACRS Review Purposes Only 1 Public Protection Notification 2 The NRC may not conduct or sponsor, and a person is not required to respond to, a collection 3 of information unless the document requesting or requiring the collection displays a currently 4 valid OMB control number.

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For ACRS Review Purposes Only 1 CONTENTS 2 ABSTRACT ................................................................................................................... III 3 CONTENTS ................................................................................................................... VI 4 LIST OF FIGURES......................................................................................................... IX 5 LIST OF TABLES .......................................................................................................... XI 6 ACKNOWLEDGMENTS .............................................................................................. XIII 7 ABBREVIATIONS AND ACRONYMS ......................................................................... XV 8 GLOSSARY ................................................................................................................ XIX 9 1 INTRODUCTION ............................................................................................... 1-1 10 1.1 Background ........................................................................................................ 1-1 11 1.2 Fuel Cladding Performance and Staffs Review Guidance................................. 1-2 12 1.3 Cladding Creep .................................................................................................. 1-4 13 1.4 Effects of Hydrogen on Cladding Mechanical Performance ............................... 1-5 14 1.5 Hydride Reorientation ......................................................................................... 1-7 15 1.5.1 Hydride Dissolution and Precipitation ..................................................... 1-8 16 1.5.2 Fuel Cladding Fabrication Process ....................................................... 1-10 17 1.5.3 End-Of-Life Rod Internal Pressures and Cladding Hoop 18 Stresses ................................................................................................ 1-11 19 1.5.4 Ring Compression Testing ................................................................... 1-15 20 1.5.5 Staffs Assessment of Ring Compression Testing Results ................... 1-22 21 2 ASSESSMENT OF STATIC BENDING AND FATIGUE STRENGTH 22 RESULTS ON HIGH BURNUP SPENT NUCLEAR FUEL ............................... 2-1 23 2.1 Introduction ......................................................................................................... 2-1 24 2.2 Cyclic Integrated Reversible Fatigue Tester....................................................... 2-1 25 2.3 Application of the Static Test Results ................................................................. 2-6 26 2.3.1 Spent Fuel Rod Behavior in Bending...................................................... 2-7 27 2.3.2 Composite Behavior of a Spent Fuel Rod .............................................. 2-7 28 2.3.3 Calculation of Cladding Strain from CIRFT Static Bending Data .......... 2-10 29 2.3.4 Calculation of Cladding Strain Using Factored Cladding-Only 30 Properties ............................................................................................. 2-13 31 2.3.4.1 Two Alternatives for Calculating Cladding Stress and 32 Strain During Drop Accidents ................................................. 2-16 33 2.3.5 Applicability to Dry Storage and Transportation ................................... 2-17 vi

For ACRS Review Purposes Only 1 2.3.5.1 Use of Static Test Results to Evaluate Safety Margins 2 in an HAC Side Drop Event .................................................... 2-20 3 2.3.5.2 Dynamic Response of a Fuel Rod .......................................... 2-22 4 2.3.5.3 Seismic Response of a Fuel Rod ........................................... 2-23 5 2.3.5.4 Thermal Cycling during Loading Operations .......................... 2-23 6 2.4 Application of Fatigue Test Results .................................................................. 2-23 7 2.4.1 Lower Bound Fatigue S-N Curves ........................................................ 2-23 8 2.4.2 Fatigue Cumulative Damage Model ..................................................... 2-27 9 2.4.3 Applicability to Storage and Transportation .......................................... 2-27 10 2.4.3.1 Seismic Events ....................................................................... 2-28 11 2.4.3.2 Thermal Cycling during Loading Operations .......................... 2-28 12 3 DRY STORAGE OF HIGH BURNUP SPENT NUCLEAR FUEL ...................... 3-1 13 3.1 Introduction ......................................................................................................... 3-1 14 3.2 Uncanned Fuel (Intact and Undamaged Fuel) ................................................... 3-4 15 3.2.1 Leaktight Confinement ............................................................................ 3-6 16 3.2.2 Non-Leaktight Confinement .................................................................... 3-7 17 3.2.3 Dry Storage Up To 20 Years ................................................................ 3-10 18 3.2.4 Dry Storage Beyond 20 Years .............................................................. 3-11 19 3.2.4.1 Supplemental Results from Confirmatory 20 Demonstration ........................................................................ 3-11 21 3.2.4.1.1 Initial Licensing or Certification .......................... 3-12 22 3.2.4.1.2 Renewal Applications ........................................ 3-12 23 3.2.4.2 Supplemental Safety Analyses ............................................... 3-12 24 3.2.4.2.1 Materials and Structural .................................... 3-13 25 3.2.4.2.2 Confinement ...................................................... 3-13 26 3.2.4.2.3 Thermal ............................................................. 3-13 27 3.2.4.2.4 Criticality ............................................................ 3-15 28 3.2.4.2.5 Shielding............................................................ 3-16 29 3.3 Canned Fuel (Damaged Fuel) .......................................................................... 3-19 30 4 TRANSPORTATION OF HIGH BURNUP SPENT NUCLEAR FUEL ............... 4-1 31 4.1 Introduction ......................................................................................................... 4-1 32 4.2 Uncanned Fuel (Intact and Undamaged Fuel) ................................................... 4-4 33 4.2.1 Leaktight Containment ............................................................................ 4-7 34 4.2.2 Non-Leaktight Containment .................................................................... 4-7 vii

For ACRS Review Purposes Only 1 4.2.3 Direct Shipment from the Spent Fuel Pool and Shipment of 2 Previously Dry-Stored Fuel (Up To 20 Years Since Fuel Was 3 Initially Loaded)..................................................................................... 4-11 4 4.2.4 Shipment of Previously Dry-Stored Fuel (Beyond 20 Years 5 Since Fuel Was Initially Loaded) .......................................................... 4-12 6 4.2.4.1 Supplemental Data from Confirmatory Demonstration ........... 4-12 7 4.2.4.2 Supplemental Safety Analyses ............................................... 4-12 8 4.2.4.2.1 Materials and Structural .................................... 4-13 9 4.2.4.2.2 Containment ...................................................... 4-13 10 4.2.4.2.3 Thermal ............................................................. 4-14 11 4.2.4.2.4 Criticality ............................................................ 4-15 12 4.2.4.2.5 Shielding............................................................ 4-18 13 4.3 Canned Fuel ..................................................................................................... 4-20 14 5 CONCLUSIONS ................................................................................................ 5-1 15 6 REFERENCES .................................................................................................. 6-1 16 viii

For ACRS Review Purposes Only 1 LIST OF FIGURES 2 Figure 1-1 Average Hydride Content [H] and Distribution in HBU SNF Cladding (from Billone 3 et al., 2013). ....................................................................................................... 1-6 4 Figure 1-2 Dissolution (Cd) and Precipitation (Cp) Concentration Curves ........................... 1-9 5 Figure 1-3 Publicly-Available Data Collected by EPRI for PWR End-Of-Life Rod Internal 6 Pressures at 25°C (77 °F) (Reproduction of Figure 2-1 from Machiels (2013)) .. 1-7 12 8 Figure 1-4 Fuel Cladding Tube with Stress Element Displaying Hoop Stress (),

9 Longitudinal Stress (z), Internal Pressure (Pi), Cladding Thickness (hm),

10 External Pressure (Po), Circumferential Coordinate (), and Inner Cladding 11 Diameter (Dmi) .................................................................................................. 1-14 12 Figure 1-5 RCT of a Sectioned Cladding Ring Specimen in ANLs Instrons 8511 Test 13 Setup. ............................................................................................................. 1-16 14 Figure 1-6 Effective Ductility vs. RCT for Two PWR Cladding Alloys Following Slow Cooling 15 from 400°C (752 °F) at Peak Target Hoop Stresses of 110 Mpa (1.6 x 104 psia) 16 and 140 Mpa (2.0 x 104 psia) (From Billone et al., 2013) ................................. 1-18 17 Figure 1-7 Ductility Data, as Measured by RCT, for As-Irradiated Zircaloy-4 and Zircaloy-4 18 Following Cooling from 400 °C (752 °F) Under Decreasing Internal Pressure and 19 Hoop Stress Conditions (From Billone et al., 2013) ......................................... 1-19 20 Figure 1-8 Ductility Data, as Measured by RCT, for as-Irradiated ZIRLO and ZIRLO 21 Following Cooling from 400 °C (752 °F) Under Decreasing Internal Pressure and 22 Hoop Stress Conditions (From Billone et al., 2013) ......................................... 1-20 23 Figure 1-9 Ductility Data, as Measured by RCT, for As-Irradiated M5 and M5 Following 24 Cooling from 400 °C (752 °F) under Decreasing Internal Pressure and Hoop 25 Stress Conditions (From Billone et al., 2013) ................................................... 1-21 26 Figure 1-10 Geometric Models for Spent Fuel Assemblies in Transportation Packages 27 (Reproduction, in Part, Of Figure 10 from Sanders et al., 1992) ...................... 1-23 28 Figure 2-1 Horizontal Layout of ORNL U-Frame Setup (Top), Rod Specimen and Three 29 Lvdts for Curvature Measurement (Middle), and Front View of CIRFT Installed in 30 ORNL Hot Cell (Bottom) (Figure 4 from NUREG/CR-7198, Revision 1 (NRC, 31 2017a)) ............................................................................................................... 2-3 32 Figure 2-2 Schematic Diagram of End and Side Drop Accident Scenarios (Revised 33 Figure 5-168 from Patterson and Garzarolli (2015))........................................... 2-7 34 Figure 2-3 Typical Composite Construction of a Bridge ...................................................... 2-9 35 Figure 2-4 Influence of cg Position on Composite Beam Stiffness .. 2-10 36 Figure 2-5 Images of Cladding-Pellet Structure in HBU SNF Rod .................................... 2-11 37 Figure 2-6 Approximate Extreme Fiber Tensile Stresses Between Pellet-Pellet Crack..... 2-12 38 Figure 2-7 Comparison of CIRFT Static Bending Results with Calculated PNNL Moment 39 Curvature (Flexural Rigidity) Derived from Cladding-Only Stress-Strain Curve ......

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For ACRS Review Purposes Only 1 Figure 2-8 Characteristic Points on Moment-Curvature Curve. A, B, C, and D are Points on 2 the Curve. ....................................................................................................... 2-14 3 Figure 2-9 High Magnification Micrograph Showing Radial Hydrides of a HBR HBU SNF 4 Hydride-Reoriented Specimen Tested Under Phase II ................................... 2-18 5 Figure 2-10 Representative Conditions Used for Radial Hydride Treatment for Preparation of 6 HBR HBU SNF Hydride-Reoriented Specimens Tested Under Phase II ......... 2-19 7 Figure 2-11 Plots of Half of the Cladding Strain Range (/2) and the Maximum Strain (//max) 8 as a Function of Number of Cycles to Failure .................................................. 2-25 9 Figure 2-12 CIRFT Dymanic (Fatigue) Test Results for As-Irradiated and Hydride- Reoriented 10 H.B. Robinson Zircaloy-4 HBU Fuel Rods. The Calculated Lower-Bound 11 Fatigue Endurance Curve is also Shown ......................................................... 2-26 12 Figure 3-1 Example Licensing and Certification Approaches for Dry Storage of High Burnup 13 Spent Nuclear Fuel ............................................................................................. 3-3 14 Figure 3-2 First Approach for Evaluating Design-Bases Drop Accidents During Dry Storage..

15 ...........................................................................................................................3-5 16 Figure 3-3 Second Approach for Evaluation of Design-Bases Drop Accidents During Dry 17 Storage ............................................................................................................... 3-6 18 Figure 4-1 Example Approaches for Approval of Transportation Packages with High Burnup 19 Spent Nuclear Fuel ............................................................................................. 4-3 20 Figure 4-2 First Approach for Evaluation of Drop Accidents During Transport .................... 4-5 21 Figure 4-3 Second Approach for Evaluation of Drop Accidents During Transport............... 4-6 22 Figure 4-4 Evaluation of Vibration Normally Incident to Transport ...................................... 4-7 23 x

For ACRS Review Purposes Only 1 LIST OF TABLES 2 Table 1-1 End of Life Rod Internal Pressures (MPa) at a Peak Temperature of 400 °C (752 3 °F) ..................................................................................................................... 1-13 4 Table 1-2 Maximum Cladding Hoop Stresses (MPa) at a Peak Temperature of 400 °C (752 5 °F) .................................................................................................................... 1-14 6 Table 1-3 End of Life Rod Internal Pressures at Room Temperature (25 °C (77 °F)) and 7 Atmospheric Conditions (1.0 x 10-1 MPa (1.5 x 101 psia)) (From FRAPCON Code 8 Predictions in Richmond and Geelhood, 2018) ................................................ 1-15 9 Table 2-1 Specifications of Rod Specimens used in NRC-Sponsored HBU SNF Test 10 Program .............................................................................................................. 2-4 11 Table 2-2 Comparison of Average Flexural Rigidity Results Between CIRFT Static Testing 12 and PNNL Cladding-Only Data ( ...................................................................... 2-15 13 Table 2-3 Characteristic Points and Quantities Based on Moment-Curvature Curves ... 2-15 14 Table 2-4 PWR 15 x 15 SNF Assembly Parameters ........................................................ 2-21 15 Table 2-5 Summary of CIRFT Dynamic Test Results for As-Irradiated and Hydride-16 Reoriented HBR HBU SNF............................................................................... 2-24 17 Table 2-6 Coordinates for Lower-Bound Enveloping S-N Curve for the HBR HBU SNF 18 Rods ................................................................................................................. 2-25 19 Table 3-1 Fractions of Radioactive Materials Available for Release from HBU SNF Under 20 Conditions of Dry Storage .................................................................................. 3-8 21 Table 4-1 Fractions Of Radioactive Materials Available for Release from HBU SNF Under 22 Conditions of Transport ...................................................................................... 4-9 23 xi

For ACRS Review Purposes Only 1 ACKNOWLEDGMENTS 2 The working group is very grateful to M. Billone (Argonne National Laboratory) for providing 3 valuable input for the writing of the report, to Olivier Lareynie (French Nuclear Safety Authority, 4 ASN) for assisting in the preparation of responses to comments on this report, and to J. Wang 5 (Oak Ridge National Laboratory) for providing valuable insights, observations, and 6 recommendations.

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For ACRS Review Purposes Only 1 ABBREVIATIONS AND ACRONYMS ADAMS Agencywide Documents Access and Management System AMP aging management program ANL Argonne National Laboratory ANS American Nuclear Society ANSI American National Standards Institute b width BWR boiling-water reactor Cd concentration at dissolution Cp concentration at precipitation CFR Code of Federal Regulations cg center of gravity CoC Certificate of Compliance CIRFT cyclic integrated reversible-bending fatigue tester CRUD Chalk River unknown deposit CWSRA cold worked stress relieved annealed p/Dmo offset strain Tdp temperature hysteresis (dissolution-precipitation)

Dmi inner (metal) cladding diameter Dmo outer (metal) cladding diameter DLF dynamic load factor DTT ductility transition temperature DOE U.S. Department of Energy DSS dry storage system average tensile strain

-N strain per number of cycles E elastic modulus Ec elastic modulus of the cladding Ep elastic modulus of the fuel pellet EOL end-of-life EPRI Electric Power Research Institute GBC general burnup credit GTCC greater-than-Class-C waste h height hm cladding (metal) thickness HAC hypothethical accident conditions (transportation)

HBR H. B. Robinson HBU high burnup HRT hydride reorientation treatment Hz hertz I moment of inertia Ic moment of inertia of the cladding xv

For ACRS Review Purposes Only Ip moment of inertia of the fuel pellet IAEA International Atomic Energy Agency IFBA integral fuel burnable absorber ISFSI independent spent fuel storage installation ISG Interim Staff Guidance curvature

-N curvature per number of cycles keff k-effective l rod length between spacers LBU low burnup LVDT linear variable differential transformer M bending moment ni number of strain cycles at strain level i Ni number of strain cycles to produce failure at i NCT normal conditions of transport NRC U.S. Nuclear Regulatory Commission ORNL Oak Ridge National Laboratory Pi rod internal pressure Po rod external pressure PNNL Pacific Northwest National Laboratory PWR pressurized-water reactor r outer radius RCT ring compression testing RHCF radial hydride continuity factor RIP rod internal pressure RXA recrystallized annealed average tensile stress cladding hoop stress z cladding longitudinal stress SNF spent nuclear fuel SRP standard review plan SSC structure, system, and component Td dissolution temperature Tp precipitation temperature w uniform applied load ymax distance to the neutral axis xvi

For ACRS Review Purposes Only Units of Measure C Celsius F Fahrenheit ft foot g 9.806 m/s2 GWd/MTU gigawatt-days per metric ton of uranium h hour in. inch lb pound m meter m micrometer, 1 x 10-6 meter mm millimeter, 0.001 meter MPa megapascal, 1 x 106 pascals N newton N*m newton meter Pa pascal psia pounds per square inch (absolute) s second Torr Torr (unit of pressure) wppm parts per million by weight 1

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For ACRS Review Purposes Only 1 GLOSSARY Accident condition of The extreme level of an event or condition, which has a specified storage resistance, limit of response, and requirement for a given level of continuing capability, which exceeds off-normal events or conditions.

Accident conditions include both design-basis accidents and conditions caused by natural and manmade phenomena.

Aging Management See Title 10 of the Code of Federal Regulations (10 CFR) 72.3, Program Definitions.

Amendment of a An application for amendment of a license or a CoC must be submitted license or certificate of whenever a holder of a specific license or CoC desires to change the compliance (CoC) license or CoC (including a change to the technical specifications that accompany the license or CoC conditions). The application must fully describe the desired change(s) and the reason(s) for such change(s),

and following as far as applicable the form prescribed for original applications. See 10 CFR 72.56, Application for Amendment of License, and 10 CFR 72.244, Application for Amendment of a Certificate of Compliance.

Assembly defect Any change in the physical as-built condition of the spent fuel assembly except for normal in-reactor changes such as elongation from irradiation growth or assembly bow. Examples of assembly defects include: (1) missing rods; (2) broken or missing grids or grid straps (spacers); and (3) missing or broken grid springs.

Breached spent A spent nuclear fuel (SNF) rod with cladding defects that permit the nuclear fuel rod release of gases or solid fuel particulates from the interior of the fuel rod. SNF rod breaches include pinhole leaks, hairline cracks, and gross ruptures.

Burnup The measure of thermal power produced in a specific amount of nuclear fuel through fission, usually expressed in gigawatt-day per metric ton uranium (GWd/MTU). For the purpose of assessing the allowable contents, the maximum burnup of the fuel is generally specified in terms of the average burnup of the entire fuel assembly (i.e., assembly average). For the purpose of assessing fuel cladding integrity in the materials and structural review, the rod with the highest burnup within the fuel assembly is generally specified in terms of peak rod average burnup.

Can for damaged fuel A metal enclosure that is sized to confine damaged SNF contents. A can for damaged fuel must satisfy fuel-specific and dry storage system/package-related functions for undamaged SNF, as required by the applicable regulations.

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For ACRS Review Purposes Only Canister (in a dry A metal cylinder that is sealed at both ends and may be used to storage system) perform the function of confinement. Typically, a separate overpack performs the radiological shielding and physical protection function.

Certificate of The certificate issued by the U.S. Nuclear Regulatory Commission compliance (CoC) (for (NRC) that approves the design of a spent fuel storage cask in a dry storage system) accordance with the provisions of 10 CFR Part 72, Licensing Requirements for the Independent Storage of Spent Nuclear Fuel, High-Level Radioactive Waste, and Reactor-Related Greater Than Class C Waste, Subpart L, Approval of Spent Fuel Storage Casks.

See 10 CFR 72.3.

Certificate of The certificate issued by the NRC that approves the design of a Compliance (CoC) (for package for the transportation of radioactive material in accordance a transportation with the provisions of 10 CFR Part 71, Packaging and Transportation package) of Radioactive Material, Subpart D, Application for Package Approval.

See 10 CFR 71.4, Definitions.

Certificate holder (for a A person who has been issued a CoC by the NRC for a spent fuel dry storage system) storage cask design under 10 CFR Part 72. See 10 CFR 72.3.

Certificate holder (for a A person who has been issued a CoC or other package approval by the transportation package) NRC under 10 CFR Part 71. See 10 CFR 71.4.

Certificate of The general licensee that has loaded a dry storage system, or compliance user (CoC purchased a dry storage system (DSS) and plans to load it, in user) accordance with a CoC issued under 10 CFR Part 72.

Confinement (in a dry The ability to limit or prevent the release of radioactive substances into storage system for the environment.

spent nuclear fuel)

Confinement systems Those systems, including ventilation, that act as barriers between areas containing radioactive substances and the environment. See 10 CFR 72.3.

Containment system The assembly of components of the packaging intended to retain the radioactive material during transport. See 10 CFR 71.4.

Controlled area See 10 CFR 72.3 and 10 CFR 20.1003, Definitions. The definition in 10 CFR 20.1003 is broader in scope and allows for, or includes, establishment of access controls to areas within the site for any reason (for radiation protection).

Criticality The condition wherein a system or medium is capable of sustaining a nuclear chain reaction.

Damaged spent Any spent fuel rod or spent fuel assembly that cannot meet the nuclear fuel pertinent fuel-specific or system-related regulations for the xx

For ACRS Review Purposes Only transportation package (10 CFR Part 71) or dry storage system (10 CFR Part 72).

Degradation Any change in the properties of a material that adversely affects the performance of that material; adverse alteration. See NUREG-1536, Revision 1, Standard Review Plan for Spent Fuel Dry Storage Systems at a General License Facility., issued July 2010.

Design bases (storage) Information that identifies the specific function(s) to be performed by structures, systems, and components (SSCs) (both important-to-safety and not important-to-safety) of a facility or of a spent fuel storage cask and the specific values or ranges of values chosen for controlling parameters as reference bounds for design. These values may be (1) restraints, derived from generally accepted state-of-the-art practices for achieving functional goals, or (2) requirements, derived from analysis (based on calculation, experiments, or both) of the effects of a postulated event under which SSCs must meet their functional goals. See 10 CFR 72.3.

Dry storage The storage of SNF in a DSS, which typically involves drying the DSS cavity and backfilling with an inert gas.

Dry storage system A system that typically uses a cask or canister in an overpack as a (DSS) component in which to store SNF in a dry environment. A DSS provides confinement, radiological shielding, sub-criticality control, structural support, and passive cooling of its SNF during normal, off-normal, and accident conditions. A DSS design may be approved under a CoC, as listed in 10 CFR 72.214, List of Approved Spent Fuel Storage Casks, or licensed under a specific license for an independent spent fuel storage installation.

g-load The acceleration experienced by an object with mass under its own self weight.

General license Authorizes the storage of spent fuel in an ISFSI at a power reactor site (storage) to persons (see definition of person in 10 CFR 72.3) authorized to possess or operate nuclear power reactors under 10 CFR Part 50 (Domestic Licensing of Production and Utilization Facilities) or 10 CFR Part 52 (Licenses, Certifications, and Approvals for Nuclear Power Plants). The general license is limited to (1) that spent fuel which the general licensee is authorized to possess at the site under the specific 10 CFR Part 50 or 10 CFR Part 52 license for the site, and (2) storage of spent fuel in casks approved under the provisions of 10 CFR Part 72, Subpart L and listed in 10 CFR 72.214. See 10 CFR 72.210 (General License Issued) and 72.212(a)(1)-(2).

Gross breach A breach in the spent fuel cladding that is larger than either a pinhole leak or a hairline crack and allows the release of particulate matter from the spent fuel rod.

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For ACRS Review Purposes Only Hairline crack A minor SNF cladding defect that will not permit significant release of particulate matter from the spent fuel rod and therefore presents a minimal as low-as-is-reasonably-achievable concern during fuel handling operations.

High burnup (HBU) SNF with assembly average burnup (see Burnup) generally exceeding spent nuclear fuel 45 GWd/MTU.

Hoop stress The tensile stress in cladding wall in the circumferential orientation of the fuel rod.

Important to safety See SSCs important to safety.

(storage)

Independent spent fuel A complex designed and constructed for the interim storage of spent storage installation nuclear fuel, solid reactor-related greater-than-Class-C (GTCC) waste, (ISFSI) and other radioactive materials associated with spent fuel and reactor-related GTCC waste storage. See 10 CFR 72.3.

Intact spent nuclear A subset of undamaged SNF. Any fuel rod or fuel assembly that can fuel meet the pertinent fuel-specific or system-related regulations for the transportation package (10 CFR Part 71) or dry storage system (10 CFR Part 72). Intact SNF rods may not contain pinholes, hairline cracks, or gross ruptures. Intact SNF assemblies may have assembly defects if able to meet the pertinent fuel-specific or system-related regulations.

Intended function A design-basis function defined as either (1) important to safety or (storage) (2) the failure of which could impact a safety function.

Interim staff guidance Supplemental information that clarifies important aspects of regulatory (ISG) requirements. An ISG provides review guidance to NRC staff in a timely manner until standard review plans are revised accordingly.

keff k-effective Effective neutron multiplication factor including all biases and uncertainties at a 95-percent confidence level for indicating the level of subcriticality relative to the critical state. At the critical state, keff = 1.0.

This has also been used to represent effective thermal conductivity.

The degree of package containment that, in a practical sense, Leaktight precludes any significant release of radioactive materials. This degree of containment is achieved by demonstration of a leakage rate less than or equal to 1 x 10-7 ref*cm3/s, of air at an upstream pressure of 1 atmosphere (atm) absolute (abs), and a downstream pressure of 0.01 atm abs or less.

Low-burnup (LBU) Spent nuclear fuel with an assembly average burnup (see Burnup) spent nuclear fuel generally less than 45 GWd/MTU.

M5 (M5) AREVA-trademarked fuel cladding alloy, which contains zirconium and niobium xxii

For ACRS Review Purposes Only Non-fuel hardware Hardware that is not an integral part of a fuel assembly. This is the term used to identify what the regulation refers to as other radioactive materials associated with fuel assemblies (see SNF definition in 10 CFR 72.3). While not integral to the assembly, it includes those items that are designed to operate and are positioned or operated within the envelope of the fuel assembly during reactor operation and are stored within the assembly envelope in the storage container. Typical examples of non-fuel hardware include: burnable poison rod assemblies (BPRAs), control element assemblies, thimble plug assemblies, and boiling-water reactor (BWR) fuel channels. Examples of items that do not meet this definition include boron sources, BWR in-core instruments, and BWR control blades.

Non-mechanistic event An event, such as cask tip-over, which should be evaluated for (dry storage) acceptable system capability, although a cause for such an event is not identified in the analyses of off-normal and accident events and conditions.

Normal events or Conditions that are intended operations, planned events, and conditions of storage environmental conditions that are known or reasonably expected to occur with high frequency during storage operations. Normal refers to the maximum level of an event or condition that is expected to routinely occur (similar to Design Event I as defined in American National Standards Institute/American Nuclear Society (ANSI/ANS) 57.9, Design Criteria for an Independent Spent Fuel Storage Installation (Dry Storage Type)). The DSS or ISFSI SSCs are expected to remain fully functional and to experience no temporary or permanent degradation of that functionality from normal operations, events, and conditions.

Specific normal conditions to be addressed are evaluated for the DSS or ISFSI and are documented in a safety analysis report for that system or facility.

Normal means The ability to move a fuel assembly with a crane and grapple used to (dry storage) move undamaged assemblies at the point of cask loading. The addition of special tooling or modifications to the assembly to make the assembly suitable for lifting by crane and grapple does not preclude the assembly from being considered movable by normal means.

Off-normal events or An event or condition that, although not occurring regularly, can be conditions of storage expected to occur with moderate frequency and for which there is a corresponding maximum specified resistance, limit of response, or requirement for a given level of continuing capability. Off-normal events and conditions are similar to a Design Event II in ANSI/ANS 57.9. A DSS or ISFSI SSC is expected to experience off-normal events and conditions without permanent degradation of capability to perform its full function (although operations may be suspended or curtailed during off-normal conditions) over the full storage term (the license period for a specific license facility or the storage period equivalent to the certificate term for a DSS). Off-normal events or conditions are referred to as anticipated occurrences in 10 CFR 72.104, Criteria for xxiii

For ACRS Review Purposes Only Radioactive Materials in Effluents and Direct Radiation from an ISFSI or MRS.

Package The packaging together with its radioactive contents as presented for (transportation) transport. See 10 CFR 71.4.

Packaging The assembly of components necessary to ensure compliance with the (transportation) packaging requirements of 10 CFR Part 71. It may consist of one or more receptacles, absorbent materials, spacing structures, thermal insulation, radiation shielding, and devices for cooling or absorbing mechanical shocks. The vehicle, tie-down system, and auxiliary equipment may be designated as part of the packaging. See 10 CFR 71.4.

Pinhole leak A minor cladding defect that will not permit significant release of particulate matter from the spent fuel rod and therefore present a minimal as low-as-is-reasonably-achievable concern during fuel handling operations.

Renewal of a license A certificate holder may apply for renewal of the design of a spent fuel or CoC (dry storage) storage cask for a term not to exceed 40 years. In the event that the certificate holder does not apply for a cask design renewal, any licensee using a spent fuel storage cask, a representative of the licensee, or another certificate holder may apply for a renewal of that cask design for a term not to exceed 40 years. See 10 CFR 72.240, Conditions for Spent Fuel Storage Cask Renewal. Specific licenses may be renewed by the Commission at the expiration of the license term upon application by the licensee for a period not to exceed 40 years. See 10 CFR 72.42, Duration of License; Renewal. The current regulatory framework for storage of spent fuel allows for multiple license or CoC renewals, subject to an aging management analysis and planning.

Ready retrieval The ability to safely remove the spent fuel from storage for further (dry storage) processing or disposal.

Recovery The capability of returning the stored radioactive materials from an (dry storage) accident to a safe condition without endangering public health and safety or causing significant or unnecessary exposure to workers. Any potential release of radioactive materials during recovery operations must not result in doses or radiation exposures that exceed the limits in 10 CFR Part 20, Standards for Protection against Radiation. Doses during recovery operations are included in the dose estimates for accidents, the total of which must not exceed the limits in 10 CFR 72.106, Controlled Area of an ISFSI or MRS.

Retrievability (dry See definition of ready retrieval above. Storage systems must be storage) designed to allow ready retrieval of SNF, high-level radioactive waste, and reactor-related GTCC waste for further processing or disposal.

See 10 CFR 72.122(l).

xxiv

For ACRS Review Purposes Only Safety analysis report The report submitted to the NRC staff by an applicant for a CoC for a (SAR) (dry storage) DSS design, or for a specific license for an ISFSI, to present information on the design and operations of the system or facility. This document provides the justification and analyses to demonstrate that the design meets regulatory requirements and acceptance criteria (10 CFR 72.24, Contents of Application: Technical Information, and 10 CFR 72.230(a)). The SAR is submitted for approval of the ISFSI or DSS design. The final SAR is as defined in 10 CFR 72.48(a)(5).

Safety function (dry The functions that DSS and DSF SSCs important to safety (see 10 CFR storage) 72.3) are designed to maintain/perform, including the following:

• protection against environmental conditions

• content temperature control

• radiation shielding

• confinement

• sub-criticality control, and

• retrievability.

Specific license (dry A license issued by the NRC to authorize the receipt, handling, storage, storage) and transfer of spent fuel, high-level radioactive waste, or reactor-related GTCC waste at an ISFSI or MRS facility. The NRC issues the license to a named person (see definition of person in 10 CFR 72.3) after the NRC has reviewed an application filed under regulations in 10 CFR Part 72, Subpart B, License Application, Form, and Contents (see 10 CFR 72.6 License Required; Types of Licenses.)

Spent nuclear fuel Nuclear fuel that has been withdrawn from a nuclear reactor after (SNF) or spent fuel irradiation, has undergone at least a 1-year decay process since being used as a source of energy in a power reactor, and has not been chemically separated into its constituent elements by reprocessing.

Spent fuel includes the special nuclear material, byproduct material, source material, and other radioactive materials associated with fuel assemblies. See 10 CFR 71.4 and 10 CFR 72.3.

For purposes of this report, spent nuclear fuel refers to high burnup SNF unless otherwise noted.

Structures, systems, See 10 CFR 72.3. Those features of the ISFSI and spent fuel storage and components cask whose functions are at least one of the following:

(SSCs) important to

• to maintain the conditions required to safely store spent fuel, safety (storage) high-level radioactive waste, or reactor-related GTCC waste

• to prevent damage to the spent fuel, the high-level radioactive waste, or reactor-related GTCC waste container during handling and storage

• to provide reasonable assurance that spent fuel, high-level radioactive waste, or reactor-related GTCC waste can be xxv

For ACRS Review Purposes Only received, handled, packaged, stored, and retrieved without undue risk to the health and safety of the public.

Undamaged spent Any fuel rod or fuel assembly that can meet the pertinent fuel-specific or nuclear fuel system-related regulations for the transportation package (10 CFR Part 71) or dry storage system (10 CFR Part 72). Undamaged (non-intact) SNF rods may contain pinholes or hairline cracks, but may not contain gross ruptures. Undamaged SNF assemblies may have assembly defects if they are still able to meet the pertinent fuel-specific or system-related regulations.

Zircaloy An alloy of zirconium, tin, and other metals, used chiefly as cladding for nuclear reactor fuel.

ZIRLO' (ZIRLO) Westinghouse-trademarked fuel cladding alloy, which contains zirconium, tin and niobium.

xxvi

For ACRS Review Purposes Only 1 1 INTRODUCTION 2 1.1 Background 3 As required by Title 10 of the Code of Federal Regulations (10 CFR) 72.44(c), a specific license 4 for dry storage of spent nuclear fuel (SNF) is to include technical specifications that, among 5 other things, define limits on the fuel and allowable geometric arrangements. Further, as 6 required by 10 CFR 72.236(a), a Certificate of Compliance (CoC) for a dry storage system 7 (DSS) design must include specifications for the type of spent fuel (i.e., boiling water reactor 8 (BWR), pressurized water reactor (PWR), or both), maximum allowable enrichment of the fuel 9 prior to any irradiation, burn-up (i.e., megawatt-days/MTU), minimum acceptable cooling time of 10 the spent fuel before storage in the spent fuel storage cask, maximum heat designed to be 11 dissipated, maximum spent fuel loading limit, condition of the spent fuel (i.e., intact assembly or 12 consolidated fuel rods), and inerting atmosphere requirements. These specifications ensure 13 that the loaded SNF assemblies remain within the bounds of the safety analyses in the 14 approved design basis.

15 The regulations in 10 CFR Part 72, Licensing Requirements for the Independent Storage of 16 Spent Nuclear Fuel, High-Level Radioactive Waste, and Reactor-Related Greater Than Class C 17 Waste, include a number of fuel-specific and DSS-specific requirements that may be 18 dependent on the design-basis condition of the fuel cladding. As required by 19 10 CFR 72.122(h)(1), the SNF cladding is to be protected against degradation that leads to 20 gross ruptures, or the fuel must be otherwise confined such that degradation of the fuel during 21 storage will not pose operational safety problems with respect to its removal from storage. In 22 addition, 10 CFR 72.122(l) states that the DSS must be designed to allow ready retrieval of the 23 SNF. According to Interim Staff Guidance1 (ISG)-2, Revision 2, Fuel Retrievability in Spent 24 Fuel Storage Applications, issued in April 2016 (NRC, 2016a), this may be demonstrated by 25 either (A) removing individual or canned SNF assemblies from wet or dry storage; (B) removing 26 a canister loaded with SNF assemblies from a DSS cask or overpack; or (C) removing a DSS 27 cask loaded with SNF assemblies from its storage location. The ready retrieval requirement is 28 defined by the approved design basis for the DSSs Certificate of Compliance or the 29 Independent Spent Fuel Storage Installations specific license. Therefore, the integrity of the 30 cladding is an important consideration for demonstrating ready retrieval under option A. The 31 condition of the fuel cladding may also impact the safety analyses used to demonstrate 32 compliance with DSS-specific requirements in 10 CFR 72.124(a), 10 CFR 72.128, and 10 CFR 33 72.236(m).

34 Similarly for transportation, the regulations in 10 CFR Part 71, Packaging and Transportation of 35 Radioactive Material, also include a number of fuel-specific and package-specific requirements.

36 The regulations in 10 CFR 71.31, Contents of application and 10 CFR 71.33, Package 37 description, requires an application for a transportation package to describe the proposed 38 package in sufficient detail to identify the package accurately and provide a sufficient basis for 39 evaluation of the package, which includes a description of the chemical and physical form of the 40 allowable contents. The regulations in 10 CFR Part 71 also require that (1) the geometric form 41 of the package contents not be substantially altered under the tests for normal conditions of 42 transport (NCT) (10 CFR 71.55(d)(2)) and (2) a package used for the shipment of fissile material 1 The current revisions of all ISG documents will be rolled into revised standard review plans (SRPs) for dry storage and transportation of SNF, as appropriate, and will then be removed from the public domain. The revised SRPs will be issued for public comment prior to being finalized.

1-1

For ACRS Review Purposes Only 1 is to be designed and constructed and its contents so limited that under the tests for 2 hypothetical accident conditions (HAC) specified in 10 CFR 71.73, Hypothetical accident 3 conditions, the package remains subcritical (10 CFR 71.55(e)). The requirement assumes that 4 the fissile material is in the most reactive credible configuration consistent with the damaged 5 condition of the package and the chemical and physical form of the contents 6 (10 CFR 71.55(e)(1)).

7 To comply with the requirements mentioned above, the fuel cladding generally serves a design 8 function in both DSSs and transportation packages for ensuring that the configuration of 9 undamaged and intact fuel remains within the bounds of the reviewed safety analyses.2 10 Therefore, an application should address potential degradation mechanisms that could result in 11 gross cladding ruptures during operations. To assist the safety review of potential degradation 12 mechanisms, the U.S. Nuclear Regulatory Commission (NRC) staff (the staff) has historically 13 issued guidance on acceptable storage and transport conditions that limit SNF degradation 14 during operations and ensure that the reviewed safety analyses remain valid.

15 1.2 Fuel Cladding Performance and Staffs Review Guidance 16 Time-dependent (i.e., age-related, not event-related) mechanisms resulting in changes to the 17 fuel cladding performance are all primarily driven by the fuels temperature, rod internal 18 pressure (and corresponding pressure-induced cladding hoop stresses), and the environment 19 during dry storage or transport operations. Contrary to the hoop stresses experienced by the 20 fuel cladding during reactor operation, which are generally compressive because of the high 21 reactor coolant pressure, the hoop stresses during drying-transfer, dry storage, and transport 22 operations are tensile because of the low pressure external to the cladding. For instance, the 23 pressure of the environment surrounding the fuel in the reactor can be 1.6 x 107 Pa 24 (2.3 x 103 psia) while the environment surrounding the fuel in the DSS confinement cavity may 25 be as low as 4.0 x 102 Pa (5.8 x 102 psia) at the end of vacuum drying and 5 x 105 Pa 26 (7.3 x 101 psia) during dry storage. The magnitude of the cladding hoop stresses will depend 27 on the differential pressure across the cladding wall and thus the rod internal pressure at a 28 given time. Various factors determine the rod internal pressure, including the fuels fabrication 29 and irradiation conditions (i.e., fabrication rod gas fill pressure, rod void (plenum) volume, 30 cladding thickness, presence of burnable absorbers, burnup) and the average gas temperature 31 within the fuel rods. The average gas temperature within the fuel rods has a first-order effect on 32 the hoop stress in the cladding and thus cladding performance. Therefore, an important 33 consideration for demonstrating adequate cladding performance is to control the peak cladding 34 temperature of the fuel rods during vacuum drying and storage/transport operations to 35 temperatures demonstrated to preserve cladding integrity.

36 To assist in the safety review of DSS and transportation packages, the staff has developed 37 guidance with a supporting technical basis for setting adequate fuel conditions, including 38 acceptable peak cladding temperatures during short-term loading operations so that the 39 cladding meets the pertinent regulations. Historically, guidance has been issued as ISG-11, 40 Cladding Considerations for the Transportation and Storage of Spent Fuel, which has been 41 revised multiple times to incorporate new data and lessons learned from the staffs review 42 experience. Initial standard review plans (SRPs) prior to ISG-11 stated that DSSs and 43 transportation packages needed to be dried to a level where galvanic corrosion could be ruled 2 If the fuel is classified as damaged, a separate canister (e.g., a can for damaged fuel) that confines the assembly contents to a known volume may be used to provide this assurance.

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For ACRS Review Purposes Only 1 out as a fuel degradation mechanism. The guidance specified moisture levels only for low 2 burnup (LBU) fuel (i.e., burnup below 45 gigawatt-day per metric ton uranium (GWd/MTU))

3 because of the lack of degradation data at higher burnup values. In 1999, the staff first issued 4 ISG-11 to supplement the SRPs by addressing potential degradation of high burnup (HBU) fuel 5 (i.e., burnup exceeding 45 GWd/MTU).

6 In 2000, the staff issued ISG-11, Revision 1 to incorporate new data, but also to give the 7 applicant the responsibility for demonstrating that the cladding was adequately protected. ISG-8 11, Revision 1 stated that cladding oxidation should not be credited as load-bearing in the fuel 9 cladding structural evaluation and also defined a 1-percent creep strain limit on the cladding. It 10 also discussed the use of damaged fuel cans for confining fuel with gross ruptures. ISG-11, 11 Revision 1, accounted for Zircaloy-clad fuel rods and not for advanced cladding alloys (e.g.,

12 ZIRLO' (ZIRLO) and M5 (M5)).

13 In 2002, the staff issued ISG-11, Revision 2, to change the definition of damaged fuel, remove 14 the 1-percent creep strain limit, and discuss criteria to limit hydride reorientation in the cladding.

15 It also made the guidance applicable to all zirconium-based claddings and all burnup levels.

16 The revision described onerous calculations, dependent on the characteristics of the fuel to be 17 stored, to determine the maximum cladding temperature for the design-basis fuel per a justified 18 creep strain limit. Gruss et al. (2004) discuss in more detail the data used for supporting ISG-19 11, Revision 2. Historically, ISG-11 has not discussed the use of an inert atmosphere to 20 mitigate fuel degradation. Research has shown that the uranium dioxide (UO2) in the fuel pellet 21 may oxidize (U4O9) at temperatures less than 230 °C (446 °F) (McEachern and Taylor, 1998; 22 Jung et. al, 2013). Therefore, ISG-22, Potential Rod Splitting Due to Exposure to an Oxidizing 23 Atmosphere during Short-Term Cask Loading Operations in LWR or Other Uranium Oxide 24 Based Fuel, issued May 2006 (NRC, 2006), addressed the use of an inert atmosphere for 25 loading operations.

26 In November 2003, the staff issued ISG-11, Revision 3, Cladding Considerations for the 27 Transportation and Storage of Spent Fuel (NRC, 2003a). The guidance was eventually 28 incorporated into NUREG-1536, Revision 1, Standard Review Plan for Spent Fuel Dry Storage 29 Systems at a General License Facility, issued in July 2010 (NRC, 2010), although not yet 30 incorporated into a revision of NUREG-1567, Standard Review Plan for Spent Fuel Storage 31 Facilities, issued in March 2000 (NRC, 2000a) (i.e., the standard review plan for specific 32 licenses under 10 CFR Part 72). ISG-11, Revision 3 replaced the calculation of the maximum 33 cladding temperature per a justified creep strain limit with a generic 400 °C (752 °F) peak 34 cladding temperature limit applicable to normal conditions of storage and transportation, as well 35 as short-term loading operations (e.g., drying, backfilling with inert gas, and transfer of the DSS 36 cask or canister to the storage pad). ISG-11, Revision 3 also defined a higher short-term 37 temperature limit applicable to LBU fuel if the applicant demonstrated by calculation that the 38 cladding hoop stress would not exceed 90 MPa (1.3 x 104 psi) for the proposed temperature 39 limit. The guidance also defined a generic maximum cladding temperature limit of 570 °C 40 (1,058 °F) for off-normal and accident conditions applicable to all burnups. This is discussed 41 further in Section 1.3.

42 In addition to creep, ISG-11, Revision 3 (NRC, 2003a), also considered minimizing hydride 43 reorientation. At the time of its issuance, the technical basis discussed in ISG-11, Revision 3 44 supported the staffs conclusion that hydride reorientation would be minimized by maintaining 45 cladding temperatures below 400 °C (752 °F) and restricting the change in cladding 46 temperatures during drying-transfer operations to less than 65 °C (149 °F). This temperature 47 change limit was based on the temperature drop required to obtain the degree of 1-3

For ACRS Review Purposes Only 1 supersaturation required for the precipitation of radial hydrides in a short thermal cycle 2 (see Section 1.5.1). Therefore, ISG-11, Revision 3, states that the cladding should not 3 experience more than 10 thermal cycles, each not exceeding 65 °C (149 °F), which provided 4 assurance that hydride reorientation would be limited.

5 Research results obtained since the ISG-11, Revision 3, have shown that hydride reorientation 6 can still occur below the generic 400 °C (752 °F) peak cladding temperature limit (Aomi et al, 7 2008; Billone et al., 2013; Billone et al., 2014; Billone et al., 2015). To better understand 8 hydride reorientation, both the NRC and the U.S. Department of Energy (DOE) have obtained 9 additional data on the performance of HBU SNF cladding with reoriented hydrides to determine 10 if the guidance in ISG-11, Revision 3, ought to be revised. This is discussed further in Section 11 1.5 12 1.3 Cladding Creep 13 Creep is the time-dependent deformation of a material under stress. The main driving force for 14 cladding creep at a given temperature is the hoop stress caused by internal rod pressure. The 15 internal rod pressure results from the initial fill gas pressure condition and, to a smaller extent, 16 from fission and decay gases released to the gap between the fuel and cladding during dry 17 storage operations (Ito, at al., 2004). Fuel pellet swelling may also result in localized stresses 18 on the cladding due to the mechanical interaction between the cladding and the fuel. Pellet 19 swelling may occur due to: (1) the incorporation of soluble and insoluble solid fission products in 20 the fuel matrix, (2) the formation of intra- and intergranular fission gas bubbles, particularly in 21 the hot interior region of a fuel pellet, and (3) the formation of a large number of small gas 22 bubbles in the fine-grained ceramic structure that builds inward from the outer pellet surface for 23 HBU fuel. If excessive creep of the cladding were to occur during dry storage, it could lead to 24 thinning, hairline cracks, or gross ruptures (Hanson et al, 2012) and potentially compromise the 25 ability to safely retrieve by normal means the HBU fuel on a single-assembly basis (if required 26 by the design basis).

27 The appendix to ISG-11, Revision 3 (NRC, 2003a) reviewed the data used by the staff to obtain 28 reasonable assurance that creep will not result in gross ruptures for peak cladding temperatures 29 below 400 °C (752 °F). The design and materials used for fabrication of fuel rods are such that 30 the creep of the cladding is self-limited. As the average gas temperature of the fuel rod 31 increases during drying-transfer and storage/transport operations, the gas pressure within the 32 fuel column increases (with a corresponding increase in cladding hoop stresses). If the 33 increase of gas pressure is sufficient to result in cladding creep, the internal volume of the rod 34 will increase, which will, in turn ,reduce the gas pressure within the fuel column (with a 35 corresponding decrease in cladding hoop stresses). The net effect is a slow decrease in 36 pressure and hoop stress with increasing creep strain. The stress also decreases with 37 increasing storage or transport time due to the decrease in rod internal pressure with 38 decreasing temperature. ISG-11, Revision 3, concluded the following:

39 1. deformation caused by creep will proceed slowly over time and will decrease the rod 40 pressure, 41 2. the decreasing cladding temperature also decreases the hoop stress, and this too will 42 slow the creep rate so that during later stages of dry storage, further creep deformation 43 will become exceedingly small, and 44 3. in the unlikely event that a breach of the cladding due to creep occurs, it is believed that 45 this will not result in gross rupture.

1-4

For ACRS Review Purposes Only 1 These conclusions are considered applicable to fuel at all burnups because the relatively small 2 differences in creep rate as a function of materials and burnup are not expected to have a 3 significant impact on the maximum creep strains in the rod. The technical basis in ISG-11, 4 Revision 3 (NRC, 2003a) has provided reasonable assurance to the staff that creep strains 5 during dry storage and transportation will not result in fuel failures nor compromise the assumed 6 fuel configuration in the safety analyses. However, the staff recognizes the uncertainties 7 associated with extrapolating short-term accelerated test data to extended periods of dry 8 storage. The staff further recognizes the separate-effects nature of the accelerated creep 9 testing conducted to date, which would not account for potential combined effects with other 10 phenomena occurring during dry storage (e.g., annealing of irradiation hardening, hydride 11 reorientation). Therefore, the staff considers it prudent that long-term observation of HBU SNF 12 stored in a deployed DSS be used to confirm the conclusions of the accelerated short-term 13 testing. To aid users in demonstrating adequate creep performance during storage periods 14 beyond 20 years, in June 2016, the staff issued guidance in NUREG-1927, Revision 1, 15 Standard Review Plan for Renewal of Specific Licenses and Certificates of Compliance for Dry 16 Storage of Spent Nuclear Fuel (NRC, 2016b), which discusses the use of an Aging 17 Management Program using a surrogate surveillance and monitoring program to provide this 18 confirmatory long term data.

19 1.4 Effects of Hydrogen on Cladding Mechanical Performance 20 During irradiation, hydrogen is generated due to water-coolant corrosion (i.e., oxidation) of the 21 cladding, which diffuses into the zirconium-based material. As the solubility limit of hydrogen in 22 the cladding is exceeded, circumferential hydrides precipitate (Figure 1-1). The preferential 23 circumferential precipitation of the hydrides during reactor operation results from the texture of 24 cladding, which is determined by the manufacturing process. The number density of these 25 circumferential hydrides varies across the cladding wall due to the temperature drop from the 26 fuel side (hotter) to the coolant side (cooler) of the cladding. When the cladding absorbs 27 significant hydrogen, precipitation of dissolved hydrogen into the coolant side of the cladding 28 can result in the formation of a rather dense hydride rim just below the outer coolant-side 29 cladding oxide layer, with a higher concentration of hydrides occurring in the outer 1/3 of the 30 cladding. The hydride number density and thickness of this hydride rim depend on cladding 31 design and reactor operating conditions for a given fuel type. For example, fuel rods operated 32 at high linear heat ratings (heat fluxes) to high burnup generally have a very dense hydride rim 33 that is less than 10 percent of the cladding wall thickness. Conversely, fuel rods operated at low 34 linear heat ratings (heat fluxes) to high burnup have a more diffuse hydride distribution that 35 could extend as far as 50 percent across the cladding wall (Adamson, et al., 2007). Therefore, 36 the distribution of hydrides varies across the thickness of the cladding, as shown in Figure 1-1, 37 and is a consideration in the mechanical performance of the fuel cladding.

1-5

For ACRS Review Purposes Only 1 Figure 1-1 Average Hydride Content [H] and Distribution in HBU SNF Cladding (from 2 Billone et al., 2013).

3 The staff concluded in ISG-11, Revision 3 (NRC, 2003a), that the hydride rim, along with any 4 cladding metal oxidized during reactor operation, should not be considered as load bearing 5 when determining the effective cladding thickness for the structural evaluation of the assembly 6 in the DSS or transportation package. However, the staff recognizes that there is no reliable 7 predictive tool available to calculate this rim thickness, which varies along the fuel-rod length, 8 around the circumference at any particular axial location, from fuel rod to fuel rod within an 9 assembly, and from assembly to assembly. Moreover, recent data generated by Argonne 10 National Laboratory (ANL) have shown that, for the full range of gas pressures anticipated 11 during drying and storage, the hydride rim remains intact following cooling under conditions of 12 decreasing pressure (Billone et al., 2013; Billone et al., 2014; Billone et al., 2015). The results 13 suggest that hydride rims have some load bearing capacity and, therefore, it may be appropriate 14 to include the hydride rim in the effective cladding thickness calculation. Therefore, the staff 15 considers as acceptable the inclusion of the hydride rim thickness in the calculation of the 16 effective cladding thickness when mechanical test data referenced in the structural evaluation 17 have adequately accounted for its presence. Historically, this has been the case during the 18 review of DSS and transportation packages, as applicants have provided mechanical property 19 data generated from tests with irradiated cladding samples with an intact hydride rim. These 20 data include test results derived from uniaxial tensile tests or pressurized tube tests of samples 21 that do not have a machined gauge section.

1-6

For ACRS Review Purposes Only 1 Applicants have generally relied on a public database of materials properties for Zircaloy-4, 2 Zircaloy-2 and ZIRLO to analyze the behavior of as-irradiated cladding (Geelhood et al, 2008; 3 Geelhood et al, 2014) during dry storage and transportation. Additional data for engineering 4 properties (e.g., yield stress, ultimate tensile stress, and uniform elongation) can be found in the 5 open literature for ZIRLO (Cazalis et al., 2005; Pan et al., 2013), Optimized ZIRLO (Pan et al.,

6 2013), and M5 (Cazalis et al., 2005; Fourgeaud et al., 2009; Bouffioux et al., 2013). These 7 references are provided for informational purposes. The applicant for a DSS or transportation 8 package should adequately justify the use of any of these properties and the associated 9 experimental methods for the relevant fuel designs cited. Any use of mechanical properties 10 from uniaxial-tension and ring-expansion tests on cladding specimens with machined gauge 11 sections, where some of the hydride rim would have been inadvertently removed during outer-12 surface oxide removal, should be adequately justified. The mechanical property data from 13 these specimens are still valuable, but characterization of their remaining rim thickness, post-14 test determination of their hydrogen concentration, or both may be needed.

15 1.5 Hydride Reorientation 16 As discussed in Section 1.4, hydrogen is picked up by the cladding during reactor operation.

17 The excess hydrogen (i.e., hydrogen exceeding the solubility limit in the cladding) precipitates 18 primarily in the circumferential-axial direction. However, under temperature and stress 19 conditions experienced during vacuum drying and storage/transport operations, some of these 20 hydrides may redissolve and subsequently reprecipitate as new hydrides. During this process, 21 the orientation of these precipitated hydrides may change from the circumferential-axial to the 22 radial-axial direction.

23 The technical basis discussed in ISG-11, Revision 3 (NRC, 2003a) has supported the staffs 24 conclusion that if peak cladding temperatures are maintained below 400 °C (752 °F) or the 25 pressure-induced hoop stresses in the cladding were maintained below 90 MPa (1.3 x 104 psia),

26 then hydride reorientation would be minimized. The database used for this determination (see 27 Figure 3 in Chung, 2004) had a mixture of results from irradiated and non-irradiated material, 28 high and low hydrogen concentrations, different cladding types, different cooling rates, and 29 other variables. In addition, the methods to determine if there were radial hydrides varied 30 considerably from researcher to researcher. Since the issuance of ISG-11, Revision 3, 31 research results generated at ANL (Billone et al., 2013; Billone et al., 2014; Billone et al., 2015) 32 and in Japan (Aomi et al., 2008) have shown that hydride reorientation can still occur at lower 33 temperatures and stresses than those assumed in ISG-11, Revision 3. Because of the number 34 of variables involved, the staff agreed that it would not be practical to precisely determine the 35 temperature and stress conditions to prevent reorientation. Rather, the critical question was 36 what effect hydride reorientation would have on the mechanical behavior of the cladding, 37 particularly since the design-basis structural evaluation of the SNF assembly generally assumes 38 as-irradiated cladding mechanical properties (i.e., properties not accounting for hydride 39 reorientation). If hydride reorientation had an observable effect on the mechanical behavior of 40 the cladding (i.e., it decreased the failure strain limit of the cladding in response to stresses 41 during operations), then the failure limits as defined in the design-basis structural evaluations 42 would have to be modified.

43 Because both circumferential and radial hydrides are oriented in the planes parallel to the 44 principal normal tensile stress during bending loading, the staff has expected that HBU SNF 45 fatigue strength and bending stiffness would not be sensitive to the hydride orientation under 1-7

For ACRS Review Purposes Only 1 bending moments that produce longitudinal tensile stresses in the rod (Tang et al, 2015).3 2 Experimental confirmation of this expectation was prudent. Therefore, the NRC and DOE 3 conducted complementary research programs to investigate the cyclic fatigue and bending 4 strength performance of HBU SNF cladding in both as-irradiated and reoriented conditions 5 (Wang et al., 2016; NRC, 2017).

6 Even with the expectation that hydride orientation would not have a significant impact on the 7 fatigue strength and bending stiffness of HBU SNF under bending moments that produce 8 longitudinal tensile stresses in the rod, the staff expressed concern that hydride orientation 9 could impact the failure stresses and strains under pinch-type loads. Pinch-type loads could 10 potentially occur during postulated drop accidents in storage, normal conditions of transport 11 (NCT), or hypothetical accident conditions (HAC) during transportation. The staff was 12 particularly concerned with reduced cladding ductility during the HAC 9-m (30-ft) side drop or a 13 tip-over handling accident, where pinch loads could occur due to rod-to-grid spacer contact, rod-14 to-rod contact, or rod-to-basket contact. If the fuel temperature were to be sufficiently low at the 15 time of the accident, these pinch loads could compromise the analyzed fuel configuration.

16 Thus, research was conducted in the United States and Japan to study the ductility of cladding 17 with reoriented hydrides under diametrically-opposed pinch loads. Ring compression testing 18 (RCT) was used to assess residual ductility of de-fueled HBU SNF cladding specimens 19 subjected to hydride reorientation (see Section 1.5.4). This testing led to the establishment of a 20 ductility transition temperature (DTT) (i.e., a temperature at which the tested cladding segments 21 were determined to lose ductility relative to as-irradiated cladding). The following section 22 discusses important parameters affecting the DTT and provides the staffs conclusion on its 23 relevance for future licensing and certification actions involving HBU SNF.

24 1.5.1 Hydride Dissolution and Precipitation 25 During drying-transfer operations, the cladding temperature increases, which causes some of 26 the circumferential hydrides to dissolve as hydrogen. The amount of hydrogen dissolved 27 depends on the temperature (Td) and increases according to the solubility curve (Cd) for 28 zirconium-based alloys (Kammenzind et al., 1996; Kearns, 1967; McMinn, et al., 2000).

29 Zirconium-based alloys are materials that can have hydrogen in a supersaturated solution 30 because of the extra energy (strain, thermal) required to precipitate zirconium hydrides in the 31 cladding matrix. This results in a hysteresis in the solubility-precipitation curves as shown in 32 Figure 1-2.

3 Hydrides are essentially two-dimensional features since their thickness is relatively small compared to the other two dimensions. Radial hydrides span in the longitudinal and radial directions, and circumferential hydrides span in the longitudinal and circumferential directions. The bending tensile stresses are in the longitudinal direction.

Therefore, the bending tensile stresses are parallel to the plane of both the radial and circumferential hydrides.

1-8

For ACRS Review Purposes Only 450 Kammenzind Cd (Zry-4) 400 Kammenzind Cp (Zry-4)

Hydrogen Content (wppm) 350 Kearns Cd (Zr,Zry-2,Zry-4) 300 250 Tdp 200 Precipitation 150 100 50 Dissolution 0

0 100 200 300 400 500 Temperature (°C) 1 Figure 1-2 Dissolution (Cd) and Precipitation (Cp) Concentration Curves Based on the 2 Data of Kammenzind et al. (1996) for Non-Irradiated Zircaloy-4 (Zry-4) 3 (Revised Figure 1 from Billone, et al., 2014). Also Shown Is the Best Fit to 4 the Dissolution Curve (Cd) for Zirconium (Zr), Zircaloy-2 (Zry-2), and 5 Zircaloy-4, Which Includes the Zircaloy-2 and Zircaloy-4 Data Generated by 6 Kearns (1967). Tdp = Td - Tp Refers to the Temperature Drop Required for 7 Precipitation, where Td and Tp are the Corresponding Temperatures in the 8 Solubility and Precipitation Curves for the Same Hydrogen Content 9 The solubility curves (Cd) plotted in Figure 1-2 indicate that the amount of hydrogen that 10 dissolves increases with increasing temperature, but it is relatively independent of alloy 11 composition and fabricated microstructure (recrystallized annealed (RXA) and cold worked 12 stress relieved annealed (CWSRA)) (Kearns, 1967). Both Kammenzind et al (1996) and Kearns 13 (1967) used diffusion couples, with one sample containing excess hydrogen and the other 14 sample containing essentially no hydrogen, exposed to long annealing times (e.g., 2 days at 15 525 °C (977 °F) and 10 days at 260 °C (500 °F)). As shown in Figure 1-2, Kearns dissolution 16 correlation for Zircaloy-2 and Zircaloy-4 is in excellent agreement with the correlation of 17 Kammenzind et al. (e.g., 207 wppm versus 210 wppm at 400 °C (752 °F), and 127 wppm versus 18 133 wppm at 350 °C (662 °F)) and is well within experimental error. In terms of precipitation, 19 the temperature drop (Tdp = Td - Tp, where Td and Tp are the corresponding temperatures in 20 the solubility and precipitation curves at the same hydrogen content) required for precipitation is 21 approximately 65 °C (149 °F). That is, for irradiated cladding that contains no radial hydrides 22 prior to heating, the 65 °C (149 °F) temperature decrease is necessary to initiate precipitation of 1-9

For ACRS Review Purposes Only 1 radial hydrides.4 However, if circumferential hydrides are present at the peak cladding 2 temperature, some hydrogen will precipitate by growth of the existing circumferential hydrides 3 during this 65 °C (149 °F) temperature drop because of the lower energy required to grow rather 4 than to initiate precipitation of new hydrides (Colas et al., 2014). The strain field remaining from 5 the regions of the hydrides that dissolved during heating also facilitates the growth of existing 6 hydrides.

7 McMinn et al. (2000) used a different method (differential scanning calorimetry) to generate an 8 independent data set for dissolution-precipitation curves per non-irradiated and lightly-irradiated 9 Zircaloy-2 and Zircaloy-4 samples with low hydrogen content ( 77 wppm with most data at 60 10 wppm) exposed to temperatures less than 320°C (608 °F). The data show the effects of 11 irradiation (increase in solubility), as well as pre-annealing time and temperature (decrease in 12 solubility). The increase in hydrogen solubility for irradiated materials is likely the result of 13 hydrogen trapped in irradiation-induced defects. However, it is not clear yet whether the 14 trapped hydrogen is available for precipitation unless the temperature is high enough to anneal 15 out some of these defects. Extrapolation of the dissolution correlation of McMinn et al. (2000) 16 for non-irradiated cladding alloys gives only 172 wppm of dissolved hydrogen at 400 °C (752 °F) 17 and 102 wppm at 350 °C (662 °F), while the data for irradiated cladding agree quite well with 18 the correlations of Kammenzind et al (1996) and Kearns (1967). The staff considers these two 19 sources to be reasonably representative of dry storage and transportation because the long 20 annealing times used to achieve equilibrium for dissolution are more applicable to drying-21 storage than the much shorter times used for measurements taken by differential scanning 22 calorimetry. Further, the staff considers these data to provide an upper bound for non-irradiated 23 cladding and close to a best fit for irradiated cladding.

24 The amount of hydrogen dissolved will depend on the peak cladding temperature during drying-25 transfer, dry storage, and transport operations. This temperature is typically achieved during 26 vacuum drying, which generally takes about 8 to 40 hours4.62963e-4 days <br />0.0111 hours <br />6.613757e-5 weeks <br />1.522e-5 months <br /> depending on the DSS or transport 27 package design and loading parameters. Figure 1-2, along with an assessment of the axial 28 hydrogen content of the fuel rods and the peak cladding temperature, can be used to estimate 29 the amount of dissolved hydrogen for a given allowable fuel in a DSS or transportation package.

30 The degree of reorientation will depend on the fuel cladding fabrication process, as well as the 31 cladding hoop stresses and temporal thermal profile of the fuel during operations. The following 32 discussions provide additional details on these parameters.

33 1.5.2 Fuel Cladding Fabrication Process 34 The cladding alloy and corresponding fabrication process are important factors for determining 35 the extent of hydride reorientation. Two predominant cladding microstructures are produced 36 during fabrication of zirconium-based cladding: CWSRA and RXA. Zircaloy-4 and ZIRLO are 37 generally CWSRA, whereas Zircaloy-2 and M5 are RXA. Because hydrides tend to precipitate 38 in the grain boundaries, RXA claddings are more susceptible to hydride reorientation, since 39 these cladding types have a larger fraction of grain boundaries in the radial direction (equiaxed 40 grains) relative to CWSRA claddings (which have more elongated grains). However, RXA 41 claddings also have lower hydrogen uptakes during reactor operation than CWSRA claddings 42 (Patterson and Garzarolli (2015)).

4 This hysteresis is the basis for the guidance in ISG-11, Revision 3 (NRC, 2003a), to limit repeated thermal cycling (repeated heatup/cooldown cycles) during loading operations to less than 10 cycles, with cladding temperature variations that are less than 65 °C (149 °F) each.

1-10

For ACRS Review Purposes Only 1 1.5.3 End-Of-Life Rod Internal Pressures and Cladding Hoop Stresses 2 Most rods are initially backfilled with a pressurized inert helium atmosphere to improve thermal 3 conductivity during irradiation and to decrease the rate of cladding creep-down onto the fuel.

4 During the course of irradiation, fission gases are generated in the fuel pellets. Some of the 5 fission gas will be released to the void volume within the fuel column and plenum. Observations 6 of commercial power fuel indicate that the fission gas released is about 1 to 3 percent for PWR 7 fuel rods irradiated under low-to-moderate conditions up to a burnup of about 45 GWd/MTU, at 8 which point the rate of release increases gradually to about 5 to 7 percent for a burnup of 65 9 GWd/MTU. For BWR fuel rods, the fission gas release can be in the range of 10 to 15 percent 10 at burnups exceeding 45 GWd/MTU. PWR fuel rods with internal burnable poisons (e.g., boron-11 10 in zirconium-diboride coating on fuel pellets) can also release decay gases (e.g., helium) 12 within the fuel rod. The pressure of these gases in PWR fuel rods increases with burnup due to 13 the increase in fission gas generation, the decay gases generated from the burnable poisons, 14 and the decrease in void volume resulting from cladding creep down and fuel swelling.

15 The internal pressure of the rod exerts hoop and axial stresses in the cladding, which increase 16 with burnup because of the increase in internal pressure and the decrease in cladding thickness 17 due to waterside corrosion (i.e., oxidation). For BWR fuels, increased cladding oxidation and 18 hydrogen pickup are observed at burnups exceeding 50 GWd/MTU.5 In PWRs, hydrogen 19 pickup is usually correlated to the oxide thickness, which varies depending on the alloy. The 20 condition of the fuel as it is removed from the reactor is described more fully in the International 21 Atomic Energy Agency (IAEA) Nuclear Energy Series NF-T-3.8 (IAEA, 2011).

22 Post-irradiation examination of cladding specimens subjected to representative drying-transfer 23 and dry storage operations has shown that the degree of radial hydride precipitation is very 24 sensitive to the peak cladding hoop stresses. The range of relevant cladding hoop stresses 25 depends on the range of end-of-life (EOL) rod internal pressures (RIPs), the range of average 26 gas temperatures within fuel rods during drying-transfer and storage/transport operations, and 27 fuel design and operational parameter used to assess the pressure difference across the 28 cladding. Therefore, an understanding of EOL RIPs is important for assessing the extent of 29 hydride reorientation in each fuel design.

30 The publicly-available database for EOL RIPs for PWR fuel rods is sparse relative to the 31 number of rods that have been irradiated. In addition, the RIP data in this database are for 32 standard fuel rods, mostly those clad in zirconium-tin alloy (Zircaloy-4) with older (1975-1985) 33 fuel designs and reactor operating conditions.6 Thus, the database is heavily populated with 34 data from what are generally called legacy fuel rods. Figure 1-3 shows the publicly-available 35 empirical data for standard fuel rods, as collected by the Electric Power Research Institute 36 (EPRI) (Machiels, 2013). The EOL RIP data in Figure 1-3 are evaluated at 25 °C (77 °F), and 37 are identified by the reactor, the assembly design, and the as-fabricated helium fill pressure at 5 Zirconium liners in Zircaloy-2 cladding used in BWR fuel are located at the claddings inner surface and occupy about 10% of the wall thickness. The liners are metallurgically bonded to the Zircaloy-2 tube and consist of zirconium alloyed with varying amounts of iron (Fe). The addition of Fe improves corrosion resistance during reactor operations. In Zircaloy-2 cladding with a zirconium liner, hydrogen is observed to diffuse preferentially to the liner as cooling rates decrease. Such preferential diffusion results from the lower solubility of hydrogen in pure zirconium relative to the solubility in Zircaloy-2.

6 Publicly-available empirical EOL RIP data are available for ZIRLO-clad SNF rods but not for M5-clad SNF rods.

1-11

For ACRS Review Purposes Only 1 25°C (77 °F). Data points labeled as ENUSA are for fuel rods irradiated in the Vandellos Unit 2 2 reactor in Spain.

3 The public database consists of 92 data points:

4

• 27 at 45 GWd/MTU (24 Zircaloy-4 and 3 ZIRLO) 5

• 35 in the range of >45 GWd/MTU to 60 GWd/MTU (25 Zircaloy-4 and 10 ZIRLO) 6

• 30 in the range of >60 GWd/MTU to 74 GWd/MTU (15 each of Zircaloy-4 and ZIRLO) 7 Helium fill pressures at fabrication range from 2.00 MPa (290 psia) 3.45 MPa (500 psia).

8 However, some of the older legacy fuel designs have initial helium fill pressures as high as 2.52 9 MPa (365 psia). As shown in Figure 1-3, the EOL RIP data appear to be relatively flat between 10 about 40 GWd/MTU and 65 GWd/MTU.

11 Figure 1-3 Publicly-Available Data Collected by EPRI for PWR End-Of-Life 12 Rod Internal Pressures at 25°C (77 °F) (Reproduction of Figure 2-1 from 13 Machiels (2013))

14 Publicly-available empirical EOL RIP data are not available for ZIRLO-clad integral fuel burnable 15 absorber (IFBA) rods (zirconium diboride-based), which would have the highest EOL RIP values 16 due to the production of helium from the B-10 neutron reaction. Given the sparsity of the 17 database and the absence of publicly available data for standard M5-clad rods and ZIRLO-clad 18 IFBA rods, predictions are needed for a wide range of advanced cladding alloys, advanced fuel 19 designs, and more current operating conditions.

20 Recent public reports have provided EOL RIP values for ZIRLO-clad IFBA rods from 21 calculations performed with FRAPCON, an NRC-sponsored fuel performance code. The 1-12

For ACRS Review Purposes Only 1 FRAPCON fuel performance code is well-validated for standard BWR and PWR rod 2 predictions, as well as for IFBA PWR rod predictions. Oak Ridge National Laboratory (ORNL) 3 published a set of calculations for over 68,000 Zircaloy-4 and ZIRLO fuel rods irradiated 4 during the first 10 cycles of the Watts Bar Nuclear Plant Unit 1 reactor (Bratton et al, 2015).

5 FRAPCON was used to predict RIPs for standard rods and IFBA rods irradiated for one cycle, 6 two cycles, and three cycles, with each cycle consisting of 18 months. The ORNL report 7 analyzed rods with an isothermal temperature profile. However, an isothermal temperature 8 profile is not a realistic scenario and as such is of limited use in comparing internal pressure 9 and hoop stress results. Additionally, the ORNL report did not use FRAPCONs validated 10 IFBA He-release model and therefore did not adequately capture the interrelated effects of 11 RIP on fuel rod deformation and fission gas release. Therefore, the ORNL report overpredicts 12 the EOL RIP for IFBA rods.

13 More recently, Pacific Northwest National Laboratory (PNNL) used FRAPCON to calculate EOL 14 RIP for three modern fuel designs with three representative dry storage thermal transients, each 15 involving drying operations with a peak cladding temperature of 400 °C (752 °F) (Richmond and 16 Geelhood, 2018). The power histories and axial profiles used were realistic limiting cases 17 meant to give maximum rod internal pressure thus bounding the hoop stress predictions. PNNL 18 generated each power history from a survey of typical maximum power histories for each 19 reactor type. The rod average burnup was 53.23 GWd/MTU for a representative 10 x 10 BWR 20 assembly. The PWR rod average burnup was 55.24 GWd/MTU for the 17 x 17 PWR assembly 21 and 57.71 GWd/MTU for the 17 x 17 IFBA PWR assembly. Although these burnups are lower 22 that the rod average burnup allowed for reactor operation in the United States, experience has 23 shown that rods run with high power are more pressure limited than rods run at low power to 24 higher burnups.

25 PNNLs analyses characterized the effects of fuel design and initial fill gas pressure for 26 determining reasonably bounding cladding hoop stresses (see Section 2 of Richmond and 27 Geelhood, 2018, for additional details on the FRAPCON model and assumptions). The report 28 provides code predictions for maximum EOL RIP for both standard and IFBA rods (Table 1-1),

29 which account for the effects of different canister fill gas pressure on cladding hoop stress 30 (vacuum, medium flow, high flow). EOL RIP values are absolute pressure.

31 Table 1-1 End of Life Rod Internal Pressures (MPa) at a Peak Temperature of 400 °C 32 (752 °F) (From FRAPCON Code Predictions in Richmond and Geelhood, 33 2018) 34 Vacuum Medium Flow High Flow 4.1 x 10-4 MPa 1.0 x 10-1 MPa 6.9 x 10-1 MPa Profile (5.9 x 10-2 psia) (1.5 x 101 psia) (1.0 x 102 psia) 10 x 10 BWR Assembly 5.4 6.1 6.4 17 x17 PWR Assembly 6.2 6.8 7.0 17 x17 PWR Assembly 10.6 11.1 11.5 (IFBA Rods) 35 The cladding hoop stress () is a function of the gas pressure difference across the cladding 36 wall (Pi - P0), where Pi is the rod internal pressure and Po is the external pressure to the rod, the 1-13

For ACRS Review Purposes Only 1 cladding inner diameter (Dmi), and the cladding metal wall thickness (hm), as shown in Eqn. 1-1 2 for the average hoop stress across the cladding wall (Figure 1-7).

3 = [Dmi / (2

• hm)] (Pi - Po) (Eqn. 1-1) 4 The geometrical parameter Dmi/(2

• hm) will tend to increase with burnup due to waterside 5 corrosion of the cladding outer surface, which reduces hm. PNNLs FRAPCON calculations 6 were adjusted for clad thinning due to inner and outer diameter cladding oxidation. Table 1-2 7 provides the results for the maximum cladding hoop stresses for the various corresponding 8 cases in Table 1-1.

9 10 Figure 1-4 Fuel Cladding Tube with Stress Element Displaying Hoop Stress (),

11 Longitudinal Stress (z), Internal Pressure (Pi), Cladding Thickness (hm),

12 External Pressure (Po), Circumferential Coordinate (), and Inner Cladding 13 Diameter (Dmi) 14 15 16 Table 1-2 Maximum Cladding Hoop Stresses (MPa) at a Peak Temperature of 400 °C 17 (752 °F) (From FRAPCON Code Predictions in Richmond et al., 2018)

Vacuum Medium Flow High Flow 4.1 x 10-4 MPa 1.0 x 10-1 MPa 6.9 x 10-1 MPa Profile (5.9 x 10-2 psia) (1.5 x 101 psia) (1.0 x 102 psia) 10 x 10 BWR Assembly 40.0 43.8 41.7 17 x17 PWR Assembly 49.9 53.4 50.5 17 x17 PWR Assembly 84.4 88.1 86.3 (IFBA Rods) 1-14

For ACRS Review Purposes Only 1 PNNL compared their FRAPCON code predictions to the previously-discussed EPRI empirical 2 database by analyzing EOL RIPs and rod void volumes at atmospheric conditions (1.0 x 10-1 3 MPa (1.5 x 101 psia)) and room temperature (25 °C (77 °F)). PNNLs EOL RIP values at these 4 conditions are listed in Table 1-3. Comparison of these results to the EOL RIP values shown in 5 Figure 1-3 demonstrate that PNNLs results fall within EPRIs empirical database. Further, 6 PNNLs code predictions for rod void volume also lie within the EPRIs empirical dataset 7 indicating that the mechanical response of the fuel was accurately modeled. These 8 comparisons give confidence that although PNNLs code predictions evaluated a relatively small 9 number of cases, the results are still considered representative for current LWR designs.

10 Table 1-3 End of Life Rod Internal Pressures at Room Temperature (25 °C (77 °F)) and 11 Atmospheric Conditions (1.0 x 10-1 MPa (1.5 x 101 psia)) (From FRAPCON 12 Code Predictions in Richmond and Geelhood, 2018)

Profile End of Life Rod Internal Pressure (MPa) 10 x 10 BWR Assembly 2.9 17 x17 PWR Assembly 3.1 17 x17 PWR Assembly (IFBA Rods) 5.4 13 PNNLs FRAPCON code predictions support that the maximum cladding hoop stresses remain 14 below 90 MPa (1.3 x 104 psia) for the ZIRLO-clad IFBA rods, even at a peak cladding 15 temperature of 400 °C (752 °F). Therefore, in the absence of publicly-available empirical data 16 on EOL RIPs for IFBA rods and with the evidence provided by the code-predicted values 17 (validated by non-publicly available empirical data), the staff concludes that the EOL RIPs in 18 both standard and IFBA rods result in cladding hoop stresses below the 90 MPa (1.3 x 104 psia) 19 level that has been shown to be capable of producing hydride reorientation in ZIRLO fuel rod 20 cladding (see Section 1.5.4). This would suggest that the mechanical properties of the cladding 21 during drying-transfer, storage and transport operations, would not be meaningfully different 22 from the as-irradiated condition. The above discussion provides a technical basis used by the 23 staff for determining that the radial hydride treatment used for testing of HBU SNF mechanical 24 performance in NRC independent test program used conservative bounding cladding hoop 25 stress conditions (see Section 2.3.4 of this report). The staff notes that the U.S. Department of 26 Energy has sponsored additional empirical measurements on end-of-life rod internal pressures 27 at both ORNL and PNNL. However, these laboratories have not yet publicly-issued their final 28 reports on these data.

29 1.5.4 Ring Compression Testing 30 Ring compression testing (RCT) has been conducted in the United States and Japan to assess 31 effective ductility of cladding with reoriented hydrides following pinch loads (Aomi et al., 2008; 32 Billone et al., 2013; Billone et al., 2014; Billone et al., 2015). The term effective ductility is 33 used throughout this report to differentiate the RCT-measured ductility from the material 34 property elongation (i.e., the classically-defined ductility typically tabulated in the technical 35 literature). RCT of zirconium-based cladding alloys has shown reduced effective ductility when 36 subjected to pinch loads at a sufficiently low temperature; this temperature has been generally 37 referred to as a ductile-to-brittle transition temperature or ductility transition temperature (DTT).

38 In previous NRC-sponsored research, Argonne National Laboratory (ANL) sectioned rings from 39 pressurized-and-sealed rodlets fabricated with cladding from ZIRLO-clad and Zircaloy-4-clad 1-15

For ACRS Review Purposes Only 1 fuel rods irradiated to high burnup (beyond the NRCs peak rod licensing limit in 2 commercial PWRs) (Billone et al., 2013) (Figure 1-10). These rodlets had been heated to a 3 peak temperature of 400 °C (752 °F) (consistent with the guidance limit in ISG-11, Revision 3 4 (NRC, 2003a) and held at this temperature for 1 to 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> with variable target hoop stresses 5 (110 MPa (1.6 x 104 psia), 140 MPa (2.0 x 104 psia)), and then cooled at 5 °C/h (9 °F/h) under 6 conditions of decreasing pressure and hoop stress. This cooling rate does not allow for 7 sufficient time at temperature for appreciable annealing of irradiation hardening to occur, thus 8 allowing a separate assessment of the effects of hydride reorientation. Metallographic 9 examination of one cladding ring surface per rodlet was used to quantify the degree of radial 10 hydride precipitation in terms of the average length of radial hydrides. Several other rings were 11 used to determine the average hydrogen content of the rodlet, along with circumferential and 12 axial variations in hydrogen content. Up to four rings were subjected to RCT to induce pinch 13 loads at test temperatures from 20 °C (68 °F) to 200 °C (392 °F).

14 Figure 1-5 RCT of a Sectioned Cladding Ring Specimen in ANLs Instrons 8511 Test 15 Setup. Tests Were Conducted in the Displacement-Controlled Mode to a 16 1.7-mm Maximum Displacement in a Controlled Temperature Environment (

17 p = RCT Offset Displacement at 12 Oclock Position Relative to Static 18 Support at 6 oclock; Dmo = Outer Diameter of Cladding Metal; p/Dmo =

19 RCT Offset Strain (Percent)) (Reproduction of Figure 6 From Billone et al.,

20 2012))

21 RCT load-displacement curves were used to determine the offset displacement (normalized to 22 the pretest sample outer diameter to give offset strain) as a function of test temperature. The 23 offset strain was plotted against test temperature for each rodlet to determine the DTT 24 (see Figure 1-11). Post-RCT metallographic examinations were also performed to determine 25 the number and extent of cracks that had formed, as well as to generate additional data for the 26 degree of radial hydride precipitation (Billone, et al., 2013).

27 To define an effective ductility for RCT, a 2-percent offset strain (p/Dmo) before a crack 28 extended through more than 50 percent of the cladding wall thickness was chosen to define the 29 transition between ductile and brittle behavior (Billone et al., 2013). In other words, if the 1-16

For ACRS Review Purposes Only 1 sample exhibited more than 2% offset strain before significant cracking occurred (i.e., crack 2 extension exceeding 50% of the cladding thickness), ANL was confident that the samples had 3 adequate effective ductility. For temperatures at which the offset strains dropped below 2%,

4 ANL concluded that the effective ductility was too low to be measured with confidence by the 5 RCT.

6 Figure 1-11 shows representative deformation (i.e., offset strain) curves as a function of the 7 alloy, peak hoop stress at a 400 °C (752 °F) peak cladding temperature, and actual RCT 8 temperature. The figure also shows the radial hydride continuity factor (RHCF), which 9 represents the effective radial length of continuous radial-circumferential hydrides normalized to 10 the wall thickness. ANL used the RHCF for determining the degree and severity of radial 11 hydride precipitation. The radial hydrides in Zircaloy-4 HBU SNF ring specimens were relatively 12 short (i.e., RHCF of 9 percent for a peak hoop stress of 110 MPa (1.6 x 104 psia), and 16 13 percent for a peak hoop stress of 140 MPa (2.0 x 104 psia)) and the effective ductility increased 14 gradually with temperature. In ZIRLO-clad HBU SNF ring specimens, the radial hydrides were 15 longer (i.e., RHCF of 30 percent for a peak hoop stress of 110 MPa (1.6 x 104 psia), and 65 16 percent for a peak hoop stress of 140 MPa (2.0 x 104 psia)) and the effective ductility increased 17 sharply with the increase in RCT temperature. ANL fit the limited ZIRLO data points with S-18 shaped curves (hyperbolic tangent functions) typical of materials that exhibit a ductile-to-brittle 19 transition. The data show that the DTT shifted from around room temperature in a cladding 20 material with short radial hydrides to higher values in a cladding material with longer radial 21 hydrides. The limited data also indicates a trend of lower DTTs for materials with lower peak 22 cladding stresses.

1-17

For ACRS Review Purposes Only 14 9+/-5% RHCF; Zry-4 @ 110 MPa 12 16+/-4% RHCF; Zry-4 @140 MPa Offset Strain (percent) 30+/-12% RHCF; ZIRLO' @ 110 MPa 10 65+/-17% RHCF; ZIRLO' @140 MPa 8

6 4

Ductile 2

Brittle 0

0 25 50 75 100 125 150 175 200 225 RCT Temperature (°C) 1 Figure 1-6 Effective Ductility vs. RCT for Two PWR Cladding Alloys Following Slow 2 Cooling from 400°C (752 °F) at Peak Target Hoop Stresses of 110 MPa 3 (1.6 x 104 psia) and 140 MPa (2.0 x 104 psia) (From Billone et al., 2013) 4 ANL also conducted RCT research under DOE sponsorship. It obtained results for the following 5 (Billone et al., 2014; Billone et al., 2015):

6

• HBU Zircaloy-4 in the as-irradiated condition with moderate-to-high hydrogen content 7

• HBU ZIRLO in the as-irradiated condition and following simulated drying-storage at peak 8 temperatures of 400 °C (752 °F) and 350 °C (662 °F) with peak hoop stresses from 9 80 MPa (1.2 x 104 psia) to 94 MPa (1.4 x 104 psia) 10

• HBU M5 in the as-irradiated condition and following simulated drying-storage at 11 400 °C (752 °F) with peak hoop stresses of 90 MPa (1.3 x 104 psia), 110 MPa 12 (1.6 x 104 psia), and 140 MPa (2.0 x 104 psia) 13 ANL conducted two additional tests with HBU ZIRLO cladding subjected to three drying cycles 14 (e.g., from 400 °C (752 °F) to 300 °C (572 °F) and from 350 °C (662 °F) to 250 °C (482 °F)) at 15 peak hoop stress of about 90 MPa (1.3 x 104 psia). The latter results suggest that multiple 16 drying cycles have no effect on the length of radial hydrides or the DTT at this low stress level.

17 Figures 1-12 through 1-14 show results for Zircaloy-4, ZIRLO, and M5 in both as-irradiated and 18 hydride-reoriented condition following cooling from 400°C (752 °F) (Billone et al., 2014; Billone 19 et al., 2015).

1-18

For ACRS Review Purposes Only 14 300+/-15 wppm H High Burnup 12 640+/-140 wppm H Zircaloy-4 520+/-90 wppm H Offset Strain (percent) 10 615+/-82 wppm H 113 MPa at 400°C 8

6 145 MPa As-Irradiated at 400°C 4

Ductile 2

Brittle 0

0 25 50 75 100 125 150 175 200 225 RCT Temperature (°C) 1 Figure 1-7 Ductility Data, as Measured by RCT, for As-Irradiated Zircaloy-4 and 2 Zircaloy-4 Following Cooling from 400 °C (752 °F) Under Decreasing 3 Internal Pressure and Hoop Stress Conditions (From Billone et al., 2013) 1-19

For ACRS Review Purposes Only 14 530+/-70 wppm H High Burnup ZIRLO 12 535+/-50 wppm H As-Irradiated 530+/-115 wppm H Offset Strain (percent) 480+/-131 wppm H 10 385+/-80 wppm H 8

111 MPa 6 at 400°C 4 80 MPa 89 MPa at 400°C at 400°C 2 Ductile Brittle 0

0 20 40 60 80 100 120 140 160 180 RCT Temperature (°C) 1 Figure 1-8 Ductility Data, as Measured by RCT, for as-Irradiated ZIRLO and ZIRLO 2 Following Cooling from 400 °C (752 °F) Under Decreasing Internal Pressure 3 and Hoop Stress Conditions (From Billone et al., 2013) 1-20

For ACRS Review Purposes Only 14 90 MPa High Burnup M5 12 at 400°C 10 76+/-5 wppm H Offset Strain (%)

58+/-15 wppm H (90 MPa) 8 72+/-10 wppm H (111 MPa)

As-Irradiated 94+/-4 wppm H (142 MPa) 6 111 MPa 4 at 400°C 142 MPa at 400°C Ductile 2

Brittle 0

0 25 50 75 100 125 150 175 200 225 RCT Temperature (°C) 1 Figure 1-9 Ductility Data, as Measured by RCT, for As-Irradiated M5 and M5 Following 2 Cooling from 400 °C (752 °F) under Decreasing Internal Pressure and Hoop 3 Stress Conditions (From Billone et al., 2013) 4 The staff recognizes the uncertainties associated with the ductility curve fits of ANLs RCT data 5 because of the limited number of data points. However, the limited results appeared to support 6 the following general conclusions: (1) the DTT generally increases with increasing hoop 7 stresses (i.e., the ductility transition shifts to higher cladding temperature), (2) both the 8 susceptibility to radial hydride precipitation and ductility changes depend on cladding type and 9 initial hydrogen content, and (3) depending on the cladding and test conditions, the DTT can 10 occur at temperatures in the range of 20 °C (68 °F) to 185 °C (365 °F). The results for as-11 irradiated Zircaloy-4 are consistent with studies by Wisner and Adamson (1998) and Bai et al 12 (1994). The staff considered these conclusions when defining limiting conditions for inducing 13 radial hydrides and conducting fatigue and bending testing of HBU SNF (see Chapter 2).

14 It is important to note that the DTT is not an intrinsic property of a cladding alloy material with a 15 given homogeneous composition, in the classical metallurgical sense, but it is highly dependent 16 on the composite microstructure (hydride-zirconium matrix, as determined by reactor operating 17 conditions), fabrication conditions (degree of cold working, recrystallization) and the operating 18 conditions during drying-transfer, storage or transportation (peak cladding temperature, peak 19 hoop stress, temporal cooling profile). Further, the DTT was established based on an arbitrarily-20 defined performance criterion (e.g., 50 percent cladding through-wall crack prior to 2-percent 21 offset strain deformation), and based on a limited number of data points for each cladding alloy.

22 It is also important to note that, due to the radial and axial temperature gradients in a DSS or 23 transportation package, it is highly likely that only a small fraction of the cladding in a given 24 assembly will reach high enough temperatures and hoop stresses to have sufficient hydride 25 reorientation during cooling. Those hotter axial locations of the cladding will likely be the last to 26 reach a DTT during transport.

1-21

For ACRS Review Purposes Only 1 1.5.5 Staffs Assessment of Ring Compression Testing Results 2 As previously discussed, the staff has long expected that hydride reorientation would not 3 compromise cladding integrity due to fuel rod bending (i.e., bending expected during normal 4 conditions of storage and transport), since the principal tensile stress field associated with rod 5 bending caused by lateral inertia loads is parallel to both radial and circumferential hydrides 6 (Tang et al., 2015). The staff has considered that any reduced cladding ductility due to hydride 7 reorientation could only potentially compromise the analyzed fuel configuration for pinch loads 8 experienced during drop accident scenarios, if the fuel had significantly cooled during the 9 transportation period. More specifically, the staff had expressed concern that reorientation 10 could decrease failure stresses and strains in response to transportation-induced pinch loads 11 during a 9-m (30-ft) drop scenario as a result of rod-to-grid spacer contact, rod-to-rod contact, or 12 rod-to-basket contact.

13 To address the concern of reduced ductility during drop accidents, the staff previously proposed 14 varied approaches to demonstrate that the failure limits for as-irradiated cladding (as used in the 15 design-basis structural evaluations) would continue to be adequate even if hydride reorientation 16 occurred. One of these approaches was based on justifying an RCT-measured DTT for each 17 cladding alloy in the proposed fuel contents, and demonstrating that the minimum cladding 18 temperature remained above the RCT-measured DTT for the entire duration of transport. The 19 minimum cladding temperature assumed for transport operations would need to be bounding to 20 the contents upon consideration of the cold temperature requirement in 10 CFR 71.71(c)(2), i.e.

21 an ambient temperature of -40 °C (-40 °F) in still air and shade. If these conditions were met, 22 then mechanical properties of the as-irradiated cladding material (i.e., material that did not 23 account for the precipitation of radial hydrides), would be considered adequate for the structural 24 evaluation.

25 As an alternative approach, if the applicant could not reasonably demonstrate that sections of 26 the fuel cladding remained above the RCT-measured DTT during the entire duration of 27 transport, the staff proposed that the application provide additional safety analyses assuming 28 hypothetical reconfiguration of the HBU fuel contents. If neither of these two approaches is 29 satisfactory for demonstrating compliance with 10 CFR Part 71 regulations, then the staff would 30 expect that the fuel would be canned and classified as damaged.

31 Since proposing these approaches, the staff has reevaluated whether results from RCT of 32 defueled specimens are accurately representative or if they are overly conservative relative to 33 the actual hoop-loading conditions experienced by the fuel during a 9-m (30-ft) drop. During 34 RCT, the circumferential (hoop) tensile bending stress is perpendicular to the plane of the radial 35 hydrides, which is different from the relative orientation of the applied stress and hydrides under 36 axial tensile bending where the longitudinal (axial) tensile bending stress is always parallel to 37 the plane of both the circumferential and radial hydrides. The orientation of the tensile stress is 38 expected to make a difference in the response of the cladding.

39 The RCT defined a DTT used to determine cladding failure due to pinch loads. However, it is 40 necessary to consider the importance of this failure mode in the determination of cladding 41 integrity in the event of a drop accident. To do this, the RCT must be examined for what it is, a 42 test in which diametrically-opposed concentrated compressive forces are applied to a fuel 43 cladding longitudinal segment that does not contain fuel. During NCT and HAC side drops, the 44 fuel rod is loaded by lateral inertia loads that are resisted by distributed loads applied to the 45 bottom of the rod at the flexible grid spacer springs (Figure 1-15). Further, the inertia load in the 1-22

For ACRS Review Purposes Only 1 rod is transferred to the grid spacer support as a shear force in the cladding (and pellets) not as 2 a concentrated load at the top of the rod.

Single Rod Model Single Assembly Model 3 Figure 1-10 Geometric Models for Spent Fuel Assemblies in Transportation Packages 4 (Reproduction, in Part, Of Figure 10 from Sanders et al., 1992) 5 Given that the forces and displacements in the RCT are measurably different from the actual 6 forces and displacements applied to the rod at the grid spacer support, it is not likely that the 7 pinch-mode of failure will play a significant role in undermining cladding integrity. To quantify 8 the difference between these loading cases, the staff analyzed two ring segments for different 9 loading conditions and the change in diameter calculated. In the first case the ring segment 10 was loaded by diametrically-opposed compressive forces like those of RCT (Case 1, Table 17, 11 Roark and Young (1975)). In the second case the ring segment was supported at the bottom by 12 a concentrated reaction and loaded by a downward load uniformly distributed around the 13 circumference of the ring to simulate a shear loading as in a side drop (Case 13, Table 17, 14 Roark and Young (1975)). In both cases the total applied load was the same. The ratio of the 15 change in diameter of the second case to the first case is 0.48. Thus, the diametrically-opposed 16 compressive forces produced more than twice the displacement when compared to the 17 circumferentially distributed load. In addition, the gap at the pellet-cladding interface is 18 generally closed at rod segments irradiated to high burnup due to pellet expansion during 19 irradiation. The closed gap will limit the deflection of the cladding before experiencing 20 mechanical resistance by the pellet. Thus, the staff considers that, under a pinch load, 21 ovalization of the cladding cross-section is very unlikely and any circumferential bending stress 22 that does exist will be negligible. The RCT conducted to date does not account for the rods 23 resistance to ovalization provided by the pellet.

24 Based on the RCT load-displacement data, ANL defined the effective cladding ductility (i.e.,

25 the transition between ductile and brittle behavior) to be a 2-percent offset strain prior to a 26 crack extending through more than 50 percent of the cladding wall (Billone et al., 2013). If the 27 strains experienced during RCTs diametrically-opposed loads result in twice those that would 28 be experienced during lateral inertial loads, then the DTT is likely to shift to lower 29 temperatures (potentially room temperature or lower). Therefore, the staff considers that the 30 DTT defined by RCT experiments is overly conservative and not representative of actual fuel 31 and stress conditions during NCT and HAC drop scenarios. The DOE is planning on 32 sponsoring a research program in which 25 HBU fuel rods will undergo testing to determine 33 their characteristics, material properties, and rod performance following representative drying-34 transfer and cooldown (Hanson et al., 2016). The staff expects that material property testing 1-23

For ACRS Review Purposes Only 1 conducted under this program will provide confirmation that the cladding displacements 2 experienced by fueled cladding specimens during RCT will be lower than those measured in 3 defueled specimens and that ductility during accident drop scenarios is not compromised.

4 Results from the static and fatigue bend testing discussed in Chapter 2 further justify the 5 staffs conclusion that the pellet imparts structural support to the mechanical performance of 6 the fuel rod, as previously evaluated by finite element analysis (Machiels, 2005).

1-24

For ACRS Review Purposes Only 1 2 ASSESSMENT OF STATIC BENDING AND FATIGUE STRENGTH 2 RESULTS ON HIGH BURNUP SPENT NUCLEAR FUEL 3 2.1 Introduction 4 The sealed canister, cask cavity, or overpack generally serves as the primary barrier in a dry 5 storage system (DSS) or transportation package for protecting against the release of radioactive 6 solid particles or gases from the loaded spent nuclear fuel (SNF) to the atmosphere. The spent 7 fuel cladding also serves as a confinement or containment barrier for preventing radioactive 8 solid particles and fission gasses from being released into the interior cavity of the DSS or 9 transportation package. The cladding not only provides a barrier for preventing the release of 10 radioactive material but also prevents fuel reconfiguration during storage and transport 11 operations. Therefore, the integrity of the cladding is an essential component of a defense-in-12 depth strategy to protect the public health and safety.

13 Until recently, research to understand the structural behavior of spent fuel rods during 14 transportation and storage has focused entirely on obtaining mechanical and strength properties 15 of spent fuel cladding. As a result, the flexural rigidity and structural response of fuel rods 16 during normal and accident events have been based on the mechanical and strength properties 17 of only the cladding. The contribution of the fuel pellets to increasing the flexural rigidity of the 18 rod has been neglected. However, recent research discussed in NUREG/CR-7198, Revision 1, 19 Mechanical Fatigue Testing of High-Burnup Fuel for Transportation Application, issued 20 October 2017 (NRC, 2017a), on the static bending response and fatigue strength of fuel rods 21 considered as a composite system of cladding and fuel pellets, has begun to provide some of 22 the necessary data to allow a more accurate assessment of the structural behavior of the 23 composite fuel rod system under normal conditions of transport (NCT) and hypothetical accident 24 conditions (HAC), as well as DSS drop and tip-over events.

25 The following discussion assesses the results from the NRCs independent test program on the 26 mechanical performance of high burnup (HBU) SNF under static and dynamic bending 27 conditions. Section 2.2 discusses the available fuel rod composite static and dynamic bending 28 empirical data and its acquisition. Section 2.3 describes the application of the static bending 29 empirical data for the evaluation of design-basis drop accidents in storage and transportation, 30 and the development of a composite rod analytical model. Section 2.4 discusses the application 31 of the dynamic bending empirical data to the evaluation of fatigue during transportation.

32 2.2 Cyclic Integrated Reversible Fatigue Tester 33 In 2009, the U.S. Nuclear Regulatory Commission (NRC) tasked Oak Ridge National Laboratory 34 (ORNL) with investigating the flexural rigidity and fatigue life of high burnup (HBU) SNF 35 (NRC, 2017a). The testing was designed to evaluate the fuel rod as a composite system, 36 including the presence of intact fuel inside the cladding and any pellet/cladding bonding effects.

37 The project proceeded in two phases. Phase I involved testing HBU SNF in the as-irradiated 38 state, where hydrides are expected to be predominantly in the circumferential-axial orientation.

39 Phase II involved testing HBU SNF segments subjected to a treatment designed to reorient the 40 hydrides in the cladding to be predominantly in the radial-axial orientation. All testing was 41 conducted at room temperature, which is expected to result in the most limiting cladding 42 ductility.

2-1

For ACRS Review Purposes Only 1 In response to the NRC tasking, in 2011, ORNL proposed a bending fatigue system for testing 2 HBU SNF rods. The system is composed of a U-frame equipped with load cells for imposing 3 pure bending loads on the SNF rod test specimen and measuring the in-situ curvature of the 4 fuel rod during bending using a set-up of three linear variable differential transformers (LVDT) 5 (Figure 2-1). Pure bending is a condition of stress in which a bending moment is applied to a 6 beam without the simultaneous presence of axial, shear, or torsional forces.

Universal testing machine links Rigid arms Connecting plates (top)

Load cell Rod specimen Three LVDTs for curvature measurement (middle)

LVDT clamp SNF rod segment End-blocks (bottom) 2-2

For ACRS Review Purposes Only 1 Figure 2-1 Horizontal Layout of ORNL U-Frame Setup (Top), Rod Specimen and Three 2 Lvdts for Curvature Measurement (Middle), and Front View of CIRFT 3 Installed in ORNL Hot Cell (Bottom) (Figure 4 from NUREG/CR-7198, 4 Revision 1 (NRC, 2017a))

2-3

For ACRS Review Purposes Only 1 On August 19, 2013, a testing system was installed in a hot cell at ORNLs Irradiated Fuels 2 Examination Laboratory and formally named the cyclic integrated reversible-bending fatigue 3 tester (CIRFT). After tuning of the test system and performance of benchmark testing in 4 September 2013, testing began on HBU SNF rod segments with intact Zircaloy-4 cladding 5 irradiated in the H.B. Robinson Steam Electric Plant (HBR) Unit 2. The rod-average fuel burnup 6 for the 15 x 15 PWR assembly was 67 GWd/MTU. Table 2-1 identifies the burnup for each 7 tested rod segment.

8 Table 2-1 Specifications of Rod Specimens used in NRC-Sponsored HBU SNF Test 9 Program (Reproduced in Part from Table 2, NUREG/CR-7198, Revision 1 (NRC, 2017a))

Burnup Estimated Hydrogen of Span Specimen Label (GWd/MTU) (wppm)

Static Tests S1 66.8 550 - 750 S2 66.5 360 - 550 S3 66.5 550 - 750 S4 66.5 550 - 750 Dynamic Tests D0 66.5 360 - 550 D1 63.8 550 - 750 D2 63.8 550 - 750 D3 66.5 550 - 750 D4 66.5 360 - 550 D5 66.5 360 - 550 D6 66.5 550 - 750 D7 66.5 550 - 750 D8 66.8 550 - 750 D9 66.5 550 - 750 D10 66.8 550 - 750 D11 63.8 550 - 750 D12 63.8 550 - 750 D13 66.5 750 - 800 D14 66.5 750 - 800 D15 66.5 750 - 800 HR1 63.8 360 - 400 HR3 63.8 360 - 400 2-4

For ACRS Review Purposes Only Burnup Estimated Hydrogen of Span Specimen Label (GWd/MTU) (wppm)

HR4 63.8 360 - 400 1 Under Phase 1 testing, ORNL completed four static tests under displacement control at the rate 2 of 0.1 mm/s to a maximum displacement of 12.0 mm. In early November 2013, the benchmark 3 and static test results were critically reviewed at a meeting between representatives from the 4 NRC and ORNL. Dynamic testing was then initiated, and 16 cyclic tests were completed in the 5 Irradiated Fuels Examination Laboratory. Load ranges applied to the CIRFT varied, to produce 6 bending moments in the rod, from +/-5.08 to +/-35.56 N*m. There were 12 dynamic tests with rod 7 fracture and 4 tests without rod fracture. One of the cyclic tests reached 1.3 x 107 cycles with 8 no rod fracture. The test was terminated as higher cycles would not be expected during actual 9 transport.

10 Phase II testing began in 2016, again using HBR HBU SNF rods with intact Zircaloy-4 cladding, 11 which had been subjected to an aggressive hydride reorientation treatment (HRT) (see 12 Section 2.3.4). ORNL completed testing on four specimens in the CIRFT following an HRT:

13 one in static loading (hereafter referred to as HR2), and three in dynamic loading (hereafter 14 referred to as HR1, HR3, and HR4). The fatigue lifetime and flexural rigidity of these samples 15 were compared to the results obtained in Phase I for as-irradiated samples.

16 The following observations can be made about the results of the static testing:

17

• The HBR HBU SNF rods in the as-irradiated state exhibited a multiple-stage constitutive 18 response, with the two linear stages followed by a nonlinear stage. The flexural rigidity at 19 the initial stage was 63 to 78 Nm2, corresponding to an elastic modulus of 101 to 20 125 GPa. The flexural rigidity at the second stage was 55 to 61 N*m2, and the 21 corresponding elastic modulus was 88 to 97 GPa.

22

• Most HBR HBU SNF rods in the as-irradiated state under static unidirectional loading 23 fractured at a location coincident with the pellet-to-pellet interface, as validated by the 24 posttest examinations showing pellet end faces in most of the fracture surfaces.

25 Fragmentation of the pellets also occurred to a limited degree, along with cladding 26 failure.

27

• The static CIRFT results indicate a significant increase in a fueled SNF rods flexural 28 rigidity compared to a calculated response for cladding only. This applied to both as-29 irradiated and HRT SNF rods.

30

• For the HBR HBU SNF rods, the static CIRFT test results show that at bending moments 31 less than 30 N*m the flexural rigidities of the as-irradiated rods and the HRT HR2 rod are 32 essentially the same.

33

• The sample subjected to an HRT and tested under a static bending load showed 34 reduced flexural rigidity at higher loads compared to as-irradiated samples.

35 Nevertheless, material tested in the as-irradiated and HRT state both had higher flexural 36 rigidity than the calculated cladding-only response.

37

• The static CIRFT test result for HR2 supports the pretest expectation (hypothesis) that 38 because the tensile bending stress in the cladding is parallel to the plane of both the 2-5

For ACRS Review Purposes Only 1 radial and circumferential hydrides, the presence of radial hydrides would not 2 significantly alter the flexural response when compared to the case where only 3 circumferential hydrides are present.

4

• The CIRFT test methodology and the methodology developed in NUREG-2224 for 5 calculating cladding stress and strain are applicable to all current commercial power fuel 6 rod types, and the use of cladding-only properties to calculate cladding stress and strain 7 is always conservative.

8

• The HBR HBU SNF rods in the as-irradiated state survived static unidirectional bending 9 to a maximum curvature of 2.2 to 2.5 m-1, or a maximum moment of 85 to 87 Nm. The 10 maximum static unidirectional bending values were bounded by the CIRFT device 11 displacement capacity. The maximum equivalent strain was 1.2 to 1.4 percent.

12

• Based on the static CIRFT test results, the lower-bound safety margin against fuel rod 13 failure during an HAC side drop event is 2.35 assuming the side drop imparts a 50-g 14 load to the package body (see Section 2.3.4.2).

15 The following observations can be made about the results of the dynamic testing:

16

• The fatigue life of HBR HBU SNF rods in the as-irradiated state in the cyclic tests 17 depended on the level of loading. Under loading with moments of +/-8.20 to +/-33.67 18 Nmnamely +/-0.066 to +/-0.335percent3335 percent strain the fatigue life ranged from 19 5.5 x 103 to 2.3 x 106 cycles.

20

• The -N curve of the HBR HBU SNF rods in the as-irradiated state can be described by 21 a power function of y = 3.839*x-0.298, where x is the number of cycles to failure, and y is 22 the strain amplitude (percent).

23

• The failure of HBR HBU SNF rods under cyclic loading often occurred near pellet-to-24 pellet interfaces.

25 The following sections provide an assessment by the NRC staff (the staff) of ORNLs CIRFT 26 data and present conclusions as to the expected structural performance of HBU SNF during dry 27 storage and transportation.

28 2.3 Application of the Static Test Results 29 When evaluating the HAC 9-m (30-ft) drop test, as required by Title 10 of the Code of Federal 30 Regulations (10 CFR) 71.73(c)(1), two drop orientations produce distinctly different structural 31 behaviors in the fuel rods. These orientations are the side drop and the end drop (Figure 2-2).

32 In the side drop, lateral inertia loads are applied to the fuel rods, and bending dominates the 33 structural response. In the end drop, axial compression and the associated buckling of the fuel 34 rod dominates the structural response. For a side-drop event, the CIRFT static bending test 35 results from NUREG/CR-7198, Revision 1 (NRC, 2017a), can be directly applied to quantify the 36 fuel rod structural response. For the end drop, the presence of axial compression in the fuel rod 37 represents a force component that was not present in the CIRFT static bending tests. This, 38 however, does not pose a problem since the CIRFT static test results can be used to 39 conservatively quantify the effect of the fuel pellets on increasing the flexural rigidity of the rods 40 to resist buckling.

2-6

For ACRS Review Purposes Only 1 Figure 2-2 Schematic Diagram of End and Side Drop Accident Scenarios 2 (Revised Figure 5-168 from Patterson and Garzarolli (2015))

3 2.3.1 Spent Fuel Rod Behavior in Bending 4 The behavior of a fuel rod in bending generally depends on three things: (1) the type of loading, 5 (2) the bond between the cladding and fuel, and (3) the behavior of the pellet-pellet interface.

6 Fundamentally, there are two types of bendingbending without shear and bending with shear.

7 Bending without shear is pure bending (i.e., constant moment or curvature, as exhibited in the 8 ORNL CIRFT tests) and produces no shear stress at the interface between the cladding and 9 fuel pellet. Pure bending is a special case that does not often occur in practice. What occurs 10 more often is the case of a laterally-supported fuel rod subjected to a transverse inertial loading, 11 as in a side drop, where the rod is subjected to both bending and shear forces.1 Although both 12 bending and shear are acting, the structural response would be expected to be different, 13 depending on whether the cladding is bonded to the fuel pellet.

14 2.3.2 Composite Behavior of a Spent Fuel Rod 15 Until recently, experimental testing on the structural behavior of SNF rods during transportation 16 and storage has focused primarily on obtaining mechanical properties that consider only the 17 material strength of the cladding. Historically, the fuel pellets contribution to the flexural rigidity 18 and structural response of the fuel rod during normal and accident conditions has been ignored 19 because of the lack of experimental bending test data, although it has been previously 20 evaluated by finite element analysis to improve the composite rods mechanical response 21 (Machiels, 2005)). Recent research sponsored by the U.S. Nuclear Regulatory Commission 22 (NRC) on the static bending response and fatigue strength of HBU SNF rods with the presence 23 of the fuel pellets has provided data necessary to more accurately assess the structural 1 Because the fuel behaves in a brittle manner while the cladding behaves in a ductile manner, all of the bending tensile stresses will occur in the cladding. The cladding and fuel will resist the shear forces, but for simplicity, it can be conservatively assumed that all of the shear is resisted by the cladding. A simple calculation shows that during a side drop event, the uniformly loaded fuel rod spanning over multiple grid spacers will have maximum tensile stresses due to bending that are more than an order of magnitude greater than the maximum tensile stresses due to shear. Therefore, bending dominates the response of the fuel rod, and this is why the CIRFT tests can accurately represent the behavior of an actual fuel rod during a side drop event.

2-7

For ACRS Review Purposes Only 1 behavior of the composite HBU SNF rod system (NRC, 2017a). These results have provided an 2 opportunity for the NRC to assess the conservatism associated with only assuming the 3 mechanical strength of the cladding in the design-basis structural evaluations of DSSs and 4 transportation packages.

5 A spent fuel rod is considered to be a composite system consisting of cladding and fuel. The 6 structural response of the fueled-rod composite system is usually explained as follows.

7 On one hand, if the pellet is not bonded to the cladding, displacement compatibility is not 8 maintained at the pellet-cladding interface, and composite action does not occur. In this case, 9 the flexural rigidity is given by the following equation, where the fuel is assumed to be a 10 homogeneous solid:

11 EI = EcIc + EpIp (Eqn. 2-1) 12 That is, the flexural rigidity is equal to the sum of the individual flexural rigidities of the cladding 13 and fuel pellets, where Ec and Ic are the elastic modulus and moment of inertia of the cladding, 14 respectively, and Ep and Ip are the elastic modulus and moment of inertia of the pellet, 15 respectively.

16 On the other hand, if the pellet is bonded to the cladding, displacement compatibility is 17 maintained at the pellet-cladding interface and composite action occurs. In this case, the 18 flexural rigidity is calculated by transforming the pellet properties into equivalent cladding 19 properties (i.e., by multiplying the pellet moment of inertia by Ep/Ec). This is the same technique 20 commonly used for reinforced concrete (Winter and Nelson, 1979).

21 The remainder of this section will explain the behavior of composite systems, in general, and 22 then specifically address the spent fuel rod composite system by assuming the fuel material is a 23 homogenous uncracked solid. To fully understand the unique behavior of this composite 24 system, the bending behavior of a more general composite beam will be discussed. Consider a 25 composite concrete and steel I-beam where a concrete slab, rectangular in cross-section, is 26 poured onto the top flange of a steel I-beam (Figure 2-3). This type of composite beam is 27 commonly found in highway bridge construction. Assume the concrete and steel beam are 28 simply supported and a concentrated load is applied at mid-span. If the concrete slab and steel 29 beam are not bonded to each other, no shear transfer takes place at the interface between the 30 steel and concrete, and the flexural rigidity (EI) is equal to the sum of the individual flexural 31 rigidities of the concrete slab and steel beam taken separately.

2-8

For ACRS Review Purposes Only 1 Figure 2-3 Typical Composite Construction of a Bridge 2 On the other hand, if the concrete slab and steel beam are bonded to each other, as typically 3 done using shear studs, then shear transfer takes place and the concrete slab and steel beam 4 act as a composite section. In this case, the flexural rigidity of the composite beam will be 5 significantly greater than the sum of the individually flexural rigidities taken separately. This 6 example of a concrete slab bonded to the top flange of a steel beam illustrates the behavior of a 7 composite system where the centers of gravity of each of the two components (i.e., concrete 8 slab and steel I-beam) are not coincident.

9 For the special case where the centers of gravity of the two components are coincident, the 10 flexural rigidity of the composite section is always equal to the sum of the flexural rigidities of the 11 individual components regardless of whether the components are bonded or unbonded. The 12 following example illustrates this concept. Consider a simply supported span composed of two 13 beams, each with a rectangular cross-section 2 in. wide, and 6 in. deep (i.e., a 2 x 6). Let the 14 2 x 6s be configured one on top of the other, where the centers of gravity (cgs) are not 15 coincident as shown in Figure 2-4a. If the beams are unbonded, the moment of inertia of the 16 section (I = bh3/12 per beam), is equal to: 2 x 2 in. x (6 in.)3/12 = 72 in.4. If they are bonded, 17 then the moment of inertia of the section is equal to: 2 in. x (2 x 6 in.)3/12 = 288 in.4.

2-9

For ACRS Review Purposes Only 1 Figure 2-4 Influence of cg Position on Composite Beam Stiffness:

2 (a) cgs Are Not Coincident, (b) cgs Are Coincident 3 Now let the 2 x 6s be configured as shown in Figure 2-4b, where the cgs are aligned on the 4 same bending axis (i.e., they are coincident). If they are unbonded, the moment of inertia of 5 the section is: 2 x 2 in. x (6 in.)3/12 = 72 in.4. If they are bonded I = 2 x 2 in. x (6 in.)3/12 = 72 6 in.4. Thus, when the cgs of the 2 x 6s are "coincident" the flexural rigidity of the beam is the 7 sum of the individual flexural rigidities of the 2 x 6s regardless of whether the 2 x 6s are bonded 8 or unbonded. While previously unrecognized, this is the situation with a spent fuel rod, where 9 the cladding cylindrical tube and the spent fuel cylindrical solid section have coincident cgs.

10 Thus, for a spent fuel rod, where the fuel is assumed to be a homogeneous solid, the flexural 11 rigidity is given by Equation 2-1, regardless of whether or not the fuel is bonded to the cladding.

12 All moments of inertia are taken about the neutral axis of the fuel rod.

13 2.3.3 Calculation of Cladding Strain from CIRFT Static Bending Data 14 The objective of this section is to develop a simple methodology that uses the CIRFT static test 15 data for fully-fueled composite spent fuel rods to evaluate spent fuel rod cladding strain. The 16 methodology presented here to determine cladding response (i.e., cladding stresses and 17 strains) is based on a set of assumptions that are consistent with those made by ORNL in its 18 presentation of CIRFT results in NUREG/CR-7198, Revision 1 (NRC, 2017a). These 19 assumptions, which are discussed in greater detail below, are based on the integrated average 20 response of the fuel rod along its gauge length. Further, the methodology recognizes the actual 21 behavior of the fuel rod where the fuel is no longer a homogenous solid, as previously 22 discussed in Section 2.3.2 (i.e., the fuel pellets crack at their interface during bending).

2-10

For ACRS Review Purposes Only 1 Figure 2-5 Images of Cladding-Pellet Structure in HBU SNF Rod (66.5 GWd/MTU, 2 40 - 70 µm Oxide Layer, 500 wppm Hydrogen Content in Zircaloy-4):

3 (a) Overall Axial Cross Section and (b) Enlarged Area (Revised Figure 33 4 from NUREG/CR-7198, Revision 1 (NRC, 2017a))

5 The fuel rod composite system (Figure 2-5) is composed of cladding, which exhibits ductile 6 behavior, and the fuel pellet, which exhibits brittle behavior. In a spent fuel rod subject to 7 bending, where the fuel is a homogenous solid, the neutral axis is at the center of the rod cross-8 section, provided that the brittle fuel does not crack in tension. Once the fuel cracks, the neutral 9 axis will shift toward the compression side of the cross-section. The ORNL tests show that the 10 region of the fuel weakest in tension is at the pellet-pellet interface. When the pellet-pellet 11 interface cracks, the tensile stress in the cladding at the crack face will increase significantly.

12 On either side of the crack face the shear stress between the cladding and fuel is high and 13 decreases parabolically with distance from the crack (Figure 2-6). The high tensile stress in the 14 cladding at the crack face also decreases parabolically with distance from the crack. Thus, the 15 cladding tensile stresses will vary significantly along the length of the rod; they are highest at the 16 crack face and much lower away from the crack face. Even though this behavior is known to 17 occur, only the average tensile bending stress can be calculated from the static test results 18 because the measured curvature is the integrated average curvature over the measurement 19 length (gauge length) of the rod.

2-11

For ACRS Review Purposes Only 1 Figure 2-6 Approximate Extreme Fiber Tensile Stresses Between Pellet-Pellet Crack 2 The LVDTs measure displacements at three locations on the test specimen. The distance 3 between the first and third probes is the gauge length of the specimen. Because the bending 4 moment is constant along the gauge length, it would be expected that several pellet-pellet 5 interface cracks would develop within the gauge length. That being the case, the cladding 6 tensile stresses and strains along the gauge length will vary significantly. However, this 7 variation in strain along the gauge length was not, and cannot be, measured. What was 8 measured is the average curvature along the gauge length. Therefore, only the average tensile 9 strain (i.e., the smeared tensile strain) can be calculated. The average tensile strain, , along 10 the gauge length is equal to the curvature, , multiplied by the distance to the neutral axis, ymax:

11 =

• ymax (Eq. 2-2) 12 However, ymax can vary significantly along the gauge length. At a section where the fuel has not 13 cracked, ymax is equal to the outer radius, r. At a pellet-pellet interface crack, ymax would be 14 greater than the radius but less than the diameter. However, because the measured and 15 calculated results are averages over the gauge length, a convention must be adopted for 16 calculating cladding strain and this convention must be consistently applied throughout. The 17 convention used in NUREG/CR-7198, Revision 1 (NRC, 2017a), and adopted in this document 18 to convert average curvature to average cladding strain, is to assume that the distance from the 19 tensile face of the cladding to the neutral axis is equal to the outside radius, r.

20 Average cladding tensile stress, , should be calculated directly from average cladding strain 21 using the following equation:

22 =

• Ec (Eq. 2-3) 23 Equation 2-3 provides a consistent and compatible relationship between stress and strain.

2-12

For ACRS Review Purposes Only 1 2.3.4 Calculation of Cladding Strain Using Factored Cladding-Only Properties 2 The following discussion describes a methodology that can be easily implemented to calculate 3 the cladding tensile strain and stress and fuel rod flexural rigidity using only cladding-only 4 properties. Section 4.2.2 of NUREG/CR-7198, Revision 1 (NRC, 2017a), presents analyses 5 comparing the measured flexural rigidity from the CIRFT static test results to the calculated 6 flexural rigidity values using the validated cladding-only mechanical property models developed 7 by PNNL (Geelhood et al., 2008). The purpose of the comparison was to investigate the effect 8 of fuel pellets on the fuel rod's flexural rigidity and cladding strain.

9 10 Figure 2-7 Comparison of CIRFT Static Bending Results with Calculated PNNL 11 Moment Curvature (Flexural Rigidity) Derived from Cladding-Only Stress-12 Strain Curve (Reproduction of Figure 22 from NUREG/CR-7198, Revision 1 13 (NRC, 2017a)). S1, S2, S3, and S4 Represent the Experimental Results for 14 HBR HBU SNF As-Irradiated Specimens, HR2 Represents the Experimental 15 Results for HBR HBU SNF Hydride-Reoriented Specimen, and PNNL 16 Represents the Results Calculated Using the Validated Cladding-Only 17 Mechanical Property Models Developed by PNNL (From Geelhood et al.,

18 2008) 19 20 The CIRFT static test results plotted in Figure 2-7 show the moment-curvature response of the 21 four HBR HBU SNF as-irradiated specimens S1, S2, S3, and S4 and the hydride-reoriented 22 specimen HR2. The loading portion of the moment-curvature response begins at 0 N*m and 23 reaches a maximum at about 80 N*m, at which point the specimens begin to unload. The 24 moment-curvature responses of the four HBR HBU SNF as-irradiated specimens during loading 25 were similar up to a moment of 35 N*m. They are characterized by two distinct linear 26 responses, EI1 and EI2, followed by a nonlinear response during the loading and a linear 27 response upon unloading (EI3) (Figure 2-8).

28 Also shown in Figure 2-7 is the cladding-only moment-curvature loading curve constructed 29 using the PNNL cladding-only mechanical property models. The static test results for both as-2-13

For ACRS Review Purposes Only 1 irradiated and hydride-reoriented specimens show much higher bending moment resistance 2 during loading compared to the PNNL cladding-only data. The slopes, EI1 and EI2, of the four 3 HBU fuel rods are greater than the slope of the PNNL data for the cladding-only rod.

4 5 Figure 2-8 Characteristic Points on Moment-Curvature Curve. A, B, C, and D are 6 Points on the Curve. EI1 is the Slope of the Loading Curve Between 0 and 7 A. EI2 is the Slope of the Loading Curve Between A and B. EI3 is the Slope 8 of the Unloading Curve Between D and 0. The Cladding-Only Moment-9 Curvature Loading Curve Constructed Using the PNNL Cladding-Only 10 Mechanical Property Models is not Shown (Reproduction of Figure 21 from 11 NUREG/CR-7198, Revision 1 (NRC, 2017a))

12 13 Figure 2-7 also shows that at bending moments during loading less than 35 N*m, the flexural 14 rigidities of the four as-irradiated rods, which have only circumferential hydrides, and HR2, 15 which has both circumferential and radial hydrides, are essentially the same. This result 16 supports the pretest expectation that, because the bending tensile stress in the cladding is 17 parallel to the plane of both the radial and circumferential hydrides, the presence of radial 18 hydrides would not significantly alter the flexural response from the case where only 19 circumferential hydrides are present. The results of tests currently being conducted by the U.S.

20 Department of Energy (DOE) will further confirm this hypothesis as it applies to other cladding 21 types.

22 23 In the CIRFT static test results for HBR HBU SNF rods shown in Figure 2-7, no failures 24 occurred. The lower-bound maximum moment achieved in the tests is approximately 80 N*m.

25 In addition, it is important to point out that a bending moment of 80 N*m is significantly greater 26 than the bending moment an HBR HBU SNF rod will experience during an HAC 9-m (30-ft) side 27 drop (see Section 2.3.5.1). This means that fuel rod integrity is expected to be maintained 28 during an HAC drop scenarios, and therefore, fuel rod reconfiguration is very unlikely.

2-14

For ACRS Review Purposes Only 1 For the as-irradiated HBR HBU SNF rods, Table 2-1 shows that in the EI1 region of the 2 moment-curvature results, the average flexural rigidity is 2.66 (i.e., 71.58 N*m 2/26.93 N*m 2) 3 times greater than the cladding-only case, and in the EI2 region the average flexural rigidity is 4 2.16 (i.e., 58.10 N*m 2/26.93 N*m 2) times greater than the cladding-only case. For the 5 hydride-reoriented fuel rod, HR2, Table 2-1 shows that in the EI1 region, the average flexural 6 rigidity is 2.33 (i.e., 62.77 N*m2 / 26.93 N*m2) times greater than the cladding-only case, and in 7 the EI2 region, the average flexural rigidity is 1.54 (i.e., 41.52 N*m2 / 26.93 N*m2) times greater 8 than the cladding-only case.

9 Table 2-2 Comparison of Average Flexural Rigidity Results Between CIRFT Static 10 Testing and PNNL Cladding-Only Data (From Validated Mechanical 11 Property Models in Geelhood et al., 2008)

EI1/ EI2/

Test Specimen EI1 (N*m2) EI2 (N*m2) EI3 (N*m2) EIcladding EIcladding As-Irradiated 71.576 58.099 48.133 2.66 2.16 (S1, S2, S3, and S4)

Hydride-Reoriented 62.769 41.517 43.333 2.33 1.54 (HR2)

Cladding-Only (Validated PNNL 26.933 26.933 -

Models) 12 Table 2-3 Characteristic Points and Quantities Based on Moment-Curvature Curves 13 (Reproduction, in Part, of Table 4 from NUREG/CR-7198, Revision 1 14 (NRC, 2017a))

EI1 EI2 EI3 A B C D MA MB MC MD Spec label (Nm2) (Nm2) (Nm2) (m-1) (m-1) (m-1) (m-1) (Nm) (Nm) (Nm) (Nm)

S1 78.655 57.33 51.027 0.202 0.968 2.009 2.166 16.695 60.599 83.595 85.413 S2 73.016 60.848 52.699 0.32 1.009 2.001 2.154 20.18 62.133 85.914 87.294 S3 71.517 59.369 47.101 0.311 0.933 2.149 2.308 22.338 59.288 83.728 85.235 S4 63.117 54.849 41.704 0.503 0.862 2.329 2.507 28.54 48.244 81.656 85.02 As-irradiated 71.576 58.099 48.133 0.334 0.943 2.122 2.284 21.938 57.566 83.723 85.741 Avg.

As-irradiated 6.422 2.603 4.886 0.125 0.062 0.154 0.164 4.977 6.322 1.741 1.048 Std. Dev.

HR2 62.769 41.517 43.333 0.487 1.007 1.585 2.158 30.301 51.884 66.809 79.606 15 In developing a simplified methodology using cladding-only mechanical properties, the staff 16 considers it conservative to use the flexural rigidity ratio from the EI2 data. More specifically, 17 using the average minus two standard deviations of the EI2 data from Table 2-2 is 52.90 N*m2 18 (i.e., 58.10 Nm2 - 2 (2.60 Nm2)), which results in an EI2 ratio of an HBU fuel rod to a 19 cladding-only rod of 1.96 (i.e., 52.90 Nm2 / 26.93 Nm2). The average minus two standard 20 deviations has a 98 percent exceedance probability, which means there is a 98 percent chance 21 that the actual value of the EI ratio will be greater than 1.96. To account for the effects of 2-15

For ACRS Review Purposes Only 1 hydride reorientation, this result is reduced by 0.713 (i.e., 1.54/2.16), which is the ratio of the 2 reoriented hydride results to the as-irradiated results that were calculated in the previous 3 paragraph. Multiplying 1.96 by 0.713 results in a factor of 1.40. However, recognizing the 4 limited test data available to calculating the 1.40 factor, the factor has been further reduced to 5 1.25 to account for the additional uncertainty associated with using limited data. Thus, for the 6 purpose of calculating lateral displacements in the simplified methodology, the flexural rigidity of 7 the HBU fuel rod is equal to the flexural rigidity of the cladding-only rod multiplied by the factor 8 1.25:

9 (EI)HBU rod = 1.25 (EI)clad only (Eq. 2-4) 10 The curvature, , of the HBU fuel rod is given by:

11 = M/(EI)HBU rod (Eq. 2-5) 12 or:

13 = M/[1.25 * (EI)clad only] (Eq.2-6) 14 where M is the bending moment in the rod.

15 The tensile strain is given by:

16 =

• ymax (Eq. 2-7) 17 where ymax is equal to the outer radius, r, of the rod, and the maximum tensile stress is given by:

18 =

• Ec (Eq. 2-8) 19 The methodology described above for using cladding-only properties to calculate cladding 20 strains while accounting for the increased flexural rigidity imparted by the fuel pellet can also be 21 applied to cladding alloys other than Zircaloy-4. Once CIRFT static bending results for other 22 HBU SNF rods (i.e., ZIRLO' (ZIRLO)-clad and M5 (M5)-clad rods) are obtained under 23 planned DOE-sponsored research (Hanson et al., 2016), this methodology can be replicated to 24 obtain a numerical factor that allows for crediting the flexural rigidity of the fuel pellet in those 25 fuel types. Until those results are available, the staff considers the use of cladding-only 26 mechanical properties to calculate cladding stress and strain to be conservative. The staff 27 expects that CIRFT static bending results for other HBU SNF rods obtained by the DOE-28 sponsored research will confirm this conclusion.

29 2.3.4.1 Two Alternatives for Calculating Cladding Stress and Strain During Drop 30 Accidents 31 Two alternatives can be used to calculate cladding stress and strain, and cladding flexural 32 rigidity, for the evaluation of drop accident scenarios. The first alternative is to use cladding-33 only mechanical properties from as-irradiated cladding (which has only circumferential hydrides) 34 or from hydride-reoriented cladding (which would account for radial hydrides precipitated after 35 the drying process). As discussed in Section 2.3.3, the staff considers that the orientation of the 36 hydrides is not a critical consideration when evaluating the adequacy of cladding-only 37 mechanical properties. The properties necessary to implement this alternative are derived from 38 cladding-only uniaxial tensile tests and include modulus of elasticity, yield stress, ultimate 2-16

For ACRS Review Purposes Only 1 tensile strength and uniform strain, and the strain at failure (i.e., the elongation strain).

2 Additional considerations for acceptable cladding-only mechanical properties (i.e., alloy type, 3 burnup, and temperature) may be found in either of the current standard review plans (SRPs) 4 for dry storage of SNF (NUREG-1536, Revision 1, Standard Review Plan for Spent Fuel Dry 5 Storage Systems at a General License Facility, issued in July 2010 (NRC, 2010) for the review 6 of applications for Certificates of Compliance under 10 CFR Part 72; and NUREG-1567, 7 Standard Review Plan for Spent Fuel Storage Facilities, issued in March 2000 (NRC, 2000a) 8 for the review of applications for specific licenses under 10 CFR Part 72) or transportation 9 (NUREG-1617, Standard Review Plan for Transportation Packages for Spent Nuclear Fuel, 10 issued in March 2000 (NRC, 2000b)) - hereafter these documents will be referred to as the 11 current SRPs for dry storage or transportation for SNF.

12 The second alternative is to use cladding-only mechanical properties that have been modified 13 by a numerical factor to account for the increased flexural rigidity imparted by the fuel pellet.

14 This numerical factor is obtained from static CIRFT static bending results for fully-fueled rods for 15 the particular HBU SNF cladding type and fuel type, as previously discussed. This second 16 alternative would be necessary only if the structural evaluation using cladding-only mechanical 17 properties is unsatisfactory, although an applicant may choose to implement this alternative 18 even if the first alternative were to yield satisfactory results. The acceptance criteria for cladding 19 performance following dry storage and transport-related drop accident scenarios can be found 20 in the current SRPs for dry storage and transportation of SNF, respectively.

21 2.3.5 Applicability to Dry Storage and Transportation 22 As discussed in Section 1.5.3, the end-of-life rod internal pressures in both standard and IFBA 23 rods result in cladding hoop stresses below the 90 MPa (1.3 x 104 psia) level that has been 24 shown to be capable of producing significant hydride reorientation in HBU SNF rod cladding.

25 However, the staff chose a highly conservative testing approach (radial hydride treatment under 26 a pressure of 140 MPa (2.0 x 104 psia) to maximize the fraction of cladding radial hydrides 27 precipitated in the test specimens. The approach was designed to produce specimens that, 28 when tested, would provide the most limiting mechanical response and therefore would be 29 reasonably bounding for assessing the mechanical performance of modern HBU SNF.

30 During the radial hydride treatment, each test specimen was pressurized to induce a maximum 31 hoop stress of 140 MPa (2.0 x 104 psia) at a target temperature of 400 °C (752 °F) for 3 hours3.472222e-5 days <br />8.333333e-4 hours <br />4.960317e-6 weeks <br />1.1415e-6 months <br />, 32 cooled at 1 °C/min to 170 °C (under conditions of decreasing pressure and hoop stress), and 33 then heated at 1 °C/min to the hold temperature of 400 °C (752 °F) (under conditions of 34 increasing pressure and hoop stress). This thermal cycling was repeated for five cycles2 to 35 further induce a higher fraction of radial hydrides. The specimen was then furnace-cooled from 36 170 °C (338 °F) to room temperature after the last cycle and the pressure was released.

37 Argonne National Laboratory defined the radial hydride continuity factor (RHCF) as the ratio of 38 the maximum length of continuous radial-circumferential hydrides projected in the radial 39 direction to the cladding thickness within a 150-m arc length (see Section 1.5.4). This metric 40 can be used to quantify the degree of reorientation induced in the hydride-reoriented specimen 41 that was static-bend tested in the CIRFT instrument (specimen HR2). Figure 2-9 shows a 2 A condition that HBU SNF assemblies would not experience in practice, if drying operations are performed according to the guidance in ISG-11, Revision 3, Cladding Considerations for the Transportation and Storage of Spent Fuel, issued November 2003 (NRC, 2003a)see prior Section 1.2 of this report.

2-17

For ACRS Review Purposes Only 1 metallographic image of the hydride microstructure of test specimen HR1 (used for CIRFT 2 dynamic testing) after the aggressive hydride reorientation procedure used for HBR HBU SNF 3 rod specimens.3 The HR2 specimen underwent the same radial hydride treatment (Figure 2-10) 4 as HR1.

5 The aggressive hydride reorientation treatment used for the preparation of the CIRFT test 6 specimens is evidenced by the high radial hydride fraction observed by metallography following 7 testing. As Figure 2-9 shows, the conservative conditions of the radial hydride treatment 8 induced a RHCF exceeding 50% in part of the cladding thickness.

9 Figure 2-9 High Magnification Micrograph Showing Radial Hydrides of a HBR HBU 10 SNF Hydride-Reoriented Specimen Tested Under Phase II (Specimen HR1 11 Results Shown; Hydrogen Content 360-400 wppm) (Reproduction of 12 Figure 35a in NUREG/CR-7198, Revision 1 (NRC, 2017a))

3 Section 3.4.1 of NUREG-7198, Revision 1 (NRC, 2017a), presents a more detailed discussion of the radial hydride treatment used for preparation of the Phase II test specimens.

2-18

For ACRS Review Purposes Only 1 Figure 2-10 Representative Conditions Used for Radial Hydride Treatment for 2 Preparation of HBR HBU SNF Hydride-Reoriented Specimens Tested Under 3 Phase II. The HBU SNF Specimen Was Pressurized to (2.0 x 104 psia) at 4 400 °C (752 °F) with Five Thermal Cycles (Reproduction of Figure 14 from 5 NUREG/CR-7198, Revision 1 (NRC, 2017a))

6 The static test results for the hydride-reoriented Zircaloy-4 fuel rod (specimen HR2; Figure 2-7) 7 show minimal difference in the flexural response compared to the as-irradiated rods up to the 8 bending moments pertinent to a 9-m (30-ft) drop accident (i.e., bending moments below 35 N*m 9 - see Section 2.3.4.2 for pertinent calculation). More importantly, the flexural rigidity of the 10 hydride-reoriented specimen is still markedly higher than the calculated cladding-only response 11 according to validated PNNL mechanical property models. The major difference between the 12 response of the hydride-reoriented HR2 specimen and the as-irradiated rods is the slightly lower 13 flexural resistance of HR2 at higher loads. The slightly lower flexural resistance at higher loads 14 may be the result of the higher density of hydrides in HR2 or the greater extent to which 15 debonding occurred between the cladding and pellet away from the pellet-to-pellet crack 16 interface. However, those loads would not be expected during transportation or dry storage 17 operations.

18 The static test results for the hydride-reoriented HR2 and the as-irradiated HBR HBU SNF 19 Zircaloy-4-clad fuel rods support the staffs conclusion that the use of cladding-only mechanical 20 properties is adequate for the structural evaluation of HAC and NCT drop events. Further, the 21 HAC drop events required for transportation packages apply inertia loads to the fuel rods that 22 bound the design basis storage drops (e.g., drops during transfer operations and non-23 mechanistic tip over). Therefore, this conclusion based on the CIRFT static test results of 24 Zircaloy-4 can be applied to both transportation and storage.

25 The cladding strains that control the static response of an intact fuel rod are the high tensile 26 strains at the face of the crack at the pellet-pellet interface. If a pinhole or hairline crack were to 27 be present at this location, it could have an effect on the static test results because of the strain 28 concentrations they may create. However, the staff considers the probability that a pinhole or 29 hairline crack is at the pellet-pellet crack face simultaneously longitudinally and circumferentially 2-19

For ACRS Review Purposes Only 1 to be low. Therefore, it is reasonable that the CIRFT static test results for intact fuel rods can 2 also be applied to undamaged fuel with pinholes or hairline cracks.

3 The staff expects that a similar mechanical response should be observed by other modern 4 commercial cladding alloy types that may experience hydride reorientation (i.e., Zircaloy-2, 5 ZIRLO and M5) since:

6

• the hydride reorientation treatment used for Zircaloy-4 test specimen preparation was 7 based on highly conservative parameters that would bound operating conditions during 8 dry storage and transportation, which is evidenced by the high RHCF per metallography 9 of the samples. These conditions are:

10 - bounding peak cladding temperature of 400 °C (752 °F) 11 - conservative cladding hoop stresses of 140 MPa (2.0 x 104 psia), well exceeding 12 the maximum cladding hoop stresses for PWR IFBA rods of 90 MPa (1.3 x 104 13 psia) - see Section 1.5.3, and 14 - conservative thermal transients equivalent to five reflooding cycles during loading 15 operations.

16

• the rod-average burnup of the tested hydride-reoriented Zircaloy-4-clad HBU SNF 17 specimens is conservative per the HBU SNF irradiated in commercial reactors in the 18 United States, and 19

• the average hydrogen content of the tested hydride-reoriented Zircaloy-4-clad HBU SNF 20 specimens is bounding to other M5-clad HBU SNF irradiated in commercial reactors in 21 the United States, and conservative to the average hydrogen content of other Zircaloy-2, 22 Zircaloy-4 and ZIRLO-clad HBU SNF irradiated in commercial reactors in the United 23 States 24 The staffs expectation is that future DOE-sponsored CIRFT static testing conducted on other 25 cladding alloy types, beyond that already obtained (see Wang et al., 2016 for additional CIRFT 26 data as obtained under DOE-sponsorship) will provide confirmation of this conclusion (Hanson 27 et al., 2016).

28 2.3.5.1 Use of Static Test Results to Evaluate Safety Margins in an HAC Side Drop 29 Event 30 The CIRFT static test results can be used to determine a lower bound safety margin against fuel 31 rod failure during an HAC side drop event. The safety margin is calculated by dividing the load 32 (or moment) at rod failure by the maximum applied load (or moment) occurring during the side 33 drop event.

34 Figure 2-7 shows that static testing of the HBR HBU SNF rods did not result in rod failures. The 35 lower bound maximum moment achieved in the tests is approximately 80 Nm. Based on the 36 slope of the curves at 80 Nm, it is reasonable to assume that rod failure probably occurs at a 37 moment at or below 100 Nm. Therefore, using 80 Nm provides a conservative basis for 38 calculating safety margin. To quantify the safety margin it is necessary to know the bending 39 moment in the fuel rod as a function of the g-load acting on the rod due to a side drop event.

40 Each fuel rod in the fuel assembly is supported by grid spacers at multiple locations along the 2-20

For ACRS Review Purposes Only 1 rod. Therefore, for the purpose of calculating the maximum bending moment, the rod can be 2 idealized as a uniformly loaded continuous beam.

3 Relationship Between Applied G-Load and Bending Moment 4 For the purpose of evaluating a safety margin, two different fuel rods are initially considered.

5 The first is a fuel rod from a PWR 15 x 15 fuel assembly, and the second is an HBR fuel rod 6 that was tested by ORNL in the CIRFT testing device and reported in NUREG/CR-7198, 7 Revision 1 (NRC, 2017a).

8 The properties of the PWR 15 x 15 fuel bundle (Table 2-3) are taken from NUREG-1864, A 9 Pilot Probabilistic Risk Assessment of a Dry Cask Storage System at a Nuclear Power Plant, 10 Appendix C, Table C.1, issued March 2007 (NRC, 2007a).

11 Table 2-4 PWR 15 x 15 SNF Assembly Parameters Total fuel rod weight 7.011 lb Fuel length 154 in.

Number of grid spacers 8 Rod length between grid spacers (l) 20.5 in.

Uniform applied load (w = 7.011 lb / 154 in.) 0.0455 lb/in.

12 The maximum moment in a uniformly-loaded continuous beam can be approximated by the 13 maximum moment in a uniformly loaded three-span continuous beam as shown in Eqn. 2-9:

14 Mmax = 0.100

• w

• l2 (Eqn. 2-9) 15 i.e., Mmax = (0.100)(0.0455 lb/in.)(20.5 in.)2 = 1.91 lb*in. = 0.216 N*m 16 This is the moment resulting from a 1 g-loading. The g-load necessary to produce a moment of 17 1 N*m = 1 g / 0.216 N*m = 4.63 g / N*m.

18 For the HBR HBU SNF rod, the weight per unit length is calculated from the weight density of 19 fuel and the weight density of cladding, which can be determined from the information in 20 NUREG-1864, Table C.1 (NRC, 2007a) for a BWR 7 x 7 fuel rod.

21 Fuel density = 0.34 lb / in.3 22 (i.e., 9.60 lb / [()(0.25)2(144)] = 0.34) 23 Cladding density = 0.234 lb / in.3 24 (i.e., 1.98 / [()(0.535)(0.035)(144)] = 0.234) 25 The diameter (outer, inner) and thickness of the cladding of an HBR HBU SNF rod as given in 26 NUREG/CR-7198, Revision 1 (NRC, 2017a) are:

27 Outer diameter = 10.743 mm = 0.423 in.

2-21

For ACRS Review Purposes Only 1 Cladding thickness = 0.748 mm = 0.0294 in.

2 Inner diameter = 0.364 in.

3 From the HBR HBU SNF rod cross-sectional dimensions and the fuel and cladding densities 4 calculated using the data for the BWR 7 x 7 fuel rods, the fuel and cladding weight per unit 5 length can be calculated as follows:

6 HBR fuel weight = 0.0354 lb / in.

7 HBR cladding weight = 0.0085 lb / in.

8 w = 0.0354 + 0.0085 = 0.0439 lb / in.

9 l = distance between HBR SNF assembly grid spacers = 26.2 in.

10 Mmax = (0.100)(0.0439)(26.2)2 = 3.01 lb*in = 0.340 N*m 11 This is the moment resulting from a 1 g-loading. The g-load necessary to produce a moment of 12 1 N*m = 1 g / 0.340 N*m = 2.94 g / N*m.

13 This example illustrates the fact that the static transverse g-load necessary to produce a 14 bending moment of 1 N*m in a fuel rod supported by multiple grid spacers varies from rod to 15 rod. For the two rods in this example, the static transverse g-load required to produce a 16 bending moment of 1 N*m varied from 2.9 to 4.6 g depending on the rod cross sectional 17 dimensions and assembly geometry.

18 2.3.5.2 Dynamic Response of a Fuel Rod 19 During a HAC 9-m (30-ft) side drop of a transportation package with impact limiters, the cask 20 body will typically experience inertia loads on the order of 50 g. However, the fuel rod is flexible, 21 as are the intervening components that support the rod between the cask body and the rod.

22 Therefore, the rigid body deceleration of the cask body will be amplified during a side drop event 23 by the flexibility of the rod and intervening components, resulting in a g-load in the rod that is 24 higher than the g-load acting on the cask body. This increase in g-load is expressed by a 25 dynamic load factor (DLF), which is the ratio of the deflection due to a dynamically applied load 26 to the deflection that would have resulted from the static application of the load. The DLF will 27 depend on the rod's natural frequency, the duration of the loading, and the shape of the load 28 time history.

29 Since natural frequency, load duration and load time history shape all depend on the physical 30 characteristics of the fuel assembly, the rod and the cask, including impact limiters, a 31 conservative approach will be used to calculate safety margin by using a maximum DLF of 2.0 32 (Biggs, 1964).

33 Thus, the statically equivalent g-load the fuel rod is subjected to is:

34 (DLF) * (50 g) = 2.0 * (50 g) = 100 g 35 which produces a bending moment in the rod of:

36 100 g / (2.94 g/N*m) = 34.0 N*m 2-22

For ACRS Review Purposes Only 1 The safety margin (SM) against fuel rod bending failure during a side drop event (upon 2 assuming the lower-bound maximum bending moment achieved in the CIRFT static bending 3 tests discussed in Section 2.3.4) is then:

4 SM = (80 N*m)/(34.0 N*m) = 2.35 5 2.3.5.3 Seismic Response of a Fuel Rod 6 The seismic response of a fuel rod can be determined using a variety of structural models.

7 These range from simple idealized models, for which hand calculation methods could be used, 8 to very detailed finite element models. The seismic loads can be applied to these models using 9 either the response spectrum method or a time history analysis method. However, regardless 10 of whether the fuel rod is in a DSS or transportation package, seismic loads will not dominate 11 fuel rod response, because the g-loads produced by a seismic event are not large enough. In 12 storage the g-loads applied to the fuel are dominated by the non-mechanistic tipover event and 13 in a transportation package the g-loads applied to the fuel rod are dominated by the HAC. Both 14 of these events produce g-loads on the fuel rod that are approximately an order of magnitude 15 larger that the g-loads produced by a seismic event. In addition, these two events do not occur 16 coincidently with a seismic event and therefore the seismic event does not add to either of these 17 two events.

18 2.3.5.4 Thermal Cycling during Loading Operations 19 The staff recognizes that the thermal cycling criterion in ISG-11, Revision 3 limits the 20 operational options for a licensee if there is a need for reflooding of HBU SNF during loading 21 operations. The results discussed in NUREG-CR/7198, Revision 1 (NRC, 2017a), and 22 evaluated in this technical report, provide reasonable assurance that intact HBU SNF can be 23 subjected to at least one thermal cycle exceeding 65 °C (117°F) (e.g., during reflooding) without 24 compromising the safety analyses for design-basis drop accidents of a transportation package 25 or dry storage system. The staffs conclusion applies to HBU SNF with cladding demonstrated 26 to be free of hairline cracks and pinholes, as well as other larger defects (i.e., this conclusion 27 applies to HBU SNF with cladding material in a condition equivalent to that tested under the 28 NRC-sponsored program as discussed in NUREG-CR/7198, Revision 1 (NRC, 2017a)). An 29 applicant may provide a justification, on a case-by-case basis, for the effects of reflooding on 30 potential oxidation of the fuel pellet during reflooding operations if the cladding is not 31 demonstrated to be intact (e.g., undamaged cladding with hairline cracks and pinholes).

32 2.4 Application of Fatigue Test Results 33 2.4.1 Lower Bound Fatigue S-N Curves 34 Fatigue strength data are commonly presented in the form of an S-N curve, where S is a 35 strength parameter, such as stress or strain, and N denotes the number of cycles to failure at a 36 specific value of the strength parameter. The objective of this section is to develop a lower-37 bound fatigue S-N curve that envelopes the HBR HBU Zircaloy-4 fuel rod fatigue data and 38 includes both as-irradiated rods and rods with reoriented hydrides. The lower-bound curve 39 serves as an example that applicants may replicate for HBU SNF with other cladding alloys.

40 Table 2-4 presents the fatigue test data for the HBR HBU fuel rods. In Figure 2-11, half of the 41 cladding strain range (/2, which is in Table 2-4) and the maximum strain (//max) are plotted 42 against the number of cycles required to produce cladding failure at a particular strain 2-23

For ACRS Review Purposes Only 1 amplitude. The strain range is the average of the strains caused by positive and negative 2 bending moments, which produce different values of curvature and hence strain. The maximum 3 strain is the maximum of these two strains.

4 Table 2-5 Summary of CIRFT Dynamic Test Results for As-Irradiated and Hydride-5 Reoriented HBR HBU SNF (Reproduction of Table 6 in NUREG/CR-7198, 6 Revision 1 (NRC, 2017a))

Load Moment Spec amp. amp. Number a llmax a a llmax label Seg. ID (N) Nm of cycles Failure (m-1) m1) (MPa) (percent) (percent)

D0 605D1F 250 24.068 2.50E+04 Yes 0.439 0.444 206.109 0.236 0.239 D1 607C4B 150 14.107 1.10E+05 Yes 0.215 0.24 117.26 0.117 0.13 D2 608C4B 50 4.207 6.40E+06 No 0.046 0.067 35.496 0.025 0.036 D3 605C10A 100 9.17 1.00E+06 Yes 0.125 0.171 77.938 0.067 0.092 D4 605D1C 75 6.726 1.10E+07 No 0.089 0.12 57.596 0.048 0.065 D5 605D1B 90 8.201 2.30E+06 Yes 0.114 0.123 69.706 0.061 0.066 D6 609C4 125 11.624 2.50E+05 Yes 0.205 0.218 99.546 0.11 0.117 D7 609C3 200 18.923 6.50E+04 Yes 0.351 0.37 160.835 0.189 0.199 D8 606C3E 87.5 7.743 1.28E+07 No 0.107 0.118 66.309 0.057 0.063 D9 609C7 350 33.667 7.10E+03 Yes 0.576 0.624 288.308 0.31 0.335 D10 606C3A 125 11.552 1.80E+05 Yes 0.174 0.213 98.185 0.094 0.115 D11 607C4A 300 29.021 5.50E+03 Yes 0.469 0.564 241.223 0.254 0.306 D12 608C4A 110 9.986 3.86E+05 Yes 0.144 0.171 83.617 0.078 0.092 D13 606B3E 135 12.551 1.29E+05 Yes 0.151 0.199 106.677 0.081 0.107 D14 606B3D 87.5 7.842 2.74E+05 Yes 0.112 0.135 66.652 0.06 0.073 D15 606B3C 75 6.639 2.24E+07 No 0.087 0.125 56.426 0.047 0.067 HR1 607D4C 150 15.152 4.19E+04 Yes 0.424 0.433 128.788 0.228 0.233 HR3 608D4A 100 8.982 2.44E+05 Yes 0.219 0.233 76.342 0.118 0.125 HR4 608D4C 160 14.759 5.47E+04 Yes 0.323 0.344 125.449 0.174 0.185 2-24

For ACRS Review Purposes Only 1 Figure 2-11 Plots of Half of the Cladding Strain Range (/2) and the Maximum Strain 2 (//max) as a Function of Number of Cycles to Failure. Markers with Arrows 3 Indicate that the Tests Were Stopped Without Failure. (Reproduction of 4 Figure 31b in NUREG/CR 7198, Revision 1 (NRC, 2017a))

5 The lower bound enveloping S-N curve for the HBR HBU SNF rods is composed of three 6 straight line segments when plotted on a linear-log scale. To account for uncertainty with 7 respect to future test results (including the uncertainty associated with higher test 8 temperatures), the equivalent strain amplitude of all segments has been reduced by a factor of 9 0.9. The 0.9 is justified to account for uncertainty with respect to future test results. Each 10 segment's beginning and end point labels from Table 2-4 coordinates (equivalent strain 11 amplitude percent, number of cycles to failure) are given in Table 2-5 and plotted in Figure 2-12.

12 Table 2-6 Coordinates for Lower-Bound Enveloping S-N Curve for the HBR HBU SNF 13 Rods (Equivalent Strain Amplitude Percent, Number of Cycles to Failure)

Segment Beginning Point End Point 1 (D11 to D13) (0.275, 5.50E+3) (0.096, 1.29E+5) 2 (D13 to D14) (0.096, 1.29E+5) (0.066, 2.74E+5) 3 (D14 to D15) (0.066, 2.74E+5) (0.060, 2.24E+7) 2-25

For ACRS Review Purposes Only 1 Figure 2-12 CIRFT Dymanic (Fatigue) Test Results for As-Irradiated and Hydride-2 Reoriented H.B. Robinson Zircaloy-4 HBU Fuel Rods. The Calculated 3 Lower-Bound Fatigue Endurance Curve is also Shown 4 The fatigue data plotted in Figure 2-11 show that at the same number of cycles all of the 5 Zircaloy-4 fuel rods with reoriented hydrides failed at nearly the same strains as the as-6 irradiated Zircaloy-4 fuel rods. Rod specimen D2, which did not fail, was tested at a very low 7 moment amplitude resulting in a very low maximum strain amplitude. The test was also 8 terminated prematurely at 6.4 x 106 cycles. Based on the results for the other test specimens 9 that did not fail, it would be expected that specimen D2 would not have failed until 1 x 108 cycles 10 or beyond. Therefore, rod specimen D2 is not included in the development of the lower bound 11 curve since it would have inappropriately skewed the results. Therefore, the staff considers that 12 a lower-bound fatigue curve developed from as-irradiated data for other cladding alloys is 13 adequate for assessing the fatigue life of alloys with reoriented hydrides.

14 With respect to a fatigue endurance limit for irradiated zirconium alloy, it should be pointed out 15 that some materials, like steel, have a fatigue endurance limit. However, other materials, like 16 aluminum, do not have a fatigue endurance limit. At the present time, there is not sufficient test 17 data to determine whether the various irradiated zirconium-alloys used in HBU SNF (i.e.,

18 Zircaloy-2, Zircaloy-4, ZIRLO', M5) have a fatigue endurance limit.

19 Fatigue data for reoriented cladding alloys other than Zircaloy-4 (e.g., Zircaloy-2, ZIRLO',

20 M5) may not yet be available - see Wang et al., 2016 for additional CIRFT data as obtained 21 under DOE-sponsorship. However, the staff believes the methodology described above for 22 developing a lower-bound fatigue curve can be used to construct a lower-bound fatigue curve 23 for other cladding alloys once the as-irradiated fatigue data become available. Further, the staff 24 notes that an applicant may be able to demonstrate a generic lower-bound fatigue curve for 25 various modern cladding alloys if an adequate safety margin is incorporated.

2-26

For ACRS Review Purposes Only 1 2.4.2 Fatigue Cumulative Damage Model 2 During NCT if a fuel rod were to vibrate at a constant strain amplitude, all that would be 3 necessary to predict the fatigue life of the rod is the S-N curve. However, fuel rod vibration 4 during NCT is expected to have a series of many cycles encompassing a range of strain 5 amplitudes and with each cycle, damage to the fuel rod cladding is continuously accumulating.

6 A fatigue damage model can be used to express how damage from these cycles accumulates.

7 To date, more than 50 fatigue damage models have been proposed, but unfortunately none of 8 these models enjoys universal acceptance, and the applicability of each model varies from case 9 to case. Unlike the aerospace industry, which has conducted extensive research on the 10 accumulation of fatigue damage to materials, such as steel, aluminum, and titanium, no 11 research has been conducted on fatigue damage to HBU spent fuel cladding. Nevertheless, for 12 many metals, the simple linear damage rule developed by Miner (Gaylord and Gaylord, 1979) 13 appears to provide a simple and reasonably reliable prediction of fatigue behavior under random 14 loadings, and therefore, will be used to evaluate fatigue damage accumulation in HBU SNF rods 15 during NCT.

16 For failure, the linear damage rule is, the following:

17 i ni/Ni = n1/N1 + n2/N2 + n3/N3 + ... = 1 (Eqn 2-9) 18 where:

19 ni = number of strain cycles at strain level i 20 Ni = number of strain cycles to produce failure at i.

21 To apply this simple linear damage rule it is assumed that the NCT loading history can be 22 reduced to a series of different strain levels where the number of cycles associated with each 23 strain level, i, is, n. To account for uncertainty in using a simple linear damage rule to describe 24 the accumulated fatigue damage in HBU fuel, the right side of the above equation should be set 25 equal to 0.7. This value is considered an approximate lower bound for the uncertainty in Miners 26 damage model (Hashin, 1979).

27 2.4.3 Applicability to Storage and Transportation 28 The CIRFT fatigue tests were conducted under conditions that produced a uniform bending 29 moment in the fuel rod. Thus, these results apply only to loading conditions that produce 30 longitudinal bending stresses in the cladding of the fuel. Such loading conditions occur when 31 fuel rods vibrate during NCT. Fluctuating loads can also occur during storage when the 32 cladding experiences thermal cycles because of daily and seasonal fluctuations in ambient 33 temperature. These thermal cycles will induce cyclic stresses on the cladding due to changes in 34 fission and decay gas pressure, which will result in fluctuations in cladding hoop stresses. As 35 explained above, however, the fatigue test results apply only to loading conditions that produce 36 longitudinal bending stresses in the cladding of the fuel. The fatigue test results are not 37 applicable to loading conditions that produce fluctuations in hoop stress. Therefore, the fatigue 38 test results cannot be applied to thermal fatigue during dry storage (see NUREG-2214 (NRC, 39 2019) for discussion of thermal fatigue of SNF cladding during dry storage).

40 In the CIRFT static and fatigue tests the fuel rods were subjected to a constant bending moment 41 which resulted in a longitudinal bending stress in the cladding. However, in an actual spent fuel 2-27

For ACRS Review Purposes Only 1 rod there is internal gas pressure, which creates hoop stresses on the order of 90 MPa 2 (1.3 x 104 psia) or less - see Section 1.5.3. The presence of the hoop stresses creates a non-3 proportional biaxial stress state in the cladding. The stress state is non-proportional because 4 the hoop stress remains constant while the longitudinal bending stress fluctuates. Recent 5 research on the effect of proportional biaxial stress fields on fatigue crack growth shows no 6 significant effect of the biaxial stress field on fatigue crack propagation behavior (Pickard, 2015).

7 It is expected that the same result would also hold for non-proportional biaxial stress fields.

8 Based on these results, the staff considers that the presence of a biaxial stress field in a spent 9 fuel rod does not need to be considered Therefore, only the longitudinal bending stresses in the 10 cladding need to be considered when using the ORNL static and fatigue test data.

11 2.4.3.1 Seismic Events 12 During storage or transportation, it is possible that a seismic event could occur. Typically, the 13 strong motion duration of a seismic event is approximately 10 seconds. A fuel rod generally 14 responds to seismic input in the 10 to 30 hertz (Hz) frequency range. This means that the 15 number of fatigue cycles associated with a seismic event would be no more than about 300 16 cycles (10 seconds x 30 Hz = 300 cycles). In addition, it is expected that the seismic load 17 applied to the rod would be less than 10-g. Based on the results summarized at the end of 18 Section 2.3.4.1, a 10-g load would produce a bending moment in the rod of about 3.5 N*m.

19 From Table 2-4, a bending moment of 3.5 N*m would result in a maximum cladding strain of 20 about 0.03%. From an event that produced 300 bending cycles at a maximum strain of 21 0.03%, Figures 2-11 and 2-12 show that virtually no fatigue damage would be expected. For 22 example, extrapolating the lower bound curve in Figure 2-12 to 300 cycles shows that it would 23 require a strain of more than 0.45% to cause a fatigue failure. This is 15 times greater than 24 the 0.03% caused by a seismic event. Therefore, seismic events during storage or 25 transportation are not expected to compromise the fuel integrity.

26 2.4.3.2 Thermal Cycling during Loading Operations 27 The staff recognizes that the thermal cycling criterion in ISG-11, Revision 3 limits the 28 operational options for a licensee if there is a need for reflooding of HBU SNF during loading 29 operations. The results discussed in NUREG-CR/7198, Revision 1 (NRC, 2017a), and 30 evaluated in this technical report, provide reasonable assurance that intact HBU SNF can be 31 subjected to at least one thermal cycle exceeding 65 °C (117°F) (e.g., during reflooding) without 32 compromising the lower-bound curve for the evaluation of HBU SNF rod fatigue in a 33 transportation package. The staffs conclusion applies to HBU SNF with cladding demonstrated 34 to be free of hairline cracks and pinholes, as well as other larger defects (i.e., this conclusion 35 applies to HBU SNF with cladding material in a condition equivalent to that tested under the 36 NRC-sponsored program as discussed in NUREG-CR/7198, Revision 1 (NRC, 2017a)). An 37 applicant may provide a justification, on a case-by-case basis, for the effects of reflooding on 38 potential oxidation of the fuel pellet during reflooding operations if the cladding is not 39 demonstrated to be intact (e.g., undamaged cladding with hairline cracks and pinholes).

2-28

For ACRS Review Purposes Only 1 3 DRY STORAGE OF HIGH BURNUP SPENT NUCLEAR FUEL 2 3.1 Introduction 3 The U.S. Nuclear Regulatory Commission (NRC) staff (the staff) has developed example 4 licensing and certification approaches for dry storage of high burnup (HBU) spent nuclear fuel 5 (SNF). Applicants may use these approaches to provide reasonable assurance of compliance 6 with Title 10 of the Code of Federal Regulations (10 CFR) Part 72, Licensing Requirements for 7 the Independent Storage of Spent Nuclear Fuel, High-Level Radioactive Waste, and Reactor 8 Related Greater Than Class C Waste, during normal, off-normal and accident conditions of 9 storage. The staff developed these example approaches according to the conclusions of the 10 engineering assessment in Chapter 2. Figure 3-1 provides a high-level diagram of these 11 approaches, which vary based on (1) the condition of the fuel (undamaged or damaged), and 12 (2) the length of time the fuel has been in dry storage. Section 3.2.2 discusses considerations 13 for additional analyses expected for non-leaktight dry storage system (DSS) designs. An 14 applicant may consider and demonstrate other approaches that may be acceptable.

15 As required by 10 CFR 72.24(b) and 10 CFR 72.236(a), an application for a specific license for 16 an independent spent fuel storage installation (ISFSI) or an application for a Certificate of 17 Compliance (CoC) for a DSS design, respectively, should identify the allowable SNF contents 18 and condition of the assembly and rods per the design bases. The allowable cladding condition 19 for the SNF contents is generally defined in the Technical Specifications of the specific license 20 (10 CFR 72.44(c)) or CoC (10 CFR 72.236(a)), and the nomenclature may vary between 21 different DSS designs. For example, the terms intact and undamaged have both been 22 historically used to describe cladding without any known gross cladding breaches. In 23 accordance with 10 CFR 72.212(a)(1) and 10 CFR 72.212(b)(3), users of DSSs (general 24 licensees) are to comply with the Technical Specifications of the CoC by selecting and loading 25 the appropriate fuel, and are to maintain records that reasonably demonstrate that loaded fuel 26 was adequately selected, in accordance with their approved site procedures and Quality 27 Assurance Program.

28 Interim Staff Guidance (ISG)-1, Revision 2, Classifying the Condition of Spent Nuclear Fuel for 29 Interim Storage and Transportation Based on Function, issued in May 2007 (NRC, 2007b),

30 provides guidance for developing the technical basis supporting the conclusion that the SNF 31 (both rods and assembly) to be loaded in a DSS are intact or undamaged. 1 This would include 32 considering whether the material properties, and possibly the configuration, of the SNF 33 assemblies can be altered during the requested dry storage period. If the alteration is 34 significant enough to prevent the fuel or assembly from performing its intended functions, then 35 the fuel assembly should be classified as damaged.

36 Damaged SNF is generally defined in terms of the characteristics needed to perform functions 37 to ensure compliance with fuel-specific and DSS-related regulations. A fuel-specific regulation 38 defines a characteristic or performance requirement of the SNF assembly. Examples of such 39 regulations include 10 CFR 72.122(h)(1) and 10 CFR 72.122(l). A DSS-related regulation 40 defines a performance requirement placed on the fuel so that the DSS can meet its regulatory 41 requirements. Examples of such regulations include 10 CFR 72.122(b) and 10 CFR 72.124(a).

1 The current revisions of all ISG documents will be rolled into revised standard review plans (SRPs) for dry storage and transportation, as appropriate, and will then be removed from the public domain. The revised SRPs will be issued for public comment prior to being finalized.

3-1

For ACRS Review Purposes Only 1 The glossary in this report provides the staffs definitions of intact, undamaged, and damaged 2 fuel. For additional information, refer to the current standard review plans (SRPs) for dry 3 storage of SNF (NUREG-1536, Revision 1, Standard Review Plan for Spent Fuel Dry Storage 4 Systems at a General License Facility, issued in July 2010 (NRC, 2010) for the review of 5 applications for Certificates of Compliance under 10 CFR Part 72, and NUREG-1567, Standard 6 Review Plan for Spent Fuel Storage Facilities, issued in March 2000 (NRC, 2000a) for the 7 review of applications for specific licenses under 10 CFR Part 72) - hereafter, these documents 8 will be referred to as the current SRPs for dry storage SNF.

9 3-2

3-3 For ACRS Review Purposes Only Figure 3-1 Example Licensing and Certification Approaches for Dry Storage of High Burnup Spent Nuclear Fuel

For ACRS Review Purposes Only 1 Consistent with the guidance in (ISG)-1, Revision 2 (NRC, 2007b), HBU SNF assemblies with 2 any of the following characteristics, as identified during the fuel selection process, are generally 3 classified as damaged unless an adequate justification is provided for not doing so:

4

• There is visible deformation of the rods in the HBU SNF assembly. This does not refer 5 to the uniform bowing that occurs in the reactor; instead, this refers to bowing that 6 significantly opens up the lattice spacing.

7

• Individual fuel rods are missing from the assembly. The assembly may be classified as 8 intact or undamaged if the missing rod(s) do not adversely affect the structural 9 performance of the assembly, or radiological and criticality safety (e.g., there are no 10 significant changes to rod pitch). Alternatively, the assembly may be classified as intact 11 or undamaged if a dummy rod that displaces a volume equal to, or greater than, the 12 original fuel rod is placed in the empty rod location.

13

• The HBU SNF assembly has missing, displaced, or damaged structural components 14 such that either:

15 - Radiological and/or criticality safety is adversely affected (e.g., significant change 16 in rod pitch),

17 - The structural performance of the assembly may be compromised during normal, 18 off-normal, and accident conditions of storage, or 19 - The assembly cannot be handled by normal means (i.e., crane and grapple), if 20 the design bases relies on ready retrieval of individual fuel assemblies.

21

• Reactor operating records or fuel classification records indicate that the HBU SNF 22 assembly contains fuel rods with gross rupture.

23

• The HBU SNF assembly is no longer in the form of an intact fuel bundle (e.g., consists 24 of, or contains, debris such as loose fuel pellets or rod segments).

25 Defects such as dents in rods, bent or missing structural members, small cracks in structural 26 members, and missing rods do not necessarily render an assembly as damaged, if the intended 27 functions of the assembly are maintained; i.e., the performance of the assembly does not 28 compromise the ability to meet fuel-specific and DSS-related regulations.

29 3.2 Uncanned Fuel (Intact and Undamaged Fuel) 30 Undamaged HBU SNF can be stored in the DSS without the need for a separate fuel can (i.e., a 31 separate metal enclosure sized to confine damaged fuel particulates) to maintain a known 32 configuration inside the DSS confinement cavity. This fuel includes rods that are either intact 33 (i.e., no breaches of any kind) or that contain small cladding defects (i.e., pinholes or hairline 34 cracks) that may permit the release of gas from the interior of the fuel rod. Cladding with gross 35 ruptures that may permit the release of fuel particulates may not be considered undamaged. The 36 configuration of undamaged HBU SNF may be demonstrated to be maintained if loading and 37 transport operations are designed to prevent and/or mitigate degradation of the cladding and 38 other assembly components, as discussed in ISG-22, Potential Rod Splitting Due to Exposure to 39 an Oxidizing Atmosphere during Short-Term Cask Loading Operations in LWR or Other Uranium 40 Oxide Based Fuel, issued May 2006 (NRC, 2006).

3-4

For ACRS Review Purposes Only 1 Following the approaches delineated in Figure 3-1, an application for dry storage of undamaged 2 HBU SNF would include a structural evaluation of the fuel rods under design-bases drop 3 accident scenarios. The evaluation serves to demonstrate that the uncanned fuel remains in a 4 known configuration after a drop accident scenario.

5 Two alternatives may be used to calculate cladding stress and strain, and cladding flexural 6 rigidity, for the aforementioned evaluation of drop accident scenarios. The first alternative, 7 shown in Figure 3-2, is to use cladding-only mechanical properties from as-irradiated cladding 8 (i.e., cladding with circumferential hydrides, primarily), or hydride-reoriented cladding (i.e.,

9 cladding that accounts for radial hydrides precipitated after the drying process).

10 Figure 3-2 First Approach for Evaluating Design-Bases Drop Accidents During Dry 11 Storage 12 As discussed in Section 2.3.3, the staff considers the orientation of the hydrides not to be critical 13 when evaluating the adequacy of cladding-only mechanical properties. Therefore, the properties 14 necessary to implement this first alternative may be derived from cladding-only uniaxial tensile 15 tests and include modulus of elasticity, yield stress, ultimate tensile strength and uniform strain, 16 and the strain at failure (i.e., the elongation strain). Refer to the current SRPs for dry storage of 17 SNF for additional considerations for acceptable cladding-only mechanical properties (i.e., alloy 3-5

For ACRS Review Purposes Only 1 type, burnup, and temperature) and the acceptance criteria for cladding performance during dry 2 storage operations.

3 A second alternative, shown in Figure 3.3, is to use cladding-only mechanical properties that 4 have been modified by a numerical factor to account for the increased flexural rigidity imparted 5 by the fuel pellet. This numerical factor can be obtained from static test data from the cyclic 6 integrated reversible-bending fatigue tester (CIRFT) for fully-fueled rods for the particular 7 cladding type and fuel type (see Section 2.3.3). The second alternative would be necessary 8 only if the structural evaluation using cladding-only mechanical properties is unsatisfactory, 9 although an applicant may choose to implement it even if the first alternative were to yield 10 satisfactory results. Refer to the current SRP for dry storage of SNF for acceptance criteria on 11 cladding performance during dry storage operations.

12 Figure 3-3 Second Approach for Evaluation of Design-Bases Drop Accidents During 13 Dry Storage 14 3.2.1 Leaktight Confinement 15 Consistent with the guidance in the current SRPs for dry storage of SNF, an application for a 16 DSS for HBU SNF is expected to define the maximum allowable leakage rate for the entire 17 confinement boundary. The maximum allowable leakage rate is based on the quantity of 3-6

For ACRS Review Purposes Only 1 radionuclides available for release and is evaluated to meet the confinement requirements for 2 maintaining an inert atmosphere within the DSS confinement cavity and compliance with the 3 regulatory limits of 10 CFR 72.104, Criteria for Radioactive Materials in Effluents and Direct 4 Radiation from an ISFSI or MRS, and 10 CFR 72.106, Controlled Area of an ISFSI or MRS.

5 Leakage rate testing is performed on the entire confinement boundary (over the course of 6 fabrication and loading) and ensures that the package can maintain a leak rate below the 7 maximum allowable leakage rate per ANSI N14.5 (2014).

8 If the entire DSS confinement boundary, including its closure lid, is designed and tested to be 9 leaktight as defined in American National Standards Institute (ANSI) N14.5 - 2014, American 10 National Standard for Radioactive MaterialsLeakage Tests on Packages for Shipment and the 11 current SRPs for dry storage of SNF, then the application is not expected to include additional 12 dose calculations based on the allowable leakage rate that demonstrate compliance with the 13 regulatory limits of 10 CFR 72.104(a) and 10 CFR 72.106(b). In addition, the structural analysis 14 of the package is to demonstrate that the confinement boundary will not fail under the postulated 15 drop scenarios and that the confinement boundary will remain leaktight under all conditions of 16 storage. Refer to the current SRPs for dry storage of SNF for additional guidance on 17 demonstrating compliance with the leaktight criterion.

18 3.2.2 Non-Leaktight Confinement 19 For those DSS designs not tested to a leaktight confinement criterion, the application is 20 expected to include dose calculations based on the allowable leakage rate to demonstrate 21 compliance with the regulatory limits of 10 CFR 72.104(a) and 10 CFR 72.106(b). Leakage rate 22 testing is performed on the entire confinement boundary (over the course of fabrication and 23 loading) and ensures that the package can maintain a leak rate below the maximum allowable 24 leakage rate, which can be calculated using the methodology in ANSI N14.5 (2014).

25 To determine the dose rate for the confinement boundary, an application for a non-leaktight 26 DSS is expected to provide a technical basis for the assumed bounding HBU fuel failure rates 27 for normal, off-normal, and accident conditions of storage. If an application is not able to 28 provide and justify its bounding fuel failure rates, then the fuel failure rates below can be 29 assumed as bounding values for normal, off-normal, and accident conditions of storage:

30

• Normal conditions of storage: 1 percent 31

• Off-normal conditions of storage: 10 percent 32

• Accident conditions of storage: 100 percent 33 Bounding Release Fractions for High Burnup Spent Nuclear Fuel 34 HBU SNF fuel has different characteristics than low burnup (LBU) SNF with respect to cladding 35 oxide thickness, hydride content, radionuclide inventory and distribution, heat load, fuel pellet 36 grain size, fuel pellet fragmentation, fuel pellet expansion and fission gas release to the rod 37 plenum [See Appendix C.5 to NUREG/CR-7203, A Quantitative Impact Assessment of 38 Hypothetical Spent Fuel Reconfiguration in Spent Fuel Storage Casks and Transportation 39 Packages, issued September 2015 (NRC, 2015) for additional details on HBU SNF]. These 40 characteristics may affect the mechanisms by which the fuel can breach and the amount of fuel 41 that can be released from failed fuel rods. Hence, the staff evaluated open literature on HBU 42 fuel rod failure rates and release fractions of Chalk River unknown deposits (CRUD), fission 3-7

For ACRS Review Purposes Only 1 gases, volatiles, and fuel fines to assist in the review of applications for non-leaktight 2 confinement boundaries. Table 3-1 provides release fractions that may be considered 3 reasonably bounding for HBU SNF. If these release fractions are not used, other release 4 fractions may be used in the analysis provided the applicant properly justifies the basis for their 5 usage. Justification of the proposed release fractions of the source terms should consider an 6 adequate description of burnup for the test specimen, number of tests, collection method for 7 quantification of release fractions, test specimen pressure at the time of fracture, and source 8 collection system.

9 Table 3-1 Fractions of Radioactive Materials Available for Release from HBU SNF 10 Under Conditions of Dry Storage (for both Pressurized Water Reactor and 11 Boiling Water Reactor Fuels)

Accident-Normal Off-Normal Accident-Fire Impact Variable Conditions Conditions Conditions Conditions Fraction of Fuel Rods 0.01 0.1 1.0 1.0 Assumed to Fail Fraction of Fission Gases Released Due to 0.15 0.15 0.15 0.35 a Cladding Breach Fraction of Volatiles Released Due to a 3 x 10-5 3 x 10-5 3 x 10-5 3 x 10-5 Cladding Breach Mass Fraction of Fuel Released as Fines Due 3 x 10-5 3 x 10-5 3 x 10-3 3 x 10-5 to a Cladding Breach Fraction of CRUD 0.15 0.15 1.0 1.0 Spalling Off Cladding 12 CRUD 13 The average CRUD thickness in HBU SNF cladding has been estimated to be similar to that 14 observed on LBU SNF cladding. A review of data in the literature (NRC, 2000c; Einziger and 15 Beyer, 2007) indicates that a release (spalling off) of 15 percent of cladding CRUD may be 16 assumed as reasonably bounding to both normal and off-normal conditions of storage, and a 17 release of 100 percent of the cladding CRUD is conservatively bounding to both postulated fire 18 and impact accidents during storage (NRC, 2014).

19 Fission Gases 20 The NRCs FRAPCON steady-state fuel performance code has been previously used to assess 21 release fractions of fission gases during transportation (NRC, 2011). The seven most common 22 fuel designs were evaluated using FRAPCONs modified Forsberg-Massih model (8x8, 9x9, 23 and 10x10 fuel for boiling water reactors (BWRs) and 14x14, 15x15, 16x16, and 17x17 for 24 pressurized-water reactors (PWRs). For each fuel design, a number of different power histories 25 aimed at capturing possible realistic reactor irradiations were modeled. The fission gas content 26 within the free volume of the rods was evaluated for a total of 243 different cases (39 for each of 3-8

For ACRS Review Purposes Only 1 the BWR fuel designs; 37 for 14x14 and 16x16 PWR fuel designs, and 26 for 15x15 and 17x17 2 PWR fuel designs). A review of the results indicates that a release of 15 percent of fission 3 gases may be assumed as reasonably bounding to normal conditions of transport scenarios for 4 rod average burnups up to 62.5 GWd/MTU. The same release fraction may be reasonably 5 assumed for both normal and off-normal conditions of storage.

6 During a fire accident scenario in storage, the fuel is not expected to reach temperatures high 7 enough that fission gases can diffuse out of the pellet matrix or grain boundaries to the rod 8 plenum. The thermal rupture tests showed that release occurred at higher temperatures than 9 those experienced during a transportation fire accident (NRC, 2000c). The same behavior is 10 expected during a postulated fire accident condition of storage. Therefore, the same release 11 fraction of 15 percent of fission gases during normal/off-normal conditions of storage may be 12 assumed to be reasonably bounding to the fire scenario under accident conditions of storage.

13 In the case of postulated impact accident (drop) scenarios (e.g., during transfer or retrieval 14 operations), the pellet may be conservatively assumed to crumble. In this scenario, fission 15 gases retained within the pellet grain boundaries may be released in addition to those already 16 released from the fuel rod free volume (i.e., from the fuel-cladding gap and plenum). The 17 FRAPFGR model in FRAPCON may be used to predict the location of the fission gases within 18 the fuel pellet (NRC, 2011). The model has been validated with experimental data obtained 19 using an electron probe micro analyzer. The FRAPFGR model was used to calculate the 20 maximum fraction of the pellet-retained fission gases that may be released during a drop 21 impact, which was determined to be 20 percent. Therefore, assuming all fission gases within 22 the pellet grain boundaries are released, a 35 percent (15 percent + 20 percent) maximum 23 release fraction may be assumed to be reasonably bounding to a postulated accident fire 24 scenario during storage. This value accounts for the 15 percent maximum fission gases 25 released from the fuel rod free volume (as calculated with the modified Forsberg- Massih model) 26 and the 20 percent maximum fission gases released from the fuel pellet grain boundaries (as 27 calculated with the FRAPFGR model). These release fraction estimates are consistent with 28 previous NRC estimates (NRC, 2000c; NRC, 2007; Einziger and Beyer, 2007).

29 Volatiles 30 Most of the volatile release fractions originate from cesium-based compounds in the form of 31 oxides or chlorides (NRC, 2000c; NRC, 2014). These volatiles exhibit a different release 32 behavior in comparison to fission gases. Volatiles tend to migrate and aggregate at the rim on 33 the outer surface of the fuel pellet during reactor irradiation, which is characteristic of burnups 34 near or exceeding 60 GWd/MTU. The pellet rim is characterized by a fine crystalline grain 35 structure (0.1 - 0.3 µm or submicron in characteristic size) (Spino et al., 2003; Einziger and 36 Beyer, 2007), a high porosity that may exceed 25 percent, and a high concentration of actinides 37 relative to the inner pellet matrix.

38 Sandia National Laboratories assessed the maximum release fraction of volatiles (cesium and 39 other ruthenium-based compounds) under drop and fire accident scenarios of transportation, 40 and determined it to be 0.003 percent (3x10-5) (NRC, 2000c). This assessment included 41 modeling and analyses using various data from the literature. The volatile release fraction 42 during a fire accident scenario was determined to be lower than the release fraction during a 43 drop accident scenario (NRC, 2014; NRC, 2000c). Therefore, a volatile release fraction of 44 0.003 percent (3 x 10-5) may be assumed to be reasonably bounding to normal, off-normal, and 45 accident conditions of storage. This release fraction estimate is also consistent with an 46 independent estimate by Einziger and Beyer (2007).

3-9

For ACRS Review Purposes Only 1 Fuel Fines 2 Release fractions from SNF fines during storage and transportation have been previously 3 documented (NRC, 2000c; NRC, 2007; Benke et al., 2012; NRC, 2014). HBU SNF has a 4 different pellet microstructure than LBU SNF, which is characterized by an inner matrix and an 5 outer pellet rim layer. The thickness of the outer pellet rim layer increases with higher fuel 6 burnup. Therefore, differences in microstructure between the inner pellet matrix and the outer 7 pellet rim should be considered when evaluating release fractions of fuel fines from HBU SNF.

8 Although there is no reported literature on HBU SNF rim fracture as a function of impact energy, 9 other data can be used to indirectly assess the contribution of the rim layer to the release 10 fractions of fuel fines. Spino et al (1996) estimated the fracture toughness of the rim layer from 11 micro-indentation tests. Compared to the inner SNF matrix, the rim layer showed an increase of 12 fracture toughness. The increase of fracture toughness implies a decrease of release fraction.

13 Hirose et al (2015) also discussed results of axial dynamic impact tests simulating accident 14 conditions during transport, which are expected to be bounding to postulated drop scenarios 15 during dry storage. The dispersed particles from pellet breakage following impact were 16 collected and correlated to impact energy. The staff has compared the measured release 17 fraction of fuel fines from Hirose et al (2015) with previous NRC estimates of release fraction 18 versus impact energy for SNF and other brittle materials (depleted UO2, glass and Synroc) (see 19 Figure 3 of NUREG 1864, A Pilot Probabilistic Risk Assessment of a Dry Cask Storage System 20 at a Nuclear Power Plant (NRC 2007)). Based on these analyses, the staff concludes that 21 there is no indication that pellet rim layer contributes to increased release fractions.

22 Since the outer HBU fuel pellet rim does not appear to contribute to additional release fractions, 23 previous NRC estimates for release fractions of fuel fines may continue to be used (NRC, 24 2000c; NRC, 2007; Benke, et al., 2012; Ahn et al., 2012; NRC, 2014). Per the range of 25 estimates in the literature, a release fraction for fuel fines of 0.003 percent (3x10-5) may be 26 assumed to be reasonably bounding to normal, off-normal, and accident (drop impact) 27 conditions of storage. During a fire accident scenario, fuel oxidation is conservatively assumed 28 to increase the release fraction of fuel fines by a factor of 100 (NRC, 2000c; Ahn et al 2012).

29 Therefore, a 0.3 percent (3x10-3) release fraction of fuel fines may be assumed as reasonably 30 bounding to fire accident conditions of storage.

31 The staff recognizes that various international cooperative research programs are currently 32 investigating release fractions from HBU SNF. Once those data are available to the public, the 33 staff will review and determine whether the conservative estimates in the above discussion 34 should be revisited.

35 3.2.3 Dry Storage Up To 20 Years 36 Section 1.2 discussed the staffs review guidance for the licensing and certification of dry 37 storage of HBU SNF for a period of up to 20 years. The technical basis referenced in that 38 guidance supports the staffs conclusion that creep is not expected to result in gross rupture if 39 cladding temperatures are maintained below 400 °C (752 °F).

40 Chapter 2 also provided an assessment of the effects of hydride reorientation per static and 41 fatigue bending test results on HBU SNF specimens. Those test results provide a technical 42 basis for the staffs conclusion that the use of cladding mechanical properties (with either as-43 irradiated or hydride- reoriented microstructure) is adequate for the structural evaluation of HBU 44 SNF when evaluating postulated drops during dry storage (e.g., drops during transfer 3-10

For ACRS Review Purposes Only 1 operations, non-mechanistic DSS cask tipover). Refer to the current SRPs for dry storage of 2 SNF for staff review guidance on additional considerations for acceptable cladding-only 3 mechanical properties (i.e., alloy type, burnup, temperature), on acceptable references for 4 cladding mechanical properties and on acceptance criteria for the structural evaluation of the 5 HBU fuel assembly for the drop accident scenarios. As indicated in Figure 3-1, supplemental 6 safety analyses are not expected for HBU SNF in dry storage for periods not exceeding 20 7 years.

8 3.2.4 Dry Storage Beyond 20 Years 9 As indicated in Figure 3-1, to address age-related uncertainties related to the extended dry 10 storage of HBU SNF (i.e., dry storage beyond 20 years), the application is expected to be 11 supplemented with either results from a surrogate demonstration program or supplemental 12 safety analyses assuming justified hypothetical fuel reconfiguration scenarios. The results from 13 a surrogate demonstration program are meant to provide field-obtained confirmation that the 14 fuel has remained in the analyzed configuration after 20 years of dry storage. If confirmation is 15 not provided, the safety analyses for the DSS should be supplemented to assume reconfigured 16 fuel. Consistent with the requirements in 10 CFR Part 72, the supplemental information may be 17 provided in either the initial license or CoC application (per 10 CFR 72.40(a) and 18 10 CFR 72.238, Issuance of an NRC Certificate of Compliance) or in a renewal application 19 (10 CFR 72.42(a) and 10 CFR 72.240(a)).

20 The NRC has approved the licensing and certification of HBU SNF for an initial 20-year-term per 21 the technical basis in the staffs review guidance, as discussed in Section 1.2. However, the 22 staff has recognized that the technical basis is based on short-term accelerated creep testing 23 (i.e., laboratory scale testing up to a few months), which results in increased uncertainties when 24 extrapolated to long periods of dry storage - see Appendix D to NUREG-1927, Revision 1 25 (NRC, 2016b). Although the staff has confidence based on this short-term testing that creep-26 related degradation of the HBU fuel will not adversely affect its analyzed configuration for 27 storage periods beyond 20 years, there is no operational field-obtained data to confirm this 28 expectation, as was done in the prior demonstration on LBU fuel described in NUREG/CR-6745, 29 Dry Cask Storage Characterization ProjectPhase 1; CASTOR V/21 Cask Opening and 30 Examination, issued September 2001 (NRC, 2001),; and NUREG/CR 6831, Examination of 31 Spent PWR Fuel Rods after 15 Years in Dry Storage, issued September 2003 (NRC, 2003b).

32 In addition, the staff also acknowledges that while the CIRFT results obtained to-date (as 33 discussed in Chapter 2) provide an adequate technical basis for assessing the separate effects 34 of hydride reorientation, the results do not account for potential synergistic effects of various 35 physical and chemical phenomena occurring during extended dry storage (e.g., cladding creep, 36 hydride reorientation, irradiation hardening, oxidation, hydriding caused by residual water 37 hydrolysis, etc. - see NUREG-2214, Managing Aging Processes in Storage (MAPS) Report, 38 Final Report issued August 2019 (NRC, 2019) for discussions on these phenomena).

39 Therefore, the staff considers it prudent to gather and review evidence that HBU fuel in dry 40 storage beyond 20 years has maintained its analyzed configuration be gathered and reviewed.

41 3.2.4.1 Supplemental Results from Confirmatory Demonstration 42 A demonstration program, like that conducted for LBU SNF (NRC, 2003; NRC, 2001; NRC, 43 2003b), may be used to confirm the results from separate-effects testing, which has provided 44 the technical bases for dry storage of HBU SNF beyond 20 years.

3-11

For ACRS Review Purposes Only 1 3.2.4.1.1 Initial Licensing or Certification 2 Consistent with 10 CFR 72.42(a) and 10 CFR 72.238, an applicant may request approval for dry 3 storage of HBU SNF for periods up to 40 years. These applications are not required to provide 4 aging management programs (AMPs), as these programs are expected only in renewal 5 applications. Instead, for initial licenses and CoC approvals for dry storage beyond 20 years (up 6 to 40 years), the application may describe the activities to obtain and evaluate confirmatory data 7 from a demonstration program under the aegis of a maintenance plan. The maintenance plan 8 would be implemented after the initial 20 years of dry storage. Applicants may refer to 9 Appendices B and D to NUREG-1927, Revision 1 (NRC, 2016b) when developing the 10 description of activities to assess data from the confirmatory demonstration.

11 3.2.4.1.2 Renewal Applications 12 Consistent with 10 CFR 72.42(a) and 10 CFR 72.240(a), a renewal application for a specific 13 license or CoC, may describe the activities to obtain and evaluate confirmatory data to be 14 performed under the aegis of an AMP. Applicants may refer to Appendices B and D to NUREG-15 1927, Revision 1 (NRC, 2016b) when developing the description of activities to assess data 16 from the confirmatory demonstration.

17 3.2.4.2 Supplemental Safety Analyses 18 As an alternative approach to a confirmatory demonstration for HBU SNF, an application may 19 supplement the design bases with safety analyses that demonstrate the DSS can still meet the 20 pertinent regulatory requirements by assuming hypothetical reconfiguration of the HBU fuel 21 contents into justified geometric forms. This alternative approach would demonstrate that the 22 design-bases fuel, even if reconfigured, can still meet the 10 CFR Part 72 requirements for 23 thermal, confinement, criticality safety and shielding during normal, off-normal, and accident 24 conditions. For renewal applications, a separate license amendment or CoC amendment may 25 be required if the changes in the supplemental safety analyses do not meet the acceptance 26 criteria in 10 CFR 72.48, Changes, Tests, and Experiments.

27 In NUREG/CR-7203 (NRC, 2015), ORNL Oak Ridge National Laboratory (ORNL) evaluated the 28 impact of a wide range of postulated fuel reconfiguration scenarios under non-mechanistic 29 causes of fuel assembly geometry change with respect to criticality, shielding (dose rates),

30 containment, and thermal. The study considered three fuel reconfiguration categories , which 31 were characterized by either category 1, cladding failure; category 2, rod/assembly deformation 32 without cladding failure; or category 3 changes to assembly axial alignment without cladding 33 failure. Within configurations in both Category 1 and Category 2, the study identified various 34 scenarios:

35

• Category 1: cladding failure 36 - Scenario 1(a): breached rods 37 - Scenario 1(b): damaged rods 38 39 3-12

For ACRS Review Purposes Only 1

• Category 2: rod/assembly deformation without cladding failure 2 - Scenario 2(a): configurations associated with side drop 3 - Scenario 2(b): configurations associated with end drop 4

• Category 3: changes to assembly axial alignment without cladding failure 5 The analyses in NUREG/CR-7203 (NRC, 2015) considered representative SNF transportation 6 packages, and a range of fuel initial enrichments, discharge burnup values, and decay times.

7 Two package designs were analyzed: a general burnup credit (GBC)-32 package containing 32 8 PWR fuel assemblies and a GBC-68 package containing 68 BWR fuel assemblies. Although 9 NUREG/CR-7203 did not evaluate reconfiguration in DSSs, the scenarios and analytical 10 methods may also be applicable to those designs, as the loads experienced during transport 11 conditions (normal, hypothetical accident) are expected to bound those experienced during 12 storage (normal, off-normal and accident). The results in NUREG/CR-7203 should not be 13 assumed to be generically applicable as fuel reconfiguration may have different consequences 14 for a DSS design other than the generic models evaluated in the study. However, the following 15 sections discuss considerations in developing supplemental safety analyses for other DSS 16 designs according to the reconfiguration scenarios considered in NUREG/CR-7203.

17 3.2.4.2.1 Materials and Structural 18 An application relying on supplemental safety analyses based on hypothetical reconfiguration of 19 the HBU SNF contents is expected to provide a structural evaluation for the package and its fuel 20 contents using any of the approaches discussed in Section 3.2. The staff will review the 21 structural evaluation and the assumed material mechanical properties, including any changes 22 due to higher temperatures resulting from fuel reconfiguration, in a manner consistent with the 23 guidance in the current SRP for dry storage of SNF.

24 3.2.4.2.2 Confinement 25 An applicant may demonstrate that a DSS design meets the regulatory requirements for 26 confinement for periods beyond 20 years by assuming hypothetical reconfiguration of the HBU 27 SNF into a bounding geometric form. However, if the thermal, structural, and material analyses, 28 together with aging management activities for the DSS subcomponents supporting confinement, 29 are used to provide assurance that the integrity of the confinement boundary is maintained even 30 after hypothetical reconfiguration of the fuel under normal, off-normal and accident-level 31 conditions, supplemental safety analysis for the confinement performance of the DSS design are 32 not expected. Thermal analyses demonstrate that all DSS subcomponents supporting 33 confinement (i.e., confinement boundary) will be able to withstand their maximum operating 34 temperatures and pressures under normal, off-normal and accident-level conditions.

35 3.2.4.2.3 Thermal 36 Fuel reconfiguration can affect the efficiency of heat removal from the fuel because of changes 37 in (1) thermo-physical properties of the canister gas space stemming from release of fuel rod 38 inert gas and fission product gases, (2) heat source location within the canister, and (3) changes 39 in flow area (convection), conduction lengths (conduction) and radiation view factors (thermal 40 radiation). As part of a defense-in-depth approach for addressing age-related uncertainties for 41 uncanned and undamaged HBU fuel in dry storage beyond 20 years, the thermal analyses 3-13

For ACRS Review Purposes Only 1 would be expected to analyze scenarios for normal, off-normal, and accident conditions of 2 storage by assuming the fuel may become substantially altered. NUREG/CR-7203 (NRC, 2015) 3 describes the impact on the DSS canister pressure and the fuel cladding and DSS component 4 temperatures for various scenarios of fuel geometry changes. These are examined below. In 5 general, the results in NUREG/CR-7203 should not be considered generically applicable. The 6 thermal analyses of the application are expected to consider scenarios discussed in 7 NUREG/CR-7203 to determine consistency in the analytical methods, scenario phenomena, 8 and results. The thermal analyses are expected to assess the impact of the fuel reconfiguration 9 on the fuel cladding and DSS component temperatures and the canister pressure for the 10 particular DSS design.

11 For Scenario 1(a) in Category 1 (see Section 3.2.4.2) , the fuel rods are assumed to breach in 12 such a manner that the cladding remains in its nominal geometry (no fuel reconfiguration), but 13 depending on the canister orientation (horizontal or vertical), the release of fuel rod fill gas and 14 fission product gases may have an effect on heat transfer which can cause a change to 15 maximum component temperatures. For Scenario 1(b) in Category 1, for configurations where 16 an assembly (or assemblies) is represented as a debris pile(s) inside its basket cell, fuel 17 reconfiguration has a larger impact on the component temperatures for the vertical orientation 18 than for the horizontal orientation, but the packing fraction of the debris bed has minor impact on 19 the component temperatures. For both Scenarios 1(a) and 1(b), release of the fuel rod gaseous 20 contents increases the number of moles of gas and therefore increases the canister pressure.

21 The canister pressure is expected to increase with the increased fuel rod release fractions.

22 For Scenarios 2(a) and 2(b), the fuel rods are assumed to remain intact without gaseous 23 leakage into the canister space. The changes of the fuel assembly lattice (contraction in 24 Scenario 2(a) and expansion in Scenario 2(b)) could cause either an increase or decrease in 25 the component temperatures of the storage system depending on the initial assembly geometry 26 and whether the storage system relies on convection for heat transfer. In general, scenarios 27 Scenario 2(a) and Scenario 2(b) have minor impact on the fuel cladding and DSS component 28 temperatures and canister pressure. For Category 3, the fuel rods are assumed to remain intact 29 without gaseous leakage into the canister space, but the axial shifting of the assembly changes 30 the heat source location within the canister. Changes in assembly axial alignment within the 31 basket cells are expected to have minor impact on the component temperatures and the 32 canister pressure.

33 Normal, Off-Normal, and Accident Conditions of Storage 34 Based on the thermal phenomena described above and NUREG/CR-7203 (NRC, 2015), an 35 approach acceptable to staff would evaluate the impact of Scenarios 1(a) and 1(b) on the 36 canister pressure and the fuel cladding and package component temperatures assuming 37 rupture of 1 percent, 10 percent and 100 percent of the fuel rods for normal, off-normal, and 38 accident conditions, respectively.

39 Although Scenarios 2(a) and 2(b) in Category 2 and Category 3 are not expected to have a 40 significant impact on DSS thermal performance under normal, off-normal and accident 41 conditions, because the fuel rods in Scenarios 2(a), 2(b) and 3 are assumed to remain intact 42 without gaseous leakage into the canister space, the applicant may need to provide a thermal 43 evaluation due to specifics of the DSS design.

3-14

For ACRS Review Purposes Only 1 3.2.4.2.4 Criticality 2 An application may demonstrate that a DSS design meets the regulatory requirements for 3 criticality safety for periods beyond 20 years by assuming hypothetical reconfiguration of the 4 HBU SNF into a bounding geometric form. This approach is one way to ensure compliance with 5 10 CFR 72.124, Criteria for Nuclear Criticality Safety, or 10 CFR 72.236(c) during normal, off-6 normal, and accident conditions, if the structural evaluation does not adequately define the 7 mechanical properties of the cladding.

8 As mentioned previously, ORNL examined hypothetical fuel reconfiguration for various 9 scenarios and the impacts on the criticality safety of a DSS and documented the results in 10 NUREG/CR-7203. This study considers burnup up to 70 GWd/MTU for criticality evaluations.

11 NUREG/CR-7203 provides some insight into the reactivity trends for various reconfiguration 12 scenarios; however, the results in NUREG/CR-7203 (NRC, 2015) should not be considered 13 generically applicable with respect to criticality safety analyses.

14 Criticality is not a concern for dry SNF systems, as SNF requires moderation to reach criticality.

15 Although DSS casks are expected to remain dry while in storage, cask users may be allowed to 16 load and unload a cask in a wet environment. The criticality analyses in NUREG/CR-7203 are 17 performed with an assumption of fully flooded conditions and any conclusions adopted are 18 applicable to analyses that support wet loading and unloading. The following considerations for 19 criticality evaluations for reconfigured fuel are applicable only to DSS scenarios where there 20 may be flooding within the canister. Otherwise, the staff does not find reconfiguration to pose a 21 criticality safety concern for a dry system.

22 All of the criticality safety analyses presented in NUREG/CR-7203 take credit for burned fuel 23 nuclides (burnup credit) and the conclusions may not be applicable to criticality analyses that 24 assume a fresh fuel composition. In its review of the burnup credit methodology and code 25 benchmarking used to support a criticality safety evaluation, the staff will follow the guidance in 26 ISG-8, Revision 3, Burnup Credit in the Criticality Safety Analyses of PWR Spent Fuel in 27 Transportation and Storage Casks, issued in September 2012 (NRC, 2012) to review the 28 burnup credit analyses. ISG-8, Revision 3, does not endorse any particular methodology for 29 BWR fuel burnup credit. The staff does not necessarily endorse the methodology described in 30 NUREG/CR-7203 for BWR fuel DSS, and considers it to be for illustration only.

31 For criticality safety analyses using burnup credit, NUREG/CR-7203 (NRC, 2015) shows that 32 reactivity increases for longer decay times (e.g., analyses supporting storage beyond 20 years);

33 therefore the application would need to use an appropriate decay time within the criticality 34 evaluations. The enrichment and burnup values assumed within the criticality evaluations in 35 NUREG/CR-7203 may differ from those allowed within another storage system. However 36 NUREG/CR-7203 states that no significant differences were observed in trends between 37 configurations that evaluated fuel at 44.25 GWd/MTU and 70 GWd/MTU.

38 The following sections discuss information from NUREG/CR-7203 that may be applicable when 39 performing reconfiguration analyses within a criticality evaluation for HBU fuel under normal, off-40 normal, and accident conditions of storage.

41 Normal Conditions of Storage 42 In an approach acceptable to the staff, the applicants criticality safety analyses would consider 43 the reactivity impact of 1-percent fuel failure during normal conditions of storage. The most 3-15

For ACRS Review Purposes Only 1 applicable scenario from NUREG/CR-7203 (NRC, 2015) is Scenario 1(a) (See Section 3.2.4.2 2 above for a description of the scenarios).

3 ORNL created Scenario 1(a) to represent breached rods. ORNL assumed that a percentage of 4 the rods were breached and that cladding from these rods failed completely and then removed 5 this percentage of fuel rods from the system. This is conservative as SNF systems are 6 undermoderated and replacing fuel with moderator typically causes reactivity to increase. Using 7 a fresh fuel composition for PWR fuel, ORNLs models in NUREG/CR-7203 showed that 8 reactivity decreases when removing rods. Therefore, this type of analysis may not be 9 appropriate for PWR analyses that assume a fresh fuel composition. The location assumed for 10 failed or removed rods can significantly affect reactivity. ORNL showed in Section A.1.1 of 11 NUREG/CR-7203 that removing rods from the center of the assembly causes reactivity to 12 increase the most.

13 In NUREG/CR-7203, ORNL also showed the number of rods removed that produces the 14 maximum reactivity. For the systems studied, NUREG/CR-7203 shows that the maximum 15 reactivity occurs when a number of rods far greater than 1-percent is removed from the system.

16 NUREG/CR-7203 also presents the results of a sensitivity study showing that reactivity increases 17 even more for Scenario 1(a) when it is assumed that the failed fuel relocates to a location outside 18 of the absorber plate. This is based on the generic systems modeled for the study. A different 19 system may allow relocation of the failed rod material outside of the absorber plate material to a 20 different extent.

21 Off-Normal Conditions of Storage 22 In an approach acceptable to the staff, the applicants criticality safety analyses would consider 23 the reactivity impact of 10-percent fuel failure under off-normal conditions of storage. The 24 methods discussed in the previous section on normal conditions of storage also apply to off-25 normal conditions of storage; however, the applicant would consider fuel failure up to 10 percent 26 rather than 1 percent. Scenario 1(a) can be used to represent rod failure via removing rods 27 from the system. In this case an applicant would remove 10-percent of the rods rather than 1-28 percent. The applicant would remove rods in such a way that it produces maximum reactivity 29 and consider relocation of the fuel to outside of the absorber plates.

30 Accident Conditions of Storage 31 In an approach acceptable to the staff, the applicants criticality safety analyses would consider 32 the reactivity impact of 100-percent fuel failure under accident conditions of storage. The 33 damaged fuel models in Section A.1.2 for Scenario 1(b) from NUREG/CR-7203 are applicable 34 when representing 100 percent failed fuel.

35 Scenario 1(b) from NUREG/CR-7203 considers reconfiguration of damaged fuel. With 100-36 percent compromise in cladding integrity, reconfiguration is considered to the maximum extent.

37 Section A.1.2 of NUREG/CR-7203 shows that a model assuming an ordered pellet array is 38 more reactive than a homogenous mixture of fuel, cladding materials and water.

39 3.2.4.2.5 Shielding 40 An application may demonstrate that a DSS continues to meet the regulatory dose limits for 41 periods beyond 20 years by assuming hypothetical reconfiguration of the HBU SNF into a 3-16

For ACRS Review Purposes Only 1 justified bounding geometric form under normal, off-normal, and accident conditions. This 2 method is one way to demonstrate compliance with 10 CFR 72.104, 10 CFR 72.106, or 10 CFR 3 72.236(d).

4 To assess the impacts of various fuel geometry changes on the shielding designs of DSSs and 5 ISFSIs, ORNL analyzed various scenarios of fuel geometry changes and the impact on the 6 annual dose at the ISFSI boundary and dose rates near the cask and presented the results in 7 NUREG/CR-7203 (NRC, 2015).

8 Appendix B to NUREG/CR-7203 provides some insight into the effects on external dose for 9 various reconfiguration scenarios; however, the results in NUREG/CR-7203 should not be 10 considered generically applicable with respect to external dose and dose rate evaluations. A 11 DSS designer would assess the impacts of fuel reconfiguration on external dose and dose rates 12 for its particular design using insights from NUREG/CR-7203 for reconfigured geometry.

13 This section discusses an approach acceptable to the staff for addressing the impacts on 14 external dose and dose rates when considering possible reconfiguration of HBU fuel for a period 15 of storage beyond 20 years. This discusses the scenarios from NUREG/CR-7203 most 16 applicable to the reconfiguration under normal, off-normal, and accident conditions of storage as 17 well as the analytical assumptions likely to result in bounding dose and dose rates based on the 18 results from NUREG/CR-7203. The NUREG has considered burnup up to 65 GWd/MTU within 19 its dose and dose rate evaluations. As discussed in Section B.5 of NUREG/CR-7203, different 20 nuclides become important to external dose and dose rate based on the decay time.

21 Since reconfiguration is to be considered after 20 years of storage, and this length of cooling 22 time is generally much longer than cooling times used to establish loading tables, applicants 23 may be able to make the justification that increases to external dose due to reconfiguration are 24 bounded by the additional cooling time the assemblies will experience.

25 NUREG/CR-7203 also indicates that fuel assembly type, (i.e., PWR vs BWR), may have a 26 significant impact on the surface dose rate and controlled area boundary dose under fuel 27 reconfiguration scenarios. Tables 13 and 14 of NUREG/CR-7203 show the difference in dose 28 rate increase for BWR and PWR SNF. A DSS system may permit storage of other fuel 29 assemblies, with different allowable burnup and enrichments to which the results of 30 NUREG/CR-7203 (NRC, 2015) do not apply. The burnup profile and depletion parameters used 31 to create the source term within NUREG/CR-7203 may also not be generically applicable.

32 Normal Conditions of Storage 33 In an approach acceptable to the staff, the applicants external dose and dose rate evaluation 34 would consider the impact of 1-percent fuel failure during normal conditions of storage. The 35 most applicable scenario from NUREG/CR-7203 is Category 1, fuel failure, Scenario, 1(a). If 36 cladding is breached and the fuel fails, this could lead to source relocation or change of the 37 geometric shape of the source. Based on NUREG/CR-7203, the impact on the controlled-area 38 boundary dose caused by source relocation resulting from 1-percent fuel failure is insignificant.

39 For a different DSS, the application may need to discuss potential fuel failure and source 40 reconfiguration and the potential impact on controlled-area boundary doses as required by 10 41 CFR 72.104 and 10 CFR 72.106.

42 Depending on the DSS and the resultant fuel geometry, the dose rate may increase significantly 43 as the detector moves close to the cask. Although it may not cause a significant change to the 3-17

For ACRS Review Purposes Only 1 dose far away from the cask and therefore may not constitute a significant concern for people at 2 the controlled area boundary, the changes of source term geometry will affect the doses of 3 occupational workers who need to perform necessary work around the casks. In general, an 4 application should consider the impact of HBF failure on the near cask dose rate and potential 5 impacts on radiation protection associated with ISFSI surveillance and maintenance operations.

6 Off-Normal Conditions of Storage 7 In an approach acceptable to the staff, the applicants external dose and dose rate evaluation 8 for HBF would consider the impact of 10-percent fuel failure under off-normal conditions of 9 storage. If cladding is breached and fails, the fuel, and hence the source, may relocate to 10 different parts of the fuel basket. The impact of HBF failure on dose at the controlled-area 11 boundary for storage under off-normal conditions of dry storage operations should be examined.

12 A 10-percent fuel failure is similar to Scenario 1(a) in NUREG/CR-7203 (NRC, 2015). For 13 Scenario 1(a), breached rods, ORNL assumed the rods turned to rubble and calculated the 14 dose rate when the fuel mixture relocated to the bottom of the fuel assembly. ORNL assumed 15 failure of 10-percent of fuel rods collected into the available free volume within the assembly 16 lower hardware region. Section B.4.1 of NUREG/CR-7203 discusses the implementation in 17 detail. ORNL reduced the source strength and density of the active fuel zone by the failure 18 percentage and relocated this source to the bottom of the fuel assembly and increased the 19 source strength and density accordingly. The storage system in NUREG/CR-7203 is modeled 20 as a vertically-oriented storage system. Fuel would likely not relocate this way in a horizontal 21 storage system, and the model is not necessarily applicable to a horizontal system.

22 In Section B.5.5 of NUREG/CR-7203, ORNL discuss the results of the study performed on the 23 individual DSS, which shows that there could be significant increases in the dose rate near the 24 cask. It concludes that fuel configuration changes can cause significant dose rate increases 25 relative to the nominal intact fuel configuration in the cask outer regions that face air vent 26 locations. NUREG/CR-7203 states that the change in radiation dose rate away from air vent 27 locations is either small or negligible.

28 Similar to normal conditions of storage, the changes in source term geometry will impact the 29 doses of occupational workers who need to perform necessary surveillance and maintenance 30 work around the casks. To assess the impacts on radiation protection, an applicant may need 31 to evaluate the surface dose rate increase resulting from reconfiguration.

32 Accident Conditions of Storage 33 In an approach acceptable to the staff, the applicants external dose and dose rate evaluation 34 for HBF would consider the impact of 100-percent fuel failure during accident conditions of 35 storage. If cladding is breached and the fuel fails, this may cause the fuel, and hence the 36 source, to relocate to different parts of the fuel basket. Based NUREG/CR-7203 (NRC, 2015),

37 the impacts on the controlled-area boundary dose caused by source relocation resulting from 38 100 percent fuel failure will result in significant increases in the dose rate near the cask and 39 annual dose at the controlled area boundary. Scenarios 1(b) and 2 in NUREG/CR 7203 can 40 represent 100-percent fuel failure.

41 At the controlled area boundary, 100-percent fuel reconfiguration can have a significant impact 42 on the annual dose. It can also significantly affect the dose rate near the cask and the radiation 43 protection associated with ISFSI remediation operations. Tables B.9 and B.10 of Appendix B to 3-18

For ACRS Review Purposes Only 1 NUREG/CR-7203 (NRC, 2015) show the relative changes in dose rates at 1 meter from a 2 sample PWR fuel cask and a sample BWR fuel cask, respectively. Table B.11 of Appendix B to 3 NUREG/CR-7203 shows the estimated relative impact on controlled-area boundary dose from 4 fuel reconfiguration. The data presented in these tables show that the impacts on the dose 5 rates at the cask side, particularly the dose rate near the vent ports are significant.

6 In Scenario 1(b), ORNL assumed that the assembly and basket plate material is homogenized, 7 placed it at the bottom of the cask, and determined that the limiting packing fraction is 0.58.

8 This scenario did not produce an increase in site boundary dose; however, it did show an 9 increase in local dose rates. The location of the bottom of the cask would depend on whether 10 the DSS is vertical or horizontal. Homogenizing the basket material with the fuel rubble may be 11 overly conservative for a horizontal configuration, and applicants may choose to maintain basket 12 integrity similar to the Scenario S2 model in Section B.4.2 of NUREG/CR-7203 when evaluating 13 dose or dose rates for a horizontal system or a tip-over scenario.

14 For Scenario 1(b), ORNL also assumed that the fuel and basket material forms a homogenized 15 rubble that is distributed throughout the canister cavity. This scenario produced an increase in 16 site boundary dose.

17 3.3 Canned Fuel (Damaged Fuel) 18 10 CFR 72.122(h)(1) requires SNF, including HBU, with gross ruptures (i.e., classified as 19 damaged) be placed in a can designed for damaged fuel or in an acceptable alternative. The 20 staff will follow the guidance in the current SRPs for dry storage of SNF in its review of an 21 application for a DSS with damaged HBU SNF contents.

3-19

For ACRS Review Purposes Only 1 4 TRANSPORTATION OF HIGH BURNUP SPENT NUCLEAR FUEL 2 4.1 Introduction 3 The U.S. Nuclear Regulatory Commission (NRC) staff (the staff) has developed example 4 approaches for approval of transportation packages with high burnup (HBU) spent nuclear fuel 5 (SNF). Applicants may use these approaches to provide reasonable assurance of compliance 6 with Title 10 of the Code of Federal Regulations (10 CFR) Part 71, Packaging and 7 Transportation of Radioactive Material, during normal conditions of transport and hypothetical 8 accident conditions. The staff developed these example approaches based on the conclusions 9 of the engineering assessment in Chapter 2. Figure 4-1 provides a high-level diagram of these 10 approaches, which vary based on (1) the condition of the fuel (undamaged or damaged), and 11 (2) the length of time the fuel has been in prior dry storage. Considerations for additional 12 analyses expected for non-leaktight transportation packages are also provided (see Section 13 4.2.2). An applicant may consider and demonstrate other approaches to be acceptable.

14 As required by 10 CFR 71.33(b), an application for a transportation package should identify 15 allowable SNF contents and condition of the assembly and rods. The allowable cladding 16 condition for the SNF contents is generally defined in the certificate of compliance (CoC), and 17 the nomenclature may vary between different transportation packages. For example, the terms 18 intact and undamaged have both been used to describe cladding without any known gross 19 cladding breaches. In accordance with 10 CFR 71.17(c)(2) (for NRC licensees) and 20 49 CFR 173.471 (for non-NRC licensees), users of transportation packages must comply with 21 the CoC by selecting and loading the appropriate fuel, and, in accordance with 10 CFR 71.91, 22 Records, must maintain records that reasonably demonstrate that loaded fuel was adequately 23 selected, in accordance with their approved site procedures and Quality Assurance Program.

24 Interim Staff Guidance (ISG)-1, Revision 2, Classifying the Condition of Spent Nuclear Fuel for 25 Interim Storage and Transportation Based on Function, issued in May 2007 (NRC, 2007b),

26 provides guidance for developing the technical basis supporting the conclusion that the HBU 27 SNF (both rods and assembly) to be shipped are intact or undamaged.1 This would include 28 considering whether the material properties, and possibly the configuration, of the SNF 29 assemblies may have been altered during prior dry storage. If the alteration is not within the 30 bounds of the approved contents for the transportation package, then an application must be 31 submitted to revise the CoC. This application must show that, with the altered condition of the 32 SNF, the package can still meet the regulations in 10 CFR Part 71.

33 Damaged SNF is generally defined in terms of the characteristics needed to perform functions 34 to assure compliance with fuel-specific and package-related regulations. A fuel-specific 35 regulation defines a characteristic or performance requirement of the SNF assembly (e.g., 10 36 CFR 71.55(d)(2)). A package-related regulation defines a performance requirement placed on 37 the fuel so that the transportation package can meet a regulatory requirement (e.g., 10 CFR 38 71.55(e)). The glossary provides the staffs definitions of intact, undamaged, and damaged fuel.

39 For additional information, refer to the current standard review plan (SRP) for transportation of 40 SNF (NUREG-1617, Standard Review Plan for Transportation Packages for Spent Nuclear 1 The current revisions of all ISG documents will be rolled into revised standard review plans (SRPs) for dry storage and transportation, as appropriate, and will then be removed from the public domain. The revised SRPs will be issued for public comment prior to being finalized.

4-1

For ACRS Review Purposes Only 1 fuel, issued in March 2000 (NRC, 2000b)) - hereafter referred to as the current SRP for 2 transportation SNF.

4-2

4-3 For ACRS Review Purposes Only Figure 4-1 Example Approaches for Approval of Transportation Packages with High Burnup Spent Nuclear Fuel

For ACRS Review Purposes Only 1 Consistent with the guidance in (ISG)-1, Revision 2 (NRC, 2007b), SNF assemblies with any of 2 the following characteristics, as identified during the fuel selection process, are generally 3 classified as damaged unless an adequate justification is provided that shows otherwise:

4

• There is visible deformation of the rods in the HBU SNF assembly. This is not referring 5 to the uniform bowing that occurs in the reactor; instead, this refers to bowing that 6 significantly opens up the lattice spacing.

7

• Individual fuel rods are missing from the assembly. The assembly may be classified as 8 intact or undamaged if the missing rod(s) do not adversely affect the structural 9 performance of the assembly, and radiological and criticality safety (e.g., there are no 10 significant changes to rod pitch). Alternatively, the assembly may be classified as intact 11 or undamaged if a dummy rod that displaces a volume equal to, or greater than, the 12 original fuel rod is placed in the empty rod location.

13

• The HBU SNF assembly has missing, displaced, or damaged structural components 14 such that either of the following occurs:

15 - Radiological and/or criticality safety is adversely affected (e.g., significantly 16 changed rod pitch) 17 - The structural performance of the assembly may be compromised during normal 18 conditions of transport (NCT) or hypothetical accident conditions (HAC).

19

• Reactor operating records or fuel classification records indicate that the HBU SNF 20 assembly contains fuel rods with gross ruptures.

21

• The HBU SNF assembly is no longer in the form of an intact fuel bundle (e.g., it consists 22 of, or contains, debris such as loose fuel pellets or rod segments).

23 Defects such as dents in rods, bent or missing structural members, small cracks in structural 24 members, and missing rods do not necessarily render an assembly damaged, if the intended 25 functions of the assembly are maintained (i.e., if the performance of the assembly does not 26 compromise the ability to meet fuel-specific and package-related regulations).

27 4.2 Uncanned Fuel (Intact and Undamaged Fuel) 28 Undamaged HBU SNF can be transported without the need for a separate can for damaged fuel 29 (i.e., a separate metal enclosure sized to confine damaged fuel particulates) to maintain a 30 known configuration inside the package containment cavity. This fuel includes rods that are 31 either intact (i.e., there are no breaches of any kind) or that contain small cladding defects (i.e.

32 pinholes or hairline cracks), which may permit the release of gas from the interior of the fuel rod.

33 Cladding with gross ruptures that may permit the release of fuel particulates may not be 34 considered undamaged. The configuration of undamaged HBU SNF may be demonstrated to 35 be maintained if loading and transport operations are designed to prevent or mitigate 36 degradation of the cladding and other assembly components, as discussed in ISG-22, Potential 37 Rod Splitting Due to Exposure to an Oxidizing Atmosphere during Short-Term Cask Loading 38 Operations in LWR or Other Uranium Oxide Based Fuel, issued May 2006 (NRC, 2006)..

39 As the approaches delineated in Figure 4-1 show, an application for a CoC for a package that 40 includes undamaged HBU SNF would include a structural evaluation of the fuel rods under NCT 4-4

For ACRS Review Purposes Only 1 and HAC drop accident scenarios. The evaluation serves to demonstrate that the uncanned fuel 2 remains in a known configuration after a drop accident scenario.

3 Two alternatives may be used to calculate cladding stress and strain, and cladding flexural 4 rigidity, for the aforementioned evaluation of drop accident scenarios. The first alternative, 5 shown in Figure 4-2, is to use cladding-only mechanical properties from as-irradiated cladding 6 (i.e., cladding with circumferential hydrides, primarily), or hydride-reoriented cladding (i.e, 7 cladding that accounts for radial hydrides precipitated after the drying process). As indicated in 8 the discussion in Section 2.3.3, the staff considers that the orientation of the hydrides is not 9 critical in evaluating the adequacy of cladding-only mechanical properties during drop accident 10 scenarios. The properties necessary to implement this alternative may be derived from 11 cladding-only uniaxial tensile tests and include modulus of elasticity, yield stress, ultimate 12 tensile strength and uniform strain, and the strain at failure (i.e., the elongation strain). Refer to 13 the current SRP for transportation of SNF for additional considerations on acceptable cladding-14 only mechanical properties (i.e., alloy type, burnup, and temperature) and the acceptance criteria 15 for cladding performance during transport operations can be found in are described in.

16 Figure 4-2 First Approach for Evaluation of Drop Accidents During Transport 4-5

For ACRS Review Purposes Only 1 The second alternative, outlined in Figure 4-3, is to use cladding-only mechanical properties that 2 have been modified by a numerical factor to account for the increased flexural rigidity imparted 3 by the fuel pellet. This numerical factor can be obtained from static test data from the cyclic 4 integrated reversible-bending fatigue tester (CIRFT) for fully-fueled rods for the particular 5 cladding type and fuel type (see Section 2.3.3). The second alternative would be necessary only 6 if the structural evaluation using cladding-only mechanical properties is unsatisfactory, although 7 an applicant may choose to implement it even if the first alternative were to yield satisfactory 8 results. Refer to the current SRP for transportation of SNF for acceptance criteria on cladding 9 performance following NCT and HAC drop scenarios.

10 Figure 4-3 Second Approach for Evaluation of Drop Accidents During Transport 11 In addition to the structural evaluation for NCT and HAC drop accident scenarios, the 12 application would contain a fatigue evaluation for NCT using the cumulative damage approach 13 described in Section 2.3. The satisfactory performance under fatigue would serve to 14 demonstrate compliance with the requirement in 10 CFR 71.71(c)(5).

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For ACRS Review Purposes Only 1 Figure 4-4 Evaluation of Vibration Normally Incident to Transport 2 4.2.1 Leaktight Containment 3 An application for a transportation package CoC with HBU SNF as contents is expected to 4 define the maximum allowable leakage rate for the entire containment boundary. The maximum 5 allowable leakage rate is based on the quantity of radionuclides available for release and is 6 evaluated to meet the containment requirements for maintaining an inert atmosphere within the 7 containment cavity and compliance with the regulatory release limits of 10 CFR 71.51, 8 Additional Requirements for Type B Packages. The leakage rate testing is performed on the 9 entire containment boundary (over the course of fabrication and loading) and ensures that the 10 package can maintain a leakage rate below the maximum allowable leakage rate per ANSI 11 N14.5-2014.

12 If the entire containment boundary of the transportation package, including its closure lid, is 13 designed and tested to be leaktight as defined in American National Standards Institute (ANSI) 14 N14.5-2014, American National Standard for Radioactive Materials Leakage Tests on 15 Packages for Shipment, and the current SRP for transportation of SNF, then the application is 16 not expected to include release calculations that demonstrate compliance with the regulatory 17 release limits of 10 CFR 71.51. In addition, the structural analyses of the package 18 demonstrates that the containment boundary will not fail under the tests for NCT and HAC and 19 that the containment boundary will remain leaktight under all conditions of transport. Refer to 20 the current SRP for transportation of SNF for additional guidance on demonstrating compliance 21 with the leaktight criterion.

22 4.2.2 Non-Leaktight Containment 23 Transportation packages certified to transport HBU SNF must satisfy the release limits of 24 10 CFR 71.51. For those packages not tested to a leaktight criterion, the application is 4-7

For ACRS Review Purposes Only 1 expected to include release calculations and identify the allowable NCT and HAC volumetric 2 leakage rates in accordance with ANSI N14.5. The standard provides an acceptable method to 3 determine the maximum permissible volumetric leakage rates based on the allowed regulatory 4 release limits under both NCT and HAC. Refer to the current SRP for transportation of SN for 5 additional guidance on demonstrating compliance with 10 CFR 71.51 for non-leaktight packages.

6 The leakage rate testing is performed on the entire containment boundary (over the course of 7 fabrication and loading) and ensures that the package can maintain a leakage rate below the 8 maximum allowable leakage rate, which can be calculated using the methodology in ANSI N14.5 9 (2014). In order to determine the release rates for the primary containment boundary, an 10 application for certification of a non-leaktight package should provide a technical basis for the 11 assumed bounding HBU fuel failure rates for both NCT and HAC. If an application is not able to 12 provide and justify its bounding HBU fuel failure rates, then the fuel failure rates below may be 13 assumed as bounding values for NCT and HAC:

14

• NCT: 3 percent 15

• HAC: 100 percent 16 Bounding Release Fractions for High Burnup Fuel 17 HBU SNF has different characteristics relative to low burnup (LBU) SNF with respect to cladding 18 oxide thickness, hydride content, radionuclide inventory and distribution, heat load, fuel pellet 19 grain size, fuel pellet fragmentation, fuel pellet expansion and fission gas release to the rod 20 plenum (See Appendix C.5 to NUREG/CR-7203, A Quantitative Impact Assessment of 21 Hypothetical Spent Fuel Reconfiguration in Spent Fuel Storage Casks and Transportation 22 Packages, issued September 2015 (NRC, 2015) for additional details on HBU SNF).

23 Differences in these characteristics affect the mechanisms by which the fuel can breach and the 24 amount of fuel that can be released from failed fuel rods. Hence, the staff evaluated open 25 literature on HBU fuel rod failure rates and release fractions (Chalk River unknown deposits 26 (CRUD), fission gases, volatiles, and fuel fines) to assist in the review of applications for non-27 leaktight containment boundaries. Table 4-1 provides release fractions that may be considered 28 reasonably bounding for HBU SNF. If these release fractions are not used, other release 29 fractions may be used in the analysis provided the applicant properly justifies the basis for their 30 usage. Justification of the proposed release fractions of the source terms should consider an 31 adequate description of burnup for the test specimen, number of tests, collection method for 32 quantification of release fractions, test specimen pressure at the time of fracture, and source 33 collection system.

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For ACRS Review Purposes Only 1 Table 4-1 Fractions of Radioactive Materials Available for Release from HBU SNF 2 Under Conditions of Transport (for Both Pressurized Water Reactor and 3 Boiling Water Reactor Fuels)

HAC-Fire HAC-Impact Variable NCT Conditions Conditions Fraction of Fuel Rods 0.03 1.0 1.0 Assumed to Fail Fraction of Fission Gases Released Due to 0.15 0.15 0.35 a Cladding Breach Fraction of Volatiles Released Due to a 3 x 10-5 3 x 10-5 3 x 10-5 Cladding Breach Mass Fraction of Fuel Released as Fines Due 3 x 10-5 3 x 10-3 3 x 10-5 to a Cladding Breach Fraction of CRUD 0.15 1.0 1.0 Spalling Off Cladding 4 CRUD 5 The average CRUD thickness on HBU SNF cladding has been estimated to be similar to that 6 observed on LBU SNF cladding. A review of data from the literature (NRC, 2000c; Einziger and 7 Beyer, 2007) indicates that a release (spalling off) of 15 percent of cladding CRUD may be 8 assumed as reasonably bounding to NCT scenarios, and a release fraction of 100 percent of 9 the cladding CRUD, that spalls off, is conservatively bounding to HAC scenarios (NRC, 2014).

10 Fission Gases 11 NRCs FRAPCON steady-state fuel performance code has been previously used to assess 12 release fractions of fission gases during transportation (NRC, 2011). The seven most common 13 fuel designs were evaluated using FRAPCONs modified Forsberg-Massih model (8 x 8, 9 x 9, 14 and 10 x 10 fuel for BWR; and 14 x 14, 15 x 15, 16 x 16, and 17 x 17 for PWR). For each fuel 15 design, a number of different power histories aimed at capturing possible realistic reactor 16 irradiations were modeled. The fission gas content within the free volume of the rods was 17 evaluated for a total of 243 different cases (39 for each of the BWR fuel designs; 37 for 14 x 14 18 and 16 x 16 PWR fuel designs, and 26 for 15 x 15 and 17 x 17 PWR fuel designs). A review of 19 the results indicates that a release of 15 percent of fission gases may be assumed as 20 reasonably bounding to NCT scenarios for rod average burnups up to 62.5 GWd/MTU.

21 During an HAC fire scenario, per 10 CFR 71.73(c)(4), the fuel is not expected to reach 22 temperatures high enough that fission gases can diffuse out of the pellet matrix or grain 23 boundaries to the rod plenum. The thermal rupture tests showed that release occurred at 24 higher temperatures than those experienced during HAC (NRC, 2000c). Therefore, the same 25 release fraction of 15 percent of fission gases during NCT scenarios may also be assumed to 26 be reasonably bounding to the HAC fire scenario.

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For ACRS Review Purposes Only 1 In the case of HAC drop (impact) conditions, the pellet may be conservatively assumed to 2 crumble. In this scenario, fission gases retained within the pellet grain boundaries may be 3 released in addition to those already released from the fuel rod free volume (i.e., from the fuel-4 cladding gap and plenum). The FRAPFGR model in FRAPCON may be used to predict the 5 location of the fission gases within the fuel pellet (NRC, 2011). The model has been validated 6 with experimental data obtained using an electron probe micro analyzer. The FRAPFGR model 7 was used to calculate the maximum fraction of the pellet-retained fission gases that may be 8 released during a drop impact, which was determined to be 20 percent. Therefore, assuming all 9 fission gases within the pellet grain boundaries are released, a 35-percent (15-percent + 20-10 percent) maximum release fraction may be assumed to be reasonably bounding to the HAC fire 11 scenario. This value accounts for the 15-percent maximum fission gases released from the fuel 12 rod free volume (as calculated with the modified Forsberg- Massih model) and the 20-percent 13 maximum fission gases released from the fuel pellet grain boundaries (as calculated with the 14 FRAPFGR model). These release fraction estimates are consistent with previous NRC 15 estimates (NRC, 2000c; NRC, 2007; Einziger and Beyer, 2007).

16 Volatiles 17 The majority of the volatile release fractions originate from cesium-based compounds in the 18 form of oxides or chlorides (NRC, 2000c; NRC, 2014). These volatiles exhibit a different 19 release behavior in comparison to fission gases. Volatiles tend to migrate and aggregate at the 20 rim on the outer surface of the fuel pellet during reactor irradiation, which is characteristic of 21 burnups near or exceeding 60 GWd/MTU. The pellet rim is characterized by a fine crystalline 22 grain structure (0.1 - 0.3 µm in characteristic size) (Spino et al., 2003; Einziger and Beyer, 23 2007), a high porosity that may exceed 25-percent, and a high concentration of actinides 24 relative to the inner pellet matrix.

25 Sandia National Laboratories determined the maximum release fraction of volatiles (cesium and 26 other ruthenium-based compounds) under HAC drop and fire scenarios to be 0.003 percent (3 27 x10-5) (NRC, 2000c). The assessment included modeling and analyses using various data from 28 the literature. The volatile release fraction during an HAC fire scenario was determined to be 29 lower than the release fraction during an HAC impact scenario (NRC, 2014; NRC, 2000c).

30 Therefore, a volatile release fraction of 0.003 percent (3 x 10-5) may be assumed to be 31 reasonably bounding to NCT, HAC fire, and HAC impact scenarios. This release fraction 32 estimate is also consistent with an independent estimate by Einziger and Beyer (2007).

33 Fuel Fines 34 Release fractions from SNF fines during storage and transportation have been previously 35 documented (NRC, 2000c; Benke et al., 2012; NRC, 2007; NRC, 2014). HBU SNF has a 36 different pellet microstructure relative to LBU SNF, which is characterized by an inner matrix 37 and an outer pellet rim layer. The thickness of the outer pellet rim layer increases with higher 38 fuel burnup. Therefore, differences in microstructure between the inner pellet matrix and the 39 outer pellet rim should be considered when evaluating release fractions of fuel fines from HBU 40 SNF.

41 Although there is no reported literature on HBU SNF rim fracture as a function of impact energy, 42 other data can be used to indirectly assess the contribution of the rim layer to the release 43 fractions of fuel fines. Spino et al (1996) estimated the fracture toughness of the rim layer from 44 micro-indentation tests. Relative to the inner SNF matrix, the rim layer showed an increase of 45 fracture toughness. The increase of fracture toughness implies a decrease of release fraction.

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For ACRS Review Purposes Only 1 Hirose et al (2015) also discussed results of axial dynamic impact tests simulating accident 2 conditions during transport. The dispersed particles due to pellet breakage following impact 3 were collected and correlated to impact energy. The staff has compared the measured release 4 fraction of fuel fines from Hirose et al (2015) with previous NRC estimates of release fraction 5 versus impact energy for SNF and other brittle materials (depleted UO2, glass and Synroc) (see 6 Figure 3 of NUREG-1864 (NRC (2007)). Based on these analyses, the staff concludes that 7 there is no indication that pellet rim layer contributes to increased release fractions for HBU 8 SNF.

9 Since the outer HBU fuel pellet rim does not appear to contribute to additional release fractions, 10 previous NRC estimates for release fractions of fuel fines may continue to be used (NRC, 11 2000c; NRC, 2007; Benke et al., 2012; Ahn et al., 2012; NRC, 2014). Based on the range of 12 estimates in the literature, a release fraction for fuel fines of 0.003 percent (3 x10-5) may be 13 assumed to be reasonably bounding to both NCT and HAC (drop impact) scenarios. During an 14 HAC fire scenario, fuel oxidation is conservatively assumed to increase the release fraction of 15 fuel fines by a factor of 100 (NRC, 2000c; Ahn et al, 2012). Therefore, a 0.3 percent (3 x 10-3) 16 release fraction of fuel fines may be assumed as reasonably bounding to an HAC fire scenario.

17 The staff recognizes that various international cooperative research programs are currently 18 investigating release fractions from HBU SNF. Once the data is available to the public, the staff 19 will review and determine if the conservative estimates in the above discussion should be 20 revisited.

21 4.2.3 Direct Shipment from the Spent Fuel Pool and Shipment of Previously Dry-22 Stored Fuel (Up To 20 Years Since Fuel Was Initially Loaded) 23 Section 1.2 discussed the staffs review guidance for the licensing and certification of dry 24 storage of HBU SNF for a period up to 20 years. The technical basis referenced in that 25 guidance has supported the staffs conclusion that creep is not expected to result in gross 26 ruptures if cladding temperatures are maintained below 400 °C (752 °F). Creep is a time-27 dependent mechanism. Therefore, the short transportation period (relative to dry storage) is not 28 expected to compromise the integrity of HBU SNF, if the cladding temperatures remain below 29 400 °C (752 °F).

30 Chapter 2 also provided an assessment of the effects of hydride reorientation per static and 31 fatigue bending test results on HBU SNF specimens. Those results provide a technical basis 32 for the staffs conclusion that the use of best-estimate cladding mechanical properties (with 33 either as-irradiated or hydride-reoriented microstructure) is adequate for the structural 34 evaluation of HBU SNF. This finding applies to the evaluation of the drop tests for NCT (per 35 10 CFR 71.71(c)(7)) and HAC (per 10 CFR 71.73(c)(1)). Refer to the current SRP for 36 transportation of SNF for staff review guidance on additional considerations for acceptable 37 cladding-only mechanical properties (i.e., alloy type, burnup, temperature), on acceptable 38 references for cladding mechanical properties and on acceptance criteria for the structural 39 evaluation of the HBU fuel assembly following the drop tests. As Figure 4-1 shows, 40 supplemental safety analyses are not expected for dry storage of HBU SNF directly loaded from 41 the spent fuel pool or HBU SNF that has previously been in dry storage for periods not 42 exceeding 20 years.

4-11

For ACRS Review Purposes Only 1 4.2.4 Shipment of Previously Dry-Stored Fuel (Beyond 20 Years Since Fuel Was 2 Initially Loaded) 3 To address age-related uncertainties related to the transportation of HBU SNF previously in dry 4 storage for extended periods (i.e., periods of storage exceeding 20 years), the application 5 should be supplemented with either results from a surrogate demonstration program or 6 supplemental safety analyses assuming justified hypothetical fuel reconfiguration scenarios (see 7 Figure 4-1). The results from a surrogate demonstration program can provide field-obtained 8 confirmation that the fuel has remained in the analyzed configuration after 20 years of dry 9 storage, if that is the approved configuration for the transportation package. If confirmation is 10 not provided, the safety analyses for the transportation package should be revised to assume 11 reconfigured fuel.

12 The licensing and certification of storage containers for HBU SNF has been approved for an 13 initial 20-year-term per the technical basis for the evaluation of creep, as discussed in Chapter 14 1. However, the staff has recognized that the technical basis is based on short-term 15 accelerated creep testing (i.e., laboratory scale testing up to a few months), which results in 16 increased uncertainties when extrapolated to long periods of dry storage (see Appendix D to 17 NUREG-1927, Revision 1, Standard Review Plan for Renewal of Specific Licenses and 18 Certificates of Compliance for Dry Storage of Spent Nuclear Fuel, issued June 2016 (NRC, 19 2016b). Although the staff has confidence based on this short-term testing that creep-related 20 degradation of the HBU fuel will not adversely affect its analyzed configuration for storage 21 periods beyond 20 years, there is no operational field-obtained data to confirm this expectation, 22 as in the prior demonstration for LBU fuel (NRC, 2001; NRC, 2003b).

23 In addition, the staff also acknowledges that, while the CIRFT results obtained to date (as 24 discussed in Chapter 2) provide an adequate technical basis for assessing the separate effects 25 of hydride reorientation, the results do not account for potential synergistic effects of various 26 physical and chemical phenomena occurring during extended dry storage (e.g., cladding creep, 27 hydride reorientation, irradiation hardening, oxidation, hydriding caused by residual water 28 hydrolysis (see NUREG-2214, Managing Aging Processes in Storage (MAPS) Report, issued 29 August 2019 (NRC, 2019) for discussions on these phenomena). Therefore, evidence that HBU 30 fuel in dry storage beyond 20 years has maintained its analyzed configuration is expected prior 31 to transport, if that is the approved configuration for the transportation package.

32 4.2.4.1 Supplemental Data from Confirmatory Demonstration 33 One example approach for approval of a transportation package with HBU SNF previously in 34 dry storage for periods exceeding 20 years (e.g., 40 years) involves supplementing the 35 application with results from a surrogate demonstration program. Such a program could provide 36 field-obtained confirmation that the fuel configuration has been maintained before transport.

37 The applicant may refer to Appendices B and D to NUREG-1927, Revision 1 (NRC, 2016b),

38 which describe attributes and acceptance criteria of an acceptable surrogate demonstration 39 program.

40 4.2.4.2 Supplemental Safety Analyses 41 As an alternative approach to relying on a surveillance and monitoring program for the 42 transportation of HBU SNF previously in dry storage for longer than 20 years, an application 43 may demonstrate that a transportation package can still meet the pertinent regulatory 44 requirements by assuming hypothetical reconfiguration of the fuel contents into justified 4-12

For ACRS Review Purposes Only 1 geometric forms. This alternative approach would include supplemental safety analyses to 2 demonstrate that the HBU SNF contents, even if reconfigured, can still meet the pertinent 10 3 CFR Part 71 regulations for containment, thermal performance, criticality safety and shielding 4 after the required tests for NCT and HAC.

5 In NUREG/CR-7203 (NRC, 2015), Oak Ridge National Laboratory (ORNL) evaluated the impact 6 of a wide range of postulated fuel reconfiguration scenarios under non-mechanistic causes of 7 fuel assembly geometry change with respect to criticality, shielding (dose rates), containment, 8 and thermal performance. The study considered three fuel reconfiguration categories, which 9 were characterized by either (1) cladding failure, (2) rod/assembly deformation without cladding 10 failure or (3) changes to assembly axial alignment without cladding failure. Within 11 configurations in both Category 1 and Category 2, various scenarios were identified:

12

• Category 1: cladding failure 13 - Scenario 1(a): breached spent fuel rods 14 - Scenario 1(b): damaged spent fuel rods 15

• Category 2: rod/assembly deformation without cladding failure 16 - Scenario 2(a): configurations associated with side drop 17 - Scenario 2(b): configurations associated with end drop 18

• Category 3: changes to assembly axial alignment without cladding failure 19 The analyses in NUREG/CR-7203 (NRC, 2015) considered representative SNF transportation 20 packages, and a range of fuel initial enrichments, discharge burnup values, and decay times.

21 The analyses examined two package designs: a general burnup credit (GBC)-32 package 22 containing 32 PWR fuel assemblies and a GBC-68 package containing 68 BWR fuel 23 assemblies. The results in NUREG/CR-7203 should not be assumed to be generically 24 applicable, as fuel reconfiguration may have different consequences for a transportation 25 package other than the generic models evaluated in NUREG/CR-7203, however, the following 26 sections discuss considerations in developing supplemental safety analyses for other packages 27 according to the reconfiguration scenarios considered in NUREG/CR-7203.

28 4.2.4.2.1 Materials and Structural 29 An application for package certification relying on supplemental safety analyses based on 30 hypothetical reconfiguration of the HBU SNF contents should still provide a structural evaluation 31 for the package and its fuel contents using any of the approaches discussed in Section 4.2. The 32 staff will follow the guidance in the current SRP for transportation of SNF in its review of the 33 structural evaluation and the assumed material mechanical properties, including any changes 34 caused by higher temperatures resulting from fuel reconfiguration.

35 4.2.4.2.2 Containment 36 An application relying on supplemental safety analyses based on hypothetical reconfiguration of 37 the HBU SNF is expected to demonstrate that the transportation package design meets the 38 regulatory requirements for containment if data from a surrogate demonstration program, used 4-13

For ACRS Review Purposes Only 1 for confirmatory demonstration consistent with the guidance in NUREG-1927 (NRC, 2016b), are 2 not available before shipment of fuel in prior dry storage for periods longer than 20 years.

3 Thermal, structural, and material analyses, together with aging management activities for the 4 DSS subcomponents supporting confinement (i.e., confinement boundary) during prior dry 5 storage,1 serve to provide assurance that the allowable leak rate is maintained even after 6 hypothetical reconfiguration of the fuel under NCT and HAC. Supplemental thermal analyses 7 should demonstrate that the containment boundary will be able to withstand their maximum 8 operating temperatures and pressures under NCT and HAC. If the canister serves as the 9 confinement boundary at the future storage location, then the canister is expected to be leak-10 tested while it is within the transportation package after it reaches its new storage location.

11 4.2.4.2.3 Thermal 12 Fuel reconfiguration can affect the efficiency of heat removal from the fuel because of changes 13 in (1) thermo-physical properties of the container gas space resulting from the release of fuel 14 rod fill gas and fission product gases, (2) heat source location within the container, and (3) 15 changes in flow area (convection), conduction lengths (conduction) and radiation view factors 16 (thermal radiation). As part of a defense-in-depth approach to addressing age-related 17 uncertainties for uncanned or undamaged HBU SNF in shipment for fuel previously in dry 18 storage for periods longer than 20 years, the thermal analyses would be expected to analyze 19 the spent fuel at NCT and HAC by assuming the fuel has become substantially altered.

20 NUREG/CR-7203 (NRC, 2015) describes impacts on canister pressure and fuel cladding, and 21 package component temperatures for various scenarios of fuel geometry changes. These 22 impacts are examined below. In general, the results in NUREG/CR-7203 (NRC, 2015) should 23 not be considered generically applicable. The thermal analyses of the application should 24 consider scenarios discussed in NUREG/CR-7203 to determine consistency in the analytical 25 methods, scenario phenomena, and results. The thermal analyses would be expected to 26 assess the impact of fuel reconfiguration on the fuel cladding and component temperatures and 27 the internal pressure for the particular transportation package design.

28 For Scenario 1(a) of Category 1 (see list of scenarios in Section 4.2.4.2 of this report) from 29 NUREG/CR-7203, the fuel rods are assumed to breach in such a manner that the cladding 30 remains in its nominal geometry (no fuel reconfiguration), but the release of fuel rod backfill gas 31 and fission product gases can cause a change to the package component peak temperatures.

32 For Scenario 1(b) of Category 1, for configurations where an assembly (or assemblies) is 33 represented as a debris pile(s) inside its basket cell, fuel reconfiguration has a larger impact on 34 the component temperatures for the vertical orientation than for the horizontal orientation, but 35 the packing fraction of debris bed has minor impact on the component temperatures. For both 36 Scenario 1(a) and Scenario 1(b), release of the fuel rod gaseous contents increases the number 37 of moles of gas and thus the package container pressure. The canister pressure is expected to 38 increase with the increased fuel rod failure fractions.

39 For Category 2 (Scenarios 2(a) and 2(b)), the fuel rods are assumed to remain intact without 40 gaseous leakage into the canister space. The changes of the fuel assembly lattice (contraction 41 in Scenario 2(a) and expansion in Scenario 2(b)) could cause either an increase or decrease in 42 the package component temperatures depending on the initial assembly geometry and whether 3 Aging management activities may be conducted under the aegis of an NRC-approved AMP (for renewal applications) or a maintenance plan (for initial license or CoC applications requesting approval for periods exceeding 20 years).

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For ACRS Review Purposes Only 1 the package relies on convection for heat transfer. In general, the impact from Scenarios 2(a) 2 and 2(b) is expected to be minor for the package component temperatures and canister 3 pressure.

4 For Category 3, the fuel rods are assumed to remain intact without gaseous leakage into the 5 canister space, but the axial shifting of the assembly changes the heat source location within 6 the packaging. It is expected that changes in assembly axial alignment within the basket cells 7 have minor impact on the component temperatures and canister pressure.

8 Normal Conditions of Transport 9 Based on the thermal phenomena described in Section 4.2.4.2.3 and NUREG/CR-7203 (NRC, 10 2015), an application should evaluate the impact of Scenarios 1(a) and 1(b) of Category 1 on 11 the canister pressure and the fuel cladding and package component temperatures for 3-percent 12 fuel rod failure for NCT thermal evaluation.

13 For Scenarios 2(a) and 2(b) in Category 2 and Scenario 3 in Category 3, although the impact of 14 hypothetical fuel reconfiguration on package thermal performance (e.g., temperature and 15 pressure) is not expected to be significant because the fuel rods are assumed to remain intact 16 without gaseous leakage into the canister space, the applicant may need to provide thermal 17 analyses due to the specifics of the package design.

18 Hypothetical Accident Conditions 19 Based on thermal phenomena described in Section 4.2.4.2.3 and NUREG/CR-7203 (NRC, 20 2015), an application should evaluate the impact of Scenarios 1(a) and 1(b) of Category 1 on 21 the canister pressure and the fuel cladding and package component temperatures for 100 22 percent fuel rod failure for HAC thermal evaluation.

23 For Scenarios 2(a) and 2(b) in Category 2 and Scenario 3 in Category 3, although the impact of 24 fuel reconfiguration on package thermal performance (e.g., temperature and pressure) is not 25 expected to be significant because the fuel rods are assumed to remain intact without gaseous 26 leakage into the canister space, the applicant may need to provide thermal analyses due to 27 specifics of the package design.

28 4.2.4.2.4 Criticality 29 An application may demonstrate that a transportation package meets the regulatory 30 requirements for criticality safety by assuming hypothetical reconfiguration of the HBU SNF into 31 justified bounding geometric forms. If data from a surrogate demonstration program are not 32 available before the shipment of fuel previously dry-stored for periods longer than 20 years, this 33 approach is one way to provide additional assurance of compliance with 10 CFR 71.55, 34 General Requirements for Fissile Material Packages, and 10 CFR 71.59, Standards for 35 Arrays of Fissile Material Packages, during NCT and HAC.

36 To assess the impacts of hypothetical fuel reconfiguration, ORNL performed criticality safety 37 analyses for various scenarios and examined the impacts on the reactivity of a package. The 38 results were described in NUREG/CR-7203 (NRC, 2015) which considers burnup up to 70 39 GWd/MTU for criticality evaluations. The study characterized the assumed hypothetical 40 reconfiguration scenarios were categorized depending on the nature of the assembly damage, 41 as described previously.

4-15

For ACRS Review Purposes Only 1 With respect to criticality safety analyses, NUREG/CR-7203 (NRC, 2015) provides some insight 2 on the reactivity effects of some reconfiguration scenarios; however, the values in the results 3 are not generically applicable. Fuel reconfiguration may have different reactivity effects on a 4 transportation package other than the generic models used in NUREG/CR-7203.

5 Criticality is not a concern for dry SNF transportation packages, as SNF requires moderation to 6 reach critical. The criticality analyses in NUREG/CR-7203 (NRC, 2015) assume fully-flooded 7 conditions, and any conclusions adopted are applicable only to analyses that include moderator 8 intrusion. The staff will follow the guidance in ISG-19, Moderator Exclusion under Hypothetical 9 Accident Conditions and Demonstrating Subcriticality of Spent Fuel under the Requirements of 10 10 CFR 71.55(e), issued in May 2003 (NRC, 2003), to review an application for moderator 11 exclusion. The following considerations for criticality evaluations for reconfigured fuel apply 12 only to transportation packages that do not employ moderator exclusion.

13 All of the criticality safety analyses presented in NUREG/CR-7203 (NRC, 2015) take credit for 14 burned fuel nuclides (burnup credit) and the results may not apply to analyses that assume a 15 fresh fuel composition. To review the staff will follow the guidance in ISG-8, Revision 3, Burnup 16 Credit in the Criticality Safety Analyses of PWR Spent Fuel in Transportation and Storage 17 Casks, issued in September 2012 (NRC, 2012), to review the the burnup credit methodology 18 and code benchmarking used to support a criticality safety evaluation. ISG-8, Revision 3 does 19 not endorse any particular methodology for BWR fuel burnup. The staff does not necessarily 20 endorse the methodology used to perform the study presented in NUREG/CR-7203 for BWR 21 fuel DSS, and considers it to be for illustration only.

22 For criticality safety analyses using burnup credit, NUREG/CR-7203 (NRC, 2015) shows that 23 reactivity increases for longer decay times. Therefore, analyses supporting storage beyond 20 24 years would need to use an appropriate decay time in the criticality evaluations. The 25 enrichment and burnup values assumed in the criticality evaluations in NUREG/CR-7203 may 26 differ from the values allowed in another transportation package. However, NUREG/CR-7203 27 states that no significant differences were observed in trends between configurations that 28 evaluated fuel at 44.25 GWd/MTU and 70 GWd/MTU.

29 The following sections discuss an approach acceptable to the staff for addressing increases in 30 reactivity resulting from the potential reconfiguration for HBU fuel under NCT and HAC. These 31 sections identify the most applicable information from NUREG/CR-7203 to address each of 32 these specific conditions.

33 Normal Conditions of Transport 34 In an approach acceptable to the staff, the applicants criticality safety evaluations would 35 consider the reactivity impact of 3-percent fuel failure under NCT. Based on NUREG/CR-7203 36 (NRC, 2015), the impacts on the package keff resulting from 3-percent fuel failure may become 37 significant. Applicants for transportation packages may need to consider the 3-percent fuel 38 failure for both single package and array analyses under NCT.

39 The scenario most applicable to 3-percent fuel failure under NCT is Category 1, Scenario 1(a) 40 from NUREG/CR-7203. ORNL created this scenario to represent breached rods. ORNL 41 assumed that a percentage of the rods were breached, and that cladding from these rods failed 42 completely. ORNL then removed this percentage of fuel rods from the system. This is 43 conservative as SNF is under moderated and replacing fuel with moderator typically causes 44 reactivity to increase. Using a fresh fuel composition for PWR fuel, NUREG/CR-7203 shows 4-16

For ACRS Review Purposes Only 1 that reactivity decreases when removing rods and therefore this type of analysis may not be 2 appropriate for PWR analyses that assume a fresh fuel composition. The location assumed for 3 failed or removed rods can have a significant effect on reactivity. NUREG/CR-7203 shows in 4 Section A.1.1 that removing rods from the center of the assembly causes reactivity to increase 5 the most.

6 In NUREG/CR-7203 (NRC, 2015), ORNL also determines the number of rods removed that 7 produces the maximum reactivity. For the systems studied in NUREG/CR-7203 the maximum 8 reactivity occurs when more than 3-percent of the rods are removed from the system.

9 NUREG/CR-7203 (NRC, 2015) also presents the results of a sensitivity study that shows 10 increased reactivity for an alternative Category 1, Scenario 1(a), which assumed that the failed 11 fuel relocates to a location outside of the absorber plate. This is based on the generic system 12 modeled in NUREG/CR-7203. A different package may allow relocation of the failed rod 13 material outside of the absorber plate material to a different extent, and an applicant would 14 evaluate an alternative scenario for the specific transportation package being evaluated.

15 Hypothetical Accident Conditions 16 In an approach acceptable to the staff, the applicants criticality safety evaluations would 17 consider the reactivity impact of 100-percent fuel failure under HAC. Based on NUREG/CR-18 7203 (NRC, 2015), the impacts on the package keff resulting from 100-percent fuel failure may 19 be significant. Applicants for transportation packages may need to consider the 100 percent 20 fuel failure for both single package and array analyses under HAC.

21 The applicable scenarios from NUREG/CR-7203 (NRC, 2015) for the hypothetical case of 100 22 percent fuel failure are a combination of Category 1 Scenario 1(b), Category 2 Scenarios, 23 Category 3 Scenarios.

24 In Scenario 1(b) in Section A.1.2 of NUREG/CR-7203 (NRC, 2015), ORNL considered 25 reconfiguration of damaged fuel. With 100-percent compromise in cladding integrity, 26 reconfiguration is considered to the maximum extent. Section A.1.2 of NUREG/CR-7203 shows 27 that a model assuming an ordered pellet array is more reactive than a homogenous mixture of 28 fuel, cladding materials, and water.

29 In Scenario 2 in Section A.2 of NUREG/CR-7203 (NRC, 2015), ORNL considered rod/assembly 30 deformation from side and end impact events. ORNL investigated the effects on birdcaging and 31 bottlenecking by changing the pitch uniformly and non-uniformly. For all pitch contraction 32 cases, ORNL calculated a decrease in keff from the nominal pitch. For the uniform pitch 33 expansion ORNL found that the maximum pitch increase possible within the basket cell resulted 34 in the highest keff. For the non-uniform pitch expansion, ORNL increased the pitch of the inner 35 fuel rods/pins by decreasing the space between the outer rods/pins. The results in NUREG/CR-36 7203 show that non-uniform pitch expansion produces keff values higher than uniform pitch 37 expansion for all cases except the unchanneled BWR fuel.

38 In Scenario 3 in Section A.3 of NUREG/CR-7203 (NRC, 2015), ORNL considered reactivity 39 effects of changes in assembly axial alignment. Neutron absorber panels may not extend the 40 full length of the basket and it may be possible for fuel to reconfigure outside of the neutron 41 absorber panels. ORNL investigated the change in reactivity resulting from the displacement of 42 intact fuel assemblies outside of the neutron absorber panels. NUREG/CR-7203 shows that the 43 maximum reactivity increase results when displacing the assemblies to the maximum extent at 4-17

For ACRS Review Purposes Only 1 the top, versus the bottom, because there is less burnup at the top of the assembly. The 2 amount of displacement possible depends on the particular transportation package and may be 3 different from that of the package(s) analyzed in NUREG/CR-7203. Higher burnup assemblies 4 show the largest change in keff upon displacement; however, the increase in keff caused by the 5 displacement may be bounded by the keff from a non-displaced lower burned assembly.

6 4.2.4.2.5 Shielding 7 An application may demonstrate that a transportation package meets the regulatory 8 requirements for shielding safety by showing that, with reconfiguration of the HBU SNF, the 9 package meets the dose rate limits under NCT and HAC. If a confirmatory demonstration is not 10 applicable or available, this approach is one way to provide additional assurance of compliance 11 with 10 CFR 71.47, External Radiation Standards for All Packages; 10 CFR 71.51(a)(1) for 12 NCT, and 10 CFR 71.51(a)(2) under HAC.

13 To assess the impacts of various fuel geometry changes on the calculated external dose rates 14 of an SNF transportation package, ORNL evaluated the external dose rate for various scenarios 15 of fuel geometry changes and show the results in NUREG/CR-7203 (NRC, 2015) for example 16 BWR and PWR transportation packages.

17 With respect to external dose rate analyses, the results in NUREG/CR-7203 (NRC, 2015) 18 should not be considered generically applicable. The impacts of fuel reconfiguration on the 19 maximum external dose rates may be different based on the package design.

20 Since reconfiguration is to be considered for transportation packages shipped after 20 years of 21 storage, and this length of cooling time is generally much longer than cooling times used to 22 establish loading tables, applicants may be able to justify that increases to external dose 23 resulting from reconfiguration are bounded by the additional cooling time the assemblies will 24 experience. As discussed in Section B.5 of NUREG/CR-7203 (NRC, 2015), based on the decay 25 time different nuclides become important in the evaluations.

26 NUREG/CR-7203 (NRC, 2015) also indicates that fuel assembly type (i.e., PWR vs BWR) may 27 have a significant impact on the external dose rate under fuel reconfiguration scenarios. Tables 28 9-12 of NUREG/CR-7203 shows the difference in dose rate increase for BWR and PWR SNF.

29 In addition, a transportation package may allow transport of other fuel assemblies, with different 30 allowable burnup and enrichments. The burnup profile and depletion parameters used to create 31 the source term within NUREG/CR-7203 may also not be generically applicable. Appendix B to 32 NUREG/CR-7203 presents details of the analyses.

33 The following sections discuss an approach acceptable to the staff for addressing increases in 34 external dose rate resulting from the potential reconfiguration of HBU fuel under NCT and HAC.

35 These sections identify the most applicable information from NUREG/CR-7203 (NRC, 2015) to 36 address each of these specific conditions.

37 Normal Conditions of Transport 38 In an approach acceptable to the staff, the applicants external dose rate evaluations would 39 evaluate the impact of 3-percent fuel failure under NCT. Based on NUREG/CR-7203 (NRC, 40 2015), source relocation resulting from 3 percent fuel failure may have a significant impact on 41 the dose rates prescribed in 10 CFR 71.47(b). The most applicable scenario from NUREG/CR-42 7203 is Category 1 (fuel failure), Scenario, 1(a). The results show that the dose rate changes 4-18

For ACRS Review Purposes Only 1 are sensitive to the number of fuel rod breaches and available space for fuel to move in the 2 cavity.

3 For Category 1 Scenario 1(a), breached rods, ORNL assumed that when the cladding is 4 breached, the rods turn to rubble and calculated the dose rate when the rubbleized fuel mixture 5 relocated within the fuel assembly. ORNL assumed failure of 10 and 25-percent of PWR fuel 6 rods and 11-percent of BWR fuel rods failed. Section B.4.1 of NUREG/CR-7203 (NRC, 2015) 7 discusses the implementation in detail. ORNL reduced the source strength and density of the 8 active fuel zone by the failure percentage, relocated this source to a different part of the fuel 9 assembly and increased the source strength and density accordingly. ORNL calculated 10 external dose rates using models with the fuel rubble mixture relocated to varied locations of the 11 package (top, middle, bottom). The limiting location for the relocated fuel rubble would be 12 based on the characteristics of the transportation package being analyzed.

13 Hypothetical Accident Conditions 14 In an approach acceptable to the staff, the applicants external dose rate evaluations would 15 consider the impact of 100-percent fuel failure under HAC. The applicable scenarios from 16 NUREG/CR-7203 (NRC, 2015) are Category 1 Scenarios, Category 2 Scenarios and Category 17 3 Scenarios. ORNL assumed that there was no neutron shield present for the HAC models.

18 This is a typical assumption in HAC dose rate evaluations as it is difficult to predict the condition 19 of the neutron shield after the HAC fire event. Therefore source terms with high neutron 20 radiation, such as HBU fuel, tend to be limiting for HAC.

21 Based on NUREG/CR-7203 (NRC, 2015), source relocation resulting from 100-percent fuel 22 failure can have a significant impact on external dose rates under HAC. Tables 11 and 12 of 23 NUREG/CR-7203 show the relative changes for the example packages under HAC. These 24 dose rate change ratios are for dose rates at 1 m from the package as required by 10 CFR 25 71.51(a)(2).

26 For Category 1 Scenarios (cladding failure), ORNL assumed in the analyses in NUREG/CR-27 7203 (NRC, 2015) that when the cladding fails the rods turn to rubble, and created a model with 28 a homogenized fuel and basket material. ORNL determined that the limiting mass packing 29 fraction for rubbleized fuel and basket material is 0.58. When evaluating dose rates for a 30 package in the vertical orientation, the damaged fuel model from the Category 1 Scenario 1(b) 31 in NUREG/CR-7203 is applicable. For a package in a horizontal orientation, the Category 2 32 Scenario from NUREG/CR-7203 would be more applicable. In this scenario, ORNL analyzed 33 the dose rates when the fuel is kept within its respective basket cell but pushed to the side walls 34 as shown in Section B.4.2 of NUREG/CR-7203. The limiting scenarios for any given 35 transportation package would depend on the specific characteristics of that package.

36 In the Category 3 Scenario in NUREG/CR-7203 (NRC, 2015), ORNL evaluates the dose rate 37 increase when an intact fuel assembly is pushed to the bottom or top of the package, thus 38 increasing dose rates at the bottom or top, or radially if the source becomes aligned with an 39 area of the package where there is streaming. The results from NUREG/CR-7203 generally 40 show a smaller increase in dose rates for this scenario than for Category 1 Scenarios and the 41 Category 2 Scenario and are likely to be bounded by the results for those situations. However, 42 there may be specific features from a particular package that may cause this scenario to be 43 worth considering.

4-19

For ACRS Review Purposes Only 1 4.3 Canned Fuel 2 HBU SNF that has been classified as damaged should be placed in a can designed for 3 damaged fuel or in an acceptable alternative. The staff will follow the guidance in the current 4 SRP for transportation of SNF when reviewing an application for a transportation package with 5 damaged HBU SNF contents.

4-20

For ACRS Review Purposes Only 1 5 CONCLUSIONS 2 The information in this report provides technical background information on the mechanical 3 performance of high burnup (HBU) spent nuclear fuel (SNF) after drying operations for storage 4 and transportation. The report also provides an engineering assessment of the test results for 5 HBU SNF discussed in NUREG/CR-7198, Revision 1 , Mechanical Fatigue Testing of 6 High-Burnup Fuel for Transportation Applications, issued October 2017 (NRC, 2017a), and 7 proposes example approaches for licensing and certification of HBU SNF for dry storage (under 8 Title 10 of the Code of Federal Regulations (10 CFR) Part 72, Licensing Requirements for the 9 Independent Storage of Spent Nuclear Fuel and High-Level Radioactive Waste, and Reactor-10 Related Greater Than Class C Waste, and transportation (under 10 CFR Part 71, Packaging 11 and Transportation of Radioactive Material) based on the engineering assessment.

12 Until recently, experimental testing on the structural behavior of SNF rods during transportation 13 and storage has focused primarily on obtaining mechanical properties that consider only the 14 material strength of the cladding. Historically, the fuel pellets contribution to the flexural rigidity 15 and structural response of the fuel rod during normal and accident conditions has been ignored 16 because of the lack of experimental bending test data. Recent research sponsored by the U.S.

17 Nuclear Regulatory Commission (NRC) on the static bending response and fatigue strength of 18 HBU SNF rods (i.e., rods with burnup exceeding 45 GWd/MTU) with the presence of the fuel 19 pellets, has provided some of the data necessary to more accurately assess the structural 20 behavior of the composite HBU SNF rod system (NRC, 2017a). The staff has examined the 21 results from this research to assess the expected behavior of HBU SNF under normal 22 conditions of transport (NCT) and hypothetical accident conditions (HAC), as well as DSS drop 23 and tip-over accident scenarios.

24 The results in NUREG/CR-7198, Revision 1 (NRC, 2017a) for static bend testing of 25 aggressively hydride-reoriented Zircaloy-4 HBU SNF rods supports the staffs conclusion that 26 the use of best-estimate cladding mechanical properties that do not account for the presence of 27 the fuel pellet continues to be adequate for assessing the structural performance of HBU SNF 28 rods during a hypothetical 9-m (30-ft) drop accident, per the requirement in 10 CFR 71.73(c)(1).

29 The same conclusion applies to the lower loads experienced during a 0.3-m (1-ft) drop, per the 30 requirement in 10 CFR 71.71(c)(7), and postulated drop and cask tip-over accident scenarios 31 during dry storage operations, per the requirement in 10 CFR 72.122(b). Further, the staff finds 32 that the orientation of the hydrides is not a critical consideration when evaluating the adequacy of 33 cladding-only mechanical properties. Therefore, the use of mechanical properties for cladding in 34 either the as-irradiated or hydride-reoriented condition is considered acceptable for the 35 evaluation of drop and cask tip-over accident scenarios. If an applicant is unable to demonstrate 36 satisfactory performance of the HBU SNF rod by assuming cladding-only mechanical properties, 37 the staff has proposed an alternative approach for using the results from static bend testing to 38 account for the increased flexural rigidity imparted by the fuel pellet.

39 After considering the aggressive hydride reorientation treatment used for the Zircaloy-4 HBU 40 SNF rods, the staff concludes that the same response is expected for all modern commercial 41 cladding alloy types that may experience hydride reorientation (i.e., Zircaloy-2, ZIRLO' and 42 M5). The staff has also reviewed proprietary and non-proprietary data on end-of-life rod internal 43 pressures for fuel rods with boron-based integral fuel burnable absorbers (see Section 1.5.3) 44 and considers these rods to be reasonably bound by the maximum rod internal pressure used in 45 the radial hydride treatment of the Zircaloy-4 HBU SNF rods. The staffs expectation is that 46 additional static bend testing and fatigue testing of HBU SNF composite rods with other 47 claddings will provide confirmation of this conclusion. The U.S. Department of Energy is 5-1

For ACRS Review Purposes Only 1 currently planning to conduct these tests, which the NRC will evaluate when available (Hanson 2 et al., 2016).

3 In addition, the results in NUREG/CR-7198, Revision 1 (NRC, 2017a), on the fatigue testing of 4 aggressively hydride-reoriented Zircaloy-4 HBU SNF rods have provided an adequate technical 5 basis for establishing a reasonable lower-bound fatigue curve and endurance limit for tensile 6 axial-bending loads experienced during transport. Therefore, the staff finds that applicants can 7 use a cumulative damage approach and the curve mentioned above in support of their structural 8 evaluation to assess vibration normally incident to transport of Zircaloy-4 HBU SNF, per the 9 requirement in 10 CFR 71.71(c)(5). Fatigue test data for other cladding alloy types would be 10 needed to develop their respective lower-bound fatigue curves and endurance limits. The U.S.

11 Department of Energy is currently planning to conduct additional fatigue strength testing of HBU 12 SNF composite rods with other claddings, which will provide the necessary data to develop 13 those curves and define the respective endurance limits (Hanson et al., 2016).

14 This report also presents example licensing and certification approaches for HBU SNF to 15 address age-related uncertainties associated with conclusions based on accelerated separate-16 effects testing. One of these approaches, the use of a surveillance and monitoring program for 17 confirmation of design basis HBU SNF configuration, is consistent with the guidance in NUREG-18 1927, Revision 1, Standard Review Plan for Renewal of Specific Licenses and Certificates of 19 Compliance for Dry Storage of Spent Nuclear Fuel, issued June 2016 (NRC, 2016b).

20 Alternatively, the staff has proposed an example approach based on demonstrating compliance 21 with the pertinent regulatory requirements even if hypothetical reconfiguration of the design 22 basis fuel were to occur. This example approach considers lessons learned from an NRC-23 sponsored generic consequence assessment for transportation packages, as discussed in 24 NUREG/CR-7203, A Quantitative Impact Assessment of Hypothetical Spent Fuel 25 Reconfiguration in Spent Fuel Storage Casks and Transportation Packages, issued 26 September 2015 (NRC, 2015).

5-2

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6-1

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28 NRC. 2000a. "Standard Review Plan for Spent Fuel Storage Facilities NUREG-1567, 29 Washington D.C. ADAMS Accession No. ML003686776.

30 NRC. 2000b. "Standard Review Plan for Transportation Packages for Spent Nuclear Fuel, 31 NUREG-1617, Washington D.C. ADAMS Accession No. ML003696262.

32 NRC. 2000c. "Reexamination of Spent Fuel Shipment Risk Estimates, NUREG/CR-6672, 33 SAND2000-0234, Washington D.C. ADAMS Accession No. ML003698324.

34 NRC. 2001. Dry Cask Storage Characterization ProjectPhase 1; CASTOR V/21 Cask 35 Opening and Examination. NUREG/CR-6745, Washington D.C. ADAMS Accession No.

36 ML013020363.

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For ACRS Review Purposes Only 1 NRC. 2003a. Cladding Considerations for the Transportation and Storage of Spent Fuel, 2 Interim Staff Guidance 11, Revision 3, Washington, DC. ADAMS Accession No. ML033230335.

3 NRC. 2003b. Examination of Spent PWR Fuel Rods after 15 Years in Dry Storage.

4 NUREG/CR-6831, Washington D.C. ADAMS Accession No. ML032731021.

5 NRC. 2003c. Moderator Exclusion under Hypothetical Accident Conditions and 6 Demonstrating Subcriticality of Spent Fuel under the Requirements of 10 CFR 71.55(e), Interim 7 Staff Guidance 19, Washington, DC. ADAMS Accession No. ML031250639.

8 NRC. 2006. Potential Rod Splitting Due to Exposure to an Oxidizing Atmosphere During 9 Short-term Cask Loading Operations in LWR or Other Uranium Oxide Based Fuel, Interim Staff 10 Guidance 22, Washington, DC. ADAMS Accession No. ML061170217.

11 NRC. 2007a. A Pilot Probabilistic Risk Assessment of a Dry Cask Storage System at a 12 Nuclear Power Plant, NUREG-1864, Washington D.C. ADAMS Accession No. ML071340012.

13 NRC. 2007b. Classifying the Condition of Spent Nuclear Fuel for Interim Storage and 14 Transportation Based on Function, Interim Staff Guidance 2, Revision 1, Washington, D.C.

15 ADAMS Accession No. ML071420268.

16 NRC. 2010. Standard Review Plan for Spent Fuel Dry Storage Systems at a General License 17 Facility, Revision 1. NUREG-1536, Revision 1, Washington D.C. ADAMS Accession No.

18 ML101040620.

19 NRC. 2011. FRAPCON-3.4: A Computer Code for the Calculation of Steady-State Thermal-20 Mechanical Behavior of Oxide Fuel Rods for High Burnup, NUREG/CR-7022, Vol. 1, 21 Washington D.C. ADAMS Accession No. ML11101A005.

22 NRC. 2012. Burnup Credit in the Criticality Safety Analyses of PWR Spent Fuel in 23 Transportation and Storage Casks, Interim Staff Guidance 8, Revision 3, Washington, D.C.

24 ADAMS Accession No. ML122261A433.

25 NRC. 2014. Spent Fuel Transportation Risk Assessment - Final Report. NUREG-2125, 26 Washington D.C. ADAMS Accession No. ML14031A323.

27 NRC. 2015. A Quantitative Impact Assessment of Hypothetical Spent Fuel Reconfiguration in 28 Spent Fuel Storage Casks and Transportation Packages. NUREG/CR-7203, Washington D.C.

29 ADAMS Accession No. ML15266A413.

30 NRC. 2016a. Fuel Retrievability in Spent Fuel Storage Applications, Interim Staff Guidance 2, 31 Revision 2, Washington, D.C. ADAMS Accession No. ML16117A080.

32 NRC. 2016b. Standard Review Plan for Renewal of Specific Licenses and Certificates of 33 Compliance for Dry Storage of Spent Nuclear Fuel, NUREG-1927, Revision 1, Washington, 34 DC. ADAMS Accession No. ML16179A148.

35 NRC. 2017a. Mechanical Fatigue Testing of High-Burnup Fuel for Transportation 36 ApplicationApplications. NUREG/CR-7198, Revision 1, Washington DC. ADAMS Accession 37 No. ML17292B057.

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For ACRS Review Purposes Only 1 NRC. 2019. Managing Aging Processes in Storage (MAPS) Report. NUREG-2214, 2 Washington DC. ADAMS Accession No. ML19214A111.

3 Pan, G., A.M. Garde, and A.R. Atwood. 2013. Performance and Property Evaluation of High-4 Burn-up Optimized ZIRLO' Cladding, Proc. 17th ASTM Sym. on Zirconium in the Nuclear 5 Industry, Feb. 3-7, 2013, Hyderabad, India.

6 Patterson, C. and F. Garzarolli. 2015. Dry Storage Handbook: Fuel Performance in Dry 7 Storage, A.N.T. International: Mlnlycke, Sweden.

8 Peehs, M. 1998. Assessment of Dry Storage Performance of Spent LWR Fuel Assemblies 9 with Increasing Burnup Proc 1st RCM meeting, Washington, DC., April 20-24, 1998.

10 Pickard, A. 2015. Fatigue Crack Propagation in Biaxial Stress Fields," J Strain Analysis for 11 Engineering Design. 50(1): 25-39.

12 Richmond, D. J. and K. J. Geelhood.. 2018. FRAPCON Analysis of Cladding Performance 13 during Dry Storage Operations. Pacific Northwest National Laboratory, PNNL--27418, April 14 2018.

15 Roark, R. and W. Young, "Formulas for Stress and Strain," McGraw-Hill, 5th Edition, 1975.

16 Sanders, T., Seager, K., Rashid, Y., Barrett, P., Malinauskas, A., Einziger, R., Jordan, H.,

17 Duffey, T., Sutherland, S., Reardon, P., 1992, A Method for Determining the Spent Fuel 18 Contribution to Transport Cask Containment Requirements, SAND90-2406, Sandia National 19 Laboratories, Albuquerque, New Mexico.

20 Spino, J., M. Coquerelle and D. Baron. 1996. Microstructure and Fracture Toughness 21 Characterization of Irradiated PWR Fuels in the Burnup Range of 40-67 GWd/MTU.

22 Proceedings of the Technical Committee Meeting IAEA on Advances in Fuel Technology.

23 Spino, J., J. Cobos-Sabate and F. Rousseau. 2003. Room-temperature Microindentation 24 Behavior of LWR-fuels, Part 1: fuel microhardness. J. Nucl. Mater. 322:204-216.

25 Tang, D., A. Rigato, and R.E. Einziger. 2015. Flaw Effects and Flaw Reorientation on Spent 26 Fuel Rod Performance, a Simulation with Finite Element Analysis Proceedings of the ASME 27 2015 Pressure Vessels and Piping Conference, July 19-23, 2015, Boston, MA, USA.

28 Wang, J.-A., H. Wang, H. Jiang, Y. Yan, B. B. Bevard, J. M. Scaglione. 2016. FY 2016 Status 29 Report: Documentation of All CIRFT Data including Hydride Reorientation Tests. Oak Ridge 30 National Laboratory, ORNL/SR-2016/424, September 14, 2016.

31 Winter, G. and A. Nelson. 1979. "Design of Concrete Structures," McGraw-Hill, 9th Edition, 32 1979.

33 Wisner, S. and R. Adamson. Combined Effects of Radiation Damage and Hydrides on the 34 Ductility of Zircaloy-2. Nuclear Engineering and Design. Vol. 185. pp. 33-49. 1998.

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For ACRS Review Purposes Only 1 NUREG-2224 2

NUREG 3

Dry Storage and Transportation of High Burnup Spent Fuel - Draft for Comment July 2018 Type Title and Subtile here February 2016 T. Ahn, H. Akhavannik, G. Bjorkman, F.C. Chang, W. Reed, A. Rigato, Select & delete if n/a Technical D. Tang, R.D. Torres, B.H. White, V. Wilson Please list Authors here. Technical dates Division of Spent Fuel Management Office of Nuclear Material Safety and Safeguards U.S. Nuclear Regulatory Commission Washington, DC 20555-0001 Please type Organization and address here.

Street Same as above Division of

. Office of The potential Please for changes inhere.

type supplementary the cladding performance of high burnup (HBU) spent nuclear fuel (SNF) to compromise the analyzed fuel configuration in dry storage systems and transportation packages has been historically addressed through safety review guidance. The guidance defines adequate fuel conditions, Please typepeak including or cutcladding and paste Abstract information temperatures during here. Pleaseloading short-term note that this sectiontocant operations be over prevent 200 words.

or mitigate degradation of the cladding. The purpose of this report is to expand the technical basis in support of that guidance, as it pertains to the mechanism of hydride reorientation in HBU SNF cladding.

This report also provides an engineering assessment of the results of NRC-sponsored research on the mechanical performance of HBU SNF following hydride reorientation and, per the conclusions of that assessment, provides example approaches for licensing and certification of HBU SNF for dry storage (under 10 CFR Part 72) and transportation (under 10 CFR Part 71).

High Burnup Spent Fuel Dry Storage Systems Transportation Key Packages words/descriptors Hydride Reorientation Please leave blank

For ACRS Review Purposes Only For ACRS Review Purposes Only For ACRS Review Purposes Only For ACRS Review Purposes Only