Spent Fuel Transportation 

Package Response to the 

MacArthur Maze Fire Scenario

Final Report

Office of Nuclear Material Safety and Safeguards

NUREG/CR-7206

PNNL-23701

 

Spent Fuel Transportation 

Package Response to the 

MacArthur Maze Fire Scenario

Final Report 

Manuscript Completed: July 2016

Date Published: October 2016

Prepared by:

H. E. Adkins, Jr.1, J. M. Cuta1, N. A. Klymyshyn1, 

S. R. Suffield1, C. S. Bajwa2, K. B. McGrattan3, 

C. E. Beyer1, and A. Sotomayor-Rivera4

1Pacific Northwest National Laboratory

P. O. Box 999

Richland, WA 99352

2International Atomic Energy Agency

Vienna International Centre

P.O. Box 100

 A-1400 Vienna, Austria

3National Institute of Standards and Technology

Engineering Laboratory

100 Bureau Drive, Stop 8600

Gaithersburg, MD 20899-8600

4U.S. Nuclear Regulatory Commission

J. Piotter, NRC Project Manager

NRC Job Code J5710

Office of Nuclear Material Safety and Safeguards

NUREG/CR-7206

PNNL-23701

iii

ABSTRACT

The U.S. Nuclear Regulatory Commission has established requirements for packaging and 

transportation of spent nuclear fuel assemblies under normal conditions of transport and for 

hypothetical accident conditions. Real-world accidents of greater severity are possible, but are 

of much lower probability, and the probability of such an accident involving a spent nuclear fuel 

(SNF) package is even lower. However, because of the potential consequences, the U.S. 

Nuclear Regulatory Commission has undertaken the examination of specific accidents to 

determine the potential consequences to an SNF package. The MacArthur Maze accident of 

April 2007, which did not involve SNF, was selected for evaluation because of the severity of the 

fire and the unusual structural consequences, in which the heat from the fire caused the 

overhead roadway segments to collapse onto the roadway where the fire was burning. 

The General Atomics GA-4 legal weight truck transportation package was selected for this 

investigation. Based on fire modeling with the Fire Dynamics Simulator code, and physical 

examination of material samples obtained onsite, a bounding fire scenario was defined for this 

accident. The complex and dynamic fire conditions are represented as a fully engulfing pool fire 

at 2012°F (1100°C) prior to the overhead roadway collapse, and as a smaller and less severe 

fully engulfing pool fire at 1652°F (900°C) afterward. 

Thermal models of the GA-4 package were constructed for the ANSYS and COBRA-SFS 

codes, to determine the response of the package to the fire scenario, including the long post-fire 

cooldown transient. Additional detailed structural and thermal-structural models were 

developed using ANSYS and LS-DYNA for the roadway and package, which showed that the 

falling overhead segments could impose only relatively innocuous loads on the stainless steel 

body and DU gamma shield, compared to the hypothetical accident conditions structural loading 

that the package is designed to withstand. 

Thermal evaluations of the package response to this fire scenario predict that the peak cladding 

temperature would exceed the short-term limit of 1058°F (570°C) long before the end of the fire. 

Maximum cladding temperatures on all rods in the package are predicted to exceed this 

temperature limit in the course of the transient, and remain above this limit for several hours. 

The maximum peak cladding temperature in the transient is predicted to be in the range of 

1350-1400°F (732-760°C), and occurs approximately 3 hours after the end of the postulated fire 

accident. Temperatures in the regions of the package seals exceed the seal material limits for 

most of the fire duration.

The FRAPTRAN-1.4 code was used to estimate a fuel rod burst rupture temperature of 1097°F 

(592°C). Together with the temperature histories of the fuel, this suggests that there is the 

potential for all rods in the package to rupture in this fire scenario. The package seals are 

assumed to fail. However, a detailed thermo-structural model showed that the lid closure bolts 

maintain a positive clamping force throughout the transient, thus limiting the release. Using 

conservative and bounding modeling assumptions, the total possible release was estimated at 

approximately one-fourth of the mixture A2. Since the regulatory limit is specified as an A2 

quantity per week for accident conditions, the estimated release is below the prescribed limit for 

safety. Therefore this very conservative estimate indicates that the potential release from this 

package, were it to be involved in a fire accident as severe as the MacArthur Maze fire scenario, 

would not pose a risk to public health and safety.

v 

TABLE OF CONTENTS

ABSTRACT ...................................................................................................................iii

LIST OF FIGURES........................................................................................................vii

LIST OF TABLES ..........................................................................................................xi

EXECUTIVE SUMMARY .............................................................................................xiii

ACKNOWLEDGMENTS.............................................................................................xxv

ABBREVIATIONS AND ACRONYMS......................................................................xxvii

1.0 INTRODUCTION.................................................................................................. 1-1

2.0 THE MACARTHUR MAZE FIRE ......................................................................... 2-1

2.1 Description of the MacArthur Maze Fire................................................................. 2-1

2.2 Maximum Material Temperatures in the MacArthur Maze Fire............................... 2-7

3.0 NUMERICAL MODELING OF THE MACARTHUR MAZE FIRE......................... 3-1

3.1 FDS Model Geometry ............................................................................................ 3-1

3.2 FDS Model Fire ..................................................................................................... 3-3

3.3 FDS Fire Model Output.......................................................................................... 3-5

3.4 Fire Model Results for Pre-Collapse Configuration ................................................ 3-9

3.5 Bounding Assumptions for Post-Collapse Fire..................................................... 3-10

4.0 THE MACARTHUR MAZE FIRE SCENARIO...................................................... 4-1

4.1 Thermal Conditions in MacArthur Maze Fire Scenario ........................................... 4-1

4.2 Structural Loads in MacArthur Maze Fire Scenario................................................ 4-2

5.0 ANALYTICAL MODELS FOR THE MACARTHUR MAZE FIRE SCENARIO..... 5-1

5.1 GA-4 Legal Weight Truck Spent Fuel Shipping Package ....................................... 5-1

5.2 ANSYS Model of GA-4 Package............................................................................ 5-3

5.2.1 GA-4 Package Representation.................................................................. 5-3

5.2.2 GA-4 Package beneath Collapsed Roadway ............................................ 5-9

5.2.3 Material Properties for GA-4 Package in ANSYS Model.......................... 5-15

5.3 COBRA-SFS Model of GA-4 Package ................................................................. 5-17

5.3.1 COBRA-SFS Model of GA-4 Package in Fully Engulfing Pool Fire.......... 5-18

5.3.2 COBRA-SFS Model of GA-4 Package beneath Collapsed Upper 

Roadway................................................................................................. 5-19

5.4 Structural Models for the MacArthur Maze Fire Scenario..................................... 5-20

5.4.1 Structural Models for Roadway Drop Scenarios...................................... 5-20

5.4.2 Preliminary Models for GA-4 Bolt Thermal Expansion Calculations......... 5-27

5.4.3 Detailed Models for GA-4 Closure Lid Bolt Evaluations........................... 5-31

6.0 ANALYSIS METHOD .......................................................................................... 6-1

6.1 NCT and HAC Fire for GA-4 Package.................................................................... 6-1

6.2 Thermal Modeling Assumptions............................................................................. 6-2

6.3 Thermal Boundary Conditions for GA-4 Package Models ...................................... 6-4

6.3.1 Thermal Boundary Conditions................................................................... 6-5

6.4 Structural Impact Model Assumptions and Analysis Method .................................. 6-8

6.4.1 Overpass Temperature Evaluation.......................................................... 6-10
vi

7.0 ANALYSIS RESULTS .........................................................................................7-1

7.1 GA-4 Package: Thermal Results for Fire Transient................................................7-1

7.1.1 Pre-collapse Fire (2012°F [1100°C]) .........................................................7-2

7.1.2 Post-collapse Fire (1352°F [900°C])..........................................................7-5

7.2 GA-4 Package: Thermal Results for Post-Fire Cooldown Transient....................... 7-9

7.3 GA-4 Package: Structural Evaluation................................................................... 7-15

7.3.1 Structural Case #1: Package Perpendicular to Upper Roadway 

Girders....................................................................................................7-19

7.3.2 Structural Case #2: Package Parallel to Upper Roadway Girders ........... 7-22

7.3.3 Structural Case #3: Impact Localized on the Package Lid....................... 7-24

7.3.4 Structural Case #4: Impact Localized on the GA-4 Trunnions ................. 7-26

7.3.5 Structural Impact Modeling Summary ..................................................... 7-29

7.3.6 Structural Issues Related to Bolt Thermal Expansion.............................. 7-29

8.0 POTENTIAL CONSEQUENCES .........................................................................8-1

8.1 Potential for Loss of Shielding ...............................................................................8-1

8.1.1 Neutron Shielding .....................................................................................8-1

8.1.2 Gamma Shielding .....................................................................................8-2

8.2 Performance of Package Containment Seals.........................................................8-2

8.2.1 Seal Temperatures in the MacArthur Maze Fire Scenario ......................... 8-2

8.2.2 Seal Performance Testing.........................................................................8-7

8.3 Potential Release Issues .......................................................................................8-8

8.3.1 Fuel Rod Cladding Performance ...............................................................8-8

8.3.2 Potential Release to GA-4 Package Cavity ............................................. 8-13

8.3.3 Potential Release from GA-4 Package in MacArthur Maze Fire 

Scenario..................................................................................................8-17

9.0 OVERALL SUMMARY AND CONCLUSIONS ....................................................9-1

10.0 REFERENCES...................................................................................................10-1

MATERIAL PROPERTIES FOR COBRA-SFS MODEL OF GA4 

PACKAGE ........................................................................................... A-1

MATERIAL PROPERTIES FOR ANSYS MODEL OF GA4 

PACKAGE ........................................................................................... B-1
vii

LIST OF FIGURES

1.1 Roadway Configuration after the MacArthur Maze Fire ............................................... 1-2

2.1 MacArthur Maze Fire at +39 Seconds .........................................................................2-1

2.2 MacArthur Maze Fire at +6.8 Minutes..........................................................................2-2

2.3 MacArthur Maze Fire at +16.7 Minutes........................................................................2-3

2.4 MacArthur Maze Fire at +19.8 Minutes........................................................................2-4

2.5 MacArthur Maze Fire at +37.3 Minutes........................................................................2-5

2.6 MacArthur Maze Fire at +72.3 Minutes........................................................................2-6

2.7 MacArthur Maze Fire at +107.3 Minutes......................................................................2-6

2.8 Approximate Locations of Collected Specimens for Materials Evaluation of Effects of 

MacArthur Maze Fire ...................................................................................................2-8

2.9 Approximate Locations of Collected Specimens Obtained from the Remains of the 

Tanker Truck Following the MacArthur Maze Fire .......................................................2-9

2.10 Typical Structural Steel Yield Strength Variation with Temperature ........................... 2-11

3.1 Simplified Model Geometry..........................................................................................3-1

3.2 Transverse View of Model Geometry...........................................................................3-2

3.3 Longitudinal Views of Model Geometry........................................................................3-2

3.4 Maximum Fire Pool Size, Based on Areas of Concrete Spalling on I-880 Surface....... 3-3

3.5 Predicted AST Values 1 Meter above the Lower Roadway Surface............................. 3-9

3.6 Predicted AST Values 3 Meters above the Lower Roadway Surface......................... 3-10

3.7 Estimated Fire Pool for Post-collapse Portion of the MacArthur Maze Fire ................ 3-11

3.8 Estimated Bounding Maximum Fire Temperatures for the MacArthur Maze Fire ....... 3-12

5.1 GA-4 Package: Exploded View....................................................................................5-2

5.2 Cross Section of ANSYS® Model of GA-4 Package Near Midplane ............................. 5-5

5.3 GA-4 Package Geometry, Including Impact Limiters....................................................5-6

5.4 GA-4 Package Geometry Model: Impact Limiter Details .............................................. 5-6

5.5 GA-4 Package Geometry Model: Closure Assembly Details........................................ 5-7

5.6 Locations Where an SNF Package Could be Covered by Collapsed I-580 Spans ..... 5-10

5.7 Diagram of Collapsed Roadway Configuration over SNF Package............................ 5-11

5.8 ANSYS Analysis of GA-4 Package: Model Element Plot (Axial View)....................... 5-12

5.9 Space Node for Convection Heat Transfer ................................................................ 5-13

5.10 Zones for Convection Computations for Package Assembly Surfaces....................... 5-14

5.11 Zones for External Heat Transfer from Roadways and External Environment............ 5-14

5.12 Cross-section of COBRA-SFS Model of GA-4 Package, Including Fuel Assemblies, 

Basket, Package Body, and Neutron Shield .............................................................. 5-18

5.13 Impact Model Geometry ............................................................................................5-21

5.14 Impact Model Mesh ...................................................................................................5-22

5.15 Impact Model Mesh Density in Impact Region........................................................... 5-22

5.16 Model Mesh of Bottom Roadway (I-880) Segment .................................................... 5-23

5.17 Model Mesh of the GA-4 Package Body.................................................................... 5-26

5.18 Model Mesh of the DU Gamma Shield....................................................................... 5-26

5.19 Diagram of Impact Limiter, Illustrating Attachment Bolts............................................ 5-28

5.20 Package Closure Region, Showing Package Closure Lid and Bolts .......................... 5-28
viii

5.21 Model Mesh of Impact Limiter Bolt.............................................................................5-30

5.22 Finite Element Model Geometry for Closure Lid, Flange, and Bolt, including Thread 

Insert .........................................................................................................................5-33

5.23 Detailed FEA Mesh in Thread Insert Region, Showing Bonded Surfaces .................. 5-34

6.1 Fire Boundary Temperatures.......................................................................................6-5

6.2 Girder Temperatures, (°C [°F]), at 18 Minutes Predicted with ANSYS Model............. 6-11

7.1 Initial Conditions for Fire Transient Analyses: GA-4 at NCT......................................... 7-1

7.2 ANSYS Results: Thermal Cross-section (°F) of GA-4 Package at NCT....................... 7-2

7.3 ANSYS Model Results: Axial and Central Radial Thermal Cross-sections (°F) of 

GA-4 Package at End of 2012°F (1100°C) Fully Engulfing Fire (37 minutes)............... 7-3

7.4 Peak Component Temperatures Predicted with ANSYS Model for 37-minute Fully 

Engulfing 2012°F (1100°C) Fire...................................................................................7-4

7.5 Peak Component Temperatures Predicted with COBRA-SFS Model for 37-minute 

Fully Engulfing 2012°F (1100°C) Fire ..........................................................................7-5

7.6 ANSYS Model Results: Axial and Central Radial Thermal Cross-sections (°F) of 

GA-4 Package at the End of the Fire at 108 Minutes...................................................7-6

7.7 Peak Component Temperatures Predicted with ANSYS Model at 108 Minutes ........... 7-7

7.8 Peak Component Temperatures Predicted with COBRA-SFS Model at 108 Minutes .. 7-8

7.9 Peak Concrete Surface Temperatures Predicted with the COBRA-SFS Model for 

the 37-minute Fire at 2012°F (1100°C) and the 71-minute Fire at 1652°F (900°C)...... 7-9

7.10 Peak Component Temperatures Predicted with COBRA-SFS Model for Post-fire 

Cooldown to 6.5 Hours ..............................................................................................7-11

7.11 Peak Component Temperatures Predicted with ANSYS Model for Post-fire 

Cooldown to 6.5 Hours ..............................................................................................7-11

7.12 Axial Temperature Evolution on Hottest Fuel Rod Predicted with the COBRA-SFS 

Model for Post-fire Cooldown with the Package under Concrete ............................... 7-12

7.13 ANSYS Model Predictions of Temperature Distributions (°F) in GA-4 Package 

during Post-fire Cooldown .........................................................................................7-13

7.14 ANSYS Model Predictions of Temperature History during the Post-fire Cooldown at 

Locations in the GA-4 Package Shielded by the Top Impact Limiter.......................... 7-14

7.15 Impact Case #1: GA-4 Package Perpendicular to Upper Roadway Girders; Main 

Impact on Package Center ........................................................................................7-16

7.16 Impact Case #2: GA-4 Package Parallel to Upper Roadway Girders; Main Impact 

along Axial Length of Package ..................................................................................7-16

7.17 Impact Case #3: GA-4 Package Oriented to Yield Main Impact on Package Lid........ 7-17

7.18 Impact Case #4: GA-4 Package Oriented for Trunnion Impact .................................. 7-17

7.19 Illustration of Actual and Calculated Failure of Overhead Roadway Girders .............. 7-18

7.20 Case #1: Pre-impact Geometry—Girders Perpendicular to SNF Package................. 7-19

7.21 Case #1: Deformation of I-580 Span after Impact...................................................... 7-20

7.22 Case #1: Effective Plastic Strain in Package Body Wall ............................................ 7-20

7.23 Case #1: Maximum Plastic Strain in Package Body Wall........................................... 7-21

7.24 Case #1: Effective Plastic Strain in DU Gamma Shield.............................................. 7-21

7.25 Case #2: SNF Package and Girder Orientation ......................................................... 7-22

7.26 Case #2: Effective Plastic Strain in Steel Body Wall of SNF Package ....................... 7-23

7.27 Case #2: Local Plastic Strain in Containment End..................................................... 7-23

7.28 Case #2: Effective Plastic Strain in DU Gamma Shield.............................................. 7-24
ix

7.29 Case #3: Girder Deformation and Effective Plastic Strain.......................................... 7-24

7.30 Case #3: Package Effective Plastic Deformation ....................................................... 7-25

7.31 Case #3: Package End Plastic Strain ........................................................................ 7-25

7.32 Case #3: DU Gamma Shield Effective Plastic Deformation ....................................... 7-26

7.33 Case #4: Girder Deformation.....................................................................................7-27

7.34 Case #4: Package Wall Plastic Strain........................................................................ 7-27

7.35 Case #4: Through-wall Plastic Strain......................................................................... 7-28

7.36 Case #4: Plastic Strain in DU Gamma Shield ............................................................ 7-28

7.37 Average Cavity Gas Temperature from ANSYS Thermal Model Calculated to 

14.5 Hours, with Power-law Extrapolation to 400 Hours ............................................ 7-35

7.38 Cavity Pressure Calculated from Average Cavity Gas Temperature, Assuming No 

Leakage ....................................................................................................................7-36

7.39 Bolt Force and Clamping Force Predicted using Classic Bolt Equations.................... 7-37

7.40 Strengths of Thread Inserts and Bolts Compared to Bolt Force Predicted using 

Classic Bolt Equations...............................................................................................7-38

7.41 Estimated Effect of Yielding Thread Inserts Using Classic Bolt Equations ................. 7-39

7.42 Diagram of Detailed Mesh in Thread Insert Region ................................................... 7-40

7.43 Stress Distribution Predicted with Elastic-only Assumptions in FEA Model................ 7-41

7.44 Diagram of Compression Volume for (a) Bolt and (b) Screw Fasteners..................... 7-42

7.45 Stress-Strain Curves used for Thread Inserts in Multi-linear FEA Model.................... 7-43

7.46 Plastic Strain in Thread Insert Predicted with Multi-linear FEA Model........................ 7-44

8.1 Operating Temperature Limit as a Function of Exposure Time for Ethylene 

Propylene Seal Material ..............................................................................................8-3

8.2 Closure Lid Seal Temperatures Predicted with the ANSYS Mode (with Impact 

Limiters) for the MacArthur Maze Fire Scenario...........................................................8-4

8.3 Gas Sample Valve/Port Seal Temperatures Predicted with the ANSYS Model (with 

Impact Limiters) for the MacArthur Maze Fire Scenario ...............................................8-5

8.4 Drain Valve/Port Seal Temperatures Predicted with the ANSYS Model (with Impact 

Limiters) for the MacArthur Maze Fire Scenario...........................................................8-5

8.5 Seal Temperatures Predicted with ANSYS Model (with Impact Limiters) for 

MacArthur Maze Fire Scenario and Post-fire Cooldown ..............................................8-6

8.6 Experimental Results for Rod Burst Rupture Testing at Low Heating Rates .............. 8-11

8.7 Summary of Activity in Radionuclides Released to GA-4 Package Cavity from WE 

15x15 (35GWd/MTU, 10-yrs-cooled fuel) for Bounding Release Fractions Specified 

in NUREG-1617 ........................................................................................................8-16

8.8 Summary of Activity in Radionuclides Released to GA-4 Package Cavity from WE 

15x15 (45GWd/MTU, 15-yrs-cooled fuel) for Bounding Release Fractions Specified 

in NUREG-1617 ........................................................................................................8-17

8.9 Equivalent Gap between Closure Lid and Package Body Flange after Seal Failure .. 8-20

8.10 Volumetric Leak Rate for GA-4 Package after Seal Failure ....................................... 8-21

xi

LIST OF TABLES

2.1 Material Temperatures Estimated from Evaluation of Samples.................................... 2-9

3.1 Summary of Heat Release Rates for Large Pool Fires ................................................3-5

5.1 Summary of Elements in ANSYS Model of GA-4.........................................................5-4

7.1 Key Bolt Parameters..................................................................................................7-34

7.2 Results of Classical Bolt Analysis ..............................................................................7-36

7.3 Best Estimate Results from Non-Linear FEA Model................................................... 7-44

8.1 Results of Fuel Performance Analyses in the MacArthur Maze Fire Scenario............ 8-12

8.2 Time above Predicted Rod Rupture Temperatures in the MacArthur Maze Fire 

Scenario....................................................................................................................8-13

8.3 Bounding Values of Release Fractions from Ruptured Fuel Rods ............................. 8-14

8.4 Radionuclide Inventory for a Single Assembly in the GA-4 Package ......................... 8-14

xiii

EXECUTIVE SUMMARY

The U.S. Nuclear Regulatory Commission (NRC) has established requirements for packaging 

and transportation of spent nuclear fuel (SNF) assemblies under normal conditions of transport 

(NCT) and for hypothetical accident conditions (HAC). These requirements (10 CFR 71) 

conservatively bound fire conditions that an SNF package might credibly encounter. However, 

real-world accidents of greater severity are certainly possible, and the NRC has undertaken the 

examination of such accidents, to determine what the potential consequences might be for a

spent nuclear fuel package. Two previous studies of transportation accidents, one resulting in a 

fire in a railroad tunnel (NUREG/CR-6886 2009) and one in a highway tunnel (NUREG/CR-6894

2007) were undertaken with three different SNF package designs. Based on conservative 

scenarios constructed from these real-world fire conditions, the results of these studies have 

shown that the design basis for SNF packages is sufficiently robust for them to survive such 

beyond-design-basis conditions without adverse consequences to public safety. In all cases 

evaluated, the modeling results showed that the various SNF packages would be expected to 

maintain required shielding for ionizing radiation, and also would maintain the integrity of the 

containment boundary sufficiently to limit potential release of radioactive material from the 

packages to within regulatory bounds for accident conditions.

The MacArthur Maze accident of April 29, 2007 was selected as a third study in this series of 

evaluations of real-world accidents because of the severity of the fire and the unusual structural 

consequences, in which the heat from the fire caused the overhead roadway segments to 

collapse onto the lower roadway where the fire was burning. Since this was a highway 

accident, the only type of SNF package that could potentially be involved would be a legal 

weight truck (LWT) package. The General Atomics GA-4 LWT transportation package was 

selected for this investigation, mainly because it can carry a relatively large payload for an overthe-road transportation package, and therefore the potential consequences of package failure 

could be more severe than for packages with smaller payload capacities. The GA-4 package is 

designed to transport up to four intact pressurized water reactor spent fuel assemblies, with a 

maximum total package decay heat load of 2.5 kW.

The MacArthur Maze accident involved a gasoline tanker truck and trailer that overturned and 

caught fire on the I-880 connector of the MacArthur Maze interchange in Oakland, CA. The fire 

lasted approximately 108 minutes, consuming the tanker’s entire load of 8,600 gallons of 

gasoline. The heat from the fire caused two sections of the overhead I-580 freeway to collapse 

onto the lower roadway, the first falling at approximately 17 minutes into the fire, the second 

collapsing on only one end, and reaching its final configuration by about 37 minutes. Figure S.1 

shows an image of the fire just prior to the collapse of the first overhead roadway section to fall. 

Figure S.2 shows the configuration of the collapsed roadway, in an image taken in daylight the 

next day, after the fire was out. (Note that these images were captured from opposite sides of 

the freeway, and therefore the left-right orientation of the sections of roadway is reversed in the 

two images.)
xiv

Figure S.1. MacArthur Maze Fire at +16.7 Minutes (WTP Video Image at 03:54:24.61 PDT, 

photo from MAIT Report, CHP 2007, reprinted with permission.)

Figure S.2. Roadway Configuration after the MacArthur Maze Fire (photo from MAIT Report, 

CHP 2007, reprinted with permission.)
xv

Based on fire modeling with the Fire Dynamics Simulator (FDS) code, and physical examination 

of material samples obtained from the damaged highway girders and the remnants of the tanker 

truck, a bounding fire scenario was defined for the thermal and structural evaluations of the 

potential effects of this fire on an SNF package. The complex and dynamic fire conditions are 

represented as a fully engulfing pool fire at 2012°F (1100°C) prior to the overhead roadway 

collapse, and as a slightly smaller and less severe fully engulfing pool fire at 1652°F (900°C) 

after the roadway collapse. These temperatures represent conservative bounding values for 

open pool hydrocarbon fires for any possible configuration of both the pre-collapse and postcollapse fire pools in this accident. 

As an additional simplifying conservatism in the definition of the scenario, it is assumed that the 

pre-collapse pool fire (at 2012°F [1100°C]) lasts for the full 37 minutes required for the 

completion of the collapse of the overhead segments. The smaller fire size is assumed as a 

step change to 1652°F (900°C), after 37 minutes, and this smaller pool fire is assumed to 

persist unchanged until the end of the fire, at 108 minutes. The fire scenario for modeling 

purposes also assumes that in the post-fire configuration, the fallen overhead roadway segment

completely covers the SNF package, resulting in an additional barrier to heat transfer from the 

package during the cooldown phase of the transient.

Thermal and Structural Modeling Approach and Summary of Results

Detailed thermal models of the GA-4 package were constructed for the ANSYS and COBRASFS codes, for transient evaluations to determine the temperature response of the package to 

the fire scenario, including the long post-fire cooldown transient. Figure S.3 shows an exploded 

view of the GA-4 package, illustrating its main design features. Figure S.4 shows an axial

cross-sectional diagram of the ANSYS model of the package. 

Figure S.3. GA-4 Package: Exploded View (General Atomics 1998)
xvi

Figure S.4. Axial Cross-Section of ANSYS Model of GA-4 Package

Figure S.5 shows a cross-sectional diagram of the COBRA-SFS model of the package. The 

initial condition of the package at the start of the fire scenario was defined as steady-state NCT. 

Additional detailed structural and thermal-structural models were also developed using ANSYS 

and LS-DYNA for the roadway and package, for evaluation of the package response to the 

effect of the roadway falling on it.

Figure S.5. Cross-Section of COBRA-SFS Model of GA-4 Package

fuel rod

control rod 

thimble

fuel assembly

fuel assembly

fuel assembly

fuel assembly
xvii

Considerable effort was given to defining bounding and conservative estimates of the possible 

configurations of the package on the roadway that could produce the “worst case” structural 

loading of the SNF package due to the overhead roadway spans falling on it. However, the 

results of these evaluations clearly showed that the most adverse possible impacts of the 

overhead spans imposed relatively innocuous loads on the stainless steel body and depleted 

uranium (DU) gamma shield of the package, compared to the HAC structural loading that the 

package is designed to withstand. At a nominal fully loaded weight of approximately 55,000 lb

(nearly 28 tons), the package itself falling from a height of 30 ft (9 meters) – the HAC package 

drop scenario (see 10 CFR 71) – would be expected to do far more damage, even with the 

added impact of the projecting “blades” of the steel girders. 

The only real challenge of the overhead roadway drop in the fire scenario is that the impact is 

postulated to occur with the package at higher temperatures than are typically assumed in the 

structural analyses for HAC scenarios. (The HAC drop is postulated to occur before the HAC 

fire [10 CFR 71]). This could potentially make the package more vulnerable to structural 

damage, due to the reduction in the strength of steel with increasing temperatures. However, 

the steel girders of the overhead span suffer more from this effect, and the weight of the 

overhead roadway concrete is not sufficient to impart significant loading to damage the package 

in any way. Figure S.6 (a) and (b) illustrates the results of the most severe case of dropping the 

overhead roadway span onto the GA-4 package, and the resulting plastic strain in the package 

body wall.
xviii

(a)

(Note: image of upper roadway shows girders only; concrete roadway omitted from image for 

clarity.)

(b)

Figure S.6. Deformation of I-580 Span after Impact with Package Body, (a) Predicted 

Deformation Due To Impact of Upper Roadway on GA-4 Stainless Steel Body, and 

(b) Predicted Effective Plastic Strain in Package Body Wall
xix

Much more interesting structural analyses were undertaken to investigate in detail the response 

of the bolts attaching the impact limiters to the package, and the package lid closure bolts. 

Complete evaluation of bolt performance was further complicated by the use of thread inserts in 

all bolt attachments in the package, in which helical coils of Type 304 stainless steel fill the 

interface between the bolt threads and the threaded holes in the package body. Differential 

thermal expansion of the Inconel bolts relative to the XM-19 stainless steel package body, and 

different strength-versus-temperature properties of the three metals involved, results in a timeand temperature-dependent history of force on the bolts that raised the possibility that the 

impact limiters might detach from the package. These material issues also raised the possibility 

that there could be a loss of clamping force between the lid and the package body during the 

post-fire cooldown. Figure S.7 illustrates the finite element analysis (FEA) model meshes 

constructed for detailed modeling of the impact limiter bolts and for the closure lid, flange and 

bolt structure, including thread inserts.

 (a) (b)

Figure S.7. Detailed Model Mesh Diagrams for (a) Impact Limiter Bolt Evaluations with ANSYS 

and (b) Closure Lid Bolt Evaluations with LS-DYNA

Detailed evaluations of the structural and thermal response of the impact limiter bolts to the 

conditions of the MacArthur Maze fire scenario with FEA modeling using ANSYS show 

definitively that the impact limiter bolts will not fail under these conservative and bounding 

thermal and structural loading conditions. Loss of the impact limiters is not a credible 

consequence of this fire scenario for the GA-4 package. Additional detailed evaluations of the 

response of the lid closure bolts to the fire scenario undertaken with LS-DYNA show 

unambiguously that the lid closure bolts maintain a positive clamping force between the 

package lid and body flange during all phases of the fire scenario, including the fire duration 

(108 minutes) and the very long cooldown period of approximately 400 hours, back to post-fire 

steady-state ambient conditions. This means that there is at all times forced metal-to-metal 

contact between the lid and the package body. This is particularly important to assessing the 

response of the GA-4 package to this fire scenario, because the thermal evaluations show that 

the seals exceed their rated temperature limits within the first hour or so of the transient. The

metal-to-metal contact with positive clamping force constitutes the main containment boundary 

of the package in this scenario.
xx

Thermal evaluations of the package response to this fire scenario predict that the peak cladding 

temperature would be expected to exceed the short-term limit of 1058°F (570°C) long before the 

end of the fire. Maximum cladding temperatures on all rods in the package are predicted to 

exceed this temperature limit in the course of the transient, and remain above this limit for 

several hours. In addition, the thermal inertia of the package and the insulating effect of the 

fallen overhead roadway, which is assumed to blanket the package during the post-fire 

cooldown means that fuel cladding temperatures continue to rise for many hours after the end of 

the fire. The insulating effect of the impact limiters, which shield the package ends from direct 

heating by the fire, results in the cooler ends of the rods continuing to heat up for several hours 

after the end of the fire, as heat in the hot central region of the rods redistributes throughout the 

package. Figure S.8 illustrates the thermal response of the package from the end of the fire to 

approximately 12 hours into the cooldown transient, which at that point is far from over.
xxi

 (a) at end of fire (1.8 hours)

 (b) 2.2 hours after end of fire

 (c) 6.2 hours after end of fire

 (d) 12.2 hours after end of fire

Figure S.8. ANSYS Model Predictions of Temperature Distributions (°F) in GA-4 Package 

during Post-fire Cooldown
xxii

The maximum peak cladding temperature in the transient is predicted to be in the range of 

1350-1400°F (732-760°C), and occurs approximately 3 hours after the end of the fire. Based on 

the predicted fuel cladding temperatures from the COBRA-SFS modeling of the complete 

MacArthur Maze fire scenario, fuel performance was evaluated by direct comparison to fuel rod 

burst data as a function of cladding hoop stress and temperature. In addition to comparison to 

relevant data, predicted fuel rod rupture temperatures were obtained using the burst rupture 

model in the FRAPTRAN-1.4 code (NUREG/CR-7023 2011). Creep rupture is considered a 

possible alternative mechanism of failure for spent fuel rods. To evaluate this possibility, a 

separate analysis was performed with a creep rupture model, using the FRAPCON-3.4 code 

(NUREG/CR-7022 2011) in conjunction with the DATING code (Simonen and Gilbert 1988; 

Gilbert et al. 2002).

Fuel performance analyses for peak temperatures on the hottest rod in the MacArthur Maze fire 

scenario as predicted with the COBRA-SFS model, predict cladding rupture temperatures of 

1097°F (592°C) using LOCA burst strain modeling (FRAPTRAN) and 1229°F (665°C) using 

creep rupture modeling (FRAPCON/DATING). Applicable experimental data (NUREG/CR0344) yields measured rupture temperatures in the range 1205-1256°F (652-680°C). The burst 

rupture and creep rupture models both predict that the hottest fuel rod would rupture if subjected 

to the temperatures predicted in this fire scenario. Furthermore, the peak temperature on the 

hottest rod at the time of rupture is eventually exceeded by all rods in the package during the 

transient, which suggests that there is the potential for all rods in the package to rupture in this 

fire scenario.

Potential Radiological Consequences

Neutron and gamma radiation dose rates from the GA-4 package as a result of the postulated 

conditions of the MacArthur Maze fire scenario would not exceed the design basis of the 

package, which is well within the regulatory limits for hypothetical accident conditions. The 

neutron shielding is lost very early in the transient, but loss of the neutron shield tank is a 

design-basis assumption for this package in all HAC analyses. The more severe conditions of 

the MacArthur Maze fire can do no more damage to the GA-4 package neutron shield than is 

assumed a priori in the HAC analyses. The gamma shielding for the GA-4 is provided by a 

layer of DU within the stainless steel package body. The shielding function of this material is 

not affected by the higher temperature it is predicted to reach in the MacArthur Maze fire 

scenario. There is no credible scenario in this fire accident that could result in neutron and 

gamma dose rates from the design-basis GA-4 package exceeding the regulatory limits for 

accident conditions.

Loss of the package seals due to exceeding seal material thermal limits means that there is the 

potential for radioactive material to escape from the package. Rupture of all rods in the 

package, as is predicted by the fuel performance analyses, based on the calculated thermal 

response of the fuel, means that fission gases and fuel particulate would be released to the 

package cavity. In addition, the assumption of 100% spalling of CRUD from the external 

surfaces of the fuel rods is assumed for all accident conditions for SNF packages, per NRC 

guidance. Therefore, it must be assumed that there is material available in the package cavity 

that could be released through the failed seals. But because the lid closure bolts maintain

positive clamping force throughout the transient, it is not physically possible for very much of it 

to actually escape. Conservative and bounding modeling assumptions yield an estimate of the
xxiii

maximum possible release as 0.24 of the mixture A2 quantity1 determined for the design basis 

contents of the package. The HAC regulatory limit specifies a maximum allowable release rate

of an A2 per week. The estimated value is for the total potential release from the package in this 

scenario. This predicted release estimate is below the prescribed limit, and indicates that the 

potential release from this package in the MacArthur Maze fire scenario would not pose a risk to 

public health and safety.

 1 An A2 quantity is defined in 49 CFR 173.403 as the maximum activity of a Class 7 (radioactive) material 

permitted in a Type A package, which does not require an accident resistant design. The amount of 

material that constitutes an A2 quantity depends on its specific activity and other radiological properties. 

Appendix A of 10 CFR 71 specifies the specific A2 quantities for a large number of radioactive materials, 

and defines methods for calculating values for materials not listed in the table. Spent nuclear fuel 

requires a Type B package, which can carry more than an A2 quantity of radioactive material, but must 

retain the integrity of containment and shielding under normal conditions of transport (as per 49 CFR 173) 

and meet the release limits of less than an A2 per week for hypothetical accident conditions.

xxv 

ACKNOWLEDGMENTS 

This evaluation of the potential thermal and structural consequences to an SNF package of a 

transportation accident of the severity of the MacArthur Maze fire and highway collapse has 

been the work of many contributors over several years. The successful completion of this work 

has been due in large measure to Chris Bajwa, the NRC Project Manager initially guiding this 

effort, and who has continued to support and encourage the work after leaving NRC to work 

with the IAEA. The work of Kevin McGrattan (NIST), was an invaluable contribution, providing 

realistic assessment of the behavior of the fire itself, and defining the bounding fire scenario 

used in the analysis. The California Highway Patrol (CHP), in particular the members of the 

Multi-Disciplinary Accident Investigation Team (MAIT), provided invaluable assistance and 

detailed information, through personal communications and also in the detailed and 

comprehensive MAIT Report published by CHP describing the accident. CHP officers 

personally assisted the NRC in obtaining physical samples of fire-damaged girders, concrete, 

and truck remnants, used to provide direct estimates of temperatures reached in the fire. The 

work of Carl Beyer of PNNL, and his expertise related to nuclear fuel and cladding behavior

under severe thermal and mechanical stresses that could lead to burst rupture was vital to the 

assessment of the potential consequences of exposing spent nuclear fuel to an accident of the 

severity of the MacArthur Maze scenario. The efforts of technical editors Colleen Winters, 

Cornelia Brim, and Susan Tackett were absolutely essential to making this work accessible, 

readable, and understandable to the wide audience for which it has been prepared. 

xxvii

ABBREVIATIONS AND ACRONYMS

ASME American Society of Mechanical Engineers

AST adiabatic surface temperature

ASTM American Society of Mechanical Engineers

BCL Battelle Columbus Laboratory

CFD computational fluid dynamics

CG center of gravity

CHP California Highway Patrol

CNWRA Center for Nuclear Waste Regulatory Analyses

CRUD Chalk River Unknown Deposit, a generic term for corrosion and 

wear products (rust particles, etc.) that become radioactive (i.e., 

activated) when exposed to radiation.

DU depleted uranium

FDS Fire Dynamics Simulator

FEA finite element analysis

FSS fuel support structure

HAC hypothetical accident conditions

I-580 Interstate 580

I-880 Interstate 880

IAEA International Atomic Energy Agency

ILSS impact limiter support structure

LOCA loss-of-coolant accident

LWT legal weight truck

MAIT Multi-Disciplinary Accident Investigation Team

NCT normal conditions of transport

NIST National Institute of Standards and Technology

NRC U.S. Nuclear Regulatory Commission

NS neutron shield

ORNL Oak Ridge National Laboratory

PWR pressurized water reactor

SAR safety analysis report

SARP Safety Analysis Report for Packaging

SFST Spent Fuel Storage and Transportation

SNF spent nuclear fuel

TBq Terabecquerel (SI unit for radioactivity; equal to 27 Curies (Ci))

WTP East Bay Municipal Utility District Wastewater Treatment Plant

1-1

1.0 INTRODUCTION

Current U.S. Nuclear Regulatory Commission (NRC) regulations specify that spent nuclear fuel 

(SNF) transportation 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 

1475°F (800°C)” (10 CFR 71). The package1 must maintain containment, shielding, and 

criticality functions throughout the fire event and post-fire cooldown in order to meet NRC 

requirements. The performance of spent fuel packages in severe accidents has been examined 

in previous studies by the NRC, as documented in NUREG-0170 (Final Environmental 

Statement on the Transportation of Radioactive Material by Air and Other Modes), NUREG/CR4829 (Shipping Container Response to Severe Highway and Railway Accident Conditions, also 

known as the “Modal Study”), and NUREG/CR-6672 (Re-examination of Spent Fuel Shipment 

Risk Estimates). These studies evaluated a broad range of hypothetical transportation 

accidents involving collisions, fires, and collisions followed by fires. However, these studies did 

not specifically examine the effects of an actual transportation accident involving a severe fire 

that included a roadway collapse.

NRC has undertaken the examination of real-world accidents of greater severity than postulated 

in the hypothetical accident conditions (HAC) fire, to determine what the potential consequences 

might be, were such an accident ever to involve an SNF package. Two previous studies of 

transportation accidents have been performed; the first was of the 2001 fire in the Howard 

Street railroad tunnel in Baltimore, MD (NUREG/CR-6886 2009) and the second was of the 

1982 fire in the Caldecott Tunnel on California State Route 24 near Oakland, California

(NUREG/CR-6894 2007). Based on conservative scenarios constructed from these real-world 

fire conditions, the results of these studies have shown that the design basis for SNF packages 

is sufficiently robust for them to survive such beyond-design-basis conditions without adverse 

consequences to public safety. In all cases evaluated, the modeling results showed that the 

various SNF packages would be expected to maintain required shielding for ionizing radiation, 

and also would maintain the integrity of the containment boundary sufficiently to limit potential 

release of radioactive material from the packages to within regulatory bounds for accident 

conditions.

The MacArthur Maze accident of April 29, 2007 was selected as a third study in this series of 

evaluations of real-world accidents because of the severity of the fire and the unusual structural 

consequences, in which the heat from the fire caused the overhead roadway spans to collapse 

onto the roadway where the fire was burning. On April 29, 2007 at approximately 3:37 a.m., a 

tanker truck and trailer carrying 8,600 gallons (32,554 liters) of gasoline overturned and caught 

fire on the Interstate 880 (I-880) connector of the MacArthur Maze interchange located in 

Oakland, California. The intense heat from the fire weakened the steel girders of the Interstate 

580 (I-580) roadway above the fire, collapsing two adjacent spans (approximately 156 feet 

[47.55 m]) of the elevated roadway onto the section of freeway below. A surveillance camera 

from the monitoring system of the East Bay Municipal Utility District Wastewater Treatment 

Plant (WTP) adjacent to the roadway captured a video of almost the entire fire duration. This 

video shows the first I-580 roadway span beginning to sag by about 10 minutes into the fire and 

collapsing completely at approximately 17 minutes. The video also shows a second span of the 

I-580 roadway descending slowly to the lower (I-880) roadway, beginning at about 17 minutes 

and reaching its final (partially collapsed) configuration by about 37 minutes. The video shows 

 1 The term “package” refers collectively to the contents (in this case spent nuclear fuel), and the protective 

enclosure into which the contents are placed.
1-2

that the collapse of the second span greatly reduced the size of the fire, but it continued to burn 

intensely until about 102 minutes. As a fire management decision, the first responders on the 

scene allowed the fire to burn unchecked until the hydrocarbon fuel was fully consumed. At that 

point, the fire began to noticeably decrease in brightness, diminishing to a small glowing spot by 

approximately 108 minutes after the start of the fire. On the video, there is no visible glow from 

the fire after about 120 minutes. 

Figure 1.1 shows a post-fire aerial view of the collapsed spans, extracted from the California 

Highway Patrol Multi-Discipline Accident Investigation Team (MAIT) report (CHP 2007). The 

staff of the NRC Division of Spent Fuel Storage and Transportation (SFST) undertook an 

investigation of the fire and roadway collapse to determine what impact this event might have on 

the risk associated with SNF transportation on public roadways. This evaluation included an 

assessment of the fire exposure temperatures of the upper roadway girders and tanker truck 

(NRC 2008), computational fluid dynamics (CFD) modeling of the open pool (pre-collapse) 

portion of the fire, and an analytical evaluation of the response of a representative NRC certified 

SNF transportation package to boundary conditions simulating temperatures predicted for the 

MacArthur Maze fire.

Figure 1.1. Roadway Configuration after the MacArthur Maze Fire (photo from MAIT Report, 

CHP 2007, reprinted with permission.)

This report provides a description of the analytical evaluation of the transportation package 

response and a detailed presentation of the results of the evaluation. Section 2.0 contains a 

summary description of the MacArthur Maze fire, and Section 3.0 describes the numerical 

modeling of the fire. Section 4.0 describes the fire scenario developed for this evaluation in 

detail, based on the known accident conditions and the numerical modeling of the fire. 

Analytical models of the SNF transportation package are described in Section 5.0. Section 6.0

presents the analytical approach, including detailed description of modeling assumptions. 

Analysis results are presented in Section 7.0. Section 8.0 addresses potential consequences of 

the fire scenario, with respect to the SNF transportation package. Results and conclusions of 

this study are summarized in Section 9.0, and references are listed in Section 10.0.
2-1

2.0 THE MACARTHUR MAZE FIRE

This section presents a detailed description of the fire and summarizes the results of analyses 

of material samples obtained to characterize temperatures reached by structures in or near the 

fire. Section 2.1 contains the fire description. Section 2.2 summarizes the analyses undertaken 

to determine estimates of peak temperatures reached in sampled materials from the roadway 

and tanker truck exposed to the fire. 

2.1 Description of the MacArthur Maze Fire

Documentation of the MacArthur Maze fire is unusual in that nearly the entire fire duration was 

captured on video by the surveillance camera system of the nearby East Bay Municipal Utility 

WTP. The WTP video shows the rapid development of a large, openly burning fire on the I-880 

roadway, ignition of ground fires below the roadway, the collapse of the overhead I-580 spans, 

and the post-collapse fire near Bent1 19. Key points in the fire duration are illustrated with 

video-capture images in Figure 2.1 through Figure 2.7.

Figure 2.1. MacArthur Maze Fire at +39 Seconds (WTP video image at 03:38:22.93 PDT,

photo from MAIT Report, CHP 2007, reprinted with permission.)

 1 The term “bent” is used by the California highway authority (CalTrans) to refer to the structures 

consisting of a horizontal beam supported by two pillars, used to hold up elevated freeway segments. 
2-2

The image in Figure 2.1 shows that the fire developed extremely rapidly, attaining nearly full 

size within minutes of a sudden bright flash from the direction of the freeway, which was 

captured in the video footage and occurs at 3:37 a.m. By the time the camera was turned to 

view the fire, approximately 10 seconds later, flame extended for nearly the full length of the 

lower roadway segment between Bent 18 and Bent 19, as shown in the video capture image in 

Figure 2.1. The east pillar of Bent 19 appears to be engulfed in flame below the level of the I880 roadway, indicating that fuel is spilling off the roadway through the bridge scuppers for 

rainwater run-off near this location. 

By approximately 7 minutes, the fire had reached full size, as illustrated in Figure 2.2. At this 

point, flame entirely fills the space between the upper and lower roadway and is impinging on 

the girders on the underside of I-580. In this image, brush on the ground below the roadway is 

burning vigorously, but the flames around the Bent 19E pillar have self-extinguished. The video 

shows that the fire configuration illustrated in Figure 2.2 persisted with little significant change 

until the collapse of the overhead roadway spans. 

Figure 2.2. MacArthur Maze Fire at +6.8 Minutes (WTP video image at 03:44:31.96 PDT, 

photo from MAIT Report, CHP 2007, reprinted with permission.)
2-3

The sag in the span between Bent 19 and Bent 202 begins to be discernable in the video at 

about 9.3 minutes. Figure 2.3 shows the deep sag in the span at 16.7 minutes, moments before 

total collapse. (This image shows a rare glimpse of Bent 20, illuminated by the ground fires.) 

Figure 2.3. MacArthur Maze Fire at +16.7 Minutes (WTP video image at 03:54:24.61 PDT,

photo from MAIT Report, CHP 2007, reprinted with permission.)

Moments after the complete collapse of the span between Bent 19 and Bent 20, the roadway 

span between Bent 18 and Bent 19 was also visibly sagging, as can be seen in Figure 2.4. In 

this image, fire is no longer visible to the right of Bent 19 in the region of the fallen roadway. To 

the left of Bent 19, the sagging portion of the span between Bent 18 and Bent 19 is intruding 

into the fire, significantly affecting flame shape and distribution.

 2 Bent 20 is generally not visible in the video images, mainly because of the camera angle; it is located to 

the right of Bent 19, from the perspective of the WTP.
2-4

Figure 2.4. MacArthur Maze Fire at +19.8 Minutes (WTP video image at 03:57:33.19 PDT,

photo from MAIT Report, CHP 2007, reprinted with permission.)

The slow gradual descent of the Bent 19 end of the second I-580 span to the lower roadway is 

clearly shown in the WTP video. The partial collapse of this span appears to be complete by 

37.3 minutes into the fire, as illustrated in Figure 2.5. This image shows that the fire has selfextinguished in the region between Bent 18 and the point where the upper span contacts the 

roadway. The fire is confined to a relatively narrow region near Bent 19, between the ends of 

the two fallen I-580 spans.
2-5

Figure 2.5. MacArthur Maze Fire at +37.3 Minutes (WTP video image at 04:15:00.02 PDT,

photo from MAIT Report, CHP 2007, reprinted with permission.)

The remainder of the WTP video shows that after the two upper roadway spans reached their 

final collapsed configurations, the fire continued to burn until the available fuel supply was 

consumed. Figure 2.6 shows an image of the fire at +72.3 minutes, and there is almost no 

change in brightness or configuration of the fire compared to the image in Figure 2.5 at 

+37.3 minutes. This behavior is observed on the WTP video until essentially the end of the fire. 

Figure 2.7 shows the relatively abrupt end of the fire, with an image captured at approximately 

+107 minutes.
2-6

Figure 2.6. MacArthur Maze Fire at +72.3 Minutes (WTP video image at 04:50:00.33 PDT, 

photo from MAIT Report, CHP 2007, reprinted with permission.)

Figure 2.7. MacArthur Maze Fire at +107.3 Minutes (WTP video image at 05:25:57.00 PDT, 

photo from MAIT Report, CHP 2007, reprinted with permission.)
2-7

The WTP video shows that the fire characteristics changed significantly over the duration of the 

fire. Initially, there was a large, intense, openly burning fire because of the fuel spill. The 

relatively straightforward “open pool” nature of this portion of the fire was complicated by the 

presence of the I-580 roadway overhead, which partially confined and channeled the hot fire 

gases rising from the fire along the steel support girders on the underside of the upper roadway. 

This caused the fire to behave more like a very well ventilated tunnel fire than a classic open 

pool fire, and probably led to higher fire temperatures than would typically be obtained in an 

open pool fire of similar size. 

The complete collapse of the overhead span between Bent 19 and Bent 20 at approximately 

17 minutes, and the protracted partial collapse of the span between Bent 18 and Bent 19, 

beginning at about 18 minutes, resulted in a significant reduction in the overall fire size, while 

also substantially changing the configuration of the steel and concrete structures affected by the 

fire. By +18 minutes, the fire was extinguished on the I-880 roadway past Bent 19 in the 

direction of Bent 20, because of the fallen I-580 span. Beyond the contact point of the second 

partially fallen I-580 span, in the direction of Bent 18, the fire on the lower roadway was greatly 

reduced in size and intensity, and was completely extinguished in this region by +37 minutes 

into the fire. From that point on, the fire was confined to a narrow strip approximately 12-13 ft 

(5 m) wide along the roadway between the Bent 19 East and West pillars.

The WTP video shows very clearly what happened during the fire and provides evidence that 

can be used to bound specific features of the fire, but it cannot provide the detailed temperature 

information needed for analytical evaluations of the potential consequences of the fire. The 

following section discusses the results of evaluations to determine the temperatures reached by 

materials exposed to the MacArthur Maze fire.

2.2 Maximum Material Temperatures in the MacArthur Maze Fire

The dramatic failure of the steel girders supporting the upper I-580 roadway spans in the 

MacArthur Maze fire prompted early speculation in the media that the fire could have produced 

temperatures in excess of 3000°F (1650°C), the melting point of steel. This speculation failed to 

take into account two crucial factors: the maximum temperatures achievable in an open 

hydrocarbon-fueled pool fire, and the temperature-dependent nature of the strength of structural 

steel. Although there was considerable evidence of fire damage to the roadway structures in 

the post-fire investigations and clean-up, no evidence of melting of steel was found at any 

location.

Based on experimental and analytical evaluations of large pool fires (Society of Fire Protection 

Engineers 2008), a consistent estimate of the bounding flame temperature for these types of 

fires is approximately 1832°F (1000°C). Higher temperatures may be achievable if the fire is 

confined in a manner that does not restrict the flow of oxygen to the fire or remove significant 

heat from the fire by means of conduction or evaporation. However, the upper limit is only 

about 2462°F (1350°C), based on tunnel fire testing (Brekelmans et al. 2008; NFPA 2008). 

Section 3.0 presents a discussion of the results of modeling the open pool, pre-collapse portion 

of the MacArthur Maze fire using CFD with the Fire Dynamics Simulator (FDS) code (McGrattan 

et al. 2008), developed at the National Institute of Standards and Technology (NIST). The FDS 

analysis (as discussed in Section 3.4) predicts fire temperatures in the range 1472-1922°F (800-

1050°C), with an overall peak of 1994°F (1090°C), at a location 3.3 ft (1 m) above the lower 

roadway surface and fully engulfed in the fire. These results show that a bounding value of 

2012°F (1100°C) is a reasonable and conservative estimate of the flame temperature for the 
2-8

open burning portion of the MacArthur Maze fire. Temperature estimates based on evaluation 

and testing of material samples obtained from the damaged roadway structures are somewhat 

lower than this bounding value.

These physically based bounding values for the fire temperatures show that the steel girders 

could not have melted at any point during the fire, and no evidence of melted steel was found in 

the post-fire evaluation and clean-up. An estimate of the approximate range of temperatures 

that the structures of the highway spans and the tanker truck were exposed to was obtained by 

evaluation of physical samples taken from the accident site (NRC 2008). These included 

samples of the structural steel girders, welds, surface paint and paint flakes, and selected 

materials from the remains of the tanker truck. Figure 2.8 illustrates the locations on the fallen 

roadway spans where physical samples were obtained for this evaluation. (Sample numbers 

correspond to numbering in Table 2.1.) Figure 2.9 shows the remains of the tanker truck at the 

accident site, labeling some of the few recognizable components from which material samples 

could be obtained.

Figure 2.8. Approximate Locations of Collected Specimens for Materials Evaluation of Effects 

of MacArthur Maze Fire (photo from MAIT Report, CHP 2007, reprinted with 

permission.)
2-9

Figure 2.9. Approximate Locations of Collected Specimens Obtained from the Remains of the 

Tanker Truck Following the MacArthur Maze Fire (photo from MAIT Report, CHP 

2007, reprinted with permission.)

Based on analysis of temperature-dependent physical changes in the materials examined, the 

maximum steel temperatures were estimated to be in the range 1796-1868°F (980-1020°C). 

Table 2.1 lists temperatures estimated from the condition of specific samples. These results 

indicate that material temperatures were generally below 1832°F (1000°C), and varied 

significantly with location in the fire. For example, there were unmelted segments of the 

tanker’s aluminum tank, and only partial melting of at least one of the truck’s aluminum wheels.

Table 2.1. Material Temperatures Estimated from Evaluation of Samples

Sample ID

Estimated

Fire Exposure 

Temperature

(°F) (°C) Description of Sample Location

NRC 9 392 200 plate girder 12 with stiffener, near Bent 18 (no evidence of 

overheating)

1292 700 from paint damage 10 ft from NRC-9

NRC 3 1652 900 Girder 5 (showing significant distortion of stiffener plate)

NRC 4 1562 850 Girder 5 (minimum temperature for sample)

NRC 5 572 300 Box Girder 7 lower plate with side and weld (portion of beam that 

fell on tanker truck)

NRC 7 752 400 rivet head in Box Girder 8; orange discoloration of paint

NRC 1S 1472 800 Plate Girder 3 with stiffener near Bent 19; with welds

NRC 10 752 400 flakes peeled off plate girder angles on Box Beam Cap 8 

Truck S-13 1328 720 radiator (aluminum screen)
2-10

Table 2.1. (continued)

Sample ID

Estimated

Fire Exposure 

Temperature

(°F) (°C) Description of Sample Location

Truck S-6 1382 750 3 large bolts on frame near engine

Truck S-11 1292 700 brass fitting on engine

Truck S-12 1657 903 bolt on engine, passenger side (includes steel wire and melted 

aluminum)

Truck S-8, 9, 

10 1094 590 copper (grounding strap on frame, battery cable, electrical system)

Truck S-17 1000 538 stainless steel mirror support bracket

Truck S-7 932 500 bolt on frame

Truck S-15 1328 720 aluminum tank section

1058 570 unmelted tank segments

The material evaluations suggest that the steel girders experienced maximum temperatures in 

the range of 1472-1652°F (800-900°C) at locations where flames directly impinged on the 

girders. This is not sufficient to melt steel, but exposure to these temperatures would 

significantly reduce the strength of the load-bearing girders. Figure 2.10 shows the extremely 

rapid drop in yield strength beginning at about 800°F (427°C) for A36 steel (equivalent to the 

type of steel in the girders3 supporting the I-580 roadway). The yield strength of this material at 

the estimated maximum temperatures experienced during the fire is less than 20% of its normal 

room-temperature value. With such a reduction in strength, the girders could not support the 

overhead spans.

 3 According to Caltrans reports 580 Damage – Preliminary Report (5-3-07) and 580 Damage –

Supplementary Report (5-15-07), (Caltrans tracking number CAL0422, from the Office of Structural 

Materials), as-built plans from the seismic retrofit project in 1994 indicate that the steel girders are ASTM 

A709 (Grade 50). The original as-built plans from 1953 specify ASTM A7. OSM tested girder samples to 

“ASTM A7 Spec.” The ASTM web site (www.astm.org) shows A7 as “withdrawn 1967, replaced with 

A36/A36M.”
2-11

Figure 2.10. Typical Structural Steel Yield Strength Variation with Temperature (plot based on 

data from Brockenbrough and Merritt 1999.)

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

1.1

0 200 400 600 800 1000 1200 1400 1600 1800 2000

Temperature (oF)

ratio of yield strength at high temperature 

to value at room temperature

A36 steel

estimated maximum temperature on steel girders in MacArthur Maze fire scenario

3-1

3.0 NUMERICAL MODELING OF THE

MACARTHUR MAZE FIRE

This section describes numerical simulations of the MacArthur Maze fire using the FDS code, a 

computational fluid dynamics model developed specifically to study fire behavior (McGrattan 

et al. 2008). Various versions of the software were used during the development of the model. 

The final calculations were performed with FDS version 5.4.3.

A preliminary model of this fire was developed for NRC at the Center for Nuclear Waste 

Regulatory Analyses (CNWRA), Southwest Research Institute, San Antonio, Texas under 

contract NRC-02-07-006, and provided an initial scoping analysis of the fire. The model was 

then refined and final calculations were performed by NIST.

3.1 FDS Model Geometry

The computational domain of the model was initially defined as 131 ft wide by 197 ft long by 

59 ft high (40 m by 60 m by 18 m). In subsequent refinement of the model, the height was 

doubled to 118 ft (36 m). Eight mesh blocks consisting of square cells 1.64 ft (0.5 m) or 0.82 ft 

(0.25 m) on a side divide the volume into uniform sub-volumes. Figure 3.1 shows the overall 

geometry of the model, in relation to the roadway configuration. Figure 3.2 and Figure 3.3

illustrate the simplified representation of structural components, including the roadways.

Figure 3.1. Simplified Model Geometry (photo from MAIT Report, CHP 2007, reprinted with 

permission.)
3-2

Figure 3.2. Transverse View of Model Geometry

Figure 3.3. Longitudinal Views of Model Geometry

The location and size of the major structural components (e.g., girders and columns) were 

based on a review of documentary photographs and drawings, primarily from the MAIT report

(CHP 2007). The dimensions of the roadway components were approximated to within 1.64 ft 

(0.5 m), due to of the relatively coarse numerical grid used in the model to solve the governing 

flow equations. The base model used 1.64 ft (0.5 m) resolution, and sensitivity studies were 

performed with a mesh resolution of 0.82 ft (0.25 m). All of the steel structural components 

were modeled with a 0.4 inch (1 cm) thickness at the surface, to allow the FDS code to provide

estimates of surface temperatures of these components. As a conservatism in the fire 

modeling, the FDS code does not evaluate heat transfer within the structural interior of these 

structures, and therefore does not remove heat from the fire that would in reality be absorbed by 

them. A more detailed heat conduction calculation for the steel girders was performed with a 

finite-element model, as part of the development of appropriate boundary conditions for analysis 

of the structural collapse of the overhead spans. This analysis is discussed in Section 6.4.1.
3-3

The concrete roadways were represented in the FDS model as slabs 1.64 ft (0.5 m) thick. Each 

roadway included concrete side barriers modeled as 3.3 ft (1 m) high by 1.64 ft (0.5 m) thick. 

The tanker truck shown in the model diagrams was represented as a simple obstruction. The 

contribution to the fire of its combustible interior was considered negligible in comparison to that 

of the spilled gasoline, and was not included explicitly in the total heat release. In the model 

simulation, the fire was started by assuming that the truck was initially burning and radiating 

thermal energy to the road surface, which was covered with gasoline over the area defining the 

fuel pool. The fuel was assumed to begin burning when the local roadway temperature reached 

a specified threshold value of 140°F (60°C). This resulted in the fire spreading radially from the 

truck location, and encompassing the entire pool area in approximately 3 minutes. This gives a 

reasonable and conservative estimate of the development of the fire, since the WTP video 

shows that the fire reached full size within approximately 6 minutes.

3.2 FDS Model Fire

The WTP video shows that the fire was located mainly between Bent 18 and Bent 19, and 

extended some distance beyond Bent 19 in the direction of Bent 20. The extent of the fire 

across the width of the roadway could not be determined from the video, but post-fire 

examination of the I-880 roadway surface (Wiss, Janney, Elstner Associates, Inc. 2007)

suggests that the fire probably spanned much of the width of the deck, as indicated by extensive 

spalling of the surface concrete. This is shown on the roadway diagram in Figure 3.4(a). Based 

on the spalling patterns and the video images, the modeled fire was assumed to span the full 

width of the roadway and extend from Marker 4 to Marker 12, as illustrated in Figure 3.4(b). 

(a) (b)

Figure 3.4. Maximum Fire Pool Size, Based on Areas of Concrete Spalling on I-880 Surface

(Wiss, Janney, Elstner Associates, Inc. 2007, photo from MAIT Report, CHP 2007, 

reprinted with permission.)

Sensitivity studies were performed in which the area of the fire was increased to twice the 

estimated maximum size, and then decreased to half of this value. Depending on the specified 

burning rate, variations in the fire area changed the overall heat release rate of the fire, but 
3-4

within the burning region, there was essentially no change in the predicted fire temperatures. In 

all cases, the calculations predicted that the upper roadway would be exposed to direct flame 

impingement during the pre-collapse portion of the fire. The results of these analyses indicate 

that within the range of uncertainty in the actual fire area, the characteristic temperatures do not 

change significantly with assumed fire size. 

The total heat release rate for a fire is a function of the burning rate (i.e., mass loss rate, which 

is also a function of the availability of oxygen), fuel properties (including heat of combustion and 

density), and the geometry of the fire. Although the total amount of fuel consumed in the 

MacArthur Maze fire is known, the actual distribution of the fuel during the various phases of the 

fire is not known. Some fraction poured out over the I-880 roadway to create the large pool fire, 

some burned within the partially destroyed tanks of the truck and trailer, and some spilled off the 

roadway through the bridge scuppers to feed the ground fires. The ground fires burned 

vigorously and for much of the total fire duration, resulting in a considerable amount of spalling 

at the base of Column 35L supporting the I-880 roadway (Wiss, Janney, Elstner Associates, Inc.

2007), which is elevated approximately 6.5 m (21.3 ft) above the ground. The burning rate 

would have varied significantly with location, and it is therefore impossible to directly calculate 

the heat release rate for this fire.

The burning rate of the gasoline fire was estimated based on evaluations of burning rate for 

pools formed by realistic fuel spills on concrete1 (Society of Fire Protection Engineers 2008), 

rather than prepared fuel pools in test tanks. The burning rate, or mass loss rate, can be as 

high as 0.045 kg/m2

/s to 0.065 kg/m2/s for a deep gasoline pool fire larger than a few meters in 

diameter, but the depth of the fire pool is an important parameter. Shallow pools formed by 

liquid fuels spilled on concrete burn at approximately 0.01 kg/m2

/s, which is significantly lower 

than the burn rate of a deep pool carefully prepared for a fire experiment. During the MacArthur 

Maze fire, the depth as well as the extent of the pool formed by the gasoline spilling from the 

tanker varied both with location and time, and in addition, some amount of fuel was lost from the 

pool on the roadway as it spilled through the scupper drains. 

Because of these unknowns, the heat release rate for the MacArthur Maze fire can only be 

approximated based on an average burning rate over the known fire duration and estimated fuel 

supply. Table 3.1 summarizes heat release rates for a range of burning rate values 

corresponding to different assumptions about the fire pool configuration. The heat release rates 

are calculated assuming a nominal effective heat of combustion of 43,700 kJ/kg (Society of Fire 

Protection Engineers 2002) for gasoline, which typically has a density of 760 kg/m3

. A value of 

1000 kW/m2 was chosen for the heat release rate specified in the final FDS calculations 

representing the initial 17-minute phase of the fire. This was based on the assumption that the 

depth of the spilled gasoline would have been relatively shallow over most of the roadway. 

Near the truck, however, the spill could have been deeper, resulting in a local burning rate for 

the gasoline that may have been greater than that of the rest of the pool. The value of 

1000 kW/m2 represents a conservative estimate for a large pool that could have deep spots, but 

for the most part is relatively shallow.

 1 Specifically, the chapter “Liquid Fuel Fires”, by D.T. Gottuk and D.A. White, in SFPE Handbook of Fire 

Protection Engineering, 4th ed.
3-5

Table 3.1. Summary of Heat Release Rates for Large Pool Fires

Burning Rate Heat Release Rate Pool Characteristics

0.010 kg/m2/s 437 kW/m2 shallow pool; spill on concrete

0.0229 kg/m2/s 1000 kW/m2 estimated average for shallow pool with some deep spots

0.045 kg/m2/s 2404 kW/m2 deep pool; lower bound for burning rate

0.065 kg/m2/s 2840 kW/m2 deep pool; upper bound for burning rate

The assumed average burning rate of 0.0229 kg/m2

/s for the maximum estimated pool size was 

used for the entire fire duration of 108 minutes in the FDS simulation. This is a conservative 

assumption, since the fire would have consumed the entire 8,600 gallons of gasoline in about 

54 minutes at this burning rate. In addition, the pre-collapse pool fire would have consumed

about 5,900 gallons, or approximately 68% of the gasoline fuel supply, in the first 37 minutes 

(assuming no change in the pool size), leaving only 32% of the fuel for the remaining two-thirds 

of the total fire duration. This indicates that the average burning rate of 0.0229 kg/m2

/s, 

extended over the entire 108 minutes of the fire duration, is a conservative estimate of the 

average burning rate for the MacArthur Maze fire.

3.3 FDS Fire Model Output

The primary purpose of performing the FDS simulations was to determine appropriate 

temperature boundary conditions for evaluating the potential effect of the MacArthur Maze fire 

scenario on an SNF transportation package. A significant output of the fire model for this 

purpose is the quantity referred to as the adiabatic surface temperature (AST). This is a 

potentially misleading term, since the surface in this context is a virtual surface, not an actual 

physical surface in the model. The surface referred to in this term is a hypothetical thermal 

concept defined in fire temperature measurement calculations to represent a perfect, nonintrusive measurement at a specific location within the fire. An AST defines the source 

temperature at a given location in the fire for thermal radiation and convective heat transfer from 

flames and hot gases to actual solid surfaces that see the fire. An AST can be obtained for any 

point in the fire, and is determined in the manner described below. 

The net total heat flux seen by an actual surface exposed to fire is composed of two 

components; thermal radiation and convection. This can be defined simply as:

tot rad con q" = q" + q" (3-1)

where 

q"tot = net total local heat flux

q"rad = local heat flux due to thermal radiation

q"con = local heat flux due to convection

The thermal radiation term in Equation (3-1) is the difference between the absorbed incident 

thermal radiation and that emitted from the surface. The heat transmitted through the surface is 

neglected, and the absorptivity and emissivity are assumed equal, neglecting any dependence 

on wavelength. With these simplifications, the net heat received by the surface as thermal 

radiation can be written as: 
3-6

q" (q" T )

4

rad inc −σ s = ε (3-2)

where

q"inc = incident thermal radiation heat flux

ε = emissivity of the surface 

σ = Stefan-Boltzmann constant

Ts = local surface temperature 

The emissivity (or absorptivity) is a material property of the surface that can be determined by 

measurement. However, in most cases of structural materials exposed to fire, it can be 

assumed that the initial emissivity will change rapidly to a very high value due to sooting of the 

surface. A conservative estimate is 0.9 for highly sooted surfaces. A minimum value of 0.8 for 

absorptivity of exterior surfaces of an SNF package in the HAC fire is specified in 10 CFR 71.

Because fires are characterized by widely varying temperature distributions in space and time, 

the incident thermal radiation heat flux should ideally include all contributions from nearby 

flames, hot gases, and other surfaces. The incident thermal radiation may therefore be written 

as the sum of the contributions from all of the radiating sources:

4

i

i

q"inc = ∑εiFi σT (3-3)

where

q"inc = total local incident thermal radiation heat flux on a given surface from all 

sources

εi = emissivity of the i

th source surface 

σ = Stefan-Boltzmann constant

Fi = dimensionless geometric viewfactor between the local surface and the i

th

source surface

Ti = local surface temperature of the i

th source

FDS includes an algorithm for calculating the incident thermal radiation heat flux using Eq. (3-3), 

based on the local surface temperatures and the geometry of the mesh.

The convective heat flux depends on the difference between the surrounding gas temperature 

and the surface temperature, and on local fluid dynamics. The relationship between heat flux 

and temperature difference is generally characterized with a heat transfer coefficient, which is 

determined from an empirical heat transfer correlation, such that: 

q" h(T T ) con = g − s (3-4)

where

h = local heat transfer coefficient 

Tg = gas temperature adjacent to the exposed surface

Ts = local surface temperature
3-7

Substituting Eq. (3-2) and Eq. (3-4) into Eq. (3-1), the total net heat flux to a surface can 

therefore be expressed as

q" (q" T ) h(T T ) g s

4

tot = ε inc −σ s + − (3-5)

The relationship in Eq. (3-5) can be used to determine the AST at the location of an actual 

surface in the model. The “virtual” surface at this (and any other) location is by definition a 

perfect insulator, and since the total net heat flux to this idealized perfect insulator “surface” is 

by definition zero, Eq. (3-5) reduces to

(q" T ) h(Tg TAST ) 0 4 ε inc − σ AST + − = (3-6)

Numerically, the adiabatic surface temperature is a very useful quantity because it provides a 

natural interface between models that represent fire behavior and models that represent thermal 

and mechanical behavior of structures. A fire model in this context is any calculation method 

used to predict the temperature and species concentrations of a fire-driven flow. A structural 

model is any calculation method used to predict temperatures or stress/strain responses in an 

object exposed to the fire. The fire model may compute the evolving temperature of the 

bounding surfaces out of necessity, but it does not generally include a detailed representation of 

the thermal response of solid objects. Even a computational fluid dynamics model may only 

approximate a bounding solid as an infinitely thick slab for the purpose of estimating its surface 

temperature. 

If the results of the fire model are to be used to perform a more detailed heat transfer calculation 

of the thermal response of a solid object within or near the fire, then some sort of interface is 

required to transfer information at the gas-solid interface. The most obvious quantity for this 

purpose is the heat flux at the surface, but in practice, this leads to major computational 

difficulties. The net heat flux to a surface computed by the fire model is dependent on the 

corresponding surface temperature, which is also computed by the fire model. Depending on 

the model, this surface temperature might not be of the desired accuracy. In addition, it is

common in many popular solid phase heat transfer programs to input a prescribed thermal 

boundary based on external gas temperature and calculated surface temperature (as in Eq. 3-9) 

rather than as a prescribed heat flux. Both of these problems can be circumvented by using the 

adiabatic surface temperature TAST as the intermediary between the fire and structural models. 

The interface is fairly simple. At every surface point at which the fire model computes an 

incident thermal radiation heat flux and a corresponding gas temperature adjacent to that 

surface, the following implicit equation can be solved for the adiabatic surface temperature, 

assuming that the emissivity and convective heat transfer coefficient are effectively constant at 

that location. 

(q" T ) h(Tg, TAST ) 0 4 ε inc,FM − σ AST + FM − = (3-7)
3-8

where

q"inc,FM = incident thermal radiation heat flux computed by the fire model at the 

exposed surface 

Tg,FM = gas temperature computed by the fire model adjacent to the exposed 

surface

TAST = adiabatic surface temperature computed by the fire model

A key feature of Equation (3-7) is that the fire model does not require any assumptions to 

compute the incident thermal radiation heat flux. This equation merely serves as the definition 

of the adiabatic surface temperature, but it does not imply that the fire model calculates the heat 

flux in any particular way. Most importantly, it does not imply that the fire model uses a fixed 

heat transfer coefficient, h. The values of TAST for any location in the fire model can be stored in 

a file according to a user-specified time interval and length increment appropriate for the 

application. 

For the structural model, the heat flux to an object’s surface and its temperature due to the fire 

conditions computed by the fire model can be calculated by the relationship:

q" (q" T ) h(T T ) g, s,SM

4

tot,SM = ε inc,FM − σ s,SM + FM − (3-8)

Subtracting Eq. (3-8) from Eq. (3-7) yields the total net heat flux to the surface of an object as:

q" (T T ) h(T T ) s,

4

s,

4

tot,SM σ AST − SM + AST − SM = ε (3-9)

where

q"inc,FM = incident thermal radiation heat flux computed by the structural model at the 

exposed surface 

Tg,FM = gas temperature computed by the structural model adjacent to the 

exposed surface

TAST = adiabatic surface temperature computed by the fire model

The AST is interpreted by the structural model as an effective black body radiation temperature 

for the purpose of computing the incident thermal radiation at an actual surface in the structural 

model, and as a gas temperature for the purpose of computing the convective heat flux at the 

given surface. The advantage of this approach is that it requires transfer of only one quantity, 

the AST, from a fire model to a structural model, rather than bringing over a heat flux, surface 

temperature, and convective heat transfer coefficient. A side benefit is that the structural model 

need not be reconfigured to accept a heat flux as its boundary condition. It needs only to be 

modified to accept a temporally and spatially varying “exposing” temperature (i.e., the AST), 

which it can use to calculate the heat flux based on that temperature and the surface 

temperature calculated in the structural model. Most models of this type are already configured 

to accept a time-varying “exposing” temperature curve as an external boundary condition. 
3-9

3.4 Fire Model Results for Pre-Collapse Configuration

The FDS analysis using the model described in Sections 3.1 and 3.2 was used to determine the 

fire behavior during the pre-collapse phase of the fire, which lasted approximately 17 minutes. 

This was the hottest and most intense portion of the fire, and would be expected to produce the 

most severe conditions for the thermal analysis of the effect of the fire on a SNF transportation 

package. The upper bound on the peak predicted temperatures from the FDS model during the 

first phase of the fire is in the neighborhood of 2012°F (1100°C). This is illustrated in Figure 3.5

with the predicted ASTs for a “target” object located 3.3 ft (1 m) above the roadway at a point in 

the fire near the final position of the tanker truck. The target is a rectangular block consisting of 

surfaces facing up, down, north, south, east, and west, and the temperature exposure for a 

given surface on this hypothetical object depends on orientation. The different curves of the plot 

are labeled in reference to the direction the surface faces. Figure 3.6 shows a similar plot of 

AST values for the same location at a higher elevation, approximately 3.3 ft (1 m) below the 

overhead girders. 

Figure 3.5. Predicted AST Values 1 Meter above the Lower Roadway Surface

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18

Elapsed Time (minutes)

Temperature (oC)

(above)

(below)

North

South

East

West

bounding value: 1100oC

collapse of Bent 19-20 span
3-10

Figure 3.6. Predicted AST Values 3 Meters above the Lower Roadway Surface (1 meter

below the Upper Roadway)

Based on these CFD analysis results and on material analyses of physical samples taken from 

the steel girders, roadway concrete, and remains of the destroyed tanker truck, (see discussion 

in Section 2.1), the initial portion the fire transient is conservatively bounded by a fully engulfing 

fire with uniform flame temperature of 2012°F (1100°C). The fire emissivity is assumed to be 

0.9, which is characteristic of a sooty, optically dense hydrocarbon pool fire.

3.5 Bounding Assumptions for Post-Collapse Fire

In the MacArthur Maze fire, the collapse of the overhead spans greatly reduced the size and 

intensity of the fire. The abrupt fall of the span between Bent 19 and Bent 20 effectively 

extinguished the portion of the fire extending to the right of Bent 19. The slow descent of the 

Bent 19 end of the second span, which occurred over approximately 20 minutes, introduced this 

structure into the middle of the large pool fire on the lower roadway. The I-580 span eventually 

made contact with the lower roadway surface along a line extending approximately the full width 

of the roadway, at a distance of about 5 m (15 ft) from Bent 19. The effect of this intrusion was 

a fairly rapid decrease in the extent of the fire in the region beneath the descending roadway 

span. Within approximately 20 minutes (from 17 to 37 minutes after ignition), the fire was 

completely extinguished in the direction of Bent 18, and the extent of the fire was reduced to the 

gap between the fallen roadway spans near Bent 19. An estimate of the maximum fire area 

during this second phase is illustrated in Figure 3.7. 

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18

Elapsed Time (minutes)

Temperature (oC)

(above)

(below)

North

South

East

West

bounding value: 1100oC

collapse of Bent 19-20 span
3-11

Figure 3.7. Estimated Fire Pool for Post-collapse Portion of the MacArthur Maze Fire

The area of this region is about one-fifth of the estimated maximum area of the first phase of the 

fire. CFD codes such as FDS generally do not have the capability to significantly alter the 

geometry of the physical structure of the fire environment or make large changes in the physical 

extent of the fire (other than changes caused by consuming available fuel.) The FDS analysis is 

not able to track the changing fire conditions during this portion of the fire, and estimates of the 

bounding fire temperature must rely on alternative methods. 

The intrusion of the partially fallen span into the post-collapse fire would have the effect of 

wicking heat out of the fire, thereby reducing the flame temperature. In addition, the reduction in 

physical extent of the fire is sufficient to categorize this portion of the fire as only partially 

engulfing, which means that a large object (on the order of the size of a small automobile) would 

not be completely immersed in flames. In the parlance of thermal radiation heat transfer, it 

would “see” the colder surroundings. Experimental work has shown (Society of Fire Protection 

Engineers 2002)1 that partially engulfing fires typically range from 1472°F (800°C) to 1652°F 

(900°C). 

The post-collapse portion of the MacArthur Maze fire, therefore, is conservatively bounded by 

assuming a flame temperature of 1652°F (900°C). A further level of conservatism is added if 

the fire is assumed to be fully engulfing for a SNF transportation package. This approach 

neglects the effect of heat absorbed from the fire by the structural mass of the concrete 

roadways, and is a conservative and bounding representation of the fire behavior in this portion 

of the transient.

 1 Specifically, the chapter “Fire Hazard Calculations for Large, Open Hydrocarbon Fires,” by Beyler, in 

SFPE Handbook of Fire Protection Engineering, 2nd ed.
3-12

The appropriate bounding temperature for the fire during the 20-minute transition from a large 

pool fire at 2012°F (1100°C) to a smaller pool fire at 1652°F (900°C) is not readily apparent from 

the information available on the MacArthur Maze fire. As a conservative simplification, this 

transition interval is modeled as if the large open pool fire had persisted for the full 20 minutes 

required for the partial collapse of the Bent 18-19 span, out to 37 minutes of the fire duration. At 

37 minutes, the boundary temperature representing the fire is abruptly reduced from 2012°F 

(1100°C) to 1652°F (900°C), and remains at this value out to 108 minutes, to conservatively 

bound the end of the fire duration. These bounding fire temperatures are illustrated in Figure 

3.8.

Figure 3.8. Estimated Bounding Maximum Fire Temperatures for the MacArthur Maze Fire

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10 20 30 40 50 60 70 80 90 100 110

Elapsed Time (minutes)

Temperature (oC)

bounding flame temperature

abrupt collapse of Bent 19-20 span

Bent 18-19 span reaches final configuration

end of fire
4-1

4.0 THE MACARTHUR MAZE FIRE SCENARIO

There are several different aspects of the MacArthur Maze fire that could expose an SNF 

transportation package to conditions potentially more severe than the HAC fire specified in 

10 CFR 71. These can be summarized as follows:

1. exposure of the package to the large fully engulfing fire prior to the collapse of the overhead 

I-580 roadway span between Bent 19 and Bent 20, which is at a higher engulfing flame 

temperature and longer duration than the HAC fire

2. subsequent exposure of the package to the relatively long duration of the fire following the 

collapse of the overhead spans, which is also at a higher engulfing flame temperature and 

significantly longer duration than the HAC fire

3. physical impact of a collapsing overhead span on the package 

4. post-fire cooldown with the package assumed to be covered by the concrete “blanket” of a 

collapsed overhead span.

As the above list shows, this fire scenario subjects the SNF package to both structural and 

thermal conditions that could adversely affect its performance. To assure that bounding 

conditions were considered for all aspects of this complex scenario, the thermal and structural 

analyses were performed separately. To conservatively bound the worst that the MacArthur 

Maze fire and roadway collapse could do to the SNF package, the scenario selected for 

analysis evaluated the most adverse thermal conditions and the most adverse structural 

configuration. The package was assumed to be positioned in the most adverse location for the 

different portions of the thermal analyses and the structural analyses, without realistic 

constraints on how the package could possibly relocate from one place to another during the fire 

scenario. 

The thermal analyses for this fire scenario, which cover items 1, 2, and 4 of the above list, are 

discussed in Section 4.1. The structural analyses (item 3 above) are discussed Section 4.2.

4.1 Thermal Conditions in MacArthur Maze Fire Scenario 

The thermal analyses were performed assuming that the SNF package would be fully engulfed 

in the fire for the full 108 minutes of the total fire duration. The effects of the physical impact of 

the collapsing span(s) were ignored for this portion of the analysis, for two specific reasons. 

First, the impact of the overhead span on the SNF package would completely smother the fire (if 

the package had been located beneath the first span to fall) or greatly decrease the intensity of 

the fire in the vicinity of the package (if it had been located beneath the second span to fall). In 

the first case, the fire would have lasted less than 20 minutes; in the second case, the fire would 

no longer be fully engulfing after approximately 25 minutes, and the concrete roadway of the 

fallen span would absorb a significant portion of the energy from the fire. In either case, the 

potential thermal load on the package would be greatly diminished. 

For the thermal analysis, the package is assumed to be in the following locations:

• the package is on the lower I-880 roadway, fully engulfed in fire for 37 minutes, exposed to a 

flame temperature of 2012°F (1100°C)
4-2

• after 37 minutes, the package is still on the lower I-880 roadway, fully engulfed in fire, but the 

flame temperature is assumed to drop to 1652°F (900°C) for the remaining 71 minutes of the 

smaller post-collapse fire, resulting in a total fire exposure duration of 108 minutes

• after 108 minutes of fire exposure, the package is still on the lower I-880 roadway, but is 

enclosed in a concrete ‘tunnel’ simulating the collapsed roadway, which is cooled only by 

natural convection from the exposed concrete surfaces of the upper and lower roadways.

4.2 Structural Loads in MacArthur Maze Fire Scenario

The primary structural load considered in the MacArthur Maze fire scenario is the impact of a 

section of the upper roadway falling onto the package. If a tractor-trailer rig carrying a legal 

weight truck (LWT) SNF package were to be involved in an accident of the severity of the 

MacArthur Maze fire scenario, the package could be subjected to additional impact loads, such 

as collisions with other vehicles caught in the fire or impact with road barriers or debris. 

Potential scenario-dependent impact loads of this nature are not considered in this study. 

Regardless of the path the package might have to take to reach the postulated locations within 

the MacArthur Maze fire scenario, the loads it would be subjected to are assumed to be minor 

and within the realm of regulatory accident scenarios that provide the design bases for all SNF 

transportation packages. 

SNF transportation packages are designed to survive and maintain containment under severe 

impact scenarios. The HAC drop scenario, as specified in 10 CFR 71, is a 30-ft drop onto an 

unyielding surface, with the package impact occurring at the most adverse orientation. The 

large mass1 of a typical SNF package leads to a relatively large amount of kinetic energy at 

impact, which would be much greater than the kinetic energy that could be imparted to the 

package by the falling overhead section. On that basis alone, the falling overpass section would 

not be expected to be a more damaging scenario than the HAC drop scenario. However, the 

falling overhead section impact is postulated to occur at elevated temperatures, which reduce 

the package material strength, compared to the assumptions for the HAC drop scenario (which 

is assumed to occur at normal operating conditions). To evaluate the package performance in 

the more challenging thermal environment of the MacArthur Maze fire scenario, a range of 

postulated package locations and orientations on the lower roadway beneath the falling 

overhead section were selected for analysis. The goal was to select a limited number of 

analyses that would bound potential package response in this scenario. 

A realistic location for the package to receive the maximum impact force from the collapsing 

overhead span would be near the edge of the large pool fire, or possibly outside the fire pool 

entirely. That is, a location that would result in maximum impact loading and post-fire blanketing 

by the fallen overhead roadway would be a location likely to receive minimum fire exposure. 

Conversely, if the package were positioned to receive maximum fire exposure (i.e., fully 

engulfed for both the pre-collapse and post-collapse fire conditions) it would have to be located 

near the middle of the area encompassed by the smaller post-collapse fire pool (see Figure 

3.7), where it could not be struck at all by either of the two collapsed spans.

As a bounding assumption, the peak temperatures predicted in the thermal analysis for the fully 

engulfing 2012°F (1100°C) fire conditions (see Section 4.1 above) were imposed on the 

package in the structural analysis. The package was positioned at a location where it would 

 1 Total weight of the GA-4 package when fully loaded is 55,000 lb (27.5 T).
4-3

receive the maximum force of impact from the collapse of the I-580 overhead span between 

Bent 19 and Bent 20. As a further conservatism, the package was positioned on the lower 

roadway, rather than on the truck conveyance transporting it. There is no credible means in this 

accident for the package to have become separated from its conveyance, but the I-880 roadway 

provides the stiffest possible platform for the package when it receives the impact force of the 

collapsing overhead span. If the package were on the conveyance, much of the force of the 

impact would be absorbed by the compliance of the overall system. In addition, the impact 

forces would be smaller than those obtained with the package on the roadway, since the freefall distance of the collapsing span before impact with the package on the conveyance would be 

reduced by about 30%.

The package is assumed to be at the maximum temperature obtained from the thermal analysis 

after 37 minutes of exposure to the fire. The temperature of the descending roadway span (in 

particular, the temperature of the steel girders) is estimated from a separate thermal analysis of 

the upper roadway response to the fire, and corresponds to the conditions at 18 minutes into the 

fire. This disparity between the time-stamp for the package and the roadway girders is to obtain 

conservatively bounding values for each component. For the package, the hotter it is at the 

time of impact, the more vulnerable the steel and other material would be to deformation or 

damage. It is therefore assumed that the package has the longest possible time to heat up (i.e., 

37 minutes) before the overhead span falls on it. For the roadway girders, the cooler they are at 

the time of impact, the stiffer they would be, and the more damage they could potentially do to 

the package. It is therefore assumed that the falling roadway has the shortest possible time to 

heat up (i.e., 18 minutes) before falling on the package. This timing difference could not occur 

in reality, but as a modeling assumption, it yields a conservative bound for the possible range of 

behavior for both components.

The orientation of the SNF package with respect to the falling roadway determines how many of 

the girders actually strike the package, where they strike, and at what angle, and therefore has 

a significant effect on the potential consequences of the impact. The spacing of the girders 

relative to the length of the package is such that no more than two girders could strike the 

package in any given configuration. Possible orientations include:

• a single girder striking along the full axial length of the package

• a single girder striking across the package, perpendicular to the long axis

• two girders striking across the package at an angle, with one impact near the package lid.

It is not obvious which impact orientation could do the most damage to the package, and the 

analysis considers the full range of possibilities, including impacts on the package lid flange 

region. The analysis also evaluated the effect of a girder striking directly onto one of the 

package lifting trunnions, which would impart an extremely localized load to the package body. 

This requires the highly improbable configuration of the package standing on the roadway on 

one side pair of its lifting trunnions, to obtain a geometry in which such an impact could actually 

occur. 

The assumptions defining components of the package model can affect the magnitude of the 

loads that are imparted to the package by the accident scenario. To maximize the impact load

on the stainless steel body of the package, the effect of impact limiters was neglected in all 

cases, even though the impact limiters are not expected to detach from the package during the 

accident. Similarly, the thin outer shell and liquid contents of the neutron shield are neglected, 

on the assumption that the liquid would be gone by the time the overpass collapsed and that the 

thin steel plates of the neutron shield would absorb only a negligible amount of impact energy 
4-4

through deformation. In addition to conservatively neglecting energy absorption mechanisms, 

modeling choices were made to simplify the impact model and reduce the computational cost as 

much as reasonably possible. These modeling choices result in the falling overpass structure 

making direct contact with the stainless steel body of the package or its trunnions, rather than 

being first intercepted by the impact limiters and neutron shield tank. Detailed descriptions of 

the overpass and package structural models are presented in Section 5.4.1 and Section 5.4.2,

respectively. Section 6.4 describes the assumptions and analysis method in more detail.

In addition to the overpass and package impact modeling, structural evaluations were

performed for key package bolts under thermal expansion loading. The differences in thermal 

expansion coefficients of the package structural materials and the fastener bolts lead to 

increased bolt tension, even under uniform temperature loading. Temperature results from the 

thermal models were used to develop temperature load histories for the bolts and surrounding 

material. These bolt evaluations are assumed to be independent of the impact events. In all 

credible load cases, the key temperature states occur much later in the transient than the 

overpass collapse at 37 minutes. Section 5.4.3 discusses the bolt thermal expansion models in 

more detail.
5-1

5.0 ANALYTICAL MODELS FOR THE MACARTHUR MAZE

FIRE SCENARIO

This section describes the analytical models developed to investigate the potential effects of the 

MacArthur Maze fire scenario on a typical over-the-road spent fuel transportation package

design. This analysis evaluates the transient thermal response of the package and roadway 

structure from the initial steady-state conditions through the fire scenario, and many hours into 

the post-fire cooldown. The models appropriately capture the thermal inertia of the SNF 

package and the transient temperature response of the system. 

Because of the severity and complexity of this fire scenario, multiple thermal and structural 

models were developed. This included two independent thermal modeling approaches using 

the finite element analysis (FEA) code ANSYS and the finite difference thermal-hydraulics code 

COBRA-SFS and structural models using ANSYS and LS-DYNA. The ANSYS thermal model 

represented the complete package, including the impact limiters. The COBRA-SFS thermal 

model was developed to evaluate the potential effects of loss of the impact limiters, and to 

provide a detailed, best-estimate evaluation of the thermal response of the fuel rod cladding to 

this fire scenario. The structural models were used to evaluate the potential effects of the 

impact of the falling overhead span on the package, and assumed a range of possible 

orientations of the package on the lower roadway. Additional detailed models of the bolted lid 

and flange were also developed, to evaluate potential effects on the package containment 

boundary as a consequence of the thermal and structural loading imposed by this fire scenario.

The basic design of the package selected for this analysis is described in Section 5.1. The 

models representing this package for analysis with ANSYS and COBRA-SFS are presented in 

Sections 5.2 and 5.3, respectively. Structural models developed using LS-DYNA and ANSYS 

are presented in Section 5.4.

5.1 GA-4 Legal Weight Truck Spent Fuel Shipping Package

The General Atomics GA-4 LWT transportation package was selected for this investigation to 

evaluate the potential effects of an accident of the magnitude and severity of the MacArthur 

Maze fire on an NRC-certified SNF transportation package. This package can carry a relatively 

large payload for an over-the-road transportation package, and therefore the potential 

consequences of package failure could be more severe than for packages with smaller payload 

capacities. The GA-4 package is designed to transport up to four intact pressurized water 

reactor (PWR) spent fuel assemblies with a maximum decay heat load of 2105.4 Btu/hr 

(0.617 kW) per assembly, for a total package decay heat load of 8423 Btu/hr (2.468 kW). 

The GA-4 can carry zircaloy-clad UO2 fuel with maximum initial enrichment of 3.15% 235U, in 

14x14 assemblies with maximum average burnup of 35 GWd/MTU (minimum cooling time of 

10 years), or 15x15 assemblies with maximum average burnup of 45 GWd/MTU (minimum 

cooling time of 15 years). This package is not licensed to carry high burnup fuel (i.e., fuel with 

average burnup greater than 45 GWd/MTU). There are packages permitted to carry high 
5-2

burnup fuel pins, but their contents are less than a complete fuel assembly. In addition, 

transportation of high burnup fuel (>45 GWd/MTU) by road is currently evaluated on a case-bycase basis, pending development of general guidance1

. 

For the purpose of this analysis, the package was assumed to contain four WE 14x14 PWR 

spent nuclear fuel assemblies at the maximum decay heat load. This is the limiting design basis 

configuration for thermal analysis of the package. Figure 5.1 shows an exploded view of the 

package, illustrating the main design features. The payload capacity is 6648 lb (3015 kg), and 

the fully loaded package weighs approximately 55,000 lb (24,948 kg). The package 

containment boundary is provided by the following structures:

• stainless steel package body wall 

• stainless steel bottom plate 

• stainless steel package closure lid secured by Inconel fasteners 

• dual O-ring seals for the closure lid, gas sample port, and drain valve. 

The stainless steel package body encloses the gamma shield, which consists of an inner shell 

of depleted uranium. Neutron shielding is provided by a stainless steel neutron shield tank 

external to the package body, containing a water/propylene glycol mixture. Aluminum 

honeycomb impact limiters, completely enclosed in a thin stainless steel outer skin and inner 

housing, are attached to each end of the package. Configuration details, including design 

drawings, are provided in the safety analysis report (SAR) for this transport package (General 

Atomics 1998).

Figure 5.1. GA-4 Package: Exploded View (General Atomics 1998)

 1 Transportation of high-burnup fuel is specifically addressed in Revision 2 of NRC Interim Staff Guidance 

11 (ISG-11, Rev. 2). A summary of current status of this issue is provided in the Electric Power Research 

Institute (EPRI) report Transportation of Commercial Spent Nuclear Fuel, Regulatory Issues Resolution, 

EPRI, Palo Alto, CA. 2010. 1016637.
5-3

5.2 ANSYS Model of GA-4 Package

A detailed three-dimensional representation of the GA-4 package was constructed using 

ANSYS® (ANSYS 2003). To simulate the effects of the different segments of the MacArthur 

Maze fire scenario (as described in Section 4.0), the ANSYS model also includes elements 

representing a segment of the lower roadway beneath the package and the collapsed upper 

roadway, which is assumed to cover the package at 37 minutes into the fire scenario. However, 

the model conservatively neglects the thermal effect of direct contact between the package and 

the roadway. During the fire, the package is treated as fully engulfed in flame. The model 

considers forced convection due to the hot fire gases flowing past the package, and thermal 

radiation to the optically dense fire environment. During the post-fire cooldown, the model 

includes thermal radiation from the outer surfaces of the package to the roadway surfaces, and 

conduction-only heat transfer to the surrounding air. The roadway surfaces were included to 

evaluate the potential effect of the collapsed roadway on the rate of cooldown of the package 

after the end of the fire. (A detailed description of the thermal boundary conditions for this multistep scenario is presented in Section 6.3.)

Section 5.2.1 describes the detailed ANSYS model of the GA-4 package. Section 5.2.2

describes the approach used to represent the package buried under the collapsed roadway in 

the post-fire analysis. Section 5.2.3 presents the material properties used to represent the 

different elements of the package in the fire and post-fire cooldown transients.

5.2.1 GA-4 Package Representation

The package is assumed to be oriented horizontally throughout the fire scenario, including the 

actual fire duration, for maximum heat input into the package from the fire. The horizontal 

orientation of the package beneath the fallen roadway also represents the most adverse 

conditions for heat removal from the package in the post-collapse fire environment. The 

conveyance carrying the package is omitted from the model as a conservative representation 

for both the thermal and structural modeling of this fire scenario. In the thermal model, the fire 

is treated as fully engulfing, such that the package is subjected to a uniform bounding flame 

temperature in all directions. In effect, the package is treated as suspended in the fire, and 

thermal effects of contact with the surface of the road (e.g., heat conduction losses and potential 

thermal shielding of portions of the package) are neglected. Including the conveyance in a 

realistic manner would have the effect of partially shielding the package from the fire burning on 

the roadway below. These assumptions constitute a significant conservatism in the overall 

modeling approach, since the conveyance and the roadway beneath the package could provide 

substantial limitations on the rate of heat deposition to the package in this fire scenario.

The model geometry was developed from engineering drawings provided in the SAR for the GA4 package (General Atomics 1998). Table 5.1 summarizes the ANSYS model element types 

used for the various components of the package and surrounding roadway. The structure of the 

package is represented in fine detail, including the lifting trunnions and impact limiters. 

Convection and thermal radiation heat transfer is represented for specific interior and exterior 

surfaces, including thermal radiation between the outer surfaces of the package and the 

external environment. During the fully engulfing fire scenario, the package sees only the 

bounding fire temperature. During the post-fire cooldown, thermal radiation exchange is 

calculated between the package and the enclosing roadway concrete. Surface elements were 
5-4

also generated along the exterior of the package to account for solar insolation loads to 

calculate the normal conditions of transport, which defines the initial temperature distribution for 

the package.

Table 5.1. Summary of Elements in ANSYS Model of GA-4

Number of 

Elements Element Type Modeled Structure(s) or Connections

1,851,067 SOLID70 8-node 

brick elements

fuel assembly, fuel spacer, fuel support structure (FSS) inner frame, 

helium gap, FSS liner, DU gamma shield, package body, neutron 

shield, stiffener ring, impact limiter support structure, outer shell, 

trunnion assembly, closure assembly, and honeycomb structure of 

the impact limiters

45,240

SHELL57 4-node 

quadrilateral 

thermal elements

exterior surface of the impact limiters

761

LINK33 3-D 

conduction bar 

elements

package closure bolts; impact limiter attachment bolts

25,331 CONTA173 contact 

elements connecting impact limiters, closure assembly, and lifting trunnions to 

appropriate package assembly surfaces 27,893 TARGE170 contact 

elements

232,980 SURF152 elements convective heat transfer and solar insolation loads at the outer 

surfaces of the package

218 MATRIX50 

elements

radiative heat exchange between internal package surfaces, and 

between the external surfaces of the package and the environment

A cross-sectional view of the ANSYS model is shown in Figure 5.2, with the major components 

of the GA-4 package indicated. All components illustrated in Figure 5.2 were modeled using 

brick elements. The square blocks shown in red are homogeneous regions representing the 

four fuel assemblies within the package. The fuel assemblies are contained within the cruciform 

stainless steel fuel support structure (FSS) and FSS liner. The helium gas in the gaps between 

the homogenized fuel assembly regions and the FSS plates was explicitly modeled with solid 

elements. The model includes a composite representation of the layers of the cruciform inner 

frame of the FSS, which consists of thin sheets of stainless steel enclosing boron carbide rods. 

The thin steel of the FSS liner is represented with a single layer of nodes (illustrated in light blue 

in the diagram in Figure 5.2).

The GA-4 gamma shield (represented by three layers of elements illustrated in multiple colors in 

Figure 5.2), consisting of a rectangular tube of depleted uranium (DU), encloses the FSS liner. 

The DU gamma shield is in non-loadbearing contact with the square cross-section of the FSS

liner, and has rounded outer corners, in order to fit within the cross-sectional geometry of the 

steel package body. The rectangular stainless steel package body forms the inner surface of 

the liquid neutron shield (NS) tank. The liquid neutron shield tank contains a 56% propylene 

glycol/water mixture that is modeled as a solid material using the elements shown between the 

steel package body and the outer wall of the tank. The outer wall of the NS tank is a thin 

cylindrical stainless steel shell, and is represented in the model as a single layer of elements, as 

shown in Figure 5.2. This layer constitutes the outer surface of the package assembly. 
5-5

Figure 5.2. Cross Section of ANSYS® Model of GA-4 Package Near Midplane

The diagram in Figure 5.2 shows a cross-section of the package near the center of the axial 

length of the cask cavity. In this region, there is only liquid in the region between the cask body 

and the NS tank outer shell. At either end of the package, in the regions covered by the impact 

limiters, the NS tank is structurally supported by 36 radially distributed stainless steel ribs

designated as the impact limiter support structure (ILSS). These ribs extend radially from the 

thick steel shell of the package body to the thin outer stainless steel shell, and provide additional

pathways for conduction heat transfer from the cask body to the NS tank outer shell, in addition 

to structural support. This region is explicitly modeled in detail in the ANSYS model, but for 

clarity is omitted from the diagram in Figure 5.2. 

A slice through the long axis of the model is illustrated in the diagram in Figure 5.3, and shows

the modeling of the ends of the package, including the impact limiters, which consist of an 

internal aluminum honeycomb structure enclosed within a stainless steel skin. The stainless 

steel shell of each impact limiter was modeled with shell elements. All other components were 

modeled using brick elements. A detailed representation of the model in the region of the top 

impact limiter and package closure is illustrated in Figure 5.4. This diagram shows the impact 

limiter stainless steel skin and a thin air gap between the impact limiter and the external surface 

of the package. This gap, which conservatively accounts for the tolerance of the fit of the 

impact limiter onto the package, was represented in the model geometry using SOLID70 brick 

elements.
5-6

Figure 5.3. GA-4 Package Geometry, Including Impact Limiters

Figure 5.4. GA-4 Package Geometry Model: Impact Limiter Details

For this analysis, the air gap between the closure lid and the impact limiter steel liner was 

assumed to remain open during the fire and in the post-fire cooldown, even though deformation 

impact limiter honeycomb NS outer shell

 package closure

 package cavity

 FSS steel liner

 DU gamma shield

 stainless steel body wall

 NS 

liquid
5-7

or warping of the impact limiter in response to the fire conditions could potentially reduce or 

eliminate this gap. During the fire, this assumption would tend to slow the rate of heat input to 

the package through the impact limiters, but because very little heat from the fire can enter the 

package through the highly insulating material of the impact limiters, this assumption would be 

expected to have a negligible effect on the thermal response of the package. In the post-fire 

cooldown, however, this assumption would tend to slow the rate of heat removal from the 

package, by increasing the insulating effect of the impact limiters. 

The thermal inertia of an SNF package can result in significantly higher temperatures being 

reached on some components in the post-fire cooldown, compared to temperatures reached 

during the fire, particularly for temperatures in nominally cooler regions of the package. It was 

therefore deemed more important to capture the effect of retaining the air gap throughout the 

fire scenario, particularly since heat transfer in the package end regions would not be expected 

to affect the peak component temperatures during the fire, which occur near the package 

midplane, due to direct heat input from the fire.

The lower end of the package consists of a thick stainless steel bottom plate welded to the steel 

inner and outer walls of the package. The upper end of the package is sealed with a stainless 

steel closure assembly that attaches to a stainless steel flange on the steel body wall. Figure 

5.5 shows the detailed representation of the closure assembly developed for this model. 

Helium-filled gaps between the closure assembly and the FSS, and between the stainless steel 

flange and the closure assembly, were included in the model geometry. These gaps were 

represented with solid brick elements.

Figure 5.5. GA-4 Package Geometry Model: Closure Assembly Details

Sections 6.2 and 6.3 provide detailed discussions of the modeling assumptions and boundary 

conditions for the fire analysis. The detailed representation of the package internals was 

designed to capture all three possible modes of heat transfer (i.e., conduction, convection, and 

thermal radiation) between all of the components of the model. Conduction is handled 

inherently in ANSYS by the elements and corresponding material properties representing each 

component, but convective and radiative mechanisms must be carefully implemented to 
5-8

properly capture the physical behavior of the system. The representation of the fuel assemblies 

is particularly important in appropriately modeling the thermal response of the fuel rods and 

predicting the peak cladding temperature. Heat transfer within the fuel assemblies is primarily 

by conduction and thermal radiation, with convection only a relatively minor contributor.

The fuel assemblies were modeled as homogeneous regions with an effective radial 

conductivity determined using an effective conductivity model (Bahney and Lotz 1996) that is 

widely used in the nuclear industry in safety analysis for SNF packages. In this model, the 

combined effect of thermal radiation and conduction is characterized using an effective 

conductivity that is a function of assembly geometry and decay heat. The application of the fuel 

effective conductivity model developed for this analysis introduces a modification to more 

accurately account for the temperature gradient between the outermost row of rods in the 

assembly and the enclosing wall. This is accomplished by including a helium gap between the 

homogenized material region representing the fuel assembly and the wall of the enclosing

basket (in this case the FSS cruciform and liner, as shown in the diagram in Figure 5.2), rather 

than extending the homogenous region to the wall, as is the approach normally used in the 

effective conductivity model. An additional feature of this modified representation is that it more 

directly takes into account the effect of the non-uniform wall temperature distribution around the 

fuel assembly, which can be of particular significance in modeling fire scenarios. 

Axial conduction within the fuel assembly region was modeled only in the fuel cladding and 

backfill gas, to be consistent with typical applications of the fuel effective conductivity model, 

conservatively neglecting axial conduction in the uranium oxide fuel. The axial effective 

conductivity was determined with a cross-sectional area weighting scheme based on the total 

cross-sectional area of the assembly. However, to appropriately capture the thermal inertia of 

the fuel assemblies for the transient response in the fire scenario, the effective density and heat 

capacity for the fuel region was defined based on volumetric averages of the corresponding 

properties of the helium gas, fuel rod cladding, and uranium oxide fuel pellets. 

An average volumetric heat generation of 2105 Btu/hr (617 W) was applied over the active fuel 

length for each fuel assembly. The axial distribution of decay heat was represented by dividing 

the active fuel length into 16 separate zones, and the local heat load was determined by 

multiplying the average by an appropriate peaking factor for that particular zone. The peaking 

factor was determined based on the bounding axial power profile presented in the SAR, which 

has a normalized peaking factor of 1.1.

The helium gas filling the 0.5075-inch gap between the nominal fuel assembly cross-section and 

FSS was modeled with solid elements and used standard helium thermal properties for 

conduction, density, and specific heat. Convection across the gap was accounted for by 

multiplying the local gas conduction values by an empirically derived1 Nusselt number of 3.66. 

Thermal radiation exchange across the gap was modeled with MATRIX50 super elements. 

These were created by using SHELL57 elements to designate the discrete enclosure. The 

AUX-12 hidden ray-tracing method was used to compute view factors for each element within 

the super-element. All other gaps in the package assembly, such as between the closure 

assembly and FSS, or the impact limiters and package skin were modeled in a similar manner, 

which included thermal radiation and conduction across surfaces but assumed negligible 

convection. 

 1 This value is based on thermal measurements in full-scale spent fuel storage systems. See Michener 

et al. (1995) and Creer et al. (1987).
5-9

Other potential gaps not explicitly modeled within the geometry, such as between the gamma 

shield and package FSS, and between the gamma shield and stainless steel wall were 

accounted for by modifying the material properties of the adjacent materials to include the 

calculated effective properties for the material and gap. For very small gaps, the calculations 

were based on the following assumptions:

• the thermal radiation view factor is specified as 1.0 (gap completely enclosed) 

• the temperature difference across the gap is small 

• convection heat transfer across the gap can be neglected. 

For the pre-fire steady-state and post-fire transient cooldown analyses, nominal gap distances 

were used to determine the effective thermal conductivity. During the fire transient, the gaps 

were assumed to close due to thermal expansion of the package materials, such that the 

calculated effective thermal conductivity across a gap reduced to the thermal conductivity of the 

adjacent solid material without the gap. This ensured a conservative approach with respect to 

the effect of heat transfer across the gaps throughout the entire analysis.

The impact limiter attachment bolts and the closure assembly bolts were represented as line 

elements within the model. Small variations in the overall length of individual bolts were

accounted for by calculating an equivalent cross-sectional area, which was specified in the real 

constant properties for the line elements.

5.2.2 GA-4 Package beneath Collapsed Roadway

The location of the package and configuration of the concrete roadway overlay, as represented 

in the fire scenario modeling, is defined to maximize coverage of the package, within the 

physical geometry of the collapsed structures. The overall length of the GA-4 package without 

impact limiters is 15.65 ft (4.77 m), and the outer diameter is 3.3 ft (1 m). With impact limiters, 

the overall length is 19.5 ft (5.94 m) and the maximum diameter is 7.5 ft (3.39 m). Figure 5.6

shows block “footprints” of the package (with impact limiters) approximately to scale in the 

locations where an object of that size could be completely covered by the fallen roadway spans. 

The conveyance is not considered in this evaluation, as noted in Section 5.2.1 as a modeling 

conservatism for the fire exposure. It is also a conservative assumption for consideration of the 

effect of the roadway collapsing onto the package. Due to the size of the package and the

conveyance, it would be virtually impossible to completely cover them both with the fallen 

roadway in this accident scenario. 
5-10

Figure 5.6. Locations Where an SNF Package Could be Covered by Collapsed I-580 Spans

(photo from MAIT Report, CHP 2007, reprinted with permission.)

Modeling of the structural response of the package due to the collapse of the roadway spans is 

discussed in Section 5.4. From the standpoint of thermal performance of the GA-4 package, the 

important effect of the collapse is the additional thermal barrier imposed by a concrete “blanket”

over the package in the post-fire cooldown. Of the two possible locations where the package 

would be completely covered by a fallen span, the location beneath the fully collapsed Bent 19-

20 span clearly imposes the more severe conditions. At the location beneath the partially 

collapsed Bent 18-19 span, only the top of the package would be covered, due to the limited 

area of contact between the upper span and the lower roadway. Beneath the completely 

collapsed Bent 19-20 span, the package would be entirely covered by the fallen roadway.

Figure 5.7 shows a cross-section diagram of the assumed configuration for the GA-4 package 

beneath the collapsed I-580 span. As a conservative simplification, it is assumed that the 

package is not actually in physical contact with the concrete structure enclosing it, and the 

neutron shield tank outer shell is not deformed by contact with the fallen span. This preserves 

the air gap of the evacuated shield tank, which insulates the package body during cooldown. 

The crushed girders on the underside of the fallen overhead span are assumed to hold the 

roadway 9.4 inches (0.24 m) away from the package body. This distance is estimated from the 

collapsed height of the severely deformed girders of the actual roadway structure that were 

exposed to the intense heat of the large pool fire. In addition, the package is assumed to stand 

on its lifting trunnions, so that there is a 5-inch (0.13-m) gap between the package body and the 

lower roadway. A similar gap is assumed between the package outer shell and the surface of 

the concrete edge barrier. The touch-point of the fallen I-580 span on the I-880 roadway is 

assumed to be approximately one-quarter of the lower roadway width, at 4.66 ft (1.42 m) from 

the package centerline. Given the stiffness of the concrete-and-steel roadbed, it is reasonable 

to suppose that the presence of the package would tend to hold the upper span away from 

contact with the lower roadway over much of its width, which makes this an extremely 

conservative assumption.
5-11

Figure 5.7. Diagram of Collapsed Roadway Configuration over SNF Package

In the ANSYS model, the configuration shown in Figure 5.7 is represented by enclosing the 

entire SNF package within a simplified structure representing the external environment, as 

illustrated in Figure 5.8. During the fire, this structure provides boundary conditions for the 

thermal analysis simulating a fully engulfing fire surrounding the package. (See Section 6.4 for 

discussion of the modeling of the structural response of the package in this fire scenario.) After 

the end of the fire, the surrounding enclosure models the configuration of the package 

positioned on the lower I-880 roadway adjacent to the side barrier, beneath the fallen upper I580 roadway. Heat transfer between the package and the enclosing concrete was assumed to 

consist of conduction and convection through the surrounding air within the enclosure, and 

thermal radiation exchange between the package external surfaces and the concrete surfaces 

representing the roadway above and below. This provides a conservative representation of the 

postulated package configuration during the post-fire cooldown.

NOTE: Diagram not to scale
5-12

Figure 5.8. ANSYS Analysis of GA-4 Package: Model Element Plot (Axial View)

Thermal radiation interaction between the package and concrete roadway surfaces was defined

by coating the exterior surfaces with SHELL57 elements having appropriately specified emissive 

material properties. The SHELL57 elements were then used to produce highly structured AUX12 generated MATRIX50 super elements. A total of twenty-four MATRIX50 super elements 

were defined to capture the radiation interaction between the package and roadway surfaces. 

The twenty-four MATRIX50 super elements represent two sets of thermal radiation conditions; 

half of the super elements represented pre-fire thermal radiation emissivities, and the other half 

represented the fire and post-fire emissivities. Only one set of external MATRIX50 superelements was active at a time during the calculation, depending on the stage of the analysis 

(i.e., pre-fire, fire, or post-fire).

The concrete enclosure is essentially the same for the ANSYS and the COBRA-SFS models, 

except for the dimensions assumed for the post-collapse structure. For the ANSYS model, the 

size of the enclosure is expanded to accommodate the impact limiters, while in the COBRA-SFS 

model, the concrete overlays the package body (as described in Section 5.3). In reality, of 

course, the falling roadway would crush the impact limiters against the package body, but this 

was conservatively neglected in the ANSYS model. Accounting for the effect on thermal 

behavior due to crushing of the impact limiters would add unnecessary complication to model,

and in addition, would be a less conservative approach, as it would tend to enhance conduction 

heat transfer from the package. Assuming the stainless steel shells of the impact limiters retain 

their original configuration imposes a conservative representation of the insulating barrier these 

structures present to heat transfer from the package during the cooldown phase of the transient.
5-13

Convection heat transfer on the external surfaces of the package was determined using

empirical correlations, which are described in Section 6.3. Free convection was assumed 

before and after the fire, while forced convection was assumed during the fire. Convection was 

implemented using SURF152 elements. These elements are placed on the exterior surface of a 

body and communicate with the designated sink temperature assigned to a single node (called 

the “space node”) to compute the heat flux. Figure 5.9 shows a representation of a space node 

relative to the package assembly.

Figure 5.9. Space Node for Convection Heat Transfer

In order to represent variation in convection heat transfer over different regions of the package 

surface, due to variation in the surface geometry, the external surfaces were partitioned into 

axial and radial zones. The package was divided into four different zones along the exposed 

outer surfaces, corresponding to the top, bottom, and left and right sides of the horizontal 

package, as illustrated by the colored segments in Figure 5.10, to appropriately model variations 

in convection along the exterior of the package. On the package body, these zones correspond 

to the radial positions of the four fuel assemblies within the package. The outer surfaces of the 

impact limiters were also split into four zones in the same manner as the package body, as

illustrated in Figure 5.10. In addition, the four lifting trunnions were split into two different zones 

each, with two located along the top package boundary and the remaining two located along the 

bottom boundary of the package. Within each zone different heat transfer coefficient 

correlations were used, corresponding to the geometry of the surface (e.g., vertical flat surface, 

horizontal curved surface), by defining subzones within each zone. A sink node was defined for 

each zone and subzone, which defines the ambient air temperature seen by the package in that 

zone or subzone, and is used to calculate the local convection coefficient at the package 

surface. 
5-14

Figure 5.10. Zones for Convection Computations for Package Assembly Surfaces

The external environment was split into eight different zones corresponding to the upper 

roadway and lower roadway, with the ends open to the ambient, as illustrated in Figure 5.11. 

(As shown in Figure 5.11, these modeling structures are tilted and rotated for clarity; when 

properly oriented, the cross-section of the defined zones match the cross-section of the model 

diagram shown in Figure 5.8.) The convective heat transfer coefficient is assigned to the 

package and external elements based on the temperature difference between the surface and 

sink temperature, and on the surface geometry, as described in Section 6.0. The heat 

exchange between these surfaces and the space node is then computed by ANSYS as part of 

the overall solution for a given time-step in the transient.

Figure 5.11. Zones for External Heat Transfer from Roadways and External Environment
5-15

During the fire, the sink node temperature for the SURF152 elements is set to the bounding 

flame temperature representing the fully engulfing fire. The external convection coefficient is 

computed using a forced convection relation derived from gas temperatures and velocities 

predicted in the fire simulation with the FDS code (as discussed in detail in Section 3.3). These 

boundary conditions are provided for the top, sides, and bottom of the enclosure, corresponding 

to the four zones on the package, as shown in Figure 5.10. During the fire, the nodes 

representing the external environment are also set to the bounding flame temperature.

5.2.3 Material Properties for GA-4 Package in ANSYS Model

The specific thermal material properties used to represent the components of the GA-4 package 

and roadway structures in the ANSYS model are listed in detail in Appendices A and B. For 

elements of the model representing the major components of the package, the specified 

properties are those of the single material comprising that component. However, for efficiency 

of meshing, the thin plates of the FSS and enclosed neutron absorber rods, the complex 

honeycomb structures of the impact limiters, and the fuel assemblies are represented using 

effective thermal properties defined specifically for the overall region. In addition, the effect of 

the fire on the integrity of the liquid-filled neutron shield tank was also explicitly modeled with 

changes in material properties in the transient calculation.

The neutron absorber plates of the FSS are composed of boron carbide rods sandwiched 

between thin stainless steel (XM-19) panels, with helium surrounding the boron carbide rods. 

Homogeneous material properties were defined for the elements representing the FSS plates, 

based on volumetric averaging of the material properties for XM-19 stainless steel, boron 

carbide, and helium. It was assumed that convection in the helium gas would be negligible in 

the narrow enclosed space within the FSS plates, and the effective thermal conductivity was 

calculated based on conduction and thermal radiation heat transfer only. Anisotropic properties 

were defined for this material, assuming conduction only along the axial length of the FSS, with 

conduction and thermal radiation through the thickness of the composite plate. Thermal 

radiation was modeled assuming that the helium-filled space between the boron carbide rods 

and the enclosing steel plates was very small, completely enclosed within the stainless steel 

panels, with a very small temperature difference between them. 

The stainless steel inner support structure and outer shell of the impact limiters was explicitly 

modeled using elements with properties of XM-19 stainless steel. Composite material 

properties were used to model the aluminum honeycomb material enclosed within the steel 

shell. The design of the impact limiters is defined in the package SAR (General Atomics 1998) 

as a standard non-reinforced hexagonal aluminum structure, and includes specific regions with 

differing densities, which are bonded together and to the stainless steel shell with adhesive 

foam. Effective properties for these regions were determined based on material data for 

aluminum honeycomb from HEXEL Composites (1999), using a volumetric averaging scheme. 

This approach included the properties of the adhesive foam as well as the air-filled aluminum

honeycomb. The effective thermal conductivity values for the honeycomb regions were 

calculated assuming the material was isotropic within a region, as indicated by the HEXEL 

Composites data for the honeycomb.

In the course of the transient calculation, the material properties of the impact limiters were 

modified to account for structural configuration changes and effects of the fire. Portions of the 

aluminum honeycomb in the impact limiters are assumed to melt during the fire, due to the 

extremely high temperatures predicted in this transient. For the aluminum honeycomb material 

in the impact limiters, local melting would be expected to significantly increase in the void (air) 
5-16

volume compared to the intact honeycomb material. This would tend to increase the insulating 

effect of the impact limiters, reducing the rate of heat transfer through this material. During the 

fire portion of the transient, the impact limiters were conservatively assumed to remain intact, 

allowing the maximum heat transfer to the package through these components during the fire. 

However, the assumption of intact impact limiters is no longer conservative in the post-fire 

cooldown portion of the transient. With larger air regions within the impact limiter structure due 

to local melting of the honeycomb, the damaged impact limiters would tend to further slow the 

rate of heat removal from the package during the cooldown transient, compared to the effect of 

intact impact limiters. The material properties of elements in the ANSYS model representing the 

honeycomb material were therefore modified in the post-fire portion of the calculation to account 

for the effects of melting. 

Fire damage to the impact limiters was determined from the predicted temperature distribution 

within these regions at the end of the fire. The percentage of honeycomb nodes above the 

melting point of the aluminum alloy (approximately 1100°F [593°C]) was used to calculate the 

total volume of melted aluminum, and the volume of “lost” honeycomb. It was assumed that the 

molten aluminum would flow due to gravity to the lowest point on the horizontal side of the 

impact limiters. Therefore, elements in this region encompassing a volume corresponding to 

the volume of melted aluminum were modified to have the properties of aluminum alloy, rather 

than the honeycomb mesh. The remaining volume of the impact limiter was assumed to be a 

mixture of air (corresponding to the volume of the melted mesh) and unmelted intact 

honeycomb. The thermal conductivity of the elements representing this volume within the 

impact limiters was modified using an effective thermal conductivity calculated based on 

volume-averaging of the thermal properties of air and the unmelted honeycomb mesh material. 

The effect of the fire on the material properties of the liquid neutron shield was also explicitly 

represented in the transient calculation. The neutron shield liquid temperature is calculated to 

exceed its boiling point very early in the fire transient. Prior to rupture, heat transfer through the 

liquid in the tank is represented with an effective conductivity relationship based on an empirical 

correlation (Guyer and Brownell 1989) for convection and conduction heat transfer across a gap 

between two long, horizontal concentric cylinders at different temperatures. The fluid thermal 

conductivity used in this relationship was determined based on material property data for 

propylene glycol and water mixtures provided in the GA-4 SAR (General Atomics 1998).

(Appendix B contains a detailed description of this correlation, and verification of its applicability

to the geometry of the GA-4 neutron shield tank.) 

The neutron shield tank is assumed to rupture when the peak temperature in the liquid is 

predicted to exceed the boiling point of the water-glycol mixture. After rupture, the neutron 

shield tank contents are assumed to consist only of air, with heat transfer by conduction and 

convection. Thermal radiation between the inner walls of the empty tank is also accounted for, 

by direct calculation between the elements on the inner surface of the tank outer shell and the 

outer surface of the package body. 

The effective conductivity of the material within the neutron shield tank was determined as a 

function of the average tank temperature and the radial temperature difference between the 

tank inner and outer surfaces. The radial temperature difference was calculated separately 

along the flats and corners of the neutron shield, to account for the effect of the non-uniform gap 

due to the square cross-section of the tank inner surface within the circular outer tank shell.

Material properties for the tank were updated between each time step during the transient 

solution. The affected nodes were assumed to consist of a 56% propylene glycol solution up to 

the point where the maximum temperature reached the mixture’s boiling point of 276°F (136°C).
5-17

The boiling point for the tank contents, and hence the time of assumed tank rupture, was 

calculated based on the maximum normal operating pressure of the neutron shield tank

(General Atomics 1998), and data for vapor pressure versus temperature of aqueous solutions 

of propylene glycol (Dow Chemical Company 2003). When the maximum temperature in the 

tank exceeded the boiling point, it was assumed that rupture had occurred and all the liquid in 

the tank instantly vaporized. The effective conductivity was then computed using dry air as the 

medium. This calculation extended through the remainder of the fire and was also continued 

during the cooldown period. This approach conservatively neglects energy absorbed by the 

phase change (i.e., the heat of vaporization for the liquid), but this is mainly as a matter of 

convenience, since this would constitute a very small deduction from the total energy imparted 

to the package. 

5.3 COBRA-SFS Model of GA-4 Package

The GA-4 package was also analyzed with COBRA-SFS (Michener et al. 1995), a thermalhydraulic code developed for analysis 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 transfer, storage, and transportation systems, and fuel 

assemblies under forced and natural circulation flow conditions. It is applicable to both steadystate and transient conditions in single-phase gas-cooled spent fuel packages with thermal 

radiation, convection, and conduction heat transfer. 

The COBRA-SFS model was developed to provide detailed temperature distributions on the 

individual fuel rods within the package in this fire scenario. In addition, the COBRA-SFS model 

was used to investigate the potential effects on the package if it is assumed that the impact 

limiters have detached from the package in the accident prior to the fire. The impact limiters 

have a very large effect on the thermal response of the package internals to fire conditions, 

since these structures act as thermal shields on the package ends. The impact limiters are 

designed to remain attached to the package during design-basis accident scenarios, but the 

MacArthur Maze fire involves fire temperatures exceeding that of the HAC fire defined in 

10 CFR 71. In addition, this scenario includes structural impacts due to the collapse of the 

overhead highway spans after the fire exposure, which is the reverse of the impact-then-fire 

scenario prescribed for hypothetical accidents in 10 CFR 71. It is unlikely that the impact 

limiters would come off, even in this extraordinary scenario, but if it were to happen, it would 

significantly influence the thermal behavior of the package. The effect of the loss of the impact 

limiters on the package response to this fire is therefore included in this analysis.

As in the evaluations with the ANSYS model, evaluations with COBRA-SFS consider two 

configurations for the package during the fire scenario. Prior to and during the collapse of the 

two overhead roadway spans (as discussed in Section 2.1), the package is assumed to be in a 

fully engulfing fire at 2012°F (1100°C), and sees a uniform bounding fire temperature in all 

directions. After the collapse, the package is assumed to be exposed to a fully engulfing 

1652°F (900°C) fire. In a separate independent calculation, the concrete structure of the 

collapsed roadway configuration is exposed to these fully engulfing fire conditions, in 

preparation for the post-fire cooldown calculation. The concrete surface temperatures obtained 

in this calculation are used as boundary conditions for both the ANSYS model and the COBRASFS model evaluations of the post-fire cooldown conditions. For the post-fire analysis with 

COBRA-SFS, the model includes the GA-4 package, the lower roadway, side barrier, collapsed 

upper roadway, and the air space between the package and the concrete structures. The 
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model of the package is described in Section 5.3.1. The package-under-concrete configuration 

for the post-fire scenario is discussed in Section 5.3.2.

5.3.1 COBRA-SFS Model of GA-4 Package in Fully Engulfing Pool Fire

The GA-4 package was modeled for the COBRA-SFS calculations in sufficient detail to capture 

the thermal response of the system components in the radial and axial directions. Material 

properties used in the model are listed in Appendix A. The four fuel assemblies within the 

basket are each modeled as rod and subchannel arrays (Michener et al. 1995), for appropriate 

representation of thermal radiation heat transfer as well as conduction and convection. The 

basket separating and containing the fuel assemblies is represented using multiple layers of 

solid conduction nodes, to capture the effect of the B4C poison rods stacked within the steel 

plates forming the arms of the cruciform structure. 

The steel inner liner, DU gamma shield, and steel package body are also represented with 

multiple layers of solid conduction nodes, to appropriately resolve the temperature gradients 

through these relatively thick components. In addition, these structures are also divided radially, 

to capture the effects of non-uniform external conditions surrounding the package. A crosssection diagram illustrating the noding for the COBRA-SFS model is shown in Figure 5.12. 

Figure 5.12. Cross-section of COBRA-SFS Model of GA-4 Package, Including Fuel 

Assemblies, Basket, Package Body, and Neutron Shield

The impact limiters on the package ends are assumed lost, and are not included in the COBRASFS model of the GA-4. The impact limiters are designed to remain in place through the 

sequence of accidents specified in 10 CFR 71 for HAC; however the MacArthur Maze fire 

scenario involves fire temperatures exceeding that of the HAC fire. Although it is unlikely that 

the impact limiters would detach from the package even in this severe accident scenario, the 

fuel rod

control rod 

thimble

fuel assembly

fuel assembly

fuel assembly

fuel assembly
5-19

impact limiters would be expected to suffer severe degradation, due to the long fire duration and 

the postulated impact of the overhead span on the package. The impact limiters provide 

significant thermal shielding for this package in any fire scenario, and therefore can have a large 

effect on the response of the system to the fire. Therefore, this assumption is included in the 

COBRA-SFS model, for completeness in the overall analysis of the effects on the package of

the MacArthur Maze fire scenario. However, evaluation results shown in Section 7.3.6.1 show 

that the impact limiters would remain attached to the GA-4 package. Results obtained with the 

COBRA-SFS model assuming loss of the impact limiters are therefore bounding and 

conservative. 

The neutron shield tank initially contains a liquid 56% propylene glycol/water mixture with 

maximum design pressure of 150 psig (1.135 MPa), which gives a boiling temperature of 276°F 

(135.6°C). However, the tank is not an ASME pressure vessel, and in the SAR analyses for the 

HAC fire at 1472°F (800°C), it is conservatively assumed that in the initial steady-state, the tank 

has already ruptured and contains only air. In the COBRA-SFS analysis, the tank is assumed to 

remain intact up to the point in the transient where the maximum temperature in the neutron 

shield region exceeds the boiling point of the glycol/water mixture. Heat transfer across the 

neutron shield tank is treated in the same manner as in the ANSYS analysis, described in 

Section 5.2.3, with an effective conductivity correlation defined for the neutron shield tank 

contents.

The maximum temperature in the region modeling the tank contents is predicted to occur within 

sixty seconds of the initiation of the fire transient in this scenario, at which point the medium 

within the tank is assumed to be dry air, and thermal radiation between the tank inner surfaces 

is added to the model. The internal surfaces of the shield tank are specified with a uniform 

emissivity 0.9 after the assumed loss of liquid contents, to conservatively represent the effect of 

sooting, on the assumption that highly sooted fire gas could enter the ruptured and fully vented 

tank. Additional package modeling details related to boundary conditions and assumptions are 

included in Section 6.3.

5.3.2 COBRA-SFS Model of GA-4 Package beneath Collapsed Upper Roadway

The representation of the GA-4 package for the post-fire cooldown with the package beneath 

the collapsed upper roadway is the same as the model developed for the fully engulfing fire 

portion of the scenario. For the post-fire cooldown, the model is expanded beyond the package 

exterior shell to include additional nodes representing the concrete of the upper and lower 

roadways in relation to the package. The location of the package and configuration of the 

concrete covering it is defined to maximize coverage of the package, within the physical 

geometry of the collapsed structures in this fire scenario, as discussed in Section 5.2.2 and 

illustrated by Figure 5.7. The concrete enclosure is essentially the same for the ANSYS and the 

COBRA-SFS models, except for the interior dimensions of the post-collapse structure. For the 

ANSYS model, the size of the enclosure is expanded to accommodate the diameter of the 

impact limiters, while in the COBRA-SFS model, the concrete overlays the package body. 

As illustrated in Figure 5.7, the space between the package and the concrete structure is 

modeled as an air channel, roughly analogous to the air flow pathway in a ventilated concrete 

storage module. The main differences are that the orientation is horizontal, and the geometry of 

the concrete “annulus” is decidedly irregular. Heat transfer to the environment is modeled as 

free convection from the exposed surfaces of the concrete roadways and side barrier. For 

simplicity, the cross-section geometry shown in Figure 5.7 is assumed to extend uniformly over 

the axial length of the package, and the “edge” of the model where the upper span touches the 
5-20

lower roadway is treated as an adiabatic boundary. Both assumptions are conservative, as they 

limit or eliminate potential paths for heat to be transferred from the package by conduction 

through the concrete.

5.4 Structural Models for the MacArthur Maze Fire Scenario

The structural evaluations undertaken to determine the potential consequences of the 

MacArthur Maze fire scenario focus on two main issues. The first issue is to evaluate the effect 

of dropping a section of the I-580 roadway onto a SNF package that has been subjected to a 

fully engulfing fire. This accident scenario differs significantly from the HAC drop scenarios 

postulated in 10 CFR 71, and it is necessary to determine whether or not this real-life accident 

scenario is bounded by the postulated accidents defined in the regulations. The second issue is 

the structural response of the package to the extraordinary thermal load imposed on the 

package closure lid bolts and on the impact limiter attachment bolts in this fire scenario. The 

severity and duration of the MacArthur Maze fire scenario exceed the postulated HAC fire in 

10 CFR 71. Retention of the impact limiters on the package body and the response of the 

closure lid bolts greatly affect the potential consequences of this fire scenario, if an SNF 

package were to be subjected to such severe conditions. 

Evaluation of these two important issues requires modeling approaches on two different scales. 

The drop scenario evaluations require large-scale modeling of the package and roadway 

structures. These models are described in detail in Sections 5.4.1, and include both the 

roadway and the SNF package. The bolt response evaluations require detailed modeling of the 

regions of the package containing the impact limiter attachment bolts and the lid closure bolts. 

These evaluations use two independent approaches to evaluate this issue, initially using

classical bolt modeling studies, and then developing detailed FEA modeling of these structures 

in the GA-4, with boundary conditions provided from the thermal modeling of the package. The 

classical modeling and preliminary FEA modeling is described in Section 5.4.2 and the detailed 

FEA modeling is described in Section 5.4.3.

5.4.1 Structural Models for Roadway Drop Scenarios

This modeling study is a three-dimensional impact analysis evaluating the effect of dropping a 

section of the I-580 roadway onto the SNF package, which has been subjected to a fully 

engulfing fire at 2012°F (1100°C). The structural impact models were developed using the LSDYNA code (Livermore Software Technology Company 2007) to simulate the impact of the 

falling overhead I-580 span on the SNF package positioned beneath it on the lower I-880 

roadway. This includes appropriate models of the roadway segments involved, and of the SNF 

package. These models were used to evaluate the potential for gross structural failure of the 

package, which could affect the containment boundary. The structural models of the roadway 

are described in Section 5.4.1.1, and the model of the SNF package is described in 

Section 5.4.1.2.

5.4.1.1 Modeling of Roadway Segments

The potential effects of the collapse of the overhead roadway onto the SNF package is 

investigated in this analysis by modeling a free-fall drop of the I-580 roadway span between 

Bent 19 and Bent 20 onto a structural model of the GA-4 package on the lower I-880 roadway. 

The partial collapse of the span between Bent 18 and Bent 19 was not modeled, as it would 

result in a much less severe impact on the SNF package. Based simply on the physical forces 
5-21

involved, the slower partial collapse of the span between Bent 18 and Bent 19 could not impart 

as much energy to the SNF package as the rapid free-fall drop of the first span to collapse, and 

therefore would not have the potential to do as much damage. 

For analysis of the drop scenario, the structural model in LS-DYNA consists of the upper I-580 

roadway span between Bent 19 and Bent 20, the lower I-880 roadway segment beneath the 

overhead span, and the SNF package at rest on the I-880 roadway. Figure 5.13 shows a 

conceptual diagram of the impact model geometry. Each object is represented with a minimum 

amount of detail to make the calculation size manageable. The upper I-580 span and the I-880 

span are modeled with similar components, but the upper I-580 span includes more detail, 

higher element resolution, and represents the steel girders with material properties that reflect 

the much higher temperature reached by the girders in the fire scenario. 

Figure 5.13. Impact Model Geometry

The level of detail in each component reflects its function in the model. The falling I-580 girders 

need greater element resolution because they impact and deform around the package and the 

roadway side barriers. The fine mesh is illustrated in Figure 5.14 and in Figure 5.15, which 

more clearly shows the increased mesh density in the two girders that impact directly on the 

SNF package. The girders of the I-880 roadway (also shown in Figure 5.14) are represented 

with a much coarser mesh, as they are needed only to provide a supporting surface with 

realistic stiffness beneath the package.
5-22

Figure 5.14. Impact Model Mesh

Figure 5.15. Impact Model Mesh Density in Impact Region
5-23

The upper I-580 roadway span between Bent 19 and Bent 20 is modeled as a deformable 

impact object, using geometry information from the original plate girder design drawings. The 

plate girders are the most important components of the overpass system for the impact 

modeling because these are the roadway structures that would contact the package directly in 

any potentially damaging drop scenario. The longitudinal girders are modeled with the major Isection plates connected as continuous material. Small lateral stiffening webs and the lateral 

plates and crossing members that connect to the longitudinal girders are neglected. This 

modeling approach is based on post-fire photographs (CHP 2007) showing that the main girder 

plates remained attached to the underside of the roadway but the stiffening ribs and cross 

members tended to fail or separate. Under localized impact, these components would be 

expected to fail without significantly influencing the impact response of the system. 

Figure 5.16 shows a diagram of the mesh for the lower roadway (I-880) portion of the impact 

model. The finite element representation is similar to the upper roadway span, but with a 

coarse mesh density. A broad rigid surface (omitted from the diagram, for clarity) is also 

included in the model at ground level, to catch the end of the upper roadway that hangs over the 

edge of the lower roadway. In the impact analysis, the results show that one end of the falling 

roadway impacts the package while the other end continues falling until it hits the ground. This 

results in a significant bending of the falling span along the edge of the lower roadway. The 

purpose of this modeling effort does not extend to matching the final resting state of the fallen 

span, but it is a useful verification of the modeling fidelity that the overall behavior predicted with 

this model is consistent with the post-fire images of the fallen span, such as the one shown in

Figure 1.1.

Figure 5.16. Model Mesh of Bottom Roadway (I-880) Segment
5-24

The concrete and rebar of the I-580 roadway is modeled as a homogenized linear-elastic 

material with an artificially low modulus of elasticity. Actual concrete behavior during the impact 

would be more complex, with concrete fracturing and breaking apart, but these nonlinear effects 

absorb energy and would reduce the amount of energy available to do damage to the package. 

The low elastic modulus for the upper roadway concrete results in the weight of the upper 

roadway acting on the plate girders through the full period of impact. Treatment of the roadway 

as a linear-elastic material is conservative in that it contributes to higher and more prolonged 

impact forces than would be obtained with a more physically realistic nonlinear model. 

The steel girders of the upper I-580 roadway span are represented as elastic-plastic material, 

with temperature-dependent mechanical properties to account for effects of the elevated 

temperatures due to the fire. A uniform temperature distribution is assumed for the final set of 

analyses, so the material properties do not vary spatially. Tiebreak contact definitions are used 

in the model to attach the roadway to the plate girders. These are virtual links in the model that 

connect a node to a surface until a specified reaction force is exceeded, at which point the node 

is free to move away from the surface. However, even in cases when the tiebreaks fail due to 

excessive normal or shear force, the plate girders and roadway still detect contact with each 

other and all other components in the model through an all-inclusive general contact definition.

The falling span is subjected to constant acceleration due to gravity and a specified initial 

downward velocity that depends on the placement of the cask on the lower roadway. It is 

assumed that the overpass section falls straight down and the cask is resting on the I-880 

roadway surface. Sagging of the girders before collapse is not considered, as it would reduce 

the potential freefall distance and associated impact velocity. The selected location and 

orientation of the cask dictates the distance the overpass can fall before it makes contact with 

the cask. That freefall distance and the initial location of the bridge span in the model dictates 

the initial velocity, which averages 30 ft/s (9.1 m/s) for the four impact scenarios considered. 

If it were assumed that the package remained on the conveyance, the drop distance would be 

about 3.3 ft (1 m) less, resulting in a lower impact velocity (approximately 15 ft/s [4.41 m/s]), and 

therefore lower kinetic energy upon impact. In addition, the crushing of the conveyance would 

absorb some of the impact energy. The modeled drop assumes an uninterrupted fall onto the 

package on the roadway surface, neglecting all factors that might mitigate the force of the 

impact, including drag effects. These conservative assumptions result in a much more 

damaging drop scenario than could occur in reality in this accident scenario.

The support columns initially holding up the upper I-580 roadway span and the lateral box beam 

spanning the Bent 19 support columns are conservatively neglected in the model, since they 

have essentially no effect on the impact scenario. Based on post-fire photographs (CHP 2007), 

the box beam remained attached to the partially collapsed span between Bent 18 and Bent 19. 

As noted above, the partial collapse of that span was not analyzed in the impact model because 

the impact loads would be much smaller than in the case of the freefall of the span between 

Bent 19 and Bent 20.

The lower I-880 roadway is conservatively modeled to provide a foundation of appropriate 

stiffness beneath the package. It consists of a set of plate girders supporting the concrete 

roadway, with the ends of the plate girders fixed in space. The model also includes the 

concrete barriers on either side of the roadway deck. All materials of the lower roadway are 

treated as linear-elastic, and are represented with properties at 80°F (27°C) nominal 

temperature. This conservatively neglects temperature-dependence of the material properties 

of the lower roadway, which would be more compliant at higher temperatures. In the model, the 
5-25

girders and side barriers are connected to the concrete roadway by tied surface definitions. 

Tied surface constraints effectively bond two surfaces by forcing the nodes of the “slave” side to 

maintain their relative spacing with the nodes of the “master” side. This type of connection 

persists throughout the analysis, unlike the tiebreak connections used in the upper roadway that 

allow the connection to end when the forces required to maintain the connection exceed a 

prescribed limit.

Conservative estimates of temperatures were obtained for the girders in the drop calculations, 

based on the results of thermal analyses to determine fire effects on the overhead girder 

temperatures. For the girders, a conservative temperature would be as low a temperature as 

can be reasonably postulated; at lower temperatures, the steel is stronger and therefore able to 

impart more energy to the package. Based on sensitivity studies of the effect of the fire on the 

overhead span prior to falling, a uniform temperature of 1800°F (982°C) was assumed for the 

girders and the upper I-580 roadway. The development of this important boundary condition is

discussed in detail in Section 6.4.1. The peak temperature on the I-580 girders may have been 

higher than 1800°F (982°C), possibly as high as 2462°F (1350°C), but for the purposes of this 

analysis, the lower estimate of the girder temperatures is conservative, compared to assuming 

higher temperatures. Because of the temperature-dependent properties of steel (see Figure 

2.10 for yield strength-versus-temperature), the lower temperature assumption results in a more 

rigid overhead span, and hence a more severe impact on the package. 

5.4.1.2 GA-4 Package Model for Structural Analysis

The GA-4 package model is simplified to represent only the components that give it structural 

integrity. These are the steel package body, the DU gamma shield, and the lifting trunnions 

attached to the steel body. The fuel assemblies and FSS structure within the package cavity 

are neglected, as they would have limited effect on the package response to an external impact.

The impact limiters and the fragile steel shell of the neutron shield tank are also ignored in this 

simulation, as a conservative simplifying assumption, since the effect of these structures would 

be to absorb energy as they crumpled under the impact of the overhead span. 

The XM-19 stainless steel package body is a long rectangular box with rounded corners, with 

four lifting trunnions attached near either end, on the rounded corner edges. Figure 5.17 shows 

a diagram of the mesh used to represent the package for the calculations with LS-DYNA. The 

lid and flange region is treated as a solid volume of steel, without the physical interface between 

the two components, resulting in a fully enclosed package body. Realistic interaction between 

the lid, flange, and lid bolts is not captured in this model. The behavior of these components is 

investigated separately, using models described in Sections 5.4.2 and 5.4.3.
5-26

Figure 5.17. Model Mesh of the GA-4 Package Body

The DU gamma shield is modeled inside the package as a single free-moving body, since this 

component is enclosed within the XM-19 stainless steel cask body but is not physically attached 

to it. Figure 5.18 shows a diagram of the DU gamma shield mesh in the model. A small gap 

exists between the inner package wall and the gamma shield, on the order of 0.04 inches 

(1 mm), as represented in the design drawings from the SAR (General Atomics 1998). 

Dimensional changes in the package due to thermal expansion are neglected in this analysis, 

as their effect is minimal in a model with this level of detail.

Figure 5.18. Model Mesh of the DU Gamma Shield
5-27

The connection between the trunnions and the package wall are represented with tied surface 

definitions. In the model evaluations, the outer faces of the trunnions are constrained to prevent 

the package from rolling under the impact of the overhead span. This forces the impact with the 

descending span to occur on the uppermost corner of the package cross-section, as a 

conservative assumption to localize the impact interaction.

Conservative estimates of temperatures of the GA-4 package were obtained for the drop 

calculations, based on the results of the thermal analyses of the package. For the package, a 

conservative temperature would be as high a temperature as can be reasonably estimated, as 

the material of the package will then be weaker, and therefore less able to absorb the force of 

the impact without damage. A temperature significantly above the hottest steel body 

temperatures calculated in the thermal analysis of the GA-4 package in the MacArthur Maze fire 

scenario was specified for the GA-4 package body, to obtain a conservative representation of 

the potential effect of the impact. A detail description of boundary conditions for the structural 

calculations is included in Section 6.4.

5.4.2 Preliminary Models for GA-4 Bolt Thermal Expansion Calculations

Preliminary models were developed to investigate the response of the package bolted 

connections to the extremely high temperatures of this fire scenario, since failure of these bolts 

could potentially affect containment integrity of the package. There are two critical concerns for 

the GA-4 package in this fire scenario; the retention of the impact limiters throughout the entire 

transient, and performance of the package closure lid bolts. 

The impact limiter attachment bolts were evaluated to determine if these fasteners would be 

expected to hold at the elevated temperatures of this fire scenario. Failure of these bolts could 

potentially result in separation of the impact limiters from the package, which (for the top impact 

limiter) would subject the lid closure bolts to direct exposure to the engulfing fire. This would 

result in significantly higher temperatures for the lid closure and closure bolts than are predicted 

during the fire if the top impact limiter remains in place. 

The evaluations of the behavior of the impact limiter connection bolts and of the package lid 

closure bolts consisted primarily of studies of the effect of differential thermal expansion on bolt 

loading. Temperatures for the bolts and surrounding structures were determined from results of

the detailed transient simulations of the fire scenario with the ANSYS and COBRA-SFS thermal 

models. A diagram of the impact limiter geometry, with the locations of the attachment bolts 

indicated, is shown in Figure 5.19. (Note that this diagram is not to scale; the size of the bolts 

relative to the impact limiter is greatly exaggerated for clarity of illustration.) These bolts thread 

into an anchor plate that is attached to the steel body wall of the GA-4 package. Figure 5.20

shows a close-up view of the package closure lid and bolts, from the exploded view diagram of 

the entire package (shown in Figure 5.1).
5-28

Figure 5.19. Diagram of Impact Limiter, Illustrating Attachment Bolts

Figure 5.20. Package Closure Region, Showing Package Closure Lid and Bolts
5-29

The performance of the lid closure bolts was evaluated assuming that the impact limiters would 

be retained throughout the fire scenario, since the results of the evaluations of the performance 

of the impact limiter attachment bolts (as discussed in Section 7.0) shows conclusively and 

conservatively that the impact limiter attachment bolts would not fail in this fire scenario. The 

impact limiters therefore act as “friend and foe” to the closure lid during this fire scenario. 

During the fire, the impact limiters shield the package lid from the extreme temperatures of the 

fire, and therefore limit the temperature rise on this structure during the fire. In the post-fire 

cooldown, however, the impact limiter is an extremely effective insulator, and package 

temperatures in this region continue to rise for many hours after the end of the fire, as the 

extreme heat developed in the package body and internal structures during the fire naturally 

conducts toward the lower-temperature ends of the package. Evaluations of the closure bolts 

were undertaken to determine the response of the closure connection to the long exposure to 

elevated temperatures in the complete fire scenario, since the effect of the fire scenario on

these bolts could potentially result in failure of package containment.

Both sets of attachment bolts are Inconel 718, a nickel alloy (General Atomics 1998). The 

closure lid is secured with 12 bolts; 3 per side of the square cross-section of the lid. The impact 

limiter is attached with 8 bolts, uniformly spaced in a circular ring. The package body flange that 

receives the lid closure bolts is XM-19 stainless steel, as are the anchor plates that the impact 

limiter bolts thread into. In both sets of connections, steel helical thread inserts are used to 

make the connection between the bolt threads and the matching threads of the package body 

flange (for the lid closure bolts) or anchor plate (for the impact limiter bolts). 

The threaded inserts are ASTM Type 304 stainless steel. Inconel 718 and XM-19 stainless 

steel both retain their strength better than Type 304 stainless steel as temperature increases, 

which makes the threaded inserts a weak link in the threaded connection at elevated 

temperatures. In addition, the coefficient of thermal expansion for the nickel alloy of the bolts is 

significantly less than that of XM-19 stainless steel, so elevated temperatures will cause an 

increase in bolt tension and thread shear load. This issue is further complicated by the large 

thermal gradients predicted in the impact limiter attachment bolts, based on the results of the 

thermal analyses of the GA-4 package in the MacArthur Maze fire scenario.

For both the closure bolts and the impact limiter bolts, a preliminary evaluation was performed 

using classical mechanics equations to calculate bolt tension at a series of uniform temperature 

states. The coefficient of thermal expansion mismatch between the bolt material and the flange 

or anchor material results in an increase in tension in the bolt shank, which must be supported 

by an opposing shear force on the bolt threads. The strength of the materials is a function of

temperature, so the yield and failure thresholds vary with time and changing temperature in this 

fire scenario. One concern in each case is the yield and failure limits of the bolt shank. The 

shear strength of the bolt threads are not a concern because the Type 304 thread inserts are 

weaker for all temperatures, and therefore would be expected to yield before the bolt threads. A 

classic thread-stripping equation is used to relate bolt tension to total shear force on the 

threaded interface. 

The closure bolts and the impact limiter bolts can yield without actually failing in this scenario. 

Bolt materials are expected to have sufficient ductility to reduce the design-basis tensile load to 

zero without failing, as a general design criterion. Only if the tensile load could not be relieved 

sufficiently by yielding, and the residual load after yielding was still high enough to exceed the 

ultimate strength of the bolt material, would the bolt be expected to fail. For the closure bolts, a 

potential consequence of bolt failure is an unattached package closure lid. For the impact 

limiter bolts, potential consequences of bolt failure include detachment of the impact limiter from 
5-30

the cask. In either case, failure of the bolts would indicate significant potential consequences in 

the overall evaluation of the MacArthur Maze fire scenario. 

The preliminary evaluations provided an initial indication of the bolt response to the MacArthur 

Maze fire scenario, but additional evaluations were also undertaken using finite element models, 

to obtain a more complete understanding of the potential consequences of this fire scenario. 

For the impact limiter bolts, the uniform temperature assumption used in the initial evaluations 

was determined to be overly simplistic, due to large temperature gradients along the length of 

the bolt and surrounding material. To investigate the effect of these temperature gradients, a 

fully three-dimensional FEA model of the impact limiter bolt as it connects into the package was 

developed using ANSYS, to analyze the structural and thermal response of a bolt and the local 

surrounding material. Figure 5.21 shows meshing diagrams of the impact limiter bolt model 

developed for this analysis. Temperature distributions along the axial length of the bolts, 

obtained from the detailed ANSYS thermal model of the package, were incorporated into 

evaluations with this impact limiter bolt model. 

Figure 5.21. Model Mesh of Impact Limiter Bolt

In the evaluation of the results obtained with the preliminary models for the response of the 

closure bolts in the fire scenario, it was determined that the evaluation using mechanics 

equations was excessively conservative, and did not yield a sufficiently detailed picture of the 

consequences of the bolt response to this fire scenario. Therefore, additional, highly detailed 

FEA modeling was undertaken for the end region of the package, including the closure lid, steel 

body flange, and closure bolts. The modeling developed for this evaluation is described in 

Section 5.4.3.

cut-away view
5-31

5.4.3 Detailed Models for GA-4 Closure Lid Bolt Evaluations

The results of the preliminary models evaluating the behavior of the bolts attaching the impact 

limiters to the package show conclusively that the impact limiters would be expected to remain 

in place during the severe conditions of this fire scenario. In addition, the preliminary models 

show that even with extremely conservative assumptions, the bolts fastening the closure lid will 

not fail due to stresses related to differential thermal expansion of dissimilar materials. 

However, the simplifications and conservatisms in the preliminary models do not provide results 

with a sufficient level of detail to determine the detailed history of the clamping force the bolts 

impart to the lid/flange interface throughout the fire scenario. These analyses therefore cannot 

provide a means of obtaining a reliable estimate of the size of any potential leakage pathway for 

the contents of the package to escape to the environment.

This is crucial to determining the potential consequences of the MacArthur Maze fire scenario,

and an additional detailed study was undertaken to develop a realistic evaluation of the closure 

bolt response over time. Evaluation of the closure force on the flange for the entire duration of 

the fire scenario, including the post-fire cooldown to a new ambient steady-state, can be used to 

determine a meaningful evaluation of the potential effect on package containment for the 

conditions postulated in this fire scenario.

Section 5.4.3.1 describes the modeling approach used in the preliminary evaluations. 

Section 5.4.3.2 presents the detailed FEA model developed for this most important evaluation in 

the overall study of the potential consequences to an SNF package of an accident of the 

magnitude of the MacArthur Maze fire scenario.

5.4.3.1 Modeling for Preliminary Closure Lid Bolt Evaluations

Classic bolt equations and analysis methods were applied to evaluate the response of the 

closure lid, bolts, and flange to the conditions of the MacArthur Maze fire scenario. In this 

context, “classic” refers to the set of equations that define bolt and flange loading, and can be 

found in any comprehensive engineering textbook on the subject, or as closed form equations 

generally used in the literature. The primary reference relied upon in this study was 

Fundamentals of Machine Component Design by Juvinall and Marshek (Juvinall and Marshek 

1991).

The classic bolt calculations for the GA-4 closure bolts were used to determine estimates of the 

bolt tension and clamping force over time, extending from initial steady-state normal conditions 

of transport (NCT), through the fire transient, and on out through the long post-fire cooldown. 

These evaluations considered all three types of loads that these fasteners would be expected to 

experience in the MacArthur Maze scenario, which consist of the following: 

• the preload caused by the initial torque on the bolts (as per the package manufacturer’s 

technical specifications), 

• the external load caused by the internal gas pressure within the package cavity, and 

• the thermal load due to the dissimilar coefficients of thermal expansion of the materials of the 

XM-19 stainless steel closure lid and flange, and the Inconel bolts. 
5-32

The effect of the Type 304 stainless steel thread inserts was also taken into account, with 

sensitivity studies of the effect of this interfacing component on joint stiffness.

5.4.3.2 Modeling for Detailed FEA Evaluations of Closure Lid Bolt Response

In this evaluation, it is important to note a particular feature of the lid closure bolts of the GA-4 

package. They are referred to as “bolts” in the SAR, and to avoid confusion this terminology is 

also used in the current study, but these fasteners are in fact used as screws, since they are 

threaded into tapped holes in the steel flange of the package body. In order for these fasteners 

to function as bolts, they would have to pass through the flange of the steel package body and 

thread into nuts on the other side. The design of the GA-4 package body does not allow this 

configuration for the fasteners holding the lid in place.

The distinction between “bolts” and “screws” for the GA-4 package closure lid is not particularly 

relevant to the preliminary evaluation using classic bolt equations. Both types of threaded 

fasteners generally perform the same function and are subject to the same family of equations 

for determining thread stripping, initial torque to a specified pre-load, and sizing to withstand an 

external load. The details of the physical configuration of the fasteners only becomes important 

in the FEA modeling undertaken to obtain a more accurate and detailed evaluation of the 

response of these fasteners to the extreme conditions of the MacArthur Maze fire scenario.

As with the FEA model of the impact limiter bolt, described in Section 5.4.2 above, radial 

symmetry was assumed in the loading of the closure lid bolts throughout the MacArthur Maze 

fire scenario. Significant gradients are present only in the axial direction, parallel to the axis of 

the bolt shank. This allowed the modeling simplification of representing only a single bolt. To 

reduce the model to a more tractable computational size, it was further assumed that an 

individual bolt was essentially radially symmetric, and could be appropriately modeled with a 

one-quarter section of symmetry, as shown in Figure 5.22, illustrating the model geometry of the 

bolt, lid and flange. This illustration also includes a close-up image of the modeling of the 

helical coil of the Type 304 stainless steel thread insert filling the thread patterns of the 

conjoined bolt and flange. The lid and flange are represented as an annulus with an outer 

diameter equal to 2.7 inches. This corresponds to a minimum annulus area of about 4.6 square 

inches, which is effectively equivalent to the estimate used in the classic bolt equations.
5-33

Figure 5.22. Finite Element Model Geometry for Closure Lid, Flange, and Bolt, including 

Thread Insert

The mesh density in the region of the threaded interface and thread insert is shown in detail in 

Figure 5.23. The yellow ovals indicate the surfaces where the threads and inserts are joined 

with degree-of-freedom couples. This forms a perfect bond between the materials (in all three 

directions) at the load-bearing surfaces and permits the other surfaces to separate. The use of 

more advanced contact surface behavior was considered to be unnecessary based on the 

results obtained using this simpler method. The octagonal shape of the thread inserts is 

primarily defined by the thread flanks, which are assumed to have spacing and angles defined 

by standard thread characteristics (Oberg et al. 2004). The thread flanks end at the minimum 

diameter of the external bolt thread and the maximum diameter of the internal flange thread. 

The bolt-side threads and flange-side threads are also assumed to be sized as per standard bolt 

threads, so the two sides have a mating geometry rather than an identical geometry. The bolt 

threads have a shear area that is proportional to 3/4-pitch. The flange threads have a shear 

area that is proportional to 7/8-pitch. This is consistent with the preliminary calculations using 

the classical equations for bolt modeling, in which it was assumed that the critical shear area for 

the insert occurs at the maximum diameter of the bolt thread and corresponds to 7/8-pitch. 
5-34

Figure 5.23. Detailed FEA Mesh in Thread Insert Region, Showing Bonded Surfaces

This detailed FEA model was used to perform evaluations to determine the effect of the 

MacArthur Maze fire scenario on the integrity of containment of the GA-4 package, as 

determined by the clamping force on the closure lid throughout the transient. Sensitivity studies 

were performed on the effect of assuming linear or non-linear material properties for the thread 

inserts. Additional cases were developed, to evaluate the causes of differences between results 

obtained with the detailed FEA model and those obtained with the classical bolt equations, to 

verify that differences were due to more realistic modeling, and assess the conservatism of the 

simpler modeling approach.
6-1

6.0 ANALYSIS METHOD

The detailed analysis models described in Section 5.0 were developed for evaluation of the GA4 package performance if exposed to the severe conditions of the MacArthur Maze fire. For the 

thermal analyses, the models account for all significant heat transfer paths to and from the 

package by means of conduction, convection, and thermal radiation during the fire and in the 

post-fire cooldown. For the structural analysis, the models provide a conservative 

representation of the response of the package to the impact of the overhead roadway span that 

collapsed at 17 minutes into the fire. The structural analysis also includes evaluation of the 

effects of the thermal expansion mismatch between package closure flange and the closure 

bolts, and between the impact limiter anchor plates and the impact limiter bolts. The 

temperatures and temperature distributions in the bolts and associated package components 

are based on thermal model results. All of these various analyses using different analysis 

codes and physics regimes taken together provide an overall evaluation of the GA-4 package 

response to the Macarthur Maze fire scenario.

This section presents the initial conditions, modeling assumptions, and boundary conditions 

used to predict the response of the GA-4 package to the MacArthur Maze fire scenario. The 

first three sections focus on the package thermal model conditions. Section 6.1 presents the 

steady-state temperature results for NCT predicted with the ANSYS and COBRA-SFS models 

of the GA-4 package, with and without impact limiters, respectively. Section 6.2 describes the 

significant assumptions and simplifications used in developing the thermal models. Section 6.3

describes the fire scenarios assumed for the transient analyses, and the boundary conditions 

used for the calculations. The final sections deal with the assumptions, boundary conditions, 

and methodology particular to the structural modeling. Section 6.4 discusses the package 

impact analysis.

6.1 NCT and HAC Fire for GA-4 Package

Steady-state analyses for NCT with the ANSYS and COBRA-SFS models of the GA-4 package 

predict similar peak component temperatures for NCT, within the capabilities of their respective 

models for the fuel assemblies. The peak cladding temperature predicted with the ANSYS 

model is 306°F (152°C), and the COBRA-SFS model predicts 293°F (146°C). This is a 

difference of about 4 percent, and is consistent with the expected differences between the 

results obtained with a detailed thermal-hydraulic model of the fuel assemblies compared to the 

results obtained with the k-effective model for the fuel. The k-effective model for the 

homogenized fuel assembly is designed to yield results that are 5-15% conservative, compared 

to results obtained with a detailed CFD model of a fuel assembly (Bahney and Lotz 1996).

As a verification test of the accuracy of the geometry and material properties input and 

boundary condition specification logic, the ANSYS and COBRA-SFS models of the GA-4 

package were exercised for the design-basis HAC fire transient defined in 10 CFR 71. This 

consists of an engulfing fire with uniform flame temperature of “1475°F (800°C)” and fire 

emissivity of 0.9, lasting 30 minutes. In the SAR analysis for the HAC fire, the initial steadystate assumes an ambient temperature of 100°F (38°C), which is the same as for NCT, but 

solar insolation is neglected. The solar heat flux is trivial in comparison to the heat input from 

the fire, and as a modeling convenience, it was included in the analyses with the ANSYS and 

COBRA-SFS models. The steady-state initial conditions for all transient calculations are NCT; 

i.e., assuming standard solar insolation (24-hr average solar heat flux) at an ambient 

temperature of 100°F (38°C). 
6-2

Boundary conditions for the 30-minute HAC fire were specified for the ANSYS and COBRA-SFS 

models in essentially the same manner as in the SAR calculation. In general, the results 

obtained with the ANSYS and COBRA-SFS models are in very good agreement, with

differences due mainly to the alternative fuel modeling approaches used in the two analyses. In 

addition, the omission of the impact limiters from the COBRA-SFS model and their inclusion in 

the ANSYS model affect the location of the maximum cladding temperature during the fire and 

influence the post-fire cooldown behavior. 

The ANSYS model predicts 455°F (235°C) for the maximum peak clad temperature during the 

transient, and the COBRA-SFS model predicts 469°F (242°C) near the end of the hottest rod, 

since this model neglects the impact limiters. In the ANSYS model, the maximum temperature 

occurs near the axial center of the package. Both of these values are slightly higher than the 

maximum fuel cladding temperature of 442°F (228°C) reported the GA-4 SAR (General Atomics 

1998) for the HAC analysis. The difference is due mainly to simplifications in the SAR analysis 

that tend to smear out temperature gradients, and thereby reduce the conservatism of the 

analysis somewhat. However, for this package, the temperatures are low enough in the HAC 

fire that the difference can be considered insignificant. 

6.2 Thermal Modeling Assumptions

Computational modeling requires simplifying assumptions for even the most detailed 

representation of a physical system. The assumptions used in developing the detailed 

geometry models of the GA-4 package are discussed in Section 5.0 above. This section 

summarizes the major assumptions relevant to analysis of the response of this package if it 

were exposed to the conditions of the MacArthur Maze fire scenario. These assumptions apply 

to both the ANSYS and the COBRA-SFS models, unless specifically noted otherwise.

1. Initial conditions for the package are defined as steady-state NCT at 100°F (38°C) ambient 

with insolation, as defined in 10 CFR 71.71. This assumption conservatively neglects the 

effect of the actual conditions at the time of the MacArthur Maze accident (i.e., at night, with 

ambient temperature of 50°F [10°C]).

2. The decay heat load in the GA-4 package is assumed to be at its maximum design basis 

value of 2,105.4 Btu/hr (0.617 kW) per assembly, with a total package decay heat load of 

8,423 Btu/hr (2.468 kW). This is a bounding assumption, as the actual decay heat load of 

an SNF package is typically lower than the design basis configuration.

3. Material properties of package components specified as inputs to the thermal models are 

listed in Appendix A. These were obtained from the GA-4 SAR (General Atomics 1998), 

with the following exceptions;

a. The temperature-dependent thermal conductivity values used in the SAR for XM-19 

stainless steel are lower-bounding values based on properties of high alloy steels1

. At 

NCT, the thermal conductivity values from the SAR are approximately 20% below values 

published in material data sheets for XM-19 stainless steels. This is conservative for the 

NCT analysis, but is non-conservative for the fire analysis, since the lower bounding 

thermal conductivity would result in a lower rate of heating of the package during the fire. 

 1 The SAR values used for thermal conductivity of XM-19 steel are from Material Group E “high alloy 

steels” in Table I-4.0 of the ASME code, 1986.
6-3

Therefore, thermal conductivity values specific to XM-19 steel1 were used in the fire 

analyses.

b. The thermal conductivity for DU reported in the SAR is for a temperature of 

approximately 100°F (68°C), and does not take into account the significant increase in 

thermal conductivity with increasing temperature for this material. As with XM-19, this is 

a conservative approximation for NCT, but is non-conservative for fire analysis. 

Therefore, temperature-dependent thermal conductivity values were used for the DU in 

the thermal analyses, as documented in Appendix A.

4. Clearance gaps within the package (e.g., between the steel inner liner and the DU gamma 

shield, between poison rods and the steel plates of the cruciform basket) are modeled at “as 

built” values, based on design drawings.

a. Gaps are assumed closed due to thermal expansion during the fire transient, to 

conservatively maximize heat transfer into the package.

b. Gaps are assumed open, and at nominal “cold” values during the cooldown portion of 

the transient, to conservatively limit heat removal from the package.

5. The content of the neutron shield tank is conservatively represented to maximize heat 

transfer through this region during the fire, and minimize it during the post-fire cooldown.

a. Initial steady-state is represented with the effective conductivity model from the SAR, to 

account for natural circulation of the neutron shield liquid. This model is used in the fire

transient until the peak liquid temperature reaches 276°F (136°C), the saturation 

temperature corresponding to the maximum operating pressure for the tank.

b. The liquid is assumed lost when the predicted peak temperature in the neutron shield 

region exceeds 276°F (136°C). Thermal energy absorbed in the vaporization of the 

liquid is conservatively neglected.

c. After loss of the liquid, heat transfer between the inner surface of the NS tank outer shell 

and outer surface of the package body is assumed to consist of thermal radiation and 

conduction through air for the remainder of the fire and post-fire cooldown transient. The 

inner surfaces of the tank are assumed to be affected by soot, and the emissivity is 

conservatively specified at 0.9. Mainly because of the high thermal radiation heat flux at 

the elevated fire temperatures, this results in a higher heat transfer rate into the package 

through the neutron shield during the fire than would be achieved with only conduction 

and natural convection heat transfer through the neutron shield liquid, if it were assumed 

that the neutron shield tank did not rupture during the fire.

6. The exterior surface of the neutron shield tank is assumed to have an emissivity of 0.15, as 

specified in the SAR, for the initial pre-fire steady-state calculation. At the start of the fire, 

the package surface emissivity is set to 0.9, to represent the effect of sooting of the outer 

surface of the package and impact limiters. This value is also used throughout the post-fire 

cooldown. (This is slightly more conservative than the value of 0.8 to 0.85 documented in 

the SAR for the package surfaces in the HAC fire.)

7. Convection heat transfer during the fire is conservatively modeled assuming forced 

convection to the package from the hot external environment. (See Section 6.3.1 for 

discussion of the specific correlations used.)

 1 Values used are for Allegheny Ludlum ATI 50TM Alloy (UNS S20910), Type XM-19. See the Technical 

Data Sheet in Appendix A.
6-4

8. For the post-collapse configuration (after 37 minutes), the package is assumed to be fully 

covered by the fallen I-580 span between Bent 19 and Bent 20, while simultaneously 

exposed to a fully engulfing pool fire with an uniform temperature of 1652°F (900°C). This 

modeling approach conservatively places the GA-4 package in the most adverse 

configuration possible relative to the collapsed overhead spans, and at the same time yields 

a bounding fire exposure for the package.

9. Convection heat transfer from the upper surface of the fallen I-580 roadway in the postcollapse fire scenario is treated as free convection from a horizontal plate with heated 

surface facing upward. For the lower surface of the I-880 roadway beneath the package, 

heat transfer is treated as free convection from a horizontal plate with heated surface facing 

downward. (See Section 6.3.1 for discussion of the specific correlations used.)

10. For the post-fire cooldown portion of the transient, the air temperature is assumed to be 

100°F (38°C) with insolation, to conservatively bound long-term ambient conditions.

11. During the fire, the aluminum impact limiters in the ANSYS model are assumed to remain 

intact within their stainless steel outer shells, and are represented with effective thermal 

material properties for the honeycomb material, based on bulk density and thermal 

conductivity of the component materials. This assumption maximizes heat input to the 

package during the fire by conduction through the impact limiters. 

12. After the fire, the elements representing the honeycomb material of the impact limiters in the 

ANSYS model were modified to account for melting of the aluminum. This assumption 

maximizes the thermal resistance to heat removal from the package by conduction through 

the impact limiters. Unmelted portions were treated as a combination of aluminum 

honeycomb and air, and melted portions were assumed to have the thermal properties of 

aluminum alloy 5052. It was also assumed that the molten aluminum would settle to the 

bottom of the impact limiters. The effective thermal material properties of the various 

elements of the impact limiters affected by melting were calculated using a volumeaveraging scheme. (Section 5.2 discusses this modeling approach in detail.)

13. The latent heat absorbed by the honeycomb material in the phase change due to melting 

was conservatively neglected. 

14. The effect of the conveyance carrying the GA-4 package is conservatively neglected, both 

for the thermal analysis of the effects of the fire and the structural analysis of the effects of 

the collapse of the upper highway span. In the thermal analysis, the fire is assumed fully 

engulfing, and any shielding effect that the conveyance might have is neglected. In the 

structural analysis, the conveyance would tend to increase the compliance of the system, 

and would also decrease the drop distance between the falling roadway and the GA-4 

package. Both factors would mitigate the effects of the impact of the upper roadway falling 

on the package. The impact is assumed to occur with the package lying on the lower 

roadway, to maximize the drop distance and the overall stiffness of the system.

6.3 Thermal Boundary Conditions for GA-4 Package Models

The boundary conditions for the thermal analysis define the external environment that the GA-4 

package experiences during the fire and post-fire cooldown. These are specifically defined in 

Section 6.3.1. The boundary conditions for the structural analysis define the physical behavior 

of the overhead I-580 span that falls onto the SNF package, the temperature of the falling 

roadway and girders, and the temperature of the package at the time of impact. These 

boundary conditions are discussed in Section 6.4.
6-5

6.3.1 Thermal Boundary Conditions

As described in the detailed description of the fire in Section 2.0, the MacArthur Maze fire began 

as a large open pool fire that lasted for approximately 17 minutes, until the collapse of the 

overhead roadway span between Bent 19 and Bent 20. During the following 20 minutes (17 to 

37 minutes of the total fire duration), the partial collapse of the span between Bent 18 and 

Bent 19 reduced the size and extent of the large open pool fire to a much smaller fire in the 

approximately 12-15 ft (3.7-4.6 m) gap between the collapsed spans at Bent 19. This smaller 

fire burned steadily until the end of the fire, at approximately 108 minutes. 

This fire scenario is conservatively modeled as a fully engulfing fire lasting 108 minutes, with 

two different constant bounding fire temperatures, as shown in Figure 6.1. For the first 

37 minutes, the fire is represented with a uniform flame temperature of 2012°F (1100°C). For 

the remaining 71 minutes, the fire is represented with a uniform flame temperature of 1652°F 

(900°C). The fire emissivity is specified at 0.9. This value is also applied to all external 

surfaces of the package and roadway exposed to the fire, including the inner surfaces of the 

failed neutron shield tank, to conservatively represent the effect of sooting.

Figure 6.1. Fire Boundary Temperatures

As noted in Section 6.2 in the presentation of modeling assumptions, convection heat transfer at 

the SNF package surface during the fire was treated in both models as forced convection. 

Along the package body (for both the ANSYS and COBRA-SFS models) and the sides of the 

impact limiters (in the ANSYS model), the Nusselt number is defined using a correlation (Kreith 

and Bohn 2001)1 for axial flow over a flat or slightly curved surface at zero angle of attack, and 

has the form

 1 Source reference for this correlation is Kreith; primary references cited in Kreith are Rohsenow, 

Patankar and Spalding, and Bejan.

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

1300

1400

1500

1600

1700

1800

1900

2000

2100

0 10 20 30 40 50 60 70 80 90 100 110

Elapsed Time (minutes)

Temperature (oF)

bounding flame temperature

abrupt collapse of Bent 19-20 span

Bent 18-19 span reaches final configuration

end of fire
6-6

b d

L Nu = a(Re + c) Pr (turbulent regime; ReL > 5.0x(105

), Pr > 0.5) (6-1)

where 

a = 0.036

b = 0.8

c = -23,200

d = 0.3333

L = characteristic length (in this case, the exposed package body or impact 

limiter side)

Pr = Prandtl Number

ReL = ρU∞ L / μ

where 

ρ = fluid density

U∞ = free-stream external velocity

μ = fluid viscosity

On the package ends (in the COBRA-SFS model), and on the flat ends of the impact limiters (in 

the ANSYS model), the Nusselt number is defined using a correlation for forced convection from 

an isothermal disk or circular plate with axis perpendicular to the flow direction (Kreith and Bohn 

2001). The correlation has the form

b c

D Nu = a Re Pr (900 < ReD < 30,000) (6-2)

where 

a = 0.591

b = 0.564

c = 0.3333

Pr = Prandtl Number

ReD = ρU∞D / μ

where 

ρ = fluid density

U∞ = free-stream external velocity

D = diameter of disk or plate

μ = fluid viscosity

In the above correlations, fluid properties are evaluated at the near-wall temperature, defined as 

the average of the wall surface temperature and the ambient temperature, which in this 

application is the fire temperature. The free-stream external velocity was specified at a 

bounding value of 12 ft/s (3.7 m/s), based on the velocities predicted in the FDS simulation of 

the open pool fire.
6-7

In the post-fire cooldown, with the package beneath the concrete “blanket” of the collapsed 

upper roadway span, the boundary conditions on the external roadway surfaces were specified 

as free convection to ambient still air. For the COBRA-SFS model, the boundary conditions on 

the SNF package were determined by solving for the flow and heat transfer within the irregular 

annulus between the package and the concrete roadway surfaces. The air velocity within the 

annulus was assumed to be essentially stagnant in the post-fire cooldown. 

In typical open pool fires, it is conservative to assume free convection to ambient for the postfire cooldown when evaluating heat transfer from an object such as an SNF package which had 

been engulfed in the fire. For a horizontal cylinder, correlations typically yield Nusselt number 

values on the order of 100-200, which for a cylindrical object the size of this SNF package 

produces heat transfer coefficient values on the order of 0.8 to 1.5 Btu/hr-ft2

-°F (4.5 to 8.5 W/m2-

°K). However, for the package beneath the concrete roadway, this assumption is neither 

conservative nor correct, since the presence of the roadway would severely inhibit the 

development of natural convection flow patterns around the package. 

In the COBRA-SFS analysis, convection heat transfer in the annulus channel between the 

package outer surface and interior surfaces of the concrete was represented with a bounding 

value of Nu = 3.66 for laminar flow, based on experimental studies (Creer et al. 1987) of heat 

transfer in enclosed spaces within a horizontal package. This approach conservatively treats 

the “blanket” of the collapsed roadway as an enclosure. After the end of the fire, the hot air in 

the irregular annular flow channel is assumed to be optically transparent, allowing thermal 

radiation heat transfer via gray body exchange factors between the package surface and interior 

surfaces of the overlying upper roadway, side barrier, and lower roadway. The temperature of 

air trickling into the annulus due to the changing temperature and pressure gradients in the flow 

channel is assumed to be 100°F (38°C). 

In the analysis with the ANSYS model, thermal radiation exchange between the package and 

the enclosing concrete was calculated directly, as described in Section 5.2.2 above. However, it 

was assumed that the impact limiters would effectively block air flow within the annulus, so 

convection heat transfer between the package exterior surfaces and the enclosing concrete 

surfaces was conservatively neglected. This was modeled by specifying Nu = 1.

The temperatures of the concrete surfaces surrounding the package at the beginning of the 

cooldown portion of the transient were determined assuming that these structures had also 

been exposed to the same fire boundary temperatures of 2012°F (1100°C) and 1652°F (900°C), 

for the same time intervals as the SNF package. This approach establishes bounding concrete 

surface temperatures for the post-fire cooldown portion of the transient. The effect of spalling of 

the concrete, which dissipates thermal energy and therefore would tend to result in lower 

concrete temperatures, was conservatively neglected. 

Heat transfer to the external environment from the concrete surfaces of the upper and lower 

roadways was represented with free convection from these surfaces (Kreith and Bohn 2001)1

and thermal radiation to ambient. For the upper surface of the fallen upper roadway span, the 

correlation used is for free convection from a horizontal heated surface facing upward, and is of 

the form

 1 Source reference is Kreith; primary references are MacAdams, plus Incropera and DeWitt.
6-8

b Nu = aRa L (6-3)

where

RaL = Rayleigh Number (GrLPr)

where 

L = (surface area)/perimeter

For laminar flow (105 < RaL < 107

), the coefficients are

a = 0.54

b = 0.25

For turbulent flow (107 < RaL < 1012), the coefficients are

a = 0.15

b = 0.3333

For the lower surface of the lower roadway, the correlation is of the same form, but with 

coefficients for free convection from a horizontal heated surface facing downward

b Nu = aRa L (105 < RaL < 1010) (6-4)

where 

a = 0.27

b = 0.25

In the above correlations, fluid properties are evaluated at the near-wall temperature, defined as 

the average of the wall surface temperature and the ambient temperature. The ambient 

temperature external to the fire is conservatively assumed to be 100°F (38°C), as a convenient 

bounding value.

6.4 Structural Impact Model Assumptions and Analysis Method

Because of the complexity of the impact portion of the MacArthur Maze fire scenario, the 

structural model of the fallen overhead span was developed in progressive stages, starting with 

a simple approximation using only a truncated section of a single plate girder for scoping 

calculations, and evolved to include representation of the complete overpass section. The 

purpose of the preliminary structural modeling was to determine the level of complexity and 

sophistication in the final set of impact models needed to assure that all potentially significant 

structural consequences to the containment boundary were considered. The initial modeling 

work included mesh sensitivity studies and evaluation of the effects of including or neglecting 

the conveyance. A primary conclusion of the initial modeling phase was that the thin plate 

girders would tend to deform around the stronger package wall instead of causing significant 

local deformation to the package, making a failure of the containment boundary unlikely. The 

more realistic final impact models described in this report were carefully crafted to include the 
6-9

upper bound of potential package response without resorting to incredible conservatism. The 

basic features of the structural models are described in Section 5.4.1.

Fire exposure raises the package temperature, which lowers the structural material strength and 

makes the package more susceptible to damage from impact loading. Therefore, a 

conservative estimate of the maximum package temperature was obtained from the thermal 

analyses, to appropriately bound the package temperature at the time of the collapse of the 

overhead span. As described in Section 2.0, the collapse of the first span occurred at 

approximately 17 minutes, and the partial collapse of the second span was complete at about 

37 minutes after the start of the fire. The thermal analysis (see Section 6.3) conservatively 

assumed that the large fully engulfing (pre-collapse) pool fire lasted for the full 37 minutes. For 

the purposes of the structural analysis, the peak temperature on the package surface predicted 

at 37 minutes was used to define the overall package temperature, conservatively assuming 

that the entire package was at the peak temperature. This approach results in an estimated 

bounding temperature of 1800°F (982°C) for the package in the structural analyses. 

The results of the FDS model of the pre-collapse fire (see Section 3.0) and an additional 

ANSYS thermal model of the I-580 roadway were used to determine a reasonably conservative 

uniform temperature for the falling steel girders of the overpass structure. In the case of the 

falling girders, the conservative condition errs on the cool side, so the strength and stiffness in 

the girders is as high as possible, within the thermal constraints of the scenario. In this case, 

1800°F (982°C) was chosen as the representative girder temperature for the drop scenario. 

Coincidentally, this value is the same as the assumed package temperature, which was chosen 

using a different rationale, as described in Section 6.4.1. The package temperature was chosen 

to be bounding with as high a value as possible; the girder temperature was chosen to be as 

conservatively low as possible, based on realistic package locations. 

The temperature of the lower I-880 span at the location of the GA-4 package is assumed to be a 

nominal 80°F (27°C). This conservatively neglects elevated temperature effects on the material 

properties of the lower roadway. The roadway would have been much hotter during the fire, 

and would have been more compliant at a higher temperature than the assumed bounding 

value. Assuming higher, more realistic temperatures for the lower roadway would allow it to 

dissipate more of the energy of the impact, mitigating to some degree the effect of impact on the 

package. Similarly, air drag forces and reduction in fall distance resulting from sagging of the 

girders before collapse of the span are conservatively neglected in the impact analysis. 

Yield strength and the plastic behavior of the materials are critical parameters in the impact 

model. The evaluation of the package containment is based on the amount of plastic strain 

developed in the XM-19 stainless steel package body. The significant capacity of the plate 

girders to deform during impact because of the elevated temperature, visible in the post-fire 

photographs of the girders (CHP 2007), is a fundamental physical phenomenon captured in the 

analysis with this model. 

The DU gamma shield is not a structural component of the containment boundary, but it does 

offer potential support to the steel body wall. The DU is modeled as a bilinear (elastic/plastic) 

material with a near-zero tangent modulus. If the yield stress limit is exceeded, the material is 

allowed to deform with a negligible increase in load. This behavior results in the DU layer 

providing realistic support to the package as long as it remains in the elastic range, but the 

support drops off quickly if the stress in the material exceeds the elastic range. The tangent 

modulus is six orders of magnitude lower than the elastic modulus for this material, resulting in 

calculated plastic strains that are conservatively higher than they would be if a realistic stress-
6-10

strain curve were implemented. However, the response of the DU gamma shield is not 

specifically evaluated for failure in this analysis, since it is not critical to the containment 

boundary.

In the calculations, contact between a falling girder and the package body or any other 

component of the model was defined with an automatic single surface contact definition. This 

contact rule prohibits the components from occupying the same space, and applies a coefficient 

of static and dynamic friction of 0.5. Dry steel-to-steel friction is typically reported in the range of 

0.7-0.8 for static friction and 0.4-0.6 for dynamic friction, but there is significant variation and 

instances of data reported outside those ranges. The 0.5 was chosen as a middle range 

dynamic friction value, based on sensitivity studies that showed friction was not a major force in 

this scenario. The relative fraction of frictional energy to total system energy was in the range of 

1-2%, while the fraction of frictional energy to internal energy (due to material deformation) was 

in the range of 1-5%.

Global mass damping of 5% was applied to the model as an additional source of energy 

dissipation. This level of damping is typically used in dynamic impact modeling for numerical 

stability and to account for natural sources of energy dissipation that are not specifically 

modeled. The amount of energy this damping dissipates in the model is comparable to the 

amount of energy dissipated by friction. Global damping does not have a significant effect on 

the results obtained in this analysis.

6.4.1 Overpass Temperature Evaluation

The temperature of the I-580 overpass at the time of impact is an important assumption that 

affects its material characteristics and the overall impact response. The elastic modulus and the 

yield strength of the girder material are two material properties that affect the force and energy 

transmission during impact, and both material properties decrease at elevated temperatures. 

As discussed in Section 2.2, evidence gathered from the post-collapse material indicated that 

the plate girders experienced temperatures at least as high as 1472-1652°F (800-900°C) at 

locations near the middle of the fire. Material data in the literature for this temperature range 

predict a reduction in the yield strength of steel to values roughly 10% of the room-temperature 

yield strength (see Figure 2.10), and a substantial increase in ductility. 

To obtain a more detailed picture of the temperature distribution on the girders than could be 

determined from the relatively small range of sampling locations, additional analysis was 

performed with an ANSYS model of the I-580 upper roadway span between Bents 18 and 20. 

ASTs from the FDS model of the pre-collapse fire (discussed in Section 3.0), were used as 

boundary conditions for thermal analysis of the girders. The AST boundary values were applied 

as time-history thermal loads over the interval from initiation of the fire to 18 minutes, 

encompassing the time frame of the complete collapse of the overhead span between Bent 19 

and Bent 20. The lower temperature corresponding to the actual time-frame of the collapse of 

this span is more conservative for the purposes of the structural analysis. However, a higher 

temperature is more conservative for the package temperature, and the package temperature 

for the impact analysis was based on a fire exposure time of 37 minutes, corresponding to the 

time of final collapse of both overhead spans.

The thermal model of the upper spans predicts a rapid heating of the steel girders directly above 

the fire. By 5 minutes into the fire, the peak girder temperature exceeds 2192°F (1200°C). At 

18 minutes, the temperature distribution reaches the state shown in Figure 6.2, with peak 

temperatures up to 2462°F (1350°C). The left half of the thermograph shows the hotter side of 
6-11

the overpass, which is directly above most of the fire area, beneath the span between Bents 18 

and 19. The right side is generally cooler, and represents the span that fell straight down and is 

modeled in the impact analyses. This is consistent with the physical distribution of the fire pool 

(as shown in the fire video images in Figure 2.2, Figure 2.3, and Figure 2.4). Over much of the 

girder length exposed to the fire, the ANSYS results show girder temperatures that are in the 

range determined from the material analysis. (This corresponds to the light green to yellow 

segments of the color thermograph in Figure 6.2). However, this calculation also predicts much 

higher temperatures for the girders in locations that were not sampled for the material analysis. 

Figure 6.2. Girder Temperatures, (°C [°F]), at 18 Minutes Predicted with ANSYS Model

Sensitivity studies on thermal processes that would carry heat away from the girders, such as 

conduction heat transfer longitudinally down the girders and conduction into the concrete above 

the girders, showed that these processes do not play a major role in determining the girder 

temperature or temperature distributions. The local heat input to the girders due to the fire, as 

defined in the ANSYS model using the AST values from the FDS fire model, dominates all other 

thermal processes, even when conservatively high heat dissipation through the girders is 

assumed, by such means as enhanced convection coefficients between the upper roadway 

surface and the cool air above it. 

As illustrated in Figure 6.2, the peak girder temperatures are predicted to be as high as 2462°F 

(1350°C). This very high temperature zone includes girder segments above potential impact 

locations for the SNF package. However, from the standpoint of the drop scenario, higher 

girder temperatures would result in a less severe impact, due to the greatly decreased yield 

strength of steel with increasing temperature. Using these very high temperatures to define the 

girder temperatures for the impact analysis would result in a less severe drop scenario than 

would be obtained by assuming lower temperatures for the girders. The goal in assigning a 

temperature to the girder structure was to ensure the temperature was conservatively low, 

giving the structural steel of the girders conservatively high strength and stiffness, but remaining 

consistent with the range of temperatures that could reasonably be expected from the accident 

conditions. 
6-12

A uniform overpass girder temperature of 1800°F (982°C) was chosen as a conservatively low 

temperature estimate for the hot side of Bent 19. The geometry of overhead roadway for the 

impact models actually corresponds to the right half of Figure 6.2, which is the cooler side, 

because that is the span that fell relatively cleanly onto the lower roadway in the MacArthur 

Maze accident (i.e., the span between Bents 19 and 20), and offers the potentially most 

damaging mechanical loading. The thermal state of the impact model is defined based on 

conditions from the hot side, to ensure the package is at its hottest and structurally weakest.
7-1

7.0 ANALYSIS RESULTS

This section presents the results of the structural and thermal analyses of the GA-4 package 

exposed to the conditions of the MacArthur Maze fire scenario. The thermal analysis of the fire 

portion of the transient is presented in Section 7.1, for both the ANSYS model and the COBRASFS model. The thermal analysis of the post-fire cooldown transient with these models, in 

which the package is covered by the concrete “blanket” of the collapsed upper roadway, is 

presented in Section 7.2. The results of the structural analysis are discussed in Section 7.3.

7.1 GA-4 Package: Thermal Results for Fire Transient

The starting point for the fire transient thermal modeling was assumed to be steady-state NCT. 

The initial temperature distribution in the package is summarized in Figure 7.1, which shows the 

peak component temperatures predicted with the ANSYS and COBRA-SFS models of the GA-4 

at NCT. Figure 7.2 shows a color thermograph illustrating the ANSYS model temperature 

results for the package cross-section for this initial steady state evaluation. As discussed in 

Section 6.1 above, the peak fuel region temperature predicted with the ANSYS model is 306°F 

(152°C), and the peak cladding temperature predicted with the COBRA-SFS model is 293°F 

(146°C). This is consistent with the expected differences between the results obtained with a 

homogeneous representation of the fuel region using a k-effective model when compared to a 

detailed thermal-hydraulic model of the fuel assemblies. The two models give essentially the 

same results for NCT, within modeling uncertainty, and are in reasonable agreement with the 

results reported in the GA-4 SAR (General Atomics 1998).

Figure 7.1. Initial Conditions for Fire Transient Analyses: GA-4 at NCT

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Peak Fuel

Clad

basket inner steel

liner

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Shield

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Shield

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Peak Component Temperature (°F)

GA-4 SAR

ANSYS model

COBRA-SFS model (without impact limiters)
7-2

Figure 7.2. ANSYS Results1

: Thermal Cross-section (°F) of GA-4 Package at NCT

The MacArthur Maze fire was simulated for the GA-4 package by imposing boundary conditions

representing the large (pre-collapse) engulfing fire, followed by the smaller (post-collapse) 

engulfing fire. To simulate the pre-collapse fire, the package model was subjected to an 

ambient boundary temperature of 2012°F (1100°C) for 37 minutes. The detailed results for the 

fully engulfing fire at 2012°F (1100°C) are described in Section 7.1.1. To simulate the postcollapse fire, the fire boundary temperature was reduced to 1652°F (900°C) for the remaining 

71 minutes of the fire portion of the transient. The results obtained for the post-collapse portion 

of the fire are described in Section 7.1.2. 

7.1.1 Pre-collapse Fire (2012°F [1100°C])

The initial fully engulfing pool fire of the MacArthur Maze fire scenario at 2012°F (1100°C) was 

assumed to extend to 37 minutes of the total fire duration, to conservatively encompass the 

collapse of both of the upper roadway segments. Figure 7.3 shows a color thermograph of the 

package central cross-section illustrating the temperature results obtained with the ANSYS 

model at the end of this portion of the fire. (Note that the TIME reported on the plots includes 

the initial pre-fire pseudo-transient of 0.5 hr for NCT.)

 1 To ensure a smooth transition within the ANSYS calculation between the NCT steady-state analysis and 

the transient fire analysis, the NCT analysis was run as a transient solution with an arbitrary time-step, 

updating temperature-dependent material properties and external convection coefficients until the solution 

did not change significantly between time-steps. Time-stamps on graphics produced using ANSYS 

include the arbitrary 0.5 hrs of the NCT analysis, and therefore are off-set by 0.5 hr compared to other 

plots referenced to the start of the fire as time zero.
7-3

Figure 7.3. ANSYS Model Results: Axial and Central Radial Thermal Cross-sections (°F) of 

GA-4 Package at End of 2012°F (1100°C) Fully Engulfing Fire (37 minutes)
7-4

Peak temperatures predicted with the ANSYS model for the main components of the GA-4 

package are shown as a function of time in Figure 7.4. The peak temperature on the outer 

surface of the package, which consists of the thin stainless steel shell of the neutron shield tank, 

rises rapidly in the initial minutes of the fire, paralleling the rapid rise in the exterior ambient 

temperature. Within approximately 5 minutes, the surface temperature of the package is 

approaching the fire temperature on an asymptotic curve. In a similar manner, the peak 

temperature on the outer surface of the impact limiters very rapidly approaches the fire 

boundary temperature. This temperature occurs on the thin stainless steel outer surface of the 

impact limiter, which is effectively insulated on the inner side by the relatively low thermal 

conductivity of the honeycomb structure that the steel is there to protect. The temperature of 

the outer stainless steel surface therefore approximates an adiabatic surface temperature at the 

local fire boundary condition.

Figure 7.4. Peak Component Temperatures Predicted with ANSYS Model for 37-minute Fully 

Engulfing 2012°F (1100°C) Fire

The peak fuel region temperature does not begin to show any increase until about 9 minutes 

into the fire transient. Such a delay is typical of the peak fuel temperature response to a fire 

transient, because of the thermal inertia of the massive body of an SNF package. In addition, 

the location of the maximum temperature in the fuel region shifts toward the outer edge of the 

hottest fuel assembly as the fire transient progresses, and part of the delay is due to the time 

required to heat up the region of the fuel assembly that is initially cooler than the central region 

in the pre-fire steady state at NCT. 

Peak temperatures predicted with the COBRA-SFS model for the major components of the GA4 package in response to the fire conditions are shown in Figure 7.5. The plots in Figure 7.4

and Figure 7.5 show that the two models are predicting similar responses for the GA-4 package 

in this severe fire transient. The significant difference between them is in the response of the 

fuel region. The ANSYS model uses the homogeneous k-effective model for the fuel region, 

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0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37

Elapsed Time (minutes)

Peak Temperature (°F)

fuel region basket stainless steel liner

DU gamma shield stainless steel body neutron shield contents

outer surface impact limiter surface end of 1100°C fire

short-term PCT limit bounding fire temperatures
7-5

which is a steady-state model that is by design a conservative representation of heat transfer 

within the rod array. The more detailed rod-and-subchannel representation of the fuel assembly 

in the COBRA-SFS model results in a more realistic prediction of the rate of fuel cladding 

temperature rise. 

Figure 7.5. Peak Component Temperatures Predicted with COBRA-SFS Model for 37-minute 

Fully Engulfing 2012°F (1100°C) Fire

The COBRA-SFS model predicts that the peak cladding temperature near the ends of the rods, 

which in this model are not protected by the impact limiters, rises to 955°F (513°C) by 

37 minutes. The COBRA-SFS model predicts 677°F (358°C) by this time for the peak cladding 

temperature at the midplane of the package, the location where the ANSYS model predicts the 

peak temperature in the fuel region. Because of the more conservative homogeneous keffective model for the fuel region, the ANSYS model predicts a somewhat higher peak 

temperature of 1023°F (551°C) by 37 minutes. The cladding peak temperatures predicted with 

both of these models have not exceeded the limit of 1058°F (570°C) for short-term operations 

by the end of the 37-minute bounding representation of the pre-collapse fire. However, the total 

fire duration has 71 more minutes to run, at the lower fire temperature of 1652°F (900°C), 

simulating the post-collapse portion of the fire. 

7.1.2 Post-collapse Fire (1352°F [900°C])

The post-collapse portion of the MacArthur Maze fire, in which the physical extent of the fire was 

greatly reduced by the intrusion of the fallen upper roadway spans, was modeled as a fully 

engulfing fire at the lower boundary temperature of 1652°F (900°C). The temperatures 

predicted with the ANSYS model for the end of the fire, after 37 minutes of exposure to 2012°F 

(1100°C) followed by 71 minutes of exposure to 1652°F (900°C), are shown in Figure 7.6.

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stainless steel liner DU gamma shield stainless steel body

neutron shield contents outer surface end of 1100°C fire

short-term PCT limit bounding fire temperatures
7-6

Figure 7.6. ANSYS Model Results: Axial and Central Radial Thermal Cross-sections (°F) of 

GA-4 Package at the End of the Fire at 108 Minutes
7-7

Peak temperatures predicted with the ANSYS model for the main components of the GA-4 

package are shown as a function of time in Figure 7.7. Figure 7.8 shows the response of major 

components of the package to the fully engulfing post-collapse fire predicted with the COBRASFS model. By the end of the 37-minute fire at 2012°F (1100°C), both models predict peak 

temperatures on the outer surface of the package that are significantly above the post-collapse 

fire temperature of 1652°F (900°C). As a result, the initial response of these components during 

the post-collapse fire is to cool down toward the new ambient fire temperature. In the first 

13 minutes of this portion of the transient, the peak temperatures on the outboard components 

drop rapidly in response to the sudden change in the boundary fire temperature, then stabilize 

to a slow asymptotic rise toward the new fire temperature.

Figure 7.7. Peak Component Temperatures Predicted with ANSYS Model at 108 Minutes

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fuel region basket stainless steel liner

DU gamma shield stainless steel body neutron shield contents

outer surface impact limiter surface end of 1100°C fire

end of fire short-term PCT limit bounding fire temperatures
7-8

Figure 7.8. Peak Component Temperatures Predicted with COBRA-SFS Model at 108 Minutes

Similarly, the peak fuel temperatures respond to the lower boundary temperature by slightly 

slowing their rate of rise, but these components are also on an asymptote approaching the new 

fire temperature. By the end of the fire transient at 108 minutes, the peak fuel region 

temperature predicted with the ANSYS model has reached 1433°F (779°C), exceeding 1382°F 

(750°C), the temperature at which burst rupture of zircaloy cladding has been assumed in 

previous SNF package transportation studies (NUREG/CR-6672). The peak fuel cladding 

temperature predicted with the COBRA-SFS model also exceeds this temperature, reaching 

1388°F (753°C) on the end of the hottest fuel rod. The mid-plane peak fuel cladding 

temperature predicted with this model is not far behind, at 1248°F (675°C). 

These predicted peak temperatures are in response to a conservative and bounding 

representation of the MacArthur Maze fire. However, even after the fire is over, the transient is 

not finished, as the package is evaluated in the post-fire cooldown, with the conservative 

assumption that it is blanketed by the concrete structure of the fallen overhead roadway span. 

Experience with modeling of SNF packages in long-duration fires (NUREG/CR-6487; 

NUREG/CR-6886 2009) has shown that the maximum fuel cladding temperature can occur well 

after the end of the fire, during the post-fire cooldown of the package. In addition to the rise in 

temperature on the fuel rods in response to heat input from the fire, some portion of the 

temperature rise is due to the high ambient fire temperature preventing decay heat removal 

from the fuel rods during the fire and for some time after the fire while the outboard components 

of the package are above the maximum fuel temperature.

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fuel cladding fuel cladding (midplane) basket

stainless steel liner DU gamma shield stainless steel body

neutron shield contents outer surface end of 1100°C fire

end of fire short-term PCT limit bounding fire temperatures
7-9

7.2 GA-4 Package: Thermal Results for Post-Fire Cooldown 

Transient

For the cooldown portion of the transient, in which the GA-4 package is enclosed by the fallen 

roadway, the roadway surface temperatures at the end of the fire were determined by a 

separate calculation with COBRA-SFS. This calculation represented the roadway as described 

in Section 5.2.2 above (see Figure 5.7), exposing the roadway surfaces to the same fire 

conditions as the GA-4 package. Figure 7.9 shows the peak temperatures predicted for the 

roadway concrete surfaces in response to the pre-collapse fire at 2012°F (1100°C) and the 

post-collapse fire at 1652°F (900°C). These temperatures exhibit the same asymptotic behavior 

as the outer surface of the SNF package. 

Figure 7.9. Peak Concrete Surface Temperatures Predicted with the COBRA-SFS Model for 

the 37-minute Fire at 2012°F (1100°C) and the 71-minute Fire at 1652°F (900°C)

In Figure 7.9, the upper roadway temperature is the peak temperature on the under surface of 

the upper roadway span that is assumed to cover the SNF package in the post-fire scenario. 

The lower roadway temperature shown in the plot is the peak temperature predicted on the 

surface of the lower roadway beneath the SNF package. The concrete barrier temperature is 

the peak temperature predicted for the inner surface of the lower roadway side barrier, which 

the SNF package is assumed to be lying beside. The peak temperature of 1774°F (968°C) 

predicted for the roadway surfaces occurs at the end of the 37-minute fully engulfing fire. The 

side barrier is predicted to reach approximately 1369°F (743°C) at this time. The temperatures 

on the roadway surfaces drop somewhat during the 71-minute interval of exposure to the lower 

temperature of the post-collapse fire, and by the end of the fire, all of the concrete surfaces 

exposed to the fire are at approximately the same temperature of 1609°F (876°C).

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lower roadway, upper surface

end of 1100C fire

end of fire

MacArthur Maze fire scenario
7-10

The results in Figure 7.9 show that it is extremely conservative to assume that the concrete 

structures would be exposed to the bounding fire temperatures for the full duration of the precollapse and post-collapse fires. As with all other components exposed to these extreme 

boundary conditions, the concrete temperatures begin to asymptotically approach the fire 

temperature. Post-fire evaluation of the damaged roadway, including detailed assessment of 

the lower roadway deck and side barriers (Wiss, Janney, Elstner Associates, Inc. 2007), found 

that the concrete surfaces were exposed to temperatures in the range 446°F (230°C) to 842°F 

(450°C). The maximum temperature exposure at the concrete surfaces was estimated as 

1472°F (800°C), over a relatively small localized area. The bounding approach used in this 

model sets up an extremely conservative environment for the SNF package in the post-fire 

cooldown transient beneath the collapsed roadway. 

The roadway temperatures at the end of the fire define the initial conditions for the cooldown 

transient calculated with the ANSYS and COBRA-SFS models for the GA-4 package beneath 

the concrete ‘blanket’ of the upper roadway. Figure 7.10 shows the response predicted with the 

COBRA-SFS model for the major components of the system in the post-fire cooldown in the first 

5 hours after the end of the fire. Figure 7.11 shows the response predicted with the ANSYS 

model for this portion of the transient. In both models, all major components show an initial 

rapid decrease in peak temperature, followed by a more gradual decline as the temperature 

gradients flatten. 

The results for the COBRA-SFS model show the effects of the rapid cooling of the ends of the 

package, where the peak temperatures are predicted during the fire transient. By about 

2.3 hours, the ends have cooled sufficiently for the location of the peak temperatures to shift

back toward the axial center of the package, and the effects of thermal inertia are seen in the 

peak fuel cladding temperature and the peak basket temperature. For these components, the 

temperatures near the center of the package continue to rise after the end of the fire, and 

eventually exceed the peak temperatures on the ends. This is illustrated in Figure 7.12, which 

shows the evolution of the axial temperature profile on the hottest rod in the package, as 

predicted with the COBRA-SFS model.
7-11

Figure 7.10. Peak Component Temperatures Predicted with COBRA-SFS Model for Post-fire 

Cooldown to 6.5 Hours

Figure 7.11. Peak Component Temperatures Predicted with ANSYS Model for Post-fire 

Cooldown to 6.5 Hours

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outer surface end of fire short-term PCT limit
7-12

Figure 7.12. Axial Temperature Evolution on Hottest Fuel Rod Predicted with the COBRASFS Model for Post-fire Cooldown with the Package under Concrete

The peak fuel cladding temperature near the axial center of the hottest fuel rod is predicted to 

reach a maximum of 1347°F (731°C) at approximately 250 minutes, which is nearly 2.4 hours 

after the end of the fire. The results obtained with the ANSYS model also show the effect of 

thermal inertia for the fuel, sustaining a peak fuel region temperature of nearly 1400°F (760°C) 

for approximately 3 hours after the end of the fire, as shown in Figure 7.11. 

Because the ANSYS model includes the impact limiters, the ends of the package do not 

experience the high temperatures predicted with the COBRA-SFS model, which omits the 

impact limiters. However, in the post-fire cooldown, the impact limiters act as insulators, 

slowing the rate of heat removal from the ends of the package. As the package cools, the steep 

temperature gradients within the package, due to the heat input from the fire and the decay heat 

trapped within the fuel region, result in the cooler ends of the package continuing to increase in 

temperature for some time after the fire, even as the peak temperatures near the center 

decrease. This is illustrated in Figure 7.13, with color thermographs of the package axial crosssection. In these thermographs, the scale of the color coding is held constant, to clearly 

illustrate the increasing temperature on the ends and interior components, while the peak 

temperatures steadily decrease. 

1000

1050

1100

1150

1200

1250

1300

1350

1400

1450

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180

Axial Location (inches)

Temperature (oF)

short-term pct limit at end of fire; rod #196 (assm 3) 149 min.; rod #183 (assm 4)

161 min.; rod #183 (assm 4) 170 min.; rod #183 (assm 4) 180 min.; rod #183 (assm 4)

250 min.; rod #183 (assm 4)
7-13

 (a) at end of fire (1.8 hours)

 (b) 2.2 hours after end of fire

 (c) 6.2 hours after end of fire

 (d) 12.2 hours after end of fire

Figure 7.13. ANSYS Model Predictions of Temperature Distributions (°F) in GA-4 Package 

during Post-fire Cooldown
7-14

As shown in Figure 7.13, at the end of the fire the temperatures at the ends of the fuel region 

(beneath the impact limiters) are relatively cool, but continue to increase in temperature for 

many hours after the fire is over. The temperature increase in the end region of the fuel is 

summarized below.

By 12.2 hours, the impact limiters and outer shell of the package are at temperatures in the 

range 1034°F to 1122°F (557°C to 606°C), slightly cooler than the end of the fuel region. The 

package begins to experience a uniformly decreasing temperature at all points, including the 

sheltered locations within the package beneath the impact limiters, only after about 12.5 hours

beyond the end of the fire. 

This behavior is more precisely illustrated with the temperature plot in Figure 7.14 for three 

specific locations beneath the top impact limiter in the ANSYS model of the package. This 

figure shows the temperature histories during the cooldown for a point in the fuel region near the 

lid end of the package, a point in the top impact limiter immediately adjacent to the package lid, 

and a point near the upper end of the DU gamma shield.

Figure 7.14. ANSYS Model Predictions of Temperature History during the Post-fire Cooldown 

at Locations in the GA-4 Package Shielded by the Top Impact Limiter

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

1300

1400

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

Elapsed Time (hours)

Temperature (oF)

fuel region near lid end

Impact limiter, near lid surface

end of DU, near lid end

end of fire

Timeframe Temperature Range at End of Fuel Region

end of fire (108 minutes) 328°F to 416°F (164°C to 213°C)

2.2 hours after end of fire 725°F to 1122°F (385°C to 606°C)

6.2 hours after end of fire 1122°F to 1255°F (606°C to 679°C)

12.2 hours after end of fire 1167°F to 1255°F (630°C to 680°C)
7-15

The overall results of the thermal analysis with the ANSYS and COBRA-SFS models show that 

a fire of the severity of the MacArthur Maze fire scenario would result in extremely high 

component temperatures on an SNF package such as the GA-4, and that those temperatures 

would be sustained over a long period of time. In particular, fuel cladding temperatures are 

predicted to exceed the short-term limit of 1058°F (570°C) for several hours, over essentially the 

full axial length of the fuel assemblies. In addition, the temperatures in the regions of the 

package seals exceed the seal material operating temperature limits for most of the fire 

transient and for many hours of the post-fire cooldown transient. The potential consequences of 

this extended period at elevated temperatures are discussed in Section 8.0.

7.3 GA-4 Package: Structural Evaluation

The positioning of the SNF package on the lower roadway has a significant effect on the 

potential consequences of the overpass collapse. It affects the freefall distance for the 

overhead span, and the location(s) at which the girder(s) can strike the package. In addition, 

the alignment of the center of gravity (CG) of the package and the CG of the falling span affect 

the total energy of impact imparted to the package. All other things being equal, the more 

closely the two CGs are to the same vertical line, the greater the amount of energy transmitted 

by the impact. 

It is not obvious what impact orientation could do the most damage to the package, and in this 

analysis, four different configurations are considered, to cover a range of possible impact 

behavior and to explore the effects of the most significant variables. In all cases considered, the 

package location on the lower roadway is as close as possible to the center of gravity of the 

falling span, within the constraints of the geometry of the lower and upper roadway structure. 

The cases evaluated are summarized as follows:

• Case #1: a girder strikes directly across the center of the package (Figure 7.15)

• Case #2: a girder strikes along the axial length of the package (Figure 7.16; the package is 

barely visible beneath the third girder from the right)

• Case #3: a girder strikes directly on the package lid (Figure 7.17)

• Case #4: a girder strikes directly onto the package lifting trunnions, which are assumed to be 

oriented vertically, such that the impact applies a highly localized load to the package wall 

(Figure 7.18).
7-16

Figure 7.15. Impact Case #1: GA-4 Package Perpendicular to Upper Roadway Girders; Main 

Impact on Package Center

Figure 7.16. Impact Case #2: GA-4 Package Parallel to Upper Roadway Girders; Main Impact 

along Axial Length of Package
7-17

Figure 7.17. Impact Case #3: GA-4 Package Oriented to Yield Main Impact on Package Lid

Figure 7.18. Impact Case #4: GA-4 Package Oriented for Trunnion Impact

The results of the analysis for all four of these cases show that the steel plate girders of the 

overhead roadway are more severely affected by the impact than the SNF package. The 

calculations show that the impact would cause significant plastic strains in the girders, such that 

the metal plates would deform under the impact, while the SNF package is relatively unaffected 
7-18

by the impact force. Localized plastic strains are predicted in the package wall and the DU 

gamma shield, but these are much smaller than the strains predicted for the girders. 

For the conditions of this scenario, it is expected that failure mode for the steel of the girders 

would be ductile fracture. Uniaxial tensile test data for A36 steel suggest that for conditions in 

the temperature range of the MacArthur Maze fire scenario, the elongation capacity for this 

material is over 50% and the reduction of area is over 95%. With this amount of available 

ductility, the calculated effective plastic strains below 20% appear to be well within the material’s 

capacity. The girders would be severely deformed, but would be unlikely to fracture. This is 

illustrated conceptually in Figure 7.19, comparing a photo of a damaged girder from the actual 

roadway collapse following the fire to a graphic image from the LS-DYNA calculation for 

Case #1.

 

Figure 7.19. Illustration of Actual and Calculated Failure of Overhead Roadway Girders

 (photo from MAIT Report, CHP 2007, reprinted with permission.)

The maximum peak effective plastic strain in the XM-19 stainless steel of the package wall is 

calculated as 11 percent, and occurs in Case #2. An industry source lists short-term elevated 

temperature elongation data (Nitronic 50 Product Data Bulletin) for this material as having

maximum elongation values between 41% and 59%t at 1500°F (816°C). Based on this data, 

the local plastic strains are not sufficient to cause structural failure of the steel components of 

the GA-4 package. In each impact orientation, the plastic strain regions are further evaluated 

based on relative size and penetration into the volume of the material or through the wall. In all 

cases, the localized plastic strains represent surface damage, not gross deformation of the steel 

package body.

The DU gamma shield is included in the model to give the steel package body realistic support. 

It is not considered a structural member in the GA-4 design, but the package SAR (General 

Atomics 1998, Section 2.1.2.2.2) notes that the DU may transfer compressive loads and provide 

backing for the containment wall. It is modeled to perform this function in a conservative 

manner. Since the DU is not a part of the containment boundary, it is not specifically evaluated 

for failure. Plastic deformations in the DU are reported and plotted to show the locations where 

the containment wall required maximum support from the DU structure. 
7-19

7.3.1 Structural Case #1: Package Perpendicular to Upper Roadway Girders

The axial length of the GA-4 package and the spacing of the I-580 girders permit direct 

perpendicular impact on the package from only two of the girders simultaneously, with a narrow 

miss from a third. If one of the impacting girders strikes the package near its mid-point, the 

second one strikes on one end. Figure 7.20 shows a diagram of the pre-impact configuration 

for this scenario. 

Figure 7.20. Case #1: Pre-impact Geometry—Girders Perpendicular to SNF Package

Figure 7.21 shows the deformation of the upper roadway girders after 0.205 s, which is after the 

maximum load has been transferred to the package. The calculation continues to 0.4 s, and the 

plastic strain in the package body at this time is shown in Figure 7.22. Plastic strain is localized 

at the contact points with the girders on the upper edge of the package, and with the lower 

roadway on the lower edge of the package. The maximum through-wall plastic strain is plotted 

in Figure 7.23. The maximum strain is predicted to be 2.4%. Plastic strain drops to zero on the 

inside surface. 
7-20

Figure 7.21. Case #1: Deformation of I-580 Span after Impact

Figure 7.22. Case #1: Effective Plastic Strain in Package Body Wall
7-21

Figure 7.23. Case #1: Maximum Plastic Strain in Package Body Wall

Figure 7.24 shows the plastic strain in the DU gamma shield for this scenario. The steel shell 

wall deforms across the 1-mm gap between it and the DU, imparting to the DU material plastic 

strains up to 1.5%. The plastic strain region in the middle of the length shows where the 

containment boundary requires the most support in this impact scenario.

Figure 7.24. Case #1: Effective Plastic Strain in DU Gamma Shield
7-22

7.3.2 Structural Case #2: Package Parallel to Upper Roadway Girders

This case considers the effect of an impact with one of the upper roadway girders aligned along 

the spine of the GA-4 package. Figure 7.25 shows the positioning of the package and girders 

before impact. (The roadway concrete has been omitted from this image, for clarity.) 

Figure 7.25. Case #2: SNF Package and Girder Orientation

Figure 7.26 shows that plastic strain in the package is below 1.1% for most of the axial length, 

with higher values occurring only near the end of the package. In this region, plastic strains are 

predicted to be up to 11%. This value is the highest predicted in the four cases. Figure 7.27

shows the plastic strain is localized, but it does penetrate into the third layer of elements with 

plastic strains below 1%. This is the end region, at either the base plate or closure end, and not 

through the package shell wall. While the closure plate and bolts are not modeled, the closure 

end is structurally robust because the thick closure plate extends into the flange for additional 

structural support. Figure 7.28 shows the predicted plastic strain in the DU gamma shield, 

within the steel body wall. Plastic strain in the DU is more widespread but also at a relatively 

low magnitude, up to about 3%. 
7-23

Figure 7.26. Case #2: Effective Plastic Strain in Steel Body Wall of SNF Package

Figure 7.27. Case #2: Local Plastic Strain in Containment End
7-24

Figure 7.28. Case #2: Effective Plastic Strain in DU Gamma Shield

7.3.3 Structural Case #3: Impact Localized on the Package Lid

In this case, the package was positioned such that only one girder made contact with the end of 

the package. As in all the other cases, the simplified model of the GA-4 package does not 

include detailed representation of the closure bolts, but the purpose of this case is to investigate 

the amount of potential damage to the flange area. (See Section 7.3.6 for analysis and 

discussion of the effect of the fire scenario on the closure bolts.) Figure 7.29 shows the girder 

deformation around the package for this drop scenario.

 

Figure 7.29. Case #3: Girder Deformation and Effective Plastic Strain
7-25

Figure 7.30 shows the plastic strain in the package wall for this drop scenario, which is localized 

at the impact area with a magnitude of about 5%. Figure 7.31 shows the extent of the plastic 

strain region. Plastic strains barely penetrate into the third layer of elements, with plastic strain 

magnitudes less than 0.1%. Figure 7.32 shows the plastic strain in the DU gamma shield, 

which remains below 3%.

Figure 7.30. Case #3: Package Effective Plastic Deformation

Figure 7.31. Case #3: Package End Plastic Strain
7-26

Figure 7.32. Case #3: DU Gamma Shield Effective Plastic Deformation

7.3.4 Structural Case #4: Impact Localized on the GA-4 Trunnions

This case investigates the potential for trunnion impact to localize the load on the package wall. 

Figure 7.33 shows the girder impact location and deformation. Figure 7.34 shows the plastic 

deformation in the package, which is very low and localized. Since peak plastic strains are 

below 1%, this case is the least damaging of the four considered here. Figure 7.35 shows the 

through-wall plastic strains in the maximum location. Figure 7.36 shows the plastic strain in the 

DU gamma shield, which is spread over a few areas but remains below 1.5%. 
7-27

Figure 7.33. Case #4: Girder Deformation

Figure 7.34. Case #4: Package Wall Plastic Strain
7-28

Figure 7.35. Case #4: Through-wall Plastic Strain

Figure 7.36. Case #4: Plastic Strain in DU Gamma Shield
7-29

7.3.5 Structural Impact Modeling Summary

The GA-4 package meets all regulatory requirements for over-the-road spent fuel transportation 

packages, and has a current Certificate of Compliance1 issued by the NRC. Although the 

conditions of the MacArthur Maze fire, as conservatively modeled in these analyses, constitute 

conditions that are not specifically defined in the design basis for this package (as per 

10 CFR 71), the results of the analyses with the structural models show that this package 

design is robust in this scenario. The structure of the GA-4 package wall is predicted to remain 

largely undamaged if subjected to this fire and roadway collapse scenario. The plastic strains

imparted to the package wall by the girders are well within the expected ductility limit. Plastic 

strain tends to be localized on the outside surface, with through-wall plastic strains tapering off 

to zero or near zero on the inside surface. Some permanent deformation of the package wall 

would be expected, but there would be no gross failure or rupture of the package. 

The most important mechanisms at work in this impact scenario appear to be the fundamental 

characteristics of the girder structure and the package structure. The girders are thin and 

comparatively weak at high temperatures, so tend to deform on impact instead of imparting 

energy to the stronger package. Based on the physics, minor variations in temperature or 

geometry are not likely to change the broad conclusion that the package would maintain its 

structural integrity during this scenario. 

The results of Case #2 show the potential for localized plastic deformation of the cask if the 

impact were to occur in the closure region. The package model used in these impact analyses 

represents the closure lid and flange as a single undifferentiated structure. In reality, the 

closure lid is inset into the top flange, extending 11 inches into the package cavity. 

Approximating the region as a solid slab of material is reasonable for determining the general 

package response, but the individual response of the two components is not evaluated. In the 

cases with impact at the cask lid region, plastic strain predicted at the flange end could cause 

localized deformations that could potentially affect the seating and seal of the closure lid. 

A complete representation of the impact response of this region would require more detailed 

modeling of the lid and top flange. However, the results of the thermal analyses of the package, 

as discussed in Section 7.1 and 7.2 above, show that the seals would be expected to fail in this 

fire scenario, so it is not necessary to examine additional potential modes of seal failure, such 

as the possibility of localized deformation due to impact loading. The issue of retaining the 

seating of the closure lid is investigated in Section 7.3.6, which reports the results of evaluations 

of the consequences of differential thermal expansion of the closure bolts, package lid, and top 

flange using a range of approaches, from classic bolt analysis to detailed FEA modeling of the 

components. These evaluations show that these crucial fasteners continue to perform their 

function, and would be expected to hold the closure lid in place even in the severe thermal 

environment of the MacArthur Maze fire scenario. 

7.3.6 Structural Issues Related to Bolt Thermal Expansion

Evaluations with the thermal and structural models of GA-4 package have identified two critical 

sets of bolts that could potentially affect its performance under the conditions of the MacArthur 

 1 Docket No. 71-9226, Certificate No. 9226, originally issued in 1998. Updated with Revision 1 in 2003, 

Revision 2 in 2008, Revision 3 in 2009, and is currently licensed under Revision 4. Expiration date for 

Revision 4 is October 31, 2018.
7-30

Maze accident scenario. These are the closure bolts fastening the lid to the package body, and 

the bolts fastening the top impact limiter to the anchor plate attached to the package body. Fire 

temperatures cause a significant amount of thermal expansion in the bolts and the package 

structure in general. The bolts are Inconel, and expand less than the XM-19 top flange and lid. 

The stresses caused by this thermal expansion mismatch are high, and add to the initial bolt 

tension. The closure bolts affect the containment boundary, while the impact limiter bolts affect 

the thermal response of the package, especially in the closure region, in that they determine 

whether or not the impact limiters would remain attached to the package in this accident 

scenario.

 

Inconel and XM-19 stainless steel are materials that retain high strength at elevated 

temperatures, but the Inconel experiences significantly less thermal expansion than the XM-19. 

This leads to increased tension in the bolt, beyond the initial tension due to the bolt torque 

applied at normal operating temperature. As the temperature increases during the fire scenario, 

the stresses from thermal expansion increase, and material strength begins to decline. There is 

a temperature limit for both sets of bolts, above which there could be failure of the connection, 

either through the tensile load on the bolt shank or the shear load on the threaded interface. 

This effect is independent of the postulated mechanical impulse load due to the overpass falling 

on the package, which occurs early in this transient scenario.

A critical feature of the GA-4 package design relative to its potential performance in this fire 

scenario is that the Inconel bolts are not threaded directly into the XM-19 structure of the 

package, either for the lid or the impact limiters. They are instead threaded into Heli-Coil 

threaded inserts, which protect the harder metals from galling. These inserts are made of Type 

304 stainless steel, which decreases in strength much more rapidly with increasing temperature 

than does XM-19. As a result, the insert threads can become the weakest link in the fastener, 

which is not consistent with standard bolted fastener design practice1

. The dissimilar materials 

involved in this design, coupled with the elevated temperatures resulting from the MacArthur 

Maze fire scenario, require detailed evaluation. The approach used in this analysis is based on 

calculations with engineering mechanics equations, using temperature results from the FEA 

thermal model. The effects of elevated temperature and forces due to differential thermal 

expansion are also evaluated with the structural and thermal FEA models of these components. 

The radial symmetry in the package geometry and the time-varying but uniform ambient 

boundary temperature throughout the fire scenario result in a symmetrical temperature 

distribution around the ends of the package. The uniform boundary temperature, which during 

the fire portion of the transient is a bounding temperature based on the FDS analysis presented 

in Section 3.0, is a reasonable approximation for the actual conditions seen by a surface with an 

assumed bounding uniform view of ambient. The radial symmetry of the temperature 

distribution extends to the ring of closure bolts, and also to the ring of impact limiter attachment 

bolts. As a result, all of the bolts in a given ring would be expected to have essentially the same 

temperature distribution. The response of any given bolt or insert can be treated as 

representative of the response of all bolts and inserts of that ring. 

 1 Bolts are generally designed to fail (yield) in the shank before an unexpectedly high load damages the 

threads. Analyses reported in the GA-4 SAR (General Atomics 1998) show that this would not be an 

issue for the maximum temperatures and impact loadings predicted for the design basis of this SNF 

package (10 CFR 71). 
7-31

One important unknown in the bolt thermal expansion evaluation is the post-yield behavior of 

the threaded inserts. The inserts carry a shear load that maintains the tension in the bolts due 

to initial bolt torque and differential thermal expansion. If the threaded interface were to yield, it 

would reduce tension in the bolt, and thereby reduce the potential for the bolt to yield. However, 

if the threaded interface were to fail, it could potentially result in a complete loss of bolt restraint 

and the possibility of detachment of the closure lid or impact limiter from the package. These 

possibilities are evaluated by considering two thresholds for the shear strength of the threaded 

inserts: shear yield, based on the material’s yield strength, and shear failure, based on the 

material’s ultimate tensile strength. The total shear force on the interface is compared to these 

thresholds over the effective thread shear area. It is assumed that the shear area remains 

constant between yield and failure, which is reasonable in this case, since the inserts fill the void 

between thread patterns and have little physical room to deform.

The presence of the impact limiters was treated as a variable in the thermal cask analyses by 

including them in the ANSYS thermal model but omitting them from the COBRA-SFS model. 

The thermal response of the package is affected by this variable, in that the location of the 

maximum fuel temperatures is different for the two cases, but the maximum fuel temperature 

values are not substantially different. In the context of the closure bolt evaluation, however, this 

becomes a critical issue. 

The thermal models of this study consider two possibilities: the impact limiters detach before the 

fire or they remain attached through the fire duration and post-fire cooldown. A case where the 

fire or the bridge collapse causes the impact limiter to detach at some point during the fire is not 

explicitly modeled, but the thermal results of that case would be bounded by the assumption 

that the impact limiters are lost before the beginning of the fire. This treatment provides a 

bounding thermal evaluation for the package and the spent fuel assemblies contained within it 

(as discussed in Section 7.2), but carrying this uncertainty forward into the closure bolt 

evaluation is unnecessary. As discussed below in Section 7.3.6.1, loss of the impact limiters is 

not a credible consequence of the MacArthur Maze accident scenario. Therefore, the impact 

limiters are assumed to remain attached to the package for the evaluations of the closure lid 

bolts presented in Section 7.3.6.2 and Section 7.3.6.3. 

7.3.6.1 Results of Impact Limiter Bolt Evaluations

A 3D structural and thermal finite element model of one impact limiter bolt and the surrounding 

material was developed, using the ANSYS thermal model results as a baseline temperature 

distribution. (See Section 5.4.2 for model description details.) This model considered the 

effects on bolt temperature of an additional heat transfer path for the bolts that was neglected in 

the ANSYS thermal model of the GA-4 package. The detailed single-bolt model was used to 

evaluate bolt thermal expansion at the predicted temperatures.

The GA-4 package impact limiter attachment bolt heads are located at the bottom of long holes 

extending the axial depth of the impact limiter cap. (See diagram in Figure 5.19.) The bolt 

heads are by design uncovered, for access when attaching or removing the impact limiters. 

This configuration results in the bolt heads being directly exposed to ambient conditions. The 

deep holes would essentially preclude convection heat transfer to the bolt heads, but they would 

be exposed to thermal radiation exchange with ambient conditions. 

This pathway for heat to enter the system was not captured in the ANSYS thermal model of the 

GA-4 package, since it would not be expected to have a significant effect on the total package 

thermal response to the fire scenario. However, this omission could have an effect on the local 
7-32

temperature distribution and affect the thermal expansion calculations for the impact limiter 

bolts. Therefore, additional evaluations were performed with the impact limiter bolt model, to 

determine the potential effect of the additional thermal energy that could be entering the region 

by means of thermal radiation. There are two possible ways the extra energy could affect the 

thermal response of the impact limiter bolts to the fire conditions. It could increase the 

temperature in the threaded interface, or it could increase tension in the bolt, due to changing 

the temperature distribution in the bolt shank. Both of these possibilities are explored using the

ANSYS thermal model of the impact limiter bolt region.

The single bolt model used in the thermal expansion structural evaluation was extended to the 

thermal-physics environment by changing element types, degree of freedom constraints, and 

other ANSYS model parameters. The thermal radiation load was modeled with ideal heat 

transfer between the fire environment and the top bolt head surface, with all other surfaces 

adiabatic. The initial temperature was set to 124°F (51°C) throughout the bolt and local 

package region, which is the minimum initial temperature in the region, based on the ANSYS 

thermal model of the GA-4 package at NCT. The single bolt model was then subjected 

sequentially to the two phases of the MacArthur Maze fire scenario. As described above, the 

first phase consists of an ambient temperature of 2012°F (1100°C) lasting 37 minutes, and the 

second phase consists of an ambient fire temperature of 1652°F (900°C), lasting for another 

71 minutes. Evaluation of the post-fire cooldown period was found to be unnecessary.

The results obtained with this thermal model show that including thermal radiation to the top of 

the bolt slightly increases the temperature of the bolt shank at the end closest to the fire, but has 

almost no effect on the temperature at the other end of the shank, in the threaded region of the 

bolt. This response is due to the relatively low thermal conductivity of the bolt nickel alloy, and 

the relatively small amount of radiant thermal energy that can be absorbed through the small 

cross-section of the bolt heads. The net effect of including thermal radiation at the bolt head is 

to slightly increase the bolt shank temperature relative to the package steel wall, which slightly 

reduces overall bolt tension, because of the corresponding small increase in thermal expansion. 

This is because thermal radiation helps reduce the component of bolt tension caused by the 

non-uniform temperature distribution in the bolt shank and box region. The results of this 

analysis show that it is appropriate and conservative to neglect the effect of thermal radiation on 

the bolt head in this analysis. 

These evaluations show that the fire conditions defined to bound the MacArthur Maze fire could 

be damaging to the threaded inserts, but the potential for failure of the threaded interface is 

limited by the bolt shank’s capacity to yield and relieve tension. The results of these analyses 

support the conclusion that the attachment of the impact limiters to the package would not fail 

due to thermal loads imposed by this fire scenario. The mechanical loads of the accident are 

within the regulatory realm and do not require additional evaluation. This fire scenario 

represents a more challenging fire accident than that required in the design basis for the GA-4 

package, but a detailed thermal evaluation of the system shows that the connection formed by 

the impact limiter bolts would survive even under these conditions.

The baseline temperature history case evaluates the thermal expansion of the impact limiter bolt 

using a temperature history from the results obtained with the ANSYS thermal model of the GA4 package. (See Section 7.2 for package thermal modeling results.) This model yields detailed 

nodal temperature predictions in the region of the bolts, but the mesh does not directly align with 

the nodal mesh of the bolt model. The predicted temperatures were mapped onto the bolt 

model mesh using a simple averaging scheme at locations along the axial length of the bolt. 

The appropriateness of this approach was verified by developing a mapping scheme based on 
7-33

the maximum temperatures at a given axial location. The temperature history considered here 

includes only the 108-minute fully engulfing fire transient. The results show that considering the 

post-fire cooldown period is not necessary.

Total force on the thread interface is extracted from the ANSYS model at each solution time 

step, and the tension due to the bolt preload is added to obtain the total shear force. The initial 

preload due to bolt torque is a tension of 21,657 lbs (96.3 kN) at room temperature (General 

Atomics 1998). To account for temperature-dependent material properties, the preload force is 

scaled by the modulus at the average shank temperature. For example, with an average shank 

temperature of 754°F (401°C), the preload is reduced to 88% of its initial value. 

The bolt will yield wherever the temperature is hottest along its length, since yield strength 

decreases with temperature, but the force load on the cross section remains constant along the 

axial length. The shear yield behavior of the threaded inserts is not well quantified, but the 

tensile yield behavior of the bolt shank would closely resemble a typical tensile test. In the 

temperature range of the conditions predicted for the MacArthur Maze fire scenario, the 

elongation capacity of the bolt material is approximately 20% or greater. The amount of 

elongation required to relieve all tension in the bolt at these temperatures is less than 1%. This 

indicates that the bolt shank has enough ductility to accommodate the yield without failure. 

For all temperatures in the range of interest, including post-fire temperatures, these evaluations 

show that bolt yield is reached before insert shear failure. Shear failure of the insert before the 

bolt shank yields and releases tension could occur only with the insert at a much higher 

temperature than the bolt shank. This is not physically possible in this scenario, because of 

package geometry considerations, and is therefore not of concern.

7.3.6.2 Results of Lid Closure Bolt Evaluations (Classical Approach)

The results of classical bolt calculations for the bolts securing the closure lid of the GA-4 

package are presented in this section. The analysis considers the bolt preload (due to the initial 

torque on the bolts), the external load (due to the internal pressurization of the package cavity), 

and the thermal load (due to differential thermal expansion of the Inconel bolts and XM-19 

stainless steel closure lid and flange of the package body). The bolt preload is defined based 

on the manufacturer’s requirements (General Atomics 1998). The external load due to package 

pressurization is calculated based on the average gas temperature within the package cavity, as 

predicted with the ANSYS thermal model. The thermal load is determined in this analysis, 

based on the detailed temperature predictions from the ANSYS thermal model.

Many parameters in the classic bolt analysis are based on the geometry of the system, and 

therefore are straightforward to determine, such as the diameter of the bolt shank and its crosssectional area. Less readily defined parameters include the effective joint area, which affects 

the representative stiffness of the lid and flange combined. One common method (Juvinall and 

Marshek 1991) uses the average area of a conical volume between clamped members. An 

alternative approach using bolt mechanics gives a more approximate estimate, based on bolt 

diameter. For the GA-4 geometry, the effective areas estimated from these two methods are 

4.60 in2 and 3.18 in2

, respectively. The difference in estimated area results in less than a 3% 

difference in predicted bolt tension, which is within the uncertainty in the overall approach, and 

therefore either estimate would be considered reasonable.

The thread inserts are a complicating feature for this analysis. Although thread inserts are 

commonly used to prevent thread galling, there is little information in the literature concerning 
7-34

their effect on the mechanics of a bolted joint. In typical applications, it can generally be 

assumed that the thread inserts have negligible effect on the bolt tension, particularly if the 

thread insert material is similar to the surrounding material. In the GA-4 package, however, the 

bolts are nickel alloy (Inconel), the closure lid and flange is XM-19 stainless steel, and the 

inserts are Type 304 stainless steel thread inserts. At NCT, temperatures are in a range where 

the XM-19 and Type 304 stainless steels have similar mechanical and thermal properties. 

Potential problems arise in the fire scenario as the predicted temperatures on these 

components rise to ranges in which the yield strength of the Type 304 insert decreases much 

more than the XM-19. At high temperatures and high tensions, the classic bolt calculations 

demonstrate that the inserts could have a dramatic effect on the clamping force between the 

closure lid and the package body flange.

As part of an initial assessment of the bolt behavior due to the elevated thermal loads of the fire 

scenario, the effect of the thread insert was neglected. Classical bolt calculations were used to 

determine the bolt tension and clamping force over time, assuming the thread insert would have 

no effect on the bolt or joint stiffness. The joint effective area was assumed to be 4.5 square 

inches, which is near the high end of the area estimates described above. The key parameters 

in this calculation are listed in Table 7.1.

Table 7.1. Key Bolt Parameters

Parameter Specified Value

Critical Bolt Cross Section (in2) 0.605

Effective Bolt Shank Length (in) 4.62

Bolt Pitch Diameter (in) 0.9100 - 0.9168

Bolt Major Diameter (in) 0.9830 - 0.9980

Bolt Shank Diameter (in) 0.878

Lid Bolt Hole Diameter (in) 1.08

Flange Bolt Hole Maximum Diameter (in) 1.21

Bolt Head Diameter (in) 2.0

Effective Joint Area (in2) 4.5

The significant loads on the bolt are determined from the initial tension specified by the design 

and the temperature and internal pressure response of the package to thermal conditions of the 

MacArthur Maze fire scenario, as determined from the thermal modeling. The ANSYS thermal 

model calculated the transient out to 14.5 hours, which is approximately an hour beyond the 

point at which all components of the package were decreasing in temperature. This is more 

than 12 hours after the end of the actual fire duration, but average component material 

temperatures within the package are still above 1000°F (538°C), and the average cavity 

pressure is more than 100 psi above the initial pressure at the start of the transient. Given the 

large thermal inertia of the GA-4 package, it would take a very long time for the system to cool 

back down to its starting conditions after exposure to a fire of the severity of this bounding 

model of the MacArthur Maze fire scenario. 
7-35

Figure 7.37 shows the extrapolated cavity gas temperature calculated with a reasonable 

assumed power-law cooldown rate, based on the physical behavior of a large object with high 

thermal inertia cooling in air. The corresponding cavity pressure from the ideal gas law 

(conservatively assuming no leakage), is shown in Figure 7.38. Based on this extrapolated 

cooldown rate, the time required for the closure lid bolt region to drop from its peak temperature 

of 1223°F down to 500°F would be on the order of 50 hours. It would require approximately

400 hours for the system to return to its original temperature and pressure state. The long-term 

temperature response is important to this analysis, because it is in the cooldown phase that the 

clamping force has potentially the greatest challenge in this scenario. 

Figure 7.37. Average Cavity Gas Temperature from ANSYS Thermal Model Calculated to 

14.5 Hours, with Power-law Extrapolation to 400 Hours
7-36

Figure 7.38. Cavity Pressure Calculated from Average Cavity Gas Temperature, Assuming 

No Leakage

The results of the classical bolt analysis are tabulated in Table 7.2, with the three components 

of bolt force separated for comparison. Bolt Force is the tension on each of the twelve bolts, 

considering all loads and the stiffness of the bolt and joint material. Clamp Force is expressed 

on a per-bolt basis, so the surfaces in contact would experience a total force that is twelve times 

the value given in the table. Figure 7.39 shows plots of these force values, on a per-bolt basis. 

Table 7.2. Results of Classical Bolt Analysis

Time

(hrs)

Temperature 

(°F)

Pressure 

(psi)

Components of Bolt Force

Bolt 

Force

(lbs)

Clamp 

Force

(lbs)

Initial Bolt 

Preload

Fi

(lbs)

Elongation

Fe

(lbs)

Thermal 

Expansion

Ft

(lbs)

0 129.9 73.7 20,655 2,806 0 20,998 18,192

0.17 153.4 75.7 20,655 5,297 13,356 34,612 29,314

2 857.3 169.9 20,655 6,470 18,779 40,424 33,953

5 1184.6 199.4 20,655 7,594 26,953 48,810 41,216

10 1199.9 197.3 20,655 7,512 27,262 49,103 41,590

14.5 1114.9 185.4 20,655 7,062 25,512 47,300 40,239

52 482.8 120.6 20,655 4,592 9,768 31,052 26,460

400 129.9 73.7 20,655 2,806 0 20,998 18,192
7-37

Figure 7.39. Bolt Force and Clamping Force Predicted using Classic Bolt Equations

The peak value of 49,103 lbs for the bolt force calculated from the classical bolt analysis is 

significantly high, considering the strength of materials involved. Figure 7.40 shows the

predicted strength of the thread insert and bolt materials over time (in response to the 

temperature transient), compared to the predicted bolt force during this transient. The strengths 

are determined based on the thread insert minimum shear area, which is defined at the major 

diameter of the bolt thread. As the plot in Figure 7.40 shows, the bolt is not in any danger of 

yielding under the load of the predicted bolt force in this analysis. By contrast, the weaker 

thread insert is predicted to exceed its yield limit at approximately three hours into the transient. 

However, the ultimate shear strength of the insert is not exceeded at any time during the 

transient, which suggests that the insert would yield and release some tension without 

completely failing. 
7-38

Figure 7.40. Strengths of Thread Inserts and Bolts Compared to Bolt Force Predicted using 

Classic Bolt Equations

This is an important conclusion for the release calculations, and provides some indication that 

the GA-4 package could withstand the MacArthur Maze fire scenario without significant failure of 

the containment boundary. However, the yielding in the inserts and the effect this would have 

on the clamping force needs to be more precisely quantified, so that it can be used in 

evaluations of potential consequences of subjecting an SNF package to this accident scenario. 

It is therefore necessary to expand this analysis with the classic bolt equations to include an 

estimate of the effect of insert yielding on the closure lid clamping history.

If the thread insert were to yield, this would be expected to result in a reduction of tension in the 

bolt. Furthermore, such yielding would be expected to cause a permanent reduction of the 

initial tension load, which would persist after the system returned to its starting temperature and 

pressure. This is truly a complex scenario, but a few simplifying assumptions can be made to 

adjust the classic bolt equations to estimate the potential reduction in clamping force.

The first assumption is that the force that can be carried by the thread insert is limited to the 

yield strength at each particular temperature. This is a conservative assumption because in 

reality the material would be expected to strain-harden, which means that the actual load that 

the insert can support would tend to increase as the material starts to yield. A second 

assumption is that the effect of plastic deformation in the thread insert corresponds to a poundfor-pound reduction in the initial bolt tension. In other words, the initial preload force (Fi in Table 

7.2) is permanently reduced by the amount of bolt force that exceeds the thread insert yield 

limit. This is potentially a non-conservative assumption because the yielding occurs at elevated 

temperatures where the stiffness is lower than the starting condition. Yielding at high 
7-39

temperatures might contribute to a proportionally higher reduction of bolt tension when the 

system cools. Both of these assumptions are made to simplify the problem enough to obtain an 

estimate of the potential effect on the clamping force. 

Figure 7.41 shows the effect on bolt force and clamping force when yielding of the thread insert 

is taken into account in this analysis. The plot in this figure compares these results from the 

predicted forces shown in Figure 7.39, calculated assuming no yielding of the thread insert. The 

plot in Figure 7.41 shows that accounting for yielding of the thread insert results in a dramatic 

change in the overall results. By the time the system returns to its initial state, the clamping 

force is significantly reduced from its starting value. This is a very different outcome than is 

predicted in the elastic case, which assumes the bolt will ultimately return to its starting state of 

tension. 

Figure 7.41. Estimated Effect of Yielding Thread Inserts Using Classic Bolt Equations

Since the package ethylene propylene seals are not expected to survive the MacArthur Maze 

fire temperatures, the clamping force on the metal-to-metal contact surfaces of the closure lid 

and package body flange is all that will be holding the pressurized contents inside the package 

cavity. In this configuration, the leak rate from the package cavity would be directly related to 

the clamping force. Basing estimates of leak rates on this type of simplified modeling evaluation 

would lead to large uncertainties in any estimate of potential release rate from the package. 

Because of the sensitive nature of the leak rate calculations, a more sophisticated estimate of 

the clamping force history is needed. This is provided by the results of the detailed FEA 

modeling approach, as described in Section 7.3.6.3.
7-40

7.3.6.3 Results of Lid Closure Bolt Evaluations (FEA Modeling Approach)

As described in Section 5.4.3.2, the FEA modeling of the lid closure bolt region consisted of a 

detailed representation of a single bolt, due to the radial symmetry of the package and assumed 

boundary conditions for the MacArthur Maze fire scenario. Figure 7.42 shows a diagram of the 

model, including a close-up of the detailed mesh in the thread insert region.

Figure 7.42. Diagram of Detailed Mesh in Thread Insert Region

In the classical bolt analysis presented in Section 7.3.6.2, the critical shear area for the thread 

insert was assumed to occur at the maximum diameter of the bolt thread, corresponding to 7/8-

pitch. In reality, the critical shear area of the insert would be more likely to occur on the flange 

thread. However, in this geometry, the area of the two possible shear areas would not be 

significantly different. The finite element model shows that peak stress concentrations occur on 

the bolt side, but it is not obvious which shear plane would have the higher average stress. This 

is illustrated in Figure 7.43 which shows the peak von Mises stress (psi) state in the thread 

inserts at the peak temperature in the evaluation assuming elastic materials. The high stresses 

would certainly cause plastic strain to occur, with a corresponding redistribution of the stress. 

This is further evidence that the evaluation requires a multilinear elastic-plastic material model.
7-41

Figure 7.43. Stress Distribution Predicted with Elastic-only Assumptions in FEA Model

With elastic materials, the bolt force and clamping force agree with the classic bolt calculations 

to within 1% at the initial temperature and pressure. But as temperature and pressure increase 

in the transient, the bolt force and clamping force predicted with the FEA model are 

approximately 20% lower than the values predicted by the classic bolt calculations. Sensitivity 

evaluations comparing the FEA model results and the classic bolt calculations showed that the 

two can be brought in good agreement using assumptions and parameter adjustments that 

simplify the FEA model to approximate the level of detail captured in the classical approach. 

These evaluations demonstrated the significance of the stiffness of the thread insert material in 

determining the behavior of the system, and show that the classic hand calculations 

overestimate the bolt tension and clamping force, and would offer a conservative prediction of

the amount of plastic deformation that would occur in the thread insert.

This is consistent with the distinction between a bolt and a screw, as was noted previously as an 

important factor in this evaluation. The nut-and-bolt configuration assumes the clamping force 

is applied by the head of the bolt on one side and the axial face of the nut on the other side of 

the joint. The clamping area estimate for a true bolt (Juvinal and Marshek 1991) assumes that 

the conical volume extends from the bolt head to the nut, as illustrated in Figure 7.44(a). The 

clamping area estimate for the corresponding screw configuration, as shown in Figure 7.44(b)

would result in an effective clamped volume that would be smaller than in the bolt-and-nut 

configuration. The angle defining the volume for the screw fastener cannot readily be 

determined in the classical approach, and for this reason alone, the elastic FEA results would 

be expected to provide a better estimate of the effective joint area of this screw configuration 

(2.0 in2

) than the values obtained in the classic bolt estimates (3.18-4.60 in2). 
7-42

 (a) (b)

Figure 7.44. Diagram of Compression Volume for (a) Bolt and (b) Screw Fasteners

Even when considering only elastic behavior of the system, the FEA model provides a more 

realistic prediction than the classic bolt calculations. Further refining the FEA model to include 

multilinear temperature-dependent material properties for the thread insert material permits a 

best-estimate of bolt tension and clamping force for the conditions assumed in the MacArthur 

Maze fire scenario.

For the nonlinear analysis, the only change in the FEA model was in the definition of the 

material properties for the thread insert. The multi-linear kinematic hardening model was used 

to define stress-strain curves for a range of temperatures between 70°F (21°C) and 1200°F

(649°C). Figure 7.45 shows the stress-strain curves in units of psi versus strain (in/in). The bolt 

material and flange material were both defined as elastic material. The nickel alloy bolt retains 

such high strength that even localized yielding is not expected. 
7-43

Figure 7.45. Stress-Strain Curves used for Thread Inserts in Multi-linear FEA Model

The model predicts small plastic strains throughout the thread inserts, particularly along a line 

from the bolt outer diameter to the flange inner diameter, crossing from one shear plane to the 

other. Figure 7.46 shows contours of equivalent plastic strain at the peak temperature and 

pressure. The peak localized plastic strain is about 5%, while the maximum average along a 

plane is roughly 2% in the bottom segment. At the top of the thread pattern, the plastic strain is 

generally much less, and the maximum average plane drops to less than 1%. However, even 

this relatively small plastic strain is significant when strain hardening is considered. Considering 

a plastic strain of 1% along the critical shear plane at 1200°F (649°C), strain hardening would 

increase the yield strength by roughly 50%. This suggests that the predictions obtained with the 

classic calculations are highly conservative, even when considering the effect of yielding thread 

inserts (as shown in Figure 7.41), since that approach does not consider the effect of strain 

hardening. 
7-44

Figure 7.46. Plastic Strain in Thread Insert Predicted with Multi-linear FEA Model

The bolt tension and clamping force history, on a per-bolt basis predicted with the nonlinear 

finite element model is summarized in Table 7.3. These results can be considered a best 

estimate evaluation, within the conservative assumptions of the MacArthur Maze fire scenario. 

Comparing the initial conditions and final cooldown conditions at 400 hours, the ultimate 

reduction in clamping force due to plastic deformation of the thread inserts is predicted to be 

less than 8%. In absolute numbers, the loss of initial load (Fi from the classic bolt calculations) 

is only 1358 lbs. 

Table 7.3. Best Estimate Results from Non-Linear FEA Model

Time

(hrs)

Temperature 

(°F)

Pressure 

(psi)

Bolt Force

(lbs)

Clamp Force

(lbs)

0 129.9 73.7 20,920 18,191

0.167 153.4 75.7 29,995 24,876

2 857.3 169.9 33,923 27,680

5 1184.6 199.4 39,210 31,885

10 1199.9 197.3 39,399 32,149

14.5 1114.9 185.4 38,208 31,391

52 482.8 120.6 26,452 22,013

400 129.9 73.7 19,557 16,833

The nonlinear finite element model provides a reasonable best estimate prediction of the 

clamping force on the flange throughout the MacArthur Maze scenario and its eventual 

cooldown. These results are used as the basis for leak rate calculations for the GA-4 package, 

presented in Section 8.3.3.
8-1

8.0 POTENTIAL CONSEQUENCES

Potential adverse consequences of a severe accident involving an SNF transportation package 

fall into two general categories; a loss of shielding, which could pose an exposure risk to 

members of the public, and a failure of the containment boundary of the package, which could 

lead to a release of radioactive material to the environment. 

Loss of either the neutron or gamma shielding of the package could potentially result in a direct 

radiation dose to an individual in close proximity to the package. Failure of any of the 

components (e.g., package seals) that make up the containment boundary could result in a 

release of radioactive material from inside the package, potentially resulting in a direct 

radioactive dose to first responders at the scene of an accident, or possibly to members of the 

public in the surrounding area. Loss of shielding as a potential consequence of the MacArthur 

Maze fire scenario is discussed in Section 8.1. Section 8.2 discusses the performance of 

package seals in this fire scenario. Section 8.3 presents detailed evaluations of the potential for 

a release from the GA-4 package as a result of the severe conditions predicted for this

extremely conservative representation of the MacArthur Maze fire scenario.

8.1 Potential for Loss of Shielding

The potential for increased neutron and gamma radiation dose rates from the GA-4 as a result 

of exposure to the MacArthur Maze fire scenario was evaluated. Direct radioactive dose rate 

limits are specified in 10 CFR 71 for NCT and HAC conditions. As a licensed transportation

package, the design basis of the GA-4 complies with the regulatory limits for all conditions of 

transport.

Section 8.1.1 describes the consequences of loss of neutron shielding for the GA-4 in the 

MacArthur Maze fire scenario. Section 8.1.2 discusses the potential for loss of gamma 

shielding.

8.1.1 Neutron Shielding

Neutron shielding for the GA-4 package is provided by neutron-absorbing liquid in an annular 

tank surrounding the steel body of the package (see Sections 5.1 and 5.2 for details of package 

geometry). The neutron shielding material is a mixture of 56% propylene glycol and water, with 

1% dissolved boron. The neutron shield tank is not generally expected to survive the 

hypothetical accident conditions prescribed in 10 CFR 71 for SNF transportation packages, 

which include a 30-minute fully engulfing fire at “1475°F (800°C).” 

The GA-4 package is designed to be in compliance with the regulatory limits for all conditions of 

transport. Loss of the neutron shield tank contents is a design-basis assumption for HAC, and 

analyses presented in the SAR (General Atomics 1998) for this package assume loss of the 

neutron shield in all accident scenarios, including the HAC fire. The conditions of the MacArthur 

Maze fire, although more severe than the HAC fire, can do no more damage to the neutron 

shield of the GA-4 than is assumed a priori in the HAC fire. Therefore, the GA-4 package would 

be expected to remain below the regulatory dose limits after loss of neutron shielding in the 

MacArthur Maze fire scenario. 
8-2

8.1.2 Gamma Shielding

Gamma shielding for the GA-4 is provided by a 2.64-inch-thick (6.7 cm) layer of DU encased 

within the stainless steel body of the package. The DU layer extends a few inches beyond the 

full axial length of the package inner cavity to assure complete coverage of the active fuel 

length, and is positioned between the stainless steel inner liner and the 1.5-inch-thick (3.81 cm) 

stainless steel body of the package.

The DU material experiences a significant increase in temperature, but the performance of the 

DU gamma shield is unaffected by the MacArthur Maze fire scenario. As discussed in 

Section 7.3, structural analysis of a range of postulated scenarios in which the overhead I-580 

highway span is assumed to impact the GA-4 package shows that the package would be 

expected to survive without structural failure. The thermal analyses of the package response to 

the fire scenario show that the peak temperature in the DU material could reach up to about 

1480°F (805°C). This is significantly below this material’s melting temperature of 2070°F 

(1132°C). 

These results show that the gamma shielding of the GA-4 can be expected to remain intact and 

functional even if subjected to the severe conditions of the MacArthur Maze fire scenario. 

Therefore, the GA-4 package would be expected to remain below the limits specified in 10 CFR

71 for accident conditions.

8.2 Performance of Package Containment Seals

Based on the results of the thermal analysis (as discussed in Sections 7.1 and 7.2), there is a 

possibility of a release from the package because of failure of components that make up the 

containment boundary of the package. Calculated temperatures in the region of the lid closure 

seal, drain valve/port, and gas sample valve/port seals during the transient far exceed the 

continuous use temperature limits for the seal material. Therefore, the potential exists for the 

release of contents from the package in this fire scenario.

Section 8.2.1 provides a detailed discussion of seal temperatures predicted for the GA-4 

package in the MacArthur Maze fire scenario. Section 8.2.2 presents a discussion of seal 

performance at elevated temperatures, based on experimental data relevant to the conditions of 

the MacArthur Maze fire. 

8.2.1 Seal Temperatures in the MacArthur Maze Fire Scenario

The containment boundary for the GA-4 package is maintained by the seals on the package lid, 

drain valve and port, and gas sample valve and port. The package lid seal consists of primary 

and secondary O-rings at the interface of the lid and the package stainless steel body. The gas 

sample valve is located within the package lid, and the drain valve is located in the steel base of 

the package. The gas sample valve is sealed with primary and secondary O-rings, and for 

transport conditions, the outer face of the port is fitted with a steel plug that is threaded to a

specified torque of 20 ft-lb. The drain valve is sealed within its access port with primary,

secondary, and tertiary O-rings. The drain valve cover and drain port plug are also sealed with 

O-rings, in addition to being threaded, and are torqued to 20 ft-lb. 

The O-ring seals at all locations are ethylene propylene, which has a continuous-use 

temperature limit of 302°F (150°C). Figure 8.1 shows a graph of the temperature limit on this 
8-3

material as a function of exposure time. As exposure temperature increases, the time limit for 

allowed exposure decreases. (Note that the horizontal axis, Exposure Time, in this plot is on a 

logarithmic scale.) The maximum temperature this material is rated to withstand without

effectively immediate failure is 790°F (421°C), but it will tolerate this exposure for only six 

minutes. 

Figure 8.1. Operating Temperature Limit as a Function of Exposure Time for Ethylene 

Propylene Seal Material (Based on Data Presented in the GA-4 SAR [General 

Atomics 1998])

All of the containment boundary seals in the GA-4 package are in locations that are covered by 

either the top or bottom impact limiter assembly. Without the impact limiters in place, the seal 

locations are directly exposed to the fire conditions, and as a result, exceed all operating 

temperature limits within minutes, and high temperatures persist at these locations throughout 

the fire transient and into the post-fire cooldown. However, as discussed above in Section 7.0, 

structural and thermal evaluations show that loss of the impact limiters is not a credible 

consequence of this fire scenario. Therefore, the seal performance evaluations consider only 

the case with the impact limiters in place.

In the HAC fire analysis (30 minutes at 800°C) presented in the package SAR, the predicted 

peak temperatures on the seal components do not exceed the bounds of the operating

temperature limit curve shown in Figure 8.1. This is verified by results obtained for this transient 

with the ANSYS and COBRA-SFS models, as noted in Section 6.1. It is therefore justifiable to 

assume that the seals would survive intact in the HAC fire. In the conditions of the MacArthur 

Maze fire scenario, however, the higher fire temperatures and longer duration of the fire 

exposure result in temperatures that far exceed the rated temperatures of this hydrocarbon seal 

material. Figure 8.2 shows the temperature predicted with the ANSYS model (with impact 

limiters in place) for the closure lid seal location during the fire scenario. 

0

100

200

300

400

500

600

700

800

900

1000

0.1 1 10 100 1000

Exposure Time (hours)

Temperature (oF)
8-4

Figure 8.2. Closure Lid Seal Temperatures Predicted with the ANSYS Mode (with Impact 

Limiters) for the MacArthur Maze Fire Scenario

The plot in Figure 8.2 also shows the time in the transient when the closure lid seal location is 

predicted to exceed specific exposure temperature limits for the seal material. This plot 

indicates that the seal material could survive these conditions for possibly as long as 1.7 hours; 

almost the entire duration of the fire portion of the transient. However, temperatures in the seal 

location continue to rise for many hours after that point, and by approximately 2 hours into the 

transient, exceed all operating temperature limits defined by the graph in Figure 8.1. Similar 

plots in Figure 8.3 for the gas sample valve/port seals and in Figure 8.4 for the drain valve/port 

seals show that all operating temperature limits for these seals are exceeded, in approximately 

the same time-frame.
8-5

Figure 8.3. Gas Sample Valve/Port Seal Temperatures Predicted with the ANSYS Model (with 

Impact Limiters) for the MacArthur Maze Fire Scenario

Figure 8.4. Drain Valve/Port Seal Temperatures Predicted with the ANSYS Model (with Impact 

Limiters) for the MacArthur Maze Fire Scenario
8-6

The plots in Figure 8.2, Figure 8.3, and Figure 8.4 incidentally confirm the SAR prediction, 

showing that these seals could probably survive a fire lasting only 30 minutes, followed by a 

cooldown at normal ambient conditions. However, if exposed to the conditions of the MacArthur 

Maze fire scenario, the seal material would exceed its rated long-term temperature limit within 

1.5 to 2.5 hours at all locations. Complete evaluation of the effect of the fire scenario on the 

seal region requires evaluating the seal region temperatures during the post-fire cooldown 

transient, as well as during the fire itself. As discussed in Section 7.2, the insulating effect of the 

impact limiters results in the temperatures of components near the ends of the package 

continuing to increase long after the end of the fire. Figure 8.5 shows that the peak 

temperatures in the seal region locations continue to increase for more than 4 hours after the 

fire, reaching approximately 1150°F (621°C), and after 14.5 hours are still above 1000°F 

(538°C). 

Figure 8.5. Seal Temperatures Predicted with ANSYS Model (with Impact Limiters) for 

MacArthur Maze Fire Scenario and Post-fire Cooldown

The results in Figure 8.2 through Figure 8.5 show that the seal region temperatures would 

exceed all exposure temperature limits for a long period of time during the post-fire cooldown. 

Seal failure, which is defined as the inability of the seal material to maintain a stable differential 

between the internal pressure within the package cavity and the external ambient pressure, 

seems inevitable under these conditions. In general, if seal materials exceed their rated 

temperature limits, they are treated for safety evaluation purposes as having failed entirely. 

However, this is a bounding assumption, since it does not attempt to quantify the degree of seal 

failure or resulting leak rate as a function of temperature or time-at-temperature. Section 8.2.2

summarizes results of experimental measurements of the performance of elastomer seals at 

elevated temperatures, which show that seal failure is a complex process, not a simple pass/fail 

test. 

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

1300

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

Elapsed Time (hours)

Peak Temperature (°F)

Closure Seal Gas Sample Port Seal

Drain Valve Seal maximum seal exposure temperature for 6 minutes

maximum seal exposure temperature for 12 minutes maximum seal exposure temperature for 1 hour

maximum seal exposure temperature for 10 hours long-term seal temperature limit

end of 1100°C fire end of fire
8-7

8.2.2 Seal Performance Testing

The NRC and NIST conducted testing of seals in thermal conditions simulating fire 

environments that exceeded the rated temperatures for the seals tested (NUREG/CR-7115

2012). These tests evaluated the performance of one type of metallic seal and two different 

polymeric compound seals typically used in SNF transportation packages, using a small

stainless steel cylinder with a single O-ring seal. The test vessel was pressurized at 75°F 

(24°C) with helium to 73.5 psia (5 bar) for the tests with metallic seals, and to 29.4 psia (2 bar) 

for the tests with polymeric seals. The fire was simulated using an electric furnace that could 

maintain a controlled thermal environment for a specified duration, which was varied in different 

tests from several hours to 24 hours, and in some cases up to 72 hours. (These tests were 

designed to simulate accident conditions. Typical seal tests for long-term normal operating 

conditions are performed for a minimum of 1000 hours.) Following the simulated fire exposure 

duration, the test vessel was cooled to room temperature within the electric furnace. 

A total of 15 tests were conducted in this study, including the initial shake-down test for which 

results were not recorded, due to instrumentation failure. Of the 14 tests for which 

measurements were recorded, 11 tests were with a metallic seal, 2 tests were with an ethylene

propylene seal, and 1 test was with a polytetraflouroethylene (PTFE) seal. In terms of the 

applicability of this testing to the evaluation of the GA-4 package in the MacArthur Maze fire 

scenario, the two tests with ethylene propylene seals are of significance, since this is the seal 

material used in the GA-4 package for the lid closure and the gas sampling port valve and the 

drain valve. 

The most severe exposure for ethylene propylene seals in the testing was at 842°F (450°C). 

Based on the performance curve in Figure 8.1, the exposure time limit for this seal material at 

this temperature is less than 6 minutes. The seal material failed in this test within the first three 

hours of the simulated fire transient, but exhibited a much slower leak rate than would be 

expected for the test vessel with no seals at the test conditions. The second test with ethylene 

propylene seals reached a much lower peak temperature, and simulated a fire environment with

incremental heating from 302°F (150°C) to 572°F (300°C). Based on the performance curve in 

Figure 8.1, the exposure time limit at the maximum heating temperature in this test is less than 

30 minutes. The total duration of the simulated fire was more than 20 hours, but in this test, the 

seal held with no measurable leakage.

The results of two tests in a small (not-to-scale) test vessel are not sufficient data on which to 

base a general evaluation of ethylene propylene seal material at temperatures above rated 

performance values. However, this testing does show that even when exposed to temperatures 

above rated time-at-temperature exposure limits, the seal material may retain sealing capability. 

Additional testing is needed to obtain a broader technical basis for determining the performance 

of seal materials at elevated temperatures, to evaluate the influence of such variables as the 

size of pressure vessel, the magnitude of the pressure differential at the high temperature 

exposure, and the effectiveness of double versus single O-ring seals. (SNF packages typically 

use double O-ring seals, and this testing used only a single O-ring in the test vessels.) These 

seal tests demonstrate that ethylene propylene seals such as those that are used for the GA-4 

package may have a performance envelope that far exceeds the conservative temperature 

limits indicated in the ratings for long-term performance provided by the seal manufacturer. 

Seals will not necessarily exhibit catastrophic leakage when exposed to temperatures beyond 

design basis, even at the extreme thermal exposure postulated for the MacArthur Maze fire 

scenario.
8-8

The results of the thermal analysis indicate that for the GA-4 package in the MacArthur Maze 

fire scenario, all seals would be significantly challenged. Based on the seal testing conducted 

by NRC discussed here, it is remotely possible that some sealing capability would remain for the 

seals even after exposure to excursion temperatures well above their rated temperatures. 

However, for the purposes of determining the potential release from the GA-4 package in this 

fire scenario, a simple pass/fail criterion is used to evaluate potential seal performance. If the 

temperatures predicted in the transient exceed the maximum recommended service 

temperature for the seal material, the seal is assumed to fail. This evaluation shows that the 

assumption of complete seal failure is bounding for the performance of the GA-4 package seals, 

and may be quite conservative. Nevertheless, this assumption is the basis for determining that 

a release is possible from the GA-4 package in this fire scenario due to package containment 

boundary leakage. The analyses presented in the following sections determine the character 

and amount of material that could be released.

8.3 Potential Release Issues

NRC staff evaluated the potential for release of radioactive material from the GA-4 package as a 

consequence of the MacArthur Maze fire scenario. Based on the results of the thermal analysis 

(as discussed in Sections 7.1 and 7.2), there is a possibility of a release from the package 

because of failure of components that make up the containment boundary of the package. 

Calculated temperatures in the region of the lid, drain valve/port, and gas sample valve/port 

seals during the transient far exceed the continuous use temperature limits for the seal material, 

as discussed in Section 8.2 above. In addition, the peak temperatures predicted for all fuel rods 

in the package reach the range where burst rupture of zircaloy cladding can occur, as discussed 

in the presentation of results of the thermal analyses in Sections 7.1 and 7.2. Therefore, the 

potential exists for the release of fission products and spent fuel particles, as well as particulate 

resulting from CRUD1 detaching from the fuel rod surfaces.

Results of analysis of fuel performance for the conditions encountered in this fire scenario are 

presented in Section 8.3.1. Evaluations of the potential for release from fuel rods to the GA-4 

package cavity are presented in Section 8.3.2. Evaluations of the potential for release from the 

GA-4 package to the surrounding environment are presented in Section 8.3.3.

8.3.1 Fuel Rod Cladding Performance

Spent fuel has two potential sources of radioactive material that could serve as source terms for 

a release from an SNF transportation package; the CRUD on the rod outer surface, and the 

radioactive material (fission products and fuel fragments) confined within the metal cladding. 

The Standard Review Plan (NUREG-1617 2000) specifies the assumption of 100% spallation of 

CRUD from fuel rod surfaces for HAC analyses. For consistency, this assumption is also 

applied to the MacArthur Maze scenario. Determining the amount of material that could 

potentially be released from within the rods, however, requires additional analysis of fuel rod 

behavior for the conditions of the fire scenario. If it can be shown that the fuel rods remain 

intact throughout the fire scenario, there would be no release of material from within the rods. If 

 1 Chalk River Unknown Deposit; generic term for material deposited on the rod surface from the coolant 

during reactor operations. The significant activated element is Cobalt-60. Regulatory guidance specifies 

a bounding value of 140 μCi/cm2 for spent fuel rods in PWR assemblies. A bounding estimate for total 

activity due to CRUD can be calculated from the total fuel rod surface area and the age of the fuel (i.e., 

time out of the reactor).
8-9

conditions are such that the fuel rods could fail, the nature and severity of the potential failure 

must be determined.

The predicted fuel cladding temperatures obtained with the COBRA-SFS model of the GA-4

were used to evaluate the potential for rod failure in the severe conditions of the MacArthur 

Maze fire scenario. The results obtained with the ANSYS model are more conservative, as 

discussed in Section 7.1 and 7.2, due to the more conservative representation of the fuel using 

the homogeneous k-effective model. The COBRA-SFS model, with a more detailed 

representation of the fuel region using the rod-and-subchannel approach, produces predictions 

of fuel and cladding temperatures with a more accurate evaluation of thermal radiation and the 

transient thermal inertia of the fuel and cladding. However, the many conservatisms in the 

modeling approach and fire scenario definition that apply equally to the ANSYS model and the 

COBRA-SFS model ensure that the cladding temperature results obtained with this code are 

also very conservative, even though they do not include the specific conservative effect of the 

effective conductivity model in the representation of the fuel assemblies. 

Based on the predicted fuel cladding temperatures from the COBRA-SFS modeling results for 

the MacArthur Maze fire scenario, fuel performance was evaluated using the burst rupture 

model in the FRAPTRAN-1.4 code (NUREG/CR-7023 2011). In the FRAPTRAN code, cladding

rupture is evaluated with a burst stress/strain model developed from test data obtained for loss 

of coolant accident (LOCA) analysis and reactivity insertion accident (RIA) evaluations. Burst 

rupture is the expected mechanism of failure for fuel rods in the reactor core when subjected to 

severe accident conditions, and is a potential failure mode for spent fuel at high temperatures. 

Creep rupture is considered a possible alternative mechanism of failure for spent fuel rods. To 

evaluate this possibility, a separate analysis was performed with a creep rupture model, using 

the FRAPCON-3.4 code (NUREG/CR-7022 2011) in conjunction with the DATING code 

(Simonen and Gilbert 1988). The version of the code used in this analysis has been updated 

with creep coefficients from creep tests on irradiated cladding (Gilbert et al. 2002), for the 

temperatures in the range predicted for the hottest rod in the MacArthur Maze fire scenario.

Fuel performance evaluations used the same design basis fuel configuration assumed in the 

thermal analysis. The fuel rod cladding initial conditions and cladding temperatures assumed 

during the fire scenario are summarized as follows:

• Westinghouse 14x4 fuel design operating at a rod average linear power rating of 5.7 kW/ft,

up to 33 GWd/MTU rod average burnup in-reactor prior to discharge

• Initial as-fabricated rod pressure of 460 psig at room temperature 

• Rod pressure at room temperature after irradiation to 33 GWd/MTU is 651 psig. 

– The increase in pressure from the as-fabricated pressure of 460 psig is due mainly to a 

reduction in rod void volume because of fuel swelling and cladding creepdown. There is 

also a small increase in pressure due to release of fission gas (0.0024 release fraction) 

from the pellets to the fuel rod void space. Fission gas release fractions and 

consequently rod pressure would be significantly higher (between 0.05 to 0.20) for a 

peak power rod operating rod in the core. A release fraction of 0.002 to 0.005 would be 

representative of the majority of rods (but not the peak operating rods) in a core at this 

discharge burnup.

• Initial conditions at the start of the fire are assumed to be NCT, which results in an initial 

steady-state peak cladding temperature of 293°F (145°C) with a rod pressure of 922 psig

(hoop stress approximately 50 MPa)
8-10

• Cladding temperature as a function of time for the hottest rod in the package, as predicted in 

the COBRA-SFS thermal analysis for the fire conditions (see Figure 7.8, which shows the 

peak cladding temperatures; however, the complete axial temperature distribution as a 

function of time was provided as input to the burst rupture calculations, not simply the peak 

temperature)

The cladding temperatures from the fire, as calculated with COBRA-SFS, and rod pressures 

calculated by FRAPCON-3.4 (NUREG/CR-7022 2011) assuming the spent fuel had been 

subjected to normal reactor operation at 5.7 kW/ft, were input into FRAPTRAN-1.4 to calculate 

the cladding stresses. The FRAPTRAN-1.4 cladding burst model was also used to calculate the 

rupture temperature during the fire. The calculated cladding temperatures during the fire from 

the COBRA-SFS analysis, and the calculated hoop stresses obtained from FRAPTRAN-1.4 for 

the fire conditions were input into FRAPCON-DATING to calculate cladding rupture based on 

the out-of-reactor creep relationship in the DATING subroutine. 

8.3.1.1 Fuel Rod Cladding Performance: FRAPTRAN Evaluation

The peak cladding temperatures calculated with COBRA-SFS for the MacArthur fire were 293°F 

(145°C) at the start of the fire and reached a peak cladding temperature of 1388°F (753°C) in 

the fire transient. Based on these temperatures, the calculated cladding hoop stress is 50 MPa 

at the start of the fire and reaches a peak of 121 MPa just prior to predicted cladding rupture at 

1098°F (592°C), as predicted with the burst strain model in FRAPTRAN-1.4. This relatively low 

rupture temperature reflects the conservatism in the cladding temperature history predicted in 

the thermal analysis, and the uncertainty in the FRAPTRAN predictions at the relatively low 

heating rate for the cladding in this fire scenario. 

The FRAPTRAN code was designed to predict nuclear fuel behavior during reactor accidents.

In particular, failure models have been developed to provide reasonably accurate predictions for 

Reactivity Initiated Accident (RIA) and Loss-Of-Coolant Accident (LOCA) failures. For the case 

of a fire accident scenario in transportation of spent nuclear fuel, potential cladding failure can 

occur as a result of temperature increase and the associated rod internal pressure increase. 

These concurrent temperature and pressure increases can result in sufficient stress to cause 

ballooning and rupture of the cladding due to rapid high temperature creep. In a FRAPTRAN 

calculation of such an event, if the temperature and stress in the cladding are such that the 

cladding deforms to its uniform elongation in the hoop/circumferential direction, fuel rod 

ballooning is predicted. If the stress exceeds a correlated temperature-dependent level, 

cladding rupture will be predicted. In the less severe conditions of the HAC fire, the stress in the 

cladding is not expected to exceed the elastic limit, and therefore burst rupture would not be 

predicted with the models in the FRAPTRAN code.

Ballooning and rupture models such as those in the FRAPTRAN code have been developed 

with the specific intent to accurately predict cladding failures during a LOCA, where the 

temperature increase rate is typically much higher (on the order of 10°C/s or higher) than in the 

case of fire scenarios (typically on the order of 0.2°C/s for the HAC fire, and for the MacArthur 

Maze fire scenario, conservatively estimated to be 0.27°C/s). For a given stress level, a slower 

heatup rate will generally tend to result in a lower rupture temperature, but there is very little 

data in the FRAPTRAN validation database that has heatup rates below 1°C/s. There is much 

more data at higher heating rates, ranging from 5°C/s to >30°C/s, as fully documented in the 

FRAPTRAN code manual (NUREG/CR-7023 2011). Figure 8.6 shows a summary of burst 

temperature data in the FRAPTRAN modeling database that are from tests with heating rates in 

the range 0.08 to 5°C/s. (This data set is from NUREG/CR-0344.) The lowest burst (i.e.,
8-11

rupture) temperatures observed in this limited subset of the validation database are around 

667°C (1232°F). Due to the sparseness of the data in this low temperature range, burst rupture 

temperatures predicted with FRAPTRAN for heatup rates below 1°C/s have a greater 

uncertainty than predictions obtained for higher heating rates, where the database is more fully 

populated. In particular, predictions of burst rupture temperatures lower than 667°C (1232°F) 

should be evaluated as indicative of the possibility of rupture, rather than absolute indicators of 

rupture, since these results are outside the available validation database.

Figure 8.6. Experimental Results for Rod Burst Rupture Testing at Low Heating Rates (data 

from NUREG/CR-0344)

Based on the validation range of the models in FRAPTRAN, and the conservative assumptions 

in the thermal modeling that impose an extraordinarily severe temperature transient on the fuel 

rods within the GA-4 package in this fire scenario, the predicted cladding rupture at 1098°F 

(592°C) obtained in the FRAPTRAN analysis can be considered an extremely conservative 

result. However, the predicted peak cladding temperature obtained in the thermal modeling is 

1388°F (753°C) in this fire scenario, and this temperature is well within the range where the data 

in Figure 8.6 shows burst rupture temperatures at low heating rates. The specific temperature 

value for burst rupture predicted with FRAPTRAN for these conditions may be quite 

conservative, and may have a fairly large uncertainty, but there is little uncertainty that the 

cladding would at some point fail by burst rupture if subjected to the severe conditions predicted 

for the fuel in the GA-4 package in the MacArthur Maze fire scenario. 
8-12

8.3.1.2 Fuel Rod Cladding Performance: FRAPCON/DATING Evaluation

The cladding failure temperature predicted with the creep model in the DATING code is 1229°F 

(665°C), which is significantly higher than the burst rupture temperature of 1098°F (592°C) 

obtained in the FRAPTRAN analysis. The DATING code was designed to predict creep failures 

and temperature limits for dry storage of spent fuel, based on creep failures of the cladding.

The range of applicability of the DATING code is for lower temperatures, resulting in much 

longer times to failure, when compared to the FRAPTRAN ballooning models. The databases 

used to develop creep rate correlations and creep rupture models for the DATING code span 

temperature ranges that are in general lower than the temperatures and heating rates typically 

encountered in fire scenarios. 

The DATING code is a more general creep prediction tool than FRAPTRAN, with its ballooning 

and rupture models, which are effectively high temperature creep models. However, it must be 

noted that, as with FRAPTRAN, the DATING code is being applied outside its validation 

databases when used to evaluate cladding response to the conditions of the MacArthur Maze 

fire scenario. However, the results obtained with both modeling tools show that although there 

might be some uncertainty as to the exact temperature at which it would occur, fuel cladding 

could and probably would fail, if subjected to the severe conditions postulated for the MacArthur 

Maze fire scenario. For the purposes of this evaluation, the predicted temperatures from these 

codes are taken at face value, and treated as conservative estimates of rupture temperature in 

the evaluation of potential consequences of this fire scenario. 

Table 8.1 summarizes the results of the fuel performance modeling analyses for temperatures 

on the hottest rod in the MacArthur Maze fire scenario, as predicted with the COBRA-SFS 

model. The burst rupture and creep rupture models both predict that the hottest fuel rod would 

rupture if subjected to the temperatures predicted in this fire scenario. Furthermore, the peak 

temperature on the hottest rod at the time of rupture is eventually exceeded by all rods in the 

package during the transient, which suggests that there is the potential for all rods in the 

package to rupture in this fire scenario. 

Table 8.1. Results of Fuel Performance Analyses in the MacArthur Maze Fire Scenario

LOCA Burst Strain Model (FRAPTRAN) Creep Rupture Model (FRAPCON/DATING)

Cladding Temperature Rupture Conditions Cladding Temperature Rupture Conditions

1097°F (592°C) rod rupture in end region 1229°F (665°C) rod rupture near end

The burst rupture and creep rupture models predict cladding failure at a single location along 

the axial length of a fuel rod. Based on the temperature predictions obtained with the COBRASFS model, which omits the impact limiters, the fuel performance models predict rod rupture in 

the end region of the rod. Temperature distributions obtained with the ANSYS model, which 

assumes the impact limiters remain in place throughout the transient, result in the highest 

temperatures occurring near the axial center of the fuel region, and rod rupture would be 

expected near the middle of the rod for this package configuration. Since the design basis fuel 

for the GA-4 is low burnup (i.e., no more than 45 GWd/MTU), the degree of pellet-clad 

interaction would be relatively limited, and a single rod breach would be expected to effectively 

depressurize the fuel rod. Therefore, no additional ruptures are predicted on a given rod, and 

potential release calculations are based on the assumption of one rupture per rod. 
8-13

The peak fuel cladding temperatures predicted with the ANSYS model are somewhat higher 

than the peak temperatures on the rod ends predicted with COBRA-SFS (see Sections 7.1 and 

7.2). Furthermore, the rod temperatures in both analyses remain much higher than the 

predicted rupture temperatures for an extended period of time. It is therefore reasonable to 

conclude that for the temperature distribution predicted with the ANSYS model (with the impact 

limiters in place), rod ballooning and rupture would also be expected to occur, but in the central 

region of the rod, rather than at an end. Table 8.2 summarizes the elapsed time and time 

duration that the hottest rod peak temperatures are predicted to exceed the calculated burst 

rupture temperatures. 

Table 8.2. Time above Predicted Rod Rupture Temperatures in the MacArthur Maze Fire 

Scenario

Rod Condition

PCT at time 

of rupture

COBRA-SFS model ANSYS model

Max PCT in 

fire transient

1388°F

(753°C)

Max PCT in 

fire transient

1433°F 

(779°C)

Elapsed Time 

(hours)

Time Above 

Rupture 

Temperature 

(hours)

Elapsed 

Time (hours)

Time Above 

Rupture 

Temperature 

(hours)

rod rupture 

(burst strain 

model)

1097°F 

(592°C) 0.8 16 0.69 >14.5

rod rupture 

(creep model)

1229°F 

(665°C) 1.15 10.5 0.97 11.5

PCT = Peak Cladding Temperature

Based on the burst strain model, the fuel rods are expected to rupture before the end of the fire. 

Based on the creep rupture model, the fuel rods would also be expected to begin rupturing 

before the end of the fire, but slightly later in the transient. Furthermore, the peak temperatures 

remain significantly above these predicted rupture temperatures for more than 10 hours. The 

results presented in Sections 7.1 and 7.2 show that the fuel rod temperatures continue to 

increase even after the end of the fire, because of thermal inertia and build-up of decay heat 

that is not removed from the package during and immediately after the fire.

By the time of the secondary peak of 1348°F (731°C) in cladding temperature predicted with the 

COBRA-SFS model, which occurs at 250 minutes elapsed time (142 minutes after the end of 

the fire), the peak temperature on every rod in the package exceeds the highest temperature 

predicted for rod rupture (1229°F [665°C]). The peak temperature of 1343°F (728°C) predicted 

with the ANSYS model is at essentially the same value as that predicted with the COBRA-SFS 

model at this point in the cooldown transient. More significantly, at this time the lowest peak rod 

temperature is 1285°F (696°C) in the COBRA-SFS model results, and the lowest axial peak 

temperature predicted in the fuel region in the ANSYS model is approximately 1134°F (612°C). 

Based on these results, it is assumed that all of the rods in each of the four assemblies within 

this package would rupture in the MacArthur Maze fire scenario. 

8.3.2 Potential Release to GA-4 Package Cavity

Determining potential release quantities from an SNF package involves first determining the 

amount of material that is available for release from the fuel rods, and then determining the 

amount of this material that can be released from the package. This section presents analyses 
8-14

performed by NRC staff to determine the total amount of activity that could be released from the 

four assemblies defining the design-basis payload for the package, as described in the GA-4 

SAR. Analyses to determine the potential for release from the package to the environment are 

presented in Section 8.3.3.

Typically, release quantities are expressed in terms of release fractions, a ratio calculated as 

the amount of material actually released divided by the total amount available for release. 

Regulatory guidance for determining the releasable source term for SNF transportation 

packages is provided in the Standard Review Plan for Transportation Packages for Spent 

Nuclear Fuel: Final Report, NUREG-1617 and in Containment Analysis for Type B Packages 

Used to Transport Various Contents, NUREG/CR-6487. The release fractions specified in 

these documents are listed in Table 8.3. These release fractions define bounding values for the 

fraction of material that is assumed to be released from the fuel rods to the package under NCT 

and HAC. 

Table 8.3. Bounding Values of Release Fractions from Ruptured Fuel Rods

Radionuclide Group Release Fraction

(NCT) (HAC)

non-reactive gases (e.g., Kr-85) 0.3 0.3

volatile gases (e.g., cesium and iodine compounds) 0.0002 0.0002

particulate (fuel fragments or fines) 0.00003 0.00003

CRUD spallation fraction 0.15 1.0

The potential release from the GA-4 package corresponding to the release fractions in Table 8.3

is a function of the contents of the package. The radionuclide inventories for the two design 

basis fuel configurations for the GA-4 package were obtained using ORIGEN-ARP (Gauld et al. 

2009). The source term inventories obtained in these calculations are listed in Table 8.4. 

Consistent with the criticality and shielding calculations in the GA-4 package SAR (General 

Atomics 1998), WE14x14 fuel at 35 GWd/MTU burnup and 10-years cooling is bounded by 

WE15x15 at 35 GWd/MTU and 10-years cooling. Therefore, all source term and potential 

release calculations are performed assuming 10-year-old WE 15x15 fuel at 35 GWd/MTU, even 

though the thermal analysis is based on WE 14x14 fuel geometry. This is a conservative 

assumption, since the temperatures obtained with WE 14x14 fuel would be slightly higher than 

those predicted for WE 15x15 fuel for the same design basis decay heat loading. Table 8.4 lists 

the calculated source terms for a single assembly. The total inventory within the GA-4 is four 

times the quantities listed in this table, since this package can carry up to four fuel assemblies.

Table 8.4. Radionuclide Inventory for a Single Assembly in the GA-4 Package

Nuclide

Activity (Ci)

WE 15x15

(45 GWD/MTU; 15 yrs cooled)

WE 15x15

(35 GWD/MTU; 10 yrs cooled)

Ag-110m 2.44E-03 2.42E-01

Am-241 5.56E+03 3.54E+03

Am-242 1.59E+01

Am-242m 2.15E+01 1.60E+01

Am-243 9.68E+01 4.74E+01
8-15

Table 8.4. (continued)

Nuclide

Activity (Ci)

WE 15x15

(45 GWD/MTU; 15 yrs cooled)

WE 15x15

(35 GWD/MTU; 10 yrs cooled)

Ba-137m 1.78E+05 1.57E+05

Ce-144 3.21E+00 2.84E+02

Cm-242 1.77E+01 1.32E+01

Cm-243 5.71E+01 3.22E+01

Cm-244 9.37E+03 3.99E+03

Cm-245 1.64E+00

Cs-134 2.78E+03 1.01E+04

Cs-137 1.88E+05 1.66E+05

Eu-152 6.17E+00 7.38E+00

Eu-154 5.41E+03 5.54E+03

Eu-155 8.50E+02 1.23E+03

H-3 6.42E+02 6.37E+02

Kr-85 8.28E+03 9.80E+03

Np-239 9.68E+01 4.74E+01

Pm-147 6.01E+03 2.27E+04

Pr-144 2.84E+02

Pr-144m 3.97E+00

Pu-238 8.78E+03 5.31E+03

Pu-239 6.51E+02 6.41E+02

Pu-240 1.26E+03 1.04E+03

Pu-241 1.50E+05 1.60E+05

Pu-242 6.98E+00 4.22E+00

Rh-106 1.17E+03

Ru-106 4.74E+01 1.17E+03

Sb-125 4.41E+02 1.28E+03

Sm-151 7.47E+02 7.02E+02

Sn-119m 2.47E-02

Sn-121 1.80E+01

Sn-121m 2.94E+01 2.32E+01

Sr-90 1.19E+05 1.14E+05

Tc-99 3.38E+01 2.74E+01

Te-125m 1.08E+02 3.13E+02

U-234 1.91E+00 1.98E+00

U-237 3.84E+00

Y-90 1.19E+05 1.14E+05

Zr-93 4.26E+00 3.46E+00
8-16

The bounding values for release fractions defined in Table 8.3 were applied to the source terms 

listed in Table 8.4 to determine a bounding estimate of the activity that could be released from 

the four fuel assemblies to the GA-4 package interior in the MacArthur Maze fire scenario. 

Figure 8.7 shows the activity released to the package for the source term inventory from Table 

8.4 for the bounding configuration of 10-year-cooled WE 15x15 fuel at 35 GWd/MTU. Figure 

8.8 shows the activity released to the package for the source term inventory from Table 8.4 for 

15-year-cooled WE 15x15 fuel at 45 GWd/MTU.

Figure 8.7. Summary of Activity in Radionuclides Released to GA-4 Package Cavity from WE 

15x15 (35GWd/MTU, 10-yrs-cooled fuel) for Bounding Release Fractions Specified 

in NUREG-1617
8-17

Figure 8.8. Summary of Activity in Radionuclides Released to GA-4 Package Cavity from WE 

15x15 (45GWd/MTU, 15-yrs-cooled fuel) for Bounding Release Fractions Specified 

in NUREG-1617

8.3.3 Potential Release from GA-4 Package in MacArthur Maze Fire Scenario

Release rates from SNF packages are typically calculated for NCT and HAC using models 

based on guidance in NUREG/CR-6487, which contains models that reference ANSI standards 

for leakage tests on packages for shipment of radioactive materials (ANSI N14.5 1997). The 

analyses presented in the GA-4 SAR show that as long as the package seals remain intact, the 

package can be expected to meet all containment requirements, and potential releases from the 

package would be well below regulatory limits. However, as discussed in Section 8.2.1, the GA4 package seals are predicted to exceed operational temperature limits after approximately 1.5 

to 2.5 hours of exposure to the thermal conditions of the MacArthur Maze fire scenario. In 

addition, as discussed in Section 8.3.1, all fuel rods in all four assemblies contained within the 

GA-4 package are predicted to exceed temperatures at which burst rupture or creep rupture of 

the zircaloy cladding would be expected to occur. Therefore, there is the potential for leakage 

of radioactive material from the GA-4 package after this point in the fire transient.

Determining an appropriate leak rate for the package in the conditions predicted for the 

MacArthur Maze fire scenario presents an interesting challenge. The models for leak rates 

derived from the ANSI standard ANSI N14.5 are not based on the assumed seal conditions in 

this fire scenario, and there is very little information in the literature on leak rates associated with 

failed seals. In typical engineering applications, the leak rates of failed seals are unacceptable 

by definition, and their potential magnitude is of no practical interest. What little information to 

be found tends to focus rather narrowly on special applications where time-to-failure could be a 

critical design parameter (e.g., equipment that will be sent into orbit). In these types of studies, 
8-18

the focus is on the time interval to the point where the seal begins to leak, not on the leak rate 

itself, and the work is mainly interested in modes of seal failure or seal behavior prior to failure.

It was therefore necessary to develop a modeling approach to determine a reasonable bounding 

leak rate for the GA-4 package for the long portion of the transient following the time after 

assumed seal failure due to exceeding thermal operating limits. Section 8.3.3.1 presents the 

model developed for this analysis, and describes its application to the GA-4 package in the 

MacArthur Maze fire scenario. Section 8.3.3.2 presents the potential release calculations for 

the GA-4 package, based on the leak rate determined with this model.

8.3.3.1 Leak Rate Model for GA-4 Package without Seals

For leak rate modeling, the interface between the closure lid and end flange of the package 

body is of greatest significance. (There is also the potential for leakage paths through the gas 

sample valve/port and the drain valve/port; this is discussed in Section 8.3.3.2.) Failure of the 

seals in the fire scenario is conservatively treated in this evaluation as if the seals simply cease 

to exist after one hour of the fire duration. This timeframe conservatively bounds the interval of 

the estimated time when all seals are predicted to have exceeded operating temperature limits. 

The possibility of damaged seal material affecting the geometry of the leakage path is ignored. 

If it is assumed that there is no O-ring seal material remaining in the seal grooves of the lid and 

flange, and the only barrier to flow through the interface is the actual physical contact between 

these two components.

The closure lid and body flange both have smooth metal surfaces where the two components 

are in contact, and the closure bolts are torqued to a specified pre-load, such that there is a 

positive and essentially uniform clamping force at the interface. The evaluations presented in 

Section 7.3, investigating the response of the lid closure bolts to the extreme thermal 

environment of the MacArthur Maze fire scenario, show definitively that the bolts maintain a 

positive clamping force throughout the fire transient, including the long cooldown back to 

ambient conditions.

Flow of gas through the very narrow space between the closure lid and body flange can be 

treated as analogous to fluid flow through fractured material in which the local scale of motion 

can be approximated by the cubic law for flow between parallel plates. This is a simplified form 

of the momentum conservation equation, and is a function of the geometry of the flow path and 

the driving pressure difference between the package interior and the external environment 

(Brown 1987). A formulation of this relationship, expressed in cylindrical coordinates, is given 

by

QQLLLL = 2ππ � ddee

3

12μμ

� � ∆PP

ln( rroo⁄rrii)

�

where:

QLR = volumetric flowrate through the leakage path

de = equivalent gap between surfaces in contact

μ = viscosity of flowing gas

ΔP = driving pressure difference

ro = outer radius

ri = inner radius
8-19

The equivalent gap between the surfaces in contact is the critical unknown in the above 

equation, since all other parameters can be readily determined from the geometry of the GA-4 

package closure lid and flange, and the conditions calculated for the MacArthur Maze fire 

scenario with the thermal models. The actual gap is a function of the surface roughness of the 

components in contact and the clamping force holding them together. This gap cannot readily 

be estimated with any degree of certainty without knowing the exact microscale geometry of the 

surfaces involved. 

Therefore, an alternative approach was developed by considering another much simpler 

physical process in which the gap between two surfaces in direct contact has an important 

effect on physical behavior; the flow of heat between two surfaces in direct or very close 

contact. The thermal resistance between two such components is a strong function of the 

contact pressure and surface texture of the two surfaces, and is typically expressed in terms of 

the overall thermal contact resistance, as

RRtttt = ddee

kk

where:

Rt c = thermal contact resistance (m2

-K/W) as a function of contact pressure and 

surface texture for two surfaces in direct contact

de = equivalent gap between surfaces in contact

k = thermal conductivity of gas in spaces between contacting surfaces

Using thermal contact resistance data for stainless steel surfaces (Shajaefard and Goudarzi 

1987) as a function of contact pressure and surface roughness at the interface, and assuming 

helium gas in the very constrained spaces between the contacting surfaces, the above 

relationship can be used to determine an equivalent gap for the closure lid and package flange 

in the MacArthur Maze fire scenario.

The results presented in Section 7.3 for evaluations of the lid closure bolt response to the 

MacArthur Maze fire scenario using detailed FEA modeling provide a history of the clamping 

force between the closure lid and package body flange. Figure 8.9 shows the estimated 

equivalent gap after seal failure, based on the contact resistance as a function of the lid/flange 

contact pressure due to the clamping force, and the thermal conductivity of helium gas.
8-20

Figure 8.9. Equivalent Gap between Closure Lid and Package Body Flange after Seal Failure

As discussed in Section 7.3, the clamping force on the lid increases during the post-fire 

cooldown, due to differential thermal contraction between the nickel alloy closure bolts and the 

stainless steel lid and package body. The effect is to essentially close the gap entirely, for all 

practical purposes, by about 3.75 hours into the fire transient. This effectively limits the 

“window” of time in which material could leak out of the package to less than 3 hours. This has 

the effect of greatly reducing the potential for a substantial release of radioactive material from 

the package, as shown by the release evaluations in Section 8.3.3.2. 

8.3.3.2 Bounding Estimate of Potential Release from GA-4 Package

Using the leak rate model and equivalent gap width relationship presented in Section 8.3.3.1, a 

conservative bounding estimate was obtained for potential release of radioactive material from 

the GA-4 package in the MacArthur Maze fire scenario. The fluid viscosity of pure helium was

used for this calculation, rather than attempting to quantify the viscosity of the mixture of helium 

and fission gases that would actually be in the package following the rod ruptures. This is a 

conservative assumption, since the viscosity of the mixture would be higher than the viscosity of 

pure helium. The difference between the mixture property and that of pure helium would in any 

case be expected to be small, since the gas released from the fuel rods would consist mainly of 

helium.

The pressure difference driving the volumetric flow through the interface between the package 

cavity and ambient was calculated assuming a constant external ambient pressure of 1 atm. 

The internal cavity pressure was calculated using the ideal gas law, based on the average gas 

temperature predicted with the ANSYS thermal model. The initial pressure in the cavity was 

assumed to be at the Maximum Normal Operating Pressure (MNOP) for the GA-4 package. 
8-21

This is a conservative initial pressure, as it corresponds to the pressure effect of 100% of the 

fuel rods in the package having ruptured, and the density change is determined for B&W 15x15 

fuel1

, which is the most limiting fuel configuration for the maximum operating pressure. This 

approach provides a bounding estimate of the cavity internal pressure throughout the fire 

transient, and avoids the complication of changing the gas density in the package at the 

predicted time of rod rupture in this analysis.

Figure 8.10 shows the predicted leak rate as a volumetric flow of helium gas through the 

equivalent gap. Two leak rate calculations were performed; a bounding case in which the 

package gas density was assumed to remain constant throughout the transient, and a more 

realistic case in which the change in gas density (and hence pressure) due to outflow of gas 

from the package was accounted for. The difference between the two cases is relatively small, 

due to the small leak rates predicted for this configuration with the closure lid clamped tightly to 

the package body flange throughout the transient.

Figure 8.10. Volumetric Leak Rate for GA-4 Package after Seal Failure

The release calculations were performed assuming the bounding leak rate over time (as shown 

in Figure 8.10), providing a bounding estimate of potential release from the package. The 

activity within the package cavity was assumed to be uniformly distributed within the gas, with 

 1 For the thermal analysis, the most limiting fuel is WE 14x14, and this is the fuel configuration 

represented in the thermal models, as described in Sections 5 and 6. However, for maximum cavity 

pressure evaluations, as presented in the SAR, B&W 15x15 is the limiting configuration, due to the fuel 

rod design of this fuel assembly. Therefore, the cavity pressure obtained assuming 100% rod rupture (for 

four assemblies) with this fuel design was used, as a conservatism in the leak rate evaluations.
8-22

all particulate (i.e., fuel fines and spalled CRUD) suspended in the gas as an aerosol. It is 

assumed that all solid particles remain suspended in the gas. The total release of each 

component was calculated simply as the activity of that component times the volumetric fraction 

of gas escaping from the package. 

A number of additional conservatisms were incorporated into the release calculation, including 

the following assumptions

• the entire quantity of fuel and CRUD particulate was assumed to remain suspended in the 

gas within the cavity; the possibility of particulate settling or plating out on internal package 

structures was ignored

• the filtering effect of the equivalent gap size was neglected; the maximum size of the 

equivalent gap is only about 2 micrometers, and the release calculations do not consider that

a large percentage of the fuel and CRUD particulate simply could not escape from the 

package, due to the small size of the gap

These assumptions result in a very conservative estimate of the amount of activity that could 

escape from the package in the approximately 2.7 hours that the package could sustain a 

significant leakage.

The activity of the large number of radionuclides comprising the estimated release can be more 

conveniently expressed in combined form, as a function of their combined isotopic A2 limit1

values from 10 CRF 71, Appendix A.

The A2 value for a mixture of normal form material can be determined using the following 

relationship from 10 CFR 71, (Appendix A, Section IV.d), as

AA2 ffff mmmmmmmmmm mm = 1

�∑ ff( )

AA2( ) nn

ii=1 �

where 

n = number of radionuclides in mixture

f(i) = fraction of total mixture activity due to the i

th component

A2(i) = A2 value for the i

th component

 1 An A2 quantity is defined in 49 CFR 173.403 as the maximum activity of a Class 7 (radioactive) material 

permitted in a Type A package, which does not require an accident resistant design. The amount of 

material that constitutes an A2 quantity depends on its specific activity and other radiological properties. 

Appendix A of 10 CFR 71 specifies the specific A2 quantities for a large number of radioactive materials, 

and defines methods for calculating values for materials not listed in the table. Spent nuclear fuel 

requires a Type B package, which can carry more than an A2 quantity of radioactive material, but must 

retain the integrity of containment and shielding under normal conditions of transport (as per 49 CFR 173) 

and meet the release limits of less than an A2 per week for hypothetical accident conditions.
8-23

Using this approach, the A2 for the mixture of radionuclides in the estimated potential release 

from the GA-4 package is calculated as 88 Ci (3.25 TBq) for WE 15x15 fuel at 45 GWd/MTU, 

15 yrs cooling. The corresponding result for WE 15x15 fuel at 35 GWd/MTU, 10-yrs cooling is a 

mixture A2 of 143 Ci (5.3 TBq). The calculation of the mixture A2 for each fuel configuration 

includes all fission gas and particulate released from the fuel, plus the CRUD assumed to spall 

from the exterior surfaces of the rods.

Based on the leak rate model, the total release from the package is estimated as 21 Ci

(0.78 TBq) for the higher burnup fuel, and as 24.5 Ci (0.91 TBq) for the lower burnup fuel. 

Expressed as an A2 fraction, relative to the mixture A2 for each configuration, these release 

rates are 0.24 and 0.17, respectively. Therefore, the bounding estimate of the total release from 

the package is 0.24 of the mixture A2 calculated assuming WE 15x15 fuel at 45 GWd/MTU, 

15 yrs cooling. As mentioned above, if the effect of particulate settling and the restriction of 

large particulate from passing through a small gap were taken into account, the release 

estimate would be significantly reduced. 

The evaluations of potential release from the GA-4 package assume that the estimated release 

by way of the closure lid is sufficiently conservative to be bounding on the possible contribution 

of leakage through the drain valve/port and gas sampling valve/port, which form part of the 

containment boundary of the GA-4 package. These components also contain seals that would 

be expected to exceed their operating temperature limits in this fire scenario, as discussed in 

Section 8.2 above. However, these penetrations of the package are less than an inch in 

diameter, compared to the approximately 2-ft diameter of the closure lid rim, and therefore do 

not provide a significant increase in the area available for potential leakage. In addition, the 

ports consist of long and convoluted flow paths that would tend to filter any particulate that might 

be carried through them to the ambient environment. The gas sample port is effectively blocked 

by the sample valve itself, which in addition to having primary and secondary O-ring seals, is 

threaded into place over a length of several inches. Also, for transport conditions, the outer face 

of the gas sample port is plugged with a threaded cover that extends to a depth of more than an 

inch. Similarly, the drain port is plugged by the drain valve, and capped with a threaded drain 

valve cover and port plug.

Based on the geometry of the valve/ports in this package, it is reasonable to assume that 

leakage from the package at these locations due to failed seals would be much less likely to be 

significant compared to leakage through for the much larger area and more direct flow path of 

the closure lid seal region. The conservative assumptions regarding the amount of material that 

could be transported out of the package through the lid closure/flange equivalent gap are 

sufficient to bound any possible contribution of the valve/port leakage paths. It is therefore 

justifiable to neglect the effect of the valve/ports, without compromising the conservatism of the

estimated leak rate and total package release calculations.

9-1

9.0 OVERALL SUMMARY AND CONCLUSIONS

The U.S. Nuclear Regulatory Commission has established requirements for packaging and 

transportation of spent nuclear fuel assemblies under NCT and for HAC. These requirements

(10 CFR 71) conservatively bound the expected range of conditions that an SNF package might 

be subjected to in the course of its service life. However, real-world accidents of greater 

severity are certainly possible, and rare as they may be, the NRC has proactively undertaken 

the examination of such accidents, to determine what the potential consequences might be, 

were such an accident ever to involve an SNF package. 

Two previous studies of transportation accidents, one resulting in a fire in a railroad tunnel 

(NUREG/CR-6886 2009) and one in a highway tunnel (NUREG/CR-6894 2007) were 

undertaken with three different SNF package designs. Based on conservative scenarios 

constructed from these real-world fire conditions, the results of these studies have shown that 

the design basis for SNF packages is sufficiently robust for them to survive such beyonddesign-basis conditions without adverse consequences to public safety. In all cases evaluated, 

the modeling results showed that the various SNF packages would be expected to maintain 

required shielding for ionizing radiation, and also would maintain the integrity of the containment 

boundary sufficiently to limit potential release of radioactive material from the packages to within 

regulatory bounds for accident conditions.

The MacArthur Maze accident of April 29, 2007 was selected as a third study in this series of 

evaluations of real-world accidents because of the severity of the fire and the unusual structural 

consequences, in which the heat from the fire caused the overhead roadway segments to 

collapse onto the roadway where the fire was burning. Since this was a highway accident, the 

only type of SNF package that could potentially be involved would be a LWT package. The 

General Atomics GA-4 LWT transportation package was selected for this investigation, mainly 

because it can carry a relatively large payload for an over-the-road transportation package, and 

therefore the potential consequences of package failure could be more severe than for 

packages with smaller payload capacities. The GA-4 package is designed to transport up to 

four intact PWR spent fuel assemblies, with a maximum total package decay heat load of 

2.5 kW.

Bounding Scenario for the MacArthur Maze Accident

The MacArthur Maze accident involved a gasoline tanker truck and trailer that overturned and 

caught fire on the I-880 connector of the MacArthur Maze interchange in Oakland, CA. The fire 

lasted approximately 108 minutes, consuming the entire load of gasoline fuel. The heat from 

the fire caused two sections of the overhead I-580 freeway to collapse onto the lower roadway, 

the first falling at approximately 17 minutes into the fire, the second collapsing on only one end, 

and reaching its final configuration by about 37 minutes. (Refer to Figure 1.1, which shows the 

configuration of the collapsed roadway, in an image taken in daylight the next day, after the fire 

was out. See Section 2.0 for a detailed discussion of the fire scenario, with images from a video 

of the fire.)

Based on fire modeling with the FDS code, and physical examination of material samples 

obtained from the damaged highway girders and the remnants of the tanker truck, a bounding 

fire scenario was defined for the thermal and structural evaluations of the potential effects of this 

fire on an SNF package. The complex and dynamic fire conditions are represented as a fully 

engulfing pool fire at 2012°F (1100°C) prior to the overhead roadway collapse, and as a slightly 

smaller and less severe fully engulfing pool fire at 1652°F (900°C) after the roadway collapse. 
9-2

These temperatures represent conservative bounding values for open pool hydrocarbon fires for 

any possible configuration of both the pre-collapse and post-collapse fire pools in this accident. 

As an additional simplifying conservatism in the definition of the scenario, it is assumed that the 

pre-collapse pool fire (at 2012°F [1100°C]) lasts for the full 37 minutes required for the 

completion of the collapse of the overhead segments. The smaller fire size is assumed as a 

step change to 1652°F (900°C), after 37 minutes, and this smaller pool fire is assumed to 

persist unchanged until the end of the fire, at 108 minutes. The fire scenario for modeling 

purposes also assumes that in the post-fire configuration, the fallen overhead roadway segment

completely covers the SNF package, resulting in an additional barrier to heat transfer from the 

package during the cooldown phase of the transient.

Thermal and Structural Modeling Approach and Summary of Results

Detailed thermal models of the GA-4 package were constructed for the ANSYS and COBRASFS codes, for transient evaluations to determine the temperature response of the package to 

the fire scenario, including the long post-fire cooldown transient. The initial condition of the 

package at the start of the fire scenario was defined as steady-state NCT. Additional detailed 

structural and thermal-structural models were also developed using ANSYS and LS-DYNA for 

the roadway and package, for evaluation of the package response to the effect of the roadway 

falling on it.

Considerable effort was given to defining bounding and conservative estimates of the possible 

configurations of the package on the roadway that could produce the “worst case” structural 

loading of the SNF package due to the overhead roadway segment falling on it. These 

evaluations showed that the worst that the overhead spans could do to the package imposed 

relatively innocuous loads on the stainless steel body and DU gamma shield compared to the 

HAC structural loading that the package is designed to withstand. At a nominal fully loaded 

weight of approximately 55,000 lb (nearly 28 tons), the package itself falling from a height of 

30 ft (9 meters) – the HAC package drop scenario – would be expected to do far more damage

than the roadway falling on it, even with the added impact of the projecting “blades” of the steel 

girders. The only real challenge of the overhead roadway drop in the fire scenario is that the 

impact is postulated to occur with the package at higher temperatures than are typically 

assumed in the structural analyses for HAC scenarios. This could potentially make the package 

more vulnerable to structural damage, due to the reduction in the strength of steel with 

increasing temperatures. However, the steel girders of the overhead span suffer more from this 

problem, and the weight of the overhead roadway concrete is not sufficient to impart significant 

loading to damage the package in any way.

Much more interesting structural analyses were undertaken to investigate in detail the response 

of the bolts attaching the impact limiters to the package, and the package lid closure bolts. 

Issues of bolt performance were further complicated by the use of thread inserts in all bolt 

attachments in the package, in which helical coils of Type 304 stainless steel fill the interface 

between the bolt threads and the threaded holes in the package body. Differential thermal 

expansion of the Inconel bolts relative to the XM-19 stainless steel package body, and different 

strength-versus-temperature properties of the three metals involved, results in a time-andtemperature-dependent history of force on the bolts that raised the possibility that the impact 

limiters might detach from the package. These material issues also raised the possibility that 

there could be a loss of clamping force between the lid and the package body during the postfire cooldown.
9-3

Detailed evaluations of the structural and thermal response of the impact limiter bolts to the 

conditions of the MacArthur Maze fire scenario with FEA modeling using ANSYS show 

definitively that the impact limiter bolts will not fail under these conservative and bounding 

thermal and structural loading conditions. Loss of the impact limiters is not a credible 

consequence of this fire scenario for the GA-4 package. Additional detailed evaluations of the 

response of the lid closure bolts to the fire scenario were undertaken with LS-DYNA, and the 

modeling accuracy was verified with classic bolt equation methods. These evaluations show 

unambiguously that the lid closure bolts maintain a positive clamping force between the 

package lid and body flange during all phases of the fire scenario, including the fire duration 

(108 minutes) and the very long cooldown period of approximately 400 hours, back to post-fire 

steady-state ambient conditions. This means that there is at all times forced metal-to-metal 

contact between the lid and the package body. This is particularly important to assessing the 

response of the GA-4 package to this fire scenario, because the thermal evaluation shows that 

the seals exceed their rated temperature limits within the first hour or so of the transient, and 

this metal-to-metal contact becomes the containment boundary of the package.

Thermal evaluations of the package response to this fire scenario predict that the peak cladding 

temperature would be expected to exceed the short-term limit of 1058°F (570°C) long before the 

end of the fire. Maximum cladding temperatures on all rods in the package are predicted to 

exceed this temperature limit in the course of the transient, and remain above this limit for 

several hours. In addition, the thermal inertia of the package and the insulating effect of the 

fallen overhead roadway, which is assumed to blanket the package during the post-fire 

cooldown, means that fuel cladding temperatures continue to rise for many hours after the end 

of the fire. The insulating effect of the impact limiters, which shield the package ends from 

direct heating by the fire, results in the cooler ends of the rods continuing to heat up for several 

hours after the end of the fire, as heat in the hot central region of the rods redistributes 

throughout the package. 

The maximum peak cladding temperature in the transient is predicted to be in the range of 

1350-1400°F (732-760°C), and occurs approximately 3 hours after the end of the fire. In 

addition, temperatures in the regions of the package seals exceed the seal material operating 

temperature limits for most of the fire transient and for several hours of the post-fire cooldown 

transient.

Fuel Rod Performance Evaluation

Based on the predicted fuel cladding temperatures from the COBRA-SFS modeling, fuel 

performance was evaluated by direct comparison to fuel rod burst data as a function of cladding 

hoop stress and temperature. In addition to comparison to relevant data, predicted fuel rod 

rupture temperatures were obtained using the burst rupture model in the FRAPTRAN-1.4 code 

(NUREG/CR-7023 2011). Creep rupture is considered a possible alternative mechanism of 

failure for spent fuel rods. To evaluate this possibility, a separate analysis was performed with a 

creep rupture model for the temperatures predicted for the hottest rod in the MacArthur Maze 

fire scenario, using the FRAPCON-3.4 code (NUREG/CR-7022 2011) in conjunction with the 

DATING code (Simonen and Gilbert 1988), which has been updated with creep coefficients 

from creep tests on irradiated cladding (Gilbert et al. 2002).

Fuel performance analyses for peak temperatures on the hottest rod in the MacArthur Maze fire 

scenario, as predicted with the COBRA-SFS model, predict cladding rupture temperatures of 

1097°F (592°C) using LOCA burst strain modeling (FRAPTRAN) and 1229°F (665°C) using 

creep rupture modeling (FRAPCON/DATING). Applicable experimental data (NUREG/CR-
9-4

0344) yields measured rupture temperatures in the range 1205-1256°F (652-680°C). The burst 

rupture and creep rupture models both predict that the hottest fuel rod would rupture if subjected 

to the temperatures predicted in this fire scenario. Furthermore, the peak temperature on the 

hottest rod at the time of rupture is eventually exceeded by all rods in the package during the 

transient, which suggests that there is the potential for all rods in the package to rupture in this 

fire scenario. 

Potential Radiological Consequences

Neutron and gamma radiation dose rates from the GA-4 package as a result of the postulated 

conditions of the MacArthur Maze fire scenario will not exceed the design basis of the package, 

which is well within the regulatory limits for hypothetical accident conditions. The neutron 

shielding is lost very early in the transient, but loss of the neutron shield tank is a design-basis 

assumption for this package in all HAC analyses. The more severe conditions of the MacArthur 

Maze fire can do no more damage to the GA-4 package neutron shield than is assumed a priori

in the HAC analyses. The gamma shielding for the GA-4 is provided by a layer of DU within the 

stainless steel package body. The shielding function of this material is not affected by the 

higher temperature it is predicted to reach in the MacArthur Maze fire scenario. There is no 

credible scenario in this fire accident that could result in neutron and gamma dose rates from 

the design-basis GA-4 package exceeding the regulatory limits for accident conditions.

Loss of the package seals due to exceeding seal material thermal limits means that there is the 

potential for radioactive material to escape from the package. Rupture of all rods in the 

package, as is predicted by the fuel performance analyses, based on the calculated thermal 

response of the fuel, means that fission gas and fuel particulate would be released to the 

package cavity. In addition, 100% spalling of CRUD from the external surfaces of the fuel rods 

is an assumed for all accident analyses for SNF packages, per NRC guidance. Therefore, it 

must be assumed that there is material available in the package cavity that could be released 

through the failed seals. But because the lid closure bolts maintain positive clamping force 

throughout the transient, it is not physically possible for very much of it to actually escape. 

Conservative and bounding modeling assumptions show that the maximum possible release

total release is 0.24 of the A2 quantity calculated for total activity of the mixture of radionuclides 

(comprised of fission gases, fuel particulate and CRUD) released from the package. The 

regulatory limit specifies a maximum allowable release rate of an A2/week. The predicted total 

release estimate of approximately one-fourth of a mixture A2 is below the prescribed limit for 

safety, and indicates that the potential release from this package in the MacArthur Maze fire 

scenario would not pose a risk to public health and safety.
10-1

10.0 REFERENCES

10 CFR 71. 2003. “Packaging and Transportation of Radioactive Material.” Code of Federal 

Regulations, U.S. Nuclear Regulatory Commission, Washington, D.C.

49 CFR 713.403. 2011. “Subpart 1 – Class 7 (Radioactive) Materials - Definitions.” Code of 

Federal Regulations, U.S. Nuclear Regulatory Commission.

ANSYS, Inc. 2003. ANSYS Users Guide for Revision 8.0. ANSYS, Inc., Canonsburg, 

Pennsylvania.

ANSI N14.5. 1997. American National Standard for Radioactive Materials—Leakage Tests on 

Packages for Shipment. American National Standards Institute.

Bahney RH III and TL Lotz. 1996. Spent Nuclear Fuel Effective Thermal Conductivity Report, 

BBA000000-01717-5705-00010 Rev. 00, TRW Environmental Safety Systems, Inc., Fairfax, 

Virginia.

Brekelmans J, R ven den Bosch, and K Both. 2008. Summary of Large Scale Fire Tests in the 

Runehamar Tunnel in Norway. UPTUN Research Program, official deliverable, Workpackage 2 

Fire development and mitigation measures, D213. Available from 

http://www.vegvesen.no/_attachment/61890/binary/15101.

Brockenbrough RL and FS Merritt, eds. 1999. Structural Steel Designer’s Handbook. 3rd

edition, MacGraw-Hill, Inc., New York.

Brown R. 1987. “Fluid Flow through Rock Joints: The Effect of Surface Roughness.” Journal 

of Geophysical Research, 92(B2):1337-1347.

Burian RJ. 1985. Response of Spent LWR Fuel to Extreme Environments. SAND85-7213.

Sandia National Laboratory, Albuquerque, New Mexico.

CHP - California Highway Patrol. 2007. California Highway Patrol Multi-Disciplinary Accident 

Investigation Team Report, MacArthur Maze Accident, April 29, 2007. CHP 558D (REV 6-84) 

OPI 065, Oakland, California.

Creer JM, TE Michener, MA McKinnon, JE Tanner, ER Gilbert, and RL Goodman. 1987. The 

TN-24P PWR Spent-Fuel Storage Cask: Testing and Analysis. EPRI-NP-5128/PNL-6054, 

Electric Power Research Institute, Palo Alto, California.

Dow Chemical Company. 2003. A Guide to Glycols. Available at 

http://www.dow.com/webapps/lit/litorder.asp?filepath=propyleneglycol/pdfs/noreg/117-

01682.pdf. The Dow Chemical Company, Midland, Michigan.

Einziger RE, SD Atkin, DE Stellrecht, and V Pasupathi. 1982. “High Temperature PostIrradiation Materials Performance Spent Pressurized Water Reactor Fuel Rods Under Dry 

Storage Conditions.” Nuclear Technology 57(1):65 – 80.
10-2

Gauld IC, SM Bowman, and JE Horwedel. 2009. ORIGEN-ARP: Automatic Rapid Processing 

for Spent Fuel Depletion, Decay, and Source Term Analysis. ORNL/TM-2005/39, Oak Ridge 

National Laboratory, Oak Ridge, Tennessee.

General Atomics. 1998. GA-4 Legal Weight Truck Spent Fuel Shipping Cask Safety Analysis 

Report for Packaging (SARP). Document No. 910469, Revision G, General Atomics. (Public

version available in ADAMS: ML070020004, ML070020006, and ML070020007.)

Gilbert ER, CE Beyer, EP Simonen, and PG Medvedev. 2002. “Update of CSFM Creep and 

Creep Rupture Models for Determining Temperature Limits for Dry Storage of Spent Fuel.”

International Congress on Advanced Nuclear Power Plants (ICAPP) ANS Embedded Topical 

Meeting, June 9-13, Hollywood, Florida.

Guyer EC and DL Brownell, eds. 1989. Handbook of Applied Thermal Design. McGraw-Hill, 

Inc., New York.

HEXCEL Composites. 1999. HexWebTM Honeycomb Attributes and Properties. Available at 

http://www.hexcel.com/Resources/DataSheets/Brochure-DataSheets/Honeycomb_Attributes_and_Properties.pdf.

Johnson AB and ER Gilbert. 1983. Technical Basis for Storage of Zircaloy-clad Spent Fuel in 

Inert Gases. PNL-4835. Battelle Pacific Northwest Laboratory, Richland, Washington.

Johnson AB, ER Gilbert, and RJ Guenther. 1983. Behavior of Spent Nuclear Fuel and 

Components in Dry Interim Storage. PNL-4189. Pacific Northwest Laboratory, Richland, 

Washington.

Juvinall RC and KM Marshek. 1991. Fundamentals of Machine Component Design, 2nd

Edition. John Wiley & Sons, New York.

Kreith F and MS Bohn. 2001. Principles of Heat Transfer. 6th ed. Brooks/Cole, Forest Grove, 

California.

Livermore Software Technology Company. 2007. LS-DYNA Keyword User’s Manual, Version 

971. Livermore Software Technology Company, Troy, Michigan.

McGrattan KB, B Klein, S Hostikka, and J Floyd. 2008. Fire Dynamics Simulator (Version 5), 

User’s Guide. Special Publication 1019-5, National Institute of Standards and Technology, U.S. 

Department of Commerce, Washington, D.C.

Michener TE, DR Rector, JM Cuta, RE Dodge, and CW Enderlin. 1995. COBRA-SFS: A 

Thermal-Hydraulic Code for Spent Fuel Storage and Transportation Casks. PNL-10782. 

Pacific Northwest Laboratory, Richland, Washington.
10-3

NFPA - National Fire Protection Association. 2008. NFPA 502: Standard for Road Tunnels, 

Bridges, and Other Limited Access Highways. National Fire Protection Association. Nitronic 50 

Product Data Bulletin1

. High Performance Alloys, 444 Wilson St. Tipton, IN 46072. Available at 

http://www.hpalloy.com/Alloys/brochures/Nitronic50book.pdf.

NRC – U.S. Nuclear Regulatory Commission. 2008. Analysis of Structural Material Exposed to 

a Severe Fire Environment – Final Letter Report, Contract NRC-02-07-006, prepared by 

Southwest Research Institute and Center for Nuclear Waste Regulatory Analyses, San Antonio, 

Texas, published by U.S. Nuclear Regulatory Commission, Washington, D.C.

NUREG-0170. 1977. Final Environmental Statement on the Transportation of Radioactive 

Material by Air and Other Modes. U.S. Nuclear Regulatory Commission, Washington, D.C. 

NUREG-1617. 2000. Standard Review Plan for Transportation Packages for Spent Nuclear 

Fuel: Final Report. U.S. Nuclear Regulatory Commission, Washington, D.C.

NUREG/CR-0344. 1980. Embrittlement Criteria for Zircaloy Cladding Applicable to Accident 

Situations in Light-Water Reactors: Summary Report. U.S. Nuclear Regulatory Commission, 

Washington, D.C.

NUREG-0630. 1980. Cladding Swelling and Rupture Models for LOCA Analysis. U.S. Nuclear 

Regulatory Commission, Washington, D.C.

NUREG/CR-0772. 1980. Fission Product Release from Highly Irradiated LWR Fuel. U.S. 

Nuclear Regulatory Commission, Washington, D.C.

NUREG/CR-4829. 1987. Shipping Container Response to Severe Highway and Railway 

Accident Conditions. U.S. Nuclear Regulatory Commission, Washington, D.C.

NUREG/CR-6487. 1996. Containment Analysis for Type B Packages Used to Transport 

Various Contents. BL Anderson, RW Carlson, and LE Fischer. U.S. Nuclear Regulatory 

Commission, Washington, D.C.

NUREG/CR-6672. 2000. Re-examination of Spent Fuel Shipment Risk Estimates. U.S. 

Nuclear Regulatory Commission, Washington, D.C. 

NUREG/CR-6886, Rev. 2. 2009. Spent Fuel Transportation Package Response to the 

Baltimore Tunnel Fire Scenario. U.S. Nuclear Regulatory Commission, Washington, D.C. 

NUREG/CR-6894, Rev. 1. 2007. Spent Fuel Transportation Package Response to the 

Caldecott Tunnel Fire Scenario. U.S. Nuclear Regulatory Commission, Washington, D.C.

NUREG/CR-7022, Vol. 1. 2011. FRAPCON-3.4: A Computer Code for the Calculation of 

Steady-State Thermal-Mechanical Behavior of Oxide Fuel Rods for High Burnup. U.S. Nuclear 

Regulatory Commission, Washington, D.C.

 1 This bulletin has no specific publication date documented. The pdf file was downloaded March 26, 

2014, and verified to be the same document as used in the work performed here.
10-4

NUREG/CR-7023, Vol. 1. 2011. FRAPTRAN 1.4: A Computer Code for the Transient Analysis 

of Oxide Fuel Rods. U.S. Nuclear Regulatory Commission, Washington, D.C.

NUREG/CR-7115. 2012. Performance of Metal and Polymeric O-Ring Seals in BeyondDesign-Basis Temperature Excursions. U.S. Nuclear Regulatory Commission, Washington, 

D.C.

Oberg E, FD Jones, HL Horton, and HH Ryffel. 2004. Machinery’s Handbook, 27th Edition, and 

Guide to Machinery’s Handbook. Industrial Press. On-line version at

http://app.knovel.com/hotlink/toc/id:kpMHEGMH0B/machinerys-handbook-27th.

Sanders TL, KD Seager, YR Rashid, PR Barrett, AP Malinauskas, RE Einziger, H Jordan, 

TA Duffey, SH Sutherland, and PC Reardon. 1992. A Method for Determining the Spent-Fuel 

Contribution to Transport Cask Containment Requirements, Appendix III, Spent Fuel Response 

to Transport Environments. SAND2000-0324. Sandia National Laboratory, Albuquerque, New 

Mexico. Accessed on http://pbadupws.nrc.gov/docs/ML0427/ML042710347.pdf.

Shajaefard ML and K. Goudarzi. 2008. “The Numerical Estimation of Thermal Contact 

Resistance in Contacting Surfaces.” American Journal of Applied Sciences 5(11):1566-1571.

Simonen EP and ER Gilbert. 1988. DATING – A Computer Code for Determining Allowable 

Temperatures for Dry Storage of Spent Fuel in Inert and Nitrogen Gases. PNL-6639, UC-85, 

Pacific Northwest Laboratory, Richland, Washington.

Society of Fire Protection Engineers. 2002. SFPE Handbook of Fire Protection Engineering.

2nd ed., National Fire Protection Association, Quincy, Massachusetts. (Specifically, the chapter 

Fire Hazard Calculations for Large, Open Hydrocarbon Fires, by Beyler, in SFPE Handbook of 

Fire Protection Engineering, 2nd ed.)

Society of Fire Protection Engineers. 2008. SFPE Handbook of Fire Protection Engineering.

4th ed. National Fire Protection Association, Quincy, Massachusetts. (Specifically, the chapter 

Liquid Fuel Fires, by DT Gottuk and DA White, in SFPE Handbook of Fire Protection 

Engineering, 4th ed.)

Wilmot EL. 1981. Transportation Accident Scenarios for Commercial Spent Fuel. SAND80-

2124. Sandia National Laboratory, Albuquerque, New Mexico.

Wiss, Janney, Elstner Associates, Inc. 2007. Route 580/880 Fire and Collapse: Evaluation of 

Concrete. Prepared by Wiss, Janney, Elstner Associates, Inc., for the Office of Rigid Pavement 

Materials and Structural Concrete, Caltrans, 4 May 2007. WJE No. 2005.2002.3.3.
MATERIAL PROPERTIES FOR COBRA-SFS MODEL

OF GA4 PACKAGE

A-1

APPENDIX A

MATERIAL PROPERTIES FOR COBRA-SFS MODEL OF GA4 

PACKAGE

Table A.1. Internal Fill Gas—Helium at Atmospheric Pressure

Temperature

(°F)

Enthalpy

(Btu/lbm)

Thermal 

Conductivity

(Btu/hr-ft-°F)

Specific Heat

(Btu/lbm-°F)

Specific 

Volume

(ft3/lbm)

Viscosity

(lbm/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

(°F)

Enthalpy

(Btu/lbm)

Thermal 

Conductivity

(Btu/hr-ft-°F)

Specific Heat

(Btu/lbm-°F)

Specific 

Volume

(ft3/lbm)

Viscosity

(lbm/hr-ft)

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 0.2586 64.214 0.1242

4000 1522 0.0671 0.4524 116.428 0.1242
A-2

Table A.3. Material Properties

Specific Heat

(Btu/lbm-°F)

Density

(lbm/ft3)

Thermal 

Conductivity

(Btu/hr-ft-°F) Emissivity Description Source

0. 11 492.5 see Eq. (A-1) see 

Table A.4

SA-240, Type XM19 stainless steel, 

for basket plates, 

inner liner, package 

body, and neutron 

shield tank outer 

shell

Density and specific heat 

from GA-4 SAR (General 

Atomics1998); thermal 

conductivity from ATI 50TM 

Technical Data Sheet (see 

below)

0.065 1185.4 14.8 0.5 Depleted uranium 

for gamma shielding

Specific heat from Table 3.2-1 

of GA-4 SAR; density from 

SAR Section 2.3, p. 2.3-1;

Thermal conductivity from 

W21 SAR (see Appendix B)

0.29 151 15.0 0.8 Boron carbide rods 

within basket plates

Table 3.2-1 of GA-4 SAR, p. 

3.2-2

0.787 61.72

kNsliq. = 0.186

keff = 5.92 N/A

60% propylene

glycol and water 

mixture (neutron 

shield)

Table 3.2-2 of GA-4 SAR 

(selected value at 194°F), and 

correlation for keff of liquid 

(see Eq. [A-3])

0.210 150 1.0

0.63 (prefire) Concrete roadways 

and side barrier

From material exposure 

analysis report (NRC 2008) 0.90 (fire 

and postfire)

Emissivity values for thermal radiation exchange were obtained from Table 3.2-3 of the GA-4 

SAR (General Atomics 1998). However, the emissivity of package surfaces exposed to the fire 

was conservatively represented with a value of 0.9, rather than the “0.8 or 0.85” listed in the 

SAR. Table A.4 summarizes the emissivity values used for the XM-19 stainless steel 

components during the various phases of the fire scenario.

Table A.4. Emissivity Values for XM-19 Stainless Steel Components

Emissivity Component Transient Conditions

0.20

steel inner liner 

basket plates

package body inner surface

pre-fire steady state, fire, and post-fire cooldown

0.20 package body outer surface 

NS tank shell inner surface pre-fire steady state

0.15 NS tank shell outer surface pre-fire steady state

0.9

package body outer surface

NS tank shell inner surface

NS tank shell outer surface

fire and post-fire cooldown

Temperature-dependent thermal conductivity (in units Btu/hr-ft2

-°R) for XM-19 stainless steel 

was evaluated in the COBRA-SFS model using a linear regression fit to ATI 50 thermal 

conductivity data (see Appendix B for the material data sheet). The relationship from this fit is
A-3

k XM −19 = a0 + a1T (A-1)

where

a0 = 5.4446

a1 = 0.0047

T = material temperature (°R)

The relationship in Eq. (A-1) is a polynomial curve fit to the same data used to derive the linear 

equation presented in the GA-4 SAR, which has the form

ks = a0 + a1T (A-2)

where 

a0 = 3.6 (empirical coefficient)

a1 = 0.00532 (empirical coefficient)

T = material temperature (°R)

These two equations give essentially identical results for temperatures below about 1000°F 

(538°C), but Equation (A-1) is more conservative by 15-20% at the high temperatures 

encountered in the MacArthur Maze fire scenario.

The formula for the effective conductivity used to model natural convection in the liquid neutron 

shield is documented in the GA-4 SAR as

( ) 2 Pr

Pr

1

0 .

a

NSliq D

eff

a

a k Gr

k

+ = (A-3)

where

a0 = 0.135 (empirical coefficient)

a1 = 1.36 (empirical coefficient)

a2 = 0.278 (empirical coefficient)

kNSliq. = thermal conductivity of NS liquid (propylene glycol/water mixture)

Pr = Prandtl Number

GrD = Grashoff number, using thickness of neutron shield tank as characteristic 

length
A-4

© 2009 ATI Allegheny Ludlum

Figure A.1. Source for Thermal Conductivity of XM-19 Stainless Steel
A-5

Figure A.2. Source for Thermal Conductivity of Depleted Uranium

MATERIAL PROPERTIES FOR ANSYS MODEL

OF GA4 PACKAGE

B-1

APPENDIX B

MATERIAL PROPERTIES FOR ANSYS MODEL OF GA4

PACKAGE

Table B.1. ASME SA-240 Grade XM-19

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific Heat

(Btu/lbm-°F) Description

50 0.65333

0.2850 0.1150

Used for FSS liner, package 

body, ILSS, bottom plate, outer 

shell, trunnions, closure

100 0.67333

300 0.75167

500 0.83000

700 0.90833

900 0.98667

1100 1.0650

1300 1.1433

1500 1.2217

1700 1.3000

1900 1.3783

2100 1.4567

Table B.2. FSS Inner Frame (XM-19, Helium, and Boron Carbide Composite)

Temperature

(°F)

Thermal Conductivity

(Btu/hr-in-°F) Density

(lbm/in3)

Specific Heat

Kxx Kyy Kzz (Btu/lbm-°F) Description

0 0.32494 0.40625 0.33828 0.19272 0.19893

Calculated composite 

properties of XM-19 steel, 

helium, and boron carbide 

pellets

100 0.33972 0.42913 0.35146 0.19272 0.19893

200 0.36773 0.46354 0.38037 0.19272 0.19893

300 0.39439 0.49628 0.40806 0.19272 0.19893

400 0.42001 0.52716 0.43482 0.19272 0.19893

500 0.44263 0.55498 0.45892 0.19272 0.19893

600 0.46443 0.58131 0.48228 0.19272 0.19893

700 0.48302 0.60502 0.50280 0.19272 0.20954

800 0.50101 0.62760 0.52280 0.19272 0.22016

900 0.51884 0.65034 0.54269 0.19272 0.22494

1000 0.53611 0.67167 0.56210 0.19272 0.22971

1100 0.55338 0.69302 0.58153 0.19272 0.23821
B-2

Table B.3. Homogeneous Fuel Region for Westinghouse 14x14 OFA

Temperature

(°F)

Thermal Conductivity

(Btu/hr-in-°F) Density

(lbm/in3)

Specific 

Heat

Kxx Kyy Kzz (Btu/lbm-°F) Description

0 - - 0.05923

0.1446 0.0747

Used for active fuel 

assembly region (WE 

14x14)

75 0.01688 0.01688 -

100 0.01815 0.01815 0.05923

150 0.02069 0.02069 -

200 0.02323 0.02323 0.05923

250 0.02576 0.02576 -

300 0.02865 0.02865 0.06163

350 0.03173 0.03173 -

400 0.03498 0.03498 0.06436

450 0.03848 0.03848 -

500 0.04220 0.04220 0.06706

550 0.04628 0.04628 -

600 0.05061 0.05061 0.06998

650 0.05525 0.05525 -

675 0.05768 0.05768 -

700 0.06011 0.06011 0.07344

725 0.06266 0.06266 -

750 0.06545 0.06545 -

800 - - 0.07689

900 - - 0.08033

1000 - - 0.08143

Table B.4. Helium

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific Heat

(Btu/lbm-°F) Description

0 0.00650 0.6900 E-5 1.240

Used for gaps within package 

assembly

200 0.00808 0.4810 E-5

400 0.00958 0.3690 E-5

600 0.01075 0.2990 E-5

800 0.01150 0.2520 E-5

1400 0.01370 0.1710 E-5
B-3

Table B.5. Air

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific Heat

(Btu/lbm-°F) Description

0 0.001092 0.4994 E-4 0.2396

Used for trunnion air pockets, 

outer closure assembly gap, and 

the impact limiter to outer shell 

gap

32 0.001159 0.5039 E-4 0.2398

100 0.001297 0.4103 E-4 0.2400

200 0.001483 0.3484 E-4 0.2411

300 0.001661 0.3021 E-4 0.2427

400 0.001833 0.2674 E-4 0.2448

500 0.002001 0.2390 E-4 0.2473

600 0.002163 0.2164 E-4 0.2504

800 0.002469 0.1823 E-4 0.2567

1000 0.002769 0.1574 E-4 0.2631

1200 0.003060 0.1383 E-4 0.2688

1400 0.003331 0.1233 E-4 0.2740

Table B.6. ASME SA-479 S21800, Nitonic 60

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific Heat

(Btu/lbm-°F) Description

- 1.00 0.2750 0.1150 Used for trunnion sleeves

Table B.7. ASTM A-276 GR 304

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific 

Heat

(Btu/lbm-°F) Description

- 0.8333 0.2836 0.1100 Used for stiffener ring

Table B.8. Aluminum Honeycomb 220 psi

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific 

Heat

(Btu/lbm-°F) Description

0 0.22856

0.0024 0.210 Used for honeycomb Section 2 of 

impact limiters

100 0.28238

200 0.34957

300 0.40339
B-4

Table B.9. Aluminum Honeycomb 725 psi

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific 

Heat

(Btu/lbm-°F) Description

0 0.59172

0.0046 0.210 Used for honeycomb Section 3 of 

impact limiters

100 0.73086

200 0.90488

300 1.04410

Table B.10. Aluminum Honeycomb 1400 psi

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific 

Heat

(Btu/lbm-°F) Description

0 1.0322

0.0061 0.210 Used for honeycomb Section 1 of 

impact limiters

100 1.2751

200 1.5787

300 1.8216

Table B.11. Stainless Steel 304L

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific 

Heat

(Btu/lbm-°F) Description

- 0.8333 0.2836 0.110 Used for fuel spacer tube

Table B.12. ASTM A-412 Grade XM-11

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific Heat

(Btu/lbm-°F) Description

-99.4 0.52500

0.2830 0.1150 Used for impact limiter shell

203.0 0.65777

401.0 0.77777

599.0 0.87500

797.0 0.97223

1200.0 1.18750

1600.0 1.39580
B-5

Table B.13. SB-637 Alloy N07718

Temperature

(°F)

Thermal 

Conductivity

(Btu/hr-in-°F)

Density

(lbm/in3)

Specific Heat

(Btu/lbm-°F) Description

- 0.5493 0.2960 0.1040 Used for assembly bolts

Table B.14. Emissivity Values for Thermal radiation Heat Transfer

Component Material

Emissivity 

Before Fire

Emissivity 

During/After Fire

Solar 

Absorptivity

Inner Steel Surfaces stainless steel 0.35 0.35 -

Outer Cask Skin stainless steel 0.15 0.9 0.4

Outer Impact Limiter Shell steel 0.85 0.9 0.6

Depleted Uranium depleted 

uranium

0.5 0.5 -

Fuel Assembly - 0.7 0.7 -

Boron Carbide Pellets boron carbide 0.8 0.8 -

Surface Exposed to Fire - 0.9 0.9 -

Ambient Environment - 0.9 0.9 -

Inside of Cask Skin stainless steel 0.9 0.9 -

Outer Surface of Package 

Body

stainless steel 0.9 0.9 -

Depleted Uranium – See Figure A.2 in Appendix A

Neutron Shield – Effective Conductivity Calculations:

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. The 

effective conductivity of the fluid within the tank is based on heat transfer between two 

concentric cylinders. Sensitivity studies were performed on this correlation to verify that it is 

applicable to the neutron shield tank geometry, in which the inner surface of the tank has a 

square cross-section with rounded corners, rather than a circular cross-section. This correlation 

produces reasonable values of keff, and the transient conditions are generally within its 

applicable range. The correlation relates the Nusselt number to the ratio of the effective 

conductivity over the actual conductivity, and is expressed as

keff

kc

= Nu = 0.386 Pr

0.861+ Pr

 

   

  

0.25

Rac

0.25 (B-1)

where

keff = effective thermal conductivity of material in node

kc = thermal conductivity of motionless fluid in node

Nu = Nusselt number

Pr = Prandtl number

Rac = modified Rayleigh number
B-6

The modified Rayleigh number is defined as:

Rac = [ln(Do Di)]

4

L3 Di

−0.6 + Do

−0.6

[ ]

5 Ra (B-2)

where

Do = annulus outer diameter

Di = annulus inner diameter

Ra = Rayleigh number

L = (Do − Di)/2

The Rayleigh number is based on the temperature difference across the annular gap and is 

expressed as:

Ra = gβ(Ti − To )L3

να

(B-3)

where

g = acceleration of gravity

Ti = inner surface temperature

To = outer surface temperature

β = thermal expansion coefficient

α = thermal diffusivity of fluid

ν = kinematic viscosity of fluid

Using the correlations listed above, a macro was written to calculate the effective conductivity 

after each solution step within the transient model. For conditions below 276°F, the properties 

of 56% propylene glycol and water were used to calculate the effective conductivity. Once the 

maximum temperature within the tank exceeded 276°F, the properties of air were used to 

determine the thermal conductivity. 

Verification of Effective Conductivity Model for GA-4 Neutron Shield 

Configuration

The effective conductivity model described above is based on experimental data for natural 

convection mixing of fluid between horizontal concentric cylinders. The neutron shield tank of 

the GA-4 package consists of an inner surface formed by the package body, which is a square 

with rounded corners, and an outer cylindrical shell. To verify that this empirical model could be 

applied to the GA-4 package neutron shield geometry, the correlation predictions were 

compared to results from a computational fluid dynamics model.

Calculations were performed with Star-CD1

, for a 2-D “slice” model at the midplane of the 

package, using two basic configurations to model the GA-4 neutron shield tank. In one model, 

 1 STAR-CD, Version 4.14 Methodology, Computational Dynamics Ltd. 2010.
B-7

the neutron shield fluid region is represented as a solid material with thermal conductivity 

determined using the relationship for the effective thermal conductivity, as defined in Eq. (B-1) 

above. In the other model, the neutron shield fluid region is represented as a liquid, with the 

fluid properties of the propylene-glycol/water mixture reported in the GA-4 SAR [11].

The results of this evaluation are summarized in Table B.15, with comparisons of the maximum 

and minimum predicted temperatures obtained with the Star-CD model for all cases considered. 

All calculations in this evaluation were performed at NCT. The maximum temperature is the 

peak temperature in the fuel region1

, and the minimum temperature is the minimum temperature 

on the package outer shell surface. As shown by the results in Table B.15, a computation mesh 

that was appropriate for the neutron shield represented as a solid was not sufficient resolution 

for the CFD model. The number of computational elements required was approximately two 

orders of magnitude larger. 

Star-CD results for the case with the neutron shield represented as a solid material and for the 

case with the neutron shield represented as a fluid (with an appropriately refined mesh) are 

shown graphically with color thermographs in Figure B.1. Overall, this evaluation has shown 

that the effective conductivity model predicts temperatures that are results are consistent with 

the CFD model results. There is also some indication that the effective conductivity model may 

yield results that are slightly conservative.

Table B.1. Summary of STAR-CD Model Results

Case

Description

Peak Fuel Region 

Temperature, °F 

(°C)

Minimum Outer 

Shell Temperature, 

°F (°C)

Number of 

Computational 

Elements

Number of 

Fluid 

Elements

Effective conductivity 

model 302 (150) 194 (90) 3,664 0

Baseline CFD model 312 (156) 194 (90) 3,664 1232

CFD model

(2x2 refine, all) 307 (153) 194 (90) 14,596 4928

CFD model 

(4x4 refine, all) 303 (151) 192 (89) 58,384 19,712

CFD model 

(5x5 refine) 2x solids 301 (149) 189 (87) 46,936 30,800

CFD model 

(8x8 refine, all) 300 (149) 189 (87) 233,536 78,848

 1 Note that the 2-D “slice” model used in this study oversimplifies features captured in the fully 3-D 

ANSYS model used for the MacArthur Maze fire calculations. As a result, temperatures predicted for 

NCT with the fully 3-D ANSYS model differ slightly from the temperatures reported in this study with StarCD. With the ANSYS model, the peak fuel region temperature is 306°F (152°C) and the minimum outer 

shell temperature is 188°F (87°C).
B-8

 (a) (b)

Figure B.1. Mid-plane Temperature Distributions Predicted with Star-CD Model of GA-4 

Package at NCT: (a) solid material neutron shield with effective conductivity model, 

and (b) liquid neutron shield with (8x8) refined mesh




NUREG/CR

-7206

Final

Spent Fuel Transportation Package Response to the MacArthur 

Maze Fire Scenario