NUREG- 1864

A Pilot Probabilistic

Risk Assessment

Of a Dry Cask Storage System

At a Nuclear Power Plant

U.S. Nuclear Regulatory Commission

Office of Nuclear Regulatory Research

Washington, DC 20555-0001
AVAILABILITY OF REFERENCE MATERIALS

IN NRC PUBLICATIONS

NRC Reference Material

As of November 1999, you may electronically access

NUREG-series publications and other NRC records at

NRC's Public Electronic Reading Room at

http://www.nrc.qov/readinq-rm.html.

Publicly released records include, to name a few,

NUREG-series publications; Federal Register notices;

applicant, licensee, and vendor documents and

correspondence; NRC correspondence and internal

memoranda; bulletins and information notices;

inspection and investigative reports; licensee event

reports; and Commission papers and their

attachments.

NRC publications in the NUREG series, NRC

regulations, and Title 10, Energy, in the Code of

Federal Regulations may also be purchased from one

of these two sources.

1. The Superintendent of Documents

U.S. Government Printing Office

Mail Stop SSOP

Washington, DC 20402-0001

Internet: bookstore.gpo.gov

Telephone: 202-512-1800

Fax: 202-512-2250

2. The National Technical Information Service

Springfield, VA 22161-0002

www.ntis.gov

1-800-553-6847 or, locally, 703-605-6000

Non-NRC Reference Material

Documents available from public and special technical

libraries include all open literature items, such as

books, journal articles, and transactions, Federal

Register notices, Federal and State legislation, and

congressional reports. Such documents as theses,

dissertations, foreign reports and translations, and

non-NRC conference proceedings may be purchased

from their sponsoring organization.

Copies of industry codes and standards used in a

substantive manner in the NRC regulatory process are

maintained atThe NRC Technical Library

Two White Flint North

11545 Rockville Pike

Rockville, MD 20852-2738

These standards are available in the library for

reference use by the public. Codes and standards are

usually copyrighted and may be purchased from the

originating organization or, if they are American

National Standards, fromAmerican National Standards Institute

11 West 4 2nd Street

New York, NY 10036-8002

www.ansi.org

212-642-4900

I

A single copy of each NRC draft report for comment is

available free, to the extent of supply, upon written

request as follows:

Address: U.S. Nuclear Regulatory Commission

Office of Administration

Mail, Distribution and Messenger Team

Washington, DC 20555-0001

E-mail: DISTRIBUTION@,nrc.,ov

Facsimile: 301-415-2289

Some publications in the NUREG series that are

posted at NRC's Web site address

htto:l/www.nrc.aov/readina-rmldoc-collectionslnureas

Legally binding regulatory requirements are stated

only in laws; NRC regulations; licenses, including

technical specifications; or orders, not in

NUREG-series publications. The views expressed

in contractor-prepared publications in this series

are not necessarily those of the NRC.

The NUREG series comprises (1) technical and

administrative reports and books prepared by the

staff (NUREG-XXXX) or agency contractors

(NUREG/CR-XXXX), (2) proceedings of

conferences (NUREG/CP-XXXX), (3) reports

resulting from international agreements

(NUREG/IA-XXXX), (4) brochures

(NUREG/BR-XXXX), and (5) compilations of legal

decisions and orders of the Commission and

Atomic and Safety Licensing Boards and of

Directors' decisions under Section 2.206 of NRC's

regulations (NUREG-0750).

are updated periodically and may differ from the last

printed version. Although references to material found

on a Web site bear the date the material was

accessed, the material available on the date cited may

subsequently be removed from the site.
NUJREG-1864

A Pilot Probabilistic

Risk Assessment

Of a Dry Cask Storage System

At a Nuclear Power Plant

Manuscript Completed: June 2006

Date Published: March 2007

A. Malliakos, NRC Project Manager

Division of Risk Assessment and Special Projects

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Spent Fuel Project Office

Office of Nuclear Material Safety and Safeguards

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001
NUREG-1864, has been reproduced from

the best available copy.
ABSTRACT

In response to a request from the U.S. Nuclear Regulatory Commission (NRC), Office of Nuclear Material

Safety and Safeguards (NMSS), the Office of Nuclear Regulatory Research (RES) and the NMSS

Spent Fuel Project Office (SFPO) have jointly developed and applied a methodology for performing

a pilot probabilistic risk assessment (PRA) of a dry cask storage system at a nuclear power plant site

(i.e., an independent spent fuel storage installation). This RES/NMSS report documents the pilot PRA

for a specific dry cask system (Holtec International HI-STORM 100) at a specific boiling-water reactor

(BWR) site. The methodology developed can serve as a guide for performing similar PRAs in the future.

The pilot study can provide guidance for assessing the risk to the public and identifying the dominant

contributors to that risk. The cask system consists of a multipurpose canister (MPC) that confines the fuel,

a transfer overpack that shields workers from radiation while the cask is being prepared for storage,

and a storage overpack that shields people from radiation and mechanically protects the MPC during

storage. The study covers various phases of the dry cask storage process, from loading fuel from

the spent fuel pool, preparing the cask for storage and transferring it outside the reactor building, moving

the cask from the reactor building to the storage pad, and storing the cask for 20 years on the storage pad.

The study develops and assesses a comprehensive list of initiating events, including dropping the cask

during handling and external events during onsite storage (such as earthquakes, floods, high winds,

lightning strikes, accidental aircraft crashes, and pipeline explosions). Potential cask failures from

mechanical and thermal loads are modeled. The study estimates the annual risk for one cask in terms of

the individual probability of a prompt fatality within 1.6 km (1 mile) and a latent cancer fatality within

16 km (10 miles) of the site.

iii

FOREWORD

The spent fuel pools of commercial nuclear power plants are becoming filled with spent fuel assemblies

and many utilities have been removing older fuel and storing it onsite in dry casks. For this purpose, the

fuel first is loaded from the spent fuel pool into a cask which is then prepared for storage. Then the cask

is transferred to and stored on the storage pad. Given the increasing trend to onsite storage, it is

necessary to assess the risk to the public and to identify the dominant contributors to that risk from this

operation. For this reason, the U.S. Nuclear Regulatory Commission (NRC), Office of Nuclear

Regulatory Research (RES), has developed and applied a methodology for performing a pilot

probabilistic risk assessment (PRA) of a dry cask storage system. On that basis, RES and the NRC's

Office of Nuclear Material Safety and Safeguards (NMSS) have jointly performed the pilot PRA as

documented in this report, for a specific cask system at a specific boiling-water reactor site.

To analyze the risk in this report, the staff developed a comprehensive list of initiating events, and

evaluated the risk associated with each initiating event. Initiating events considered include dropping the

cask inside the secondary containment building during transfer operations, as well as external events

during onsite storage (such as earthquakes, floods, high winds, lightning strikes, accidental aircraft

crashes, and pipeline explosions). Potential cask failures from mechanical and thermal loads, including

thermal loads caused by mis-loading events, were modeled. In the event of a cask failure/breach, the fuel

inventory available for release was based on 10-year-old fuel. Weather conditions and the population

distribution in the vicinity of the site were also considered.

The risk to the public is measured in terms of the individual probabilities of a prompt fatality within 1.6

km (1 mi) and a latent cancer fatality within 16 km (10 mi) of the site. No prompt fatalities are expected.

The resulting calculated risk is extremely small, with an individual probability of a latent cancer fatality

of 1.8x 10- 2 during the first year of service, and 3.2x 1014 per year during subsequent years of storage.

The methodology developed in this study can be used as a guide for performing other similar PRAs.

Moreover, the results of this study can be used in conjunction with the methodology selected to

determine the need for other PRAs, improvements in data gathering and analysis, and additional

engineering design analysis. It should be noted that the focus of this pilot study was solely on the

methodology and its limited (i.e., case-specific) application. Thus, no inferences or conclusions should be

drawn with regard to the study's regulatory implications.

Bri W. Sheron, irecto

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

v

CONTENTS

Page

ABSTRACT....................................................................1iii

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

EXECUTIVE SUMMARY ........................................................ xi

ACKNOWLEDGMENTS ........................................................ xiii

ABBRIEVIIATIONS ............................................................. xv

SYMBOLS................................................................... xvii

GLOSSARY OF SELECTED TERMS.............................................. xix

1. INTRODUCTION ........................................................ 1-I

1.1 Purpose.................................... ... 1-1

1.2 Scope ............................................................ 1-1

1.2.1 Issues Within the Scope ........................................ 1-1

1.2.2 Issues Beyond the Scope......................................... 1-2

1.3 Overview of the Risk Calculations....................................... 1-5

1.4 Elements of the PRA................................................. 1-7

2. DRY CASK AND SECONDARY CONTAINMENT SYSTEMS.....................2-1

2.1 Dry Cask System.................................................... 2-1

2.2 Stages of the Dry Cask Storage Operation................................. 2-2

2.3 Secondary Containment Isolation System.................................. 2-9

3. INITIATING EVENTS.................................................... 3-1

3.1 List of Initiating Events............................................... 3-1

3.2 Initiating Events That Cannot Affect the Subject Plant ....................... 3-6

3.2.1 Floods ..................................................... 3-6

3.2.2 Tsunamis ................................................... 3-7

3.2.3 Volcanic Activity............................................. 3-7

3.2.4 Intense Precipitation........................................... 3-7

3.2.5 Storage Tanks, Transformers, Barges, Trucks, Railcars,

and Nearby Industrial Facilities................................... 3-8

3.3 Frequencies of Initiating Events ........................................ 3-8

3.3.1 Dropped Fuel Assembly ........................................ 3-8

3.3.2 Dropped Transfer Cask......................................... 3-9

3.3.3 Seismic Events.............................................. 3-11

3.3.4 High Winds................................................. 3-13

3.3.5 Meteorites.................................................. 3-14

3.3.6 Lightning Strikes............................................. 3-15

vii
3.3 .7 A ircraft ...................................................... 3-16

4. MULTIPURPOSE CANISTER AND FUEL ..................................... 4-1

4.1 M echanical Loads ..................................................... 4-1

4.1.1 M echanical Load M odels ......................................... 4-1

4.1.2 Response of the Transfer Cask to Loads ............................. 4-2

4.1.3 Response of the Storage Cask to Mechanical Loads .................... 4-6

4.2 Therm al Loads ........................................................ 4-9

4.2.1 Heating During Normal Operation, Blocked Vents, and Fires ............. 4-9

4.2.2 Heating During Lightning Strikes .................................. 4-16

4.2.3 Response of the Fuel Rods to Thermal Loads ........................ 4-18

4.3 M PC Integrity ....................................................... 4-19

4.3.1 Integrity M odels ............................................... 4-19

4.3.2 Probability of M PC Failure ...................................... 4-23

4.4 Fuel Failure ......................................................... 4-25

4.4.1 Fuel Integrity M odel ............................................ 4-25

4.4.2 Fuel Rod Cladding Failure ....................................... 4-25

5. SECONDARY CONTAINMENT ISOLATION .................................. 5-1

5.1 Containment Isolation Reliability M odel ................................... 5-1

5.2 Probability of Secondary Containment Isolation Failure ....................... 5-3

6. CONSEQUENCES .................................................... 6-1

6.1 Source T erm ............................. ............................ 6-1

6.1.1 Source Term Estim ates ........................................... 6-2

6.2 C onsequences ......................................................... 6-3

6.2.1 Consequence M odel ............................................. 6-3

6.2.2 Consequence M easures ........................................... 6-5

7. RISK A SSESSM EN T ........................................................ 7-1

7.1 R isk M odels .......................................................... 7-1

7.2 R isk R esults .......................................................... 7-4

8. OTHER ASPECTS OF DRY CASK STORAGE RISK ............................ 8-1

8.1 M is-Loading of Spent Fuel into the MPC ................................... 8-1

8.2 Potential for Criticality ................................................. 8-2

8.3 C orrosion ............................................................ 8-4

8.4 Pressurization ......................................................... 8-4

8.5 Hydrogen Generation in the M PC .............. ........................... 8-4

9. C O N C L U SIO N S ............................................................ 9-1

10. R E FERE N C E S ............................................................ 10-1

viii
Appendices

Pacie

A . M echanical Loads .......................................................... A - 1

B. Cask Response to Mechanical and Thermal Loads ................................ B - 1

C . Fuel Response to Loads ..................................................... C - 1

D. Methodology to Determine Source Term for a Dry Storage Cask ..................... D - 1

E . Consequence A nalysis ...................................................... E - 1

F. Q uestions and A nsw ers ....................................................... F - 1

Figures

Paqe

1. Structure of the Risk Calculations and Organization of the Report ..................... 1-6

2. M PC and Transfer Overpack .................................................. 2-1

3. M PC and Storage Overpack ................................................... 2-2

4. Movement of the Transfer Cask and Storage Cask

Through the Stages of the Dry Cask Storage Operation .............................. 2-5

5. Vertical View of the Equipment Hatch and Detail of the Ground Floor ................. 2-6

6. Lift Yoke with Stays Attached (A) Prior to and (B) after Engaging the Yoke Arms

on the T runnions ............................................................ 2-7

7. Replacing the Pool Lid with the Transfer Lid ..................................... 2-8

8. Annual Probability of Exceedance as a Function of Ground Acceleration Due to a

Seismic Event at the Subject Plant .............................................. 3-12

9. Probability of Tornado W inds Exceeding Speeds .................................. 3-13

10. Geometrical Configuration and the Computational Grid Generated for the Storage

C a sk . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4 -10

11. Centerline Fuel Temperature .................................................. 4-12

12. Storage Overpack Concrete Temperature During Normal Operation ................... 4-13

13. Storage Overpack Concrete Temperature with Bottom Vents Blocked .................. 4-14

14. Temperature of Storage Cask after 3-hour Fire with an Energy of 82.5 MW ............. 4-15

15. Maximum MPC Shell Temperature in a Fire Scenario .............................. 4-16

16. W elds of the M PC ........................................................... 4-19

17. Secondary Containment HVAC M odel .......................................... 5-1

18. Secondary Containment Isolation Event Tree ..................................... 5-2

19. Upper 95' Percentile Confidence Interval Eigenvalue for MPC-68

in Pure W ater at 27 'C (81 'F) ................................................. 8-3

ix
Tables

Page

1. Stages of the Dry Cask Storage Operation ........................................ 2-4

2. Detectors in the Secondary Containment System ................................... 2-10

3. Volumetric Flow Rates of Secondary Containment Ventilation ....................... 2-11

4. Initiating Events During the Handling Phase

and the Sections Where Frequency, MPC Failure, and Fuel Failure Are Discussed ........ 3-2

5. Initiating Events During the Transfer Phase and

Section Where Frequency, MPC Failure, and Fuel Failure Are Discussed .............. 3-2

6. Initiating Events During the Storage Phase and

Section Where Frequency, MPC Failure, and Fuel Failure Are Discussed ............... 3-3

7. Very Heavy Load Lifts at Representative Sites in 1968 - 2002 ........................ 3-10

8. Probability of Exceeding Selected Tornado Wind Speeds (Reference 18) ............... 3-13

9. Airfields near the Subject Plant Site ............................................. 3-16

10. Storage Cask and ISFSI Parameters for Load and Stress Calculations .................. 4-2

11: Comparison of Results for Normal Storage Conditions .............................. 4-12

12. Summary of MPC Failure Probabilities for Various Mechanical Impact Loads ........... 4-24

13. Probability of Fuel Cladding Failure for Various Drop Scenarios ...................... 4-26

14. Secondary Containment Isolation Failure Types ................................... 5-4

15. Sequences Where the Secondary Containment Fails to Isolate ...................... 5-5

16. Release Fraction Values for a Transfer Cask Drop of 30.5 Meters (100 Feet) ............ 6-3

17. M ajor Inputs for Consequence Calculations ....................................... 6-4

18. MACCS2 Consequence Calculation Results ...................................... 6-4

19. Summary Table of Dry Cask Storage Risk Calculations ............................. 7-6

20. Conditions Potentially Affecting MPC Response to Initiating Events ................... 8-1

x
EXECUTIVE SUMMARY

The spent fuel pools of commercial nuclear power plants are becoming filled with spent fuel assemblies.

To avoid having to cease operations when the pools are full, many utilities have been removing older fuel

from the pools and storing it onsite in dry casks. In this study, sponsored by the U.S. Nuclear Regulatory

Commission (NRC), the Office of Nuclear Regulatory Research (RES) and the Spent Fuel Project Office

(SFPO) in the Office of Nuclear Material Safety and Safeguards (NMSS) have jointly developed

a methodology for performing a probabilistic risk assessment (PRA) of a dry cask storage system

and applied the methodology by performing a pilot PRA of a specific cask system at a specific boilingwater reactor (BWR) site. Although the study results do not necessarily apply to other cask systems

or sites, the methodology can serve as a guide for performing other such PRAs. Nonetheless, it should be

noted that the focus of this pilot study is on the methodology and its limited (i.e., case-specific)

application. Consequently, no inferences or conclusions should be drawn with regard to the regulatory

implications of this study.

The pilot PRA assesses the risk to the public and identifies the dominant contributors to risk associated

with dry cask storage involving a single cask at a specific BWR site. Among the items that were beyond the

scope of the study were subsequent versions of the specific cask studied in this report, unloading of the

cask, offsite transportation, repository storage, uncertainty analysis, worker risk, human reliability,

fabrication errors, mis-loading of spent nuclear fuel, aging effects, and combinations of factors that could

impact the probability of MPC failure.

The cask system analyzed is the Holtec International HI-STORM 100. This cask system, which is used

for onsite storage only, consists of three major components, including a multipurpose canister (MPC),

a transfer overpack, and a storage overpack. The MPC confines spent fuel assemblies for the duration

of the storage period (typically 20 years). While the MPC is loaded with spent fuel assemblies from the

spent fuel pool and is prepared for storage, it is inside the transfer overpack to shield workers from

radiation. During onsite storage, the storage overpack shields people from radiation, provides a path to cool

the MPC by convection though vents, and protects the MPC during storage. In this study, the transfer

overpack containing the MPC is referred to as the transfer cask, and the storage overpack containing the

MPC is referred to as the storage cask.

The dry cask storage operation is divided into the three phases of handling, transfer, and storage.

During the handling phase, spent fuel assemblies are placed into the MPC within a transfer overpack.

After the MPC is dried, inerted with helium, and sealed, it is removed from the transfer cask and placed

in the storage overpack. The transfer phase takes place when the storage cask is moved out of

the secondary containment to the storage pad. The storage phase takes place after the storage cask

is set down on the storage pad. To analyze the risk of the dry cask storage operation, the phases are

divided into 34 stages. A stage is defined as a discrete part of the operation that is convenient for

the risk analysis. For example, because the height at which the cask is moved and the composition

of the surface over which it is moved determine the mechanical load on the MPC if it were dropped,

these aspects of the operation are used to define some of the stages.

xi
A comprehensive list of initiating events is developed, and the risk associated with each initiating event

is evaluated. Analyses are performed to determine the cask's response to the mechanical and thermal

loads imposed by the initiating events. For example, the following analyses evaluate the consequences

of dropping the transfer cask during the handling phase:

A mechanical load analysis determines the stresses on the fuel cladding and MPC as a function

of drop height.

Given the stress on the fuel cladding and MPC from dropping the transfer cask or MPC,

failure analyses determine the probability of the fuel cladding and MPC failing.

Source term analysis is used to determine the releases from the fuel cladding and the MPC.

An engineering analysis determines the reliability of the secondary containment to isolate

if there is an accidental release of radioactivity inside the secondary containment.

To quantify the risk, the study uses "best available" point estimates. When there is insufficient

information or data, conservative bounding assumptions or estimates are used. Because no uncertainty

analysis was performed, the identification of the dominant contributors was based on the point estimates

developed by the study. The dominant contributors might possibly change if uncertainty is considered.

The study measures the risk to the public in terms of the individual probabilities of a prompt fatality

and a latent cancer fatality within 16 km (10 mi) of the site. In calculating the risks, the study

also considers weather conditions and the population distribution in the vicinity of the site.

The results of this analysis indicate that the risk is solely from latent cancer fatalities, and no prompt

fatalities are expected. The risk is dominated by accident sequences occurring in three stages of the

handling phase. These involve the drop of the transfer cask through the equipment hatch (Stage 18) and

drops of the MPC into the storage overpack (Stages 20 and 21). The aggregated risk values are quite

low. Some of the scenarios have zero risk because either their initiating events cannot occur at the

subject plant or no radioactive release will result. The overall risk of dry cask storage was found to be

extremely low. That is, the estimated aggregate risk is an individual probability of a latent cancer fatality

of 1.8x 1012 during the first year of service, and 3.2x 10"4 per year during subsequent years of storage.

Questions pertinent to this study may arise from the public. Thus, a list of potential questions and their

answers as related to this study has been prepared and is presented in Appendix F.

xii
ACKNOWLEDGMENTS

Team Members Contributions

U.S. Nuclear Regulatory Commission

Ronaldo Jenkins

Alan Rubin

Edward Rodrick

Lee Abramson

John Thorp

Asimios Malliakos

Christopher Ryder

Christina Antonescu

Gordon Bj orkman

Kim Hardin

Tze-Jer Chuang

Robert Einziger

Jason Piotter

Monideep Dey

Brad Hardin

Douglas Kalinousky

Jocelyn Mitchell

Carlos Navarro

Cayatano Santos

Shah Malik

Jason Schaperow

Syed Shaukat

Ghani Zigh

Tony Ulses

Overall management

Overall management and technical oversight (through March 2006)

Team Leader (through September 2001)

Team Leader, data analysis, technical review, report editing

(through September 2004)

Team Leader, technical review, report editing (through August 2005)

Team Leader and final risk assessment calculations, technical review, and

report editing

Technical coordination among technical disciplines, report compilation,

and review; flooding analysis and project management of secondary

containment analyses at Lawrence Livermore National Laboratory;

and peer review at Sandia National Laboratories (through September 2003)

Effects of lightning, and project management of lightning analysis

at Oak Ridge National Laboratory

Structural mechanics, stress analysis and buckling, LS-DYNA model

modifications and analyses, weld failure methodology, MPC integrity

evaluation and failure probability, fuel rod integrity methodology

and analyses, and aircraft impact analysis

SFPO Team Leader, coordination among SFPO technical disciplines,

technical review, and interface with RES (2004-2005)

SFPO Team Leader, coordination among SFPO technical disciplines,

technical review and report editing, and interface with RES (2005-2006)

Source term methodology and source term estimates

Implemented LS-DYNA model modifications and LS-DYNA analyses

(March 2006 to February 2006)

Fires (through November 2004)

Frequencies of aircraft crashes and natural events

Failure analysis

Consequences

Consequences

Failure analysis

Failure analysis

Source term and consequences

Structural mechanics and project management of structural analyses

at Brookhaven National Laboratory

Thermal-hydraulic analyses of fires and blocked vents

Criticality analysis

xiii
Team Members Contributions

Brookhaven National Laboratory

Joseph Braverman Mechanical load calculations

Charles Hofmayer Mechanical load calculations

Charles Miller Mechanical load calculations

Richard Morante Mechanical load calculations

Jim Xu Mechanical load calculations

Jinsuo Nie Mechanical load calculations

Nikolaos Simos Mechanical load calculations

Lawrence Livermore National Laboratory

Peter Prassinos Reliability of secondary containment isolation

Oak Ridge National Laboratory

James Yugo Electromagnetic analysis of lightning effects on dry cask

Sandia National Laboratories

Douglas Ammerman Cask and fuel rod response for source term estimates

Robert Kalan Cask and fuel rod response for source term estimates

Vincent Luk Seismic load calculations

Jeremy Sprung Source term estimates

Comment Review Group Group Leaders

SFPO Thermal Group Jorge Solis

SFPO Structural Group Gordon Bjorkman, Henry Lee, Bob Tripathi,

Jason Piotter, and David Tang

SFPO Materials Group Robert Einziger, Tze-Jer Chuang,

and Chuck Interrante

SFPO Criticality Group Carl Withee, Kim Hardin, and Bernie White

SFPO Containment and Radiation Protection Kim Hardin, Steve Baggett, Mike Waters,

Michel Call

Transportation and Storage Safety & Inspection Rob Temps

Section

Spent Fuel Licensing Section Steve O'Connor

NMSS Risk Task Group (no longer in existence) Bret Leslie

Specialists Dennis Damon and Robert Einziger

xiv
ABBREVIATIONS

ASME

BNL

BWR

CFD

CRUD

CFR

DBE

DCF

EAB

EPS

FAA

FEM

FSAR

GFD

American Society of Mechanical Engineers

Brookhaven National Laboratory

boiling-water reactor

computational fluid dynamics

Chalk River unidentified deposits

Code of Federal Regulations

design-basis earthquake

dose conversion factor

exclusion area boundary

effective plastic strain

Federal Aviation Administration

finite element model

final safety analysis report

ground flash density

HI-STORM

HI-TRAC

HRA

HVAC

IFT

ISFSI

KE

LWR

Holtec International Storage and Transfer Operation Reenforced Module

Holtec International Transfer Cask

human reliability analysis

heating, ventilation, and air conditioning

isolation failure type

independent spent fuel storage installation

kinetic energy

light-water reactor

MACCS2 MELCOR Accident Consequence Code System

MPC multipurpose canister

NG noble gases

NLDN National Lightning Detection Network

NMSS Office of Nuclear Material Safety and Safeguards (NRC)

NRC U.S. Nuclear Regulatory Commission

NSSS nuclear steam supply system

ORNL Oak Ridge National Laboratory

OT other than noble gases

xv
PE potential energy

PMP probable maximum precipitation

PRA probabilistic risk assessment

PT dye penetrant (test) examination

PWR pressurized-water reactor

Ref reference

RES Office of Nuclear Regulatory Research (NRC)

RT radiographic (test) examination

SA submerged arc (weld)

SGTS standby gas treatment system

SIMPLE Semi-Implicit Method for Pressure-Linked Equations

SME seismic margin earthquake

SNL Sandia National Laboratories

SRP Standard Review Plan

TIG tungsten-inert-gas (process)

TO transfer overpack

UT ultrasonic testing (nondestructive test process)

xvi
SYMBOLS

atm

C

cfm

cm

F

ft

g

hr

in.

kg

km

kPa

kW

lb

m

MPa

mi

min

mph

MW

MWd/MTU

psi

S

yr

atmospheric pressure

Centigrade/Celsius

cubic feet per minute

centimeter

Fahrenheit

feet

gravitational constant

hour

inch

kilogram

kilometer

kilo-Pascal

kilowatt

pound

meter

mega-Pascal

mile

minute

miles per hour

megawatts

megawatt-days per metric ton

pound per square inch

second

year

xvii

GLOSSARY OF SELECTED TERMS

Probabilistic Risk Assessment (PRA): A systematic method for addressing the risk triplet, as it relates

to the performance of a complex system, to understand likely outcomes, sensitivities, areas of importance,

system interactions, and areas of uncertainty. The risk triplet is the set of three questions that the NRC

uses to define "risk": (1) what can go wrong? (2) how likely is it? and (3) what are the consequences?

The NRC identifies important scenarios from such an assessment [from "Staff Requirements -

SECY 98-144 - White Paper on Risk-Informed and Performance-Based Regulation" (Reference 1)].

Storage Cask: A Holtec International Storage and Transfer Operation Reenforced Module

(HI-STORM) containing a multipurpose canister (MPC) loaded with fuel assemblies.

Storage (HI-STORM) Overpack: The long-term storage container designed and fabricated to allow

safe movement of an MPC containing spent fuel assemblies to a storage pad and to provide shielded,

safe storage of the MPC in the independent spent fuel storage installation (ISFSI). When the MPC

containing fuel assemblies is transferred into the storage overpack, it is referred to as the "storage cask."

Transfer Cask: A Holtec International Transfer Cask (HI-TRAC) containing an MPC loaded with fuel

assemblies.

Transfer (HI-TRAC) Overpack: The shielded container designed and fabricated to allow safe loading

of fuel assemblies into the MPC, preparation of the MPC for storage, and transfer of the MPC

into the storage (HI-STORM) overpack. When the transfer overpack has been loaded with an MPC

and associated spent fuel assemblies, it is referred to as the "transfer cask."

xix

1. INTRODUCTION

The spent fuel pools of commercial nuclear power plants are becoming filled with spent fuel assemblies.

To avoid having to cease operations when the pools are full, many utilities have been removing older fuel

from the pools and storing it in dry casks in an onsite independent spent fuel storage installation (ISFSI).

In this study, a methodology for performing a probabilistic risk assessment (PRA) of a dry cask storage

system was developed and applied by performing a pilot PRA of a specific cask system at a specific

boiling-water reactor (BWR) site. Although the study results do not necessarily apply to other cask

systems or sites, the methodology can serve as a guide for performing other such PRAs. Nonetheless,

it should be noted that this study does not endorse the use (or non-use) of a dry cask storage system.

at any particular site.

The U.S. Nuclear Regulatory Commission (NRC), Office of Nuclear Regulatory Research (RES), jointly

with the Spent Fuel Project Office (SFPO), performed this study in response to a request by the NRC's

Office of Nuclear Material Safety and Safeguards (NMSS), to support its efforts to risk-inform NMSSregulated activities. It is expected that NMSS will use the results of this study in conjunction with

the methodology to develop a basis to determine the need for other PRAs, improvements in data gathering

and analysis, and additional engineering design analysis, and to identify general program areas that

may be candidates for increased or decreased staff review or inspection focus. Because the focus of this

pilot study is solely on the methodology and its limited (i.e., case specific) application, no inferences

or conclusions should be drawn with regard to the regulatory implications of this study.

1.1 Purpose

The purpose of this study is to develop and apply a methodology for performing a PRA to assess the risk

to the public and to identify the dominant contributors to risk associated with dry cask storage involving

a single cask at a specific site. The risk is evaluated using "best available" point estimates; when there is

insufficient information or data, conservative bounding assumptions or estimates are used. The results

of this study can be helpful for future studies to determine the extent and depth of any uncertainty analysis

that may be needed. Similar to PRAs for nuclear power plants, this PRA includes a list of initiating

events, response of system (i.e., dry cask) to initiating events, secondary containment analysis,

and consequence analysis. Contrary to PRAs for nuclear power plants, this PRA did not include

a human reliability analysis (HRA).

1.2 Scope

1.2.1 Issues Within the Scope

The cask system analyzed is the Holtec International Storage and Transfer Operation Reinforced Module

(HI-STORM) 100 (Reference 2). This cask system, which is used for onsite storage only, consists of

three major components, including a multipurpose canister (MPC), a Holtec International Transfer Cask

(HI-TRAC overpack) which in this report is called a "transfer overpack," and a HI-STORM (overpack)

called the "storage overpack." The MPC confines spent fuel assemblies for the duration of the storage

period (typically 20 years). When the MPC is removed from the spent fuel pool and prepared for storage,

it is already inside the transfer overpack to shield workers from radiation.

1-1
In this study, the transfer overpack containing the MPC is referred to as the "transfer cask,"

and the storage overpack containing the MPC is referred to as the "storage cask." During onsite storage,

the storage cask shields people from radiation, provides a path to cool the MPC by convection through

vents, and mechanically protects the MPC.

The study's assessment begins when spent nuclear fuel assemblies are loaded into the MPC, within

the transfer overpack. The transfer cask is then moved to an area on the refueling floor, where the MPC

is prepared for storage. The transfer cask is lowered through the equipment hatch to the storage overpack

on the ground level, where the MPC is moved from the transfer cask into the storage overpack.

The loaded storage cask is then moved to the ISFSI, where it is set on a concrete pad.

For source term and release fractions, this study considered spent BWR fuel. The fuel considered

is high-burnup, having been stored and cooled for 10 years, respectively, in the spent fuel pool.

1.2.2 Issues Beyond the Scope

1.2.2.1 Version of the Cask

The version of the cask used in this study is that documented in Reference 2. Any subsequent changes

to the design, documented as amendments or performed in accordance with Title 10, Section 72.48,

of the Code of Federal Regulations (10 CFR 72.48), are beyond the scope. Nonetheless, changes

to the cask would have to be fairly significant to impact the main conclusion of this report

(that the overall risk is extremely low).

1.2.2.2 Unloading, Offsite Transportation, and Repository Storage

A storage cask may have to be unloaded. The storage overpack is designed for onsite transfer and storage;

it cannot be moved offsite. If spent fuel is moved to a permanent repository, at a minimum, the MPC

would have to be removed from the storage overpack and placed in a transport overpack. If, for some

reason, the spent fuel needed to be reconfigured (e.g., because incorrect assemblies were mistakenly loaded),

the MPC would need to be opened. Neither circumstance is analyzed in this study. The storage overpack

is designed for onsite transportation and storage. Thus, offsite transportation and repository storage

are not analyzed in this study.

1.2.2.3 Damage to the Plant

The effects of dropping the transfer cask on the floors of the load path from the cask pit (the alcove

of the spent fuel pool where the transfer cask is loaded) to the equipment hatch were estimated

where details of the floors were obtained. Any damage to underlying plant systems, equipment,

and components, or capabilities to shut down the reactor were not analyzed and are beyond the scope

of this study. Where the floor has supporting beams, girders, or concrete walls located underneath,

the floor is expected to hold the cask, were it to drop from a distance typical of the height to which it is

lifted while being moved. It is important to note that the crane carrying the transfer cask along the load

path between the cask pit and equipment hatch is single-failure-proof.

1-2
1.2.2.4 Uncertainty

Although there are inherent uncertainties in the initiating event frequencies, conditional probabilities,

and consequences used to calculate the risk, this study does not consider those uncertainties. This is

because the purpose and scope of this study were limited, as were the resources for developing

and applying a methodology to assess the risk to the public and to identify the dominant contributors

to risk associated with dry cask storage involving a single cask at a specific site. An essential element

of an uncertainty analysis is a sensitivity analysis whereby the possible ranges of the input variables are

determined and their effects on the output measures are systematically evaluated. With a few exceptions,

this study did not involve a sensitivity analysis of its input variables. This study uses best available

point estimates, when available; conservative or bounding assumptions are used when best estimates

are not available. However, because no uncertainty analysis is performed, the degree of conservatism

in the risk estimates cannot be determined.

1.2.2.5 Worker Risk

The PRA methodology developed in this study is patterned after that used for nuclear power plants,

which considers public risk. PRAs for reactors, such as those in Reference 3, were performed, in part,

to evaluate the extent to which five plants met the NRC's safety goals (Reference 4) for risk to the public.

Like PRAs for reactors, this study does not consider worker risk.

1.2.2.6 Human Reliability

A human reliability analysis (HRA) of loading spent fuel into the MPC, lifting the transfer cask during

the handling phase, and welding the MPC when preparing it for storage, was not incorporated into this

pilot PRA.

1.2.2.7 Fabrication of the MPC, Transfer Overpack, and Storage Overpack

This study assumes that the cask is fabricated as designed (Reference 2).

0 Assumption: The MPC, transfer overpack, and storage overpack are constructed as designed.

Fabrication includes manufacturing and assembly of cask components. While fabrication errors could be

modeled and analyzed, this was not done in this study.

Welds made according to accepted standards inherently contain small flaws that depend on the material

being welded, the welding materials, and the welding process. Although this study does not consider

fabrication errors, the weld failure evaluation of the MPC does take into account normal flaws that exist

in weld-deposited austenitic stainless steel. Statistical distributions describing the densities and sizes

of flaws were obtained from results documented in Reference 5.

1-3
1.2.2.8 Mis-Loading Spent Nuclear Fuel

An HRA to determine the probability of mis-loading spent nuclear fuel into the MPC is beyond the scope of

this study for the reasons discussed in Section 1.2.2.6. Even though the frequency of mis-loading is not

estimated, deterministic calculations were performed to investigate the effects of mis-loading on thermal

loads, the failure probability of the MPC (Section 8.1), and the possibility for criticality (Section 8.2).

1.2.2.9 Aging Effects

To evaluate possible aging effects, a CASTOR-V/21 dry storage cask was examined for degradation

(Reference 6). This cask was produced for use in testing aging effects on long-term dry cask storage.

(This cask design was not put into production.) The examination consisted of remote indirect visual

examination [cask internal, lid seals, and pressurized-water reactor (PWR) spent fuel assemblies]

and temperature measurements of selected spent fuel assemblies. Interior crud samples were taken,

and surface gamma and neutron dose rates were measured. Selected fuel rod assemblies were removed

from One assembly, visually examined, and subjected to nondestructive, destructive, and mechanical

examination. The helium inerting gas was sampled for analysis.

After 14 years of storage, no evidence of degradation that would affect the performance of the cask

or integrity of the fuel was evident. The fuel was intact; there was no evidence of creep or rod bow.

Crud found on the fuel rods was attributed to oxidized steel while the fuel was in the reactor; none of

the oxidation is attributed to dry storage. Of 16 stitch welds that attached borated plates to the basket,

15 were cracked. The stitch weld cracking was attributed to the tight fit of the assemblies in the basket,

not to storage. Small amounts of air mixed in with the helium were attributed to the process of inerting

the cask.

While the results for the CASTOR-V/21 dry cask suggest that there will be no significant aging effects

for the subject cask, it is unclear whether the two casks experience similar conditions. Accordingly,

except for possible cask and fuel corrosion (see Section 8.3), aging effects are beyond the scope

of this study.

1.2.2. 10 Combinations of Factors That Could Impact the Probability of MPC Failure

Section 8 discusses several factors that can influence the probability of MPC failure. Individual factors

are investigated one at a time with sensitivity studies. It is not expected that combinations of factors

will significantly impact the risk estimates.

1.2.2.11 Military Missiles, Sabotage, and Terrorism

The impact of military missiles, sabotage, and terrorism was not incorporated into this pilot PRA.

1-4
1.3 Overview of the Risk Calculations

The structure of the risk calculations, as illustrated in Figure 1, closely mirrors the structure of this report.

In the figure, boxes indicate analyses, models, and results. The associated section numbers indicate

where this report discusses the results needed to estimate risk. Circled numbers are keyed to

the discussion below.

The system that is to be modeled consists of the dry cask (MPC, transfer overpack, storage overpack)

and the plant layout (load path, floor structure, crane). The structure of the dry cask components

is determined from engineering drawings and cask specifications (1). The next step is to describe

the secondary containment isolation system (2).

Challenges to the transfer cask and the storage cask are determined from other studies (3), such as

Reference 7. The list of initiating events in Reference 7 is a general list that was developed for risk

assessments of commercial nuclear power plants. Along with information about the cask system (1)

and the subject plant (2), the initiating events are evaluated (4) to derive a list of initiating events (5)

that is relevant to the subject cask at the subject site. The frequency of each initiating event in the list

is then determined (6) from other reference information (3).

With information about the cask system (1), and information about the plant layout (2), engineering

analyses (7) are performed to determine the mechanical loads (8) on the cask that result from initiating

events that can occur at the subject site and affect the cask. A mechanical load model (9) is developed

with engineering analyses (7). With the mechanical loads for each initiating event, the mechanical load

model is used to determine the mechanical stresses (10) in the MPC. The mechanical stresses are then

used in a failure model of the MPC (11) to assess the probability of MPC failure (12). The mechanical

loads are also used in the fuel failure model (16) to determine the fraction of fuel rods that fail (17).

With information about the cask system (1) and a list of initiating events (5) that can occur at the subject

site and affect the cask, engineering analyses (7) are performed to determine the thermal loads (13).

A thermal load model (14) is developed with engineering analyses (7). Given the thermal loads (13),

the thermal load model is used to determine the thermal stresses (15) on the MPC and the fuel rods.

The thermal stresses are then used in the failure model of the MPC (11) to assess the probability of MPC

failure (12), and in the fuel failure model (16) to determine the fraction of fuel rods that fail (17).

From the plant information (2), engineering analyses (7) are performed to assess the systems reliability

to isolate the secondary containment in the event of a release of radioactive material inside the secondary

containment. A reliability model (18) is then used to determine the conditional probability

that the secondary containment will fail to isolate (19) if the MPC is dropped, fails, and releases

radioactive material into the secondary containment.

1-5
StUd LS

EI \Iltuatiolis

c alIN ii Sis

c\ Clils

sccII~iol

specificaos I 1

7

Eii ,iiieeii O,,

aiiulvses

Mech'anlicallT ellnl

S I S loads loadS

ýcI Sect4o2

Scini4.1 'ciii 4

i 14

NIeloa iil duIT nIW

mtodel 1odeLC

S•ciuol 4 1. 11cciion 4. 21

N e caloIru i I 1ThermaliSon'i

S( esses StI elsses tir)

Scciiois 4.1 1 SCeijous 4 I.2 Secni

4 1C,- & 41 ý 4.2

I II I lC' ISI ' 11

nio0ckel modcel Retiabhlý e~il

8ccimoi 4. Swimol 4.4A ~ ~

17I)

Pirobahi lities Era ctio os a l, Pioha hili I lvcu C on seq

RZisk- IeSultS

Scciiouu 7.2

20

I-requencies

SCcioll

Figure 1. Structure of the Risk Calculations and Organization of the Report

1-6
From other studies (3), source terms (20) are determined. The source terms, adjusted to account for

the inventory of radionuclides in the MPC and other information (3) about the plant site (e.g., population

distribution), are then used in a consequence model (21) to estimate the consequences (22) of an accident.

The values needed to assess risk (23) are the frequencies of the initiating events (6), the probability

of MPC failure (12), the fractions of fuel rods that fail (17), the probability that the secondary

containment will fail to isolate (19), and the consequences (22). These five values are combined

to calculate the risk to the public. To quantify the risk, best available point estimates are used.

When there is insufficient information or data, conservative bounding assumptions or estimates are used.

1.4 Elements of the PRA

The dry cask storage operation is divided into the three phases of handling, transfer, and storage. During

the handling phase, spent fuel assemblies are placed into the MPC within the transfer cask. After the MPC

is dried, inerted with helium, and sealed, the transfer cask is lowered through the equipment hatch

to the storage overpack on the ground level. The MPC is then moved from the transfer cask into

the storage overpack. The transfer phase takes place when the storage cask is moved out of the secondary

containment to the storage pad. The storage phase takes place after the storage cask is set down

on the storage pad. To analyze the risk of the dry cask storage operation, the phases are divided into

34 stages. A stage is defined as a discrete part of the operation that is convenient for the risk analysis.

For example, because the height at which the cask is moved and the composition of the surface over which

it is moved determine the mechanical load on the MPC if it were dropped, these aspects are used to define

some of the stages.

A comprehensive list of initiating events is developed, and the risk associated with each initiating event

is evaluated. Analyses are performed to determine the cask's response to the mechanical and thermal

loads imposed by the initiating events. For example, the following analyses evaluate the consequences

of dropping the transfer cask during the handling phase:

* A mechanical load analysis determines the stresses on the fuel cladding and MPC as a function

of drop height.

* Given the stress on the fuel cladding and MPC from dropping the transfer cask or MPC,

failure analyses determine the probability of the fuel cladding and MPC failing.

* Source term analysis is used to determine the releases from the fuel cladding and the MPC.

* An engineering analysis determines the reliability of the secondary containment to isolate

if there is an accidental release of radioactivity inside the secondary containment.

The annual risk to the public of handling, transfer, and storage of a single cask is estimated. The risk

measures are the individual probabilities of a prompt fatality within 1.6 km (1.0 mi) and a latent cancer

fatality within 16 km (10 mi) of the site. In calculating the annual risks, the PRA considers weather

conditions and population distribution in the vicinity of the site. (Appendix E to this report provides more

detailed discussion and basis for emergency planning assumptions.) These are the consequence measures

that are typically calculated in the risk assessments of nuclear power plants (e.g., Reference 4). Risks to

the environment and property are not typically calculated in reactor PRAs and were not considered in this

study.

1-7

2. DRY CASK AND SECONDARY CONTAINMENT SYSTEMS

2.1 Dry Cask System

The Holtec HI-STORM 100 dry cask system consists of an MPC that confines the fuel, a transfer

overpack that shields workers from radiation while the cask is being prepared for storage, and a storage

overpack that shields people from radiation and protects the MPC during storage. When the transfer

overpack contains the MPC, the unit is referred to as the transfer cask. When the storage overpack

contains the MPC, the unit is referred to as the storage cask. Details of the cask system are illustrated

in Figures 2 and 3. These and other figures are based on drawings in Sections 1 and 8 of Reference 2.

Figure 2 is a composite sketch of the MPC

and transfer overpack. The MPC is 4.8 m ])S iI I C I-Il1'_r

(15.8 feet) high and 1.73 meter (5.7 feet) in - tl co'cl iCIa tc

diameter. When loaded with BWVR fuel- in p,,iI

assemblies, it weighs 36 metric tons (40 MvC lid

tons). The transfer cask is 5.0 meters (16.4 !/

feet) high and 2.3 meters (7.6 feet) in ML'c\

diameter. When loaded with BWR fuel T.UC

assemblies, it weighs 91 metric tons -buskc

(100 tons). The fuel basket is free standing, 4.8

but held in position inside the MPC by basket r151 I M1C 'shell

supports. Boral plates along the walls of --- l)uin

each cell prevent the spent fuel from 1

becoming critical, while upper and lower fuel 157 1 ' [p)cer

spacers keep the fuel assemblies vertically 4€ [11run1it0n

positioned in the basket. Except for the boral Wuacijuckct

plates and heat conduction elements, all MPCG

components are made of stainless steel. The I InnCr steel shell

transfer overpack consists of an inner steel O-Radial leed sLteel shield sell

shell, lead shield, and outer steel shell, and a lsiL \ter jacket

0\ 4 C 1) puk ( Rdi alchannels water jacket provides additional shielding. 1. Shellc auialt lcanel

The pool lid at the bottom of the transfer

overpack seals the inner cavity, thereby

retaining clean demineralized water

to prevent the exterior of the MPC from 2Poo •IL

being contaminated. A lift yoke (not shown) t7.6 ti

attaches to lift trunnions on the transfer

overpack. Figure 2. MPC and Transfer Overpack

Figure 3 is a composite sketch of the storage

cask with the MPC partially inserted. The overpack consists of inner and outer steel shells connected by

full-length steel radial plates that extend from the bottom to the top. The volume between the shells is

filled with unreinforced concrete. The shells and the radial plates form the structural member. The

concrete is for radiation shielding. The lid is bolted to the overpack at anchor blocks, which are attached

to the radial plates.

2-1
Channels between the overpack and the MPC allow air to circulate, entering the inlet vent at the bottom

of the overpack and exiting the outlet vent at the top of the overpack. The channels are designed

to progressively collapse during a severe impact, absorbing some of the impact that could be transmitted

to the MPC. Shield cross-plates reduce radiation streaming through the vents. Screen covers prevent

insects, birds, and animals from entering the vents to build nests that might block the vents.

)uIIct \CI1I

Aunchor bhlock -

19.

ShCII -

- ~I 1'2 IH'

lnlct

Figure 3. MPC and Storage Over-pack

2.2 Stacies of the Dry Cask Storaqe Operation

As previously noted, the dry cask storage operation is divided into the three phases of handling, transfer,

and storage:

(1) During the handling phase, activities take place on the refueling floor and ground level.

From the refueling floor, about 29.9 meters (98.1 feet) above ground level, the transfer cask

is lowered about 13 meters (42.7 feet) to the bottom of the cask pit next to the spent fuel pool.

There, the dry cask storage operation begins when spent fuel assemblies are loaded into the MPC.

After the MPC basket is full, the lid is installed and the transfer cask is lifted out of the spent fuel

pool and moved to a preparation area on the refueling floor, where the MPC is drained, dried,

inerted, and sealed. Additional preparations are made, and the transfer cask is lowered

to the storage cask, still inside the secondary containment. The MPC is then lowered

from the transfer cask to the storage cask.

2-2
(2) After the Kevlar stays and the alignment device are removed, the storage cask (with the MPC inside)

is moved through an airlock to outside the secondary containment to begin the transfer phase.

The storage cask is moved away from the secondary containment, where final preparations

are made before moving the storage cask to the storage pad with a specially designed tractor,

called an overpack (cask) transporter. There, the storage cask is set down in a predetermined

location (footprint), which is indicated by paint marks on the storage pad. This predetermined

storage footprint ensures that the storage casks are adequately spaced and their weight is properly

distributed over the storage pad.

(3) During the storage phase, storage casks remain on the storage pad for 20 years. The storage cask

does not have (and does not require) instruments or other mechanisms to detect heat loads

or leaks in the MPC. Routine surveillance ensures that the vents remain unblocked.

The handling, transfer, and storage phases of the dry cask storage operation are divided into 34 stages,

as listed in Table 1. Figure 4 is a schematic diagram illustrating the movement of the transfer cask

and storage cask through the stages in Table 1 as the MPC is moved from the spent fuel pool to the ISFSI,

where the storage pads are located. The numbers along the paths in Figure 4 correspond to the stages

in Table 1.

A stage is defined as a discrete part of the operation that is convenient for risk analysis. Factors used

to define the various stages include the following examples:

* height at which the cask is carried

* direction in which the cask is moved

• rigging of the cask

* surface (e.g., concrete, asphalt, gravel) over which the cask is moved

The phases described above reflect the process used at the particular plant site that served as the basis

for this study. Terminology may vary to some extent from plant to plant, so in such cases, the reader

is advised to compare the given plant's terminology with the way the terms are used in this report.

2-3
Table 1. Stages of the Dry Cask Storage Operation

Height (A)

Stages In ft

1 Loading fuel assemblies into the MPG B 4.8 16

2 Placing the MPG lid onto the MPG and engaging the lift yoke on the transfer cask 0 0

3 Lifting the transfer cask out of the cask pit 13 42.5

4 Moving the transfer cask over a railing of the spent fuel pool 0.9 3

5 Moving the transfer cask to the preparation area (Is' segment) 0.3 1

6 Moving the transfer cask to the preparation area (2 "~ segment) 0.3 1

7 Moving the transfer cask to the preparation area (3 rd segment) 0.3 1

8 Lowering the transfer cask onto the preparation area (D) 0.3 1

9 Preparing (draining, drying, inerting, and sealing) the MPG for storage 0 0

10 Installing the short stays and attaching the lift yoke ()0 0

11 Lifting the transfer cask 0.6 2

12 Moving the transfer cask to exchange bottom lids of the transfer cask (1St segment) 0.6 2

13 Moving the transfer cask to exchange bottom lids of the transfer cask (2 nd segment) 0.6 2

14 Replacing the pool lid with the transfer lid 0.1 0.25

15 Moving the transfer cask near the equipment hatch 0.6 2

16 Holding the transfer cask 0.6 2

17 Moving the transfer cask to the equipment hatch 0.6 2

18 Lowering the transfer cask to the overpack through the equipment hatch 24.4 80

19 Preparing (remove short stays, disengage lift yoke, attach long stays) to lower the MPG 0 0

20 Lifting the MPG and opening doors of transfer lid 5.8 19

21 Lowering the MPG through the transfer cask into the storage cask 5.8 19

22 Moving the storage cask into the airlock on Heiman rollers 0 0

23 Moving the storage cask out of the airlock on Heiman rollers 0 0

24 Moving the storage cask away from the secondary containment on Heiman rollers 0 0

25 Preparing (installing lid, vent shield cross-plates, vent screens) the storage cask for storage 0 0

26 Lifting the storage cask above the Heiman rollers with the overpack transporter 0.1 0.25

27 Moving the storage cask above a cushion on the preparation area <0. 1 <0.25

28 Holding the storage cask above the cushion while attaching a Kevlar belt <0. 1 <0.25

29 Moving the storage cask above the concrete surface of the preparation area 0.3 1

30 Moving the storage cask above the asphalt road 0.3 1

31 Moving the storage cask above the gravel surface around the storage pads 0.3 1

32 Moving the storage cask above the concrete storage pad 0.3 1

33 Lowering the storage cask onto the storage pad 0.3 1

34 Storing the storage cask on the storage pad for 20 years 0 0

(A) Height is the distance the cask would fall if the support system failed.

(B) Prior to Stage 1, the MPG is inserted into the transfer cask, and after other preparations, lowered into the cask pit.

The storage overpack is placed on Heiman rollers and moved under the equipment hatch.

r) The lift yoke attaches to the trunnions of the transfer cask.

(D) Stays attach to the lift yoke on one end and cleats of the MPG on the other end.

2-4
I I dC"Cci~d t spcill ftile stol-Ia_'C

- - - - - - - - - - - -

[ito i liiiý5

C3 C3 - -

El chaid iI,

id

- - - - - - - - - -

pit

1,1floor floor

Loadijiia

spell[

linicear

RICI

Secondarn

co •l [li 111 c l I

Figure 4. Movement of the Transfer Cask and Storage Cask

Through the Stages of the Dry Cask Storage Operation

Figure 5 shows a vertical view of stages where the transfer cask is in the equipment hatch. Except for

the floor thicknesses, the vertical axis is drawn to scale; the inset shows the wall supporting the overpack

at ground level. The steel beams and concrete walls supporting the refueling floor along the path

where the transfer cask is carried are not shown. The lift height is kept to the minimum necessary

to clear obstacles on the floor (such as railings, pipes, and raised edges of floors).

2-5
Variations in the lift height and path over which

the transfer cask and storage cask are normally I,,

carried are considered uncertainties. -,k _ ,,

Uncertainty is beyond the scope of this study I

(see Section 1.2.2.4). ,1 111 I III 

L'.IkIIl' C~

-L - 1 I- -I - - -- - - - - -- - - - -- - -- -

Pre-stage conditions Sta and e l•,i..... preparations prior to .. = = .".i

Stage I\I\'l

C:,<l. & '-1411)"' b ':

The secondary containment is closed because i

the reactor is at full power. The normal slightly Iil..Il

negative pressure inside the secondary -l

containment is established. Wl ,,d " .....

The storage overpack is on Helman rollers on the z......

ground floor underneath the equipment hatch.

Figure 5. Vertical View of the Equipment Hatch The MPG is inserted into the transfer overpack

(transfer cask). After the MPC is filled with water,

the transfer cask is lowered into the cask pit

(an alcove of the spent fuel pool).

Gates between the spent fuel pool and cask pit, above the level of the fuel, have been removed to allow

movement of spent fuel assemblies from the spent fuel pool into the MPC inside the empty transfer cask

in the cask pit. After the MPC is filled with 68 spent fuel assemblies, the gates are replaced and the lid

is placed on the MPC.

The analysis uses the following assumptions:

* The secondary containment remains closed during the handling phase.

* The transfer cask and storage cask are carried at the heights specified in the formal procedures

(see Table 1).

* The transfer cask is carried along the load path over the refueling floor of the secondary

containment.

* The storage cask is carried along the designated route to its placement on the storage pad

in the ISFSI.

Stages 1-3

In Stage 1, fuel assemblies are placed inside the MPC basket. Specific assemblies are loaded into specific

spaces of the basket to take advantage of the shielding, criticality control, and heat rejection capabilities

of the cask system. After the MPC is loaded with spent fuel assemblies, the lid is installed to shield

workers from radiation emanating from the top (the transfer cask shields workers from radiation

emanating from the sides and bottom). The lift yoke is engaged on the trunnions of the transfer cask

(Stage 2). The transfer cask is lifted out of the cask pit to about 13 meters (43 feet) above the bottom

of the cask pit, where it is rinsed with clean water (Stage 3).

2-6
Stages 4-7

The transfer cask is moved over the cask pit to clear a railing and a pipe (Stage 4), then lowered to about

0.3 meters (1 foot) as it is moved along a predetermined path to the preparation area. The load path

of the refueling floor is supported by steel beams and walls such that it can hold a static load

of 127 metric tons (125 tons). The load path to the preparation area consists of three perpendicular

segments.

Stages 8-13

At the preparation area, the transfer cask is lowered to the refueling floor (Stage 8) where the MPC

is partially drained and prepared for storage (Stage 9). After the lid is welded to the shell of the MPC,

the MPC is filled with water to hydrostatically test the weld. After the water is drained, the MPC is dried,

purged, and filled with helium. A port cover and drain cover are then welded to the lid, and the closure ring

is welded to the lid and the shell for redundant sealing. The top lid of the transfer cask is then installed

and bolted to the transfer overpack.

After the MPC is prepared,

the transfer cask is rigged to move

(Stage 10). While the yoke is over

the cask with the holes of the arms

below the trunnions, rigging personnel

on scaffolds install the short Kevlar

support stays, which are attached

from the lift yoke through an access

hole in the top lid of the transfer cask

to cleats on the MPC. In Figure 6,

the body of the yoke is close to

the transfer cask so the stays are slack

while being attached to the cleats.

As the yoke is raised, the stays

become taut, lifting the MPC off of

the pool lid. When the holes

of the arms are next to the trunnions,

the yoke arms are placed over

the trunnions, and the lift yoke

(A) (13)

Figure 6. Lift Yoke with Stays Attached (A) Prior to and (B)

after Engaging the Yoke Arms on the Trunnions

is attached. The transfer cask is then lifted about 0.6 meters (2 feet) above the refueling floor (Stage 11),

and is moved along a load path to another area of the refueling floor (Stages 12 and 13).

2-7
Stages 14-22

The pool lid at the bottom

of the transfer cask is replaced

by a transfer lid (Stage 14) for use

in Stage 20. The transfer lid is on

a inSae2.The carriage. pooife Short unbolted... stays lid is hold The the MPC ,, ,,,.,,,,_.,,,

above the pool lid inside the transfer ,'i: I.:',,,t•,••..,, 

cask. The transfer cask is set down I., , ']T Ion "d,

in the empty side of the carriage. -------------

The pool lid is unbolted. The,,,,.---

transfer cask is then raised slightly,

and the carriage is moved to bring the

transfer lid underneath the transfer

cask. The transfer cask is lowered

and the transfer lid is bolted to ----------------

the transfer cask shell. The bottom .... -- -

lid exchange is illustrated in Figure 7. - - ----- --

With the transfer lid attached, { . , }

the transfer cask is lifted to a height _________-_______

of about 0.6 meters (2 feet) above the

refueling floor and moved (Stage 15) Figure 7. Replacing the Pool Lid with the Transfer Lid

to a holding area near the equipment

hatch (Stage 16).

The transfer cask is held while preparations are made on the ground floor to receive the transfer cask.

The transfer cask is then moved to the equipment hatch (Stage 17). The height is sufficient to clear

a raised lip around the hatch on the refueling floor. The transfer cask is then lowered through

the equipment hatch of the refueling floor, 24.4 meters (80 feet) to the top of the storage overpack

(storage cask) at ground level (Stage 18). Two workers assist in guiding the transfer cask onto

the storage cask. An alignment device (i.e., collar on the storage cask) has receptacles that determine

the correct position of the transfer cask on the storage cask. The lift yoke is removed and the short stays

are replaced with long stays, which are used to lift the MPC off the transfer lid (Stage 19). After the MPC

is lifted several inches off the transfer door by the Kevlar stays (Stage 20), the transfer doors are opened,

and the MPC is lowered into the storage cask (Stage 21). As it is lowered, the MPC is guided by

the transfer cask and, when in the storage cask, by channels in the storage cask. Once lowered, the MPC

comes to rest on the pedestal inside the storage cask. The storage cask is moved along tracks on Helman

rollers into the airlock of the secondary containment (Stage 22). The movement is accomplished by

pulling the storage cask with a diesel-powered rail vehicle called the trackmobile. The airlock is needed

to maintain the negative pressure inside the secondary containment, which is developed by the ventilation

system.

2-8
Stages 23-28

The trackmobile continues to pull the storage cask from the airlock, outside the secondary containment

(Stage 23). The storage cask is moved a short distance outside the secondary containment, still on

the Helman rollers (Stage 24), to an area where it is prepared for storage (Stage 25). The storage cask lid

is installed. Shield cross-plates are inserted into the vents on the top and bottom of the storage cask,

and screens are installed over the vents. Cleats are bolted to the top of the storage cask. A transporter

is moved up to the storage cask, and the lift yoke is attached to the cleats. The lift yoke is engaged

and the storage cask is lifted about 0.3 meters (1 feet) above the ground (Stage 26). Four pins are inserted

into the hydraulic lifting mechanism of the storage cask transporter to lock them. The hydraulic mechanism

is deactivated, bringing the weight of the storage cask onto the four pins. The storage cask is moved

over a cushion (Stage 27), where it is held (Stage 28) while a Kevlar belt is wrapped around the midline

of the storage cask to keep it from swaying while being moved to the ISFSI.

Stages 29-33

The storage cask is moved by the transporter about 0.6 km (about 0.4 mi) from the reactor building

at a speed of about 0.64 km/hr (0.4 mph). This travel begins on concrete (Stage 29), but the road from

the preparation area to the ISFSI is asphalt (Stage 30), and the road between the storage pads inside

the ISFSI is gravel (Stage 31). The transporter travels along the gravel road and pivots next to the place

on the storage pad where the storage cask will be placed. The storage cask is then moved onto

the concrete storage pad (Stage 32), where it is set down in a predetermined location (Stage 33).

Stage 34

Storage casks are kept in the ISFSI. The ISFSI of the subject plant has four concrete pads, and each pad

holds 12 storage casks. Storage casks are placed on the pads in a pattern that distributes their weight.

Paint marks on the concrete indicate where each storage cask is to be placed.

2.3 Secondary Containment Isolation System

Stages 1-22 (Table 1) occur inside the secondary containment, which is a large concrete structure

that encloses the refueling and spent fuel pool areas, both reactor units, and associated equipment rooms

and areas. The secondary containment has a total free volume of approximately 84,951 m3 (3 x 106 ft3).

The atmosphere inside is cooled by four heating, ventilation, and air conditioning (HVAC) systems.

If a release of radioactive material were to occur inside the secondary containment following a drop

of the transfer cask, three distinct functions would occur to (1) detect radioactive material, (2) isolate

the secondary containment, and (3) operate the standby gas treatment system (SGTS). Each of these

functions is accomplished by redundant and independent trains of systems. The release of radioactive

material is detected by one or more of six detector systems that signal the containment to isolate.

Each of these detector systems consists of two trains, with each train having two detectors.

Both detectors in a train must detect radiation to isolate the containment. Only one of the 12 trains

is needed to isolate the secondary containment.

2-9
The radiological monitoring system for the secondary containment is part of the process radiation

monitoring system. Table 2 shows the process and effluent radiological monitors of the secondary

containment. This table provides a brief description of each radiological monitor in these systems.

Table 2. Detectors in the Secondary Containment System

Location Purpose Principal Radionuclides Detected

Reactor building ventilation stack

Main stack

Reactor building ventilation exhaust

Reactor building ventilation filter

discharge

Refueling floor zone ventilation exhaust

Refueling floor ventilation filter

Standby gas treatment system

Monitor discharge

to environment

Monitor discharge

to environment

Isolate building

and initiate SGTS

Monitor filter exhaust

Isolate building

and initiate SGTS

Monitor filter exhaust

Monitor process duct

Ar-41, Xe-133

Kr-85m, Kr-87, Kr-88, Xe-135

Kr-85m, Kr-87, Kr-88, Xe-135

Kr-85m, Kr-87, Kr-88, Xe-135

Kr-85m, Kr-87, Kr-88, Xe-135

Kr-85m, Kr-87, Kr-88, Xe-135

Ar-41, Xe-133

Normally, the HVAC systems ventilate the secondary containment, maintaining a slight negative pressure

inside with respect to the outside environment. If the containment is opened to atmosphere, fans will trip

on overload as they throttle to maintain the negative pressure, and indicators will alarm in the control room.

Table 3 shows that total exhaust flow rates are greater than total supply flow rates, to maintain a negative

pressure in the secondary containment. The difference in the flow rates is attributable to leakage through

cracks and joints into the secondary containment. The SGTS is normally in a standby mode. If a small

amount of radioactive material is released into the accessible area (where people are working), the exhaust

of this area isolates, and flow is redirected to the inaccessible area where it is filtered before being

exhausted. This arrangement utilizes the more effective filters on the exhausts of the inaccessible area

ventilation system and avoids isolating the entire secondary containment when small amounts

of radioactive material are released in the secondary containment, as would be expected during normal

plant operations. If a larger amount of radioactive material passes through the filter, radiation monitors

signal the HVAC systems to isolate the entire secondary containment.

2-10
Table 3. Volumetric Flow Rates of Secondary Containment Ventilation

HVAC

Supply Exhaust [ SGTS

m3/min ft3/min m3/min t/minI m3

/min ft 3/min

Unit 1 K

Unit 2

Refueling Area 850 30000 905 31970

Rx Bldg Accessible 1699 60000 850 30000 85 3000

SRx Bldg Inaccessible _950 31970

Refueling Area 850 30000 905 31970

Rx Building 170 6000 170 60001 113 4000

Total: 1 3568 J 126000 1 3735 1 131910 1 198 1 7000

In the event of a breach of the MPC and release of radioactive material into the secondary containment

building after a transfer cask drop, the HVAC system is designed to isolate the secondary containment

to prevent the release of radioactive material to the environment. The containment is isolated by stopping

the HVAC fans and closing one of two serial isolation valves located on each HVAC supply and exhaust duct.

One isolation valve is located inside, and the other is located outside the secondary containment.

While all HVAC fans receive an isolation signal, 6 of the 12 detector trains generate an isolation signal

for the inside isolation valves, while the other 6 detector trains generate an isolation signal for the outside

isolation valves. This arrangement allows for independence between the two serial isolation valves

and their isolation trains.

Each SGTS consists of two trains of filters and exhaust fans. Both trains of both SGTSs receive

a start signal from any of the 12 detector trains. While the capacity of both SGTSs is not enough

to overcome the flow generated by a single operating HVAC fan (supply or exhaust), only one SGTS

is needed to mitigate an uncontrolled and unfiltered release with all HVAC fans off. Even if the isolation

valves fail to close, the containment effectively isolates if all fans are off and either SGTS operates.

2-11

3. INITIATING EVENTS

3.1 List of Initiating Events

An initiating event is a disturbance in the normal operation of the dry cask storage system, which could

potentially lead to a release of radioactive material to the environment. For this study, a comprehensive

list of initiating events was developed. Those initiating events that could not affect the subject plant,

the transfer/storage cask(s), or the MPC were eliminated from further consideration. For the remaining

initiating events, estimates of frequencies, probability of MPC failure, and consequences were developed.

Tables 4-6 list initiating events at the subject plant during the handling, transfer, and storage phases,

respectively. These initiating events were identified from Reference 7, as well as design and operational

data for the specific cask and plant being studied. Information concerning the cask system design

was obtained from licensing documents (e.g., Reference 2). Analysts visited the subject plant to observe

the operation and equipment used during the handling, transfer, and storage phases. Written descriptions

of the procedures were also obtained and studied, and additional details were obtained through discussions

with plant personnel.

The lists of initiating events were reviewed by NRC staff who had also reviewed and licensed the dry cask

storage system. This review drew upon extensive knowledge and diverse perspectives regarding the cask

system. Based on these reviews and the process used to construct the list, Tables 4-6 are believed

to comprise a complete list of all initiating events that could conceivably affect the cask system.

The initiating events represent potential challenges to the cask's confinement function. They are grouped

by operational phase (i.e., handling, transfer, storage) and type of load placed on the cask (i.e., mechanical

and thermal).

In Tables 4-6, initiating events that are marked with an X in the status column have a risk of zero,

because the frequency is zero for the subject site, the probability of MPC failure is zero, or the fraction

of the fuel failing is zero. That is, the risk is zero if any one of these parameters is zero.

3-1
Table 4. Initiating Events During the Handling Phase

and the Sections Where Frequency, MPC Failure, and Fuel Failure Are Discussed

Mechanical Events Status Frequency MPC Failure Fuel Failure

Drop of the Transfer Cask

Into the cask pit 3.3.2 4.1.2.1 4.4

Cask tip-over 3.3.2 4.1.2.2 4.4

Onto storage overpack 3.3.2 4.1.2.3 4.4

Onto concrete floor 3.3.2 4.1.2.4 4.4

Onto refueling floor 3.3.2 4.1.2.5 4.4

Drop of the MPC into the storage overpack 3.3.2 4.1.2.6 4.4

Thermal Events

Fire from diesel fuel in the trackmobile X 4.2.1.2

LEGEND: X indicates that the initiator does not affect the subject plant or does not breach the MPC. ]

Table 5. Initiating Events During the Transfer Phase and

Section Where Frequency, MPC Failure, and Fuel Failure Are Discussed

Mechanical Events Status Frequency MPC Failure Fuel Failure

Drop of the Storage Cask

On concrete X 4.1.3.1 4.4

On asphalt X 4.1.3.1 4.4

On gravel X 4.1.3.1 4.4

Tip-over of the Storage Cask

While the cask is being moved X 4.1.3.1 4.1.3.2 4.4

Cask impacted by a vehicle X 4.1.3.1 4.1.3.2 4.4

Thermal Events

Fire from diesel fuel in trackmobile X 4.2.1.2

Fire from diesel fuel in cask transporter X 4.2.1.2

LEGEND: X indicates that the initiator does not affect the subject plant or does not breach the MPC.

3-2
Table 6. Initiating Events During the Storage Phase and

Section Where Frequency, MPC Failure, and Fuel Failure Are Discussed

Mechanical Events Status Frequency MPC Failure Fuel Failure

Water currents during a flood X 3.2.1

Tsunamis X 3.2.2

Seismic events 3.3.3 4.1.3.2

Volcanic activity X 3.2.3

High winds X 3.3.4 4.1.3.2

Soil erosion from intense precipitation X 3.2.4

Strikes from heavy objects

Aircraft 3.3.7 4.1.3.3

Wind-driven objects X 4.1.3.3

Heavy objects in flood waters X 3.2.1

Vehicle X 4.1.3.3

Meteorite 3.3.5 3.3.5

Shockwaves from explosions

Pipelines X 4.1.3.4

Nearby trucks and railcars X 4.1.3.4

Transformers X 3.2.5

Nearby barge X 3.2.5

Storage tanks X 3.2.5

Military missile X 3.2.5

Other facilities X 3.2.5

Thermal Events

Vent Blockage

Flood water over vents but below MPC X 3.2.1

Flood water over vents contacting MPC X 3.2.1

Flood water over vents submerging MPC X 3.2.1

Snow X 4.2.1.2

Ice X 4.2.1.2

Hail X 4.2.1.2

Intense precipitation X 4.2.1.2

Dirt & debris from winds X 4.2.1.2

Dirt from landslide due to heavy rains X 4.2.1.2

Dirt from landslide due to volcanic activity X 4.2.1.2

Volcanic ash deposits X 4.2.1.2

Accumulations or biological intrusions X 4.2.1.2

Lightning X 3.3.6 4.2.2.2

Fire causes pressurization or differential heating

Fire from aircraft fuel X 3.3.7 4.2.1.2

Fire from burning barge contents X 4.2.1.2

Fire from site-specific materials X 4.2.1.2

Fire from a gas main X 4.2.1.2

Forest fire X 4.2.1.2

LEGEND: X indicates that the initiator does not affect the plant or does not breach the MPC.

3-3
The lists of initiating events exclude those that are not predicted to occur as a result of physical

limitations. For example, water level from the hypothetically largest possible storm, along with three

dams failing, and 35 mph winds causing waves, is not predicted to reach the ISFSI. Because this is the

largest conceivable flood height, flooding is excluded as an initiating event.

A zero probability of MPC failure means that based on the weld failure data and methodology discussed

in Appendix B, and the MPC maximum stress (strain) levels calculated in Appendix A, the probability

of weld failure is less than 1 x 106 for the mechanical load event.

Reference 8 describes the modeling of the individual risk of "prompt fatality," using a two parameter

Weibull function. In the model, zero probability of prompt fatality is typically assumed at doses below

a threshold of 150 Rem to the red bone marrow and 500 Rem to the lungs. A prompt fatality can be

described as any death that occurs as a result of a large acute total body exposure sufficient to cause

one or more of three major classes of fatal syndromes (Reference 9) identified as cerebrovascular syndrome

causing death within 30-50 hours from exposures of about 100 Gy (10,000 Rads), gastrointestinal syndrome

causing death within about 9 days from exposures of about 10 Gy (1000 Rads), and hematopoietic

syndrome (bone marrow death) causing death within several weeks from exposures of 2.5-8 Gy

(250-800 Rads). While a single threshold value is not easily cited as the exposure below which

there is zero probability of prompt fatality, the assumption discussed above is consistent with the findings

discussed in Reference 9. The levels of exposure that could result from the events considered in this study

do not approach those that would be of concern for prompt fatality.

Relative to the probability of latent cancer fatalities, two general models were considered. The linear

threshold model predicts that doses below a specific threshold do not cause latent cancer fatality.

By contrast, the linear no-threshold model predicts a probability of a latent cancer fatality that is

proportional to dose. The current study used the linear no-threshold model.

If an uncertainty analysis were performed, some of the risks shown as zero in Tables 4-6, might become

positive, but would still be extremely small. However, such an analysis is beyond the scope of this study.

The initiating events that can occur during the handling phase are shown in Table 4. When the cask

is lifted out of the cask pit (Stage 3) and moved (Stages 4 through 8) to the preparation area, the MPC lid

is on, but not welded. MPC failure is not an issue because the MPC is unsealed. Also, since the MPC

is filled with water, for any fuel cladding failure that may occur during these stages, the fuel particulate

would be scrubbed since the gap release passes through the water in the MPC. In addition, there is no

cladding failure in a case of a tip-over. In Stage 8, the transfer cask is lowered to the preparation area

where the MPC is prepared for storage (Stage 9). In Stages 11-18, the transfer cask is prepared

and lowered to the storage overpack on the ground floor. In Stage 18, the drop height from the equipment

hatch to the top of the storage cask is about 24.4 meters (80 feet). The probability of the MPC failing

when the transfer cask is dropped from this height onto the storage overpack is 0.0002, based on a weld

failure probability distribution and strain values produced by LS-DYNA analyses of drops from various

heights. (See Appendix A for details.) In Stage 19, the transfer cask is resting on the storage cask, where

the rigging is changed to lower the MPC into the storage cask. Notably, the transfer cask cannot be dropped

while it is resting on the storage cask, and momentum calculations show that it cannot be tipped over.

3-4
After rigging, the MPC is lifted off the transfer lid doors (Stage 20) and lowered into the storage cask

(Stage 21). The drop height is as high as 5.8 meters (19 feet), and the probability of MPC failure

when dropped from this height into the storage cask is 0.28, based on a weld failure probability distribution

and strain values produced by LS-DYNA analyses of drops from various heights. This drop of the MPC

(by itself) into the storage overpack produces a much higher calculated probability of MPC failure

compared to the 24.4-meter (80-foot) drop of the transfer cask containing the MPC; this difference indicates

the protection that the transfer cask affords to the MPC during an accidental load drop. This protection

derives from the fact that the mass and ruggedness of the transfer cask locally crushes any target it impacts.

This crushing acts as an "impact limiter," which absorbs large amounts of kinetic energy and, in turn,

significantly limits the strain levels in the MPC. By contrast, a drop of the MPC into the storage overpack

is a very hard impact, in which virtually all of the MPC's kinetic energy is absorbed by its shell.

This energy absorption results in the higher strains in the MPC and a higher probability of failure.

The storage cask is on rollers when it receives the MPC, and those rollers allow the storage cask to be

moved inside an airlock of the secondary containment (Stage 22). The storage cask cannot be dropped

during this process. Stage 22 is the last stage in the handling phase.

In assessing the risk from the dry cask storage operation, only one initiating event at a time is postulated

to occur during an otherwise normal operation. That is, the assessment does not account for multiple

initiating events, such as the following examples:

aircraft crashes that might damage the secondary containment during an initiating event

in the handling phase

seismic events during the handling phase

Moreover, the analysis does not consider certain other circumstances, such as the following:

* The containment is normally closed (Section 5.1), and the analysis does not consider the effect

of the containment inadvertently being opened.

Even though such simultaneous occurrences could increase the consequences, the incremental risk

is negligible because of the minuscule incremental frequency of such events occurring simultaneously.

The initiating events during the transfer phase are shown in Table 5. Stage 23 begins the transfer phase

when the storage cask is moved out of the secondary containment airlock and away from the secondary

containment. As in Stage 22, the storage cask is on rollers and cannot be dropped. Similarly, in Stage 25,

the storage cask is being prepared for storage before it is moved to the ISFSI. During Stages 26-32,

the storage cask is lifted above the rollers, over a cushion, and then transferred over concrete, asphalt,

and gravel surfaces before being set down on the concrete storage pad (Stage 33). The transfer phase

ends in Stage 33 with the storage cask being set down on the storage pad. During the transfer phase,

the cask is lifted no more than 0.3 meter (1 foot) above the surface level. At this height, the MPC has

zero failure probability.

The initiating events that can occur during the storage phase are shown in Table 6. The storage phase

comprises only one stage (Stage 34), but many postulated initiating events.

3-5
3.2 Initiating Events That Cannot Affect the Subject Plant

3.2.1 Floods

The flooding analysis at the subject plant is based on the probable maximum precipitation (PMP),

which is defined as the "theoretically greatest depth of precipitation for a given duration that is physically

possible over a particular drainage area at a certain time of the year." The National Weather Service

determines the maximum precipitation by maximizing the parameters of a hypothetical storm

over a drainage area. This maximum is called "probable" because there is a chance, albeit unknown,

that a storm could be more severe. However, to the extent that the maximum can be determined, the PMP

is the most severe case.

To obtain the PMP for the subject site, the rainfall from a hurricane that occurred in Florida during 1916

was considered to occur at the drainage basin at the subject site. The storm was analytically positioned

within meteorological limits over the drainage basin in order to result in the maximum volume

of precipitation at the subject site. The 48.3 centimeters (19 inches) of rain produced an estimated

volumetric flow rate of 17,330 m3/s (612,000 ft3

/s). A stage-discharge model was developed from

the limited amount of data. From the models, a peak discharge of 17,330 m3/s (612,000 ft3

/s) is expected

to result in a flood stage of 32 meters (105 feet).

Hydrologists also determined the effects of winds to create waves. The maximum sustained wind

was taken to be 72 km/hr (45 mph). From these winds, a maximum wave height of 2.0 meters (6.5 feet)

(crest to trough) was predicted. The waves crests would add 1.0 m (3.3 feet) to the maximum flood stage

of 32 meters (105 feet).

In the drainage basin, two dams are in series and a third dam is in parallel with the other two. A break

of the lower serial dam would add about 2,832 m3

/s (100,000 ft3

/s) to the flood waters at the subject site.

This would increase the water level by about 1.2 meter (4 feet). If both serial dams failed, the predicted

wave would be 1.5 meters (5 feet). If the parallel dam failed instead, a wave of 0.3 meter (1 foot)

is predicted. Combining all of these effects results in a maximum flood height of 35 meters (114 feet).

Reference 10 requires updating plants flood analyses using revised maximum precipitation estimates

from Reference 11. The largest hypothetical storm in the drainage area for the subject site has a duration

of 72 hours and produces 63.0 centimeters (24.8 inches) of rain. Given that a storm of 48.3 centimeters

(19 inches) produced a volumetric flow rate of 17,330 m3

/s (612,000 ft3

/s), 63.0 centimeters (24.8 inches)

of precipitation would produce a volumetric flow rate of 22,625 m3

/s (799,000 ft3

/s). According to

the previously determined relation of flood stage to flow rate, a flood stage of 33 meters (110 feet)

would result. Sustained winds of 28 km/hr (45 mph) would raise the flood stage by about 1 meter

(3.3 feet), while dam failure would result in waves at the subject site of 0.3-1.5 meters (1-5 feet).

Therefore, the maximum flood height is 36 meters (119 feet).

The ISFSI at the subject site is at an elevation of 38 meters (126 feet). Thus, the flood waters from

a combined maximum storm, sustained winds, and dam failures would be insufficient to reach

storage casks on the storage pad.

3-6
3.2.2 Tsunamis

Tsunamis, the large ocean waves resulting from earthquakes originating near or in the ocean, occur

mainly in the Pacific Ocean where they can threaten shoreline areas such as the west coast of Washington,

Oregon, and California in the United States. Historically, there have been some tsunamis in the Carribean,

but none on the East Coast that could conceivably affect the subject storage site. The U.S. National

Oceanic and Atmospheric Administration states that tsunamis typically lose their energy within 8.5 km

(5 mi) as a result of frictional losses and impacts with trees, vertical land masses, and other structures.

The subject plant is far enough inland that it will not be affected by a tsunami.

3.2.3 Volcanic Activity

According to the U.S. Geological Survey, volcanic regions are located in California, Oregon,

Washington, Alaska, and Hawaii. The hazards from volcanoes are as follows:

Lava flows are streams of molten rock from an erupting vent during either nonexplosive

or explosive activity.

Pyroclastic flows are high-density mixtures of hot, dry rock fragments and hot gases that move

away from the erupting vent at high speeds. Most pyroclastic flows consist of a basal flow of

coarse fragments that moves along the ground, and a turbulent cloud of ash that rises above the

basal flow. Ash may fall from the cloud over a wide area downwind from the pyroclastic flow.

Tephra are fragments of volcanic rock and lava that are blasted into the air by explosions

or carried upward by hot gases in eruption columns or lava fountains. The fragments can be

in the form of ash or large rocks. Large-sized tephra typically falls back to the ground close to

the volcano, but ash can travel hundreds to thousands of kilometers downwind from a volcano.

Lahar is a hot or cold mixture of water and rock fragments flowing down the slopes of a volcano

or river valleys.

Landslides are large masses of rock and soil that fall, slide, or flow rapidly under the force

of gravity.

Eruption-induced atmospheric shock waves are strong compressive waves driven by rapidly

moving volcanic ejecta. Although most volcanic eruptions are not associated with such waves,

examples are known. Air-shock waves can be sufficiently energetic to damage structures

far from their source. Air shocks can couple to the ground strongly enough to cause damage

to buildings at 100 km (62.1 miles) away from an eruption.

Floods can be produced by melting snow and ice during eruptions of ice-clad volcanoes,

by heavy rains that may accompany eruptions, and by transformation of lahars to stream flow.

Floods caused by an eruption can occur suddenly and have a large volume if other flood

conditions preexist.

The subject plant is far from these volcanic regions, and well out of the influence of volcanic activity.

3.2.4 Intense Precipitation

Intense precipitation can erode soil. However, the ISFSI is designed so that graded land and drains

conduct water away from storage pads. Accordingly, intense precipitation will not affect the storage cask.

3-7
3.2.5 Storage Tanks, Transformers, Barges, Trucks, Railcars,

and Nearby Industrial Facilities

* There are no storage tanks near the route of the storage cask to the storage pad.

* There are no transformers near the route of the storage cask to the storage pad.

* There is no commercial traffic on the nearby river in the vicinity of the site. The only barge

traffic on the river in the vicinity of the storage pad is a snagging barge operated by the Army

Corps of Engineers. This barge passes the site about two times per year, once going upstream,

and the other time going downstream. The barge does not carry explosive or flammable

materials.

Other than a spur track from a railway line, there are no railway lines near the route

of the storage cask from the reactor building to the storage pad or at the storage pad.

Within a radius of 8 km (5 miles) of the subject plant, there are no manufacturing or chemical

plants, refineries, storage facilities, mining operations, military bases, military bombing ranges,

aircraft low-level flight patterns, missile sites, transportation facilities, oil or gas wells,

or underground storage facilities. The area within this radius of the subject plant is mostly rural;

the land is either residential or agricultural. The area is not expected to change significantly

in the foreseeable future.

3.3 Frequencies of Initiating Events

3.3.1 Dropped Fuel Assembly

Data from 1968 - 2002 show that there have been 11 events in which reactor fuel assemblies have been

dropped during movement of the assemblies either prior to loading into a reactor, during loading, or after

removal as spent fuel (Reference 12). None of the 11 dropped fuel assemblies failed.

Data from the U.S. Department of Energy (Reference 13) for the same time period show that the projected

number of permanently discharged spent fuel assemblies is approximately 159,600 (including both BWR

and PWR fuel assemblies). Each permanently discharged fuel assembly has been moved at least twice

during fuel loading and for storage in the spent fuel pool. Given the number of fuel assemblies typically

found in the core of reactors (Reference 14) and the number of reactors in operation (Reference 15),

the number of fuel assemblies in the reactor cores is at least 24,800. (It may be as large as 40,800,

depending on how many assemblies are assumed in the reactors of Reference 14.) The assemblies

in the cores have been lifted at least once. Thus, the total number of lifts is at least 344,000 during

the subject time period (1968 - 2002). Therefore, the probability of a fuel assembly being dropped

is estimated to be 11/344,000 or 3.2x 105.

3-8
3.3.2 Dropped Transfer Cask

The frequency of dropping the transfer cask depends on the number of lifts and the probability of dropping

the transfer cask given a lift. There are two approaches to estimating the drop probability. The first

approach is to perform a reliability analysis of the crane used to lift the transfer cask and an HRA

of workers' actions to rig the cask and operate the crane. The second approach is to obtain an empirical

estimate based on experience with lifting heavy loads. This study used the second approach.

Although the first approach provides more insight and is possibly more accurate, it is much more complex

than the second approach. It must account for both the reliability of the lifting equipment (e.g., crane, yoke)

and the reliability of workers to rig the transfer cask and operate the crane. A fault tree analysis

of the crane equipment must be based on detailed design and operational information (e.g., lift heights,

lift speeds, lift times, movements of the bridge, movements of the trolley). While the fault tree analysis

can be performed with standard methods, the HRA requires further evaluation of human performance issues

relevant to dry cask storage operations and, possibly, further development of HiRA methods. For example,

the kinds of actions that could result in dropping the transfer cask, such as the potential for human error

in attaching the lift yoke to the trunnions at the subject plant, are not well understood, and not every

erroneous action would necessarily cause the transfer cask to fall.

The probability of dropping the transfer cask can be estimated from data on lifts of very heavy loads

at U.S. nuclear power plants. A very heavy load is defined as weighing 27 metric tons (30 tons) or more.

The data used in Reference 16 were evaluated for use in this study. The database of very heavy loads

in Reference 16 is considered relevant to the transfer cask.

Industry-wide experience on moving very heavy loads can be used to make inferences about the risk

at a given nuclear power plant. The application of industry-wide data is not a matter of having identical

plants, but of having relevant aspects of lifting a cask similar enough to analytically treat them the same.

Reference 12 describes the results of a survey of crane operating experience at U.S. nuclear power plants

during 1968 - 2002, and includes the industry's responses. Of 74 respondents to that bulletin, 18

reported that they have a single failure-proof crane, 7 reported that they do not have such a crane, 39 did

not specify which type of crane they have, and another 10 indicated their cranes meet the heavy load

requirements of NUREG-0612 (Reference 16). All crews are qualified according to accepted standards.

Some plants use professional riggers, while other plants use trained plant staff, but all follow the

guidance discussed in Reference 16.

In contrast to some plants in the industry database, the subject plant has a single failure-proof crane,

and professional rigging personnel are contracted to move the cask. Therefore, the industry's estimated

drop frequency is a conservative estimate of the drop frequency at the subject plant.

The drop height probability used in this study is determined from empirical data that reflect lifts

of various heights. In some cases, the probability of dropping the transfer cask may depend on the height

of the lift. Two drop scenario examples are a drop that results when the block of the lift yoke is forced

against the block of the crane boom ("two-blocking") or a drop attributable to a failure of a flawed cable.

In this study, the drop probability is over all possible lift heights.

3-9
As reported in Reference 12, nine sites were visited to collect operational data on very heavy load lifts.

Each site was chosen to represent one type of plant [i.e., nuclear steam supply system (NSSS)]. The type

of NSSS reflects the types of loads that have to be lifted. Although some sites were selected for convenience,

the nine sites are broadly representative of all nuclear power plant sites in the United States.

The 9 selected sites have a total of 19 plants. At each of the 19 plants, the number of very heavy load lifts

was determined from maintenance logs. Table 7 lists the number of lifts, from both refueling and power

operation, at the representative sites in 1968 - 2002. Most, but not all of those lifts were done during

refueling outages. Thus, the number of refueling outages is used to determine the total number of lifts

throughout the population of plants in the United States.

Table 7. Very Heavy Load Lifts at Representative Sites in 1968 - 2002

Representative Number of Very Heavy

Group Type of Plant Site and Units Load Lifts

1 BWR, Mark I, G4 Browns Ferry 1, 2, 3 980

2 PWR, Westinghouse, 4-loop Comanche Peak 1, 2 230

3 PWR, Westinghouse, 4-loop Diablo Canyon 1, 2 344

4 BWR, Mark-I, G3 Dresden 2, 3 554

5 BWR, Mark-III, G6 Grand Gulf 118

6 BWR, Mark-II, G4 Limerick 1, 2 950

7 PWR, B&W Oconee 1, 2, 3 1656

8 BWR, Mark-I, G2 Oyster Creek 504

9 PWR, CE80 Palo Verde 1, 2, 3 2277

The estimated total number of heavy lifts for all operating plants in United States (54,000) was obtained

from Reference 12 and is used as a reasonable surrogate value in determining the load drop frequency.

The number of times that drops occurred was obtained from docketed information and licensee event

reports of all plants in the United States (not just the representative plants). In the subject time period

(1968 - 2002), there were three related drop events. Based on these events, a conservative estimate

of the drop probability in 1968 - 2002 is obtained by taking the number of drops of very heavy loads

as three. The probability of a drop given a lift is then estimated by Equation 1.

Pr{drop} - Ndrops

Ltotal

where: Ndrop, = number of drops of a very heavy load

Ltotal = number of lifts of a very heavy load

Solving Equation 1 with Ndrps = 3 and Ltotai = 54,000, the probability of dropping a cask given a lift

is 5.6 x 10-5.

3-10
The estimated drop probability is considered to be a conservative estimate for two reasons:

The subject plant has a single failure-proof crane to move its transfer cask. Most plants use

non-single failure proof cranes, which are not as reliable as single failure-proof cranes.

The collective experience during the subject period (1968 - 2002) should, therefore, reflect

more unreliability than would be expected from a single failure-proof crane.

The initiating events of dropping a cask in Table 4 are drops of very heavy loads. During

the study period of Reference 12, only one free-fall drop occurred. Two of the three events

used in the numerator of Equation 1 were not drops, but were uncontrolled descents of loads.

Because these events had the potential to damage the transfer cask if it were being lifted,

they were included in the numerator of Equation 1.

3.3.3 Seismic Events

The storage cask will not see significant stresses attributable to a seismic event unless it tips over.

For any given ground acceleration, the likelihood of a tip-over can only increase as the coefficient

of friction between the storage cask and the storage pad increases. From Section 4.1.3.2, the maximum

coefficient of friction is 0.53. With this maximum coefficient of friction and at an acceleration of 1.35 g

[nine times the design-basis earthquake (DBE) of the subject plant], the storage cask will slide,

but will not tip-over. Therefore, only seismic events with ground accelerations exceeding 1.35g

have the potential to tip-over the storage cask.

Figure 8 shows the annual probability of exceedance as a function of ground acceleration attributable to

a seismic event at the subject site in terms of the mean, 15 th, and 8 5 th percentiles of the uncertainty

distribution. The curves in Figure 8 are based on information in Reference 17, which provides frequency

estimates for ground accelerations up to 1.0g at nuclear power plant sites in the Eastern United States.

These estimates are linearly extrapolated on a semi-log scale as the dotted portion of the curves in Figure 8.

From the extrapolation of the curve for the mean of the uncertainty distribution, the annual probability

of an earthquake with ground acceleration exceeding 1.35g is 7x 10-7/year. Since even this ground

acceleration will not cause a tip-over, a conservative estimate of the frequency of an earthquake

that will cause the storage cask to tip-over is 7x 10-7/year.

3-11
AIoLM Probabeii ty

10-

Priobabilities of'

Z111 eýIICtrthqAk ol'

0. 15 ý'Or ureater.

flrohahi Iit\ 0 I* all

CZIatihqulKkC Of

I 35iý Or urcatcr.

10-

10-

IV(-

I ()-II

Figure 8. Annual Probability of Exceedance as a Function of Ground Acceleration Due to a Seismic

Event at the Subject Plant

The potential for soil liquefaction at the site was evaluated. The portion of soil susceptible to liquefaction

at the subject site is within the range of 13.7-27.4 meters (45-90 feet) below grade. Thus, the region

of potential liquefaction at the site is too far below the surface to expect appreciable dispersal of soil

into the cask vents.

3-12
3.3.4 High Winds

The effect of a tornado could be to slide the storage cask on the storage pad, tip it over, or propel a heavy

object into it. The weight of the storage cask that is loaded with BWR fuel assemblies is about 163,293 kg

(360,000 lb). Analysis (Section 4.1.3.2) indicates that a wind speed of approximately 644 km/hr (400 mph)

would be required to cause sliding of a storage cask on the storage pad, while a wind speed of 966 km/hr

(600 mph) would be required to cause the storage cask to tip-over, and winds in excess of 1448 km/hr

(900 mph) would be required to propel a heavy object into a storage cask with enough force to cause

significant damage.

Table 8. Probability of Exceeding Selected Tornado Wind Speeds (Reference 18)

Wind Speed Exceedance Probability

(mph)

73

113

158

207

261

0.746

0.336

0.0766

0.0138

0.00118

The probability of a tornado affecting a storage

cask is determined as follows. Reference 18

gives the probabilities that any given tornado

occurring in the United States will equal or

exceed a specific wind speed (exceedance

probability). The exceedance probabilities for

selected wind speeds were deduced from the

tornado data for 1954- 1983 (Reference 18).

The reported speeds were derived from inspection

of damage caused by tornados and the assignment

of intensity ratings using the Fujita system

of classification (Reference 19). In Figure 9,

the data in Table 8 is plotted on a lognormal

scale. On the ordinate, the wind speeds needed to

affect the storage cask are also indicated.

1 6 'A Win

\\ d *Srý'd

Figure 9. Probability of Tornado Winds

Exceeding Streeds

There is no recorded evidence of tornado wind

speeds beyond 482.8 km/hr (300 mph) from 1955

to the present, the period for which reliable

records have been kept. The highest recorded wind speed reported in Reference 18 is 420 km/hr

(261 mph). There is no experimental or empirical information supporting a physical limit of tornado

wind speeds, however, some scientists believe there may be a limit. Accordingly, for this study,

the maximum wind speed is taken to be 482.8 km/hr (300 mph). The wind speeds required to affect

the cask are significantly larger than that maximum wind speed. Thus, high winds are treated as having

no effect on the storage cask.

3-13
3.3.5 Meteorites

Meteorites continually enter the Earth's atmosphere with initial speeds of tens of thousands of miles

per hour. However, most disintegrate to a large extent as they are slowed by friction with the Earth's

atmosphere. Over 90% of the meteorites are of a stony composition and tend to break apart upon entering

the atmosphere. A smaller number (approximately 6%) are of an iron-like composition and are more

likely to reach the ground with less breakup than the stony meteorites (References 20-22).

A dynamic structural analysis was performed to determine the size, weight, and velocity of a meteorite

that would have just enough energy to fail the storage cask. It was assumed that a meteorite striking

the cask vertically with sufficient kinetic energy to cause the concrete to reach an ultimate strain of 0.003

would result in failure of the cask (Section 10.3 of Reference 23). Meteorites weighing less than 10 tons are

slowed by the Earth's atmosphere such that they strike the earth with terminal velocities varying between

321.9 and 643.7 km/hr (200 and 400 mph), depending on the mass and cross-sectional area

of the meteorite. By iteratively solving the equations used to predict terminal velocity and the kinetic

energy (various combinations of mass and velocity) predicted to cause failure, it was determined that

a stony meteorite with a mass less than 1,542 kg (3,400 lb) would not be expected to cause cask failure.

It was similarly predicted that an iron-like meteorite with a mass less than 1,088.6 kg (2,400 lb) would not

result in failure. Knowing the meteorite diameters associated with the failure masses, the frequency

of meteorites striking the earth that could fail the cask can be estimated. For stony meteorites, for which

there is the most data, meteorites having a diameter corresponding to a mass of about 1,542 kg (3,400 lb)

would strike the earth intact approximately once per year. Considering the differences in density

and frequency of occurrence for stony and iron-like meteorites, iron meteorites weighing 1,088.6 kg

(2,400 lb) would strike the earth intact about once each 1 V2 year. Given the sparseness of data for

iron-like meteorites, it is believed that this latter prediction is less robust than that for stony meteorites.

Based on the above estimates, it is conservatively assumed that the total frequency of strikes from

meteorites that could fail the storage cask is about twice per year. Thus, the following assumption

was used in this assessment:

A meteorite capable of breaching the storage cask strikes the earth about twice per year.

The surface area of the earth is 5.08x 108 km2 (1.96x 108 mi2). Thus, a meteorite with sufficient energy

to fail the cask is predicted to strike the earth with a frequency per unit of surface area of 4.0x 10`

events/yr-km2 (1L.0x 108 events/yr-mi2).

The storage cask is 3.35 m (11 ft) in diameter, giving a surface area on the top of the storage cask

of 8.8 m2 (95 ft2

) or 8.8x 106 km2 (3.41 x 10-6 mi2). Considering a meteorite approaching the storage cask

from directly above, the additional surface area from the sides of the storage cask can be ignored. Using

this as the target area, AT, the impact frequency estimated above as fM, and assuming that a meteorite

that strikes the storage cask breaches it, the annual frequency of a breach of a single storage cask

by a meteorite is estimated by Equation 2:

Fc= fM AT (2)

where: Fc = annual frequency of meteorites breaching a storage cask

fm = annual frequency of meteorite strike which could breach the cask (events/km2)

AT = target area of a storage cask (km2)

3-14
Substituting fM < 4.0x 10 9/yr-km2 and AT = 8.8x 10-6 km2 into Equation 2 results in a frequency

of meteorites breaching the storage cask of Fc < 3.5 x 10-14events/yr.

3.3.6 Lightning Strikes

Lightning activity in the United States is monitored by the National Lightning Detection Network

(NLDN). Data from the monitoring system is collected and disseminated by Global Atmospherics, Inc.

The geographical region of the subject plant has a moderately high occurrence of lightning strikes

(compared to other areas of the United States) (Reference 24). Data from the NLDN taken over the past

10 years indicate that there has been an average lightning strike density of about 4.3 strikes/yr-km2

(11 strikes/yr-mi2). This parameter is often referred to in the literature (Reference 25) as the ground

flash density (GFD). Knowing the GFD and an effective target area for the lightning to strike, the annual

number of strikes on the target may be estimated from Equation 3 (Reference 25):

FL = PGF AET (3)

where: FL = frequency of lightning strikes in a target area per year

PGF = ground flash density, frequency of strikes per year per km2 in the target area

AET = equivalent target area (km2)

There are no structures in the vicinity of the storage pad that are close enough to affect either the strike

location or the frequency of strikes on the casks. The storage casks are the tallest objects in the area

of the storage pad, and their height is accounted for in the calculation of the equivalent area.

The equivalent target area must include an adjustment to account for the height of the target structure

because higher structures present a larger attractive target area than those of less height. AET is defined

as an area of ground surface that has the same annual frequency of direct lightning strikes as the target

structure. For an isolated single structure with a height less than 55 meters (180 feet), the equivalent area

can be determined using the standard approach in Reference 25. This was estimated as a circular area

surrounding the target, with a radius determined by extending a line with 1:3 slope (3 units of horizontal

distance for each vertical unit) from the top of the structure to the ground. The storage cask height is

5.7 meters (18.7 feet); the extended radius of the equivalent target area is then 17.1 meters (56.1 feet)

and, therefore, the equivalent target area is AET = rT (17.1)2 = 918.6 m2 = 9 .2 x 10-4 km2 .

As indicated by NLDN data, the average lightning strike density for the past 10 years in the vicinity

of the subject plant is PGF = 4.3 strikes/yr-km2 (11 strikes/yr-mi 2

). From Equation 3, the lightning strike

frequency is 4.0x 10'3 strikes/year.

3-15
3.3.7 Aircraft

Two categories of aircraft flights that pose a crash threat to the storage cask are (1) crashes of planes

landing or taking off from airfields in the vicinity of the subject plant, and (2) crashes of planes flying

en route in the vicinity of the subject plant (i.e., overflights).

There are four airfields in the vicinity of the subject plant. These airfields are listed in Table 9,

along with other information needed to determine the frequency of crashes attributable to landing

and taking off into the storage cask.

Table 9. Airfields near the Subject Plant Site

Airfield Distance from Subject Plant (mi) Annual Number of Flights

I 16 10,500

II 18 6,100

III 18 6,000

IV 29 6,000

Reference 26 describes a standard approach for estimating the frequency of an aircraft crashing into a

target during takeoff or landing by the use of Equation 4.

n

F A 1 JCj Nj (4)

j=1

where: Ftl = frequency of crashes during takeoffs or landings

A = effective target area for a plane to strike the target on takeoff or landing

Cj = probability per square kilometer of a crash per aircraft movement at airfield j

Nj = number of movements per year at airfield j

n = number of airfields near the subject plant

The summation is over all airfields in the vicinity of the storage cask. The crash parameter, Ci,

is a decreasing function of the distance from the end of the runway of each airfield to the storage cask.

From Reference 26 (Section 3.5.1.6), values of Cj for different values of distance from the runway

are given for distances of 16 km (10 mi) and less. Since the closest airfield to the subject plant

is approximately 25.7 km (16 mi), the values in Reference 26 were extrapolated using a logarithmic plot

and regression analysis to determine values of Cj for the subject plant. The extrapolated value of Cj

for the nearest airfield is estimated to be 3.9x 10's crashes/km2 -movement. This value of CJ

is conservatively used for all four airfields.

3-16
The effective target area, A,,, depends on the dimensions of the storage cask and the plane and the length

of the skid path. The diameter of the storage cask is 3.35 m (11 ft). The largest aircraft which is able to

land at any of the four local airfields is the Gulfstream IV. The Gulfstream IV has two engines on either

side of the fuselage, spaced about 4.2 m (13.8 ft) from centerline to centerline, and each engine is about

1.6 m (5.2 ft) in diameter. In order for the crash to possibly cause cask failure, one of the plane's two

engines must strike the cask. The engines are spaced such that both engines shafts cannot fully strike

the cask together. Evaluating aircraft trajectories that would result in impact by either engine, the lateral

dimension (width) of the target area (perpendicular to the plane's trajectory) is the sum of the centerline

spacing between the engines, the diameter of an engine, and the diameter of the storage cask.

This dimension is 9.2 meters (30.3 feet). A reasonable value for shadow and skid length is 60.8 meters

(200 feet), which takes into account the terrain surrounding the storage cask and the angular window

through which a plane must pass to strike the storage cask. The impact area is considered to be

a rectangle with length equal to the shadow and skid length, 60.8 meters (200 feet), and width equal to

9.2 meters (30.3 feet). Using these values, A, = 5.6x 10.- km2 for the Gulfstream IV.

The frequency, Ft,, is conservatively estimated using the values of parameters from Table 9 and the value

of A,, for the Gulfstream IV. Because Cj is the same for every airfield, Ft, depends only on the total

number of flights, which is 28,600/yr. Substituting into Equation 4 values of A,1 = 5.6x10-4, Cj

3.9x 10s-, and Nj = 28,600 yields a conservative value for frequency of crashes during takeoffs

or landings at airfields in the vicinity of the site of Ft, = 6.2x 10.9 crashes/yr that could impact the cask.

A large aircraft could potentially crash into the storage cask during overflights. To analyze the crash

frequency for overflights, Fof, the method given in Reference 27 is used. In this approach, the frequency

of crashes during overflights is estimated using Equation 5:

F-z = Co Aof (5)

where: Fof = frequency of overflight crashes

Cof = overflight crash rate per square kilometer per year

Aof = equivalent target area for overflight crashes (ki 2)

The value of Cof to be used in Equation 5 for air carriers is 1.6x 10-7 crashes/km2-yr (4x 10-7

crashes/miZ-yr).

Similar to the analysis for takeoffs and landings, the impact area is taken to be a rectangle with length

equal to the skid length, 60.8 meters (200 feet), and width equal to 9.2 meters (30.3 feet). The value of

Aof is 5.6x 10-4 km 2 (2.2x 10-4 mi2).

Substituting in Equation 5 values of Cof = 1.6x 10-7 crashes/km2-yr, Aof = 5.6 x 1 0 1 km2, yields

a conservative value for the frequency of overflight crashes from air carriers that could impact

the storage cask of 9.0x 10" crashes/year.

3-17
The total frequency of aircraft crashes that could impact the cask is the sum of the crash frequency

for takeoffs and landings and the crash frequency for overflight related crashes, or 6.3 x 10-9 crashes/year.

As noted in Section 4.1.3.3, "Strikes on the Storage Cask from Heavy Objects," and Section 7.0, "Risk

Assessment," the probability of MPC failure is assumed to be directly related to the overflight crash rate

for commercial aviation, since only commercial aircraft larger than a Gulfstream IV and traveling at high

velocity, could possibly cause a breach of the MPC on impact. For convenience, it is assumed that all

overflights consist of large commercial aircraft and impact the cask at high velocity. Therefore,

the probability of MPC failure and release, if struck by an aircraft, is, for the purpose of the PRA, equal to

the frequency of overflight crashes (9.0xl0-`) divided by the total frequency of aircraft crashes (6.3x10- 9

)

which is 0.0 14.

An alternative approach for estimating the overflight crash rate is given in Reference 26. The approach

allows for a more location-specific analysis in that it utilizes flight information on traffic in flight

corridors that are approved by the Federal Aviation Administration (FAA). Four such corridors are in

the vicinity of the subject plant, but information on the cumulative number of flights in these corridors

is unavailable because it is not collected by the FAA.

3-18
4. MULTIPURPOSE CANISTER AND FUEL

4.1 Mechanical Loads

4.1.1 Mechanical Load Models

To evaluate the structural behavior of the transfer cask and storage cask for the postulated initiating

events, simplified and conservative analyses were used in many cases. Table 10 presents key parameters

of the storage cask and ISFSI. The analysis methods used included hand calculations based on first

principles, common analytical methods and industry recognized approaches, and solution of the

differential equations of motion for which closed form solutions were obtained. When an analysis

required sophisticated computer codes and large amounts of resources, existing calculations performed

by Holtec and reviewed by the staff were used in addition to independent analyses described in this

report. In some cases the LS-DYNA computer code was used to perform non-linear dynamic impact

analyses (Reference 28). The loads and stresses calculated for all of the examined initiating events were

used in subsequent analyses to determine the probability of failure of the MPC and fuel cladding.

The following assumptions were made in this assessment:

The transfer cask falls in a nearly vertical orientation.

* In most simplified analyses, the transfer cask and the storage cask are considered rigid.

* A drop of the transfer cask in the fuel-handling building occurs on a reinforced concrete floor

supported by a 76-cm (30-in) thick concrete shear wall beneath the floor. The worst case impact

occurs if the cask falls on the area of the floor that is supported by the wall because the stiffness

provided by the wall maximizes the impact acceleration and forces in the MPC.

During the transfer phase, the storage cask is carried by the cask transporter at a height

of 0.3 meter (1 foot) above the ground surface.

If the storage cask were dropped from the cask transporter during the transfer phase, the cask

transporter is traveling in the direction of the open end of the cask transporter.

An analysis was done to evaluate the response of the MPC in the cask to the DBE at the ISFSI of the

subject plant; this was set equal to one-half of the subject plant's seismic margin earthquake (/2 SME)

ground response spectrum. Seismic events were evaluated for different levels of seismic forces with

different coefficients of friction between the bottom steel plate of the cask and the concrete pad using

three-dimensional (3-D) coupled finite element models of the cylindrical cask, a flexible concrete pad,

and an underlying soil foundation. Two coefficients of friction are considered at the interface between

the cask bottom steel plate and the concrete pad. A lower-bound friction coefficient of 0.25 is used in

investigating the sliding of the storage cask, and an upper-bound friction coefficient of 0.53 is used in

examining the possibility of storage cask tip-over.

4-1
Table 10. Storage Cask and ISFSI Parameters for Load and Stress Calculations

Storage Cask

Parameter Value

Diameter of the storage cask 3.4 meters (11 feet)

Height of the storage cask 5.9 meters (19 feet)

Height of the center of gravity 3.0 meters (10 feet)

Weight of storage cask, MPC, BWR fuel assemblies 163,293 kg (360,000 lb)

ISFSI

Parameter Value

Thickness of concrete 0.61 m (2.0 ft)

Concrete compression strength 20.7 MPa (3,000 psi)

Yield strength of reinforcement top and bottom 413,685 kPa (60,000 psi)

Soil effective modulus of elasticity 193,058 kPa (28,000 psi)

For missiles caused by tornado or flood, the Spectrum II missiles identified in the Standard Review Plan

(SRP, Reference 29) are considered. Based on the SRP, the most severe missile that has the potential to

cause sliding or tip-over of the storage cask is an automobile weighing 1,810 kg (3,990 lb) at a velocity

of 59 m/s (132 mph). An automobile of this weight is comparable to one of today's common sport utility

vehicles. Based on studies of the behavior of heavy objects in high winds such as tornadoes, it has been

determined that the maximum velocity of an automobile propelled by a tornado may be approximated by

taking 0.2 times the maximum velocity of the tornado wind speed (SRP Section 3.5.1.4). An automobile

velocity of 59 m/s (132 mph) would then correspond to a maximum tornado wind speed of about 295 m/s

(660 mph) indicating that the analysis based on the SRP model is quite conservative. The frequency of

tornado winds in excess of 295 m/s (600 mph) at the storage site is estimated to be 4x 10' events/yr.

4.1.2 Response of the Transfer Cask to Loads

Analyses were performed for the initiating events to determine the stresses in the MPC while being

moved in the transfer cask. Although the transfer cask can be carried on its side or in the upright position

when the MPC is sealed, at the subject plant it is always carried in the upright position. All transfer cask

drop analyses are conducted assuming an upright cask position. Thus, the following assumption was

used in this assessment:

The transfer cask is always carried in the upright position.

4-2
4.1.2. 1 Drop of the Transfer Cask into the Cask Pit

In Stage 3, the transfer cask, which is filled with spent nuclear fuel assemblies, is lifted from the cask pit.

The depth of the cask pit is 11.4 meters (37.5 feet). When the transfer cask is out of the water, it is lifted

an additional 1.5 m (5 ft) to a height of 13 meters (42.5 feet). It takes 0.5 m (1.5 ft) to get to the level of

the refueling floor and an additional 1.1 m (3.5 ft) to get high enough above the refueling floor to clear a

railing on its way to the preparation area (Stage 4).

If the transfer cask were accidentally dropped while being lifted in Stage 3, the cask would fall a total of

1.5 meter (5 feet) through the air, drop through 11.4 meters (37.5 feet) of water, and impact the concrete

floor of the pit. The analysis included buoyancy and drag effects, which slightly reduce the velocity and

the energy at impact. Because the subject plant has a single-failure-proof crane and professional rigging

personnel are contracted to move the cask, the consequences of a drop of the transfer cask on the

structural integrity of the cask pit do not have to be evaluated. However, should the cask be accidently

dropped, based on information available for design of the reinforced concrete spent fuel cask pit, the

results of the impact analysis show that the shear stresses developed in the concrete exceed the ultimate

capacity of the concrete floor. Therefore, cracking and possibly more severe damage or failure of the

spent fuel cask pit floor may occur if the cask is dropped from the maximum height of 13 meters

(42.5 feet) above the floor of the cask pit. It is assumed that the consequences to the integrity of the

MPC and fuel from such a drop are the same as for a free drop in air of 12.2 meters (40 feet) regardless

of whether the cask pit floor fails or not. For a drop onto concrete from this height, Table 13 shows that

100% of the fuel cladding is breached. Since the MPC is not sealed during this stage, the probability of a

release from the MPC is 1.0.

4.1.2.2 Tip-Over of the Transfer Cask

In Stages 4-17, the crane is moving the transfer cask over the refueling floor. Analyses were performed

to determine the effects of this movement at the maximum speeds; the bridge moves at 15 meters/min

(50 feet/min); the trolley moves at 3 meters/min (10 feet/min); the hoist moves at 1.3 meters/min

(4.2 feet/min). These velocities are very small and will have negligible effect on the conditions required

for the cask to tip-over.

If the crane were not moving, the transfer cask would tip-over only if it is dropped at an angle such that

the center of gravity of the cask is over the point of first contact with the floor. This angle requires that

the cask axis be oriented at less than 67' with the horizontal. This angle cannot be reached unless the

drop height is more than 40 cm (16 in.) measured at the center of the transfer cask base. From Table 1,

the transfer cask is lifted more than 40 cm (16 in.) in Stages 4, 11, 12, 13, 15, 16, and 17. Accordingly,

the transfer cask may tip-over if it is dropped in any of those stages at an angle such that the center of

gravity is over the point of first impact. It will not tip-over if it is dropped in Stages 5-10 or Stage 14.

A discussion of the consequences associated with cask tip-over is found in Section 4.1.3.2.

4-3
4.1.2.3 Drop of the Transfer Cask onto the Storage Cask

The accidental vertical drop of the transfer cask is analyzed during Stage 18 if the cask were dropped

while being lowered from the operating floor, striking the storage cask. The maximum vertical drop in

this case is 24.4 meters (80 feet). For this analysis, a detailed finite element model is used that includes

the transfer cask, MPC, fuel assemblies, fuel basket, fuel basket support, and storage cask. Drop heights

of 1.5, 12.2, and 24.4 meters (5, 40, and 80 feet) were selected for analysis. Analysis of this model using

the LS-DYNA computer code calculated the maximum effective plastic strain in the MPC which was

used in the weld failure analysis to calculate the probability of MPC failure. In Section 4.3.2, Table 12

summarizes the MPC failure probabilities for all of the drop scenarios evaluated. For the drop of the

transfer cask onto the storage overpack, the probability of MPC failure is very small and, as expected,

varies with drop height. For a drop of 24.4 meters (80 feet), the probability of MPC failure is 0.0002.

This low probability of failure is attributable to the fact that the drop of the rugged transfer cask onto the

storage overpack results is a relatively soft impact, with most of the impact energy being absorbed by the

storage overpack.

For the purpose of the PRA, fuel cladding failure is linked to a 1.0% failure strain, which is a typical

lower bound value for high-burnup fuel with circumferential hydrides. Table 13 summarizes fuel

cladding failure for various drop heights and scenarios. For the 24.4-meter (80-foot) drop of the transfer

cask onto the storage overpack, 100% of the fuel cladding is assumed to fail. Therefore, the probability

of release of radionuclides from the MPC is set equal to the probability of MPC failure.

4.1.2.4 Drop of the Transfer Cask onto a Concrete Floor

The accidental vertical drop of the transfer cask while it is being lowered from the operating floor to the

ground floor were analyzed. If the transfer cask were to drop from the operating floor and strike the

concrete floor at ground level, the maximum drop would be 30.5 meters (100 feet). To perform the

analysis, a detailed finite element model was developed that included the transfer cask, MPC, fuel

assemblies, fuel basket, fuel basket support, concrete floor, and the wall beneath the floor. Drop heights

of 1.5, 12.2, 21.3, and 30.5 meters (5, 40, 70, and 100 feet) were selected for analysis. Analysis of this

model using the LS-DYNA computer code calculated the maximum effective plastic strain in the MPC

which were used in the weld failure analysis to calculate the probability of MPC failure. In Section

4.3.2, Table 12 summarizes the MPC failure probabilities for all of the drop scenarios evaluated. For the

drop of the transfer cask onto the concrete floor, the probability of MPC failure is small and, as expected,

varies with drop height. For a drop of 30.5 meters (100 feet), the probability of MPC failure is 0.02. The

probability of failure is relatively low because the drop of the rugged transfer cask onto the concrete floor

results in a soft impact, with most of the impact energy being absorbed by the concrete floor and wall

beneath the floor.

For the purpose of the PRA, fuel cladding failure is linked to a 1.0% failure strain, which is a typical

lower bound value for high-burnup fuel with circumferential hydrides. Table 13 summarizes fuel

cladding failure for various drop heights and scenarios. For the 30.5-meter (100-foot) drop of the

transfer cask onto the concrete floor, 100% of the fuel cladding is assumed to fail. Therefore, the

probability of release of radionuclides from the MPC is set equal to the probability of MPC failure.

The relative deformation between the transfer cask top lid and the top of the MPC was reviewed to

ensure that there is no contact/impact between the transfer cask and the MPC. A review of the relative

displacement, throughout the time history for the maximum drop (worst case), demonstrated that a gap is

maintained, and therefore, the transfer cask top lid does not impact the top of the MPC.

4-4
4.1.2.5 Drop of the Transfer Cask onto the Refueling Floor

When the transfer cask is moved from the fuel pool to the handling area, the transfer cask may fall from

a height of 0.3 meter (1 foot). For this drop, the refueling floor is expected to hold the transfer cask since

the floor has beams, girders, or concrete walls underneath that lie along the load path. For a drop of

0.3 meter (1 foot), Tables 12 and 13 show that the probability of MPC failure is 1x 106 , and there is

no fuel cladding failure.

4.1.2.6 Drop of the MPC While Being Lowered into the Storage Cask

When the MPC is lowered from the transfer cask into the storage cask (Stages 20 and 21), it could

possibly fall a maximum of 5.8 meters (19 feet). This fall is analyzed using a finite element model that

includes the MPC, storage cask inner shell, storage cask vertical channel sections, and the storage cask

bottom lid (which serves as the target plate). The MPC model, its material properties model, fuel

assemblies, and fuel basket and fuel basket support model are the same as those utilized in the drop of the

transfer cask onto the concrete floor. The top plate of the storage cask bottom lid is modeled as a rigid

target plate for the vertical impact of the MPC and its internal components. The storage cask vertical

wall is modeled using a rigid shell to represent the storage cask inner shell, which surrounds the MPC

shell wall.

A dynamic analysis using LS-DYNA was performed to calculate the maximum effective plastic strain in

the MPC for input into the weld failure analysis to calculate the probability of MPC failure. From Table

12 in Section 4.3.2, the probability of the MPC failing when it is dropped during Stages 20 and 21 is

0.28. For the purpose of the PRA, fuel cladding failure is linked to a 1.0% failure strain, which is a

typical lower bound value for high-burnup fuel. Table 13 summarizes fuel cladding failure for various

drop heights and scenarios. For the 5.8-meter (19-foot) drop of the MPC into the storage overpack,

100% of the fuel cladding is assumed to fail. In contrast, for the 6. 1-meter (20-foot) drop of the transfer

cask onto concrete, fuel cladding failure is not expected to occur. These different outcomes for

essentially the same drop height are due to the fact that the 5.8-meter (19-foot) drop is a much harder

impact and results in a very different mode of fuel rod buckling, as explained in Appendix C. Because

fuel failure occurs for the 5.8-meter (19-foot) drop, the probability of release of radionuclides from the

MPC is equal to the probability of MPC failure, which is 0.28.

4-5
4.1.3 Response of the Storage Cask to Mechanical Loads

4.1.3. 1 Drop of the Storage Cask onto Concrete, Asphalt, or Gravel Surfaces

During Stages 29-33, the storage cask is moved by the cask transporter to the ISFSI at a height of

0.3 meter (1 foot). A drop of the storage cask from this height onto concrete, asphalt, and gravel, was

also evaluated. The two bounding cases considered were a drop without any rotation (i.e., end drop) and

a drop causing the maximum rotation.

During an end drop, calculations show that the frictional resisting force of the storage cask

sliding against the concrete was much less than the overturning force; thus, the storage cask

would slide and not tip-over regardless of the velocity of the cask transporter.

Calculations demonstrate that it is geometrically impossible for the rotation of a dropped storage

cask to be large enough to cause a tip-over. The maximum amount of rotation possible is 5.140

(with respect to the ground), which occurs when one of the two vertical supports holding the

storage cask at the top breaks and the storage cask makes contact with the ground at one edge.

Since this angle is less than 29.5', the storage cask will not tip-over. Furthermore, the force

required to tip-over the storage cask from the rotated orientation was calculated to be greater than

the frictional resisting force. Therefore, the storage cask would slide but not tip-over.

In both the end drop and the rotation drop, the cask transporter can cause the storage cask to slide, not

tip-over. The maximum stress in the MPC shell for a drop of 0.3 meter (1 foot) onto a concrete pad is

calculated to be 53.3 MPa (7,732 psi). This stress produces extremely small strains in the MPC and in

Section 4.3.2, Table 12 shows that the probability of failure is less that 1x10 6 .

If the cask transporter is traveling at the maximum operating speed of 0.64 km/hr (0.4 mph) and the

storage cask that it is carrying strikes another storage cask on the storage pad, the storage cask on the

storage pad will slide but not tip-over. If the cask transporter does not stop and continues pushing the

struck storage cask, the storage cask on the storage pad could slide into another storage cask or be pushed

off the storage pad onto the surrounding gravel area. In any case, for the struck storage cask, the

maximum stresses in the base plate, lid, and the MPC shell will be much less than the stresses of a drop

of 0.3 meter (1 foot).

The stresses resulting from a drop onto the asphalt or gravel surface are bounded by a drop onto the

ISFSI concrete pad. This is because asphalt and gravel are more flexible and have a lower strength and

modulus than concrete, and neither of these surfaces have steel reinforcing bars. When the storage cask

is dropped from a height of 0.3 meter (1 foot) onto concrete, asphalt, or gravel, the failure probability of

the MPC is less than 1.0x106.

4-6
4.1.3.2 Tip-Over of the Storage Cask

Hypothetically, if the storage cask were to tip-over because the center of gravity of the cask passes over

the point of rotation without an initial force or velocity, the maximum circumferential stress in the MPC

shell is approximately 421 MPa (61,044 psi) at 4.53 meters (14.9 feet) above the bottom of the MPC.

This is based on Reference 2 where the stress in the MPC shell was calculated using a finite element

model of a slice of the MPC and fuel baskets subjected to an acceleration of 45g. Under this loading, the

MPC and fuel baskets deform outward until portions of the MPC shell reach the rigid boundary of the

storage cask inner steel shell. This analysis is consistent with the methodology of the American Society

of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code and is based on a linear elastic

approach. The maximum calculated stress in the MPC shell is less than the allowable stress intensity of

450 MPa (65,200 psi) based on the Level D Service Condition of the ASME Code (Reference 30).

ASME Code Division 1, Subsection NG, Paragraph NG-3225 and Appendix F, Paragraph F-1331, limit

the primary membrane (general or local) plus primary bending stress intensity to 150% of the general

primary membrane stress intensity. For the purpose of evaluating MPC weld failure, the allowable stress

limit of 450 MPa (65,200 psi) was used.

Since the weld failure criterion is based on plastic strain, the elastically calculated maximum stress must

be converted to a plastic strain that could have reasonable resulted from a non-linear analysis of the same

event. Also, since the "damage" that can be inflicted on the MPC during a tip-over event is energylimited, an energy balance approach can be used to estimate the maximum plastic strain. In this

approach, the strain energy per unit volume absorbed at the point of the elastically calculated maximum

stress is equated to the strain energy per unit volume absorbed through elastic-plastic deformation of the

material using the idealized engineering stress-strain curve in the LS-DYNA program that was used for

the drop analyses.

Using this approach, and the data in Appendix A, Table A.2, the calculated maximum plastic strain is

0.0031 in/in. Since the maximum membrane plus bending stress must have occurred on the boundary,

the maximum value that the triaxiality factor can be is 2.0 (see Appendix B, Section B. 1.4). The

maximum adjusted plastic strain is, therefore, 0.0062 in/in. Assuming that the maximum plastic strain

occurs at one of the axial or circumferential welds, the probability of weld crack initiation from

Appendix B, Table B.2 is less that lxl0-6.

Two initiating events, namely seismic events and high winds, could potentially tip-over the storage cask:

Seismic events. A seismic event is evaluated for various levels of seismic forces with different

coefficients of friction between the bottom steel plate of the cask and the concrete pad using 3-D

coupled finite element models of the cylindrical cask, a flexible concrete pad, and an underlying

soil foundation. When higher levels of seismic excitations are used in the analyses, the

coefficient of friction at the cask/pad interface plays an increasingly important role in the sliding

and rotational behavior of the storage cask. For a minimum friction coefficient, the storage cask

exhibits a translational motion without much rotation. If the friction coefficient is 0.25 and an

earthquake occurs equal to 11 times the DBE of the ISFSI (i.e., 11 DBE), the seismic excitations

will cause the storage cask to slide 0.93 meter (36.6 inches). Even though the sliding storage

cask may hit a neighboring storage cask, neither storage cask will tip-over. Assuming a

maximum friction coefficient, the storage cask experiences more rotational movement, but does

not tip-over. When the friction coefficient is 0.53 and an earthquake occurs equal to nine times

the DBE of the ISFSI (i.e., 9 DBE), the storage cask top will move 1.19 meter (47 inches), and

the storage cask base will slide 0.84 meter (33 inches). Even though the sliding storage cask may

hit a neighboring storage cask, neither storage cask will tip-over.

4-7
High winds. The wind speed required to slide the storage cask is at least 644 km/hr (400 mph),

and the wind speed required to tip-over the storage cask is at least 966 km/hr (600 mph). The

design wind speed for the plant is 580 km/hr (360 mph). Detailed calculations are discussed in

Section A.5.9.6.

4.1.3.3 Strikes on the Storage Cask from Heavy Objects

Vehicle impact. A car or a truck crashing into the cask while stored on the ISFSI pad is not analyzed

directly. However, by extrapolating from the analyses for sliding and tip-over attributable to missile

impact, we can conclude that a vehicle weighing 4,536 kg (10,000 lb) and traveling at a speed of

241 km/hr (150 mph) speed will slide the cask less than 66 cm (26 inches), and will not cause tip-over.

This conservatively considers that the impact for a tip-over occurs at the top of the cask for tip-over and

the lowest coefficient of sliding friction applies for sliding. The MPC remains intact.

Accidental crash by an aircraft. Based on the results of the aircraft impact analysis (Section A.4.9.3),

the probability of breaching the MPC, if struck by Gulfstream IV aircraft during landing or takeoff,

is extremely small and is assumed to be zero. However, many commercial aircraft overflying the site

would be larger than a Gulfstream IV and could impact the cask at high velocity. Since the makeup

and characteristics of commercial aircraft overflying the site are not known and have not been evaluated,

it is conservatively assumed for the purpose of the PRA that all commercial aircraft overflying the site

are larger than a Gulfstream IV and that the probability of MPC breach due to the impact of an overflying

commercial aircraft is, for the purpose of the PRA, 1.0. Therefore, the probability of MPC failure and

release, if struck by an aircraft, is, for the purpose of the PRA, equal to the frequency of overflight

crashes (9.0xl01 1, see Section 3.3.7) divided by the total frequency of aircraft crashes (6.3x 109 , see

Section 3.3.7), which is 0.014.

Wind-driven missiles. Missiles generated by a tornado have been analyzed. Two different cases of

wind-driven missiles are discussed in Section 4.1.1 of this report. The first case analyzed is based on the

SRP section addressing missiles generated by natural phenomena (Reference 26). This involves an

impact from an automobile weighing 1,810 kg (3,990 lb), traveling with a velocity of 59 m/s (132 mph).

Such missiles will not penetrate the storage cask, and will not cause sliding or tip-over. The concrete

barrier thickness required to prevent penetration is 72.4 cm (28.5 inches). The entire wall (concrete and

steel liners) of the storage cask is 75 cm (29.5 inches) thick. However, an automobile as a missile

[1,810 kg (3,990 lb) with a velocity of 212 km/hr (132 mph)] may slide the storage cask if the coefficient

of friction between the concrete pad and the storage cask were less than or equal to 0.26. This is near the

low end of the coefficient of friction range. If the coefficient of friction were greater than 0.26, the

storage cask will not slide. It is estimated the stresses in the cask attributable to missile impact will be

much smaller than the tip-over case discussed in Section 4.1.3.2.

In addition, a second study assumes a larger vehicle with a weight of 3,251 kg (7,160 lb), propelled by a

very severe tornado having a wind speed of 134 m/s (300 mph). This wind speed is slightly higher than

any tornado wind speed recorded to date. To assess the potential impact of such a missile striking a

storage cask, the kinetic energy of this missile is compared to that of an aircraft crash. Since the kinetic

energy of the large vehicle wind-driven impact would be less than 1% of the kinetic energy of a large

aircraft impact, it is concluded that the impact of a very large vehicle traveling as a tornado missile at a

relatively high velocity will not result in cask failure.

4-8
4.1.3.4 Shockwaves on the Storage Cask from Explosions

Tanker trailer. The nearest public highway is at least 914 meters (3000 feet) from the ISFSI. The

shockwaves caused by a tanker trailer containing explosive materials (e.g., gasoline, liquid, natural gas)

at this distance from the storage cask will not jeopardize the integrity of the MPC. Reference 31 was

utilized to evaluate this event. The cask is designed for a 69 kPa (10 psi) peak transient external

pressure. At 914 meters (3,000 feet), the peak pressure attributable to the explosion would be less than

10 percent of the pressure magnitude for which the cask is designed. Therefore, the MPC will remain

intact.

Gas pipeline. The nearest gas pipeline (natural gas) is 7.2 km (4.5 miles) from the subject site. The

shockwaves caused by the pipeline explosion will not jeopardize the integrity of the MPC. The methods

described in References 31 and 32, which analyzed a tanker trailer explosion, were utilized to evaluate

this event. To compare a pipeline explosion with the tanker trailer explosion, Equation 1 in Reference 31

was used for a distance of 7.2 km (4.5 miles). This resulted in a 6.9 kPa (1 psi) transient pressure on the

storage cask, which is negligible. Based on these assessments, it is concluded that the storage cask will

remain intact, and a pipeline explosion will not jeopardize the integrity of the MPC.

4.2 Thermal Loads

4.2.1 Heating During Normal Operation, Blocked Vents, and Fires

4.2. 1.1 Heatup Model

The thermal analysis employed the use of Fluent (Reference 33), a computational fluid dynamics (CFD)

program, to model the relevant physical phenomena resulting from both the fire and blocked vents

scenarios of the dry cask. Fluent is a general-purpose CFD program with the ability to model a wide

range of practical problems by solving the conservation equations for mass, momentum, energy and

chemical species using a control volume based, finite difference method. Complex two-dimensional

(2-D), axisymmetric and 3-D geometries can be modeled. Fluent has been validated for a wide range of

flow conditions, including laminar and turbulent flow in various geometries, natural and forced

convection problems, heat exchange flow, and combustion phenomena. Additionally, the NRC staff

reviewed Holtec International's confirmation of Fluent code's capability to reliably predict temperature

fields in the dry storage application using an independent full-scale test on a loaded cask (Reference 34).

CFD analyzes fluid flow, heat and species transport in two or three dimensions. In the present analysis,

the computational domain includes the storage cask and the ambient conditions. The domain is divided

into control volumes, and the governing conservation equations are solved for each control volume,

providing detailed information of all the flow variables. The conservation equations are reduced to a set

of algebraic equations by discretization, which are then solved by numerical techniques.

4-9
Figure 10 shows the geometrical configuration and computational grid generated for the storage cask.

The MPC is inside the storage cask. The fuel basket containing fuel assemblies is inside the MPC. A

combination of heat source and porous media was used to model the MPC. The chosen grid was refined

until a grid independent solution was obtained. Figure 10 shows two exploded views to depict the

airflow channel around the MPC. Arrows indicate the airflow direction. The sides and top boundaries of

the computational domain were extended far away from the dry cask structure to safely apply pressure

boundary conditions at the top and the sides. Axisymetry was used in this model. Thus, the vents in the

computational model appear narrower than the physical model so that the total airflow cross-sectional

area in both models are identical.

Stolage

Caslk C'omputational

Grid

M Ft

6.07

5.9

5.0-

4.0-

3.0-

2.0-

1.0-

-20.019.3

-18.0

-16.0

-14.0

-12.0

-10.0

- 8.0

- 6.0

- 4.0

- 2.0

-0

3.4 m111 _

(11.2 tft)

Figure 10. Geometrical Configuration and the Computational Grid Generated for the Storage Cask

4-10
Normal storage conditions were obtained first using steady state analysis to be used as initial conditions

for the transient analysis. The MPC-68 design holds 68 BWR fuel assemblies. It consists of a fuel

basket inside a cylindrical shell. Within the fuel basket, a grid work of stainless steel plates forms an

array of square cross-section compartments, each holding a single fuel assembly. The fuel basket is

positioned within the shell by basket supports. To allow thermal expansion of the fuel basket, small gaps

exist between the fuel basket and the basket supports. Heat is continuously transported from the MPC

interior to the periphery of the fuel basket by conduction through the stainless steel plates. Materials

present in the MPC include stainless steels (Alloy X), Boral neutron absorber, aluminum Alloy 1100 heat

conduction elements, and helium. Materials present in the HI-STORM storage cask include carbon steels

and concrete.

Effective values of thermal conductivity, density and heat capacity were taken from References 34 and 35.

These effective thermal properties were evaluated using ANSYS (a finite element code). The analysis

looked at the combined effect of conduction and radiation inside the MPC. For accurate analysis, all the

properties in Fluent were supplied as a function of the temperature and mass fraction. Kinetic theory was

used to derive the thermo-physical properties of pure substances like air or helium. The normal storage

condition as well as the blocked vent scenario were modeled in a steady mode. The fire scenario was

modeled in a transient fashion using a second-order time marching scheme for 3 hours.

In all the cases, the convective terms in the transport equations used a second order forward differencing

scheme. Semi-Implicit Method for Pressure-Linked Equations (SIMPLE) was used to bridge the

momentum to the continuity equation. All modes of heat transfer (i.e., conduction, radiation, and

convection) were taken into account. Porous medium was used to model the axisymmetric inner basket

region and the peripheral gap between the basket and the MPC shell. Turbulence was modeled using low

Reynolds k-epsilon model. In this model, the conservation equations for turbulence kinetic energy and

its dissipation were solved. A full buoyancy effect was added in the turbulence production and its

dissipation. Standard wall functions were not used to bridge the sub-laminar layer near the walls to the

fully turbulent flow in the core region as in widely used standard k-epsilon model. Instead, full

integration to the wall was carried out using fine mesh near the wall. The discrete ordinates method was

used to model radiation. Only surface-to-surface radiation was considered. Absorption and scattering in

the gas phase was neglected (no participating media was considered). All walls were modeled as thick

walls. Conduction was modeled through each wall in all directions. The computational domain was

extended away from the cask boundaries. No boundary conditions were assigned at the inlet and outlet

vents for the airflow. The external walls of the cask were allowed to exchange heat with the surrounding

atmosphere through convection and radiation.

The heat source representing the decay heat from the fuel rods was assumed to not be uniformly

distributed. Instead, it was distributed as a function of the axial burnup peaking factor. The axial burnup

peaking factor for BWR fuel is provided in Reference 34. The total heat decay used was 21.4 kW.

In the fire scenario, the combustion of the jet fuel was considered. An air-fuel ratio of 14.67 was used.

Product gases (flue gas) arise from combustion. As a result, an additional conservation equation was

solved to track the species generated in the combustion. The solution of the problem involved a total of

eight conservation equations to be solved simultaneously.

4-11
4.2.1.2 Response of the Storage Cask to Heat Loads

Normal Steady-State Operation

The results of the steady state for normal storage operation were compared to the analysis performed in

Reference 34, and shown in Table 11.

Table 11: Comparison of Results for Normal Storage Conditions

Reference 34 Recalculated

Parameter Analysis CFD Analysis

Maximum cladding temperature (*C) 393 298

Average temperature inside the MPC (°C) 250 179

Maximum MPC shell (°C) 175 180

Normal operating pressure inside the MPC (kPa) 564 564

The maximum fuel temperature

values in Reference 34 are more Teipeirttuie

conservative. In Reference 34, to be (c)

conservative, atmospheric pressure 500

was used inside the canister. This 400,

lower pressure results in low helium

density and, thus, less mass to cool 300

the inside of the canister.

Additionally, at higher pressure, the 200

hotter helium is at the top of the 100o

canister instead of the middle as seen

in Reference 34. This upward 0o Metrs 1 l•ight

temperature profile shift as seen in 4 6 Alon, the

Figure 11 increases the temperature . ..8..... 1 2ask . U 4 8 1) 6 2

of the MPC wall at the top as seen in 1 1 2o

Table 11. ---- ----

II

MPC centerline fuel temperatures asa function of axial direction are

shown in Figure 11. A maximum

temperature of 298 'C (568 'F) was

reached in the fuel region with the Figure 11. Centerline Fuel Temperature

highest bumup peaking factor. • Normal Operating Temperature

w Temperature with Blocked Vents

4-12
Figure 12 illustrates the temperature

variation as a function of the axial

position at three radial positions across

the cask concrete overpack. As

expected the inner surface of the

concrete will be at higher temperature

for the normal storage operation.

Figure 12 also reflects the increase of

the air temperature as it flows upward

in the cooling passage between the

MPC and the cask. The maximum

temperature of the MPC shell under

normal conditions is 180 °C as shown

by the time zero temperature in

Figure 15.

["cnpciaturc

(- C)

1007-

80-

60-

40-

20-

0-

605tý

1 2 3 4 5 ~6 Ao~ii mcý01"ý I Iciu,111

Alo7io thc

() 4 I I'- 2 16 2(. storauc

I - L

Figure 12. Storage Overpack Concrete Temperature During

Normal Operation

• Inner Boundary Temperature

* Middle of Concrete

A Outer Border Temperature

4-13
Blocked Vents

Steady-state and transient analyses 250

were performed, assuming that all

the vents were closed. The transient 200-

analysis showed that it will take 150-

approximately 25 days to reach

steady-state values. The analysis 100-

was performed using the Fluent 50_

CFD program (Reference 33).

0 I Meter Nit I Icight

1 2 3 4 M Aomig thc

In Appendix C, Figure C.21 shows Storag._e

that a maximum temperature of '4 S ' I_ 2.

461 0C (862 'F) was reached in the __

fuel region with the highest burnup

peaking factor.

In Appendix B, Figure B. 12 shows

that the maximum temperature in

the MG shell for the blocked vents Figure 13. Storage Overpack Concrete Temperature with

scenario was 283 'C (542 'F). Bottom Vents Blocked

Figure 13 shows the storage cask Inner Boundary Temperature m Middle of Concrete

concrete temperature variation as

function of the axial direction. A A Outer Border Temperature

maximum temperature of 216 'C

(421 'F) is reached in the concrete

overpack. None of the blocked vent or other thermal scenarios resulted in conditions that could fail

the MPC.

4-14
Fires

An accident scenario was

developed that takes into account

a crash of a Gulfstream IV

aircraft from a nearby airport, and

used, as a key assumption, the

volume of fuel this aircraft

carries. An 82.5 MW (heat

release rate) engulfing external

fire was, therefore, modeled in a

transient mode. The intensity of

the fire is based on a conservative

bounding estimate using the

amount of fuel that could

combust over a 3-hour period as a

result of the crash of a

Gulfstream IV, the largest

private/commercial aircraft

that takes off and lands at any

of the four airfields located

around the site; however,

a credible fire scenario will be

less than 30 minutes.

As the fire started, hot gases

emanating from the fire

surrounded the entire cask. The

stream introduced to the cooling

passage is now hot and consists

of combustion product gases as

well as air. The analysis was

lVempet-auic

400 -

350 -

300 -

250 -

200 -

50-

0-

- 7l•._

I I I II

0 1 2 3 4 5 6

Metens 1Ie60h1

A~ion the

stoiahee I I I I I I I I I I I F ee, 0 4 8 12 16 20 I i I

I I I

I I I

I I

I I I

I-----------

()Oerpack

Figure 14. Temperature of Storage Cask after 3-hour Fire

with an Energy of 82.5 MW

* Inner Boundary Temperature

* Middle of Concrete

A Outer Border Temperature

performed using the Fluent CFD program (Reference 33). A peak temperature of 1,200 'C (2,433 IF)

was reached in the fire region. In the fire region immediately surrounding the cask, temperatures closer

to the outer wall of the cask are significantly lower than the maximum temperature reached in the fire

zone. As cooler air mixes with hot gasses, it flows upward along the surface of the cask.

Figure 14 shows the temperature variation as function of the axial direction in the cask concrete structure

at three radial locations at the last time step of the analysis (180 min). The concrete wall inner border

temperature (middle curve) reflects the temperature change of the hot mixture as it flows up the cooling

channel. The mixture will lose heat to the MPC as it flows upward. The external border of the cask

concrete layer (upper curve-triangles) and the middle of the concrete (lower curve-squares) are more

affected by the hot gases bordering the dry cask from the outside (References 36-38).

4-15
Figure 15 depicts the rise in MPC shell temperature during a 3-hour (180-minute) fire scenario. The

maximum temperature reached in 180 minutes was 352 'C (666 'F). Note that steady state (i.e.,

equilibrium temperature) had not been reached at the 3-hour point. In Appendix C, Figure C. 19 shows

that the maximum fuel temperature reached was 298 'C (568 'F).

Temperature (OC)

400-

350-

300-

250-

200

150

100

50

0 50 100 150 200

Figure 15. Maximum MPC Shell Temperature in a Fire Scenario

4.2.2 Heating During Lightning Strikes

4.2.2.1 Lightning Dissipation Model

Reference 2 states that lightning would cause no damage to the storage cask. The underlying assumption

is that the lightning energy is discharged to ground. Reference 39 (on requirements of the lightning

protection code addressed in Reference 40) states that a direct connection to a grounding system is a

necessary element of lightning protection for metal towers and tanks, the category best matching the

storage cask. However, the storage cask at the subject plant does not have direct grounding. Therefore, a

detailed analysis of the potential event scenarios resulting from a lightning strike was conducted to more

fully characterize the potential consequences of a lightning strike on the storage cask in both grounded

and ungrounded configurations.

4-16
For the pilot PRA, the material representing the MPC contents was assumed for calculation purposes to

be homogenized by using average properties for electrical and thermal analyses. A model with uniformly

distributed, homogenized material will have continuous electrical paths throughout the structure. While

the material has average properties to account for the spatial dilution of the material, the possible paths

for currents to flow are continuous in all directions. In contrast, the MPC is made up of discrete

components with voids and discontinuities between them, making their electrical contact imperfect.

Because the homogeneous model allows current to flow uniformly throughout the MPC and contents, the

predicted average power deposition would likely be higher than in the physical situation where the

current is along much more restricted paths. The homogeneous model gives average predictions and

cannot give information about local maxima and minima. A much more detailed and complex model

would be required to obtain accurate local values for heat deposition or temperature. However, this also

means that locally, the power deposition can be much higher than the average values calculated with the

simple models. Analysis of both grounded and ungrounded models of the storage cask predict that power

deposited within the MPC is insufficient to cause temperature increases that could cause internal damage

or rupture of the MPC. The models predict a power deposition to cause only a small average temperature

increase in the MPC of approximately 2 'C (4 'F).

4.2.2.2 Response of the Storage Cask to Lightning Strikes

Power Deposition in Concrete Radiological Shielding. Substantially more power is deposited in the

concrete radiological shielding within the-storage cask when the cask is ungrounded compared to when

grounded. The ratio of power deposition in grounded and ungrounded casks varies during the excitation

pulse from a factor of 2 to a factor of 13 more power deposited in the concrete in an ungrounded cask.

As an overall average, about 3 times as much power is deposited in the concrete shielding in an

ungrounded cask than in a grounded cask. In either case, the peak and average temperature rises in the

concrete were small. In a 2-D calculation, with an initial temperature of 66 'C (151 'F) (Reference 2),

the average temperature rise was found to be about 0.2 'C (0.4 'F). The peak temperature rise was found

to be 5.9 'C (10.6 'F) in the grounded cask and 25.3 'C (45.5 'F) in the ungrounded cask. The 2-D

results for an ungrounded cask indicate the possibility of a local temperature excursion sufficient for

some damage to the concrete (e.g., fracture or cracking), but that is not considered to be sufficient to

cause rupture of the steel liner encasing the concrete.

Power Deposition in the Storage Cask Lid. The largest power deposition density occurs at the lightning

strike location, which was assumed to be on the storage cask lid in these simulations. Because of the

wide variation of spatial resolution between various models, different simulation models predicted

widely varying peak temperatures at the lightning strike location. A failure of the storage cask lid does

not appear to be a credible outcome in any of the models.

Voltage on the Outer Shell of the Storage Cask. A major difference between grounded and ungrounded

storage casks is the voltage profile on the outer shell of the storage cask. With the cask grounded, the

cask voltage at the interface with the pad will be close to zero. With the cask ungrounded, the voltage at

the concrete interface is several megavolts, causing concern for any personnel, if they were near the cask

during a lightning strike. In a real lightning strike, voltages will also be higher in the concrete pad and

soil for both grounded and ungrounded casks because the currents cannot be spread and dissipated as

quickly as in an idealized case. However, these voltages should cause no structural damage to the storage

cask itself (Reference 41).

4-17
4.2.2.3 Conclusions

The lack of a ground connection to the cask allows high electric potentials to develop on the cask near

the concrete pad. These electrical potentials may result in arcing to nearby structures. Without a ground

connection, voltages on the cask near the concrete pad can reach several megavolts and could pose a

significant electrical hazard to workers if they were near the cask. Grounding the cask significantly

modifies the voltage distribution on the outer shell of the cask, reducing it to near zero at the concrete pad.

Although lightning currents are predicted to penetrate both the storage cask and MPC for both the

grounded and ungrounded configurations, calculations indicate that the MPC will remain intact.

Grounding the cask appears to have little effect on power deposition in the interior of the MPC, but

results in a threefold decrease in the average power deposited in the concrete shielding within the cask.

Although the concrete immediately below the simulated lightning strike location on the cask lid is heated

by current penetrating through the steel lid, the deposited energy is insufficient to cause structural

damage. The temperature rise in all regions of the cask, except for the metal at the strike location itself,

is well below the level where damage to the integrity of the storage cask would occur. Localized heating

of the metal lid at the lightning strike location is predicted to be sufficient to melt and vaporize some

material. However, this occurs only in a very localized region at the lightning strike location and a

failure of the storage cask does not appear to be a credible outcome.

4.2.3 Response of the Fuel Rods to Thermal Loads

The thermal models described in Section 4.2.1.1 do not model the spent fuel assemblies in the MPC;

therefore, the cladding temperatures are not explicitly known. However, the temperature of the helium

fill gas within the MPC is modeled for these accident scenarios (blocked vents and 3-hour fire).

One of the key assumptions made in this assessment is as follows:

* The cladding temperature is equal to the temperature of the helium.

For thermal loads, spent fuel rod failure is determined by comparing the cladding temperature to the

temperature limit imposed by Reference 2. The reference lists 394 'C (742 'F) as the Zircaloy

temperature limit for BWR spent fuel assemblies that are 5 years old. This temperature limit is based on

a methodology developed by Pacific Northwest National Laboratory for calculating permissible peak

clad temperatures for generic commercial spent fuel rods. More recent Interim Staff Guidance

Memorandum No. 11, "Cladding Considerations for the Transportation and Storage of Spent Fuel"

(November 2003, ADAMS Accession No. ML022110372) identifies the temperature limit as 400 °C

(752 'F) for normal conditions of storage and short-term loading operations. For off-normal and accident

conditions like fire and blocked vent scenarios, the maximum cladding temperature should not exceed

570°C (1,058 °F). Fuel damage does not occur at this temperature. However, events that result in

temperatures exceeding this limit are, for the purposes of this PRA, conservatively assumed to cause damage

to portions of the fuel.

Analysis performed using the Fluent CFD program (Reference 33) shows that once steady-state is

achieved, the maximum helium temperature in the MPC under normal conditions is 298 'C (568 'F).

If vent blockage occurs, this temperature will increase until a new equilibrium condition is reached.

Appendix C, Figure C. 12 shows that the cladding temperature reaches a steady temperature of 462 'C

4-18
(862 'F) after vent blockage has occurred. The cladding temperature for the blocked vents scenario did

not exceeded the accident temperature limit.

The analysis was performed using the Fluent CFD program (Reference 33). During a 3-hour fire,

the Zircaloy temperature only reaches 298 'C (568 'F), which is less than both the short-term

temperature limit of 570 'C (1,058 'F) and the long-term temperature limit of 400 'C (752 'F).

Even though the long-term temperature limit is approached, the fuel will not fail as a result of exceeding the

long-term temperature limit during this short-term event because the long-term temperature limit

was based on a change in material properties, rather than failure of the cladding. The time needed

for the fuel temperature to equal temperature limits specified in Reference 2 is 3.7 hours.

4.3 MPC Integrity

4.3.1 Integrity Models

The MPC is the confinement boundary C k~urc ri , shcfl i eld -

for the spent fuel. A drawing of the d--- ... -------

welds joining the shell and plate '

components of the MPC is shown in

Figure 16. The MPC is made entirely of P -

austenitic stainless steel. The 1.27-cm MC'd

(0.5-inch) thick cylindrical shell is

constructed with one circumferential M I"",,...

and four axial seam welds. The two ,,:.

axial welds on each side of the MPC are i

slightly offset at the circumferential xcl

weld. The cylindrical shell is then M • hel

welded to a 6.35-cm (2.5-inch) thick MI( A\ial - IL I, 1,,

baseplate. These are full-penetration i i

submerged arc welds which undergo dye d.

penetrant (PT) and radiographic

examinations. They are not post-weld .' H- ,-

heat treated to remove residual stresses. LBasIcpT tIshe mll

After the MPC is loaded with spent fuel

assemblies, the MPC lid is welded to the Figure 16. Welds of the MPG

-shell. This 0.75-inch (1.9-cm) thick

weld undergoes PT examinations after

the root, intermediate, and final weld

passes.

4-19
The vent and drain ports used for hydrostatic testing, drying, and inerting are welded closed with cover

plates. These welds undergo a final penetrant examination only. The MPC lid-to-shell weld and the vent

and drain cover plate welds form the primary seal of the MPC lid. A closure ring is then placed on the

MPC lid and welded to the shell and lid to form a redundant seal. The vent and drain cover plate welds

and the ring welds are also performed using the tungsten-inert gas (TIG) process. This design ensures

that the MPC environment is separated from the outside atmosphere by two independent lid welds. The

MPC lid welds were not considered in the MPC failure assessments for the following reasons:

The TIG process used on the lid produces is expected to produce a much tougher weld than the

submerged arc (SA) weld used in the shell of the MPC.

The redundancy of the lid design requires at least two welds to fail in order to compromise MPC

confinement.

The applied stresses at these locations are less than the stresses in the MPC shell for all events

analyzed.

The following failure mechanisms were postulated for the MPC under mechanical and thermal loads:

* weld fracture

* exceeding the limit load

* creep rupture

All three failure mechanisms were considered for thermal events, while only weld fracture was

postulated for mechanical impacts. For the purposes of this PRA, failure mechanisms are assumed to be

independent. The probability of MPC failure is given as the sum of the failure probabilities from each

failure mechanism.

The failure models used in this PRA do not attempt to predict the hole size in the MPC for a given failure

probability. Any failure is assumed to create a hole in the MPC large enough to allow the particles and

gases within the MPC to be released. This assumption was made because of the difficulty inherent in

trying to predict the equivalent hole size produced from a particular MPC failure mode.

The thermal analyses described in Section 4.2 indicated steady state conditions of the MPC under normal

conditions would be attained prior to the various handling and transfer events used to move fuel out to

the storage pad. These conditions are an internal pressure of 0.56 MPa (81 psi) with an MPC

temperature that varies from 73 'C (163 'F) to 180 'C (356 'F).

Several mechanical impact accident scenarios are considered:

* Drop of the transfer cask onto the refueling floor (Stages 4-8, 11-13, 15-17). The transfer cask

is dropped on a concrete floor from heights less than 0.91 m (3 feet). All such drops are bounded

by the 1.5-meter (5-foot) drop onto the concrete floor.

* Drop of the transfer cask onto the storage cask (Stage 18). This is a handling accident, in which

the transfer cask is being lowered from a height of up to 24.4 m (80 feet) and falls on the storage

cask.

4-20
Drop of the MPC while moving the MPC into the storage cask (Stages 20-21). The MPC falls

from a height of 5.8 m (19 ft) while being lowered from the transfer cask to the storage cask.

Drop of the transfer cask onto the concrete floor (Stage 18). This is a handling accident, in

which the transfer cask is being lowered from a height of up to 30.5 m (100 feet) and, rather than

striking the storage cask, falls onto the concrete floor.

Drop of the storage cask during the transfer phase (Stages 23, 24, and 26-33). The storage cask

is dropped from a height of 0.30 m (1 feet) on concrete, gravel, or asphalt.

It was assumed that the MPC is under normal steady-state conditions during the mechanical load impact

scenarios listed above. In addition, all of the fuel is assumed to be intact during these events. Fuel -

cladding failure would cause the internal pressure of the MPC to increase. If a mechanical impact event

did cause fuel cladding failure, the resulting increase in pressure would not occur during the accident

scenario. The increase in internal pressure due to fuel cladding failure would be a result of the

mechanical loading, but because of the extremely short duration of the impact, it would not affect the

MPC internal pressure until after the mechanical loading is complete.

There are six stages (Stages 11-13, and 15-17) in which the transfer cask could be dropped 0.6 meters

(2 feet) onto the refueling floor. Stress levels at this drop height are not sufficient to cause failure of the

MPC, and are bounded by the analysis of a 1.5-meter (5-foot) drop onto the concrete floor. Therefore, a

drop from this height was not pursued in the mechanical impact scenarios.

4.3.1.1 Weld Fracture Model

The failure analysis model for weld fracture was based on test data of Type 308 weld deposited stainless

steel specimens, and specimens taken from the stainless steel weldments of Process Water Piping of

nuclear production reactors constructed in the 1950s at the Savannah River Site (Reference B. 19). The

mean and standard deviation of the reduction in cross sectional area of these specimens was calculated

from the data and converted to true strain at failure. The true strain at failure data was then adjusted for

the effects of strain rate, temperature and state of stress. These results provided a normal distribution of

failure data, which was used directly to determine the probability of weld fracture given the maximum

strain calculated for each impact event.

4.3.1.2 Limit Load Model

For the case of an MPC with no flaws, a flow stress model was used to predict MPC failure: Failure

occurs if the primary membrane stress equals or exceeds the flow stress of the material. --

The probability of failure was calculated by performing Monte Carlo simulations at various temperatures.

The probability of MPC failure is the number of MPC failures divided by the total number of

simulations. At a given temperature, the applied stresses were calculated. The yield and ultimate tensile

strengths of the material were sampled so the critical flow stress values could be determined. If the

circumferential or axial stresses of the MPC were greater than the membrane flow stress, failure was

predicted. Similarly, if the bending and membrane stress of the MPC shell-to-baseplate weld exceeded

the bending plus membrane flow stress, failure was also predicted. If either of these limits was exceeded,

the MPC was considered failed.

4-21
4.3.1.3 Creep Rupture Model

The methodology used to predict creep rupture of the MPC is based on the model developed by Argonne

National Laboratory for creep rupture of steam generator tubes (Reference 42). In this model, the

structure accumulates a fraction'of creep damage after exposure to any specific time/temperature/stress

condition. If the sum of these creep damage fractions exceeds 1, failure is predicted.

The probability of creep rupture failure as a function of temperature was determined as follows. The

MPC stresses at any time-temperature condition were determined. For this stress, the Larson-Miller

parameter required to cause creep rupture was sampled from a distribution. Given this temperature and

Larson-Miller parameter, the time required to cause creep rupture failure was determined. The

incremental creep damage was calculated by dividing the actual time spent in these conditions by the

time needed to cause creep rupture. This procedure was repeated for the next time-temperature-stress

condition. Once the summation of creep damage fractions exceeds 1, MPC failure would be predicted at

the corresponding temperature.

Over time, the heat load in the storage cask would decrease because of the reduced decay heat. This heat

reduction would cause a corresponding reduction in MPC temperature and internal pressure even if the

cooling vents were blocked. The following conservative assumption is made:

The internal heat load of the MPC does not decrease over time.

Monte Carlo simulations were performed to determine the probability of creep rupture failure. A

simulated MPC consists of both flawed and unflawed regions. The flawed region consists of the four

axial welds, one circumferential weld, and one shell-to-baseplate weld. Since the occurrence of large

crack-like flaws in these welds is small, relatively few failures are predicted. Only welds are analyzed

using this model, since plates are not expected to contain any crack-like flaws. It is assumed that the only

flaws of potential concern are produced during welding. If any of the regions were predicted to fail, the

MPC was considered to fail. The probability of MPC failure is the number of failures divided by the

number of simulations.

Under thermal loads the only significant MPC stresses are attributable to internal pressure. Stresses are

calculated using the classical shell theory equations for pressurized cylinders. The highest stress

associated with internal pressure loads occurs at the shell-to-baseplate connection. Section 3.4.4.3.1.2 of

Reference 2 analyzed the MPC stresses attributable to an internal pressure of 689 kPa (100 psi) at various

locations. The results indicated that the membrane plus bending stress at the MPC shell-to-baseplate

connection was 303.3 MPa (43,986 psi). The axial membrane stress for this situation is 23.6 MPa

(3,419 psi). The difference in these values gives the bending stress acting on the shell-to-baseplate weld

as 280.0 MPa (40,567 psi). The total stress at other pressures was assumed to be directly proportional to

those at 0.69 MPa (100 psi). Therefore, at a pressure of 0.565 MPa (82 psi), the total stress at this

location is only 248.7 MPa (36,068 psi). Subtracting the axial membrane stress of 19.05 MPa (2762 psi)

gives a bending stress of 229.6 MPa (33,306 psi).

4-22
Appendix B (Figure B. 14) shows the temperature variation as a function of axial height for the

steady-state conditions of the MPC under normal operating conditions with internal pressure of 0.56 MPa

(81 psi). If all four cooling vents of the cask become blocked, the MPC temperature and internal pressure

* would increase until a new steady-state condition was reached. In this case, the potential of the MPC

failing due to creep rupture is introduced. See Appendix B, Sections B.1.7 and B.1.8 for a more detailed

discussion of MPC creep rupture response to blocked vents. In summary, none of the blocked vent or

other thermal scenarios resulted in conditions that could fail the MPC.

4.3.2 Probability of MPC Failure

Mechanical Loads

Appendix A evaluated the structural response of the MPC for various drop scenarios. The objective was

to determine the maximum effective plastic strain on the MPC confinement boundary for each of these

scenarios. The analyses showed that the most highly stressed regions of the MPC are near the base of the

cylindrical shell and in the weld joining the shell to the 6.4-cm (2.5-inch) thick baseplate.

To determine if the integrity of the confinement boundary is compromised, the maximum effective

plastic strain in the MPC must be compared to an appropriate strain failure criterion. For a valid

comparison, the conditions under which the maximum effective plastic strain is calculated, and the

conditions under which the failure strain limit is measured, should be consistent. The comparison,

therefore, must account for how the strain is measured, as well as the effects of strain rate, temperature

and state of stress. Appendix B, Section B. 1.4 establishes a valid basis for this comparison, and

estimates the probability of exceeding the failure strain limit for a given drop event.

Table 12 provides a summary of MPC failure probabilities for various mechanical impact loads. Those

stages where the MPC is sealed are briefly discussed below.

If the transfer cask were dropped during Stages 11-17, the probability of MPC failure is <1x10-6. During

these stages the transfer cask is sealed and carried 0.3-0.6 meter (1-2 feet) above the refueling floor. In

Stage 18, the transfer cask is lowered 24.4 meters (80 feet) to the storage cask. If the transfer cask is

dropped from this height onto the storage overpack, the probability of MPC failure is very small

(0.0002). On-the-other-hand, should the transfer cask drop 30.5 meters (100 feet) and impact the

concrete floor, the probability of failure is approximately 0.02.

During Stages 20 and 21, the MPC can potentially be dropped 5.8 m (19 ft) into the storage overpack. If

this were to occur, the probability of MPC failure is 0.28.

During Stages 29-33 if the storage cask were to be dropped from a height of 0.3 m (1 ft) onto concrete,

asphalt, or gravel, the failure probability of the MPC is <1 xl 06. For both the end drop and rotation drop,

the cask transporter can cause the storage cask to slide, but not tip-over. The stresses generated during

sliding events are much lower that for the drop events or tip-over and would not cause the MPC to fail.

4-23
The probability of MPC failure in the event of tip-over during storage is calculated to be <xlx 06. The

dry storage cask system SAR also analyzed a tip-over and concluded that it would not lead to MPC

failure. The probability of cask tip-over attributable to various potential events, such as seismic events,

high winds, collision by vehicles, tornado-driven missile impacts, and shock waves from tanker or

pipeline explosions was evaluated, and for all such events it was concluded that there would be no tipover and, therefore, zero probability of damage to the MPC.

Table 12. Summary of MPC Failure Probabilities for Various Mechanical Impact Loads,

Impact Surface Drop Height MPC Failure Event Scenario (Target) (feet) Probability(')

5 < 0.000001

40 0.000360

Concrete Floor

70 0.002600

Transfer Cask Vertical Drop 100 0.019600

5 0.000002(2)

Storage Cask 40 0.000014

80 0.000203

Drop of MPC from Transfer Cask into Storage Overpack

Storage Overpack during Transfer Pedestal19 0.282000

Operations

Concrete 1 < 0.000001

Storage Cask Vertical Drops on Asphalt 1 <0.000001

Various Surfaces Asphalt_1_<_0.000001

Gravel 1 < 0.000001

Storage Cask Tip-over Concrete Pad NA < 0.0000001

(1) See Note 1 of Table B.3 and Section B.1.4 of Appendix B.

(2) This result is not consistent with other results presented in the table (i.e., it should be <0.000001). This is

because the response results for this case were based on the original (unmodified) LS-DYNA analysis, which, due to

a number of factors, produced very conservative results at low drop heights.

4-24
Thermal Loads

Evaluations of the probability of MPC failure attributable to fracture, load limit, and creep rupture from

thermal loads are detailed in Appendix B, Sections B. 1.3 and B.1.5-B. 1.7. Failure models were created

in Microsoft Excel with an add-on module called @RISK (wwwjpalisade.com), which allows MonteCarlo simulations to be run in Excel by allowing values in the spreadsheet cells to be sampled from

statistical distributions. The spreadsheet cell specified as output identifies each iteration as either an

MPC success or failure. This process is repeated many times in a Monte Carlo simulation. The

probability of failure was calculated by dividing the total number of MPC failures by the total number of

iterations performed. If normal steady-state conditions are maintained for the full 20-year license period

of the cask, no MPC failures are expected. If the storage cask is exposed to an external fire for 3 hours or

all four cooling vents become blocked, the MPC temperature and pressure will increase, but no MPC

failures are expected.

Even with the yield and tensile strengths below the minimums, the limit load model was never the

limiting failure mechanism in any of the heat-up scenarios. None of the blocked vent or other thermal

scenarios resulted in conditions that could fail the MPC.

4.4 Fuel Failure

4.4.1 Fuel Integrity Model

The peak g loads imposed on a fuel rod from the vertical drop scenarios considered in the PRA far

exceed the elastic buckling load of the fuel rods in the assemblies. Therefore, to determine the strain

demand on the fuel cladding, the inelastic buckling capacity of the fuel rods must be considered. The

approach taken in Appendix C was to analyze the fuel rod as an elastic-plastic beam-column with initial

curvature under dynamic impact. Failure of the fuel rod is determined by comparing the maximum strain

in the cladding to a strain limit based on experimental data.

Since the impact event entails the interaction of potentially hundreds of rods, their spacer grids, tie plates,

fuel spacers, basket, cask, and target under dynamic loading, computational efficiency was achieved by

modeling a single fuel rod. Thus, an explicit finite element analysis model of a single fuel rod with

lateral displacement constraints was used to study the inelastic behavior of a fuel rod under dynamic

impact loads. (See Appendix C for details.)

4.4.2 Fuel Rod Cladding Failure

For each drop height, Table 13 lists the maximum principal strain in the cladding, the failure strain limit

and whether or not fuel cladding failure occurs based on the limit strain. For the drop of the transfer cask

(TC) onto the concrete floor, additional drops of 6.1 and 16.8 meters (20 and 55 feet) were added to the

standard drop heights of 0.3, 1.5, 12.2, 21.3, and 30.5 meters (1, 5, 40, 70, and 100 feet) in order to

highlight transitions in buckling behavior or differences in response attributable to target stiffness.

4-25
Table 13. Probability of Fuel Cladding Failure for Various Drop Scenarios

Maximum Strain Limit Selected for Probability of

Event Scenario Impact Surface Drop Height Principal Strain Highcue u Fue Cladi ng

(Target) (feet) in Cladding High-Burnup Fuel Cladding

(ni)Fuel Failure

(in/in) (in/in)

1 0.0043 0.010 0.0

5 0.0062 0.010 0.0

20 0.0072 0.010 0.0

Concrete Floor 40 0.011. 0.010 1.0

Transfer Cask 55 0.025 0.010 1.0

Vertical Drop 70 0.037 0.010 1.0

100 0.052 0.010 1.0

5 (1) na 0.0

Storage Overpack 40 (1) na 1.0

80 (1) na 1.0

MPC Drop into Storage Overpack.

Storage Pedestal19 0.090 0.010 1.0

Overpack

Storage Cask Concrete Pad na 0.0055 0.010 0.0

Tip-Over

(1) This drop was not evaluated for cladding response. Results for the transfer cask drop onto the concrete

floor from the same height were used to determine cladding failure.

The results show that the cladding yield strain of 0.0084 in/in is not exceeded until drops of the transfer

cask are in the range of 6.1-12.2 meters (20-40 feet). The results also show that target- stiffness has a

significant influence on cladding strain. For example, the 6.1-meter (20-foot) drop of the transfer cask on

to the concrete floor results in a strain less than yield, whereas, the 5.8-meter (19-foot) drop of the MPC

into the storage overpack, which is a much harder impact, produces cladding strains that are 10 times

greater than yield. This is because the buckling modes for these two impacts are completely different, as

discussed in Appendix C.

4-26
The cladding strain at failure (rupture or breach) of high-burnup fuel, with circumferential hydrides,

which is fuel with a burnup greater than 45 GWd/MTU, is expected to range from 1.0% to 3.0%.

For the purpose of the PRA, the lower value of 1.0% strain is selected as the strain limit at failure.

Appendix D of the PRA shows that the particulate release fraction from a spent fuel rod is a function of

three variables: (1) the fraction of the fuel rim layer that fractures, (2) the percentage of particulate

entrained by gasses during release, and (3) the number of tears (breaches) per rod. In addition, the

fraction of the rim layer fractured and the number of breaches per rod are in turn functions of the g loads

experienced by the fuel and the severity of the buckling that takes place during a given drop.

In a detailed PRA each of these variables would be associated with a probability distribution and'sampled

as part of an overall simulation. However, because of the level of effort inherent in such an approach,

Appendix D simplifies the effort by recommending an upper bound value for release fractions to be used

in the consequence analysis. This in turn simplified the effort in Appendix C by not requiring an

estimate of the number of breaches per rod, but only whether or not the rod breaches at all. Also, since

the analysis of fuel rod failure is based on a single pin model, it was assumed that when one rod fails, all

rods fail, although this is very unlikely.

A tip-over of the storage cask will cause a peak lateral load of 45g on the fuel assemblies. Based on an

evaluation of eight different laterally loaded BWIR and PWR fuel assemblies ranging from 7x7 to 17x1 7

arrays, Reference 43 concludes that, for the particular cladding properties used, the cladding yield strain

of 0.0077 in/in is reached for a lateral loading of 63g. Since this is a linear elastic analysis, the strain at

45g would be 0.0055 in/in, which is much less than the cladding failure strain limit of 0.01 in/in.

The failure probability of fuel rods if a single assembly were dropped, as in Stage 1, is estimated by using

data on individual assemblies. To estimate a failure probability when no failures have been observed, the

50% confidence limit can be used. For a binomial distribution, the 50% confidence limit when no

failures have been observed is calculated by dividing 0.7 by the number of cases. Thus, the probability

of fuel assembly failure is estimated as 6.4x 10.2 (0.7/11 = 0.064). (See Section 3.3.1.)

4-27

5. SECONDARY CONTAINMENT ISOLATION

5.1 Containment Isolation Reliability Model

Given a release of radioactive material inside the secondary containment following a drop of the transfer

cask, the three distinct functions that occur are to detect radioactive material, isolate the containment, and

operate the SGTS. Each of these functions is accomplished by redundant and independent trains of

systems. The release of radioactive material will be detected by one or more of six detector systems that

signal the secondary containment to isolate. Each of these detector systems consists of two trains, each

train having two detectors. Both detectors in a train must detect radiation to isolate the secondary

containment. Only 1 of the 12 trains is needed to isolate the secondary containment.

The secondary containment is

modeled as one large volume as

shown in Figure 17. The letters ©(

in circles next to the names of the , , RCCDI

HVAC systems relate these R ea kch,, ,C

systems to the event tree that is ,l\W

illustrated in Figure 18. The

success criteria for the secondary 0 R,,•,o,

containment model are as Ic .. i

f o l l o w s : o p e r t i o ,.. ...... I '- ... ... . ,-B ik , .,

During normal operation,

all direct paths to the I I V,

external environment are

closed.

SGTIT I SiT

A negative differential ( - A,.1 ,

pressure is maintained II.( ©

between the inside of the @ @

secondary containment Figure 17. Secondary Containment HVAC Model

and the environment.

When an isolation signal

occurs, the SGTSs of

Unit 1 and Unit 2 take suction from the inside of the secondary containment, filtering the intake

and directing the exhaust to the plant stack (elevated release).

The event tree used to determine the probability of failing to isolate the secondary containment is shown

in Figure 18. The top events of the event tree are the failure of valves along the ventilation paths to close

and the failure of fans to stop moving air through the secondary containment. Each top event is a valve

or fan of the ventilation system, except for the last two top events, which are the SGTSs. The valves and

fans are paired in their respective system for brevity and clarity. The systems have letters (A) - (K)

corresponding to the same systems illustrated in Figure 17. The top events are quantified with fault trees

that model the fans, valves, and radiation sensors to initiate the isolation of the secondary containment if

radioactive material were released from a transfer cask. Inputs to the fault trees are determined from

published sources of reliability data, such as References 44 and 45.

5-1
l'op ENs ws

olation \ilk Cs close t ( Retfuelingt area unit I I\IVAC supply

F ms isolate

I [intul us N . ik i s Close IrV Fits isolate .jG Refueline aen unit eshaUst -HVAC 

Fis isolate © ( Refluelitta areaC unit 2 I VAt' supply

l OI5la®io eR line area unit 2 exhaust

Fi res isolate u

oI 'lilion ,iahcs clo se t ® Fanes isolate ItQ RCeCtolr hutildfing unit I accessible area I \iAC suppi y

Falls is o lat Reactor building unit I accessible are'a HVAC Cxhau1tlSt

I i\,l' C Closes (D ReCtCttor iLlilding unit I tilrcra I I\A ' Cxltust

Fl ns isolata

I \'ttl \.'ahe s s oeiise clse J "

F isoF tc j (D React '" b utdiiitnt uI it 2I IV\\ C Cs i up s t

I)peatcs ý (D (il's ,,,it 1

Qlperatcs Q@ Suit2

Fi d stlit. ('ha'acelristis

Lh n~CS~itrssuri'.d I C~ll,

I InC'ssttrtzcCd leak JItLIClIC'Cnlrc s l' cdi clc

NunlhlCr l

I (eqeti(el 'm irelease

Isohlaihon 0.I00t cf11 relelase

I lsilm- RttiiinI I UII). IIItlrTce I-' i Rl ls h tll-tte i l -. ,

KType I i Filtered

- - IJ

I t!

3 -~ It---I [ 0- ,ttX "-II ;- 

N- " !

_~ {}[ l ] EndvI- .)j i- -• •• - --{3c]-i ! St.-te

Figure 18. Secondary Containment Isolation Event Tree

5-2
The event tree delineates the ways in which the secondary containment isolation can fail by modeling the

failure of valves to close, fans to isolate, and the SGTSs to operate. It does not model the progression of

events that would occur if the secondary containment receives an isolation signal. If the secondary

containment fails to isolate, thereby opening a release path for radioactive material, other release paths

are not modeled. The analysis conservatively treats all the material that is released from the transfer cask

as being released to the environment.

The 20 top events delineate 43 sequences in the event tree analysis. Of these 43 sequences, 22 result in

successful isolation of the secondary containment. The secondary containment successfully isolates

when one or both SGTS(s) functions, even if there is an open path between the secondary containment

and the outside atmosphere. SGTSs filter the release through an elevated plant stack. The other 21

sequences result in a failure to isolate the secondary containment. Eight types of secondary containment

isolation failures are delineated in Figure 18. The isolation failure types (IFTs) are defined in Table 14.

The following assumptions were made in developing and evaluating the event tree model:

Doors, hatches, and personnel locks are secured according to procedures before the cask

operation begins.

The possibility of an open door or other access opening during the cask operation is negligible.

The justifications for this assumption are as follows. First, the HVAC system automatically

adjusts to provide a negative differential pressure across the secondary containment boundary.

Second, if a door or other access opening were inadvertently left open, the HVAC would not be

able to draw a vacuum. An open door in the secondary containment would alarm in the control

room. Workers would be directed to close the open door. The handling phase of the operation

would cease until the negative pressure in the secondary containment could be reestablished.

Because the SGTS filters are passive components, they have a negligible probability of failing to

reduce the radioactive material content of the gas passing through them.

If the HVAC systems isolate and both SGTSs fail to operate, the elevated temperature inside the

secondary containment, with respect to the outside air, provides a driving pressure for a release

of radioactive material from the failed MPC. However, as compared to other releases from

reactor accidents, the flow rate is small.

5.2 Probability of Secondary Containment Isolation Failure

Table 14 shows the results of probability calculations for sequences in Figure 18. To calculate the

probability of the secondary containment failing to isolate, only the sequence probabilities are summed as

shown in Table 15. The small total probability of 1.5 x 10' is attributable to the large amount of

redundancy in the secondary containment isolation systems.

5-3
Table 14. Secondary Containment Isolation Failure Types

Isolation

Failure

Type (IFT) Probability, Sequences Description

I 1.4x 10-4 3 Unpressurized leak. All systems have isolated (automated design response of the isolation

system or over-current trip of HVAC fan due to closure of the isolation valves) and both

SGTS fail. Leak occurs through cracks, around doors, etc.

II 3.1 x 10-' 6, 18, 30, 38 Unpressurized filtered release through exhaust, Release through HVAC exhaust system due

to exhaust isolation valves failing to close. Supply and exhaust fans isolate. Exhaust filtered

release through reactor building vent plenum.

III 2.7 x 10-' 12, 26, 34, Unpressurized unfiltered release. Release through HVAC supply system due to supply

isolation valves failing to close. Supply and exhaust fans isolate. Unfiltered release through 42 supply louvers.

IV 8.0x 10. 14, 15, 27, Pressurized-leak. All HVAC exhaust systems isolate but the secondary containment is

pressurized by failure of HVAC supply system to isolate with fan continuing to run. 35, 43 Leak occurs through cracks, and around doors.

V 1.3 x l0. 8, 9 Filtered release at 6,000 cfm. Reactor building Unit 2 exhaust system fails to isolate with fan

running and either SGTS fails. Exhaust filtered release through reactor building vent plenum.

VI 6.9 x 10.6 19, 31, 39 Filtered release at 30.000 efm. Refueling area Unit 1, refueling area Unit 2 or reactor

building Unit I inaccessible area HVAC exhausts fail to isolate, associated fan continues to

run at 30,000 cfmr. Exhaust filtered release through reactor building vent plenum.

VII 2.3 x 10-6 23 Unfiltered release at 30,000 cfm. Reactor building accessible area HVAC exhaust fails to

isolate. Fan continues to run at 30,000 cfmn. Unfiltered release occurs through reactor

building vent plenum.

VIII 7.6 x 10.9 22 Unpressurized unfiltered release. Reactor building accessible area 1HVAC exhaust fails to

isolate. Fan isolates. Unfiltered release through reactor building vent plenum.

I Values in this table are shown as they are computed in the SAPHIRE sottware.

5-4
Table 15. Sequences Where the Secondary Containment Fails to Isolate

Sequence Sequence

Number Probability Number Probability

3 1.4xl04 23 2.3x106

6 7.6x109 26 2.7x10 8

8 2.8x10 9 27 8.0x 10.6

9 1.0xl08 30 7.6x10 9

12 4.9x10-12 31 2.3x106

14 1.5x101 2 34 4.9x10 1 2

15 6.2x101 2 35 1.6x10 9

18 7.6x10 9 38 7.6x10 9

19 2.3x 10-6 39 2.3x 10.6

22 7.6x10 9 42 4.9x10 1 2

43 1.6x10 9

Total 1.5x10-4

5-5

6. CONSEQUENCES

6.1 Source Term

This consequence assessment addresses the severe accident scenario related to the inadvertent drop

[30.5-meter (100-foot) drop] of the HI-STORM 100 cask system while in the Stage 18 of the transfer

process. This consequence assessment was performed to support development of a PRA for dry cask

storage for the specific BWR reactor facility analyzed in the report. In order to examine radiological

consequences, release fractions for this accident were developed. Release fractions were estimated as the

product of a rod-to-canister release fraction and a canister-to-environment release fraction. Input was

obtained from Sandia National Laboratories (SNL) for the development of source term. Detailed reanalysis on source term analysis is documented in Appendix D.

This assessment was performed using the MELCOR Accident Consequence Code System (MACCS2)

(Reference 8 and Appendix E, Reference E.2) for a BWR facility. Specific data needed to model a

30.5-meter (100-foot) HI-STORM 100 drop accident scenario for the MACCS2 consequence calculation

were collected and used. The important parameters/variables required to model the BWR site are the

radionuclide inventory, the source term (i.e., release fraction, release start time, and release duration), the

initial plume dimensions (related to the system geometry), the plume heat content, the population

density/distribution, and the subject site weather. Other input parameters and models necessary for a

MACCS2 calculation were taken from the NUREG-1 150 study MACCS2 input file prepared for the

Surry nuclear power plant. The input file is documented in Appendix C of the MACCS2 code manual

and is referred there as Sample Problem-A (SP-A)'. This SP-A MACCS2 input file was used because of

the availability and pedigree of this input file. Additionally, the SP-A input parameters are suitable to

model the subject site's dry cask storage consequence calculation. A summary of the results obtained by

MACCS2 is presented in Table 18. A discussion of the parameters used for the subject site's variables,

shown in Table 18, can be found in Appendix E, Sections E.3 - E. 11.

The consequence measures chosen for this study were the mean values given by MACCS2 for individual

risk of early fatality within 1.6 km (1 mi) and the individual risk of cancer fatality within 16 km (10 mi)

of the site boundary.2 The 1.6 km (1 mi) value is the distance from the exclusion area boundary (EAB),

assumed to be 0.5 km (0.3 mi), out to a distance of 2.1 km (1.3 mi). These distances and consequence

measures were chosen to allow comparison to the NRC's quantitative health objectives for reactor

accident risk. Since cask accident consequences in terms of individual risk of early fatality and cancer

fatality are small, the mean consequence values calculated for individual lifetime dose commitment (the

mean values of the variable known in MACCS2 as "peak dose found on spatial grid") may also be of

interest. Therefore, this section includes the estimated mean individual lifetime dose between 1.2 and 1.6

km (0.75 and 1.0 mi) as measured from the release point. Also, close-in consequences are generally

more representative of the source term, while consequences further away tend to be controlled by

assumptions on long-term relocation.

I Evacuation was not considered in the MACCS2 consequences calculations for this study.

2 All consequence measures given in Appendix E are mean values as calculated by the MACCS2 code unless stated

otherwise. The mean values in MACCS2 are defined as the average (expected) consequences over all weather

conditions (Reference E.2).

6-1
An impact on the cask will fail a fraction of the fuel rods (Frods) and make their contents available for

release. A portion of this available material will be released from the rod to the cask environment (FRC).

Part of the inventory that has been released to the cask will be transported through the breach in the cask

to the environment (FCE). The fractional release from the cask (F,,,) can be approximated by the equation:

FreI = Frods XFRC xFCE

The fuel rods can have radionuclides on the cladding surface in the form of solid Chalk River

unidentified deposits (CRUD), inside the fuel rod as gases and volatile fission products, and within U0 2

fuel grains as both solids and gases. Due to the different physical forms of the radionuclides, the above

equation is analyzed separately for each group of similar radionuclides.

This release fraction model (Appendix D) accounts for the different properties of the rim region formed

in high-bumup fuel and the reduction in ductility of the cladding due to higher bumup and hydrogen

absorption. The methodology, as presented, is only applicable to impact type events and not those that

result in an elevated temperature attributable to a long-duration fire.

The fraction of the rods fractured (Frods), calculated in Appendix C, is dependent on the energy imparted

to the rods, which in most cases is a function of drop height. The release from the rods to the cask makes

use of the properties of high-bumup fuel and analyzes the particulate release from the rim and body

regions of the fuel separately. Fracture of the fuel into fines is based on a modification of the equations

in the DOE Handbook (see Appendix D), which relate the fraction of fuel that fragments into a respirable

size to the impact energy that the fuel experiences. FRC depends on the number of fracture sites in a rod,

the entrainment of the fines in the gas stream during the depressurization of the rod, and the extent to

which the rim region actually fractures. Uncertainty in these parameters is analyzed. The release of

radionuclides from the cask to the environment is based on the fraction of gas that is ejected when the

cask depressurizes to atmospheric pressure. An ideal gas law relationship is assumed. An analysis of the

propensity for settling and plating out of particulate, by a number of mechanisms, so they are not in the

gas stream during depressurization is also conducted. Details can be found in Appendix D.

6.1.1 Source Term Estimates

In addressing source term estimates, the discussion necessarily includes consideration of release fraction,

release start time and release duration.

6.1.1.1 Release Fraction

The release fractions calculated in Appendix D are based on the number of fuel fractures calculated in

Appendix C.

6.1.1.2 Release Start Time

In the MACCS2 calculation for the cask drop event, it was conservatively assumed that as soon as the

accident occurs, the release to the environment occurs immediately.

6-2
6.1.1.3 Release Duration

The probability of failure of the MPC, in the event of a drop of 30.5 meters (100 feet), is discussed

elsewhere in this report. While the likelihood of failure of the MPC in this type of event is in reality

quite low, in order to examine radiological consequences, release fractions for this accident were

developed. Release fractions were estimated as the product of a rod-to-canister release fraction and a

canister-to-environment release fraction. The release fractions presented in Table 16 for a transfer cask

dropped 30.5 meters (100 feet) are the fraction of inventory that can be released by depressurization from

the cask at the time of the accident. Details are provided in Appendix D.

Table 16. Release Fraction Values for a Transfer Cask Drop of 30.5 Meters (100 Feet)

Radionuclide Release Fraction Release Fraction with Filtration

Nobel Gases 0.12 0.12

Particles 7x10 . 6 to 1.2x 10 . 7x×10 7 to 1.2x 10-4

CRUD 1.5x10 "3 1.5x10-4

6.2 Consequences

6.2.1 Consequence Model

Table 17 shows the major inputs for the calculation, including a brief description. Details are given in

Appendix E.

Offsite consequences from radioactive material released from a transfer cask or storage cask during a

severe accident are modeled with the MACCS2 code (Reference 46). A plume of radioactive gases and

aerosols are released from a point and dispersed in the atmosphere, exposing the surrounding population.

The characteristics of the release, such as the amount of material, the type of material, and the heat

content, are determined from other studies (see Appendix E)._ The consequence measure that is used in

risk estimates is an expected value of the consequences from the dose calculations given a release of

radioactive material (i.e., source term). The calculated consequence measures are early fatalities and

latent cancer fatalities as seen in Table 18. Sheltering and evacuation were not considered in the

MACCS2 consequence analysis.

6-3
Table 17. Major Inputs for Consequence Calculations

Input Description

Radionuclides Inventory:

Source Term:

Meteorological Data:

Population Data:

Emergency Response:

Specifies the radionuclides that are to be modeled.

Specifies the characteristics of the radioactive plume such as the number of

releases (i.e., plume segments), the heat content, the time of the release, and the

duration of the release.

Hourly information of 1 year of meteorological data at the subject site. Each

hourly data set includes the stability of the atmosphere, the wind speed, the wind

direction, and the precipitation.

Population distribution surrounding the release point developed from census

data for the subject site location.

Emergency response actions modeled in the region surrounding the release

point. Emergency response actions modeled was relocation. Relocation defines

the actions taken at the appropriate time to move individuals residing in zones

where dose criteria are expected to exceed critical values, and the duration they

will be kept in a relocated status.

Table 18. MACCS2 Consequence Calculation Results

Release Fraction Individual Risk of Individual Risk of Individual

Prompt Fatality Cancer Fatality Peak Dose

Noble Particles CRUD Release Height within 16 km within 16 km (1.2-1.6 km

Gases meters (feet) (10 miles) (10 miles) Sv (rem)

.12 1.2xl03 1.5xl03 50(164) 0 3.6xl04 1.85(185)

.12 1.2x103 1.5xl0- 120 (393) 0 2.1x10-4 0.14 (14)

.12 1.2x10-4 1.5x104 501(164) 0 5.2x10' 0.22(22)

.12 7x10-6 1.5x10-3 50(164) 0 4.3x106 0.026(2.6)

.12 7x10-6 1.5x10-3 120 (393) 0 2.6x10"6 0.0036 (0.36)

.12 7x10-7 1.5x10"4 501 (164) 0 4.3x 10" 0.0028 (0.28)

Results corresponding to a release fraction with filters operable.

6-4
6.2.2 Consequence Measures

The dose conversion factors (DCFs) used in this study are the same as those used in the NUREG-1 150

study (DOSDATA.INP), as provided by K. Eckerman from Oak Ridge National Laboratory (ORNL)

(Reference E.4). These DCFs for 60 radionuclides important in reactor accidents include definitions for

19 organs. The DOSDATA.INP DCF file is appropriate to generate long-term health effects calculations,

but is not suitable to generate acute health effect calculations (e.g., early fatalities) for the dry storage

cask PRA study because the MACCS2 calculation with the DOSDATA.INP DCF generate overestimated

values for acute health effects. Even with this limitation for the DCF file with acute health effects, the

MACCS2 calculations predicted zero early fatalities.

It should be noted the estimated consequences (individual probability of a latent cancer fatality) from a

release of the inventory of the MPC, for example, for high fuel burnup and a release height of 50 meters

(164 feet), include the essentially negligible contribution from noble gases (-10.10), and the contribution

from radionuclides other than noble gases is 3.6x 10'. These estimated values are based on the output of

the MACCS2 code calculations.

6-5

7. RISK ASSESSMENT

7.1 Risk Models

Reference 1 defines risk for NRC use as the risk triplet, in which the elements of risk are the scenarios,

the frequencies of the scenarios, and the consequences of the scenarios. Reference 1 also states that a

measure of risk is the consequences multiplied by the frequency of the consequences. The measure of

risk is given by Equation 6 for one initiating event:

m

R= fZ PnKn (6)

n=1

where: R aggregate risk

f frequency of the initiating event (1/time)

Pn = conditional probability of the nth accident scenario given the initiating event

K, = consequence of the nth accident scenario

The aggregate risk is the risk measure used in this study.

To determine the risk of all plausible accident scenarios in the stages of the dry cask storage operation,

Equation 6 can be rewritten as Equation 7:

R= Z1 fj,mZ Pjm,nKj,m,n (7)

j m n

where: R = aggregate risk

fj,r, = frequency of the mth initiating event while in the j

t1

h stage

Pj,m,n = conditional probability of the nth accident scenario given the mtih initiating event

while in the j stage

Kjmn = consequence of the nth accident scenario from the mth initiating event while in the

jth stage

The aggregate risk of storing spent fuel in a dry cask is the risk incurred during the service life of a cask.

As shown in Table 1, the service life is divided into three phases:

(1) handling within the refueling building

(2) transfer to the storage pad

(3) storage on the storage pad

An important consideration is that the risk of handling and transfer is demand-related while the risk of

storage is time-related. Thus, the risks of the three phases must be expressed in a manner that permits

them to be added to obtain the total risk. For each phase, the units of risk are consequences per year,

expressed as either prompt or individual probability of a latent cancer fatality.

7-1
Consider a cask in the first year of its service life at the subject plant. During this year, the cask passes

through the handling and transfer phases and begins the storage phase. In this study, the handling phase

is treated as a demand-related activity because of its short duration. Because the handling phase lasts no

longer than a few days, time-related external initiating events that could disturb the lifts of the transfer

cask or otherwise challenge the integrity of the transfer cask are unlikely to occur. For example, an

earthquake is unlikely to occur during the short period of lifting.. The transfer cask is not subject to other

external events, such as high winds or floods, because all the handling is done inside the containment.

For these reasons, the time-related risk during the handling phase is negligible relative to the

demand-related risk. In the handling phase, the only demand-related initiating event is a drop while the

cask is lifted, moved, or lowered by the crane.

The only initiating event that can occur in the handling phase is a drop of the transfer cask or a drop of

the MPC (Stages 20 and 21). The risk associated with drops in the handling phase is given by Equation

7. Therefore, there is no need to sum over the initiating events. Equation 7 is rewritten as Equation 8,

where the first summation is over all stages during the handling phase:

R H f.HZP.HK I H(8 i~n(8) J in 

J n

where: RH = risk of the handling phase

= frequency of dropping the transfer cask or MPC during the jth stage in the

handling phase

PHPj, probability of a release of radioactive material to the environment during the nth

sequence given that the cask is dropped in the jth stage of the handling phase Hh K,, = consequence of a release of radioactive material during the nth sequence the transfer cask is dropped during the jth stage of the handling phase

The frequency of each lift is expressed as one per year during the first year of service life. The handling

risk isexpressed as the individual probability of a prompt or latent cancer fatality per year, but it applies

only to the first year of service life.

The risk of the transfer phase is determined analogously to the risk of the handling phase. Although the

transfer phase occurs outside of the secondary containment where time-related events such as high winds

and floods can occur, the transfer phase activities are performed only under favorable weather conditions.

An earthquake is very unlikely to occur during the brief transfer period. A fire from the diesel fuel in the

overpack transporter will not lead to a radionuclide release. For these reasons, the time-related risk

during transfer is negligible relative to the demand-related risk. In the transfer phase, the only

demand-related initiating events are drops.

7-2
The risk in the transfer phase is given by Equation 9. The first summation is over the stages during the

transfer phase. The second summation is over all of the demand-related initiating events that can occur

during the transfer phase:

R T = = 11 f fTm mL JPT Pj,min KT j,m,n

j m n

(9)

where: RT

P j, m, n

T

Kj, m,n

= risk of the transfer phase

= frequency of the mth initiating event during the j' stage of the transfer phase

= probability of a release of radioactive material during the nt" sequence given the

m" initiating event in the jth stage of the transfer phase.

consequence given a release of radioactive material during the nth sequence

given the mti initiating event in the jth stage of the transfer phase

During the storage phase, the overpack cask is subject to time-related initiating events (e.g., earthquakes,

high winds) that can challenge the integrity of the overpack and the MPC. Because the storage phase has

only one stage, the risk in the storage phase is given by Equation 9 without the first summation. This

equation is rewritten as Equation 10, where the first summation is over all initiating events that can occur

in the storage phase:

Rs = I sZPs Ks

m n

(10)

where: Rs

.,n

H

KM. n

risk of the storage phase

= frequency of the m' initiating event while in the storage phase

= probability of a release of radioactive material during the nth sequence given the

m!' initiating event while in the storage phase

consequence of a release of radioactive material during the nth sequence given

the mth initiating event while in the storage phase

The storage risk is expressed as the individual probability to the public of a prompt or latent cancer

fatality per year.

The aggregate risk during the first year of operation is the sum of the risks of the three operational

phases. During the first year of operation, the MPC passes through the handling and transfer phases and

moves into the storage phase.. The handling and transfer phases last a small fraction of a year, so that the

MPC is in storage for almost a full year. Accordingly, the aggregate risk of the transfer cask and

overpack cask for the first year of its service life is given by Equation 11:

R(1) = RH+ R T + Rs (11)

7-3
After the first year of operation, the overpack cask is no longer handled or transferred. Only the storage

risk is incurred for each of the remaining 19 years of the MPC's service life.

The total risk of the cask during 20 years of its service life is equal to the first year risk plus 19 times the

storage risk. The contribution of handling and transfer to risk is much larger than the risk incurred during

storage.

7.2 Risk Results

For this pilot PRA, the risk calculations are summarized in Table 19. Each row of the table represents an

initiating event in a stage. The initiating events in Table 19 are those listed in Tables 4, 5, and 6 that

have the potential to affect the subject plant or breach the MPC. Each initiating event has only one

accident scenario - the failure of the MPC and fuel rods. Aggregate risk is the sum of the risks of the

accident scenarios. Table 19 shows the risk for the first year of service and the annual risk for

subsequent years. The risk during the first year consists of the risk from the handling and transfer phases

and part of the storage phase. The risk of subsequent years consists only of the annual risk of storage on

the ISFSIL

The risk measure in Table 19 is the annual individual risk of a latent cancer fatality. Because all offsite

releases result in doses to individuals that are below the threshold for prompt fatalities, no prompt

fatalities are expected.

The components of risk (frequencies of initiating events, probabilities of subsequent events following the

initiating events, and the consequences of the scenarios) are shown in the columns of Table 19. These

components are consistent with the definition of risk in Reference 1 where the frequency of the scenario

is the product of the initiating event frequency and the event probabilities. The accident scenarios consist

of MPC failures and, in the handling phase, releases from the secondary containment. The numbers in

the last column of Table 19 are the products of the entries in the previous columns and are the risks of the

scenarios.

The notes for each scenario in Table 19 state the sections in this study that discuss the corresponding

initiating event frequency, the probability of MPC failure, the probability of a radioactive release from

the containment (where applicable), and the consequences as measured by the individual probability of a

latent cancer fatality given the occurrence of the scenario. The notes also identify which radionuclides

are released from the containment noble gases only (NG), radionuclide inventory other than noble gases

(OT), or all radionuclides in the inventory of the MPC (All). The distinction is necessary because the

noble gases are not filtered by the containment isolation system while the remainder of the radionuclide

inventory is filtered.

7-4
For example, consider the initiating event of a drop of the transfer cask in Stage 3 while it is being lifted

out of the cask pit. The drop frequency is 5.6x10 5/yr because the transfer cask is lifted once in Stage 3

and the drop probability is 5.6x10 5 /yr (Section 3.3.2). Because the drop height is as much as 13.0 meters

(42.5 feet), all of the fuel rods are assumed to breach the cladding (Section 4.4.2) and the probability of

MPC failure is 1.0 because the lid has not yet been welded to the MPC shell (Section 4.1.2.1); thus, the

probability of a release from the fuel rods and the MPC is 1.0. The release consists of both noble gases

and the radionuclide inventory in the MPC. However, only the noble gases that are released pass through

the water of the spent fuel pool. Because the noble gases are not filtered by the containment isolation

system, the probability of noble gases release is 1.0. The consequences of a release of noble gases is

1 x 101 ° probability of a latent cancer fatality (Section 6.2.2). The total risk for this scenario is 5.6x 10-15

annual individual probability of a latent cancer fatality.

7-5
18: Lowering transfer cask thru

equipment hatch 100 ft drop

19

Subsequent Years of Storage

tipped by seismic event 7.0x107 1.0xl0 6 All X 3.6x10 4 2.5x10' 6

34 struck by aircraft 6.3x10 9 L 0.014 All X 3.6x10 4 3.2x1014

struck by meteorite 3.5x1014 1 All X 3.6x10-4 1.3x10l

heated by aircraft fuel I 3.7x10 9 0 [ X N/A 0

Notes: "NG" = noble gases

"OT" = radionuclides other than noble gases

"All" = noble gases and all other types of radionuclides released

"X" = probability of release from containment in the storage phase is not applicable

7-6
Notes For Table 19

Stage 1: Loading fuel assemblies into the MPC. The probability of dropping a fuel assembly, 2.2 x 10'-, is

the product of the probability of dropping a single assembly, 3.2 x 10-5 (Section 3.3.1), and the number of

spent fuel assemblies loaded into a cask, 68. The drop probability is discussed in Section 3.3.1. The

probability of the fuel failing given that it is dropped is 6.4x 10.2 (Section 4.4.2).

The noble gases that are released pass through the water of the spent fuel pool and the secondary

containment filters. The probability of release from the containment is 1. Other radionuclides are retained

by the water. This is discussed in Section 3.1. Because this release is from a cask holding 68 assemblies,

the consequence for one assembly is 1/68 of this release, or 1.5x 1012. The risk of Stage 1 is the product of

the fuel assembly drop frequency, the probability of the fuel failing, the probability of a release of nobles

gases from the containment, and the consequences of the noble gas release, or 1.9x 10-16.

2. Stage 2: Placing the MPC lid on the MPC and engaging the lift yoke on the transfer cask. The transfer cask

is on the floor of the cask pit where it cannot fall any farther. There are no initiating events that can occur

in*this stage. Therefore, the risk of Stage 2 is zero.

3. Stage 3: Lifting the transfer cask out of the cask pit. The probability of dropping the transfer cask is

5.6x 10` (Section.3.3.2). All of the fuel rods fail (Section 4.4.2). The MPC is unsealed, thereby rendering

it ineffective in confining radionuclides. Section 4.1.2.1 discusses cracking and possibly more severe

damage or failure of the spent fuel cask pit floor that may occur if the cask is dropped from the maximum

height of 13 meters (42.5 feet) above the floor of the cask pit. Because the subject plant has a'singlefailure-proof crane and professional rigging personnel are contracted to move the cask, the consequences of

a drop of the transfer cask on the structural integrity of the cask pit do not have to be evaluated. It is

assumed that the consequences to the integrity of the MPC and fuel from such a drop are the same as for a

free drop in air of 12.2 meters (40 feet). For a drop onto concrete from this height, Table 13 shows that

100% of the fuel cladding is breached. Since the MPC is not sealed during this stage, the probability of a

release from the MPC is conservatively assumed to be 1.0. This assumption is conservative because it is

expected that the preponderance of the radioactive inventory will remain in the cask under a drop. The

water is too cold and the time is too short to have breaching other than the inventory in the pellet-cladding

gap. The water may wash the particulate that would have settled to the cask interior. The noble gases that

are released pass through the water of the spent fuel pool. Other radionuclides are retained by the water.

Because the noble gases are not removed by the filters, the probability of noble gasses release from

containment is 1. The consequences of a release of noble gases is 1.OxlO0

1 (Section 6.2.2). The total risk

for this scenario is 5.6x 1015 annual individual probability of a latent cancer fatality.

4. Stage 4: Moving the transfer cask over a railing of the spent fuel pool. The probability of dropping the

transfer cask is 5.6x 105 (Section 3.3.2). The MPC is unsealed and remains in the upright position when

dropped (Section 4.1.2.2). No fuel rods are expected to fail if the cask drops 0.9 meter (3 feet) during this

stage. Even if the rods don't fail, the drop will dislodge CRUD from the rod surface. This CRUD is

available for release. For the calculation of risk, it was assumed that all radionuclides are available for

release. The risk of Stage 4 is 3.0x 10"8.

5. Stage5: Moving the transfer cask to the preparation area (1st segment). The difference between Stages 4

and 5 is that the transfer cask is lowered from about 0.9 meter (3 feet) above the refueling floor to about 0.3

meter (1 foot). The possibility of dropping the transfer cask stems from the lowering. The other

constituents of risk are as discussed in the note for Stage 4.

6. Stage 6: Moving the transfer cask to the preparation area (2nd segment). The possibility of dropping the

transfer cask stems from the changed direction of travel along the load path of the refueling floor. The

other constituents of risk are as discussed in the note for Stage 4.

7-7
7. Stage 7: Moving the transfer cask to the preparation area (3rd segment). The possibility of dropping the

transfer cask stems from the changed direction of travel along the load path of the refueling floor. The

other constituents of risk are as discussed in the note for Stage 4.

8. Stage 8: Lowering the transfer cask onto the preparation area. The difference between Stage 8 and Stage 5

is that the transfer cask is lowered from about 0.3 meter (1 foot) above the refueling floor to the refueling

floor. The possibility of dropping the transfer cask stems from the lowering. The other constituents of risk

are as discussed in the note for Stage 4.

9. Stages 9 and 10: Stage 9: Preparing (drain, dry, inert, seal) the MPC for storage. Stage 10: Installing the

short stays and attaching the lift yoke. The transfer cask is on the refueling floor and cannot fall. There are

no other initiating events in Stages 9 and 10. The risk of Stages 9 and 10 is zero.

10. (Stages 11-17) Stage 11: Lifting the transfer cask. Stage 12: Moving the transfer cask to exchange the

bottom lids of the transfer overpack (1st segment). Stage 13: Moving the transfer cask to exchange the

bottom lids of the transfer overpack (2nd segment). Stage 14: Replacing the pool lid with the transfer lid.

Stage 15: Moving the transfer cask near the equipment hatch. Stage 16: Holding the transfer cask. Stage

17: Moving the transfer cask to the equipment hatch. The frequency of dropping the transfer cask in any

one of the stages is 5.6x105 (Section 3.3.2). No fuel rods are expected to fail. These stages are represented

by one row in Table 19 because the risk of each stage is 3.Ox 10'-.

11. Stage 18: Lowering the transfer cask to the storage overpack through the equipment hatch. The probability

of dropping the transfer cask is 5.6xl0.- (Section 3.3.2). The probability of MPC failure is 0.0002 for a

drop of 24.4 meters (80 feet) onto the storage overpack, and 0.02 (0.0196) for a drop of 30.5 meters

(100 feet) onto the concrete floor. The more conservative 30.5-meter (100-foot) drop results are used

(Sections 4.1.2.3 and 4.1.2.4). The entire inventory of radionuclides is considered to be released. Noble

gases are not removed by the filters and their probability of release from the containment is 1.0. The

consequences of a release of noble gases is 1.OxlO0

1' (Section 6.2.2). The total risk for this scenario is

1.1 xl 016 annual individual probability of a latent cancer fatality. The inventory of other radionuclides will

be released only if the containment isolation system fails, which has a probability of 1.5x 10-4 . The

probability of latent cancer fatality if the filters are off is 3.6x 104 . For the release of other nuclides, the

product of the risk components is 6.Ox 1

0 4 .

12. Stage 19: Preparing (remove short stays, disengage lift yoke, attach long stays) to lower the MPC. The

transfer cask is on the storage overpack where it cannot drop. The doors of the transfer lid of the transfer

cask are closed, hence, the MPC itself cannot drop. There are no initiating events in this stage. The risk of

Stage 19 is zero.

13. Stage 20: Lifting the MPC and opening the doors of the transfer lid.

Stage 21: Lowering the MPC through the transfer cask into the storage overpack.

The frequency of dropping the MPC in any one of the stages is 5.6x 10. (Section 3.3.2). This frequency is

the same as that in the preceding stages because the crane used to lift the transfer cask and the MPC itself is

the same and the rigging is similar. For this drop, 100% of the fuel is assumed to fail (Section 4.4.2). The

probability of the MPC failing is 0.28 (Sections 4.1.2.6 and 4.3.2). For this site, the risk associated with

these two stages is 8.5 x 10'3. For sites where the transfer of the MPC from the transfer cask to the storage

cask is done at the ISFSI pad instead of inside the 10 CFR Part 50 facility, the risk is approximately

5.6x 10', since the secondary containment is not available. In addition, since for these sites the transfers are

done using a hydraulic lift gantry instead of the plant's single failure proof crane, the drop frequency could

be higher which would then increase the final risk above 5.6x 10'.

7-8
14. Stage 22: Moving the storage cask into the airlock on Helman rollers. Stage 23: Moving the storage cask

out of the airlock on Helman rollers. Stage 24: Moving the storage cask away from the secondary

containment on Helman rollers. The storage cask is on rollers and cannot be dropped because it is not

lifted. The base of the storage cask on the platform of the rollers is sufficiently wide so that it cannot tip.

There are no other initiating events that occur in these stages. These stages are represented by one row in

Table 19 because the risk of each stage is zero.

15. Stage 25: Preparing (install lid, vent shield cross-plates, vent screens) the storage cask for storage. The

storage cask is stationary while it is being prepared and cannot be dropped. The base of the storage cask is

sufficiently wide that it cannot tip. There are no other initiating events that can occur. Hence, the risk is

zero.

16. Stage 26: Lifting the storage cask above the Helman rollers with the overpack transporter. Stage 27:

Moving the storage cask above a cushion on the preparation area. Stage 28: Holding the storage cask

above the cushion while attaching a Kevlar belt. Stage 29: Moving the storage cask above the concrete

surface of the preparation area. Stage 30: Moving the storage cask above the asphalt road. Stage 31:

Moving the storage cask above the gravel surface around the storage pads. Stage 32: Moving the storage

cask above the concrete storage pad. Stage 33: Lowering the storage cask onto the storage pad. The MPC

failure probability and the fuel failure probability are both approximately zero if the storage cask is dropped

from an elevation of 0.3 meter (1 foot) onto concrete, asphalt, or gravel (Section 4.1.3.1). Because the risk

is so small, the initiating event frequencies and the consequences are not determined. These stages are

represented by one row in Table 19 because the risk of each stage is approximately zero.

17-24. Stage 34: Storing the storage cask on the storage pad for 20 years. To calculate the risk, this stage is

divided into substages depending on the initiating event. The risk is also calculated separately for the first

year of service and for the subsequent 19 years on the storage pad. However, because the duration of the

handling and transfer phases is much less than 1 year, the frequencies of all initiating events in the first year

of service are essentially the same as the corresponding frequencies for subsequent years on the storage pad.

17. Storage cask tipped by a seismic event during the first year of service. The initiating event frequency is

S7x×10 7 (Section 3.3.3). If the storage cask tips over no fuel rods are expected to fail (Sections 4.1.3.2 and

4.4.2). Even if the rods don't fail, a tip-over will dislodge CRUD from the rod surface. This CRUD is

available for release. For the calculation of risk, it was assumed that all radionuclides are available for

release. The risk at this stage is 2.5x 10" 6 , because the MPC has such a low probability of failure.

18. Storage cask struck by an aircraft during the first year of service. The frequency of the strikes that could

impact the storage cask is 6.3x 10. (Section 3.3.7). Based on the results in Section A.4.9.3, the probability

of breaching the MPC, if struck by an aircraft during landing or takeoff is extremely small and is assumed to

be zero. However, many commercial aircraft overflying the site would be larger than a Gulfstream IV and

could impact the cask at high velocity. Since the makeup and characteristics of commercial 'ircraft

overflying the site are not known and have not been evaluated, it is conservatively assumed that all

commercial aircraft overflying the site are larger than a Gulfstream IV and that. the probability of MPC

breach due to the impact of an overflying commercial aircraft is, for the purpose of the PRA, 1.0.

Therefore, the probability of MPC failure and release, if struck by an aircraft, is, for the purpose of the

PRA, equal to the frequency of overflight crashes (9.Ox1-W', see Section 3.3.7) divided by the total

frequency of aircraft crashes (6.3x10 9 , see Section 3.3.7), which is 0.014. Because the ISFSI is outside the

containment, the radionuclides are released to the environment. Because all of the fuel rods are assumed to

fail, the consequences are 3.6x 10. (Section 6.3.1, Table 18). The risk is the product of the aircraft strike

frequency, the MPC failure probability and the consequences, or 3.2x 10- 4

.

7-9
19. Storage cask struck by a meteorite during the first year of service. The frequency of meteorites that could

breach the storage cask is 3.5 x 10"4 (Section 3.3.5). The probability of a release from the MPC is

conservatively set to 1, for the purposes of this PRA. Because the ISFSI is outside the secondary

containment, the radionuclides are released to the environment. Because all of the fuel rods are assumed to

fail, the consequences are 3.6x 104 (Section 6.3.1, Table 18). The risk is the product of the meteorite strike

frequency, the MPC failure probability and the consequences, or 1.3 x 107.

20. Storage cask heated by burning fuel when an aircraft crashes during the first year of service. It is assumed

that a fire will result whenever an aircraft hits the storage cask. The frequency of this is the sum of the

frequencies of takeoff or landing crashes plus overflight crashes. From Section 3.3.7, the frequency of the

storage cask being struck by an aircraft and heated by burning jet fuel is 6.3x 10'. From Appendix B,

Section B. 1.8, any fire caused by a takeoff, landing or overflight crash will not fail the cask. Hence, the

probability of a release and the risk are both zero.

21-24. These notes are the same as Notes 17-20, respectively, for each of the subsequent 19 years of storage.

7-10
8. OTHER ASPECTS OF DRY CASK STORAGE RISK

Table 19 identifies conditions that could potentially affect the MPC's response to loads. Other possible

conditions are described in the paragraphs of this section and summarized in Table 20. None of these

conditions is expected to affect the integrity of the MPC.

Table 20. Conditions Potentially Affecting MPC Response to Initiating Events

Conditions Affecting the MPC Storage Section

Excessive hot fuel assemblies mistakenly loaded 8.1

Criticality 8.2

Corrosion from coastal atmosphere 8.3

Corrosion from long-term industrial releases 8.3

Pressurization from corroding fuel 8.4

Pressurization from failed fuel 8.4

Hydrogen gas accumulation when the MPC lid and shell are welded 8.5

8.1 Mis-Loading of Spent Fuel into the MPC

For the purposes of the PRA, sensitivity analyses were performed to determine MPC failure probabilities

for a gross mis-loading of wrong fuel assemblies into the MPC. The scenario of a gross series of errors,

resulting in every fuel assembly loaded into the MPC being incorrect/insufficiently cooled is not

considered credible, but useful for the purposes of exploring the impact of such an event. In these

scenarios the heat load in the MPC was increased from 21 kW to 33 kW, 66 kW, and 126 kW. These

heat loads correspond to fuel that has cooled in the spent fuel pool for 5 years, 2.5 years, 1 year, and 0.5

year, respectively. Even if every single fuel assembly loaded into the MPC were incorrect, no MPC

failures are predicted for heat loads of 21 kW, 33 kW, and 66 kW. But if the heat load were 126 kW, the

MPC could be expected to fail from creep rupture. The probability of failure in such an unlikely case

would be 0.02 after 20 years of storage. No fuel failures are expected for heat loads of 21 kW and 33

kW. In these extremely unlikely scenarios, fuel failure could be expected to occur when the heat load is

6.6 kW and 126 kW. The fuel temperature would reach 543 'C (1,010 °F) and 809 'C (1,488 °F),

respectively when the cask reaches thermal equilibrium. The long-term temperature limit for dry fuel

storage is 400 'C (752 °F). [It is highly unlikely that rods at 543 °C (1,010 OF) would fail. Also, 400 'C

(752 °F) is not a rod failure criterion. It was chosen to limit the hydride reorientation problem. It also

limits creep.] Based on the extremely low likelihood of a gross mis-loading scenario, made all the more

unlikely by the significant procedural requirements, fuel assembly verifications and instrumentation used

in the process, it is concluded MPC integrity will not be affected by any occurrences of fuel mis-loading.

8-1
8.2 Potential for Criticality

A criticality study was performed on the transfer cask in Stage 3, during which the MPC is loaded with

spent nuclear fuel while it is in the cask pit. This study was intended to provide insights as to degree of

sub-criticality expected during limiting (not conservative) MPC loading operations. The reader is

encouraged to keep in mind that these assumptions, while possible, are highly unlikely because we are

effectively assuming that reactor operators will be discharging fuel well before its operational lifetime.

Further depletion in the core will reduce the eigenvalues shown in this section and, therefore, the results

of this study can be considered limiting. This study is not a rigorous nuclear criticality analysis of MPC

operation. Since the results demonstrated roughly a 20% sub-critical margin, a more detailed analyses

was deemed unnecessary. The analysis is only of events and conditions that are within the scope of the

PRA as discussed in Section 1.2. The possibility of criticality was also studied in other stages, such as

Stage 34, during which the storage cask is in the ISFSI. Given the assumptions of this analysis, the

results are applicable to only BWR fuel in the subject cask. A BWVR fuel assembly designed by General

Electric was studied because it is highly enriched and contains large amounts of burnable poisons that

will tend to maximize the reactivity increase effect. The BWR assembly was the most reactive assembly

found in available fuel design references. Burnable poisons were modeled as four concentric annuli, and

the assembly was burned pin-by-pin in steps of 1000 megawatt-days per metric ton (MWd/MTU). The

enrichment was used as designed; the fuel depletion was modeled using the as designed assembly, rather

than by assuming an initial homogeneous enrichment.

The SCALE suite of codes (Reference 47) was used to perform this analysis. The fuel was depleted with

the TRITON module and criticality analysis were performed with KENO-VI using the 238 group

ENDF/B-V library.

The fuel in the canister was modeled to the same level of detail as in the burnup calculation. The model

was essentially 2-D. A 15.2-cm (6-inch) "slice" of the fuel using its most reactive axial zone and applied

reflective top and bottom boundary conditions was modeled. One-quarter of the MPC basket was

modeled with two lines of symmetry. The basket can be modeled as being surrounded by an effectively

infinite annulus of water because an infinite reflector should yield the most reactive configuration.

To allow for fuel compression/expansion studies, it was confirmed that an isolated cell of the basket

model yields effectively equivalent answers as the entire model. By this we mean that an isolated cell

gave approximately the same eigenvalue as the complete MPC model. The fuel compression/expansion

study assumed a reduction (or increase) of the pin pitch to ensure that there is not a compression (or

expansion) accident condition that could lead to a significant increase in reactivity which would cause

the eigenvalue to approach a 5% sub-critical margin. The MPC poisons were assumed to stay in the '

cask. This model is sufficient for trending studies, given the expectation that the results will be highly

sub-critical.

Four concerns of criticality are the following:

(1) excessive heat generation can overpressurize the MPC

(2) excessive gamma radiation may be insufficiently shielded

(3) excessive neutron radiation may be insufficiently shielded

(4) radionuclides could be released if the fuel cladding were to fail

8-2
Several calculations were performed to evaluate MPC performance. The first observation is that the

compression/expansion study demonstrates that there are no credible compression/expansion scenarios

which lead to an increase in the eigenvalue to anywhere near the 5% sub-critical margin. Similar

analyses are shown in Reference 48. The second observation is that there are no low density, optimal

moderation configurations which means that pool draining events will not lead to a reduction in the subcritical margin. Figure 19 and the work in Reference 49 confirm this by showing the predicted

eigenvalue at the upper bound 95th percentile confidence interval with reduced water density. As is

evident from the figure, even when changes in water density are considered, the fuel in the MPC is highly

sub-critical. In summary, the criticality analyses performed for this study, while limited in scope, show

similar trends as other studies by confirming that the configuration is under-moderated and highly subcritical.

When the MPC is dry, criticality is not

physically possible under any conditions Meani

in the absence of other neutron sources. IL!ef\aIUe

This is true because the original fuel 0.84

enrichment was limited to 5 weight

percent, which is a limit based in part on 0.83

the requirement that unmoderated 0.S2.

criticality is not possible, even for BWR

fuel at its most reactive point in life. 0.81

This position is based on experimental

data and calculations in Reference 50. 0.80

0.79'

The boral plating in the MPC-68 basket

is an effective means of criticality 0.78

control; criticality is not allowed whether

moderated or unmoderated. No credible 0.77.

scenario which would eliminate the boral 0.76 1 I I I I

and leave the geometry of the fuel intact 0.7 0.8 0.9 1.0

has been identified. Even if the boral w81c1 i)ClSj.\

plates were to separate from the basket t i'

structure, they would have limited room

to move and, once the basket is sealed, Figure 19. Upper 95th Percentile Confidence Interval

there is no credible means by which they Eigenvalue for MPC-68 in Pure Water at 27 °C (81 'F)

could fall out of the basket. If the

contents of the MPC-68 were to

somehow relocate, the boral would relocate with the debris. In all likelihood, this postulated

configuration would be highly subcritical because of the boral and because a bed of debris is not an

optimal geometry. At the enrichments used in BWR fuel, a square pitch lattice is the optimal geometry

for criticality. In the absence of moderation, experiments and calculations have demonstrated that

criticality is not possible at the enrichments currently used in light-water reactor (LWR) fuel.

8-3
8.3 Corrosion

The MPC, which acts as the confinement boundary for the HI-STORM dry cask storage system, is

constructed entirely from austenitic stainless steel Types 304, 316, 304L, or316L. All of these stainless

steel grades are corrosion resistant in high-humidity and industrial environments. Therefore, coastal and

industrial atmospheres should have no effect on the confinement ability of the MPC (Reference 51).

The MPC is drained, dried, and filled with helium. Helium is an inert gas; it does not react with the fuel

cladding or the internal structures of the MPC.

8.4 Pressurization

BWR fuel rods are normally pressurized with helium to 0.69 MPa (100 psi) (Reference 14). The fuel

failure analysis of Appendix C uses a conservative assumption of 15.5 MPa (2250 psi) for internal fuel

rod pressure. If the fuel rods were to breach, such as by creep rupture, the pressure in the rods would

pressurize the MPC. An incremental increase in pressure from all of the fuel rods in the MPC rupturing

is addressed in the Appendix B assessment of the response of the MPC to mechanical and thermal loads,

Section B. 1.8, "Results." Fuel failure would cause the internal pressure of the MPC to increase. If a

mechanical impact event did cause fuel failure, the resulting increase in pressure would not occur during

the drop accident scenarios. The increase in internal pressure due to fuel failure would be a result of the

mechanical loading, but because of the extremely short duration of the impact, it would not affect the

MPC conditions until after the mechanical loading is complete.

The failure analysis of the vent blockage scenario assumed bulk internal pressure takes into account the

increased pressure from fuel failure. These results are shown in Appendix B, Table B4. The failure

analyses show that even after 20 years of vent blockage and 20% fuel failure, the added pressure from

failed fuel does not result in any expected MPC failures.

Calculations discussed in more detail in Appendix C determined that for the vent blockage scenario, all

fuel rods located within a radius of 0.38 meter (1.25 foot) from the cask center will have a maximum clad

temperature exceeding the long-term temperature limits. Using a ratio of affected areas, the fraction of

fuel expected to exceed long-term temperature limits is 0.20.

8.5 Hydrogen Generation in the MPC

In Stage 9, the transfer cask is in the preparation area where the MPC lid is welded to the MPC shell.

The water in the MPC is lowered to about 0.3 meter (1 foot) below the lid. While welding the lid to the

shell, the root pass, several intermediate passes, and the final pass are tested with dye penetrant. The

water level is then raised to the lid elevation, and the MPC is pressurized to hydrostatically test the weld.

Should no leakage occur, the MPC is drained, dried, and filled with helium. The cover plates and closure

ring are then installed and welded.

8-4
When the MPC contains water, reactions can occur that generate hydrogen gas. When the water level is

below the MPC lid prior to hydrostatic testing, hydrogen gas can accumulate in the space between the lid

and the surface of the water while the lid is welded to the shell. To prevent accumulation, the subject

plant vents the space by pumping argon into the MPC through the drain line. The argon percolates

through the water to the top of the MPC and exits through the port, purging hydrogen from the space.

Another reason for purging the space between the water surface and the lid is to eliminate any possible

adverse effect that hydrogen could have on the lid-to-shell weld. Hydrogen in weldmetal can cause

cracks if the concentration is high enough and the material being welded is carbon steel or alloy steel.

The type of stainless steel presently used for spent fuel canisters is highly resistant to hydrogen induced

cracking. Despite this resistance to hydrogen induced cracking, this step to preclude hydrogen from the

weld area is always implemented to further ensure high-quality welds.

The lid/shell and the closure ring/shell welds are not evaluated in the PRA because the lid/shell weld is

more robust than the other welds of the MPC. Less-ductile welds, such as the baseplate weld, are

evaluated in the PRA. For the following reasons the lid/shell weld is not evaluated:

* Because of the welding process, the lid/shell weld and the closure ring/shell weld is expected to

be tougher than other welds of the cask.

The MPC lid has redundant welding, consisting of the. lid/shell weld and the closure ring/shell

weld.

Stress in the lid/ shell weld are low for all drop events.

The baseplate weld is modeled in the failure models of the MPC.

8-5

9. CONCLUSIONS

Starting with a comprehensive list of initiating events (Tables 3-5), the annual risk to the public from the

onsite handling, transfer, and storage of a single transfer cask and storage cask at a specific BWR site is

estimated. First, the frequencies of all initiating events which could affect the subject cask at the subject

site are estimated. Second, the responses of the MPC and fuel to the mechanical and thermal loads

induced by the initiating events are determined. Third, the consequences of a release of radioactive

material to the environment is calculated in terms of the individual probability of a latent cancer fatality.

(No prompt fatalities are predicted.) Finally, these results are combined to estimate the risk to the public

in terms of the annual probability of a latent cancer fatality.

Because no uncertainty analysis was performed, the identification of the dominant contributors was

based on the point estimates developed by this study. The dominant contributors might possibly change

if uncertainty were considered.

The risk estimates are extremely low. Some of the scenarios have zero risk because either their initiating

events cannot occur at the subject plant or no radioactive release will result. As Table 19 shows, the

risks of the scenarios with nonzero risk are extremely low because they are the products of two or three

very small numbers.

For example, when lowering the transfer cask through the equipment hatch, the initiating event of

dropping the transfer cask has a frequency of 5.6x 10' per year and a probability of 0.02 failure of the

MPC. If only noble gases are released, the consequence probability is 1.0x 10"1, if other radioactive

materials are released, the consequence probability is 3.6x 10-4, and the probability of the secondary

containment failing to isolate is 1.5 x 104. Thus, the risk of the noble gas release is:

(5.6x 104)(0.02)(1.Ox 10-'°) = 1.1 x1016

and the risk of the release of other radionuclides is:

(5.6x 10-)(0.02)(1.5 x 10-4)(3.6x 10-4) = 6.Ox 10-14

The handling phase Stages 4-8 have negligible risk due to drops of small heights. The probability of

release of radioactive material from the MPC is lxi 0-6. This produces extremely low risk and is shown

in the risk columns of Table 19 as 3.0x1018.

There are two other stages (Stages 1 and 3) in the handling phase that have negligible risk. These stages

involve release of noble gases if the containment fails to isolate. The releases occur because the lid of

the MPC is not welded shut until Stage 9. The risk of a noble gas release is 1.9x 10-16 in Stage 1 and

5.6x10-15 in Stage 3.

9-1
The estimated aggregate risk is an individual probability of a latent cancer fatality of 1.8x 10-12 during the

period encompassing the initial cask loading and first year of service, and 3.2x 10-1' per year during

subsequent years of storage. The only significant contributions to the risk during the handling phase are

from~drops of the transfer cask. There is zero risk associated with the transfer phase. During the transfer

phase (movement of the storage cask out to the storage pad), drops of the storage cask are not high

enough to induce sufficient stresses to fail the MPC or fuel cladding. During the storage phase, the

dominant contribution to the risk is from aircraft striking the storage cask, although the risk is still

extremely small.

9-2
10. REFERENCES

1. Memorandum to William D. Travers, Executive Director for Operations, from Annette L.

Vietti-Cook, Secretary, "Staff Requirements - SECY-98-144 - White Paper on Risk-Informed

and Performance-Based Regulation," Washington, DC, March 1, 1999.

2. HOLTEC International, "HI-STORM Topical Safety Analysis Report," HI-951312, Revision 8,

Marlton, New Jersey, June 25, 1999.

3. U.S. Nuclear Regulatory Commission, "Severe Accident Risks: An Assessment for Five U.S.

Nuclear Power Plants," NUREG-1 150, Washington, DC, December 1990.

4. U.S. Nuclear Regulatory Commission, "Safety Goals for the Operation of Nuclear Power Plants"

(Corrections and Republication of Policy Statement), Federal Register, Volume 51, No. 162,

pp. 30028-30033, Washington, DC, August 21, 1986.

5. Chapman, O.J.V., and F.A. Simonen, "RR-PRODIGAL Model for Estimating the Probabilities of

Defects in Reactor Pressure Vessel Welds," NUREG/CR-5505, U.S. Nuclear Regulatory

Commission, Washington, DC, 1998&

6. R Kenneally, J. Price, and D. Koelsch, "Dry Cask Storage Characterization Project: Final

Report," Electric Power Research Institute, Palo Alto, California; U.S. Nuclear Regulatory

Commission, Washington, DC; U.S. Department of Energy, OCRWM, North Las Vegas,

Nevada; and U.S. Department of Energy, Idaho Operations, Idaho Falls, Idaho, EPRI Doc. No.

1002882, September 2002.

7. U.S. Nuclear Regulatory Commission, "PRA Procedures Guide," NUREG-2300, Washington,

DC, January 1983.

8. U.S. Nuclear Regulatory Commission, "Code Manual for MACCS2," NUREG/CR-6613, Vol. 1,

Washington, DC, May 1998.

9. Hall, Eric J., Radiobiologyfor the Radiologist, 4 th Edition, J.B. Lippencott Co., Philadelphia,

Pennsylvania, 1994.

10. Generic Letter 89-22, "Resolution of Generic Safety Issue No. 103, 'Design for Probable

Maximum Precipitation'," U.S. Nuclear Regulatory Commission, Washington, DC, October 19,

1989.

11. National Oceanic and Atmospheric Administration, National Weather Service,

Hydrometeorological Report (HMR) No. 51, "Probable Maximum Precipitation Estimates,

United States East of the 1 0 5th Meridian," Washington, DC, June 1978.

12. U.S. Nuclear Regulatory Commission, "A Survey of Crane Operating Experience at U.S. Nuclear

Power Plants from 1968 through 2002," NUREG-1 774, Washington, DC, July 2003.

10-1
13. U.S. Department of Energy, "Integrated Database Report - 1994: U.S. Spent Nuclear Fuel and

Radioactive Waste Inventories, Projections, and Characteristics," DOE/RW-006, Rev. 11,

Washington, DC, September 1995.

14. U.S. Nuclear Regulatory Commission, "A New Comparative Analysis of LWR Fuel Designs,"

NUREG-1754, Washington, DC, December 2001.

15. U.S. Nuclear Regulatory Commission, "Information Digest," NUREG-1350, Vol. 11,

Washington, DC, 1999.

.16. U.S. Nuclear Regulatory Commission, "Control of Heavy Loads at Nuclear Power Plants:

Resolution of Generic Technical Activity A-36," NUREG-0612, Washington, DC, July 1980.

17. U.S. Nuclear Regulatory Commission, "Revised Livermore Seismic Hazard Estimated for

Sixty-Nine Nuclear Power Plant Sites East of the Rocky Mountains," NUREG-1488,

Washington, DC, April 1994.

18. U.S. Nuclear Regulatory Commission, "Tornado Climatology of the Contiguous United States,"

NUREG/CR-4461 (PNL-5697), Washington, DC, 1986.

19. Fuj ita, Theodore, U.S. Tornadoes, Part One, 70-Year Statistics, Satellite and Mesometeorology

Research Project (SMRP), Department of the Geophysical Sciences, The University of Chicago,

SMRP Research Paper Number 218, Library of Congress Catalog Card Number 86-51637,

National Technical Information Service PB 87-127742, Chicago, Illinois, Copyright 1987.

20. Kring, D.A., Meteorites and Their Properties, Lunar and Planetary Laboratory, University of

Arizona, Tucson, Arizona, 2nd Edition, 1998.

21. French, Bevan M., Traces of Catastrophe.: A Handbook of Shock-Metamorphic Effects in

Terrestrial Meteorite Impact Structures, Lunar and Planetary Institute, 1998.

22. Cassidy, William A., Estimated Frequency of a Meteorite Striking An Aircraft, University of

Pittsburgh, Department of Geology and Planetary Science, Pittsburgh, Pennsylvania,

December 17, 1997.

23. HOLTEC International, HI-STORM 100 LAR 1014-1, Revision 1, Supplement 1, Marlton, NJ,

October 2000.

24. Watson, A.I., and R.L. Holle, "An Eight-Year Lightning Climatology of the Southeast United

States Prepared for the 1996 Summer Olympics," Bulletin of the American Meteorological

Society, Vol. 77, No. 5, Boston, Massachusetts, May 1996.

25. International Electrotechnical Commission, "Protection of Structures Against Lightning - Part

1: General Principles, Section 10: Guide A: Selection of Protection Levels for Lightning

Protection Systems," IEC 61024-1-1, Ed. 1.0 b: 1993.

10-2
26. U.S. Nuclear Regulatory Commission, "Standard Review Plan for the Review of Safety Analysis

Reports for Nuclear Power Plants," NUREG-0800, Rev. 2, Section 3.5.1.6, "Aircraft Hazards,"

Washington, DC, July 1981.

27. U.S. Department of Energy, "Guidance and Data for Impact Frequency Calculation," Appendix

B to DOE Standard 3014-96, Washington, DC, October 1996.

28. Lawrence Livermore Software Technology, LS-DYNA Computer Code, Version 960, Livermore,

California, March 2001.

29. U.S. Nuclear Regulatory Commission, "Standard Review Plan for the Review of Safety Analysis

Reports for Nuclear Power Plants," NUREG-0800, Rev. 2, Section 3.5.1.4, "Missiles Generated

by Natural Phenomena," Washington, DC, July 1981.

30. American Society of Mechanical Engineers, ASME Boiler and Pressure Vessel Code, Section III,

New York, New York, 1989.

31. U.S. Nuclear Regulatory Commission, "Evaluation of Explosions Postulated to Occur on

-Transportation Routes Near Nuclear Power Plants," Regulatory Guide 1.91, Revision 1,

Washington, DC, February 1978.

32. Mechanics Research, Inc., for the Tennessee Valley Authority, "Nuclear Power Plant Risks from

a Natural Gas Pipeline," Knoxville, Tennessee, 1974.

33. Fluent Incorporated, Fluent (Computational Fluid Dynamics Software), Release 6, Lebanon,

New Hampshire, 2000.

34. HOLTEC International, HI-STORM 100 Thermal Calculations, HOLTEC Report No: HI981892, Marlton, NJ, November 2000.

35. American Concrete Institute, ACI 349-01, "Code Requirement for Nuclear Safety Related

Concrete Structures," Farmington Hills, Michigan, 2001.

36. Society of Fire Protection Engineers (SFPE), The SFPE Handbook of Fire Protection

Engineering, 2nd Edition, Chapter 3-11, Section 3, pp. 3-202, and 3-199, Bethesda, Maryland,

1995.

37. Kramer, M.A., et al., "Measurements of Heat Transfer to a Massive Cylindrical Object Engulfed

in a Regulatory Pool Fire," Proceedings of NHTC '01: 3 5th National Heat Transfer Conference,

Anaheim, California, June 10-12, 2001.

38. Drysdale, D., An Introduction to Fire Dynamics, John Wiley and Sons, Hoboken, New Jersey,

1996.

39. National Fire Protection Association, "Standard for the Installation of Lightning Protection•

Systems," NFPA-780, Quincy, Massachusetts, 1995.

10-3
40. U.S. Nuclear Regulatory Commission, "Standard Review Plan for Dry Cask Storage Systems,"

NUREG-1536, Chapter 2 "Guidance: Principal Design Criteria," Washington, DC, February 1,

1996.

41. Brauer, J.R., and B.S. Brown, EMAS User's Manual, Version 4, Ansoft Corporation, Pittsburgh,

Pennsylvania, July 1997.

42. U.S. Nuclear Regulatory Commission, "Failure Behavior of Internally Pressurized Flawed and

Unflawed Steam Generator Tubing at High Temperatures: Experiments and Comparison with

Model Predictions," NUREG/CR-6575, ANL-97/17, Washington, DC, March 1998.

43. Chun, R., Witte, M., Schwartz, M., "Dynamic Impact Effects on Spent Fuel Assemblies,"

UCID-21246, Lawrence Livermore National Laboratory, Livermore, California, 1987.

44. General Electric Corporation, "ABWR Standard Safety Analysis Report," Chapter 19D,

"Probabilistic Evaluations," Amendment 31, July 1993.

45. Combustion Engineering, Inc., "CESSAR Design Certification, System 80+ Standard Design

FSAR, Appendix 19.5A - Data Calculations for Generic Component Data," Amendment M

(March 15, 1993) and Amendment W (June 17, 1994).

46. Jow, J.N., et al., "MELCOR Accident Consequence Code System (MACCS), Model

Description," NUREG/CR-4691, SAND-1562, Vol. 2, U.S. Nuclear Regulatory Commission,

Washington, DC, February 1990.

47. U.S. Nuclear Regulatory Commission, "SCALE: A Modular Code System for Performing

Standardized Computer Analyses for Licensing Evaluation," Vols. I-III, NUREG/CR-0200,

Rev. 6 (ORNL/NUREG/CSD-2/R6), Washington, DC,.May 2000.

48. U.S. Nuclear Regulatory Commission, "Technical Study of Spent Fuel Pool Accident Risk at

Decommissioning Nuclear Power Plants," NUREG-1 738, Washington, DC, February 28, 2001.

49. U.S. Nuclear Regulatory Commission, "Reexamination of Spent Fuel Shipment Risk Estimates,"

NUREG/CR-6672, Washington, DC, March 2000.

50. Paxton, H.C., and N.L. Pruvost, "Critical Dimensions of Systems Containing U235, Pu239, and

U233," Report LA-10860-MS, Los Alamos National Laboratory, Los Alamos, New Mexico, July

1967 (1986 Revision).

51. American Society for Metals, "Corrosion Resistant Materials," Metals Handbook Desk Edition,

Materials Park, Ohio, 1985.

10-4
APPENDIX A

MECHANICAL LOADS

APPENDIX A

MECHANICAL LOADS

A.1 Purpose

While conducting the dry cask storage PRA, specific questions were identified that related to the

handling, transfer, and storage of the cask system and the effects of natural phenomena such as wind,

flood, and earthquake. This appendix contains summaries of the results and corresponding calculations

to address each of those questions.

The purpose of this appendix is to examine several initiating events during the handling, transfer, and

storage of the cask system and to determine the effects of natural phenomena (wind, flood, and earthquake).

The results of this evaluation will provide the stresses in the multi-purpose canister (MPC), which contains

the spent fuel that will be used in subsequent evaluations to determine the probability of breaching the MPC.

A.2 Scope

The scope of this evaluation is limited to performing calculations in response to specific questions.

These questions all relate to the transfer cask and the storage cask used at the subject plant. The

questions are as follows:

Handling Phase Questions

(1) If the fuel handling building crane is traveling at maximum speeds [bridge 15 m/min (50 ft/min),

trolley 3 m/min (10 ft/min), and hoist 1.3 m/min (4.2 ft/min)] with an MPC and its lid secured by

gravity in the transfer overpack, what drop height causes the transfer cask to tip-over onto the

concrete floor?

(2) Same question as Question 1, but the MPC lid is "sealed."

(3) Analyze the vertical drop of the transfer cask containing a sealed MPC falling on the concrete

floor at ground elevation from heights of 1.5, 12.2, 21.3, and 30.5 meters (5, 40, 70, and 100

feet). Determine the magnitude and location of maximum stresses/strains in the MPC.

(4) Analyze the vertical drop of the transfer cask containing a sealed MPC falling onto the storage

overpack at ground elevation from heights of 1.5, 12.2, and 24.4 meters (5, 40, and 80 feet).

Determine the magnitude and location of maximum stresses/strains in the MPC.

(5) Determine the magnitude and location of maximum stresses/strains in the MPC when a sealed

MPC falls vertically 5.8 meters (19 feet) into the storage overpack.

A-1
Transfer Phase Question Topics

(6) If the storage cask drops onto the independent spent fuel storage installation (ISFSI) concrete

pad, asphalt area, or gravel area, while it is hung from the top trunnion and being transported by

the crawler, will it slide and/or tip-over? If tip-over can occur, what velocity of the crawler

would be required to tip it over? How much time does the operator have to stop the crawler

before tip-over occurs? What are the stresses in the MPC?

(7) If the crawler vehicle weighing 71,440 kg (157,500 lb) is traveling at 0.64 km/h (0.4 mi/h) while

carrying the storage cask weighing 163,293 kg (360,000 lb) and hits another storage cask on a

concrete pad, will the impact cause the struck cask to slide or tip-over?

(a) If it slides, will a sliding storage cask hit another storage cask? If it hits another storage

cask, what are the maximum stresses in the MPC?

(b) If it tips over, what are the maximum stresses in the MPC?

(c) If it does not slide or tip-over, what travel velocity of the crawler would be necessary to

strike the storage cask and cause it to slide and/or tip-over?

Storage Phase Question Topics

(8) What seismic forces will cause the storage cask to slide and/or tip-over?

(9) What impact forces attributable to a given small aircraft, tornado missiles (beyond design basis)

and heavy objects in flood waters will cause the storage cask to slide and/or tip-over? If the

impacting object penetrates the storage overpack, what will be the stresses in the MPC?

(10) If the storage cask tips over on the concrete pad, what will be the stresses in the MPC?

(11) If the storage cask tips over, will the impact of the cask tipping over on another storage cask or

on the concrete pad cause the overpack lid to be dislodged such that the MPC could slide out?

(12) Will the structural integrity of the cask system be jeopardized by shock waves caused by

postulated explosions? If not, provide the supporting arguments.

A.3 Method

To evaluate the structural response of the cask system to the event scenarios posed by the twelve

questions, several methods of analysis were used, which included hand calculations based on first

principles, common analytical methods and industry recognized approaches, and solution of the

differential equations of motion for which closed form solutions were obtained. Wherever possible

existing analyses, such as the Holtec finite element analysis of the MPC, fuel basket and fuel basket

supports (Reference A. 1), that had been performed for a number of impact events were utilized by

scaling the stress results in Reference A. 1 by the ratio of the maximum acceleration calculated in

response to the question divided by the acceleration in Reference A. 1. For the cask drop impact

scenarios, the Holtec finite element models were modified and expanded, as discussed in the appropriate

sections below.

A-2
A.4 Calculation Summary Descriptions

A.4.1 Question 1

If the fuel handling building crane is traveling at maximum speeds [bridge 15 m/min (50 fl/min), trolley 3

mr/min (10 ft/min), and hoist 1.3 m/min (4.2 ft/min)] with an MPC and its lid secured by gravity in the

transfer overpack, what drop height causes the transfer cask to tip-over onto the concrete floor?

A.4.2 Question 2

Same as Question 1 but the MPC lid is "sealed."

The response to Questions 1 and 2 are combined into one evaluation. Solutions are obtained for various

cask orientations when it is dropped. When the transfer cask impacts the concrete floor it is assumed to

have the same orientation and horizontal velocity as when it was dropped. The vertical velocity is

increased (positive measured down) by the velocity added during free fall. At impact the horizontal and

vertical velocity components are transferred to an equivalent rotational velocity. The transfer cask then

rotates about the point of impact until it either tips over (when the center of gravity intersects a vertical

line immediately above the point of impact) or rotates back to a vertical orientation. The question is

resolved by determining the rotational velocity needed to tip-over the transfer cask. It should be noted

that the transfer cask drops at an acceleration of 1 g so that the lid secured by gravity would not separate

from the MPC during the fall. Therefore, the answers to Questions 1 and 2 are the same. Detailed

analyses are first summarized and then followed by a discussion of the answer.

With No Initial Velocity

The specific geometry of the transfer cask is shown in Figure A. 1. The weight of the loaded transfer cask

is assumed to be 86,636 kg (191,000 lb) but as will be seen the weight does not enter into the solution.

Note that the transfer cask is dropped at an angle (a) with the vertical and the drop height to the center on

the base of the cask is H. When the angle (a) equals 90 - D = 670, the transfer cask is oriented so that the

center of gravity of the cask is located above the point on the cask that first impacts the slab. It is

obvious that the transfer cask will tip-over if it is dropped at angles less than 670.

A-3
D

Figure A.1. Geometry of Transfer Cask

Note that it is not possible to drop the transfer cask at angles smaller than ca =67' unless the drop height

(Hm) is greater than that calculated from the following expression:

Hm

Cos - (A.1) D/2

where: D = 2.06 m (81.25 inches)

L = 4.84 m (190.5 inches)

h = 2.39 m (93.95 inches)

C = 2.64 m (104 inches)

S =230

Hm = / * D * Cos a

Therefore, with no initial velocity a tip-over will occur only if the drop height is greater than 40 cm (15.9

inches) and the angle a is less than 67'.

A-4
With Initial Velocity

At impact the horizontal (VO) and vertical (Vv) velocities of the cask are transformed to a rotational

velocity of the transfer cask (o').

The subsequent response of the transfer cask is developed next. A sketch of the transfer cask defining

the response parameters used is shown in Figure A.2.

ac

I).

C Cos ()T-U~4+0)

Figure A.2. Rotation of Transfer Cask after Impact

The response of the cask is then governed by:

01=

VhSin(z - ¢- a) - VvSin(o + 6a -;/2)

C

(A.2)

J&"+WC Cos(z- -a±O)=O (A.3)

with the initial conditions, 0 (0) = 0 ; and 0'(0)= 0o' and where J is the rotational inertia of the transfer

cask.

A-5
A.4.3 Question 3

Analyze the vertical drop of the transfer cask containing a sealed MPC falling on the concrete floor at

ground elevation from heights of 1.5, 12.2, 21.3, and 30.5 meters (5, 40, 70, and 100 feet). Determine the

magnitude and location of maximum stresses/strains in the MPC.

Model Development

To respond to this question, a finite element model was developed that includes the MPC, fuel

assemblies, fuel basket, fuel basket supports, transfer overpack, and concrete floor and wall (Reference

A.24). These components are shown in Figures A.3 through A.5. The finite element model (FEM)

shown in Figure A.3 contains all of the structural components mentioned above. The model is a

quarter-model representation of the cask system and concrete floor/wall, which takes advantage of

symmetry about two orthogonal vertical planes through the centerline of the MPC. Figure A.4 shows a

close-up view of the transfer overpack (TO) and MPC resting on the concrete floor, and Figure A.5

shows another close-up view at the MPC connection to the MPC baseplate along one plane of symmetry.

The fuel basket plate on the plane of symmetry has been removed so that the fuel assemblies within the

basket compartments can be seen.

MPC[FCISLAB 100'DROP

Time s 0

Transfer Overpack

MPC with fuel

Concrete Floor

Concrete Wall

Figure A.3. Transfer Cask Vertical Drop onto Concrete Floor/Wall Finite Element Model

A-7
MPCITCISLAB 1 O'DROP

Time 0

Top Flange

Upper Trunnion

Outer ShellILeadIlnner Shell

Water JacketlWater

Lower Trunnion

Concrete Floor

MPC Top Plate

Fuel Basket

\ Transfer LidlDoor

Figure A.4. Transfer Cask Vertical Drop onto Concrete Floor/Wall.

Close-Up View of the Cask Resting on the Concrete Floor.

A-8
Figure A.5. Transfer Cask Vertical Drop onto Concrete Floor/Wall

Showing MPC Connection to the MPC Baseplate at Plane of Symmetry (Time = 0)

The TO model used in this analytical study is the 125 ton HI-TRAC Model with the MPC-68 canister

(full of fuel) and bottom transfer lid, as received from HOLTEC International. The major revisions made

are as follows (A complete listing of changes can be found in Reference A.24.):

(1) To minimize the significant computational effort, a quarter model representation of the TO,

MPC, and concrete floor/wall beneath the floor was developed. This quarter model was created

considering the two vertical planes of symmetry that exist for the TO and the MPC. The TO

quarter model was developed from the HOLTEC half model by deleting half of the nodes and

element definitions and creating a node set to define the two planes of symmetry.

(2) Contact surfaces within various components of the TO and between the TO and the concrete

floor were defined to properly transfer impact forces.

(3) The HOLTEC TO model represents the combined MPC/fuel region modeled as brick type finite

elements. This single brick representation of the MPC/fuel was deleted and replaced by a

detailed finite element model of the MPC and internal components, as described later. (Removal

of the single brick representation of the fuel and basket and replacing it with an explicit

representation of the fuel and basket is not necessary for the analysis of the MPC. This was only

done in the event that the model would be used to predict fuel cladding strains.)

A-9
The dynamic impact analyses were performed using the LS-DYNA computer code, Version 970,

Rev. 3858 (Reference A.4). LS-DYNA is a general-purpose finite element code for analyzing the large

deformation dynamic response of structures. The computer code utilizes an explicit solution method to

solve for the dynamic impact response. Strains in all elements are calculated at each time step, and if the

effective plastic strain (EPS) reaches the user specified failure strain for a particular element, that

element is automatically eliminated from the model and the analysis continues.

The material model for the Type 304 stainless steel components in the TO and MPC are based on a

bilinear elastic plastic stress strain curve, which requires the elastic modulus, yield strength, tangent

modulus, and failure strain as input. The modulus of elasticity was taken from the ASME Code, Division 1 -

Appendices (Reference A.5). The material properties for the yield strength and ultimate strength were

based on the data contained in Appendix B, and then adjusted for strain rate effects using the information

in Reference A.6. The engineering strain at failure was based on the data contained in Reference A.6 at a

high strain rate. The tangent modulus was calculated from the yield strength, ultimate strength and

engineering failure strain. All of these material properties were developed at the appropriate temperature

for each component of the TO and MPC.

The design temperatures for the various TO and MPC components are presented in Table A. 1, and are

consistent with HOLTEC's design (Reference A. 1). The material properties (modulus of elasticity, yield

strength, tangent modulus, and rupture strain) at the temperatures of the TO and MPC components are

presented in Table A.2.

Table A.1. Transfer Overpack, MPC, and Fuel Cladding Temperatures

Components Temperature IC (IF)

Fuel and fuel basket 393 (740)

MPC Shell 237 (459)

Transfer overpack inner shell and water jacket inner 162 (323)

surface

Water jacket outer surface 106 (223)

Top lid, transfer lid, trunnion, water jacket ribs, water,

and transfer lid vertical plates 162 (323)

A-10
Table A.2. Material Properties

Temp Modulus of Yield Tangent Engineering Minimum

°C (

0F) Elasticity Strength Modulus Strain at True Strain at

(psi) (psi) (psi) Failure Failure (1)

Material (in/in) (in/in)

393 (740) 2.35x107 23386 1.03x10 0.377 0.92

Stainless 237 (459) 2.61x107 38344 5.85x104 0.456 0.92

Steel 162 (323) 2.69x107 44719 4.60x 104 0.48 0.92

(Type 304)

106 (223) 2.75x10' 51012 4.03x104 0.498 0.92

Zircaloy-2 393(740) 9.80x106 43807 1.41x10 NA NA

Lead 162 (323) 1.90X106 380 0 NA NA

1 psi = 6.895 kPa

(1) Includes the effects of strain rate and temperature up to 100/sec and 1 OOOF respectively.

Although stainless steel material Types 304LN, 316, and 316LN might also be used to construct the TO,

MPC and other stainless steel components, the material properties (yield and ultimate strength) for

stainless steel Type 304 are equal to or less (i.e., more conservative) than the values for Type 304LN,

316, and 316LN for the temperature range of interest and, therefore, were used in the analysis.

The material model used in the analysis assumes that the stress-strain curve and failure criterion (rupture

strain), which are derived from tensile test data, are also applicable to elements in compression. Using

the same stress-strain curve in tension and compression is realistic for steel materials.

Since the LS-DYNA computer program considers the reduction in the area of the finite elements in the

computation of stress during the execution of the analysis, a true stress strain curve is expected to be

input for the material properties. In the analyses, however, the more conservative engineering stress

strain curve was used. It is well known that engineering stress strain values are very close to the true

stress strain values during the initial stages of yielding, and that in the plastic region the engineering

stress values are always lower than the true stress values. Thus, at any given deformation level the total

energy absorbed by the MPC will be underestimated. This results in a conservative response because it

overestimates the total deformation (strain). For comparison, the minimum true strain at failure of Type

304 stainless steel is also shown in Table A.2.

The MPC consists of the 1.27-cm (1⁄2-inch) thick vertical cylindrical shell, a 25.4-cm (10-inch) thick top

shield plate, and a 6.35-cm (2.5-inch) thick baseplate. The top plate and baseplate are modeled using

brick elements. The MPC cylindrical shell is modeled using six brick elements through the thickness

from the top of the baseplate to 30 cm (11.75 inches) above the baseplate. Beyond this elevation, the

vertical shell wall is modeled using shell elements. Shell elements are sufficiently accurate for the region

above 30 cm (11.75 inches), and this transition to shell type elements reduces the execution time for

solution. Even with this transition to shell type elements,. the entire model is quite large with 79,973

nodes and 66,704 elements.

A-1I
The fuel basket is modeled as a grid of 0.635-cm (1⁄4-inch) thick vertical plates spaced 16.5 cm (6.5

inches) apart to form the square cells that hold the fuel assemblies. The fuel basket is freestanding on top

of the MPC baseplate. The LS-DYNA model of the fuel basket is shown in Figures A.4 and A.5. In

Figure A.5, the basket plates at the plane of symmetry have been removed to show the fuel assemblies

within the basket cell walls. The mass density of the basket plates was increased to account for the

additional boral material attached to the sides of the basket plates. The material model used for the TO

and MPC was also utilized for the stainless steel fuel basket plates. The specific material properties used

for the fuel basket are shown in Tables A. 1 and A.2.

To maintain a reasonable execution time, a single equivalent beam model of a typical fuel assembly was

developed and placed inside each of the basket cell compartments. Each fuel assembly model consisted

of a fuel support region (at the bottom of the fuel assembly), fuel region, and a handle (at the top of the

fuel assembly). To capture the interaction between the fuel assemblies and the MPC baseplate, it was

necessary for the fuel beam to have plastic capability. To incorporate plasticity in a beam element,

LS-DYNA requires that the beam be defined with explicit dimensions. To obtain the correct assembly

stiffness and nonlinear behavior, the fuel region was modeled as a single circular tube with

cross-sectional dimensions equal to an individual fuel rod, and the yield stress, modulus of elasticity,

tangent modulus, and mass density were multiplied by the total number of fuel rods in the assembly.

In this analysis, the primary purpose of the fuel assembly model is to capture the overall effects of the

fuel assembly on the structural response of the MPC in a vertical drop. Therefore, the most important

parameters are the mass, vertical stiffness, and energy absorption capability of the fuel assemblies. An

assumption made in the fuel assembly model is that the mass of the fuel is added to the mass of the fuel

cladding, but only the fuel cladding provides stiffness. This assumption is consistent with the

methodology utilized by HOLTEC in its design (Reference A.7), and complies with References A.25 and

A.26.

The same material model and material properties (elastic modulus, yield strength, tangent modulus, and

rupture strain) used for the fuel basket described above, were also utilized for the stainless steel fuel

support (beneath the fuel region) and the handle (above the fuel region). The material properties used for

the fuel region were based on the properties of Zircaloy-2 cladding material at a temperature of 483 'C

(902 °F) (References A.8 and A.9). Table A.2 presents the various material properties used for the fuel

assembly members.

Based on calculations performed by HOLTEC (Reference A.7), the fuel region was expected to buckle

elastically. However, LS-DYNA does not check for classical bifurcation buckling directly; but it can

achieve an accurate prediction of static buckling by introducing a tiny bias (initial deflection) in the

middle of the fuel rod. Therefore, a small bias was introduced in the middle of each equivalent fuel beam

(between spacer grids) to allow large lateral deformation to take place.

Also included in the fuel assembly model are nine spacer plates (grids), at the elevations specified in the

fuel assembly design drawings. The spacers, modeled with shell elements, are constrained to move

vertically and horizontally with the equivalent fuel beams. A small gap is modeled between these spacer

plates and the basket cell walls. Once the gap is closed, the basket cell walls provide lateral restraint for

the spacer plates, which provide lateral restraint to the fuel rods. The fuel assemblies, like the fuel

basket, are also freestanding on top of the MPC baseplate.

A-12
A finite element model of the concrete floor and wall beneath the floor Was developed. The model of the

concrete wall is 12.5 meters (41 feet) high from the bottom of the floor, 3.96 meters (13 feet) long and 38

cm (15 inches) thick. The concrete floor is 3.96 meters (13 feet) wide, 3.96 meters (13 feet) long and 61

cm (24 inches) thick. This concrete floor/wall model is a quarter model representation of the actual

structure and is consistent with the quarter models of the TO and MPC. Although the actual floor and

wall do not have two well-defined planes of symmetry, this representation is considered reasonable.

The concrete floor/wall model was developed with 10 separate "parts" using the same concrete material

model, each of which can be adjusted independently when needed for hourglass control or to change

element formulation types. A finer mesh was developed near the impact region, while a coarser mesh

was created away from this critical region. The concrete floor/wall has fixed boundary conditions at the

ends of the wall and at the edges of the floor, excluding the two planes of symmetry.

The concrete material model and the material properties utilized in the analyses are those used by the

cask manufacturer in their analyses for design certification and are presented in Appendix C to Reference

A. 11. A comparison analysis using an alternate concrete material model was also performed to

determine whether LS-DYNA Material Model Type 16, which was used in the analysis, is reasonable

and comparable to other more recently developed material models. The results indicate that the

alternative material model gives comparable results to Material Model Type 16.

The transfer cask rests on top of the concrete floor at the start of the analysis, and the drop height was

simulated by introducing an initial velocity corresponding to the desired drop height. The initial velocity

was calculated using the following expression:

Velocity = 2gh , where g is the acceleration of gravity and h is the drop height.

Four different drop heights at 1.5, 12.2, 21.3, and 30.5 meters (5, 40, 70, and 100 feet) were considered in

the vertical drop analyses.

The design of the TO specified a small gap between the MPC shell wall and the inner shell of the TO. If

after the fall the MPC lands at the center of the bottom lid of the TO, there would be a uniform gap equal

to 0.5 cm (0.203 inches) based on average construction tolerances. Under this configuration, when the

MPC shell deforms radially by an amount greater than the uniform radial gap, the TO provides resistance

to further radial deformation. For the analysis reported herein, a worst-case assumption was made that

during the accidental drop of the MPC/TO, the MPC may land completely to one side of the TO. This

creates a maximum gap on the opposite side equal to approximately 1.03 cm (0.406 inches) which is

twice the average. This off-center drop location is simulated in the analysis by circumferentially

enlarging the gap between MPC shell and TO inner shell to 1.03 cm (0.406 inches). In this manner, the

analysis can provide an estimate of the effect of an off-center drop, while keeping the quarter model

representation of the actual casks.

A-13
Analysis Strategy and Modifications to the Finite Element Model

The strategy behind the development of a finite element model and subsequent analysis must reflect the

intended use of the results. In this case the results will ultimately be used to predict the probability of

failure of the MPC confinement boundary. As such, the results from the finite element analysis must

support the methodology used to predict material failure in the MPC shell and weldments.

Following receipt of the report and models (Reference A.24), the engineering failure strain limit was

eliminated which had been set for the deletion of "failed" elements. This was done as a result of the a

revision of the weld failure methodology in Appendix B so that the actual true strain, which is the result

that is directly linked to the probability of weld failure in Appendix B, could be continuously computed

until the end of the event. By running the models this also allowed to calculate the Triaxiality Factor at

the location of maximum strain. As explained in Appendix B, the Triaxiality Factor is an important

parameter for predicting failure, because it accounts for loss of material ductility in regions where a

triaxial state of stress may constrain plastic flow.

As part of accepting the drop analysis report and input files, shakedown evaluations of the model were

performed. Based on those results the following modifications were made:

(1) All analyses were run using double-precision.

(2) The basket supports' attachment to the MPC shell was modified to more accurately reflect the

Holtec drawings.

(3) The MPC baseplate mesh was slightly modified at the junction with the shell to remove a local

hard spot in the mesh.

These modifications were included in all analyses of the drops of the transfer cask onto the concrete floor

and the 5.8 meters (19 foot) drop of the MPC into the storage overpack. They were not included in the

analysis of the drops of the transfer cask onto the storage overpack because the strains were relatively

low for these drops.

Analysis Results

The drop heights selected for analysis were 1.5, 12.2, 21.3 and 30.5 meters (5, 40, 70, and 100 feet),

which provide results at the maximum drop height and at intermediate levels. The structural response of

the MPC for the vertical end drop of the transfer cask onto the concrete floor are summarized in Table

A.3. For each drop height, the maximum EPS in the MPC shell, the element producing the maximum

EPS and the true strain at failure are tabulated. Only the 30.5-meter (100-foot) drop is discussed in

detail.

A-14
Table A.3. Structural Response of the MPC for Various Drop Heights

Maximum EPS Element Number with Minimum True Strain

Drop Height cm/cm (in/in) Maximum EPS at Failure

m (ft) cm/cm (in/in)

1.52 (5) 0.024 16539 0.92

12.2 (40) 0.195 15973 0.92

21.3 (70) 0.24 15997 0.92

30.5 (100) 0.256 15997 0.92

Figures A.6 and A.7 show the deformation of the MPC shell and basket s .pports for the 30.5-meter

(100-foot) drop case at the end of the event (0.022 seconds) when the kinetic energy has reached a

minimum and the transfer overpack and MPC begin rebounding from the impact. These figures show the

response of the structural components at the connection of the MPC cylindrical shell to the baseplate

adjacent to the plane of symmetry through the center of the concrete wall. They clearly show that the

deformation is dominated by buckling in the lower region of the shell. The figures also show that the

MPC shell wall deforms primarily in a vertical downward direction, and radially outward near the

baseplate. In addition, Figure A.7 shows that there is some rotation of the MPC baseplate at the junction

with the shell wall. This helps relieve some of the strain in the shell at this intersection location. Figure

A'7 also shows the separation between the basket support and MPC shell in the region between the stitch

welds. (Note that the actual location of the one inch stitch weld is on the far side of the support since the

centerline of the support is on the plane of symmetry.) The figure shows that the shell meridional

curvature changes several times in this region. This is primarily attributable to the constraint of radial

displacement caused by the radial contact of the MPC shell with the inner shell of the transfer cask.

A-15
Figure A.6. Deformation of the MPC Shell and Basket Supports at 0.022 seconds

for the 100-foot Drop of the Transfer Cask onto the Concrete Floor

A-16
Figure A.7. Deformation of the MPC Shell and Basket Supports at 0.022 seconds

for the 100-foot Drop of the Transfer Cask onto the Concrete Floor

Figure A.8 is a plan view of a horizontal section through the MPC shell showing the location of the

basket supports. Figure A.9 is an elevation view of the MPC shell, baseplate and basket supports in their

undeformed position, and Figure A. 10 is the same view at the time of maximum response (0.022

seconds). Figure A. 11 shows contours of EPS at the time of maximum response, and when viewed

together with Figure A. 10 shows that the maximum EPS occurs in the shell at the bottom of the supports.

Figure A.12 shows the EPS contours along the back of the shell. Both Figures A. 11 and A.12 show that

the maximum strains occur in the MPC shell region between the bottom of the basket supports and the

baseplate. Away from the bottom region of the MPC shell, strains diminish rapidly to very small values.

A-17
Figure A.8. Plan view of a horizontal section through the MPC shell

showing the location of the basket supports

A-18
Figure A.9. Elevation View of the MPC Shell, Baseplate, and Basket Supports

In Their Undeformed Positions

A-19
Figure A.10. Elevation View of the MPC Shell, Baseplate, and Basket Supports at 0.022 seconds

(100-Foot Drop)

A-20
Figure A.11. Contours of EPS on the Interior of the MPC Shell

at the Time of Maximum Response (0.022 seconds) (100-Foot Drop)

A-21
Figure A.12. Contours of EPS on the Exterior of the MPC Shell

at the Time of Maximum Response (0.022 seconds) (100-Foot Drop)

Figure A.13 is a close up view of Figure A.1 1 showing the location of the element with the maximum

EPS (Element 15997) and Element 9421 at the junction of the shell and baseplate on the plane of

symmetry. Figure A. 14 is a vertical section through the MPC shell at the location of Element 15997 and

the maximum EPS. Element 15997 is on the inside surface of the shell and Element 15577, also shown

in the figure, is on the outside surface at the same elevation. Figure A. 15 shows the time history of EPS

in each element through the thickness at this location. The figure shows how rapidly the plastic strains

diminish from a maximum of 0.256 in/in on the inner surface to a minimum of 0.040 in/in at the outer

surface. Figure A. 16 shows the time history of EPS in each element through the thickness between

Elements 9445 and 9025 at the bottom of the shell where it joins the baseplate on the same vertical

section as shown in Figure A. 14. Figure A. 16 shows that the EPS in Element 9025, which is at the

location of the MPC shell to baseplate weld, is 0.220 in/in, which is close to the maximum EPS.

A-22
Figure A.13. Close-up View of Figure A.11 Showing the Location of the Element

With the Maximum EPS (Element 15997) and Element 9421 at the Junction of the Shell

And Baseplate on the Plane of Symmetry (100-Foot Drop)

A-23
Figure A.14. Vertical Section through the MPC Shell at the Location of Element 15997

and the Maximum EPS (100-Foot Drop)

A-24
MPC/TC/SLAB 100'DROP 30106

0. 3

-, t

0.:

,~ a IU,

U

.~0. Cu

0.

w

LU

25-

F

.2-

15-

D D D

= aR

SA~ AElement no.

_A_15577

-B-15661

Q_15745

JOL15829

-E-15913

F 15997

0.!

1 1 -1- 1- -~

0 0.005 0.01 0.015 0.02 0.025

Time

Figure A.15. Time History of EPS in Each Element Through the Thickness

At the Location of the Maximum EPS (100-Foot Drop)

A-25
MPC/TC/SLAB 100'DROP 30106

Element no.

-A-9025

0..9109

.f,9193

J.fl9277

-.E-9361

F 9445

.S

w

Time

Figure A.16. Time History of EPS in Each Element Through the Shell

Thickness At the Location of the MPC Baseplate on the Vertical

Section Shown in Figure A.14 (100-Foot Drop)

In Table A.3, it is tempting to compare the maximum EPS in the element to the true strain at failure.

Such a comparison, however, is not an accurate predictor of margin against element failure because these

two strain measures, (i.e., calculated maximum EPS and true strain at failure), are each derived from

different states of stress and must be adjusted by a Triaxiality Factor before a meaningful comparison can

be made. The derivation of the Triaxiality Factor and its application to predicting material failure in the

MPC is discussed in Appendix B.

A.4.4 Question 4

Analyze the vertical drop of the transfer cask containing a sealed MPC falling onto the storage overpack

at ground elevation from heights of 1.5, 12.2, and 24.4 meters (5, 40, and 80 feet). Determine the

magnitude and location of maximum stresses/strains in the MPC.

To respond to this question, dynamic analyses using LS-DYNA were performed for the accidental

vertical drop of the transfer cask onto the storage overpack, while the transfer cask is being lowered from

the operating floor to the ground floor. The maximum drop height corresponds to the distance between

the operating floor and the top of the storage overpack, which could be up to 24.4 meters (80 feet).

Several drop heights up to the maximum height were analyzed.

A-26
The finite element model of the transfer cask (transfer overpack, MPC, and internals), utilized for this

drop case, is the same model used in the response to Question 3 except for the MPC, which did not

include the modifications stated above. A finite element model of the storage overpack was also

developed and placed beneath the transfer cask model. A quarter model representation of the storage

overpack was created to be consistent with the quarter model transfer cask. The storage overpack model

consists of a 68 cm (26.8 inches) thick concrete cylinder with steel plates on the top, inside, and outside

surfaces of the concrete wall. The concrete and steel plates are modeled using brick type finite elements.

The material model and properties for the concrete cylinder are the same as those described in the

response to Question 3. The material model and properties for the steel plates are taken at room

temperature. The bottom of the storage overpack is restrained in the model. This is slightly

conservative, since it is actually resting on the concrete floor that is supported by a concrete wall beneath

the floor.

The drop heights selected for analysis were 1.52, 12.2, and 24.4 meters (5, 40, and 80 feet). The MPC

stainless steel material properties correspond to a temperature of 237 'C (459 'F), which is the

temperature used by the manufacturer for the design of the MPC. The structural response of the MPC is

presented in Table A.4.

Table A.4. Structural Response of MPC

Drop Height m (ft) 'Max. EPS at Element 9421 (Azimuth = 0 ) cm/cm (in/in)

1.52 (5) 0.0363 2

12.2 (40) 0.060 1 2

24.4 (80) 0.0976 2

EPS is the effective plastic strain; Element 9421 is the first MPC vertical brick element at the plane of

symmetry. The drop of the transfer cask onto the storage overpack is a softer impact than the drop of the transfer cask

onto the concrete floor. Therefore, the resulting strain in the MPC for this drop is less at the higher (12.2

and 24.4 meters ) drop heights than for the drop onto the concrete floor. The result for the 1.52 meters drop

is larger than the result for the 1.52 meters drop onto the concrete floor because (a) the difference in target

stiffness has less effect at lower drop heights, and (b) the modifications to the model discussed in response

to Question 3 were not made to this model because it produced very conservative results that were bounded

by the 12.2, 21.3, and 30.5 meters drops onto the concrete floor.

A.4.5 Question 5

Determine the magnitude and location of maximum stresses/strains in the MPC when a sealed MPC

vertically falls 5.8 meters (19 feet) into the storage overpack.

Such an event could occur during the transfer of the MPC from the transfer overpack to the storage

overpack.

When the MPC is moved from the transfer overpack to the storage overpack (Stage 21), it could possibly

fall a maximum of 5.8 meters (19 feet). This load drop is analyzed using a finite element model.

A-27
Model Development

The finite element model includes the MPC, storage overpack inner shell, storage overpack vertical

channel sections, and the storage overpack bottom lid (which serves as the target plate). The analysis

strategy and the finite element models of the MPC, fuel assemblies, fuel basket and fuel basket support

are the same as those used in response to Question 3. These components are shown in Figures A. 17 and

A.18. The FEM shown in Figure A.17 is a quarter-model representation of the cask system (due to

symmetry about two vertical planes). Figure A. 18 shows a close-up view at the MPC connection to the

MPC bottom baseplate at the plane of symmetry. The fuel basket plate on the plane of symmetry has

been removed, so that the fuel assemblies within the basket compartments can be seen.

03-29-04 QUARTER MPC MODEL

Time - 0

MPC Shell

MPC Baseplate

Rigid Target Plate (Top of

Storage Overpack Bottom Lid)

MPC Top Plate

Fuel Basket

Storage Overpack

Inner Shell

x

Z'-ý V

Figure A.17. Finite Element Model of the MPC for 19-Foot Drop into Storage Overpack

A-28
03-29-04 QUARTER MPC MODEL Thme= 0

MPC (Shell

Element Region)

Fuel Assembly

Fuel Basket

B a s k e t S u p p o r t s W e l d e d -' ---- .-- -- -

to Inside of MPC

Vertical Channels Welded

to Storage Overpack Inner

Shell (Half Channel Shown

at Plane of Symmetry)

Storage Overpack

Inner Shell

MPC Shell (Brick

Element Region)

MPC Baseplate

x

j Rigid Target Plate (Top

of Storage Overpack

Bottom Lid)

Figure A.18. MPC 19-Foot Drop into Storage Overpack.

View of Model at MPC Shell Wall Connection to Baseplate at Plane of Symmetry

The top plate of the storage overpack bottom lid is modeled as a rigid target plate for the vertical impact

of the MPC and its internal components. The target plate is fully restrained in the model for the

following reasons:

The storage overpack bottom lid is considered to be rigid since it is constructed from a 12.7-cm

(5-inch) top plate, a 5.1-cm (2-inch) bottom plate, and 43.2 cm (17 inches) of confined concrete

between the plates.

The storage overpack rests on a reinforced concrete floor supported by a reinforced concrete wall

beneath the floor. The assumption of a rigid target plate for this drop scenario is conservative.

The storage overpack vertical wall is modeled using a rigid shell to represent the storage overpack inner

steel shell, which surrounds the MPC shell wall. The vertical channels, which are welded to the storage

overpack inner shell, are included in the model using shell elements. The gap between the channel webs

and the MPC shell wall is included in developing the model.

A-29
For this drop scenario, a dynamic analysis was performed using LS-DYNA. The MPC and internal

components rest on the target plate at the start of the analysis, and the drop height was simulated by

introducing an initial velocity that was calculated using the following expression:

V= ý2gh (A.7)

where: V velocity

g - gravitational acceleration

h drop height.

To permit the insertion of the MPC into the storage overpack, the cask design specified a gap between

the MPC shell wall and the inner radius of the vertical channel sections welded to the storage overpack

inner shell. If the MPC impacts the bottom lid of the storage overpack in the center, there would be a

uniform gap equal to 1.67 cm (0.657 inches) based on average construction tolerances or a uniform gap

of 2.18 cm (0.859 inches) based on maximum construction tolerances. Under this configuration, When

the MPC shell deforms radially by an amount greater than the uniform radial gap, the channel sections

attached to the storage overpack provide resistance to further radial deformation. For the analysis

reported herein, a worst case assumption was made that during the accidental drop of the MPC, it may

land completely to one side of the storage overpack bottom lid. This creates a maximum gap on the

opposite side equal to approximately 3.34 cm (1.313 inches) based on average construction tolerances.

This off-center drop location is simulated in the analysis by removing from the computer model the

contact definition between the MPC shell and the storage overpack channel sections. This permits the

MPC shell to deform radially outward without the resistance provided by the channel sections. In this

way, the analysis can provide an estimate of the effect of an off-center drop.

Analysis Results

Figure A. 19 shows the displacement of the MPC shell and basket support on the xz plane of symmetry at

the time of maximum displacement (0.005 seconds). Figures A.20 and A.21 are plots of radial (z) and

vertical (x) displacements at the three nodes located on the MPC shell in Figure A. 19. The maximum

radial displacement is less than the maximum gap of 3.34 cm (1.313 inches) and, therefore, no contact

occurs between the MPC shell and the channel sections attached to the inside of the storage overpack.

The curvature of the shell and basket support, as well as the contact and separation between the shell and

support can be clearly seen. In contrast to Figure A.7, where the radial displacement was limited by

contact with the transfer overpack, Figure A. 19 shows a smoother variation in curvature because radial

displacement is allowed to occur freely.

A-30
Figure A.19. Displacement of the MPC Shell and Basket Support

Along the xz Plane of Symmetry at the Time of Maximum Displacement (0.005 seconds)

A-31
Node no.

._A42321

._._23281

C 9001

0 0.01 0.02 0.03

Time

Figures A.20. Plots of Radial (z) Displacement at the Three Nodes

Located on the MfPC Shell in Figure A.19

A-32
0.4

Node no.

-A-42321

B1-23281

_C_.9001

0.2

2

-0-I

-0.I

0 0.01 0.02 0.03

Time

Figures A.21. Plots of Vertical (x) Displacement at the Three Nodes

Located on the MPC Shell in Figure A.19

Figure A.22 shows the contours of EPS in the MPC shell, basket supports and baseplate at the time of

maximum response, which is the time in Figure A.23 when the kinetic energy of the system is a

minimum. Figure A.22 shows that the EPS is 0.459 in/in and occurs in Element 9035. With the basket

supports removed, Figures A.24 and A.25 show the location of Element 9035 where the maximum EPS

occurs. This element is located at the bottom of the shell on the outside surface within the weld joining

the shell to the baseplate. Figure A.26 is the same view as Figure A.25 except the basket support is

included and the element mesh is defined.

A-33
Figturs fEfie Plasti Co torsafi Sinh P Sel aktSpors n aelt

Mxptvatlte Tieo4aiu epne hc steTm nFgr .593e0

When the Kinti Energy Is88Miimu

A-34
03-03-06 MPC 19' DROP 6

Global

A1

LUJ

LU

C

0

Time

Figure A.23. Global System Kinetic Energy

A-35
03-03-06 MPC 19' DROP

finle - 0.023 Fringe Levels

Contours of Effective Plastic Strain

rnax ipt. value 4.593e-01

min ýý 0, at elemg 38803

rnax=(J.4b9303, at clenill 9035 4.1 34e-01

3.674e-01

3.21 5e-01

2,756c-O 1

2.297c 01

1.837e-Ol

1.3/8e-01

9. 1 8be-02

4.593c 02

0.000e+00 H 9035

X

y

Figure A.24. With the Basket Supports Removed from Figure A22,

Figure A.24 Shows the Location of Element 9035 Where the Maximum EPS of 0.459 Occurs

A-36
Figure A.25. Closeup View of the Location of Element 9035

On the Bottom Outside Edge of the MPC Shell

A-37
Figure A.26. Same View as Figure A.25

Except the Basket Support Is Included and the Element Mesh Is Defined

Figure A.27 is a vertical section through the shell and baseplate at the location of Element 9035. The

figure shows the relative displacement between the shell and basket support. Because the stiffness of the

basket support at this location is so much greater than the basket support on the plane of symmetry

(Figure A. 19) the shell does not separate as much from the basket support between the stitch welds.

Figure A.28 is the displacement history of node 20340 shown in Figure A.27, and shows that the

maximum radial displacement is about the same as it is on the plane of symmetry.

A-38
4

4 ____________________________4

4

4

4-

4-

4-

4,

4-

Figure A.27. Vertical Section Through the Shell and Baseplate at the Location of Element 9035

Showing the Relative Displacement Between the Shell and Basket Support

A-39
03-03-06 MPC 19' DROP

1.51

Node no.

_AX-displacement

_BY-di splacement-20340

__CZ-di splacement-20340

LDResultant Displacement-20340

1

0.5

0

"z 0

-0.5

-1

0 0.01 0.02 0.03

Time

Figure A.28. Displacement History of Node 20340 Shown in Figure A.27

Figure A.29 is a vertical slice through the MPC shell and baseplate showing the location of Element 9035

and the local distribution of EPS within the shell. Figure A.30 labels the elements through the shell

thickness at the junction of the shell and baseplate, and Figure A.31 plots the time history of EPS within

each element. These plots show the significant decrease in the maximum EPS away from the outside

surface of the shell. The 5.8 meters (19 feet) drop of the MPC into the storage overpack produces the

highest EPS (0.459 in/in) of any of the cask and MPC drop events. This event, and the probability of

breaching the MPC confinement boundary are discussed in greater detail in Appendix B.

A-40
03-03-06 MPC 19' DROF

Time = 0.023

Contours of Effective PK

max Ipt, value

min= , at elem# 38803

max=0.459303, at elen

Fringe Levels

4.593e-Ol

4.1 34c-01_

3.674e-01

3.215e-01

2.756e-01

2.297e-01

1.837e-01

1.378e-01

9.186e-02

4.593e-02

Figure A.29. Slice Through the MPC Shell and Baseplate

Showing the Location of Element 9035 and the Local Distribution of EPS Within the Shell

A-41
Figure A.30. Labels the Elements Through the Shell Thickness at the Junction of the Shell

And Baseplate at the Location of Maximum EPS

A-42
03-03-06 MPC 19' DROP

Ug.l

0-4

~0-3

wi0-1

FA A

_J _ F

Element no.

_A_9035

JBL9119

_C9203

J.9287

-F-9371 P F 9455 C r __ __ _ __ _ .

F F

"I P n I I 11 " )gz fl F..4-' t I, t II1f

0 0

0 0.01 0.02 0.03

Time

Figure A.31. Time History of EPS Within Each Element Through the Thickness

as Shown in Figure A.30

A.4.6 Question 6

If the storage cask drops onto the ISFSI concrete pad, asphalt area, or gravel area, will it slide and/or tipover? If tip-over can occur, what velocity of the crawler would be required to tip it over? How much

time does the operator have to stop the crawler before tip-over occurs? What are the stresses in the

MPC?

The storage cask is hung from the top while being transported by the crawler. To perform this

calculation a few assumptions were made based on photos and a drawing of the crawler. The height of

the storage cask above the ground surface before the drop occurs is assumed to be 30.5 cm (12 inches).

The crawler is traveling in the direction of the open end of the crawler. The crawler horizontal square

tube bumper support members that make contact with the storage cask are 1.85 m (73 inches) off the

ground based on a telephone conversation with Lift Systems (crawler manufacturer).

A-43
Two bounding cases were evaluated for a drop of the storage cask onto the ISFSI concrete pad:

Case 1. vertical drop without any rotation ("end drop")

Case 2. drop causing maximum possible rotation

For Case 1, the calculation showed that the frictional resisting force of the storage cask sliding against

the concrete was much less than the tip-over force and, thus, the storage cask would slide and not tip-over

regardless of the velocity of the crawler. The maximum acceleration of the storage cask for an "end

drop" case was calculated to be 44.13 g by Holtec (Reference A. 1) assuming a drop of 27.9 cm (11

inches). This acceleration was increased by a factor of 1.045 to account for the drop height of 30.5 cm

(12 inches) being considered. This load factor was calculated using the Law of Conservation of Energy

which equated the kinetic energy at impact to the work done in compressing the concrete pad and soil.

Applying this load factor, the acceleration for the drop of 30.5 cm (12 inches) was calculated to be 46.1

g. To obtain stresses for this end drop case use the stress results presented in the calculations for

Question 10 (end drop case) and scale the results by the ratio 46.1 g / 60 g (= 0.768 factor).

For Case 2, the calculation demonstrated that it is geometrically impossible for the rotation of the

dropping storage cask to be large enough to cause a tip-over. The maximum amount of rotation possible

is 5.140 (with respect to the ground) which occurs when one of the two vertical supports holding the

storage cask at the top breaks and the cask makes contact with the ground at one edge. Since this angle is

less than the "tip-over" angle of 29.5' (angle which would cause the storage cask to tip-over), the storage

cask will not tip-over. Furthermore, the force required to tip-over the storage cask from the rotated

orientation was calculated to be greater than the frictional resisting force. Therefore, the storage cask

would slide but not tip-over. The acceleration and stresses under case 1 would also apply to this case.

In both cases, it is demonstrated that if the crawler has sufficient force, it would cause the storage cask to

slide (not tip-over). However, if one object is pushing against another object (with both objects having

the same coefficient of friction), the lighter object will slip or slide regardless of the magnitude of the

coefficient of friction. Since the crawler weight is 71,440 kg (157,500 lb) and the storage overpack with

MPC and fuel is 163,290 kg (360,000 lb), the crawler "tank treads" will slip and the storage cask will not

slide or tip-over.

The drop of the storage cask onto an asphalt or gravel surface was also considered. The g values

corresponding to this drop and the resulting stresses would be bounded by the drop onto the ISFSI

concrete pad. This occurs because the asphalt and gravel have a lower strength and Young's modulus

than concrete and both of these surfaces do not have steel reinforcing bars. Regarding sliding or tip-over,

the same conclusion for the drop onto a concrete pad applies here that the crawler will slip and will not

be able to slide or tip-over the storage cask. This assumes that both the crawler and storage cask are on

the same surface (i.e., gravel or asphalt).

A. 4.6. 1 Refined Evaluations and Results for Question 6

Subsequently, refined calculations were performed for the 30.5-cm (12-inch) drop of the storage cask

onto the three surfaces (the asphalt, gravel, and ISFSI concrete pad). These refined nonlinear impact

analyses were performed on the computer code LS-DYNA (Reference A.2) using existing computer

models of the storage overpack obtained from Holtec International.

A-44
A. 4.6.2 Storage Cask Drop onto Reinforced Concrete Storage Pad

A 30.5-cm (12-inch) vertical end drop of the storage cask impacting a 61-cm (24-inch) thick reinforced

concrete pad, resting on soil was analyzed. Three analyses were performed corresponding to the lower

bound, best estimate, and upper bound elastic soil modulii, which are equal to 324 MPa, 641 MPa, and

1,579 MPa (47,000 psi, 93,000 psi, and 229,000 psi), respectively. Poison's ratio is equal to 0.4 and the

soil density is 2,003 kg/m3 (125 lb/ft3).

The maximum acceleration values for the 30.5-cm (12-inch) drop of the storage cask onto the 61-cm

(24-inches) concrete pad resting on soil are shown in Table A.5.

Table A.5. Storage Cask Drop A alyses

Impact Thickness of Elastic Soil Maximum

Surface Target Modulus* (psi) Acceleration (g)

Layer (in) Comment

24 47000 40.4 Lower bound soil property

Concrete 24 93000 41.2 Best estimate soil property

24 229000 44.5 Upper bound soil property

Best estimate asphalt property at 4.4 7C

(40 'F) and best estimate soil property

Best estimate asphalt property at

Asphalt 12 93000 23.2 23.9 'C (75 7F) and best estimate soil

property

Bounded by Best estimate asphalt property at

12 93000 23.9 °C 43.3 'C (110 17F) and best estimate soil

(75 'F Case) property

24 93000 21.9 Best estimate gravel and soil properties

Gravel 12 93000 15.8 Best estimate gravel and soil properties

12 93000 19.0 Upper bound gravel property and best

estimate soil property

* Elastic soil modulus was based on the shear modulus calculated from the shear wave velocity of the free field

The maximum acceleration at the bottom center of the storage cask is 41.2 g for the best estimate soil

property. The acceleration time history at this location is shown in Figure A.32. The acceleration time

history plot is initially flat at 1.0 g for a period of time because the storage cask was dropped from a

height of 30.5 cm (12 inches) above the concrete pad and then impacted the pad at about 0.248 seconds.

As shown in Table A.5, the variation in soil property does not have a significant effect on the maximum.

acceleration of the storage cask for the configuration and parameters defined for this load case.

A-45
CU

N

HISTORM END DROP H=12" V=o (DROP = 12") FC=3000 TP=24" ES=93

50-

elol

30-

,If

j' r" S' , 

10-

J'~ii

1 0 ...................................................... ...............

-10 ,,

Node Ids

._Ab2533

m

0.22

min=-2.6433

max=41.234

0.24 0.26

Time

0.28

Figure A.32. Acceleration Time History: 30.5-cm (1-Foot) Drop on Concrete Pad

Since the MPC shell was not discretely modeled, the stresses in the MPC were calculated using the

stresses from Table A.6 [Reference 9] for the end drop case and scaling the results in proportion to the

new calculated g values.

Using this approach and the appropriate dynamic load factor, the stresses in the MPC shell were

calculated to be 53.3 MPa (7,732 psi). This stress value is well below the elastic buckling stress and

yield for the MPC stainless steel material.

A.4.6.3 Storage Cask Drop onto Asphalt Surface

A 30.5-cm (12-inch) vertical end drop of the storage cask impacting a 30.5-cm (12-inch) thick layer of

asphalt resting on soil was analyzed. Because asphalt material properties are sensitive to temperature,

three loading cases were considered. The three cases correspond to material properties of asphalt at

4.4 'C, 23.9 'C, and 43.3 'C (40 'F, 75 'F, and 110 'F). This approximates the expected average

temperature at the subject facility corresponding to winter, spring/fall, and summer. The properties for

the asphalt layer were determined from a literature review of available information for a "good" asphalt

mix. The best estimate values for the elastic modulus of asphalt are 10,342 MPa (1,500,000 psi) at

4.4 °C (40 °F), 4,482 MPa (650,000 psi) at 23.9 'C (75 'F), and 1,379 MPa (200,000 psi) at 43.3 'C

(110 °F). At 4.4 -C (40 'F) the modulus of asphalt is approximately one-half the modulus of concrete.

A-46
Table A.6. MPC Stress Summary

Safety Factor

(SF)

Drop Ref. Max. Direction/

MPC Orien- A.1 Stress Type of Acceleratio

Case Component tation Value Table # (psi) Stress n (g) Notes

End Fuel Basket Vertical 3.591 3.4.3 -3739 longitudinal 60 (A)

Drop Fuel Basket Vertical N.A. (B)

Support

MPC Shell Vertical 1.21 3.4.4 -11260 longitudinal 60 (C)

MPC Lid Vertical 2.8 3.4.4 ± 2,723 bending 60 (D)

MPC Vertical 1.282 3.4.4 N.A. 60 (E)

Baseplate

Tip- Fuel Basket 45 Deg. 1.283 3.4.3 & ± 38,3354 PL + Pb 45 (F)

over 3.4.6

Fuel Basket 0 Deg. 1.185 3.4.4 & ± 15,1304 direct + 45 (G)

Support 3.4.6 bending

MPC Shell 0 Deg. 1.1 3.4.4 + 58,9594 circum- 45 (H)

- 45,2834 ferential

MPC Lid N.A. (I)

MPC N.A. (I)

Baseplate

Negative stress = compression

Positive stress = tension

PL = Local membrane stress intensity

Pb = Primary bending stress intensity

1. SF of 3.59 is for lateral loading (not longitudinal)

2. Corresponds to ± 84.8 MPa (12,299 psi)

3. Corresponds to MPC-24

4. For tip-over case, stresses must be multiplied by a dynamic load factor of 1.04 (Ref A. 1)

5. Based on ANSYS Finite Element Model analysis

Notes column:

(A) Reference A. 1, Sec. 3.4.4.3.1.3

(B) Basket support provides lateral not vertical restraint

(C) Reference A.21, App. 3.H Case 5 - Buckling

(D) Reference A.21, App. 3.E.8.2. 1.1

(E) MPC baseplate supported by storage overpack baseplate.

(F) Reference A.1, App. 3.T.32

(G) Reference A.1, App. 3.Y.3 for MPC-68

(H) Reference A.1, App. 3.T.34; includes stress due to D.W. & internal pressure [a of +47.1 MPa (+6,838 psi)]

(I) Bounded by End Drop Case

A-47
The maximum acceleration values for a 30.5-cm (12-inch) drop of the storage cask for the 4.4 'C and

23.9 'C (40 'F and 75 'F) are shown in Table A.5. The 43.3 'C (110 'F) drop case was not analyzed

because at this high temperature, the acceleration would be lower than the 23.9 'C (75 'F) case. The

maximum accelerations at the bottom center of the storage cask, for the best estimate asphalt property are

25.4 g and 23.2 g for the 4.4 'C and 23.9 'C (40 'F and 75 'F) cases, respectively.

The stresses in the MPC were calculated using the stresses from Table A.6 for end drop case and scaling

the results in proportion to the new calculated g values. Using this approach and the appropriate dynamic

load factor, the vertical stresses in the MPC shell were calculated to be 50.0 MPa (7,245 psi) and 45.6

MPa (6,618 psi) for the 4.4 'C and 23.9 'C (40 'F and 75 'F) cases, respectively. Both of these stress

values are well below the elastic buckling stress and yield for the MPC stainless steel material.

A.4.6.4 Storage Cask Drop onto Gravel Surface

Two cases were considered for the storage cask drop on a gravel surface. A-30.5 cm (12-inch) vertical

end drop of the cask impacting 30.5-cm (12-inch) and 61-cm (24-inch) thick gravel layers resting on soil

were evaluated. The 61-cm (24-inch) layer corresponds to the gravel region surrounding the concrete

pad where the storage casks are stored, while the 30.5-cm (12-inch) layer corresponds to the gravel

region away from the concrete pad where the gravel layer tapers to smaller thicknesses. The properties

for the gravel layer were determined from a literature review of available information on "well graded"

crushed stone aggregate. The best estimate value of 345 MPa (50,000 psi) for the elastic modulus would

tend to be somewhat higher than what might be expected at the subject facility. This would be

conservative because a higher modulus would result in higher impact forces and accelerations. To gauge

the effect of varying the modulus of the gravel material, an analysis was performed for the best estimate

value 345 MPa (50,000 psi) and for the upper bound estimate 689 MPa (100,000 psi) for the 30.5-cm

(12-inch) thick gravel layer case.

The maximum acceleration values for a 30.5-cm (12-inch) drop of the storage cask onto the 61-cm and

30.5-cm (24-inch and 12-inch) gravel layers are shown in Table A.5. The maximum accelerations at the

bottom center of the storage cask, for the best estimate gravel property are 21.9 g and 15.8 g for the 61-

cm and 30.5-cm (24-inch and 12-inch) gravel layer cases, respectively. As a measure of the sensitivity of

the gravel material property, an upper bound case for the gravel property was performed for the 30.5-cm

(12-inches) gravel layer. The maximum acceleration for this upper bound gravel property was 19.0 g.

The stresses in the MPC were calculated using the stresses from Table A.6 for end drop case and scaling

the results in proportion to the new calculated g values. Using this approach and the appropriate dynamic

load factor the vertical stresses in the MPC shell, for the 30.5-cm (12-inch) drop case, were calculated to

be 39.7 MPa (5,754 psi) and 28.6 MPa (4,151 psi) for the 61-cm (24-inch) and 30.5-cm (12-inch) gravel

layers, respectively. Both of these stress values are well below the elastic buckling stress and yield for

the MPC stainless steel material.

A-48
A.4.7 Question 7

If the crawler vehicle weighing 71,440 kg (157,500 lb) is traveling at 0.64 kmi/h (0.4 mi/h) while carrying

the storage cask weighing 163,293 kg (360,000 lb), hits another storage cask on a concrete pad, will this

impact slide or tip-over the "struck" cask?

(1) If it slides, will a sliding storage cask hit another storage cask? If it hits another storage cask,

what would be the maximum stresses (e.g., axial, circumferential, tensile, and compressive) in

the MPC?

(2) If it tips over, what would be the maximum stresses in the MPC?

(3) If it does not slide or tip-over, what travel velocity of the crawler would be necessary to strike the

storage cask and cause it to slide and/or tip-over?

The evaluation first determined whether a horizontal force would cause the storage cask to tip-over or

slide. For tip-over, a horizontal force of 149,732 kg (330,103 lb) applied at 1.85 meters (73 inches) off

the ground would be needed. This dimension off the ground corresponds to the height of the horizontal

square tube bumper supports on'the crawler. These supports show up on the photos and Lift System

drawings and this dimension was confirmed by Lift Systems (crawler manufacturer). The force needed

to overcome friction ranges from 40,823 kg (90,000 lb) for minimum coefficient of friction of 0.25 to

86,545 kg (190,800 lb) for maximum coefficient of friction of 0.53. For the range of coefficient of

friction used see Question 8. Since the force necessary to overcome friction is much lower than the force

needed to cause tip-over, the struck cask will slide and not tip-over.

Since the struck cask will slide, only Part (A) of Question 7 is applicable. The initial velocity of the

struck cask was calculated using the principle of Conservation of Momentum. This led to the initial

velocity of the struck cask:

mV

A(.8)

where: m = mass of crawler + mass of storage cask being transported

V - velocity of crawler

W weight of storage cask

When the crawler carrying the storage cask strikes another storage cask, the resulting velocity will be

reduced. The calculated resulting velocity x'o of the combined crawler, storage cask being transported,

and struck storage cask is 0.38 km/h (0.24 mi/h).

The equation of motion of the storage cask after impact is as follows:

M3x + /u W- Fc = 0 (A.9)

where: M 3 = mass of crawler + 2 storage casks

F = horizontal force that the crawler can generate, 139,603 kg (307,772 lb)1

Source: Lift Systems Spec. Sheet and confirmed by telephone discussion with C. Perkins of Lift Systems

A-49
The solutions for velocity and displacement are found by integrating Equation A. 10 and evaluating the

two constants of integration from the initial velocity of Equation A.9 and zero initial displacement.

Equations A. 1I and A. 12 for velocity and displacement are as follows:

x'= Fc-,uW t mV

X ~M3 )LT gm (A.10)

x=<c -J1W>t mV

W (A.1)

M+

The maximum displacement occurs when the velocity equals zero and is found to be as follows:

1 mV _

x (A.12)

M3 m+--

This equation tells us that as the value Fc approaches g W, the displacement goes to infinity. By

inserting the appropriate numbers, Fc is calculated to be greater than g W. Therefore, it is concluded that

the crawler will continually push the storage cask. The sliding storage cask will eventually hit another

storage cask or slide off the pad onto the surrounding area. This assumes that the crawler does not stop

but continues pushing the struck cask.

Part (A) of Question 7 also asks, what would be the maximum stresses in the MPC? Equation A. 10 can

be used to determine the acceleration of the storage cask. Solving forx" provides the acceleration as

follows:

x" = (F, - jt W) / M 3 = 0.25 g

Therefore, the maximum acceleration of the storage cask would be 0.25 g. The positive value of

acceleration indicates that the storage cask would accelerate under the constant force (not decelerate).

The maximum stresses in the MPC could be approximated by multiplying the stresses presented in the

calculations for Question 10 for the tip-over case by the ratio of accelerations 0.25g / 45g for the fuel

basket, basket supports, and the MPC shell. For the MPC lid and baseplate, the end drop case bounds

and the stresses would be multiplied by 0.25g / 60g. Note that there is also the dynamic load factor of

1.04 to be applied to the tip-over case (see calculations for Question 10).

A-50
A.4.8 Question 8

What seismic forces will cause the storage cask to slide and/or tip-over?

A.4.8. 1 Simplified Seismic Evaluation

Simplified analyses without soil structure interaction were performed for a storage cask (containing a

fully loaded MPC-68) subjected to earthquake forces. The storage cask was assumed to be resting on the

ISFSI concrete pad. The seismic analyses used the equivalent static load method which simply multiplies

the mass times the seismic equivalent static acceleration to arrive at the seismic forces. To determine

whether any amplification occurs as a result of the flexibility of the storage cask, the natural frequency of

the storage cask was calculated. The fundamental frequency of the storage cask was calculated to be

well above 20.6 Hz which is above the amplified region of the ground spectra or the spectra that would

exist on the ISFSI concrete pad. The storage cask frequency was calculated by treating the inner and

outer steel cylindrical shells of the storage overpack as cantilever beams fixed at the base.

Conservatively, the stiffness contribution of the concrete inside the overpack and the MPC shell was not

included. The calculation considered flexural and shear effects in calculating the beam frequency. The

seismic analysis also assumed no amplification attributable to soil-structure interaction. Based on the

above discussion, the use of the equivalent static method without any amplification would be appropriate

and the resulting acceleration calculated for sliding and tip-over would correspond to the maximum

ground acceleration.

Two cases are considered in this evaluation, a 3-D earthquake (2 horizontal and 1 vertical) and a 2-D

earthquake (1 horizontal and 1 vertical). Both are being considered because the subject plant FSAR

indicates that the three directional responses are combined by the square-root-of-sum-of-squares method

or by the absolute sum of the worst horizontal with the vertical. The vertical acceleration is assumed for

this calculation to be equal to 2/3 of the horizontal acceleration. This is consistent with most nuclear

power plants in the United States and is consistent with the FSAR for the subject plant.

To determine whether seismic forces cause sliding or tip-over, the seismic force that would initiate

sliding was calculated and compared to the force needed to cause tip-over of the storage cask. Sliding

would occur when the seismic force exceeds the frictional resistance of the storage cask attributable to

*dead weight. The frictional resistance (Ff) is defined in Equation A.13.

Ff = uN (A.13)

where: t = coefficient of friction between steel and concrete

N dead weight of the fully loaded storage cask

Since the coefficient of friction varies, a range of values for jI was utilized as follows:

• g. (minimum) is 0.25, based on a conservative value (for lower bound) in Reference A. 1, Section

3.4.6 & 3.4.7.

* ! (maximum) is 0.53, based on Reference A.12, Section 3.4.

* j (mean) is 0.43, based on Reference A. 13. This paper presents results of tests conducted with

anchored baseplates which confirm that the range of g shown above is reasonable and also

indicates that the standard deviation of 4 is 0.09.

A-51
Using these coefficients of friction for a 3 D and 2 D earthquake, the seismic forces, in terms of

individual horizontal acceleration, that would cause sliding are calculated. To obtain the corresponding

vertical accelerations, the horizontal accelerations are multiplied by 2/3. The resulting governing case

(sliding) is presented in Table A.7 for the 2-D and 3-D earthquake cases.

Horizontal acceleration required for tip-over was calculated to be 0.529 g for 3-D case, and 0.410 g for

2-D case. Since the accelerations shown in Table A.7 are less than the accelerations that would cause

tip-over, it was concluded that the storage cask would slide and not tip-over.

Table A.7. Seismic Maximum Accelerations for Sliding

2 - D Earthquake Plant 3 - D Earthquake Plant

Horizontal Vertical Horizontal Vertical

Coefficient of Acceleration Acceleration Acceleration Acceleration

Friction (4) For Sliding (g) (g) For Sliding* (g) (g)

0.25 (min) 0.214 0.143 0.243 0.162

0.43 (mean) .334 0.223 0.413 0.275

0.53 (max) 0.392 0.261 0.5 0.333

* For the 3 D earthquake, the horizontal accelerations correspond to each of the two orthogonal directions.

It should be noted that these two peak accelerations would not occur at the same time which is accounted

for in the square-root-of-sum-of-squares method used in this calculation.

A.4.8.2 Refined Seismic Evaluation

Calculations for the seismic question were independently done with the consideration of soil-structure

interaction. A three-dimensional coupled finite element model of a cylindrical storage cask, a flexible

concrete pad, and an underlying soil foundation was analyzed using the ABAQUS/Explicit code,

(Reference A. 14). The cylindrical storage cask is partitioned into four horizontal sections with six rows

of solid elements in each section and 64 elements around the outside perimeter. The density of solid

elements in each horizontal section is calculated and distributed in such a manner that the center of

gravity of the storage cask is located at the correct design position. The storage cask and its internals are

modeled as elastic bodies.

In the coupled model, the base of the concrete pad was assumed rigidly bonded to the top surface of the

soil foundation. At the storage cask/pad interface, a sliding contact condition was assumed with an

assigned friction coefficient of 0.25, or 0.53. The top of the pad was designated to be the "master"

surface and the underside of the module to be the "slave" surface to prevent any portion of the module

from entering the pad. The four vertical sides of the soil foundation were represented by edge columns

that allow horizontal shear deformation only in order to simulate infinite boundary conditions.

A-52
The earthquake excitation information was treated with a deconvolution procedure to produce a modified

time history of deconvoluted accelerations with properly adjusted frequencies and magnitudes to be

applied at the base of soil foundation. The concept of deconvolution is a mathematically rigorous

solution process that applies the wave propagation equation of the free-field surface along with the

boundary conditions. It has been proven that the solution would be unique and rigorously correct for a

linear representation of the soil mass (that is, linear shear modulus and viscous damping model). The

deconvolution procedure has been discussed in References A, 15 and A. 16. The input motions of

deconvoluted seismic accelerations were applied to all nodes at the base of the soil foundation. The

following data were used for the storage overpack with MPC-68 fully loaded with fuel:

* Weight = 163,293 kg (360,000 lb)

* Outside Diameter = 3.37 m (11.04 ft)

* Height = 5.87 m (19.27 ft)

* Height of center of gravity above pad = 3.00 m (9.86 ft)

A section of the concrete pad that holds 4 (2x2) storage casks was selected as the optimal dimension for

the pad in the coupled model, even though only one storage cask was simulated. The selected pad was

9.45 meters (31 feet) square. The pad thickness was 0.61 meter (2 feet),

The size of soil foundation model plays an important role in assessing the soil-structure-interaction

effect. Sensitivity studies on the size were performed to demonstrate that its chosen model size could

simulate the behavior of a semi-infinite soil foundation underneath the concrete pad. Three different

lateral dimensions were selected having 1.0, 1.22, and 1.33 times a baseline lateral geometry of a square

of 85 meters (279 feet), which is nine times the pad dimension. It should be noted that the outside layer

of elements on the four vertical sides of soil foundation model, with width equal to the pad dimension of

9.45 meters (31 feet), are represented by edge columns. This model setup indicates that the true model

size is defined by the nodes at the inner row of the layer with degrees of freedom constrained to those at

the outside row. Therefore, the baseline geometry of soil foundation model is only seven times the pad

dimension. This selection of the lateral dimension of soil foundation model follows the soil-structureinteraction modeling guidelines in Reference A. 17.

The sensitivity studies on the model sizes for the horizontal displacements of the storage cask base with

respect to the pad showed closer displacement solutions for the cases of 1.22 and 1.33 times the baseline

geometry, indicating that a convergence trend of soil foundation model sizes was observed. Therefore,

the final soil foundation model, which was selected to be 1.22 times the baseline geometry, was a square

with length of 103.94 meters (341 feet). In addition, a depth of 56.39 meters (185 feet) ,which was

partitioned into eight horizontal layers was selected for soil foundation model.

The storage cask and the concrete pad are assumed to behave elastically when subjected to seismic

excitations. Therefore, their elastic material properties were chosen as shown in Table A.8.

Table A.8 Elastic Material Properties of Storage Cask and Pad

Structural Element Young's Modulus, E (MPa) (x 10') Poisson's Ratio, ±t Density, p (kg/m3)

6408 (Section 1)*

Storage cask 27.789 0.2 2600 (Section 2)*

2600 (Section 3)*

3407 (Section 4)*

Pad 24.856 0.2 2403

A-53
The geometric definitions of the horizontal sections are as follows:

* Section 1: from storage cask base to 0.2 m (8 inches) above base

* Section 2: from 0.2 m (8 inches) above base to 0.6 m (2 ft) above base

* Section 3: from 0.6 m (2 ft) above base to 3 m (9.86 ft) above base

* Section 4: from 3 m (9.86 ft) above base to storage cask top

The 56.39-meter (185-foot) depth of soil foundation model is partitioned into eight horizontal layers.

The depth variation of soil properties such as shear wave velocity and damping profiles was developed to

incorporate the site specific soil profiles. The best-estimated strain-compatible soil properties were

selected to be used in the seismic analyses. The weight density of the top 56.39 meters (185 feet) of soil

foundation was estimated to be 2,162 kg/m3 (135 lb/ft3

).

The final seismic analyses used the mass proportional soil damping, x" for each layer such that

S a0 whra 2 ¢r

= n where aT (A.14)

For shear wave propagation model of horizontal motion site response, site period T is assumed as 0.53

second. For compression wave propagation model of vertical motion site response, site period T is

assumed as 0.33 second.

The time history utilized was the one used by the applicant for the design basis. The selected PGA level

is 0.15 g for the two horizontal directions and 0.1 g for the vertical direction. Two horizontal

components and one vertical component of the seismic accelerations, which are "surface" motions, were

applied simultaneously to the coupled model. A frequency domain deconvolution procedure was applied

to the time histories of seismic surface accelerations. This procedure uses the Fourier transforms to

adjust simultaneously their magnitudes and frequency contents. The net outcome is that when

deconvoluted seismic motions are applied at the base of soil foundation model, the dynamic

characteristics of the original seismic motions is preserved and the desired surface shaking intensity can

be achieved.

For seismic excitation analyzed with horizontal PGA of 0.15g, and vertical PGA of 0.1g using 3-D

coupled nonlinear finite element model of the storage cask, pad and soil underneath, the analyses show

that the storage cask may slide less than 2.5 mm (0.1 inch) assuming lower bound coefficient of friction

0.25, and will not tip-over with any coefficient of friction from 0.25 to 0.53.

Sensitivity analyses were then performed for different level of seismic excitations with different

coefficients of friction between the bottom steel plate of the storage overpack and the concrete pad using

three-dimensional coupled finite element models of the cylindrical storage cask, a flexible concrete pad,

and an underlying soil foundation. When higher levels of seismic excitations are used in the analyses,

the coefficient of friction at the storage cask/pad interface plays an increasingly important role in the

sliding and rotational behavior of the storage cask.

For a minimum friction coefficient, the storage cask exhibits a translational motion without much

rotation. When the friction coefficient is 0.25, eleven times the design-basis earthquake of the ISFSI pad

(i.e., 11 DBE) is needed to cause the storage cask to slide 0.93 m (36.6 inches); this is more than half the

separation distance of 1.36 m (53.5 inches) between neighboring casks.

A-54
For a maximum friction coefficient, the storage cask experiences more rotational movement, but not tipover. When the friction coefficient is 0.53 and the seismic excitations are increased to nine times the

design-basis earthquake (i.e., 9 DBE), the neighboring storage casks may collide as the cask top moves

1.19 m (46.7 inches) and the cask base slides 0.84 m (33.1 inches). In these cases, the storage cask will

slide but not tip-over. Collision may occur at the seismic excitation level of 9 DBE or higher. This level

of earthquake is so intense that the plant may have serious consequences, and the consequences

attributable to cask collision will be of no significance compared to the plant.

A.4.9 Question 9

What impact forces attributable to given small aircraft, tornado missiles (beyond design basis) and heavy

objects in flood waters will cause the storage cask to slide and/or tip-over? If an object of such an impact

penetrates the storage overpack, what would be the stresses in the MPC?

The aircraft impact to be evaluated is from a small, non-commercial aircraft, Gulfstream IV, accidentally

hitting the storage cask either by a free fall or with its engine running.

The tornado missiles to be considered are the Spectrum II missiles identified in Reference A. 18. The

calculations below use the tornado missile which most likely will cause the storage cask to slide/tip

and/or penetrate the storage overpack and strike the MPC. If this initial analysis shows no penetration or

negligible stresses in the MPC, less severe missiles need not be evaluated; otherwise, analyze them in the

order of severity (i.e., the most severe one first).

A.4.9. 1 Missile Impact Analysis

This calculation addresses the effects of the missiles impact such as automobile and aircraft crashing into

the storage overpack. The missile and storage cask are assumed to be rigid. Missile impact on the

storage cask can result in either tip-over or sliding the storage cask. The tip-over and sliding modes are

considered independently and it is assumed that the storage cask either slides or tips over with no

interaction between the two response modes. An approach is first developed to evaluate the effects of

any type of rigid missile impact. Then, this approach is used to analyze the impact attributable to an

automobile and small aircraft crash.

A.4.9.1.1 Tip-Over Due to Missile Impact

The tip-over mode is considered first. Consider the storage cask with the configuration shown in Figure

A.33. A missile of mass (in) and velocity (V) impacts the storage cask at a distance (h) above the base.

The horizontal velocity of the missile is first converted to an angular velocity (wr) about point (0) using

the following.equation:

V Cosý (A.15)

m - C

A-55
ý4 D

Missile (in) "[rT

q4A

h

hcg

Y

0

Figure A.33. Tip-Over of Storage Cask

Conservation of momentum results in a combined angular velocity of the storage cask and missile equal

to the following:

MC V Cosý

ý' ýA 2 + M C2ý (A.16) A.6

Conservation of energy principles are then used to determine the initial angular velocity (cow) required to

tip-over when the center of gravity is directly above point 0:

2 0)c =

2[W A(1- Coso)+ mg C(1- Cos )]

WA + mC2/

g

(A.17)

Combining Equations A. 16 and A. 17 and solving for (V) results in the following equation:

2[WA(1- Cos 0) + mgC(1- Cos A 2g + mC2I

v2 = (A.18)

(m C Cos )

2

A-56
A spreadsheet is then used to evaluate, V, given a missile mass, m, where:

W 163,290 kg (360,000 lb)

D = 3.4 m (133.875 inches)

g = 3 m (118.4 inches)

S = 29.50

A = 3.45 m (136 inches); C = (h2 + D2

)0.5

= tan' (D/h); the value of (h) is taken as 1, 0.75, 0.5, and 0.25 times the height of

the storage cask-- 5.87 m (231.25 inches)

The resulting velocities required to tip-over the storage cask are shown in Figure A.34.

7--

7--

550

500

450

400

350

300

2150

200

150

100

50 I I

10 15

Missile Weight (kilo-lb) 1 kilo-lb - 453.6 - k--. - mi/l =

0 20 25

1.61 kmii/h

Figure A.34. Missile Velocity for Tip-Over of Storage Cask

A-57
A.4.9.1.2 Sliding Due to Missile Impact

Sliding of the storage cask is considered next. A sketch of the problem is shown in Figure A.35.

Missile (m)

w

•L*W

Figure A.35. Sliding of Storage Cask

The height of the missile impact is not important for this problem, because here we are concerned with

the sliding mode only. Conservation of momentum results in the following initial velocity (x',) of the

storage cask after impact:

mV

XI(m++ W (A.19)

The equation of motion for the storage cask after impact is as follows:

x" +/I g = 0 (A.20)

Integrating Equation A.20 and evaluating the two constants of integration from the initial velocity of

Equation A. 19 and zero initial displacement results in the following solutions for velocity and

displacement:

mVt )Ugt 2

2 (A.21)

where t = time after impact

A-58
The maximum displacement occurs when the velocity equals zero and is found to be as follows:

mV

m/Wgt (A.22)

2

(A.23 )

Xmax= 2ug

The clear spacing between the storage casks on the ISFSI pad is at least 1.3 meters (52.1 inches) and the

closest clear distance from the edge of the storage cask to the edge of the pad is 66 cm (26 inches).

Solutions are therefore obtained for xm. = 66 cm (26 inches). Three values of the coefficient of friction

are used (see calculations for Question 8); a lower bound of 0.25, a mean value of 0.43, and an upper

bound of 0.53. Equation A.23 is solved for the velocity required to slide the storage cask 66 cm (26

inches) given the missile mass. These solutions are shown in Figure A.36.

A-59
550

500

450

400

><• 300 250 : ~cint c:o:: Fr~~~ictinCe'ent 0.5 3 3

- 150 '

F~~~1riction ;of~i 0.25l

0510 0 5 20 25

Missile Weight (kilo-0b)

1 kilo-lb =453.6 1,:g 1 mi/h = 1.61 k~m/lh

Figure A.36. Missile Velocity Required to Slide the Storage Cask by 66 cm (26 inches)

The following conclusions can be drawn by comparing Figures A.34 and A.36:

(1) Impact velocities required to cause the storage cask to slide are about the same as those required

to cause it to tip-over if the coefficient of friction = 0.25 and the impact is at the top of the

storage cask. The storage cask will slide rather than tip-over if the impact is below the top of the

cask and the friction coefficient = 0.25.

(2) The storage cask will tip-over rather than slide if the friction coefficient =0.43 and the impact

height is larger than• about 80 % of the height. It will slide if the impact is below this height.

(3) The storage cask will tip-over rather than slide if the friction coefficient = 0.53 and the impact

height is larger than about 62.5 % of the height. It will slide if the impact is below this height.

A.4.9.2 Vehicle Impact Analysis

A car or a truck crashing into the storage casks while stored on the ISFSI pad is not analyzed directly for

a driver's accidental maneuver, physical ailment or mistake.. However, by extrapolation from the

analyses done for tip-over and sliding attributable to missile impact (from Figures A.34, 35, and 36

respectively), it can be concluded that a 4,536 kg (10,000 lb) vehicle traveling at 241 km/h (150 mi/h)

speed will slide the storage cask less 66 cm (26 inches), and will not cause tip-over. This conservatively

assumes that the impact occurs at the top of the storage cask (for tip-over) and the lowest Coefficient of

friction applies (for sliding).

A-60
A.4.9.3 Aircraft Impact Analysis

The analyses described below are included in this appendix to provide a methodology, for determining the

response of a dry storage cask to aircraft impact. The results of the aircraft impact analyses are used in

the PRA to evaluate aircraft impact events on overall risk.

As discussed in Section 3.3.7 of the main report, the largest aircraft that is able to land at any of the four

local airfields near the ISFSI site is a Gulfstream IV. This aircraft's geometric characteristics were used

to estimate the frequency of an aircraft crash into a storage cask during takeoff or landing. The data and

equations presented in Sections A.4.9.3.1 and 2 are used to evaluate the global and local structural

response of a storage cask to impact by a Gulfstream IV. Section A.4.9.3.3 summarizes the results and

discusses the probability of breaching the MPC confinement boundary.

A.4.9.3.1 Global Response

The Gulfstream IV aircraft weighs 74,600 lbs on takeoff and is powered by two Roll-Royce Tay 611-8

turbofan jet engines each weighing 3,135 lbs. The DOE Standard on aircraft impact (Reference A. 19)

recommends that the gross response of a target can be evaluated by considering the impact of a rigid

missile having a mass equal to twice the engine weight. Thus, the appropriate missile weight for the

analysis is 2 x 3,135 = 6,270 lbs. Figures A.34 and A.36 are used to obtain the impact velocity required

for tip-over and sliding. The results are shown in Table A.9a.

Table A.9a. Global Response of Storage Cask to Aircraft Impact

Tip-Over Sliding

h/L Impact Velocity Required Impact Velocity Required

(mph) (mph)

1 250 0.25 260

0.75 340 0.43 330

0.5 500 0.53 370

A.4.9.3.2 Local Response

Another mode of response to missile impact is local penetration or perforation of the shell and lid of the

storage overpack. The DOE Standard (Reference A. 19) recommends that the local damage evaluation be

performed using relatively nondeformable components of the aircraft. Such a component is the engine

shaft. The Rolls-Royce Tay 611-8 engine is 94.8 inches long and 45.0 inches in diameter. Exact

dimensions of the engine shaft were not available from the manufacturer, so dimensions were estimated

from available photographs and cross sections ofjet engines of this class.

A typical turbofan engine has multiple concentric shafts, but for the purpose of this analysis a single solid

shaft is assumed. The shaft is assumed to have an average diameter of 4.0 inches over 80% of the engine

length. The resulting weight of the shaft is (p x 22 x 94.8 x 0.80 x 490/1728 = 270) 270 lbs.

The shell of the storage overpack consists of an outer steel shell 0.75 inches thick, a poured concrete

shell 26.75 inches thick and an inner steel shell 1.25 inches thick. The lid consists of a top steel plate 2.0

inches thick, a bottom steel plate 2.0 inches thick, a concrete shield plug 10.5 inches thick and a bottom

shield plate 1.25 inches thick.

A-61
The DOE Standard on aircraft impact analysis (Reference A. 19) recommends the Chang formula

(Equation 24a) to compute the perforation thickness of concrete targets and the Ballistic Research

Laboratory (BRL) formula (Equation A.24b) to compute the penetration thickness of steel targets:

0.5(MV 2 0.5

t = (A.24a)

where:

U = reference velocity = 200 ft/sec

V = missile impact velocity (ft/sec)

M = mass of the missile = W/g

W = missile weight (lb), g=32.2 ft/sec 2

D = effective missile diameter (ft)

f'c = ultimate compressive strength of concrete (lb/ft2)

= mean perforation thickness (ft)

T15 = 0.5MV2 (A.24b)

17,4OOKS D'

where

T = predicted thickness to just perforate a steel plate (in.)

M - W/g missile mass (lb-sec2/ft)

V - missile impact velocity (ft/sec)

K= constant depending on the grade of steel (usually z 1)

D = missile diameter (in.)

Both formulae in their standard form assume the missile is rigid. To prevent perforation of the target,

Reference A. 19 recommends that the thickness computed using the Chang and BRL formulae be reduced

by a factor of 1.2 and 1.25 respectively.

The storage overpack shell and lid consist of multiple layers of steel and concrete. The approach used to

determine the impact velocity that will just perforate all the layers of the storage overpack shell and lid is

to separately calculate the kinetic energy associated with the impact velocity required to just perforate

each of the steel and concrete layers, and then add the kinetic energies from each layer to determine the

total kinetic energy and resulting impact velocity. This methodology is based on the conservative

assumption that each layer behaves independently and that there is no interaction between layers. Since

multiple layers resist perforation, it would be overly conservative to reduce the thickness of each layer by

the reduction factors cited above, because these factors are intended to prevent perforation. To establish

a realistically conservative best estimate of the impact velocity that will just perforate all the layers, only

the layer contributing most to perforation resistance is reduced in thickness.

A-62
For each storage overpack shell and lid layer, Table A.9b shows the material and thickness of the layer,

the reduced thickness, the kinetic energy required to just perforate the layer, the total kinetic energy from

all layers and the resulting impact velocity of the aircraft. This methodology produces conservative

results. More accurate results can be obtained using explicit dynamics found in several commercially

available computer codes. Before using such codes, however, material models, elements and meshes

should be benchmarked against the results of actual test data.

Table A.9b. Aircraft Maximum Impact Velocity that Does Not Perforate

the Storage Overpack Shell or Lid.

Kinetic Maximum

Reduced Eneto Total Kinetic Impact

Thickness Thickness Energy to Energy Velocty Event Target Layer (in) (in) Perforate (ft-lbs) city\,

(ft-lbs) (mph)

Outer Steel 0.75. 0.75 (3)

Horizontal Shell

Impact into Concrete Shell 26.75 22.3 (3) (3) (3)

Overpack

Shell Inner Steel 1.25 1.25 (3)

Shell

Top Lid Plate 2 2 (3)

Bottom Lid

Vertical Plate

Impact into Shield (3) (2)(3)

Overpack Lid Concrete 10.5 10.5 (3)

BottomShield 1.25 1.25 (3)

Plate

(1) The methodology used here produces conservative results. More accurate results can be obtained using

explicit dynamics found in several commercially available finite element computer codes. Before using

such codes, material models, elements and meshes should be benchmarked against the results of impact test

data.

(2) Assumes an impact angle of 50 degrees above the horizontal.

(3) Text withheld under 10 CFR 2.390.

A.4.9.3.3 Results of Aircraft Impact Analysis

Aircraft speeds during landing and takeoff generally range from 140 to 200 mph (Reference A. 19).

Table A.9a shows that cask tip-over or sliding attributable to aircraft impact would not be expected to

occur until impact speeds exceed 250 mph, and then, only if the aircraft struck near the top of the cask.

The response of the MPC to cask tip-over is far more significant than to cask sliding and, therefore, only

cask tip-over need be considered. As reported in Table B.6, the probability of MPC failure attributable to

cask tip-over is less than 1xIl 06 . Table A.9b shows that perforation of the storage overpack would not

occur until impact speeds exceed (text withheld under 10 CFR 2.390). For the largest aircraft that can

land or takeoff from any of the four airports near the site, these results show that the storage overpack

provides a rugged barrier that makes the likelihood of an MPC breach from aircraft impact exceedingly

small.

A-63
Based on these results, the probability of breaching the MPC, if struck by an aircraft during, landing or

takeoff is extremely small and is assumed to be zero. However, many commercial aircraft overflying the

site would be larger than a Gulfstream IV and could impact the cask at velocities much higher than those

associated with landing and takeoff. Since the makeup and characteristics of commercial aircraft

overflying the site are not known, and have not been evaluated, it is conservatively assumed that all

commercial aircraft overflying the site are larger than a Gulfstream IV and that the probability of MPC

breach attributable to the impact of an overflying commercial aircraft is, for the purpose of the PRA, 1.0.

Therefore, the probability of MPC failure and release, if struck by an aircraft, is, for the purpose of the

PRA, equal to the frequency of overflight crashes (9.0xl0' 11, see Section 3.3.7) divided by the total

frequency of aircraft crashes (6.3x 10-9, see Section 3.3.7), which is 0.014.

A.4.9.4 Tornado Missile Impact Analysis

This calculation considers the local effects of tornado missiles hitting the storage cask and addresses the

question of whether the missiles penetrate the storage overpack.

Reference A. 1 states that the storage overpack (HI-STORM 100) system was checked for a constant wind

speed of 579 km/h (360 mi/h), a pressure drop of 20.7 kPa (3.0 psi) and three tornado-generated missiles.

The three missiles are the same as the Spectrum I missiles identified in Reference A. 18 and are

characterized as follows:

(1) automobile with a mass of 1,800 kg (3,970 lb) and a velocity of 203 km/h (126 mi/h)

(2) rigid solid steel cylinder having a diameter of 20.3 cm (8 inches) with a mass of 125 kg (276 lb)

and a velocity of 203 km/h (126 mi/h)

(3) solid sphere having a diameter of 2.5 cm (1 inches) with a mass of 0.22 kg (0.49 lb) and a

velocity of 203 km/h (126 mi/h)

The analysis in Reference A. 1 demonstrates that the combination of the large tornado missile

(automobile) plus either steady tornado wind or instantaneous pressure drop causes a maximum

horizontal excursion of the storage cask midpoint (approximately equal to the cask center of gravity) of

less than 14 cm (5.5 inches) or 2.7 degrees rotation. In order for the storage cask to tip-over, Reference

A. I reports the centroid must undergo a horizontal displacement of 168 cm (5.5 feet). Therefore, the

storage cask will not tip-over as a result of the automobile impact. Furthermore, it is shown in Reference

A. 1 that the other two postulated missiles will not penetrate the storage overpack. The Spectrum II

missiles in Reference A. 18 are characterized in Table A. 10.

Table A.1O. Characteristics of Spectrum II Missiles

Missile Mass (kg) Dimensions (m) Velocity mn/s (mi/h)

(A) Wood Plank 52 0.092 x 0.289 x 3.66 83 (186)

(B) 6" Sch 40 pipe 130 0.168D x 4.58 52(116)

(C) 1" Steel rod 4 0.0254D x 0.915 51(114)

(D) Utility pole 510 0.343D x 10.68 55 (123)

(E) 12" Sch 40 pipe 340 0.32D x 4.58 47 (105)

(F) Automobile 1810 5x2x 1.3 59 (132)

A-64
A.4.9.4.1 Automobile as Tornado Missile

The Spectrum HI automobile has about the same mass as the Spectrum I automobile and the velocity is

only slightly higher, i.e., 212 km/h versus 203 km/h (132 mi/h versus 126 mi/h). Therefore, the Spectrum

II automobile missile when combined with the tornado wind or pressure drop will not cause the storage

cask to tip-over.

The Reference A. 1 report does not provide an evaluation of the potential for the storage cask to as a

result of the combined effects of an automobile impact and tornado wind. The potential for this effect

was evaluated in this appendix. The Principle of Conservation of Momentum was used to calculate the

initial velocity of the storage cask attributable to the impact of the automobile. The initial velocity of the

storage cask was determined to be 2.33 kmih (1.45 mi/h). The force attributable to tornado wind effects

was taken from Reference A. 1 which showed that the force attributable to pressure drop governed over

the tornado wind pressure.

The equation of motion for the struck storage cask is as follows:

x" +,u W- FTornado = 0

g (A.24)

where: W

FTorado

= weight of storage cask

= force attributable to tornado wind effect

Solving for x" results in the accelerations shown in Table A. 11 for the range of coefficient of friction.

Table A.11. Acceleration for Range of gi

Coefficient of Friction (p) Acceleration(g)

minimum 0.005

mean -0.18

maximum -0.28

For a coefficient of friction of pmin = 0.25, the storage cask will have a positive acceleration of 0.005 g.

This means that the storage cask will accelerate and continue to displace without stopping until it slides

off the pad or hits another storage cask. For a coefficient of friction greater than 0.255, the storage cask

acceleration is negative which means it will decelerate until it stops.

A-65
A.4.9.4.2 Other Spectrum II Tornado Missiles

The potential for the other Spectrum II missiles to penetrate the storage overpack was evaluated using the

procedures given in Reference A. 19 on pages C-6 through C-9. The penetration depth (x) is given as

follows:

=4K WD V if-<2 x (A.25)

x= KNWD ( +D if D-> 2 (A.26)

where: x penetration depth (in) K = concrete penetrability factor = 180/(')5= 2.846

N = shape factor = 0.72 for blunt missile

W = weight of missile - V = velocity of missile in ft/sec

D = diameter of missile (in.)

When the missile first strikes the target, there is a tendency for the front face of the concrete to scab. The

barrier thickness (td,) required to prevent this is given as follows:

tdsl 1.1(2.12D+ 1.36x) for0.65< x/D <11.75 (A.27)

tds- 1.1 7.91x-5.06 forx/D < 0.65 (A.28)

If this scabbing thickness is not satisfied the penetration depths are likely to be larger than computed with

Equations A.25 and A.26. Note that the 1.1 factors in Equations A.27 and A.28 are safety factors added

to curve fits of empirical data. Of course, the outer steel plate on the storage overpack prevents this

scabbing and, therefore, is not considered in this analysis.

The barrier thickness required to prevent the missile from perforating the shield (tdp) is then as follows:

tdp=1.2(1.32D+ 1.24x) for 1.35 < x/D < 13.5 (A.29)

tdp=l.2 3.19x -- - forx/D < 1.35 (A.30)

These calculations are conservative for the storage overpack for several reasons:

* The wall of the storage overpack includes inner and outer steel plates with the concrete contained

within the plates. The above analysis includes only the effects of the concrete barrier so that

lower penetration depths would be expected in the actual storage overpack.

* The above analysis was developed from data collected for plane barriers. The circular shape of

the storage overpack would increase its strength as compared with a plane structure.

* The circular shape of the outer boundary of the storage overpack will tend to deflect missiles

and, thereby, reduce the probability that a missile can penetrate the barrier.

A-66
The above calculations are carried out for the Spectrum II missiles. All of the missiles are included -

except for the automobile which was treated separately as discussed above because it is a large flexible

missile and is more likely to cause gross response (sliding or tip-over) of the storage cask rather than to

penetrate the storage overpack. The results of the calculations are shown in Table A. 12.

The barrier thickness required to prevent penetration is less than the 68 cm (26.75 inches) of concrete

provided except for the 12 inch Schedule 40 pipe where 72 cm (28.5 inches) is predicted as the thickness

required to prevent penetration. Because of the conservatism cited above it is not likely that any of the

missiles will penetrate the storage overpack.

Table A.12. Penetration of Missiles Through Reinforced Concrete Wall of Storag Overpack

Missile Weight Missile Velocity Missile Diameter V/1000D x Req'd

Missile fc (psi) K N (kg) (lb) (m/s) (ft/s) (cm) (in) **1.8 (in) x/D t (in)

Wood 4000 2.846 0.72 52 114.64 83 272.3 18.4 7.244 0.00272 4.31 0.59 14.3

Plank

6" Sch 40 

40 pip pipe

4000 2.846 0.72 130 286.6 52 170.6 16.8 6.614 0.00138 4.64 0.7 14.9

I" Steel 1' 4000 2.846 0.72 4 8.82 51 167.3 2.54 1 0.04003 1.7 1.7 4.1

Rod

Utility 4000 2.846 0.72 510 1124.3 55 180.4 34.3 13.504 0.0004 7.26 0.54 26.6

Pole

12"Sch 40 pipe 4000 2.846 0.72 340 749.56 47 154.2 32 12.598 0.0004 5.29 0.42 28.5

Note: 1 inch = 2.54 cm; I psi = 6.895 kPa

A. 4.9.5 Flood Water/Debris Impact Analysis

This calculation considers whether the effects of heavy objects in floodwaters will cause the storage

cask to slide and/or tip-over.

The subject plant FSAR states that none of the safety-related facilities would be exposed to river

flooding by the most severe flood at the site. It is further reported that the probable maximum discharge

concurrent with a severe wind of 72 km /h (45 mi/h) would develop a wave height of 2.5 meters

(6.5 feet). The elevation of the wave crest attributable to a sustained wind of this magnitude occurring

during the Probable Maximum Flood would be well below the elevation of the top of the dry cask storage

pad. Thus, it would be an incredible event for flood waters to even reach the bottom of the storage cask

and even more incredible for the storage cask to be impacted by heavy objects in flood waters.

The Reference A. 1 report demonstrates that the accident condition design external pressure of 414 kPa

(60 psi) for the MPC bounds the maximum hydrostatic pressure on the storage cask in a flood where the

water level is conservatively set at 38 meters (125 feet). The Reference A. 1 report also evaluated the

factor of safety against tip-over and sliding for a design basis flood velocity of 4.6 m/s (15 ft/s). They

calculated a sliding factor of safety of 2.09 and a tip-over factor of safety of 4.68.

A calculation is performed to evaluate what forces/conditions attributable to floating debris would cause

the storage cask to slide and/or tip-over. For this calculation, a floating tree moving at varying velocities

is considered. The mass of the tree was approximated as three times the weight of the Spectrum II utility

pole which results in a total weight of 1,530 kg (3,370 lb) that is close to the weight of the automobile.

Conservatively, the location of impact was assumed to occur at the top of the storage cask.

A-67
The drag force attributable to the flowing flood water was also included for the same varying velocities

as used for the debris impact. The velocities considered are 1.5 mr/s (5 ft/s), 3.0 m/s (10 ft/s), 6.1 m/s (20

ft/s), and 9.1 m/s (30 fl/s).

A.4.9.5.1 Tip-Over Evaluation

The same methodology developed for the aircraft impact tip-over analysis was utilized (with an

additional term for the constant water drag force). In this calculation, the kinetic energy attributable to

impact and drag force was compared to the potential energy required to cause tip-over. This comparison

led to the conclusion that for velocities less than approximately 10.7 m/s (35 ft/s), the storage cask will

not tip-over. The calculation also concluded that the contribution of the debris impact is negligible and

can be dropped from the flood analysis of the storage cask. This occurred because the mass of the debris

is a very small fraction of the storage cask mass and it is traveling at a very slow velocity.

A.4.9.5.2 Slidinq Evaluation

Sliding would occur if the horizontal drag force is greater than the frictional resistance (the effect of

debris impact can be neglected - see above). The calculation demonstrates that for velocities less than

approximately 7.6 m (25 ft/s), the storage cask will not slide for the entire range of coefficients of

friction. For velocities greater than that, the storage cask will slide depending on the velocity and the

specific coefficient of friction. As an example, for a velocity of flowing water at 9.1 mn (30 ft/s), the

storage cask will slide for mmi = 0.25 but will not slide for gmean =0.43 and ýImax =0.53.

A.4.9.5.3 Analysis for Floods

An analysis for floods was performed with the following assumptions:

* Surface area of the storage cask subject to floods is 231.25 x 132.5 / 144 = 212.78 ft2 or 19.8 M2.

* Storage cask is rigid.

* Storage overpack has ventilation openings so it is assumed that the hydrostatic pressure from

flood submergence acts only on the MPC.

• Water flow velocity is assumed to be uniform over the height of the storage overpack.

• Cd, the drag coefficient is 0.5.

* p, the water weight density is assumed to be 1,000 kg/mi3 (62.4 lbs/ft3).

• g., the coefficient of friction between steel plate and concrete pad is assumed to be 0.25.

Kmin, the buoyancy factor = 0.64 (Reference A. 1)

A-68
Sliding Evaluation

Applied Force (lbs) = Cd • Velocity head • Area

= 0.5 x (p(lb/ft3 ) . V(ft/s) 2 / 2 g(ft/s2 ) x 212.78ft2 106.39 (62.4 V2 / (2 x 32.2))

= 103.09 V2 (where V is in ft/sec; g is in ft/sec2

)

Resisting Force = L" K - W (lbs)

where: t = coefficient of friction

K buoyancy

W weight (lbs)

= 0.25 x 0.64 x W

= 0.16W

For equilibrium, 103.09 V2 = 0.16 W

V2= 0.001552 W

V = (0.001552 W)" (ft/sec)

For overpack with minimum MPC Weight (lbs) V = (0.001552 x 303000)= 21.7 ft/s (6.6 m/s)

For overpack with full MPC Weight V =( 0.001552 x 360000)'/= 23.6 ft/s (7.2 m/s)

Tip-Over Evaluation

Applied Force (ibs) =103.09 V2 (where V is in ft/sec)

Applied Moment = Force • Height of A.G.

= 103.09 V2 x (118.46 inches)/12 inches/foot = 1018 V2 lb-ft

Resisting Moment = K • W • Diameter / 2 where W is weight

= 0.64 x (132.5/24) W

= 3.533 W

For equilibrium, 1018 V2 = 3.533 W

V2= 0.00347 W (lbs)

V (ft/sec) = (0.00347 W (lbs))w

For overpack with minimum MPC Weight (lbs) V = (0.00347 x 303000)' = 32.4 ft/s (9.9 m/s)

For overpack with full MPC Weight (lbs) V = (0.00347 x 360000)' = 35.3 ft/s (10.8 m/s)

To summarize, the flood required to slide the storage cask is at least 23.6 ft/s (7.2m/s), and the flood

required to tip-over the storage cask is at least 35.3 ft/s (10.8 m/s).

A-69
A.4.9.6 Analysis for High Winds

An analysis for high winds (tornado) was performed with the following assumptions:

* Storage cask is assumed to be subjected to a constant wind force for full height.

* Storage cask is assumed to be rigid.

* j•, the coefficient of friction between steel plate and concrete pad is assumed to be 0.25.

* Storage cask is assumed to pivot about a fixed point at the outer edge of the contact surface.

• Surface area of the storage cask subject to wind = 231.25 x 132.5 / 144 = 212.78 ft2 or 19.8 m2

• Cd, the drag coefficient = 1.0

Sliding Evaluation

Applied Force (lbs) = Cd " Velocity head • Area = 1.0 x(p -V2 /2 g) xArea Where p = air density = 0.07651 lb/ft3 @ 15 0C;

V is in ft/sec; g is in ft/sec 2

= 0.07651 / (2 x 32.2) x V2 x Area

= 0.00119 x (5280/3600)2 x V2 x Area Where V is converted to mi/h

= 0.00256 x V2 x Area

= 0.00256 x 212.78 x V2

= 0.5447 V2

Resisting Force = t - W where i' is coefficient of friction; W is weight (lbs)

= 0.25 W (lbs)

For equilibrium, 0.5447 V2 = 0.25 W

V2 = 0.459 W

V (mi/hr) = (0.459 W)'J

For overpack with minimum MPC Weight (lbs) V = (0.459 x 303000)4 = 373 mi/h or 600 km/h

For overpack with full MPC Weight (lbs) V = (0.459 x 36 0000)' = 406 mi/h or 653 km/h

Tip-Over Evaluation

Applied Force = 0.5447 V2 Where V = wind in mi/h

Applied Moment = Force -Height of A.G.

= 0.5447 V 2 x 118.46/12

= 5.377 V2 lb-ft

Resisting Moment = W • Diameter / 2 where W is weight

= (132.5/24) W = 5.52 W lb-ft

For equilibrium, 5.377 V2 = 5.52 W

V2= 1.02675 W

V = (1.02678 W)'/'

For overpack with minimum MPC Weight V = (1.02675 x 303000)' = 558 mi/h (898 km/h)

For overpack with full MPC Weight V = (1.02675 x 36 00 0 0)' - 608 mi/h (978 km/h)

A-70
To summarize, the wind speed required to slide the storage cask is at least 653 km/hr (406 mi/hr), and the

wind speed required to tip-over the storage cask is at least 978 km/hr (608 mi/hr). The design wind speed

for the plant is 580 km/hr (360 mph).

A.4.10 Question 10

If the storage cask tips over on the concrete pad, what would be the stresses in the MPC?

To calculate accurate stresses in the MPC would require more than simplified analytical methods. Since

the Reference A. 1 report has already performed refined analyses, often using detailed finite element

models, it would be best to utilize their results and scale the stresses to suit the scenario/case being

evaluated. Therefore, the stress results from the Reference A. 1 report will be used to tabulate the stresses

in the MPC attributable to the tip-over case, as well as the vertical "end drop" case. The end drop case is

also included in this calculation because other calculations refer to this calculation to respond to some of

the other questions. For example, stresses in the MPC attributable to transfer cask drop (Questions 3 &

4) refer to this calculation for a more complete listing of stresses.

For each type of loading, the acceleration and stresses are provided. If another calculation refer to this

calculation for stresses, then the stresses presented herein should be scaled by the ratio of accelerations

(ANew Ques./AQues. 10). Scaling these stresses to higher acceleration values determined in other calculations is

conservative.

Stresses in the MPC are provided for the fuel basket, basket supports, MPC shell, lid, and baseplate

based on Reference A. 1, Tables 3.4.3, 3.4.4, and 3.4.6. Maximum stresses are presented for the end drop

case and the tip-over case for the MPC components listed above. The acceleration values corresponding

to these stresses are also provided to permit scaling if needed in other calculations. For the tip-over case,

two orientations of the basket walls with respect to the ground were considered, 0 degree and 45 degrees.

The stresses and corresponding accelerations are summarized in the Table A.6.

A.4.10.1 Non-Mechanistic Tip-Over

Non-mechanistic tip-over refers to the case where the storage cask tips over attributable to the center of

gravity of the cask passing over the point of rotation without any initial force or velocity. This case is

non-mechanistic because there is no design basis type loading that would cause the cask to tip-over. It is

presented to have a sense of what could be expected from a scenario that causes storage cask to be tilted

to the point where it tips over. The maximum stress in the MPC, from Table A.6 for the tip-over case, is

406.5 MPa (58,959 psi). This is based on the Reference A. 1 report which calculated the stress inthe

MPC shell using a finite element model of a slice of the MPC and fuel baskets subjected to an

acceleration of 45 g. Under this loading, the MPC and fuel baskets deform outward until portions of the

MPC shell reach the rigid boundary of the storage overpack inner steel shell. This analysis is considered

conservative because it is based on a linear elastic approach. This stress of 406.5 MPa (58,959 psi) needs

to be multiplied by a dynamic load factor of 1.04. Before applying the dynamic load factor, stresses

attributable to internal pressure of 0.69 MPa (100 psi) must be removed, and then added after applying

the dynamic load factor.

Stress attributable to 689 kPa (100 psi) internal pressure for the MPC shell thickness of 1.27 cm (0.5

inch) and an average diameter of 1.74 meter (68.375 inches or 5.7 feet) is calculated by p • r / t = (100•

68.375/0.5)/0.5 = 6,838 psi or 47.1 MPa.

A-71
Stress in the MPC for tip-over = (58959-6838) • 1.04 + 6838 =61,044 psi or 421 MPa

So, the maximum circumferential stress in the MPC shell is approximately 421 MPa (61,044 psi) at

4.53 m (178.5 inches or 14.9 feet) above the bottom of the MPC. The maximum calculated stress is less

than the allowable stress intensity of 450 MPa (65,200 psi) under Level D Service Condition of the

ASME Code. Therefore, it does not cause the failure of the MPC.

The stress intensity of the MPC shell at the junction with the baseplate is the combination of stress

attributable to internal pressure and the stress attributable to g value caused by tip-over. The stress

intensity attributable to internal pressure of 0.69 MPa (100 psi), is 303 MPa (43,986 psi), based on the

finite element analysis from Reference A. 13. The stress intensity attributable to the tip-over is negligible

at the MPC shell junction with the baseplate.

The total stress intensity at that junction is nearly the same as the stress caused by the internal pressure

alone, which is less than the allowable stress intensity for Level D Service Condition of the ASME Code.

Therefore, it does not cause the failure of the MPC.

If the storage cask tips over on another storage cask or on the ISFSI pad, calculations demonstrate that

the studs retaining the lid on the overpack will resist theimpact loads. In addition, the struck cask(s) will

not tip-over.

A.4.11 Question 11

If the storage cask tips over, will the impact of the cask tipping over on another storage cask or on the

concrete pad cause the overpack lid to be dislodged such that the MPC could slide out?

The bounding case is when the storage cask falls onto the concrete pad because this causes a larger

acceleration. An analysis of the overpack lid integrity has already been performed by Reference A. 1,

Section 3.4.4.3.2.2, Appendix 3.K, and Appendix 3.L. Section 3.4.4.3.2.2 of Reference A. 1 provides a

summary of the overpack lid evaluation. Reference A. 1, Appendix 3.K provides the stress analysis of the

overpack lid. The Reference A. 1, Appendix 3.L provides the calculation that demonstrates that the 4

studs hold the lid in place. For the reader's benefit the relevant portions of Appendix 3.K and Appendix

3.L are reproduced below.

A.4. 11.1 Appendix 3.K, Lid Stress Analysis

Appendix 3.K describes the analysis of the lid attributable to the tip-over of the storage cask. It states

that the maximum acceleration at the top of the overpack lid is 49 g as determined in Reference A. 1,

Appendix 3.A. Using the 49 g loading, the following lid components were checked in Reference A. 1:

(1) Lid Top Plate Stress Analysis (Plate Bearing Stress at Bolt Holes)

(2) Lid Shell-To-Lid Top Plate Weld

(3) Shield Block Shells-To-Lid Top Plate Weld

(4) Lid Shell Stress Evaluation

Reference A. 1, Section 3.4.4.3.2.2 summarizes the limiting stress of the lid to be as follows:

,u = 61.6 MPa (8,940 psi)

7 (allowable) = 202.7 MPa (29,400 psi)

F actor of Safety = 3.29

A-72
A.4.11.2 Appendix 3.L, HI-STORM Lid Top Plate Bolting

The Appendix 3.L calculation shows that the four studs of 8.9 cm (3.5 inches) diameter holding the lid in

place have sufficient strength. The calculation uses 48.5 g as the acceleration attributable to the tip-over

based on Reference A.1, Table 3.A.4. Using a weight of 10,433 kg (23,000 lb) for the lid times 48.5 g

divided by the four studs result in a shear load on each stud equal to 126,099 (278,000 lb). This results

in a shear stress as follows:

'r (bolt) = 231.8 MPa (33,620 psi) per bolt

-r (allowable) = (min. of 0.42Su or 0.6 Sy) = 420 MPa (60,900 psi)

Factor of Safety= 1.81

Appendix 3.L also calculated the tensile stress attributable to (1) bending of the stud as a result of the

lateral load acting on the stud over the thickness of the lid plate and (2) the pre-load of 407 N'm (300 lbft). The total tensile stress was calculated to be as follows:

a = 98,260 (bending) + 1,124 (pre-load) = 99,384 psi (685 MPa)

o (allowable) = S. = 1000 MPa (145,000 psi)

Factor of Safety = 1.46

Appendix 3.L then calculated the combined effect of tension and shear using an interaction equation

which resulted in a factor of safety of 1.29.

It appears that the Reference A. 1 report did not consider the effect of centrifugal force attributable to

rotation of the storage cask. The centrifugal force would cause the MPC to push against the overpack lid

imposing loads on the lid and studs. To determine the significance of this, the tensile stress on the studs

attributable to centrifugal force was evaluated. The stress was calculated to be as follows:

o (centrifugal force) =19.7 MPa (2,857 psi)

This stress would not be in addition to the pre-load but in-place of the pre-load because this load would

relieve the pre-load. Considering tension only and re-calculating the factor of safety, it reduces from

1.46 to 1.43 ( a very small reduction). Considering shear and tension interaction, the factor of safety also

shows a slight reduction from 1.29 to 1.26.

Based on the above discussion, it can be concluded that the overpack lid and studs are structurally

adequate to resist the loads attributable to a tip-over of the storage cask and, thereby, will prevent the

MPC from sliding out of the storage cask.

A.4.12 Question 12

The explosion-related shock waves that should be evaluated are those generated from (1) an explosion

from a typical tanker trailer containing explosive materials (gasoline, liquid/natural gas) at a distance of

914 meters (3,000 feet) from the storage cask, and (2) an explosion of a natural gas pipeline at 7.2 km

(4.5 mi) from the storage cask. Would the structural integrity of the cask system be in jeopardy if the

shock waves caused by the postulated explosions are applied? If not, provide the supporting arguments.

A-73
Event A - Tanker Trailer at 914 meters (3,000 feet)

Evaluation of Event A was conducted utilizing Reference A.22. Using the recommended Trinitrotoluene

(TNT) equivalence factor of 240% and 22,680 kg (50,000 Ib) cargo weight, the effective TNT weight is

54,431 kg (120,000 lb). From Equation (1) of Reference A.22, the peak transient pressure would be 1 psi

at 677 meters (2,220 feet). At 914 meters (3,000 feet), the peak pressure would be less than 6.9 kPa (1

psi). The HI-STORM 100 cask (storage cask) is designed for a 69 kPa (10 psi) peak transient external

pressure. Therefore, at 914 meters (3,000 feet), the peak pressure attributable to the explosion would be

less than 1/10 the pressure magnitude that the storage overpack is designed for.

Event B - Pipeline at 7.2 km (4.5 mi)

Evaluation of Event B was conducted utilizing References A.22 and A.23, which developed the

probability of exceeding 3 psi external overpressure on the containment building of a proposed plant

from a 21-inch ID pipeline containing methane at 4.9 MPa (720 psi) at a distance of 808 meters (2,650

feet). The best estimate annual probability of exceedance was 6.Ox 107/yr. Considering uncertainties, the

range was 3.8 x I08/yr to 2.3 x 106/yr. Considering the higher allowable external pressure 69 kPa (10 psi)

versus 21 kPa (3 psi), the much greater distance of 7.2 km (4.5 mi) versus 0.8 km (0.5 mi), and the

smaller pipe ID (12-inch versus 21-inch) for the present case, the annual probability of accedence should

be significantly lower, perhaps by two to three orders of magnitude.

Reference A.22 does not address pipelines because of uncertainty associated with the formation of

combustible gas clouds. However, for comparison to Event A, the equivalent of 6 river vessels

[(27,215,542 kg (60,000,000 lb cargo)] simultaneously exploding at 7.2 km (4.5 mi) would produce a

transient pressure of 6.9 kPa (1 psi) on the storage cask, based on Equation 1 of Reference A.22.

A.5 Conclusions

The events analyzed for the transfer cask are presented in response to Question 1 through 4. The

conclusion reached from the analyses is that the strains in the MPC shell are well below the failure strain

of the material. The MPC may yield and undergo some plastic deformation depending on the drop

height, however, the MPC shell is not expected to rupture.

The event of MPC transfer from the transfer overpack to the storage overpack is addressed in response to

Question 5. The results of the analysis demonstrate that the strains in the MPC are well below the failure

strain of the shell material. The probability of the MPC shell rupture is discussed in Appendix B.

The events analyzed for the storage cask are presented in response to Questions 6 through 12. The

analyses demonstrate that the storage cask will not slide significantly nor tip-over. For the Gulfstream IV

impact analysis (Question 9) this conclusion is valid for an aircraft velocity less that 250 mph assuming

worst case impact parameters. Furthermore, there will be no penetration of the storage overpack

attributable to the Gulfstream IV impact or tornado missiles.

A-74
A.6 References

A. 1. NRC Docket No. 72-1014, "Topical Safety Analysis Report for the Holtec International Storage

and Transfer Operation Reinforced Concrete Module Cask System (HI-STORM 100 Cask

System)," Holtec Report HI-951312, Volumes I and II, Revision 8, June 25, 1999.

A.2. LS-DYNA Computer Code and associated manuals, Version 960, March 2001, Livermore

Software Technology Corporation.

A.3. American Concrete Institute, "Code Requirements for Nuclear Safety Related Concrete

Structures (ACI 349)."

A.4. LS-DYNA Manual, Version 970, Rev. 3858, LSTC, April 2003.

A.5. ASME Code, Section 111, 1989, American Society of Mechanical Engineers.

A.6. Holt, J., H. Mindlin, 0. C, Structural Alloys Handbook, CINDAS/Purdue University Press, 1996.

7. NRC Docket No. 72-1014, "HI-STORM Final Safety Analysis Report," Holtec International

Report HI-2002444, Revision 0.

A.8. DOE Fundamentals Handbook, Material Science, Volume 2 of 2, DOE-HDBK-1017/2-93,

January 1993, U.S. Department of Energy, Washington, DC.

A.9. LLNL report, Dynamic Impact Effects on Spent Fuel Assemblies, by Ramsey Chun et al.,

Lawrence Livermore National Laboratory, Livermore, California, October 1987.

A. 10. TrueGrid Manual, Version 2.1.0, XYZ Scientific Applications, Inc., September 7, 2001.

A. 11. NUREG/CR-6608, Summary and Evaluation of Low-Velocity Impact Tests of Solid Steel Billet

Onto Concrete Pads, by LLNL, February 1998.

A. 12. U.S. Nuclear Regulatory Commission, Certificate of Compliance No. 1014, Appendix B,

"Approved Contents and Designed features for the HI-STORM 100 Cask System."

A. 13. Cook, R., and R. Klingner, "Behavior of Ductile Multiple-Anchor Steel to Concrete Connections

with Surface-Mounted Baseplates," ACI Publication SP-130, 1999.

A.14. ABAQUS/Explicit: User's Manual, Version 5.8-19, Hibbitt, Karlsson, and Sorensen, Inc.,

Pawtucket, Rhode Island, 1998.

A.15. Idriss, I.M., and H.B. Seed, "Seismic Response of Horizontal Layers," Journal of the Soil

Mechanics and Foundation Division, ASCE, Vol. 94, No. SM 4, July 1968.

A. 16. Schnabel, P.B., J. Lysmer, and H.B. Seed, "SHAKE: A Computer Program for Earthquake

Response Analysis of Horizontally Layered Sites," Report No. UCB/EERC-72/12, Earthquake

Engineering Research Center, University of California, Berkeley, California, December 1972.

A-75
A. 17. U.S. Army Corps of Engineers, Engineer Technical Letter No. 1110-2-339, March 1993.

A. 18. U.S. Nuclear Regulatory Commission, NUREG-0800, "Standard Review Plan," Section 3.5.1.4,

"Missiles Generated by Natural Phenomena," Rev. 2, July 1981.

A. 19. DOE Standard 3014-96, "Accident Analysis for Aircraft Crash into Hazardous Facilities,"

October 1996.

A.20. U.S. Nuclear Regulatory Commission, NUREG/CR-5042, "Evaluation of External Hazards to

Nuclear Power Plants in the United States," December, 1987.

A.21. NRC Docket No. 72-1008, "HI-STAR 100 Final Safety Analysis Report," Holtec Report HI2012610, Rev. 0.

A.22. U.S. Nuclear Regulatory Commission, Regulatory Guide 1.91, "Evaluation of Explosions

Postulated to Occur on Transportation Routes Near Nuclear Power Plants," Revision 1, February

1978 (For Comment).

A.23. Mechanics Research Inc., Report No. MRI-2687-75, "Nuclear Power Plant Risks from a Natural

Gas Pipeline," prepared for Tennessee Valley Authority, Knoxville, Tennessee, August 2, 1974.

A.24. Braverman, J.I., J. Nie, R.J. Morante, N. Simos, and C.H. Hofmayer, "Accidental Drop Analyses

of a Spent Fuel Dry Cask System," Brookhaven National Laboratory, BNL-Y6423-032905,

March 2005.

A.25. Bjorkman, G.S., "The Buckling of Fuel Rods in Transportation Casks under Hypothetical

Accident Conditions," Proceedings of PATAM 2004, Berlin, Germany, 2004.

A.26. U.S. Nuclear Regulatory Commission, "Buckling of Irradiated Fuel under Bottom End Drop

Conditions," ISG-12, Revision 1, 1999.

A-76
APPENDIX B

CASK RESPONSE TO MECHANICAL AND THERMAL LOADS

APPENDIX B

CASK RESPONSE TO MECHANICAL AND THERMAL LOADS

BA Purpose and Scope

This appendix provides a more detailed analysis and discussion of the response of the multi-purpose

canister (MPC) to the mechanical and thermal load events that could occur as a result of the sequence of

actions that occur during the handling and transfer process, and conditions that could be experienced

during storage. Toward that end, this appendix provides information in addition to that presented in the

main body of the probabilistic risk assessment (PRA) report.

The scope of this appendix encompasses the range of possible responses to mechanical and thermal load

scenarios and events that could pose challenges to the MPC. It is assumed that the MPC is built as designed

(so fabrication errors are not included in the scope), and that materials are procuredand components

constructed to the quality standards set forth in the safety analysis report (SAR, Reference B. 1) and the

American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel (B&PV) Code.

Although weld metal typically has a higher strength than base metal, it is generally known that

weld-deposited austenitic stainless steels, while very ductile, have a tendency to be less ductile than the

wrought product. At the strain levels that may be encountered in beyond-design-basis events, such as

severe drop impacts, this potential reduction in ductility of the deposited weld metal is important and,

therefore, is considered in evaluating the structural integrity of the MPC.

B.1.1 MPC Description

In the HI-STORM dry storage cask system, the MPC acts as the confinement boundary. A drawing of the

major welds of the MPC and its components is shown in Figure B.1. The MPC is made entirely of

austenitic stainless steel. The cylindrical shell is constructed from one circumferential and four axial

seam welds that join the 1.27-cm (0.5-inch) thick plates. The two axial welds on each side of the MPC

are slightly offset at the circumferential weld. This cylindrical shell is then welded to a 6.35-cm

(2.5-inch) thick baseplate. The MPC seam welds and shell-to-baseplate weld are then subjected to a

hydrostatic test of 0.86 MPa (125 psi). The welds are fabricated in accordance with the specifications of

ASME Code, Section UI, Subsection NB. These are full-penetration submerged arc welds, which

undergo dye-penetrant (PT), radiographic (RT), or ultrasonic (UT) examinations in accordance with

Section V of the ASME Code, (Articles 2, 6, and 5 respectively). The acceptance criteria for the welds

are those in Articles NB-5320, NB-5350, and NB-5360 for PT, RT, and UT examinations, respectively.

They are not post-weld heat-treated to remove residual stresses.

After the cask is loaded with fuel, the MPC lid is welded to the shell. This 1.9-cm (0.75-inch) thick weld

undergoes PT examinations after the root, intermediate and final weld passes. These lid welds are

performed using the tungsten-inert-gas (TIG) process.

The vent and drain ports used for drying and inerting are welded closed with cover plates. These welds

undergo a final PT examination and leak test. The MPC lid-to-shell weld and the vent and drain cover

plate welds form the primary seal to the MPC lid. A closure ring is then placed on the MPC lid and

welded to the shell and lid to form a redundant cask seal. These vent and drain cover plate welds and the

ring welds are also made using the TIG process. This design ensures the cask environment is separated

from the outside atmosphere by two independent lid welds.

B-1
Closure rinQ / shell weld

Lid/shell weld

"',,,Closure rino.

Port CoverMPC lidS I MPC

A x ial . weld

I",Vent

Circumfiential Drai , weld

PMPC

MP Shell

Axial 

wel I 

5

d~ Water 7

Ie\-eI

W e I d iLIII

Baseplate/shell weld

Figure B.1. Major Welds of the MPC

The MPC lid welds were not considered in the MPC failure assessments for the following reasons:

(1) The TIG process used on the lid generally produces a tougher weld than the SA weld used in the

shell of the MPC.

(2) The redundancy of the lid design requires at least two welds to fail in order for MPC

confinement to be compromised.

(3) The applied stresses at these locations are significantly less than the stresses in the MPC shell for

all events analyzed.

B-2
B.1.2 MPC Stresses

Mechanical Loads

In the PRA, several mechanical impact accident scenarios were considered:

(1) Transfer Cask Drop onto Concrete Floor. This is a handling accident in which the transfer

overpack (HI-TRAC) with the MPC inside is dropped on a concrete floor from various heights.

(2) Transfer Cask Drop onto Storage Cask. This is a handling accident in which the transfer

overpack (HI-TRAC) with the MPC inside is being lowered from a height of up to 24.4 meters

(80 feet) and fails on the storage overpack.

(3) MPC Drop. This is also a handling accident in which the MPC falls from a height of 5.8 meters

(19 feet) when being lowered from the transfer overpack (HI-TRAC) to the storage overpack

(HI-STORM).

(4) Storage Cask Drop. This is a transport accident in which the storage overpack (HI-STORM)

with the MPC inside is dropped from a height of 0.30 meter (1 foot) onto concrete, gravel, and

asphalt.

(5) Storage Cask Tip-Over. This is a storage accident in which the storage cask with the MPC inside

is tipped over on the concrete storage pad.

The analyses used to calculate the MPC stresses for these events are described in Appendix A to this

report.

The thermal analyses described in Section 4.2 indicated that steady-state conditions within the MPC

under normal conditions would be attained prior to the initiation of the various handling and transfer

events that move fuel out to the storage pad. These steady-state conditions are an internal pressure of

0.564 MPa (82 psi) with an MPC temperature that varies axially from 73 'C (163 'F) to 180 'C (356 'F).

It was assumed that the MPC is under normal steady-state conditions during the mechanical impact

scenarios listed above. In addition, all of the fuel is assumed to be intact during these events. Fuel

cladding failure would cause the internal pressure of the MPC to increase. If a mechanical impact event

did cause fuel cladding failure, the resulting increase in pressure would not occur during the accident

scenario. The increase in internal pressure attributable to fuel cladding failure would be a result of the

mechanical loading; however, because of the extremely short duration of the impact, it would not affect

the MPC internal pressure until after the mechanical loading is complete.

Thermal Loads

Only two thermal accident scenarios are considered in this PRA:

(1) Storage Cask External Fire. This is a storage accident in which the storage overpack

(HI-STORM) with the MPC inside is exposed to an external fire resulting from a Gulfstream IV

aircraft crash. A credible fire scenario will be less than 30 minutes. However, the analysis was

performed for a fire duration of 3 hours.

(2) Storage Cask Vent Blockage. This is a storage accident in which the cooling vents of the storage

overpack (HI-STORM) become completely blocked for a period of 20 years.

B-3
Under thermal loads, the only significant MPC stresses are attributable to internal pressure. Membrane

stresses were calculated using the classical shell theory equations for pressurized cylinders:

Per OCaxial 2- (Bt1 2.t

Por (B.2) C'hoop t

t

where:- axia -= axial stress in psi

Ohoop hoop stress in psi

P = internal pressure in psi

r = MPC internal radius in inches (= 33.68.75 in. or 85.5663 cm)

t = MPC wall thickness in inches (= 0.5 in, or 1.27 cm)

The initial steady-state pressure of the 300.

MPC is 0.565 MPa (82 psi). Using

Equations B. 1 and B.2, the resulting

axial and hoop stresses are 19.05 MPa 250

(2,762 psi) and 38.1 MPa (5,525 psi),

respectively. 200+

The highest stress attributable to -- *--Hoop stress (membrane)

internal pressure loads occurs at the 150- A Axial Stress (membrane)P

shell-to-baseplate connection. U-Shell to BP Weld Stress (bending)

Section 3.4.4.3.1.2 of Reference B.1 • 100.

analyzed. the MPC stresses at various

locations attributable to an internal

pressure of 0.69 MPa (100 psi). 50.* * 9 9

These results showed that the total - A - ; -

(membrane plus bending) stress at 0 : " • . : . : ....

the MPC shell-to-baseplate connection 0.50 0.55 0.60 0.65 0.70

was 303 MPa (43,986 psi). The axial Pressure (MPa)

membrane stress at this pressure is

23.6 MPa (3,419 psi). The difference Figure B.2. MPG Stresses as a Function of Internal Pressure

in these values gives the bending stress acting on the shell-to-baseplate weld as 280 MPa (40,567 psi).

The total stress at other pressures was assumed to be directly proportional to those at 0.69 MPa (100 psi).

Therefore, at a pressure of 0.565 MPa (82 psi), the total stress at this location is only 248.7 MPa

(36,068 psi). Subtracting the axial membrane stress of 19.05 MPa (2,762 psi) gives a bending stress of

229.6 MPa (33,306 psi).

B4
The thermal accident scenarios listed above cause the MPC temperature and internal pressure to increase.

The thermal analyses used to calculate the new MPC temperatures and pressures are described in

Section 4.2.1.1, "Heatup Model," of the Dry Cask Storage PRA report. Figure B.2 shows the increase

in axial membrane, hoop membrane, and shell-to-baseplate bending stress with increasing pressure.

B.1.3 MPC Failure Mechanisms

The following failure mechanisms were postulated for the MPC under mechanical and thermal loads:

(1) weld fracture

(2) exceeding the limit load

(3) creep rupture

All three failure mechanisms were considered for thermal events, while only weld fracture was

postulated for mechanical impacts. For the purposes of this PRA, failure mechanisms are assumed to be

independent. The probability of MPC failure is given as the sum of the failure probabilities from each

failure mechanism.

The failure analysis for weld fracture was based on test data of Type 308 weld deposited stainless steel

specimens, and specimens taken from the weldments of process water piping of nuclear production

reactors constructed in the 1950s at the Savannah River Site. The mean and standard deviation of the

failure strain data was calculated and adjusted for the effects of strain rate, temperature and state of

stress. These results were used directly to determine the probability of weld fracture given the maximum

strain calculated for each impact event. (Note that to simplify the evaluation, the adjustment for state of

stress was applied to the calculated strain, rather than the failure strain.)

Failure analyses for limit load and creep rupture were conducted differently from the failure analysis for

weld fracture attributable to mechanical loads. Failure models 'for limit load and creep rupture were

created in a spreadsheet with an add-on module called @RISK® (from www.palisade.com). This module

allows Monte Carlo simulations to be run in the spreadsheet by allowing values in the spreadsheet cells

to be sampled from statistical distributions. The user specifies which cells are considered outputs, and

the number of iterations to perform. For each iteration, the program samples values from all cells in

which a statistical distribution is specified, recalculates the entire spreadsheet, and records the values in

the specified output cells. The distribution of output values can then be plotted and analyzed.

Separate spreadsheets were developed for the limit load and creep rupture mechanisms. The entire MPC

was modeled through a series of spreadsheets representing the different parts of the MPC. Input cells for

the spreadsheets defined the material properties, flaw parameters, and applied stresses. The only inputs

not sampled from statistical distributions were the applied stresses on the MPC. The statistical

distributions used to describe the material properties and flaw parameters are described in Section B. 1.5.

The specific inputs and calculations performed in each failure model spreadsheet are described in detail

in Sections B. 1.6 and B. 1.7. The spreadsheet cell specified as output identifies each iteration as either an

MPC success or failure. This process is repeated many times in a Monte Carlo simulation. The

probability of failure was calculated by dividing the total number of MPC failures by the total number of

iterations performed.

B-5
The failure models used in this PRA do not attempt to predict the hole size in the MPC for a given failure

probability. Any failure is assumed to create a hole in the MPC large enough to allow the particles and

gases within the MPC to be released. This assumption was made because of the difficulty inherent in

trying to predict the equivalent hole size produced from a particular MPC failure mode.

B.1.4 Weld Fracture Attributable to Mechanical Loads

Appendix A evaluated the structural response of the MPC for various end drop scenarios. The objective

was to determine the maximum effective plastic strain (EPS) on the MPC confinement boundary for each

of these scenarios. The analyses showed that the most highly stressed regions of the MPC are near the

base of the cylindrical shell and in the weld joining the shell to the 2.5-inch-thick baseplate.

To determine if confinement boundary integrity is compromised, the maximum EPS in the MPC must be

compared to an appropriate strain failure criterion. For a valid comparison, the conditions under which

the maximum EPS is calculated, and the conditions under which the failure strain limit is measured,

should be consistent. The comparison, therefore, must account for how the strain is measured, as well as

the effects of strain rate, temperature and state of stress. It is the objective of this section to establish a

valid basis for this comparison, and to estimate the probability of exceeding the failure strain limit for a

given drop event.

Strain Measure

The LS-DYNA computer program used in the analysis for Appendix A to calculate MPC stresses and

strains considers the reduction in cross-sectional area of the finite elements in the computation of stress

within the element. Therefore, the program expects the user to input a true stress strain curve, which, by

definition, considers the reduction in area in the computation of stress. Since the calculation of EPS

within LS-DYNA is based on true strain, the failure criterion should also be based on a true strain

measure.

The true strain at failure is calculated from the reduction in area (RA) within the necked-down region of

a failed specimen as given in the following equation:

Et (B.3)

where, Ft = true strain of failure, and

RA = reduction in area

The MPC shell is fabricated from Type 304 stainless steel. (See the discussion in Section A.4.3.) Based

on a review of References B. 15 and B. 16, a typical RA of Type 304 stainless steel at room temperature

and static loading is 70%. Substituting RA equal to 0.70 into Equation B.3, the true strain at failure is

1.20 in/in or 120% strain. This value is consistent with typical true stress strain curves that can be found

in Reference B. 17 for Type 304 stainless steel.

B-6
Weld Metal Failure Data

In the analyses performed in Appendix A, no distinction is made between the base metal and weld metal

in the MPC shell. Procurement of the weld material, which is typically Type 308 stainless steel, and

fabrication of the MPC in accordance with the requirements of the ASME Code ensures that the weld

strength is equal to or greater than the base metal. However, it is generally known that weld-deposited

austenitic stainless steels (Type 308), although very ductile, have a tendency to be less ductile than the

wrought (Type 304) product. At the high strain levels that may be encountered in beyond-design-basis

events, such as severe drop impacts, this potential reduction in ductility of the deposited weld metal can

be important to the structural integrity of the MPC shell.

In Reference B. 18, 24 Type 308 weld-deposited metal specimens cut from fabricated weldments were

statically tested at room temperature. The specimens were taken from both transverse and longitudinal

orientations within the weldments. The mean value of the RA was 59.7% with a standard deviation of

9.1%. Using Equation B.3, this converts to a mean true strain at failure of 0.91 in/in. The mean true

strain at failure minus one and two standard deviations are 0.71 in/in and 0.54 in/in, respectively.

In Reference B. 19, the mechanical properties of archival process water system (primary coolant) piping

and weld materials having approximately 6 years of service were measured. The process water system

piping of the nuclear production reactors constructed in the 1950s at the Savannah River Site was

composed primarily of Type 304 stainless steel with Type 308 stainless steel weld filler. Tensile

properties were measured for base metal, weld metal and heat-affected zone (HAZ) materials. The test

specimens represented both ASTM L-C and C-L orientations to allow comparison of the mechanical

response for the cases of flaws oriented parallel and perpendicular to the pipe axis or rollingdirection.

The tensile properties of the archival piping material were found to be typical of recently produced

commercial melts of Type 304 stainless steel piping. The values of the mean RA and standard deviation

of the base metal, weld metal and HAZ material for static loading at room temperature are shown in

Table B. 1.

Table B.1. Reduction in Area Mean and Standard Deviation Test Results from Reference B.19

Stainless Steel Material Mean Reduction in Area (RA) Standard Deviation

Type 304 Base Metal 72% 9.0%

Type 308 Weld Metal 61% 10.0%

Heat-Affected Zone (HAZ) 71% 1.0%

The mean RA for the base metal of 72% in Table B. 1 compares well to the typical values of 70% and

77% from References B.15 and B.16 respectively. Also, the mean RA for the weld metal of 61% and

standard deviation of 10% compare well to the values of 59.7% and 9.1% cited in Reference B.18,

discussed above. This data clearly shows that the 308 stainless steel weld metal is very ductile, although

its ductility is slightly less than the ductility of the Type 304 base metal.

B-7
The RA data for Type 308 stainless steel weld metal at room temperature and static loading used in the

evaluation herein includes all of the data in References B. 18 and B. 19 discussed above, as well as

dynamic test data at low strain rates (0.05/sec) from Reference B. 19. Combining this data of 37 tests

yields a mean RA of 59.0% and a standard deviation of 9.7%. These values establish the basis for

determining the probability of weld failure and assumes that the weld failure data is normally distributed.

Temperature and Strain Rate Effects

In all drop scenarios evaluated in Appendix A, the most highly stressed region of the MPC occurs within

the bottom foot of the shell. (See Appendix A.) While the MPC is being moved in the HI-TRAC

transfer cask, the temperature profile of the MPC shell, varies from 302 OF at the bottom of the MPC to

445 °F at the top (Reference B.1, Holtec HI-STORM FSAR, Table 4.5.4). The test data used to calculate

the mean RA and standard deviation, discussed above, is for static loading at room temperature. Typical

RA values for Type 308 weld metal as a function of elevated temperature are reported in Figure 7 of

Reference B. 18. The trend curve in the figure shows that at 302 OF, the temperature at the bottom of the

MPC shell, the ratio of RA at temperature to RA at room temperature is 0.90. This factor is be used to

adjust the RA room temperature values to those that would be expected at the higher temperature at the

bottom of the MPC.

Strain rate can also have an effect on the ductility of stainless steel weldments. Typical strain rates for

the drop impact scenarios considered in Appendix A are on the order of 100/sec. Reference B.20 shows

that for strain rates of 1200/sec the ductility of stainless steel weldments deceases to about 88% of the

static value. For a strain rate of 100/sec for Type 304 base metal, Reference B. 15 shows that at 1,000 OF

RA is approximately 7% greater than the value at room temperature. These results are insufficient to

draw any firm conclusions, but in the interest of conservatism, a strain rate reduction factor of 0.98 was

selected based on the information in Reference B.20.

To account for both strain rate and temperature effects, the mean RA of 59.0% for weld metal was

reduced by the factor 0.88 (0.90 x 0.98 = 0.88). This results in a mean RA of 52%, accounting for both

strain rate and temperature. The standard deviation is assumed to remain the same at 9.7%. From

Equation B.3, the true strain at failure of stainless steel weld metal in a uniaxial tension test, considering

the effects of strain rate and temperature, can be estimated to be 0.73 in/in with a mean minus one and

two standard deviations of 0.55 in/in and 0.40 in/in respectively. Table B.2 below shows the mean true

strain at failure minus several standard deviations, along with the probability that the weld metal's true

strain at failure is less than the tabulated value.

The probability of failure of stainless steel weld metal at the plastic strains given in Table B.2 are

approximate, since the amount of test data capturing true strain at failure (or RA) of stainless steel

weldments at elevated temperature and high strain rates is limited. Nonetheless, Table B.2 provides a

reasonable estimate of weld failure probability in uniaxial tension at the strain rates and temperatures

within the most highly stressed region of the MPC.

B-8
Table B.2. Mean True Strain at Failure minus Several Standard Deviations, and the Probability

That the Weld Metal's True Strain at Failure Is less than the Tabulated Value

Standard Deviation True Strain at Failure (TSF) Probability the TSF is Less than

from the Mean (in/in) the Tabulated Value [x < TSF]

0.0 (Mean) 0.73 0.5

0.5 0.64 0.3085

1 0.55 0.1587

1.5 0.47 0.0668

2 0.4 0.0228

2.5 0.32 0.0062

3 0.26 0.0013

3.5 0.2 0.00023

4 0.14 0.000032

4.5 0.087 0.000003

5 0.036 < 0.000001

Effect of State of Stress

For each drop scenario, the maximum EPS in the MPC shell was calculated in Appendix A. However,

cracking of a stainless steel weldment cannot be determined by simply comparing the calculated

maximum EPS to the true strain at failure even when temperature and strain rate have been considered.

This is because the EPS was calculated from a complex three-dimensional state of stress, while the true

strain at failure was derived from a one-dimensional state of stress. Unlike a one-dimensional state of

stress, a three-dimensional state of stress may constrain plastic flow in the material and lower the EPS at

which failure occurs. This loss of ductility can be accounted for by the use of a triaxiality factor (TF),

which is the ratio of normal stress to shear stress on the octahedral plane, normalized to unity for simple

tension, as given by Equation B.4 (References B.21 and B.22):

Triaxiality Factor = "1 + U2 + C3

[C - C + (2 - C 3 )2 + - (B.4)

where, oa = maximum principal stress

02 = intermediate principal stress

T3 = minimum principal stress

B-9
The reduction in ductility associated with the triaxial stress state is approximately equal to the true strain

at failure in uniaxial tension divided by the TF. More precisely, the ratio of EPS at failure to the true

strain at failure in uniaxial tension, which is called the ductility ratio (DR), is given by Equation B.5

(Reference B.23):

DR = 2 (1-TF) (B.5)

where, DR = the ductility ratio

TF = the triaxiality factor

Thus, to determine the actual strain at which weld cracking will initiate requires that either the calculated

EPS or the true strain at failure be adjusted to account for the triaxial nature of the state of stress. [Note

that TF = 1.0 represents uniaxial tension. TFs greater than 1.0 are produced by states of stress that

constrain plastic flow, reduce ductility and result in a higher likelihood of failure for a given EPS. TFs

less than 1.0 are produced by states of stress that enhance plastic flow, increase ductility and result in a

lower likelihood of failure for a given EPS.]

Probability of Weld Failure

The probability that the true strain at which a weld fails is less than the calculated maximum EPS cannot

be determined directly from Table B.2. To use Table B.2 to determine the probability of weld failure, the

calculated maximum EPS must first be divided by the DR to adjust it for the effects of the triaxial stress

state at the location where the EPS occurs. However, the location (element) where the maximum EPS

occurs may not be the location most susceptible to failure. The location where the probability of failure

is greatest is where the quotient of EPS divided by the DR results in the highest value. For example, in

the case of the 100-foot drop of the transfer cask onto the concrete floor a maximum EPS of 0.256 in/in

occurred in element 15997 on the inside surface of the MPC shell beneath the basket support, as shown

in Figures A. 13 and A. 14. For this element, the TF is negative, indicating a very favorable state of stress,

and the EPS adjusted for triaxiality is 0.13 in/in. For this lower EPS, the probability of failure decreases

considerably. Thus, the element with the calculated maximum EPS does not necessarily control failure.

To determine the location where the maximum EPS adjusted for triaxiality occurs, a set of elements with

high values of EPS and attributes likely to result in a high TF were selected for evaluation. For these

elements, EPS, TF, DR, and adjusted EPS were calculated at each output time. This was done for the

30.5, 21.3, and 12.2 meter (100, 70, and 40 foot) drops of the transfer cask onto the concrete floor and for

the 5.8 meters (19 foot) drop of the MPC into the storage overpack. For the 1.52 meter (5 foot) drop of

the transfer cask onto the concrete floor and for drops of the transfer cask onto the storage overpack, the

maximum EPS together with the maximum TF for a biaxial stress state of 2.0 were used to conservatively

calculate the maximum EPS adjusted for triaxiality.

B-10
Transfer Cask and MPC Drop Events

For drops of the transfer cask falling onto the concrete floor, Table B.3 gives the calculated maximum

EPS, the maximum EPS adjusted for triaxiality, and the probability of weld crack initiation. It is

important to emphasize that the probability of weld crack initiation is calculated using the maximum EPS

adjusted for triaxiality when using Table B.2. Table B.4 provides the same information for the drop of

the transfer cask onto the storage overpack, and Table B.5 shows the results for the 19-foot drop of the

MPC into the storage overpack.

Table B.3. Probability of MPC Weld Crack Initiation from the Drop of the Transfer Cask

onto the Concrete Floor

Drop Height Maximum EPS Maximum EPS Probability of Weld Crack

(ft) (in/in) Adjusted for Triaxiality Initiation (1)

(in/in)

5 0.024 0.048 (2) < 0.000001

40 0.195 0.213 0.00036

70 0.24 0.285 0.0026

100 0.256 0.385 0.0196

(1) The term "weld crack initiation" or "cracking of the MPC weld" is used in place of "weld failure" or

"failure of the MPC weld" to distinguish between a "crack" in the weld, that may not propagate through the

entire thickness, and "failure" of the weld, which would imply a through thickness rupture of the weld and

breach of the confinement boundary. When developing a weld failure criterion using uniaxial tension tests

the terms failure, rupture and breach are synonymous. However, when the maximum EPS occurs on the

surface of a region of high bending curvature, as shown in Figures A. 14 and A.29, the initiation of a

failure- a crack - does not necessarily result in a rupture or breach. Thus, some caution is required when

using the word "failure" to describe what happens when a failure limit is reached at a point on the surface of

a region with high curvature (strain gradient), especially when the energy input is limited.

(2) A maximum TF of 2.0 was used.

Table B.4. Probability of Weld Crack Initiation

for the Drop of the Transfer Cask onto the Storage Overpack

Drop Height Maximum BPS Maximum BPS Probability of Weld (fi) at Element 9421 (1) Adjusted for Triaxiality (2) Crack Initiation

(in/in) (in/in)

5 0.0363 0.0726 0.000002

40 0.0601 0.1202 0.000014

80 0.0976 0.1952 0.000203

(1)

(2)

See Table A.4, Note 2.

A maximum TF of 2.0 was used.

B-1I
Table B.5. Probability of Weld Crack Initiation

for the 19-Foot Drop of the MPC into the Storage Over ack

Maximum EPS Drop Height Maximum EPS Adjusted for Triaxiality Probability of Weld

(fi) (in/in) (in/in) Crack Initiation (1)

19 0.456 0.623 0.282

(1) See Note 1 from Table B.3 and discussion below related to tearing of the basket support fillet welds for the

19-foot drop.

The 19-foot drop of the MPC into the storage overpack results in the highest probability (0.282) that a

weld on the MPC confinement boundary will crack on impact. This result, as well as the results for the

other drop events, must be viewed as a very conservative estimate of the probability of MPC confinement

boundary breach or rupture. There are two reasons for this:

(1) The probability of weld crack initiation is the probability that the most highly stressed element of

the six elements through the thickness of the MPC shell will fail. Because this occurs in a region

of high curvature in a highly indeterminate structure, the failure of one element on the surface of

the shell will result in the redistribution of internal forces to the stiffer (in bending) adjacent

elements on the surface, with little of the internal forces redistributing to the lower stressed

interior elements'through the thickness. This makes it more likely that additional surface

elements will fail, and less likely for the failure to propagate through the thickness. In addition,

the process itself is contained, since the total kinetic energy is limited, as in all impact events.

(2) The maximum EPS of 0.456 in/in and the adjusted maximum EPS of 0.623 in/in result from

constraint imposed on the MPC shell. This constraint is created by the discontinuous basket

support "stiffeners" that terminate 3.8 cm (1.5 inch) above the MPC baseplate. The constraint on

the shell imposed by the support is maintained by the 2.5 cm (1-inch) long 3.2 mm (1

/8-inch) fillet

weld that attaches the bottom portion of the support to the shell. Figure B.3 shows the

deformation and contours of EPS in this region, and Figure B.4 shows the shear deformation and

the very evident shear-lag in the elements of the support at the location of the fillet weld between

Nodes 18469 and 21444. (This basket support is the same one shown in Figure A.26.) In the

likely event that this weld tears during impact, the local constraint on the shell is relieved and the

shell will behave locally as though it were unstiffened. A published analysis (Reference B.24) of

a 7.6 meters (25 foot) drop of the MPC onto a rigid surface, in which the basket supports were

not included, produced a maximum EPS of 0.21 in/ in, which is a significant reduction from the

value of 0.456 in/in when the attachment welds remain intact. Thus, the potential tearing of

these small attachment welds can significantly reduce the probability that a weld on the

confinement boundary may crack.

B-12
03-n3-U) MPC 1 q* DROF'

lime - 0.0?3 f ringe Levels

Contours of Effective PIwAic

max i 1. value 4"Me-01

minT at elernlt 3880"')

at el(- 4.134e 01

3.674e 01

IN lie-01

2.75be-01

2.297c-Bl

1.837e-01

1.378e-01

9. 1 f3be-02

4.593e-02

0.000ei 00

MAL,

X

71 1

Figure B.3. Deformation and Contours of EPS

in the Vicinity of the Stitch Weld Attaching the Basket Support to the Shell

B-13
Figure B.4. Shear Deformation and Shear-Lag Are Evident in the Elements of the Support

at the Location of the Fillet Weld Between Nodes 18469 and 21444

Storage Cask Tip-Over Event

The storage cask tip-over event is discussed in Section A.4.6. The maximum stresses in the MPC for this

event are given in Table A.6 and come from the HI-STORM 100 SAR (Reference B. 1). The maximum

stress (stress intensity) is approximately 406.8 MPa (59,000 psi), which is less than the allowable stress

intensity for primary membrane plus bending of 449.5 MPa (65,200 psi) for this service level. For the

purpose of evaluating weld fracture, the allowable stress of 449.5 MPa (65,200 psi) is used herein. In

accordance with the ASME B&PV Code, the maximum stress would have been calculated from a linearelastic analysis, and in this case the analysis produced stresses that exceed the yield stress of 262.0 MPa

(38,000 psi) at temperature (Table A.2). Therefore, the maximum stress of 449.5 MPa (65,200 psi)

cannot be converted to a strain simply by dividing by the elastic modulus.

B-14
Since the weld failure criterion is based on plastic strain, the elastically calculated maximum stress must

be converted to a plastic strain that could have reasonably resulted from a non-linear analysis of the same

event. Also, since the "damage" that can be inflicted on the MPC during a tip-over event is energylimited, an energy balance approach can be used to estimate the maximum plastic strain. In this

approach, the strain energy per unit volume absorbed at the point of the elastically calculated maximum

stress is equated to the strain energy per unit volume absorbed through elastic-plastic deformation of the

material using the idealized engineering stress-strain curve in the LS-DYNA program that was used for

the drop analyses.

Using this approach and the data in Table A.2, the calculated maximum plastic strain is 0.0031 in/in.

Since the maximum membrane plus bending stress must have occurred on the boundary, the maximum

value of the TF is 2.0. Therefore, the maximum adjusted plastic strain is 0.0062 in/in. Assuming that the

maximum plastic strain occurs at one of the axial or circumferential welds, the probability of weld .crack

initiation from Table B.2 is less than 0.000001. This result is summarized in Table B.6.

Table B.6. Probability of Weld Crack Initiation from Storage Cask Tip-Over

Estimated Maximum Maximum Plastic Strain Probability of Weld Crack

Plastic Strain Adjusted for Triaxiality Initiation (1)

(in/in) (in/in) Initiation_(1)

Storage Cask

Tip-Over 0.0031 0.0062 < 0.000001

Event

(1) See Note 1 from Table B.3.

B.1.5 Failure Analysis Inputs for Limit Load and Creep Rupture

In order to determine the probability of MPC failure for the limit load and creep rupture, distributions of

the following material properties are needed at various temperatures:

(1) yield strength

(2) tensile strength

(3) creep rupture strength

In addition to the material properties, stress concentration factors are also needed.

The MPC may be constructed from several different types of austenitic stainless steels. In general, the

tensile and creep rupture properties of Type 304 stainless steel tend to be more limiting regarding MPC

failure than other types of austenitic steels. Therefore, yield, tensile and creep rupture strengths used in

the failure analyses are based on Type 304 stainless steel. Comparisons between the material properties

of Type 304 and 316 are given in the sections that follow.

B-15
Yield and Tensile Strength

The yield and tensile strength of austenitic stainless

steel decreases as the material temperature increases.

Figure B.5 compares representative yield and tensile ,-"-. Y i

strengths for Type 304 and 316 stainless steel over a S0o0 M -

wide range of temperatures (References B.2 and B.3). .M, z 

Even though the required minimum room temperature a

yield and tensile strengths for these two materials are 300

identical, this figure shows that typically Type 316 has

higher yield and tensile strengths between 24 'C

(75 -F) and 816 °C (1,500 'F). Therefore, the tensile -A

properties used in the failure analysis of the MPC are

for Type 304 stainless steel. 0

0 200 400 600 800 1000

Table B.7 describes the reduction in yield and tensile Temperature (C)

strength of Type 304 stainless steel with temperature as Figure B.5. Yield and Tensile Strength

a fraction of the yield and tensile strength at room Behavior of Type 304 and 316 Austenitic

temperature (Reference B.3). This same reference cites Stainless Steels at Elevated Temperatures

mean room temperature yield and tensile strengths for (References B.2-B.3)

this material as 254 MPa (36,900 psi) and 578 MPa

(83,900 psi), respectively. The standard deviations are 44 MPa (6,400 psi) for yield strength and 30 MPa

(4,300 psi) for tensile strength.

Best estimate yield and tensile strengths at various temperatures were determined using these reduction

ratios and room temperature properties. Table B.7 lists these best estimate values at temperatures up to

871 -C (1,600 7F).

The ASME Code places minimum allowable yield and tensile strengths for this material at temperatures

up to 427 'C (800 'F). These minimums are also shown in Table B.7. At temperatures above 427 'C

(800 'F), the minimum allowable yield and tensile strengths were detenmined by applying the reduction

ratios to the ASME minimums at room temperature.

B-16
Table B.7. Reduction in Yield and Tensile Strengths of Type 304 Stainless Steel

as a Function of Temp erature

Best Best

Yield Estimate Minimum Tensile Estimate Minimum

Strength Yield Yield Strength Tensile Tensile

Temperature Reduction Strength' Strength2 Reduction Strength3 Strength4

°C (OF) Ratio Mpa (psi) Mpa (psi) Ratio Mpa (psi) Mpa (psi)

24 (75) 1 254 (36900) 207 (30000) 1 578 (83900) 517 (75000)

38 (100) 0.96 244 (35424) 207 (30000) 0.97 561 (81383) 517 (75000)

93 (200) 0.83 211 (30627) 172 (25000) 0.86 497 (72154) 490 (71000)

149 (300) 0.75 191 (27675) 155 (22500) 0.80 463 (67120) 455 (66000)

204 (400) 0.69 176 (25461) 143 (20700) 0.78 451 (65442) 444 (64400)

260 (500) 0.65 165 (23985) 134 (19400) 0.77 445 (64603) 438 (63500)

316 (600) 0.61 155 (22509) 125 (18200) 0.77 445 (64603) 435 (63100)

371 (700) 0.59 150 (21771) 119 (17300) 0.77 445 (64603) 431 (62500)

427 (800) 0.56 142 (20664) 114 (16600) 0.76 440 (63764) 425 (61700)

482 (900) 0.54 137 (19926) 112 (16200) 0.74 428 (62086) 383 (55500)

538 (1000) 0.52 132 (19188) 108 (15600) 0.70 405 (58730) 362 (52500)

593 (1100) 0.49 125 (18081) 101 (14700) 0.63 364 (52857) 326(47250)

649(1200) 0.47 120(17343) 97(14100) 0.55 318(46145) 284(41250)

704 (1300) 0.44 112(16236) 91 (13200) 0.46 266(38594) 238 (34500)

760 (1400) 0.39 99(14391) 81(11700) 0.35 202 (29365) 181 (26250)

816 (1500) 0.31 79 (11439) 64 (9300) 0.25 145 (20975) 129 (18750)

871 (1600) 0.2 51(7380) 41 (6000)

Notes:

I1. Best estimate yield strengths based on reduction ratios and room temperature yield strength of 36,900 psi

(Reference B.3).

2. Minimum yield strengths from 24 'C (75 OF) to 427 OF (800 OF) taken from ASME Code. Minimum yield

strengths from 482 °C (900 OF) to 871 'C (1,600 OF) calculated using reduction ratios and ASME

minimum at 24 'C (75 "F).

3. Best estimate tensile strengths based on reduction ratios and room temperature tensile strength of 578 MPa

(83,900 psi) (Reference B.3).

4. Minimum tensile strengths from 24 "C (75 "F) to 427 "F (800 "F) taken from ASME Code. Minimum

tensile strengths from 482 "C (900 "F) to 871 "C (1,600 "F) calculated using reduction ratios and ASME

minimum at 24 "C (75 "F).

B-17
700

Figure B.6 shows the decrease in the best -W-Meas Tensile

estimate and minimum tensile properties of 600 -W Tesl

Type 304 stainless steel at increased M 500 Q -*"e C- -Wn Yield

temperatures. The minimum allowable tensile 400

strengths are very close to the best estimate 40

values. Above 427 'C (800 OF) this difference ' 300

increases. For the yield strengths, the a

difference between the minimum and best 200

estimate values slowly decrease at higher 100

temperatures.

0 200 400 600 800 1,000

Temperature (C)

Figure B.6. Yield and Tensile Strength Behavior of Type 304

Austenitic Stainless Steels at Elevated Temperatures

(Reference B.3)

It was assumed that both the yield and tensile properties of Type 304 stainless steel at any given

temperature could be described by a normal distribution with a mean equal to the best estimate value at

that temperature. In addition, the standard deviations defined above were assumed to apply at all

temperatures. Tensile properties at temperature between those listed in Table B.7 were calculated by

interpolation.

Type 304 stainless steel base metals are typically welded using Type 308 as the weld filler metal. The

minimum required room temperature yield strengths and tensile strengths are identical for both of these

materials. Table B.8 lists some elevated temperature tensile properties for Type 308 austenitic stainless

steel.

Table B.8. Tensile Properties for Type 308 Stainless Steel at Elevated Temperatures

Temperature Yield Strength Tensile Strength

°C (OF) MPa (Psi) MPa (Psi)

482 (900) 310 (45,000) 444 (64,400)

593(1100) 261 (37,800) 301 (43,700)

649 (1200) 222 (32,200) 239 (34,700)

Comparing Tables B.7 and B.8 shows that the expected weld material (Type 308) is expected to have

superior tensile properties to that of the base metal (Type 304). Therefore, the properties listed in Table

B.1 can conservatively be used to describe the MPC welds and plates.

B-1 8
Creep Ruture Strengh

Creep rupture strength of Type 304 stainless steel was compiled from data in the open literature

(References B.2 - B.7). It should be noted that this creep rupture data is not comprehensive. The time

and temperature conditions for this data were converted to the Larson-Miller parameter (Reference B.7):

LarsonMiller - To (C + log(t,)) (B.6)

where: T

C

tr

temperature in *R (= 'F + 459.67)

constant (18 for austenitic stainless steel)

time to rupture in hours

Figure B.7 shows the stress at creep rupture

data plotted as a function of this LarsonMiller parameter for Types 304, 316, and 308

austenitic stainless steels. For a given

Larson-Miller parameter value, Type 316

tends to have a higher creep rupture stress.

The data for Type 304 and 308 are similar.

Therefore, the creep rupture data used for the

MPC failure analysis is based upon Type 304

material properties. This data is used to

model both the plate and weld material

properties.

Figure B.8 is typically how creep rupture data

is presented. However, for the purposes of

developing a creep rupture model, the

Larson-Miller parameter is needed as a

function of the applied stress. Figure B.7

shows the Larson-Miller parameter plotted as

a function of the log of applied stress for only

Type 304 materials. The data was fit to a line

with an R2 value of 0.98. The equation of this

line has a slope of -11,032 and a y-intercept of

81,404. The best fit, 95% upper bound, and

95% lower bound trend lines are also shown in

Figure B.8.

At a given stress, there is some variability in

the Larson-Miller parameter required to cause

creep rupture of Type 304 stainless steel. It is

assumed that this variability in Larson-Miller

parameter is normally distributed about a mean

value given by the best-fit line described

1000

0~

4)

1~

U,

100

10

CXType 316

AType 304

#Type 308

x

x

x

1 +-a

25,000

: : : : . . . . .. : ..... . .... .. .. ...

35,000 45,000

Larson Miller Parameter

55,000

Figure B.7. Creep Rupture Properties of Type 304, 308, and 316

Austenitic Stainless Steels (References B.2 - B.7)

~MJ

45-

.- 40'

30~

265

-9-% Bounds

-- Best Fit Lin~e

Ys=.1 1032x+ 81404

W =0.9825

I I I I I I I

3 3.5 4

Log (stress in psi)

Figure B.8. Best Estimate and Bounding Larson-Miller

Parameters as a Function of Log (Stress)

for Type 304 Austenitic Stainless Steel

B-19
above. It is also assumed that the 95% bounds represent 2 standard deviations about this mean value.

Therefore, both the mean and standard deviation of the Larson-Miller parameter vary with the applied

stress.

Flaw Distribution

An analysis of the MPC welds was performed using the PRODIGAL computer code to obtain flaw

distributions for the failure analyses. The PRODIGAL code is described in several of the references

listed at the end of this report (References B.8-B.11). These references also describe a number of

applications of the code to nuclear pressure vessels and piping. Other references describe studies that

have collected data on observed flaws in welded construction and have compared the numbers and sizes

of the observed flaws with predictions made with the PRODIGAL code (References B. 12 and B. 13).

The PRODIGAL code is an expert system model that simulates the occurrence of flaws in nuclear grade

welding. An additional aspect of the model is the simulation of the radiography and dye-penetrant

inspections performed to detect flaws so that defective welds can be rejected and/or repaired. The focus

of PRODIGAL is on crack-like flaws that have the potential to impact structural integrity. Specific

limitations of the model are that it excludes structurally benign flaws such as pores and rounded slag.

The model does not address certain inspections such as UT and leak testing. It was designed to predict

the expected defect distributions in an average weld but does not include a consideration of uncertainties

or the expected weld-to-weld variations.

PRODIGAL outputs the size, number, and J.E0W

location of flaws per meter of weld. The total I.•o

number of flaws per meter of weld with a depth Be,, Flw

greater than a given percentage of weld

thickness is shown in Figure B.9. This figure 1.&.03

also distinguishes between surface breaking and

embedded flaws in the weld. The results show I

that MPC welds have more embedded flaws 1.E-o5

than surface breaking flaws. The largest flaw & F

predicted by PRODIGAL is approximately 80%

of the weld thickness.

There are six shell welds in each MPC. This 0 10 20 30 40 5oD s 70 s0 90

includes four axial welds, one circumferential Flaw D•h (% 6•iw.ss)

weld, and one shell-to-baseplate weld. The Figure B.9. Prodigal Flaw Distribution Results for MPC Welds:

length of two axial welds on each side of the Total Number of Flaws with Depth Greater than a Given Depth

MPC is 4.77 meters (15.65 feet). These two

axial welds are modeled as one continuous

weld extending the entire height of the MPC. The length of the circumferential and shell-to-baseplate

weld is 5.46 meters (17.9 feet) each. The total number of flaws in each of the MPC welds can be

calculated by taking the PRODIGAL results and multiplying by these weld lengths. There are

approximately 0.04 flaws per axial weld extending the entire height of the MPC. The circumferential

and shell-to-baseplate welds each have approximately 0.05 flaws per weld. In other words, 1 in 25 (0.04)

axial welds will contain a flaw while 1 in 20 (0.05) circumferential or shell-to-baseplate welds are

expected to contain a fabrication flaw.

If a flaw is present in an MPC weld, the location of the flaw is also based on PRODIGAL results. The

flaws in the MPC welds were separated into three locations, inner surface breaking, outer surface

B-20
breaking, and embedded. Analysis of PRODIGAL results showed that 98.44% of the flaws were

expected to be embedded, 0.83% were inner surface breaking, and 0.73% were expected to be outer

surface breaking.

Given a flaw in a particular location, a 0.

probability distribution of flaw depths was

determined as follows. In the PRODIGAL -.*-ESbed king

analysis, 1,000,000 meters of weld were 0.3. --- Outer Surface Breaking

simulated. The PRODIGAL output for a given

flaw depth was multiplied by 1,000,000 meters

to give the number of simulated flaws of that 0.

flaw depth. Summing these values for all flaw

depths gives the total number of simulated

flaws. The probability of obtaining a flaw of a

given depth was calculated by dividing the 0.1

number of simulated flaws of that depth by the

total number of simulated flaws. This process

was performed for inner surface breaking .0 10 20 30 40 50 60 70 80 90

flaws, outer surface breaking flaws, and Flaw Depth (% Thickness)

embedded flaws. The resulting probability

distributions are shown in Figure B. 10. The Figure B.10. Probability Distributions of Flaw Depth

most likely flaw depth present in these welds is for MPC Welds Based on PRODIGAL Results

approximately 1/3 of the wall thickness. This is

expected, since the flaw depths predicted by PRODIGAL are a function of the size of the welding bead

and the welds are fabricated in three passes.

For a given flaw location and depth,

PRODIGAL also provides a range of flaw

aspect ratios (flaw length/depth). PRODIGAL

reports these aspect ratios in 14 bins ranging

from less than 2 to greater than 14. A

probability distribution of aspect ratios was

determined for each flaw depth and location

based on these results. An example of the

aspect ratio probability distribution is shown in

Figure B. 11 for an embedded flaw with a depth

of 31% of the weld thickness. The probability

of the last bin (>14) is much larger than the

other bins in this histogram because this bins

includes the entire tail of the distribution

representing all flaws with an aspect ratio

greater than 14.

0.2 --

0.20•

0.15.1,

0.10.1

MEmbedded Flaw, Depth=31%

<2 2-5 3.5 4.5 &.5 6.5 7.5 8.5 9.5 10.6 11.5 12.5 13.5 >14

Aspect Ratio (Length/Depth)

0,064

0.00.-

Figure B.11. Distribution of Aspect Ratios for an Embedded

Flaw with a Depth of 31% of the Weld Thickness

B-21
Stress Concentration

A stress concentration is present at the shell-to-baseplate weld because of the discontinuity formed where

the MPC shell and baseplate connect. The maximum stress concentration occurs at the surface and

quickly approaches a value of 1 away from the shell surface. The equation used to estimate this stress

concentration is as follows:

C= 1 +0.5 + 2.5 _(B.7)

where: C = bending stress concentration in the shell-to-baseplate weld

r = fillet radius (taken as 0.0254 cm or 0.01 inch)

x = depth from the shell surface

This equation is based on the stress distribution for a circular hole of radius, r, in a plate under tension,

where the maximum stress concentration factor has been increased to 4 in order to be more representative

of a small radius fillet.

For the shell-to-baseplate weld, the bending stress at a given thickness location was determined by

multiplying the nominal bending stress by this stress concentration.

B.1.6 Limit Load Model

A flow stress model was also used to predict MPC failure. Failure occurs if the applied membrane stress

attributable to internal pressure equals or exceeds the flow stress of the material:

membrane (ayild " uts) (B.8) °"flow (13.8

/10W 2

where: cfow = flow stress

Gyid = yield strength

Ou~ts ,tensile strength

The Level D stress limits in this material for membrane plus primary bending as specified in Section Ell

of the ASME Code is 150% of the primary membrane limit. Using the above definition for flow stress as

the primary membrane limit, the membrane plus bending stress limit is given below:

membrane+ bending (B.9

-flow _ 0.75* (-yld + outs)

At a given temperature, the applied stresses were calculated using the equations in Section B. 1.2. The

yield and ultimate tensile strengths of the material were sampled so the critical flow stress values could

be determined. If the circumferential or axial stresses of the MPC were greater than the membrane flow

stress, failure was predicted. Similarly, if the bending and membrane stress of the MPC shell-tobaseplate exceeded the membrane plus bending flow stress, failure was also predicted. If either of these

limits was exceeded, the entire MPC was considered failed.

B-22
Again, the probability of failure attributable to exceeding critical flow stress was calculated by

performing Monte Carlo simulations at various temperatures. The probability of MPC failure was

defined as the number of MPC failures divided by the total number of iterations.

The sampled values were correlated such that if a mean plus 1 standard deviation value of yield stress

was sampled then a mean plus 1 standard deviation tensile strength was sampled. The yield and tensile

strengths sampled in this model were allowed to fall below the minimum values defined in Table B. 1.

Even with the yield and tensile strengths below the minimum allowables, the limit load model never

contributed to failure for any of the heat-up scenarios.

B.1.7 Creep Rupture Model

The methodology used to predict creep rupture of the MPC is based on the model developed by Argonne

National Laboratory for creep rupture of steam generator tubes (Reference B. 14). In this model, the

structure accumulates a fraction of creep damage after exposure to any specific time/temperature/stress

condition. If the sum of these creep damage fractions exceeds 1, failure is predicted. In equation form,

this failure criteria is as follows:

f dt (B.10)

t, (T, Co)

where: tr(T,O) = time to cause rupture at temperature, T, and stress, a

dt = time spent at temperature, T, and stress, a

tf = total time

The above description applies to unflawed structures. If cracks are present, the above failure criteria is

modified by multiplying the stress in an unflawed structure by an appropriate magnification factor, rp:

tff dt

(tr(T,mp .-a) (B.1)

B-23
The stress magnification factors for part through axial and circumferential flaws are given below.

For axial flaws, the magnification factors are given by Equations B. 12:

1-a* a

mot I- a

1---

t

m- 0.614+ 0.481 9A + 0.386. exp(-1.259 A.)

1.82*c

Fe ran • t

(B.12a)

(B.1 2b)

(B.12c)

(B.12d)

where: nmp

a

c

t

rmean

magnification factor

flaw depth

flaw length

MPC shell thickness (= 0.5 in. or 1.27 cm)

mean radius of MPC (- 33.9375 in. or 86.2 cm)

For circumferential flaws, the magnification factors are given by Equations B. 13:

1

(B. 13a)

(B.13b) a

Not

(B. 13c) N--1+0.2. 2

where: a

0

t

flaw depth

one half angular length of flaw

MPC shell thickness = 0.5 in (1.27 cm)

B-24
The probability of creep rupture failure as a function of temperature was determined as follows. The

MPC stresses at any time-temperature condition were determined using the equations in Section B. 1.2.

For this stress, the Larson-Miller parameter required to cause creep rupture was sampled from the

distribution described in Section B.1.5. Given this temperature and Larson-Miller parameter, the time

required to cause creep rupture failure, tr(T,y), can be determined. The incremental creep damage is

calculated by dividing the actual time spent in these conditions, dt, by the time needed to cause creep

rupture, tr(T,a). This procedure is repeated for the next time-temperature-stress condition. If, at the

temperature of interest, the summation of creep damage fractions is greater than 1, MPC failure is

predicted.

Over time, the heat load in the cask would slowly drop. This heat reduction would cause a corresponding

reduction in MPC temperature and internal pressure. This effect was not modeled in the thermal

analyses, so it was assumed that the internal heat load remained constant for the entire 20-year license

period of the cask.

Monte Carlo simulations were performed on MPCs to determine the probability of creep rupture failure.

A simulated MPC consisted of both flawed and unflawed regions. The flawed regions consisted of the

axial welds, one circumferential weld and one shell-to-baseplate weld. The unflawed regions were made

up of the MPC shell plates. The creep rupture analysis assumed that all flaws were located on the weld

surface. If any of the regions were predicted to fail, the entire MPC was counted as a failure. The

probability of MPC failure is the number of failures divided by the number of simulations performed.

B.1.8 Results

Mechanical Loads

Table B.9 shows the probability of the failure of a single finite element (i.e., the probability of weld crack

initiation) through the thickness of the MPC shell for (1) the drop of the transfer cask onto the concrete

floor from 30.5, 21.3, 12.2, and 1.52 meters (100, 70, 40, and 5 feet); (2) the drop of the transfer cask

onto the storage overpack from 24.4, 12.2, and 1.5 meters (80, 40, and 5 feet); (3) the drop of the MPC

into the storage overpack from 5.8 meters (19 feet); and (4) the storage cask tip-over event. As discussed

in Section B. 1.4, the probability of a single element failure through the thickness of the shell is not the

same as the probability of a breach of the confinement boundary. However, for the purpose of this PRA,

the probability of a single element failure through the thickness will be conservatively equated to the

probability of a breach.

B-25
Table B.9. Summary of the Probability of MPC Weld Crack Initiation for Various Impact Events

Event Scenario Impact Surface Drop Height Probability of Weld Crack

(Target) (ft) Initiation

5 < 0.000001

40 0.00036

Concrete Floor

70 0.0026

Transfer Cask Vertical Drop 100 0.0196

5 0.000002

Storage Cask 40 0.000014

80 0.000203

Drop of MPC into Storage Overpack Storage Overpack 19 0.282

during Transfer Operations Pedestal

Storage Cask Tip-Over Concrete Pad NA < 0.000001

This value should be <0.000001. See Table A.4, Note 2.

Thermal Loads

The creep, limit load, and fracture models were used to analyze the MPC failure probability for the

external fire and blocked vent accident scenarios. Table B. 10 shows the MPC temperature and pressure

conditions for these events.

Under normal steady-state conditions, the bulk MPC internal pressure is 0.565 MPa (82 psi). The MPC

temperature varies with axial location as shown in Figure B. 12. In this figure, height is measured from

the base of the storage overpack such that the bottom of the helium gas and fuel rods are at a height of

0.67 meter (2.2 feet). The minimum temperature of 73 'C (163 'F) occurs at the shell-to-baseplate weld

at the bottom of the MPC. The maximum MPC temperature of 180 'C (356 'F) is located at the top of

the MPC. Even though the axial weld temperature varied between these two extremes, the failure

analyses assumed a uniform temperature equal to the maximum value. Even if these conditions were

maintained for the full 20-year license period of the cask, the failure analysis shows that no MPC failures

are expected.

If the storage cask was exposed to an external fire for 3 hours, as a result of a Gulfstream IV aircraft

crash, the MPC pressure will increase to 0.634 MPa (92 psi). However, as mentioned in section B.1.2, a

credible fire scenario will be much less than 30 minutes. As a result, the values resulting from the 3-hour

fire analysis are very conservative. The MPC temperature will also increase ,as shown in Table B. 10 and

Figure B. 12. Since creep failure is a "time at temperature" phenomenon, it was assumed that the 3-hour

fire occurred after the 20-year license period of the cask. Even under these conditions, the failure

analysis shows that no MPC failure is expected.

B-26
Table B.10. MPC Temperature and Pressure Conditions

Following External Fires and Blocked Vent Scenarios

100% Vent 1000/ Vent

Quantity, Normal Steady-State 3-Hour Blockae, Blockae, SlUnt (.S)CodiinsFieBlockage, Blockage, SI Units (U.S.) Conditions Fire No Fuel Rupture 50% Fuel Rupture

Internal Pressure, 0.565 (82) 0.634 (92) 0.745 (108) 0.9 (130.5)

MPa (psi)

Maximum Temperature 172(342) 352(666) 283 (542) 283(542)

in Axial Weld, 'C (°F)

Temperature of

Circumferential Weld, 133 (271) 331 (628) 279 (534) 279 (534)

°C (OF)

Temperature of Shell to Bemperature Baseplate W Weld, f Seld to 73 (163) 335 (635) 206 (403) 206 (403) 'C (°F)II

If all four cooling vents become blocked, the

MPC temperatures will increase to new steadystate values shown in Table B. 10 and Figure

B. 12. The bulk MPC pressure would become

0.745 MPa (108 psi). The results in

Appendix C, "Response of Fuel to Loads,"

show that for vent blockage, some of the fuel

within the cask is expected to fail. This will

result in an increase in MPC pressure, which is

not considered in the thermal analysis.

Table 4.4.14 of the HI-STORM SAR

(Reference 1) shows the normal MPC pressure

with no fuel failure as 0.496 MPa (72.0 psi).'

Assuming 10% and 100% fuel rod rupture, these

pressures increase to 0.517 MPa (75.0 psi) and

.,00

240

ISO

,.... ....... . ............. .

............. .....

-- - - ----------

LBlockcd Veni~

I 1.5 2 2.5 3 3.5 4 4.5 5

I Icighb• Ira

Figure B.12. MPC Wall Temperatures

as a Function of Axial Height

0.704 MPa (102.1 psi), respectively. These represent pressure increases by factors of 1.042 and 1.418,

respectively. Interpolating between these values gives an MPC pressure increase of 1.208 for the case of

50% fuel failure. Results of the thermal analysis performed for this PRA show MPC pressure of 0.745

MPa (108 psi). Increasing this value by a factor of 1.208 results in a pressure of 0.9 MPa (130.5 psi).

The failure analysis of the vent blockage scenario assumed this bulk internal pressure to take into account

the increased pressure from fuel failure. These results are shown in Table B. 10. The failure analyses

show that even after 20 years of vent blockage and 50% fuel failure, no MPC failures are expected.

Gauge pressures listed in the HI-STORM SAR (Reference 1) converted to actual pressures.

B-27
B.2 References

B. 1. HOLTEC International, "HI-STORM Final Safety Analysis Report," HI-2002444, Revision 1,

Marlton, New Jersey, September 6, 2002.

B.2. American Iron and Steel Institute, "High-Temperature Characteristics of Stainless Steels,"

Publication No 9004, Nickel Institute, Washington, DC, 1979.

B.3. Smith, G.V., "An Evaluation of the Yield, Tensile, Creep, and Rupture Strengths of Wrought

304, 316, 321, and 347 Stainless Steels at Elevated Temperatures," ASTM Data Series DS 5S2,

American Society for Testing and Materials, West Conshohocken, Pennsylvania, 1969.

B.4. Klueh, R.L., and D.A. Canonico, "Elevated Temperature Tensile Strength and Microstructure

of a Weld-Overlaid Type 304 Stainless Steel Forging," Symposium on Elevated Temperature

Properties of Austenitic Stainless Steels, June 24-28, American Society of Mechanical

Engineers, 1974.

B.5. National Research Institute for Metals, "Data Sheets on the Elevated-Temperature Properties

for Base Metals, Weld Metals, and Welded Joints of 1 8Cr-8Ni Stainless Steel Plates,"

NRIM Creep Data Sheet No 32A, Tokyo, Japan, 1995.

B.6. Vacari, J.A., "Standard Wrought Austenitic Stainless Steels," Materials Engineering, Reinhold

Publishing Company, December 1974.

B.7. Conway, J.B., Stress-Rupture Parameters. Origin, Calculation and Use, Gordon and Breach,

Science Publishers, New York, New York, 1969.

B.8. Chapman, O.J.V., "Simulation of Defects in Weld Construction," Reliability and Risk in

Pressure Vessels and Piping, ASME PVP Vol. 251, pp. 81-89, American Society of Mechanical

Engineers, New York, New York, 1993.

B.9. Chapman, O.J.V., M.A. Khaleel, and F.A. Simonen, "A Simulation Model for Estimating

Probabilities of Defects in Welds," ASME PVP Vol. 323, pp. 375-391, "Fatigue and Fracture,"

Volume 1, American Society of Mechanical Engineers, New York, New York, 1996.

B.10. Chapman, O.J.V., and F.A. Simonen, "RR-PRODIGAL: Model for Estimating the Probabilities

of Defects in Reactor Pressure Vessel Welds," NUREG/CR-5505, U.S. Nuclear Regulatory

Commission, Washington, DC, 1998.

B.1 1. Chapman, O.J.V., M.A. Khaleel, F.A. Simonen, and D.O. Harris, "A Method for Estimating

the Probabilities of Defects in Piping Welds," ASME PVP Vol. 386, pp. 127-144, "Probabilistic

and Environmental Aspects of Fracture and Fatigue," American Society of Mechanical

Engineers, New York, New York, 1999.

B.12. Schuster, G.J., S.R, Doctor, and F.A. Simonen, "A Methodology for Determining Fabrication

Flaws in a Reactor Pressure Vessel," Proceedings of ASME-JSME International Conference on

B-28
Nuclear Engineering, ICONE4, Volume 1, Part A, "Basic Technological Advances,"

pp. 187-194, American Society of Mechanical Engineers, New York, New York, 1996.

B. 13. Schuster, G.J., S.R. Doctor, and A.F. Pardini, "Validation of Reactor Pressure Vessel Fabrication

Flaws," Proceedings of NDE Damage Assessment Workshop, October 6-7, 1997, Electric Power

Research Institute, La Jolla, California, 1997.

B.14. Majumdar, S., W.J. Shack, D.R. Diercks, K. Mruk, J. Franklin, and L. Knoblich, "Failure

Behavior of Internally Pressurized Flawed and Unflawed Steam Generator Tubing at High

Temperatures: Experiments and Comparison with Model Predictions," NUREG/CR-6575,

U.S. Nuclear Regulatory Commission, Washington, DC, March 1998.

B. 15. Holt, J., H. Mindlin, and C. Ho, Structural Alloys Handbook, Purdue University, Center for

Information and Numerical Data Analysis and Synthesis (CINDAS), West Lafayette, Indiana,

1996.

B. 16. Peckner, D., and I. Bernstein, Handbook of Stainless Steel, McGraw-Hill, Columbus, Ohio, 1977.

B.17. Atlas of Stress-Strain Curves, Second Edition, American Society of Metals, Material Park, Ohio,

2003.

B.18. Ward, A., and Blackburn, L., "Elevated Temperature Tensile Properties of Weld-Deposited

Austenitic Stainless Steels," ASME Journal of Engineering Materials and Technology, American

Society of Mechanical Engineers, New York, New York, July 1976.

B.19. Stoner, K., R. Sindelar, and G. Caskey, "Reactor Materials Program: Baseline Material Property

Handbook - Mechanical Properties of 1950s Vintage Stainless Steel Weldment Components,"

TN: 89-023-A-1, WSRC-TR-91-10, Westinghouse Savannah River Company, Aiken,

South Carolina, April 1991.

B.20. Lee, W., C. Lin, C. Liu, and C. Cheng, "The Effects of Strain Rate and Welding Current Mode

on the Dynamic Impact Behavior of Plasma-Arc-Welded 304L Stainless Steel," Metallurgical

and Material Transactions, Volume 35A, May 2004.

B.21. Davis, E., and F. Connelly, "Stress Distribution and Plastic Deformation in Rotating Cylinders

of Strain-Hardening Materials," ASME Journal of Applied Mechanics, American Society of

Mechanical Engineers, New York, New York, March 1959.

B.22. Manjoine, M., "Ductility Indices at Elevated Temperature," Transactions ofASME, American

Society of Mechanical Engineers, New York, New York, April 1975.

B.23. Manjoine, M., "Creep-Rupture Behavior of Weldments," Welding Journal.- Research

Supplement, Vol. 61, No. 2, American Welding Society, Miami, Florida, February 1982.

B.24. Singh, K.P., J. Zhai, "Multi-Purpose Canister: A Bulwark of Safety in the Post-9/11 Age,"

International High-Level Radioactive Waste Management Conference, Las Vegas, Nevada, 2003.

B-29

APPENDIX C

FUEL RESPONSE TO LOADS

APPENDIX C

FUEL RESPONSE TO LOADS

C.1 Fuel Response to Mechanical Loads

C.1.1 Fuel Rod Buckling

For cask vertical drop accidents, the fuel rods in the assemblies are analyzed as long, slender columns

with intermediate supports provided by the assembly's spacer grids. Since the section of the fuel rod

between the bottom two spacer grids supports the weight of the entire rod, only this section of the rod

is typically analyzed. For this segment, the critical buckling load, acr, is given by Equation C. 1 from

Reference C. 1:

ccr - 2 W W(.1)

where: acr = critical inertia load magnitude for rod buckling (g)

C = axial length of the fuel rod between two neighboring transverse spacer grids

E, I, = flexural rigidity of the cladding tube column

W, = total weight of the fuel rod cladding

Wf = total weight of the fuel pellets in the fuel rod

Under static and quasi-static loading, where load duration is relatively long, unrestrained lateral buckling

and cladding failure (rupture) can be considered synonymous. By contrast, for dynamic impact events,

where load duration may be on the order of only a few milliseconds, exceeding the critical buckling load

will not, by itself, cause failure of the cladding. Rather it is the combination of load and load duration

that determines whether the cladding reaches its failure strain. However, regardless of whether the

analysis is static, quasi-static, or dynamic; whether the fuel is bonded or unbonded to the cladding; or

whether the fuel is in granular or solid form, all (100%) of the fuel weight must be included in the

buckling process (Reference C. 1).

Table C. 1 lists the specifications of fuel rods in BWR 7x7, 8x8, and lOxl 0 assemblies that could be

loaded into the MPC at a BWR site. Table C.1 also lists the specifications of a fuel rod in a PWR 15x15

assembly. Even though a PWR assembly would not be contained in an MPC at a site, it was selected to

illustrate the methodology for fuel response to drop impact, because PWR fuel rods are generally more

vulnerable to impact events than BWR fuel rods (Reference C.5). This can be seen by comparing the

static buckling loads of the fuel rods in Table C.1. Therefore, to bound the BWR assemblies, a PWR

15x15 fuel rod was used to illustrate the methodology.

C-1
C.1.2 Methodology and Modeling

The peak g loads imposed on a fuel rod from the vertical drop scenarios considered in the PRA far

exceed the elastic buckling load of the fuel rods in the assemblies listed in Table C. 1. Therefore, to

determine the strain ductility demand on the fuel cladding, the inelastic buckling capacity of the fuel rods

must be considered. The approach taken here is to analyze the fuel rod as an elastic-plastic beam-column

with initial curvature under dynamic impact. Failure of the fuel rod is determined by comparing the

maximum strain in the cladding to a strain limit based on experimental data.

Table C.1. BWR 7x7, 8x8, and 10xl0 Fuel Bundle Parameters

Used to Calculate Buckling Loads for Vertical Drop Scenarios

Rod Array BWR 7x7 BWR 8x8 BVR 10x101 PWR 15x15

Fuel Length, in (in) 144 (3.66) 144 (3.66) 154 (3.91)

# Spacers 7 7 8

Fuel Rod Length between 19.62 (0.50) 19.62 (0.50) 20.5 (0.52)

spacers, in

Modulus, Psi (Pa) 9.8x106 9.8x106 9.8x106 9.8x106

(6.79x 10'0) (6.79x101°) (6.79x 10) (6.79x10l1)

Thickness, in (m) 0.035 0.035 0.0265 (0.0007)

(8.89x 104) (8.89x 104)

Outer Radius, in (in) 0.285 (0.00724) 0.247 (0.00626) 0.215 (0.005)

Inner Radius, in (m) 0.250 (0.00635) .0.212 (0.00537) 0.189 (0.004)

Cross-Sectional Area, in' 0.059 0.050

(m2) (3.8x 10-5) (3.25x 10-)

I (Cladding Moment .00211 .00133 0.00068

of Inertia), in4 (M 4) (8.8x10- I) (5.53x 1010) (2.83x 10'0)

Cladding Weight, lbf (N) 1.98 (8.82) 1.70 (7.55) 7.0112 (31.19)

Fuel Weight, lbf (N) 9.60 (42.70) 7.46 (33.18)

P.r (Critical Buckling Load), 531.09 333.72 138.00

lbf (N) (2362.42) (1484.45) (613.86)

a (Critical Buckling Load), g 46 (46) 36 (36) 24 (24) 22 (22) (g)

2

Nvussimg iii. rmatuon is proprietary.

Total fuel rod weight.

C-2
Analysis Approach

Since the impact event entails the interaction of potentially hundreds of rods, their spacer grids, tie plates,

fuel spacers, basket, cask, and target under dynamic loading, the necessity for computational efficiency

can be achieved by modeling a single fuel rod, as was done in some previous studies (References C.5 and

C.8). Therefore, an explicit finite element analysis model of a single fuel rod with lateral displacement

constraints was used to study the inelastic behavior of a fuel rod under dynamic impact loads. The

LS-DYNA model used herein was initially developed by Pacific Northwest National Laboratory (PNNL)

and the NRC staff (Reference C.3) to study the behavior of fuel rods under the hypothetical accident

conditions described in Title 10, Part 71, of the Code of Federal Regulations (10 CFR Part 71), and was

adapted by the NRC staff to the target characteristics of the drop scenarios evaluated in the PRA.

The model used in this evaluation is based on the following assumptions:

(1) All individual rods of a fuel assembly displace with the same deformed shape. This suggests that

the single rod response is characteristic of all rods and can be studied independently. The center

fuel rod in the assembly is considered representative of the typical response of the cross-section

of rods through the whole assembly.

(2) All bending deformations occur in a single plane, although an actual assembly could exhibit

bending deflections and/or twisting within the fuel compartment. This assumption also allows

only half of the fuel rods to be modeled.

(3) The three-dimensional (3-D) finite element model uses a shell element to model the mass and

stiffness properties of the cladding cross-section, but no comparable representation of the fuel

was included. This assumption, which is consistent with standard fuel rod buckling analysis,

provides a realistic conservative representation of the fuel by discounting any contribution of the

fuel pellets to the fuel rod's flexural rigidity, but, at the same time, including 100% of their mass

contribution by lumping fuel mass at each shell node of the cladding. An accurate account of the

pellet-to-cladding interactions would be difficult to quantify considering the limited information

available. However, to overcome unrealistic ovalization of the open cross-section, nonlinear

single degree-of-freedom springs were included to represent a Hertzian-type contact between the

cladding and the fuel during lateral deformation.

(4) Spent fuel rods will typically exhibit deviations from straightness after exposure to a neutron flux

(Reference C.5). For rods mounted within an assembly using spacer grids, the unsupported

sections of rod tend to bow out as a result of the assembly's position in the core and the fuel rods

location within the assembly (i.e., proximity to a moderator rod, etc. Accordingly, a small

amount of rod bowing was introduced to facilitate computation. (See discussion below.)

(5) The spacer grids are composed of a small array of leaf springs that hold the rods within the

assembly. The spacer grid springs are assumed to provide lateral support to the fuel rods with a

prescribed nonlinear force-deflection curve. No rotational support is accounted for, and the grids

cannot slide axially relative to the rod.

(6) The cask was included in the model to capture the fuel rod and cask dynamics. This was

achieved by defining a point mass to represent the cask (equal to the total physical mass of the

transfer overpack and MPC divided by the total number of fuel rods accommodated in the cask).

(7) The target stiffness is modeled with a cask-to-ground spring defined by a constant

force-deflection curve whose magnitude is determined through an iterative process to match

the deformation characteristics of the MPC baseplate for each drop scenario.

C-3
Model Geometry and Details

Figure C. I illustrates the single fuel rod model, which is composed of a single fuel rod, a lumped cask

mass, springs representing the spacer grids, contact surfaces representing the basket compartment wall,

and a spring representing the target stiffness (cask-to-ground/impact limiter). Several views of the actual

finite element mesh are shown in Figure C.2. In this figure, the views shown are (a) the entire model,

(b) top of rod with basket compartment walls and spacer grid spring, (c) top of rod with fuel pellet springs,

and (d) bottom of rod with nodes representing the cask and target (impact limiter).

Pin

Spacer Grid

Springs

Lateral

Gap

Compartment

Wall

I

Initial

Velocity

of Pin

and Cask

Initial

Bowing

Pin-Cask

SprAg

Cask-GTound

Spring

(Impact Limiter)

Figure C.1. Single Fuel Rod Model

C-4
Slvacer Grid Sprmws

Fuel Pin Spring

N Cladding

Compartment Walls/ •"

Cladding

N".• Pin-Cask Springd

Cask Mass

Impact Limiter Spring

Ground

Figure C.2. Single Rod Model Showing (a) the Entire Model, (b) Top of Rod with Basket Compartment Walls and Spacer Grid Spring, (c) Top of Rod with Fuel Pellet Springs, and (d) Bottom of Rod with Nodes Representing the Cask and Target (Impact Limiter)

The typical fuel assembly rod is composed of cylindrical fuel pellets within a tubular cladding material

that has reinforced end thimbles and an internal spring to minimize gap between the pellets at all times.

The rod was modeled using shell elements with appropriate thickness and diameter to represent the

cladding, and included about 20,000 elements and 10,000 nodes. The shell element sizing provided sufficient resolution to accurately capture the stress state, as demonstrated through sensitivity studies.

C-5
Fuel Rod

The rod was modeled with an initial bow geometry of 0.25 mm (0.010 inch) between the spacer grids.

The bow between grid spacers was calculated from the overall assembly bow based on the guidance in

Reference C.7. It was assumed that the rod had an initial constant curvature in these sections. At the rod

ends, it was assumed that they had no initial rotation and that this condition would exist as a result of the

constraint provided by the upper and lower tie plates.

Grid Spacers

The spacer grids are arrayed structures that provide intermediate support to the fuel rods in the assembly.

The grids are composed of small leaf springs that ensure proper spacing of the rods in the assembly array.

The grids were assumed to provide lateral support to the rod at eight intermediate locations (Figure C. 1)

according to the fuel geometry considered. This contribution was represented by sets of compression-only

springs attached to each side of the cladding cross-section. These springs behave such that a lateral rod

deflection will compress one spring and generate a compressive force, while the spring on other side is

unloaded to simulate loss of contact.

Fuel Compartment

Deflection of the fuel rods during an impact event will lead to rod-to-rod and rod-to-compartment wall

contact. For an assembly with all rods deforming identically, the outer rods will be the first to contact

the compartment walls. Additional deformation causes the next set of rods to seat against these outer

rods. This stacking continues until all rods are compressed against the compartment wall. To implement

this behavior in a single rod model, the total lateral distance over which a rod could deform must be

determined. The gap was determined by considering the rod diameter, number of rods, and compartment

width to determine the total free space in which the center rods in an assembly could unilaterally deform,

which in the case of the PWR 15x15 assembly was 3.0 cm (1.2 inch). The compartment walls were then

modeled as rigid planes fixed in space at the proper lateral gap (Figure C.2). Contact elements were

implemented to automatically capture contact between the cladding and the compartment walls.

Rod-Cask-Target Interaction

The cask and target stiffness (impact limiter) contribute significantly to rod dynamics. To include the

effects of cask mass on system dynamics, the cask was modeled as a point mass. A compression-only

spring was then used between the rod bottom and the cask. A second compression-only spring was then

used to simulate the target stiffness and ground contact, where the ground was assumed to be rigid. This

model construction permitted the system to impact the target surface; however, any rebounding and loss

of contact between the rod and the cask or the cask and the ground would be captured.

C-6
Material Properties

Cladding

The temperature-dependent mechanical properties for the Zircalloy cladding material were obtained from

PNNL data on high-burnup fuels published in Reference C.4. Since fuel rod buckling under vertical

impact is expected to initiate at the bottom of the assembly where the fuel is coolest, a uniform

temperature of 572 'F (300 'C) was assumed for the rod. It was also assumed that the high-burnup fuel

had been irradiated to a neutron flux (fluence) of 12x 1025 n/m2 and would be subjected to a strain rate

during buckling of 10/s. The following material property estimates correspond to these conditions:

* Elastic Modulus = 11.0x10 6 psi

* Yield Strength = 92.4x10 3 psi

• Yield Strain = 0.0084 in/in

• Tangent Modulus = 14.5x 104 psi

* Poisson's Ratio = 0.40

Fuel Pellet Spring

To simulate the fuel pellet's ability to limit gross cladding ovalization when the cladding cross-section

is subjected to high bending curvature, the fuel pellet was modeled as an equivalent spring. The

force-deflection curve for this spring was adapted from the solution of an elastic cylinder compressed

between two plates under diametrical Hertzian contact as given in Roark (Reference C.9).

Grid Spacer Spring

The force-deflection curves for the spacer grid springs were taken from Reference C.5. The curves

represent elastic deformations from a number of coupled events that take place during the lateral

movement of fuel rods in an assembly, including (1) compression of the spacer grid leaf spring,

(2) compression of the spacer grid frame, (3) buckling of consecutive spacer grid cells, and (4) crushing

of the fuel rod cross-section. The spacer grids at the top and bottom of the assembly are attached to the

end plates and, consequently, are stiffer, so they are represented by a different curve.

C-7
Cask-Ground Sprinq

The cask-ground spring represents the behavior

of the target impacted by the cask. This is the

major source of energy absorption in the

system. The target (impact limiter) is

characterized by constant force deflection

curve, where the force magnitude is determined

through an iterative process to approximately

match the deformation characteristics of the

displacement time-history of the MPC

baseplate for each drop scenario. Since

30.5 cm (1 foot) is the most frequent height

from which a cask could be dropped, it is

selected to compare the actual MPC baseplate

displacement time-history to the displacement

time-history used in the fuel rod buckling

model. For the 30.5 cm (1 foot) drop, Figure

C.3 shows the displacement time-history

response of five nodes evenly spaced along the

top of the baseplate from the centerline of the

MPC baseplate to the MPC shell. For

comparison, Figure C.4 shows the

displacement time-history of the cask node in

the fuel rod buckling model for the 30.5 cm

(1 foot) drop.

Boundary and Initial Conditions

The boundary and initial conditions for the

model are summarized below:

(1) Out-of-plane translations and rotations

were constrained along the symmetry

plane of the rod.

07-09-04 MPCUTCISLAB I'IROP U.

-0.1

-0.25

-0.2-

-0.35,

_n A

Node no.

A&101218

B 101183

._,.10109B

D 100965

_F_100596

0.002 0.004 0.006 0.008

Time

Figure C.3. Displacement of MPC Baseplate Nodes

for 1-Foot Drop

UW15Xl5. MPC .C SLAa .ROP 1 FT ------

Node no.

-A10088

-0.

V

-l

.1,

°w

A

wUJ•

0 0.002 0.004 0.006 0.008

Time

Figure C.4. Displacement of Fuel Rod Model Cask

Node for 1-Foot Drop

(2) In-plane rotation and lateral translation was constrained at the bottom of the rod. This is based

on the assumption that the rods will likely be "jammed" into the bottom tie-plate and restrict

rotation and lateral translation.

(3) Vertical translation only was enforced for the rod-cask spring, cask, and cask-ground spring.

(4) All translations were fixed at the free ends of the spacer grid springs.

(5) All translations were fixed at the ground surface.

(6) An initial velocity of corresponding to the drop height was applied to the rod and cask nodes.

C-8
Fuel Rod Internal Pressure

After a fuel assembly has been exposed to a neutron flux and burned for high-burnup durations

(>45 GWd/MTU), internal rod pressures attributable to fission gas generation can range from 1,000 to

2,200.psi (6.89 to 15.16 MPa). A pressure loading of 1,400 psi (9.65 MPa) was applied to the inner

cladding surface to simulate the effects of this fission gas generation. This was applied as a preload to

the system using a stress initialization procedure prior to solution of the impact event.

C.1.3 Fuel Rod Response and Buckling Behavior

Table C.2 lists the maximum principal strain in the cladding for each drop height evaluated. For the drop

of the transfer cask (TC) onto the concrete floor, additional drops of 6.1 and 15.2 meters (20 and 55 feet)

were added to the standard drop heights of 0.3, 1.5, 12.2, 21.3, and 30.5 meters (1, 5, 40, 70, and 100 feet)

in order to highlight transitions in buckling behavior or differences in response due to target stiffness.

The results show that the cladding yield strain of 0.0084 in/in is not exceeded until drops of the transfer

cask are in the range of 6.1-12.2 meters (20-40 feet). The results also show that target stiffness has a

significant influence on cladding strain. For example, the 6.1 -meter (20-foot) drop of the TC on to the

concrete floor results in a strain less than yield. However, the 5.8-meter (19-foot) drop of the MPC into

the storage overpack, which is a much harder impact, produces cladding strains that are ten times greater

than yield. This is because the buckling modes resulting from these two impacts are completely

different, as will be illustrated later in the discussion.

Figures C.5 and C.6 show the response of the fuel rod for the 6.1-meter (20-foot) drop at the time when

the maximum principal strain in the cladding is reached. Because the material response is entirely elastic

and because of the interaction with the grid spacer springs, the maximum response occurs between the 5h

and 6' grid spacers during the elastic rebound phase. The maximum lateral displacement of 3.0 cm

(1.2 inch) is equal to the maximum allowable displacement imposed by the basket compartment wall as

can be seen in the displacement plot in Figure C.7. It can also be seen from Figure C.6 that no flattening

of the rod occurs as it touches the wall.

Figures C.8 and C.9 show the response of the fuel rod for the 12.2-meter (40-foot) drop at the time when

the maximum principal strain in the cladding is reached. As expected, the maximum lateral displacement

of 3.0 cm (1.2 inch) is equal to the maximum allowable displacement imposed by the basket

compartment wall, but a slight flattening of the fuel rod against the compartment wall can be seen in

Figure C.9.

C-9
I arle .2. IvILaximum rrmcipai Ntram in 1he iladmg for faacn urop Height Evaluatea

Maximum

Event Scenario Impact Surface (Target) Drop Height Principal Strain

_ _ _(in/in) _t) 

1 0.0043

5 0.0062

20 0.0072

Transfer Cask Vertical Concrete Floor 40 0.011

Drop

55 0.025

70 0.037

100 0.052

MPC Drop into Storage Storage Overpack 19 0.09

Overpack Pedestal

BW15X15, MPC TC SLAB 20 FT DROP

Tim. : n 0 flu1 - -- Frnpp I ijl

Contours of Mid Surface Max Prin Strain

min=O. at elemi 19339 7.204e-03

max=0.O0720393o at elemnX 5557 I 8

6.484e-03 f

5.763e-035

5.043e-03_

4.322e-03

3.602e-0J3_

2.882e-03-

2.1 61 e-03_

1.441 e-03_

7.204e-04

0,300e+00

3.0e0

L.8e1

Figure C.5. Maximum Principal Strain at Time of Maximum Response (20-Foot Drop)

C-10
BW16SXh. MPC TC SLAB 20 FTDROP

Time = 0.044A1 - Contours of Mid Surface Max Prin Strain

miln=0 at clemt 19339

max=0.00720393, at elem# 55S7

Y

6-X

j Fringe Levels

7.204e-03-

6.484e-03

5.763e-03 U

5.043e-03 _

4.322e-03-

3.602e-03 _

2.882e-03 _

2.161 e-03 3

1.441 e-83

0.000e+00_

Figure C.6. Maximum Principal Strain at Time of Maximum Response (20-Foot Drop)

BW15X15- MPC TC SLAB 20 FT DROP t~j.

-0.

4~ U.

5

~ -0.

Cu

~-0.

2-

A8-

Node no.

A .5759

o

-t

-1 2 0 0.01 0.02 0.03 0.04 0..05

Figure C.7.

Time

Fuel Rod Lateral Displacement at Point of Maximum Response (20-Foot Drop)

C-1I
B1S 6A . MPC TC SLAB 40 FT DROP

gI 111 - u.uI.

Contours of Mid Surfaes Max Prin Strain mln*Q at slemrn 19339 mox-d.01 10281. at olemi 2310

Lx

Minlie LCVGI

1.1030-02

7.720e-03

6.617oe-03

5.514"03

4.411 "-3

3.300a-03

2.206"~3

1 .103"13

0.000e+00

I ; 2310

ii

/,

Figure C.8. Maximum Principal Strain at Time of Maximum Response (40-Foot Drop)

BW15X'15 M PC T SLAB 40 FrIDR( Time , 0.01 - Contours of Mid Surface Max Prin SI

min-I) at Clem# 19339

maxaQ.0110281, at eclm# 2310

Lx

i

ani I

p

Fringe Levels

1.103e-02 --

9.S250-03 ~f

8.023e0431

7.7200m03

6.61 7e-03

5.51 40-03-

4.411 -03

3.308"-3-

2.206"-3 -

1.1 03"-3

0.000c0+..E

2310

Figure C.9. Maximum Principal Strain at Time of Maximum Response (40-Foot Drop)

C-12
Contrast the 15.2-meter (40-foot) drop response with the response for a 16.8 meter (55-foot) drop.

Figures C.10-C.12 show the progression of lateral deformation up to the maximum response at three

different times (16,0, 16,5, and 17.0 msec) for the 16,8 meter (55-foot) drop, and Figure C.13 shows the

response over two grid spacers at 17.0 msec. Figure C. 14 shows the extreme flattening of the rod against

the compartment wall and the concentration of curvature at the location where the rod loses contact with

the wall. Figure C.1 5 is the same as Figure C. 14, with the exception that it shows the longitudinal

(meridional) strain. The response for the 21.3 and 30.5 meter (70 and 100 foot) drops are similar.

-. 1.

me u 0.016- -

Cotoulrs of Mid Surface Max Prhn Strkin

minagO at Coando 19339

max*O.O! 1 Z629. at allmI 1903

L

i

K

Fringe Levels

9101 Go03 ~f

6.750043-

5.631 "-3-

3.379c-43

2.253t-03-m

1.126".35

0.0000 +80

Figure C.10. Maximum Principal Strain at 0.0160 seconds (55-Foot Drop)

C-13
Id

Time 0.016 _

Contour$ of Mid Surface Max Prin Strain

misO, at eleml 19339 max- 6.0123501, at Clem# 661

Fringe Levels

1.235"42,

1.11204)2 1

9.880"43

91645"43 - ,

7.41 0"43

6.1 75"03

3,705"-3

2.470"30

1 .2350043

p

0.O00eop

Figure C.11. Maximum Principal Strain at 0.0165 seconds (55-Foot Drop)

+o00

y

kU I ~ I _________________________________________________________________________________ O~rIO•IB, M"L; I' •., AV • rI1 UHUI" Time w 0.01T -

Centauri of Mid Surface Max Prin Strain mlnwO, at Clem# 19339

max-O.0250808. at eleml 763 I,1 Fringe Levels

2.S08a-02 -

2.2570-02.

2.0060"2

1.7560-02 -

1,505C-02

1.254"2-02.

1.003C-.2

7.524"-3

6.01 6-03

2.080,03 -i

0.0000+

Figure C.12. Maximum Principal Strain at 0.0170 seconds (55-Foot Drop)

+00 i

Y

C-14
ffr~i w.a-WE O of-Anrr- =F nP

Time = 5.015r - Contours of Mid Sudace Max Prin Strain

min=0. at eleml 19339

max=0.0250808, at elemn 763

I

Fringe Levels

2.508e-02

2.257e-02_

2.006e-02

1.756e-02 _

1.505e-02 _

1.254e-02 -

1.003e-02

/7.524e-0 5.01 6e-O

2.508e-0

0.O000c+

6-x

Figure C.13. Maximum Response Shown Over the Span of Two Grid Spacers

(55-Foot Drop)

3

3

3

01

Strain

2.508e-02-

2.257e-02 f

2.006"-2

1 .756e-02

1 .5iJse-02_

1.254C-02_

1.003C3-02_

7.524c-03_

5.01 6t-03_

2.508"1035

0.1101e+00U

y

6-X

Figure C.14. Extreme Flattening of the Rod Against the Compartment Wall

and Concentration of Curvature at the Location Where the Rod Loses Contact

with the Wall (55-Foot Drop)

C-15
BWW5XI5, MPC TC SLAB 55 FT DROP

Time = 0.01T - Fringe Levels

Contours of Mid Surtace Y-strain

minz-0.O302726. at eleml 1380 2.469"-2-

maxO0.0246949. at atorm# 763

1.9205e-02

1.370e02.

2.708e-03

-2.789e-03

-1.378"42-

-1.92se-02

-2.478e-"2

-3.027e-02

LX

Figure C.15. Closer View of Figure C.14 Showing the Longitudinal Strain Distribution

(55-Foot Drop)

The response of the fuel rod for the 5.8-meter (19-foot) drop of the MPC into the storage overpack

exhibits completely different behavior than the transfer cask drops onto the concrete floor. Figure C.16

shows the global behavior just prior to the time of maximum response. Figures C. 17 and C. 18 are

closeup views showing the maximum principal strain and the distribution of longitudinal stress at the

time of maximum response. The buckled shape is indicative of the classic response of a rod that is

unrestrained laterally when impacting a rigid surface (Reference 10). It is clear that the location of the

grid spacers does not matter, because the short wavelength mode of buckling is so dominant that it

overrides the flexible lateral resistance of the grid spacers and the lateral bow. As can be seen this free

mode of buckling has a much shorter half-wavelength than the distance between grid spacers.

C-16
W1 5X15 .MPC 19 FT DROP INTO TC

Time = 0.0054997

Y

L x

Figure C.16. Buckling Response at the Rod Bottom End

Just Prior to the Time of Maximum Response

(19-Foot Drop of MPC into Storage Overpack)

BW1 5X15. MPC 19 FT OROP INTO TC Time t 0.0062497

Contours of Mid Surface Max Prin Strain min=0, at elemt 19339

max=0.0903551. at eleml! 145

Fringe Levels

9.036e-02

8.13:2e-02

7.220e-02

6.325e-02 -

5.421e-02 e4.51Be-02 -

3.614L-02 -

2.711a-02 -

1.807e-02

0.O000+O0 _

y1

Lx

Figure C.17. Maximum Principal Strain at Time of Maximum Response (19-Foot Drop)

C-17
5W15X156. MK 10 FT DROP INTO TC

Time a 0.0062497 Contour$ of Y-Stress

max Ipt. value minz-S24161.71. at aelom# 159 max-1 04381, at eleml 3

Fringe Levels

1 .044e*05

8.4?Oe*04 I

6.501 e+04 -

4.53 3e+04

2.564o+04-

5.960e'-03_

-1.372e+04j

-3.341 e+04

-5,309e+04-

-7,278c+04

-9.246e+04

Y

L-x

Figure C.18. Longitudinal Stress Distribution at Time of Maximum Response

(19-Foot Drop)

C.1.4 Fuel Rod Cladding Failure

The cladding strain at failure (rupture or breach) of high-bumup fuel, which is fuel with a burnup greater

than 45 GWd/MTU, is expected to range from 1.0% to 3.0%. For the purpose of the PRA, the lower

value of 1.0% strain is selected as the strain limit.

Appendix D of the PRA shows that the particulate release fraction from a spent fuel rod is a function of

three variables: (1) the fraction of the fuel rim layer that fractures, (2) the percentage of particulate

entrained by gasses during release, and (3) the number of tears (breaches) per rod. In addition, the

fraction of the rim layer fractured and the number of breaches per rod are, in turn, functions of the

g loads experienced by the fuel and the severity of the buckling that takes place during a given drop.

In a full-blown PRA each of these variables would be associated with a probability distribution and

sampled as part of an overall simulation. However, because of the level of effort inherent in such an

approach, Appendix D simplifies the methodology by recommending an upper bound value for release

fractions to be used in the consequence analysis. This, in turn, simplifies the effort here in Appendix C

by not requiring an estimate of the number of breaches per rod, but only whether or not the rod breaches

at all. Also, since the analysis of fuel rod failure is based on a single pin model, it will be assumed that

when one rod fails, all rods fail, although this is very unlikely.

For each drop height, Table C.3 lists the maximum principal strain in the cladding, the failure strain limit

and whether or not fuel cladding breach occurs based on that limit strain.

C-18
Table C.3. Likelihood of Fuel Cladding Breach for Various Drop Scenarios

Maximum Strain Limit

Event Scenario Impact Surface Drop Height Principal Strain Selected for Fuel Cladding (Target) (ft) in Cladding High Bumup Breach

(in/in) Fuel

1 0.0043 0.01 No

5 0.0062 0.01 No

20 0.0072 0.01 No

Concrete Floor 40 0.011 0.01 Yes

Transfer Cask 55 0.025 0.01 Yes

Vertical Drop 70 0.037 0.01 Yes

100 0.052 0.01 Yes

5 -1 na No

Storage Overpack 40 -1 na Yes

80 -1 na Yes

MPC Drop into Storage Overpack

Storage Oeestal 19 0.09 0.01 Yes

Overpack

(1) Thns drop was not evaluated for cladding response. Results tor the transfer cask drop onto the concrete

floor from the same height were used to determine cladding breach.

C.2 Fuel Response to Thermal Loads

The thermal analyses described in Section 4.2.3 of this report calculated the temperature of the fuel

cladding for both the external fire and blocked vent scenarios. For these accident scenarios, fuel

response is determined by comparing the calculated cladding temperatures to the temperature limits as

discussed in Section 4.2.3. The long-term normal and short-term accident temperature limits for

Zircalloy cladding are 400 'C (752 'F) and 570 'C (1,058 'F), respectively.

In the case of an external fire, the short-term temperature limit is used because the fire duration is only

3 hours. In the vent blockage scenario, the long-term normal temperature limit is used because vent

blockage is assumed to occur for 20 years for the purpose of the PRA only.

C-19
C.2.1 Results of Fuel Response to Thermal Loads

The thermal analyses performed for this PRA

show that under normal steady-state conditions

the bulk MPC pressure is 0.565 MPa (82 psi)

and the average fuel temperature is only 179 'C

(354 IF). If the storage cask is exposed to an

external fire for 3 hours the pressure increases

slightly to 0.634 MPa (92 psi). Figures C.19

and C.20 plot the distribution of maximum

cladding temperatures following a 3-hour fire.

Figure C. 19 shows the maximum cladding

temperature as a function of axial height while

Figure C.20 shows the maximum cladding

temperature as a function of radial position

from the cask center. The axial height is

measured from the base of the overpack such

that the bottom of the fuel rods is at a height of

0.67 meter (2.2 feet). Also shown in these

figures is the 570 'C (1,058 °F) short-term

temperature limit. These figures clearly show

that the cladding temperature limit is not

exceeded. Therefore, no fuel failures occur if

the storage cask is exposed to an external fire

for 3 hours. [Note that cladding failure would

not be expected to occur until cladding

temperatures reached 750 'C (1,382 IF.]

For the vent blockage scenario, the MPC

pressure and cladding temperature are also

expected to increase. For the purpose of this

PRA, the assumption is made that vent

blockage occurs for the entire 20 year license

period of the cask. Under these conditions, the

MPC pressure will eventually reach a new

steady-state value of 0.745 MPa (108 psi).

i0u

100)

f Max Fuel Temp Limit

-- Max Fuel Temp After 3 hr

S 1.5 2 2.5 3 3.5 A4 4.5

l I r)nih 

Figure C.19. Maximum Fuel Temperature

as a Function of Axial Height after a 3-Hour External Fire

2o

5t(} I00 Short Term Tempefature Limit

Max Fuel Temperature After 3 hr

0 02 01.4 0.6 0).8

R 111)

Figure C.20. Maximum Fuel Temperature

as a Function of Radius after a 3-Hour External Fire

C-20
Figure C.21 shows the maximum cladding temperature as a function of axial height after steady-state

vent blockage. Again the axial height is measured from the base of the overpack such that the bottom of

the fuel rods is at a height of 0.67 meter (2.2 feet). Figure C.22 plots the maximum cladding temperature

as a function of radial position after vent blockage. These figures also show that some of the fuel within

the cask will exceed the normal long-term temperature limit of 394 °C (742 OF). The fraction of fuel rods

that exceeds this temperature limit can be estimated as follows. The dashed line in Figure C.22

represents a polynomial fit (trend line) to the radial temperature distribution data. The equation of this

polynomial fit is as follows:

TF.eI = -- 294.53 * x2 + 68.731 9 x + 462.01 [C.2]

where: Tfu.j

x

= Maximum fuel cladding temperature in °C

- Radial distance from cask center in meters

The fit to the data is excellent with an R2 value of 0.994.

Based on this equation, all fuel rods located within a radius of 0.615 meter (2.02 feet) of the cask center

will have a maximum cladding temperature that exceeds the zircalloy normal long-term temperature

limit. The circular area associated with these failed fuel rods is 1.19 mr2 (12.79 ft2). The cross-sectional

area of the entire MPC interior is 2.37 in2 (25.5 ft2). The fraction of fuel rods that exceeds the

temperature limit is estimated by taking the ratio of these two areas. Using this procedure, the fraction of

fuel expected to exceed long-term temperature limits under blocked vent conditions is 0.5.1

"5

M\ l\ I id . l t

I .5 2 215 3 3,5 4 4.5 5.

I I¢iL-ht (m)

Figure C.21. Maximum Fuel Temperature

with Vent Blockage as a Function of Axial Height

50

1 300- Temp(C).-294.53x^2+68.7a1x+462.01

Avtere, x = Radial Distance(m)

200 R^2-=0945

*• Max Fuel Tomp - Blocked -Vents |

_-Long-Torm Tomp Limitt

100 . . TrendLneBloc.d Vent

0 . . . . . . . . I . . . . . . .... - , .-

0 0.1 0.2 0.3 0.4 0.5 0.8 0.7 0.8 0.9

Radial Distance (m)

Figure C.22. Maximum Fuel Temperature

as a Function of Radius After Vent Blockage

and Zircalloy Normal Long-Term Temperature Limit

One component of determining the rod stress is the gas expansion attributable to increased temperature. At least half of

the gas in the plenum is usually not the hottest part of the rod. Therefore, the stress will not be as high as if the rod was

at a uniform temperature. Calculations for creep, resulting in the 400 IC (752 'F) temperature are based on a uniform

temperature. Thus, the calculation is a conservative over-prediction of the rod failure rate.

C-21
C.3 References

C. 1. Bjorkman, G.S., "The Buckling of Fuel Rods under Hypothetical Accident Conditions,"

Proceedings of PATRAM2004, Berlin, Germany, September 2004.

C.2. HOLTEC International, "HI-STORM Topical Safety Analysis Report," HOLTEC Report HI951312, Revision 8, Marlton, NJ, 1999.

C.3. Adkins, H.E., B.J. Koeppel, and D.T. Tang, "Spent Nuclear Fuel Structural Response When

Subject to an End Drop Impact Accident," Proceedings ASME/JSME Pressure Vessels and

Piping Conference, PVP-Vol. 483, American Society of Mechanical Engineers, New York,

New York, 2004.

C.4. Geelhood, K.J., and C.E. Beyer, "Mechanical Properties of Irradiated Zircalloy," Transactions of

ANS Winter Meeting, Vol 93, American Nuclear Society, Washington, DC, November 2005.

C.5. Sanders, T.L., K.D. Seager, Y.R. Rashid, P.R. Barrett, A.P. Malinauskas, R.E. Einziger,

H. Jordan, T.A. Duffey, S.H. Sutherland, and P.C. Reardon, "A Method for Determining the

Spent-Fuel Contribution to Transport Cask Containment Requirements," SAND90-2460,

Sandia National Laboratories, Albuquerque, New Mexico, 1992.

C.6. U.S. Nuclear Regulatory Commission, "Buckling of Irradiated Fuel Under Bottom End Drop

Conditions," ISG 12, Rev. 1, Washington, DC, 1999.

C.7. U.S. Nuclear Regulatory Commission, "Buckling of Irradiated Fuel Under Bottom End Drop

Conditions" ISG 12, Rev 2 (DRAFT), Washington, DC, 2006.

C.8. ABB Combustion Engineering, Inc., "Fuel-Assembly Behavior Under Dynamic Impact Loads

Due to Dry-Storage Cask Mishandling," EPRI-NP-7419, prepared for Electric Power Research

Institute, Palo Alto, California, 1991.

C.9. Roark R.J., Formulas for Stress and Strain, 5th Ed., McGraw-Hill, New York, New York, 1975.

C. 10. Jones, N., Structural Impact, Cambridge University Press, Cambridge, England, 1989.

C-22
APPENDIX D

METHODOLOGY TO DETERMINE SOURCE TERM

FOR A DRY STORAGE CASK

APPENDIX D

METHODOLOGY TO DETERMINE SOURCE TERM

FOR A DRY STORAGE CASK

D.1 Introduction

Should a cask containing spent nuclear fuel be impacted by a projectile or strike a hard surface, the

potential exists for radioactive material to be released from the cask if the magnitude of the impact is

large enough to breach the cask and disrupt the fuel rods. The extent of the release will depend on the

magnitude of the impact, the characteristics of the payload of the cask, the configuration and

environment of the cask, and the physical details of the cask. This section will present a model to

evaluate the fraction of material released from a breached cask, and discuss the variability of the input

parameters required to exercise the model. The release fractions for a particular cask and payload,

derived by the methodology will be given.

D.2 Methodologv

A severe enough impact on the cask will fail a fraction of the fuel rods (Frods) and make their contents

available for release. A portion of this available material will be released from the rod to the cask

environment (FRC). Part of the inventory that has been released to the cask will be transported through

the breach in the cask to the environment (FCE). The fractional release from the cask (Fel) can be

approximated by the following equation:

Fre = Frods x FRc x FcE (D. 1)

The fuel rods can have radionuclides on the cladding surface in the form of solid Chalk River

Unidentified Deposits (CRUD), inside the fuel rod as gases and volatile fission products, and within U0 2

fuel grains as both solids and gases. Because of the different physical forms of the radionuclides,

Equation D. 1 must be analyzed separately for each radionuclide or group of similar radionuclides. Since

long-duration fires are not considered in this evaluation, and the maximum design storage temperature of

the fuel is 400 'C (752 'F), transport of fission products, such as Cs in the form of CsI that has a melting

point >600 'C (1,112 'F), as condensable vapors will not occur, the radionuclides in spent fuel were

assigned to one of the following three chemical element groups (k= 1 to 3): noble gases, particulates, and

CRUD. For each of these three groups, the inventory of each radionuclide that might be released from a

breached cask to the atmosphere (the source term STi for radionuclide 1) can be calculated as the product

of the release fraction Fml,k for the chemical element group k to which the radionuclide belonged and the

inventory Ii of that radionuclide in the cask. Thus, we have the following relationship:

3 (D.2)

i k-L I ik Frel, k

D-1
The major difference between the current and previous models (Reference D. 1) is that the current model

accounts for the different properties of the rim region formed in high-bumup fuel, as well as the reduced

ductility of the cladding as a result of higher bumup and hydrogen absorption. The methodology, as

presented, is only applicable to impact type events (not those that result in an elevated temperature

attributable to a long-duration fire). A long-duration fire may introduce volatile fission products as a new

class of releasable radionuclides, add oxidation of the fuel as a release mechanism (for example oxidation

of the stable RuO2 oxide to the volatile RuO4), change the particle size distribution, produce additional

rod failures by thermal burst rupture of unfailed rods, and increase the driving force for dispersal of

material.

D.2.1 Radionuclide Inventory (Ii)

The ORIGEN2 code (References D.2 and D.3) is recommended to calculate the radionuclide inventory of

the fuel. Because an ORIGEN calculation provides output for more than 800 radionuclides, the number

of radionuclides in the ORIGEN output can be reduced by dividing the 10-year cooled amount of each

radionuclide by its A2 value (References D.4 and D.5) and then selecting the smallest set of these

normalized Curie amounts that yielded a sum greater than 99.9 percent of the sum of all of these

normalized Curie amounts. This procedure reduces the ORIGEN output to a much smaller set of 12

radionuclides that are important for the estimation of radiological health effects. If fuel fresher than

10-year cooled is loaded, this list of radionuclides may change and should be reevaluated.

In addition, five radionuclides, '°6Ru, 134 Cs, 144Ce, 1

47

Pm, and 1

54

Eu, available for release in most fuels

should be added so that most of the fuel inventories contain the same set of radionuclides. In order to

have a noble gas in the inventory, although eliminated by the A2 screen, 8'Kr, the noble gas with the

largest Curie amount in the ORIGEN inventory for 10-year cooled high-burnup spent boiling-water

reactor (BWR) fuel, should also be added to the reduced set of radionuclides.

Finally, 6"Co should be added to the reduced set of radionuclides so that release of 6

"Co attributable to the

spallation of CRUD (Reference D.6) from rod surfaces could be addressed. CRUD is the deposits that

form on the surfaces of spent fuel rods while the rods are in an operating reactor (Reference D.7). The

properties of CRUD were extensively reviewed by Sandoval (Reference D.7). Recommendations for

calculating the CRUD inventory are given in the Table D. 1 below.

D-2
Table D.1. Recommendations for Calculating the CRUD Inventory

CRUD

Parameter Recommendation

Inventory Recommended values in the Standard Review Plan bound 99% of the fuel inventory. The

CRUD inventory is reduced by a factor of 10-20 if only 90% of the fuel is bounded.

Recommend using inventory bounding 90% of the fuel to be realistic or calculate release as a

function of the percent of fuel bounded.

Out-of-reactor Reduce the inventory of 6

"Co due to decay.

time

Axial CRUD does not tend to be evenly distributed along a fuel rod. The ratio of the average CRUD

Distribution thickness to maximum CRUD thickness is about a factor of 2. Reduce maximum surface

CRUD concentrations by a factor of 2.

Burnup There is no apparent burnup dependence for CRUD thickness but the inventory can be scaled

linearly with bumup for conservatism. Some newer fuels have very thick crud resulting from

water chemistry. The influence of water chemistry should be handled as a special case.

Fuel Type Use inventory appropriate for fuel type (i.e., BWR or PWR).

While the ORIGIN2 code gives overall inventories for the fuel, it must be recognized that the

radionuclides are not evenly distributed. The spatial distribution must be accounted for due to the

variation of burnup both axially along a rod, and radially within a rod. In particular, for high-bumup fuel

that contains a rim region, the actinide inventory in the rim is about twice that in the body of the fuel

pellet. As a result, any analysis of release fractions should consider the rim region and the body of the

pellet as separate entities. The axial distribution is only important if thermal gradients in the cask make

some axial rod locations more prone to fracture than others.

D.2.2 Fraction of Fractured Rods (Fods) and Fracture Sites per Rod

Only a fraction of the energy imparted to the cask system during an impact will be available for rod

fracture. The remainder will be absorbed during cask deformation, or used to spall CRUD and fracture

the fuel within the rods.

In previous studies (Reference D. 1), calculated rod strains were compared to a rod strain failure criterion

in order to estimate how many rods would be failed and how many failure sites were expected along a

rod. There is no definitive information on what the exact strain to failure should be, especially at high

bumups. Estimates ranging from 0.25 to 4% have been proposed (Reference D.8). Although this strain

to failure should be temperature-dependent, this dependence was not considered in previous studies.

Another section in this report describes a method for determining rod failure (See Section 4.4).

D-3
D. 2.2.1 Variation in Cladding Material and Material Properties

During irradiation, hydrogen is introduced into the cladding as a result of waterside corrosion. The

preponderance of this hydrogen precipitates as circumferential hydrides in the outer third of the cladding.

The mechanical properties of irradiated Zircaloy, as a function of hydrogen content, have recently been

reviewed by Geelhood, et al. (Reference D.9). During the drying of the fuel after the loading of the

storage cask, conditions might be suitable for reorientation of some of the hydrides from the

circumferential to a radial orientation, reducing the ductility of the cladding. As fuel is driven to higher

burnups, newer claddings such as M5, and Zirlo are being put into service to reduce cladding corrosion.

Cladding mechanical properties appropriate to the type of cladding and hydride orientation within the

cladding should be used in any calculation of the number of breached rods.

D.2.2.2 Number of Fracture Locations (ntear/rod) and Size of Fracture

The number of fracture locations along a rod will not affect the amount of gas available to drive

radioactive solid particulates from the cask to the environment since all the available gas is released

irrespective of the number of sites. Since the release of fuel fines to the cask environment is a localized

event, the number of fracture sites will somewhat linearly increase the release of fuel fines. An estimate

of 5 tears/failed rod has been made in the past using the strain- to-failure method, and should be used as a

default value, unless a more realistic value (such as calculated in Section 6.1) is supportable.

At the current time, there is no data supporting a distribution of cladding crack sizes under impact. For

the purpose of this model it is assumed that cracks caused by impact are half circumferential cracks with

a length to width aspect ratio of 10. For example, the cracks formed in BWR spent fuel rods with an

inner diameter of 1.07 cm by rod impact with the spacer grid will respectively have lengths (l1) and

widths (we) of 1. = rc (drod,ipne, 2 ) = 1.7 cm and w, = 0.17 cm.

D.2.3 Driving Force for Release from Rod-to-Cask and Cask-to-Environment

Rod fracture makes two main contributions to the release of radionuclides. Unless the cladding fractures

to expose the fuel pellets and fission gases, and radionuclides contained within, only CRUD is available

for release. Secondly, depressurization of the rods carry gases and entrains fuel fines from the rod to the

cask and contributes to the force pushing any radioactivity from the cask to the environment. When the

cladding breaches, the gas contained in the rods will be added to the initial 1 to 5 atmospheres of cask

inert fill gas that will be available for cask depressurization. This is a more significant contributor in

PWR systems where the rods are initially filled to a higher gas pressure, and in casks only pressurized to

1 atmosphere.

D. 2.3. 1 Rod and Cask Depressurization

The release of small particulate residing on the internal crack network of the fuel by entrainment depends

on whether, during depressurization of the rod, the gas flows predominately through the cracks in the fuel

or any fuel-cladding gap that might have opened up during cool down. In addition, the rate of

depressurization of a rod will depend on the hydraulic diameter of the depressurization path.

D-4
In-core measurement, at temperature, of the hydraulic diameters of average- and high-burnup spent fuel

rods from the Halden Reactor show that average- and high-burnup spent fuel rods have hydraulic

diameters of about 50 and 35 ýtm respectively (Reference D.10). At temperature, the fuel cladding gap is

closed. The extent to which it opens will depend on the temperature of the fuel (i.e., the cooling time of

the fuel). Estimates of the gap from photomicrographs at room temperature range from no gap to as

much as a 50 ýLm gap. In the latter case, the hydraulic diameter and the geometric diameter of the

reopened gap should be about the same and both should be substantially larger than the hydraulic

diameter of the crack network in the pellets. This raises the question of whether rods will depressurize

primarily through the reopened gap and only minimally through the pellet crack network. If this is the

case, crack gas flow rates may be so small that particles formed on the surfaces of the cracks will not

become entrained in the gas flow and, therefore, may not be released.

The depressurization time At of a volume V through an orifice with area A can be calculated (Reference

D. 11) using the following equation:

At V/A 2 (D.3)

V \PO

where Po/po = RTo/MW, PI/PO = PITO/PoTI, y = Cp/C, and PO, Po, MW, To, Cp, and Cv are pressure,

density, molecular weight, temperature, and heat capacities at constant pressure and constant volume of

the mixture of gases in the volume V respectively, R is the ideal gas constant, and p, is the density of the

gas mixture in the volume into which the volume V is venting.

This equation with Po/po replaced by RTo/MW and p,/p0 replaced by P1To/P0 T1 can be used to estimate

the depressurization time of the cask and also of the failed spent fuel rods in this cask.

For depressurization of the mixture of He and Xe from failed rods into the interior of a cask, MW = 94.2

g/mole, y = 2.5, T1 = To = cask temperature and PO = rod pressure at temperature. V is determined by

dividing the internal rod volume by the number of crack sites on the rod (See Section D.2.2.2).

Assume for the purpose of example that the gap does not open upon cooling or that the accident occurs

shortly after loading and the fuel is still relatively hot. Then, dhmds = 3.5 x 10- 3 cm and A = 7(dh,,ods/2)2.

Even though the internal free volume of the cask is large compared to the free volume of a spent fuel rod,

because the hydraulic diameter of the fuel pellets and any pellet cladding gap in the rod is only about 35

mm, the cask should depressurize much more rapidly than the rods. So, during almost all of the rod

depressurization time period, the cask pressure will be 1 atm, which means that for rod depressurization,

P1 = 1 atm. Calculations using typical BWR rod parameters and internal pressures in Equation D.3 yield

At - 679 sec for the depressurization time of the average failed rod in a typical cask. Using Equation

D.3, it can also be shown that the assumption of rapid depressurization of the cask is also satisfied for

this example.

D-5
D. 2.3.2 Depressurization and Gas Flow Rates

The flow rate of gas through failed rods to cladding failures, out of failed rods into the cask fuel tubes,

and through the cask fuel tubes to the location of the closure failure will determine whether entrainment

of particles formed on exposed pellet surfaces, inertial impact of particles with rod spacers, and

deposition of particles onto rod and fuel tube surfaces by gravitational settling and turbulent diffusion are

significant.

The rate of gas flow out of a volume can be calculated using the following relationship:

A v At = V (Pi/Pf) (Pi- Pf) / Pi = V (Pi- Pf) / Pf (D.4)

where A is the area of the flow path, v is the linear flow velocity, At is the depressurization (flow) time,

V is the volume of gas that depressurizes, Pi is the initial gas pressure in this volume, Pf is the pressure in

the volume after depressurization, V(Pi/Pf) is the volume that the gas in the volume V would occupy were

it at the depressurization pressure Pf, and (Pi - Pf)/Pi is the fraction of the gas in the volume V that will

flow out of that volume as a result of depressurization. Rearrangement yields the following equation:

v = V(Pi - Pf) / (Pf A At) (D.5)

If particles formed on pellet surfaces by impact fracturing become entrained in rod depressurization gas

flows over these surfaces, these particles could be transported from the rod through cladding failures to

the cask interior. Because of the very small hydraulic diameter of high-burnup spent fuel rods, rod gas

pressures should be fairly uniform throughout the rod's internal crack network until the depressurization

gas flow nears a cladding failure. Thus, for flow in the internal crack network, v = V/AAt, where A is

the hydraulic cross-section. Alternatively, if rod depressurization occurs predominantly through a

reopened pellet/cladding gap, a maximum value for the flow cross-sectional area is A = Tc[(d1d,inr/2) 2

-(dpen/2)2]. For the same B'WR example as above, for flow through the internal crack network Ventain.ent

=37 m/sec. Alternatively, if rod depressurization occurs predominantly through a reopened

pellet/cladding gap, a minimum value is ven•-.ent = 0.021 rn/sec.

Impingement of the gas stream out of the rod onto a grid spacer or adjacent rod during depressurization

of the rod will determine if impaction removes particulate from the gas stream. The efficiency of particle

removal by inertial impaction will depend on the relative flow rate of rod gases as they exit cladding

failures (Vod fails) to the velocity of the gas moving perpendicular to the breach down the rod. From

Equation D.5, VCod failrs = V(Pi - Pf)/(PfAAt), where V = Vrod fiel(number of crack sites thru which the

average rod depressurizes).

Because the cask depressurizes much more rapidly than the rods, when no fuel-cladding gap is present,

the fuel (basket) tubes in the failed cask will be at atmospheric pressure while the particles released from

failed rods are flowing through them to the location of the cask's closure failure. Accordingly, the flow

velocity down the fuel tube (vmbe), can be calculated as

vtube = Vrod freeFrodsnmds/assembly(Prod - Pam./(Pat.Ambe freAt)

where Vrod free = rod free volume, Frod, = the fraction of the rods in an assembly that fail, nrods/assembly = rods

per assembly, Prod is the rod pressure before the rod fails, P5tm = 1.0 atm, At = rod depressurization time,

and Atube free equals the free cross-sectional area of a basket tube, where Ambe - Arodlds/assembly = (Wmbe) 2 -

r (drd/2) 2nrdsssembly.

D-6
D.2.4 Release from Rod to Cask Environment (FRc)

In high-burnup spent fuel pellets, capture of epithermal neutrons by uranium atoms in the pellet rim

produces a friable layer of fuel fines on the outer surface of each pellet. Measurements indicate that the

thickness of this friable rim layer (tm) in high-burnup BWIR spent fuel is about 150 ltm (References

D.12, D.13, and D.14). PWR rods have a somewhat thinner rim at an equivalent bumup. Because the

body and the rim of high-burnup spent fuel pellets have quite different morphologies (References D. 12,

D. 15, and D. 16), release of noble gases and of fuel fines from these two regions of a high-burnup pellet

may be quite different. Like fresh U0 2 pellets, the body regions of high-burnup spent fuel pellets consist

of sintered 10 pm UO2 particles. In the rim layer, 0.1 to 0.3 pim subgrains are generated by the

recrystallization of U0 2 .

The volume of fuel pellets in a single rod and the volume of the rim layer on these pellets are given by

Vpelets 7t(dpeilet/2)2 Lactive and Vrim = 7idpellet trim Lactive, and thus, Wrim/Vpellets= 4 trim/dpellet. However, the

capture of epithermal neutrons, which causes the rim layer to form, doubles the concentration of

radionuclides in the rim layer compared to their concentration in the pellet body (Reference D. 13). For a

typical BWR rod about 11.3 percent of the total radionuclide inventory in the pellets resides in the friable

rim layer of these pellets and the remaining 88.7 percent of the total inventory is contained in the body of

the pellets. Accordingly, for BWR rods,

FRCk = (0.113)(Frim release, 0 + (0. 8 8 7 )(Fbody release, ) (D.6)

A similar relationship can be developed for PWR fuel.

D.2.4.1 Particulate Release from the Rod to the Cask

D.2.4.1.1 Fuel Fracture

The release fraction for fuel fine particles can be estimated as follows. First, estimate the initial fraction

of the U0 2 mass in a rod that is converted to fuel fines by abrasion and vibration, radiation

embrittlement, and impacts. Second, estimate the fraction of fuel fines that could become gas borne.

Lastly, estimatethe fraction of the gasborne particles that might be trapped by particle bed filtration as a

result of passage through particle beds during transport to the location of the rod failure.

The release fraction for particles is then calculated as shown in Equation D.7 below, using the fractions

from the above steps for particles formed in the following locations:

(1) in the friable rim of the pellet located right next to the rod failure

(2) in the friable rims of other pellets (pellets not located right next to the rod failure)

(3) on the surfaces of the network of internal pellet cracks by abrasion, vibration, and impact

fracturing

D-7
Frel,particles = 0.1 13(Fijitrim +Fjmpr.i) { [ntears/rodFtear,rim] + [Fentrjm(Fbed)] }

+ 0.887 {ntearsirod(Finitbody + Fimpbody)(Fbed)} (D.7)

where:

Fit = the fraction of the mass of the U0 2 fuel pellets in a rod that has been converted

to respirable fuel fines by abrasion and vibration before the impact event takes

place. This fraction can be calculated for either the body or the rim region.

Fimp,body = the fraction of the mass of the U0 2 in the body of a spent fuel pellet that is

converted to respirable fuel fines by brittle fracture as a result of the impacts.

Fjmp,jm = the fraction of the mass of the U0 2 in the rim layer of a spent fuel pellet that is

converted to respirable fuel fines by brittle fracture as a result of the impacts.

Ftear,r = the fraction of the mass of the U0 2 in the rim layer that is blown out of the rod

from one rod tear during rod depressurization without filtering by passage

through a particle bed.

Fent,rim = the fraction of the mass of the U0 2 in the rim layer that is entrained in the

depressurization gas flow through the rim layer-cladding gap and transported to

the rod tear and then out into the cask.

Fbed = the fraction of the respirable particles that are not captured during flow through a

particle bed.

nteas/rod = 5 = the number of tears in an average failed rod in the cask.

D.2.4.1.2 Estimation of In-Reactor Fuel Fracture (Fn1 iJ

During reactor operation, compression of fuel cladding (PWR fuel) and swelling and cracking of the fuel

pellets in a fuel rod largely eliminates the rod's fuel-cladding gap at high burnup and creates an internal

network of cracks inside of the pellets that has a cross-sectional area in the fuel column given in Table

D.2. Abrasion during insertion of the pellets into the rod and vibration during reactor operation causes

fuel fines to form on pellet external surfaces and also on pellet crack surfaces, especially at the

pellet-pellet interfaces.

Table D.2 Free Cross-Sectional Area in Fuel Column

PWR BWR

Low-Burnup Initial gap area less correction for 0.6% Fuel cladding gap area

cladding creep down

High-Burnup << LBU PWR fuel < LBU BWR fuel

Note that the factors 0.1138 and 0.887 are for very high-burnup BWR fuel and will differ slightly for lower burnups

or PWR fuel.

D-8
Lorenz heated lengths of low-burnup PWR fuel to burst and measured the amount and size of the

particulate emanating from the burst site. The mass fraction of respirable fines in the pellet body can be

estimated from the results of Lorenz (Reference D. 17) and the filtering efficiency (Ff,) of a particle bed

for respirable particles. Let

l = the particle bed length (cm) that provides a filtering efficiency of F61 =

0.99 for respirable particles

L - 101

= the particle bed length (cm) that intercepts 100 percent of the respirable

particles passing through that length

V2(L+) = the volume of the spent fuel pellets in a rod section with a length 2(L+/)

P 10 initial = M 10 2(L+/) V2(L+/

= the mass concentration of respirable fines in the volume V 2(L+/

MK 102(L+ = the mass of respirable particles in the volume V2(L+O

M•10 rl = the mass of respirable fines released during one of the Lorenz burst

rupture release experiments

M12 in= the pellet mass in one of Lorenz's 30.5 cm (12 inch) rod test segments

p = l0g/cm3

- the mass density of a U0 2 spent fuel pellet

Finit,body = P10 initial/P•

The analysis of particle release was performed (Reference D. 1) using standard aerosol physics

(Reference D. 18) to estimate that a 0.3-cm (0.1-inch) long bed of 200 tim particles (maximum size

particle released in the Lorenz experiments) would have a filtering efficiency F6 l = 0.99 for respirable

particles. Therefore, I = 0.3 cm and L = 101 = 3 cm, where 2L is the length of the rod from which

respirable particles escape but are filtered as a result of passage through particle beds, 21 is the length of

the rod from which respirable particles escape with only partial (0 to 0.99) filtering (i.e., the particle bed

length is too short to provide significant filtering), and 2(L+l) is the length of the rod beyond which no

respirable particles escape. Given these definitions, we find the following relationship:

V2(L+I) Plinitial 2 (1 -Ff,) 2+ (1.0) =M_10re ~~na42(L+l1) 2(L +1) j

where the coefficient 2 corrects for the fact that the lengths L and I occur on each side of the rod failure

and, since the filtering efficiency (Ff6 = 0.99, we have the following relationship:

2(L+ 2(L+/ L+ "

D-9
Lorenz calculated the release fraction for particles from a 30.5-cm (12-inch) long test rod sections as the

ratio of the mass of particles released divided by the mass of U0 2 in the 30.5-cm (12-inch) long rod test

section. Thus, the release fraction value of 4x 10-6 for respirable particles, which can be derived from the

results of Lorenz, equals M 10 rel/M 12 in. Therefore, for the Lorenz'experiments, we have the following:

V 2 (L+l) P!2 0 initial L(.0O1)j+ /I =- M<r Ore 4X10-6

M 12in L + l M12in

Now because M12 in = pV 12 in and, therefore, V2(L+/M12in = 2(L+I)/p(12 in), substitution and

rearrangement of the preceding equation gives the following formula:

Pinitbody -- PlO iinitial =4X10- 6 30.5cm [L(O'O1)±ll - p 2(L+l)_ L+I I

Because heating of the rod section in the Lorenz experiments caused a bulge to grow in the cladding

before a 1 mm diameter hole bursts somewhere in the bulge, the unfiltered length of the rod test section

is the length of the bulge, which is about 5 cm. Therefore, 2 1 5 cm, 1 = 2.5 cm, and because p = 10

g/cm3 and L still = 3 cm the preceding equation yields Fiit body = 2.4x 10-. A similar calculation cannot be

done for the rim region since no tests similar to Lorenz's using high-burnup fuel have been conducted to

our knowledge.

D.2.4.1.2 Expressions for Fimp,body and Firmp,rim

The DOE Handbook (Reference D. 19) gives the following formula for Fimp, the fraction of a brittle

material that is converted to respirable fines by impact fracturing:

Fimp = 2x10" pgh = 2x10-11 mgh/V = 2x10- 1 (E/V) = 2x10-1'p (E/m) (D.8)

where E/V and E/m are the energy per unit volume and per unit mass of the brittle material that is subject

to fracturing and the units of the leading coefficient are cm3

/erg.

Figure D. 1 compares this equation [Firp = 2x 10-1 p(E/m)] to experimental data for the impact fracturing

of a variety of brittle materials including depleted U0 2 and average burnup spent fuel pellets, but no

high-burnup spent fuel pellets. The figure shows that the handbook equation lies about a factor of 10

higher than the experimental data. Accordingly, in order to be somewhat conservative, for average

burnup spent fuel pellets and for the body of high-burnup spent fuel pellets; the leading coefficient in the

DOE Handbook equation is reduced by a factor of five, which gives

Fimp = 4x 10-12 E/V (D.9)

D-10
Ffi,,Y can be calculated from Equation

D.9, knowing the energy that is imparted

to the fuel during the impact. The

accumulation of noble gas atoms in the

pores in the rim layer of high-bumup spent

fuel causes these pores to become highly

pressurized, which means that substantial

amounts of energy are stored in these

pores. Because increase of gas pressure

with time in rim layer pores is relieved by

growth of pore volumes (Reference D. 12),

at any instant in time, the energy stored in

pores will by itself be too small to initiate

growth of grain boundary surfaces by

breaking of U-O bonds or by plastic

deformation of pore surfaces. However, if

propagation of grain boundary cracks

initiated by impact fracturing is supported

by the release of this stored energy, the

extent of grain boundary cracking and the

fraction of respirable particles generated

by fracturing may both be significantly

10S

Extrapolated Respirable Fraction

* , ""0t 44x

V 0

Ii M.1

+,

0.01

IAj A

aAof046

* Jadne (Gen, Syuroc 4

M Ruhmann P33 GWDfTM a &Iecaui4DOopbd U02) x Akvmz (28 QVW~MM

a~ DOE Handbookc

.r

IA* 1648

Specific Energy (ftJg

lOOM 1140.6

higher than the extent of cracking and respirable Figure D.1. Comparison of the DOE

particle production observed for average burmup Handbook Respirable fraction equation to

spent fuel pellets and also for the body portion experimental values of the specific energy

of high-burnup spent fuel pellets. Consequently, input into the brittle material

when applied to the rim layer of high-burnup

spent fuel pellets, the preceding equation needs

to be modified to reflect release of energy stored in the pressurized gas contained in rim layer pores.

None of the material used to develop Equation D.8 was similar to the rim structure in high-bumup fuel

(i.e., very fine grained material having large pores containing high-pressure gas). Using a modified

Equation D.9, and making assumptions about the pore energy, one can to calculate that the rim fracture

and the fracture of the body should be approximately equal. By contrast, anecdotal evidence from

attempts to prepare metallurgical samples of the rim indicated that this region is much more friable than

the body of the fuel. At best, it can be concluded that there is a value of high uncertainty that can range

over several orders of magnitude. This uncertainty must be accounted for in any release determination.

In the current study, the fraction of the rim is being allowed to fracture varies from a low equal to the

same fraction as the body to a high of complete fracture.

D.2.4.1.3 Estimation of F.t,,,,

Because the friable rim layer of high-burnup spent fuel pellets is comprised of recrystallized, highly

compacted, cubical U0 2 subgrains with sides between 0.1 and 0.3 pLm (Reference D. 13), plenum gases

should not be able to cause the recrystallized rim layer to slide toward the nearest rod tear. However, if

during storage the rod cladding separates slightly from the pellet rim, a small rim layer-cladding gap

might open along the length of the rod. If flow through this rim layer-cladding gap during rod

depressurization is significant, some particles formed from rim materials might become entrained in this

D-11
flow and, if not captured by particle bed filtering, be carried out of the rod to the cask interior. In

actuality, entrainment may be insignificant, since any reopened pellet-cladding gap will be clogged with

large particles formed by impact fracturing (mechanical or shock loading), i.e., the large particles will

interfere with the boundary layer flow needed to entrain particles off of surfaces.

Should the conditions for entrainment occur, we have the following relationship:

Fent,rm = RentAtArim /PVrim = Rent Atn dpL/pz dp tnmL=RentAt/Ptrm (D.10)

Data on the variation of entrainment rate Rent with the free stream velocity of the gas flowing over the

powder has been compiled for metal powders (Al, Cu, Mo, Si, and W), and Fe20 3. It also shows that the

entrainment rates of the Tungsten (References D.20 and D.21) and of Fe20 3 (Reference D.22) begin to

turn over at free stream velocities of 5 to 10 m sec"1. Since entrainment data for U0 2 powder is not

available, it was estimated ( Ren, = 5x 104 g cm-2 s-1) by interpolating between the data for Fe20 3 and the

data for Tungsten.

Based on the time to depressurize the rod (see Section D.2.3) and Gelbard's data, Fent,rim can be

calculated. Sprung found that the uncertainty in the calculations led to the physically impossible

entrainment fraction greater than 1. Because Fentn could not be larger than 1 but could be less, Fent,rim

should be set to 1 to obtain a maximum release fraction. Smaller values can be assumed depending on

the extent that the conditions required for entrainment are thought to prevail.

D.2.4.1.4 Estimation of Ftearxrim

If rod failure causes a plug of fuel fines to be blown out of the rod tear from the possibly friable rim layer

located next to the tear, we can estimate Ftearrim using the following equation:

Ftear,rim = W Q d / 7dp tim L (D. 11)

where w, e, and d are the final width, length, and depth of the ejected rim materials and dp, tm and L are

the diameter of the fuel pellets, the thickness of the pellets friable rim layer, and the active length of the

spent fuel rod (the length of the pellet stack in the rod). For a 10 xl0 fuel dp = 0.87 cm, tm= 150 im =

0.015 cm (Reference D.13), and L = 369 cm. If the rod tear has a length/width aspect ratio of 10/1 and

the tear length is equal to half of the circumference of a pellet, the plug length I = 'T dp/2 = 1.4 cm and the

plug width = 0.14 cm.

After the ejection of a rim layer plug, the depressurization flow of rod gases over the sides of the hole left

by plug ejection should entrain more rim material into that flow, which will increase the final width of

the total mass of rim material lost by plug ejection plus entrainment. Since the volume of rim layer

material that will be lost by entrainment is given by Vent = RentAt Aplug sideP and the width of the mass of

materials entrained from one side of the hole is went = Ven/Aplug side,

Went = Rent At/p

Material from both sides of the breach is being entrained, so 2 times went must be added to w to get a

maximum effective breach width (w e = w + 2 went). The actual breach width might be much smaller.

D-12
Thus,

Ftearrim -- We P trim/c dp trim L (D. 12)

For an atrium 10 x10 BWR rod, Ftearrim = 2.8x1004.

D.2.4.1.5 Rod-to-Cask Particle Release Fractions

Values have now been developed for all of the parameters in the expression for the release fraction for

rod-to-cask release of particles. Substitution of these values into Equation D.7 yields the fractional

release in the respirable size range from the rod to the cask.

D.2.4.2 Rod Surface-to-Cask Release Fraction for CRUD

Because impact spallation data for CRUD is not available, consistent with the analysis performed in

NUREG/CR-6672 (Reference D.1), the impact loading of the spent fuel rods was assumed to cause about

10 percent of the CRUD deposits on the surfaces of the spent fuel rods to spall. [Other estimates from

Sandoval are in the same range (i.e., 15%).] Thus, the release fraction for 6"Co in CRUD from the rod

surface to the interior of the cask was assumed to be 0.1. A cumulative number distribution for the

CRUD particles that formed on the surface of a Quad Cities BWR rod Was developed by examination of

a Scanning Electron Microscopy (SEM) photo of the surface of this rod (Reference D.7). Conversion of

this number distribution to a mass distribution indicates that about 15 percent of the total particles mass

is contained in respirable particles (particles with aerodynamic particle diameters dp _<, 10 tim). There are

no data to indicate if the CRUD spallation fraction is independent of the particle mass. Therefore,

FRC,CRUD = (0.1)(0.15) = 0.015.

D.2.4.3 Rod-to-Cask Release Fraction for Noble Gases

The voids and the internal crack network that form in pellet bodies cause these body regions to have

porosities of 5 to 10 percent (References D.13, D.14, and D.16). During reactor operation, about 8% of

the noble gas atoms generated by the decay of fission products from spent fuel pellets with burnups of

55-60 GWd/MTU diffuse to particle grain boundaries and then escape to the rod free volume through the

pellet's internal crack network (References D. 13 and D. 14). Additional fission gas will be released when

the body of the high-burnup spent fuel pellet fractures but the exact amount is highly speculative.

Estimates of the fission gas release upon fracturing the unrestructured material have been advanced as

high as 25%.

In the rim layer, 0.1 to 0.3 nm subgrains are generated by the recrystallization of UOa and gas atoms in

subgrain interiors migrate to subgrain boundaries, where they accumulate in micropores that were formed

by clustering of lattice vacancies (References D. 12, D. 15, and D. 16). Because distances to grain

boundaries are so short in rim layer subgrains, noble gas atoms formed by fission product decay migrate

efficiently to rim layer subgrain boundaries. Consequently, 90% of the noble gas atoms formed in the

pellet rim layer is contained in the micropores that have formed on subgrain boundaries (Reference

D. 14). Once again, the release of this gas resulting from the fracture of the rim upon impact is an

unknown quantity. If the rim disintegrates, as postulated by some, essentially all of the fission gas in the

rim region being released is a possibility.

D-13
Accordingly, the rod-to-cask release fraction for noble gases from high-bumup spent fuel pellets was

calculated as the sum of the fraction of the inventory in the pellet rim and body that is released by

diffusion plus impact fracturing with each quantity weighted by the fraction of the total inventory that is

contained in each region.

Fission gas is contained in three regions in a high-bumup rod: the void volume, the fuel rim, and the

body of the fuel. There would be no rim in a lower-burnup rod. The size of the rim and the percentage

of fission gas that is in the void volume will be a function of the burnup. Because of the neutronics in the

reactor, the concentration of actinides in the rim is about twice that in the body of the fuel. Assume a

fraction Z of the produced fission gas is released to the void volume during irradiation. Because of the

radial temperature gradient in pellets during irradiation, the fission gas accumulated in the void volume

tends to come from the hotter central body of the fuel pellet so the density of fission gases in the rim are

about twice that in the central body of the pellet. (1-Z) of the fission gas will be left in the fuel column to

be partitioned between the rim and the body. Let the following values apply:

PR = density of fission gas in the rim region

PB = the density of fission gas in the body

RR = fraction of the rim that fractures

RB = fraction of the body that fractures

RRg fraction of gas released per volume of fractured rim

RBg = fraction of gas released per volume of fractured.pellet body

X = fraction of the pellet volume (V) that is rim material; VR = X V

As indicated above, PR = 2 PB. It can be shown that the overall density of fission gas in the pellet (p) can

be given by p = PB (l+X).

Let I be the total inventory of fission gas. Then, I = IR + IB' where 'R and IB are the inventories of fission

gas in the rim and body respectively.

After irradiation, the total gas inventory is given by the following equation:

I = gas in gap + gas in rim + gas in pellet body

= ZI + 2 PB VR + (PB VB - ZI)

After an impact, the rim and the body may fracture. When the cladding is breached the total gas in the

gap, ZI is available for release as well as gas that is released from the grain boundaries and body of the

rim and pellet body. The gas released (If) is given by the following equation:

If = ZI + RR RRg 2 p X V/ (+X) + RB RBg [p V (l-X)/ (I+X) - Z I]

Let the fraction of fission gas released from the rod after fracture to the cask be given by RRcg = I/I

Since p V = I, RRcg is given by the following equation:

RRcg = Z + RR RRg 2 X / (1+X) + RB RBg [(l-X)/ (1+X) - Z]

If typical values for high-burnup fuel, Z= 0.08 and X = 0.1, are substituted, the equation becomes as

follows:

D-14
RRcg = 0.08 + 0.18 RR RRg + 0. 7 4 RB RBg

The gas release will be highly dependent on the magnitude of the event and the amount of fuel that is

fractured. If it is assumed that there are 5 breach locations in the cladding, and that ±5 cm around the

breach site are fractured and that 25% of the gas in the body and rim are released, the equation predicts a

release of -12% of the fission gas. This is similar to the 10% predicted by Sprung. By contrast, if it is

assumed that the whole rod fractures and that 80% of the gas is released from the rim region, -40% of

the gas in the rod is released. This illustrates how the uncertainty in the fracture properties of the

cladding and the rim introduce uncertainties for the driving force for the release from both the rods and

the cask. Until more definitive data are available, the choice of rim fracture fraction is at the discretion

of the analyst. This parameter will be varied over a range for this study.

D.2.5 Cask-to-Environment Release Fractions (FCE,k)

After the fuel particulate and CRUD are released from the rod to the cask atmosphere, a portion will

deposit within the cask and the remainder will be released from the cask in proportion to the volume of

gas that is released to depressurize the cask to atmospheric pressure. The fraction of the material in

chemical element group k that escapes from the cask to the environment, FCE,,k, can be estimated using the

following expression (Reference D. 1):

FCE k = (I - F deposition k ) I P P atm (I - Fdeposition ,k )Fexp (D.13)

S P rod failure

where:

Fdeposition,,k= the fraction of the materials in chemical element group k that is deposited onto

cask interior surfaces after release to the cask atmosphere from failed rods

Patm = 1.0 atm = atmospheric pressure

Prod fail = the pressure that the depressurized cask would reach upon rod failure, if release

of rod gases to the environment were not occurring

Fexp = the fraction of the gases in the cask atmosphere that is released to the

environment by cask redepressurization following rod depressurization

This equation must be used with care since it is a static approximation of a dynamic event. Two

simultaneous venting events are taking place; the breached rods into the cask, and the cask to the

environment. Because of the large pressure and volume differences (-50 atm and 20 cc for the rods, and

a few atmospheres and 6,000 liters for the cask) the size of the leak path becomes very important. In the

high-burnup rods, the leak path is governed by the hydraulic diameter that results in a very small leak

path and long depressurization time. This may not be the cask for severely shattered rods or lower

burnup rods. The cross-sectional area of the leak from the cask will depend on the size of the impact, and

cover a wide range of sizes (Reference D. 1).

D-15
If the cask depressurizes rapidly with respect to the rod depressurization, the only material available for

release will be the CRUD, and rim and body particulate initially ejected from the rod. The driving force

for the removal of the material from the cask will only be the temperature dependent fill gas pressure

differential with the cooler environment. The release of fission gas to the cask will not play a part in the

release of this material. The particulate that is removed from the rod as a result of entrainment in the gas

stream will have a much small driving force to remove it from the cask, because the pressure buildup in

the cask resulting from rod depressurization will be small. By contrast, if the cask depressurizes slowly

with respect to the rod depressurization, the driving force will include the full pressure (Prod failure )

calculated below. In addition, the entrained material will also be swept out with this depressurization.

D.2.5. 1 Fractional Release Fraction of Cask Atmosphere (Fexp)

The fraction of the cask atmosphere released to the environment is given by the following equation:

Fexp 1 - Patm ( P rod failure (D. 4)

Prod failure is calculated using the ideal gas law. Define the following values:

Prod = the pressure of He and fission product noble gases in a rod at ambient conditions

including fission product noble gases released by pellet fracturing

Vrod = the free volume of a spent fuel rod

nrod = the number of moles of He plus fission product noble gases in a 10-year cooled

high-burnup spent fuel rod

Pcask = the pressure of He fill gas in the cask at ambient conditions

Vcask = the free volume of the cask

ncask = the number of moles of He fill gas in the cask

R = the ideal gas law constant

T = the average temperature of the cask interior and the spent fuel rods in the cask at

.ambient conditions

Pful = Prod failure

Nrods = the number of spent fuel rods in the cask

Vfin -- the sum of the cask free volume plus the total free volume of the failed rods

The moles of gas in the breached rod is given by ProdVrod = nrodRT. The number of mole of gas originally

in the cask is given by PoaskVc.ak = n..cakRT. These can be combined (PfnIVI, =n, 1 RT) to determine the

amount of gas that is pressurizing the free volume in the cask and the now-breached rods. These are

given by the following equations:

nfinal = FrodsNrodsnrod + ncask

Vfinal-- FrodsNrodsVrod + Vcask

D-16
Substitution yields the final pressure in the cask after rod failure but prior to rod depressurization.

FrodsN rods Prod Vrod + Pcask Vcask

Frods N rods Vrod + Vcask

The parameters required to calculate the final cask pressure can be obtained from cask safety analysis

reports, fuel rod performance reports, and so forth.

The size of the impact event will determine the size of the opening in the cask and which of the two

above scenarios, or one in between, should prevail. An upper bound for release is obtained if one

assumes that the rods depressurize into the cask rapidly to build up pressure before significant cask

depressurization occurs. Under this scenario, Sprung used cask dimensions to calculate a final pressure

in a cask of 1.5 atm, up from the initial backfill pressure of 1 atmosphere. At a final pressure of 1 atm,

Fexp is -333%. For the HI-STORM 100 cask, with an initial backfill pressure at temperature of-5.6 atm,

depending on the number and type of failed rods, the amount of expelled gas can range between 82 and

98%.

D. 2.5.2 Deposition Mechanisms onto Cask Internal Surfaces

Because gas flow rates out of the failed rods are high, removal of particles by inertial impaction onto

spacer plates or rod surface located next to cladding failures may occur, provided particle sticking

probabilities per collision are significant at the particle impact speed. Those particles not sticking as a

result of impaction may settle before reaching the cask opening. Because rod depressurization is slow

(see Section D.2.3), gas flow velocities through the failed cask will also be slow, so particle removal

during transport through the cask to the location of the cask's closure failure will occur by Brownian and

turbulent diffusion to cask interior surfaces and by gravitational settling onto upward facing interior

surfaces. If rod depressurization occurs through a reopened pellet-cladding gap, depressurization may

not be slow (though it still may not be fast enough to make inertial processes important). Removal of the

larger respirable U0 2 particles (U0 2 particles with geometric diameters close to 3.16 jim) will occur

primarily by gravitational settling, while small particles, are removed primarily by Brownian and

turbulent diffusion.

D.2.5.2.1 Impingqement on Adjacent Rods

Sprung calculated the typical impingement distance from a rod breach to the next nearest surface for a

variety of casks and concluded that in almost every instance that the largest size respirable particulate

(3.16 ý.im diameter) released from a rod would impinge on the neighboring surface. However, the impact

velocity would be so great ( > 4.3x i04 cm/s) that the sticking coefficient for U0 2 particles is most likely

_• 1 x 10- (Reference D.23). Hardly any of the particles that impact the spacer plates will stick to the

plates, and, therefore, Finerial impaction = FswraceFstick = (1.0)(0.0) = 0.0. Because of the uncertainty of the

sticking factor, no deposition attributable to impaction is recommended.

D-17
D.2.5.2.2 Gravitational Settlinq

After a drop impact, there are two possible configurations of the cask: (1) a tip-over with the cask lying

horizontally on its side, or (2) the cask remaining in a standing vertical orientation with the breach in the

bottom weld:

(1) Horizontal: During cask depressurization, one side of each of the spent fuel assembly basket

tubes in the cask, and half of the outer surface area of the spent fuel rods in each tube will be

facing upward and, therefore there will be surfaces onto which particles can deposit by

gravitational settling. The force on the particulate in the gas stream attributable to the gas flow is

in the horizontal direction. The force attributable to gravity is orthogonal to the gas flow in the

vertical direction. If the time to traverse the distance to the adjacent rod or basket surface is less

than the time for the particle to be swept to the breach site, the particle can be considered to be

gravitationally settled.

Under quiescent conditions, the gravitational settling velocity of 3.16 Itm UO2 particles is about

0.2 cm/sec (Reference D.24). Since the diagonal separation distance between the 1Ox 10 fuel

rods in an assembly is -0.8 cm, it takes about 5 sec for deposition by gravitational settling of

3.16 ['m UO2 particles under quiescent conditions. This is short compared to the time required to

depressurize a rod (see Section 2.3). If rod depressurization occurs through a reopened pelletcladding gap this assumption might change. Since the flow conditions in the cask during rod

depressurization are relatively quiescent, and every 3.16 [Im UO2 particle will encounter a

horizontal surface in about 5 sec, removal of respirable U0 2 particles by gravitational settling

will be efficient.

Under the relatively quiescent that will prevail during the passage of rod gases and fission

products from rod failures to the cask closure failure, any respirable U0 2 particles not removed

by inertial impaction should encounter a horizontal surface before they exit the cask. Since the

sticking coefficient for 3.16 [im UO2 particles at their gravitational settling impact speed is

certainly Ž_0.9 (Reference D.23), Fgravitational settling = FsurfaCeFstick = (1.0)(0.9) = 0.9.

(2) Vertical: In this configuration, both the gas stream and gravitational force are pushing the

entrained particles down until they meet the first horizontal surface, which is the grid spacer with

its horizontal projection of the spring surfaces, and plates. The fuel assembly is made of seven

spans between grid spacers. We can assume, similar to the horizontal case, that at least 90% of

the particulate settle at the grid spacer. If the cladding breaches have an equal probability of

forming anywhere on the rod, 1/7 of the fractures will form in each span. If 90% of the entrained

particles settle on a grid spacer, the total gravitational settling can be expressed by the following

equation:

7

Fgravitational settling :1 - 0.143 x I (1 _ 0 .9 )n = 0 . 9 8 4 n=l

As a bounding case, breaches may not be equally distributed along the rod but may be expected to

concentrate on the lower grid span. The particles will settle only on the lower end plate of the assembly

* and the 90% settling factor is more probable and will be used in this analysis. This analysis does not take

into account the tortuosity of the path through the seal.

Similarly, after spallation from rod surfaces attributable to impact forces, CRUD particles will also be

removed by gravitational settling onto cask and rod surfaces.

D-18
For both the respirable fuel and CRUD particles, the fraction of particulate that escapes the cask to the

environment is given by

FCE, (CRUD, fuel)= 0.1 x (depressurization ratio for the particular situation)

FCE, (CRUD, fuel) may range from 0.03 3 to 0.1 depending on the cask design and fuel loading. For the

HI-STORM 100 cask, which has a depressurization ratio between 0.82 and 0.98, depending on the

number of breached rods, FCE (CRUD, fuel) is between 0.08 and 0.1. Since noble gases won't settle, the

escape fraction is given by the following equation:

FCE, g. = depressurization ratio

D.3 Input Parameters To Model

D.3.1 Fuel, Cladding, Cask, and Storage Environments

All of the fuel, cladding, and cask physical parameters and tolerances that are used in the model are well

known although possibly hard to locate or proprietary. Although no systematic evaluation of the

influence of the uncertainty or variation of these parameters on the release rate has been made, it is not

suspected that parameters such as cladding, thickness, internal rod volume, etc., will have a large

influence. By contrast, internal cask pressurization or cask free internal volume, for example, will

probably have a large effect, as they have a large effect on the driving force from the cask cavity to the

environment.

D.3.2 Fuel Performance Parameters

The changes that occur in the properties of the fuel and the cladding while in-reactor may introduce large

errors into the determination of the release factors, because of the uncertainty of the database. Table D.3

identifies some of the changes in the fuel attributable to burnup and briefly describes the uncertainties

introduced by these changes. No attempt has been made to quantify the degree of the uncertainties or to

determine if they are significant to the risk.

D-19
Table D.3. Fuel Characteristic Changes That Occur at High-Burnup

That Can Introduce Uncertainty into the Release Fractions

Bumun

Gap The pellet-cladding gap closes as the burnup increases. The partitioning of the gas flow and entrainment of

fuel between the rim and the body will be dependent on the gap size. Until better data on the gap at room

Gas generation, and The fission gas release from the pellet to the rod void volume increases with burnup. Circa <2% at lower

release bumups, it may rise to 8-10% at high bumups. This additional gas is available for driving particulate from

the fractured rods, and increases the pressurization of the cask resulting in more particulate release to the

environment. The release from the cask is directly proportional to the release from the pellets

Fission gas release The majority of the generated fission gas is trapped in the grains of the fuel pellet. This gas is available for

upon fracture release when the pellet fractures. There are no good measurements of the amount that is released upon fuel

fracture. Sprung uses 10% release in his calculations, but release may be as high as 25% or more in the

impacted fuel regions

Rim development At about 50 GWd/MTU average bumup or 100 GWd/MTU localized burnup, a rim forms on the exterior of

the fuel pellet. The rim has a fine grain structure (submicron) and retains much of its fission gas in large

pores

Cladding

mechanical

properties

During irradiation, the cladding thins as a result of oxidation and embrittles as a result of irradiation and

hydrogen uptake. Mechanical properties that lead to fracture will change with bumup and temperature.

The number of rods that fracture and the number of fracture sites per rod will depend on the cladding mechanical nronerties mechanical_ n--".....

Cladding type The fracture properties of the newer cladding types and newer heat treatments of older cladding types has

not been well developed

Rim

Size The size of the rim will increase exponentially with bumup. It is about 50% larger in BWR fuel at

equivalent bumup.

Inventory The actinide inventory in the rim is -twice that in he body of the fuel pellets

Fracture during Based on pore pressurization, the rim should fracture in a similar manner to the body of the pellet.

accident Anecdotal evidence and expert opinion tends to indicate that the rim should fracture rather easily and into

finer particulate than the body of the pellet. The maximum respirable fraction will occur if the entire rim

was fractured into submicron size particulate. In this case, the release fraction could increase by a factor of

3. For lower energy of impact, the difference could be considerably greater because of the small fracture of

the body of the fuel. No estimate of the energy input required to do this has been made and itis highly

unlikely complete fracture of the rim will occur.

Fracture in reactor Experiments similar to Lorenz's have not been conducted on high-bumup fuel with a rim. No data are

available on the fracture and particle size distribution of the rim as it comes out of the reactor as a result of

thermal cycling in the reactor and subsequent handling. For this exercise, it assumed that fracture similar to

that in the body of the pellet has occurred in reactor. Because of the uncertainty in the fracture of the rim

during the fall, the uncertainty in this parameter may be unimportant.

Particulate size The rim is -10% of the fuel volume. The grains are 10-100 times smaller in diameter. Most of the rim

distribution would fracture into respirable size particulate because of the grain size.

Cask Drying During drying, hydrides may reorient to radial direction further decreasing ductility leading to more

fractures and/or more fracture sites per rod.

CRUD

amount Most data is pre 1990 and based on lower burnup fuel. At that level, no bumup dependence was found;

however, this may change at higher bumups. Some assemblies have shown very large CRUD deposits,

enough to span fuel rods. Also, some rods are showing azimuthal dependence to CRUD distribution.

Size distribution It is unknown whether CRUD fractures into small particulate on the micron size range as shown in

photomicrographs or stays agglomerated in flakes. If the latter, it will settle very quickly in the casks

Spallation 10% spallation of CRUD is assumed. No data on spallation from impacts are available. Some work on

thermal spallation indicates 15%. There is no technical reason that this should be the same for impacts

Fuel pellet fracture database Little fracture size distribution data is available on irradiated fuel pellets. The DOE Handbook database

seems to indicate that both glasses and ceramics fracture in a similar manner. Newer work appears to

indicate that the Handbook values over predict the fraction of respirable by at least an order of magnitude.

D-20
D.4 Release Fractions to Environment

The model can'be applied to any cask, and fuel loading to determine the potential release fractions for

that particular situation. The model can also use ranges of input parameters and combinations, with

suitable weighting factors to determine the potential range of release fractions. Based on the model, the

release fractions for a HI-STORM 100 cask, containing 68 lOx 10 BWR fuel assemblies with a bumup of

60 GWd/MTU that is dropped 30.5, 12.2, and 1.52 meters (100, 40, and 5 feet) were calculated and

tabulated in Table D.4.

Table D.4. Values of Fr,,,, F F and Fr__k

Chemical Element Group

(k) Inventory Frods Fgc.k Fc..k * Frelk = FOr c .kFCE .k

Noble Gases ORIGN I to 0.33 0.12 to 0.4 1 4x102 to 0.4

Particles ORIGN I to 0.33 Table 5 0.1 (0.1 to 0.033) x value

in Table 5

CRUD 0.72 Ci/rod* 1.0 0.015 0.1 1.5 x 10"3

* based on 10 yr old cooled 1OxO fuel, bounding 90% of the fuel inventory

** HI-STORM 100 is pressurized to 5.6 atm at temperature resulting in 82% release. Rod breaches will

increase this to 98% release of gas. Other casks may have lower values because of lower initial cask

pressurizing.

FRCk was calculated for the above cask and fuel for a range of nlowvtear, rim fracture and the rim

entrainment. At a minimum, the rim was assumed to fracture in the same respirable fraction as the body

of the fuel. At the opposite extreme, the rim was assumed to fracture into 100% respirable fraction since

the rim consists mainly of submicron grains. The entrainment ranged from zero to complete entrainment.

Results are shown in Figures D.2 and D.3. At this point, a distribution should be assigned to both

variables and a sampling done to obtain a most likely value and uncertainty using a probabilistic

calculation. Based on these plots and others like it, four release fractions can be assigned to each drop

height for the extreme values of the rim fracture and the entrainment. These are shown for a sample case

in the 2x2 matrix in Table D.5 below. Other values can be obtained from the plots or by calculation but

since there is currently no technical justification to chose an intermediate value for the rim fracture or

entrainment, this table is made for bounding simplicity.

D-21
100 ft Drop

1.00E-01

1.OOE-02

---- 1 Tear

S1.OOE-03 -e3er -i ,••, •5 Tears

• : -x--7 Tears

1.00E-04 -• , , ....

1.00E-05 - - -

0.0001 0.001 0.01 0.1

Fraction of Rim Layer Fracture

Figure D.2. FRck for 100% Entrainment of the Rim

as a Function of the Number of Tears and Fraction of Rim Layer Fracture

D-22
5 Tears per Rod

1.00E-02

C

.IT

3a •100ft

1 .OOE-03 ---- 40 ft

5ft

1.OOE-04

1.O0E-04 1.O0E-03 1.O0E-02 1.00 E-O1 1 00E+00

Fraction of Particles Entrained

Figure D.3. FRC,k for 100% Rim Fracture and 5 Tears per Rod

as a Function of Fraction of Particles Entrained

Table D.5. Variation of FRC,k with Input Parameters for 100-Foot Drop and 5 Breaches per Rod

Entrainment, F•,L•

None = 0 Complete = 1

Rim fracture, FLp.• Same as body of fuel =1.24 x 60-4 5.6 x10-' 6.8 x 10.O

Complete fracture = 1 2.2 x 101 1.2 x 10.2

Lacking a probabilistic calculation, a deterministic evaluation is being done in this report to illustrate the

methodology. Without experimental data, it is difficult to justify the value in Table D.5 to choose for

FRCk. Entrainment requires a fuel cladding gap for the gap to flow through and pick up rim particles.

High-burnup fuel has little or no gap and, as a result the gas flow, will occur between the fractured pellet

fragments. Even with a small gap, there should be sufficient gas flow to entrain the particles; therefore,

full entrainment probably occurs. Because of the very limited information on rim fraction, no technical

justification can be given for choosing an FRCk value based on rim fracture. Plots and tables similar to

Table D.5 and Figures D.2 and D.3 were constructed for a variety of drop heights and fracture sites per

rod. The results are tabulated in Table D.6 below.

D-23
The most conservative position is to use the highest release fraction from Table D.6, but that may be

considerably higher than a realistic value. The overall release from the cask can be calculated using

Equations D. 1 and D.2. The values for use in these equations for this particular case of the HI-STORM

cask with a 30.5 meter (100-foot) drop are given in Table D.7.

*

Chemi

Gr

Noble

Parti

CR]

F

F

F

Table D.7. Values of Release Fractions Frods, FRC,k, FCE,k, and Frel,k*

(10-year-old cooled 10x10 Atrium fuel, 30.5 meter (100-foot) drop of a HI-STORM 100 cask,

all of the rods fail with an average of 5 breach Iocations per rod)

cal Element FreLk =FrodsFRckFcE,k

roup (k) Inventory Frjd. FRC k Fr, k respirable

Gases (k=l) ORIGEN 1 0.12 1 0.12

icles (k=2) ORIGEN I 7x10 5 to 1.2x10 2 0.1 7x10"6 to 1.2x10 3

JD (k=3) 0.72Ci/rod 1 0.015 0.1 0.0015

ros = the fraction of rods contributing to the release [1.0 (100% of the rods, see Table D.4)].

Ck = the fraction of material released from a single rod to the cask environment.

CE,k = the fraction of respirable radioactive material in the cask atmosphere that is released to

the environment and Frelk = fraction of the radioactive material in the cask that is released

to the environment in the respirable size range.

D.5 Conclusions

If the imparted energy is sufficient when a cask is dropped, a leak path can open in the cask and

radionuclides released from damaged fuel. The release can be in the form of noble gases, CRUD, or fuel

particulate. The driving force for release from the cask to the environment is the pressure differential

between the cask cavity and the environment outside the cask.

A linear model was built to determine the release fraction of the gas, CRUD, and particulate. The model

consists of four units: rod radionuclide inventory, number of rods that will breach to release material,

release of material from the rods to the cask cavity, and release of material from the cask cavity to the

environment. All of these components will depend on the type of cask, characteristics of the fuel loaded

into the cask, and magnitude of the impact force. A brief discussion is given on the uncertainty in these

input parameters. By far, the biggest uncertainties are the general fracture characteristics of irradiated

ceramic U0 2 fuel, the fracture characteristics of the rim region in high-bumup fuel, the variation of

mechanical properties of the irradiated cladding, and deposition characteristics of the particulate in the

cask. This model has not been benchmarked against experimental data. As a result, some of the

phenomena submodels such as particulate entrainment may be significantly off. In these cases, an

attempt was made to determine the uncertainty introduced into the final results by varying the potential

output of the submodel over a bounding range.

D-24
Output of the model was presented for a typical cask and fuel loading to illustrate the expected

magnitude of the release fractions. The quantity of radioisotope released from the cask is the product of

the inventory, contained in or on the rods, of that isotope in the cask times the fraction of the inventory

that is released from the cask. Based on Tables D.5 and D.6 above, Table D.7 values should be used for

the consequence analysis.

D.6 References

D. 1. Sprung, J.L., et al., "Reexamination of Spent Fuel Shipment Risk Estimates," NUREG/CR-6672,

U.S. Nuclear Regulatory Commission, Washington DC, March 2000.

D.2. Croff, A.G., "ORIGEN2: A Revised and Updated Version of the Oak Ridge Isotope Generation

and Depletion Code," ORNL-5621, Oak Ridge National Laboratory, Oak Ridge, Tennessee,

July 1980.

D.3. Oak Ridge National Laboratory, "ORIGEN2 Isotope Generation and Depletion Code," CCC-371,

Oak Ridge, Tennessee, 1991.

D.4. International Atomic Energy Agency, Safety Series No. 7, "IAEA Safety Guides:

Explanatory Material for the IAEA Regulations for the Safe Transport of Radioactive Material"

(1985 Edition), 2 nd Edition, Vienna, 1987.

D.5. Code of Federal Regulations, Title 49, Section 173.435 (49 CFR 173.435).

D.6. Hazelton, R.F., "Characteristics of Fuel Crud and Its Impact on Storage, Handling, and Shipment

of Spent Fuel," PNL-6273, Pacific Northwest Laboratory, Richland, Washington, September

1987.

D.7. Sandoval, R.P., et al., "Estimate of CRUD Contribution to Shipping Cask Containment

Requirements," SAND88-1358, Sandia National Laboratories, Albuquerque, New Mexico,

January 1991.

D.8. Geelhood, K.J., and C.E. Beyer, "Examination of Strain to Failure Data from Irradiated

Zircaloy," white paper prepared for the NRC's Spent Fuel Package Vulnerability Study, 2004.

D.9. Geelhood, K.J., and C.E. Beyer, "Mechanical Properties for Irradaiated Zircaloy," Transactions,

Vol. 93, ANS Winter Meeting, Washington, DC, November 2005.

D.10. Vankeerberghen, M., "Recent Gas Flow Measurements in IFA-504 Rods," HWR-433,

Halden Reactor Project, Halden, Norway, November 1995.

D. 11. Bird, R., W.E. Stewart, and E.N. Lightfoot, Transport Phenomena, John Wiley and Sons,

New York, New York, pp. 480-482, 1960.

D-25
D. 12. Manzel, R., and C.T. Walker, "EPMA and SEM of Fuel Samples from PWR Rods with an

Average Burnup of Around 100 MWd/kgHM," Journal of Nuclear Materials, 301:170,

Elsevier, Inc., Burlington, Massachusetts, 2002.

D. 13. Einziger, R.E., and C.E. Beyer, "Characteristics of High-Burnup Fuel that May Affect the

Transportation Source Term," white paper prepared for the NRC's Spent Fuel Package

Vulnerability Study, 2004.

D. 14. Manzel, R., and C.T. Walker, "High-Burnup Fuel Microstructure and its Effect on Fuel Rod

Performance," Proc. Intl. Topical Meeting on LWR Fuel Performance, Park City Utah,

April 2000.

D. 15. Thomas, L.E., C.E. Beyer, and L.A. Charlot, "Microstructural Analysis of LWR Spent Fuel at

High Burnup," Journal of Nuclear Materials, 188:80, Elsevier, Inc., Burlington, Massachusetts,

1992.

D. 16. Spino, J., K. Vennix, and M. Coquerelle, "Detailed Characterization of the Rim Microstructure

in PWR Fuels in the Burnup Range 40-67 GWd/tM," Journal of Nuclear Materials, 231:179,

Elsevier, Inc., Burlington, Massachusetts, 1996.

D. 17. Lorenz, R.A., et al., "Fission Product Release from Highly Irradiated LWR Fuel,"

NUREG/CR-0722, Oak Ridge National Laboratory, Oak Ridge, Tennessee, pp. 48-80,

February 1980.

D.1 8. Otani, Y., Aerosol Science & Technology, 10:463, American Association for Aerosol Research,

Mt. Laurel, New Jersey, 1989.

D. 19. U.S. Department of Energy, "DOE Handbook: Airborne Release Fractions/Rates and Respirable

Fractions for Non-Reactor Nuclear Facilities," DOE-HDBK-3010-94, Section 4.3.3, "Free-Fall

Spill and Impaction Stress," Washington, DC, December 1994, p. 4-52.

D.20. Iversen, J.D., "Particulate Entrainment by Wind," Project 1699, ISU-ERI-Ames-85010,

Department of Aerospace Engineering, Iowa State University, Ames Iowa, Report prepared for

Los Alamos National Laboratory, 1984.

D.21. Wright A.L., and W.L. Pattison, "Results from Simulated Upper-Plenum Aerosol Transport and

Aerosol Resuspension Experiments," Proceedings of the CSNI Specialists Meeting on Nuclear

Aerosols in Reactor Safety, 4 'h to 6 th September, 1984, Karlsruhe, FRG, Ed. W.O. Schikarski and

W. Schock, CSNI 95, KfK 3800, February 1985.

D.22. Fromentin, A., "Particle Resuspension from a Multilayer Deposit by Turbulent Flow,"

PSI-Bericht Nr. 38, Programm LWR-Sicherheit, Paul Scherrer Institute, Switzerland, 1989.

D-26
D.23. Beal, S.K., "Correlations for the Sticking Probability and Erosion of Particles," Journal of

Aerosol Science, 9:455, Elsevier, Inc., Burlington, Massachusetts, 1978.

D.24. Reardon, P.C., G.S. Brown, and Y.R. Rashid, "On the Particle Size Distribution of Crushed

Spent Fuel," 3r' International Radiological Waste Management Conference, Las Vegas, Nevada,

1992.

D-27

APPENDIX E

CONSEQUENCE ANALYSIS

APPENDIX E

CONSEQUENCE ANALYSIS

E.1 Introduction

This appendix presents a more detailed explanation of the consequence analysis introduced in Section 6.0

for the dry cask probabilistic risk assessment (PRA). This consequence analysis assesses a severe accident

scenario related to an inadvertent 30.5 meter (100-foot) drop of a HI-STORM 100 cask system during the

transfer process (Reference E. 1). The consequence measures chosen for this study were the mean values

given by the MELCOR Accident Consequence Code System (MACCS2) for the individual risk of early

fatality and the individual risk of cancer fatality within 16 km (10 miles)'. The value of 1.6 km (1 mile)

is the measured distance from the exclusion area boundary (EAB), assumed to be 0.5 km (0.3 mile), out

to a distance of 2.1 km (1.3 mile). These consequence measures were chosen to allow comparison to the

NRC's quantitative health objectives for reactor accident risk. Because cask accident consequences in

terms of individual risks of early fatality and cancer fatality are small, the consequence mean value

calculated for individual lifetime dose commitment (the mean value of the variable known in MACCS2

as "peak dose found on spatial grid") may also be of interest. Therefore, this report includes the mean

individual lifetime dose between 1.2 and 1.6 km (0.75 and 1.0 mile) as measured from the release point.

Also, close-in consequence estimates are generally more dependant on the source term and evacuation

timing, while consequences further away tend to be controlled by long-term relocation assumptions. The

MACCS2 consequence analysis performed for this study does not consider sheltering and evacuation.

E.2 Approach

This assessment was performed using MACCS2 (Reference E.2) for a representative site. Site-specific

data important to modeling a HI-STORM dry cask 30.5 meter (100-foot) drop accident scenario in the

MACCS2 consequence calculation were collected and used. The important parameters/variables

required to model the site are the population density/distribution and the site meteorology. The

radionuclide inventory, source term (i.e., release fraction, release start time, and release duration), initial

plume dimensions (related to the system geometry), and plume heat content are discussed below. Other

settings and models necessary for a MACCS2 calculation (e.g., food chain model) were taken from the

NUREG-1 150 study MACCS2 input file prepared for the Surry Power Station. The input file is

documented in Appendix C to the MACCS2 code manual and is referred to there as Sample Problem-A.

Sample Problem-A was chosen because of its ready availability and previous use by the NRC in scoping

studies of various types. Discussions of some of the values used for the MACCS2 calculations are given

in Sections E.3 - E. 11.

Unless otherwise stated, all consequence measures given in this appendix are mean values calculated using the

MACCS2 code. The mean value in MACCS2 is defined as the average (expected) consequence over all weather trials

(Reference E.2).

E-1
E.3 Dose Conversion Factor

The dose conversion factors (DCFs) used in this study are the same as those used in the NUREG-1 150

study (DOSDATA.INP). The DCFs were provided by K. Eckerman from Oak Ridge National

Laboratory (Reference E.3). These DCFs include definitions for 19 organs and 60 radionuclides

important in reactor accidents.

E.4 Radionuclide Inventory

High-burnup spent fuel is placed inside a HI-STORM cask no sooner than 10 years after being removed

from the reactor. A maximum of 68 boiling-water reactor (BWR) assemblies can be loaded in a

HI-STORM cask. The inventory per cask calculated with the ORIGEN code (References E.4 and E.5)

is presented in Table E. 1. The analysis assumes that the spent fuel placed in HI-STORM cask for the

30.5 meter (100-foot) drop accident is a 10-year-cooled high-bumup spent fuel.

Table E.1. Dry Cask Nuclide Inventory: SNL ORIGEN 50,008 MWD/MTITLM (10-yr-cooled)

Nuclide Bq Ci Nuclide Bg Ci

Co-60 1.61 X 1014 3133 Pu-238 3.98x1015 107440

Kr-85 2.77x 10"5 74800 Pu-239 1.87x 1014 5060

Y-90 3.40x 016 918000 Pu-240 3.47x 10" 9384

Sr-90 3.40x 1016 918000 Pu-241 5.23x1016 1414400

Ru-106 2.92x1014 7888 Am-241 1.20x10's 32504

Cs-134 5.13x1015 138720 Am-242m 1.97x1013 532

Cs-137 5.54x 1016 1496000 Am7243 3.07x 1013 830

Ce-144 5.08x 1013 1374 Cm-243 3.02x 1013 816

Pm-147 3.37x10" 1 91120 Cm-244 5.66x1015 153000

Eu-154 4.15x10"5 112200

The radionuclide inventory in Table E. I was generated using the ORIGEN code, corresponding to an

8x8, 50 GWd/MTU, 10-year cooled high-burnup BWR spent fuel. This inventory was used to

demonstrate the methodology of developing a consequence analysis. As recommended in Appendix D,

specific radionuclide ORIGEN calculation should be performed for the actual spent fuel loaded inside the

HI-STORM casks.

Cobalt-60 is formed in spent fuel by activation of Ni-60 in cladding, but its most important source is in

the CRUD that forms on cladding surfaces during reactor operation (References E.6 - E.8). During a

30.5 meter (100-foot) cask drop impact accident, the amount of Co-60 that was formed by activation of

Ni-60 in cladding will tend to remain trapped where it formed, while the Co-60 in CRUD can be released

by spallation. Therefore, the Co-60 inventory in Table E.4.1 represents the CRUD inventory. The value

was calculated from the Appendix D inventory of 0.72 Ci/rod for a l0x 10 fuel assembly (see Table D.7).

This value was assumed to be similar for the inventory used in this analysis. The value for CRUD in

Table E.1 is for 64 BWR rods per assembly and 68 assemblies per cask.

E-2
E.5 Chemical Element Groups

Separate release fractions need not be developed for each element in a dry cask inventory. The releases

for the cask drop accident are from mechanical means, not because of differing volatilities. Accordingly,

release fractions need to be developed only for three chemical element groups: noble gases, CRUD, and

particulates. Inspection of Table E. 1 shows that the noble gas chemical element group contains only

Kri8, the CRUD chemical element group contains only Co-60, and all other elements in the canister

inventory are members of the particulate chemical element group because these elements can only escape

from failed fuel rods as constituents of fuel fine particles. These chemical elements groups for the

30.5 meter (100-foot) drop accident scenario, are consistent with the analysis and methodology presented

in Appendix D.

E.6 Source Term

E.6.1 Release Fraction

The release fractions for the calculations are given in Appendix D. Two different source terms were

evaluated. The first represents a large release, one that assumes little reduction in fuel fines escaping

from the cask while the second represents a median estimate of fuel fine reduction in the escape path.

Further, for those cases where filters are assumed to be operable, a reduction of a factor of 10 for

particulates and CRUD has been assumed. In all cases, no credit is given for deposition within the

reactor building. Table E.3 includes the release fraction for each of the consequence results analyzed.

E.6.2 Release Start Time

The MACCS2 calculation assumed that the release to the environment occurs immediately.

E.6.3 Release Duration

The release duration used in the MACCS2 consequence calculation was assumed to be 2 minutes.

A methodology for estimating this value is presented in Appendix D.

E.7 Release Height

The dropped cask scenario occurs inside the reactor building. Two different release locations are

possible. The first one is through vents located 50 meters (164 feet) above ground. These vents contain

filters which, when operating, could reduce the release as discussed above. The second release path of

radionuclides to the environment is through the stack. The opening from the stack to the environment is

located at a height of about 120 meters (394 feet) from ground. Table E.3 include the release height and

the reduced release fraction from filters, as applicable, for each of the consequence results analyzed.

E-3
E.8 Initial Plume Dimension

The initial plume dimensions (i.e., the plume dimensions at the release point) are determined by the

width and height of the building wake. In MACCS2, the initial plume dimension is specified by the user

in terms of ay and a., which represent the lateral and vertical plume spreading from the release point

respectively. To reduce the effect of plume spreading attributable to the building wake effect, it was

decided to assume that the release is from a point source. This was implemented in the MACCS2

calculation by setting the dimensions of the structure releasing the radioactive material to be 4 meters

high by 4 meters wide (13 x 13 feet) (i.e., approximately the cask dimension). The vertical and lateral

spreading, oy, and a., is then estimated by dividing the building width and the building height by 4.3

and 2.15, respectively. These values are shown in Table E.2.

Table E.2. Plume Dimensions at the Release Point

Point Source

structure width 4 meters (13.1 feet)

structure height 4 meters (13.1 feet)

CyY 0.93

CFý 1.86

E.9 Plume Heat Content

The plume heat content determines the rise of the plume in the atmosphere. The plume heat content

for the dropped cask accident is expected to be approximately equal to the heat content of the spent fuel.

For ten-year cooled spent fuel, the maximum decay heat load is 264 watts per assembly. The MACCS2

calculations used 18 kilowatts for 68 high-bumup fuel assemblies in a cask. The plume resulting from

this release will not be hot enough to produce significant plume rise.

E.10 Population Distribution

The population distribution for the site was obtained from the SECPOP code (Reference E.9), based on

year 2000 census data.

E-4
E.11 Site Weather

MACCS2 code samples from 1 year of hourly weather data to calculate offsite radiological

consequences. This weather data includes wind speed, wind direction, atmospheric stability, and rainfall.

The weather data file in MACCS2 format was requested and obtained from the site licensee. Because

this file had some missing hourly weather data, it was modified in two ways. First, if data were missing

for a number of consecutive hours, the weather of the previous day was used. Second, if data were

missing for only 1 hour, either the hour before or the average of the hour before and the hour after was

used. In addition, the weather data provided by the licensee did not include the mixing layer height for

each season. The mixing layer height is the top of the well-mixed surface layer of air, and is used in

MACCS2 to limit both buoyant plume rise and vertical dispersion. Therefore, the Sample Problem-A

MACCS2 mixing heights were used in the calculation. The option used in MACCS2 to sample weather

data was theweather category bin sampling. This option is typically used in MACCS2 in order to ensure

adequate sampling of rain cases.

E.12 Results

Six separate MACCS2 calculations were performed. For each of the two source terms, releases at 50

meters (164 feet) with and without filter operation and at 120 meters (393 feet) were evaluated. The

results are presented in Table E.3 and are mean values as calculated by MACCS2, where mean is the

average (expected) consequence over all weather trials (Reference E.2). In all cases, the health impacts

are small.

Table E.3. MACCS2 Consequence Calculation Results

Release Fraction Individual Risk of Individual Risk of

Release Prompt Fatality Cancer Fatality Individual Peak

Noble Particles CRUD Height within 16 km within 16 km Dose at 1.2-1.6

Gases meters (feet) (10 miles) (10 miles) km Sv (rem)

.12 I.2xl03 1.5x1O"3 50(164) 0 3.6x104 1.85 (185)

.12 1.2x103- 1.5x10"3 120 (393) 0 2.1x104 0.14 (14)

.12 1.2x104 1.5x104 50' (164) 0 5.2x10s 0.22 (22)

.12 7x106 1.5x103 50 (164) 0 4.3x106 0.026 (2.6)

.12 7x106 I.5x10-3 120(393) 0 2.6x10.6 0.0032 (0.32)

.12 7xl0v 1.5x]0"4 501 (164) 0 4.3x10-7 0.0027 (0.27)

Results corresponding to a release fraction with filters operable.

E-5
E.13 References

E. 1. HOLTEC International, "HI-STORM Topical Safety Analysis Report," HOLTEC REPORT HI951312, Revision 8, Marlton, NJ.

E.2. U.S. Nuclear Regulatory Commission, "Code Manual for MACCS2," NUREG/CR-6613, Vol. 1,

Washington, DC, May 1998.

E.3. U.S. Nuclear Regulatory Commission, "DOSFAC2 User's Guide," NUREG/CR-6457,

Washington, DC, December 1997.

E.4. Croff, A.G., "ORIGEN2: A Revised and Updated Version of the Oak Ridge Isotope Generation

and Depletion Code," ORNL-5621, Oak Ridge National Laboratory, Oak Ridge, Tennessee,

July 1980.

E.5. "ORIGEN2 Isotope Generation and Depletion Code," CCC-371, Oak Ridge National Laboratory,

Oak Ridge, Tennessee, 1991.

E.6. Hazelton, R.F., "Characteristics of Fuel. Crud and Its Impact on Storage, Handling, and Shipment

of Spent Fuel," PNL-6273, Pacific Northwest Laboratory, Richland, Washington,

September 1987.

E.7. Sandoval, R.P., et al., "Estimate of CRUD Contribution to Shipping Cask Containment

Requirements," SAND88-1358, Sandia National Laboratories, Albuquerque, New Mexico,

January 1991.

E.8. Lukic, J.D., and J.S. Schmidt, "Taming the CRUD Problem: A Utility Perspective,"

Nuclear Technology, 142:283, American Nuclear Society, La Grange Park, Illinois, 2003.

E.9. Humphreys, S.L., J.A. Rollstin, and J.N. Ridgely, "SECPOP90: Sector Population,

Land Fraction, and Economic Estimation Program," NUREG/CR-6525, U.S. Nuclear Regulatory

Commission, Washington, DC, September 1997.

E-6
APPENDIX F

QUESTIONS AND ANSWERS

APPENDIX F

QUESTIONS AND ANSWERS

Question 1: Why did the NRC conduct this study?

Answer 1: The spent fuel pools of commercial nuclear power plants are becoming filled with spent fuel

assemblies. Because spent fuel pools were not typically designed to accommodate all the spent fuel

assemblies generated during a plant's lifetime, many licensees have been removing the older spent fuel

assemblies and storing them onsite in dry cask storage systems. The NRC conducted this study to

develop a methodology for assessing the potential risk to the public from the storage of spent fuel in a

dry cask storage system.

Question 2: How did the NRC assess the potential risk from dry cask storage systems?

Answer 2: The NRC developed and applied a methodology for performing a probabilistic risk

assessment (PRA) of a dry cask storage system at a boiling-water reactor site. The PRA provides

insights into the potential impact of various challenges and accidents on the dry cask storage system, as

well as how the various stages of operation contribute to risk.

Question 3: What was considered in the PRA?

Answer 3: The NRC developed a comprehensive list of challenges and accidents that could impact the

integrity of the dry cask storage system, and evaluated the risk associated with each of them. The NRC

evaluated a number of ways casks might break fiom extremes in temperature and pressure. Conditions

considered included earthquakes, floods, high winds, lightning strikes, accidental aircraft crashes, and

pipeline explosions, as well as dropping the cask during transfer operations. In evaluating the risk, the

NRC considered realistic weather conditions and a representative population distribution in the vicinity

of the site.

Question 4: Did the NRC evaluate terrorism events in this study?

Answer 4: No. The NRC has evaluated the impact of various terrorism events on dry cask storage

systems in separate studies. Collectively, these studies confirm the adequacy of the designs and

demonstrate that they provide an acceptable level of safety. Due to the sensitive nature of the studies,

they are not publically available.

Question 5: What are the results of the study?

Answer 5: The results of PRA studies are normally presented in measures such as the probability of a

prompt fatality and the probability of a latent cancer fatality. The results of this study indicated that no

prompt fatalities are expected. The resulting calculated risk for a latent cancer fatality is extremely small

F-1
(i.e., less than one in a trillion years). Due to the exceedingly low risk numbers calculated, the

conclusion that should be reached is that cask storage systems provide a safe means to store spent nuclear

fuel.

Question 6: Why a re the calculated risk measures so low?

Answer 6: A major contributor is t hat dry cask storage systems are inherently passive, i.e. they do not

rely on active support systems or human actions to achieve their safety function. In addition, the designs

of dry cask storage systems are inherently robust from a structural integrity perspective, i.e., they are

massive structures that are highly damage tolerant. To have a measurable impact on risk, the dry cask

storage system would have to be breached and spent nuclear fuel would need to be disrupted and released

to the environment.

Question 7: If the calculated risks from dry cask storage systems are so low, why shouldn't all the

spent fuel be stored in them, as opposed to a spent fuel pool?

Answer 7: There are two reasons. Fuel that has been recently removed from a reactor must be kept in a

pool for several years to cool (to remove heat generated by radioactive decay). Only after the fuel has

cooled sufficiently may it be safely stored in dry casks. Therefore, spent fuel pools are necessary.

Second, various studies have shown that spent fuel pools and dry cask storage systems are both safe and

secure. Assuming that the aged fuel in a spent fuel pool is air cooled, similar to the dry cask storage

system, and the probability of direct release from the fuel due to associated risk contributors (e.g.,

aircraft, seismic event) is similar, the overall risks should be comparable. Since the NRC has determined

that either storage system ensures that spent fuel is stored safely and securely, the decision to move spent

fuel from the spent fuel pool to dry casks is a decision for licensees.

Question 8: How can the study be used?

Answer 8: The methodology developed in NUREG-1864 can be used as a guide for performing similar

PRAs for other dry cask storage. systems at other nuclear power plant sites. The NRC will consider the

insights from this study in our continuing efforts to improve and risk-inform the licensing and inspection

activities associated with dry cask storage.

F-2
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

(9-2004) (Assigned by NRC, Add Vol., Supp., Rev.,

NRCMD 3.7 and Addendum Numbers, if any.)

BIBLIOGRAPHIC DATA SHEET

(See instructions on the reverse) NUREG-1864

2. TITLE AND SUBTITLE 3. DATE REPORT PUBLISHED

A Pilot Probabilistic Risk Assessment of a Dry Cask Storage System MONTH YEAR

At a Nuclear Power Plant March 2007

4. FIN OR GRANT NUMBER

5. AUTHOR(S) 6. TYPE OF REPORT

Gordon Bjorkman, Tze-Jer Chuang, Robert Einziger, Shah Malik, Asimios Malliakos, Technical

Jocelyn Mitchell, Carlos Navarro, Christopher Ryder, Syed Shaukat, Anthony Ulses, Ghani Zigh

7. PERIOD COVERED (Inclusive Dates)

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (If NRC, provide Division, Office or Region, U.S. Nuclear Regulatory Commission, and mailing address; if contractor,

provide name and mailing address.)

Division of Risk Assessment & Special Projects Spent Fuel Project Office

Office of Nuclear Regulatory Research Office of Nuclear Material Safety and Safeguards

U.S. Nuclear Regulatory Commission U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001 Washington, DC 20555-0001

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (If NRC, type "Same as above", if contractor, provide NRC Division, Office or Region, U.S. Nuclear Regulatory Commission,

and mailing address.)

Same as above

10. SUPPLEMENTARY NOTES

Asimios Malliakos, NRC Project Manager

11. ABSTRACT (200 words or less)

As the basis for this report, the U.S. Nuclear Regulatory Commission (NRC) has developed and applied a methodology for

performing a pilot probabilistic risk assessment (PRA) of a dry cask storage system at a nuclear power plant site (i.e., an independent spent fuel storage installation). This report documents the pilot PRA for a specific dry cask system (Holtec

International HI-STORM 100) at a specific boiling-water reactor (BWR) site. The methodology developed can serve as a guide

for performing similar PRAs in the future. The pilot study can provide guidance for assessing the risk to the public and

identifying the dominant contributors to that risk. The cask system consists of a multipurpose canister (MPC) that confines the

fuel, a transfer overpack that shields workers from radiation while the cask is being prepared for storage, and a storage

overpack that shields people from radiation and mechanically protects the MPC during storage. The study covers various

phases of the dry cask storage process, from loading fuel from the spent fuel pool, preparing the cask for storage and

transferring it outside the reactor building, moving the cask from the reactor building to the storage pad, and storing the cask for

20 years on the storage pad. The study develops and assesses a comprehensive list of initiating events, including dropping the

cask during handling and external events during onsite storage (such as earthquakes, floods, high winds, lightning strikes,

accidental aircraft crashes, and pipeline explosions). Potential cask failures from mechanical and thermal loads are modeled.

The study estimates the annual risk for one cask in terms of the individual probability of a prompt fatality within 1.6 km (1 mile)

and a latent cancer fatality within 16 km (10 miles) of the site.

12. KEY WORDS/DESCRIPTORS (List words or phrases that will assist researchers in locating the report.) 13. AVAILABILITY STATEMENT

dry cask storage system unlimited

dry cask storage operation 14. SECURITY CLASSIFICATION

HI-STORM 100 (This Page)

probabilistic risk assessment (PRA) unclassified

multipurpose canister (MPC) (This Report)

independent spent fuel storage installation (ISFSI) unclassified

dropped fuel assembly

dropped transfer cask 15. NUMBER OF PAGES

storage/transfer overpack

spent fuel assemblies 16. PRICE

NRC FORM 335 (9-2004) PRINTED ON RECYCLED PAPER
Federal Recycling Program