Spent Fuel Transportation 

Package Response to the 

MacArthur Maze Fire Scenario

Draft Report for Comment 

Office of Nuclear Material Safety and Safeguards

NUREG/CR-7206

PNNL-23701
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Spent Fuel Transportation 

Package Response to the 

MacArthur Maze Fire Scenario

Draft Report for Comment

Manuscript Completed: August 2015

Date Published: December 2015

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
COMMENTS ON DRAFT REPORT

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For any questions about the material in this report, please contact: Joseph Borowsky, Thermal 

Engineer, 301-415-7507 or by e-mail at Joseph.Borowsky@nrc.gov.

Please be aware that any comments that you submit to the NRC will be considered a public 

record and entered into the Agencywide Documents Access and Management System 

(ADAMS). Do not provide information you would not want to be publicly available.
iii

1 ABSTRACT

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

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

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

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

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

7 Nuclear Regulatory Commission has undertaken the examination of specific accidents to 

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

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

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

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

12

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

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

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

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

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

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

19

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

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

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

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

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

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

26 that the package is designed to withstand. 

27

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

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

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

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

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

33 1350-1400°F (732-760°C), and occurs approximately 3 hours after the end of the postulated fire 

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

35 most of the fire duration.

36

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

38 (592°C). Together with the temperature histories of the fuel, this suggests that there is the 

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

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

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

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

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

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

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

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

47 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

ACKNOWLEGMENTS............................................................................................................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

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

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

10.0REFERENCES ...............................................................................................................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

1 EXECUTIVE SUMMARY

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

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

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

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

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

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

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

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

10 2007) were undertaken with three different SNF package designs. Based on conservative 

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

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

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

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

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

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

17 packages to within regulatory bounds for accident conditions.

18

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

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

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

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

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

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

25 selected for this investigation, mainly because it can carry a relatively large payload for an over26 the-road transportation package, and therefore the potential consequences of package failure 

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

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

29 maximum total package decay heat load of 2.5 kW.

30

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

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

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

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

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

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

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

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

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

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

41 two images.)
xiv

1

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

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

4

5

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

7 CHP 2007, reprinted with permission.)
xv

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

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

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

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

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

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

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

8 open pool hydrocarbon fires for any possible configuration of both the pre-collapse and post9 collapse fire pools in this accident. 

10

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

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

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

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

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

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

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

18 package during the cooldown phase of the transient.

19

20 Thermal and Structural Modeling Approach and Summary of Results

21

22 Detailed thermal models of the GA-4 package were constructed for the ANSYS and COBRA23 SFS codes, for transient evaluations to determine the temperature response of the package to 

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

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

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

27

28

29 Figure S.3. GA-4 Package: Exploded View

30
xvi

1

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

3

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

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

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

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

8 effect of the roadway falling on it.

9

10

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

12

fuel rod

control rod 

thimble

fuel assembly

fuel assembly

fuel assembly

fuel assembly
xvii

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

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

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

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

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

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

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

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

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

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

11

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

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

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

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

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

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

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

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

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

21 body wall.

22
xviii

1

2

3 (a)

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

5 clarity.)

6

7 (b)

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

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

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

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

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

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

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

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

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

7 different strength-versus-temperature properties of the three metals involved, results in a time8 and-temperature dependent history of force on the bolts that raised the possibility that the 

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

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

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

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

13 bolt structure, including thread inserts.

14

15

16 (a) (b)

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

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

19

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

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

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

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

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

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

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

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

28 (108 minutes) and the very long cooldown period of approximately 400 hours, back to post-fire 

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

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

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

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

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

34 of the package in this scenario.

35
xx

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

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

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

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

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

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

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

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

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

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

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

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

13
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

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

2 during Post-fire Cooldown

3
xxii

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

2 1350-1400°F (732-760°C), and occurs approximately 3 hours after the end of the fire. Based on 

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

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

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

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

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

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

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

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

11 Gilbert et al. 2002).

12

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

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

15 1097°F (592°C) using LOCA burst strain modeling (FRAPTRAN) and 1229°F (665°C) using 

16 creep rupture modeling (FRAPCON/DATING). Applicable experimental data (NUREG/CR17 0344) yields measured rupture temperatures in the range 1205-1256°F (652-680°C). The burst 

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

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

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

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

22 fire scenario.

23

24 Potential Radiological Consequences

25

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

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

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

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

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

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

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

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

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

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

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

37 accident conditions.

38

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

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

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

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

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

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

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

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

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

48 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 1 determined for the design basis 

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

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

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

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

6 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

ACKNOWLEGMENTS

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 1.0 INTRODUCTION

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

3 (SNF) transportation packages must be designed to survive exposure to a fully engulfing fire 

4 accident lasting no less than 30 minutes with an average flame temperature of “no less than 

1475°F (800°C)” (10 CFR 71). The package 5 1 must maintain containment, shielding, and 

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

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

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

9 Statement on the Transportation of Radioactive Material by Air and Other Modes), NUREG/CR10 4829 (Shipping Container Response to Severe Highway and Railway Accident Conditions, also 

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

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

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

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

15 that included a roadway collapse.

16

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

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

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

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

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

22 1982 fire in the Caldecott Tunnel on California State Route 24 near Oakland, California

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

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

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

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

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

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

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

30 conditions.

31

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

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

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

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

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

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

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

39 580 (I-580) roadway above the fire, collapsing two adjacent spans (approximately 156 feet 

40 [47.55 m]) of the elevated roadway onto the section of freeway below. A surveillance camera 

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

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

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

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

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

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

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

 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

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

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

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

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

5 the fire after about 120 minutes. 

6

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

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

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

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

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

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

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

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

15 SNF transportation package to boundary conditions simulating temperatures predicted for the 

16 MacArthur Maze fire.

17

18

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

20 CHP 2007, reprinted with permission.)

21

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

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

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

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

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

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

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

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

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

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

1 2.0 THE MACARTHUR MAZE FIRE

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

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

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

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

6 and tanker truck exposed to the fire. 

7

8 2.1 Description of the MacArthur Maze Fire

9

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

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

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

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

and the post-collapse fire near Bent1 14 19. Key points in the fire duration are illustrated with 

15 video-capture images in Figure 2.1 through Figure 2.7.

16

17

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

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

20

 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

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

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

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

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

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

6 Figure 2.1. The east pillar of Bent 19 appears to be engulfed in flame below the level of the I7 880 roadway, indicating that fuel is spilling off the roadway through the bridge scuppers for 

8 rainwater run-off near this location. 

9

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

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

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

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

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

15 until the collapse of the overhead roadway spans. 

16

17

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

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

20

The sag in the span between Bent 19 and Bent 202 21 begins to be discernable in the video at 

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

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

 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.
2-3

1

2

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

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

5

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

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

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

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

10 into the fire, significantly affecting flame shape and distribution.
2-4

1

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

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

4

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

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

7 37.3 minutes into the fire, as illustrated in Figure 2.5. This image shows that the fire has self8 extinguished in the region between Bent 18 and the point where the upper span contacts the 

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

10 the two fallen I-580 spans.
2-5

1

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

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

4

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

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

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

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

9 +37.3 minutes. This behavior is observed on the WTP video until essentially the end of the fire. 

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

11 +107 minutes.
2-6

1

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

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

4

5

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

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

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

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

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

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

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

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

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

8 open pool fire of similar size. 

9

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

11 17 minutes, and the protracted partial collapse of the span between Bent 18 and Bent 19, 

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

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

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

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

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

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

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

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

20

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

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

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

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

25 materials exposed to the MacArthur Maze fire.

26

27 2.2 Maximum Material Temperatures in the MacArthur Maze Fire

28

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

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

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

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

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

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

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

36 location.

37

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

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

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

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

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

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

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

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

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

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

48 1050°C), with an overall peak of 1994°F (1090°C), at a location 3.3 ft (1 m) above the lower 

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

50 2012°F (1100°C) is a reasonable and conservative estimate of the flame temperature for the 
2-8

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

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

3 lower than this bounding value.

4

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

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

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

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

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

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

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

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

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

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

15 could be obtained.

16

17

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

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

20 permission.)

21
2-9

1

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

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

4 2007, reprinted with permission.)

5

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

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

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

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

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

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

12 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)

13
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

1

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

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

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

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

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

type of steel in the girders3 7 supporting the I-580 roadway). The yield strength of this material at 

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

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

10 overhead spans.

11

 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

1

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

3 data from Brockenbrough and Merritt 1999.)

4

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

1 3.0 NUMERICAL MODELING OF THE

2 MACARTHUR MAZE FIRE

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

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

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

6 The final calculations were performed with FDS version 5.4.3.

7

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

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

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

11 then refined and final calculations were performed by NIST.

12

13 3.1 FDS Model Geometry

14

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

16 59 ft high (40 m by 60 m by 18 m). In subsequent refinement of the model, the height was 

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

18 (0.25 m) on a side divide the volume into uniform sub-volumes. Figure 3.1 shows the overall 

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

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

21

22

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

24 permission.)

25
3-2

1

2 Figure 3.2. Transverse View of Model Geometry

3

4

5 Figure 3.3. Longitudinal Views of Model Geometry

6

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

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

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

10 (0.5 m), due to of the relatively coarse numerical grid used in the model to solve the governing 

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

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

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

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

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

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

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

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

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

20
3-3

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

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

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

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

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

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

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

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

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

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

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

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

13

14 3.2 FDS Model Fire

15

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

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

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

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

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

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

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

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

24

25 (a) (b)

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

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

28 reprinted with permission.)

29

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

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

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

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

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

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

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

5 change significantly with assumed fire size. 

6

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

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

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

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

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

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

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

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

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

16 2007), which is elevated approximately 6.5 m (21.3 ft) above the ground. The burning rate 

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

18 the heat release rate for this fire.

19

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

pools formed by realistic fuel spills on concrete1 21 (Society of Fire Protection Engineers 2008), 

22 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 23 /s for a deep gasoline pool fire larger than a few meters in 

24 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 25 /s, which is significantly lower 

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

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

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

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

30

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

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

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

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

35 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 36 . A value of 

37 1000 kW/m2 was chosen for the heat release rate specified in the final FDS calculations 

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

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

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

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

1000 kW/m2 42 represents a conservative estimate for a large pool that could have deep spots, but 

43 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

1 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

2

The assumed average burning rate of 0.0229 kg/m2 3 /s for the maximum estimated pool size was 

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

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

6 54 minutes at this burning rate. In addition, the pre-collapse pool fire would have consumed

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

8 (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 9 /s, 

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

11 average burning rate for the MacArthur Maze fire.

12

13 3.3 FDS Fire Model Output

14

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

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

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

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

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

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

21 concept defined in fire temperature measurement calculations to represent a perfect, non22 intrusive measurement at a specific location within the fire. An AST defines the source 

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

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

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

26

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

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

29

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

30

31 where 

32

33 q"tot = net total local heat flux

34 q"rad = local heat flux due to thermal radiation

35 q"con = local heat flux due to convection

36

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

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

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

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

41 radiation can be written as: 

42
3-6

q" (q" T )

4

rad inc −σ s = ε (3-2)

1

2 where

3

4 q"inc = incident thermal radiation heat flux

5 ε = emissivity of the surface 

6 σ = Stefan Boltzmann constant

7 Ts = local surface temperature 

8

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

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

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

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

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

14

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

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

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

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

19

4

i

i

q"inc = ∑εiFi σT (3-3)

20

21 where

22

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

24 sources

εi = emissivity of the i

th 25 source surface 

26 σ = Stefan-Boltzmann constant

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

th 27

28 source surface

Ti = local surface temperature of the i

th 29 source

30

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

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

33

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

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

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

37 determined from an empirical heat transfer correlation, such that: 

38

q" h(T T ) con = g − s (3-4)

39

40 where

41

42 h = local heat transfer coefficient 

43 Tg = gas temperature adjacent to the exposed surface

44 Ts = local surface temperature

45
3-7

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

2 therefore be expressed as

3

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

4

tot = ε inc −σ s + − (3-5)

4

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

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

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

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

9

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

10

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

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

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

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

15 model is any calculation method used to predict temperatures or stress/strain responses in an 

16 object exposed to the fire. The fire model may compute the evolving temperature of the 

17 bounding surfaces out of necessity, but it does not generally include a detailed representation of 

18 the thermal response of solid objects. Even a computational fluid dynamics model may only 

19 approximate a bounding solid as an infinitely thick slab for the purpose of estimating its surface 

20 temperature. 

21

22 If the results of the fire model are to be used to perform a more detailed heat transfer calculation 

23 of the thermal response of a solid object within or near the fire, then some sort of interface is 

24 required to transfer information at the gas-solid interface. The most obvious quantity for this 

25 purpose is the heat flux at the surface, but in practice, this leads to major computational 

26 difficulties. The net heat flux to a surface computed by the fire model is dependent on the 

27 corresponding surface temperature, which is also computed by the fire model. Depending on 

28 the model, this surface temperature might not be of the desired accuracy. In addition, it is

29 common in many popular solid phase heat transfer programs to input a prescribed thermal 

30 boundary based on external gas temperature and calculated surface temperature (as in Eq. 3-9) 

31 rather than as a prescribed heat flux. Both of these problems can be circumvented by using the 

32 adiabatic surface temperature TAST as the intermediary between the fire and structural models. 

33

34 The interface is fairly simple. At every surface point at which the fire model computes an 

35 incident thermal radiation heat flux and a corresponding gas temperature adjacent to that 

36 surface, the following implicit equation can be solved for the adiabatic surface temperature, 

37 assuming that the emissivity and convective heat transfer coefficient are effectively constant at 

38 that location. 

39

(q" T ) h(Tg, TAST ) 0 4 ε inc,FM − σ AST + FM − = (3-7)

40
3-8

1 where

2

3 q"inc,FM = incident thermal radiation heat flux computed by the fire model at the 

4 exposed surface 

5 Tg,FM = gas temperature computed by the fire model adjacent to the exposed 

6 surface

7 TAST = adiabatic surface temperature computed by the fire model

8

9 A key feature of Equation (3-7) is that the fire model does not require any assumptions to 

10 compute the incident thermal radiation heat flux. This equation merely serves as the definition 

11 of the adiabatic surface temperature, but it does not imply that the fire model calculates the heat 

12 flux in any particular way. Most importantly, it does not imply that the fire model uses a fixed 

13 heat transfer coefficient, h. The values of TAST for any location in the fire model can be stored in 

14 a file according to a user-specified time interval and length increment appropriate for the 

15 application. 

16

17 For the structural model, the heat flux to an object’s surface and its temperature due to the fire 

18 conditions computed by the fire model can be calculated by the relationship:

19

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

4

tot,SM = ε inc,FM − σ s,SM + FM − (3-8)

20

21 Subtracting Eq. (3-8) from Eq. (3-7) yields the total net heat flux to the surface of an object as:

22

q" (T T ) h(T T ) s,

4

s,

4

tot,SM σ AST − SM + AST − SM = ε (3-9)

23

24 where

25

26 q"inc,FM = incident thermal radiation heat flux computed by the structural model at the 

27 exposed surface 

28 Tg,FM = gas temperature computed by the structural model adjacent to the 

29 exposed surface

30 TAST = adiabatic surface temperature computed by the fire model

31

32 The AST is interpreted by the structural model as an effective black body radiation temperature 

33 for the purpose of computing the incident thermal radiation at an actual surface in the structural 

34 model, and as a gas temperature for the purpose of computing the convective heat flux at the 

35 given surface. The advantage of this approach is that it requires transfer of only one quantity, 

36 the AST, from a fire model to a structural model, rather than bringing over a heat flux, surface 

37 temperature, and convective heat transfer coefficient. A side benefit is that the structural model 

38 need not be reconfigured to accept a heat flux as its boundary condition. It needs only to be 

39 modified to accept a temporally and spatially varying “exposing” temperature (i.e., the AST), 

40 which it can use to calculate the heat flux based on that temperature and the surface 

41 temperature calculated in the structural model. Most models of this type are already configured 

42 to accept a time-varying “exposing” temperature curve as an external boundary condition. 

43
3-9

1 3.4 Fire Model Results for Pre-Collapse Configuration

2

3 The FDS analysis using the model described in Sections 3.1 and 3.2 was used to determine the 

4 fire behavior during the pre-collapse phase of the fire, which lasted approximately 17 minutes. 

5 This was the hottest and most intense portion of the fire, and would be expected to produce the 

6 most severe conditions for the thermal analysis of the effect of the fire on a SNF transportation 

7 package. The upper bound on the peak predicted temperatures from the FDS model during the 

8 first phase of the fire is in the neighborhood of 2012°F (1100°C). This is illustrated in Figure 3.5

9 with the predicted ASTs for a “target” object located 3.3 ft (1 m) above the roadway at a point in 

10 the fire near the final position of the tanker truck. The target is a rectangular block consisting of 

11 surfaces facing up, down, north, south, east, and west, and the temperature exposure for a 

12 given surface on this hypothetical object depends on orientation. The different curves of the plot 

13 are labeled in reference to the direction the surface faces. Figure 3.6 shows a similar plot of 

14 AST values for the same location at a higher elevation, approximately 3.3 ft (1 m) below the 

15 overhead girders. 

16

17

18 Figure 3.5. Predicted AST Values 1 Meter above the Lower Roadway Surface

19

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

1

2 Figure 3.6. Predicted AST Values 3 Meters above the Lower Roadway Surface (1 meter

3 below the Upper Roadway)

4

5

6 Based on these CFD analysis results and on material analyses of physical samples taken from 

7 the steel girders, roadway concrete, and remains of the destroyed tanker truck, (see discussion 

8 in Section 2.1), the initial portion the fire transient is conservatively bounded by a fully engulfing 

9 fire with uniform flame temperature of 2012°F (1100°C). The fire emissivity is assumed to be 

10 0.9, which is characteristic of a sooty, optically dense hydrocarbon pool fire.

11

12 3.5 Bounding Assumptions for Post-Collapse Fire

13

14 In the MacArthur Maze fire, the collapse of the overhead spans greatly reduced the size and 

15 intensity of the fire. The abrupt fall of the span between Bent 19 and Bent 20 effectively 

16 extinguished the portion of the fire extending to the right of Bent 19. The slow descent of the 

17 Bent 19 end of the second span, which occurred over approximately 20 minutes, introduced this 

18 structure into the middle of the large pool fire on the lower roadway. The I-580 span eventually 

19 made contact with the lower roadway surface along a line extending approximately the full width 

20 of the roadway, at a distance of about 5 m (15 ft) from Bent 19. The effect of this intrusion was 

21 a fairly rapid decrease in the extent of the fire in the region beneath the descending roadway 

22 span. Within approximately 20 minutes (from 17 to 37 minutes after ignition), the fire was 

23 completely extinguished in the direction of Bent 18, and the extent of the fire was reduced to the 

24 gap between the fallen roadway spans near Bent 19. An estimate of the maximum fire area 

25 during this second phase is illustrated in Figure 3.7. 

26

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

1

2 Figure 3.7. Estimated Fire Pool for Post-collapse Portion of the MacArthur Maze Fire

3

4

5 The area of this region is about one-fifth of the estimated maximum area of the first phase of the 

6 fire. CFD codes such as FDS generally do not have the capability to significantly alter the 

7 geometry of the physical structure of the fire environment or make large changes in the physical 

8 extent of the fire (other than changes caused by consuming available fuel.) The FDS analysis is 

9 not able to track the changing fire conditions during this portion of the fire, and estimates of the 

10 bounding fire temperature must rely on alternative methods. 

11

12 The intrusion of the partially fallen span into the post-collapse fire would have the effect of 

13 wicking heat out of the fire, thereby reducing the flame temperature. In addition, the reduction in 

14 physical extent of the fire is sufficient to categorize this portion of the fire as only partially 

15 engulfing, which means that a large object (on the order of the size of a small automobile) would 

16 not be completely immersed in flames. In the parlance of thermal radiation heat transfer, it 

17 would “see” the colder surroundings. Experimental work has shown (Society of Fire Protection 

Engineers 2002)1 18 that partially engulfing fires typically range from 1472°F (800°C) to 1652°F 

19 (900°C). 

20

21 The post-collapse portion of the MacArthur Maze fire, therefore, is conservatively bounded by 

22 assuming a flame temperature of 1652°F (900°C). A further level of conservatism is added if 

23 the fire is assumed to be fully engulfing for a SNF transportation package. This approach 

24 neglects the effect of heat absorbed from the fire by the structural mass of the concrete 

25 roadways, and is a conservative and bounding representation of the fire behavior in this portion 

26 of the transient.

27

 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

1 The appropriate bounding temperature for the fire during the 20-minute transition from a large 

2 pool fire at 2012°F (1100°C) to a smaller pool fire at 1652°F (900°C) is not readily apparent from 

3 the information available on the MacArthur Maze fire. As a conservative simplification, this 

4 transition interval is modeled as if the large open pool fire had persisted for the full 20 minutes 

5 required for the partial collapse of the Bent 18-19 span, out to 37 minutes of the fire duration. At 

6 37 minutes, the boundary temperature representing the fire is abruptly reduced from 2012°F 

7 (1100°C) to 1652°F (900°C), and remains at this value out to 108 minutes, to conservatively 

8 bound the end of the fire duration. These bounding fire temperatures are illustrated in 

9 Figure 3.8.

10

11

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

13

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

1 4.0 THE MACARTHUR MAZE FIRE SCENARIO

2 There are several different aspects of the MacArthur Maze fire that could expose an SNF 

3 transportation package to conditions potentially more severe than the HAC fire specified in 

4 10 CFR 71. These can be summarized as follows:

5 1. exposure of the package to the large fully engulfing fire prior to the collapse of the overhead 

6 I-580 roadway span between Bent 19 and Bent 20, which is at a higher engulfing flame 

7 temperature and longer duration than the HAC fire

8 2. subsequent exposure of the package to the relatively long duration of the fire following the 

9 collapse of the overhead spans, which is also at a higher engulfing flame temperature and 

10 significantly longer duration than the HAC fire

11 3. physical impact of a collapsing overhead span on the package 

12 4. post-fire cooldown with the package assumed to be covered by the concrete “blanket” of a 

13 collapsed overhead span.

14

15 As the above list shows, this fire scenario subjects the SNF package to both structural and 

16 thermal conditions that could adversely affect its performance. To assure that bounding 

17 conditions were considered for all aspects of this complex scenario, the thermal and structural 

18 analyses were performed separately. To conservatively bound the worst that the MacArthur 

19 Maze fire and roadway collapse could do to the SNF package, the scenario selected for 

20 analysis evaluated the most adverse thermal conditions and the most adverse structural 

21 configuration. The package was assumed to be positioned in the most adverse location for the 

22 different portions of the thermal analyses and the structural analyses, without realistic 

23 constraints on how the package could possibly relocate from one place to another during the fire 

24 scenario. 

25

26 The thermal analyses for this fire scenario, which cover items 1, 2, and 4 of the above list, are 

27 discussed in Section 4.1. The structural analyses (item 3 above) are discussed Section 4.2.

28

29 4.1 Thermal Conditions in MacArthur Maze Fire Scenario 

30

31 The thermal analyses were performed assuming that the SNF package would be fully engulfed 

32 in the fire for the full 108 minutes of the total fire duration. The effects of the physical impact of 

33 the collapsing span(s) were ignored for this portion of the analysis, for two specific reasons. 

34 First, the impact of the overhead span on the SNF package would completely smother the fire (if 

35 the package had been located beneath the first span to fall) or greatly decrease the intensity of 

36 the fire in the vicinity of the package (if it had been located beneath the second span to fall). In 

37 the first case, the fire would have lasted less than 20 minutes; in the second case, the fire would 

38 no longer be fully engulfing after approximately 25 minutes, and the concrete roadway of the 

39 fallen span would absorb a significant portion of the energy from the fire. In either case, the 

40 potential thermal load on the package would be greatly diminished. 

41

42 For the thermal analysis, the package is assumed to be in the following locations:

43

44 • the package is on the lower I-880 roadway, fully engulfed in fire for 37 minutes, exposed to a 

45 flame temperature of 2012°F (1100°C)

46
4-2

1 • after 37 minutes, the package is still on the lower I-880 roadway, fully engulfed in fire, but the 

2 flame temperature is assumed to drop to 1652°F (900°C) for the remaining 71 minutes of the 

3 smaller post-collapse fire, resulting in a total fire exposure duration of 108 minutes

4

5 • after 108 minutes of fire exposure, the package is still on the lower I-880 roadway, but is 

6 enclosed in a concrete ‘tunnel’ simulating the collapsed roadway, which is cooled only by 

7 natural convection from the exposed concrete surfaces of the upper and lower roadways.

8

9 4.2 Structural Loads in MacArthur Maze Fire Scenario

10

11 The primary structural load considered in the MacArthur Maze fire scenario is the impact of a 

12 section of the upper roadway falling onto the package. If a tractor-trailer rig carrying a legal 

13 weight truck (LWT) SNF package were to be involved in an accident of the severity of the 

14 MacArthur Maze fire scenario, the package could be subjected to additional impact loads, such 

15 as collisions with other vehicles caught in the fire or impact with road barriers or debris. 

16 Potential scenario-dependent impact loads of this nature are not considered in this study. 

17 Regardless of the path the package might have to take to reach the postulated locations within 

18 the MacArthur Maze fire scenario, the loads it would be subjected to are assumed to be minor 

19 and within the realm of regulatory accident scenarios that provide the design bases for all SNF 

20 transportation packages. 

21

22 SNF transportation packages are designed to survive and maintain containment under severe 

23 impact scenarios. The HAC drop scenario, as specified in 10 CFR 71, is a 30-ft drop onto an 

24 unyielding surface, with the package impact occurring at the most adverse orientation. The 

large mass1 25 of a typical SNF package leads to a relatively large amount of kinetic energy at 

26 impact, which would be much greater than the kinetic energy that could be imparted to the 

27 package by the falling overhead section. On that basis alone, the falling overpass section would 

28 not be expected to be a more damaging scenario than the HAC drop scenario. However, the 

29 falling overhead section impact is postulated to occur at elevated temperatures, which reduce 

30 the package material strength, compared to the assumptions for the HAC drop scenario (which 

31 is assumed to occur at normal operating conditions). To evaluate the package performance in 

32 the more challenging thermal environment of the MacArthur Maze fire scenario, a range of 

33 postulated package locations and orientations on the lower roadway beneath the falling 

34 overhead section were selected for analysis. The goal was to select a limited number of 

35 analyses that would bound potential package response in this scenario. 

36

37 A realistic location for the package to receive the maximum impact force from the collapsing 

38 overhead span would be near the edge of the large pool fire, or possibly outside the fire pool 

39 entirely. That is, a location that would result in maximum impact loading and post-fire blanketing 

40 by the fallen overhead roadway would be a location likely to receive minimum fire exposure. 

41 Conversely, if the package were positioned to receive maximum fire exposure (i.e., fully 

42 engulfed for both the pre-collapse and post-collapse fire conditions) it would have to be located 

43 near the middle of the area encompassed by the smaller post-collapse fire pool (see 

44 Figure 3.7), where it could not be struck at all by either of the two collapsed spans.

45

46 As a bounding assumption, the peak temperatures predicted in the thermal analysis for the fully 

47 engulfing 2012°F (1100°C) fire conditions (see Section 4.1 above) were imposed on the 

48 package in the structural analysis. The package was positioned at a location where it would 

49 receive the maximum force of impact from the collapse of the I-580 overhead span between 

 1 Total weight of the GA-4 package when fully loaded is 55,000 lb (27.5 T).
4-3

1 Bent 19 and Bent 20. As a further conservatism, the package was positioned on the lower 

2 roadway, rather than on the truck conveyance transporting it. There is no credible means in this 

3 accident for the package to have become separated from its conveyance, but the I-880 roadway 

4 provides the stiffest possible platform for the package when it receives the impact force of the 

5 collapsing overhead span. If the package were on the conveyance, much of the force of the 

6 impact would be absorbed by the compliance of the overall system. In addition, the impact 

7 forces would be smaller than those obtained with the package on the roadway, since the free8 fall distance of the collapsing span before impact with the package on the conveyance would be 

9 reduced by about 30%.

10

11 The package is assumed to be at the maximum temperature obtained from the thermal analysis 

12 after 37 minutes of exposure to the fire. The temperature of the descending roadway span (in 

13 particular, the temperature of the steel girders) is estimated from a separate thermal analysis of 

14 the upper roadway response to the fire, and corresponds to the conditions at 18 minutes into the 

15 fire. This disparity between the time-stamp for the package and the roadway girders is to obtain 

16 conservatively bounding values for each component. For the package, the hotter it is at the 

17 time of impact, the more vulnerable the steel and other material would be to deformation or 

18 damage. It is therefore assumed that the package has the longest possible time to heat up (i.e., 

19 37 minutes) before the overhead span falls on it. For the roadway girders, the cooler they are at 

20 the time of impact, the stiffer they would be, and the more damage they could potentially do to 

21 the package. It is therefore assumed that the falling roadway has the shortest possible time to 

22 heat up (i.e., 18 minutes) before falling on the package. This timing difference could not occur 

23 in reality, but as a modeling assumption, it yields a conservative bound for the possible range of 

24 behavior for both components.

25

26 The orientation of the SNF package with respect to the falling roadway determines how many of 

27 the girders actually strike the package, where they strike, and at what angle, and therefore has 

28 a significant effect on the potential consequences of the impact. The spacing of the girders 

29 relative to the length of the package is such that no more than two girders could strike the 

30 package in any given configuration. Possible orientations include:

31

32 • a single girder striking along the full axial length of the package

33 • a single girder striking across the package, perpendicular to the long axis

34 • two girders striking across the package at an angle, with one impact near the package lid.

35

36 It is not obvious which impact orientation could do the most damage to the package, and the 

37 analysis considers the full range of possibilities, including impacts on the package lid flange 

38 region. The analysis also evaluated the effect of a girder striking directly onto one of the 

39 package lifting trunnions, which would impart an extremely localized load to the package body. 

40 This requires the highly improbable configuration of the package standing on the roadway on 

41 one side pair of its lifting trunnions, to obtain a geometry in which such an impact could actually 

42 occur. 

43

44 The assumptions defining components of the package model can affect the magnitude of the 

45 loads that are imparted to the package by the accident scenario. To maximize the impact load

46 on the stainless steel body of the package, the effect of impact limiters was neglected in all 

47 cases, even though the impact limiters are not expected to detach from the package during the 

48 accident. Similarly, the thin outer shell and liquid contents of the neutron shield are neglected, 

49 on the assumption that the liquid would be gone by the time the overpass collapsed and that the 

50 thin steel plates of the neutron shield would absorb only a negligible amount of impact energy 

51 through deformation. In addition to conservatively neglecting energy absorption mechanisms, 
4-4

1 modeling choices were made to simplify the impact model and reduce the computational cost as 

2 much as reasonably possible. These modeling choices result in the falling overpass structure 

3 making direct contact with the stainless steel body of the package or its trunnions, rather than 

4 being first intercepted by the impact limiters and neutron shield tank. Detailed descriptions of 

5 the overpass and package structural models are presented in Section 5.4.1 and Section 5.4.2,

6 respectively. Section 6.4 describes the assumptions and analysis method in more detail.

7

8 In addition to the overpass and package impact modeling, structural evaluations were

9 performed for key package bolts under thermal expansion loading. The differences in thermal 

10 expansion coefficients of the package structural materials and the fastener bolts lead to 

11 increased bolt tension, even under uniform temperature loading. Temperature results from the 

12 thermal models were used to develop temperature load histories for the bolts and surrounding 

13 material. These bolt evaluations are assumed to be independent of the impact events. In all 

14 credible load cases, the key temperature states occur much later in the transient than the 

15 overpass collapse at 37 minutes. Section 5.4.3 discusses the bolt thermal expansion models in 

16 more detail.
5-1

1 5.0 ANALYTICAL MODELS FOR THE MACARTHUR MAZE 

2 FIRE SCENARIO

3 This section describes the analytical models developed to investigate the potential effects of the 

4 MacArthur Maze fire scenario on a typical over-the-road spent fuel transportation package

5 design. This analysis evaluates the transient thermal response of the package and roadway 

6 structure from the initial steady-state conditions through the fire scenario, and many hours into 

7 the post-fire cooldown. The models appropriately capture the thermal inertia of the SNF 

8 package and the transient temperature response of the system. 

9

10 Because of the severity and complexity of this fire scenario, multiple thermal and structural 

11 models were developed. This included two independent thermal modeling approaches using 

12 the finite element analysis (FEA) code ANSYS and the finite difference thermal-hydraulics code 

13 COBRA-SFS and structural models using ANSYS and LS-DYNA. The ANSYS thermal model 

14 represented the complete package, including the impact limiters. The COBRA-SFS thermal 

15 model was developed to evaluate the potential effects of loss of the impact limiters, and to 

16 provide a detailed, best-estimate evaluation of the thermal response of the fuel rod cladding to 

17 this fire scenario. The structural models were used to evaluate the potential effects of the 

18 impact of the falling overhead span on the package, and assumed a range of possible 

19 orientations of the package on the lower roadway. Additional detailed models of the bolted lid 

20 and flange were also developed, to evaluate potential effects on the package containment 

21 boundary as a consequence of the thermal and structural loading imposed by this fire scenario.

22

23 The basic design of the package selected for this analysis is described in Section 5.1. The 

24 models representing this package for analysis with ANSYS and COBRA-SFS are presented in 

25 Sections 5.2 and 5.3, respectively. Structural models developed using LS-DYNA and ANSYS 

26 are presented in Section 5.4.

27

28 5.1 GA-4 Legal Weight Truck Spent Fuel Shipping Package

29

30 The General Atomics GA-4 LWT transportation package was selected for this investigation to 

31 evaluate the potential effects of an accident of the magnitude and severity of the MacArthur 

32 Maze fire on an NRC-certified SNF transportation package. This package can carry a relatively 

33 large payload for an over-the-road transportation package, and therefore the potential 

34 consequences of package failure could be more severe than for packages with smaller payload 

35 capacities. The GA-4 package is designed to transport up to four intact pressurized water 

36 reactor (PWR) spent fuel assemblies with a maximum decay heat load of 2105.4 Btu/hr 

37 (0.617 kW) per assembly, for a total package decay heat load of 8423 Btu/hr (2.468 kW). 

38

The GA-4 can carry zircaloy-clad UO2 fuel with maximum initial enrichment of 3.15% 39 235U, in 

40 14x14 assemblies with maximum average burnup of 35 GWd/MTU (minimum cooling time of 

41 10 years), or 15x15 assemblies with maximum average burnup of 45 GWd/MTU (minimum 

42 cooling time of 15 years). This package is not licensed to carry high burnup fuel (i.e., fuel with 

43 average burnup greater than 45 GWd/MTU). There are packages permitted to carry high 

44
5-2

1 burnup fuel pins, but their contents are less than a complete fuel assembly. In addition, 

2 transportation of high burnup fuel (>45 GWd/MTU) by road is currently evaluated on a case-bycase basis, pending development of general guidance1 3 . 

4

5 For the purpose of this analysis, the package was assumed to contain four WE 14x14 PWR 

6 spent nuclear fuel assemblies at the maximum decay heat load. This is the limiting design basis 

7 configuration for thermal analysis of the package. Figure 5.1 shows an exploded view of the 

8 package, illustrating the main design features. The payload capacity is 6648 lb (3015 kg), and 

9 the fully loaded package weighs approximately 55,000 lb (24,948 kg). The package 

10 containment boundary is provided by the following structures:

11

12 • stainless steel package body wall 

13 • stainless steel bottom plate 

14 • stainless steel package closure lid secured by Inconel fasteners 

15 • dual O-ring seals for the closure lid, gas sample port, and drain valve. 

16

17 The stainless steel package body encloses the gamma shield, which consists of an inner shell 

18 of depleted uranium. Neutron shielding is provided by a stainless steel neutron shield tank 

19 external to the package body, containing a water/propylene glycol mixture. Aluminum 

20 honeycomb impact limiters, completely enclosed in a thin stainless steel outer skin and inner 

21 housing, are attached to each end of the package. Configuration details, including design 

22 drawings, are provided in the safety analysis report (SAR) for this transport package (General 

23 Atomics 1998).

24

25

26 Figure 5.1. GA-4 Package: Exploded View

27

28

 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

1 5.2 ANSYS Model of GA-4 Package

2

3 A detailed three-dimensional representation of the GA-4 package was constructed using 

ANSYS® 4 (ANSYS 2003). To simulate the effects of the different segments of the MacArthur 

5 Maze fire scenario (as described in Section 4.0), the ANSYS model also includes elements 

6 representing a segment of the lower roadway beneath the package and the collapsed upper 

7 roadway, which is assumed to cover the package at 37 minutes into the fire scenario. However, 

8 the model conservatively neglects the thermal effect of direct contact between the package and 

9 the roadway. During the fire, the package is treated as fully engulfed in flame. The model 

10 considers forced convection due to the hot fire gases flowing past the package, and thermal 

11 radiation to the optically dense fire environment. During the post-fire cooldown, the model 

12 includes thermal radiation from the outer surfaces of the package to the roadway surfaces, and 

13 conduction-only heat transfer to the surrounding air. The roadway surfaces were included to 

14 evaluate the potential effect of the collapsed roadway on the rate of cooldown of the package 

15 after the end of the fire. (A detailed description of the thermal boundary conditions for this multi16 step scenario is presented in Section 6.3.)

17

18 Section 5.2.1 describes the detailed ANSYS model of the GA-4 package. Section 5.2.2

19 describes the approach used to represent the package buried under the collapsed roadway in 

20 the post-fire analysis. Section 5.2.3 presents the material properties used to represent the 

21 different elements of the package in the fire and post-fire cooldown transients.

22

23 5.2.1 GA-4 Package Representation

24

25 The package is assumed to be oriented horizontally throughout the fire scenario, including the 

26 actual fire duration, for maximum heat input into the package from the fire. The horizontal 

27 orientation of the package beneath the fallen roadway also represents the most adverse 

28 conditions for heat removal from the package in the post-collapse fire environment. The 

29 conveyance carrying the package is omitted from the model as a conservative representation 

30 for both the thermal and structural modeling of this fire scenario. In the thermal model, the fire 

31 is treated as fully engulfing, such that the package is subjected to a uniform bounding flame 

32 temperature in all directions. In effect, the package is treated as suspended in the fire, and 

33 thermal effects of contact with the surface of the road (e.g., heat conduction losses and potential 

34 thermal shielding of portions of the package) are neglected. Including the conveyance in a 

35 realistic manner would have the effect of partially shielding the package from the fire burning on 

36 the roadway below. These assumptions constitute a significant conservatism in the overall 

37 modeling approach, since the conveyance and the roadway beneath the package could provide 

38 substantial limitations on the rate of heat deposition to the package in this fire scenario.

39

40 The model geometry was developed from engineering drawings provided in the SAR for the GA41 4 package (General Atomics 1998). Table 5.1 summarizes the ANSYS model element types 

42 used for the various components of the package and surrounding roadway. The structure of the 

43 package is represented in fine detail, including the lifting trunnions and impact limiters. 

44 Convection and thermal radiation heat transfer is represented for specific interior and exterior 

45 surfaces, including thermal radiation between the outer surfaces of the package and the 

46 external environment. During the fully engulfing fire scenario, the package sees only the 

47 bounding fire temperature. During the post-fire cooldown, thermal radiation exchange is 

48 calculated between the package and the enclosing roadway concrete. Surface elements were 

49
5-4

1 also generated along the exterior of the package to account for solar insolation loads to 

2 calculate the normal conditions of transport, which defines the initial temperature distribution for 

3 the package.

4

5 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

6

7 A cross-sectional view of the ANSYS model is shown in Figure 5.2, with the major components 

8 of the GA-4 package indicated. All components illustrated in Figure 5.2 were modeled using 

9 brick elements. The square blocks shown in red are homogeneous regions representing the 

10 four fuel assemblies within the package. The fuel assemblies are contained within the cruciform 

11 stainless steel fuel support structure (FSS) and FSS liner. The helium gas in the gaps between 

12 the homogenized fuel assembly regions and the FSS plates was explicitly modeled with solid 

13 elements. The model includes a composite representation of the layers of the cruciform inner 

14 frame of the FSS, which consists of thin sheets of stainless steel enclosing boron carbide rods. 

15 The thin steel of the FSS liner is represented with a single layer of nodes (illustrated in light blue 

16 in the diagram in Figure 5.2).

17

18 The GA-4 gamma shield (represented by three layers of elements illustrated in multiple colors in 

19 Figure 5.2), consisting of a rectangular tube of depleted uranium (DU), encloses the FSS liner. 

20 The DU gamma shield is in non-loadbearing contact with the square cross-section of the FSS

21 liner, and has rounded outer corners, in order to fit within the cross-sectional geometry of the 

22 steel package body. The rectangular stainless steel package body forms the inner surface of 

23 the liquid neutron shield (NS) tank. The liquid neutron shield tank contains a 56% propylene 

24 glycol/water mixture that is modeled as a solid material using the elements shown between the 

25 steel package body and the outer wall of the tank. The outer wall of the NS tank is a thin 

26 cylindrical stainless steel shell, and is represented in the model as a single layer of elements, as 

27 shown in Figure 5.2. This layer constitutes the outer surface of the package assembly. 
5-5

1

Figure 5.2. Cross Section of ANSYS 2 ® Model of GA-4 Package Near Midplane

3

4

5 The diagram in Figure 5.2 shows a cross-section of the package near the center of the axial 

6 length of the cask cavity. In this region, there is only liquid in the region between the cask body 

7 and the NS tank outer shell. At either end of the package, in the regions covered by the impact 

8 limiters, the NS tank is structurally supported by 36 radially distributed stainless steel ribs

9 designated as the impact limiter support structure (ILSS). These ribs extend radially from the 

10 thick steel shell of the package body to the thin outer stainless steel shell, and provide additional

11 pathways for conduction heat transfer from the cask body to the NS tank outer shell, in addition 

12 to structural support. This region is explicitly modeled in detail in the ANSYS model, but for 

13 clarity is omitted from the diagram in Figure 5.2. 

14

15 A slice through the long axis of the model is illustrated in the diagram in Figure 5.3, and shows

16 the modeling of the ends of the package, including the impact limiters, which consist of an 

17 internal aluminum honeycomb structure enclosed within a stainless steel skin. The stainless 

18 steel shell of each impact limiter was modeled with shell elements. All other components were 

19 modeled using brick elements. A detailed representation of the model in the region of the top 

20 impact limiter and package closure is illustrated in Figure 5.4. This diagram shows the impact 

21 limiter stainless steel skin and a thin air gap between the impact limiter and the external surface 

22 of the package. This gap, which conservatively accounts for the tolerance of the fit of the 

23 impact limiter onto the package, was represented in the model geometry using SOLID70 brick 

24 elements.
5-6

1

2 Figure 5.3. GA-4 Package Geometry, Including Impact Limiters

3

4

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

6

7 For this analysis, the air gap between the closure lid and the impact limiter steel liner was 

8 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

1 or warping of the impact limiter in response to the fire conditions could potentially reduce or 

2 eliminate this gap. During the fire, this assumption would tend to slow the rate of heat input to 

3 the package through the impact limiters, but because very little heat from the fire can enter the 

4 package through the highly insulating material of the impact limiters, this assumption would be 

5 expected to have a negligible effect on the thermal response of the package. In the post-fire 

6 cooldown, however, this assumption would tend to slow the rate of heat removal from the 

7 package, by increasing the insulating effect of the impact limiters. 

8

9 The thermal inertia of an SNF package can result in significantly higher temperatures being 

10 reached on some components in the post-fire cooldown, compared to temperatures reached 

11 during the fire, particularly for temperatures in nominally cooler regions of the package. It was 

12 therefore deemed more important to capture the effect of retaining the air gap throughout the 

13 fire scenario, particularly since heat transfer in the package end regions would not be expected 

14 to affect the peak component temperatures during the fire, which occur near the package 

15 midplane, due to direct heat input from the fire.

16

17 The lower end of the package consists of a thick stainless steel bottom plate welded to the steel 

18 inner and outer walls of the package. The upper end of the package is sealed with a stainless 

19 steel closure assembly that attaches to a stainless steel flange on the steel body wall. 

20 Figure 5.5 shows the detailed representation of the closure assembly developed for this model. 

21 Helium-filled gaps between the closure assembly and the FSS, and between the stainless steel 

22 flange and the closure assembly, were included in the model geometry. These gaps were 

23 represented with solid brick elements.

24

25

26 Figure 5.5. GA-4 Package Geometry Model: Closure Assembly Details

27

28 Sections 6.2 and 6.3 provide detailed discussions of the modeling assumptions and boundary 

29 conditions for the fire analysis. The detailed representation of the package internals was 

30 designed to capture all three possible modes of heat transfer (i.e., conduction, convection, and 

31 thermal radiation) between all of the components of the model. Conduction is handled 

32 inherently in ANSYS by the elements and corresponding material properties representing each 

33 component, but convective and radiative mechanisms must be carefully implemented to 
5-8

1 properly capture the physical behavior of the system. The representation of the fuel assemblies 

2 is particularly important in appropriately modeling the thermal response of the fuel rods and 

3 predicting the peak cladding temperature. Heat transfer within the fuel assemblies is primarily 

4 by conduction and thermal radiation, with convection only a relatively minor contributor.

5

6 The fuel assemblies were modeled as homogeneous regions with an effective radial 

7 conductivity determined using an effective conductivity model (Bahney and Lotz 1996) that is 

8 widely used in the nuclear industry in safety analysis for SNF packages. In this model, the 

9 combined effect of thermal radiation and conduction is characterized using an effective 

10 conductivity that is a function of assembly geometry and decay heat. The application of the fuel 

11 effective conductivity model developed for this analysis introduces a modification to more 

12 accurately account for the temperature gradient between the outermost row of rods in the 

13 assembly and the enclosing wall. This is accomplished by including a helium gap between the 

14 homogenized material region representing the fuel assembly and the wall of the enclosing

15 basket (in this case the FSS cruciform and liner, as shown in the diagram in Figure 5.2), rather 

16 than extending the homogenous region to the wall, as is the approach normally used in the 

17 effective conductivity model. An additional feature of this modified representation is that it more 

18 directly takes into account the effect of the non-uniform wall temperature distribution around the 

19 fuel assembly, which can be of particular significance in modeling fire scenarios. 

20

21 Axial conduction within the fuel assembly region was modeled only in the fuel cladding and 

22 backfill gas, to be consistent with typical applications of the fuel effective conductivity model, 

23 conservatively neglecting axial conduction in the uranium oxide fuel. The axial effective 

24 conductivity was determined with a cross-sectional area weighting scheme based on the total 

25 cross-sectional area of the assembly. However, to appropriately capture the thermal inertia of 

26 the fuel assemblies for the transient response in the fire scenario, the effective density and heat 

27 capacity for the fuel region was defined based on volumetric averages of the corresponding 

28 properties of the helium gas, fuel rod cladding, and uranium oxide fuel pellets. 

29

30 An average volumetric heat generation of 2105 Btu/hr (617 W) was applied over the active fuel 

31 length for each fuel assembly. The axial distribution of decay heat was represented by dividing 

32 the active fuel length into 16 separate zones, and the local heat load was determined by 

33 multiplying the average by an appropriate peaking factor for that particular zone. The peaking 

34 factor was determined based on the bounding axial power profile presented in the SAR, which 

35 has a normalized peaking factor of 1.1.

36

37 The helium gas filling the 0.5075-inch gap between the nominal fuel assembly cross-section and 

38 FSS was modeled with solid elements and used standard helium thermal properties for 

39 conduction, density, and specific heat. Convection across the gap was accounted for by 

multiplying the local gas conduction values by an empirically derived 40 1 Nusselt number of 3.66. 

41 Thermal radiation exchange across the gap was modeled with MATRIX50 super elements. 

42 These were created by using SHELL57 elements to designate the discrete enclosure. The 

43 AUX-12 hidden ray-tracing method was used to compute view factors for each element within 

44 the super-element. All other gaps in the package assembly, such as between the closure 

45 assembly and FSS, or the impact limiters and package skin were modeled in a similar manner, 

46 which included thermal radiation and conduction across surfaces but assumed negligible 

47 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

1 Other potential gaps not explicitly modeled within the geometry, such as between the gamma 

2 shield and package FSS, and between the gamma shield and stainless steel wall were 

3 accounted for by modifying the material properties of the adjacent materials to include the 

4 calculated effective properties for the material and gap. For very small gaps, the calculations 

5 were based on the following assumptions:

6

7 • the thermal radiation view factor is specified as 1.0 (gap completely enclosed) 

8 • the temperature difference across the gap is small 

9 • convection heat transfer across the gap can be neglected. 

10

11 For the pre-fire steady-state and post-fire transient cooldown analyses, nominal gap distances 

12 were used to determine the effective thermal conductivity. During the fire transient, the gaps 

13 were assumed to close due to thermal expansion of the package materials, such that the 

14 calculated effective thermal conductivity across a gap reduced to the thermal conductivity of the 

15 adjacent solid material without the gap. This ensured a conservative approach with respect to 

16 the effect of heat transfer across the gaps throughout the entire analysis.

17

18 The impact limiter attachment bolts and the closure assembly bolts were represented as line 

19 elements within the model. Small variations in the overall length of individual bolts were

20 accounted for by calculating an equivalent cross-sectional area, which was specified in the real 

21 constant properties for the line elements.

22

23 5.2.2 GA-4 Package beneath Collapsed Roadway

24

25 The location of the package and configuration of the concrete roadway overlay, as represented 

26 in the fire scenario modeling, is defined to maximize coverage of the package, within the 

27 physical geometry of the collapsed structures. The overall length of the GA-4 package without 

28 impact limiters is 15.65 ft (4.77 m), and the outer diameter is 3.3 ft (1 m). With impact limiters, 

29 the overall length is 19.5 ft (5.94 m) and the maximum diameter is 7.5 ft (3.39 m). Figure 5.6

30 shows block “footprints” of the package (with impact limiters) approximately to scale in the 

31 locations where an object of that size could be completely covered by the fallen roadway spans. 

32 The conveyance is not considered in this evaluation, as noted in Section 5.2.1 as a modeling 

33 conservatism for the fire exposure. It is also a conservative assumption for consideration of the 

34 effect of the roadway collapsing onto the package. Due to the size of the package and the 

35 conveyance, it would be virtually impossible to completely cover them both with the fallen 

36 roadway in this accident scenario. 

37
5-10

1

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

3 (photo from MAIT Report, CHP 2007, reprinted with permission)

4

5 Modeling of the structural response of the package due to the collapse of the roadway spans is 

6 discussed in Section 5.4. From the standpoint of thermal performance of the GA-4 package, the 

7 important effect of the collapse is the additional thermal barrier imposed by a concrete “blanket”

8 over the package in the post-fire cooldown. Of the two possible locations where the package 

9 would be completely covered by a fallen span, the location beneath the fully collapsed Bent 19-

10 20 span clearly imposes the more severe conditions. At the location beneath the partially 

11 collapsed Bent 18-19 span, only the top of the package would be covered, due to the limited 

12 area of contact between the upper span and the lower roadway. Beneath the completely 

13 collapsed Bent 19-20 span, the package would be entirely covered by the fallen roadway.

14

15 Figure 5.7 shows a cross-section diagram of the assumed configuration for the GA-4 package 

16 beneath the collapsed I-580 span. As a conservative simplification, it is assumed that the 

17 package is not actually in physical contact with the concrete structure enclosing it, and the 

18 neutron shield tank outer shell is not deformed by contact with the fallen span. This preserves 

19 the air gap of the evacuated shield tank, which insulates the package body during cooldown. 

20 The crushed girders on the underside of the fallen overhead span are assumed to hold the 

21 roadway 9.4 inches (0.24 m) away from the package body. This distance is estimated from the 

22 collapsed height of the severely deformed girders of the actual roadway structure that were 

23 exposed to the intense heat of the large pool fire. In addition, the package is assumed to stand 

24 on its lifting trunnions, so that there is a 5-inch (0.13-m) gap between the package body and the 

25 lower roadway. A similar gap is assumed between the package outer shell and the surface of 

26 the concrete edge barrier. The touch-point of the fallen I-580 span on the I-880 roadway is 

27 assumed to be approximately one-quarter of the lower roadway width, at 4.66 ft (1.42 m) from 

28 the package centerline. Given the stiffness of the concrete-and-steel roadbed, it is reasonable 

29 to suppose that the presence of the package would tend to hold the upper span away from 

30 contact with the lower roadway over much of its width, which makes this an extremely 

31 conservative assumption.
5-11

1

2 Figure 5.7. Diagram of Collapsed Roadway Configuration over SNF Package

3

4

5 In the ANSYS model, the configuration shown in Figure 5.7 is represented by enclosing the 

6 entire SNF package within a simplified structure representing the external environment, as 

7 illustrated in Figure 5.8. During the fire, this structure provides boundary conditions for the 

8 thermal analysis simulating a fully engulfing fire surrounding the package. (See Section 6.4 for 

9 discussion of the modeling of the structural response of the package in this fire scenario.) After 

10 the end of the fire, the surrounding enclosure models the configuration of the package 

11 positioned on the lower I-880 roadway adjacent to the side barrier, beneath the fallen upper I12 580 roadway. Heat transfer between the package and the enclosing concrete was assumed to 

13 consist of conduction and convection through the surrounding air within the enclosure, and 

14 thermal radiation exchange between the package external surfaces and the concrete surfaces 

15 representing the roadway above and below. This provides a conservative representation of the 

16 postulated package configuration during the post-fire cooldown.

17

NOTE: Diagram not to scale
5-12

1

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

3

4

5 Thermal radiation interaction between the package and concrete roadway surfaces was defined

6 by coating the exterior surfaces with SHELL57 elements having appropriately specified emissive 

7 material properties. The SHELL57 elements were then used to produce highly structured AUX8 12 generated MATRIX50 super elements. A total of twenty-four MATRIX50 super elements 

9 were defined to capture the radiation interaction between the package and roadway surfaces. 

10 The twenty-four MATRIX50 super elements represent two sets of thermal radiation conditions; 

11 half of the super elements represented pre-fire thermal radiation emissivities, and the other half 

12 represented the fire and post-fire emissivities. Only one set of external MATRIX50 super13 elements was active at a time during the calculation, depending on the stage of the analysis 

14 (i.e., pre-fire, fire, or post-fire).

15

16 The concrete enclosure is essentially the same for the ANSYS and the COBRA-SFS models, 

17 except for the dimensions assumed for the post-collapse structure. For the ANSYS model, the 

18 size of the enclosure is expanded to accommodate the impact limiters, while in the COBRA-SFS 

19 model, the concrete overlays the package body (as described in Section 5.3). In reality, of 

20 course, the falling roadway would crush the impact limiters against the package body, but this 

21 was conservatively neglected in the ANSYS model. Accounting for the effect on thermal 

22 behavior due to crushing of the impact limiters would add unnecessary complication to model,

23 and in addition, would be a less conservative approach, as it would tend to enhance conduction 

24 heat transfer from the package. Assuming the stainless steel shells of the impact limiters retain 

25 their original configuration imposes a conservative representation of the insulating barrier these 

26 structures present to heat transfer from the package during the cooldown phase of the transient.

27
5-13

1 Convection heat transfer on the external surfaces of the package was determined using

2 empirical correlations, which are described in Section 6.3. Free convection was assumed 

3 before and after the fire, while forced convection was assumed during the fire. Convection was 

4 implemented using SURF152 elements. These elements are placed on the exterior surface of a 

5 body and communicate with the designated sink temperature assigned to a single node (called 

6 the “space node”) to compute the heat flux. Figure 5.9 shows a representation of a space node 

7 relative to the package assembly.

8

9

10 Figure 5.9. Space Node for Convection Heat Transfer

11

12 In order to represent variation in convection heat transfer over different regions of the package 

13 surface, due to variation in the surface geometry, the external surfaces were partitioned into 

14 axial and radial zones. The package was divided into four different zones along the exposed 

15 outer surfaces, corresponding to the top, bottom, and left and right sides of the horizontal 

16 package, as illustrated by the colored segments in Figure 5.10, to appropriately model variations 

17 in convection along the exterior of the package. On the package body, these zones correspond 

18 to the radial positions of the four fuel assemblies within the package. The outer surfaces of the 

19 impact limiters were also split into four zones in the same manner as the package body, as

20 illustrated in Figure 5.10. In addition, the four lifting trunnions were split into two different zones 

21 each, with two located along the top package boundary and the remaining two located along the 

22 bottom boundary of the package. Within each zone different heat transfer coefficient 

23 correlations were used, corresponding to the geometry of the surface (e.g., vertical flat surface, 

24 horizontal curved surface), by defining subzones within each zone. A sink node was defined for 

25 each zone and subzone, which defines the ambient air temperature seen by the package in that 

26 zone or subzone, and is used to calculate the local convection coefficient at the package 

27 surface. 
5-14

1

2 Figure 5.10. Zones for Convection Computations for Package Assembly Surfaces

3

4 The external environment was split into eight different zones corresponding to the upper 

5 roadway and lower roadway, with the ends open to the ambient, as illustrated in Figure 5.11. 

6 (As shown in Figure 5.11, these modeling structures are tilted and rotated for clarity; when 

7 properly oriented, the cross-section of the defined zones match the cross-section of the model 

8 diagram shown in Figure 5.8.) The convective heat transfer coefficient is assigned to the 

9 package and external elements based on the temperature difference between the surface and 

10 sink temperature, and on the surface geometry, as described in Section 6.0. The heat 

11 exchange between these surfaces and the space node is then computed by ANSYS as part of 

12 the overall solution for a given time-step in the transient.

13

14

15 Figure 5.11. Zones for External Heat Transfer from Roadways and External Environment
5-15

1 During the fire, the sink node temperature for the SURF152 elements is set to the bounding 

2 flame temperature representing the fully engulfing fire. The external convection coefficient is 

3 computed using a forced convection relation derived from gas temperatures and velocities 

4 predicted in the fire simulation with the FDS code (as discussed in detail in Section 3.3). These 

5 boundary conditions are provided for the top, sides, and bottom of the enclosure, corresponding 

6 to the four zones on the package, as shown in Figure 5.10. During the fire, the nodes 

7 representing the external environment are also set to the bounding flame temperature.

8

9 5.2.3 Material Properties for GA-4 Package in ANSYS Model

10

11 The specific thermal material properties used to represent the components of the GA-4 package 

12 and roadway structures in the ANSYS model are listed in detail in Appendices A and B. For 

13 elements of the model representing the major components of the package, the specified 

14 properties are those of the single material comprising that component. However, for efficiency 

15 of meshing, the thin plates of the FSS and enclosed neutron absorber rods, the complex 

16 honeycomb structures of the impact limiters, and the fuel assemblies are represented using 

17 effective thermal properties defined specifically for the overall region. In addition, the effect of 

18 the fire on the integrity of the liquid-filled neutron shield tank was also explicitly modeled with 

19 changes in material properties in the transient calculation.

20

21 The neutron absorber plates of the FSS are composed of boron carbide rods sandwiched 

22 between thin stainless steel (XM-19) panels, with helium surrounding the boron carbide rods. 

23 Homogeneous material properties were defined for the elements representing the FSS plates, 

24 based on volumetric averaging of the material properties for XM-19 stainless steel, boron 

25 carbide, and helium. It was assumed that convection in the helium gas would be negligible in 

26 the narrow enclosed space within the FSS plates, and the effective thermal conductivity was 

27 calculated based on conduction and thermal radiation heat transfer only. Anisotropic properties 

28 were defined for this material, assuming conduction only along the axial length of the FSS, with 

29 conduction and thermal radiation through the thickness of the composite plate. Thermal 

30 radiation was modeled assuming that the helium-filled space between the boron carbide rods 

31 and the enclosing steel plates was very small, completely enclosed within the stainless steel 

32 panels, with a very small temperature difference between them. 

33

34 The stainless steel inner support structure and outer shell of the impact limiters was explicitly 

35 modeled using elements with properties of XM-19 stainless steel. Composite material 

36 properties were used to model the aluminum honeycomb material enclosed within the steel 

37 shell. The design of the impact limiters is defined in the package SAR (General Atomics 1998) 

38 as a standard non-reinforced hexagonal aluminum structure, and includes specific regions with 

39 differing densities, which are bonded together and to the stainless steel shell with adhesive 

40 foam. Effective properties for these regions were determined based on material data for 

41 aluminum honeycomb from HEXEL Composites (1999), using a volumetric averaging scheme. 

42 This approach included the properties of the adhesive foam as well as the air-filled aluminum

43 honeycomb. The effective thermal conductivity values for the honeycomb regions were 

44 calculated assuming the material was isotropic within a region, as indicated by the HEXEL 

45 Composites data for the honeycomb.

46

47 In the course of the transient calculation, the material properties of the impact limiters were 

48 modified to account for structural configuration changes and effects of the fire. Portions of the 

49 aluminum honeycomb in the impact limiters are assumed to melt during the fire, due to the 

50 extremely high temperatures predicted in this transient. For the aluminum honeycomb material 

51 in the impact limiters, local melting would be expected to significantly increase in the void (air) 
5-16

1 volume compared to the intact honeycomb material. This would tend to increase the insulating 

2 effect of the impact limiters, reducing the rate of heat transfer through this material. During the 

3 fire portion of the transient, the impact limiters were conservatively assumed to remain intact, 

4 allowing the maximum heat transfer to the package through these components during the fire. 

5 However, the assumption of intact impact limiters is no longer conservative in the post-fire 

6 cooldown portion of the transient. With larger air regions within the impact limiter structure due 

7 to local melting of the honeycomb, the damaged impact limiters would tend to further slow the 

8 rate of heat removal from the package during the cooldown transient, compared to the effect of 

9 intact impact limiters. The material properties of elements in the ANSYS model representing the 

10 honeycomb material were therefore modified in the post-fire portion of the calculation to account 

11 for the effects of melting. 

12

13 Fire damage to the impact limiters was determined from the predicted temperature distribution 

14 within these regions at the end of the fire. The percentage of honeycomb nodes above the 

15 melting point of the aluminum alloy (approximately 1100°F [593°C]) was used to calculate the 

16 total volume of melted aluminum, and the volume of “lost” honeycomb. It was assumed that the 

17 molten aluminum would flow due to gravity to the lowest point on the horizontal side of the 

18 impact limiters. Therefore, elements in this region encompassing a volume corresponding to 

19 the volume of melted aluminum were modified to have the properties of aluminum alloy, rather 

20 than the honeycomb mesh. The remaining volume of the impact limiter was assumed to be a 

21 mixture of air (corresponding to the volume of the melted mesh) and unmelted intact 

22 honeycomb. The thermal conductivity of the elements representing this volume within the 

23 impact limiters was modified using an effective thermal conductivity calculated based on 

24 volume-averaging of the thermal properties of air and the unmelted honeycomb mesh material. 

25

26 The effect of the fire on the material properties of the liquid neutron shield was also explicitly 

27 represented in the transient calculation. The neutron shield liquid temperature is calculated to 

28 exceed its boiling point very early in the fire transient. Prior to rupture, heat transfer through the 

29 liquid in the tank is represented with an effective conductivity relationship based on an empirical 

30 correlation (Guyer and Brownell 1989) for convection and conduction heat transfer across a gap 

31 between two long, horizontal concentric cylinders at different temperatures. The fluid thermal 

32 conductivity used in this relationship was determined based on material property data for 

33 propylene glycol and water mixtures provided in the GA-4 SAR (General Atomics 1998).

34 (Appendix B contains a detailed description of this correlation, and verification of its applicability

35 to the geometry of the GA-4 neutron shield tank.) 

36

37 The neutron shield tank is assumed to rupture when the peak temperature in the liquid is 

38 predicted to exceed the boiling point of the water-glycol mixture. After rupture, the neutron 

39 shield tank contents are assumed to consist only of air, with heat transfer by conduction and 

40 convection. Thermal radiation between the inner walls of the empty tank is also accounted for, 

41 by direct calculation between the elements on the inner surface of the tank outer shell and the 

42 outer surface of the package body. 

43

44 The effective conductivity of the material within the neutron shield tank was determined as a 

45 function of the average tank temperature and the radial temperature difference between the 

46 tank inner and outer surfaces. The radial temperature difference was calculated separately 

47 along the flats and corners of the neutron shield, to account for the effect of the non-uniform gap 

48 due to the square cross-section of the tank inner surface within the circular outer tank shell.

49 Material properties for the tank were updated between each time step during the transient 

50 solution. The affected nodes were assumed to consist of a 56% propylene glycol solution up to 

51 the point where the maximum temperature reached the mixture’s boiling point of 276°F (136°C).
5-17

1 The boiling point for the tank contents, and hence the time of assumed tank rupture, was 

2 calculated based on the maximum normal operating pressure of the neutron shield tank

3 (General Atomics 1998), and data for vapor pressure versus temperature of aqueous solutions 

4 of propylene glycol (Dow Chemical Company 2003). When the maximum temperature in the 

5 tank exceeded the boiling point, it was assumed that rupture had occurred and all the liquid in 

6 the tank instantly vaporized. The effective conductivity was then computed using dry air as the 

7 medium. This calculation extended through the remainder of the fire and was also continued 

8 during the cooldown period. This approach conservatively neglects energy absorbed by the 

9 phase change (i.e., the heat of vaporization for the liquid), but this is mainly as a matter of 

10 convenience, since this would constitute a very small deduction from the total energy imparted 

11 to the package. 

12

13 5.3 COBRA-SFS Model of GA-4 Package

14

15 The GA-4 package was also analyzed with COBRA-SFS (Michener et al. 1995), a thermal16 hydraulic code developed for analysis of multi-assembly spent fuel storage and transportation 

17 systems. The code uses a lumped-parameter finite-difference approach for predicting flow and 

18 temperature distributions in spent fuel transfer, storage, and transportation systems, and fuel 

19 assemblies under forced and natural circulation flow conditions. It is applicable to both steady20 state and transient conditions in single-phase gas-cooled spent fuel packages with thermal 

21 radiation, convection, and conduction heat transfer. 

22

23 The COBRA-SFS model was developed to provide detailed temperature distributions on the 

24 individual fuel rods within the package in this fire scenario. In addition, the COBRA-SFS model 

25 was used to investigate the potential effects on the package if it is assumed that the impact 

26 limiters have detached from the package in the accident prior to the fire. The impact limiters 

27 have a very large effect on the thermal response of the package internals to fire conditions, 

28 since these structures act as thermal shields on the package ends. The impact limiters are 

29 designed to remain attached to the package during design-basis accident scenarios, but the 

30 MacArthur Maze fire involves fire temperatures exceeding that of the HAC fire defined in 

31 10 CFR 71. In addition, this scenario includes structural impacts due to the collapse of the 

32 overhead highway spans after the fire exposure, which is the reverse of the impact-then-fire 

33 scenario prescribed for hypothetical accidents in 10 CFR 71. It is unlikely that the impact 

34 limiters would come off, even in this extraordinary scenario, but if it were to happen, it would 

35 significantly influence the thermal behavior of the package. The effect of the loss of the impact 

36 limiters on the package response to this fire is therefore included in this analysis.

37

38 As in the evaluations with the ANSYS model, evaluations with COBRA-SFS consider two 

39 configurations for the package during the fire scenario. Prior to and during the collapse of the 

40 two overhead roadway spans (as discussed in Section 2.1), the package is assumed to be in a 

41 fully engulfing fire at 2012°F (1100°C), and sees a uniform bounding fire temperature in all 

42 directions. After the collapse, the package is assumed to be exposed to a fully engulfing 

43 1652°F (900°C) fire. In a separate independent calculation, the concrete structure of the 

44 collapsed roadway configuration is exposed to these fully engulfing fire conditions, in 

45 preparation for the post-fire cooldown calculation. The concrete surface temperatures obtained 

46 in this calculation are used as boundary conditions for both the ANSYS model and the COBRA47 SFS model evaluations of the post-fire cooldown conditions. For the post-fire analysis with 

48 COBRA-SFS, the model includes the GA-4 package, the lower roadway, side barrier, collapsed 

49 upper roadway, and the air space between the package and the concrete structures. The 
5-18

1 model of the package is described in Section 5.3.1. The package-under-concrete configuration 

2 for the post-fire scenario is discussed in Section 5.3.2.

3

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

5

6 The GA-4 package was modeled for the COBRA-SFS calculations in sufficient detail to capture 

7 the thermal response of the system components in the radial and axial directions. Material 

8 properties used in the model are listed in Appendix A. The four fuel assemblies within the 

9 basket are each modeled as rod and subchannel arrays (Michener et al. 1995), for appropriate 

10 representation of thermal radiation heat transfer as well as conduction and convection. The 

11 basket separating and containing the fuel assemblies is represented using multiple layers of 

12 solid conduction nodes, to capture the effect of the B4C poison rods stacked within the steel 

13 plates forming the arms of the cruciform structure. 

14

15 The steel inner liner, DU gamma shield, and steel package body are also represented with 

16 multiple layers of solid conduction nodes, to appropriately resolve the temperature gradients 

17 through these relatively thick components. In addition, these structures are also divided radially, 

18 to capture the effects of non-uniform external conditions surrounding the package. A cross19 section diagram illustrating the noding for the COBRA-SFS model is shown in Figure 5.12. 

20

21

22

23 Figure 5.12. Cross-section of COBRA-SFS Model of GA-4 Package, Including Fuel 

24 Assemblies, Basket, Package Body, and Neutron Shield

25

26

27 The impact limiters on the package ends are assumed lost, and are not included in the COBRA28 SFS model of the GA-4. The impact limiters are designed to remain in place through the 

29 sequence of accidents specified in 10 CFR 71 for HAC; however the MacArthur Maze fire 

30 scenario involves fire temperatures exceeding that of the HAC fire. Although it is unlikely that 

31 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

1 impact limiters would be expected to suffer severe degradation, due to the long fire duration and 

2 the postulated impact of the overhead span on the package. The impact limiters provide 

3 significant thermal shielding for this package in any fire scenario, and therefore can have a large 

4 effect on the response of the system to the fire. Therefore, this assumption is included in the 

5 COBRA-SFS model, for completeness in the overall analysis of the effects on the package of

6 the MacArthur Maze fire scenario. However, evaluation results shown in Section 7.3.6.1 show 

7 that the impact limiters would remain attached to the GA-4 package. Results obtained with the 

8 COBRA-SFS model assuming loss of the impact limiters are therefore bounding and 

9 conservative. 

10

11 The neutron shield tank initially contains a liquid 56% propylene glycol/water mixture with 

12 maximum design pressure of 150 psig (1.135 MPa), which gives a boiling temperature of 276°F 

13 (135.6°C). However, the tank is not an ASME pressure vessel, and in the SAR analyses for the 

14 HAC fire at 1472°F (800°C), it is conservatively assumed that in the initial steady-state, the tank 

15 has already ruptured and contains only air. In the COBRA-SFS analysis, the tank is assumed to 

16 remain intact up to the point in the transient where the maximum temperature in the neutron 

17 shield region exceeds the boiling point of the glycol/water mixture. Heat transfer across the 

18 neutron shield tank is treated in the same manner as in the ANSYS analysis, described in 

19 Section 5.2.3, with an effective conductivity correlation defined for the neutron shield tank 

20 contents.

21

22 The maximum temperature in the region modeling the tank contents is predicted to occur within 

23 sixty seconds of the initiation of the fire transient in this scenario, at which point the medium 

24 within the tank is assumed to be dry air, and thermal radiation between the tank inner surfaces 

25 is added to the model. The internal surfaces of the shield tank are specified with a uniform 

26 emissivity 0.9 after the assumed loss of liquid contents, to conservatively represent the effect of 

27 sooting, on the assumption that highly sooted fire gas could enter the ruptured and fully vented 

28 tank. Additional package modeling details related to boundary conditions and assumptions are 

29 included in Section 6.3.

30

31 5.3.2 COBRA-SFS Model of GA-4 Package beneath Collapsed Upper Roadway

32

33 The representation of the GA-4 package for the post-fire cooldown with the package beneath 

34 the collapsed upper roadway is the same as the model developed for the fully engulfing fire 

35 portion of the scenario. For the post-fire cooldown, the model is expanded beyond the package 

36 exterior shell to include additional nodes representing the concrete of the upper and lower 

37 roadways in relation to the package. The location of the package and configuration of the 

38 concrete covering it is defined to maximize coverage of the package, within the physical 

39 geometry of the collapsed structures in this fire scenario, as discussed in Section 5.2.2 and 

40 illustrated by Figure 5.7. The concrete enclosure is essentially the same for the ANSYS and the 

41 COBRA-SFS models, except for the interior dimensions of the post-collapse structure. For the 

42 ANSYS model, the size of the enclosure is expanded to accommodate the diameter of the 

43 impact limiters, while in the COBRA-SFS model, the concrete overlays the package body. 

44

45 As illustrated in Figure 5.7, the space between the package and the concrete structure is 

46 modeled as an air channel, roughly analogous to the air flow pathway in a ventilated concrete 

47 storage module. The main differences are that the orientation is horizontal, and the geometry of 

48 the concrete “annulus” is decidedly irregular. Heat transfer to the environment is modeled as 

49 free convection from the exposed surfaces of the concrete roadways and side barrier. For 

50 simplicity, the cross-section geometry shown in Figure 5.7 is assumed to extend uniformly over 

51 the axial length of the package, and the “edge” of the model where the upper span touches the 
5-20

1 lower roadway is treated as an adiabatic boundary. Both assumptions are conservative, as they 

2 limit or eliminate potential paths for heat to be transferred from the package by conduction 

3 through the concrete.

4

5 5.4 Structural Models for the MacArthur Maze Fire Scenario

6

7 The structural evaluations undertaken to determine the potential consequences of the 

8 MacArthur Maze fire scenario focus on two main issues. The first issue is to evaluate the effect 

9 of dropping a section of the I-580 roadway onto a SNF package that has been subjected to a 

10 fully engulfing fire. This accident scenario differs significantly from the HAC drop scenarios 

11 postulated in 10 CFR 71, and it is necessary to determine whether or not this real-life accident 

12 scenario is bounded by the postulated accidents defined in the regulations. The second issue is 

13 the structural response of the package to the extraordinary thermal load imposed on the 

14 package closure lid bolts and on the impact limiter attachment bolts in this fire scenario. The 

15 severity and duration of the MacArthur Maze fire scenario exceed the postulated HAC fire in 

16 10 CFR 71. Retention of the impact limiters on the package body and the response of the 

17 closure lid bolts greatly affect the potential consequences of this fire scenario, if an SNF 

18 package were to be subjected to such severe conditions. 

19

20 Evaluation of these two important issues requires modeling approaches on two different scales. 

21 The drop scenario evaluations require large-scale modeling of the package and roadway 

22 structures. These models are described in detail in Sections 5.4.1, and include both the 

23 roadway and the SNF package. The bolt response evaluations require detailed modeling of the 

24 regions of the package containing the impact limiter attachment bolts and the lid closure bolts. 

25 These evaluations use two independent approaches to evaluate this issue, initially using

26 classical bolt modeling studies, and then developing detailed FEA modeling of these structures 

27 in the GA-4, with boundary conditions provided from the thermal modeling of the package. The 

28 classical modeling and preliminary FEA modeling is described in Section 5.4.2 and the detailed 

29 FEA modeling is described in Section 5.4.3.

30

31 5.4.1 Structural Models for Roadway Drop Scenarios

32

33 This modeling study is a three-dimensional impact analysis evaluating the effect of dropping a 

34 section of the I-580 roadway onto the SNF package, which has been subjected to a fully 

35 engulfing fire at 2012°F (1100°C). The structural impact models were developed using the LS36 DYNA code (Livermore Software Technology Company 2007) to simulate the impact of the 

37 falling overhead I-580 span on the SNF package positioned beneath it on the lower I-880 

38 roadway. This includes appropriate models of the roadway segments involved, and of the SNF 

39 package. These models were used to evaluate the potential for gross structural failure of the 

40 package, which could affect the containment boundary. The structural models of the roadway 

41 are described in Section 5.4.1.1, and the model of the SNF package is described in 

42 Section 5.4.1.2.

43

44 5.4.1.1 Modeling of Roadway Segments

45

46 The potential effects of the collapse of the overhead roadway onto the SNF package is 

47 investigated in this analysis by modeling a free-fall drop of the I-580 roadway span between 

48 Bent 19 and Bent 20 onto a structural model of the GA-4 package on the lower I-880 roadway. 

49 The partial collapse of the span between Bent 18 and Bent 19 was not modeled, as it would 

50 result in a much less severe impact on the SNF package. Based simply on the physical forces 
5-21

1 involved, the slower partial collapse of the span between Bent 18 and Bent 19 could not impart 

2 as much energy to the SNF package as the rapid free-fall drop of the first span to collapse, and 

3 therefore would not have the potential to do as much damage. 

4

5 For analysis of the drop scenario, the structural model in LS-DYNA consists of the upper I-580 

6 roadway span between Bent 19 and Bent 20, the lower I-880 roadway segment beneath the 

7 overhead span, and the SNF package at rest on the I-880 roadway. Figure 5.13 shows a 

8 conceptual diagram of the impact model geometry. Each object is represented with a minimum 

9 amount of detail to make the calculation size manageable. The upper I-580 span and the I-880 

10 span are modeled with similar components, but the upper I-580 span includes more detail, 

11 higher element resolution, and represents the steel girders with material properties that reflect 

12 the much higher temperature reached by the girders in the fire scenario. 

13

14

15 Figure 5.13. Impact Model Geometry

16

17

18 The level of detail in each component reflects its function in the model. The falling I-580 girders 

19 need greater element resolution because they impact and deform around the package and the 

20 roadway side barriers. The fine mesh is illustrated in Figure 5.14 and in Figure 5.15, which 

21 more clearly shows the increased mesh density in the two girders that impact directly on the 

22 SNF package. The girders of the I-880 roadway (also shown in Figure 5.14) are represented 

23 with a much coarser mesh, as they are needed only to provide a supporting surface with 

24 realistic stiffness beneath the package.

25
5-22

1

2 Figure 5.14. Impact Model Mesh

3

4

5 Figure 5.15. Impact Model Mesh Density in Impact Region

6
5-23

1 The upper I-580 roadway span between Bent 19 and Bent 20 is modeled as a deformable 

2 impact object, using geometry information from the original plate girder design drawings. The 

3 plate girders are the most important components of the overpass system for the impact 

4 modeling because these are the roadway structures that would contact the package directly in 

5 any potentially damaging drop scenario. The longitudinal girders are modeled with the major I6 section plates connected as continuous material. Small lateral stiffening webs and the lateral 

7 plates and crossing members that connect to the longitudinal girders are neglected. This 

8 modeling approach is based on post-fire photographs (CHP 2007) showing that the main girder 

9 plates remained attached to the underside of the roadway but the stiffening ribs and cross 

10 members tended to fail or separate. Under localized impact, these components would be 

11 expected to fail without significantly influencing the impact response of the system. 

12

13 Figure 5.16 shows a diagram of the mesh for the lower roadway (I-880) portion of the impact 

14 model. The finite element representation is similar to the upper roadway span, but with a 

15 coarse mesh density. A broad rigid surface (omitted from the diagram, for clarity) is also 

16 included in the model at ground level, to catch the end of the upper roadway that hangs over the 

17 edge of the lower roadway. In the impact analysis, the results show that one end of the falling 

18 roadway impacts the package while the other end continues falling until it hits the ground. This 

19 results in a significant bending of the falling span along the edge of the lower roadway. The 

20 purpose of this modeling effort does not extend to matching the final resting state of the fallen 

21 span, but it is a useful verification of the modeling fidelity that the overall behavior predicted with 

22 this model is consistent with the post-fire images of the fallen span, such as the one shown in

23 Figure 1.1.

24

25

26 Figure 5.16. Model Mesh of Bottom Roadway (I-880) Segment

27
5-24

1 The concrete and rebar of the I-580 roadway is modeled as a homogenized linear-elastic 

2 material with an artificially low modulus of elasticity. Actual concrete behavior during the impact 

3 would be more complex, with concrete fracturing and breaking apart, but these nonlinear effects 

4 absorb energy and would reduce the amount of energy available to do damage to the package. 

5 The low elastic modulus for the upper roadway concrete results in the weight of the upper 

6 roadway acting on the plate girders through the full period of impact. Treatment of the roadway 

7 as a linear-elastic material is conservative in that it contributes to higher and more prolonged 

8 impact forces than would be obtained with a more physically realistic nonlinear model. 

9

10 The steel girders of the upper I-580 roadway span are represented as elastic-plastic material, 

11 with temperature-dependent mechanical properties to account for effects of the elevated 

12 temperatures due to the fire. A uniform temperature distribution is assumed for the final set of 

13 analyses, so the material properties do not vary spatially. Tiebreak contact definitions are used 

14 in the model to attach the roadway to the plate girders. These are virtual links in the model that 

15 connect a node to a surface until a specified reaction force is exceeded, at which point the node 

16 is free to move away from the surface. However, even in cases when the tiebreaks fail due to 

17 excessive normal or shear force, the plate girders and roadway still detect contact with each 

18 other and all other components in the model through an all-inclusive general contact definition.

19

20 The falling span is subjected to constant acceleration due to gravity and a specified initial 

21 downward velocity that depends on the placement of the cask on the lower roadway. It is 

22 assumed that the overpass section falls straight down and the cask is resting on the I-880 

23 roadway surface. Sagging of the girders before collapse is not considered, as it would reduce 

24 the potential freefall distance and associated impact velocity. The selected location and 

25 orientation of the cask dictates the distance the overpass can fall before it makes contact with 

26 the cask. That freefall distance and the initial location of the bridge span in the model dictates 

27 the initial velocity, which averages 30 ft/s (9.1 m/s) for the four impact scenarios considered. 

28

29 If it were assumed that the package remained on the conveyance, the drop distance would be 

30 about 3.3 ft (1 m) less, resulting in a lower impact velocity (approximately 15 ft/s [4.41 m/s]), and 

31 therefore lower kinetic energy upon impact. In addition, the crushing of the conveyance would 

32 absorb some of the impact energy. The modeled drop assumes an uninterrupted fall onto the 

33 package on the roadway surface, neglecting all factors that might mitigate the force of the 

34 impact, including drag effects. These conservative assumptions result in a much more 

35 damaging drop scenario than could occur in reality in this accident scenario.

36

37 The support columns initially holding up the upper I-580 roadway span and the lateral box beam 

38 spanning the Bent 19 support columns are conservatively neglected in the model, since they 

39 have essentially no effect on the impact scenario. Based on post-fire photographs (CHP 2007), 

40 the box beam remained attached to the partially collapsed span between Bent 18 and Bent 19. 

41 As noted above, the partial collapse of that span was not analyzed in the impact model because 

42 the impact loads would be much smaller than in the case of the freefall of the span between 

43 Bent 19 and Bent 20.

44

45 The lower I-880 roadway is conservatively modeled to provide a foundation of appropriate 

46 stiffness beneath the package. It consists of a set of plate girders supporting the concrete 

47 roadway, with the ends of the plate girders fixed in space. The model also includes the 

48 concrete barriers on either side of the roadway deck. All materials of the lower roadway are 

49 treated as linear-elastic, and are represented with properties at 80°F (27°C) nominal 

50 temperature. This conservatively neglects temperature-dependence of the material properties 

51 of the lower roadway, which would be more compliant at higher temperatures. In the model, the 
5-25

1 girders and side barriers are connected to the concrete roadway by tied surface definitions. 

2 Tied surface constraints effectively bond two surfaces by forcing the nodes of the “slave” side to 

3 maintain their relative spacing with the nodes of the “master” side. This type of connection 

4 persists throughout the analysis, unlike the tiebreak connections used in the upper roadway that 

5 allow the connection to end when the forces required to maintain the connection exceed a 

6 prescribed limit.

7

8 Conservative estimates of temperatures were obtained for the girders in the drop calculations, 

9 based on the results of thermal analyses to determine fire effects on the overhead girder 

10 temperatures. For the girders, a conservative temperature would be as low a temperature as 

11 can be reasonably postulated; at lower temperatures, the steel is stronger and therefore able to 

12 impart more energy to the package. Based on sensitivity studies of the effect of the fire on the 

13 overhead span prior to falling, a uniform temperature of 1800°F (982°C) was assumed for the 

14 girders and the upper I-580 roadway. The development of this important boundary condition is

15 discussed in detail in Section 6.4.1. The peak temperature on the I-580 girders may have been 

16 higher than 1800°F (982°C), possibly as high as 2462°F (1350°C), but for the purposes of this 

17 analysis, the lower estimate of the girder temperatures is conservative, compared to assuming 

18 higher temperatures. Because of the temperature-dependent properties of steel (see 

19 Figure 2.10 for yield strength versus temperature), the lower temperature assumption results in 

20 a more rigid overhead span, and hence a more severe impact on the package. 

21

22 5.4.1.2 GA-4 Package Model for Structural Analysis

23

24 The GA-4 package model is simplified to represent only the components that give it structural 

25 integrity. These are the steel package body, the DU gamma shield, and the lifting trunnions 

26 attached to the steel body. The fuel assemblies and FSS structure within the package cavity 

27 are neglected, as they would have limited effect on the package response to an external impact.

28 The impact limiters and the fragile steel shell of the neutron shield tank are also ignored in this 

29 simulation, as a conservative simplifying assumption, since the effect of these structures would 

30 be to absorb energy as they crumpled under the impact of the overhead span. 

31

32 The XM-19 stainless steel package body is a long rectangular box with rounded corners, with 

33 four lifting trunnions attached near either end, on the rounded corner edges. Figure 5.17 shows 

34 a diagram of the mesh used to represent the package for the calculations with LS-DYNA. The 

35 lid and flange region is treated as a solid volume of steel, without the physical interface between 

36 the two components, resulting in a fully enclosed package body. Realistic interaction between 

37 the lid, flange, and lid bolts is not captured in this model. The behavior of these components is 

38 investigated separately, using models described in Sections 5.4.2 and 5.4.3.

39
5-26

1

2 Figure 5.17. Model Mesh of the GA-4 Package Body

3

4

5 The DU gamma shield is modeled inside the package as a single free-moving body, since this 

6 component is enclosed within the XM-19 stainless steel cask body but is not physically attached 

7 to it. Figure 5.18 shows a diagram of the DU gamma shield mesh in the model. A small gap 

8 exists between the inner package wall and the gamma shield, on the order of 0.04 inches 

9 (1 mm), as represented in the design drawings from the SAR (General Atomics 1998). 

10 Dimensional changes in the package due to thermal expansion are neglected in this analysis, 

11 as their effect is minimal in a model with this level of detail.

12

13

14 Figure 5.18. Model Mesh of the DU Gamma Shield

15
5-27

1 The connection between the trunnions and the package wall are represented with tied surface 

2 definitions. In the model evaluations, the outer faces of the trunnions are constrained to prevent 

3 the package from rolling under the impact of the overhead span. This forces the impact with the 

4 descending span to occur on the uppermost corner of the package cross-section, as a 

5 conservative assumption to localize the impact interaction.

6

7 Conservative estimates of temperatures of the GA-4 package were obtained for the drop 

8 calculations, based on the results of the thermal analyses of the package. For the package, a 

9 conservative temperature would be as high a temperature as can be reasonably estimated, as 

10 the material of the package will then be weaker, and therefore less able to absorb the force of 

11 the impact without damage. A temperature significantly above the hottest steel body 

12 temperatures calculated in the thermal analysis of the GA-4 package in the MacArthur Maze fire 

13 scenario was specified for the GA-4 package body, to obtain a conservative representation of 

14 the potential effect of the impact. A detail description of boundary conditions for the structural 

15 calculations is included in Section 6.4.

16

17 5.4.2 Preliminary Models for GA-4 Bolt Thermal Expansion Calculations

18

19 Preliminary models were developed to investigate the response of the package bolted 

20 connections to the extremely high temperatures of this fire scenario, since failure of these bolts 

21 could potentially affect containment integrity of the package. There are two critical concerns for 

22 the GA-4 package in this fire scenario; the retention of the impact limiters throughout the entire 

23 transient, and performance of the package closure lid bolts. 

24

25 The impact limiter attachment bolts were evaluated to determine if these fasteners would be 

26 expected to hold at the elevated temperatures of this fire scenario. Failure of these bolts could 

27 potentially result in separation of the impact limiters from the package, which (for the top impact 

28 limiter) would subject the lid closure bolts to direct exposure to the engulfing fire. This would 

29 result in significantly higher temperatures for the lid closure and closure bolts than are predicted 

30 during the fire if the top impact limiter remains in place. 

31

32 The evaluations of the behavior of the impact limiter connection bolts and of the package lid 

33 closure bolts consisted primarily of studies of the effect of differential thermal expansion on bolt 

34 loading. Temperatures for the bolts and surrounding structures were determined from results of

35 the detailed transient simulations of the fire scenario with the ANSYS and COBRA-SFS thermal 

36 models. A diagram of the impact limiter geometry, with the locations of the attachment bolts 

37 indicated, is shown in Figure 5.19. (Note that this diagram is not to scale; the size of the bolts 

38 relative to the impact limiter is greatly exaggerated for clarity of illustration.) These bolts thread 

39 into an anchor plate that is attached to the steel body wall of the GA-4 package. Figure 5.20

40 shows a close-up view of the package closure lid and bolts, from the exploded view diagram of 

41 the entire package (shown in Figure 5.1).

42
5-28

1

2 Figure 5.19. Diagram of Impact Limiter, Illustrating Attachment Bolts

3

4

5

6 Figure 5.20. Package Closure Region, Showing Package Closure Lid and Bolts

7
5-29

1 The performance of the lid closure bolts was evaluated assuming that the impact limiters would 

2 be retained throughout the fire scenario, since the results of the evaluations of the performance 

3 of the impact limiter attachment bolts (as discussed in Section 7.0) shows conclusively and 

4 conservatively that the impact limiter attachment bolts would not fail in this fire scenario. The 

5 impact limiters therefore act as “friend and foe” to the closure lid during this fire scenario. 

6 During the fire, the impact limiters shield the package lid from the extreme temperatures of the 

7 fire, and therefore limit the temperature rise on this structure during the fire. In the post-fire 

8 cooldown, however, the impact limiter is an extremely effective insulator, and package 

9 temperatures in this region continue to rise for many hours after the end of the fire, as the 

10 extreme heat developed in the package body and internal structures during the fire naturally 

11 conducts toward the lower-temperature ends of the package. Evaluations of the closure bolts 

12 were undertaken to determine the response of the closure connection to the long exposure to 

13 elevated temperatures in the complete fire scenario, since the effect of the fire scenario on

14 these bolts could potentially result in failure of package containment.

15

16 Both sets of attachment bolts are Inconel 718, a nickel alloy (General Atomics 1998). The 

17 closure lid is secured with 12 bolts; 3 per side of the square cross-section of the lid. The impact 

18 limiter is attached with 8 bolts, uniformly spaced in a circular ring. The package body flange that 

19 receives the lid closure bolts is XM-19 stainless steel, as are the anchor plates that the impact 

20 limiter bolts thread into. In both sets of connections, steel helical thread inserts are used to 

21 make the connection between the bolt threads and the matching threads of the package body 

22 flange (for the lid closure bolts) or anchor plate (for the impact limiter bolts). 

23

24 The threaded inserts are ASTM Type 304 stainless steel. Inconel 718 and XM-19 stainless 

25 steel both retain their strength better than Type 304 stainless steel as temperature increases, 

26 which makes the threaded inserts a weak link in the threaded connection at elevated 

27 temperatures. In addition, the coefficient of thermal expansion for the nickel alloy of the bolts is 

28 significantly less than that of XM-19 stainless steel, so elevated temperatures will cause an 

29 increase in bolt tension and thread shear load. This issue is further complicated by the large 

30 thermal gradients predicted in the impact limiter attachment bolts, based on the results of the 

31 thermal analyses of the GA-4 package in the MacArthur Maze fire scenario.

32

33 For both the closure bolts and the impact limiter bolts, a preliminary evaluation was performed 

34 using classical mechanics equations to calculate bolt tension at a series of uniform temperature 

35 states. The coefficient of thermal expansion mismatch between the bolt material and the flange 

36 or anchor material results in an increase in tension in the bolt shank, which must be supported 

37 by an opposing shear force on the bolt threads. The strength of the materials is a function of

38 temperature, so the yield and failure thresholds vary with time and changing temperature in this 

39 fire scenario. One concern in each case is the yield and failure limits of the bolt shank. The 

40 shear strength of the bolt threads are not a concern because the Type 304 thread inserts are 

41 weaker for all temperatures, and therefore would be expected to yield before the bolt threads. A 

42 classic thread-stripping equation is used to relate bolt tension to total shear force on the 

43 threaded interface. 

44

45 The closure bolts and the impact limiter bolts can yield without actually failing in this scenario. 

46 Bolt materials are expected to have sufficient ductility to reduce the design-basis tensile load to 

47 zero without failing, as a general design criterion. Only if the tensile load could not be relieved 

48 sufficiently by yielding, and the residual load after yielding was still high enough to exceed the 

49 ultimate strength of the bolt material, would the bolt be expected to fail. For the closure bolts, a 

50 potential consequence of bolt failure is an unattached package closure lid. For the impact 

51 limiter bolts, potential consequences of bolt failure include detachment of the impact limiter from 
5-30

1 the cask. In either case, failure of the bolts would indicate significant potential consequences in 

2 the overall evaluation of the MacArthur Maze fire scenario. 

3

4 The preliminary evaluations provided an initial indication of the bolt response to the MacArthur 

5 Maze fire scenario, but additional evaluations were also undertaken using finite element models, 

6 to obtain a more complete understanding of the potential consequences of this fire scenario. 

7 For the impact limiter bolts, the uniform temperature assumption used in the initial evaluations 

8 was determined to be overly simplistic, due to large temperature gradients along the length of 

9 the bolt and surrounding material. To investigate the effect of these temperature gradients, a 

10 fully three-dimensional FEA model of the impact limiter bolt as it connects into the package was 

11 developed using ANSYS, to analyze the structural and thermal response of a bolt and the local 

12 surrounding material. Figure 5.21 shows meshing diagrams of the impact limiter bolt model 

13 developed for this analysis. Temperature distributions along the axial length of the bolts, 

14 obtained from the detailed ANSYS thermal model of the package, were incorporated into 

15 evaluations with this impact limiter bolt model. 

16

17

18 Figure 5.21. Model Mesh of Impact Limiter Bolt

19

20 In the evaluation of the results obtained with the preliminary models for the response of the 

21 closure bolts in the fire scenario, it was determined that the evaluation using mechanics 

22 equations was excessively conservative, and did not yield a sufficiently detailed picture of the 

23 consequences of the bolt response to this fire scenario. Therefore, additional, highly detailed 

24 FEA modeling was undertaken for the end region of the package, including the closure lid, steel 

25 body flange, and closure bolts. The modeling developed for this evaluation is described in 

26 Section 5.4.3.

27

cut-away view
5-31

1 5.4.3 Detailed Models for GA-4 Closure Lid Bolt Evaluations

2

3 The results of the preliminary models evaluating the behavior of the bolts attaching the impact 

4 limiters to the package show conclusively that the impact limiters would be expected to remain 

5 in place during the severe conditions of this fire scenario. In addition, the preliminary models 

6 show that even with extremely conservative assumptions, the bolts fastening the closure lid will 

7 not fail due to stresses related to differential thermal expansion of dissimilar materials. 

8 However, the simplifications and conservatisms in the preliminary models do not provide results 

9 with a sufficient level of detail to determine the detailed history of the clamping force the bolts 

10 impart to the lid/flange interface throughout the fire scenario. These analyses therefore cannot 

11 provide a means of obtaining a reliable estimate of the size of any potential leakage pathway for 

12 the contents of the package to escape to the environment.

13

14 This is crucial to determining the potential consequences of the MacArthur Maze fire scenario, 

15 and an additional detailed study was undertaken to develop a realistic evaluation of the closure 

16 bolt response over time. Evaluation of the closure force on the flange for the entire duration of 

17 the fire scenario, including the post-fire cooldown to a new ambient steady-state, can be used to 

18 determine a meaningful evaluation of the potential effect on package containment for the 

19 conditions postulated in this fire scenario.

20

21 Section 5.4.3.1 describes the modeling approach used in the preliminary evaluations. 

22 Section 5.4.3.2 presents the detailed FEA model developed for this most important evaluation in 

23 the overall study of the potential consequences to an SNF package of an accident of the 

24 magnitude of the MacArthur Maze fire scenario.

25

26 5.4.3.1 Modeling for Preliminary Closure Lid Bolt Evaluations

27

28 Classic bolt equations and analysis methods were applied to evaluate the response of the 

29 closure lid, bolts, and flange to the conditions of the MacArthur Maze fire scenario. In this 

30 context, “classic” refers to the set of equations that define bolt and flange loading, and can be 

31 found in any comprehensive engineering textbook on the subject, or as closed form equations 

32 generally used in the literature. The primary reference relied upon in this study was 

33 Fundamentals of Machine Component Design by Juvinall and Marshek (Juvinall and Marshek 

34 1991).

35

36 The classic bolt calculations for the GA-4 closure bolts were used to determine estimates of the 

37 bolt tension and clamping force over time, extending from initial steady-state normal conditions 

38 of transport (NCT), through the fire transient, and on out through the long post-fire cooldown. 

39 These evaluations considered all three types of loads that these fasteners would be expected to 

40 experience in the MacArthur Maze scenario, which consist of the following: 

41

42 • the preload caused by the initial torque on the bolts (as per the package manufacturer’s 

43 technical specifications), 

44

45 • the external load caused by the internal gas pressure within the package cavity, and 

46

47 • the thermal load due to the dissimilar coefficients of thermal expansion of the materials of the 

48 XM-19 stainless steel closure lid and flange, and the Inconel bolts. 

49
5-32

1 The effect of the Type 304 stainless steel thread inserts was also taken into account, with 

2 sensitivity studies of the effect of this interfacing component on joint stiffness.

3

4 5.4.3.2 Modeling for Detailed FEA Evaluations of Closure Lid Bolt Response

5

6 In this evaluation, it is important to note a particular feature of the lid closure bolts of the GA-4 

7 package. They are referred to as “bolts” in the SAR, and to avoid confusion this terminology is 

8 also used in the current study, but these fasteners are in fact used as screws, since they are 

9 threaded into tapped holes in the steel flange of the package body. In order for these fasteners 

10 to function as bolts, they would have to pass through the flange of the steel package body and 

11 thread into nuts on the other side. The design of the GA-4 package body does not allow this 

12 configuration for the fasteners holding the lid in place.

13

14 The distinction between “bolts” and “screws” for the GA-4 package closure lid is not particularly 

15 relevant to the preliminary evaluation using classic bolt equations. Both types of threaded 

16 fasteners generally perform the same function and are subject to the same family of equations 

17 for determining thread stripping, initial torque to a specified pre-load, and sizing to withstand an 

18 external load. The details of the physical configuration of the fasteners only becomes important 

19 in the FEA modeling undertaken to obtain a more accurate and detailed evaluation of the 

20 response of these fasteners to the extreme conditions of the MacArthur Maze fire scenario.

21

22 As with the FEA model of the impact limiter bolt, described in Section 5.4.2 above, radial 

23 symmetry was assumed in the loading of the closure lid bolts throughout the MacArthur Maze 

24 fire scenario. Significant gradients are present only in the axial direction, parallel to the axis of 

25 the bolt shank. This allowed the modeling simplification of representing only a single bolt. To 

26 reduce the model to a more tractable computational size, it was further assumed that an 

27 individual bolt was essentially radially symmetric, and could be appropriately modeled with a 

28 one-quarter section of symmetry, as shown in Figure 5.22, illustrating the model geometry of the 

29 bolt, lid and flange. This illustration also includes a close-up image of the modeling of the 

30 helical coil of the Type 304 stainless steel thread insert filling the thread patterns of the 

31 conjoined bolt and flange. The lid and flange are represented as an annulus with an outer 

32 diameter equal to 2.7 inches. This corresponds to a minimum annulus area of about 4.6 square 

33 inches, which is effectively equivalent to the estimate used in the classic bolt equations.

34
5-33

1

2 Figure 5.22. Finite Element Model Geometry for Closure Lid, Flange, and Bolt, including 

3 Thread Insert

4

5 The mesh density in the region of the threaded interface and thread insert is shown in detail in 

6 Figure 5.23. The yellow ovals indicate the surfaces where the threads and inserts are joined 

7 with degree-of-freedom couples. This forms a perfect bond between the materials (in all three 

8 directions) at the load-bearing surfaces and permits the other surfaces to separate. The use of 

9 more advanced contact surface behavior was considered to be unnecessary based on the 

10 results obtained using this simpler method. The octagonal shape of the thread inserts is 

11 primarily defined by the thread flanks, which are assumed to have spacing and angles defined 

12 by standard thread characteristics (Oberg et al. 2004). The thread flanks end at the minimum 

13 diameter of the external bolt thread and the maximum diameter of the internal flange thread. 

14 The bolt-side threads and flange-side threads are also assumed to be sized as per standard bolt 

15 threads, so the two sides have a mating geometry rather than an identical geometry. The bolt 

16 threads have a shear area that is proportional to 3/4-pitch. The flange threads have a shear 

17 area that is proportional to 7/8-pitch. This is consistent with the preliminary calculations using 

18 the classical equations for bolt modeling, in which it was assumed that the critical shear area for 

19 the insert occurs at the maximum diameter of the bolt thread and corresponds to 7/8-pitch. 

20
5-34

1

2 Figure 5.23. Detailed FEA Mesh in Thread Insert Region, Showing Bonded Surfaces

3

4

5 This detailed FEA model was used to perform evaluations to determine the effect of the 

6 MacArthur Maze fire scenario on the integrity of containment of the GA-4 package, as 

7 determined by the clamping force on the closure lid throughout the transient. Sensitivity studies 

8 were performed on the effect of assuming linear or non-linear material properties for the thread 

9 inserts. Additional cases were developed, to evaluate the causes of differences between results 

10 obtained with the detailed FEA model and those obtained with the classical bolt equations, to 

11 verify that differences were due to more realistic modeling, and assess the conservatism of the 

12 simpler modeling approach.
6-1

1 6.0 ANALYSIS METHOD

2 The detailed analysis models described in Section 5.0 were developed for evaluation of the GA3 4 package performance if exposed to the severe conditions of the MacArthur Maze fire. For the 

4 thermal analyses, the models account for all significant heat transfer paths to and from the 

5 package by means of conduction, convection, and thermal radiation during the fire and in the 

6 post-fire cooldown. For the structural analysis, the models provide a conservative 

7 representation of the response of the package to the impact of the overhead roadway span that 

8 collapsed at 17 minutes into the fire. The structural analysis also includes evaluation of the 

9 effects of the thermal expansion mismatch between package closure flange and the closure 

10 bolts, and between the impact limiter anchor plates and the impact limiter bolts. The 

11 temperatures and temperature distributions in the bolts and associated package components 

12 are based on thermal model results. All of these various analyses using different analysis 

13 codes and physics regimes taken together provide an overall evaluation of the GA-4 package 

14 response to the Macarthur Maze fire scenario.

15

16 This section presents the initial conditions, modeling assumptions, and boundary conditions 

17 used to predict the response of the GA-4 package to the MacArthur Maze fire scenario. The 

18 first three sections focus on the package thermal model conditions. Section 6.1 presents the 

19 steady-state temperature results for NCT predicted with the ANSYS and COBRA-SFS models 

20 of the GA-4 package, with and without impact limiters, respectively. Section 6.2 describes the 

21 significant assumptions and simplifications used in developing the thermal models. Section 6.3

22 describes the fire scenarios assumed for the transient analyses, and the boundary conditions 

23 used for the calculations. The final sections deal with the assumptions, boundary conditions, 

24 and methodology particular to the structural modeling. Section 6.4 discusses the package 

25 impact analysis.

26

27 6.1 NCT and HAC Fire for GA-4 Package

28

29 Steady-state analyses for NCT with the ANSYS and COBRA-SFS models of the GA-4 package 

30 predict similar peak component temperatures for NCT, within the capabilities of their respective 

31 models for the fuel assemblies. The peak cladding temperature predicted with the ANSYS 

32 model is 306°F (152°C), and the COBRA-SFS model predicts 293°F (146°C). This is a 

33 difference of about 4 percent, and is consistent with the expected differences between the 

34 results obtained with a detailed thermal-hydraulic model of the fuel assemblies compared to the 

35 results obtained with the k-effective model for the fuel. The k-effective model for the 

36 homogenized fuel assembly is designed to yield results that are 5-15% conservative, compared 

37 to results obtained with a detailed CFD model of a fuel assembly (Bahney and Lotz 1996).

38

39 As a verification test of the accuracy of the geometry and material properties input and 

40 boundary condition specification logic, the ANSYS and COBRA-SFS models of the GA-4 

41 package were exercised for the design-basis HAC fire transient defined in 10 CFR 71. This 

42 consists of an engulfing fire with uniform flame temperature of “1475°F (800°C)” and fire 

43 emissivity of 0.9, lasting 30 minutes. In the SAR analysis for the HAC fire, the initial steady44 state assumes an ambient temperature of 100°F (38°C), which is the same as for NCT, but 

45 solar insolation is neglected. The solar heat flux is trivial in comparison to the heat input from 

46 the fire, and as a modeling convenience, it was included in the analyses with the ANSYS and 

47 COBRA-SFS models. The steady-state initial conditions for all transient calculations are NCT; 

48 i.e., assuming standard solar insolation (24-hr average solar heat flux) at an ambient 

49 temperature of 100°F (38°C). 
6-2

1

2 Boundary conditions for the 30-minute HAC fire were specified for the ANSYS and COBRA-SFS 

3 models in essentially the same manner as in the SAR calculation. In general, the results 

4 obtained with the ANSYS and COBRA-SFS models are in very good agreement, with 

5 differences due mainly to the alternative fuel modeling approaches used in the two analyses. In 

6 addition, the omission of the impact limiters from the COBRA-SFS model and their inclusion in 

7 the ANSYS model affect the location of the maximum cladding temperature during the fire and 

8 influence the post-fire cooldown behavior. 

9

10 The ANSYS model predicts 455°F (235°C) for the maximum peak clad temperature during the 

11 transient, and the COBRA-SFS model predicts 469°F (242°C) near the end of the hottest rod, 

12 since this model neglects the impact limiters. In the ANSYS model, the maximum temperature 

13 occurs near the axial center of the package. Both of these values are slightly higher than the 

14 maximum fuel cladding temperature of 442°F (228°C) reported the GA-4 SAR (General Atomics 

15 1998) for the HAC analysis. The difference is due mainly to simplifications in the SAR analysis 

16 that tend to smear out temperature gradients, and thereby reduce the conservatism of the 

17 analysis somewhat. However, for this package, the temperatures are low enough in the HAC 

18 fire that the difference can be considered insignificant. 

19

20 6.2 Thermal Modeling Assumptions

21

22 Computational modeling requires simplifying assumptions for even the most detailed 

23 representation of a physical system. The assumptions used in developing the detailed 

24 geometry models of the GA-4 package are discussed in Section 5.0 above. This section 

25 summarizes the major assumptions relevant to analysis of the response of this package if it 

26 were exposed to the conditions of the MacArthur Maze fire scenario. These assumptions apply 

27 to both the ANSYS and the COBRA-SFS models, unless specifically noted otherwise.

28 1. Initial conditions for the package are defined as steady-state NCT at 100°F (38°C) ambient 

29 with insolation, as defined in 10 CFR 71.71. This assumption conservatively neglects the 

30 effect of the actual conditions at the time of the MacArthur Maze accident (i.e., at night, with 

31 ambient temperature of 50°F [10°C]).

32 2. The decay heat load in the GA-4 package is assumed to be at its maximum design basis 

33 value of 2,105.4 Btu/hr (0.617 kW) per assembly, with a total package decay heat load of 

34 8,423 Btu/hr (2.468 kW). This is a bounding assumption, as the actual decay heat load of 

35 an SNF package is typically lower than the design basis configuration.

36 3. Material properties of package components specified as inputs to the thermal models are 

37 listed in Appendix A. These were obtained from the GA-4 SAR (General Atomics 1998), 

38 with the following exceptions;

39 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 40 . At 

41 NCT, the thermal conductivity values from the SAR are approximately 20% below values 

42 published in material data sheets for XM-19 stainless steels. This is conservative for the 

43 NCT analysis, but is non-conservative for the fire analysis, since the lower bounding 

44 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 steel 1 1 were used in the fire 

2 analyses.

3 b. The thermal conductivity for DU reported in the SAR is for a temperature of 

4 approximately 100°F (68°C), and does not take into account the significant increase in 

5 thermal conductivity with increasing temperature for this material. As with XM-19, this is 

6 a conservative approximation for NCT, but is non-conservative for fire analysis. 

7 Therefore, temperature-dependent thermal conductivity values were used for the DU in 

8 the thermal analyses, as documented in Appendix A.

9 4. Clearance gaps within the package (e.g., between the steel inner liner and the DU gamma 

10 shield, between poison rods and the steel plates of the cruciform basket) are modeled at “as 

11 built” values, based on design drawings.

12 a. Gaps are assumed closed due to thermal expansion during the fire transient, to 

13 conservatively maximize heat transfer into the package.

14 b. Gaps are assumed open, and at nominal “cold” values during the cooldown portion of 

15 the transient, to conservatively limit heat removal from the package.

16 5. The content of the neutron shield tank is conservatively represented to maximize heat 

17 transfer through this region during the fire, and minimize it during the post-fire cooldown.

18 a. Initial steady-state is represented with the effective conductivity model from the SAR, to 

19 account for natural circulation of the neutron shield liquid. This model is used in the fire

20 transient until the peak liquid temperature reaches 276°F (136°C), the saturation 

21 temperature corresponding to the maximum operating pressure for the tank.

22 b. The liquid is assumed lost when the predicted peak temperature in the neutron shield 

23 region exceeds 276°F (136°C). Thermal energy absorbed in the vaporization of the 

24 liquid is conservatively neglected.

25 c. After loss of the liquid, heat transfer between the inner surface of the NS tank outer shell 

26 and outer surface of the package body is assumed to consist of thermal radiation and 

27 conduction through air for the remainder of the fire and post-fire cooldown transient. The 

28 inner surfaces of the tank are assumed to be affected by soot, and the emissivity is 

29 conservatively specified at 0.9. Mainly because of the high thermal radiation heat flux at 

30 the elevated fire temperatures, this results in a higher heat transfer rate into the package 

31 through the neutron shield during the fire than would be achieved with only conduction 

32 and natural convection heat transfer through the neutron shield liquid, if it were assumed 

33 that the neutron shield tank did not rupture during the fire.

34 6. The exterior surface of the neutron shield tank is assumed to have an emissivity of 0.15, as 

35 specified in the SAR, for the initial pre-fire steady-state calculation. At the start of the fire, 

36 the package surface emissivity is set to 0.9, to represent the effect of sooting of the outer 

37 surface of the package and impact limiters. This value is also used throughout the post-fire 

38 cooldown. (This is slightly more conservative than the value of 0.8 to 0.85 documented in 

39 the SAR for the package surfaces in the HAC fire.)

40 7. Convection heat transfer during the fire is conservatively modeled assuming forced 

41 convection to the package from the hot external environment. (See Section 6.3.1 for 

42 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

1 8. For the post-collapse configuration (after 37 minutes), the package is assumed to be fully 

2 covered by the fallen I-580 span between Bent 19 and Bent 20, while simultaneously 

3 exposed to a fully engulfing pool fire with an uniform temperature of 1652°F (900°C). This 

4 modeling approach conservatively places the GA-4 package in the most adverse 

5 configuration possible relative to the collapsed overhead spans, and at the same time yields 

6 a bounding fire exposure for the package.

7 9. Convection heat transfer from the upper surface of the fallen I-580 roadway in the post8 collapse fire scenario is treated as free convection from a horizontal plate with heated 

9 surface facing upward. For the lower surface of the I-880 roadway beneath the package, 

10 heat transfer is treated as free convection from a horizontal plate with heated surface facing 

11 downward. (See Section 6.3.1 for discussion of the specific correlations used.)

12 10. For the post-fire cooldown portion of the transient, the air temperature is assumed to be 

13 100°F (38°C) with insolation, to conservatively bound long-term ambient conditions.

14 11. During the fire, the aluminum impact limiters in the ANSYS model are assumed to remain 

15 intact within their stainless steel outer shells, and are represented with effective thermal 

16 material properties for the honeycomb material, based on bulk density and thermal 

17 conductivity of the component materials. This assumption maximizes heat input to the 

18 package during the fire by conduction through the impact limiters. 

19 12. After the fire, the elements representing the honeycomb material of the impact limiters in the 

20 ANSYS model were modified to account for melting of the aluminum. This assumption 

21 maximizes the thermal resistance to heat removal from the package by conduction through 

22 the impact limiters. Unmelted portions were treated as a combination of aluminum 

23 honeycomb and air, and melted portions were assumed to have the thermal properties of 

24 aluminum alloy 5052. It was also assumed that the molten aluminum would settle to the 

25 bottom of the impact limiters. The effective thermal material properties of the various 

26 elements of the impact limiters affected by melting were calculated using a volume27 averaging scheme. (Section 5.2 discusses this modeling approach in detail.)

28 13. The latent heat absorbed by the honeycomb material in the phase change due to melting 

29 was conservatively neglected. 

30 14. The effect of the conveyance carrying the GA-4 package is conservatively neglected, both 

31 for the thermal analysis of the effects of the fire and the structural analysis of the effects of 

32 the collapse of the upper highway span. In the thermal analysis, the fire is assumed fully 

33 engulfing, and any shielding effect that the conveyance might have is neglected. In the 

34 structural analysis, the conveyance would tend to increase the compliance of the system, 

35 and would also decrease the drop distance between the falling roadway and the GA-4 

36 package. Both factors would mitigate the effects of the impact of the upper roadway falling 

37 on the package. The impact is assumed to occur with the package lying on the lower 

38 roadway, to maximize the drop distance and the overall stiffness of the system.

39

40 6.3 Thermal Boundary Conditions for GA-4 Package Models

41

42 The boundary conditions for the thermal analysis define the external environment that the GA-4 

43 package experiences during the fire and post-fire cooldown. These are specifically defined in 

44 Section 6.3.1. The boundary conditions for the structural analysis define the physical behavior 

45 of the overhead I-580 span that falls onto the SNF package, the temperature of the falling 

46 roadway and girders, and the temperature of the package at the time of impact. These 

47 boundary conditions are discussed in Section 6.4.

48
6-5

1 6.3.1 Thermal Boundary Conditions

2

3 As described in the detailed description of the fire in Section 2.0, the MacArthur Maze fire began 

4 as a large open pool fire that lasted for approximately 17 minutes, until the collapse of the 

5 overhead roadway span between Bent 19 and Bent 20. During the following 20 minutes (17 to 

6 37 minutes of the total fire duration), the partial collapse of the span between Bent 18 and 

7 Bent 19 reduced the size and extent of the large open pool fire to a much smaller fire in the 

8 approximately 12-15 ft (3.7-4.6 m) gap between the collapsed spans at Bent 19. This smaller 

9 fire burned steadily until the end of the fire, at approximately 108 minutes. 

10

11 This fire scenario is conservatively modeled as a fully engulfing fire lasting 108 minutes, with 

12 two different constant bounding fire temperatures, as shown in Figure 6.1. For the first 

13 37 minutes, the fire is represented with a uniform flame temperature of 2012°F (1100°C). For 

14 the remaining 71 minutes, the fire is represented with a uniform flame temperature of 1652°F 

15 (900°C). The fire emissivity is specified at 0.9. This value is also applied to all external 

16 surfaces of the package and roadway exposed to the fire, including the inner surfaces of the 

17 failed neutron shield tank, to conservatively represent the effect of sooting.

18

19

20 Figure 6.1. Fire Boundary Temperatures

21

22 As noted in Section 6.2 in the presentation of modeling assumptions, convection heat transfer at 

23 the SNF package surface during the fire was treated in both models as forced convection. 

24 Along the package body (for both the ANSYS and COBRA-SFS models) and the sides of the 

25 impact limiters (in the ANSYS model), the Nusselt number is defined using a correlation (Kreith 

and Bohn 2001)1 26 for axial flow over a flat or slightly curved surface at zero angle of attack, and 

27 has the form

28

 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 1 ), Pr > 0.5) (6-1)

2

3 where 

4

5 a = 0.036

6 b = 0.8

7 c = -23,200

8 d = 0.3333

9 L = characteristic length (in this case, the exposed package body or impact 

10 limiter side)

11 Pr = Prandtl Number

12 ReL = ρU∞ L / μ

13

14 where 

15

16 ρ = fluid density

17 U∞ = free-stream external velocity

18 μ = fluid viscosity

19

20 On the package ends (in the COBRA-SFS model), and on the flat ends of the impact limiters (in 

21 the ANSYS model), the Nusselt number is defined using a correlation for forced convection from 

22 an isothermal disk or circular plate with axis perpendicular to the flow direction (Kreith and Bohn 

23 2001). The correlation has the form

24

b c

D 25 Nu = a Re Pr (900 < ReD < 30,000) (6-2)

26

27 where 

28

29 a = 0.591

30 b = 0.564

31 c = 0.3333

32 Pr = Prandtl Number

33 ReD = ρU∞D / μ

34

35 where 

36

37 ρ = fluid density

38 U∞ = free-stream external velocity

39 D = diameter of disk or plate

40 μ = fluid viscosity

41

42 In the above correlations, fluid properties are evaluated at the near-wall temperature, defined as 

43 the average of the wall surface temperature and the ambient temperature, which in this 

44 application is the fire temperature. The free-stream external velocity was specified at a 

45 bounding value of 12 ft/s (3.7 m/s), based on the velocities predicted in the FDS simulation of 

46 the open pool fire.

47
6-7

1 In the post-fire cooldown, with the package beneath the concrete “blanket” of the collapsed 

2 upper roadway span, the boundary conditions on the external roadway surfaces were specified 

3 as free convection to ambient still air. For the COBRA-SFS model, the boundary conditions on 

4 the SNF package were determined by solving for the flow and heat transfer within the irregular 

5 annulus between the package and the concrete roadway surfaces. The air velocity within the 

6 annulus was assumed to be essentially stagnant in the post-fire cooldown. 

7

8 In typical open pool fires, it is conservative to assume free convection to ambient for the post9 fire cooldown when evaluating heat transfer from an object such as an SNF package which had 

10 been engulfed in the fire. For a horizontal cylinder, correlations typically yield Nusselt number 

11 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 12 -

13 °K). However, for the package beneath the concrete roadway, this assumption is neither 

14 conservative nor correct, since the presence of the roadway would severely inhibit the 

15 development of natural convection flow patterns around the package. 

16

17 In the COBRA-SFS analysis, convection heat transfer in the annulus channel between the 

18 package outer surface and interior surfaces of the concrete was represented with a bounding 

19 value of Nu = 3.66 for laminar flow, based on experimental studies (Creer et al. 1987) of heat 

20 transfer in enclosed spaces within a horizontal package. This approach conservatively treats 

21 the “blanket” of the collapsed roadway as an enclosure. After the end of the fire, the hot air in 

22 the irregular annular flow channel is assumed to be optically transparent, allowing thermal 

23 radiation heat transfer via gray body exchange factors between the package surface and interior 

24 surfaces of the overlying upper roadway, side barrier, and lower roadway. The temperature of 

25 air trickling into the annulus due to the changing temperature and pressure gradients in the flow 

26 channel is assumed to be 100°F (38°C). 

27

28 In the analysis with the ANSYS model, thermal radiation exchange between the package and 

29 the enclosing concrete was calculated directly, as described in Section 5.2.2 above. However, it 

30 was assumed that the impact limiters would effectively block air flow within the annulus, so 

31 convection heat transfer between the package exterior surfaces and the enclosing concrete 

32 surfaces was conservatively neglected. This was modeled by specifying Nu = 1.

33

34 The temperatures of the concrete surfaces surrounding the package at the beginning of the 

35 cooldown portion of the transient were determined assuming that these structures had also 

36 been exposed to the same fire boundary temperatures of 2012°F (1100°C) and 1652°F (900°C), 

37 for the same time intervals as the SNF package. This approach establishes bounding concrete 

38 surface temperatures for the post-fire cooldown portion of the transient. The effect of spalling of 

39 the concrete, which dissipates thermal energy and therefore would tend to result in lower 

40 concrete temperatures, was conservatively neglected. 

41

42 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 43

44 and thermal radiation to ambient. For the upper surface of the fallen upper roadway span, the 

45 correlation used is for free convection from a horizontal heated surface facing upward, and is of 

46 the form

47

b Nu = aRa L 48 (6-3)

 1 Source reference is Kreith; primary references are MacAdams, plus Incropera and DeWitt.
6-8

1

2 where

3

4 RaL = Rayleigh Number (GrLPr)

5

6 where 

7

8 L = (surface area)/perimeter

9

For laminar flow (105 < RaL < 107 10 ), the coefficients are

11

12 a = 0.54

13 b = 0.25

14

For turbulent flow (107 < RaL < 1012 15 ), the coefficients are

16

17 a = 0.15

18 b = 0.3333

19

20 For the lower surface of the lower roadway, the correlation is of the same form, but with 

21 coefficients for free convection from a horizontal heated surface facing downward

22

b Nu = aRa L (105 < RaL < 1010 23 ) (6-4)

24

25 where 

26

27 a = 0.27

28 b = 0.25

29

30 In the above correlations, fluid properties are evaluated at the near-wall temperature, defined as 

31 the average of the wall surface temperature and the ambient temperature. The ambient 

32 temperature external to the fire is conservatively assumed to be 100°F (38°C), as a convenient 

33 bounding value.

34

35 6.4 Structural Impact Model Assumptions and Analysis Method

36

37 Because of the complexity of the impact portion of the MacArthur Maze fire scenario, the 

38 structural model of the fallen overhead span was developed in progressive stages, starting with 

39 a simple approximation using only a truncated section of a single plate girder for scoping 

40 calculations, and evolved to include representation of the complete overpass section. The 

41 purpose of the preliminary structural modeling was to determine the level of complexity and 

42 sophistication in the final set of impact models needed to assure that all potentially significant 

43 structural consequences to the containment boundary were considered. The initial modeling 

44 work included mesh sensitivity studies and evaluation of the effects of including or neglecting 

45 the conveyance. A primary conclusion of the initial modeling phase was that the thin plate 

46 girders would tend to deform around the stronger package wall instead of causing significant 

47 local deformation to the package, making a failure of the containment boundary unlikely. The 

48 more realistic final impact models described in this report were carefully crafted to include the 

49 upper bound of potential package response without resorting to incredible conservatism. The 

50 basic features of the structural models are described in Section 5.4.1.
6-9

1 Fire exposure raises the package temperature, which lowers the structural material strength and 

2 makes the package more susceptible to damage from impact loading. Therefore, a 

3 conservative estimate of the maximum package temperature was obtained from the thermal 

4 analyses, to appropriately bound the package temperature at the time of the collapse of the 

5 overhead span. As described in Section 2.0, the collapse of the first span occurred at 

6 approximately 17 minutes, and the partial collapse of the second span was complete at about 

7 37 minutes after the start of the fire. The thermal analysis (see Section 6.3) conservatively 

8 assumed that the large fully engulfing (pre-collapse) pool fire lasted for the full 37 minutes. For 

9 the purposes of the structural analysis, the peak temperature on the package surface predicted 

10 at 37 minutes was used to define the overall package temperature, conservatively assuming 

11 that the entire package was at the peak temperature. This approach results in an estimated 

12 bounding temperature of 1800°F (982°C) for the package in the structural analyses. 

13

14 The results of the FDS model of the pre-collapse fire (see Section 3.0) and an additional 

15 ANSYS thermal model of the I-580 roadway were used to determine a reasonably conservative 

16 uniform temperature for the falling steel girders of the overpass structure. In the case of the 

17 falling girders, the conservative condition errs on the cool side, so the strength and stiffness in 

18 the girders is as high as possible, within the thermal constraints of the scenario. In this case, 

19 1800°F (982°C) was chosen as the representative girder temperature for the drop scenario. 

20 Coincidentally, this value is the same as the assumed package temperature, which was chosen 

21 using a different rationale, as described in Section 6.4.1. The package temperature was chosen 

22 to be bounding with as high a value as possible; the girder temperature was chosen to be as 

23 conservatively low as possible, based on realistic package locations. 

24

25 The temperature of the lower I-880 span at the location of the GA-4 package is assumed to be a 

26 nominal 80°F (27°C). This conservatively neglects elevated temperature effects on the material 

27 properties of the lower roadway. The roadway would have been much hotter during the fire, 

28 and would have been more compliant at a higher temperature than the assumed bounding 

29 value. Assuming higher, more realistic temperatures for the lower roadway would allow it to 

30 dissipate more of the energy of the impact, mitigating to some degree the effect of impact on the 

31 package. Similarly, air drag forces and reduction in fall distance resulting from sagging of the 

32 girders before collapse of the span are conservatively neglected in the impact analysis. 

33

34 Yield strength and the plastic behavior of the materials are critical parameters in the impact 

35 model. The evaluation of the package containment is based on the amount of plastic strain 

36 developed in the XM-19 stainless steel package body. The significant capacity of the plate 

37 girders to deform during impact because of the elevated temperature, visible in the post-fire 

38 photographs of the girders (CHP 2007), is a fundamental physical phenomenon captured in the 

39 analysis with this model. 

40

41 The DU gamma shield is not a structural component of the containment boundary, but it does 

42 offer potential support to the steel body wall. The DU is modeled as a bilinear (elastic/plastic) 

43 material with a near-zero tangent modulus. If the yield stress limit is exceeded, the material is 

44 allowed to deform with a negligible increase in load. This behavior results in the DU layer 

45 providing realistic support to the package as long as it remains in the elastic range, but the 

46 support drops off quickly if the stress in the material exceeds the elastic range. The tangent 

47 modulus is six orders of magnitude lower than the elastic modulus for this material, resulting in 

48 calculated plastic strains that are conservatively higher than they would be if a realistic stress49 strain curve were implemented. However, the response of the DU gamma shield is not 

50 specifically evaluated for failure in this analysis, since it is not critical to the containment 

51 boundary.
6-10

1 In the calculations, contact between a falling girder and the package body or any other 

2 component of the model was defined with an automatic single surface contact definition. This 

3 contact rule prohibits the components from occupying the same space, and applies a coefficient 

4 of static and dynamic friction of 0.5. Dry steel-to-steel friction is typically reported in the range of 

5 0.7-0.8 for static friction and 0.4-0.6 for dynamic friction, but there is significant variation and 

6 instances of data reported outside those ranges. The 0.5 was chosen as a middle range 

7 dynamic friction value, based on sensitivity studies that showed friction was not a major force in 

8 this scenario. The relative fraction of frictional energy to total system energy was in the range of 

9 1-2%, while the fraction of frictional energy to internal energy (due to material deformation) was 

10 in the range of 1-5%.

11

12 Global mass damping of 5% was applied to the model as an additional source of energy 

13 dissipation. This level of damping is typically used in dynamic impact modeling for numerical 

14 stability and to account for natural sources of energy dissipation that are not specifically 

15 modeled. The amount of energy this damping dissipates in the model is comparable to the 

16 amount of energy dissipated by friction. Global damping does not have a significant effect on 

17 the results obtained in this analysis.

18

19 6.4.1 Overpass Temperature Evaluation

20

21 The temperature of the I-580 overpass at the time of impact is an important assumption that 

22 affects its material characteristics and the overall impact response. The elastic modulus and the 

23 yield strength of the girder material are two material properties that affect the force and energy 

24 transmission during impact, and both material properties decrease at elevated temperatures. 

25 As discussed in Section 2.2, evidence gathered from the post-collapse material indicated that 

26 the plate girders experienced temperatures at least as high as 1472-1652°F (800-900°C) at 

27 locations near the middle of the fire. Material data in the literature for this temperature range 

28 predict a reduction in the yield strength of steel to values roughly 10% of the room-temperature 

29 yield strength (see Figure 2.10), and a substantial increase in ductility. 

30

31 To obtain a more detailed picture of the temperature distribution on the girders than could be 

32 determined from the relatively small range of sampling locations, additional analysis was 

33 performed with an ANSYS model of the I-580 upper roadway span between Bents 18 and 20. 

34 ASTs from the FDS model of the pre-collapse fire (discussed in Section 3.0), were used as 

35 boundary conditions for thermal analysis of the girders. The AST boundary values were applied 

36 as time-history thermal loads over the interval from initiation of the fire to 18 minutes, 

37 encompassing the time frame of the complete collapse of the overhead span between Bent 19 

38 and Bent 20. The lower temperature corresponding to the actual time-frame of the collapse of 

39 this span is more conservative for the purposes of the structural analysis. However, a higher 

40 temperature is more conservative for the package temperature, and the package temperature 

41 for the impact analysis was based on a fire exposure time of 37 minutes, corresponding to the 

42 time of final collapse of both overhead spans.

43

44 The thermal model of the upper spans predicts a rapid heating of the steel girders directly above 

45 the fire. By 5 minutes into the fire, the peak girder temperature exceeds 2192°F (1200°C). At 

46 18 minutes, the temperature distribution reaches the state shown in Figure 6.2, with peak 

47 temperatures up to 2462°F (1350°C). The left half of the thermograph shows the hotter side of 

48 the overpass, which is directly above most of the fire area, beneath the span between Bents 18 

49 and 19. The right side is generally cooler, and represents the span that fell straight down and is 

50 modeled in the impact analyses. This is consistent with the physical distribution of the fire pool 

51 (as shown in the fire video images in Figure 2.2, Figure 2.3, and Figure 2.4). Over much of the 
6-11

1 girder length exposed to the fire, the ANSYS results show girder temperatures that are in the 

2 range determined from the material analysis. (This corresponds to the light green to yellow 

3 segments of the color thermograph in Figure 6.2). However, this calculation also predicts much 

4 higher temperatures for the girders in locations that were not sampled for the material analysis. 

5

6

7 Figure 6.2. Girder Temperatures, (°C [°F]), at 18 Minutes Predicted with ANSYS Model

8

9 Sensitivity studies on thermal processes that would carry heat away from the girders, such as 

10 conduction heat transfer longitudinally down the girders and conduction into the concrete above 

11 the girders, showed that these processes do not play a major role in determining the girder 

12 temperature or temperature distributions. The local heat input to the girders due to the fire, as 

13 defined in the ANSYS model using the AST values from the FDS fire model, dominates all other 

14 thermal processes, even when conservatively high heat dissipation through the girders is 

15 assumed, by such means as enhanced convection coefficients between the upper roadway 

16 surface and the cool air above it. 

17

18 As illustrated in Figure 6.2, the peak girder temperatures are predicted to be as high as 2462°F 

19 (1350°C). This very high temperature zone includes girder segments above potential impact 

20 locations for the SNF package. However, from the standpoint of the drop scenario, higher 

21 girder temperatures would result in a less severe impact, due to the greatly decreased yield 

22 strength of steel with increasing temperature. Using these very high temperatures to define the 

23 girder temperatures for the impact analysis would result in a less severe drop scenario than 

24 would be obtained by assuming lower temperatures for the girders. The goal in assigning a 

25 temperature to the girder structure was to ensure the temperature was conservatively low, 

26 giving the structural steel of the girders conservatively high strength and stiffness, but remaining 

27 consistent with the range of temperatures that could reasonably be expected from the accident 

28 conditions. 

29

30 A uniform overpass girder temperature of 1800°F (982°C) was chosen as a conservatively low 

31 temperature estimate for the hot side of Bent 19. The geometry of overhead roadway for the 

32 impact models actually corresponds to the right half of Figure 6.2, which is the cooler side, 
6-12

1 because that is the span that fell relatively cleanly onto the lower roadway in the MacArthur 

2 Maze accident (i.e., the span between Bents 19 and 20), and offers the potentially most 

3 damaging mechanical loading. The thermal state of the impact model is defined based on 

4 conditions from the hot side, to ensure the package is at its hottest and structurally weakest.

5
7-1

1 7.0 ANALYSIS RESULTS

2 This section presents the results of the structural and thermal analyses of the GA-4 package 

3 exposed to the conditions of the MacArthur Maze fire scenario. The thermal analysis of the fire 

4 portion of the transient is presented in Section 7.1, for both the ANSYS model and the COBRA5 SFS model. The thermal analysis of the post-fire cooldown transient with these models, in 

6 which the package is covered by the concrete “blanket” of the collapsed upper roadway, is 

7 presented in Section 7.2. The results of the structural analysis are discussed in Section 7.3.

8

9 7.1 GA-4 Package: Thermal Results for Fire Transient

10

11 The starting point for the fire transient thermal modeling was assumed to be steady-state NCT. 

12 The initial temperature distribution in the package is summarized in Figure 7.1, which shows the 

13 peak component temperatures predicted with the ANSYS and COBRA-SFS models of the GA-4 

14 at NCT. Figure 7.2 shows a color thermograph illustrating the ANSYS model temperature 

15 results for the package cross-section for this initial steady state evaluation. As discussed in 

16 Section 6.1 above, the peak fuel region temperature predicted with the ANSYS model is 306°F 

17 (152°C), and the peak cladding temperature predicted with the COBRA-SFS model is 293°F 

18 (146°C). This is consistent with the expected differences between the results obtained with a 

19 homogeneous representation of the fuel region using a k-effective model when compared to a 

20 detailed thermal-hydraulic model of the fuel assemblies. The two models give essentially the 

21 same results for NCT, within modeling uncertainty, and are in reasonable agreement with the 

22 results reported in the GA-4 SAR (General Atomics 1998).

23

24

25 Figure 7.1. Initial Conditions for Fire Transient Analyses: GA-4 at NCT

0

50

100

150

200

250

300

350

Peak Fuel

Clad

basket inner steel

liner

DU Gamma

Shield

Body Wall Neutron

Shield

Outer Shell ambient

Peak Component Temperature (°F)

GA-4 SAR

ANSYS model

COBRA-SFS model (without impact limiters)
7-2

1

Figure 7.2. ANSYS Results1 2 : Thermal Cross-section (°F) of GA-4 Package at NCT

3

4 The MacArthur Maze fire was simulated for the GA-4 package by imposing boundary conditions 

5 representing the large (pre-collapse) engulfing fire, followed by the smaller (post-collapse) 

6 engulfing fire. To simulate the pre-collapse fire, the package model was subjected to an 

7 ambient boundary temperature of 2012°F (1100°C) for 37 minutes. The detailed results for the 

8 fully engulfing fire at 2012°F (1100°C) are described in Section 7.1.1. To simulate the post9 collapse fire, the fire boundary temperature was reduced to 1652°F (900°C) for the remaining 

10 71 minutes of the fire portion of the transient. The results obtained for the post-collapse portion 

11 of the fire are described in Section 7.1.2. 

12

13 7.1.1 Pre-collapse Fire (2012°F [1100°C])

14

15 The initial fully engulfing pool fire of the MacArthur Maze fire scenario at 2012°F (1100°C) was 

16 assumed to extend to 37 minutes of the total fire duration, to conservatively encompass the 

17 collapse of both of the upper roadway segments. Figure 7.3 shows a color thermograph of the 

18 package central cross-section illustrating the temperature results obtained with the ANSYS 

19 model at the end of this portion of the fire. (Note that the TIME reported on the plots includes 

20 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

1

2

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

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

1 Peak temperatures predicted with the ANSYS model for the main components of the GA-4 

2 package are shown as a function of time in Figure 7.4. The peak temperature on the outer 

3 surface of the package, which consists of the thin stainless steel shell of the neutron shield tank, 

4 rises rapidly in the initial minutes of the fire, paralleling the rapid rise in the exterior ambient 

5 temperature. Within approximately 5 minutes, the surface temperature of the package is 

6 approaching the fire temperature on an asymptotic curve. In a similar manner, the peak 

7 temperature on the outer surface of the impact limiters very rapidly approaches the fire 

8 boundary temperature. This temperature occurs on the thin stainless steel outer surface of the 

9 impact limiter, which is effectively insulated on the inner side by the relatively low thermal 

10 conductivity of the honeycomb structure that the steel is there to protect. The temperature of 

11 the outer stainless steel surface therefore approximates an adiabatic surface temperature at the 

12 local fire boundary condition.

13

14

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

16 Engulfing 2012°F (1100°C) Fire

17

18 The peak fuel region temperature does not begin to show any increase until about 9 minutes 

19 into the fire transient. Such a delay is typical of the peak fuel temperature response to a fire 

20 transient, because of the thermal inertia of the massive body of an SNF package. In addition, 

21 the location of the maximum temperature in the fuel region shifts toward the outer edge of the 

22 hottest fuel assembly as the fire transient progresses, and part of the delay is due to the time 

23 required to heat up the region of the fuel assembly that is initially cooler than the central region 

24 in the pre-fire steady state at NCT. 

25

26 Peak temperatures predicted with the COBRA-SFS model for the major components of the GA27 4 package in response to the fire conditions are shown in Figure 7.5. The plots in Figure 7.4

28 and Figure 7.5 show that the two models are predicting similar responses for the GA-4 package 

29 in this severe fire transient. The significant difference between them is in the response of the 

30 fuel region. The ANSYS model uses the homogeneous k-effective model for the fuel region, 

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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

1 which is a steady-state model that is by design a conservative representation of heat transfer 

2 within the rod array. The more detailed rod-and-subchannel representation of the fuel assembly 

3 in the COBRA-SFS model results in a more realistic prediction of the rate of fuel cladding 

4 temperature rise. 

5

6

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

8 Fully Engulfing 2012°F (1100°C) Fire

9

10 The COBRA-SFS model predicts that the peak cladding temperature near the ends of the rods, 

11 which in this model are not protected by the impact limiters, rises to 955°F (513°C) by 

12 37 minutes. The COBRA-SFS model predicts 677°F (358°C) by this time for the peak cladding 

13 temperature at the midplane of the package, the location where the ANSYS model predicts the 

14 peak temperature in the fuel region. Because of the more conservative homogeneous k15 effective model for the fuel region, the ANSYS model predicts a somewhat higher peak 

16 temperature of 1023°F (551°C) by 37 minutes. The cladding peak temperatures predicted with 

17 both of these models have not exceeded the limit of 1058°F (570°C) for short-term operations 

18 by the end of the 37-minute bounding representation of the pre-collapse fire. However, the total 

19 fire duration has 71 more minutes to run, at the lower fire temperature of 1652°F (900°C), 

20 simulating the post-collapse portion of the fire. 

21

22 7.1.2 Post-Collapse Fire (1352°F [900°C])

23

24 The post-collapse portion of the MacArthur Maze fire, in which the physical extent of the fire was 

25 greatly reduced by the intrusion of the fallen upper roadway spans, was modeled as a fully 

26 engulfing fire at the lower boundary temperature of 1652°F (900°C). The temperatures 

27 predicted with the ANSYS model for the end of the fire, after 37 minutes of exposure to 2012°F 

28 (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

1

2

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

4 GA-4 Package at the End of the Fire at 108 Minutes

5
7-7

1 Peak temperatures predicted with the ANSYS model for the main components of the GA-4 

2 package are shown as a function of time in Figure 7.7. Figure 7.8 shows the response of major 

3 components of the package to the fully engulfing post-collapse fire predicted with the COBRA4 SFS model. By the end of the 37-minute fire at 2012°F (1100°C), both models predict peak 

5 temperatures on the outer surface of the package that are significantly above the post-collapse 

6 fire temperature of 1652°F (900°C). As a result, the initial response of these components during 

7 the post-collapse fire is to cool down toward the new ambient fire temperature. In the first 

8 13 minutes of this portion of the transient, the peak temperatures on the outboard components 

9 drop rapidly in response to the sudden change in the boundary fire temperature, then stabilize 

10 to a slow asymptotic rise toward the new fire temperature.

11

12

13 Figure 7.7. Peak Component Temperatures Predicted with ANSYS Model at 108 Minutes

14

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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

1

2 Figure 7.8. Peak Component Temperatures Predicted with COBRA-SFS Model at 108 Minutes

3

4 Similarly, the peak fuel temperatures respond to the lower boundary temperature by slightly 

5 slowing their rate of rise, but these components are also on an asymptote approaching the new 

6 fire temperature. By the end of the fire transient at 108 minutes, the peak fuel region 

7 temperature predicted with the ANSYS model has reached 1433°F (779°C), exceeding 1382°F 

8 (750°C), the temperature at which burst rupture of zircaloy cladding has been assumed in 

9 previous SNF package transportation studies (NUREG/CR-6672). The peak fuel cladding 

10 temperature predicted with the COBRA-SFS model also exceeds this temperature, reaching 

11 1388°F (753°C) on the end of the hottest fuel rod. The mid-plane peak fuel cladding 

12 temperature predicted with this model is not far behind, at 1248°F (675°C). 

13

14 These predicted peak temperatures are in response to a conservative and bounding 

15 representation of the MacArthur Maze fire. However, even after the fire is over, the transient is 

16 not finished, as the package is evaluated in the post-fire cooldown, with the conservative 

17 assumption that it is blanketed by the concrete structure of the fallen overhead roadway span. 

18 Experience with modeling of SNF packages in long-duration fires (NUREG/CR-6487; 

19 NUREG/CR-6886 2009) has shown that the maximum fuel cladding temperature can occur well 

20 after the end of the fire, during the post-fire cooldown of the package. In addition to the rise in 

21 temperature on the fuel rods in response to heat input from the fire, some portion of the 

22 temperature rise is due to the high ambient fire temperature preventing decay heat removal 

23 from the fuel rods during the fire and for some time after the fire while the outboard components 

24 of the package are above the maximum fuel temperature.

25

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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

1 7.2 GA-4 Package: Thermal Results for Post-Fire Cooldown 

2 Transient

3

4 For the cooldown portion of the transient, in which the GA-4 package is enclosed by the fallen 

5 roadway, the roadway surface temperatures at the end of the fire were determined by a 

6 separate calculation with COBRA-SFS. This calculation represented the roadway as described 

7 in Section 5.2.2 above (see Figure 5.7), exposing the roadway surfaces to the same fire 

8 conditions as the GA-4 package. Figure 7.9 shows the peak temperatures predicted for the 

9 roadway concrete surfaces in response to the pre-collapse fire at 2012°F (1100°C) and the 

10 post-collapse fire at 1652°F (900°C). These temperatures exhibit the same asymptotic behavior 

11 as the outer surface of the SNF package. 

12

13

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

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

16

17 In Figure 7.9, the upper roadway temperature is the peak temperature on the under surface of 

18 the upper roadway span that is assumed to cover the SNF package in the post-fire scenario. 

19 The lower roadway temperature shown in the plot is the peak temperature predicted on the 

20 surface of the lower roadway beneath the SNF package. The concrete barrier temperature is 

21 the peak temperature predicted for the inner surface of the lower roadway side barrier, which 

22 the SNF package is assumed to be lying beside. The peak temperature of 1774°F (968°C) 

23 predicted for the roadway surfaces occurs at the end of the 37-minute fully engulfing fire. The 

24 side barrier is predicted to reach approximately 1369°F (743°C) at this time. The temperatures 

25 on the roadway surfaces drop somewhat during the 71-minute interval of exposure to the lower 

26 temperature of the post-collapse fire, and by the end of the fire, all of the concrete surfaces 

27 exposed to the fire are at approximately the same temperature of 1609°F (876°C).

28

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end of fire

MacArthur Maze fire scenario
7-10

1 The results in Figure 7.9 show that it is extremely conservative to assume that the concrete 

2 structures would be exposed to the bounding fire temperatures for the full duration of the pre3 collapse and post-collapse fires. As with all other components exposed to these extreme 

4 boundary conditions, the concrete temperatures begin to asymptotically approach the fire 

5 temperature. Post-fire evaluation of the damaged roadway, including detailed assessment of 

6 the lower roadway deck and side barriers (Wiss, Janney, Elstner Associates, Inc. 2007), found 

7 that the concrete surfaces were exposed to temperatures in the range 446°F (230°C) to 842°F 

8 (450°C). The maximum temperature exposure at the concrete surfaces was estimated as 

9 1472°F (800°C), over a relatively small localized area. The bounding approach used in this 

10 model sets up an extremely conservative environment for the SNF package in the post-fire 

11 cooldown transient beneath the collapsed roadway. 

12

13 The roadway temperatures at the end of the fire define the initial conditions for the cooldown 

14 transient calculated with the ANSYS and COBRA-SFS models for the GA-4 package beneath 

15 the concrete ‘blanket’ of the upper roadway. Figure 7.10 shows the response predicted with the 

16 COBRA-SFS model for the major components of the system in the post-fire cooldown in the first 

17 5 hours after the end of the fire. Figure 7.11 shows the response predicted with the ANSYS 

18 model for this portion of the transient. In both models, all major components show an initial 

19 rapid decrease in peak temperature, followed by a more gradual decline as the temperature 

20 gradients flatten. 

21

22 The results for the COBRA-SFS model show the effects of the rapid cooling of the ends of the 

23 package, where the peak temperatures are predicted during the fire transient. By about 

24 2.3 hours, the ends have cooled sufficiently for the location of the peak temperatures to shift

25 back toward the axial center of the package, and the effects of thermal inertia are seen in the 

26 peak fuel cladding temperature and the peak basket temperature. For these components, the 

27 temperatures near the center of the package continue to rise after the end of the fire, and 

28 eventually exceed the peak temperatures on the ends. This is illustrated in Figure 7.12, which 

29 shows the evolution of the axial temperature profile on the hottest rod in the package, as 

30 predicted with the COBRA-SFS model.

31
7-11

1

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

3 Cooldown to 6.5 Hours

4

5

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

7 Cooldown to 6.5 Hours

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fuel cladding basket stainless steel liner

DU gamma shield stainless steel body neutron shield contents

outer surface end of fire short-term PCT limit

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7-12

1

2 Figure 7.12. Axial Temperature Evolution on Hottest Fuel Rod Predicted with the COBRA3 SFS Model for Post-fire Cooldown with the Package under Concrete

4

5 The peak fuel cladding temperature near the axial center of the hottest fuel rod is predicted to 

6 reach a maximum of 1347°F (731°C) at approximately 250 minutes, which is nearly 2.4 hours 

7 after the end of the fire. The results obtained with the ANSYS model also show the effect of 

8 thermal inertia for the fuel, sustaining a peak fuel region temperature of nearly 1400°F (760°C) 

9 for approximately 3 hours after the end of the fire, as shown in Figure 7.11. 

10

11 Because the ANSYS model includes the impact limiters, the ends of the package do not 

12 experience the high temperatures predicted with the COBRA-SFS model, which omits the 

13 impact limiters. However, in the post-fire cooldown, the impact limiters act as insulators, 

14 slowing the rate of heat removal from the ends of the package. As the package cools, the steep 

15 temperature gradients within the package, due to the heat input from the fire and the decay heat 

16 trapped within the fuel region, result in the cooler ends of the package continuing to increase in 

17 temperature for some time after the fire, even as the peak temperatures near the center 

18 decrease. This is illustrated in Figure 7.13, with color thermographs of the package axial cross19 section. In these thermographs, the scale of the color coding is held constant, to clearly 

20 illustrate the increasing temperature on the ends and interior components, while the peak 

21 temperatures steadily decrease. 

22

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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

1

 (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

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

3 during Post-fire Cooldown

4
7-14

1 As shown in Figure 7.13, at the end of the fire the temperatures at the ends of the fuel region 

2 (beneath the impact limiters) are relatively cool, but continue to increase in temperature for 

3 many hours after the fire is over. The temperature increase in the end region of the fuel is 

4 summarized below.

5

6

7 By 12.2 hours, the impact limiters and outer shell of the package are at temperatures in the 

8 range 1034°F to 1122°F (557°C to 606°C), slightly cooler than the end of the fuel region. The

9 package begins to experience a uniformly decreasing temperature at all points, including the 

10 sheltered locations within the package beneath the impact limiters, only after about 12.5 hours

11 beyond the end of the fire. 

12

13 This behavior is more precisely illustrated with the temperature plot in Figure 7.14 for three 

14 specific locations beneath the top impact limiter in the ANSYS model of the package. This 

15 figure shows the temperature histories during the cooldown for a point in the fuel region near the 

16 lid end of the package, a point in the top impact limiter immediately adjacent to the package lid, 

17 and a point near the upper end of the DU gamma shield.

18

19

20 Figure 7.14. ANSYS Model Predictions of Temperature History during the Post-fire Cooldown 

21 at Locations in the GA-4 Package Shielded by the Top Impact Limiter

22

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Impact limiter, near lid surface

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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

1 The overall results of the thermal analysis with the ANSYS and COBRA-SFS models show that 

2 a fire of the severity of the MacArthur Maze fire scenario would result in extremely high 

3 component temperatures on an SNF package such as the GA-4, and that those temperatures 

4 would be sustained over a long period of time. In particular, fuel cladding temperatures are 

5 predicted to exceed the short-term limit of 1058°F (570°C) for several hours, over essentially the 

6 full axial length of the fuel assemblies. In addition, the temperatures in the regions of the 

7 package seals exceed the seal material operating temperature limits for most of the fire 

8 transient and for many hours of the post-fire cooldown transient. The potential consequences of 

9 this extended period at elevated temperatures are discussed in Section 8.0.

10

11 7.3 GA-4 Package: Structural Evaluation

12

13 The positioning of the SNF package on the lower roadway has a significant effect on the 

14 potential consequences of the overpass collapse. It affects the freefall distance for the 

15 overhead span, and the location(s) at which the girder(s) can strike the package. In addition, 

16 the alignment of the center of gravity (CG) of the package and the CG of the falling span affect 

17 the total energy of impact imparted to the package. All other things being equal, the more 

18 closely the two CGs are to the same vertical line, the greater the amount of energy transmitted 

19 by the impact. 

20

21 It is not obvious what impact orientation could do the most damage to the package, and in this 

22 analysis, four different configurations are considered, to cover a range of possible impact 

23 behavior and to explore the effects of the most significant variables. In all cases considered, the 

24 package location on the lower roadway is as close as possible to the center of gravity of the 

25 falling span, within the constraints of the geometry of the lower and upper roadway structure. 

26 The cases evaluated are summarized as follows:

27

28 • Case #1: a girder strikes directly across the center of the package (Figure 7.15)

29

30 • Case #2: a girder strikes along the axial length of the package (Figure 7.16; the package is 

31 barely visible beneath the third girder from the right)

32

33 • Case #3: a girder strikes directly on the package lid (Figure 7.17)

34

35 • Case #4: a girder strikes directly onto the package lifting trunnions, which are assumed to be 

36 oriented vertically, such that the impact applies a highly localized load to the package wall 

37 (Figure 7.18).

38
7-16

1

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

3 Impact on Package Center

4

5

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

7 along Axial Length of Package

8
7-17

1

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

3

4

5 Figure 7.18. Impact Case #4: GA-4 Package Oriented for Trunnion Impact

6

7 The results of the analysis for all four of these cases show that the steel plate girders of the 

8 overhead roadway are more severely affected by the impact than the SNF package. The 

9 calculations show that the impact would cause significant plastic strains in the girders, such that 

10 the metal plates would deform under the impact, while the SNF package is relatively unaffected 
7-18

1 by the impact force. Localized plastic strains are predicted in the package wall and the DU 

2 gamma shield, but these are much smaller than the strains predicted for the girders. 

3

4 For the conditions of this scenario, it is expected that failure mode for the steel of the girders 

5 would be ductile fracture. Uniaxial tensile test data for A36 steel suggest that for conditions in 

6 the temperature range of the MacArthur Maze fire scenario, the elongation capacity for this 

7 material is over 50% and the reduction of area is over 95%. With this amount of available 

8 ductility, the calculated effective plastic strains below 20% appear to be well within the material’s 

9 capacity. The girders would be severely deformed, but would be unlikely to fracture. This is 

10 illustrated conceptually in Figure 7.19, comparing a photo of a damaged girder from the actual 

11 roadway collapse following the fire to a graphic image from the LS-DYNA calculation for 

12 Case #1.

13

14 

15 Figure 7.19. Illustration of Actual and Calculated Failure of Overhead Roadway Girders

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

17

18 The maximum peak effective plastic strain in the XM-19 stainless steel of the package wall is 

19 calculated as 11 percent, and occurs in Case #2. An industry source lists short-term elevated 

20 temperature elongation data (Nitronic 50 Product Data Bulletin) for this material as having

21 maximum elongation values between 41% and 59%t at 1500°F (816°C). Based on this data, 

22 the local plastic strains are not sufficient to cause structural failure of the steel components of 

23 the GA-4 package. In each impact orientation, the plastic strain regions are further evaluated 

24 based on relative size and penetration into the volume of the material or through the wall. In all 

25 cases, the localized plastic strains represent surface damage, not gross deformation of the steel 

26 package body.

27

28 The DU gamma shield is included in the model to give the steel package body realistic support. 

29 It is not considered a structural member in the GA-4 design, but the package SAR (General 

30 Atomics 1998, Section 2.1.2.2.2) notes that the DU may transfer compressive loads and provide 

31 backing for the containment wall. It is modeled to perform this function in a conservative 

32 manner. Since the DU is not a part of the containment boundary, it is not specifically evaluated 

33 for failure. Plastic deformations in the DU are reported and plotted to show the locations where 

34 the containment wall required maximum support from the DU structure. 

35
7-19

1 7.3.1 Structural Case #1: Package Perpendicular to Upper Roadway Girders

2

3 The axial length of the GA-4 package and the spacing of the I-580 girders permit direct 

4 perpendicular impact on the package from only two of the girders simultaneously, with a narrow 

5 miss from a third. If one of the impacting girders strikes the package near its mid-point, the 

6 second one strikes on one end. Figure 7.20 shows a diagram of the pre-impact configuration 

7 for this scenario. 

8

9

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

11

12 Figure 7.21 shows the deformation of the upper roadway girders after 0.205 s, which is after the 

13 maximum load has been transferred to the package. The calculation continues to 0.4 s, and the 

14 plastic strain in the package body at this time is shown in Figure 7.22. Plastic strain is localized 

15 at the contact points with the girders on the upper edge of the package, and with the lower 

16 roadway on the lower edge of the package. The maximum through-wall plastic strain is plotted 

17 in Figure 7.23. The maximum strain is predicted to be 2.4%. Plastic strain drops to zero on the 

18 inside surface. 
7-20

1

2 Figure 7.21. Case #1: Deformation of I-580 Span after Impact

3

4

5 Figure 7.22. Case #1: Effective Plastic Strain in Package Body Wall
7-21

1

2 Figure 7.23. Case #1: Maximum Plastic Strain in Package Body Wall

3

4 Figure 7.24 shows the plastic strain in the DU gamma shield for this scenario. The steel shell 

5 wall deforms across the 1-mm gap between it and the DU, imparting to the DU material plastic 

6 strains up to 1.5%. The plastic strain region in the middle of the length shows where the 

7 containment boundary requires the most support in this impact scenario.

8

9

10 Figure 7.24. Case #1: Effective Plastic Strain in DU Gamma Shield

11
7-22

1 7.3.2 Structural Case #2: Package Parallel to Upper Roadway Girders

2

3 This case considers the effect of an impact with one of the upper roadway girders aligned along 

4 the spine of the GA-4 package. Figure 7.25 shows the positioning of the package and girders 

5 before impact. (The roadway concrete has been omitted from this image, for clarity.) 

6

7

8 Figure 7.25. Case #2: SNF Package and Girder Orientation

9

10 Figure 7.26 shows that plastic strain in the package is below 1.1% for most of the axial length, 

11 with higher values occurring only near the end of the package. In this region, plastic strains are 

12 predicted to be up to 11%. This value is the highest predicted in the four cases. Figure 7.27

13 shows the plastic strain is localized, but it does penetrate into the third layer of elements with 

14 plastic strains below 1%. This is the end region, at either the base plate or closure end, and not 

15 through the package shell wall. While the closure plate and bolts are not modeled, the closure 

16 end is structurally robust because the thick closure plate extends into the flange for additional 

17 structural support. Figure 7.28 shows the predicted plastic strain in the DU gamma shield, 

18 within the steel body wall. Plastic strain in the DU is more widespread but also at a relatively 

19 low magnitude, up to about 3%. 
7-23

1

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

3

4

5 Figure 7.27. Case #2: Local Plastic Strain in Containment End

6
7-24

1

2 Figure 7.28. Case #2: Effective Plastic Strain in DU Gamma Shield

3

4 7.3.3 Structural Case #3: Impact Localized on the Package Lid

5

6 In this case, the package was positioned such that only one girder made contact with the end of 

7 the package. As in all the other cases, the simplified model of the GA-4 package does not 

8 include detailed representation of the closure bolts, but the purpose of this case is to investigate 

9 the amount of potential damage to the flange area. (See Section 7.3.6 for analysis and 

10 discussion of the effect of the fire scenario on the closure bolts.) Figure 7.29 shows the girder 

11 deformation around the package for this drop scenario.

12 

13

14 Figure 7.29. Case #3: Girder Deformation and Effective Plastic Strain
7-25

1 Figure 7.30 shows the plastic strain in the package wall for this drop scenario, which is localized 

2 at the impact area with a magnitude of about 5%. Figure 7.31 shows the extent of the plastic 

3 strain region. Plastic strains barely penetrate into the third layer of elements, with plastic strain 

4 magnitudes less than 0.1%. Figure 7.32 shows the plastic strain in the DU gamma shield, 

5 which remains below 3%.

6

7

8 Figure 7.30. Case #3: Package Effective Plastic Deformation

9

10

11 Figure 7.31. Case #3: Package End Plastic Strain
7-26

1

2 Figure 7.32. Case #3: DU Gamma Shield Effective Plastic Deformation

3

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

5

6 This case investigates the potential for trunnion impact to localize the load on the package wall. 

7 Figure 7.33 shows the girder impact location and deformation. Figure 7.34 shows the plastic 

8 deformation in the package, which is very low and localized. Since peak plastic strains are 

9 below 1%, this case is the least damaging of the four considered here. Figure 7.35 shows the 

10 through-wall plastic strains in the maximum location. Figure 7.36 shows the plastic strain in the 

11 DU gamma shield, which is spread over a few areas but remains below 1.5%. 

12
7-27

1

2 Figure 7.33. Case #4: Girder Deformation

3

4

5 Figure 7.34. Case #4: Package Wall Plastic Strain

6
7-28

1

2 Figure 7.35. Case #4: Through-wall Plastic Strain

3

4

5 Figure 7.36. Case #4: Plastic Strain in DU Gamma Shield

6
7-29

1 7.3.5 Structural Impact Modeling Summary

2

3 The GA-4 package meets all regulatory requirements for over-the-road spent fuel transportation 

packages, and has a current Certificate of Compliance1 4 issued by the NRC. Although the 

5 conditions of the MacArthur Maze fire, as conservatively modeled in these analyses, constitute 

6 conditions that are not specifically defined in the design basis for this package (as per 

7 10 CFR 71), the results of the analyses with the structural models show that this package 

8 design is robust in this scenario. The structure of the GA-4 package wall is predicted to remain 

9 largely undamaged if subjected to this fire and roadway collapse scenario. The plastic strains 

10 imparted to the package wall by the girders are well within the expected ductility limit. Plastic 

11 strain tends to be localized on the outside surface, with through-wall plastic strains tapering off 

12 to zero or near zero on the inside surface. Some permanent deformation of the package wall 

13 would be expected, but there would be no gross failure or rupture of the package. 

14

15 The most important mechanisms at work in this impact scenario appear to be the fundamental 

16 characteristics of the girder structure and the package structure. The girders are thin and 

17 comparatively weak at high temperatures, so tend to deform on impact instead of imparting 

18 energy to the stronger package. Based on the physics, minor variations in temperature or 

19 geometry are not likely to change the broad conclusion that the package would maintain its 

20 structural integrity during this scenario. 

21

22 The results of Case #2 show the potential for localized plastic deformation of the cask if the 

23 impact were to occur in the closure region. The package model used in these impact analyses 

24 represents the closure lid and flange as a single undifferentiated structure. In reality, the 

25 closure lid is inset into the top flange, extending 11 inches into the package cavity. 

26 Approximating the region as a solid slab of material is reasonable for determining the general 

27 package response, but the individual response of the two components is not evaluated. In the 

28 cases with impact at the cask lid region, plastic strain predicted at the flange end could cause 

29 localized deformations that could potentially affect the seating and seal of the closure lid. 

30

31 A complete representation of the impact response of this region would require more detailed 

32 modeling of the lid and top flange. However, the results of the thermal analyses of the package, 

33 as discussed in Section 7.1 and 7.2 above, show that the seals would be expected to fail in this 

34 fire scenario, so it is not necessary to examine additional potential modes of seal failure, such 

35 as the possibility of localized deformation due to impact loading. The issue of retaining the 

36 seating of the closure lid is investigated in Section 7.3.6, which reports the results of evaluations 

37 of the consequences of differential thermal expansion of the closure bolts, package lid, and top 

38 flange using a range of approaches, from classic bolt analysis to detailed FEA modeling of the 

39 components. These evaluations show that these crucial fasteners continue to perform their 

40 function, and would be expected to hold the closure lid in place even in the severe thermal 

41 environment of the MacArthur Maze fire scenario. 

42

43 7.3.6 Structural Issues Related to Bolt Thermal Expansion

44

45 Evaluations with the thermal and structural models of GA-4 package have identified two critical 

46 sets of bolts that could potentially affect its performance under the conditions of the MacArthur 

47 Maze accident scenario. These are the closure bolts fastening the lid to the package body, and 

 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

1 the bolts fastening the top impact limiter to the anchor plate attached to the package body. Fire 

2 temperatures cause a significant amount of thermal expansion in the bolts and the package 

3 structure in general. The bolts are Inconel, and expand less than the XM-19 top flange and lid. 

4 The stresses caused by this thermal expansion mismatch are high, and add to the initial bolt 

5 tension. The closure bolts affect the containment boundary, while the impact limiter bolts affect 

6 the thermal response of the package, especially in the closure region, in that they determine 

7 whether or not the impact limiters would remain attached to the package in this accident 

8 scenario.

9 

10 Inconel and XM-19 stainless steel are materials that retain high strength at elevated 

11 temperatures, but the Inconel experiences significantly less thermal expansion than the XM-19. 

12 This leads to increased tension in the bolt, beyond the initial tension due to the bolt torque 

13 applied at normal operating temperature. As the temperature increases during the fire scenario, 

14 the stresses from thermal expansion increase, and material strength begins to decline. There is 

15 a temperature limit for both sets of bolts, above which there could be failure of the connection, 

16 either through the tensile load on the bolt shank or the shear load on the threaded interface. 

17 This effect is independent of the postulated mechanical impulse load due to the overpass falling 

18 on the package, which occurs early in this transient scenario.

19

20 A critical feature of the GA-4 package design relative to its potential performance in this fire 

21 scenario is that the Inconel bolts are not threaded directly into the XM-19 structure of the 

22 package, either for the lid or the impact limiters. They are instead threaded into Heli-Coil 

23 threaded inserts, which protect the harder metals from galling. These inserts are made of Type 

24 304 stainless steel, which decreases in strength much more rapidly with increasing temperature 

25 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 26 . The dissimilar materials 

27 involved in this design, coupled with the elevated temperatures resulting from the MacArthur 

28 Maze fire scenario, require detailed evaluation. The approach used in this analysis is based on 

29 calculations with engineering mechanics equations, using temperature results from the FEA 

30 thermal model. The effects of elevated temperature and forces due to differential thermal 

31 expansion are also evaluated with the structural and thermal FEA models of these components. 

32

33 The radial symmetry in the package geometry and the time-varying but uniform ambient 

34 boundary temperature throughout the fire scenario result in a symmetrical temperature 

35 distribution around the ends of the package. The uniform boundary temperature, which during 

36 the fire portion of the transient is a bounding temperature based on the FDS analysis presented 

37 in Section 3.0, is a reasonable approximation for the actual conditions seen by a surface with an 

38 assumed bounding uniform view of ambient. The radial symmetry of the temperature 

39 distribution extends to the ring of closure bolts, and also to the ring of impact limiter attachment 

40 bolts. As a result, all of the bolts in a given ring would be expected to have essentially the same 

41 temperature distribution. The response of any given bolt or insert can be treated as 

42 representative of the response of all bolts and inserts of that ring. 

43

44 One important unknown in the bolt thermal expansion evaluation is the post-yield behavior of 

45 the threaded inserts. The inserts carry a shear load that maintains the tension in the bolts due 

46 to initial bolt torque and differential thermal expansion. If the threaded interface were to yield, it 

 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

1 would reduce tension in the bolt, and thereby reduce the potential for the bolt to yield. However, 

2 if the threaded interface were to fail, it could potentially result in a complete loss of bolt restraint 

3 and the possibility of detachment of the closure lid or impact limiter from the package. These 

4 possibilities are evaluated by considering two thresholds for the shear strength of the threaded 

5 inserts: shear yield, based on the material’s yield strength, and shear failure, based on the 

6 material’s ultimate tensile strength. The total shear force on the interface is compared to these 

7 thresholds over the effective thread shear area. It is assumed that the shear area remains 

8 constant between yield and failure, which is reasonable in this case, since the inserts fill the void 

9 between thread patterns and have little physical room to deform.

10

11 The presence of the impact limiters was treated as a variable in the thermal cask analyses by 

12 including them in the ANSYS thermal model but omitting them from the COBRA-SFS model. 

13 The thermal response of the package is affected by this variable, in that the location of the 

14 maximum fuel temperatures is different for the two cases, but the maximum fuel temperature 

15 values are not substantially different. In the context of the closure bolt evaluation, however, this 

16 becomes a critical issue. 

17

18 The thermal models of this study consider two possibilities: the impact limiters detach before the 

19 fire or they remain attached through the fire duration and post-fire cooldown. A case where the 

20 fire or the bridge collapse causes the impact limiter to detach at some point during the fire is not 

21 explicitly modeled, but the thermal results of that case would be bounded by the assumption 

22 that the impact limiters are lost before the beginning of the fire. This treatment provides a 

23 bounding thermal evaluation for the package and the spent fuel assemblies contained within it 

24 (as discussed in Section 7.2), but carrying this uncertainty forward into the closure bolt 

25 evaluation is unnecessary. As discussed below in Section 7.3.6.1, loss of the impact limiters is 

26 not a credible consequence of the MacArthur Maze accident scenario. Therefore, the impact 

27 limiters are assumed to remain attached to the package for the evaluations of the closure lid 

28 bolts presented in Section 7.3.6.2 and Section 7.3.6.3. 

29

30 7.3.6.1 Results of Impact Limiter Bolt Evaluations

31

32 A 3D structural and thermal finite element model of one impact limiter bolt and the surrounding 

33 material was developed, using the ANSYS thermal model results as a baseline temperature 

34 distribution. (See Section 5.4.2 for model description details.) This model considered the 

35 effects on bolt temperature of an additional heat transfer path for the bolts that was neglected in 

36 the ANSYS thermal model of the GA-4 package. The detailed single-bolt model was used to 

37 evaluate bolt thermal expansion at the predicted temperatures.

38

39 The GA-4 package impact limiter attachment bolt heads are located at the bottom of long holes 

40 extending the axial depth of the impact limiter cap. (See diagram in Figure 5.19.) The bolt 

41 heads are by design uncovered, for access when attaching or removing the impact limiters. 

42 This configuration results in the bolt heads being directly exposed to ambient conditions. The 

43 deep holes would essentially preclude convection heat transfer to the bolt heads, but they would 

44 be exposed to thermal radiation exchange with ambient conditions. 

45

46 This pathway for heat to enter the system was not captured in the ANSYS thermal model of the 

47 GA-4 package, since it would not be expected to have a significant effect on the total package 

48 thermal response to the fire scenario. However, this omission could have an effect on the local 

49 temperature distribution and affect the thermal expansion calculations for the impact limiter 

50 bolts. Therefore, additional evaluations were performed with the impact limiter bolt model, to 

51 determine the potential effect of the additional thermal energy that could be entering the region 
7-32

1 by means of thermal radiation. There are two possible ways the extra energy could affect the 

2 thermal response of the impact limiter bolts to the fire conditions. It could increase the 

3 temperature in the threaded interface, or it could increase tension in the bolt, due to changing 

4 the temperature distribution in the bolt shank. Both of these possibilities are explored using the

5 ANSYS thermal model of the impact limiter bolt region.

6

7 The single bolt model used in the thermal expansion structural evaluation was extended to the 

8 thermal-physics environment by changing element types, degree of freedom constraints, and 

9 other ANSYS model parameters. The thermal radiation load was modeled with ideal heat 

10 transfer between the fire environment and the top bolt head surface, with all other surfaces 

11 adiabatic. The initial temperature was set to 124°F (51°C) throughout the bolt and local 

12 package region, which is the minimum initial temperature in the region, based on the ANSYS 

13 thermal model of the GA-4 package at NCT. The single bolt model was then subjected 

14 sequentially to the two phases of the MacArthur Maze fire scenario. As described above, the 

15 first phase consists of an ambient temperature of 2012°F (1100°C) lasting 37 minutes, and the 

16 second phase consists of an ambient fire temperature of 1652°F (900°C), lasting for another 

17 71 minutes. Evaluation of the post-fire cooldown period was found to be unnecessary.

18

19 The results obtained with this thermal model show that including thermal radiation to the top of 

20 the bolt slightly increases the temperature of the bolt shank at the end closest to the fire, but has 

21 almost no effect on the temperature at the other end of the shank, in the threaded region of the 

22 bolt. This response is due to the relatively low thermal conductivity of the bolt nickel alloy, and 

23 the relatively small amount of radiant thermal energy that can be absorbed through the small 

24 cross-section of the bolt heads. The net effect of including thermal radiation at the bolt head is 

25 to slightly increase the bolt shank temperature relative to the package steel wall, which slightly 

26 reduces overall bolt tension, because of the corresponding small increase in thermal expansion. 

27 This is because thermal radiation helps reduce the component of bolt tension caused by the 

28 non-uniform temperature distribution in the bolt shank and box region. The results of this 

29 analysis show that it is appropriate and conservative to neglect the effect of thermal radiation on 

30 the bolt head in this analysis. 

31

32 These evaluations show that the fire conditions defined to bound the MacArthur Maze fire could 

33 be damaging to the threaded inserts, but the potential for failure of the threaded interface is 

34 limited by the bolt shank’s capacity to yield and relieve tension. The results of these analyses 

35 support the conclusion that the attachment of the impact limiters to the package would not fail 

36 due to thermal loads imposed by this fire scenario. The mechanical loads of the accident are 

37 within the regulatory realm and do not require additional evaluation. This fire scenario 

38 represents a more challenging fire accident than that required in the design basis for the GA-4 

39 package, but a detailed thermal evaluation of the system shows that the connection formed by 

40 the impact limiter bolts would survive even under these conditions.

41

42 The baseline temperature history case evaluates the thermal expansion of the impact limiter bolt 

43 using a temperature history from the results obtained with the ANSYS thermal model of the GA44 4 package. (See Section 7.2 for package thermal modeling results.) This model yields detailed 

45 nodal temperature predictions in the region of the bolts, but the mesh does not directly align with 

46 the nodal mesh of the bolt model. The predicted temperatures were mapped onto the bolt 

47 model mesh using a simple averaging scheme at locations along the axial length of the bolt. 

48 The appropriateness of this approach was verified by developing a mapping scheme based on 

49 the maximum temperatures at a given axial location. The temperature history considered here 

50 includes only the 108-minute fully engulfing fire transient. The results show that considering the 

51 post-fire cooldown period is not necessary.
7-33

1

2 Total force on the thread interface is extracted from the ANSYS model at each solution time 

3 step, and the tension due to the bolt preload is added to obtain the total shear force. The initial 

4 preload due to bolt torque is a tension of 21,657 lbs (96.3 kN) at room temperature (General 

5 Atomics 1998). To account for temperature-dependent material properties, the preload force is 

6 scaled by the modulus at the average shank temperature. For example, with an average shank 

7 temperature of 754°F (401°C), the preload is reduced to 88% of its initial value. 

8

9 The bolt will yield wherever the temperature is hottest along its length, since yield strength 

10 decreases with temperature, but the force load on the cross section remains constant along the 

11 axial length. The shear yield behavior of the threaded inserts is not well quantified, but the 

12 tensile yield behavior of the bolt shank would closely resemble a typical tensile test. In the 

13 temperature range of the conditions predicted for the MacArthur Maze fire scenario, the 

14 elongation capacity of the bolt material is approximately 20% or greater. The amount of 

15 elongation required to relieve all tension in the bolt at these temperatures is less than 1%. This 

16 indicates that the bolt shank has enough ductility to accommodate the yield without failure. 

17

18 For all temperatures in the range of interest, including post-fire temperatures, these evaluations 

19 show that bolt yield is reached before insert shear failure. Shear failure of the insert before the 

20 bolt shank yields and releases tension could occur only with the insert at a much higher 

21 temperature than the bolt shank. This is not physically possible in this scenario, because of 

22 package geometry considerations, and is therefore not of concern.

23

24 7.3.6.2 Results of Lid Closure Bolt Evaluations (Classical Approach)

25

26 The results of classical bolt calculations for the bolts securing the closure lid of the GA-4 

27 package are presented in this section. The analysis considers the bolt preload (due to the initial 

28 torque on the bolts), the external load (due to the internal pressurization of the package cavity), 

29 and the thermal load (due to differential thermal expansion of the Inconel bolts and XM-19 

30 stainless steel closure lid and flange of the package body). The bolt preload is defined based 

31 on the manufacturer’s requirements (General Atomics 1998). The external load due to package 

32 pressurization is calculated based on the average gas temperature within the package cavity, as 

33 predicted with the ANSYS thermal model. The thermal load is determined in this analysis, 

34 based on the detailed temperature predictions from the ANSYS thermal model.

35

36 Many parameters in the classic bolt analysis are based on the geometry of the system, and 

37 therefore are straightforward to determine, such as the diameter of the bolt shank and its cross38 sectional area. Less readily defined parameters include the effective joint area, which affects 

39 the representative stiffness of the lid and flange combined. One common method (Juvinall and 

40 Marshek 1991) uses the average area of a conical volume between clamped members. An 

41 alternative approach using bolt mechanics gives a more approximate estimate, based on bolt 

42 diameter. For the GA-4 geometry, the effective areas estimated from these two methods are 

4.60 in2 and 3.18 in2 43 , respectively. The difference in estimated area results in less than a 3% 

44 difference in predicted bolt tension, which is within the uncertainty in the overall approach, and 

45 therefore either estimate would be considered reasonable.

46

47 The thread inserts are a complicating feature for this analysis. Although thread inserts are 

48 commonly used to prevent thread galling, there is little information in the literature concerning 

49 their effect on the mechanics of a bolted joint. In typical applications, it can generally be 

50 assumed that the thread inserts have negligible effect on the bolt tension, particularly if the 

51 thread insert material is similar to the surrounding material. In the GA-4 package, however, the 
7-34

1 bolts are nickel alloy (Inconel), the closure lid and flange is XM-19 stainless steel, and the 

2 inserts are Type 304 stainless steel thread inserts. At NCT, temperatures are in a range where 

3 the XM-19 and Type 304 stainless steels have similar mechanical and thermal properties. 

4 Potential problems arise in the fire scenario as the predicted temperatures on these 

5 components rise to ranges in which the yield strength of the Type 304 insert decreases much 

6 more than the XM-19. At high temperatures and high tensions, the classic bolt calculations 

7 demonstrate that the inserts could have a dramatic effect on the clamping force between the 

8 closure lid and the package body flange.

9

10 As part of an initial assessment of the bolt behavior due to the elevated thermal loads of the fire 

11 scenario, the effect of the thread insert was neglected. Classical bolt calculations were used to 

12 determine the bolt tension and clamping force over time, assuming the thread insert would have 

13 no effect on the bolt or joint stiffness. The joint effective area was assumed to be 4.5 square 

14 inches, which is near the high end of the area estimates described above. The key parameters 

15 in this calculation are listed in Table 7.1.

16 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

17

18 The significant loads on the bolt are determined from the initial tension specified by the design 

19 and the temperature and internal pressure response of the package to thermal conditions of the 

20 MacArthur Maze fire scenario, as determined from the thermal modeling. The ANSYS thermal 

21 model calculated the transient out to 14.5 hours, which is approximately an hour beyond the 

22 point at which all components of the package were decreasing in temperature. This is more 

23 than 12 hours after the end of the actual fire duration, but average component material 

24 temperatures within the package are still above 1000°F (538°C), and the average cavity 

25 pressure is more than 100 psi above the initial pressure at the start of the transient. Given the 

26 large thermal inertia of the GA-4 package, it would take a very long time for the system to cool 

27 back down to its starting conditions after exposure to a fire of the severity of this bounding 

28 model of the MacArthur Maze fire scenario. 

29
7-35

1 Figure 7.37 shows the extrapolated cavity gas temperature calculated with a reasonable 

2 assumed power-law cooldown rate, based on the physical behavior of a large object with high 

3 thermal inertia cooling in air. The corresponding cavity pressure from the ideal gas law 

4 (conservatively assuming no leakage), is shown in Figure 7.38. Based on this extrapolated 

5 cooldown rate, the time required for the closure lid bolt region to drop from its peak temperature 

6 of 1223°F down to 500°F would be on the order of 50 hours. It would require approximately

7 400 hours for the system to return to its original temperature and pressure state. The long-term 

8 temperature response is important to this analysis, because it is in the cooldown phase that the 

9 clamping force has potentially the greatest challenge in this scenario. 

10

11

12 Figure 7.37. Average Cavity Gas Temperature from ANSYS Thermal Model Calculated to 

13 14.5 Hours, with Power-law Extrapolation to 400 Hours
7-36

1

2 Figure 7.38. Cavity Pressure Calculated from Average Cavity Gas Temperature, Assuming 

3 No Leakage

4

5 The results of the classical bolt analysis are tabulated in Table 7.2, with the three components 

6 of bolt force separated for comparison. Bolt Force is the tension on each of the twelve bolts, 

7 considering all loads and the stiffness of the bolt and joint material. Clamp Force is expressed 

8 on a per-bolt basis, so the surfaces in contact would experience a total force that is twelve times 

9 the value given in the table. Figure 7.39 shows plots of these force values, on a per-bolt basis. 

10 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

11
7-37

1

2 Figure 7.39. Bolt Force and Clamping Force Predicted using Classic Bolt Equations

3

4

5 The peak value of 49,103 lbs for the bolt force calculated from the classical bolt analysis is 

6 significantly high, considering the strength of materials involved. Figure 7.40 shows the

7 predicted strength of the thread insert and bolt materials over time (in response to the 

8 temperature transient), compared to the predicted bolt force during this transient. The strengths 

9 are determined based on the thread insert minimum shear area, which is defined at the major 

10 diameter of the bolt thread. As the plot in Figure 7.40 shows, the bolt is not in any danger of 

11 yielding under the load of the predicted bolt force in this analysis. By contrast, the weaker 

12 thread insert is predicted to exceed its yield limit at approximately three hours into the transient. 

13 However, the ultimate shear strength of the insert is not exceeded at any time during the 

14 transient, which suggests that the insert would yield and release some tension without 

15 completely failing. 

16
7-38

1

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

3 Classic Bolt Equations

4

5

6 This is an important conclusion for the release calculations, and provides some indication that 

7 the GA-4 package could withstand the MacArthur Maze fire scenario without significant failure of 

8 the containment boundary. However, the yielding in the inserts and the effect this would have 

9 on the clamping force needs to be more precisely quantified, so that it can be used in

10 evaluations of potential consequences of subjecting an SNF package to this accident scenario. 

11 It is therefore necessary to expand this analysis with the classic bolt equations to include an 

12 estimate of the effect of insert yielding on the closure lid clamping history.

13

14 If the thread insert were to yield, this would be expected to result in a reduction of tension in the 

15 bolt. Furthermore, such yielding would be expected to cause a permanent reduction of the 

16 initial tension load, which would persist after the system returned to its starting temperature and 

17 pressure. This is truly a complex scenario, but a few simplifying assumptions can be made to 

18 adjust the classic bolt equations to estimate the potential reduction in clamping force.

19

20 The first assumption is that the force that can be carried by the thread insert is limited to the 

21 yield strength at each particular temperature. This is a conservative assumption because in 

22 reality the material would be expected to strain-harden, which means that the actual load that 

23 the insert can support would tend to increase as the material starts to yield. A second 

24 assumption is that the effect of plastic deformation in the thread insert corresponds to a pound25 for-pound reduction in the initial bolt tension. In other words, the initial preload force (Fi in 

26 Table 7.2) is permanently reduced by the amount of bolt force that exceeds the thread insert 

27 yield limit. This is potentially a non-conservative assumption because the yielding occurs at 

28 elevated temperatures where the stiffness is lower than the starting condition. Yielding at high 
7-39

1 temperatures might contribute to a proportionally higher reduction of bolt tension when the 

2 system cools. Both of these assumptions are made to simplify the problem enough to obtain an 

3 estimate of the potential effect on the clamping force. 

4

5 Figure 7.41 shows the effect on bolt force and clamping force when yielding of the thread insert 

6 is taken into account in this analysis. The plot in this figure compares these results from the 

7 predicted forces shown in Figure 7.39, calculated assuming no yielding of the thread insert. The 

8 plot in Figure 7.41 shows that accounting for yielding of the thread insert results in a dramatic 

9 change in the overall results. By the time the system returns to its initial state, the clamping 

10 force is significantly reduced from its starting value. This is a very different outcome than is 

11 predicted in the elastic case, which assumes the bolt will ultimately return to its starting state of 

12 tension. 

13

14

15 Figure 7.41. Estimated Effect of Yielding Thread Inserts Using Classic Bolt Equations

16

17 Since the package ethylene propylene seals are not expected to survive the MacArthur Maze 

18 fire temperatures, the clamping force on the metal-to-metal contact surfaces of the closure lid 

19 and package body flange is all that will be holding the pressurized contents inside the package 

20 cavity. In this configuration, the leak rate from the package cavity would be directly related to 

21 the clamping force. Basing estimates of leak rates on this type of simplified modeling evaluation 

22 would lead to large uncertainties in any estimate of potential release rate from the package. 

23 Because of the sensitive nature of the leak rate calculations, a more sophisticated estimate of 

24 the clamping force history is needed. This is provided by the results of the detailed FEA 

25 modeling approach, as described in Section 7.3.6.3.

26
7-40

1 7.3.6.3 Results of Lid Closure Bolt Evaluations (FEA Modeling Approach)

2

3 As described in Section 5.4.3.2, the FEA modeling of the lid closure bolt region consisted of a 

4 detailed representation of a single bolt, due to the radial symmetry of the package and assumed 

5 boundary conditions for the MacArthur Maze fire scenario. Figure 7.42 shows a diagram of the 

6 model, including a close-up of the detailed mesh in the thread insert region.

7

8

9 Figure 7.42. Diagram of Detailed Mesh in Thread Insert Region

10

11 In the classical bolt analysis presented in Section 7.3.6.2, the critical shear area for the thread 

12 insert was assumed to occur at the maximum diameter of the bolt thread, corresponding to 7/8-

13 pitch. In reality, the critical shear area of the insert would be more likely to occur on the flange 

14 thread. However, in this geometry, the area of the two possible shear areas would not be 

15 significantly different. The finite element model shows that peak stress concentrations occur on 

16 the bolt side, but it is not obvious which shear plane would have the higher average stress. This 

17 is illustrated in Figure 7.43 which shows the peak von Mises stress (psi) state in the thread 

18 inserts at the peak temperature in the evaluation assuming elastic materials. The high stresses 

19 would certainly cause plastic strain to occur, with a corresponding redistribution of the stress. 

20 This is further evidence that the evaluation requires a multilinear elastic-plastic material model.
7-41

1

2 Figure 7.43. Stress Distribution Predicted with Elastic-only Assumptions in FEA Model

3

4 With elastic materials, the bolt force and clamping force agree with the classic bolt calculations 

5 to within 1% at the initial temperature and pressure. But as temperature and pressure increase 

6 in the transient, the bolt force and clamping force predicted with the FEA model are 

7 approximately 20% lower than the values predicted by the classic bolt calculations. Sensitivity 

8 evaluations comparing the FEA model results and the classic bolt calculations showed that the 

9 two can be brought in good agreement using assumptions and parameter adjustments that 

10 simplify the FEA model to approximate the level of detail captured in the classical approach. 

11 These evaluations demonstrated the significance of the stiffness of the thread insert material in 

12 determining the behavior of the system, and show that the classic hand calculations 

13 overestimate the bolt tension and clamping force, and would offer a conservative prediction of

14 the amount of plastic deformation that would occur in the thread insert.

15

16 This is consistent with the distinction between a bolt and a screw, as was noted previously as an 

17 important factor in this evaluation. The nut-and-bolt configuration assumes the clamping force 

18 is applied by the head of the bolt on one side and the axial face of the nut on the other side of 

19 the joint. The clamping area estimate for a true bolt (Juvinal and Marshek 1991) assumes that 

20 the conical volume extends from the bolt head to the nut, as illustrated in Figure 7.44(a). The 

21 clamping area estimate for the corresponding screw configuration, as shown in Figure 7.44(b)

22 would result in an effective clamped volume that would be smaller than in the bolt-and-nut 

23 configuration. The angle defining the volume for the screw fastener cannot readily be 

24 determined in the classical approach, and for this reason alone, the elastic FEA results would 

25 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 26 ). 
7-42

1

2 (a) (b)

3 Figure 7.44. Diagram of Compression Volume for (a) Bolt and (b) Screw Fasteners

4

5 Even when considering only elastic behavior of the system, the FEA model provides a more 

6 realistic prediction than the classic bolt calculations. Further refining the FEA model to include 

7 multilinear temperature-dependent material properties for the thread insert material permits a

8 best-estimate of bolt tension and clamping force for the conditions assumed in the MacArthur 

9 Maze fire scenario.

10

11 For the nonlinear analysis, the only change in the FEA model was in the definition of the 

12 material properties for the thread insert. The multi-linear kinematic hardening model was used 

13 to define stress-strain curves for a range of temperatures between 70°F (21°C) and 1200°F

14 (649°C). Figure 7.45 shows the stress strain curves in units of psi versus strain (in/in). The bolt 

15 material and flange material were both defined as elastic material. The nickel alloy bolt retains 

16 such high strength that even localized yielding is not expected. 
7-43

1

2 Figure 7.45. Stress-Strain Curves used for Thread Inserts in Multi-linear FEA Model

3

4 The model predicts small plastic strains throughout the thread inserts, particularly along a line 

5 from the bolt outer diameter to the flange inner diameter, crossing from one shear plane to the 

6 other. Figure 7.46 shows contours of equivalent plastic strain at the peak temperature and 

7 pressure. The peak localized plastic strain is about 5%, while the maximum average along a 

8 plane is roughly 2% in the bottom segment. At the top of the thread pattern, the plastic strain is 

9 generally much less, and the maximum average plane drops to less than 1%. However, even 

10 this relatively small plastic strain is significant when strain hardening is considered. Considering 

11 a plastic strain of 1% along the critical shear plane at 1200°F (649°C), strain hardening would 

12 increase the yield strength by roughly 50%. This suggests that the predictions obtained with the 

13 classic calculations are highly conservative, even when considering the effect of yielding thread 

14 inserts (as shown in Figure 7.41), since that approach does not consider the effect of strain 

15 hardening. 
7-44

1

2 Figure 7.46. Plastic Strain in Thread Insert Predicted with Multi-linear FEA Model

3

4 The bolt tension and clamping force history, on a per-bolt basis predicted with the nonlinear 

5 finite element model is summarized in Table 7.3. These results can be considered a best 

6 estimate evaluation, within the conservative assumptions of the MacArthur Maze fire scenario. 

7 Comparing the initial conditions and final cooldown conditions at 400 hours, the ultimate 

8 reduction in clamping force due to plastic deformation of the thread inserts is predicted to be 

9 less than 8%. In absolute numbers, the loss of initial load (Fi from the classic bolt calculations) 

10 is only 1358 lbs. 

11 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

12

13 The nonlinear finite element model provides a reasonable best estimate prediction of the 

14 clamping force on the flange throughout the MacArthur Maze scenario and its eventual 

15 cooldown. These results are used as the basis for leak rate calculations for the GA-4 package, 

16 presented in Section 8.3.3.
8-1

1 8.0 POTENTIAL CONSEQUENCES

2 Potential adverse consequences of a severe accident involving an SNF transportation package 

3 fall into two general categories; a loss of shielding, which could pose an exposure risk to 

4 members of the public, and a failure of the containment boundary of the package, which could 

5 lead to a release of radioactive material to the environment. 

6

7 Loss of either the neutron or gamma shielding of the package could potentially result in a direct 

8 radiation dose to an individual in close proximity to the package. Failure of any of the 

9 components (e.g., package seals) that make up the containment boundary could result in a 

10 release of radioactive material from inside the package, potentially resulting in a direct 

11 radioactive dose to first responders at the scene of an accident, or possibly to members of the 

12 public in the surrounding area. Loss of shielding as a potential consequence of the MacArthur 

13 Maze fire scenario is discussed in Section 8.1. Section 8.2 discusses the performance of 

14 package seals in this fire scenario. Section 8.3 presents detailed evaluations of the potential for 

15 a release from the GA-4 package as a result of the severe conditions predicted for this

16 extremely conservative representation of the MacArthur Maze fire scenario.

17

18 8.1 Potential for Loss of Shielding

19

20 The potential for increased neutron and gamma radiation dose rates from the GA-4 as a result 

21 of exposure to the MacArthur Maze fire scenario was evaluated. Direct radioactive dose rate 

22 limits are specified in 10 CFR 71 for NCT and HAC conditions. As a licensed transportation

23 package, the design basis of the GA-4 complies with the regulatory limits for all conditions of 

24 transport.

25

26 Section 8.1.1 describes the consequences of loss of neutron shielding for the GA-4 in the 

27 MacArthur Maze fire scenario. Section 8.1.2 discusses the potential for loss of gamma 

28 shielding.

29

30 8.1.1 Neutron Shielding

31

32 Neutron shielding for the GA-4 package is provided by neutron-absorbing liquid in an annular 

33 tank surrounding the steel body of the package (see Sections 5.1 and 5.2 for details of package 

34 geometry). The neutron shielding material is a mixture of 56% propylene glycol and water, with 

35 1% dissolved boron. The neutron shield tank is not generally expected to survive the 

36 hypothetical accident conditions prescribed in 10 CFR 71 for SNF transportation packages, 

37 which include a 30-minute fully engulfing fire at “1475°F (800°C).” 

38

39 The GA-4 package is designed to be in compliance with the regulatory limits for all conditions of 

40 transport. Loss of the neutron shield tank contents is a design-basis assumption for HAC, and 

41 analyses presented in the SAR (General Atomics 1998) for this package assume loss of the 

42 neutron shield in all accident scenarios, including the HAC fire. The conditions of the MacArthur 

43 Maze fire, although more severe than the HAC fire, can do no more damage to the neutron 

44 shield of the GA-4 than is assumed a priori in the HAC fire. Therefore, the GA-4 package would 

45 be expected to remain below the regulatory dose limits after loss of neutron shielding in the 

46 MacArthur Maze fire scenario. 

47
8-2

1 8.1.2 Gamma Shielding

2

3 Gamma shielding for the GA-4 is provided by a 2.64-inch-thick (6.7 cm) layer of DU encased 

4 within the stainless steel body of the package. The DU layer extends a few inches beyond the 

5 full axial length of the package inner cavity to assure complete coverage of the active fuel 

6 length, and is positioned between the stainless steel inner liner and the 1.5-inch-thick (3.81 cm) 

7 stainless steel body of the package.

8

9 The DU material experiences a significant increase in temperature, but the performance of the 

10 DU gamma shield is unaffected by the MacArthur Maze fire scenario. As discussed in 

11 Section 7.3, structural analysis of a range of postulated scenarios in which the overhead I-580 

12 highway span is assumed to impact the GA-4 package shows that the package would be 

13 expected to survive without structural failure. The thermal analyses of the package response to 

14 the fire scenario show that the peak temperature in the DU material could reach up to about 

15 1480°F (805°C). This is significantly below this material’s melting temperature of 2070°F 

16 (1132°C). 

17

18 These results show that the gamma shielding of the GA-4 can be expected to remain intact and 

19 functional even if subjected to the severe conditions of the MacArthur Maze fire scenario. 

20 Therefore, the GA-4 package would be expected to remain below the limits specified in 10 CFR

21 71 for accident conditions.

22

23 8.2 Performance of Package Containment Seals

24

25 Based on the results of the thermal analysis (as discussed in Sections 7.1 and 7.2), there is a 

26 possibility of a release from the package because of failure of components that make up the 

27 containment boundary of the package. Calculated temperatures in the region of the lid closure 

28 seal, drain valve/port, and gas sample valve/port seals during the transient far exceed the 

29 continuous use temperature limits for the seal material. Therefore, the potential exists for the 

30 release of contents from the package in this fire scenario.

31

32 Section 8.2.1 provides a detailed discussion of seal temperatures predicted for the GA-4 

33 package in the MacArthur Maze fire scenario. Section 8.2.2 presents a discussion of seal 

34 performance at elevated temperatures, based on experimental data relevant to the conditions of 

35 the MacArthur Maze fire. 

36

37 8.2.1 Seal Temperatures in the MacArthur Maze Fire Scenario

38

39 The containment boundary for the GA-4 package is maintained by the seals on the package lid, 

40 drain valve and port, and gas sample valve and port. The package lid seal consists of primary 

41 and secondary O-rings at the interface of the lid and the package stainless steel body. The gas 

42 sample valve is located within the package lid, and the drain valve is located in the steel base of 

43 the package. The gas sample valve is sealed with primary and secondary O-rings, and for 

44 transport conditions, the outer face of the port is fitted with a steel plug that is threaded to a

45 specified torque of 20 ft-lb. The drain valve is sealed within its access port with primary,

46 secondary, and tertiary O-rings. The drain valve cover and drain port plug are also sealed with 

47 O-rings, in addition to being threaded, and are torqued to 20 ft-lb. 

48

49 The O-ring seals at all locations are ethylene propylene, which has a continuous-use 

50 temperature limit of 302°F (150°C). Figure 8.1 shows a graph of the temperature limit on this 
8-3

1 material as a function of exposure time. As exposure temperature increases, the time limit for 

2 allowed exposure decreases. (Note that the horizontal axis, Exposure Time, in this plot is on a 

3 logarithmic scale.) The maximum temperature this material is rated to withstand without

4 effectively immediate failure is 790°F (421°C), but it will tolerate this exposure for only six 

5 minutes. 

6

7

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

9 Propylene Seal Material (Based on Data Presented in the GA-4 SAR [General 

10 Atomics 1998])

11

12 All of the containment boundary seals in the GA-4 package are in locations that are covered by 

13 either the top or bottom impact limiter assembly. Without the impact limiters in place, the seal 

14 locations are directly exposed to the fire conditions, and as a result, exceed all operating 

15 temperature limits within minutes, and high temperatures persist at these locations throughout 

16 the fire transient and into the post-fire cooldown. However, as discussed above in Section 7.0, 

17 structural and thermal evaluations show that loss of the impact limiters is not a credible 

18 consequence of this fire scenario. Therefore, the seal performance evaluations consider only 

19 the case with the impact limiters in place.

20

21 In the HAC fire analysis (30 minutes at 800°C) presented in the package SAR, the predicted 

22 peak temperatures on the seal components do not exceed the bounds of the operating 

23 temperature limit curve shown in Figure 8.1. This is verified by results obtained for this transient 

24 with the ANSYS and COBRA-SFS models, as noted in Section 6.1. It is therefore justifiable to 

25 assume that the seals would survive intact in the HAC fire. In the conditions of the MacArthur 

26 Maze fire scenario, however, the higher fire temperatures and longer duration of the fire 

27 exposure result in temperatures that far exceed the rated temperatures of this hydrocarbon seal 

28 material. Figure 8.2 shows the temperature predicted with the ANSYS model (with impact 

29 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

1

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

3 Limiters) for the MacArthur Maze Fire Scenario

4

5 The plot in Figure 8.2 also shows the time in the transient when the closure lid seal location is 

6 predicted to exceed specific exposure temperature limits for the seal material. This plot 

7 indicates that the seal material could survive these conditions for possibly as long as 1.7 hours; 

8 almost the entire duration of the fire portion of the transient. However, temperatures in the seal 

9 location continue to rise for many hours after that point, and by approximately 2 hours into the 

10 transient, exceed all operating temperature limits defined by the graph in Figure 8.1. Similar 

11 plots in Figure 8.3 for the gas sample valve/port seals and in Figure 8.4 for the drain valve/port 

12 seals show that all operating temperature limits for these seals are exceeded, in approximately 

13 the same time-frame.
8-5

1

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

3 Impact Limiters) for the MacArthur Maze Fire Scenario

4

5

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

7 Limiters) for the MacArthur Maze Fire Scenario

8
8-6

1 The plots in Figure 8.2, Figure 8.3, and Figure 8.4 incidentally confirm the SAR prediction, 

2 showing that these seals could probably survive a fire lasting only 30 minutes, followed by a 

3 cooldown at normal ambient conditions. However, if exposed to the conditions of the MacArthur 

4 Maze fire scenario, the seal material would exceed its rated long-term temperature limit within 

5 1.5 to 2.5 hours at all locations. Complete evaluation of the effect of the fire scenario on the 

6 seal region requires evaluating the seal region temperatures during the post-fire cooldown 

7 transient, as well as during the fire itself. As discussed in Section 7.2, the insulating effect of the 

8 impact limiters results in the temperatures of components near the ends of the package 

9 continuing to increase long after the end of the fire. Figure 8.5 shows that the peak 

10 temperatures in the seal region locations continue to increase for more than 4 hours after the 

11 fire, reaching approximately 1150°F (621°C), and after 14.5 hours are still above 1000°F 

12 (538°C). 

13

14

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

16 MacArthur Maze Fire Scenario and Post-fire Cooldown

17

18 The results in Figure 8.2 through Figure 8.5 show that the seal region temperatures would 

19 exceed all exposure temperature limits for a long period of time during the post-fire cooldown. 

20 Seal failure, which is defined as the inability of the seal material to maintain a stable differential 

21 between the internal pressure within the package cavity and the external ambient pressure, 

22 seems inevitable under these conditions. In general, if seal materials exceed their rated 

23 temperature limits, they are treated for safety evaluation purposes as having failed entirely. 

24 However, this is a bounding assumption, since it does not attempt to quantify the degree of seal 

25 failure or resulting leak rate as a function of temperature or time-at-temperature. Section 8.2.2

26 summarizes results of experimental measurements of the performance of elastomer seals at 

27 elevated temperatures, which show that seal failure is a complex process, not a simple pass/fail 

28 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

1

2 8.2.2 Seal Performance Testing

3

4 The NRC and NIST conducted testing of seals in thermal conditions simulating fire 

5 environments that exceeded the rated temperatures for the seals tested (NUREG/CR-7115

6 2012). These tests evaluated the performance of one type of metallic seal and two different 

7 polymeric compound seals typically used in SNF transportation packages, using a small

8 stainless steel cylinder with a single O-ring seal. The test vessel was pressurized at 75°F 

9 (24°C) with helium to 73.5 psia (5 bar) for the tests with metallic seals, and to 29.4 psia (2 bar) 

10 for the tests with polymeric seals. The fire was simulated using an electric furnace that could 

11 maintain a controlled thermal environment for a specified duration, which was varied in different 

12 tests from several hours to 24 hours, and in some cases up to 72 hours. (These tests were 

13 designed to simulate accident conditions. Typical seal tests for long-term normal operating 

14 conditions are performed for a minimum of 1000 hours.) Following the simulated fire exposure 

15 duration, the test vessel was cooled to room temperature within the electric furnace. 

16

17 A total of 15 tests were conducted in this study, including the initial shake-down test for which 

18 results were not recorded, due to instrumentation failure. Of the 14 tests for which 

19 measurements were recorded, 11 tests were with a metallic seal, 2 tests were with an ethylene

20 propylene seal, and 1 test was with a polytetraflouroethylene (PTFE) seal. In terms of the 

21 applicability of this testing to the evaluation of the GA-4 package in the MacArthur Maze fire 

22 scenario, the two tests with ethylene propylene seals are of significance, since this is the seal 

23 material used in the GA-4 package for the lid closure and the gas sampling port valve and the 

24 drain valve. 

25

26 The most severe exposure for ethylene propylene seals in the testing was at 842°F (450°C). 

27 Based on the performance curve in Figure 8.1, the exposure time limit for this seal material at 

28 this temperature is less than 6 minutes. The seal material failed in this test within the first three 

29 hours of the simulated fire transient, but exhibited a much slower leak rate than would be 

30 expected for the test vessel with no seals at the test conditions. The second test with ethylene 

31 propylene seals reached a much lower peak temperature, and simulated a fire environment with

32 incremental heating from 302°F (150°C) to 572°F (300°C). Based on the performance curve in 

33 Figure 8.1, the exposure time limit at the maximum heating temperature in this test is less than 

34 30 minutes. The total duration of the simulated fire was more than 20 hours, but in this test, the 

35 seal held with no measurable leakage.

36

37 The results of two tests in a small (not-to-scale) test vessel are not sufficient data on which to 

38 base a general evaluation of ethylene propylene seal material at temperatures above rated 

39 performance values. However, this testing does show that even when exposed to temperatures 

40 above rated time-at-temperature exposure limits, the seal material may retain sealing capability. 

41 Additional testing is needed to obtain a broader technical basis for determining the performance 

42 of seal materials at elevated temperatures, to evaluate the influence of such variables as the 

43 size of pressure vessel, the magnitude of the pressure differential at the high temperature 

44 exposure, and the effectiveness of double versus single O-ring seals. (SNF packages typically 

45 use double O-ring seals, and this testing used only a single O-ring in the test vessels.) These 

46 seal tests demonstrate that ethylene propylene seals such as those that are used for the GA-4 

47 package may have a performance envelope that far exceeds the conservative temperature 

48 limits indicated in the ratings for long-term performance provided by the seal manufacturer. 

49 Seals will not necessarily exhibit catastrophic leakage when exposed to temperatures beyond 

50 design basis, even at the extreme thermal exposure postulated for the MacArthur Maze fire 

51 scenario.
8-8

1 The results of the thermal analysis indicate that for the GA-4 package in the MacArthur Maze 

2 fire scenario, all seals would be significantly challenged. Based on the seal testing conducted 

3 by NRC discussed here, it is remotely possible that some sealing capability would remain for the 

4 seals even after exposure to excursion temperatures well above their rated temperatures. 

5 However, for the purposes of determining the potential release from the GA-4 package in this 

6 fire scenario, a simple pass/fail criterion is used to evaluate potential seal performance. If the 

7 temperatures predicted in the transient exceed the maximum recommended service 

8 temperature for the seal material, the seal is assumed to fail. This evaluation shows that the 

9 assumption of complete seal failure is bounding for the performance of the GA-4 package seals, 

10 and may be quite conservative. Nevertheless, this assumption is the basis for determining that 

11 a release is possible from the GA-4 package in this fire scenario due to package containment 

12 boundary leakage. The analyses presented in the following sections determine the character 

13 and amount of material that could be released.

14

15 8.3 Potential Release Issues

16

17 NRC staff evaluated the potential for release of radioactive material from the GA-4 package as a 

18 consequence of the MacArthur Maze fire scenario. Based on the results of the thermal analysis 

19 (as discussed in Sections 7.1 and 7.2), there is a possibility of a release from the package 

20 because of failure of components that make up the containment boundary of the package. 

21 Calculated temperatures in the region of the lid, drain valve/port, and gas sample valve/port 

22 seals during the transient far exceed the continuous use temperature limits for the seal material, 

23 as discussed in Section 8.2 above. In addition, the peak temperatures predicted for all fuel rods 

24 in the package reach the range where burst rupture of zircaloy cladding can occur, as discussed 

25 in the presentation of results of the thermal analyses in Sections 7.1 and 7.2. Therefore, the 

26 potential exists for the release of fission products and spent fuel particles, as well as particulate 

resulting from CRUD1 27 detaching from the fuel rod surfaces.

28

29 Results of analysis of fuel performance for the conditions encountered in this fire scenario are 

30 presented in Section 8.3.1. Evaluations of the potential for release from fuel rods to the GA-4 

31 package cavity are presented in Section 8.3.2. Evaluations of the potential for release from the 

32 GA-4 package to the surrounding environment are presented in Section 8.3.3.

33

34 8.3.1 Fuel Rod Cladding Performance

35

36 Spent fuel has two potential sources of radioactive material that could serve as source terms for 

37 a release from an SNF transportation package; the CRUD on the rod outer surface, and the 

38 radioactive material (fission products and fuel fragments) confined within the metal cladding. 

39 The Standard Review Plan (NUREG-1617 2000) specifies the assumption of 100% spallation of 

40 CRUD from fuel rod surfaces for HAC analyses. For consistency, this assumption is also 

41 applied to the MacArthur Maze scenario. Determining the amount of material that could 

42 potentially be released from within the rods, however, requires additional analysis of fuel rod 

43 behavior for the conditions of the fire scenario. If it can be shown that the fuel rods remain 

44 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

1 conditions are such that the fuel rods could fail, the nature and severity of the potential failure 

2 must be determined.

3

4 The predicted fuel cladding temperatures obtained with the COBRA-SFS model of the GA-4

5 were used to evaluate the potential for rod failure in the severe conditions of the MacArthur 

6 Maze fire scenario. The results obtained with the ANSYS model are more conservative, as 

7 discussed in Section 7.1 and 7.2, due to the more conservative representation of the fuel using 

8 the homogeneous k-effective model. The COBRA-SFS model, with a more detailed 

9 representation of the fuel region using the rod-and-subchannel approach, produces predictions 

10 of fuel and cladding temperatures with a more accurate evaluation of thermal radiation and the 

11 transient thermal inertia of the fuel and cladding. However, the many conservatisms in the 

12 modeling approach and fire scenario definition that apply equally to the ANSYS model and the 

13 COBRA-SFS model ensure that the cladding temperature results obtained with this code are 

14 also very conservative, even though they do not include the specific conservative effect of the 

15 effective conductivity model in the representation of the fuel assemblies. 

16

17 Based on the predicted fuel cladding temperatures from the COBRA-SFS modeling results for 

18 the MacArthur Maze fire scenario, fuel performance was evaluated using the burst rupture 

19 model in the FRAPTRAN-1.4 code (NUREG/CR-7023 2011). In the FRAPTRAN code, cladding

20 rupture is evaluated with a burst stress/strain model developed from test data obtained for loss 

21 of coolant accident (LOCA) analysis and reactivity insertion accident (RIA) evaluations. Burst 

22 rupture is the expected mechanism of failure for fuel rods in the reactor core when subjected to 

23 severe accident conditions, and is a potential failure mode for spent fuel at high temperatures. 

24

25 Creep rupture is considered a possible alternative mechanism of failure for spent fuel rods. To 

26 evaluate this possibility, a separate analysis was performed with a creep rupture model, using 

27 the FRAPCON-3.4 code (NUREG/CR-7022 2011) in conjunction with the DATING code 

28 (Simonen and Gilbert 1988). The version of the code used in this analysis has been updated 

29 with creep coefficients from creep tests on irradiated cladding (Gilbert et al. 2002), for the 

30 temperatures in the range predicted for the hottest rod in the MacArthur Maze fire scenario.

31

32 Fuel performance evaluations used the same design basis fuel configuration assumed in the 

33 thermal analysis. The fuel rod cladding initial conditions and cladding temperatures assumed 

34 during the fire scenario are summarized as follows:

35

36 • Westinghouse 14x4 fuel design operating at a rod average linear power rating of 5.7 kW/ft,

37 up to 33 GWd/MTU rod average burnup in-reactor prior to discharge

38 • Initial as-fabricated rod pressure of 460 psig at room temperature 

39 • Rod pressure at room temperature after irradiation to 33 GWd/MTU is 651 psig. 

40 – The increase in pressure from the as-fabricated pressure of 460 psig is due mainly to a 

41 reduction in rod void volume because of fuel swelling and cladding creepdown. There is 

42 also a small increase in pressure due to release of fission gas (0.0024 release fraction) 

43 from the pellets to the fuel rod void space. Fission gas release fractions and 

44 consequently rod pressure would be significantly higher (between 0.05 to 0.20) for a 

45 peak power rod operating rod in the core. A release fraction of 0.002 to 0.005 would be 

46 representative of the majority of rods (but not the peak operating rods) in a core at this 

47 discharge burnup.

48 • Initial conditions at the start of the fire are assumed to be NCT, which results in an initial 

49 steady-state peak cladding temperature of 293°F (145°C) with a rod pressure of 922 psig

50 (hoop stress approximately 50 MPa)
8-10

1 • Cladding temperature as a function of time for the hottest rod in the package, as predicted in 

2 the COBRA-SFS thermal analysis for the fire conditions (see Figure 7.8, which shows the 

3 peak cladding temperatures; however, the complete axial temperature distribution as a 

4 function of time was provided as input to the burst rupture calculations, not simply the peak 

5 temperature)

6

7 The cladding temperatures from the fire, as calculated with COBRA-SFS, and rod pressures 

8 calculated by FRAPCON-3.4 (NUREG/CR-7022 2011) assuming the spent fuel had been 

9 subjected to normal reactor operation at 5.7 kW/ft, were input into FRAPTRAN-1.4 to calculate 

10 the cladding stresses. The FRAPTRAN-1.4 cladding burst model was also used to calculate the 

11 rupture temperature during the fire. The calculated cladding temperatures during the fire from 

12 the COBRA-SFS analysis, and the calculated hoop stresses obtained from FRAPTRAN-1.4 for 

13 the fire conditions were input into FRAPCON-DATING to calculate cladding rupture based on 

14 the out-of-reactor creep relationship in the DATING subroutine. 

15

16 8.3.1.1 Fuel Rod Cladding Performance: FRAPTRAN Evaluation

17

18 The peak cladding temperatures calculated with COBRA-SFS for the MacArthur fire were 293°F 

19 (145°C) at the start of the fire and reached a peak cladding temperature of 1388°F (753°C) in 

20 the fire transient. Based on these temperatures, the calculated cladding hoop stress is 50 MPa 

21 at the start of the fire and reaches a peak of 121 MPa just prior to predicted cladding rupture at 

22 1098°F (592°C), as predicted with the burst strain model in FRAPTRAN-1.4. This relatively low 

23 rupture temperature reflects the conservatism in the cladding temperature history predicted in 

24 the thermal analysis, and the uncertainty in the FRAPTRAN predictions at the relatively low 

25 heating rate for the cladding in this fire scenario. 

26

27 The FRAPTRAN code was designed to predict nuclear fuel behavior during reactor accidents.

28 In particular, failure models have been developed to provide reasonably accurate predictions for 

29 Reactivity Initiated Accident (RIA) and Loss-Of-Coolant Accident (LOCA) failures. For the case 

30 of a fire accident scenario in transportation of spent nuclear fuel, potential cladding failure can 

31 occur as a result of temperature increase and the associated rod internal pressure increase. 

32 These concurrent temperature and pressure increases can result in sufficient stress to cause 

33 ballooning and rupture of the cladding due to rapid high temperature creep. In a FRAPTRAN 

34 calculation of such an event, if the temperature and stress in the cladding are such that the 

35 cladding deforms to its uniform elongation in the hoop/circumferential direction, fuel rod 

36 ballooning is predicted. If the stress exceeds a correlated temperature-dependent level, 

37 cladding rupture will be predicted. In the less severe conditions of the HAC fire, the stress in the 

38 cladding is not expected to exceed the elastic limit, and therefore burst rupture would not be 

39 predicted with the models in the FRAPTRAN code.

40

41 Ballooning and rupture models such as those in the FRAPTRAN code have been developed 

42 with the specific intent to accurately predict cladding failures during a LOCA, where the 

43 temperature increase rate is typically much higher (on the order of 10°C/s or higher) than in the 

44 case of fire scenarios (typically on the order of 0.2°C/s for the HAC fire, and for the MacArthur 

45 Maze fire scenario, conservatively estimated to be 0.27°C/s). For a given stress level, a slower 

46 heatup rate will generally tend to result in a lower rupture temperature, but there is very little 

47 data in the FRAPTRAN validation database that has heatup rates below 1°C/s. There is much 

48 more data at higher heating rates, ranging from 5°C/s to >30°C/s, as fully documented in the 

49 FRAPTRAN code manual (NUREG/CR-7023 2011). Figure 8.6 shows a summary of burst 

50 temperature data in the FRAPTRAN modeling database that are from tests with heating rates in 

51 the range 0.08 to 5°C/s. (This data set is from NUREG/CR-0344.) The lowest burst (i.e.,
8-11

1 rupture) temperatures observed in this limited subset of the validation database are around 

2 667°C (1232°F). Due to the sparseness of the data in this low temperature range, burst rupture 

3 temperatures predicted with FRAPTRAN for heatup rates below 1°C/s have a greater 

4 uncertainty than predictions obtained for higher heating rates, where the database is more fully 

5 populated. In particular, predictions of burst rupture temperatures lower than 667°C (1232°F) 

6 should be evaluated as indicative of the possibility of rupture, rather than absolute indicators of 

7 rupture, since these results are outside the available validation database.

8

9

10 Figure 8.6. Experimental Results for Rod Burst Rupture Testing at Low Heating Rates (data 

11 from NUREG/CR-0344)

12

13 Based on the validation range of the models in FRAPTRAN, and the conservative assumptions 

14 in the thermal modeling that impose an extraordinarily severe temperature transient on the fuel 

15 rods within the GA-4 package in this fire scenario, the predicted cladding rupture at 1098°F 

16 (592°C) obtained in the FRAPTRAN analysis can be considered an extremely conservative 

17 result. However, the predicted peak cladding temperature obtained in the thermal modeling is 

18 1388°F (753°C) in this fire scenario, and this temperature is well within the range where the data 

19 in Figure 8.6 shows burst rupture temperatures at low heating rates. The specific temperature 

20 value for burst rupture predicted with FRAPTRAN for these conditions may be quite 

21 conservative, and may have a fairly large uncertainty, but there is little uncertainty that the 

22 cladding would at some point fail by burst rupture if subjected to the severe conditions predicted 

23 for the fuel in the GA-4 package in the MacArthur Maze fire scenario. 

24
8-12

1 8.3.1.2 Fuel Rod Cladding Performance: FRAPCON/DATING Evaluation

2

3 The cladding failure temperature predicted with the creep model in the DATING code is 1229°F 

4 (665°C), which is significantly higher than the burst rupture temperature of 1098°F (592°C) 

5 obtained in the FRAPTRAN analysis. The DATING code was designed to predict creep failures 

6 and temperature limits for dry storage of spent fuel, based on creep failures of the cladding.

7 The range of applicability of the DATING code is for lower temperatures, resulting in much 

8 longer times to failure, when compared to the FRAPTRAN ballooning models. The databases 

9 used to develop creep rate correlations and creep rupture models for the DATING code span 

10 temperature ranges that are in general lower than the temperatures and heating rates typically 

11 encountered in fire scenarios. 

12

13 The DATING code is a more general creep prediction tool than FRAPTRAN, with its ballooning 

14 and rupture models, which are effectively high temperature creep models. However, it must be 

15 noted that, as with FRAPTRAN, the DATING code is being applied outside its validation 

16 databases when used to evaluate cladding response to the conditions of the MacArthur Maze 

17 fire scenario. However, the results obtained with both modeling tools show that although there 

18 might be some uncertainty as to the exact temperature at which it would occur, fuel cladding 

19 could and probably would fail, if subjected to the severe conditions postulated for the MacArthur 

20 Maze fire scenario. For the purposes of this evaluation, the predicted temperatures from these 

21 codes are taken at face value, and treated as conservative estimates of rupture temperature in 

22 the evaluation of potential consequences of this fire scenario. 

23

24 Table 8.1 summarizes the results of the fuel performance modeling analyses for temperatures 

25 on the hottest rod in the MacArthur Maze fire scenario, as predicted with the COBRA-SFS 

26 model. The burst rupture and creep rupture models both predict that the hottest fuel rod would 

27 rupture if subjected to the temperatures predicted in this fire scenario. Furthermore, the peak 

28 temperature on the hottest rod at the time of rupture is eventually exceeded by all rods in the 

29 package during the transient, which suggests that there is the potential for all rods in the 

30 package to rupture in this fire scenario. 

31 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

32 The burst rupture and creep rupture models predict cladding failure at a single location along 

33 the axial length of a fuel rod. Based on the temperature predictions obtained with the COBRA34 SFS model, which omits the impact limiters, the fuel performance models predict rod rupture in 

35 the end region of the rod. Temperature distributions obtained with the ANSYS model, which 

36 assumes the impact limiters remain in place throughout the transient, result in the highest 

37 temperatures occurring near the axial center of the fuel region, and rod rupture would be 

38 expected near the middle of the rod for this package configuration. Since the design basis fuel 

39 for the GA-4 is low burnup (i.e., no more than 45 GWd/MTU), the degree of pellet-clad 

40 interaction would be relatively limited, and a single rod breach would be expected to effectively 

41 depressurize the fuel rod. Therefore, no additional ruptures are predicted on a given rod, and 

42 potential release calculations are based on the assumption of one rupture per rod. 

43
8-13

1 The peak fuel cladding temperatures predicted with the ANSYS model are somewhat higher 

2 than the peak temperatures on the rod ends predicted with COBRA-SFS (see Sections 7.1 and 

3 7.2). Furthermore, the rod temperatures in both analyses remain much higher than the 

4 predicted rupture temperatures for an extended period of time. It is therefore reasonable to 

5 conclude that for the temperature distribution predicted with the ANSYS model (with the impact 

6 limiters in place), rod ballooning and rupture would also be expected to occur, but in the central 

7 region of the rod, rather than at an end. Table 8.2 summarizes the elapsed time and time 

8 duration that the hottest rod peak temperatures are predicted to exceed the calculated burst 

9 rupture temperatures. 

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

11 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

12 Based on the burst strain model, the fuel rods are expected to rupture before the end of the fire. 

13 Based on the creep rupture model, the fuel rods would also be expected to begin rupturing 

14 before the end of the fire, but slightly later in the transient. Furthermore, the peak temperatures 

15 remain significantly above these predicted rupture temperatures for more than 10 hours. The 

16 results presented in Sections 7.1 and 7.2 show that the fuel rod temperatures continue to 

17 increase even after the end of the fire, because of thermal inertia and build-up of decay heat 

18 that is not removed from the package during and immediately after the fire.

19

20 By the time of the secondary peak of 1348°F (731°C) in cladding temperature predicted with the 

21 COBRA-SFS model, which occurs at 250 minutes elapsed time (142 minutes after the end of 

22 the fire), the peak temperature on every rod in the package exceeds the highest temperature 

23 predicted for rod rupture (1229°F [665°C]). The peak temperature of 1343°F (728°C) predicted 

24 with the ANSYS model is at essentially the same value as that predicted with the COBRA-SFS 

25 model at this point in the cooldown transient. More significantly, at this time the lowest peak rod 

26 temperature is 1285°F (696°C) in the COBRA-SFS model results, and the lowest axial peak 

27 temperature predicted in the fuel region in the ANSYS model is approximately 1134°F (612°C). 

28 Based on these results, it is assumed that all of the rods in each of the four assemblies within 

29 this package would rupture in the MacArthur Maze fire scenario. 

30

31 8.3.2 Potential Release to GA-4 Package Cavity

32

33 Determining potential release quantities from an SNF package involves first determining the 

34 amount of material that is available for release from the fuel rods, and then determining the 

35 amount of this material that can be released from the package. This section presents analyses 
8-14

1 performed by NRC staff to determine the total amount of activity that could be released from the 

2 four assemblies defining the design-basis payload for the package, as described in the GA-4 

3 SAR. Analyses to determine the potential for release from the package to the environment are 

4 presented in Section 8.3.3.

5

6 Typically, release quantities are expressed in terms of release fractions, a ratio calculated as 

7 the amount of material actually released divided by the total amount available for release. 

8 Regulatory guidance for determining the releasable source term for SNF transportation 

9 packages is provided in the Standard Review Plan for Transportation Packages for Spent 

10 Nuclear Fuel: Final Report, NUREG-1617 and in Containment Analysis for Type B Packages 

11 Used to Transport Various Contents, NUREG/CR-6487. The release fractions specified in 

12 these documents are listed in Table 8.3. These release fractions define bounding values for the 

13 fraction of material that is assumed to be released from the fuel rods to the package under NCT 

14 and HAC. 

15 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

16

17 The potential release from the GA-4 package corresponding to the release fractions in Table 8.3

18 is a function of the contents of the package. The radionuclide inventories for the two design 

19 basis fuel configurations for the GA-4 package were obtained using ORIGEN-ARP (Gauld et al. 

20 2009). The source term inventories obtained in these calculations are listed in Table 8.4. 

21 Consistent with the criticality and shielding calculations in the GA-4 package SAR (General 

22 Atomics 1998), WE14x14 fuel at 35 GWd/MTU burnup and 10-years cooling is bounded by 

23 WE15x15 at 35 GWd/MTU and 10-years cooling. Therefore, all source term and potential 

24 release calculations are performed assuming 10-year-old WE 15x15 fuel at 35 GWd/MTU, even 

25 though the thermal analysis is based on WE 14x14 fuel geometry. This is a conservative 

26 assumption, since the temperatures obtained with WE 14x14 fuel would be slightly higher than 

27 those predicted for WE 15x15 fuel for the same design basis decay heat loading. Table 8.4 lists 

28 the calculated source terms for a single assembly. The total inventory within the GA-4 is four 

29 times the quantities listed in this table, since this package can carry up to four fuel assemblies.

30 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

31
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

1
8-16

1 The bounding values for release fractions defined in Table 8.3 were applied to the source terms 

2 listed in Table 8.4 to determine a bounding estimate of the activity that could be released from 

3 the four fuel assemblies to the GA-4 package interior in the MacArthur Maze fire scenario. 

4 Figure 8.7 shows the activity released to the package for the source term inventory from 

5 Table 8.4 for the bounding configuration of 10-year-cooled WE 15x15 fuel at 35 GWd/MTU. 

6 Figure 8.8 shows the activity released to the package for the source term inventory from 

7 Table 8.4 for 15-year-cooled WE 15x15 fuel at 45 GWd/MTU.

8

9

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

11 15x15 (35GWd/MTU, 10-yrs-cooled fuel) for Bounding Release Fractions Specified 

12 in NUREG-1617

13
8-17

1

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

3 15x15 (45GWd/MTU, 15-yrs-cooled fuel) for Bounding Release Fractions Specified 

4 in NUREG-1617

5

6 8.3.3 Potential Release from GA-4 Package in MacArthur Maze Fire Scenario

7

8 Release rates from SNF packages are typically calculated for NCT and HAC using models 

9 based on guidance in NUREG/CR-6487, which contains models that reference ANSI standards 

10 for leakage tests on packages for shipment of radioactive materials (ANSI N14.5 1997). The 

11 analyses presented in the GA-4 SAR show that as long as the package seals remain intact, the 

12 package can be expected to meet all containment requirements, and potential releases from the 

13 package would be well below regulatory limits. However, as discussed in Section 8.2.1, the GA14 4 package seals are predicted to exceed operational temperature limits after approximately 1.5 

15 to 2.5 hours of exposure to the thermal conditions of the MacArthur Maze fire scenario. In 

16 addition, as discussed in Section 8.3.1, all fuel rods in all four assemblies contained within the 

17 GA-4 package are predicted to exceed temperatures at which burst rupture or creep rupture of 

18 the zircaloy cladding would be expected to occur. Therefore, there is the potential for leakage 

19 of radioactive material from the GA-4 package after this point in the fire transient.

20

21 Determining an appropriate leak rate for the package in the conditions predicted for the 

22 MacArthur Maze fire scenario presents an interesting challenge. The models for leak rates 

23 derived from the ANSI standard ANSI N14.5 are not based on the assumed seal conditions in 

24 this fire scenario, and there is very little information in the literature on leak rates associated with 

25 failed seals. In typical engineering applications, the leak rates of failed seals are unacceptable 

26 by definition, and their potential magnitude is of no practical interest. What little information to 

27 be found tends to focus rather narrowly on special applications where time-to-failure could be a 

28 critical design parameter (e.g., equipment that will be sent into orbit). In these types of studies, 
8-18

1 the focus is on the time interval to the point where the seal begins to leak, not on the leak rate 

2 itself, and the work is mainly interested in modes of seal failure or seal behavior prior to failure.

3

4 It was therefore necessary to develop a modeling approach to determine a reasonable bounding 

5 leak rate for the GA-4 package for the long portion of the transient following the time after 

6 assumed seal failure due to exceeding thermal operating limits. Section 8.3.3.1 presents the 

7 model developed for this analysis, and describes its application to the GA-4 package in the 

8 MacArthur Maze fire scenario. Section 8.3.3.2 presents the potential release calculations for 

9 the GA-4 package, based on the leak rate determined with this model.

10

11 8.3.3.1 Leak Rate Model for GA-4 Package without Seals

12

13 For leak rate modeling, the interface between the closure lid and end flange of the package 

14 body is of greatest significance. (There is also the potential for leakage paths through the gas 

15 sample valve/port and the drain valve/port; this is discussed in Section 8.3.3.2.) Failure of the 

16 seals in the fire scenario is conservatively treated in this evaluation as if the seals simply cease 

17 to exist after one hour of the fire duration. This timeframe conservatively bounds the interval of 

18 the estimated time when all seals are predicted to have exceeded operating temperature limits. 

19 The possibility of damaged seal material affecting the geometry of the leakage path is ignored. 

20 If it is assumed that there is no O-ring seal material remaining in the seal grooves of the lid and 

21 flange, and the only barrier to flow through the interface is the actual physical contact between 

22 these two components.

23

24 The closure lid and body flange both have smooth metal surfaces where the two components 

25 are in contact, and the closure bolts are torqued to a specified pre-load, such that there is a 

26 positive and essentially uniform clamping force at the interface. The evaluations presented in 

27 Section 7.3, investigating the response of the lid closure bolts to the extreme thermal 

28 environment of the MacArthur Maze fire scenario, show definitively that the bolts maintain a 

29 positive clamping force throughout the fire transient, including the long cooldown back to 

30 ambient conditions.

31

32 Flow of gas through the very narrow space between the closure lid and body flange can be 

33 treated as analogous to fluid flow through fractured material in which the local scale of motion 

34 can be approximated by the cubic law for flow between parallel plates. This is a simplified form 

35 of the momentum conservation equation, and is a function of the geometry of the flow path and 

36 the driving pressure difference between the package interior and the external environment 

37 (Brown 1987). A formulation of this relationship, expressed in cylindrical coordinates, is given 

38 by

39

QQLLLL = 2ππ � ddee

3

12μμ

� � ∆PP

ln( rroo⁄rrii) 40 �

41

42 where:

43 QLR = volumetric flowrate through the leakage path

44 de = equivalent gap between surfaces in contact

45 μ = viscosity of flowing gas

46 ΔP = driving pressure difference

47 ro = outer radius

48 ri = inner radius

49
8-19

1 The equivalent gap between the surfaces in contact is the critical unknown in the above 

2 equation, since all other parameters can be readily determined from the geometry of the GA-4 

3 package closure lid and flange, and the conditions calculated for the MacArthur Maze fire 

4 scenario with the thermal models. The actual gap is a function of the surface roughness of the 

5 components in contact and the clamping force holding them together. This gap cannot readily 

6 be estimated with any degree of certainty without knowing the exact microscale geometry of the 

7 surfaces involved. 

8

9 Therefore, an alternative approach was developed by considering another much simpler 

10 physical process in which the gap between two surfaces in direct contact has an important 

11 effect on physical behavior; the flow of heat between two surfaces in direct or very close 

12 contact. The thermal resistance between two such components is a strong function of the 

13 contact pressure and surface texture of the two surfaces, and is typically expressed in terms of 

14 the overall thermal contact resistance, as

15

RRtttt = ddee

kk 16

17 where:

18

Rt c = thermal contact resistance (m2 19 -K/W) as a function of contact pressure and 

20 surface texture for two surfaces in direct contact

21 de = equivalent gap between surfaces in contact

22 k = thermal conductivity of gas in spaces between contacting surfaces

23

24 Using thermal contact resistance data for stainless steel surfaces (Shajaefard and Goudarzi 

25 1987) as a function of contact pressure and surface roughness at the interface, and assuming 

26 helium gas in the very constrained spaces between the contacting surfaces, the above 

27 relationship can be used to determine an equivalent gap for the closure lid and package flange 

28 in the MacArthur Maze fire scenario.

29

30 The results presented in Section 7.3 for evaluations of the lid closure bolt response to the 

31 MacArthur Maze fire scenario using detailed FEA modeling provide a history of the clamping 

32 force between the closure lid and package body flange. Figure 8.9 shows the estimated 

33 equivalent gap after seal failure, based on the contact resistance as a function of the lid/flange 

34 contact pressure due to the clamping force, and the thermal conductivity of helium gas.
8-20

1

2 Figure 8.9. Equivalent Gap between Closure Lid and Package Body Flange after Seal Failure

3

4 As discussed in Section 7.3, the clamping force on the lid increases during the post-fire 

5 cooldown, due to differential thermal contraction between the nickel alloy closure bolts and the 

6 stainless steel lid and package body. The effect is to essentially close the gap entirely, for all 

7 practical purposes, by about 3.75 hours into the fire transient. This effectively limits the 

8 “window” of time in which material could leak out of the package to less than 3 hours. This has 

9 the effect of greatly reducing the potential for a substantial release of radioactive material from 

10 the package, as shown by the release evaluations in Section 8.3.3.2. 

11

12 8.3.3.2 Bounding Estimate of Potential Release from GA-4 Package

13

14 Using the leak rate model and equivalent gap width relationship presented in Section 8.3.3.1, a 

15 conservative bounding estimate was obtained for potential release of radioactive material from 

16 the GA-4 package in the MacArthur Maze fire scenario. The fluid viscosity of pure helium was 

17 used for this calculation, rather than attempting to quantify the viscosity of the mixture of helium 

18 and fission gases that would actually be in the package following the rod ruptures. This is a 

19 conservative assumption, since the viscosity of the mixture would be higher than the viscosity of 

20 pure helium. The difference between the mixture property and that of pure helium would in any 

21 case be expected to be small, since the gas released from the fuel rods would consist mainly of 

22 helium.

23

24 The pressure difference driving the volumetric flow through the interface between the package 

25 cavity and ambient was calculated assuming a constant external ambient pressure of 1 atm. 

26 The internal cavity pressure was calculated using the ideal gas law, based on the average gas 

27 temperature predicted with the ANSYS thermal model. The initial pressure in the cavity was 

28 assumed to be at the Maximum Normal Operating Pressure (MNOP) for the GA-4 package. 
8-21

1 This is a conservative initial pressure, as it corresponds to the pressure effect of 100% of the 

2 fuel rods in the package having ruptured, and the density change is determined for B&W 15x15 

fuel1 3 , which is the most limiting fuel configuration for the maximum operating pressure. This 

4 approach provides a bounding estimate of the cavity internal pressure throughout the fire 

5 transient, and avoids the complication of changing the gas density in the package at the 

6 predicted time of rod rupture in this analysis.

7

8 Figure 8.10 shows the predicted leak rate as a volumetric flow of helium gas through the 

9 equivalent gap. Two leak rate calculations were performed; a bounding case in which the 

10 package gas density was assumed to remain constant throughout the transient, and a more 

11 realistic case in which the change in gas density (and hence pressure) due to outflow of gas 

12 from the package was accounted for. The difference between the two cases is relatively small, 

13 due to the small leak rates predicted for this configuration with the closure lid clamped tightly to 

14 the package body flange throughout the transient.

15

16

17 Figure 8.10. Volumetric Leak Rate for GA-4 Package after Seal Failure

18

19

20 The release calculations were performed assuming the bounding leak rate over time (as shown 

21 in Figure 8.10), providing a bounding estimate of potential release from the package. The 

22 activity within the package cavity was assumed to be uniformly distributed within the gas, with 

23 all particulate (i.e., fuel fines and spalled CRUD) suspended in the gas as an aerosol. It is 

 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

1 assumed that all solid particles remain suspended in the gas. The total release of each 

2 component was calculated simply as the activity of that component times the volumetric fraction 

3 of gas escaping from the package. 

4

5 A number of additional conservatisms were incorporated into the release calculation, including 

6 the following assumptions

7

8 • the entire quantity of fuel and CRUD particulate was assumed to remain suspended in the 

9 gas within the cavity; the possibility of particulate settling or plating out on internal package 

10 structures was ignored

11

12 • the filtering effect of the equivalent gap size was neglected; the maximum size of the 

13 equivalent gap is only about 2 micrometers, and the release calculations do not consider that

14 a large percentage of the fuel and CRUD particulate simply could not escape from the 

15 package, due to the small size of the gap

16

17 These assumptions result in a very conservative estimate of the amount of activity that could 

18 escape from the package in the approximately 2.7 hours that the package could sustain a 

19 significant leakage.

20

21 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 22

23 values from 10 CRF 71, Appendix A.

24

25 The A2 value for a mixture of normal form material can be determined using the following 

26 relationship from 10 CFR 71, (Appendix A, Section IV.d), as

27

AA2 ffff mmmmmmmmmm mm = 1

�∑ ff( )

AA2( ) nn

ii=1 �

28

29

30 where 

31

32 n = number of radionuclides in mixture

f(i) = fraction of total mixture activity due to the i

th 33 component

A2(i) = A2 value for the i

th 34 component

35

 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

1 Using this approach, the A2 for the mixture of radionuclides in the estimated potential release 

2 from the GA-4 package is calculated as 88 Ci (3.25 TBq) for WE 15x15 fuel at 45 GWd/MTU, 

3 15 yrs cooling. The corresponding result for WE 15x15 fuel at 35 GWd/MTU, 10-yrs cooling is a 

4 mixture A2 of 143 Ci (5.3 TBq). The calculation of the mixture A2 for each fuel configuration 

5 includes all fission gas and particulate released from the fuel, plus the CRUD assumed to spall 

6 from the exterior surfaces of the rods.

7

8 Based on the leak rate model, the total release from the package is estimated as 21 Ci

9 (0.78 TBq) for the higher burnup fuel, and as 24.5 Ci (0.91 TBq) for the lower burnup fuel. 

10 Expressed as an A2 fraction, relative to the mixture A2 for each configuration, these release 

11 rates are 0.24 and 0.17, respectively. Therefore, the bounding estimate of the total release from 

12 the package is 0.24 of the mixture A2 calculated assuming WE 15x15 fuel at 45 GWd/MTU, 

13 15 yrs cooling. As mentioned above, if the effect of particulate settling and the restriction of 

14 large particulate from passing through a small gap were taken into account, the release 

15 estimate would be significantly reduced. 

16

17 The evaluations of potential release from the GA-4 package assume that the estimated release 

18 by way of the closure lid is sufficiently conservative to be bounding on the possible contribution 

19 of leakage through the drain valve/port and gas sampling valve/port, which form part of the 

20 containment boundary of the GA-4 package. These components also contain seals that would 

21 be expected to exceed their operating temperature limits in this fire scenario, as discussed in 

22 Section 8.2 above. However, these penetrations of the package are less than an inch in 

23 diameter, compared to the approximately 2-ft diameter of the closure lid rim, and therefore do 

24 not provide a significant increase in the area available for potential leakage. In addition, the 

25 ports consist of long and convoluted flow paths that would tend to filter any particulate that might 

26 be carried through them to the ambient environment. The gas sample port is effectively blocked 

27 by the sample valve itself, which in addition to having primary and secondary O-ring seals, is 

28 threaded into place over a length of several inches. Also, for transport conditions, the outer face 

29 of the gas sample port is plugged with a threaded cover that extends to a depth of more than an 

30 inch. Similarly, the drain port is plugged by the drain valve, and capped with a threaded drain 

31 valve cover and port plug.

32

33 Based on the geometry of the valve/ports in this package, it is reasonable to assume that 

34 leakage from the package at these locations due to failed seals would be much less likely to be 

35 significant compared to leakage through for the much larger area and more direct flow path of 

36 the closure lid seal region. The conservative assumptions regarding the amount of material that 

37 could be transported out of the package through the lid closure/flange equivalent gap are 

38 sufficient to bound any possible contribution of the valve/port leakage paths. It is therefore 

39 justifiable to neglect the effect of the valve/ports, without compromising the conservatism of the

40 estimated leak rate and total package release calculations.

41

9-1

1 9.0 OVERALL SUMMARY AND CONCLUSIONS

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

3 transportation of spent nuclear fuel assemblies under NCT and for HAC. These requirements

4 (10 CFR 71) conservatively bound the expected range of conditions that an SNF package might 

5 be subjected to in the course of its service life. However, real-world accidents of greater 

6 severity are certainly possible, and rare as they may be, the NRC has proactively undertaken 

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

8 were such an accident ever to involve an SNF package. 

9

10 Two previous studies of transportation accidents, one resulting in a fire in a railroad tunnel 

11 (NUREG/CR-6886 2009) and one in a highway tunnel (NUREG/CR-6894 2007) were 

12 undertaken with three different SNF package designs. Based on conservative scenarios 

13 constructed from these real-world fire conditions, the results of these studies have shown that 

14 the design basis for SNF packages is sufficiently robust for them to survive such beyond15 design-basis conditions without adverse consequences to public safety. In all cases evaluated, 

16 the modeling results showed that the various SNF packages would be expected to maintain 

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

18 boundary sufficiently to limit potential release of radioactive material from the packages to within 

19 regulatory bounds for accident conditions.

20

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

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

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

24 collapse onto the roadway where the fire was burning. Since this was a highway accident, the 

25 only type of SNF package that could potentially be involved would be a LWT package. The 

26 General Atomics GA-4 LWT transportation package was selected for this investigation, mainly 

27 because it can carry a relatively large payload for an over-the-road transportation package, and 

28 therefore the potential consequences of package failure could be more severe than for 

29 packages with smaller payload capacities. The GA-4 package is designed to transport up to 

30 four intact PWR spent fuel assemblies, with a maximum total package decay heat load of 

31 2.5 kW.

32 Bounding Scenario for the MacArthur Maze Accident

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

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

35 lasted approximately 108 minutes, consuming the entire load of gasoline fuel. The heat from 

36 the fire caused two sections of the overhead I-580 freeway to collapse onto the lower roadway, 

37 the first falling at approximately 17 minutes into the fire, the second collapsing on only one end, 

38 and reaching its final configuration by about 37 minutes. (Refer to Figure 1.1, which shows the 

39 configuration of the collapsed roadway, in an image taken in daylight the next day, after the fire 

40 was out. See Section 2.0 for a detailed discussion of the fire scenario, with images from a video 

41 of the fire.)

42

43 Based on fire modeling with the FDS code, and physical examination of material samples 

44 obtained from the damaged highway girders and the remnants of the tanker truck, a bounding 

45 fire scenario was defined for the thermal and structural evaluations of the potential effects of this 

46 fire on an SNF package. The complex and dynamic fire conditions are represented as a fully 

47 engulfing pool fire at 2012°F (1100°C) prior to the overhead roadway collapse, and as a slightly 

48 smaller and less severe fully engulfing pool fire at 1652°F (900°C) after the roadway collapse. 
9-2

1 These temperatures represent conservative bounding values for open pool hydrocarbon fires for 

2 any possible configuration of both the pre-collapse and post-collapse fire pools in this accident. 

3

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

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

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

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

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

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

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

11 package during the cooldown phase of the transient.

12 Thermal and Structural Modeling Approach and Summary of Results

13 Detailed thermal models of the GA-4 package were constructed for the ANSYS and COBRA14 SFS codes, for transient evaluations to determine the temperature response of the package to 

15 the fire scenario, including the long post-fire cooldown transient. The initial condition of the 

16 package at the start of the fire scenario was defined as steady-state NCT. Additional detailed 

17 structural and thermal-structural models were also developed using ANSYS and LS-DYNA for 

18 the roadway and package, for evaluation of the package response to the effect of the roadway 

19 falling on it.

20

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

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

23 loading of the SNF package due to the overhead roadway segment falling on it. These 

24 evaluations showed that the worst that the overhead spans could do to the package imposed 

25 relatively innocuous loads on the stainless steel body and DU gamma shield compared to the 

26 HAC structural loading that the package is designed to withstand. At a nominal fully loaded 

27 weight of approximately 55,000 lb (nearly 28 tons), the package itself falling from a height of 

28 30 ft (9 meters) – the HAC package drop scenario – would be expected to do far more damage

29 than the roadway falling on it, even with the added impact of the projecting “blades” of the steel 

30 girders. The only real challenge of the overhead roadway drop in the fire scenario is that the 

31 impact is postulated to occur with the package at higher temperatures than are typically 

32 assumed in the structural analyses for HAC scenarios. This could potentially make the package 

33 more vulnerable to structural damage, due to the reduction in the strength of steel with 

34 increasing temperatures. However, the steel girders of the overhead span suffer more from this 

35 problem, and the weight of the overhead roadway concrete is not sufficient to impart significant 

36 loading to damage the package in any way.

37

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

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

40 Issues of bolt performance were further complicated by the use of thread inserts in all bolt 

41 attachments in the package, in which helical coils of Type 304 stainless steel fill the interface 

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

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

44 strength-versus-temperature properties of the three metals involved, results in a time-and45 temperature dependent history of force on the bolts that raised the possibility that the impact 

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

47 there could be a loss of clamping force between the lid and the package body during the post48 fire cooldown.

49
9-3

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

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

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

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

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

6 response of the lid closure bolts to the fire scenario were undertaken with LS DYNA, and the 

7 modeling accuracy was verified with classic bolt equation methods. These evaluations show 

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

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

10 (108 minutes) and the very long cooldown period of approximately 400 hours, back to post-fire 

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

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

13 response of the GA-4 package to this fire scenario, because the thermal evaluation shows that 

14 the seals exceed their rated temperature limits within the first hour or so of the transient, and 

15 this metal-to-metal contact becomes the containment boundary of the package.

16

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

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

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

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

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

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

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

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

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

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

27 throughout the package. 

28

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

30 1350-1400°F (732-760°C), and occurs approximately 3 hours after the end of the fire. In 

31 addition, temperatures in the regions of the package seals exceed the seal material operating 

32 temperature limits for most of the fire transient and for several hours of the post-fire cooldown 

33 transient.

34 Fuel Rod Performance Evaluation

35 Based on the predicted fuel cladding temperatures from the COBRA-SFS modeling, fuel 

36 performance was evaluated by direct comparison to fuel rod burst data as a function of cladding 

37 hoop stress and temperature. In addition to comparison to relevant data, predicted fuel rod 

38 rupture temperatures were obtained using the burst rupture model in the FRAPTRAN-1.4 code 

39 (NUREG/CR-7023 2011). Creep rupture is considered a possible alternative mechanism of 

40 failure for spent fuel rods. To evaluate this possibility, a separate analysis was performed with a 

41 creep rupture model for the temperatures predicted for the hottest rod in the MacArthur Maze 

42 fire scenario, using the FRAPCON-3.4 code (NUREG/CR-7022 2011) in conjunction with the 

43 DATING code (Simonen and Gilbert 1988), which has been updated with creep coefficients 

44 from creep tests on irradiated cladding (Gilbert et al. 2002).

45

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

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

48 1097°F (592°C) using LOCA burst strain modeling (FRAPTRAN) and 1229°F (665°C) using 

49 creep rupture modeling (FRAPCON/DATING). Applicable experimental data (NUREG/CR-
9-4

1 0344) yields measured rupture temperatures in the range 1205-1256°F (652-680°C). The burst 

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

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

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

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

6 fire scenario. 

7 Potential Radiological Consequences

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

9 conditions of the MacArthur Maze fire scenario will not exceed the design basis of the package, 

10 which is well within the regulatory limits for hypothetical accident conditions. The neutron 

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

12 assumption for this package in all HAC analyses. The more severe conditions of the MacArthur 

13 Maze fire can do no more damage to the GA-4 package neutron shield than is assumed a priori

14 in the HAC analyses. The gamma shielding for the GA-4 is provided by a layer of DU within the 

15 stainless steel package body. The shielding function of this material is not affected by the 

16 higher temperature it is predicted to reach in the MacArthur Maze fire scenario. There is no 

17 credible scenario in this fire accident that could result in neutron and gamma dose rates from 

18 the design-basis GA-4 package exceeding the regulatory limits for accident conditions.

19

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

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

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

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

24 package cavity. In addition, 100% spalling of CRUD from the external surfaces of the fuel rods 

25 is an assumed for all accident analyses for SNF packages, per NRC guidance. Therefore, it 

26 must be assumed that there is material available in the package cavity that could be released 

27 through the failed seals. But because the lid closure bolts maintain positive clamping force 

28 throughout the transient, it is not physically possible for very much of it to actually escape. 

29 Conservative and bounding modeling assumptions show that the maximum possible release

30 total release is 0.24 of the A2 quantity calculated for total activity of the mixture of radionuclides 

31 (comprised of fission gases, fuel particulate and CRUD) released from the package. The 

32 regulatory limit specifies a maximum allowable release rate of an A2/week. The predicted total 

33 release estimate of approximately one-fourth of a mixture A2 is below the prescribed limit for 

34 safety, and indicates that the potential release from this package in the MacArthur Maze fire 

35 scenario would not pose a risk to public health and safety.

36

37
10-1

1 10.0 REFERENCES

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7

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29 Burian RJ. 1985. Response of Spent LWR Fuel to Extreme Environments. SAND85-7213.

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32 CHP - California Highway Patrol. 2007. California Highway Patrol Multi-Disciplinary Accident 

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40 Dow Chemical Company. 2003. A Guide to Glycols. Available at 

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47
10-2

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4

5 General Atomics. 1998. GA-4 Legal Weight Truck Spent Fuel Shipping Cask Safety Analysis 

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8

9 Gilbert ER, CE Beyer, EP Simonen, and PG Medvedev. 2002. “Update of CSFM Creep and 

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14 Guyer EC and DL Brownell, eds. 1989. Handbook of Applied Thermal Design. McGraw-Hill, 

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HEXCEL Composites. 1999. HexWeb 17 TM Honeycomb Attributes and Properties. Available at 

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20

21 Johnson AB and ER Gilbert. 1983. Technical Basis for Storage of Zircaloy-clad Spent Fuel in 

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23

24 Johnson AB, ER Gilbert, and RJ Guenther. 1983. Behavior of Spent Nuclear Fuel and 

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27

Juvinall RC and KM Marshek. 1991. Fundamentals of Machine Component Design, 2nd 28

29 Edition. John Wiley & Sons, New York.

30

Kreith F and MS Bohn. 2001. Principles of Heat Transfer.6th 31 ed. Brooks/Cole, Forest Grove, 

32 California.

33

34 Livermore Software Technology Company. 2007. LS-DYNA Keyword User’s Manual, Version 

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36

37 McGrattan KB, B Klein, S Hostikka, and J Floyd. 2008. Fire Dynamics Simulator (Version 5), 

38 User’s Guide. Special Publication 1019-5, National Institute of Standards and Technology, U.S. 

39 Department of Commerce, Washington, D.C.

40

41 Michener TE, DR Rector, JM Cuta, RE Dodge, and CW Enderlin. 1995. COBRA-SFS: A 

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43 Pacific Northwest Laboratory, Richland, Washington.

44
10-3

1 NFPA - National Fire Protection Association. 2008. NFPA 502: Standard for Road Tunnels, 

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Product Data Bulletin1 3 . High Performance Alloys, 444 Wilson St. Tipton, IN 46072. Available at 

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5

6 NRC – U.S. Nuclear Regulatory Commission. 2008. Analysis of Structural Material Exposed to 

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9 Texas, published by U.S. Nuclear Regulatory Commission, Washington, D.C.

10

11 NUREG-0170. 1977. Final Environmental Statement on the Transportation of Radioactive 

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13

14 NUREG-1617. 2000. Standard Review Plan for Transportation Packages for Spent Nuclear 

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16

17 NUREG/CR-0344. 1980. Embrittlement Criteria for Zircaloy Cladding Applicable to Accident 

18 Situations in Light-Water Reactors: Summary Report. U.S. Nuclear Regulatory Commission, 

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20

21 NUREG-0630. 1980. Cladding Swelling and Rupture Models for LOCA Analysis. U.S. Nuclear 

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23

24 NUREG/CR-0772. 1980. Fission Product Release from Highly Irradiated LWR Fuel. U.S. 

25 Nuclear Regulatory Commission, Washington, D.C.

26

27 NUREG/CR-4829. 1987. Shipping Container Response to Severe Highway and Railway 

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29

30 NUREG/CR-6487. 1996. Containment Analysis for Type B Packages Used to Transport 

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33

34 NUREG/CR-6672. 2000. Re-examination of Spent Fuel Shipment Risk Estimates. U.S. 

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36

37 NUREG/CR-6886, Rev 2. 2009. Spent Fuel Transportation Package Response to the 

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40 NUREG/CR-6894, Rev. 1. 2007. Spent Fuel Transportation Package Response to the 

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42

43 NUREG/CR-7022, Vol. 1. 2011. FRAPCON-3.4: A Computer Code for the Calculation of 

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45 Regulatory Commission, Washington, D.C.

46

 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

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5 D.C.

6

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10

11 Sanders TL, KD Seager, YR Rashid, PR Barrett, AP Malinauskas, RE Einziger, H Jordan, 

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16

17 Shajaefard ML and K. Goudarzi. 2008. “The Numerical Estimation of Thermal Contact 

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19

20 Simonen EP and ER Gilbert. 1988. DATING – A Computer Code for Determining Allowable 

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22 Pacific Northwest Laboratory, Richland, Washington.

23

24 Society of Fire Protection Engineers. 2002. SFPE Handbook of Fire Protection Engineering.

2nd 25 ed., National Fire Protection Association, Quincy, Massachusetts. (Specifically, the chapter 

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Fire Protection Engineering, 2nd 27 ed.)

28

29 Society of Fire Protection Engineers. 2008. SFPE Handbook of Fire Protection Engineering.

4th 30 ed. National Fire Protection Association, Quincy, Massachusetts. (Specifically, the chapter 

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33

34 Wilmot EL. 1981. Transportation Accident Scenarios for Commercial Spent Fuel. SAND80-

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36

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38 Concrete. Prepared by Wiss, Janney, Elstner Associates, Inc., for the Office of Rigid Pavement 

39 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

1 APPENDIX A

2

3 MATERIAL PROPERTIES FOR COBRA-SFS MODEL OF GA4 

4 PACKAGE

5

6 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

7

8 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

9
A-2

1 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)

2

3 Emissivity values for thermal radiation exchange were obtained from Table 3.2-3 of the GA-4 

4 SAR (General Atomics 1998). However, the emissivity of package surfaces exposed to the fire 

5 was conservatively represented with a value of 0.9, rather than the “0.8 or 0.85” listed in the 

6 SAR. Table A.4 summarizes the emissivity values used for the XM-19 stainless steel 

7 components during the various phases of the fire scenario.

8

9 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

10

Temperature-dependent thermal conductivity (in units Btu/hr-ft2 11 -°R) for XM-19 stainless steel 

12 was evaluated in the COBRA-SFS model using a linear regression fit to ATI 50 thermal 

13 conductivity data (see Appendix B for the material data sheet). The relationship from this fit is

14
A-3

1 k XM −19 = a0 + a1T (A-1)

2

3 where

4

5 a0 = 5.4446

6 a1 = 0.0047

7 T = material temperature (°R)

8

9 The relationship in Eq. (A-1) is a polynomial curve fit to the same data used to derive the linear 

10 equation presented in the GA-4 SAR, which has the form

11

12 ks = a0 + a1T (A-2)

13

14 where 

15

16 a0 = 3.6 (empirical coefficient)

17 a1 = 0.00532 (empirical coefficient)

18 T = material temperature (°R)

19

20 These two equations give essentially identical results for temperatures below about 1000°F 

21 (538°C), but Equation (A-1) is more conservative by 15-20% at the high temperatures 

22 encountered in the MacArthur Maze fire scenario.

23

24 The formula for the effective conductivity used to model natural convection in the liquid neutron 

25 shield is documented in the GA-4 SAR as

26

( ) 2 Pr

Pr

1

0 .

a

NSliq D

eff

a

a k Gr

k

+

27 = (A-3)

28

29 where

30

31 a0 = 0.135 (empirical coefficient)

32 a1 = 1.36 (empirical coefficient)

33 a2 = 0.278 (empirical coefficient)

34 kNSliq. = thermal conductivity of NS liquid (propylene glycol/water mixture)

35 Pr = Prandtl Number

36 GrD = Grashoff number, using thickness of neutron shield tank as characteristic 

37 length

38

39
A-4

1

2 © 2009 ATI Allegheny Ludlum

3 Figure A.1. Source for Thermal Conductivity of XM-19 Stainless Steel

4

5
A-5

1

2 Figure A.2. Source for Thermal Conductivity of Depleted Uranium

MATERIAL PROPERTIES FOR ANSYS MODEL

OF GA4 PACKAGE

B-1

1 APPENDIX B

2

3 MATERIAL PROPERTIES FOR ANSYS MODEL OF GA4

4 PACKAGE

5

6 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

7

8 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

9
B-2

1 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

2

3 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

4
B-3

1 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

2

3 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

4

5 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

6

7 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

8
B-4

1 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

2

3 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

4

5 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

6

7 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

8
B-5

1 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

2

3 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 -

4

5 Depleted Uranium – See Figure A.2 in Appendix A

6

7 Neutron Shield – Effective Conductivity Calculations:

8

9 An empirical relationship for effective conductivity incorporating the effects of both conduction 

10 and convection was used to determine heat exchange through the liquid neutron shield. The 

11 effective conductivity of the fluid within the tank is based on heat transfer between two 

12 concentric cylinders. Sensitivity studies were performed on this correlation to verify that it is 

13 applicable to the neutron shield tank geometry, in which the inner surface of the tank has a 

14 square cross-section with rounded corners, rather than a circular cross-section. This correlation 

15 produces reasonable values of keff, and the transient conditions are generally within its 

16 applicable range. The correlation relates the Nusselt number to the ratio of the effective 

17 conductivity over the actual conductivity, and is expressed as

18

keff

kc

= Nu = 0.386 Pr

0.861+ Pr

 

   

  

0.25

Rac

0.25 19 (B-1)

20

21 where

22

23 keff = effective thermal conductivity of material in node

24 kc = thermal conductivity of motionless fluid in node

25 Nu = Nusselt number

26 Pr = Prandtl number

27 Rac = modified Rayleigh number
B-6

1 The modified Rayleigh number is defined as:

2

Rac = [ln(Do Di)]

4

L3 Di

−0.6 + Do−0.6

[ ] 3 5 Ra (B-2)

4

5 where

6

7 Do = annulus outer diameter

8 Di = annulus inner diameter

9 Ra = Rayleigh number

10 L = (Do − Di)/2

11

12 The Rayleigh number is based on the temperature difference across the annular gap and is 

13 expressed as:

14

Ra = gβ(Ti − To )L3

να

15 (B-3)

16

17 where

18

19 g = acceleration of gravity

20 Ti = inner surface temperature

21 To = outer surface temperature

22 β = thermal expansion coefficient

23 α = thermal diffusivity of fluid

24 ν = kinematic viscosity of fluid

25

26 Using the correlations listed above, a macro was written to calculate the effective conductivity 

27 after each solution step within the transient model. For conditions below 276°F, the properties 

28 of 56% propylene glycol and water were used to calculate the effective conductivity. Once the 

29 maximum temperature within the tank exceeded 276°F, the properties of air were used to 

30 determine the thermal conductivity. 

31

32 Verification of Effective Conductivity Model for GA-4 Neutron Shield 

33 Configuration

34

35 The effective conductivity model described above is based on experimental data for natural 

36 convection mixing of fluid between horizontal concentric cylinders. The neutron shield tank of 

37 the GA-4 package consists of an inner surface formed by the package body, which is a square 

38 with rounded corners, and an outer cylindrical shell. To verify that this empirical model could be 

39 applied to the GA-4 package neutron shield geometry, the correlation predictions were 

40 compared to results from a computational fluid dynamics model.

41

Calculations were performed with Star-CD1 42 , for a 2-D “slice” model at the midplane of the 

43 package, using two basic configurations to model the GA-4 neutron shield tank. In one model, 

44 the neutron shield fluid region is represented as a solid material with thermal conductivity 

 1 STAR-CD, Version 4.14 Methodology, Computational Dynamics Ltd. 2010.
B-7

1 determined using the relationship for the effective thermal conductivity, as defined in Eq. (B-1) 

2 above. In the other model, the neutron shield fluid region is represented as a liquid, with the 

3 fluid properties of the propylene-glycol/water mixture reported in the GA-4 SAR [11].

4

5 The results of this evaluation are summarized in Table B.15, with comparisons of the maximum 

6 and minimum predicted temperatures obtained with the Star-CD model for all cases considered. 

7 All calculations in this evaluation were performed at NCT. The maximum temperature is the 

peak temperature in the fuel region1 8 , and the minimum temperature is the minimum temperature 

9 on the package outer shell surface. As shown by the results in Table B.15, a computation mesh 

10 that was appropriate for the neutron shield represented as a solid was not sufficient resolution 

11 for the CFD model. The number of computational elements required was approximately two 

12 orders of magnitude larger. 

13

14 Star-CD results for the case with the neutron shield represented as a solid material and for the 

15 case with the neutron shield represented as a fluid (with an appropriately refined mesh) are 

16 shown graphically with color thermographs in Figure B.1. Overall, this evaluation has shown 

17 that the effective conductivity model predicts temperatures that are results are consistent with 

18 the CFD model results. There is also some indication that the effective conductivity model may 

19 yield results that are slightly conservative.

20 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

21

22

 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

1

2 (a) (b)

3 Figure B.1. Mid-plane Temperature Distributions Predicted with Star-CD Model of GA-4 

4 Package at NCT: (a) solid material neutron shield with effective conductivity model, 

5 and (b) liquid neutron shield with (8x8) refined mesh

6




NUREG/CR

-7206

DRAFT

Spent Fuel Transportation Package Response to the MacArthur 

Maze Fire Scenario