NUREG/CR-6886

PNNL-15313

Spent Fuel Transportation

Package Response to the

Baltimore Tunnel Fire Scenario

Draft Report for Comment

Pacific Northwest National Laboratory

U.S. Nuclear Regulatory Commission

Office of Nuclear Material Safety and Safeguards

Washington, DC 20555-0001
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NUREG/CR-6886

PNNL-15313

Spent Fuel Transportation

Package Response to the

Baltimore Tunnel Fire Scenario

Draft Report for Comment

Manuscript Completed: October 2005

Date Published: November 2005

Prepared by

H.E. Adkins, Jr., J.M. Cuta, B.J. Koeppel

Pacific Northwest National Laboratory

Richland, WA 99352

C. Bajwa and A. Hansen, NRC Project Managers

Prepared for

Spent Fuel Project Office

Office of Nuclear Material Safety and Safeguards

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Job Code J5167
COMMENTS ON DRAFT REPORT

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Note: The NUREG/CR-6886 draft report for comment is identical (except for format changes) to the

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and announced in the Federal Register (page 54780) on Friday September 16, 2005.

NUREG/CR-6886 has been reproduced

from the best available copy.
ABSTRACT . ;, I

On July 18, 2001, a freight train carrying

hazardous (non-nuclear) materials derailed and

caught fire while passing through the Howard

Street railroad tunnel in downtown Baltimore,

Maryland. The United States Nuclear Regulatory Commission (USNRC), one of the agencies

responsible for ensuring the safe transportation

of radioactive materials in the United States,

undertook an investigation of the train derailment and fire to determine the possible regulatory implications of this particular event for the

transportation of spent nuclear fuel by railroad.

The USNRC met with the National Transportation Safety Board (NTSB) to discuss the details

of the accident and the ensuing fire. Following

these discussions, the USNRC assembled a team

of experts from the National Institute of

Standards and Technology (NIST), the Center

for Nuclear Waste Regulatory Analyses

(CNWRA), and Pacific Northwest National

Laboratory (PNNL) to determine the thermal

conditions that existed in the Howard Street

tunnel fire and analyze the effects on various

spent fuel transportation package designs.

The Fire Dynamics Simulator (FDS) code

developed by NIST was used to determine the

thermal environment in the Howard Street

tunnel during the fire. The FDS results were

used as boundary conditions in the COBRA-SFS

and ANSYS® computer codes to evaluate the

thermal performance of different package

designs. The staff concluded that larger transportation packages resembling the HOLTEC

Model No. HI-STAR 100 and TransNuclear

Model No. TN-68 would withstand a fire with

thermal conditions similar to those that existed

in the Baltimore tunnel fire event with only

minor damage to peripheral components. This is

due to their sizable thermal inertia and design

specifications in compliance with currently

1, imposed regulatory requirements.

For the TN-68 and the NAC LWT, the

maximum temperatures predicted in the regions

of the lid, vent and drain ports exceed the seals'

rated service temperatures, making it possible

for a small release to occur, due to CRUD that

might spall off the surfaces of the fuel rods.

However, any release is expected to be very

small due to a number of factors. These include

(1) the tight clearances maintained between the

lid and cask body by the closure bolts, (2) the

low pressure differential between the cask

interior and exterior, (3) the tendency of such

small clearances to plug, and (4) the tendency of

CRUD particles to settle or plate out.

USNRC staff evaluated the radiological consequences of the package responses to the

Baltimore tunnel fire. The results of this evaluation strongly indicate that neither spent

nuclear fuel (SNF) particles nor fission products

would be released from a spent fuel shipping

cask involved in a severe tunnel fire such as the

Baltimore Tunnel Fire. None of the three cask

designs analyzed for the Baltimore Tunnel fire

scenario (TN-68, HI-STAR 100, and NAC

LWT) experienced internal temperatures that

would result in rupture of the fuel cladding.

Therefore, radioactive material (i.e., SNF

particles or fission products) would be retained

within the fuel rods.

There would be no release from the HI-STAR

100, because the inner welded canister remains

leak tight. The potential releases calculated for

the TN-68 rail cask and the NAC LWT truck

cask indicate that any release of CRUD from

either cask would be very small - less than an A2

quantity (see footnote 6, Section 8.)

iii
DISCLAIMER

This report was prepared as an account of work sponsored by an agency of the United States Government.

Neither the United States Government nor any agency thereof, nor Battelle Memorial Institute, nor any of

their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility

for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed,

or represents that its use would not infringe privately owned rights.

iv
CONTENTS

ABSTRACT .......................................................................;.......................................................................... iii

ABBREVIATIONS..xiii

I INTRODUCTION ........................... 1.1

1.1 The Baltimore Tunnel Fire Event ........................... 1.1

2 NIST TUNNEL FIRE MODEL ..................... 2.1

3 CNWRA MATERIALS EXPOSURE ANALYSIS ....................... ................................ 3.1

4 TRANSPORTATION OF SPENT NUCLEAR FUEL ....................................................... 4.1

4.1 Transportation Packages Analyzed ....................................................... 4.1

4.1.1 TransNuclear TN-68 SNF.Transportation Package .................................................. 4.1

4.1.2 HOLTEC HI-STAR 100 SNF Transportation Package ............................... ............. 4.2

4.1.3 NAC LWT SNF Transportation Package ....................................................... 4.3

5 ANALYSIS APPROACH ....................................................... 5.1

5.1 Model of TN-68 Transportation Package ....................................................... 5.1

5.2 Model of HI-STAR 100 Transportation Package ....................................................... 5.3

5.3 Model of NAC LWT Transportation Package ............................. .......................... 5.10

6 ANALYSIS METHOD ....................................................... 6.1

6.1 Fire Transient Assumptions and Boundary Conditions .................................... 6.1

6.1.1 Boundary Temperatures from FDS ....................................................... 6.2

6.1.2 Convection and Radiative Heat Transfer Boundary Conditions ............... ................ 6.3

6.1.3 Extrapolated Boundary Conditions for Long-Term Cool Down ............................... 6.5

6.1.4 Heat Transfer through Liquid Neutron Shield ....................................................... 6.7

6.2 Initial System Component Temperatures ....................................................... 6.8

6.3 Tunnel Fire Evaluations of Rail Packages ....................................................... 6.13

7 ANALYSIS RESULTS ....................................................... 7.1

7.1 TN-68 Fire Transient Results ....................................................... 7.1

7.1.1 TN-68 During the Fire ....................................................... 7.1

v
7.1.2 TN-68 Short-Term Post-Fire Response ........................................ 7.3

7.1.3 TN-68 Long-Term Post-Fire Response ........................................ 7.6

7.2 Holtec HI-STAR 100 Fire Transient Results ......................................... 7.8

7.2.1 HI-STAR 100 During the Fire ........................................ 7.8

7.2.2 HI-STAR 100 Short-Term Post-Fire Response ................ ........................ 7.9

7.2.3 HI-STAR 100 Long-Term Post-Fire Response ........................................ 7.11

7.3 NAC LWT Fire Transient Results ......................................... 7.14

7.3.1 LWT During the Fire ........................................ 7.14

7.3.2 LWT Short-Term Post-Fire Response ........................................ 7.16

7.3.3 LWT Long-Term Post-Fire Response ........................................ 7.18

8 POTENTIAL CONSEQUENCES ........................................ . 8.1

8.1 Results for the HI-STAR 100 Cask .................................... 8.1

8.2 Results for the TN-68 Cask .................................... 8.1

8.3 Results for the NAC LWT Cask .................................... 8.3

8.4 Summary of Potential Releases .................................... 8.5

9 REFERENCES ......................................... 9.1

vi
FIGURES

1.1 Liquid Tripropylene Tank Car ............................................ 1.2

4.1 Cross-section of TN-68 Package (drawing 972-71-3 Rev. 4, "TN-68 Packaging General

Arrangement: Parts List and Details") ............................................ 4.1

4.2 TransNuclear TN-68 Spent Fuel Transportation Package ............................................ 4.2

4.3 HOLTEC HI-STAR 100 Spent Fuel Package ............................................ 4.2

4.4 Spent Fuel Transportation Package on Railcar ............................................ 4.3

4.5 NAC LWT Transport Package (without ISO container) ............................................ 4.3

4.6 NAC LWT Transport Package (with ISO container) ............................................ 4.3

5.1 COBRA-SFS Model of TN-68 Basket and Support Rails ............................................ 5.2

5.2 ANSYS HI-STAR 100 Package Analysis Model Element Plot ............................................ 5.4

5.3 Cross-section of Package, Cradle, and Rail Car Section ............................................ 5.4

5.4 Cross-section of HI-STAR Package and MPC-24 Canister ............................................ 5.5

5.5 Close-up of Package Cross-section ............................................ 5.5

5.6 Close-up of Canister Basket Structure and Fuel Compartment Configuration (without helium

elements)...............................................5.........................................................................................5.6

5.7 Close-up of Canister Basket Structure and Fuel Compartment Configuration (with helium

elements) ............................................ 5.6

5.8 Complete Impact Limiter (Except Skin) ............................................ 5.7

5.9 Impact Limiter Skin and Primary Support Structure .................................... 5.7

5.10 Boundary Sections for Tunnel and Package Model .......................................................... 5.8

5.11 Surfaces Defined for Interaction with "Top" Gas Region .5.8

5.12 Surfaces Defined for Interaction with "Side" Gas Region. 5.9

5.13 Surfaces Defined for Interaction with "Bottom" Gas Region. 5.9

5.14 ANSYS NAC LWT Package Analysis Model Element Plot ........................................................ 5.10

5.15 Cross-section of NAC LWT Package .5.10

5.16 NAC LWT Package Geometry .5.11

5.17 Zones for Convection Heat Transfer Within the ISO Container. 5.12

5.18 Zones for External Radiation Between ISO Container and Tunnel Surfaces. 5.13

6.1 BTF Peak Transient Ambient Air Temperatures (smoothed values, NIST 20-m data) .................. 6.3

6.2 BTF Peak Transient Tunnel Surface Temperatures for Floor, Walls, and Ceiling

(smoothed values, NIST 20-m data) . ............................................................ 6.3

6.3 BTF Peak Transient Horizontal Velocities near Package Surface (smoothed values,

NIST 20-m data) ............................................................. 6.3

6.4 Heat Transfer Coefficients at Package Surface from NIST 20 m Air Temperature and

Velocity Predictions ............... 6.4

6.5 BTF Peak Transient Air Temperatures for Top, Side, and Bottom Regions (NIST and

Extrapolated Data Sets). 6.5

6.6 BTF Peak Transient Surface Temperatures for Floor, Walls, and Ceiling (NIST and

Extrapolated Data Sets). 6.6

6.7 Nusselt Number for Heat Transfer in Liquid Neutron Shield .......................................... 6.8

6.8 Effective Conductivity of Neutron Shield Tank Contents .......................................... 6.8

vii
i m

6.9 Effective Conductivity of Expansion Tank Contents ................... ................................................ 6.8

6.10 HI-STAR 100 Package Hot-Normal Condition Temperature Distribution .................................... 6.9

6.11 LWT Package Hot-Normal Condition Temperature Distribution (2.5 kW Decay Heat) ............. 6.11

6.12 LWT Package Normal Condition Temperature Distribution (2.5 kW Decay Heat) .................... 6.12

7.1 Maximum Temperature Histories for TN-68 Package Components During Fire Transient .......... 7.2

7.2 Temperature Profiles Top-To-Bottom Through TN-68 Package Axis During Fire Transient ....... 7.2

7.3 Maximum Temperature Histories for TN-68 Package Components During First 30 hr

of Transient ..................................................................... 7.3

7.4 Maximum TN-68 Package Surface Temperatures Compared to NIST Boundary Condition

Temperatures ..................................................................... 7.4

7.5 Peak Fuel Cladding Temperature History in TN-68 During First 50 hr of Transient .................... 7.5

7.6 Maximum Global TN-68 Closure/ Port Seal Temperature History During First 30 hr of

Transient ..................................................................... 7.6

7.7 Temperature Profiles Top-To-Bottom Through TN-68 Package Axis: First 30 hr of Transient .... 7.6

7.8 Maximum Temperature Histories for TN-68 Package Components During 300 hr of Transient.. 7.7

7.9 Maximum Temperature Histories for HI-STAR 100 Package Components During Fire

Transient ..................................................................... 7.8

7.10 Maximum Temperature Histories for HI-STAR 100 Package Components During First 30 hr of

Transient .................................................................. 7.9

7.11 Maximum HI-STAR 100 Package Surface Temperatures Compared to NIST Boundary Condition

Temperatures .................................................................. 7.10

7.12 Peak Fuel Cladding Temperature History in HI-STAR 100 During First 50 hr of Transient ...... 7.10

7.13 Maximum Global HI-STAR 100 Closure/Port Seal Temperature History During First

30 h of rTransient .................................................................. 7.11

7.14 Maximum Temperature Histories for HI-STAR 100 Package Components During 300 hrs

of the Transient ................................................................... 7.12

7.15 NAC LWT Package Component Maximum Temperature Histories During Fire Transient ........ 7.15

7.16 Lumped Fuel Assembly Temperature Distribution 7 hr Into Transient ..................................... 7.15

7.17 NAC LWT Package Component Maximum Temperature Histories for First 30 hr of Fire

Transient ................................................................... 7.16

7.18 Maximum ISO Container Surface Temperature History Compared with

NIST Boundary Condition Temperatures ................................................................... 7.16

7.19 Peak and Average Fuel Cladding Temperature Histories for NAC LWT Package During First

30 hr of Fire Transient .................................................................... 7.17

7.20 Maximum Seal Temperature Histories for Drain/Vent Ports and Package Lid During First

30 hr of Fire Transient ................................................................... 7.17

7.21 NAC LWT Package Component Maximum Temperature Histories During 300 hr Transient .... 7.18

viii
TABLES

6.1 TN-68 Hot-Normal Component Temperatures .............................................................. 6.9

6.2 HOLTEC HI-STAR 100 Hot-Normal Component Temperatures ................................. ................ 6.10

6.3 NAC LWT Component Temperatures at Various Decay Heat Loads ...................... ..................... 6.11

6.4 NAC LWT Component Temperatures at 2.5 kW Decay Heat Load and 130'F Ambient ............. 6.12

6.5 NAC LWT Pre-Fire Component Temperatures at 2.5 kW Decay Heat Load and 1000F

Ambient.......................................................................................................................................... .6.13

7.1 TN-68 Peak Component Temperatures During Fire Transient ........................................................ 7.7

7.2 HOLTEC HI-STAR 100 Peak Component Temperatures During Fire Transient ............. ............ 7.14

7.3 NAC LWT Peak Component Temperatures During Fire Transient ............................................... 7.18

8.1 Assumptions Used for Release Estimate for TN-68 Cask ........................................................ 8.3

8.2 Potential Release Estimates for TN-68 Cask ........................................................ 8.3

8.3 Assumptions Used for Release Estimate for NAC LWT Cask ................................... 8.4

8.4 Potential Release Estimates for NAC LWT Cask ........................................................ 8.5

ix
APPENDICES

APPENDIX A - Material Properties for COBRA-SFS Model of TN-68 Package ............... ................... A.]

APPENDIX B - Material Properties for ANSYS Model of HI-STAR 100 Package ................................ B. I

APPENDIX C - Material Properties for ANSYS Model of Legal Weight Truck Package ...................... C. I

APPENDIX D - Blackbody Viewfactors for COBRA-SFS Model of TN-68 Package ............. .............. D.1

APPENDIX E - HOLTEC HI-STAR 100 Component Temperature Distributions .................................. E.I

xi
ABBREVIATIONS

APDL ANSYS` Parametric Design Language

BTF Baltimore Tunnel Fire

BWR Boiling Water Reactor

CFD Computational Fluid Dynamics

CNWRA Center for Nuclear Waste Regulatory Analyses

CRUD Chalk River Unknown Deposit (generic term for various residues deposited on fuel rod

surfaces, originally coined by Atomic Energy of Canada, Ltd. (AECL) to describe deposits

observed on fuel removed from the test reactor at Chalk River.)

FDS Fire Dynamics Simulator (computational fluid dynamics computer code)

FEA Finite Element Analysis

ISO International Organization for Standardization (The International Organization for

Standardization has decreed the use of the initials ISO for reference to the organization,

regardless of the word order of the organization's name in any given language. This

defines a uniform acronym in all languages.)

MPC Multi-Purpose Canister

NIST National Institute of Standards and Technology

NTSB National Transportation Safety Board

OFA Optimized Fuel Assembly

PNNL Pacific Northwest National Laboratory

PWR Pressurized Water Reactor

SFPO USNRC Spent Fuel Project Office

SNF Spent nuclear fuel

USNRC United States Nuclear Regulatory Commission

xiii
1 INTRODUCTION

Current USNRC regulations specify that spent

nuclear fuel shipping packages must be designed

to survive exposure to a fully engulfing fire

accident lasting no less than 30 minutes with an

average flame temperature of no less than 14750F

(802'C)[1]. The package must maintain

containment, shielding and criticality functions

throughout the fire event and post-fire cool down

in order to meet USNRC requirements.

On July 18, 2001, a CSX freight train carrying

hazardous (non-nuclear) materials derailed and

caught fire while passing through the Howard

Street railroad tunnel in downtown Baltimore,

Maryland. (A summary description of the event is

given in Section 1 .1.) The staff of the USNRC

Spent Fuel Project Office (SFPO) undertook an

investigation of the derailment and fire in order to

determine what impact this event might have had

on spent nuclear fuel transported by rail.

The severity of the Baltimore tunnel fire has raised

questions about the performance of spent fuel

transportation packages in such an accident. As

one element of on-going evaluations related to this

accident scenario, calculations were performed for

three spent nuclear fuel transportation packages.

The TransNuclear Model No. TN-68 ("TN-68")

transport package was analyzed using the

COBRA-SFS code [2]. The HOLTEC Model No.

HI-STAR 100 ("HI-STAR 100") and the NAC

International Model No. LWT ("NAC LW"')

transport package were analyzed using the

ANSYS® code [3]. The analyses were performed

in parallel to expedite the work and to provide

independent evaluation of different modeling

approaches for the two relatively similar large

package designs.

Air temperatures and temperatures of the tunnel

wall, floor and ceiling derived from fire analyses

performed by the National Institute of Standards

and Technology (NIST) were used to define the

boundary conditions for the transient calculations.

The purpose of the evaluation was to obtain an

estimate of the temperature response of the various

components of each of these packages during and

after the fire.

This report presents a detailed description of the

analyses, including boundary conditions, modeling

approach, and computational results. Section 2

describes the NIST tunnel fire model used to

develop boundary conditions for the thermal

analyses of the spent fuel transportation packages.

Section 3 briefly describes the material exposure

analysis used to verify the predicted temperatures

obtained in the fire simulations performed by

NIST. Section 4 presents a detailed description of

the spent fuel transportation casks, and the

computational models developed for the analyses

are described in Section 5. Section 6 presents a

detailed description of the analysis method.

Section 7 presents the results of the simulation,

giving a detailed evaluation of predicted response

for each cask package during and after the fire.

Section 8 provides an analysis to determine the

magnitude of any potential release of radioactive

material as a consequence of the effects of the fire

on the casks.

1.1 The Baltimore Tunnel Fire

Event

The Howard Street tunnel in Baltimore is a single

track railroad tunnel of concrete and refractory

brick. Originally constructed in 1859, later

additions extended it to its current length of 1.65

mi (2.7 kin). The tunnel has an average upward

grade of 0.8% from the west portal to the east

portal, and has no active ventilation system. The

tunnel measures approximately 22 ft (6.7 m) high

1.1
by 27 ft (8.2 in) wide in the vicinity of the

accident; however, the dimensions vary along the

length.

The freight train involved in the accident had a

total of 60 cars pulled by 3 locomotives, and was

carrying paper products and pulp board in boxcars

as well as hydrochloric acid, liquid tripropylene',

and other hazardous liquids in tank cars [4,51. As

the train was passing through the tunnel, 1 1 of the

60 rail cars d.erailed. A tank car (see Figure 1. I)

containing approximately 28,600 gallons

(108,263 liters) of liquid tripropylene had a 1.5-

inch (3.81-cm) diameter hole punctured in it by

the car's brake mechanism during the derailment.

Ignition of the liquid tripropylene led to the ensuing fire. The exact duration of the fire is not

known. Based on NTSB interviews of emergency

responders, it was determined that the most severe

portion of the fire lasted approximately 3 hours.

Other, less severe fires burned for periods of time

greater than 3 hours. Approximately 12 hours

after the fire started, firefighters were able to

visually confirm that the tripropylene tank car was

no longer burning.

Figure 1.1. Liquid Tripropylene Tank Car

Tripropylenc carrics an NFPA hazards rating of 3 for

flammability, which is the same as that of gasoline.

1.2
- -

2 NIST TUNNEL FIRE MODEL

Experts at the National Institute of Standards and

Technology (NIST) developed a model [4] of the

Baltimore tunnel fire using the Fire Dynamics

Simulator (FDS) code [6, 7]2 to predict the range

of temperatures present in the tunnel during the

fire event. FDS is a computational fluid dynamics

(CFD) code that models combustion and flow of

hot gas in fire environments. FDS solves the

mass, momentum, and energy equations for a

given computational grid, and is also able to

construct a visual representation of smoke flow for

a given fire.

To validate FDS for tunnel fire applications, NIST

developed fire models in FDS based on the

geometry and test conditions from a series of fire

experiments conducted by the Federal Highway

Administration and Parsons Brinkerhoff, Inc. as

part of the Memorial Tunnel Fire Ventilation Test

Program [8]. NIST modeled both a 6.83x 107

Btu/hr (20 MNV) and a 1.71 x 108 Btu/hr (50 MW)

unventilated fire test from the Memorial Tunnel

Test Program, and achieved results using FDS that

were within l00F (560C) of the recorded

data[6,7].

and gas behavior. The mesh size was expanded at

distances further from the fire source, where less

resolution was needed.

Maximum temperatures calculated in the FDS

model were on the order of 18001F (9820C) in the

flaming regions of the fire. The model results

showed that the hot gas layer above the railcars

within three rail car lengths of the fire was an

average of 900'F (4820C). Temperatures on the

tunnel wall surface were calculated to be in excess

of 15007F (816'C) where the fire directly

impinged on the ceiling of the tunnel. The average

tunnel ceiling temperature, within a distance of

three rail cars from the fire, was 750'F (3990C).

Additional details associated with the analyses

performed by NIST can be found in the report on

the FDS analysis of the Howard Street tunnel fire

[4].

The full-length 3-dimensional representation of

the tunnel fire scenario developed by NIST

included railcars positioned as they were found

following the derailment and fire. The source of

the fire was specified in the simulation as a pool of

burning liquid tripropylene positioned below the

location of the hole punctured in the tripropylene

tank car. The computational grid for the tunnel

fire model was relatively fine in the immediate

vicinity of the fire, in order to properly capture fire

2 Formal publication of the FDS code documentation

began in 2001 with Version 2. Continuing validation

and development of the code led to Version 3 in 2002.

Version 3 was used in the FDS analyses discussed in

this report.

2.1
3 CNWRA MATERIALS EXPOSURE ANALYSIS

Staff from the Center for Nuclear Waste

Regulatory Analysis (CNWRA), along with staff

from NRC and NIST, examined railcars and tank

cars removed from the Howard Street tunnel

approximately one year after the fire. The

examination of physical evidence provided the

staff with further insight into the fire environment

that existed in the tunnel during the accident. Staff

from CNWRA also collected material samples

from the boxcars and tank cars inspected.

estimate the fire exposure time and temperature

for the samples tested. The material time and

temperature exposures determined by the

CNWRA analyses were consistent with the

conditions predicted by the NIST FDS model of

the Howard Street tunnel fire [4].

Additional details associated with the analyses

performed by CNWRA can be found in the report

on the analysis of the rail car components [5].

By performing metallurgical analyses on the

material samples collected, including sections of

the boxcars exposed to the most severe portion of

the fire and an air brake valve from the

tripropylene tanker car, CNWRA was able to

3.1
4 TRANSPORTATION OF SPENT NUCLEAR FUEL

NRC regulations require that spent fuel

transportation packages be evaluated for a series

of hypothetical accident conditions that include a

30-ft (9 in) drop test, a 40-inch (I m) pin puncture

drop test, and a fully engulfing fire with an

average flame temperature of 14750F (8020C) for

a period of 30 minutes. These tests are followed

by the immersion of an undamaged package under

50 ft (15 m) of water [l]. The certification process

must include either an open pool fire test or an

analysis of the package for a fire exposure meeting

the aforementioned criteria. Packages must

maintain shielding and criticality control functions

throughout the hypothetical accident conditions.

4.1 Transportation Packages

Analyzed

This investigation evaluates how a fire similar to

the Howard Street tunnel fire might affect three

NRC-approved spent fuel transportation package

designs. These include the HOLTEC HI-STAR

1 00 and TransNuclear TN-68 rail transportation

packages, and NAC LWT transportation package.

The LWVT was selected because it represents a

typical truck (over-the-road) package that can also'

be transported by rail. The design of each of these

packages is briefly described below.

4.1.1 TransNuclear TN-68 SNF

Transportation Package

The TN-68 spent fuel shipping package transpoirts'

BWR spent fuel assemblies. The basic design is

similar to that of the HOLTEC HI-STAR 100,

except that the TN-68 package does not include an

inner sealed canister. The containment boundary''

is provided by the package shell and lid seals.

The TN-68 package holds up to 68 BWR

assemblies, with a maximum total decay heat load,

of 72,334 Btu/hr (21.2 kW). The fuel assemblies

are contained within a basket structure consisting

of 68 stainless steel tubes that have aluminum and

borated aluminum (or boron carbide/aluminum

composite) neutron poison plates sandwiched

between the steel tubes. The general layout of the

TN-68 package is illustrated in Figures 4.1 and

4.2. Detailed information on the design can be

found in the appropriate sections of the TN-68

Safety Analysis Report (SAR) [9].

SECTiON A-A

FIM DWG~2-71-2 MU 

Figure 4.1.. Cross-section of TN-68 Package

(drawing 972-71-3 Rev. 4, "TN-68

Packaging General Arrangement:

Parts List and Details")

The basket structure is supported by aluminum

alloy support rails bolted to the inner carbon steel

package shell, which also serves as the inner

gamma shield. This inner steel shell is shrinkfitted within an outer carbon steel shell that serves

as the outer gamma shield. The gamma shielding

is surrounded by the neutron shielding, which

4.1:
consists of a ring of aluminum boxes filled with

borated polyester resin. The outer shell of the

package is carbon steel.

Btu/hr (20.0 kW). The MPC is placed in the

transportation package (or overpack) for shipment

after it has been loaded with spent nuclear fuel and

welded shut. A diagram of the HI-STAR 100

package system (MPC and overpack) is provided

in Figure 4.3. The package inner shell is

stainless steel, and six layers of carbon steel

plates comprise the gamma shield. The next

layer is a polymeric neutron shield,

strengthened by a network of carbon steel

stiffening fins. The outer shell of the package

is carbon steel, with a painted outer surface.

Figure 4.2. TransNuclear TN-68 Spent Fuel

Transportation Package

The package bottom is carbon steel with an inner

steel shield plate. The package lid is also carbon

steel with a steel inner top shield plate. During

transport, the ends of the package are capped with

impact limiters made of redwood and balsa and

covered in stainless steel plate. The TN-68 weighs

approximately 260,400 lb (118,115 kg) when

loaded for transport.

4.1.2 HOLTEC HI-STAR 100 SNF

Transportation Package

This design provides an additional containment

boundary in the form of a welded multi-purpose

canister (MPG) enclosing the spent fuel. (The

outermost containment boundary is provided by

the package shell and lid seals.) HOLTEC has a

variety of MPC configurations designed to

accommodate three different spent fuel loading

configurations: up to 24 PWR assemblies, up to 32

PWR assemblies, or up to 68 BWR assemblies.

The MPC-24 configuration was selected for this

evaluation. This design has an integral fuel basket

that accommodates 24 PWR spent fuel assemblies

with a maximum total decay heat load of 68,240

Figure 4.3. HOLTEC HI-STAR 100 Spent

Fuel Package

Aluminum honeycomb impact limiters with

stainless steel skin are installed on the ends of the

package prior to shipping. Impact limiters protect

the closure lid, MPC, fuel basket, and contents

from damage in the event of a package drop

accident. The impact limiters also provide thermal

4.2
insulation to the lid and port cover components in

the event of a fire exposure. Figure 4.4 shows an

illustration of this package secured to a railcar,

with impact limiters installed. This package

weighs approximately 277,300 lb (125,781 kg)

when loaded for transport. Additional

configuration details are provided in the HI-STAR

100 Package System SAR [10].

for the different fuels. For the purposes of this

analysis, the package was assumed to contain a

single PWR spent nuclear fuel assembly, with a

maximum decay heat load of 8,530 Btulhr (2.5

kW). This is the highest heat load the package is

rated for with any spent fuel it is designed to carry,

and thus provides a conservative thermal load for

the fire accident scenario. The loaded package

weighs approximately 52,000 lb (23,586 kg). The

containment boundary provided by the stainless

steel package consists of a bottom plate, outer

shell, upper ring forging, and closure lid.

, ._. v,. _Ir 7 -;I e-t. T - T ,T I

A- ... 1; M 1 I-" ' -n 's ~ In .-."

Figure 4.4. Spent Fuel Transportation Package

on Railcar3 Figure 4.5. NAC LWVT Transport Package

(without ISO container)

4.1.3 NAC LWT SNF Transportation

Package

The NAC LWVT is a small transportation package

certified for transport on a standard tractor trailer

truck, but can also be transported by rail. The

NAC LWT is typically shipped within an

International Organization for Standardization

(ISO) shipping container. Figure 4.5 shows a

picture of a NAC LWT package on a flat-bed

trailer with a personnel barrier installed, but

without an ISO container. Figure 4.6 shows an

exterior view of the package within an ISO

container on a flat-bed trailer.

This package is designed to transport a variety of

commercial and test reactor fuel types with widely

varying maximum decay heat load specifications

Figure 4.6. NAC LWNT Transport Package

(with ISO container)

The package has an additional outer stainless steel

shell to protect the containment shell, and also to

3 Image courtesy of HOLTEC International.

4.3
i All

enclose the lead gamma shield. Neutron shielding

is provided by a stainless steel neutron shield tank

containing a water/ethylene glycol mixture. An

additional annular expansion tank for the mixture

is provided, external to the shield tank. This

component is strengthened internally by a network

of stainless steel stiffeners. Aluminum honeycomb

impact limiters covered with an aluminum skin are

attached to each end of the package. Additional

configuration details are provided in the SAR for

this transport package [I I].

4.4
5 ANALYSIS APPROACH

The analytical approach used to evaluate the

response of the selected transportation packages

was to construct highly detailed 3-D models

capable of accounting for all of the significant heat

transfer paths, and exposing them to boundary

conditions from the postulated Howard Street

tunnel thermal environment. All three

transportation package models were constructed in

parallel to expedite the evaluation. Two different

computer analysis codes were used for the large

multi-assembly packages, to provide independent

verification of the analytical results. The TN-68

package was modeled using the COBRA-SFS

finite-difference thermal-hydraulic analysis code

[2], while the HI-STAR 100 and NAC LWT

packages were modeled using the ANSYS [3]

general FEA code. Three-dimensional models of

each of the packages were developed for these

analyses.

These three-dimensional models were subjected to

boundary conditions obtained from the results of

the FDS simulation using the NIST model of the

Howard Street tunnel. (See Section 6 for a

detailed discussion of the analysis method.) The

boundary conditions were developed from

temperature and flow predictions for the

postulated fire and post-fire scenario, and applied

to the TN-68, HI-STAR 100, and LWT models. IPeak tunnel surface temperatures, peak gas

temperatures, and associated gas velocities

predicted in the FDS analysis were selected at 66;:

fi (20 m) down-stream from the fire source. This. r

location corresponds to the shortest possibledistance between the fire center and a SNF:

package being transported by rail. The distance

was determined based on Department of

Transportation regulations that require railcars :

carrying radioactive materials to be separated from

other cars carrying hazardous materials or

flammable liquids by at least one innocuous railcar

(referred to as a buffer car)'[12].

The package was assumed to be horizontal, with

one end of the package facing the fire source.

This orientation results in maximum possible

exposure to the fire-driven flow of hot gas along

the length of the package, and is the most adverse

position for free convection cooling of the package

during the post-fire cool down.

The FDS analysis predicted a 7-hour fire, and the

calculation was extended out to a 23-hour post-fire

cool-down, for a total simulation time of 30 hours.

To determine the packages' complete transient

temperature responses, and to explore the effects

of prolonged exposure to post-fire conditions in

the tunnel, the COBRA-SFS and ANSYS analyses

extended the post-fire duration to 300 hours.

Tunnel wall and air temperatures predicted in the

FDS analysis at 30 hours were extrapolated from

30 hours to 300 hours using a power function, to

realistically model cool-down of the tunnel

environment.

5.1 Model of TN-68 Transportation

Package

The TN-68 package was analyzed with COBRASFS, a code developed by PNNL for thermalhydraulic analyses of multi-assembly spent fuel

storage and transportation systems. The code uses

a lumped-parameter finite-difference approach for

predicting flow and temperature distributions in

spent fuel storage systems and fuel assemblies

under forced and natural circulation flow

conditions. It is applicable to both steady-state.

and transient conditions in single-phase gas-cooled

spent fuel packages with radiation, convection,

and conduction heat transfer. .The code has been

validated in blind calculations using test data from

5.1
spent fuel packages loaded with actual spent fuel

assemblies as well as electrically heated singleassembly tests [13,14,15].

The TN-68 package was modeled in COBRA-SFS

as a one-half section of symmetry. Figure 5.1

shows a diagram of the center cross-section of the

basket and support rails as represented in the

COBRA-SFS model. The fuel assemblies within

the basket are each modeled as detailed rod and

subchannel arrays, and the tubes containing the

fuel assemblies are represented using solid

conduction nodes.

ta?, hl-eoon of spamtty

(Note: not to GcWo)

Figure 5.1. COBRA-SFS Model of TN-68

Basket and Support Rails

The aluminum and borated aluminum neutron

poison plates sandwiched between the tubes are

represented as an interconnected network of solid

conduction nodes. The gamma shielding, neutron

shielding, and outer steel shell are represented

with concentric rings of interconnected solid

conduction nodes with appropriate material

properties. (For clarity, these nodes are not

included in the diagram shown in Figure 5. 1.) The

half-section of the TN-68 package is represented

with about 69,000 fluid nodes, 53,000 fuel nodes,

and over 16,000 solid conduction nodes.

The solid conduction nodes extend over 32 axial

divisions comprising the axial length of the

package. In cross-section, the stainless steel tubes

containing the fuel assemblies are represented

using two solid conduction nodes on each face of

the enclosure, for a total of eight nodes per tube.

The aluminum and borated aluminum neutron

poison plates sandwiched between the tubes are

represented as an interconnected network of solid

conduction nodes that are in intimate physical

contact with the stainless steel tubes and with each

other. A total of 272 nodes are used to represent

the 34 steel tubes in the half-section of symmetry.

The borated aluminum neutron poison plates

making up the rest of the basket are modeled with

a total of 83 solid conduction nodes. The

aluminum alloy basket rails are represented with a

total of 36 solid conduction nodes, and provide

appropriate thermal connections between the

basket and the steel inner gamma shield.

The gamma shielding, neutron shielding, and outer

steel shell are represented in the COBRA-SFS

model as concentric rings of interconnected solid

conduction nodes. The spent fuel arrays within

the basket are assumed to be 7 x 7 BWR

assemblies (the design basis fuel loading for the

TN-68, as specified in the SAR [9]). Each

assembly is modeled in detail, with 49 rods and 64

subchannels.

In cross-section, the gamma shielding is

represented with two rings of 16 nodes each,

representing the inner and outer steel shells of this

component. The neutron shield in cross-section is

5.2
represented with three rings of nodes (for a total of

48 nodes), with properties and connections defined

to represent the material properties and thermal .

interactions of the ring of aluminum boxes filled

with borated polyester. In cross-section, the outer

steel shell of the package is represented with a ring

of 16 nodes, with appropriate thermal connections

to the neutron shielding on one side and ambient

air on the other.

The COBRA-SFS model was verified by running

the steady-state case for design basis normal hot

transport conditions. The predicted peak clad

temperature for these conditions was compared

with the peak temperature reported in the SAR.

The code predicts a peak clad temperature of

485'F (2520C); the SAR gives a value of 490'F

(2540C) for these conditions (see Chapter 3, Table

3-1 in the TN-68 SAR [9]).

The steady-state solution obtained for normal hot

transport conditions was used to define the pre-fire

condition for the package in the transient

calculations simulating the Baltimore tunnel fire.

This provides a conservative estimate of the initial

temperatures throughout the package, since the

boundary conditions for normal hot transport are

specified as I 000F (380C) ambient temperature in

still air with insolation.

The external air temperatures predicted for the fire

in the NIST simulation are sufficiently high to boil,

off the borated polyester neutron shield and

completely char the wooden impact limiters. In

both cases, the normal material would be replaced

with material that would tend to insulate the.

package from the fire (i.e., air in place of the..

borated polyester, charred wood in place of the

wooden impact limiter material.)

To maximize the heat load to the package from the

fire, it was assumed for the purposes of the

calculation that these materials would persist intact

throughout the fire (rather than gradually

degrading or burning off.) Then at the end of the

fire (6.75 hours into the transient), these materials

would be instantly transformed to a degraded

condition. For the nodes modeling the neutron

shield, this was simulated in the calculation by

changing the material properties to hot air at the

end of the fire. The material properties specified

for the nodes modeling the wooden impact limiters

were changed from redwood to charcoal at the end

of the fire.

The material properties from the package vendor's

Safety Analysis Report (SAR) were verified and

then used in the analyses [9]. The material

properties used in this evaluation are given in

Appendix A.

5.2 Model of HI-STAR 100

Transportation Package

The ANSYS model of the HI-STAR 100 package

consists of a detailed three-dimensional

representation of a half-section of symmetry for

the package, its cradle support4, and the rail car

decking directly below the cradle. (The remainder

of the rail car was omitted from the model because

it would partially shield the package from thermal

radiation from the hot tunnel surfaces and block

convection heat transfer to the package due to the

flow of hot 'gas generated by the fire.) This halfsection model of the package was placed within a

complete cross-section of the surrounding tunnel.

The model developed for the HI-STAR 100

package utilized 120,412 SOLID70 and 1,542

SHELL57 thermal elements for conduction. It

used two groups of 13,573 SURF]52 surface

effect elements for handling convection states in

4Dircnsions and materials for the rail car decking and

the cradle wcre based on current (as of June 2005)

specifications from the package vendor. Cradle design

determines the height of the cask within the tunnel, the

5.3
.-a

the pre-fire steady state and the fire accident

transient. For radiation interaction, 288 highly

structured AUX-12 generated MATRIX50

superelernents were constructed using SHELL57

elements. Solar insolation (from I OCFR71 [I]) for

the pre-fire condition was assigned via heat

generation to the first group of 13,573 SURF 152

surface effect elements. A portion of the model is

shown in Figure 5.2. (In this figure most of the

tunnel has been omitted for clarity.)

The material properties from the package vendor's

Safety Analysis Report were verified and used in

the analysis [ I0]. The model explicitly represents

the geometry of the package, including the internal

geometry of the fuel basket, all gaps associated

with the basket construction, as well as the integral

neutron absorber plates. Figures 5.3 through 5.6

show cross-sections of the HI-STAR model

highlighting key features that were included.

section and rail car section were modeled as

hollow enclosures. All internal radiation and

convection influences associated with these two

enclosures were accounted for using AUX-12

generated MATRIX50 superelements (constructed

with SHIELL57 elements) and SURF 152's with the

extra node option, respectively.

AN

Figure 5.3. Cross-section of Package, Cradle,

and Rail Car Section

Conduction within the cradle and trailer material

sections was also accounted for using additional

SHIELL57 elements with thickness option applied.

Natural convection correlations and specially

constructed automated subroutines written in

ANSYS Parametric Design Language (APDL)

were used to continuously evaluate and update the

convective coefficients of heat conductance.

Fourteen separate passive computation nodes were

assigned as "extra nodes" for the SURF 152

surface effect elements used in specifying the

convection interaction within the cradle and rail

car section (seven for each not shown in Figure

5.3). Section 6 presents a discussion of the natural

convection correlations used in this analysis.

Figure 5.4 shows the cross-section of the package

and canister, with the overall basket structure. As

in Figure 5.3, all helium conduction volumes have

been removed for visualization purposes. This

Figure 5.2. ANSYS HI-STAR 100 Package

Analysis Model Element Plot

Figure 5.3 shows the cross-section of the package.

canister, cradle, and transport car section. In this

figure, all helium conduction volumes have been

removed for visualization purposes. The cradle

'coemctrv of direct conduction paths from the cask, and

can affect thermal shielding of the cask during the fire.

5.4
figure displays the general fuel compartment

layout within the MPC-24 basket structure, and

includes the inner shell (light green), gamma

shield (dark blue), and the neutron shield (purple)

components of the HI-STAR 100.

I

.I

A

I

I

.. '. I

shield. The material definition was specified such

that the thermal conductivity could be readily

degraded to represent the effect of single-sided

fillet welds (i.e., not full penetration) that are used

in this connection. Since the thermal conductivity

through single-sided fillet welds is difficult to

determine objectively, the conductivity of this

material in the pre-fire steady state and post-fire

transient was conservatively reduced to half of that

of the solid base material. The effect of this

assumption is to conservatively minimize the

rejection of internal heat. During the fire,

, however, the thermal conductivity of these

elements was assumed to be the same as the solid

base material, mimicking that of a full penetration

weld. This approach was used to conservatively

maximize the heat input into the package during

"t r4-,: b  I . M I'  " 14.. .. . the fire.

Figure 5.4. Cross-section of HI-STAR Package

and MIPC-24 Canister

U 4.- ILl

F..... L...I ___ *1

Special material definitions were also created for

the elements making up the multi-layer steel

gamma shield (dark blue in Figures 5.5 and 5.6.)

To account for the probability of gaps between the

five steel sheets due to standard manufacturing

practices, a gap of 0.01 inch was assumed to exist

between each layer. Effective material thermal

conductivities were then calculated for the gaps,

accounting for conduction through gas in the gap

and radiation across the gap. These material

property definitions were used in the pre-fire

steady state and the post-fire transient to

conservatively minimize the rejection of internal

heat. However, the conduction properties of these

elements were reassigned to that of solid material

(i.e., equivalent to assuming no gaps between

these layers during the fire), to conservatively

maximize the heat input into the package.

Figure 5.6 presents a detailed view of the basket

cross-section showing a typical basket fuel com-.

partment (purple) containing a homogenized fuel

assembly (light blue) surrounded by Boral sheets

and their associated sheathing. Also shown in this

figure are the MPC canister shell (blue-green) and

Figure 5.5. Close-up of Package Cross-section!

The model cross-section in Figure 5.5 shows that

the fillet welds joining the sections of the package

outer skin and the expansion foam in the neutron

shield area have been modeled explicitly. Special

element material definitions were created for th6eelements providing the connection between the

fins enclosing the neutron shield and the ganima

5.5
the package containment/inner shell (lime). The

same features are shown in Figure 5.7 with the

elements for the helium regions included.

-N -----

I

l

l

-

I

I i; 

II .1 i - : i  .

I I .

i T

1 ; --- . - .'

i . .1 . .

L

I,I -

Figure 5.6. Close-up of Canister Basket

Structure and Fuel Compartment

Configuration (without helium

elements)

. i i ' I I - ! . ..: . i . i : I ..

7 -I.,

:_1

MWANOW . .. 'i :.4

 il

:!I

.4

....1-

;

,,

.I

.

- I"7171,14 .1 !I 0;

I

I

I

P--.

similar to that presented by Bahney and Lotz [16],

modified to include a helium gap between the

homogenized fuel region and the fuel

compartment, and the effect of cover gas

pressurization.

Axial conductivity for the homogenized fuel

region was modeled with the cladding as the only

conduction medium, using a cross-sectional area

weighting scheme. The remaining portion of the

homogenized region was considered to be helium.

Density and heat capacity were based on

volumetric averages of the cover gas, cladding,

upper and lower end fittings, and uranium oxide

fuel. A normalized peaking factor of 1. I (from the

design basis axial power distribution in the SAR

[10]) was used to establish the volumetric heat

generation of 2,843 Btu/hr (0.833 kW) over each

assembly along the active fuel length.

Orthotropic effective conductivity properties were

developed for the Boral to include the radiation

and conduction heat transfer components through

an assumed helium gap of 0.0035 inch between

the Boral sheet and its stainless sheathing, and

between the Boral and stainless basket structure.

Modeling of the radiation interaction within the

basket, canister, and package was accomplished by

unselecting all helium regions and coating each

respective interacting set of surfaces forming

enclosures with SHELL57 elements with specified

emissive material properties. The SHELL57

elements were then used to produce highly structured AUX-12 generated MATRIX50 superelements, each defined by a enclosure. A total of

269 MATRIX50 superelements were defined to

capture the radiation interaction within the

package and canister.

Helium conduction and convection within the gas

regions inside the canister were accounted for by

computing effective conduction properties,

assuming limited internal convection and

_ mi3~ 

I- : : .. . .2

Figure 5.7. Close-up of Canister Basket

Structure and Fuel Compartment

Configuration (with helium

elements)

Westinghouse I 7x 17 OFA fuel was selected for

this evaluation. The effective fuel conductivity in

the radial direction was determined using the

approach documented in the HOLTEC SAR [10].

This approach uses a homogenization scheme

5.6
pressurization. The effective conductivity for

these regions was established to be roughly 2.25

times the thermal conductivity of helium.

Standard helium properties were used for the

helium gas between the canister and the package.

Figure 5.8 shows an element plot of the top impact

limiter honeycomb core and steel substructure of

the HI-STA R 100 package, including the HoltiteA neutron shield material sections. (The impact

limiter skin is omitted for visual clarity.) The

bottom impact limiter is similar to the top impact

limiter, except for the bolting configuration and

the extended steel ring covering the top forging,

lid, and buttress plate. Figure 5.9 shows the top

impact limiter skin and support structure, without

the honeycomb core and the neutron shield

materials included.

The impact limiters are assembled with five different types of honeycomb sections.. Gaps between

the honeycomb sections, the steel substructure,

and skin were conservatively ignored to maximize

heat input during the fire. Thermal properties for

the honeycomb sections were based on volumetric

averages of each section using properties

published by the honeycomb manufacturer [17].

Radiation interaction between the package ends"

and impact limiters was modeled by coating each

respective interacting set of surfaces with

SHELL57 elements with specified emissive

material properties. The SHELL57 elements were

then used to produce highly structured AUX-12

generated MATRIX50 superelements. A total of

16 MATRIX50 superelements were defined to

capture the radiation interaction between the

package and impact limiter surfaces.

Conduction and natural convection heat transfer

between the package and impact limiter surfaces

was handled using SURF 152 surface effect

elements. Correlations and specially constructed

automated subroutines written in APDL were used

to continuously evaluate and update the assigned

convection coefficients of heat conductance.

'I '-x'%-.

Figure 5.8. Complete Impact Limiter (Except

Skin)

.- '-

! .1 l. .. T. I i,.. *b1.i ... . ...... a: -t - . 3 :L 1.

Figure 5.9. Impact Limiter Skin and Primary

Support Structure

Sixteen separate passive computation nodes were

assigned as "extra nodes" for the SURF 152

surface effect elements used in specifying the

convection interaction within the cradle and rail

car section (eight for each package end.) (The

natural convection correlations used are presented

in Section 6.) Convection coefficients of heat conductance were conservatively boosted by a

multiplicative factor of 100 between limiter and

package during the fire to mimic enhanced heat

conduction at this interface due to rapid thermal

5.7
expansion. These values were returned to normal

after the end of the fire.

The tunnel structure for this evaluation was

represented by an enclosure approximately 22 ft

(6.7 m) high by 27 ft (8.2 m) wide and 42 ft

(12.8 m) long. The enclosure was capped at both

ends and assigned the same boundary condition on

the end caps as on the walls and ceiling. As

specified by dimensions of the rail car decking and

cradle, the transport system was located such that

the center axis was 8.2 ft (2.5 m) above the tunnel

floor, leaving 12.2 inches (31 cm) underneath the

lowest part of the rail car decking.

The tunnel enclosure was divided into three

sections; top, side, and bottom. The floor of the

enclosure was considered the bottom. The top was

conservatively considered to be all surfaces

(including the end caps) in the range from 15.8 to

22 ft (4.8 to 6.7 in). All surfaces from the floor to

15.8 ft (4.8 m) were considered to be the side

region. These divisions are shown in Figure 5. 1 0.

and floor in the FDS simulation from NIST; see

Section 6.)

To determine the convection heat transfer to the

package during the fire scenario, the exposed

surfaces of the package were also divided into

three section. The top section was defined as all

surfaces above 9.4 ft (2.9 m). The bottom section

was defined as the bottom of the rail car segment

of the model. The side surfaces of the package

were conservatively defined to be all remaining

outer surfaces of the package. The surface

elements of these sections are pictured in Figures

5.1 1 through 5.13.

-4 1;:i:p',

01I

AN

T--

ht (T.,,V.,) Figure 5.11. Surfaces Defined for Interaction

with "Top" Gas Region

The bottom surface of the rail car section was the

only surface influenced by convection heat

transfer in the bottom gas region. In actuality,

none of the "top" surfaces would be directly

exposed to the highest temperature gas region at

the top of the tunnel, because the package is not

positioned that high in the tunnel. This

assumption therefore represents an additional

conservatism in the analysis.

Forced and natural convection correlations and

specially constructed automated subroutines

written in APDL were used to continuously

evaluate and update the assigned convective

Figure 5.10. Boundary Sections for Tunnel and

Package Model

The selected boundary temperatures for each

section were the maximum in that region (top,

side, and bottom, corresponding to ceiling, wall,

5.8
coefficients of heat conductance for the surface'of

the package during pre-fire, fire, and post-fire

phases based on gas velocity.

I.. ... . . .- . - - . I . . I

Figure 5.12. Surfaces Defined for Interaction

with "Side" Gas Region

. .

Figure 5.13. Surfaces Defined for Interaction

iw'ith "Bottom" Gas Region

Radiation interaction between the transport system !

and its partial conveyance and'the tunnel was'

established by coating all respective interacting'

surfaces with SHELL57 elements'with specified

emissive material Properties. The SHELL57 ele-2

ments were then used to produce-a highly structured AUX- 12 generated MATRIX50

superelement. '

The Baltimore tunnel fire evaluation of the HlSTAR 100 was conducted in three phases. These

were the pre-fire, fire, and post-fire phases. For

the pre-fire phase, the hot-normal conditions of

transport were evaluated assuming solar insolation

and a 1007F (38 0C) ambient temperature, in

accordance with IOCFR71.71 [1]. This

conservatively established initial component

temperatures. During this phase, the fillet welds

joining the fins to the gamma shield were specified

with their realistically reduced conduction and

conservative gas gaps were assumed between the

layers of the gamma shield.

For the fire phase of the evaluation (O < t < 7 hr),

solar insolation was shut off, the tunnel surfaces

were introduced, and the transport package and

tunnel surfaces were assigned an emissivity of 0.9

to represent surfaces affected by sooting. Air gaps

originally assumed to be present between the

gamma shield plates were removed, and good

conduction contact was assumed to exist where the

heat fins attach to the gamma shield. Convection

coefficients of heat conductance-were

conservatively multiplied by a factor of 100

between the impact limiters and package body to

mimic enhanced heat conduction due to rapid

thermal expansion. In addition to these

conservative 'measures, to maximize heat input

during the fire, all aluminum honeycomb and

neutron shield resin materials were assumed to

remain intact during the full duration of the fire.

For the post-fire phase of the evaluation (t > 7 hr),

aluminum honeycomb sections exceeding an

average temperature of 1220'F (660'C) and all

neutron shield material sections were degraded to

thermal properties identical to that of air.

Conservatively, the energy that would be absorbed

due to phase change in this material was not.

subtracted from the heat input to the package. In

addition, all gamma shield gaps and reduced fin

fillet weld conduction properties were reintroduced. Finally, convection coefficients of heat

5.9
conductance between the impact limiters and

package body were returned to normal for the

remainder of the simulation.

The material properties from the package vendor's

SAR [10]were verified and then used in the

analyses. The material properties used in this

evaluation arc listed in Appendix B.

5.3 Model of NAC LWT

Transportation Package

The model for the NAC LWT package constructed

in ANSYS is similar in structure to the Ill-STAR

100 model described in Section 5.2. A detailed

three-dimensional model of a half-section of

symmetry was developed for the package and ISO

container, within the same tunnel geometry. A

diagramn of the package and shipping container

model and partial tunnel is shown in Figure 5.14.

MATRIX50 elements were used to model

radiation heat exchange between package surfaces.

The surface effect elements were also used to

generate solar insolation loads for calculation of

the initial temperature distribution for the package.

The model geometry was developed from the

vendor's engineering drawings from the package

SAR [I I]. The model cross-section is shown in

Figure 5.15. The package contains a cylindrical

solid aluminum basket that holds a single fuel

assembly. The helium gaps between the fuel and

the basket, and between the basket and package

shell, were explicitly modeled with solid elements.

t I

11

.6

II

ANM

AN

Figure 5.14. ANSYS NAC LWT Package

Analysis Model Element Plot

Figure 5.15. Cross-section of NAC LWT

Package

The package body is constructed of several

stainless steel shells to provide structural support

and gamma shielding. The innermost shell is

surrounded by a layer of lead that acts as a gamma

shield. The outermost stainless steel shell is

surrounded by an annular tank containing a

solution of ethylene glycol and water which acts as

a neutron shield. The tank is contained by an

outer stainless steel skin and an annular overflow

tank that extends approximately one-third of the

axial length of the package body. All of these

components were modeled usinig brick elements.

The model used 40,333 SOLID70 8-node brick

elements and 3,409 S EIElL57 4-node quadrilateral

thermal elements to represent the structural

components. A total of 6,931 SURF152 elements

were used to incorporate radiation and convection

heat transfer to the ISO container and tunnel

environment for the various surfaces, and 12

5.10
The tank is constructed with eight stainless steel

support ribs (in the half section) connecting the

skin to the outer shell. These structures were

modeled with shell elements. The package bottom

is constructed with a stainless steel base, a layer of

lead shielding, and a steel cover. The upper end of

the package is sealed with a stainless steel lid (see

Figure 5.16). Impact limiters attached to the ends

of the package consist of an internal aluminum

honeycomb structure covered by an aluminum

skin. The expansion tank to handle overflow of the

liquid neutron shield consists of an outer stainless

steel skin.

:7 . idFigure 5.16. NAC LWVT Package Geometry

The entire package is contained in an ISO

container constructed of stainless steel plate. This

is based on the assumption that an ISO container

would be required if the NAC LWT were shipped

by rail. (The consequences of this hypothetical

accident scenario on a package shipped without an

ISO container are discussed with the results of this

analysis in Section 7.3.) The package and

container model is oriented horizontally in the

tunnel with the center of the ISO container is 97.7

inches (248 cm) above the tunnel floor.

Heat exchange via conduction, convection, and

radiation was carefully modeled between all of the

components to provide a sound estimate of

package temperatures during the transient fire

event. Conduction is handled inherently by the

elements modeling each component, but

convective and radiation mechanisms must be

carefully implemented.

Westinghouse 17x17 OFA fuel was used in this

evaluation. The fuel assembly was modeled with

an effective conductivity determined using a

homogenization scheme similar to that presented

by Bahney and Lotz [16], modified to include a

helium gap between the homogenized fuel region

and the fuel basket. This yields a more realistic

representation of the temperature profile through

the assembly, and takes into account the effect of

the non-uniform wall temperature distribution

around the assembly.

Axial conduction in the homogeneous fuel region

was conservatively neglected in the fuel itself, and

was modeled only in the cladding, using the

conductivity of Zircaloy modified by a weighting

scheme based on the cross-sectional area. The

effective density and heat capacity for the fuel

region was based on volumetric averages of the

properties of the helium cover gas, fuel rod

cladding, and uranium oxide fuel pellets. The

design basis axial power profile from the SAR

[11], which has a normalized peaking factor of

1.2, was used to establish the volumetric heat

generation of 8,532 Btu/hr (2.5 kW) over the

assembly along the active fuel length.

The 0.225-inch (0.57-cm) gap filled with helium

cover gas between the fuel and the basket was

modeled with solid elements and used standard

helium thermal properties for conduction and

specific heat. Convection was ignored in this

small gap. Radiation exchange between the

adjacent surfaces was modeled using MATRIX50

superelements. These were created by using

SHELL57 to designate the discrete enclosure, and

the AUX-12 hidden ray-tracing method was used

to compute view factors for each element in the

superelement. The 0.25-inch (0.64-cm) gap

between the basket and the inner shell was

5.11
modeled in the same manner, also assuming

negligible convection.

The entire package model was enclosed within

elements modeling the ISO container. For the

large air volumes in the ISO container, conduction

across the gaseous medium is negligible but

significant convection currents will be created by

thc buoyant forces due to the heated surfaces.

Surfaces with unobstructed views of'other surfaces

will also experience significant radiation exchange

that is highly dependent on the surface geometry.

Therefore, heat exchange between the package

exterior and the container interior was modeled

with internal free convection and radiation

between adtjaicent surfaces.

The radiation was implemented using the

MATRIX50 superelement procedure described in

Section 5.2 tfr the Holtec FIlI-STAR 100 model.

The convection calculations were based on

empirical relations for free convection over flat

plates and cylinders (see Section 6). Convection

was implemented using SURF152 elements. These

elements are placed on the exterior surface ofa

body and communicate with the designated sink

temperature assigned to a single node (called the

"space node") to compute the heat flux.

Because convection heat transfer rates are

expected to vary in different regions throughout

the ISO container, the single volume was divided

into 17 zones. These consisted of a zone on each

end of the package, three zones representing the

top, side, and bottom radial surfaces for each

impact limiter, and similar zones for the package

for three locations along its axial length (see

Figure 5.17.)

A sink temperature wvas defined for each zone,

computed as the average surface temperature of

the participating package and container elements

for that zone. The convective heat transfer

coefficient was assigned to the package and

container elements based on the temperature

difference between the surface and sink

temperature, and the surface geometry, as

described in Section 6. The heat exchange between

these surfaces and the space node was then

computed by ANSYS during the solution.

Figure 5.17. Zones for Convection Heat

Transfer Within the ISO Container

Convection and radiation are also the two mechanisms required to model thermal exchange from the

exterior of the ISO container. In the fire analysis,

the initial temperature distribution is obtained

from a steady-state solution with conditions

specified by IOCFR71.71 [I], followed by a

transient solution representative of the fire. For

the steady-state solution, convection is handled by

SURF 152 elements with a constant convection

coefficient of 0.891 Btu/hr-ft2-°F (5.06 W/m2 -'K)

and an ambient temperature of l00F (380C).

Solar insolation is incorporated by using SURF 1 52

elements with heat generation on the outer surface

at the rate specified in lOCFR7I [1].

During the fire, the sink node temperature for the

SURF 152 elements is set and the external

convection coefficient is computed Using a forced

convection relation derived using the gas

temperatures and velocities from the results of the

NIST fire simulation using FDS. These results

were obtained for the top, side, and bottom of the

5.12
tunnel, and applied to three zones defined on the

top, sides, and bottom of the ISO container, as

illustrated in Figure 5.18. By the end of the

transient simulation using FDS (i.e., 30 hours), the

predicted gas velocities have dropped to the point

that free convection is the only significant mode of

convection heat transfer. From this point in the

transient, the convection coefficient is computed

in the same manner as described for the steadystate initial conditions.

Thermal radiation between the container and the

tunnel during and after the fire is incorporated by

the MATRIX50 elements, as described previously,

where the top, side, and bottom temperatures of

the tunnel from the NIST fire simulation are

imposed as boundary conditions. A conservative

emissivity value of 0.9 was used for the tunnel

surfaces and ISO container exterior, to account for

the effect of-sooting.

Figure 5.18. Zones for External Radiation

Between ISO Container and

Tunnel Surfaces

The material thermal properties used in the model

are listed in Appendix C, and were obtained

primarily from the vendor's SAR [11]. Some

modifications were made to the material properties

to account for structural configuration and

expected effects of the fire. For the aluminum

honeycomb material, the significant void volume

reduces the heat transfer capability compared to

solid material. The thermal conductivity assigned

to the impact limiters was scaled by the ratio of the

honeycomb density to the solid aluminum density.

Modeling of the liquid neutron shield was

complicated by the expectation that the 56%

ethylene glycol liquid will exceed its boiling point

during a fire transient, leading to tank rupture and

vaporization of the contents, which significantly

affects the heat transfer of the package. Prior to

rupture, the liquid in the tank is expected to sustain

convective currents due to temperature gradients

through the liquid between the tank surfaces. After

rupture, empirical relations were used to obtain

separate effective conductivities for the shield tank

and expansion tank.

The empirical relations were based on correlations

by Raithby and Hollands [18], as described in

Section 6. The effective conductivity was then

determined as a function of the average tank

temperature and the radial temperature difference

between the tank inner and outer surfaces. The

material properties were updated between each

time step during the transient solution using

APDL. They were computed for 56% ethylene

glycol solution up to the point where the average

temperature reached its boiling point of 350'F

(I 77-C).

When the average temperature in the tank

exceeded the boiling point, it was assumed that

rupture occurred and the liquid was immediately

vaporized. Subsequently, the effective

conductivity was computed using dry air as the

medium. This calculation was continued during

the cool down period also. This formulation

conservatively neglects energy absorbed by the

phase change (i.e., the heat of vaporization for the

liquid), although this is mainly as a matter of

convenience, since this would constitute a very

small deduction from the total energy imparted to

the package.

5.13
The general solution procedure for this model was

similar to that for the HI-STAR 100 described

previously. The steady-state temperature solution

for normal hot conditions was computed using

solar insolation and 1000F (380C) ambient

temperature per IOCFR71.71 [I ] and used as the

initial temperature state. The insolation was

removed and the tunnel was introduced for the

transient fire analysis. The transient solution was

then obtained for the 30 hours of the NIST

simulation, representing the 7-hour fire and 23-

houLr cool down. The solution was also extended

for a total simulation time of 300 hours, in the

same fashion as described in Sections 5.1 and 5.2

for the other two package models.

5.14
6 ANALYSIS

The National Institute of Standards and

Technology (NIST) performed analyses using the

FDS code, based on the type of fire that could

have been sustained during the accident in the

Howard Street tunnel and the time required to

consume the available fuel. The results were used

to define boundary conditions for COBRA-SFS

and ANSYS evaluations of the thermal response.

of the selected spent fuel transportation packages.

Section 6.1 lists the conservative assumptions

underlying the analytical approach used and

describes the boundary conditions in detail. This,

includes temperature boundary conditions and the: l

approach used to define convection and radiation

heat transfer rates. Section 6.2 describes the initial

steady-state conditions for each cask package.

Section 6.3 describes the procedure used for the

transient calculations.

6.1 Fire Transient Assumptions

and Boundary Conditions

A number of conservative assumptions were made

in developing models and performing evaluations

of the thermal response of the three spent fuel

transport packages (TN-68, HI-STAR 100, and

NAC LWT) to the Baltimore tunnel fire transient.

The assumptions of greatest impact are listed

below.

1) Boundary conditions were taken from

predictions of peak gas temperatures in the

lower, middle, and upper zones of the tunnel

and peak surface temperatures on the tunnelfloor, walls, and ceiling. The peak values in'

each region were used to define boundary

temperatures over the entire region, rather than

using the local temperature distributions

predicted in the FDS calculation. This

approach ensures a conservative estimate of the

METHOD

boundary temperatures, since the package does.

not see the peak temperatures on all surfaces,

and in some cases may not see the peak

temperature on any surface. (For example, the

top of the package is not high enough to be

directly exposed to the peak gas temperature at

the top of the tunnel, but this value was used as

the ambient temperature for convective heat

transfer to the upper surface of the package.)

2) The package cradle and the rail car section

beneath the cradle were included in the ANSYS

model of the HI-STAR 100, but the rail car

ends and honeycomb end blocks adjacent to the

impact limiters were omitted. These structures

were neglected because they would partially

shield the package from thermal radiation from

the hot tunnel surfaces and block convection

heat transfer to the package due to the flow of

hot gas generated by the fire. The rail car was

omitted from the COBRA-SFS model of the

TN-68 cask and the ANSYS model of the NAC

LWT cask within the ISO container. This

approach eliminated any potential shielding of

these packages from thermal radiation and

convective heat transfer from the tunnel

environment.

) During the fire and the short-term post-fire cool

down period (7 hr < t < 30 hr), it was assumed

that convection heat transfer at the package

surface was due to forced convection only (due

to air flow induced in the tunnel by the

temperature gradients of the fire), using the gas

velocities predicted in the NIST analysis. This

approach neglects the possible contribution of

free convection around the package (due to

non-uniform circumferential temperatures

around the package outer shell), which would

tend to remove heat from the package. The

boundary condition was switched to solely free

6.1
convection after 30 hours, in the extrapolated

portion of the transient. This conservatively

neglects any forced convection cooling of the

package during the extended cool down period,

when the gas velocities in the tunnel are

predicted to have dropped to negligible values.

4) The effects of optical densification due to

combustion products and material degradation,

which would tend to attenuate the radiation

influence between the tunnel and package

surfaces, were not taken into account in the

boundary conditions defining the fire.

Radiation views were treated as clear and

unobscured at all times. Radiation attenuation

was also neglected between the ISO container

inner surfaces and the NAC LWT package.

5) Materials that could be expected to burn, boil

off or melt during the transient were assumed to

remain intact during the fire, to maximize the

heat input into the package. At the end of the

fire, the thermal conductivity values for these

materials were reduced to that of air. As a

result, the affected components then present an

added thermal barrier to heat removal from the

package following the fire. In addition, the

energy absorbed by these materials, due to

latent heat of diffusion or vaporization, was not

subtracted from the energy input to the package

from the fire.

6) The cask package was assumed to be the

shortest possible distance away from the center

of the fire, in order to obtain the highest

possible boundary temperatures due to the fire.

Based on Department of Transportation

regulations [12] that require railcars carrying

radioactive materials to be separated by at least

one railcar (a buffer car) from other cars

carrying hazardous materials or flammable

liquids, this distance must be at least 66 ft

(20 m).

Given these assumptions, the ANSYS and

COBRA-SFS analyses constitute conservative

evaluations of the response of the spent fuel

transportation packages. The FDS simulations for

the NIST model of the Howard Street tunnel fire

produced detailed predictions of gas flow rates,

gas temperatures, and tunnel wall, ceiling, and

floor temperatures during the 7-hour fire and 23-

hour post-fire cool down.

6.1.1 Boundary Temperatures from FDS

The FDS simulations included a significant

portion of the tunnel length, from the fire location

to the tunnel entrance. The results obtained for the

radial plane of the model at the location 66 ft

(20 m) from the center of the fire were used to

determine the boundary conditions for the analyses

with COBRA-SFS and ANSYS. As a

conservative simplification of the finely detailed

noding in the FDS simulation, the tunnel radial

geometry was divided into three regions; top, side,

and bottom (refer to the diagram in Figure 5. 10.)

Within each of these regions, the predicted peak

wall temperatures and peak gas temperatures as a

function of time (with the associated gas

velocities) were taken as representative of the

transient behavior of the entire region, rather than

following the local gradients obtained in the

detailed NIST simulation with FDS.

These temperature-vs.-time and velocity-vs.-time

values were smoothed to conservatively remove

the rapid stochastic variations typical of dynamic

fire behavior, preserving only the major peaks and

troughs related to the general physical behavior of

the simulated fire. Figure 6.1 shows these

smoothed peak air temperatures for the top, sides

and bottom regions in the tunnel at 66 ft (20 m)

from the fire center. The smoothed peak surface

temperatures for the walls, floor and ceiling of the

tunnel at this location for the same fire scenario

are shown in Figure 6.2. Figure 6.3 shows the

smoothed velocities predicted in the NIST

6.2
- - -

analyses at the locations of the peak gas

temperatures in Figure 6.1.

1100

..00

1700

1.00

1500

1400

C 1300

t* 1200

100

Fi :oo

E g00

& Boo

* 700

Goo

500

400

300

200

100

...a~~ -TopAW.To.ap -

r -SDo~blTenmp

*......4...----------...

-- L - - - - - - - - -

- - - -- -- -.-- ...-...

the boundary conditions for radiation heat transfer

between the tunnel walls and the exposed surfaces

of the cask package.

The COBRA-SFS model for the TN-68 package

incorporated the effect of the tunnel walls, ceiling

and floor by calculating a radiation heat flux at the

package surface using the local package surface

temperature and the regional tunnel surface

temperatures defined in Figure 6.2. Blackbodyview factors between the package surface and the

tunnel ceiling, walls and floor were determined

using a conventional ray-tracing scheme. (The

radiation exchange values for this geometry are

listed in Appendix D.)

A. I F.. Oa,. 'I Post-F a Owm .` ("iST Og, .-

S 5 1 0 1 0 20 25 30 Elapsed Tku. (hours) ....

Figure 6.1. BTF Peak Transient Ambient Air

Temperatures (smoothed values,

NIST 20-m data)

1000 . . . .

1700 _ _ ; a - -_ I- - '- - 4 -~ - I -_ -C *IlaagT en p -

1000 . . -

r T a , n r - -F0oaTee.p 1400 T 1 r _ _ J L 1200 r - - -- r

_1_1100I J -I- _ _.__ L _ J ._J . .

10 07a I i . . . . a

300 a7\

E 100 0 T - --- - r - - -...- ...-- r -

TOO - - - -I-0- - -

o00 - 4 - - - -- - -I

SO: a a I) I _ IL 1 a :

2- - - - - - I - - -,

0 2 4 0 a 10 12 14 is IS 20 22 24 26 28 20

Elapsed TkV (hours)

0............- - - - -

4 a-IL I - - - - -a

3B f o a ~ .

2.

F/ D m ,Pos .v a a!TD 

S ... I-...1 2 2 3

Ea p a (om a 

Figure 6.2. BTF Peak Transient Tunnel

Surface Temperatures for Floor,

Walls, and Ceiling (smoothed

values, NIST 20-m data)

The NIST fire simulation results shown in Figures

6.1 through 6.3 were used to develop the boundary

conditions applied to the TN-68, HI-STAR 100,

and LWT models. The peak surface temperatures

in Figure 6.2 were applied to all corresponding -:

tunnel surfaces in the top, side, and bottom regions

of the ANSYS models for the HI-STAR 100 and -1

LWT packages. These temperature values defined

Figure 6.3. BTF Peak Transient Horizontal

Velocities near Package Surface

(smoothed values, NIST 20-m data)

6.1.2 Convection and Radiative Heat

Transfer Boundary Conditions

The NIST analyses showed that the thermal

gradients created in the tunnel due to the fire

would result in significant air flow past a body.

located near the fire. This fire-forced convection

would significantly affect heat transfer around the

package and could have a strong influence on the

package outer shell surface temperatures. The

smoothed air temperatures in Figure 6.1 and

velocities in Figure 6.3 were used to define local

6.3
time-dependent Nusselt number values at the top,

sides and bottom of the package. These values

were used to define the local surface heat transfer

coefficient for the three computational models.

To maintain consistency between the three

models, the same Nusselt number correlation was

used to define convection heat transfer at the

package surface. The selected correlation gives

the Nusselt number for gas flow over a flat or

slightly curved surface at zero angle of attack [ 1 9],

and has the form

for laminar flow (Re, < 500,000)

NUL = 0.665 ReL'2 Pr 3

for turbulent flow (Rei > 500,000)

NUL = 0.032Re" Pr'

The characteristic length, L, used to define the

Nusselt number and Reynolds number for this

application is the package body horizontal length.

In the COBRA-SFS modeling of the TN-68, the

characteristic length was specified as 160 inches,

based on the length of exposed package body. In

the ANSYS model of the HI-STAR 100, the

characteristic length of 173 inches was used, based

on the length of exposed package body. For the

NAC LWT, a value of 240 inches was used, based

on the ISO container wetted surface length.

The peak air temperatures (see Figure 6.2) from

the NIST analysis define the ambient sink

temperature around the package during the fire

and post-fire intervals. The Nusselt number

defines the rate of heat transfer from the package

surface, which allows both codes (COBRA-SFS

and ANSYS) to calculate the convection heat flux

at the package surface. Using the above

relationship, local surface temperatures, T,, are

calculated, and the convection component of the

heat flux at the surface is solved for using the

formula

q conv = NukL L(T, -Tair ) LL

where k = thermal conductivity of ambient air

L = characteristic length

T, = package surface temperature

Tair = ambient external air temperature.

Separate boundary types were defined for the top,

sides, and bottom surfaces of the package using

the external air temperatures shown in Figure 6. 1.

The velocities in Figure 6.3 were used to define

the Reynolds number so the boundary conditions

on the package could change with time as the

transient proceeded. Figure 6.4 shows the resulting local heat transfer coefficients calculated for

the COBRA-SFS evaluation during the 30 hours

of the NIST transient calculation, at the top, sides

and bottom of the TN-68 package.

0.55

0.5

02 0.25

i- c 0.35

0.1

005

.I

I I I , I I

I .. fi F I--1 -(top) - i--i

i'[-I J I .I I (side) _ I_

I I I 1 -(bottom) 1

1 - - - - - - - - - - - -

Fire D.Gon PIt-F. DIr D f I- NIS It0 2 4 6 a 10 12 14 16 18 20 22 24 26

Elapsed Tin. (bhous)

26 30

Figure 6.4. Heat Transfer Coefficients at

Package Surface from NIST 20 m

Air Temperature and Velocity

Predictions

In addition to convection heat transfer between the

transport package and the surrounding air during

the transient, radiation heat transfer between the

package surface and the tunnel ceiling, walls, and

floor was also captured. Boundary conditions to

6.4
- -

define radiation heat transfer between the package

surface and the tunnel ceiling, walls, and floorwere included in the COBRA-SFS model in the

following fashion. The total heat flux at the

surface of the package is the sum of the two

components:

q = q + q"rod

The tunnel surface temperature profiles shown in

Figure 6.2 were used to define the radiation heat

flux as an additional boundary condition at the

package surface using the relationship

q rad = AiBijSB(TpackagoC Tsuf)

where £; = emissivity of surface i

Bij = blackbody viewfactor from surface

itoj

c5s3 = Stefan-Boltzmann constant

Tsurf = tunnel ceiling, wall or floor

surface temperature

Tpackage = package surface temperature.

The blackbody view factors between the package

surface and the tunnel ceiling, walls, and floor

were determined using a Monte Carlo ray tracing

scheme based on the package diameter and a

uniform axial node length along the length of the

package. These are presented in Appendix D.

Radiation interaction between the tunnel surfaces

and the package surfaces in the HI-STAR 100

model (and the ISO container surfaces in the NAC

LWT system) was established through the use of

ANSYS superelement definitions, as described in

Section 5 above. The NIST tunnel surface

temperature predictions (see Figure 6.2) were then

used to establish the tunnel surface boundary :

condition temperatures. The emissivity of all

tunnel surfaces and the package surface was

assumed to be 0.9 for all evaluations during the

fire and post-fire transient.

6.1.3 Extrapolated Boundary Conditions

for Long-Term Cool Down

NIST's FDS analysis was carried out for a 7-hour

fire and 23-hour post-fire cool-down. To

determine the long-term temperature responses

and explore the effects of prolonged exposure to

post-fire conditions in the tunnel, the post-fire

duration was extended to 300 hours (273 hours

after fire cessation). Temperatures predicted in the

NIST analysis for 30 hours were extrapolated from

30 hours to 300 hours using a power function to

realistically model cool-down of the tunnel environment. The extrapolated predictions are

presented in Figures 6.5 and 6.6 for the air

temperatures and wall temperatures, respectively.

300

27S

250

225

C 200

1175

T 150

2. 125 E

A 100

75

s0

r 1 T' 1 - 0~T I I . I ,

I ....... IepAW..A,T.p

\ I It | I I - I a II -STopAirTsmp -

\ I I *.........L I - - - -- -- * 1 - ...- - - - I 1

I I I I I I I

,4,4 , I

I ;EwapobW DabSet e1 , ,

0 30 so t0 120 10 100 210 240

Ehpsod ThnJ (holus)

270 300

Figure 6.5. BTF Peak Transient Air

Temperatures for Top, Side, and

Bottom Regions (NIST and

Extrapolated Data Sets)

About 20 hours into the transient, the velocities

predicted in the NIST calculations have dropped to

values of I to 2 ft/s (0.3 to 0.6 m/s) or less (refer to

Figure 6.3). Heat transfer at the package surface

for these flow conditions is a complex mixture of

forced convection (due to air flow induced in the

tunnel by the wall temperature gradients of the

fire) and free convection (driven by the nonuniform circumferential temperatures of the

package outer shell). At velocities below about 3

6.5
to 5 ft/s (I to 1.5 m/s), heat transfer rates predicted

assuming forced convection are generally lower

than the heat transfer rates due to natural

convection around the package body for these

temperature conditions.

300

275

250

225

E- 200

, 175

< 155

E 125

75

so

25

Wall T-np

FI.., T-

---- - -- -

NIS TData So

Exlraolald DOtS Sel

--for flow along a vertical plane or cylinder:

--laminar flow (I04 < Grf Prf < IO)

( )

--turbulent flow (Grf-Prf > 109)

h = 1.31 (AT)"'

where

h = heat transfer coefficient, W /(m - 'C)

AT T, -Tp,-C

T= surface or wall temperature, 0C

T= ambient temperature, 'C

L = vertical or horizontal dimension, m

Grf = Grashoff number of the gas at film

temperature, Tf= (Tw + To)/2

Prj = Prandtl number of the gas at film

temperature, TV= (Tw + Th,)/2

u . r .. . .; 0 30 60 90 120 150 150 210 240 270 300

Elapsed Time (hours)

Figure 6.6. BTF Peak Transient Surface

Temperatures for Floor, Walls,

and Ceiling (NIST and

Extrapolated Data Sets)

To avoid the modeling uncertainties associated

with mixed-mode heat transfer, forced convection

only was assumed until the end of the NIST

simulation, at 30 hours into the transient. From 30

hours to 300 hours, the heat transfer was assumed

to be natural convection only. This ensured a

conservative treatment of convection heat transfer

from the package surface during the entire

calculation, since free convection to surface heat

transfer from the package is ignored in the cool

down from 7 to 30 hours, and forced convection is

neglected in the period from 30 to 300 hours.

For consistency, the natural or buoyant convection

coefficients were those utilized for determining the

pre-fire component temperature distributions (i.e.,

Hot-normal Conditions of Transport, as defined in

10 CFR 71.71(c)(1)([l].) The heat transfer

coefficients were defined for the appropriate

surface geometries using the following

relationships [20]:

--for flow over a horizontal cylinder:

--laminar flow (104 < Grj Prf < 109),

h =1.32 AT )

where

d = diameter, m

--turbulent flow (Grf Prf > 109),

h = 1.24 (AT)11 3

--for flow over a horizontal heated plate facing

upward (cool side facing downward):

-- laminar flow (104 < Grf-Prf < IO),

6.6
h =1.32( L )

- turbulent flow (Gry Prf > 109),

h =1.52 (AT)' 13

--for laminar flow (104 < Grp Prj < 109) over a

heated plate facing downward (cool side facing

upward):

h =0.59(ATJ

Definitions of material properties for use with

these correlations were taken from Table A-3 of

Kreith [21].

6.1.4 Heat Transfer through Liquid

Neutron Shield

An empirical relationship for effective

conductivity incorporating the effects of both:

conduction and convection was used to determine

heat exchange through the liquid neutron shield.

In the SAR analysis for the LWT package [1 l],'the

effective conductivity of the ethylene glycol'

mixture for conditions below 350'F was

determined using the correlation of Bucholz [22],'

which defines the ratio of the effective

conductivity to the actual thermal conductivity as!'

equal to the Nusselt number, such that'

k .2 = Nu = 0.135(Pr2 Gr/(l .36 + Pr))0 278

kc where~~ . k. . utiio

where ke=ff effective thermal conductivity of

material in node

k= thermal conductivity of motionless

fluid in node

Nu = Nusselt number

Pr = Prandtl number

Gr = Grashoff number

The BTF transient is outside the range of the

Bucholz correlation, and it yields unrealistically

large values for kff for these conditions. An

alternative correlation from Raithby and Hollands

[18], based on heat transfer between two

concentric cylinders, was used in this analysis

instead. This correlation produces reasonable

values of kjee and the transient conditions are

generally within its'applicable range.

The form of this correlation is similar to the

Bucholz correlation in that it equates the Nusselt

number to the ratio of the effective conductivity

over the actual conductivity, but in the Raithby

and Hollands formulation, the Nusselt number is

expressed as

keff = Nu = 0.386D,(Pr/(0.861 +Pr)) 25Ra02 5

k k

where Ra = Rayleigh number (Ra = Pr*Gr)

Pr = Prandtl number

Gr = Grashoff number (based on the

temperature difference across the

annular gap)

The variable D, is a dimensionless parameter

based on the geometry of the annulus, and is

defined:

D Dr d=1343(/1)35+/D31)1 ln(D. /Di) 1

where Do = annulus outer diameter

Di = annulus inner diameter

d = width of annulus.

Figure 6.7 shows a plot of the Nusselt number

predicted with these two correlations for the liquid

(56% ethylene glycol and water mixture) in the

neutron shield annulus.

6.7 '
e0 -~ - -,- -- or -- ~- - - I

- -B hola(om - SAR)

70 -Raihby & Ho.ada-1975 - (Guye,) -

Ven RScI Chy (I-coperal & DeWII)

0 So N. Iro- SAR - dT---af

= a 0o~R.by & HolIlod- a refi onojieet dT |

- - - - - - - -4T to -- -

E i40a

10

1,E102 1 E.03 1 E504 1 E,05 1 E006 1 E-07 1 00 I E.09 1 E010

PrG, (-nitless)

ra

c

I

S E102 I

, x -- dT 500

d--T= 300

-- dT 200

10E-01- - - - - -- dT= 100

dT 70

it , -dT 5n

_ -- =25

1.E.00 - - d- i 10

--07 1

I E 02I

1 E-01 - -,

10E-02L _|

200 700 1200 1700 2200 2700

Average Ann-l-s Temperature- F)

Figure 6.7. Nusselt Number for Heat Transfer

in Liquid Neutron Shield

Figures 6.8 and 6.9 show the effective

conductivity for the annulus as a function of the

average temperature and temperature difference

for the liquid neutron shield tank and expansion

tank, respectively. (The sharp discontinuity in the

1 E.02

I E001

1 E-01

- -d- -00

dT 300

4dT 200

- - - -I - - dT 100

dT 70

dT 50

25dT2

dT= 10

-dT= I

'-- I I I

= ...r=__ -

Figure 6.9. Effective Conductivity of

Expansion Tank Contents

6.2 Initial System Component

Temperatures

The normal conditions of transport described in

10 CFR 71.71 [1] were used as initial conditions

for each analysis. All three packages were

subjected to an ambient temperature of 100IF

(380C), with solar insolation. For pre-fire

conditions, the package surface was given an

emissivity value representative of its surface finish

(e.g., 0.3 for bare stainless steel, 0.85 for painted

surfaces.) In the ANSYS models for the HI-STAR

100 and NAC LWT systems, thermal radiation

heat transfer to ambient was modeled using

surface effect elements (SURF 152).

Convection from the surface of each package was

modeled with a similar set of surface effect

elements. The natural convection correlations for

buoyancy driven flow discussed above were used

to simulate convection heat transfer at the package

surface. For the COBRA-SFS model of the TN-68

package, the surface boundary conditions also

included natural convection and thermal radiation.

Heat generation rates for decay heat loads of

68,240 Btu/hr (20kW) for the HOLTEC HI-STAR

100, 8,530 Btu/hr (2.5kW) for the NAC LWT, and

72,334 Btu/hr (21.2kW) for the TN-68, were

1.E 02_ -

200 700 1200 1700 2200

Average Annulas Temperature (F)

2700

Figure 6.8. Effective Conductivity of Neutron

Shield Tank Contents

curves on both plots represents the abrupt phase

change assumed when the average temperature of

the liquid reaches the boiling point of the ethylene

glycol and water mixture.) For low values of the

temperature difference, the results approach those

for the conduction-only case.

6.8
applied, with appropriate peaking factors, over the

active fuel region.

A steady state normal condition temperature

distribution for each package was obtained to

establish pre-fire conditions. The hot-normal

condition temperatures for each package were

verified against the results reported in the relevant

SAR. Normal condition temperatures from the.

ANSYS solution for the HI-STAR 100 are

provided in Figure 6.10. (Appendix E contains

additional plots showing the detailed temperature

distributions for these conditions predicted for the

HI-STAR 100.) The peak clad temperature

predicted with ANSYS for the HI-STAR 100 is

673-F (3560 C), compared to 7010F (3720 C)

reported in the SAR[l0].

Since COBRA-SFS does not have a graphical

post-processing module, it is not possible to

produce similar color-flooded thermographs for

the TN-68 evaluation. However, the analysis

results are similarly in very good agreement with

the corresponding SAR values. The COBRA-SFS

calculations predicted a peak clad temperature of

4850F (2520 C) in the TN-68 package, compared to

490'F (2540 C) reported in the TN-68 SAR [9].

Component temperature comparisons of results

determined in this study and those published in the

applicant's SAR documentation are presented in

Table 6.1 for the TN-68 analysis with COBRASFS and Table 6.2 for the HI-STAR 100 analysis;

with ANSYS. These tables show that the

analytical results obtained for the TN-68 and HISTAR 100 are in very good agreement with the

results presented for the corresponding cases in the

respective SARs. Minor differences between the

SAR results and those obtained in the current

study are due to differences in modeling detail and

simplifying assumptions employed in the, SARi,

models. For example, the SAR analysis of the HISTAR 100 neglects the effect of the support cradle,

on component temperatures in the evaluation for

the hot-normal conditions of transport.

AKSYS L1.a APR 1 2004 15:53:37

S._AL SOLWTOtl

SUB -1 T1WN_ .100L02

TE?49 4AVG) RM-c

_ R575.5,

.' _m rowera raphscs

- _iE } oAVUS-Kate3

. I _ ol. , _,5.}22

S 672.719

UM.29 .40

67 2S1 7S1

Figure 6.10.HI-STAR 100 Package Hot-Normal

Condition Temperature

Distribution

For the TN-68 model and the HI-STAR 100

model, the results presented in Tables 6.1 and 6.2

are representative of the initial conditions for the

fire transient analysis. For the NAC LWT,

convection from the surface of the ISO container

was modeled with a similar-set of surface effect

elements. The natural convection correlations for

buoyancy-driven flow discussed above were used

to simulate the convective heat transfer within and

at the ISO container surface.

Table 6.1. TN-68 Hot-Normal Component

Temperatures

Current Study - SARValues

: (COBRA-SFS) F (°C) Component -F(°C) (Tabe 3-1 191)

Fuel Cladding 485 (252) 490 (254)

Basket plate 467 (242) 469 (243)

Basket Rail 332 (167) 319 (159)

Inner Shell 292 (144) 262 (128)

Gamma Shell 285 (141) 260 (127)

Package Bottom 261 (127) 254 (123)

Seals 260 (127) .234 (112)

Radial Neutron 256 (124) 244 (118)

Outer Shell 243 (117) 204 (96)

6.9
Table 6.2. HOLTEC HI-STAR 100 HotNormal Component

Temperatures

Current SAR Values

Study OF (OC)

(ANSYS) (Table 3.4.10

Component (C)01) [F 

Fuel Cladding 673 (356) 701 (372)

MPC Basket Centerline 667 (353) 667 (353)

MPC Basket Periphery 432 (222) 430 (221)

MPC Outer Shell 345 (174) 315 (157)

MPC/Overpack Helium

Gap Outer Surface 305 (152) 291 (144)

Radial Neutron Shield

Inner Surface 270 (132) 271 (133)

Overpack Enclosure

Shell Surface 250 (121) 222 (106)

Axial Neutron Shield 223 (106) 292 (144)

Impact Limiter

Exposed Surface 160 (71) 121 (49)

Overpack Closure Plate 250 (121) 163 (73)

Overpack Bottom Plate 375 (191) 295 (146)

A heat generation rate equivalent to a decay heat

load of 8,530 Btulhr (2.5 kW) was applied, with

appropriate peaking factor, over the active fuel

region. The NAC LWT is currently licensed for a

maximum decay heat load of only 2.1 kW. The

value of 2.5 kW was selected for this analysis

because an Amendment to the SAR that would

increase the decay heat load limit to 2.5 kW is

currently under review 11]. This approach

ensures a conservative decay heat load for the

package in the fire transient analysis.

The steady-state initial condition temperature

distribution predicted for the NAC LWT package

was verified against the results reported in the

SAR [11]. Direct comparison is not possible,

because the SAR [II] does not include any

analytical cases similar to the detailed 3-D model

used in this study. Due to the relatively low

associated decay heat load capacity, the applicant

chose to perform a series of highly conservative

evaluations to qualify the system for its Certificate

of Compliance (CoC).

The most complex models presented in the SAR

111] involve simple 2-D ANSYS cross-sections in

which the cutting plane includes the expansion

tank as well as the neutron shield tank. This

approach does not allow axial heat flow out of the

plane of the 2-D cross-section, and also assumes

that the decay heat load axial peak occurs on that

plan, placing the spent nuclear fuel peak decay

heat location under two concentric neutron shields.

This provides conservatism for a steady-state

analysis, since the expansion tank makes a longer

conduction path over which to dissipate the decay

heat. For the fire transient, however, the

assumptions in this 2-D model would have the

effect of limiting the heat input to the cask from

the fire, and would not constitute a conservative

approach.

ANSYS cross-sectional models were also used to

represent a 1.41 kW 25-rod BWR basket assembly

and a 2.1 kW high bum-up PWR assembly, with

detailed representation of the fuel pins, pin tubes,

and can weldments with the pins resting on the pin

tubes via point contact. These models included

the ISO container, with boundary conditions that

included solar insolation and 1000F (38 0C)

ambient temperature.

The design basis model presented in Amendment

34 of the SAR [11 ] for a 2.5 kW PWR assembly

also used a 2D representation of the cask. This is

a HEATING5 model, and consists of a 2-D

axisymmetric representation using effective

diameters for the basket and fuel assembly. This

model does not include an ISO container or impact

limiters, convection at the assembly end cavities is

neglected, and the ambient temperature boundary

is specified as 130'F (54 0C).

None of these cases from the SAR [11] use

assumptions or boundary conditions identical to

the initial conditions assumed for the fire transient

in this analysis, but there are sufficient similarities

to allow reasonable comparisons to be made for

6.10
verification of the 3-D ANSYS model predictions.

The results for these three cases are reported in

Table 6.3.

Table 6.3. NAC LNN'T Component

Temperatures at Various Decay

Heat Loads

2.5 ; .41k 2.1 kW.

v-:' . F (C) II-. of,(a(,) oF (off) iv

.(Talfle ! (Table (Tabl

(onipoli-en.4 3.42 ) -7 11)3.4-10 1111)

Fuel Cladding 472 (244) 358 (181) 671 (355)

Aluminum PWR

Insert 276 (136 * 394 (201)

Inner Shell 274 (134 249 (121) 385 (196)

Gamma Shicld 273 (134 248 (120 375 (191

Outer Package

Surface 229 (109) 185 (85) 308 (153

Neutron Shield 238 (114) 235 (113) .306 (152

Lid Scal 227 (108) *

Drain/Vent Ports 231 (111) * *

Impact Limiters * * *

ISO Containcr I

* value not reported by applicant

At first glance, the differences between the results

obtained with the two models appear to be rather

sizable. The peak clad temperature predicted with

the ANSYS 3-D model is 4340 F (2230C),

compared to 472IF (2440C) reported in the SAR

[11]. The other temperatures shown in the table

are also considerably lower for the ANSYS model,

compared to the corresponding SAR values.

However, this is an expected result, given the

modeling differences between the two cases. A

more significant observation for the purposes of

this comparison is to note that the differences in

peak component temperatures between the two

models are consistent.

JUS 16 :OOS

| SEMP (CAW)

P( _~vrrphl r

EFACET-1

__ SMN. 30

I'1.428

_.3167 S

3186.285

. . C:]338.213

H20S7.14

Figure 6.11.;LWT Package hlot-Normal

Condition Temperature

Distribution (2.5 kX ' Deca2 heat)

Figure 6.11 shows the predicted temperature

distribution from the ANSYS solution for the 3-D

model developed for the current study, obtained

using 130 0F,(540 C) ambient temperature with aK

2.5 kW decay heat load. This calculation was

performed in addition to the initial conditions case

at 1000F (380 C) ambient temperature, as a

verification case for comparison to the results

obtained for the 2.5 kW case reported in the SAR

[Il]. The 2-D axisymmetric model in the SAR

[11], which used an ambient temperature of 130'F

(540C), is the most similar to the initial conditions

in the fire transient for the purposes of this

comparison, despite the exclusion of the ISO

container.

Table 6.4 presents detailed component temperature

results obtained with the 3-D ANSYS model,

compared to the values published in the SAR [11]

for this decay heat load.

I

ii

The radial temperature drop from the peak fuel

cladding temperature to the outer cask surface

temperature is 2340F (1300C) for the 3-D ANSYS

model, compared to 2437F (1355C) for the

HEATING5 axisymmetric model. This close

agreement strongly suggests that the axisymmetric

model featured in the SAR predicts essentially the

components temperature distribution as the more

detailed ANSYS model, and that most of the

differences in the point-to-point temperatures

predicted with each model are due to the

6.11
differences in assumed external boundary

conditions (I OOTF vs. 130'F (380C vs. 540C)) and

differences in modeling details.

conservative, and therefore can be expected to

yield a higher estimate of the peak temperature.

Table 6.4. NAC LWT Component

Temperatures at 2.5 kW Decay

Heat Load and 130'F Ambient

SAR

Values

Current Study (Table

(ANSYS) 3A-2 1111) AT

(component C F (OC)° F (O(' )

Fuel Cladding 434 (223) 472 (244) 38 (21)

Aluminum PWR

Insert 265 (129) 276 (136) 11 (6)

Inner Shell 228 (109) 274 (134) 46 (26)

Gamma Shield 227 (108) 273 (134) 46 (26)

Outer Packagc

Surface 200 (93) 229 (109) 29 (16)

Neutron Shield 204 (96) 238 (114) 34 (19)

Lid Seal 164 (73) 227 (108) 63 (35)

Drain/Vent Ports 164 (73) 231 (111) 67 (37)

Not

Impact Limiters 167 (75) Modeled

Not

ISO Container 167 (75) Modeled

As shown in Figure 6.12, the temperature

gradients in the cask are such that heat spreads and

dissipates axially as well as radially. As a result,

the 3D geometry yields a more realistic

representation of the heat flow in the cask. The

conservative measures used in the simpler 2-D

ANSYS cross-sectional models and the

HEATING5 axisymmetric cross-sectional model

(as reported in the SAR [11]) will tend to result in

higher predicted temperatures for steady-state

conditions. Other associated modeling

assumptions and simplifications, including

boundary conditions developed by the applicant,

also tend to drive up component temperatures,

compared to what might be obtained with a

detailed 3-D representation.

ANSYS 8.0

JUi 16 2 5

15:51 * 36

MOAL SOLUTION ST.EP-i

SU8 I

TIME-. 11OE-03 TEMP (AVG) RSYS.-0

Po,.erG rphi-s EFACET. I AVWES-Mat

S6 4 -10 6 mM X417.459

11=.841

139.662

179.R

219.047 __ 2396.86

256.724

2 765S7

3186253

3386 694

357.935

377 776 ri 397 618 -417.45'

Figure 6.12 shows the temperature distribution

predicted with the ANSYS 3-D model for the

initial steady-state conditions before the fire

transient. The boundary conditions for this

calculation are from the Normal Transport

Condition case, as described in 10 CFT 71.21 [I].

The hottest fuel temperature is predicted to occur

near the center of the assembly, at a location that

corresponds closely to the cross-section with the

highest decay heat (i.e., the location of the

maximum axial peaking factor.) This location is

some distance away from the part of the cask

covered by the expansion tank. This shows that

the assumption placing the peak location under the

expansion tank, which was used in the 2-D crosssection model in the SAR [ Il], is markedly

Figure 6.12. LWT Package Normal Condition

Temperature Distribution

(2.5 kW Decay Heat)

The main concern in analyses for normal transport

conditions is to determine a conservative rate of

heat removalfrom the package. The approach

employed in the SAR should yield conservative

estimates of peak internal temperatures for the

analysis. However, for the fire transient, the main

6.12
concern is the amount of heat that the external fire

can put into the package. In the fire transient

calculations, a best estimate of component

temperature distributions and heat transfer paths is

more appropriate.

The conservative approach is to choose

assumptions that tend to enhance the heat transfer.

paths, making it easier for heat to move into the.,

package from outside. For example, the treatment

of internal gaps between components is

conservative in the SAR, in that gaps are assumed

to be as large as possible. In the ANSYS model

for the fire analysis, the shrinkage gap between the

lead shielding and cask shell is included during the

initial steady-state calculation, but is ignored

during the fire transient. This approach tends to

minimize heat loss from the package at the initial

conditions, but then maximizes the heat input into

the package internal components during the fire.

The pre-fire peak component temperature results

determined in this study for the Normal Transport

Condition case at 1000F (380C), as described in

10 CFR 71.71 [1] are shown in Table 6.5.

Table 6.5. NAC LNVT Pre-Fire Component

Temperatures at 2.5 kW Decay

Heat Load and 1001F Ambient

Component Temperature

Component OF (OC)I

Fuel Cladding 418 (214)

Aluminum PWR Insert 242 (117)

Inner Shell 205 (96)

Gamma Shield 204 (96)

Outer Package Surface 176 (80)

Neutron Shield 180 (82)

Lid Seal 138 (59)

Drain/Vent Ports 138 (59)

Impact Limiters 141 (61)

ISO Container 140 (60)

6.3 Tunnel Fire Evaluations of Rail

Packages

The Baltimore tunnel fire simulations for the three

transport packages were conducted in three

phases. These were the pre-fire steady-state (hotnormal) conditions of transport, the fire

(consisting of the first 7 hours of the transient),

and the post-fire phase. For the pre-fire steady

state, the hot-normal conditions of transport were

evaluated with solar insolation and a 1000F (380C)

ambient temperature, according to 10 CFR

71.71(c)(1) [1]. External heat transfer was

assumed to be free convection in still air with

radiation to the environment. This conservatively

established component temperatures to serve as

initial conditions for the transient.

For the fire phase of the evaluation (0 < t < 7

hours), the energy input due to solar insolation

was set to zero, the tunnel surfaces were

introduced, and the transport package and tunnel

surfaces were assign an emissivity of 0.9 to

represent surfaces affected by sooting. A forced

convection regime was assumed to exist on the

exterior of the package, based on the gas velocity

results from the analysis performed by NIST.

These results were used to determine the surface

heat transfer coefficient, and with the gas

temperatures from the NIST analysis defining the

ambient boundary temperature, the convective

heat flux at the package surface could be

determined in the solution for the local surface

temperature. Tunnel wall temperatures were also

taken as boundary conditions from the NIST

calculations, and radiation from the tunnel walls

was also accounted for in the evaluations. In

addition to these measures, all aluminum

honeycomb, neutron shield resin materials, or

wood were assumed to remain intact during the

full duration of the fire to maximize heat input

during the fire.

6.13
For the post-fire phase of the evaluation (t > 7

hours), properties of the neutron shield resin

materials, wood, and selected portions of

aluminum honeycomb sections were replaced with

thermal properties identical to those of air. This

change in material properties simulates the

degradation of the materials due to the fire and has

the effect of reducing the rate of heat release

during the post-fire phase. Conservatively, the

energy absorbed in the degradation of these

materials was not subtracted from the heat input of

the fire to the package. Other model-specific

conservatisms were also incorporated, as described

above in Section 5.

Analysis of the post-fire phase was carried out for

a duration of 293 hours. This included the 23

hours of the post-fire portion of the transient

predicted by the FDS analysis, plus an additional

270 hours in which boundary conditions at 30

hours were extrapolated to 300 hours, using a

power function (as discussed in Section 5 above.)

Purely forced convection heat transfer correlations

(based on the NIST gas velocities and

temperatures) were imposed for the post-fire phase

of the simulation from 7 hours to 30 hours. The

forced convection boundary condition at the

package surface was then transitioned to free

convection correlations, to establish the buoyant

convective coefficients of heat conductance for the

remainder of the evaluation period.

Results obtained in the evaluations of the three

packages are discussed in Section 7.

6.14
7 ANALYSIS RESULTS

Due to the temperature limits on the spent fuel

cladding,'closure seals, impact limiter core

materials, and neutron shield core materials, these,

components are the most important elements to

consider in evaluating the response of the' transport

systems to the fire scenario. The peak cladding

temperature limit is important because the

cladding is the primary fission product

containment boundary for the spent fuel. The

temperature limit for the closure seals is important

because these seals constitute the outer-most

containment boundary for the package. The

temperature limits for the neutron shield material

and impact limiters are important because these..

materials are the most vulnerable to damage or

destruction during the fire. The results of the

analyses for the three rail packages were evaluated

primarily in relation to the peak predicted

temperatures for these components in the fire

transient.

These analyses indicate that the spent fuel

cladding reaches a peak temperature of 8871F

(475'C) in the HI-STAR 100 package, 845'F

(452°C) in the TN-68 package, and 10990F

(5930C) in the NAC LWT system. Peak cladding

temperatures for the TN-68 and HI-STAR 100

packages are below the currently accepted short

term temperature limit5 of 10580F (5700C) for

Zircaloy clad spent nuclear fuel under accident

conditions [23]. The peak cladding temperature

determined for the LWT exposed to this

hypothetical accident event exceeds the currenily5 The short-term temperature limit of 10580F (5700C) is

based on creep experiments performed on two fuel

cladding test samples which remained undamaged (i.e.,:'

no significant observable damage) when held at 10580F'

(570'C) for up to 30 and 71 days [24]. This temperature

limit is a relatively conservative limit, since the

temperature at which Zircaloy fuel rods actually fail ly

burst rupture is approximately 13820F (7500C)[25]. '

accepted short term temperature limit by

approximately 41 0F (23°C).

The transient results for each of the three systems

are discussed in detail below. Section 7.1

discusses the response of the TN-68 package

during the fire. Section 7.2 presents'results for the

HI-STAR 100 package. Section 7.3 discusses the

response of the NAC LWT package.

7.1 TN-68 Fire Transient Results

The COBRA-SFS model of the TN-68 package

consists of a total of 530,228 computational nodes

that are solved for each time step. This yields an

overwhelming volume of output that must be

sorted, sifted, and processed to produce a coherent

picture of the response of the package to this fire

scenario. The following three subsections present

the peak temperatures versus tinie for selected

components, as determined with COBRA-SFS for

the TN-68 package subjected to the fire transient

conditions described in detail in Section 6. The

results are presented separately for the three main

phases of the transient. Section 7.1.1 discusses the

predicted response of the TN-68 package during'

the fire. Section 7.1.2 presents results for the postfire transient over the duration of the NIST

simulation (to 30 hours.) Section 7.1.3 discusses

the response to the postulated long-term post-fire

conditions, out to 300 hours.

7.1.1 TN-68 During the Fire

Figure 7.1 shows the initial temperature response

of the TN-68 package predicted with COBRA-SFS

during the fire, portioii of the transient. The fire

bu'rns for the first 6.75 to'7 hours of the transient

(see Figures 6.1 and 6.2 for boundary temperatures

representing the fire with the package 66 ft (20 m)

from the fire'center.) During this time, the outer

7.1
surface temperature of the package shell increases

quite rapidly. The maximum temperature of the

package surface increases at a rate of up to

I0.50 F/min. (5.80C/min.), reaching a peak

temperature of 1789'F (9760C) at about 6.3 hours

into the fire. The maximum temperature of the

neutron shield material also shows a relatively

rapid increase, reaching a peak of 13550F (7350C)

at approximately 6.9 hours into the fire.

1000 T1700 v

1600

1500 -

1400

1300 0

0 1200

1100 I

% BIOO 300-

E, 800

e700

600

So0 o

-Fuel Cbddeg (Peak) _

-Oute, Shef

- - Shield

-Garma Shield COlere) /

Gamea Sh601 SIneer)

-SuepportlRaile - t-___1

-Backer Plates - - 1

-lerpaclurLirier) (fak) - ...-------1

-DraIn Plg/Clo-re Seal (Peak) - - -- - - - -

1- - - - - - - -

, ... 

----- S~ - t- - -----= 400

300- -- - -

2tt~Ii

M nn!

0 1 2 3 4

Elapsed Tire (hour-)

Is

Figure 7.1. Maximum Temperature Histories

for TN-68 Package Components

During Fire Transient

The internal components of the package show a

very slow thermal response during the fire. The

support rails and gamma shields take more than an

hour to show any noticeable increase in

temperature. Approximately 3.5 hours elapse

before the peak fuel cladding temperature rises as

much as I 'F above the initial steady-state peak

temperature of 486 0F (252 0C). The peak

temperatures of the basket tubes and poison plates

rise only about 4 0F (2.2 0C) in the first four hours

of the fire. During this time period, the peak

temperature on the outer shell of the package is

predicted to go up to 1647 0F (897 0C), the

predicted peak temperature of the neutron shield

rises to 1042 0F (561'C), and the peak temperature

on the gamma shields increases to 549 0F (287 0C).

By the end of the fire, marking the point at which

all volatile flammables are consumed, the peak

clad temperature has risen to only 673 0

F (356 0C),

and the peak temperature of the basket tubes and

poison plates is at about 714'F (379 0

C). The outer

shell of the package is predicted to have a peak

temperature of 1599 0F (871'C) at the end of the

fire, with the neutron shield at 1347 0

F (731 0C) and

the outer gamma shield at 886 0F (474 0C).

Figure 7.2 shows midline temperature profiles

from top to bottom vertically through the package

cross-section, including the package shell, support

rails, and basket structure during the fire at

approximately hourly intervals. The large

difference in the predicted rate of increase in

temperature for the internal and external

components of the package is illustrated by these

profiles. The temperatures of the nodes modeling

the basket tubes and poison plates change very

little during the fire. The support rails and gamma

shield nodes heat up relatively slowly, while the

outer shell and neutron shield region increase

rapidly in temperature in response to the fire.

1700 -_ 7 0 hr,

1600 -... - - - - -62 hr

1500 - e--6 0 hre1400 .5- - - 0hs -e

1300 Ho 4.0 h,0

1200 - - - - - -_ - - 3.0 hC1100 - - , - - 2 0 hrs - - - - -i

a 1000 11 Oh,

E 60 ...-we -Sladyelare Before Fre

2500

30°O0

1 2 3 4 6 7 B 9 10 11 12 13 14 15 16 17 10 19 20 21 22 23 24 25 26 27

Solid Nedo Location (top to bottern)

Figure 7.2. Temperature Profiles Top-ToBottom Through TN-68 Package

Axis During Fire Transient

The neutron shield and gamma shields insulate the

basket and fuel assemblies from the fire, but the

slow response is also due in large part to the huge

thermal inertia of the package components. The

68 assemblies within the package comprise on the

order of 20 to 25 metric tons of material (mainly

7.2
uranium dioxide and Zircaloy), with roughly

8 metric tons of material in the basket (mainly

steel and borated aluminum.) The inner and outer

gamma shields consist of approximately 40 metric

tons of carbon steel, while the outer shell of the

package is approximately 5 metric tons of steel.

Even under the severe heat load imposed by the

sustained high temperatures of a fire lasting 7;

hours, it takes time to raise the temperature of such

a large mass of material, even with its internal heat

generation due to the spent fuel assemblies.

A significant detail discovered during the evaluation of the TN-68 is that during the first quarter of

the fire transient, the total heat flux associated

with radiation heat transfer from the tunnel to the

package is nearly an order of magnitude greater

than the total heat flux associated with convection

heat transfer from the tunnel environment to the

package. As a consequence, the most severe

conditions for this transient are those that result in

the package receiving the greatest exposure to

radiation heat transfer. This means that a

horizontal orientation will result in the greatest

possible heat input for a given fire scenario. Any

package orientation other than the horizontal

orientation during the fire (e.g., the package

bounced into a vertical orientation as a result of an

accident) would yield less severe heat input to the

package.

package would fail to meet the requirement of

maintaining appropriate shielding in this scenario.

This spent fuel transportation package is expected

to lose its neutron shield material in the fire

accident specified in current regulations, and

therefore the design does not rely on the neutron

shield material remaining intact in order to

maintain shielding. This package is designed to

attenuate neutron radiation to acceptable levels

(see 10 CFR 71.51 [1]) following an accident

without the assistance of the neutron shield

material. However, the loss of the shield material

means that the neutron shield's heat transfer

capability would be expected to deteriorate rapidly

during the fire.

I1o _ I to,") | s I I II-ta id 1400 - o Shot .(OA)

1300 _ S__p0_R1-.

1200 1....L =11.J. I _d I I _ 

{-.I L\, ,Xb.Ie l40)t( (P,. 

1000 - -- - - - I- . --

boo 7 ' -1 - -1- -F - -- Ed of NIST2Ta TMo, 4

200

400 j

400 J _ J _ J

100 I * I I I I 

200 - 1 PoUDuck Pa oa (FSTDOU

* 2 4 5 5 10 12 14 If IS 20 22 24 26 20 20

Elapsed Tim. (hous)

Figure 7.3. Maximum Temperature Histories

for TN-68 Package Components

During First 30 hr of Transient,

7.1.2 TN-68 Short-Term Post-Fire In the COBRA-SFS evaluation, it was assumed

Response that the borated polyester remains in place and is

unaffected during the fire, but instantly degrades at

Figure 7.3 shows the temperature response of the the end of the fire and is replaced by hot air. This

package during the first 30 hours of the COBRA-: maximizes the heat input into the package'during

SFS simulation of the transient. (This time period, the fire, then imposes an additional barrier to heat

represents the total duration of the NIST analysis transfer from the package after the fire. From the

that is the source of the boundary conditions for standpoint of the thermal response of the system,

this calculation, in which the package is positioned this is a conservative representation of the effect of

66 ft (20 m) from the fire center.) During the fire, the fire on the neutron shield. The thermal

the material in the neutron shield reaches tempera- conductivity of the borated polyester is about 140

tures that would heavily degrade the borated poly- times that of air, so extending the residence time

ester. This does not mean, however, that the ' - of the polyester to the end of the fire results in a

7.3
calculation that overestimates the rate of heat flow

into the package during the fire. This will tend to

result in higher calculated temperatures on the

package internals than would be obtained if it were

assumed that the polyester was replaced with air

earlier in the transient. In reality, the change

would be more gradual and would occur earlier in

the transient as the neutron shield burned away.

The heat absorbed in the process of melting the

polyester material is not subtracted from the heat

of the fire, as an associated conservatism.

The temperatures in Figure 7.3 show that once the

fire is over, the peak temperatures on the package

shell and neutron shield are predicted to begin to

drop precipitately. This is primarily a response to

the rapid decrease of the boundary temperatures,

as can be seen in Figure 7.4. This plot shows the

outer shell surface temperature predicted with

COBRA-SFS compared to the tunnel ceiling

temperature and the temperature of the air above

the package, which are derived from the NIST

calculations and used as boundary conditions.

outer shell because the internal components must

also absorb the thermal load from the fuel. For the

same reason, the peak temperature of the basket

shows a continuous increase even after the end of

the fire, as does the peak clad temperature.

The plot of the predicted peak clad temperature in

Figure 7.3 shows that the thermal output of the fire

itself does not have much of an effect on the fuel.

The observed rise in peak clad temperature is

mainly a response to the effect of the external

boundary conditions on the rate of heat transfer

from the package. The heat of the fire does not

result in much of an increase in the package

internal temperatures, but the increase in the

external air temperature severely compromises the

rate of heat rejection from the package and

continues to do so long after the fire is out. This is

illustrated very clearly by the plot of the global

peak clad temperature alone, shown in Figure 7.5

for 50 hours of the transient (i.e., the 30 hours of

the NIST transient, plus an additional 20 hours of

the extended cool down beyond the NIST

calculation.)

As shown by the plot in Figure 7.5, at about 4.5

hours into the fire the peak clad temperature

begins an almost adiabatic heat up (approximately

770F/hr (430C/hr)) because the fire is preventing

normal heat removal from the package (which

occurs by thermal radiation to the external

environment and natural convection at the

surface.) This adiabatic heat-up continues for

about an hour after the end of the fire, until the

package shell temperature drops low enough to

permit some heat removal from the package by

radiation to the tunnel surfaces.

The fuel cladding temperature increase observed

shortly before the end of the fire (at -4.5 hours)

occurs in fuel in the outer periphery of the basket,

because the fire is heating the outer cylinder of the

package. This causes the peak fuel cladding

temperature to shift first to the bottom assembly in

1100

1000

1700

1600

1500

1400

L 1300

1 1200

F 1100

go1000

. 90

E 600 - 700

500

S0oo 400

300

200

100

' 1 -CeilpnA Temp

/ dof Fl,.

D FO- DOnratn Posi-F-r DO -no, (NIST Ds..)

0 2 4 6 a 10 12 14 16 10 20 22 24 26 28 30

Elapsed Time (h-urs}

Figure 7.4. Maximum TN-68 Package Surface

Temperatures Compared to NIST

Boundary Condition Temperatures

Figure 7.3 also shows that the peak temperatures

on the gamma shields and support rails of the

package decrease after the fire ends. However, the

temperature decrease for these components is

much slower than for the neutron shield region and

7.4
the horizontal basket cross-section, then to the top

assembly. The rate of increase in the predicted

peak cladding temperature slows briefly to 1-

2°F/hr (0.5-10C/hr) before increasing again as the

peak temperature location shifts back to the center

assembly in the core of the basket. (The peak fuel

clad temperature location is automatically tracked

in these results during the transient.)

Io -----.-----

70 | I I I I Me

Ia I I I I I I I I I

c70D _ - -x -l__ - - r- - T. - - -1- - - T s

o l - - - - - - - - l

E a ' a1 a l a l

I 50-I s -- az ... I ... -o -f - ... a -a- -1-I--- - - - - -1 --

-.1 '; ; ; ; ; ; ; i 400 - L400 .. . . -

40 , I I 1. 1 I I I

0 * t0 Is 20 25 30 35 40 45 s0

Elapsed Tim. 1hours) .:

; I -

Figure 7.5. Peak Fuel Cladding Temperature

History in TN-68 During First 50

hr of Transient

The peak clad temperature continues to increase

after the fire, because thermal radiation from the

tunnel ceiling, walls and floor is continuing to add

heat to the package and the external ambient air

temperature slows the rate of heat removal by ;

convection. However, the rate of increase drops to

only about 20F/hr (-1 0C/hr). At about 15 hours

into the transient, the rate of increase of the peak

clad temperature begins to climb again, to a rate of

about IO0 F/hr (60C/hr), then again drops to only .

30F/hr (-2 0C/hr) after about 20 hours. This ,

behavior is due to the decrease in the rate of heat

removal via thermal radiation as the tunnel

surfaces cool down, and the more gradual decrease

in the rate at which heat is being removed from the

package by forced convection from the hot air

flowing past the package.

By the end of the NIST transient at 30 hours, the

rate of increase of the peak clad temperature has

dropped to less than 30F/hr (<20C/hr). The global

peak cladding temperature of 8450F (4521C) is

finally reached at -40 hours into the transient.

(The dashed portion of the curve in Figure 7.5

denotes results that stem from the boundary

conditions extrapolated beyond the results

obtained in the 30-hour NIST calculation.)

The maximum temperature history of the seals in

the package closure and vacuum port is shown in

Figure 7.6. The curve in this figure represents the

global peak of all seal material used in the TN-68.

As shown in this figure, the Helicoflex® seal

material is predicted to reach a maximum

temperature of 811 0F (4330C) right at the end of

the fire, then gradually begins to cool as the

transient proceeds into the post-fire duration. This

peak temperature exceeds the maximum operating

temperature of 5360F (2801C) for this material.

Bolts and other subcomponents were not explicitly

represented at the package ends in the COBRASFS model of the TN-68. However, the depicted

temperature history (see Figure 7.6) conservatively

represents the peak temperature history of the

closure bolts due to the manner in which heat must

migrate around the top impact limiter'into the

package upper forging, through the closure seal

location, and then into the closure. This is due to

the limited conduction offered by the steelencapsulated wooden impact-damping material.

The thermal response of the package after the end

of the fire is further illustrated in Figure 7.7, with

plots showing radial temperature profiles through

the package at selected time intervals through the

transient. These profiles show that the outer shell

and'former'neutron shield cool rapidly once the

fire is over, while the temperatures of the internal

nodes representing the basket tubes and poison

plates continue to rise in response to thee heat load

from'the spent nuclear fuel.

7.5
900

800

700

a* 000

';500

E 400

300

200

100

a , | | j , Gr m Pl<V~bureSeal (Peak) |

L ~ ~ l_ ~ d I i.

/ | I ' * I | F i -

- - *-a- - - - - - - -- -1 -a - - - - -I - -1 -

Fire 4 |Poa2 -Fi,4 Boraon (NIST 2a2a) 1 0 2 4 6 0 10 12 14 16 10 20 22 24 26 28 30

Elapsed Time (hours)

Figure 7.6. Maximum Global TN-68 Closure/

Port Seal Temperature History

During First 30 hr of Transient

180C

1 700

-- 30 te (End 0f NIST Tl--s)

I5 1 es

12 los

_10 to

-7 hls (EMd of FRie)

-6245 los (Peas Fire Te-pO)

of the tunnel environment. (The extrapolated

values are presented in Figures 6.5 and 6.6 for the

air and wall temperatures, respectively.)

To explore the effects of prolonged exposure to

post-fire conditions in the tunnel, the calculations

were carried out for the full 300 hours (273 hours

after fire cessation). This is equivalent to

assuming that the package will be left in the tunnel

for an extended period (days or weeks rather than

merely hours) without any emergency responder

intervention. This assumption is not very realistic

but is highly conservative in that it defines a

relatively severe long-term ambient environment.

During the first 30 hours of the transient, the heat

transfer at the package surface was assumed to be

forced convection at the velocities predicted in the

tunnel by the NIST calculation. The basis for this

approach is discussed in Section 6.0. This is a

relatively conservative assumption, particularly for

conditions after about 20 hours into the transient,

when the velocities predicted in the NIST

calculation have dropped to values of I to 2 ft/s

(0.3 to 0.6 mis) or less. For the latter portion of

the transient (t > 20 hours), heat transfer at the

package surface is a complex mixture of forced

convection (due to air flow induced in the tunnel

by the temperature gradients of the fire) and free

convection (driven by the non-uniform

circumferential temperatures of the package outer

shell.) This was conservatively approximated by

imposing a purely forced convection heat transfer

coefficient (based on the NIST air velocities and

temperatures) for the first 30 hours of the

simulation, then imposing a free convection

coefficient for the remainder of the calculation.

Figure 7.8 shows the temperature response of the

various components of the package for the long

term transient calculation to 300 hours. (As

previously indicated, the dashed portion of the

curve is used to distinguish the results that stem

from the boundary conditions that were

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 10 19 20 21 22 23 24 25 26 27

Solid Node (top to bottom)

Figure 7.7. Temperature Profiles Top-ToBottom Through TN-68 Package

Axis: First 30 hr of Transient

7.1.3 TN-68 Long-Term Post-Fire

Response

The NIST calculation used to define the boundary

conditions for the COBRA-SFS analysis simulated

the fire transient and its aftermath out to 30 hours.

However, the trends exhibited by the temperatures

of the various components of the package at the

end of the transient indicate that the system is not

yet at a new steady state by then. Temperatures

predicted by NIST for the first 30 hours were

extrapolated from 30 hours out to 300 hours using

a power function to realistically model cool-down

7.6
extrapolated from the NIST simulation.) As

shown in Figure 7.8, the highest peak clad

temperature, 8450F (4520C), is reached at

approximately 40 hours. The peak temperature for

the basket structure is also reached at about the

same time. The predicted maximum in the peak

clad temperature is below the regulatory limit of

10580F (570'C) by a difference of 2131F (1 100C).

All other temperatures in the package have been

decreasing steadily since the end of the fire.

1500- lif -- l_ -GFooaIsbldi(ltFAi). -- I-m . . . . - O. .

sso. 1100 I- 00 '"-I' ..., - 1 - -J - -B~.-- 

- -- - Go5@tol~l -P.6k) 1- -

I NO-AM-;W~-5Wg

1400 - - 4 ------ B dl

200M - - _I_ L J Li --nptU gNak)

^ 1 10 T ~Ertd9NIST20.T-i.

w ,---- .- -7 r ---- 7--r--

___1 _- NITDo 1_ _ - _ - -I- -- 4- v -- 1 - 1-

transient is essentially an accommodation to the

new higher-temperature boundary conditions

extrapolated from the conditions predicted at the

end of 30 hours in the NIST fire analysis. The

temperature of the fuel and basket is largely

unaffected by the heat input to the package from

the fire; the increase in peak clad temperature and

peak basket temperature is due almost entirely to

having no heat removal from the package during

the fire and for about an hour immediately

afterwards. After the ambient temperatures drop

enough to allow heat removal from the package,

the rate of increase of the peak clad temperature

begins to level off and then finally turn around

about 40 hours into the transient.

Viewed on the scale of 300 hours, the fire portion

of the transient appears as a relatively short-lived

spike in the boundary conditions that significantly

affects only the outer shell, neutron shield, and

impact limiters, and to a lesser extent the outer and

inner gamma shielding. These components show a

rapid temperature increase during the fire, but after

the end of the fire immediately begin to rapidly

cool down.

Peak component temperatures for the TN-68

package over the transient fire simulation are

summarized in Table 7.1.

2 50 Iy- Empdald 0*. W5.1.

a 2s so5 75 10 1251 501 75 200 225 2S0 275 30

Elapsed Tkne (hours)

Figure 7.8. Maximum Temperature Histories

for TN-68 Package Components

During 300 hr of Transient

By 100 hours, the peak clad temperature has

dropped to 7840 F (418'C), with temperatures

decreasing at rates of about -1 F/hr (-0.61C/hr).

After 200 hours, the peak clad temperature has

dropped to 719'F (3820C), and at 300 hours is

predicted to be down to 690'F (3661C) for the

specified boundary conditions. At this point in the

transient calculation, the rate of change of local

temperatures in the system is about -0.20F/hr

(-0.1 0 C/hr). The rate of cooling is very slow due

to the huge thermal mass of the package and its

fuel load. Projections of the cooling rate indicate,1

that it would take an additional 175-200 hours to

reach a new post-fire steady state.

The trends in Figure 7.8 show that the overall

thermal response of the package to the fire,,

Table 7.1. TN-68 Peak Component

Temperatures During Fire

Transient

. - -Maximum

; - - Temperature

(COBRA-SFS) Time

; Component - F(C)' - -(hours!

Fuel Cladding 845 (452) 40

Basket Plate 836 (447) 40

Basket Rail 801 (427) 8.3

Inner Shell 857 (458) 7.0

Gamma Shell 886 (474) 7.0

Package Bottom '762 (406) 7.0

Seals 811 (433) 7.0

Neutron Shield 1355 (735) 6.9

Outer Shell 1789 (976) 6.3

7.7
7.2 Holtec HI-STAR 100 Fire

Transient Results

The ANSYS model of the HI-STAR 100 package

consists of a total of 149,100 standard computational elements and 288 superelements that are

solved for each time step. Similar to the COBRASFS model of the TN-68, this model yields an

overwhelming volume of output that must be

processed to produce a coherent picture of the

package response. The following three

subsections present the peak temperatures versus

time for selected components, as determined with

ANSYS for the HI-STAR 100 subject to the

hypothetical fire transient conditions described in

Section 6.

7.2.1 HI-STAR 100 During the Fire

Figure 7.9 shows the initial temperature response

of the HI-STAR 100 package predicted with

ANSYS during the fire portion of the transient.

The maximum temperature of the HI-STAR 100

package surface increases rapidly to a peak

temperature of 1831 0F (9990C) around 6 hours

into the fire. The maximum temperature of the

inner shell material, which defines the primary

containment boundary, also shows a relatively

rapid increase, reaching a peak of 14440 F (7840C)

approximately 6.75 hours into the fire. This

corresponds to the peak boundary condition

temperatures defined by the fire. The fire

temperatures predicted in the NIST analysis peak

at 6.75 hours, and then drop off rapidly thereafter

as the fire bums itself out. The peak temperature

of the inner shell material is predicted to lead the

gamma shield material peak temperature because

the elements selected to define the primary

containment boundary include the bottom and top

forgings and lid. A large section of the top forging

is directly exposed to the fire (un-shrouded by the

gamma shield, neutron shield/fin section and

upper impact limiter.)

3 4

EI~p.ed Ti-~ (h-~)

Figure 7.9. Maximum Temperature Histories

for HI-STAR 100 Package

Components During Fire Transient

Similar to the TN-68 results with the COBRA-SFS

model, the internal components of this package

also show a very slow thermal response during the

fire. The gamma shield takes nearly an hour to

show any noticeable increase in temperature. Two

hours elapse before the internal canister shell

temperature rises as much as I 0F above the initial

steady-state peak temperature of 4940F (257 0C).

The peak temperatures of the basket structure,

poison plates, and fuel rise only by about 20F

(1°C) in the first five hours of the fire.

This is approximately an hour later than the TN-68

response and can be attributed to the additional

thermal barrier represented by the canister in this

particular design. During this time period, the

peak temperature on the outer skin surrounding the

neutron shield of this package is predicted to go up

to 1 83 1 'F (9990 C), and the peak temperature on

the gamma shields increases to 1400'F (760'C).

By the end of the fire, marking the point at which

all volatile flammables are consumed, the peak

clad temperature has risen to only 813SF (4340C).

The outer shell of the package is predicted to have

a peak temperature of 14490F (7870C) at the end

of the fire, with the outer gamma shield at 1322cF

(717 0C). This is a bit cooler than the TN-68 at this

point in time. The difference is due mainly to the

7.8
larger thermal resistance to radial heat flow in the

thinner skin of the HI-STAR 100, compared to the

TN-68. However, both packages tend to perform

similarly overall.

As with the TN-68 results, the large difference in

the predicted rate of increase in temperature for

the internal and external components of the HI-:

STAR 100 is because the neutron shield and

gamma shield insulate the basket and fuel

assemblies from the fire. The slow response is due

mainly to the huge thermal inertia of the package

components themselves. Even under the severe

heat load imposed by the sustained high

temperatures of a fire lasting nearly 7 hours, it

takes time to raise the temperature of such a large

mass of material despite its internal heat

generation component.

7.2.2 HI-STAR 100 Short-Term Post-Fire

Response

Figure 7.10 shows the temperature response of the

package during the first 30 hours of the ANSYS

transient simulation. During the fire, the material

in the neutron shield is predicted to achieve

temperatures that will heavily degrade it. As

noted in Section 7.1.1 for the predicted loss of the

TN-68 cask's neutron shield, the HI-STAR 100 is

also designed to attenuate neutron radiation to

acceptable levels (see 10 CFR 71.51 [1]) following

an accident without the assistance of the neutron

shield material. However, the neutron shield's

heat transfer capability is expected to deteriorate

rapidly during the fire. In the ANSYS evaluation,

it was assumed that the neutron shield material

(HOLTITE-A) remains in place and unaffected,

during the fire, but instantly degrades at the end of

the fire, to be replaced by hot air., This maximizes

the heat input into the package.

This is a conservative representation of the effect,

of the fire on the neutron shield from the standpoint of the thermal response of the system. The

thermal conductivity of HOLTITE-A is approximately 16 times that of air, so extending the

residence time to the end of the fire results in the

calculation somewhat overestimating the rate of

heat flow into the package during the fire. This

will result in higher calculated temperatures on the

package internals than would occur if degradation

were accounted for at a more realistic rate during

the fire. As an additional associated conservatism,

the latent heat absorbed in the degradation of the

material, which would tend to decrease the

external heat flux due to the fire, is also neglected.

-F.W Fk)

-0-4. Sh.4

-G-Sh.1d(O")

W Sh.4 (PIV-Y C.14--t 9.1.4.y)

-Bk.0 SAd..f SORAL PW.

r =UNWBSIp)

LILW-1 P*MX)- PM

Ed d F1,.

&4 (04ST 20. U-4 I

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

--------------- PWFn D-b- (ST 0j

10 I 5 ;a 25

El.psod Thm (houm)

Figure 7.10. Maximum Temperature Histories

for HI-STAR 100 Package Components During First 30 hr of

Transient

Figure 7.10 shows the peak temperatures predicted

in the ANSYS analysis for the various components

of the package during the full 30 hours of the

transient as defined by the NIST calculations.

These results show that once the fire is over, the

predicted peak temperatures on outboard

components begin to'drop rapidly (i.e., outer shell,

gamma shield, etc.). This is primarily a response

to the rapid decrease of the boundary

temperatures, as can be seen in Figure 7.1 1, which

shows the outer shell surface temperature predicted with ANSYS compared to the tunnel ceiling

temperature and the temperature of the air above

the package derived from the NIST calculations

and used as boundary conditions.

7.9
Figure 7.10 shows that the peak temperatures on

the gamma shields and inner shell of the package

decrease after the end of the fire. However, the

temperature decrease for these components is

much slower than for the outer shell because the

internal components must absorb the thermal load

from the fuel. Similarly, the peak temperature of

the basket shows a continuous increase even after

the end of the fire, as does the peak clad

temperature.

1900

1800 - rOtr Shell

1700 -opAirTemp

1500 1---: -| L . t _cll~ng e' Temp

I 'Ed of fire

_ 1300

. 120D 0I

c 1100 ' I t i I : . I

7 00D - - - - \ -- I- a -

600

5 00

4000

500 500 - I I

very nearly over before the peak clad temperatures

show a discernable increase.

Figure 7.12 shows that a little after 5 hours into

the fire, the global peak clad temperature begins

an almost adiabatic heat up (approximately

690F/hr (380C/hr)) because the fire prevents

normal heat removal from the package by natural

convection at the surface. This adiabatic heat-up

continues for about an hour after the end of the

fire, until the package shell temperature drops low

enough to permit some heat removal from the

package by radiation to the tunnel surfaces.

950 T

9000

850-

. 8000

X 750

c700

050

600

lF~F

-I - --- l- f NI 20m - T r , l

I I I I I I I ...I I i

i I I < l I I-_

I | 1Ful(Pjk

SlI OT| T - End1 of NIST 2lm raSe l ;

200

100 Fop D.o p. t . _ _ PoPIl-Fm Duration (lST D00)

.. . . _ , ..................... .. 3 0 2 4 C 6 10 12 14 16 18 20 22 24 26 26 30

Elapsed Time (ho-m)

Figure 7.11. Maximum HI-STAR 100 Package

Surface Temperature Compared to

NIST Boundary Condition

Temperatures

The plot of the predicted peak clad temperature in

Figure 7.10 shows that the thermal output of the

fire itself has little effect on the fuel or its

accommodating basket. As discussed above for

the TN-68, the observed rise in peak clad

temperature is mainly a response to the effects of

the external boundary conditions on the rate of

heat transfer from the package. The heat of the

fire does not result in much of an increase in the

package internal temperatures, but the increase in

the external air temperature severely compromises

the rate of heat rejection from the package, and

continues to do so long after the fire is out. This is

illustrated very clearly by the plot of the peak clad

temperature alone shown in Figure 7.12 for the

first 50 hours of the NIST transient. The fire is

'l a u . . . . . . . .

0 5 10 15 20 25 30 35 40 45 50

Elapsed Times (ho-s)

Figure 7.12. Peak Fuel Cladding Temperature

History in HI-STAR 100 During

First 50 hr of Transient

The initial fuel cladding temperature rise shown to

initiate shortly before the end of the fire (-5.5

hours) occurs on fuel in the outer periphery of the

basket in a portion of the package facing the top of

the tunnel. The fuel in this region is initially rising

in temperature faster than that residing in the

center of the basket. This continues until -S hours

into the transient. However, by this time the fire

has ended (at -7 hours) and the internal heat

begins spread as component temperatures

redistribute radially throughout the package,

causing the peak fuel cladding temperature to shift

from one assembly to another.

7.10
- - -

The peak cladding temperature begins to drop for

a brief period, before rising again as fuel in the

core of the basket begins to heat up and exceed the

temperature of the fuel on the outer periphery.

(Just as with the TN-68 results, the peak fuel clad

temperature is captured in the global summary as

it moves from assembly to assembly within the

fuel basket during the transient.) The peak clad

temperature continues to increase, because the hot:

air flow and hot tunnel surfaces resulting from the,'

fire are continuing to compromise heat rejection l

from the package surface.,

The package is designed to reject heat to ambient

at 1000F (380C), but the air within the tunnel

environment is still above 200'F (931C) at 30

hours, decreasing from a peak of 15570F (8470C)

at the end of the fire. However, by the end of 30

hours, the rate of increase in the peak clad

temperature has dropped to only about 20F/hour

(10C/hour), in response to the decreasing boundary

temperatures. The global peak cladding

temperature reaches a maximum of 8870F (4750C)

at approximately 32 hours into the transient.

The maximum temperature history of the seals in

the package lid closure, ports, and port covers is

shown in Figure 7.13. The curve in this figure

represents the global peak of all seal material

utilized in the HI-STAR 100. These temperatures

are gathered by querying nodes at the seals'

locations, even though the seals were not explicitly

represented in the model. As shown in this figure,

the metallic mechanical seal material reaches a

maximum temperature of 11770F (6360C) right at

the end of the fire, then gradually begins to cool as'-

the transient proceeds into the post-fire cool down. '

Despite an abrupt rise in temperature during the

fire, the peak seal temperature remains below the t

lowest reported maximum continuous-use seal

temperature limit of l2000F (649 0C) (see Table

4.1.1 of the HI-STAR 100 SAR [10].)

Bolts were not explicitly represented at the

package lid and buttress interface in the ANSYS

model of the HI-STAR 100. However, the

depicted seal temperature history conservatively

represents the peak temperature history of the

closure bolts due to the manner in which heat has'

to migrate around the top impact limiter, into the

package upper forging (between the top limiter

and neutron shield/fin section), through the closure

seal location, and then into the'closure. This is

due to the limited conduction offered by the

stainless steel-encapsulated cellular honeycomb

material.

1100.

100'

9~ 00.~

2.600

E

400'~

200'~

I - - -- - -EM MI- - - - ...-

200 I _

. P r-ood I Prw. isb (NAST D4U)

w- . - - - - - . - - - - -I -. --. .

0 S 10 is

E~apd 'Tim* (beout)

20 25

Figure 7.13.Maximum Global HI-STAR 100

Closure/Port Seal Temperature

History During First 30 hr of

Transient

7.2.3 HI-STAR 100 Lon'g-Term Post-Fire

Response

The trends exhibited by the temperatures of the

various components of the HI-STAR 100 package

at the end of the 30 hour transient indicate that the

system is'not yet at a new steady state. Boundary

temperatures predicted by NIST were extrapolated

from 30 hours out to 300 hours using a power

function to realistically model cool down of the

tunnel environmeent. (The extrapolated values are

presented in Figures 6.5 and 6.6 for the air

temperatures and wall temperatures, respectively.)

7.11
To explore the effects of prolonged exposure to

post-fire conditions in the tunnel, the calculations

were carried out for the full 300 hours (273 hours

after fire cessation). As discussed previously, this

is equivalent to assuming that the package will be

left in the tunnel for up to 12.5 days without any

emergency response. This assumption is not very

realistic, but is highly conservative in that it

defines a relatively severe long-term ambient

environment around the package.

The same relatively conservative assumptions

applied to the TN-68 evaluation for external

convection during the fire and post-fire duration

were applied to the HI-STAR 100 evaluation. A

purely forced convection heat transfer regime and

associated heat transfer coefficient was assumed

for the first 30 hours of the simulation, then a

purely free convection regime and associated

coefficient was assumed for the remainder of the

calculation (t > 30 hours). (Refer to Section 6 for

detailed discussion of the heat transfer boundary

conditions on the package surfaces.)

Figure 7.14 shows the temperature response of the

various components of the package for the long

term transient calculation to 300 hours. As

previously discussed, the dashed portion of the

curve is used to distinguish the results that stem

from the boundary conditions that were

extrapolated from the original NIST calculation.

As noted in Section 7.2.2 (see Figure 7.12), the

highest peak clad temperature is reached at

approximately 32 hours, with a value of 8870 F

(4750C). This is 171'F (950C) below the

regulatory limit. The peak temperature for the

basket/poison plate structure is reached at about

the same time.

All other temperatures in the package have been

decreasing steadily since the end of the fire. By

100 hours, the peak clad temperature has dropped

to 7570F (403'C). Similar to the results for the

TN-68, this system is not yet at a new post-fire

steady-state by this time (see Section 7.1.3, Figure

7.8). The HI-STAR 100 is nearing its new postfire steady-state at about 200 hours, with rates of

temperature change on the order of approximately

-0.3'Flhr (-0.20C/hr). After 250 hours, the peak

clad temperature has dropped to 681'F (361'C),

and at 300 hours, it is predicted to be 6740F

(3570 C) for the specified boundary conditions. At

this point in the transient calculation, the rate of

change of local temperatures in the system is less

than -0.1 F/hr (-0.060 C/hr), and the conditions can

be treated as being essentially at a new post-fire

steady state.

*900

1700-

1500-

_ 1300

,1100.

& 900*

- 700

500

300

$ ; [ | ! ; - Fuel (Peak)

- i -Out,, Shell

l | . | . GeuueG Sheld (Ouler)

A I Ilmnee Shell (Pntay Cel-Dieent B-uuduey) | ,.[ |..Cd e Ca SW

: : | ;-hahilCShuuhBORAL Plae,

-- -LmdeRBase S-ujchl, (Peak)

| X - * -Li/dN-l PuhDraie hPu

I - L -- - - End of Fu l

j Ib - -End of NIST 20m TIaueeil

N.- -, __- -S n . I-|-i ! I I

I_,NIST Data Set = - _ .i I -~ '

: OF - '1- - -- _____21 -I-I &utrdputated Data Set I1

0 25 50 75 100 125 150 175 200 225 250

Elapsed Time (hours)

275 300

Figure 7.14. Maximum Temperature Histories

for HI-STAR 100 Package Components During 300 hrs of the

Transient

The trends in Figure 7.14 show that the overall

thermal response of the package to the fire

transient is essentially an accommodation to the

new higher temperature boundary conditions

represented by the conditions predicted at the end

of 30 hours in the NIST fire analysis. Viewed on

the scale of 300 hours (i.e., from pre-fire to postfire steady state), the fire portion of the transient

appears as a relatively short-lived spike in the

boundary conditions that significantly affects only

the outer shell, impact limiters, and the neutron

shield, and to a lesser extent, the gamma shield,

inner shell, and canister. The outer shell and

7.12
neutron shield show a rapid temperature increase

during the fire, but after the end of the fire

immediately begin to rapidly cool down.

The temperature of the fuel and basket is largely

unaffected by the heat input to the package from

the fire.' The increase in peak clad temperature

and peak basket temperature is due almost

exclusively to having no heat removal from the

package during the fire and for about an hour

immediately afterward. After the ambient

temperatures drop enough to allow heat removal

from the package, the rate of increase of the peak

clad temperature begins to level off, and then

finally turns around at about 32 hours into the

transient.

The TN-68 system displays a rapid peak cladding

temperature increase during the interval from

about 6 to 8 hours of the transient, followed by a

much slower rate of increase until about 16 hours,'

at which point it begins to rise again toward its

final peak value, reached at approxirmately 40

hours. The peak cladding temperature predicted

for the HI-STAR 100 follows a similar pattern, but

with a somewhat more dynamic response. After,

reaching a peak at approximately 8 hours, the peak

cladding temperature actually decreases fo'r a time,,

until about 16 hours, at which point it begins to

rise toward its 'final peak value, reached at about

32 hours.

The difference in respons6 of the peak clad

temperature in the two packages is due to three

main factors. There are significant differences in'

construction and thickness of the finned neutron

shield regions in the two package designs. There

is about a 15% difference in the thermal inertia

associated with the spent fuel assemblies in each

package (the HI-STAR 100 contains 24 PWR fuel

assemblies, compared to 68 BWR fuel assemblies

within the TN-68 package), and the two packages

have very different basket designs. In addition,

the high thermal conductivity of the HI-STAR 100

aluminum honeycomb impact limiters aids in

ramping up component temperatures faster in the

ends of the package, compared to the effect of the

redwood impact limiters on the TN-68 package.

The TN-68 does not utilize an internal canister to

hold spent fuel. It relies instead on seals to

prevent radioactive releases from the fuel

compartment. The maximum predicted seal

temperature, which is seen by the package lid seal,

is 81 1F (4330C), and occurs at the end of the fire.

This is below the peak seal temperature predicted

for the HI-STAR 100 and is primarily due to the

relatively low conductivity of the redwood

material used in the TN-68 impact limiter design,

compared to the aluminum honeycomb in the HISTAR 100 impact limiter design.

When comparing the heating trends associated

with the HI-STAR 100 and the TN-68 (compare

results shown in Figure 7.14 and in Figure 7.8) , it

appears that the HI-STAR 100 generally heats up

faster during the fire than the TN-68. However,

this is mainly an artifact of the differences

between the initial steady-state conditions in the

two packages, different exterior packaging, and

differences in their respective fuel loading. The

HI-STAR 100 components enter the fire transient

anywhere from 1000F to 200'F (560C to 11 1C)

hotter than the corresponding components of the

TN-68. These temperature differences are due to

the redundant encapsulation provided by the MPC

canister in the HI-STAR system, the number of

fuel assemblies that the decay heat is distributed

over (24 for the Hl-STAR 100 versus 68 for the

TN-68), and the level of shrouding of the package

surface by the support device. HI-STAR 100 is

heavily shrouded by its support cradle; the TN-68

outer surface is essentially bare to ambient

conditions. ' '

In addition to these essentially incidental

differences, there are some small differences in

design that affect the rate of heat-up of the outer

7.13
shells of the two packages. The TN-68 has a 50%

thicker solid outer skin which distributes the heat

from the fire transient circumferentially to cooler

regions of the package more effectively than the

thinner outer skin of the HI-STAR 100 package.

The outer skin of the HI-STAR 100 package

consists of relatively narrow welded metal strips,

rather than a single steel sheet. The 0. 19-inch

(0.48 cm) fillet welds joining the metal strips

(which were explicitly accounted for in the

ANSYS model) present an additional barrier to

circumferential heat flow in the HI-STAR 100

package outer shell. However, because both

packages present a very large thermal mass to the

fire and have very similar overall designs, they

respond in essentially the same manner to the fire

transient. The differences shown in these two sets

of results consist mainly of minor time-shifts in

the response to the imposed boundary conditions,

and in general the behavior of the two sets of

curves make them almost indistinguishable.

Peak component temperatures for the HI-STAR

100 over the transient fire simulation are reported

in Table 7.2.

7.3 NAC LWT Fire Transient

Results

The ANSYS model of the NAC LWT package

consists of a total of 50,673 standard

computational elements and 12 superelements that

are solved for each time step. Similar to the TN68 and HI-STAR 100 models, this model yields a

large amount of output that has been processed to

characterize the package response. The following

three subsections present the peak temperatures

versus time for selected components, as determined with ANSYS for the NAC LWT package

subject to the hypothetical fire transient conditions

described in Section 6.

7.3.1 LWT During the Fire

Figure 7.15 shows the initial temperature response

for the NAC LWT package and ISO container as

predicted with ANSYS during the fire portion of

the transient. Similar to the HI-STAR 100, the

maximum temperature of the exterior surface of

the ISO container surrounding the NAC LWT

package increases rapidly to a peak temperature of

15920F (8670C) around 6 hours into the fire. This

is roughly 200'F (I I 0C) below that of the

HI-STAR 100 external surface peak temperature,

and is due to the substantial view that the hottest

portion (top) of the ISO container has of cooler

surfaces (i.e., the package body, and the sides and

bottom of the ISO container.)

The maximum temperature of the exterior surface

of the package is slightly lower, at 15260F

(8300C). The maximum temperature of the cask

inner shell material, which defines the primary

containment boundary along with the bolted lid,

shows a more gradual increase, reaching a peak of

12720 F (6890C) at about 7 hours into the fire.

Unlike the TN-68 and the HI-STAR 100, the

internal components of the LWT package,

particularly the fuel assembly, exhibit a noticeable

Table 7.2. HOLTEC HI-STAR 100 Peak

Component Temperatures

During Fire Transient

Maximum

Temperature

(ANSYS) Time

Component OF (C) (hours)

Fuel Cladding 887 (475) 32

MPC Basket 878 (470) 9

Boral 876 (469) 9

Canister Shell 1020 (549) 7.3

Inner Shell and 1444 (784) 6.8

Forgings

Gamma Shield 1400 (760) 6.8

Package Skin 1831 (999) 6

Lid/Vent/Drain Port 1177 (636) 6.8

Seals

Impact Limiter Skin 1826 (997 6

Impact Limiter 1591 (866) 6.8

Structure

7.14
thermal response during the fire. The peak fuel

cladding temperature begins to rise at about two

hours elapsed time, and the package structural -

components show a fairly rapid rise in temperature

in the first hour of the fire. This occurs primarily

because this package has considerably less thermal

inertia than the two larger multi-assembly

packages. The additional heat transfer paths

available into the LWT package are also

contributing factors, resulting from the fuel

assembly being exposed within a cavity at each

end of the cask. As the inner shell surrounding the

assembly ends heats up, radiation exchange within

the cavities generates cladding temperatures at the

ends of the fuel rods that are significantly higher

than the temperatures at the center, as illustrated in

Figure 7.16.

1700

1500-

1300-

F 1100

, 700- I500

300-

L - - -

Surface

-Lid SaW

-Drahve'Apo.tsesis

=Cask Body Lead Er.d Lead

tso Ctaiw

End of Fs*

Other components of the package, in contrast to

the peak cladding temperature, reach their peak

temperature values at or very close to the end of

the fire. This behavior closely follows the sudden

decrease in the external thermal load on the

package as the fire bums itself out. This can

readily be seen in the peak temperatures reached

on the gamma shield and the neutron shielding.

Run s8.0

21 200 21.:2220

SI,-7 .rE-cMM

SW . 47.472

Fgr 71239

:670 3

901.151

__99J3 57

154

1174

Figure 7.16. Lumped Fuel Assembly

Temperature Distribution 7 hr into

Transient

The gamma shielding is provided by a lead layer

between the inner and outer shells as well as a lead

section in the welded base (i.e., the end opposite

the package lid). The lead temperature reaches a

maximum of 13780F (7481C) at 6.75 hours

elapsed time. This is considerably greater than the

established safe operating limit of 600'F (316'C)

[IlI] for this material, and could result in reduced

gamma shielding due to melting and slumping of

the lead. However, in this calculation the lead is

assumed to remains intact, to conservatively

maximize heat input during the transient.

Neutron shielding is provided by the tanks of

ethylene glycol solution on the package exterior.

As described in Section 5.3, the temperatures of

the nodes representing the main tank and overflow

0 1 2 3 4

Elapsed 'Tints (hours)

5 6

Figure 7.15.NAC LWT Package Component

Maximum Temperature Histories

During Fire Transient

By the end of the fire (at approximately 7 hours),

the predicted peak fuel cladding temperature in the

end region of the fuel has reached 1010 0

F_(543'C)

and is still rising. This value is approaching (and

eventually will exceed-see Section 7.3.2) the

currently accepted short term temperature limit of

10580F (570'C) for Zircaloy clad spent nuclear

fuel under accident conditions [23].

7.15
tank were monitored for temperatures indicating

rupture and evaporation throughout the transient

solution. Similar to assumptions in the standard

fire analysis included in the SAR [I l], the liquid

in the tank is expected to lose its shielding

capability when the temperature exceeds its 350'F

( l770 C) boiling point. As noted for the TN-68 and

HI-STAR 1 00 casks, the NAC LWT is also

designed to attenuate neutron radiation to

acceptable levels (see 10 CFR 71.51 [I]) following

an accident without the assistance of the neutron

shield material. However, the loss of the neutron

shield affects the rate of heat transfer into and out

of the cask during and after the fire transient.

As a measure of conservatism, tank rupture was

considered to occur only after the minimum

ethylene glycol temperature for each tank

exceeded 350'F (1770C). This assumption

effectively delays rupture to a slightly later point

in the transient than might be expected, thus

maximizing heat input into the package. The

model predicted that the inner neutron shield tank

and the outer expansion tank rupture at-I .5 hours.

Following rupture, the effective conductivity of

the tank was significantly decreased, due to the

ethylene glycol volume being expelled and

replaced with air.

7.3.2 LWT Short-Term Post-Fire

Response

Figure 7.17 shows the peak temperatures predicted

for various components of the package during the

first 30 hours of the ANSYS transient simulation

based on the NIST simulation results. The

cladding peak and average temperatures continue

to rise after the fire, just as in the analyses for the

TN-68 and Hl-STAR 100 packages, and for much

the same reason. The ambient conditions in the

tunnel immediately following the fire severely

retard the rate at which the fuel decay heat can be

removed from the package.

1700

150 - --- - - UasK Juter Sura- / j / _-Impad. Llmt.er / / n] * -Lid Seal 1300 'P0r Seat2

Elapsed Tme (hours

L1100 11 V G d -¢ -E,,d ofF-r

for | - - Ed of NIST Transient 20h 

700

Hoevr once Drthe sfire isaovr, the prDictedpak

300 - l

t r s n o 1a 2c 2(

Elapsed Time (hourf )

Figure 7.17. NAC LWT Package Component

Maximum Temperature Histories

for First 30 hr of Fire Transient

However, once the fire is over, the predicted peak

temperatures on outboard components (i.e., the

ISO container and cask outer surface) begin to

drop rapidly in response to the rapid decrease in

the boundary temperatures (see Figure 7.1l 8). This

figure shows the outer shell surface temperature

predicted with ANSYS, compared to the tunnel

ceiling temperature and the temperature of the air

above the ISO container, derived from the NIST

calculations and used as boundary conditions.

30

1900

1700

1 500

, 300

t- o , 1100

&. 900

E

t700-

500

300

100

/171', -C .itng Te-pe-atu !

Top o, Te-lp-lot-,e [

, ISO co'd-1rr

- - - End ofFie FI 

-- End ot NIST 20o, Transent

P-~t-F-o D-- (-ST D-t.-

to to 20 25 30

Elapsed Time (hour)

Figure 7.18.Maximum ISO Container Surface

Temperature History Compared

with NIST Boundary Condition

Temperatures

7.16
Figure 7.17 shows that the peak temperatures for

all package components begin to decrease shortly

after the end of the fire. The peak cladding

temperature reaches its maximum value of 10991F

(593°C) at 9 hours. This is 410F (230C) above the

short term limit of 10580F (570'C), but is still

2830F (1571C) below the temperature at which

Zircaloy fuel rods actually fail by burst rupture,

which is approximately 13820F (7501C) [25]. , '

(The maximum temperature for the basket reaches

its peak of 10690F (5760C) at about 8 hours into

the transient, but this temperature curve is omitted

from Figure 7.17 for clarity.)

As a result of the low thermal inertia of this

package, peak temperatures in the various

components occur within two hours of the fire

being extinguished, rather than 25 or 33 hours

later, as in the HI-STAR 100 and TN-68,

respectively. Because of the heating of the ends of

the fuel rods due to thermal radiation as a result of

the fire, the average fuel temperature gradually

increases to a maximum of 9880F (53 10C). This

peak is reached at 9 hours elapsed time, as shown

in Figure 7.19.

The maximum temperature histories of the seals in

the drain/vent ports and the lid are shown for the

first 30 hours in Figure 7.20. (The calculated

values were gathered by querying nodes at the

seals' locations, since the seals were not explicitly

represented in the model.) The drain and vent

ports are sealed with Teflon O-rings. The bolted

lid is sealed by both metallic and Teflon O-ring

seals. The drain and vent ports reach a maximum

temperature of 14100F (7660C), and the lid seal

reaches 1350'F (7320C) at the end of the fire.

.I I -Lidse.

swo - - r- - -r- -D -, Po S -eals~g

-- Ed of MST 20m Tans*,A

i1300 - - -/ - - - - - -t - - - - - - - - -

300 ' ---- lLF - - - - -e1 - - - - - F - - - - -

F.. Dun Pcd-Fw Dln Wrabw (IST Data)

5 10 E Is 20 - 25 -30

Elapsed Tbmw (hour)

1200, I3 

IWO

gmi~ 7000

5_ 00

E iSW

SW

400

200

200

; ; I ; -Furlql(Peak)

. 9 |--w Emtrd r)

_____ F

tl ' , - - -- ' -- -- i- ---

-J-r - - --- - -- - - -I- - - - - -r -r ---

I I I I . _ _ _ _- - -_ -_- - -_ --_ _ _ - _ - _ - _ _-- _ __- _ _ _

Figure 7.20. Maximum Seal Temperature

Histories for Drain/Vent Ports and

Package Lid During First 30 hr of

Fire Transient

These materials then gradually cool as the

transient proceeds into the post-fire cool down.

The extreme rise in temperature is due to the low

thermal inertia associated with the LWT package

and the close proximity of the seals to exterior

surfaces subject to thermal radiation from the thin

ISO container.

The ISO container is itself subject to thermal

radiation and convection heat input from the

tunnel environment. If the NAC LWT cask were

not enclosed in an ISO container, the' cask

components would reach even higher temperatures

during the postulated accident scenario. However,

with or without an ISO container, the predicted

F'-- - , I i I Peel-F.n DC m I,. (NI ST Dota)t0 5 IS IS 20 , : 25 , ".S', ' 30

Figure 7.19. Peak and Average Fuel Cladding

Temperature Histories for 'NAC '

LWT Packa'ge Duririg First 30 hr''

of Fire Transient

7.17
seal temperatures are far greater than the

maximum continuous-use seal temperature limits

of 7350F (3910C) for Teflon seals and 800'F

(4270C) for the metallic seals.

7.3.3 LWT Long-Term Post-Fire

Response

As with the TN-68 and Ill-STAR 100 analyses,

the temperatures predicted in the NIST analysis

were extrapolated from 30 hours to 300 hours

using a power function in order to realistically

model cool down of the tunnel environment. (The

extrapolated values are presented in Figures 6.5

and 6.6.)

Peak component temperatures for the LWT over

the transient fire simulation are reported in

Table 7.3. To explore the effects of prolonged

exposure to post-fire conditions in the tunnel, the

calculations for the NAC LWT were carried out

for the full 300 hours (273 hours after the end of

the fire). As discussed previously, this

conservative approach is equivalent to assuming

that the package will be left in the tunnel

indefinitely, without any emergency responder

intervention. The same conservative assumptions

used in the analysis of the TN-68 and HI-STAR

100 were used to define the convection heat

transfer boundary on the NAC LWT package. A

purely forced convection heat transfer regime was

assumed for the first 30 hours of the simulation,

then a purely free convection regime was assumed

for the remainder of the calculation (t > 30 hours).

Figure 7.21 shows the temperature response of the

various components of the package for the long

term transient calculation to 300 hours. The

maximum temperatures were reached within a

short time after the end of the fire, and the LWT at

100 hours is very close to its new steady-state

condition. This behavior is consistent with its

lower thermal inertia, in comparison to the larger

multi-assembly packages.

Table 7.3. NAC LWT Peak Component

Temperatures During Fire

Transient

Maximum

Temperature

(ANSYS) Time

Component 0F (C) (hours)

Fuel Cladding 1099 (593) 9

Aluminum PWR

Insert 1069 (576) 8

Inner Shell 1272 (689) 7

Lead Gamma Shield 1378 (748) 6.75

Outer Shell 1526 (830) 6

Liquid Neutron

Shield 1525 (829) 6

Lid Seal 1350 (732) 6.9

Drain/Vent Ports 1410 (766) 6.75

Impact Limiters 1521 (827) 6

ISO Container 1592 (867) 6

1600

1400

1200

- 1000

,5 800

r 600

400

200

-F.el Claddig (Peak) Cask Inee, Surface -Cask Oule, Surface

-Impat Limiters

-Lid Seas

-0,aiisteet Pert Seats

Cask Body Lead

-- End Lead -ISO Con-amer

- -End of Fire

- - End of NIST 20m Transien

_I

50 100 150

Elapsed Time (hours

200 250 300

Figure 7.21. NAC LWT Package Component

Maximum Temperature Histories

During 300 hr Transient

Temperature distributions within the package for

the final steady state will be slightly different than

the original, due to the dissipation of the liquid

neutron shield, changes in the surface emissivities

because of the fire, and tunnel ambient conditions

that differ from the hot-normal conditions assumed

for the pre-fire steady state (i.e., lower ambient

temperature and the absence of solar insolation.)

7.18
8 POTENTIAL CONSEQUENCES

USNRC Staff evaluated the potential for a

release of radioactive material from each of the

three transportation casks (HI-STAR 100, TN-68

and NAC LWT) analyzed for the Baltimore

tunnel fire scenario. The analysis indicates that

there would be no release expected from the HISTAR 100 cask. However, the possibility of a

small release cannot be entirely ruled out for

either the TN-68 or NAC LWT casks, because

cask temperatures during the fire or cool-down

period exceed the long-term service temperature

limits for the cask lid seals.

Staff performed an analysis to determine the

magnitude of any potential release. Based on

that analysis (which is described below), it was

determined that any potential release from either

the TN-68 or NAC LWT cask would be small-.

less than an A2 quantity.6 The potential release

would not involve a release of spent fuel or

fission products, but could possibly result from

CRUD spalling off the fuel rods.

8.1 Results for the HI-STAR 100

Cask

The thermal analysis shows that the HI-STAR

100 cask design would maintain three important

barriers throughout the fire and subsequent cool

down period, which would prevent the release of

radioactive materials: The welded inner canister

remains intact and leak tight, preventing any

release from the fuel rods themselves or as a

result of CRUD spalling off the fuel rods. The

temperature of the fuel cladding would peak at

about 990'F (5320C), well below the short-term

temperature limit of 10580F (570'C) for

Zircaloy cladding and significantly below its

projected burst temperature of 1382WF (750'C).

This would prevent the release of fission

products from the fuel rods. The maximum

temperature of I 177F (6360C) predicted for the

cask's metallic 0-rings is below their rated

continuous-use service temperature of 1200'F

(6490C). Thus, the 0-rings would not be

expected to significantly degrade.

8.2 Results for the TN-68 Cask

The thermal analysis for the TN-68 cask shows

that during the Baltimore tunnel fire scenario,

this cask design would maintain the integrity of

the fuel cladding, which is the single most

important barrier to prevent the release of

radioactive materials. At approximately 40

hours elapsed time, the temperature of the fuel

cladding would peak at about 8450F (4520C),

well below the short-term temperature limit of

10581F (570'C) for Zircaloy cladding and

significantly below its projected burst

temperature of 13821F (7501C). This would

prevent the release of fission products from the

fuel rods. However, the metallic helicoflex seals

used on the TN-68 lid and vent and drain ports

reach a maximum temperature of 811 0F (4330C)

by the end of the fire (at 7 hours elapsed time.)

This exceeds the seals' rated service temperature

of 5361F (2801C) by 2751F (1531C).

6 An A2 quantity represents the threshold below

which an accident resistant package is not required.

The acceptance requirement for Type B packages is

that they release less than an A2 quantity/week after

being subjected to the hypothetical accident

conditions in 10 CFR Part 71 [1].

Exceeding the service temperature of the seals

on the 1N-68 cask lid or vent and drain ports

means that there is a potential for a release to

occur. Potential releases would be limited,

however, by the narrow convoluted flow paths

of the drain and vent ports and by the tight

8.1
clearances of the close metal-to-metal contact

between the lid and cask body. The close

contact is maintained by the pre-load created by

the initial torque on the lid bolts. Because the

fuel cladding remains intact, it is not expected

that any radioactive material would be released

from inside the fuel rods. This limits any release

from the cask to CRUD particles that may flake

off or spall from individual fuel rods.

The amount of releasable CRUD in the

TN-68 cask was estimated using data developed

by Sandia National Laboratory for analysis of

CRUD contribution to shipping cask

containment requirements [26], based on cask

contents consisting of 68 BWR fuel assemblies,

each assembly containing 49 fuel rods. An

estimate of the maximum "spot" CRUD activity

shows that for 90% of BWR spent fuel rods the

maximum activity is 300ttCi/cm2 or less [26,

Table 1-17]. The ratio of the peak to average

concentration on the rod surface (i.e., the

maximum "spot" CRUD activity over the

average value) varies by a factor of two for

BWR fuel rods [26, Table I-17].

The CRUD activity estimates [26] are based on

newly discharged spent nuclear fuel. The CRUD

activity is expected to decay by a factor of onehalf for five-year cooled fuel, based on the decay

rate for Co60. This proves to be a good

approximation because 98% of the activity for

five-year cooled BWR fuel comes from Co60.

Based on this data, the average CRUD activity

for a BWR rod with a surface area of 1600 cm2

is about 0.12 Ci for five-year cooled fuel. The

average CRUD activity for a typical 7 x 7 BWR

assembly is about 5.9 Ci.

The amount of CRUD that might flake or spall

from the surface of a BWR rod due to thermal

stresses induced by temperature change in the

fuel rods is estimated to be a maximum of 15%

[26, Table I-10]. The major driving force for

material release results from the increased gas

pressure inside the cask due to increases in

internal temperature. The temperature change in

the cask is bounded by the difference between

the maximum gas temperature predicted during

the fire transient and the gas temperature at the

time the cask is loaded. For this analysis, the

loading temperature is defined as 1000 F (38 0C),

based on the temperature reported in the SAR

[9]. The maximum gas temperature is assumed

to be the maximum peak clad temperature

predicted during the transient. This yields a

conservative estimate of the temperature change.

A deposition factor of 0.90 was used to account

for the settling and deposition of CRUD

particles on cask surfaces and fuel assemblies.

The deposition factor was developed as part of

NRC'ssecurity assessments for spent nuclear fuel

transport and storage casks, and is based on an

analysis of the gravitational settling of small

particles. The value of 0.90 is conservative

because it does not consider the effects of

particle conglomeration and plugging. It is also

consistent with the values used in other studies

[25]. The major assumptions used to estimate the

potential CRUD release are given in Table 8.1.

To estimate the potential release from the TN-68

cask, a methodology similar to that developed

by Sandia National Laboratory (for NUREG6672 [25]) was used. This methodology was

developed for evaluation of the generic risks

associated with the transport of spent fuel by

truck and rail from commercial power plants to

potential interim storage and disposal sites.

The potential release from the TN-68 cask can

be estimated by adapting the equation developed

in NUREG/CR-6672 ([25]) to estimate the

releases from a severe fire accident. The

estimated release is given by the relationship

8.2
R = CIS(1 - D)fl Twhere R = release (curies)

Cl = amount of CRUD on fuel

assemblies (curies)

S fraction of CRUD released due

to heating

D = deposition factor

Tp= peak internal temperature (

0R)

T1 = initial internal temperature (

0

R)

Table 8.1. Assumptions Used for Release

Estimate for TN-68 Cask

Parameter' Assumed:

value

Number of Assemblies in

TN-68 Cask 68 BWR

Rods per Assembly 49

Maximum "spot" CRUD

Activity on Fuel Rod 30OpCi/cm2

Peak to axial average variation 2

CRUD decay factor (5 yr)

(based on Cow) 0.5

Average surface area per rod 1600 cm2

Average CRUD Activity on

BWR Fuel Rod (5 yr cooled) 0.12 Ci

Average CRUD Activity on

BWR Assembly (5 yr cooled) 5.9 Ci

Fraction of CRUD released

due to heating 0.15

Deposition Factor 0.90

The potential release from the TN-68 cask

based on five-year cooled fuel is estimated to be

approximately 3.4 curies of Co6w. Since the A2

value for Co6o is 11 curies, the-potential release

is about 0.3 of an A2 quantity (see footnote 6).

8.3 Results for the NAC LWT

Cask i

The thermal analysis for the NAC LWT cask

shows that this cask design would also maintain

the integrity of the fuel cladding during the

Baltimore tunnel fire scenario, and thus woiild

maintain the single most important barrier to

prevent the release of radioactive materials. The

peak temperature of the fuel cladding is

conservatively predicted to reach 10990F

(5930 C), a temperature that is below the

projected burst temperature of 13820 F (750'C)

for Zircaloy cladding.' This peak temperature

occurs at approximately'9 hours after the start of

the fire (i.e., after a'7-hour fire and 2-hour cool

down period).

However, at about 6.9 hours elapsed time, the

maximum temperature predicted for the Teflon

and metallic helicoflex seals used on the NAC

LWT lid reaches 1350'F (7321C). This value

exceeds the continuous-use rated service''

temperature limits of 7350F (3910(2) for the

Teflon seals and 8000 F (4270C)'for the metallic

helicoflex seals. Similarly, the peak temperature

experienced by the vent and drain port seals

(1I410 0 F (7660C) at approximately 6.8 hours

elapsed time), exceeds the rated long-term

service temperature of the Teflon seal material.

Table 8.2 shows the results obtained when this

equation is applied using the parameter values'

from Table 8.1 and the temperatures predicted ' , ,

for the TN-68 cask in this accident scenario. -' '

Table 8.2. Potential Release Estimate for

TN-68 Cask

Initial. Peak- - Potential

temperature temperature .release .;

OF, (OR) OF (OR) '(cu"ries),.

100 (560) 845 (1305) 3.4

Exceeding the long-ierni service temperature of

the'seals on the NAC LWT cask lid or vent and

drain ports means that there is a potential for a

release to occur. Potential releases would be

limited, however, by the narrow, convoluted

flow paths of the drain and vent ports and by the

tight clearances of the close metal-to-metal

8.3
contact between the lid and cask body. The

close contact is maintained by the pre-load

created by the initial torque on the lid bolts.

Because the fuel cladding remains intact, it is

not expected that any radioactive material would

be released from inside the fuel rods. This limits

any release from the cask to CRUD particles that

may flake off or spall from individual fuel rods.

The amount of releasable CRUD in the NAC

LWT cask was based on contents consisting of

one PWR fuel assembly containing 289 fuel

rods. An estimate of the maximum "spot"

CRUD activity shows that for 90% of PWR

spent fuel rods the maximum activity is

201.Ci/cm2 or less [26, Table I-15]. The ratio of

the peak (i.e., the maximum "spot" CRUD

activity) to average concentration on the rod

surface varies by a factor of two for PWR fuel

rods [26, Table I-12]. The CRUD activity

estimates [26] are based on newly discharged

spent nuclear fuel. The CRUD activity is

expected to decay by a factor of one-half for

five-year cooled fuel, based on the decay rate for

Co60. This proves to be a good approximation

because 92% of the activity for five-year cooled

PWR fuel comes from Co60.

Based on these data, the average CRUD activity

for a PWR rod with a surface area of 1200 cm2

is about 0.006 curies for five-year cooled fuel.

The average CRUD activity for a 17 x 17 PWR

assembly is about 1.73 Ci. The amount of

CRUD that would flake or spall from the surface

of a PWR rod due to temperatures calculated for

the fuel rods in the thermal analysis is estimated

to be a maximum of 15% [26, Table 1-10].

Finally, a deposition factor of 0.90 was used to

account for the deposition of CRUD particles on

cask surfaces and fuel assemblies.

The major assumptions used to estimate CRUD

release are given in Table 8.3. The potential

release from the NAC LWT cask can be

estimated from the same equation used for the

TN-68 release estimate, as described in Section

8.2. The major driving force for material release

results from the increased gas pressure inside the

cask due to increases in internal temperature.

The temperature change is bounded by the

difference between the maximum gas

temperature predicted during the fire transient

and the gas temperature inside the cask at the

time the cask is loaded.

For this analysis, the loading temperature is

defined as 1000F (38 0C), based on the

temperature reported in the SAR [ I1]. The

maximum gas temperature is assumed to be the

maximum peak clad temperature predicted

during the transient.

Table 8.3. Assumptions Used for

Release Estimate for NAC

LWT Cask

Assumed

Parameter value

Number of Assemblies in Cask I PWR

Rods per Assembly 289

Maximum "spot" CRUD

Activity on Fuel Rod 20 .tCi/cm2

Peak to axial average variation 2

CRUD decay factor (5 yr)

(based on Co )0.5

Average surface area per rod 1200 cm2

Average CRUD Activity on

PWR Fuel Rod (5 yr cooled) 0.006 Ci

Average CRUD Activity on

PWR Assembly (5 yr cooled) 1 .73 Ci

Fraction of CRUD released due

to heating 0.15

Deposition Factor 0.90

Table 8.4 shows the results obtained when this

equation is applied using the parameter values

from Table 8.3 and the temperatures predicted

for the NAC LWT cask in this accident scenario.

8.4
- -

The potential release from the NAC LWT cask

based on five-year cooled fuel is estimated to be

approximately 0.02 curies of Co6w. Since the A2

value for Co60 is 11 curies, the potential release

is about 0.002 of an A2 quantity (see footnote 6).

Table 8.4. Potential Release Estimate for

NAC LWT Cask

Initial ; Peak Potential -

-temperature temperature rees

.F (OR) 'F (OR) (curies)

100 (560) 1099 (1559) 0.02

8.4 Summary of Potential

Releases

The FDS model developed by NIST, as verified

by the results of the NTSB investigation of the

fire and the materials exposure analysis by

CNWRA, has provided a detailed picture of the

duration and severity of the fire that occurred in

the Howard Street tunnel in Baltimore on July

18, 2001. The fire transient analyses performed

with ANSYS and COBRA-SFS using the FDS

simulation results as boundary conditions have

shown the robust nature of the larger spent fuel

transportation package designs (HI-STAR 100

and TN-68). The predicted response of the

smaller LWT package, if hauled by rail and

exposed to the same tunnel fire environment,

indicates more component degradation.

For the TN-68 and the NAC LWT, the

maximum temperatures predicted in the regions

of the lid and the vent and drain ports exceed the

seals' rated service temperatures, making it

possible for a small release to occur, due to

CRUD that might spall off the surfaces of the

fuel rods. However, any release is expected to

be very small due to a number of factors. These

include (1) the tight clearances maintained

between the lid and cask body by the closure

bolts, (2) the low pressure differential between

the cask interior and exterior, (3) the tendency of

such small clearances to plug, and (4) the

tendency of CRUD particles to settle or plate

out.

USNRC staff evaluated the radiological

consequences of the package responses to the

Baltimore tunnel fire. The results of this

evaluation strongly indicate that neither spent

nuclear fuel (SNF) particles nor fission products

would be released from a spent fuel shipping

cask involved in a severe tunnel fire such as the

Howard Street Tunnel fire in Baltimore. None

of the three cask designs analyzed for the

Baltimore Tunnel fire scenario (TN-68, HISTAR 100, and NAC LWT) experienced

internal temperatures that would result in rupture

of the fuel cladding. Therefore, radioactive

material (i.e., SNF particles or fission products)

would be retained within the fuel rods. There

would be no release from the HI-STAR 100,

because the inner welded canister remains leak

tight and all seals remain intact. The potential

releases calculated for the TN-68 rail cask and

the NAC LWT cask (as a consequence of

exceeding seal temperature limits) indicate that

any release of CRUD from either cask would be

very small - less than an A2 quantity (see

footnote 6)

8.5
- -

9 REFERENCES

1. IO CFR 71. Jan. 1, 2003. Packaging and

Transportation of Radioactive Material.

Code of Federal Regulations, U.S. Nuclear

Regulatory Commission, Washington D.C.

2. Michener TE, DR Rector, JM Cuta, RE

Dodge, and CW Enderlin. 1995. COBRASFS: A Thermal-Hydraulic Code for Spent

Fuel Storage and Transportation Casks.

PNL-10782, Pacific Northwest National

Laboratory, Richland, Washington.

3. ANSYS, Inc. 2003. "ANSYS Users Guide

for Revision 8.0," ANSYS, Inc.,

Canonsburg, Pennsylvania.

4. McGrattan KB and A Hammins. February

2003. A. Numerical Simulation of the

Howard Street Tunnel Fire, Baltimore,

Maryland, July 2001. NUREG/CR-6793,

National Institute of Standards and

Technology, Washington D.C.

5. Garabedian AS, DS Dunn, and AH

Chowdhury. March 2003. Center for

Nuclear Waste Regulatory Analysis,

Analysis of Rail Car Components Exposed

to a Tunnel Fire Environment, NUREG/CR6799, Center for Nuclear Waste Regulatory

Analysis, Washington, D.C.

6. McGrattan KB, HR Baum, RG Rehm,

GP Forney, JE Floyd, and S Hostikka.

November 2001. Fire Dynamics Simulator

(Version 2), User's Guide. NISTIR 6784,

National Institute of Standards and

Technology, Gaithersburg, Maryland.

7. McGrattan KB, HR Baum, RG Rehm,

GP Forney, JE Floyd, and S Hostikka.

November 2001. Fire Dynamics Simulator

(Version 2), Technical Reference Guide.

NISTIR 6783, National Institute of

Standards and Technology, Gaithersburg,

Maryland.

8. Bechtel/Parsons Brinkerhoff, Inc.

November 1995. Memorial Tunnel Fire

Ventilation Test Program, Comprehensive

Test Report, Prepared for Massachusetts

Highway Department and Federal Highway

Administration.

9. NRC Docket Number 71-9293. May 19,

2000. TN-68 Transport Packaging Safety

Analysis Report, Rev. 2, TransNuclear

Incorporated, New York.

10. NRC Docket Number 72- 1008. March 30,

2001. Final Safety Analysis Report for

Holtec International Storage Transport and

Repository Cask System (HI-STAR 100

Cask System). HI-210610, Vol. I and II.

HOLTEC, Marlton, New Jersey.

11. NRC Docket Number 71-9225. Legal

Weight Truck Transport (LJT) Packaging

Safety Analysis Report, July 2000, Rev. 2;

September2001, Rev. 33; Rev. 34. Nuclear

Assurance Corporation, Alanta, Georgia.

12. 49 CFR. 171-184, Subchapter C. Oct 1,

2003. Hazardous Materials Regulations.

Code of Federal Regulations, U.S.

Department of Transportation, Washington,

D.C.

13. Creer JM, TE Michener, MA McKinnon, JE

Tanner, ER Gilbert, and RL Goodman.

1987. The TN-24P PWVR Spent-Fuel Storage

Cask: Testing and Analysis. EPRI-NP5128/PNL-6054, Electric Power Research

Institute, Palo Alto, California.

9.1
14. Rector DR, RA McCann, UP Jenquin,

CM Heeb, JM Creer, and CL Wheeler.

1986. CASTOR-IC Spent Fuel Storage

Cask Decay Heat, Heat Transfer and

Shielding Analysis. PNL-5974. Pacific

Northwest Laboratory, Richland,

Washington.

15. Lombardo NJ, JM Cuta, TE Michener, DR

Rector, and CL Wheeler. 1986. COBRASFS: A Thermal-Hydraulic Analysis

Computer Code, Volume III: Validation

Assessments, PNL-6048 Vol. III. Pacific

Northwest Laboratory, Richland,

Washington.

16. Bahney RH III and TL Lotz. July 1996.

Spent Nuclear Fuel Effective Thermal

Conductivity Report, BBAOOOO0O-017 17-

5705-00010 Rev. 00. TRW Environmental

Safety Systems, Inc., Fairfax, Virginia.

17. HEXCEL. 2003. HexWeb Honeycomb

Attributes and Properties. Available at

http.//www.hexcelcomposites.com/MarketsI

Products/Honevcomb/Hexweb attrib/hwp 0

2.htm. Hexcel, Stamford, Connecticut.

18. Guyer EC and DL Brownell, editors. 1989.

Handbook ofApplied Thermal Design.

McGraw-Hill, Inc., New York, p. 1-42.

Containing 1-, 2-, 3-, 5-, 7-, or 10-year old

PWR Spent Fuel. ORNL/CSD/TM-149.

Oak Ridge National Laboratory, Oak Ridge,

Tennessee.

23. U.S. Nuclear Regulatory Commission.

January 1997. "Standard Review Plan for

Dry Cask Storage Systems." NUREG-1536,

USNRC, Washington, DC.

24. Johnson AB and ER Gilbert. September

1983. Technical Basis for Storage of

Zircaloy-Clad Spent Fuel in Inert Gases,

PNL-4835. Pacific Northwest Laboratory,

Richland, Washington.

25. Sprung JL, DJ Ammerman, NL Breivik, RJ

Dukart, and FL Kanipe. March 2000.

Reexamination of Spent Fuel Shipment Risk

Estimates, NUREG/CR-6672, Vol. I

(SAND2000-0234). Sandia National

Laboratories, Albuquerque, New Mexico.

26. Sandoval RP, RE Einziger, H Jordan, AP

Malinauskas, and WJ Mings. January 1991.

Estimate of CRUD Contribution to Shipping

Cask Containment Requirements, SAND88-

1358. Sandia National Laboratories,

Albuquerque, New Mexico.

19. Kreith F and MS Bohn. 2001. Principles of

Heat Transfer, 6 ,h Edition. Brooks/Cole,

Pacific Grove, California.

20. Holman JP. 1986. [feat Transfer, 6"t

Edition. McGraw-Hill, Inc., New York.

21. Kreith F. 1976. Principles ofHeat

Transfer, 3 rd Edition. Intext Education

Publishers, New York.

22. Bucholz JA. January 1983. Scoping Design

Analyses for Optimized Shipping Casks

9.2
Appendix A

Material Properties for COBRA-SFS Model of TN-68 Package
Table A.1. Internal Fill Gas-Helium at Atmospheric Pressure.

Temperature Enthalpy. - Thermal Conductivity Specific Heat Specific Volume Viscosity: (0F - (Btuflibm)-; " .Btu/hr-ft-F) (Btu/llbm- 0

Fl (ft3jlbm) (bm/hr-ft) -

0 100 0.078 1.24 83.33 0.0410

200 348 0.097 1.24 119.76 0.0533

400 596 0.115 - 1.24 156.25 0.0641

600 844 0.129' 1.24 192.31 0.0727

800 1092 0.138 1.24 229.36 0.0823

1000 1340 0.138 1.24 265.25 0.0907

2552 3264 0.138 1.24 549.00 0.1138

Table A.2. External Ambient Air at Atmospheric Pressure

Temperature :,'Enthalpy Thermal Conductivity Speciflc Heat Specifle'Volume ',"Viscosity,%: . (Btu/lbm) ---- (Btn/lbm- "O(Ibtm/r-ftF'- 0F) . -(Ibm/hr-fl)

60 124.5 0.0146 --- 0.24 13.5669 0.0434

300 182.1 0.0193 0.243 19.8325 0.058

400 206.5 0.0212 0.245 22.4432 0.063

500 231.1 0.0231 0.247 25.0539 0.068

600 256 0.025 0.25 27.6645 0.072

700 281.1 0.0268 0.253 30.2752 0.077

800 - 306.7 0.0286 - 0.256 32.8859 0.081

900 332.5 0.0303 0.259 35.4966 0.085

1000 358.6 0.0319 0.262 38.1072 0.0889

2000 617.2 0.0471 : 02586 -64.214 0.1242

4000 1522 0.0671 0.4524 116.428 0.1242

Table A.3. Summary of AM Solid Material Properties Pre-Fire

spe, , e, t -D. .slty S Conducllit d. - _. ,"; ' ....... ' ' -i', itsf 

:(Btu/Ibm-¢F) (lbml/f) :- -(Btu/hr-ft-GF) Esiivty - .-' Descrlpfo- ;

0.129 - 483.8 22.92 - 0.3 gamma shielding (SA-517 grade 70 carbon steel)

0.13 - 499.4 10.44 - -0.3 - - fuel tubes (SA-240 stainless steel)

0.214 165.9 41.72 - '0.3 borated aluminum poison plates

0.311 98.5 4.34 - N/A -- neutron shield (borated polyester)

0.228 165.9 99.84 - - 0.3 - - Aluminum alloy basket rails

0.118 483.8 22.92 - 0.3 cask outer shell'

_ aluminum in neutron shield and thermal shield

0.228 165.9 84.00 .- N/A between cask and bottom impact limiter0.420 23.1 0.064 N/A' wooden impact limiters (covered with sheet steel)

, thin top layer of wood on impact limiter ends

0.420 11.0 0.053 N/A (covered with sheet steel)

'Based on nominal emissivity for carbon steel. SAR analyses use emissivity of 0.9 for painted cask surface, but cask

specifications allow option for unpainted outer surface.

A.1
Table A.4. Summary of All Solid Material Properties Post-Fire

Thermal

Specific Heat Density Conductivity

(Btu/lbm- 0P) (b(Btulhr-ft-0 F) Emissivity Description

0.129 483.8 22.92 0.3 gamma shielding (SA-517 grade 70 carbon steel)

0.13 499.4 10.44 0.3 fuel tubes (SA-240 stainless steel)

0.214 165.9 41.72 0.3 borated aluminum poison plates

hot air (replaces polyresin neutron shield

0.26 0.027 0.03 N/A vaporized in fire)

0.228 165.9 99.84 0.3 aluminum alloy basket rails

steel shell (SAR value post-fire is 0.95 for

0.118 483.8 22.92 0.8 charred cask surface emissivity)

aluminum in neutron shield; inner and outer ring

0.228 165.9 84.00 0.9 after polyresin evaporates

1020.0 134.8 0.00735 0.8 charcoal (impact limiters after the fire)

0.9 tunnel wall

COBRA-SFS Material Properties Compared with Published SAR Values

Table A.5. BWR Spent Fuel Assemblies

SAR values determined using k-effective model for homogeneous representation of fuel rods and

helium gas within fuel tube.

Transverse Thermal Axial Thermal

Temperature Conductivity Conductivity Specific Heat Density

(OF_ (Btulhr-ft-0F) (Btu/hr-ft-0F) (Btu/lbm- F) (Ibm/ft

195.8 0.0157 0.055 257.5

200.0 0.058

268.4 0.0178

365.9 0.0206

400.0 _ 0.0646

463.7 0.0239 X

561.8 0.0277

600.0 0.0709

660.3 0.0319

758.9 0.0367

800.0 0.0769 0.055 257.5

COBRA-SFS input- BWR fuel rods; conservative values at nominal operating temperature and

above.

Y

Thermal Conductivity Specific Heat Density

C7omponent (Btu/hr-ft- 0F) (Btu/Ibm OF) (Ibmnft)

fuel pellet: 3.0 0.059 655.0

cladding: 10.0 0.1 409.0

A.2
Table A.6. Stainless Steel Type 304/304L (for fuel tubes)

SAR values _ _ _

T'Temperature, 'Thermal Conductivity 'Specific Heat Densi70 7.56 0.111 499.4

100 8.76

200 9.36 0.124 - 400 10.44 0.130

600 11.28 0.134 _

800 12.24 0.140

1000 '13.2 . 499.4

COBRA-SFS input-selected conservative representative values at nominal operating

temperature and above

all 1 10.44 0.13 499.4

Table A.7. Poison Plates (borated aluminum or boron carbide/aluminum matrix)

SAR values '_:_ - Temperature Thermal Conductivity: Speclflc leat -, - Density: . E r ~~~~~~~~~-(°;i- .te)+.-o '(BtuIh - - '<4/lbm-PF) bVlIm/O0 - 68 69.36 0.214 169.3

212 83.76 - '-

482 86.64

571 86.64 0.214 169.3

COBRA-SFS input-selected conservative values based on range of allowable fabrication

variations as described for cask specifications in SAR.

all 41.72 0.214 165.9

Table A.8. Aluminum Type 6060 (for basket support rails and shims)

SAR values - ' -_-_--

Temperature Thermal Conductivity . Specific Heat: '!Densi

> - ~~~~(eF ;1';';.tuhr ift-°F)^a -- ;(flh/lh'm-0F~) --"(Ibm/ft') , 70 96.12 0.218 165.9

100 96.96 0.219

150 98.04 0.223

200 99 0.225

250 99.84 0.228 -

300 - 100.56 0.23

350 101.28 0.233

400 -101.88 0234 165.9

COBRA-SFS input-selected conservative representative values at nominal operating

temperature and above.

all 99.84 0.228 165.9

A.3
Table A.9. Carbon Steel SA-516 Grade 70 (for inner and outer gamma shield and lid)

| SAR values

Temperature Thermal Conductivity Specifc Heat Density

(IF) Btu/hr-ft-4F) (Btu/lbm--OF) (ibm/ft)

70 22.92 0.109 483.8

200 23.76 0.118

400 23.88 0.129

600 22.92 0.139

800 21.6 0.152

1000 20.16 0.169

1200 18.24 0.206

1400 15.48 0.184 483.8

COBRA-SFS input-selected conservative representative values at nominal operating temperature

and above.

all I 22.92 I 0.129 1 483.8

Table A.10. Neutron Shield (polyester resin with aluminum boxes)

SAR values-properties are composite values for polyester resin and aluminum boxes

modeled as single homogeneous material.

Temperature 1 Thermal Conductivity Specific Heat Density

CF) _ (Btu/hr-ft-0F) (Btu/lbm-OF) (Ilbm/ft')

all 0.0996 0.311 98.5

COBRA-SFS input-selected conservative representative values at nominal operating

temperature and above.

I

[borated polyester | 4.34 0.311 | 98.5

aluminum | 84.00 0.228 165.9

Table A.11. Carbon Steel SA-350 grade LF3 (for cask outer shell)

SAR values

Temperature Thermal Conductivity Specific Beat Densitd

ffi (Bht/hr-ft-0F) (Btu/lbm-.0F) (Qbm/ft)

70 23.64 0.106 489.0

100 23.88 0.11

200 24.36 0.118

400 24.24 0.128

600 23.16 0.137

800 21.72 0.149

1000 20.04 0.165

1200 18.24 0.189

1400 15.36 0.406 489.0

I

COBRA-SFS input-typical values for carbon steel at nominal operating temperature and

above, based on range of allowable fabrication variations described for cask specifications

in SAR.

all 22.92 0.118 483.8

A.4
-

Table A.12. Impact Limiters (wood covered with sheet steel)

SAR values-none provided; SAR analyses assume impact limiters act as perfect insulators on

cask ends for normal, off-normal, and fire accident conditions.

COBRA-SFS input-selected conservative representative values at nominal operating

temperature and above.

-Thermal Conductivity :Ž Specific Heat Densit-

:Mterial ; (Btuhr-ft-F K (Btuilbm-0 F) -bm/hi

redwood 0.064 0.311 98.5

balsa 0.053 0.228 165.9

carbon steel 22.92 0.118 483.8

charcoal 0.00735 1020.0 134.8

Table A.13. Air (replacing neutron shield polyethylene after fire)

SAR values |

Temperture 'Thermal Conductivity Specdfic Beat' -Density,

(Btu/hr-ft.) l-,; ' (Btu/lbm-OF) 'bmft3i)

81 0.0156 0.231 0.0734

261 0.0192 0.237 0.0551

441 0.0228 0.239 0.0440

621 0.0264 0.246 0.0367

981 0.0336 0.264 0.0275

COBRA-SFS input-selected representative values at immediate post-fire temperature and

above.

all l 0.03 | 0.26 | 0.0270

A.5
Appendix B

Material Properties for ANSYS Model of HI-STAR 100 Package
Table B.14. Homogeneous Fuel Region for Westinghouse 17x17 OFA

.Thermal ; -Thermal, ',Thermal

;fConductivity'., Conductivity Co,--: .- ,-ty Temperature :E(Btulhr-ln-' OF) (Btulhr-n-F) (iibr-ln-F) Denslt1 Specific iie t , -

: L...~. K '' x - 7l-- (,;: bmtiu) '(Btolbm-Fl?) Description

0 0.04412 0.04412 0.06256 0.14353 0.05869

100 0.04412 0.04412 0.06256 0.14353 0.05869

200 0.04412 0.04412 0.06256 0.14352 0.05869 Fuel Region

300 0.05078 0.05078 0.06509 0.14352 0.05869 (2.25 multiplier against

400 0.05895 0.05895 0.06797 -0.14352 0.05869 helium contribution to

500 0.06837 0.06837 0.07082 0.14352 0.05869 account for limited

600 0.07834 0.07834 0.07391 0.14352 0.05869 convection and

700 0.08920 0.08920 0.07756 0.14352 0.05869 pressurization

800 0.09508 0.09508 0.08121 0.15352 0.05869 enhancement)

900 0.09508 0.09508 0.08484 :0.15352 0.05869

1000 0.09508 0.09508 0.08600 0.15352 0.05869

T Table B.15. Alloy-X

Trermalm Thermal -::-.rmal,T, 

Conductivily Conductivity -Condeitlvti

Temperature (Btu/hr-ln-PF)" ;(Btu/hr-ln-°F) (BtAi/hr-In-'F) Dens) -Specific lat' -

-O) (y) O bm/V . (BtuAibm-"F) Descriptlion -

200 0.70000 * Basket Plates, Basket

40 0.81667 _ _, * 0.28993 0.12000 Supports, Boral Plate

700 0.91667 Sheathing; MPC shell,

1400 1.19670 * . impact limiter skin shel

-Table B.16. -Helium -_._.

'Termal 'Thermal. Thermal --

Conductivity Coiductivity Conduictiuvity

Temperature (Btu/br-in-lF) (Btulbr-ln-F) '(Btulhrln-QF) Density, Speclficneat - , , - C J) () , (z) (b t (BtutIbm ) Descriptlon

0 0.00650 * * - 6.90E-06

200 0.00808 . * 4.81E-06

400 0.00958 * 3.69E-06 1.24000 gas conduction between

600 0.01075 * _ 2.99E-06 MPC and cask

800 0.01150 * 2.52E-06

1400 0.01370 . - 1.71E-06

Table B.17. Helium (2.25 multiplier to account for limited convection and pressurization

enhancement) - ; ';_e

T-heral Thermal -p :Tbermal I * --

, < - "Conduttivity ' Conducifvlty~t ' '66ducilvity,'-'", -> -;l' -''

Temperature (lltu/hr.I-'n-- (Btli.IF) Dti/br-lne.) Density -Specific leat-

) - - - ' )'-: - ., - .(lbm/inl (Btu/Ibm-l) |-scription

O 0.01400 __,_._______, 6.90E-06 Conduction in: central

200 0.01740 4.81E-06 core region, between

400 0.02063 ; * 3.69E-06 guide tubes and basket

600 0.02315 * . ,, 2.99E-06 .1.24000 plates, between fuel

800 0.02476 * * 2.52E-06 and compartments, and

betweenbasket and 1400 0.02950 1 .71E3-06 MPC Shell

B.1
Table B.18. Boral Plates (includes 0.004" helium gap and gap radiation on both sides of Boral)

Thermal Thermal Thermal

Conductivity Conductivity Conductivity

Temperature (Bnt/hr-in-0F) (Btu/hr-in-0F) (Btu/hr-in^°F) Density Specific Heat

_ (F) _ Y _bin W _ (Btu/lbm-*F) Description

0 0.30836 4.62020 4.62020 0.08390

100 0.34331 4.62550 4.62550 0.08390

200 0.37738 4.64850 4.64850 0.08390

300 0.40969 4.69040 4.69040 0.08390

400 0.44166 4.73250 4.73250 0.08390 parallel to thickness

500 0.46611 4.74620 4.74620 0.08390 0.24762 (switch x & y to

600 0.49024 4.75200 4.75200 0.08390 04define cross-width)

700 0.50544 4.73700 4.73700 0.08390

800 0.52053 4.72210 4.72210 0.08390

900 0.53517 4.70710 4.70710 0.08390

1000 0.54970 4.69220 4.69220 0.08390

1100 0.56438 4.68350 4.68350 0.08390

Table B.19. Carbon Steel (SA-516, Gr. 70)

Thermal Thermal Thermal

Conductivity Conductivity Conductivity

Temperature (Btu/hr-in-°lF) (Btulhr-in^.OF) (Btu/br-in-OF) Density Specific Heat

('F) (radial) (circumferential) (axial) (ibm/in3 ) (Btu/lbm-0F) Description

200 0.17409 2.03330 2.03330

450 0.22634 1.99170 1.99170 0.28299 0.10000 Gamma Shield with

700 0.28273 1.86670 1.86670 . plates

1400 0.44136 1.46670 1.46670

Table B.20. Carbon Steel (SA-515, Gr. 70)

Thermal Thermal Thermal

Conductivity Conductivity Conductivity

Temperature (1Btulhr-in-0F) (Btu/hr-ln-0 F) (Btu.hr-inoF) Density Specific Heat

( OY) . _ (Ibni(n3 (Btu/lbin-°F) Description

200 2.43330 * * For radial channels of

700 2.05000 * * 0.28299 0.10000 overpack and enclosure of

1400 1.46670 * * shells of overpack (Fins)

Table B.21. Holtite-A

Thermal Thermal

Conductivity Conductivity Thermal

Temperature (Blu/br-in-°F) (Btu/hr-in-F) Conductivity Density Specific Heat

C.F) _ . .1). (Y) (Btu/hr-in-F) (z) (ibm/in3

) (Btulbnm-°F) Description

* 0.03108 * * 0Neutron Shield/In

0.06076 0.39000 impact limiter

Table B.22. HT-870

Thermal T Thermal

Conductivity | Conductivity Thermal Conductivity

Temperature (Btu/hr-in-°F) (Btu/hr-ln-0 F) (Btulhr-in-0 F) Densit| Speciflc Heat

I .I(.I (z) (Ibm/in ) I (Btu/lbm- 0 F) Description

* 0.00340 * * 0.00868 0.39000 Foam on back

I I I I I I I side of fins

B.2
Table B.23. . Air Properties Representing Degraded Materials

Thermal Thermal ,.Termal' -Conductivity: Cnductivity odctvt

Temperature (Btu/hr4n-F) (Btuhr-ln-0F) (Btulh'-il.F) DesItI' Specilic Heat

- (J ( 2)*,.- -, (ibm/in ) (Btu/lbm-.F) ''- .Description

200 0.00148 e . :.* 3.48E-05 0.24110

450 0.00188 * 2.53E-05 0.24605 For degraded Holite700 0.00227 * 1 .99E-05 0.25355 Honecom and

1400 1400.00336 0036**13I-S 0245 Honeycomb after fire . .. 1.311EX5 0.27445

Table B.24. One-Quarter-Inch Fillet Weld - Carbon Steel (SA-515, Gr. 70)

Thermal -, Therma1 -: termal .- - - ,<U': '.

- .. Conductivity.; Conductivity Conductivity, , . '' .

Temperature (Bltulhr-lnIaF (Btlr-ln-F) (BtuIhr4nu-eF) Dentity Specific Heat

(OF) ~ L JL.... J ~ ~ (tulmF Description

200 1.21670 2.43330 2.43330 Reduced radial channel

450 1.12920 2.25830 2.25830 0.28299 0.10000 "conductivity (Fin Fillet

700 1.02500 2.05000 2.05000 Wl ot

1400 0.73333 1.46670 1.46670 Weld_________Root)________

Table B.25. Carbon Steel (SA-516, Gr. 70)

.nt,.r SThermal - - :Thermal - .

-,onConduccivitytConductivity'-

Temperature (Btulhr-n-IF) (litulhr-in-0 F) (Btihr-ln-F) Dens t' Specific Heat

.1 IF).. t!!.... (Ibm/In) (ffbim-P) Desription 200 2.03330 _ _ Gamma Shield (intimate

700 186670 * 0.28299 0.10000 contact) and impact limiter 00 1.46670 base structure 1400 1.46670

Table B.26. Aluminum Honeycomb (700 psi unidirectional w/1700 psi cross-core backing)

Thermal .- Thermal '.~

;Conductivity..- Cnductivity Theinal Conductivity i:

Temperature (Btulhr-in-IF). (Btu/hrin- F) _ :jItulhr-In-F Densit) Sp elic Heat -

e . "(Y)- ( i)' I (lbmln') (Btu/ibm-I1

E) ' Description

68 1.11710 0.47427 1.11710 0.01406

212 1.15270 OA8944 1.15270 0.01406 0.20800 Type 1: Aluminum

752 1.42620 0.59537 1.42620 0.01406 (assumed) Honeycomb

1400 1.75440 0.72248 1.75440 0.01406

Table B.27. Aluminum Honeycomb (700 psi unidirectional and 2300 psi cross-core)

Thermal Thermal 'Thermal

|Conductivity Conductivity Conductl

Temperature (BtuibrIaF) (Btu/br-In-0F) (Btlulbr-n-F) Denslty Specific Hest

- (OeF) i (1) - -bmln (. I (1tullbinOm.) Detcrdption

68 0.82721 031682 0.82721 0.00579

212 0.85369 032693 0.85369 0.00579 0.20800 Type 2&5: Aluminum

752 1.03810 0.39771 1.03810 0.00579 (assumed) Honeycomb

1400 11.25940 0.48265 1.25940 0.00579

B.3.
Table B.28. Aluminum Honeycomb (2300 psi cross-core)

Thermal Thermal Thermal

Conductivity Conductivity Conductivity

Temperature (Btluhr-lnrF) (Btu/hr-in-1F) (fltuthr-in-0F) Densit Specific Heat

(AIF Y1) _ I (lbmhl (Btu/lbm- 0 F) Description

68 1.40690 0.63172 1.40690 0.01684

212 1.45170 0.65194 1.45170 0.01684 0.20800 Type 3: Aluminum

752 1.81430 0.79302 1.81430 0.01684 (assumed) Honeycomb

1400 2.24930 0.96231 2.24930 0.01684 1

Table B.29. Aluminum Honeycomb (1100 psi unidirectional and 2390 psi cross-core)

Thermal Thermal Thermal

Conductivity Conductivity Conductivity

Temperature (Btu/hr-in-0F) (Btu/hrin-0-F (B1tuhroin-F) Density Specific Heat

°F) ( X ( ) ) (Yb m t i 3) (Btu/1bm-IF) Description

68 1.40690 0.63172 1.40690 1.40630

212 1.45170 0.65194 1.45170 1.40630 0.20800 Type 4: Aluminum

752 1.81430 0.79302 1.81430 1.40630 (assumed) Honeycomb

1400 2.24930 0.96231 2.24930 1.40630

Table B.30. Emissivity Values for Radiation Heat Transfer

Component I Material I Enissivity

Fuel Zircaloy 0.80

Basket Alloy-X 0.36

Support Bracket Alloy-X 0.36

MPC Wall Alloy-X 0.36

Borated Aluminum Plate Boral 0.55

Bare Carbon Steel Carbon Steel 0.65

Painted Surfaces 0.90

Cask and Impact Limiter Surfaces Alloy-X 0.36

Tunnel Surface 0.90

Soot Surfaces 0.90

B.4
Appendix C

Material Properties for ANSYS Model of

Legal Weight Truck Package
-

Table C.1. 304 Stainless Steel

Temperature Thermal Conductivity- . Specific Heat'-

(O) * " . (Btulbr-in 0

°I) Dejisity (Ibm/in3) (Btutlbni-°F) - Description o

70 0.7143 - 0.1141

212 0.7800 0.2888 0.1207

392 0.8592 0.2872 0.1272 Used for cask

572 0.9333 0.2855 0.1320 body, cask lid,

752 1.0042 0.2839 0.1356 spokes

932 1.0717 0.2822 0.1385

1112 1.1375 0.2805 0.1412

Table C.2. 6061-T6 Aluminum

Temperature :Thermal Conductivity . ...; ,. I . Specific Heat :

(F) .(Btuilr-ln. 0 F) Density(lbmlin3) (Btulbm-0 F) Description

32 9.7500

212 9.9167 Used for basket,

572 11.0833 . 0.0984 0.2140 Usdobsket

932 12.9167

Table C.3. 6061-T6 Aluminum Honeycomb

Temperature Thermal Conductivity - Specific Heat. : ; ,

(

0F) (Btulhrn ) ty (bm/in 3) (Btullbm-F):' - Descrption'

32 1.6965 .. , i

212 1.7255UsdfrI1 572 1.9285 0.017118056 0.214 (Honeycomb)

932 2.2475 (Honeycomb)

Table C.4. 6061-T6 Aluminum Honeycomb'

Temperature Thermal Conductivity, -Specific Hes .

-(F) - -^(ftulhr-lio) -> 'Density (Ibm/in3) (BtulIbm-;) -- 'Descrlptlon

32 1.4235

212 1.4478 Used for IL 2

572 1.6182 0.0144 . (Honeycomb)

932 1.8858 - -- . . _ 

-__-__-_Table C.5. Helium

Temperature 1Therma] Conductivity ; ' - Specific'lleat

'(OF) (Btu/hr-n-;!F) -..Denslty(Ibm/in3)' :(Btu/lliiDF) '- i Description

200 -, 0.00808 ' 4.83E-06.

400 0.00942 . 3.70E-06 124 Used for cask gap

600 0.01075. - .- 3.OIE-06 and fuel gap

800 0.0115 - 2.52E-06 . .

C.1
Table C.6. Chemical Copper Lead

Temperature Thermal Conductivity Specific Heat

('F) (Btulhr-in-OF) Density (bm/in 3 ) (Btu/Ibm-AF) Description

68 1.6651

209 1.6308

400 1.526 0.3 0.06 Used for lead

499 1.4111 regions

581 1.2096

630 1.0079

Table C.7. 56% Ethylene Glycol Solution

Avg. Thermal

Temperature Conductivity Specific Heat Density

(,F) (Btu/hr-in-OF) (Btu/Ibm- 0F) (Ibmlin 3

)

50 0.0188 0.7405 0.0391

70 0.0187 0.7522 0.0389

100 0.0185 0.7696 0.0385

150 0.0182 0.7979 0.0378

200 0.0179 0.8255 0.0370

250 0.0177 0.8522 0.0362

260 0.0176 0.8575 0.0360

270 0.0176 0.8627 0.0358

280 0.0175 0.8679 0.0357

290 0.0175 0.8731 0.0355

300 0.0174 0.8782 0.0353

310 0.0174 0.8833 0.0351

320 0.0173 0.8884 0.0349

330 0.0173 0.8934 0.0347

340 0.0172 0.8984 0.0345

350 0.0172 0.9034 0.0343

Table C.8. Air

Avg. Thermal

Temperature Conductivity Specific Heat Density

(30F) 0Btuhr-in-. (Btu/0bm-4F) (0b.0/In

350 0.0017 0.2467 0.0000283

450 0.0018 0.2494 0.0000252

550 0.0020 0.2516 0.0000227

650 0.0022 0.2533 0.0000206

750 0.0023 0.2546 0.0000189

850 0.0025 0.2556 0.0000175

950 0.0026 0.2562 0.0000162

1050 0.0027 0.2566 0.0000152

1150 0.0029 0.2568 0.0000142

C.2
Avg.,Table C.8. Air

:-- 'vg- - , 'Themual !'.', 'i' ,' ,'.A''' I '

: Temperature ',-' Conductivity. Secific heat,. Densit

ff (tulbm-'PF) - bmuin)

1250 0.0030 0.2570 0.0000134

1350 0.0031 - .0.2571 - 0.0000126

1450 0.0033 - 0.2571 0.0000120

1550 0.0034 0.2573 0.0000114

1650 0.0035 - 0.2576 0.0000108

- 1750 - 0.0036 ----- ,0.2581 0.0000104

1850 0.0038 - ----; 0.2589 0.0000099

1950 0.0039 0.2599 0.0000095

2050 0.0040 - 0.2614 0.0000091

Table C.9. 'Effective Conductivity for Liauid Neutron Shield with 10F Temperature Gradient

3''''' ;- " 56%'Eth! ene Glycol . -Air

Avg.- _ Effective Conductivity Effective Condoctivity Effective Conductivity Effective Conductivity

Temperatureo TUre Neutron Shield ,' ' ExpainsloTank;.,_' Neutron Shield., Expansion Tank

(OF) ____1__ I u(Btulhr-iln- __ __ _ _ 0

F),-

250 0364 0.149 0.003 -0.002

260 0.374 0.153 0.003 0.002

270 0.384 0.157 0.003 0.002

280 0.393 0.161 0.003 0.002

290 0.398 0.163 0.003 0.002

300 r 0.396 0.162 0.003 0.002

310 0.395 0.162 0.003 0.002

320 0.394 0.161 0.003 0.002

330 0.393 0.161 0.003 0.002

340 0.391 0.160 0.003 0.002

350 0.390 0.160 0.003 0.002

351 * e 0.003 0.002

400 * * 0.003 0.002

500 0.003 0.002

600 * * 0.003 0.002

700 * * 0.003 0.002

800 0.003 0.002

1000 * * 0.003 0.003

1200 * * 0.003 0.003

1500 0.003 0.003

2000 * * 0.004 0.004

2500 * * 0.004 0.004

C.3
Table C.10. Effective Conductivity for Liquid Neutron Shield with 10'F Temperature Gradient

56% Et_ Gl yc__ol_ Air

Avg. Effective Conductivity Effective Conductivity Effective Conductivity Effective Conductivity

Temperature Neutron Shield Expansion Tank Neutron Shield Expansion Tank

(OF) Btu/hr-W-OF) lBtn/hr-in-F) (Bltu/hr-in-OF) (Btu/hr-n-F)

250 0.654 0.268 0.006 0.002

260 0.673 0.276 0.006 0.002

270 0.691 0.283 0.006 0.002

280 0.704 0.288 0.006 0.002

290 0.705 0.289 0.006 0.002

300 0.703 0.288 0.006 0.002

310 0.701 0.287 0.006 0.002

320 0.699 0.286 0.006 0.002

330 0.697 0.286 0.006 0.002

340 0.695 0.285 0.006 0.002

350 * *0.006 0.002

351 * *0.006 0.002

400 * *0.006 0.002

500 * *0.006 0.002

600 * *0.005 0.002

700 * *0.005 0.002

800 * *0.005 0.002

1000 * * 0.005 0.003

1200 * * 0.005 0.003

1500 * * 0.004 0.003

2000 * * 0.004 0.004

2500 * * 0.004 0.004

C.4
Table CALl Effective Conductivity for Liquid Neutron Shield with 25IF Temperature Gradient -

- 56% Ethy ene G2Wcol '- _Air,

AV - Effective Conductivity IEffective Conductivfty EffectiveiConductivty Effective Conductivity

Temperature Neutron Shield Expanslon Tank NeutronShid ExpansionTank

K(IF) : It-uhr-In-oF) -- * (Btu/hr-in- 0 BtIr F) (Biu/hr-i inF)

250 0.840 0.344 0.008 0.003

260 0.863 0.353 0.008 0.003

270 0.882 0.361 0.008 0.003

280 0.888 0.364 0.008 0.003

290 0.885 . 0.363 0.007 0.003

300 0.883 0.361 0.007 0.003

310 0.880 0.360 0.007 . 0.003

320 0.877 0.359 0.007 0.003

330 0.875 0.358 0.007 0.003

340 0.872 0.357 0.007 0.003

350 * 0.007 0.003

351 * * 0.007 0.003

400 * * 0.007 0.003

500 * 0.007 0.003

600 * ' * 0.007 0.003

700 * * 0.007 0.003

800 * 0.006 0.003

1000 * _ * 0.006 0.003

1200 * * 0.006 0.003

1500 * * 0.005 0.003

2000 * * 0.005 0.004

2500 * * 0.005 0.004

C.5
Lt .

Table C.12. Effective Conductivity for Liquid Neutron Shield with 50IF Temperature Gradient

250 1.061 0.434 0.009 0.004

260 1.058 0.433 0.009 0.004

270 1.055 0.432 0.009 0.004

280 1.052 0.431 0.009 0.004

290 1.049 0.430 0.009 0.004

300 1.046 0.428 0.009 0.004

310 1.043 0.427 0.009 0.004

320 1.039 0.426 0.009 0.004

330 0.009 0.004

340 * 0.009 0.004

350 * 0.009 0.004

351 *0.009 0.004

400 * * 0.009 0.003

500 0.008 0.003

600 *0.008 0.003

700 0.008 0.003

800 **0.008 0.003

1000 0.007 0.003

1200 **0.007 0.003

1500 0.006 0.003

2000 * * 0.006 0.004

2500 * * 0.006 0.004

C.6
- .

Table C.13. Effective Conductivity for Liould Neutron Shield with 700F Temnerature Gradient

56%Ethy eneGlycol: ir-.

yAg. i CEffective Conductivity Effective Conductivity

Temperaturet ' Neutron Shield . Epnsion Ta::. Neutron Shield -Expinsion Tank

(

0 d) (Btulhr-ln-0 F) - (Bt-ln-° F). - (Btu1hr-ln (Btu/hr-in-OF)

250 1.151 0.471 0.010 0.004

260 1.148 0.470 0.010 0.004

270 1.144 0.469 0.010 0.004

280 . 1.141 OA67 0.010 0.004

290 _ 1.138 0.466 0.010 0.004

300 1.134 OA64 0.010 0.004

310 1.131 0.463 0.010 0.004

320 * . * 0.010 0.004

330 * 0.010 0.004

340 _ 0.009 0.004

350 * * 0.009 0.004

351 * * 0.009 0.004

400 ... 0.009 0.004

500 * * 0.009 0.004

Soo * 0.009 0.004

700 * 0.008 0.003

800 * * 0.008 0.003

1000 * * 0.008 0.003

1200 * . 0.007 0.003

1500 * * 0.007 0.003

2000 * * 0.006 0.004

2500 * * 0.006 0.004

C.7
_ _ W1

Table C.14. Effective Conductivity for Liquid Neutron Shield with 100IF Temperature Gradient

250 1.253 0.513 0.011 0.004

260 1.249 0.512 0.011 0.004

270 1.245 0.510 0.011 0.004

280 1.242 0.509 0.011 0.004

290 1.238 0.507 0.011 0.004

300 1.234 0.505 0.011 0.004

310 * * 0.010 0.004

320 0.010 0.004

330 * * 0.010 0.004

340 * * 0.010 0.004

350 * * 0.010 0.004

351 * * 0.010 0.004

400 * * 0.010 0.004

500 * * 0.010 0.004

600 0.009 0.004

700 * * 0.009 0.004

800 * * 0.009 0.004

1000 0.008 0.003

1200 * * 0.008 0.003

1500 * * 0.008 0.003

2000 * * 0.007 0.004

2500 * * 0.007 0.004

C.8
Table C.15. Effective Conductivity for Liquid Neutron Shield with 200'F Temperature

Gradient

56% Ethylene Glycol - .r

- -. Effective Effective .,Effective; - .- Effective: Avg. t Conductivity, Conictivit Conductivity ; Conductivity Temperature :Neutron Shield: Epanson Tank ,. Neutron Shield . Expansion Tank

(

0F) . .(Btu/h F . Btuihr-in - (Btulhr-in-0F) (Btulhr-in-0F)

250 1.468 0.601 0.013 0.005

260 * * 0.013 - 0.005

270 * , 0.013 0.005

280 * * 0.013 0.005

290 * * 0.013 0.005

300 0.012 0.005

310 * 0.012 0.005

320 * * 0.012 0.005

330 * 0.012 0.005

340 _ * 0.012 0.005

350 * * 0.012 0.005

351 * 0.012 0.005

400 * 0.012 0.005

500 * * 0.012 0.005

600 * 0.011 0.004

700 * 0.011 0.004

800 * * 0.011 0.004

1000 * * 0.010 0.004

1200 * * 0.010 0.004

1500 * * 0.009 0.004

2000 0.008 0.004

2500 * * 0.008 0.005

C.9
Table C.16. Effective Conductivity for Liquid Neutron Shield with 300IF Temperature

Gradient

,,gTej X' perawre .R<i 

250 * * 0.014 0.005

260 0.014 0.005

270 0.014 0.005

280 * * 0.014 0.005

290 * * 0.014 0.005

300 0.014 0.005

310 * * 0.014 0.005

320 0.014 0.005

330 0.014 0.005

340 * * 0.014 0.005

350 * * 0.013 0.005

351 * * 0.013 0.005

400 * * 0.013 0.005

500 0.013 0.005

600 0.012 0.005

700 __ *0.012 0.005

800 * * 0.012 0.005

1000 * * 0.011 0.004

1200 * * 0.011 0.004

1500 * * 0.010 0.004

2000 * * 0.009 0.004

2500 * 0.009 0.005

C.10
Table C.17. Effective Conductivity for Liquid Neutron Shield with 5000F Temperature

Gradient

. 5..'--:.56% Ethylene Glycol .: . A - -' Efftive EfEffectivEfective'-: 'Effective

Avg. Conductivity ' Conducvlty -Condctivity -Coiductivity'

Temperature -' Neutron Shleld i:.'o Neuto'Sheld' Eipansion Tank'-

250 0.016 0.006

260 * * 0.016 0.006

270 * * 0.016 0.006

280 * * 0.016 0.006

290 0.016 0.006

300 0.015 0.006

310 * 0.015 0.006

320 * * 0.015 0.006

330 * * 0.015 0.006

340 * * 0.015 0.006

350 * * 0.015 0.006

351 * 0.015 0.006

400 0.015 0.006

500 _ * 0.014 0.006

600 * * 0.014 0.005

700 * * 0.014 0.005

800 * * 0.013 0.005

1000 * * 0.013 0.005

1200 0.012 0.005

1500 _ _. * 0.011 0.005

2000 * * 0.011 0.004

2500 * 0.010 0.005

Table C.18. Emissivity Values for Radiation Heat Transfer

Com.onen -. ~Ia-- ;- , Emissivity Before 'Ernissivity

--- Componen-: Meril'_ _----'Fire ''_ During/After Fire '

Canister stainless steel 0.36 0.36

Cask stainless steel 0.36 0.36

OuterNeutron Shield 0.34 0.34

Inner Neutron Shield 0.34 0.34

Basket stainless steel 0.36 0.36

Fuel Clad zircaloy 0.8 0.8

Boral Plate aluminum clad 0.55 0.55

Shell Interior stainless steel 0.36 0.36

Cask Exterior stainless steel 0.85 0.9

Tunnel/ISO various 0.9

C.l1
Appendix D

Blackbody Viewfactors for COBRA-SFS Model of TN-68 Package
C TUNNEL.1-TOP TUNNEL.2-SIDE TUNNEL.3-BOTTOM

Cnode l node_2

CASX.101, TUNNEL.1,

CASK.101, TUNNEL.2,

CASK.101, TUNNEL.3,

CASK.102, TUNNEL.1,

CASK.102, TUNNEL.2,

CASX.102, TUNNEL.3,

CASX.103,. TUNNEL.1,

CASK.103, TUNNEL.2,

CASX.103, TUNNEL.3,

CASK.104, TUNNEL.1,

CASX.104, TUNNEL.2,

CASX.104, TUNNEL.3,

CASK.201, TUNNEL.1.

CASK.201, TUNNEL.2,

CASK.201, SUNNEL.3,

CASK.202, TUNNEL.1,

CASK.202, TUNNEL.2,

CASK.202, TUNNEL.3,

CASK.203, SUNNEL.1,

CASX.203, TUNNEL.2,

,CASK.203, TUNNEL.3,

CASK.204, TUNNEL.1,

CASK.204, TUNNEL.2,

CASX.204, TUNNEL.3,

CASK.301, TUNNEL.1,

CASK.301, TUNNEL.2,

CASK.301, TUNNEL.3,

CASK.302, TUNNEL.1,

CASK.302, TUNNEL.2,

CASK.302, TUNNEL.3,

CASK.303, TUNNEL.1,

CASK.303, TUNNEL.2,

CASK.303, TUNNEL.3,

CASK.304, TUNNEL.1,

CASK.304, TUNNEL.2,

CASK.304, TUNNEL.3,

CASK.401, TUNNEL.1,

CASX.401, SUNNEL.2,

CASK.401, TUNNEL.3,

CASK.402, TUNNEL.1,

CASK.402, TUNNEL.2,

CASK.402, TUNNEL.3,

CASK.403, TUNNEL.1.

CASX.403, TUNNEL.2,

CASK.403, SUNNEL.3,

CASK.404, TUNNEL.1,

CASK.404, TUNNEL.2,

CASK.404, TUNNEL.3,

CASK.501, SUNNEL.1.

CASK.501, TUNNEL.2,

CASK.501, TUNNEL.3,

CASX.502, TUNNEL.1,

CASK.502, TUNNEL.2,

CASK.502, TUNNEL.3,

CASX.503, TUNNEL.1,

CASK.503, TUNNEL.2,

CASK.503, TUNNEL.3,

CASX.504, TUNNEL.1,

CASK.504, TUNNEL.2,

CASK.504, TUNNEL.3,

CASK.601, TUNNEL.1,

CASX.601, TUNNEL.2,

CASK.601, SUNNEL.3,

CASK.602, SUNNEL.1,

CASK.602, SUNNEL.2,

CASK.602, TUNNEL.3,

CASK.603, TUNNEL.1,

CASK.603, TUNNEL.2,

CASK.603, TUNNEL.3,

CASK.604, TUNNEL.1,

CASK.604, .TUNNEL.2,

Area-e-Bij

137.54

25.477

6.3683

67.819

87.635

13.723

12.951

90.996

65.443

6.2714

31.022-

132.32

138.22

24.978

6.3135

68.135

87.392

13.599

12.937

90. 689

65.701

6.1691

31.104

132.31

137.98

25.169

6.3354

68.192

87.312

13.687

12.880

91.267

65.225

6.1435

31.213

132.07

138.03

25.096

6.3204

68.295

87.210

13.632

13.054

90.720

65.460

6.0468

30.945

132.46

138.12

25.031

6.3279

68.608

86.890

13.693

12.946

91.115

65.179

5.9701

30.931

132.45

138.18

24.944

6.2188

68.688

86.987

13.491

13.037

91.088

65.088

5.9256

31.061

S 81 Bji 1

5 0.79418,0.00045040

5 0.14711.7.5761e-005

S 0.036772,2.1839e-005

S 0.39161,0.00022209

S 0.50603.0.00026060

$ 0.079241,4.7061e-005

S 0.074781,4.2410e-005

$ 0.52544.0.00027059

S 0.37789,0.00022443

$ 0.036213,2.0538e-005

$ 0.17913,9.225le-005

S 0.76404,0.00045376

$ 0.79812,0.00045264

S 0.14423,7.4277e-005

$ 0.036456,2.1651e-005

S 0.39343,0.00022313

$ 0.50463,0.00025988

$ 0.078527,4.6637e-005

S 0.074702,4.2365e-005

S 0.52367,0.00026968

$ 0.37938,0.00022531

S 0.035622,2.0203e-005

S 0.17961,9.2494e-005

$ 0.76399,0.00045373

$ 0.79676,0.00045187

$ 0.14533,7.4844e-005

S 0.036583,2.1726e-005

$ 0.39376,0.00022331

$ -0.50417,0.00025964

$ 0.079031,4.6936e-005

$ 0.074375,4.2180e-005

$ 0.52700,0.00027140

S 0.37663,0.00022368

$ 0.035475,2.0119e-005

$ 0.18024,9.2819e-005

S 0.76261,0.00045291

S 0.79704,0.00045202

S 0.14491,7.4627e-005

S 0.036496,2.1675e-005

S 0.39436,0.00022365

$ .0.50358,0.00025934

S 0.078717,4.6750e-005

S 0.075376,4.2748e-005

$ 0.52385,0.00026977

$ 0.37799,0.00022448

S 0.034916,1.9802e-005

$ 0.17868,9.2020e-005

$ 0.76484,0.00045424

S 0.79752,0.00045230

S 0.14454,7.4434e-005

S 0.036539,2.1701e-005

S 0.39616,0.00022468

S . - 0.50173,0.00025838

$ 0.079069,4.6959s-005

$ 0.074757,4.2397e-005

S 0.52613,0.00027095

S , 0.37636,0.00022352

S 0.034474,1.9551e-005

$ S 0.17860,9.1979e-005

S. 0.76480,0.00045422

$ 0.79788,0.00045250

$ 0.14403,7.4176e-005

$ 0.035909,2.1326e-005

S 0.39663,0.00022494

$ 0.50229,0.00025867

S 0.077899,4.6264e-005

S 0.075278,4.2692e-005

S 0.52597,0.00027087

S 0.37584,0.00022321

S .,0.034216,1.9405e-005

S 0.17936,9.2367e-005

S 0.76385,0.00045365

CASK.701, TUNNEL.1, 138.29

CASK.701, SUNNEL.2, 24.847

CASX.701, TUNNEL.3, 6.2483

- CASK.702, SUNNEL.1, 68.225

CASK.702, TUNNEL.2, 87.042

CASX.702, TUNNEL.3; 13.824

CASX.703, SUNNEL.1, 13.006

CASK.703, SUNNEL.2, 90.921

CASX.703, SUNNEL.3, 65.304

CASK.704, TUNNEL.1, 5.8582

CASX.704, TUNNEL.2, 31.026

CASX.704, TUNNEL.3, 132.40

CASX.801, TUNNEL.1, 138.10

- CASK.801, TUNNEL.2, 25.102

CASK.801, TUNNEL.3, 6.2315

CASK.802, TUNNEL.1, 67.889

CASX.802, TUNNEL.2, 87.307

CASK.802, TUNNEL.3, 13.869

CASK.803, TUNNEL,1, 12.847

CASX.803, SUNNEL.2, 90.605

CASK.803, TUNNEL.3, 65.535

CASK.804, TUNNEL.1, 5.7617

CASK.804, TUNNEL.2, 31.064

CASK.804, TUNNEL.3, 132.36

CASK.901, TUNNEL.1, 138.13

CASK.901, TUNNEL.2, 25.128

CASX.901, TUNNEL.3, 6.2207

CASK.902, TUNNEL.1, 68.091

CASX.902, 'TUNNEL.2, 87.277

CASK.902, TUNNEL.3, 13.700

CASK.903, TUNNEL.1, 12.952

CASK.903, TUNNEL.2, 90.982

CASK.903, TUNNEL.3, 65.132

CASK.904, TUNNEL.1, 5.7522

CASK.904, 'TUNNEL.2, 31.124

-CASK.904, TUNNEL.3, 132.31

'CASX.1001, TUNNEL.1, 137.84

CASK.1001, TUNNEL.2, 25.412

CASK.1001, TUNNEL.3, 6.2311

CASK.1002, TUNNEL.1, 68.564

CASK.1002, TUNNEL.2, 86.856

CASK.1002, TUNNEL.3, 13.655

CASX.1003, TUNNEL.1, 12.716

CASK.1003, TUNNEL.2, 90.924

CASK.1003, TUNNEL.3, 65.500

* CASK.1004, TUNNEL.1, 5.6573

CASX.1004, TUNNEL.2, 31.197

CASK.1004, TUNNEL.3, 132.27

CASK.1101, TUNNEL.1, 138.17

CASK.1101, TUNNEL.2, 25.006

'CASK.1101, TUNNEL.3, 6.1872

CASK.1102, TUNNEL.1, 68.446

CASK.1102, TUNNEL.2, 87.049

.CASK.1102,'TUNNEL.3,' 13.505

CASK.1103, 'TUNNEL.1, 12.985

CASK.1103, TUNNEL.2, 91.145

CASK.1103, SUNNEL.3, 65.061

CASK.1104,' TUNNEL.1,, 5.6898

CASX.1104, SUNNEL.2,- 30.943

CASK.1104, -TUNNEL.3, 132.40

CASK.1201, TUNNEL.1, 137.88

.CASK.1201, TUNNEL.2, 25.412

CASX.1201, SUNNEL.3, 6.1750'

CASX.1202, 'SUNNEL.1, 67.908

CASK.1202, TUNNEL.2, 87.439

:CASK.1202, TUNNEL.3, 13.717

CASK.1203, SUNNEL.1, 12.998

,CASK.1203,~ TUNNEL.2, 90.706

'CASX.1203, TUNNEL.3,' 65.409

CASX.1204, TUNNEL.1, 5.5919

CASX.1204, TUNNEL.2, 31.129

CASX.1204, TUNNEL.3, 132.27

CASX.1301,. TUNNEL.1, 138.25

CASK.1301, -TUNNEL.2, 24.993

S

$

S

$

$

S.

S.

S

S

$

6

$

$

S

$

S

S

$

$

$

S

$

S

S

$

S

$

$

S

S

S

S

S

$

$

S

$

S

-s

$

S

S.

-S

-. 5

$

$

S. .5 S.

S

S..

S

S.

S

S.

.5

$

'.5

'-.5

S.

.5

S

$

S

S

S

S.

$

S

0.79854,0.00045288

0.14347,7.3886e-005

0.036080,2.1428e-005

0.39395,0.00022342

0.50261,0.00025884

0.079826,4.7408e-005

0.075103,4.2593e-005

*0.52501,0.00027037

*0.37709,0.00022395

0.033827,1.9184e-005

'0.17916,9.2263e-005

0.76454,0.00045406

0.79744,0.00045225

0.14495,7.4647e-005

0.035983,2.1370e-005

0.39201,0.00022232

0.50414,0.00025963

0.080083,4.7561e-005

0.074182,4.2071e-005

0.52434,0.00027003

0.37842,0.00022474

0.033270,1.8868e-005

-0.17937,9.2374e-005

0.76430,0.00045391

0.79764,0.00045236

0.14510,7.4724e-005

0.035921,2.1333e-005

0.39318,0.00022298

0.50396,0.00025953

0.079107,4.6981e-005

0.074789,4.2415e-005

0.52536,0.00027055

0.37609,0.00022336

0.033215,1.8837e-005

'0.17972,9.2553e-005

0.76400,0.00045374

0.79593,0.00045139

0.14674,7.5569e-005

0.035981,2:1369e-005

0.39591,0.00022453

0.50153,0.00025828

0.078849,4.6828e-005

0.073426,4.1642e-005

0.52502,0.00027038

0.37822,0.00022462

-0.032667,1.8526e-005

0.18014,9.2770e-005

0.76378,0.00045361

0.79782,0.00045247

0.14440,7.4362e-005

0.035727,2.1218e-005

0.39523,0.00022415

0.50265,0.00025886

0.077980,4.6312e-005

0.074982,4.2524e-005

0.52630,0.00027104

0.37568,0.00022312

0.032855,18.633e-005

0.17868,9.2016e-005

0.76453,0.00045405

0.79616,0.00045153

0.14674,7.5567e-O05

0.035657,2.1176e-005

0.39212,0.00022238

0.50490,0;00026002

0.079208,4;7041e-005

0.075055,4.2566e-005

0.52377,0.00026973

0.37769,0.00022431

0.032289,1.'8312e-005

0.17975,9.2567e-005

0.76380,0.00045362

0.79830,0.00045274 CASK.604, SUNNEL.3, 132.28 0.14432,7.4320e-005

D.l
CASK.1301,

CASK.1302,

CASK.1302,

CASK.1302,

CASK.1303,

CASK.1303,

CASK.1303,

CASK.1304,

CASK.1304,

CASK.1304,

CASK. 1401,

CASK. 1401,

CASK. 1401,

CASK.1402,

CASK.1402,

CASK.1402,

CASK.1403,

CASK.1403,

CASK.1403,

CASK.1404,

CASK.1404,

CASK.1404,

CASK. 1501,

CASK. 1501,

CASK. 1501,

CASK.1502,

CASK.1502,

CASK.1502,

CASK.1503,

CASK.1503,

CASK.1503,

CASK.1504,

CASK.1504,

CASK.1504,

CASK. 1601,

CASK. 1601,

CASK. 1601,

CASK.1602,

CASK.1602,

CASK.1602,

CASK.1603,

CASK.1603,

CASK.1603,

CASK.1604,

CASK.1604,

CASK.1604,

CASK. 1701,

CASK. 1701,

CASK. 1701,

CASK.1702,

CASK.1702,

CASK.1702,

CASK.1703,

CASK.1703,

CASK.1703,

CASK.1704,

CASK.1704,

CASK.1704,

CASK. 1801,

CASK. 1801,

CASK. 1801,

CASK.1802,

CASK.1802,

CASK.1802,

CASK.1803,

CASK.1803,

CASK.1803,

CASK.1804,

CASK.1804,

CASK.1804,

CASK. 1901,

CASK. 1901,

CASK. 1901,

CASK.1902,

CASK.1902,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

6.1782

68. 155

87.155

13. 671

12.737

91.130

65.24 5

5.6600

31.047

132.30

138.04

25. 121

6.2019

68. 195

87.073

13.807

12.669

91.365

65.069

5. 5888

31.163

132.22

138.17

25.075

6.1911

68.139

87.197

13.646

12.784

91.009

65.230

5.6144

31.011

132.37

138.38

24.861

6.1468

68.267

87.153

13. 587

12. 646

90.521

65.886

5.5843

30.824

132.56

137.89

25.339

6. 14 99

68. 793

86.707

13.529

12.704

91.173

65.193

5.5697

31.366

132.07

138.30

24.981

6.1334

68.068

87.271

13. 649

12.694

91.352

65.097

5.5500

31.344

132.06

138.28

24.841

6.1732

68.778

86. 557

0.035675,2.1187e-005

0.39355,0.00022319

0.50326,0.00025917

0.078943,4.6884e-005

0.073546,4.1710e-005

0.52622,0.00027099

0.37674,0.00022375

0.032683,1.8535e-005

0.17928,9.2325e-005

0.76396,0.00045372

0.79707,0.00045204

0.14506,7.4703e-005

0.035812,2.1269e-005

0.39378,0.00022332

0.50279,0.00025893

0.079729,4.

7 351e-005

0.073153,4.1487e-005

0.52757,0.00027169

0.37573,0.00022314

0.032272,1.8302e-005

0.17994,9.2668e-005

0.76345,0.00045341

0.79785,0.00045249

0.14479,7.4565e-005

0.035749,2.1231e-005

0.39346,0.00022314

0.50350,0.00025930

0.078797,4.6

7

98e-005

0.073818,4.1864e-005

0.52552,0.00027063

0.37666,0.00022370

0.032419,1.8386e-005

0.17907,9.2217e-005

0.76434,0.00045394

0.79906,0.00045317

0.14355,7.3928e-005

0.035493,2.1079e-005

0.39420,0.00022356

0.50325,0.00025917

0.078457,4.6595e-005

0.073025,4.1414e-005

0.52270,0.00026918

0.38045,0.00022595

0.032245,1.8287e-005

0.17799,9.1661e-005

0.76547,0,00045461

0.79623,0.00045156

0.14632,7.5351e-005

0.035511,2.1090e-005

0.39723,0,00022528

0.50067,0.00025784

0.078123,4.6397e-005

0.073359,4.1604e-005

0.52646,0.00027112

0.37645,0.00022357

0.032161,1.8240e-005

0.18112,9.32

7 4e-005

0.76260,0.00045290

0.79856,0.00045289

0.14425,7.

4 286e-005

0.035416,2.1034e-005

0.39304,0.00022291

0.50393,0.00025952

0.078815,4.6808e-005

0.073297,4.1569e-005

0.52750,0.00027165

0.37589,0.00022324

0.032047,1.8175e-005

0.18099,9.

32 08e-005

0.76255,0.00045287

0.79848,0.00045284

0.14344,7.3870e-005

0.035646,2.11

7 0e-005

0.39714,0.00022523

0.49981,0.00025739

CASK.1902, TUNNEL.3,

CASK.1903, TUNNEL.1,

CASX.1903, TUNNEL.2,

CASK.1903, TUNNEL.3,

CASK.1904, TUNNEL.1,

CASK.1904, TUNNEL.2,

CASK.1904, TUNNEL.3,

CASK.2001, TUNNEL.1,

CASK.2001, TUNNEL.2,

CASK.2001, TUNNEL.3,

CASK.2002, TUNNEL.1,

CASK.2002, TUNNEL.2,

CASK.2002, TUNNEL.3,

CASK.2003, TUNNEL.1,

CASK.2003, TUNNEL.2,

CASK.2003, TUNNEL.3,

CASK.2004, TUNNEL.1,

CASK.2004, TUNNEL.2,

CASK.2004, TUNNEL.3,

CASK.2101, TUNNEL.1,

CASK.2101, TUNNEL.2,

CASK.2101, TUNNEL.3,

CASK.2102, TUNNEL.1,

CASK.2102, TUNNEL.2,

CASK.2102, TUNNEL.3,

CASK.2103, TUNNEL.1,

CASK.2103, TUNNEL.2,

CASK.2103, TUNNEL.3,

CASK.2104, TUNNEL.1,

CASK.2104, TUNNEL.2,

CASK.2104, TUNNEL.3,

CASK.2201, TUNNEL.1,

CASK.2201, TUNNEL.2,

CASK.2201, TUNNEL.3,

CASK.2202, TUNNEL.1,

CASK.2202, TUNNEL.2,

CASK.2202, TUNNEL.3,

CASK.2203, TUNNEL.1,

CASK.2203, TUNNEL.2,

CASK.2203, TUNNEL.3,

CASK.2204, TUNNEL.1,

CASK.2204, TUNNEL.2,

CASK.2204, TUNNEL.3,

CASK.2301, TUNNEL.1,

CASK.2301, TUNNEL.2,

CASK.2301, TUNNEL.3,

CASK.2302, TUNNEL.1,

CASK.2302, TUNNEL.2,

CASK.2302, TUNNEL.3,

CASK.2303, TUNNEL.1,

CASK.2303, TUNNEL.2,

CASK.2303, TUNNEL.3,

CASK.2304, TUNNEL.1,

CASK.2304, TUNNEL.2,

CASK.2304, TUNNEL.3,

.CASK.2401, TUNNEL.1,

CASK.2401, TUNNEL.2,

CASK.2401, TUNNEL.3,

CASK.2402, TUNNEL.1,

CASK.2402, TUNNEL.2,

CASK.2402, TUNNEL.3,

CASK.2403, TUNNEL.1,

CASK.2403, TUNNEL.2,

CASK.2403, TUNNEL.3,

CASK.2404, TUNNEL.1,

CASK.2404, TUNNEL.2,

CASK.2404, TUNNEL.3,

CASK.2501, TUNNEL.1,

CASK.2501, TUNNEL.2,

CASK.2501, TUNNEL.3,

CASK.2502, TUWNEL.1,

CASK.2502, TUNNEL.2,

CASK.2502, TUNNEL.3,

CASK.2503, TUNNEL.1,

CASK.2503, TUNNEL.2,

13.637

12.841

91.151

65.209

5.6276

30.798

132.57

138.26

24.969

6.1633

67.893

87.430

13.778

12.702

90.326

66.0 63

5. 6088

30.836

132 .54

137.99

25.223

6.2208

68.122

87.248

13. 695

12.623

91.084

65.444

5.6251

30. 935

132.48

138.00

25.240

6. 173B

68.741

86. 917

13. 4 80

12.900

91.269

64.953

5. 6790

30.883

132.53

137.99

25.190

6.2267

68.414

86.942

13.674

12.930

90. 672

65. 491

5.6849

31.044

132.32

137.91

25.207

6.2003

68.065

87.281

13. 655

12.813

90.8 43

65.521

5.7547

31.095

132.32

138.26

25. 000

6.1776

68 473

87.067

13.573

12.768

91.066

0.078746,4.6767e-005

0.074146,4.2051e-005

0.52633,0.00027105

0.37654,0.00022362

0.032496,1.8429e-005

0.17784,9.1584e-005

0.76549,0.00045462

0.79837,0.00045278

0.14418,7.4252e-005

0.035589,2.1136e-005

0.39204,0.00022234

0.50485,0.00025999

0.079561,4.7251e-005

0.073347,4.1597e-005

0.52157,0.00026860

0.38147,0.00022655

0.032387,1.8368e-005

0.17806,9.1696e-005

0.76530,0.00045451

0.79682,0.00045190

0.14564,7.50OSe-005

0.035921,2.1333e-005

0.39336,0.00022308

0.50380,0.00025945

0.079082,4.

6 967e-005

0.072889,4.1338e-005

0.52595,0.00027086

0.37790,0.00022443

0.032481,1.8421e-005

0.17863,9.1992e-005

0.76497,0.00045431

0.79686,0.00045192

0.14575,7.5057e-005

0.035650,2.11

7

2e-005

0.39693,0.00022511

0.50189,0.00025847

0.077838,4.6228e-005

0.074490,4.2246e-005

0.52702,0.00027141

0.37506,0.00022275

0.032

7 93,1.8598e-005

0.17833,9.1837e-005

0.76527,0.00045449

0.79683,0.00045190

0.14546,7.4909e-005

0.035955,2.1354e-005

0.39504,0.00022404

0.50203,0.00025854

0.078957,4.6892e-005

0.074664,4.2344e-005

0.52357,0.00026963

0.37816,0.00022459

0.032826,1.

8 617e-005

0.17926,9.2315e-005

0.76406,0.00045377

0.79636,0.00045164

0.14555,7.4957e-005

0.035803,2.1263e-005

0.39303,0.00022290

0.50399,0.00025955

0.078847,4.6827e-005

0.073989,4.1961e-005

0.52456,0.00027014

0.37834,0.00022470

0.033229,1.B845e-005

0.17955,9.2466e-005

0.76405,0.00045377

0.79836,0.00045277

0.14436,7.4342e-005

0.035672,2.1185e-005

0.39539,0.00022424

0.50275,0.00025891

0.078375,4.6547e-005

0.073729,4.1814e-005

0.52584,0.00027080

D.2
CASK.2503,

CASK.2504,

CASK.2504,

CASK.2504,

CASK.2601,

CASK.2601,

CASK.2601,

CASK.2602,

CASK.2602,

CASK.2602,

CASK.2603,

CASK.2603,

CASK.2603,

CASK.2604,

CASK.2604,

CASK.2604,

CASK.2701,

CASK.2701,

CASK.2701,

CASK.2702,

CASK.2702,

CASK.2702,

CASK.2703,

CASK.2703,

CASK.2703,

CASK.2704,

CASK.2704,

CASK.2704,

CASK.2801,

CASK.2801,

CASK.2801,

CASK.2802,

CASK.2802,

CASK.2802,

CASK.2803,

CASK.2803,

CASK.2803,

CASK.2804,

CASK.2804,

CASK.2804,

CASK.2901,

CASK.2901,

CASK.2901,

CASK.2902,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3.

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

65.294

5.8355

30.930

132.43

138.10

25.102

6.2378

68.554

86.943

13.613

12.843

91.332

64.991

5. 87 96

31.149

132.23

138.10

25.040

6.3016

68.016

87.294

13.821

12.778

90.618

65.768

5.8947

31.135

132.22

137.92

25.255

6.2785

68.305

87.253

13.602

12.785

91.110

65.358

6.0104

31.037

132.37

137.95

25.208

6.3047

68.094

0.37703,0.00022391

0.033696,1.9110e-005

0.17860,9.1976e-005

0.76468,0.00045414

0.79746,0.00045226

0.14495,7.4647e-005

0.036019,2.1392e-005

0.39585,0.00022450

0.50203,0.00025854

0.078605,4.6684e-005

0.07 4160,4.2058e-005

0.52738,0.00027160

0.37528,0.00022288

0.033951,1.9254e-005

0.17987,9.2628e-005

0.76355,0.00045347

0.79741.0.00045224

0.14459,7.4462e-005

0.036387,2.1610e-005

0.39275,0.00022274

0.50407,0.00025959

0.079805,4.7396e-005

0.073785,4.1846e-005

0.52326,0.00026947

0.37977,0.00022554

0.034038,1.9304e-005

0.1 79

7 8,9.2585e-005

0.76350,0.00045344

0.79641,0.00045167

0.14583,7.510Oe-005

0.036254,2.1531e-005

0.39442,0.00022369

0.50383,0.00025946

0.078545,4.6648e-005

0.073827,4.1869e-005

0.52610,0.00027094

0.37740,0.00022414

0.034706,1.9683e-005

0.17922,9.2294e-005

0.76435,0.00045394

0.79655,0.00045175

0.14556,7.4961e-005

0.036405,2.1621e-005

0.39320,0.00022299

CASK.2902,

CASK.2902,

CASK.2903,

CASK.2903,

CASK.2903,

CASK.2904,

CASK.2904,

CASK.2904,

CASK.3001,

CASK.3001,

CASK.3001,

CASK.3002,

CASK.3002,

CASK.3002,

CASK.3003,

CASK.3003,

CASK.3003,

CASK.3004,

CASK.3004,

CASK.3004,

CASK.3101,

CASK.3101,

CASK.3101,

CASK.3102,

CASK.3102,

CASK.3102,

CASK.3103,

CASK.3103,

CASK.3103,

CASK.3104,

CASK.3104,

CASK.3104,

CASK.3201,

CASK.3201,

CASK.3201,

CASK.3202,

CASK.3202,

CASK.3202,

CASK.3203,

CASK.3203,

CASK.3203,

CASK.3204,

CASK.3204,

CASK.3204,

TUNNEL . 2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

TUNNEL.1,

TUNNEL.2,

TUNNEL.3,

87.286

13.756

12.735

91.049

65.491

6.0330

31.507

131.85

137.97

25.049

6.3339

68.734

86.655

13.815

12.791

91.080

65.432

6.1398

31.115

132.27

137.70

25.393

6.3698

68.267

87.177

13.703

12.902

91.407

65.078

6.2118

30.914

132.43

138.23

24. 820

6.3863

67. 949

87. 413

13. 866

13.070

90. 901

65.388

6.2751

31.217

132.12

0.50402,0.00025956

0.079430,4.7173e-005

0.073535,4.1704e-005

0.52575,0.00027075

0.37817,0.00022459

0.034837,1.9757e-005

0.18193,9.3691e-005

0.76133,0.00045215

0.79670,0.00045183

0.14464,7.4489e-005

0.036574,2.1721e-005

0.39689,0.00022509

0.50038,0.00025769

0.079771,4.737 6e-005

0.073857,4.1886e-005

0.52592,0.00027084

0.37783,0.00022439

0.035453,2.0106e-005

0.17967,9.2525e-005

0.76378,0.00045361

0.79510,0.00045093

0.14663,7.551le-005

0.036781,2.1844e-005

0.39420,0.00022356

0.50339,0.00025924

0.079126,4.6993e-005

0.074502,4.2252e-005

0.52781,0.00027182

0.37578,0.00022317

0.035869,2.0342e-005

0.17851,9.1928e-005

0.76470,0.00045415

* 0.79819,0.00045268

0.14332,7.380Be-005

0.036877,2.1901e-005

0.39236,0.00022252

0.50475,0.00025994

0.080068,4.7552e-005

0.075471,4.2802e-005

0.52489,0.00027031

0.37757,0.00022424

0.036235,2.0550e-005

0.18026,9.2829e-005

0.76292,0.00045309

D.3
Appendix E

HOLTEC HI-STAR 100 Component Temperature Distributions
ANISYS S.0

MAY 10 2004

14 :32:29

1NODAL SOLUTION

STEP-10

SUB -1

TIME- .lOOE-O2

TEMP (AVG)

RSYS-O

PoverGraphlcs

EFACET-1

AVRES-Ma t

SHN -150.958

SMX -163.82 _ 150.958

CM3 151.762

G- 152.465

r 1 153.269

r 1 153.972 rD 154.776

r-J 155.48

156.284

156.987 r~i 157.791

158.494

F-J 159.298

IzzI 160.002

I S160.805

' 161.509 - 162.313

163.016

E 163.82

Figure E.l. Impact Limiter Skin Temperature Distribution -

Ii

Normal Transport Conditions.

ANSYS 8.0

KAY 10 2004

14:37:50

11ODAL SOLUTION

S73P-10

SUB -1

TIME-. 10E-02

TEMP (AVG)

RSYS-O

PowerGraphics

EFACET-1

AVRES= Mat

SE2i -153.948

SMX -214.498 _ 153.948

157.7_2

r E 161.044

IJ I164.826

ED 168.139

L-1 171.924

ED 175.235

ED 179.019

i i 182.331

i 1 186.115

CEi 189.426 E 1 193.211

iD I196.522

E i 200.307

E i 203.616

i 1 207.402

r-D 210.714

_D r214.498

- Normal Transport Conditions.

u ES . . Di 

Figure E.2. Impact Limiter Structure Temperature Distribution

E.A
xl X

Figure cond~ifl(Is

: :OIS0 I

ycontllCont 1 F i¶gu re O. 3ditons. 4 C In n e r S o cA'1 

c ra tu reflstrib u tlo -

Nolcm.2

E.2
ANSYS 5.0

14AY 10 2004

14:40:13

NODAL SOLUTION

STEP-10

SUB -1

TIME-.100E-02

TEHP (AVG)

RSYS-0

PowerGraphics

EFACET-1

AVRES-Mat

S1E4 -198.447.-

SHX -300.084

.. -204.799

/ 210.357

216.71

/ >,.> - - ;<>,-_ ....... I 1222.268

' . 234.179 ~3240.531 --- 246.089

J /1252.442

EJ258 -5 264.352

269.911

EJ |276.263

- 281.821

.-E 288.174

r-I 293.732

- 300.084

Figure E.5. Gamma Shield Temperature Distribution Normal Transport Conditions.

ANSYS 8.0

MAY 10 2004

14:41:30

NODAL SOLUTION

STEP-10

SUB -1

TIIIE-.100E-02

TEHP (AVG)

RSYS-0

PowerGraphics

EFACET-1

AVRES-Mat ._ __x>=-=- ___ --- __z;d~i SMN -173.364

SMX -257.241 _ 213.364

1e.=8 2= 178.607

-D 237.583

19g3.023

198.265

208. 095

ED 2174.24

I1222.511

I 227.753

232.34

237.sa3

21 42.17

ED 247.412

1~J 251.599

___ E 257.241

Figlure E.6. Fin Structure Temperature Distribution - Normal Transport Conditions.

E.3
iNSYS 3. 0

MAY 10 2004

14:41:S2

IIODAL SOLUTION

STEP=1 0

SUB =1

TIME=. '00E-02

TEMtIP (AOG)

RSYS=0

PowerGraphics

E'FACET-'

AVRES=Hat

286.039

- 309.529

il S330.083

-573.789

EI374.126

EJ 41398.616

_ z 66441.65i EJ462.213 m485.7,02

m 529.746

550.299

EI5C94.343

I 17.832

Figure E.7. Basket Axial Temperature Distribution - Normal Transport Conditions.

ANYSYS 8 . 0

MAY 10 2004

14:42:54

iOD.l. SOLUTION

STEP=10

I, _ |SUB = I L:= TIMIE=. 100E-02 ! TEMP (AVG)

RSYS-0

_ ~PowerGraphics ___= ____ - - EFACET=i ! ._VRES=11aL

i ilSMN =400.827

i t SNIX =661.876

iM 400.827

MOi 417.143 _ _ e __ r-43' 419

x r-441.7,34

46,2 01

l . .... . 4.9 60-2[3gl

fj 't j553.73S5

! c n r-584.-'77

i | ]6v1.284

r r- 65.5E

5766;.876

Figure E.8. Basket Radial Temperature Distribution -Normal Transport Conditions.

E.4
ANSYS R. n

FEBI 9 2005

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E.6
NRC FORM 335 US. 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) NUREGICR-6886

2. TITLE AND SUBTITLE 3. DATE REPORT PUBUSHED

Spent Fuel Transportation Package Response to the Baltimore Tunnel Fire Scenario MONTH YEAR

Draft report for comment November 2005

4. FIN OR GRANT NUMBER

__ J5167

5. AUTHOR(S) 6. TYPE OF REPORT

H. E. Adkins, Jr.

J. M. Cuta Technical

B. J. Koeppel 7. PERIOD COVERED (inclusive Dates)

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (it NRC, provide Division, Offce or Region. U.S Nudear Regulatory Commission, and mailing address; If contractor, provide name and mailing address.)

Pacific Northwest National Laboratory

Richland, WA 99352

9. SPONSORING ORGANIZATION -NAME AND ADDRESS (tfNRC. type 'Same as above'; ifcontra cor, provide NRC Oivsion, Office orRegion. U.S. NuclearRegulatory Commission,

and maihing address.)

Spent Fuel Project Office

Office of Nuclear Material Safety and Safeguards

U.S. Nuclear Regulatory Commission

Washington, D.C. 20555-0001

10. SUPPLEMENTARY NOTES

11. ABSTRACT (200t words orless)

On July 18, 2001, a freight train carrying hazardous (non-nuclear) materials derailed and caught fire while passing through the

Howard Street railroad tunnel in downtown Baltimore, Maryland. NRC staff undertook an investigation of the train derailment

and fire to determine the possible regulatory implications of this particular event for the transportation of spent nuclear fuel by

railroad.

The staff concluded that large transportation casks would withstand a fire with thermal conditions similar to those that existed in

the Baltimore tunnel fire event with only minor degradation of peripheral components. The staff also concluded that some

components of smaller transportation casks could degrade.

NRC staff evaluated the radiological consequences of the cask responses to the Baltimore tunnel fire. Though components in

some casks heated up beyond their service temperatures, the staff determined that there would be no significant release as a

result of the fire for any of these and similar casks.

12. KEY WORDS/DESCRIPTORS (List words orphrases that will assist researctiers hi 1oting the repoit) 13. AVAILABILITY STATEMENT

Spent fuel transportation , unlimited

Cask thermal performance - 14. SECURIlY CLASSIFICATION

Thermal analysis : - , . (This Page)

- . unclassified

(This Repot)

unclassified

15. NUMBER OF PAGES

16. PRICE

NRIC FORM 335 (9-2004) PRINTED ON RECYCLED PAPER
ILLL Printed

on recycled

paper

Federal Recycling Program