DOT/FRA/ORD-24-05 Final Report | February 2024

Improving Tank Car Crash Energy Management (CEM) 

i 

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PLEASE DO NOT RETURN YOUR FORM TO THE ABOVE ADDRESS.

1. REPORT DATE (DD-MM-YYYY)

2/1/24 

2. REPORT TYPE

Technical Report 

3. DATES COVERED (From - To)

08/12/2020 to 08/31/2022 

4. TITLE AND SUBTITLE

Improving Crash Energy Management (CEM) of Tank Cars 

5a. CONTRACT NUMBER

693JJ620C000011 

5b. GRANT NUMBER

5c. PROGRAM ELEMENT NUMBER

6. AUTHOR(S)

Philip S. Taylor – ORCID: 0000-0002-8043-6546

Juan Londono – ORCID: 0000-0003-2467-4686l 

Pawel Woelke – ORCID: 0000-0003-0114-9635

John W. Hutchinson – ORCID: 0000-0001-6435-3612

David Tyrell – ORCID: 0000-0001-8156-0277

David Jeong – ORCID: 0000-0003-3514-5739

5d. PROJECT NUMBER

5e. TASK NUMBER

5f. WORK UNIT NUMBER

7. PERFORMING ORGANIZATION NAME(S) AND ADDRESS(ES)

Thornton Tomasetti Inc 

120 Broadway, 

New York NY 10271

8. PERFORMING ORGANIZATION REPORT 

NUMBER

NN20212.00-R-01

9. SPONSORING/MONITORING AGENCY NAME(S) AND ADDRESS(ES)

U.S. Department of Transportation 

Federal Railroad Administration 

Office of Railroad Policy and Development 

Office of Research, Development and Technology 

Washington, DC 20590

10. SPONSOR/MONITOR'S ACRONYM(S)

11. SPONSOR/MONITOR'S REPORT 

NUMBER(S) 

DOT/FRA/ORD-24-05 

12. DISTRIBUTION/AVAILABILITY STATEMENT

This document is available to the public through the FRA website. 

13. SUPPLEMENTARY NOTES

COR: Francisco González, III 

14. ABSTRACT

The Federal Railroad Administration sponsored a 2-year research effort between August 2020 and August 2022 to investigate multiple 

mitigation concepts for hazmat tank cars, aimed at improving their puncture resistance. The team first developed and validated a finite 

element (FE) model of a full-scale tank car subject to shell (i.e., side) impact. Researchers then evaluated the effectiveness of different 

mitigation solutions by exercising the validated FE model. Validated FE models were used to explore different solutions including key 

material properties of tank car steel, the effects of changing shell thickness, the effects of splitting a single thick plate into multiple plates, 

and internal liner solutions. Results indicated the key parameter to improving puncture resistance is to increase total steel thickness; 

however, a double-tank arrangement offers improved benefit over an equivalent monolithic shell because of the high strength or high 

ductility steel used for the outer layer. Moreover, FE analyses indicated the ductility of steel in plane strain bending reduces with 

increasing plate thickness; the resistance of thick shells may be over-estimated if this is not considered.

15. SUBJECT TERMS

Rolling stock research, FEA, mitigation, hazmat, tank car, puncture resistance, TC128-B 

16. SECURITY CLASSIFICATION OF: 17. LIMITATION OF 

ABSTRACT

18. NUMBER 

OF PAGES

136

19a. NAME OF RESPONSIBLE PERSON

Francisco González, III a. REPORT b. ABSTRACT c. THIS PAGE

19b. TELEPHONE NUMBER (Include area code)

202-493-6076

Standard Form 298 (Rev. 8/98)

Prescribed by ANSI Std. Z39.18
ii

METRIC/ENGLISH CONVERSION FACTORS

ENGLISH TO METRIC METRIC TO ENGLISH

LENGTH (APPROXIMATE) LENGTH (APPROXIMATE)

1 inch (in) = 2.5 centimeters (cm) 1 millimeter (mm) = 0.04 inch (in) 

1 foot (ft) = 30 centimeters (cm) 1 centimeter (cm) = 0.4 inch (in) 

1 yard (yd) = 0.9 meter (m) 1 meter (m) = 3.3 feet (ft) 

1 mile (mi) = 1.6 kilometers (km) 1 meter (m) = 1.1 yards (yd)

1 kilometer (km) = 0.6 mile (mi)

AREA (APPROXIMATE) AREA (APPROXIMATE)

1 square inch (sq in, in2) = 6.5 square centimeters (cm2) 1 square centimeter (cm2) = 0.16 square inch (sq in, in2) 

1 square foot (sq ft, ft2) = 0.09 square meter (m2) 1 square meter (m2) = 1.2 square yards (sq yd, yd2) 

1 square yard (sq yd, yd2) = 0.8 square meter (m2) 1 square kilometer (km2) = 0.4 square mile (sq mi, mi2)

1 square mile (sq mi, mi2) = 2.6 square kilometers (km2) 10,000 square meters (m2) = 1 hectare (ha) = 2.5 acres

1 acre = 0.4 hectare (he) = 4,000 square meters (m2) 

MASS - WEIGHT (APPROXIMATE) MASS - WEIGHT (APPROXIMATE)

1 ounce (oz) = 28 grams (gm) 1 gram (gm) = 0.036 ounce (oz)

1 pound (lb) = 0.45 kilogram (kg) 1 kilogram (kg) = 2.2 pounds (lb)

1 short ton = 2,000 pounds (lb) = 0.9 tonne (t) 1 tonne (t) =

=

1,000 kilograms (kg) 

1.1 short tons

VOLUME (APPROXIMATE) VOLUME (APPROXIMATE)

1 teaspoon (tsp) = 5 milliliters (ml) 1 milliliter (ml) = 0.03 fluid ounce (fl oz)

1 tablespoon (tbsp) = 15 milliliters (ml) 1 liter (l) = 2.1 pints (pt) 

1 fluid ounce (fl oz) = 30 milliliters (ml) 1 liter (l) = 1.06 quarts (qt) 

1 cup (c) = 0.24 liter (l) 1 liter (l) = 0.26 gallon (gal) 

1 pint (pt) = 0.47 liter (l) 

1 quart (qt) = 0.96 liter (l)

1 gallon (gal) = 3.8 liters (l)

1 cubic foot (cu ft, ft3) = 0.03 cubic meter (m3) 1 cubic meter (m3) = 36 cubic feet (cu ft, ft3) 

1 cubic yard (cu yd, yd3) = 0.76 cubic meter (m3) 1 cubic meter (m3) = 1.3 cubic yards (cu yd, yd3)

TEMPERATURE (EXACT) TEMPERATURE (EXACT)

[(x-32)(5/9)] °F = y °C [(9/5) y + 32] °C = x °F

QUICK INCH - CENTIMETER LENGTH CONVERSION

0 1 2 3 4 5

Inches

Centimeters 0 1 2 3 4 5 6 987 1110 1312

QUICK FAHRENHEIT - CELSIUS TEMPERATURE CONVERSIO

 -40° -22° -4° 14° 32° 50° 68° 86° 104° 122° 140° 158° 176° 194° 212°

 

°F

 °C -40° -30° -20° -10° 0° 10° 20° 30° 40° 50° 60° 70° 80° 90° 100°

For more exact and or other conversion factors, see NIST Miscellaneous Publication 286, Units of Weights and 

Measures. Price $2.50 SD Catalog No. C13 10286
iii

Acknowledgements 

The authors gratefully acknowledge the technical contributions and guidance provided by 

Michael Carolan, Volpe National Transportation Systems, and Francisco González, III FRA, 

throughout the research. In particular, previous work undertaken by Michael Carolan and 

subsequent guidance in developing a robust mechanism for developing and validating a full-scale

finite element model of tank car puncture was invaluable to the project. 
iv

Contents

Executive Summary ........................................................................................................................ 1

1. Introduction ................................................................................................................. 3

1.1 Background ................................................................................................................. 3

1.2 Objectives .................................................................................................................... 7

1.3 Overall Approach ........................................................................................................ 7

1.4 Scope ........................................................................................................................... 8

1.5 Organization of the Report .......................................................................................... 8

2. Development of a Full-Scale Finite Element Model ................................................... 9

2.1 Full-Scale Tests ........................................................................................................... 9

2.2 Tests 6 and 8 Overview ............................................................................................. 10

2.3 Full-scale FE Model .................................................................................................. 13

3. Sub-Model ................................................................................................................. 31

3.1 Quarter Symmetry Model .......................................................................................... 31

3.2 Test 6 Response ......................................................................................................... 32

3.3 Test 8 Response ......................................................................................................... 34

4. Mitigation .................................................................................................................. 36

4.1 Key Constraints ......................................................................................................... 36

4.2 Mitigation – Enhanced TC128-B .............................................................................. 37

4.3 Mitigation – Shell Thickness ..................................................................................... 44

4.4 Mitigation – Material Properties ............................................................................... 48

4.5 Mitigation – Multi-Plate Solutions ............................................................................ 58

4.6 Mitigation – Elastomeric Liners ................................................................................ 72

4.7 Mitigation – Blunting ................................................................................................ 77

5. Investigation of the Relationship between Plate Thickness and Ductility ................ 80

5.1 Gurson-Tvergaard-Needleman (GTN) Model ........................................................... 80

5.2 Verification of Plane Strain Tension and Bending Models ....................................... 80

5.3 Modified Plane Strain Tension and Bending Models ............................................... 86

5.4 Plate Thickness vs Ductility Results Summary ......................................................... 91

6. Feasibility, Scalability and Cost ................................................................................ 94

6.1 Enhanced TC128-B ................................................................................................... 94

6.2 High Strength Steel ................................................................................................... 94

6.3 High Ductility Steel ................................................................................................... 95

6.4 Thicker Tank Shell .................................................................................................... 95

6.5 Multi-Plate Solution .................................................................................................. 96

6.6 Liners ......................................................................................................................... 97

6.7 Blunting Solutions ..................................................................................................... 97

7. Conclusion ................................................................................................................. 98

8. References ............................................................................................................... 100

Appendix A. Material Properties ................................................................................................ 104
v 

Appendix B. TC128-B Material Properties ................................................................................ 109

Appendix C. Full-Scale Model, Potentiometer Results .............................................................. 113

Appendix D. Polyurea Material Model ....................................................................................... 117
vi 

Illustrations

Figure 1. Left to Right: Macdona, TX, Accident – Head Puncture [9]; Granitville, SC, Accident – 

Shell Puncture [10]; Lac-Mégantic, Canada, Car Pileup [8] .................................................. 4

Figure 2. Derailment Pileups, Left To Right: Minot, ND, 2002; Lac-Mégantic Car Pileup .......... 5

Figure 3. Generalized Impact Scenarios: (a) Head – CFR §179.16 [13], and (b) Shell – Ongoing 

Research [15] .......................................................................................................................... 6

Figure 4. Schematic Illustration of Integrated Conceptual Design [17] ......................................... 6

Figure 5. Use of Sandwich Panels to Improve Tank Car Integrity [18], [19], [22] ........................ 7

Figure 6. Example DOT-105 Tank Car Design Specification [32] .............................................. 10

Figure 7. Photograph of a Ram Car and Head Fitted with 6 x 6-inch Indenter [30] .................... 11

Figure 8. Tank Support Skid System [30] .................................................................................... 11

Figure 9. Deformed Tank Just Prior to Puncture (12 x 12-inch Indenter) [30] ............................ 12

Figure 10. Deformed Tank Just Prior to Puncture (6 x 6-inch Indenter) [30] .............................. 12

Figure 11. Overview of Full-scale FE Model, Test 6 ................................................................... 13

Figure 12. Overview of Full-scale FE Model from Behind, Test 6 .............................................. 13

Figure 13. Elevation Showing Tank Shell Geometry ................................................................... 14

Figure 14. Isometric View of Tank Car Assembly, with Sections of Jacket (blue) And Foam 

(green) Removed to Show Tank Shell with a Local Solid Region ....................................... 14

Figure 15. Indenter Part and Mesh ................................................................................................ 15

Figure 16. Cross-section Through Tank Car FE Model Showing Indenter ~0.6 inch Away ....... 15

Figure 17. Overview of Foam Insulation (with and without mesh shown) .................................. 16

Figure 18. True Plastic Stress Strain Input Data for Foam Insulation [30] .................................. 16

Figure 19. Membrane Part (with and without mesh) .................................................................... 17

Figure 20. Water Boundary (red) .................................................................................................. 18

Figure 21. Jacket (with and without mesh) ................................................................................... 19

Figure 22. Locally Refined Mesh in Jacket at Impact Location ................................................... 19

Figure 23. Overview of Tank Shell in FE Model ......................................................................... 20

Figure 24. Overview of Tank Shell in FE model with Mesh Shown ............................................ 21

Figure 25. Image Showing Solid Region Mesh and Local Refinement of Shell Mesh at the Shellto-solid Interface ................................................................................................................... 21

Figure 26. Image Showing Extent of Solid Region ...................................................................... 21

Figure 27. True Plastic Stress-Strain Input Data for ‘Actual’ TC128-B steel [30] ...................... 22

Figure 28. B-W Damage Model for ‘Actual’ TC128-B ............................................................... 23
vii 

Figure 29. Image Showing MPC Connecting Skid Beam to Tank Shell ...................................... 24

Figure 30. Tank car String Potentiometers (plan view) [30] ........................................................ 24

Figure 31. Image showing Arrangement of String Potentiometers in FE Model ......................... 25

Figure 32. Overview of Test 6 Simulation.................................................................................... 26

Figure 33. Local Puncture in Solid Region of Tank Car Shell ..................................................... 26

Figure 34. Indenter Travel vs Time, FE Model vs Test 6 ............................................................. 27

Figure 35. Indenter Speed vs Time, FE Model vs Test 6 ............................................................. 27

Figure 36. Indenter Force vs Displacement, FE Model vs Test 6 ................................................. 28

Figure 37. Average Air Pressure vs Time, FE Model vs Test 6 ................................................... 28

Figure 38. Indenter Force vs Displacement, FE Model: 13.0, 14.0, and 15.16 mph .................... 30

Figure 39. Full-Scale Model (left) and Quarter-Symmetry Model (right) .................................... 32

Figure 40. Indenter Force vs Travel, FE Quarter-Symmetry Model vs Test 6 ............................. 32

Figure 41. Average Cavity Pressure vs Time, FE Quarter-Symmetry Model vs Test 6 ............... 33

Figure 42. Overview of Quarter-Symmetry Model Following Puncture (time = 250 ms) ........... 33

Figure 43. Overview of 6 by 6-inch Indenter ............................................................................... 34

Figure 44. Indenter Force vs Travel, FE Quarter-Symmetry Model vs Test 8 ............................. 34

Figure 45. Overview of Test 8 Simulation (Quarter-Symmetry Model with Mirror Plane Used in 

X-Z Plane) ............................................................................................................................. 35

Figure 46. Schematic Cross-section Through DOT-105 Tank Car Within Outer Envelope ........ 36

Figure 47. Experimental Engineering Stress-Strain Curves for Round and Notched Tensile 

Specimens: a) Current Steel, b) New Steel [36] ................................................................... 38

Figure 48. ArcelorMittal True Stress-strain Curves Derived for Use in FE Analyses [36].......... 39

Figure 49. Damage Initiation Criteria for Current and New Steel Grades: a) Current Steel, b) 

New Steel [36] ...................................................................................................................... 39

Figure 50. True Stress-strain Curves: Current and New Steel from ArcelorMittal, Factored Steel 

is Current Steel x 0.94 ........................................................................................................... 40

Figure 51. Damage Initiation Criteria: Current and New Steel from ArcelorMittal, Factored Steel 

is Current Steel x 1.23 ........................................................................................................... 40

Figure 52. True Stress-strain Curves: Actual TC128-B and Enhanced TC128-B ........................ 41

Figure 53. Damage Initiation Criteria: Actual TC128-B and Enhanced TC128-B ...................... 42

Figure 54. Engineering Stress-strain Curves: Actual TC128-B and Enhanced TC128-B ............ 42

Figure 55. Force-displacement Response, Full-Scale Model 15.16 mph; Actual and Enhanced 

TC128-B Materials ............................................................................................................... 43

Figure 56. Force-displacement Response, Full-Scale Model: 0.775, 1.250- and 1.500-inch-thick 

shells – Survival Speeds ........................................................................................................ 45
viii

Figure 57. Cross-section through Tanks at Peak Force (left: 1.25 inch, right: 1.50 inch) ............ 45

Figure 58. DOT-105 Puncture Chart Showing Calculated Survival Speeds (12-inch Indenter, 

Side Impact Scenario) ........................................................................................................... 46

Figure 59. Overview of Response: 1.575-inch-thick Shell & Enhanced TC128-B at 24.5 mph .. 47

Figure 60. Plastic Strain Contour Plot (Upper Limit set to 0.4) Showing Inside Face of Tank 

Shell ...................................................................................................................................... 47

Figure 61. Force-displacement Response, Test 6 Baseline Results and Sensitivity with 35,000 lbf 

Added Mass .......................................................................................................................... 49

Figure 62. Flat Bar Tensile Test Setup: left: 0.775-inch Model, right: 1.5-inch Model ............... 50

Figure 63. Force-displacement Response, 0.775- & 1.5-inch-thick Specimens Using Actual 

TC128-B Material Model ..................................................................................................... 50

Figure 64. True Plastic Stress-strain Curves for Actual TC128-B & High Strength TC128-B.... 51

Figure 65. Left: Engineering Stress-strain Curves, Actual vs High Strength TC128-B, and right: 

Force-displacement response, 0.775-inch specimen with High Strength Actual TC128-B vs 

1.5-inch Specimen with Actual TC128-B ............................................................................. 51

Figure 66. Comparison of Global Deformation at 22 mph; 0.775-inch-thick High Strength vs. 

1.5-inch-thick Actual TC128-B ............................................................................................ 53

Figure 67. Force-displacement Response at 22 mph, Actual TC128-B 1.5-inch vs High Strength 

TC128-B 0.775-inch ............................................................................................................. 53

Figure 68. B-W Damage Model for Actual TC128-B & High Ductility Actual TC128-B .......... 54

Figure 69. Engineering Stress-strain Curves Produced from Round Bar (2-inch Gauge) Model 

Using Actual TC128-B & High Ductility Actual TC128-B ................................................. 55

Figure 70. Force-displacement Response, Actual TC128-B vs High Ductility TC128-B at 14 & 

20 mph .................................................................................................................................. 56

Figure 71. Section through High Ductility Model, 19 mph, at Time of Peak Indenter 

Displacement......................................................................................................................... 56

Figure 72. Local Section View through High Ductility Model, 19 mph, at Time of Peak Indenter 

Displacement......................................................................................................................... 57

Figure 73. Punch Test and Comparison of Experimental and Numerical Results ........................ 59

Figure 74. Punch Test Setup, Abaqus ........................................................................................... 60

Figure 75. Punch Test-50-3: Left, Mesh Shown; Right, Failure Progression ............................... 60

Figure 76. Punch Test: Force-displacement Response, Experimental vs. TT FEA ...................... 60

Figure 77. Test-50-3 Comparing Results with New (Double) Plate Configurations .................... 61

Figure 78. Results from 6.0- and 7.0-inch Plate Tests .................................................................. 61

Figure 79. Overview of the Flat Plate Model ............................................................................... 62

Figure 80. Force-displacement Curves, Single Flat Plates of Different Thickness ...................... 62
ix

Figure 81. Plastic Strain Contour Plot of 3 0.5-inch-thick Plates with All Fixed Outer Edges .... 63

Figure 82. Overview of Free Edge Configurations vs. Monolithic 1.5-inch-thick Plate .............. 63

Figure 83. Force-displacement Curves, Loose Upper Plate Configurations vs. Single 1.5-inchthick Plate.............................................................................................................................. 64

Figure 84. Illustration of Tank Shell Separated into Three Continuous but Separate Layers 0.5-

inch-thick .............................................................................................................................. 65

Figure 85. Force-displacement Curves 23.5mph, 3 x 0.5-inch Continuous Shells vs. Single 1.5-

inch-thick Shell ..................................................................................................................... 66

Figure 86. Overview of Implementation of Vertical Splits .......................................................... 66

Figure 87. Response of Layered-split Plate Case #2 at Time 200ms (Jacket and Foam Removed 

for Clarity) ............................................................................................................................ 67

Figure 88. Failure of 0.5-inch-thick Outer Layer (top) Preceding Failure of 1.0-inch-thick Shell 

(bottom)................................................................................................................................. 68

Figure 89. Overview of Quarter Symmetry Tank Car Model with Strengthened Jacket 

(Configuration #2a); Arrow Shows Location of the Circumferential Split .......................... 69

Figure 90. Configuration #2a – Indenter Striking Center of 50-inch-wide Split Plate Jacket ...... 69

Figure 91. Configuration #2b – Indenter Striking at Center of Plate Boundary (No Restraint 

Between Adjacent Plates) ..................................................................................................... 70

Figure 92. Overview of Puncture Zone Featuring Solid Element Representation of Both Jacket 

and Tank Shell ...................................................................................................................... 70

Figure 93. Overview of Configuration #1 – Continuous Outer Jacket ......................................... 71

Figure 94. Response of Configurations #2a and #3 at 24 mph ..................................................... 71

Figure 95. Experimental and Simulated Engineering Stress-strain Curves for Polyurea ............. 73

Figure 96. Local Cross-section through FE Model Showing Location of Liner .......................... 74

Figure 97. Internal Liner with Mesh Shown ................................................................................. 74

Figure 98. Response of Liner Model to 15-mph Impact (Right Image Shows Plastic Strain) ..... 74

Figure 99. Plastic Strain Contour Plot Showing Response of Liner to 15-mph Impact ............... 75

Figure 100. Plastic Strain Response of Liner in 16-mph Scenario: left – Polyurea Material, right 

– Reduced Ductility Polyurea ............................................................................................... 75

Figure 101. Force-displacement Curves, Steel-elastomer Bilayer Single Element Tensile Tests 77

Figure 102. Illustration of Impact by 24 x 24-inch Indenter ......................................................... 78

Figure 103. Overview of Side Impact Arrangement with 12 x 12-inch (red) and 24 x 24-inch 

(blue) Indenters ..................................................................................................................... 79

Figure 104. Far-field Plane Strain Test Setup – Only the Highlighted 2D Region with Length, L, 

and Thickness, t, Considered in Numerical Simulations [43] .............................................. 81

Figure 105. Overview of 1.4 mm Plane Strain Tension FE Model .............................................. 82
x 

Figure 106. Through-thickness Necking and Failure of 1.4 mm Plane Strain Tension Sample ... 82

Figure 107. Initiation and Failure of 1.4 mm Plane Strain Tension Sample ................................ 83

Figure 108. Normalized Force-displacement Curves for the Far-field Plane Strain Tension Tests 

for N = 0.05 and f0 = 0.02 ..................................................................................................... 83

Figure 109. The Plane Strain Bending Test Setup [43] ................................................................ 84

Figure 110. Overview of FE Model of 1.4 mm Plane Strain Bending Case ................................. 84

Figure 111. (a) Plastic Strain Contour Plots at Late Stage of Plane Strain Bending, Showing 

Surface Instabilities and Shear Band Formation on the Tension Side of Plate Compared with 

That Photographed by Dykeman Investigating Plane Strain Bending of 22mnb5 Boron Steel 

Plate with an Aluminum-Silicon Coating; (B) Crack Propagation and Failure of Plate ...... 85

Figure 112. Normalized Bending Moment-bending Angle Curves for the Plane Strain Bending 

Test ........................................................................................................................................ 86

Figure 113. Illustration of Plane Strain Tension Samples and Plane Strain Bending Arrangement

............................................................................................................................................... 87

Figure 114. Normalized Force-displacement Curves for the Far-field Plane Strain Tension Tests 

for t = 1.4, 3.0, 9.0, and 25.4 mm ......................................................................................... 88

Figure 115. Plot showing Stress Triaxiality in 9.0 mm Plane Strain Tension Sample at Onset of 

Failure ................................................................................................................................... 88

Figure 116. Plot Showing Plastic Strain in 9.0 mm Plane Strain Tension Sample When Failure 

Initiates and Starts Propagating through the Thickness ........................................................ 88

Figure 117. Comparison of the Normalized Bending Moment-bending Angle Curves for the 

Plane Strain Bending Tests at Increased Thickness .............................................................. 89

Figure 118. Calculation of Fracture Strain: (A) Initial Undeformed Shape of Gauge Region; (B) 

Deformed Shape of Gauge Region at Initiation of Fracture; (C) Plot Showing Plastic Strain 

vs. Distance Along Deformed Gauge Length (bottom surface); (D) Fracture Strain 

Calculation ............................................................................................................................ 90

Figure 119. Plastic Strain Contour Plot of 9.0 mm Plane Strain Bending Sample When Fracture 

Initiates .................................................................................................................................. 91

Figure 120. Fracture Strain Versus Plate Thickness for Far-field Plane Strain Tension and Plane 

Strain Bending Conditions .................................................................................................... 92

Figure 121. Plot Showing Stress Triaxiality in 9-mm-thick Plane Strain Bending Simulation ... 92

Figure 122. Different Mitigation Solutions with Total Plate Thickness of 1.5 inches ................. 96

Figure 123. Percentage of Damage by Location on Insulated Tank Cars in Mainline Accidents 

[46] ........................................................................................................................................ 97

Figure 124. A1011 True Stress True Plastic Strain Curve [30] .................................................. 105

Figure 125. B-W Failure Envelope [30] ..................................................................................... 106

Figure 126. Engineering Stress Strain Curve Produced by Researchers .................................... 106
xi 

Figure 127. Engineering Stress Strain Curve Produced by Carolan [30] ................................... 107

Figure 128. Comparison of Engineering Stress Strain Curves Produced using Carolan Parameters 

(0.04 inch & S4 elements) vs Research Parameters (0.1196 inch, S4R and scaled B-W 

curve) .................................................................................................................................. 107

Figure 129. Plastic Strain Plot of Tensile Tests using Calibrate Material Properties for A1011 

Material ............................................................................................................................... 108

Figure 130. TC128-B True Stress/True Plastic Strain Curves [30] ............................................ 109

Figure 131. TC128B Fracture Loci [30] ..................................................................................... 110

Figure 132. Necking of Round Bar During Tensile Test ............................................................ 110

Figure 133. Round Bar Tensile Tests Showing Plastic Strain Contours .................................... 111

Figure 134. Engineering Stress Strain Curves from Round Bar Simulations ............................. 111

Figure 135. Engineering Stress Strain Curves from Researched Round Bar Simulations Overlaid 

onto Original Curves (Test 2 and New Normalized, dotted lines) ..................................... 112

Figure 136. Engineering Stress Strain Curve from Researched Round Bar Simulation (solid black 

line) Overlaid onto Original Curve (Actual) ....................................................................... 112

Figure 137. Comparison of String Potentiometer Length vs Time, Centre of Tank .................. 113

Figure 138. Comparison of String Potentiometer Length vs Time, 24-inch Offset from Center of 

Tank .................................................................................................................................... 114

Figure 139. Comparison of String Potentiometer Length vs Time, 48 inch Offset from Center of 

Tank .................................................................................................................................... 114

Figure 140. Comparison of String Potentiometer Length vs Time, Skid ................................... 115

Figure 141. Comparison of String Potentiometer Length vs Time, Vertical .............................. 115

Figure 142. Comparison of String Potentiometer Length vs Time, Head .................................. 116

Figure 143. Polyurea, Input Engineering Stress-strain Curve .................................................... 118

Figure 144. Polyurea, Abaqus True Stress – Logarithmic Plastic Strain Curve ......................... 118

Figure 145. Displacement Response of Polyurea Single Element Test ...................................... 119

Figure 146. Engineering Stress-Strain Curve Obtained from Abaqus Polyurea Single Element 

Test (black line) Compared to Experimental Data and NLFlex Single Element Test Results

............................................................................................................................................. 119
xii 

Tables

Table 1. Selection of Hazmat Accidents ......................................................................................... 3

Table 2. Summary of Tank Car Shell Impact Tests [27] ................................................................ 9

Table 3. Mechanical Properties of Foam Insulation [30] ............................................................. 17

Table 4. Material Properties of Membrane Material .................................................................... 18

Table 5. Water Properties Used in FE Model ............................................................................... 18

Table 6. Jacket A1011 Material Properties Used in FE Model .................................................... 20

Table 7. TC128-B Baseline Material Properties Used in FE Model ............................................ 22

Table 8. String Potentiometers, FE Model vs Test 6 .................................................................... 29

Table 9. Test 6 – Puncture Speed .................................................................................................. 29

Table 10. Test 8 Comparison, FE Quarter-Symmetry Model vs Physical Test ............................ 35

Table 11. DOT-105 Potential Configurations and Space Available for Side Impact Mitigation . 37

Table 12. Fracture Energies for Current and New Steel Grades ................................................... 39

Table 13. Scaling Factors Obtained by Comparing ArcelorMittal Current and New Steels ........ 41

Table 14. Summary of Kinetic Energies ....................................................................................... 43

Table 15. Tank Shell Weights (Approximate) .............................................................................. 44

Table 16. Results – Tank Shell Thickness Study .......................................................................... 44

Table 17. Enhanced TC128-B, 1.50 Inch Shell ............................................................................ 46

Table 18. Survival Speed Summary, High Strength vs. Steel Thickness ..................................... 52

Table 19. Comparison of Steel Strengths ..................................................................................... 52

Table 20. Survival Speed Summary, High Ductility vs. Steel Thickness ..................................... 55

Table 21. ARA Punch Plate Tests................................................................................................. 59

Table 22. New Punch Plate Simulations ....................................................................................... 61

Table 23. Single Flat Plat Results Summary ................................................................................ 63

Table 24. Layered Plate Results Summary ................................................................................... 64

Table 25. Layered-Split Plate Results Summary .......................................................................... 66

Table 26. Strengthened Jacket Solution Configurations ............................................................... 69

Table 27. Strengthened Jacket Configurations - Results .............................................................. 70

Table 28. Single Element Test Configurations ............................................................................. 76

Table 29. Survival Speeds for Different Indenter Sizes ............................................................... 78

Table 30. GTN Material Model Parameters ................................................................................. 82

Table 31. Model Parameters for Plane Strain Bending................................................................. 84
xiii

Table 32. Model Parameters for Plane Strain Tension, Increased Thickness Study .................... 86

Table 33. Model Parameters for Plane Strain Bending, Increased Thickness Study .................... 87

Table 34. Results from Plane Strain Tension, Increased Thickness Study ................................... 89

Table 35. Results from Plane Strain Bending, Increased Thickness Study .................................. 90

Table 36. Summary of Plane Strain Bending Fracture Strains for Tank Car Applications .......... 93

Table 37. Air Properties .............................................................................................................. 104

Table 38. Molar Specific Heat for Air ........................................................................................ 104

Table 39. True Stress/True Plastic Stain Inputs for A1011 [30]................................................. 104

Table 40. TC128-B Material Models .......................................................................................... 109

Table 41. TC128-B Tensile Coupon Model Characteristics ....................................................... 110

Table 42. Requirements of TC128-B Material ........................................................................... 112

Table 43. Polyurea Input Engineering Stress-strain Data ........................................................... 117

Table 44. Polyurea Material Properties Used in FE Model ........................................................ 118
1 

Executive Summary

The increasing volume of rail transport inevitably leads to an increased number of accidents, 

which raises safety concerns and presents challenges related to hazmat transportation. In the 

event of an accident or derailment, the puncture resistance of the tank is key to preventing

hazmat release and escalation of the incident. 

The Federal Railroad Administration (FRA) sponsored a 2-year research effort led by Thornton 

Tomasetti (TT) to investigate multiple mitigation concepts for hazmat tank cars, aimed at 

improving their puncture resistance. The research was conducted between August 2020 and 

August 2022. The team first developed and validated a finite element (FE) model of a full-scale 

tank car subject to shell (i.e., side) impact. Researchers then evaluated the effectiveness of 

different mitigation solutions by exercising the validated FE model. 

The team used a side impact test of a water-filled DOT-105 tank car in the validation exercise to 

establish the baseline condition. The tank car featured a 0.775-inch-thick TC128-B steel shell

that was tested for puncture resistance, achieving puncture of the shell at 15.16 mph. A survival 

speed of 13.0 mph was subsequently calculated using the validated FE model. 

The team then assessed a new low-carbon version of TC128-B steel developed by ArcelorMittal 

(AM). The new steel featured higher ductility and reduced strength and AM had predicted it 

would provide 12 - 14 percent improvement in puncture energy. Using information published by 

AM, the research team evaluated the performance of this new steel in comparison to the baseline. 

The calculated improvement in puncture energy was very similar to that estimated by AM, 

representing an incremental improvement in the puncture resistance of tank cars. 

Considering the available tank car space and weight constraints, one of the simplest theoretical 

ways to address puncture resistance would be to increase the thickness of the tank shell. This

applies to the design of new tank cars only, and the tank car industry has made generational 

improvements in which head and shell thickness have been increased. The team investigated this 

option by increasing the tank shell thickness to 1.5 inches and calculating the tank car response 

to side impact. A survival speed of approximately 22 - 23 mph (an approximate threefold 

increase in energy) was achieved using around 90 percent of the available weight. 

Recognizing the challenges related to manufacturing a tank car with 1.5-inch-thick shell, 

researchers conducted a parametric study to investigate the effect of increased ductility and 

strength. In both cases, the puncture energy approximately doubled, with survival speeds 

increasing to 19 mph and 20 mph for the high ductility and high strength materials, respectively. 

These findings are consistent with those of previous researchers. 

The team also considered implementing a multi-layer (i.e., tank-within-a-tank) approach. The 

results indicated that in various configurations the survival speed could approximately match that 

of the 1.5-inch-thick shell provided that the total steel thickness was consistent. Researchers 

found that splitting the outer plates into a series of rings did not significantly change the puncture 

resistance. One significant benefit the multi-layer solution offers is the use of increased strength 

or ductility steels for the outer layers. Additionally, a numerical investigation of the effect of 

plate thickness on ductility showed that increasing the thickness can significantly reduce ductility 

in bending. Applying a multi-layer solution avoids this ductility penalty. 
2 

Team members also investigated the use of an internal tank car liner, which appeared potentially 

effective purely as a containment mechanism in the event of puncture (i.e., no improvement in 

puncture speed was observed). Researchers also considered mitigation solutions aimed at 

blunting the indenter, finding the effectiveness of this approach similar to increasing the shell 

thickness, but with significant added complexity and cost. Therefore, the team does not 

recommend pursuing mitigation solutions aimed primarily at blunting. 

In summary, the most promising mitigation approach to improve tank car puncture resistance is 

likely to be the multi-layer (i.e., tank-within-a-tank) approach, in which different steel grades can 

be used for the outer layers with the specific purpose of enhancing the overall puncture 

resistance of the tank car. These findings match one of the key findings of the Advanced Tank 

Car Collaborative Research Program (ATCCRP), although it is not yet clear what level of 

puncture resistance may be achieved using this approach. The team recommends further research 

to investigate the level of puncture resistance possible using the multi-layer approach. 
3 

1. Introduction

This report documents a 2-year research effort funded by the Federal Railroad Administration 

(FRA) to investigate multiple mitigation concepts for tank cars to improve their puncture 

resistance. The research was conducted between August 2020 and August 2022 and was led by 

Thornton Tomasetti (TT), an engineering consulting firm with extensive experience in the design 

and assessment of various structural systems subjected to extreme loads. TT was supported by 

independent project consultants Dr. David Jeong and David Tyrell, both of whom had 

contributed significantly to previous tank car integrity research efforts. The team was further 

supported by Prof. John Hutchinson, a renowned expert in the field of solid mechanics. 

1.1 Background

Increasing volume of rail transport inevitably leads to an increased number of accidents, which 

raises safety concerns and illustrates the challenges related to hazmat transportation. Most of the 

accidents that have occurred since the Minot, North Dakota, derailment on January 18, 2002, 

have involved two main classes of hazmat: (1) toxic inhalation hazmat (TIH), also referred to as 

poison inhalation hazardous (PIH) materials (e.g., anhydrous ammonia – Minot [1]); and (2) 

flammable liquids (primarily crude oil and ethanol). In the case of the latter, rapid expansion of 

oil extraction from the Bakken Shale site in North Dakota beginning in 2006 led to a more than 

tenfold increase in the volume of crude oil transportation via unit trains (i.e., trains with all cars 

carrying the same commodity, shipped from the same origin to the same destination). This led to 

a significant increase in accidents resulting in hazmat release; some key accidents are 

summarized in Table 1. 

Table 1. Selection of Hazmat Accidents 

Date Location Hazmat Reference

10/20/2006 New Brighton, PA Ethanol [2]

10/10/2007 Painesville, OH Ethanol [3]

06/19/2009 Cherry Valley, IL Ethanol [4]

10/07/2011 Tiskilwa, IA Ethanol [5]

07/11/2012 Columbus, OH Ethanol [6]

12/30/2013 Casselton, ND Crude Oil [7]

07/06/2013 Lac-Mégantic, Quebec Crude Oil [8]

A particularly disastrous accident occurred in 2013 in Lac-Mégantic, Quebec, Canada, which 

involved the derailment of an unsecured (i.e., runaway) train [8]. In total, 16 significant accidents 

(9 involving crude oil and 7 involving ethanol) occurred between October 2006 and February 

2014. The consequences of these 16 accidents included 48 fatalities, 281 DOT-111 tank cars 

derailed, 2.8 million gallons of crude oil released, and 2.0 million gallons of ethanol released. 

Selected examples of specific accident effects and hazmat release are shown in Figure 1. 
4 

Figure 1. Left to Right: Macdona, TX, Accident – Head Puncture [9]; Granitville, SC,

Accident – Shell Puncture [10]; Lac-Mégantic, Canada, Car Pileup [8]

Prior to these accidents, FRA tank car safety research had focused on maintaining tank integrity 

under normal tank car operating conditions (e.g., fatigue crack propagation). However, most of 

the hazmat releases discussed above occurred during derailment where the car coupler and/or 

broken rail impacted the tank car shell causing its puncture. This presents an entirely different 

problem than pressurization cycles and fatigue cracks and requires investigation of tank car 

crashworthiness and structural integrity when subjected to the extreme loading conditions of 

derailments and collisions. Thus, government- and industry-sponsored research shifted toward 

these topics, focusing specifically on puncture resistance and maintaining tank car integrity 

during derailments. In 2006, the Next Generation Rail Tank Car Project (NGRTCP) was begun 

as an industry research and development effort under the sponsorship of Dow Chemical 

Company, Union Pacific Railroad (UP), and Union Tank Car Company (UTLX). 

Later in 2006, FRA and the Pipeline and Hazardous Materials Safety Administration (PHMSA) 

initiated a Notice of Proposed Rulemaking (NPRM) for transporting hazmat by railroad tank cars

[11]. In 2007, FRA and Transport Canada (TC) signed a Memorandum of Cooperation (MOC) to 

participate in the NGRTCP with the understanding that the research results would be used to 

support rulemaking for North American regulatory agencies (i.e., FRA and TC) and to help 

develop improved tank car designs for the railroad tank car industry. In 2009, the Advanced 

Tank Car Collaborative Research Program (ATCCRP) was initiated as another joint industry

effort by the Association of American Railroads (AAR) and is still on-going. In 2015, the U.S. 

Department of Transportation (USDOT) released a final safety rule governing the transportation 

of flammable liquids by rail, primarily crude oil and ethanol [12]. The rule stipulated that new 

tank cars constructed after October 1, 2015, must meet design and performance requirements for 

a new USDOT specification, DOT-117 (in Canada, the new class of tank car is designated as 

TC-117). 

While all these government- and industry-led activities resulted in improved safety in hazmat 

transportation by rail, many challenges remain. One of the most important of these challenges is 

related to operation of the existing fleet of tank cars, both in the US and Canada, which were 

constructed before 2015. 

Throughout these government- and industry-led impact test activities, both head and shell (i.e., 

side) impact scenarios were investigated in detail, although the technology for head impact 

protection can be considered at a more advanced stage than that for shell impact protection. This 

is mainly due to prior research into tank car head vulnerability resulting in federal regulations 

instituted in the 1980s requiring head shields to protect railroad tank cars carrying certain classes 

of hazardous materials [13]. Introduction of the head shields has led to improved safety of 

hazmat transport by rail through the significant reduction of the conditional probability of release 
5 

for head losses (for unshielded heads) [14]. More recent efforts to improve puncture resistance

built on that experience and validated performance improvements for both head and shell 

impacts.

One of the most important steps to improve tank car puncture resistance is to identify the critical 

collision conditions and loading scenarios. As done previously for head impact, shell impact 

scenarios were identified by examining past accidents with the most severe impact conditions. 

Examples of these accidents are shown in Figure 2 [15]. 

Figure 2. Derailment Pileups, Left To Right: Minot, ND, 2002; Lac-Mégantic Car Pileup

These considerations led to the development of a shell impact scenario and full-scale testing 

protocol analogous to the head impact test stipulated in Code of Federal Regulations (CFR) 

§179.16 [13]. The regulations describe a general coupler-head impact test in which a moving ram 

car with a minimum weight of 263,000 lb strikes the head of a subject tank car below the 

centerline of the head. The subject tank car, with its brakes released, is braced by three fully 

loaded backup cars with their brakes applied. Additionally, Appendix A to Part 179 [13] 

describes a test procedure to verify the head puncture resistance and system survivability after 

coupler-to-tank impacts at a relative speed of 18 mph. 

A similar test setup has been proposed for shell impacts. In the shell impact scenario, which is a 

subject of ongoing research (and not yet subject to any existing regulation), a moving ram car 

strikes the side of a subject tank car at its centerline, with the subject car braced against a rigid 

barrier (Figure 3). This scenario was conceived as an idealized, full-scale, side impact test that 

would result in puncture and other failure modes that occur in actual accidents. Other important 

considerations of the shell impact test include safety, repeatability, controllability, and the ability 

to numerically simulate the test conditions and tank car response. Since 2007, a series of fullscale shell impact tests have been performed along with an extensive simulation effort and smallscale lab experiments to investigate the conditions experienced during possible crash and impact 

scenarios.
6 

Figure 3. Generalized Impact Scenarios: (a) Head – CFR §179.16 [13], and (b) Shell – 

Ongoing Research [15] 

Alongside full-scale impact tests, the team undertook parallel work to evaluate the behavior of 

current design equipment in these scenarios and to develop mitigation strategies aimed at 

doubling the maximum impact speed for which tank integrity is maintained and/or hazmat is 

contained. A general framework was developed for screening and evaluating potentially 

improved designs [16]. The key design functions of different mitigation strategies were: 

1. Blunting the impact loads

2. Absorption of the collision energy 

3. Tank strengthening 

4. Controlling the load path 

Researchers proposed specific design features and concepts and evaluated their ability to achieve 

these functions. These included innovative solutions such as sacrificial panels and shields 

(addressing 1 and 2 above), head and shell reinforcements (addressing 3), and carbody redesign 

to allow the support of service loads [16]. An initial integrated concept design was also 

developed, with tank strengthening, load blunting, energy dissipation measures, and an auxiliary 

structural system to carry the service loads as well as impact within the existing weight and space 

constraints [17]. The integrated concept design (Figure 4) included not only performance during 

impact, but also feasibility, scalability, and cost. 

Figure 4. Schematic Illustration of Integrated Conceptual Design [17] 
7 

To address the multiple design functions discussed above, researchers investigated engineered 

sandwich panels with multiple geometries and core features [18], [19], [20], [21]. A detailed 

experimental and numerical study was performed, culminating in full-scale testing to 

demonstrate the concept [22]. Figure 5 shows a basic design concept with a structural carbody 

including the sandwich panels and the full-scale test setup. 

Figure 5. Use of Sandwich Panels to Improve Tank Car Integrity [18], [19], [22] 

Most of the conceptual developments discussed above were accomplished through a combination 

of detailed finite element (FE) simulations and experimental testing. This required significant 

advancement of the appropriate approaches to simulate the response of the tank car components, 

including the mitigation elements [22], [23]. One of the most important components of the 

effective simulation of tank car response to localized impact is accurate representation of 

material failure on a large scale.

Significant effort was made to characterize the tank car material response (TC128-B steel) and 

calibrate the appropriate constitutive model [24], building on recent advances in understanding 

of the ductile failure of metals [25], [26]. 

1.2 Objectives

The overall objective of the research was to provide fundamental insights into the design and 

integrity of tank cars subjected to localized impact to make tank car manufacturers and rule 

makers aware of the potential puncture resistance enhancements that are achievable by changing 

different parameters of tank car design. A further objective was to evaluate potential mitigation 

measures for the existing fleet of tank cars. 

1.3 Overall Approach

The overall approach adopted in this research was to investigate different mitigation solutions 

using FE methods. Recent full-scale testing and associated numerical simulations conducted in 

related research enabled the team to efficiently develop validated full-scale FE models of the side 

impact test. These models were used to evaluate the effectiveness of various mitigation solutions 

with a high degree of confidence. The different material properties of steel that contribute to 

puncture resistance were investigated, as were the effects of varying steel thickness and splitting 

plates into multiple layers. This led to an investigation of the relationship between steel thickness 

and ductility, with ductility having been found to strongly influence puncture resistance. 
8 

1.4 Scope

The research focused on mitigation of the side impact case specifically in relation to pressurized 

hazmat tank cars; however, the findings are considered relatable to non-pressurized tank cars. 

1.5 Organization of the Report

The report is arranged as follows: 

• Section 2 provides background into the history of side impact testing and details the steps 

undertaken to develop and validate a full-scale FE model of a DOT-105 tank car. 

• Section 3 details the development of a more efficient sub-model for use in subsequent 

mitigation analyses. 

• Section 4 documents the methodology for and results from a series of mitigation 

analyses. 

• Section 5 describes an investigation into the relationship between ductility and plate 

thickness. 

• Section 6 considers the feasibility, cost, and scalability of the most promising mitigation 

solutions. 

• Section 7 summarizes the key conclusions. 

• Supplementary information is supplied in Appendices A through D. 
9 

2. Development of a Full-Scale Finite Element Model

To evaluate the effectiveness of different mitigation measures, it was necessary for the team to 

develop a well-validated FE model upon which the selected solutions could be tested. This 

section describes the selection, development, and verification of a full-scale FE model. 

2.1 Full-Scale Tests

Since 2007, FRA has been working collaboratively with industry research partners to undertake a 

series of full-scale shell (i.e., side) impact tests to improve the understanding of tank car impact 

behavior and puncture resistance. The tests were conducted at FRA’s Transportation Technology 

Center, Inc. (TTCI), in Pueblo, Colorado, and have been accompanied by FE analysis (FEA) 

primarily performed at USDOT’s National Transportation Systems Center (Volpe)1

. 

A table summarizing the tank car shell impact tests is shown in Table 2. 

Table 2. Summary of Tank Car Shell Impact Tests [27]

Test # Test Date Specification Impact Speed (mph) Impactor Size 

(inches)

1 4/26/2007 DOT-105 14.0 17 x 23

2 7/11/2007 DOT-105 15.1 6 x 6

3 5/18/2011 DOT-105

(w/ panel) 17.8 12 x 12

4 12/18/2013 DOT-111 14.0 12 x 12

5 2/26/2014 DOT-112 14.7 12 x 12

6 4/27/2016 DOT-105 15.2 12 x 12

7 9/28/2016 DOT-117 13.9 12 x 12

8 8/1/2018 DOT-105 9.7 6 x 6

9 10/30/2018 DOT-111

(CPC-1232) 13.9 12 x 12

10 11/19/2019 DOT-113 16.7 12 x 12

11 6/11/2020 DOT-113

(surrogate) 17.3 12 x 12

12 07/24/2021 DOT-113

(surrogate) 18.3 12 x 12

As shown in Table 2, researchers physically tested various tank car specifications, with 5 of the 

12 tests involving DOT-105 tank cars. DOT-105 are insulated tank cars that carry pressurized 

gases (often TIH materials) and are slightly smaller than the non-pressurized DOT-111 and 

DOT-117 tank cars that carry flammable liquids. DOT-112 is a pressurized tank car, like DOT105, while DOT-113 is a double-walled tank car designed to transport cryogenic liquids. The 

number of DOT-113 tank cars in service is very small in comparison to the other specifications

1 FEA for Tests 1, 2, and 4 was performed by Applied Research Associates, Inc.
10 

(approximately 67 in 2020 [28]), although this number is expected to increase following FRA’s 

Final Rule (July 24, 2020 [29]) permitting transportation of liquified natural gas (LNG) by rail in 

DOT-113C120W specification cars. 

The team chose to use the DOT-105 specification for the development of a full-scale FE model

after four different impactor sizes were considered. The team chose DOT-115 because it was the 

most tested specification, had been tested with sandwich panels aimed at improving puncture 

resistance (Test 3), and recent FRA tests and analyses provided great detail on the methods for 

developing a validated full-scale FE model using this specification [30], [31]. Specifically, the 

full-scale FE model developed during this research was to align with prior Tests 6 and 8. These 

tests, and the accompanying analysis reports [30], [31], were commissioned by FRA with the 

intention of generating useful data for researchers and entities seeking to develop validated and 

verified FE models. 

2.2 Tests 6 and 8 Overview

Tests 6 and 8 in Table 2 involved DOT-105 tank cars constructed between 1979 and 1980 and 

equipped with both head protection and foam insulation within the exterior jacket. The 0.775-

inch-thick tank car shell was constructed from TC-128 Grade B steel and the internal diameter 

was 100.45 inches. The tank was covered with a 4-inch-thick layer of foam enclosed within an 

11-gauge steel jacket. An example of the general arrangement for a DOT-105 tank car is shown 

in Figure 6 [32]. 

Figure 6. Example DOT-105 Tank Car Design Specification [32]

In both tests, the tank cars were water-filled to an outage of 10.6 percent. The manhole lid was 

sealed and the car was pressurized to 100 psi before the test. The indenters were positioned to 

align with the mid-length and mid-height of the target, with a 12 x 12-inch indenter used in Test 

6, and a 6 x 6-inch indenter used in Test 8. The indenters were fitted to the head of a ram car

(Figure 7), which weighed 297,125 lb in Test 6 and 296,775 lb in Test 8. The measured impact 

speeds of the ram cars were 15.2 and 9.7 mph for Tests 6 and 8, respectively. 

The tank cars were supported on two support skids, with four sections of I-beams welded to the 

underside of the tank car to facilitate sliding and minimize rollback (Figure 8). The tank cars 

were lifted into position by crane and positioned as close as possible to the rigid wall behind. 

They were instrumented with triaxial accelerometers, pressure transducers, speed sensors, and 

string potentiometers at various locations. High speed cameras were used to record the tests. 
11 

Figure 7. Photograph of a Ram Car and Head Fitted with 6 x 6-inch Indenter [30]

Figure 8. Tank Support Skid System [30]

In both tests (Figure 9 and Figure 10), the tank cars punctured, with the ram car rebounding after 

impact. The impact speeds were selected based on pre-test estimates from FE simulations; the 

occurrence of rebound indicates that the impact speed was close to the puncture speed in both 

cases. Post-test simulations that incorporated refined material models using samples extracted 

from the tank shell showed very good agreement with the physical test results. The detailed 

methodology developed in previous research ([30], [31]) for full-scale FE models in Tests 6 and 

8 was followed in this research effort to establish a baseline full-scale model that was well 

understood and verified. This is discussed in Section 3.3. 
12 

Figure 9. Deformed Tank Just Prior to Puncture (12 x 12-inch Indenter) [30]

Figure 10. Deformed Tank Just Prior to Puncture (6 x 6-inch Indenter) [30]
13 

2.3 Full-scale FE Model

An overview of the full-scale FE model is shown in Figure 11 and Figure 12. The team 

developed the model using the commercial FE software package Abaqus (2020), with the model 

geometry developed using the Abaqus/CAE preprocessor and dynamic analyses undertaken 

using the Abaqus/Explicit solver. A half-symmetry model was developed, using the same 

techniques used in [30] and [31]. 

The model comprises a 0.775-inch-thick tank shell predominantly modelled using shell elements, 

with a local region modelled using solid elements at the impact location. Internally, there is a 

fluid cavity modelled using membrane elements, with a separation between water and air cavities 

giving an outage of 10.6 percent. The air cavity is pressurized to 100 psi. External to the tank 

shell is a 4-inch-thick layer of foam modelled using solid elements, and outside of the foam is a 

0.1196-inch-thick jacket modelled using shell elements. A single rigid skid is connected to a 

section of the tank shell near the head, and directly behind the tank assembly is a discrete rigid 

wall. Further detail of the model development is provided in the following sections. 

Figure 11. Overview of Full-scale FE Model, Test 6

Figure 12. Overview of Full-scale FE Model from Behind, Test 6
14 

2.3.1 Geometry

The geometry of the tank shell is shown in Figure 13. It was developed using mid-surface shells, 

with a local region behind the impactor replaced with a solid element region, as shown in Figure 

14. The tank body (cylindrical portion) was modelled using a shell thickness of 0.775 inches, 

with the 2:1 elliptical head modelled using a shell thickness of 0.8125 inches. A 20-inch 

diameter manway was centered and extruded 11.5 inches from the top of the tank. The vertical, 

cylindrical portion of the manway had a plate thickness of 1 inch, as did the manway cover plate. 

Directly below the manway, within a circular region with a diameter of 39 inches, the shell 

thickness was doubled. The circular partition for this region can be seen in Figure 14. The tank 

shell geometry was used as the baseline from which to create offset geometries to define the 4-

inch-thick foam layer, the 0.1196-inch-thick jacket, and the internal membrane. 

Figure 13. Elevation Showing Tank Shell Geometry

Figure 14. Isometric View of Tank Car Assembly, with Sections of Jacket (blue) And Foam 

(green) Removed to Show Tank Shell with a Local Solid Region
15 

2.3.2 Indenter

The geometry and associated mesh used for the 12 x 12-inch indenter is shown in Figure 15. The 

indenter was modelled as a discrete rigid part and in Test 6 it was assigned a weight of 148,563 

lbf (i.e., half of the total mass) due to the use of a symmetry plane. The 12-inch indenter features 

a 1-inch radius on the face so the dimension of the completely flat portion of the indenter 

(indicated by the double-headed arrow in Figure 15) is 10 inches. The mesh was locally refined 

to capture the rounded corners. 

Figure 15. Indenter Part and Mesh

The team conducted the crash simulations in two steps, the initial velocity step and the crash 

step. The initial velocity step comprised a very short step, typically 1 ms, in which the indenter 

was assigned an initial velocity in the direction of travel (i.e., Z-direction), with movement in all 

other degrees of freedom fixed. In the crash step (typical duration 300 ms), the velocity boundary 

condition was released but movement in all other degrees of freedom remained fixed. The 

indenter was offset from the jacket by a small amount (approximately 0.6 inches) to ensure that it 

did not contact the jacket during the initial step (i.e., the indenter would travel 0.266 inches over 

the course of the 1 ms-long step for an assigned velocity of 15.16 mph). This arrangement is 

shown in Figure 16. 

Figure 16. Cross-section Through Tank Car FE Model Showing Indenter ~0.6 inch Away
16 

2.3.3 Foam Insulation

A 4-inch-thick layer of foam insulation is present between the tank shell and outer jacket (see 

Figure 17). 

Figure 17. Overview of Foam Insulation (with and without mesh shown)

The foam part was modelled using solid (C3D8R) elements with a characteristic element length 

of 2 inches. This resulted in a total of 44,116 elements. A simple elastic-plastic material model 

was used based on that developed by Carolan et al [30], with the plastic response shown in 

Figure 18. The basic mechanical properties are shown in Table 3. A plastic strain-to-failure of 

0.55 independent of stress triaxiality was implemented. Failure would therefore initiate at a 

plastic strain of 0.55, with a damage progression energy of 1.0 in-lbf/in2

 used to define 

subsequent element degradation to deletion phase. 

Figure 18. True Plastic Stress Strain Input Data for Foam Insulation [30]

True Plastic Strain

True Stress (psi)

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0.5 0.55

0

20

40

60

80

100

120

140

160
17 

Table 3. Mechanical Properties of Foam Insulation [30]

Parameter Value Units

Density 2.0 lbf/ft3

Density 3 x 10-6 lbf*s2/inch

Modulus of Elasticity 941.5 psi 

Collapse Stress 8.2 psi 

Plastic Strain-to-Failure 0.55 - 

Damage Progression Energy 1.0 in-lbf/in2

2.3.4 Air and Water Phases

The air and water phases were modeled using a fluid cavity approach, in which a pneumatic 

cavity was defined for the pressurized air phase and a hydraulic cavity for the water phase. This 

is a simplified approach based on a calculation of average cavity pressure, but it showed 

satisfactory results while being much more computationally efficient than equivalent methods 

such as smooth particle hydrodynamics (SPH). 

To define the respective cavities, a single part comprising a series of connected surfaces and 

referred to as the ‘Membrane’ was developed, as shown in Figure 19. The outer surface was 

offset from the tank shell mid-plane by half the shell thickness. A horizontal plane located 16.5 

inches below the 12 o’clock position of the part was used to separate the water and air cavities

and was positioned to give an outage of 10.6 percent. 

 

Figure 19. Membrane Part (with and without mesh)

The part was modelled using Membrane elements (M3D8R) with the characteristic element 

length of 3.5 inches. Membrane elements are those that can carry membrane forces but do not 

have any bending or transverse shear stiffness. As per the methodology described in [30], 

material properties were assigned to the membrane part that would minimize the stiffness 

introduced by the part while maximizing the efficiency. To ensure a reasonable stable time 

increment, the density of the membrane material model was set to that of steel; to minimize the 

stiffness increase, the modulus of elasticity was set to 0.1 percent of steel (shown in Table 4).

The membrane thickness was set to 0.05 inches (i.e., thin). 
18 

Table 4. Material Properties of Membrane Material 

Parameter Value Units

Density 7.35x10-4 lbf*s2/in

Modulus of Elasticity 30,000 psi

2.3.4.1 Hydraulic Cavity – Water

The water lading was modelled as a hydraulic cavity with the properties shown in Table 5. While 

the density of water was included, this was primarily for the purpose of calculating the bulk 

modulus. The bulk modulus of a liquid is related to its compressibility and defined as the

pressure required to cause a unit change of volume of a liquid. Most liquids are practically 

incompressible, and therefore require very large pressures to cause any significant change in 

volume. 

Since the membrane part was effectively modeled as a thin sheet (i.e., 0.05-inch-thick) with 

relatively little mass (1482.6 lbf), additional non-structural mass was assigned to the water 

boundary as shown in Figure 20 to account for the overall mass of water lading and associated 

inertial effects. 

Table 5. Water Properties Used in FE Model 

Parameter Value Units

Density 7.35x10-4 lbf*s2/in

Bulk Modulus 309,500 lbf/in2

Part Mass 1482.6 lbf

Nonstructural Mass

(Additional weight of water) 65,162.5 lbf

Temperature 55  ̊F

Figure 20. Water Boundary (red)

2.3.4.2 Pneumatic Cavity – Air

The gas phase of the lading was modelled as air at an initial gauge pressure of 100 psi. The 

average air pressure within the cavity would respond to changes in volume due to deformation of 

the tank and because the water lading would respond as a largely incompressible fluid. The 

inputs required to define the pneumatic cavity were obtained from [30], and are summarized in 

Table 37 and Table 38 of Appendix A. 
19 

2.3.5 Jacket

The 0.1196-inch-thick outer jacket was modelled using conventional shell elements with a 

characteristic global element length of 2 inches and the mesh refined locally in the region of 

impact (Figure 21 and Figure 22). The jacket was modelled using material properties for A1011 

steel (summarized Table 6). The material model featured a Bai-Wierzbicki (B-W) failure 

envelope, further detail of which is provided in Appendix A. The B-W envelope defines the 

stress triaxially dependent plastic strain at which element failure initiates. A damage progression 

energy was used to define the subsequent damage evolution phase (i.e., the degradation to 

complete failure) and the resulting element deletion. 

Figure 21. Jacket (with and without mesh)

Figure 22. Locally Refined Mesh in Jacket at Impact Location 
20 

Table 6. Jacket A1011 Material Properties Used in FE Model

Parameter Value Units

Modulus of Elasticity 3 x 107 psi

Plasticity Nonlinear - See 

Appendix A

Nonlinear - See 

Appendix A

Poisson’s Ratio 0.3 -

Mass Density 7.35 x 10-4 lbf*s2/in

Damage Initiation B-W Envelope – See 

Appendix A

B-W Envelope –

Appendix A

Damage Progression Linear, 1,500 in-lbf/in2

2.3.6 TC128-B Tank Shell

The tank shell was modelled as two separate parts – a shell element part and a solid element part

– connected at the interface using tie constraints. As shown in Figure 23, the shell element part 

comprised the 0.775-inch-thick cylindrical body, the 0.8125-inch-thick head, and the 1-inchthick manhole extrusion as a single continuous part. At the impact location, a section of the shell 

was omitted to accommodate a solid element representation. The shell element part was 

developed using reduced integration shell elements (S3R and S4R) with a global characteristic 

element length of 2 inches. Locally, where the shell part was tied to the solid part, the mesh was 

refined to ensure reasonable enforcement of tie constraints. The solid region was meshed with a 

very fine mesh (0.085 inches), resulting in 9 elements through the thickness and 652,608 

elements within the ~15 by 35-inch region. See Figure 24, Figure 25 and Figure 26. 

Figure 23. Overview of Tank Shell in FE Model
21 

Figure 24. Overview of Tank Shell in FE model with Mesh Shown

Figure 25. Image Showing Solid Region Mesh and Local Refinement of Shell Mesh at the 

Shell-to-solid Interface

Figure 26. Image Showing Extent of Solid Region 
22 

The team developed two separate material models for the tank shell, one for the shell element

region and one for the solid element region. The key difference was that the solid region material 

model featured a B-W failure envelope developed by Carolan following tensile testing of 

samples extracted from the tank shell [30], hence the ‘Actual’ term. This is discussed further 

below, with additional information found in Appendix B. 

The baseline properties applicable to both material models are summarized in Table 7, with the 

hardening curve used to define plasticity behavior shown in Figure 27. 

Table 7. TC128-B Baseline Material Properties Used in FE Model

Parameter Value Units

Modulus of Elasticity 3 x 107 psi 

Plasticity Nonlinear – See Figure 27 & 

Appendix B - 

Poisson’s Ratio 0.3 -

Mass Density 7.35 x 10-4 lbf*s2/in

Figure 27. True Plastic Stress-Strain Input Data for ‘Actual’ TC128-B steel [30]

The B-W envelope used for the solid region material model is shown in Figure 28. It defines the 

plastic strain at which failure initiates for different stress triaxialities. 

True Plastic Strain

True Stress (psi)

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0

25000

50000

75000

100000

125000

150000

175000

200000
23 

Figure 28. B-W Damage Model for ‘Actual’ TC128-B 

The B-W curve and damage progression parameters are calibrated for a specific element type 

and size. There is a maximum element size that can be reasonably used in the round bar tensile 

tests to develop the damage model and this element size must be carried through to the full-scale 

model, hence the very fine mesh in the solid element region of the full-scale model. The 

accuracy of this region of the model is critical to capturing the correct behavior of the impact and 

defining when puncture occurs. 

2.3.6.1 Mass Scaling

The element size also dictates the speed at which analyses can be performed in an explicit 

analysis, where a small element results in a low stable time increment and therefore long 

simulation time. This can and has been alleviated to an extent using mass scaling. Mass scaling 

works by increasing the mass of the smallest elements, which slows the wave propagation speed 

across the element and therefore increases the stable increment. A typical target stable time 

increment of 5E-07 s was set, resulting in the mass of the elements within the solid region being 

scaled by a factor of 4.35. The overall increase in mass of the model through mass scaling was 

low (~0.15 percent) and was observed to have no significant effect on the response of the model. 

2.3.7 Skid Beam

The skid beam was modelled as a discrete rigid part, measuring 180 x 22 x 12 inches, and 

assigned a weight of 3,502 lbf. It was connected directly to the tank shell using a multi-point 

constraint (MPC) whereby the nodes across an area measuring 12-inches-wide and covering an 

arc 120 degrees across the bottom of the shell were rigidly connected to a discrete point on the 

skid beam, as shown in Figure 29. Boundary conditions were assigned to the skid to prevent 

movement in the vertical Y direction and rotation about the X-axis. 

Stress Triaxiality - T

Equivalent Plastic Strain - PEEQ

-0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6
24 

Figure 29. Image Showing MPC Connecting Skid Beam to Tank Shell

2.3.8 Friction

The team used the General Contact algorithm within Abaqus with ‘Hard Contact’ assigned in the 

normal direction and a coefficient of friction 0.3 assigned in the tangential direction. This contact 

relationship was applicable between all parts. The hard contact label defines the pressureoverclosure relationship, which minimizes penetration of one surface into another. 

2.3.9 String Potentiometers

The physical test was instrumented with 10 string potentiometers, 9 recording translations in the 

horizontal plane (Z-plane in Figure 29) and 1 located at the center of the tank recording vertical 

deformation of the tank shell. Figure 30, from [30], shows an overview of the arrangement of the 

string potentiometers. In the Abaqus model these were implemented using connector elements, 

as shown in Figure 31. Specifically, four connectors were used to capture displacements at the 

center of the tank, three in the horizontal plane and one in the vertical. The displacements of the 

skid and tank end were recorded simply as history output. 

Figure 30. Tank car String Potentiometers (plan view) [30]
25 

Figure 31. Image showing Arrangement of String Potentiometers in FE Model

2.3.10 Results – Test 6

An overview of the response of the FE model when the simulation was run at the test speed of 

15.16 mph is shown in Figure 32. Between 0 to 50 ms the impactor penetrates the outer jacket, 

crushes the foam at the impact location, and begins to locally deform the tank shell. The tank 

then begins to displace backwards, impacting the wall and crushing the foam on the rear side, as 

shown at time 150 ms in Figure 32. Once the foam is crushed and the tank shell is essentially 

bearing directly onto the rigid wall, the indenter force increases, causing greater local 

deformation and eventual failure of the tank shell locally behind the indenter. 

An overview of the local puncture behavior in the solid element region of the tank shell is shown 

in Figure 33. It shows that a section of material just larger than the indenter face fails and is 

pushed through into the tank. In the physical test a slightly different behavior was observed, with 

a relatively large section of the tank shell ejected outward after the impact. The ram car (i.e., 

indenter) also rebounded following impact, suggesting that failure and ejection of the section of 

tank car shell may have occurred due to internal pressure build-up (i.e., the impact fractured the 

shell, but gross failure occurred during rebound, when the weakened section of tank was forced 

out by the water pressure). While the FE model did not capture this exact and quite complex 

behavior it did correlate very well against the physical test across a range of measured 

parameters. 

Figure 34 and Figure 35 compare the displacement and speed of the indenter obtained from the 

FE model with that recorded from Test 6. Both plots show very close agreement up to the point 

at which failure initiates in the tank shell, which occurs around 220 ms into the event. Figure 36

compares the force-displacement response, which also shows good agreement. Peak force in the 

FE model was predicted to occur at a displacement of 37.4 inches, versus 38.0 inches recorded 

from the test, a difference of 1.6 percent. The peak force in the FE model was calculated to be 

approximately 14.6 percent lower than that recorded during the test, however the total energy 

absorbed, which is the integral of the force-displacement curve, was calculated to be very similar

at 2.16 versus 2.27 million foot-pounds for the FE model and Test 6 respectively (a difference of 

4.8 percent). The average air pressure in the outage showed an approximate 8 percent difference 

between measured and predicted values (Figure 37). The displacements measured in the string 
26 

potentiometers also showed good general agreement, as summarized in Table 8. Comparative

plots for the string potentiometers can be found in Appendix C. 

Overall, good agreement was obtained between the FE model and the physical test. 

Figure 32. Overview of Test 6 Simulation

Figure 33. Local Puncture in Solid Region of Tank Car Shell
27 

Figure 34. Indenter Travel vs Time, FE Model vs Test 6

Figure 35. Indenter Speed vs Time, FE Model vs Test 6

Time (s)

Indenter Travel (inches)

0 0.05 0.1 0.15 0.2 0.25

0

5

10

15

20

25

30

35

40

Test 6

FE Model

Time (s)

Indenter Speed (mph)

0 0.05 0.1 0.15 0.2 0.25

0

2

4

6

8

10

12

14

16

Test 6

FE Model
28 

Figure 36. Indenter Force vs Displacement, FE Model vs Test 6

Figure 37. Average Air Pressure vs Time, FE Model vs Test 6

Indenter Travel (inches)

Indenter Force (kips)

0 5 10 15 20 25 30 35 40

0

200

400

600

800

1000

1200

1400

1600

Test 6

FE Model

Time (s)

Average Air Pressure (psi)

0 0.05 0.1 0.15 0.2 0.25

100

105

110

115

120

125

130

135

140

145

150

Test 6

FE Model
29 

Table 8. String Potentiometers, FE Model vs Test 6

Location

FE Model*

Peak Displacement

(in)

Test 6

Peak Displacement

(in)

% Difference

Center 30.6 30.6 0.0

24-in offset 23.2 24.5 5.3

24-in offset 23.2 23.6 1.7

48-in offset 17.5 15.8 -10.8

48-in offset 17.5 18.2 3.8

Vertical 11.3 17.2 34.3

Skid 11.8 10.4 -13.5

Skid 11.8 9.9 -19.2

Head 11.5 10.3 -11.7

Head 11.5 9.8 -17.3

*Half-symmetry FE model, therefore only one value obtained for certain locations

2.3.11 Puncture Speed

Having validated the response of the full-scale model, the team conducted additional simulations 

at 13.0 and 14.0 mph to calculate the survival speed of the tank car when subjected to sideimpact with a 12 x 12-inch indenter. The results, summarized in Table 9, showed that the tank 

car survived the 13-mph case but failed at 14 mph. 

Table 9. Test 6 – Puncture Speed

Speed (mph) Initial Kinetic Energy

(m-ft-lb) Result Notes

15.16 2.283 Fail Test 6 case

14.00 1.947 Fail Tank shell fails during rebound phase

13.00 1.679 Pass Survival speed

The force-displacement response for each case is shown in Figure 38. The peak force is similar 

in each case at approximately 1300 kips. The curves show that in the 13- and 14-mph cases the 

indenter rebounds after reaching peak force, whereas in the 15.16-mph case it continues to travel 

forward. In subsequent simulations where the effectiveness of different mitigation solutions have 

been considered, the results are compared to these baseline simulations, where 13 mph is the 

calculated survival speed. 
30 

Figure 38. Indenter Force vs Displacement, FE Model: 13.0, 14.0, and 15.16 mph 

Impactor Travel (inches)

Impactor Force (kips)

0 10 20 30 40 50

0

200

400

600

800

1000

1200

1400

15.16 mph

14.00 mph

13.00 mph
31 

3. Sub-Model

The overall objective of the research effort was to investigate various tank car mitigation options 

using numerical simulation. While the validated full-scale model discussed in the previous 

section could be used to assess the effectiveness of different solutions, it could not be used to 

rapidly assess different mitigation options due to its size and complexity. For example, the fullscale model took approximately 77 hours to complete the Test 6 simulation using a conventional 

desktop analysis machine running on 12 processors. This was reduced to less than 24 hours when 

using a high-powered cluster using 23 processors. However, while a 24-hour turnaround was 

potentially workable, the output files were large, and the model would only increase in 

complexity when different mitigations were considered. Therefore, the research team sought to 

develop more efficient sub-models. 

Ultimately, the team developed a scaled-down, defeatured version of the full-scale model and 

used it for most of the mitigation simulations. More simplified sub-models were also developed 

and used in some cases to investigate specific local behavior. These models are described in 

detail in the relevant sections of the report. The focus of this section of the report however is on 

describing the development of the ‘baseline’ quarter-symmetry sub-model. 

3.1 Quarter Symmetry Model

The objective of using the sub-model was to capture most of the key behaviors exhibited by the 

full-scale model while increasing its efficiency. The key features known to influence the 

response include the shape of the tank, the incompressibility of the liquid phase and the 

compressibility of the gas phase, the inertia of the fluid, the inertia of the tank structure, the 

response of the jacket, and crushing of the foam. Crucially, the sub-model should also be capable 

of reliably predicting failure of the tank shell. The full-scale model captures these features with 

varying degrees of sophistication, so the team developed and investigated a simplified version of 

that model. 

The half-symmetry full-scale model was reduced to a quarter-symmetry model as shown in 

Figure 39. The skid and manhole features were removed and the horizontal plane separating the 

water and air phases was repositioned to give an outage of 10.6 percent. This required that the 

outage at the top of the model be decreased (the horizontal plane moved up approximately 6.0 

inches), as the use of quarter-symmetry would give rise to an equivalent outage at the bottom of 

the tank. The non-structural mass assigned to represent the water lading was halved, as was the 

mass of the indenter. The final change was made to the TC128-B solid element puncture region, 

with the full-integration incompatible mode brick elements (C3D8I) replaced by the more 

computationally efficient reduced-integration elements (C3D8R). The transition to C3D8R 

elements was first verified by repeating the round bar tensile test. The model otherwise remained 

the same. 

To verify the response of the model, two simulations were undertaken: 

• Test 6: 12 x 12-inch impactor at 15.16 mph 

• Test 8: 6 x 6-inch impactor at 9.7 mph [33]
32 

Figure 39. Full-Scale Model (left) and Quarter-Symmetry Model (right) 

3.2 Test 6 Response

The results obtained from the quarter-symmetry model for Test 6 were very similar to those 

produced by the full-scale model and showed good agreement with the physical test. Figure 40

shows a comparison of the force-displacement response of the quarter-symmetry model with the 

physical test. The displacement at failure was a close match (~38 inches) while the peak force 

obtained from the quarter-symmetry model (1432 kips) was closer to the Test 6 force (1391 kips) 

than the full-scale model (1188 kips), a difference of ~3 percent versus ~14.6 percent. The 

increase in peak force is largely attributed to the shift from C3D8I to C3D8R elements within the 

puncture zone. 

Figure 40. Indenter Force vs Travel, FE Quarter-Symmetry Model vs Test 6

Impactor Travel (inches)

Impactor Force (kips)

0 10 20 30 40

0

200

400

600

800

1000

1200

1400

1600

Test 6

Quarter-Symmetry Model
33 

Figure 41 shows the average cavity air pressure calculated using the quarter-symmetry model 

compared to the Test 6 data. The FE results are like those obtained using the full-scale model and 

are approximately 8 percent below the Test 6 recorded pressure at the peak. 

Figure 41. Average Cavity Pressure vs Time, FE Quarter-Symmetry Model vs Test 6

An overview of the deformed shape of the model following puncture is shown in Figure 42. Note 

that a ‘mirror plane’ has been used to illustrate the model as if it were half-symmetric. 

Figure 42. Overview of Quarter-Symmetry Model Following Puncture (time = 250 ms) 

In summary, across various key metrics the quarter-symmetry model compared well against the 

physical test. Moreover, the simulation time was significantly reduced, with the Test 6 

simulation shown here taking approximately 4.5 hours using 23 processors on a high-powered 

cluster.

Time (s)

Average Air Pressure (psi)

0 0.05 0.1 0.15 0.2 0.25

100

105

110

115

120

125

130

135

140

145

150

Test 6

Quarter-Symmetry Model
34 

3.3 Test 8 Response

Test 8 [33] was undertaken in 2018 using a 6 x 6-inch indenter at a speed of 9.7 mph, with the 

test setup otherwise intended to replicate Test 6. Test 8 was replicated with the objective of 

further assessing and validating the quarter-symmetry model. The 6 x 6-inch indenter was 

implemented, as shown in Figure 43, with the indenter mass changed to match one quarter of that 

used in the physical test. The indenter features a 1/2 inch radius such that the flat portion of the 

indenter face measures 5 x 5 inches. 

Figure 43. Overview of 6 by 6-inch Indenter

The resulting force-displacement curve obtained from the quarter-symmetry model is compared 

to the physical test results in Figure 44 with the curves showing good agreement. Figure 45

illustrates the response of the model just prior to failure and Table 10 compares some of the key 

parameters. The greatest difference is in the peak force, but as shown in Figure 44, this occurs 

for only a very short period in the physical test. 

Figure 44. Indenter Force vs Travel, FE Quarter-Symmetry Model vs Test 8

Impactor Travel (inches)

Impactor Force (kips)

0 5 10 15 20 25 30

0

100

200

300

400

500

600

700

800

900

Test 8

Quarter-Symmetry Model
35 

Figure 45. Overview of Test 8 Simulation (Quarter-Symmetry Model with Mirror Plane 

Used in X-Z Plane)

Table 10. Test 8 Comparison, FE Quarter-Symmetry Model vs Physical Test

Parameter FE Model Test 8 % Difference

Peak Force (kip) 754.4 842.7 11.7

Indenter displacement at failure (in) 27.1 27.9 3.0

Energy absorbed at peak force (ftkip) 921.5 929.9 0.9

Peak Outage Pressure (psi) 108.0 112.9 4.3

In general, the quarter-symmetry model demonstrated good agreement with the physical test data

for both Tests 6 and 8, and in some metrics, particularly the peak force and displacement to 

failure, the model showed slightly improved behavior when compared to the results obtained 

using a larger half-symmetry model while using a fraction of the computational cost of the halfsymmetry model. This finding suggests that capturing the puncture behavior of the shell is one of 

the most important parameters in determining puncture resistance, with features such as the 

manhole and skid of secondary importance. 

In conclusion, the quarter-symmetry model provided an efficient and reliable means to 

investigate different mitigation methods, and the team was confident that findings from its use 

would transfer to the overall behavior and puncture resistance of the tank. 
36 

4. Mitigation 

This section discusses the various mitigation methods the team investigated as part of this 

research. Section 4.1 first outlines some of the key constraints affecting the implementation of 

potential mitigation solutions and Section 4.2 through Section 4.7 detail the different mitigation 

solutions considered. 

4.1 Key Constraints

4.1.1 Clearance Envelope

The AAR Plate B Clearance Standards determine the maximum height, width, and truck center 

spacing for tank cars. As illustrated in Figure 46, the height from the top of rail is set to 15 ft-1 

in, the width to 10 ft-8 in, and truck centers not to exceed 41 ft-3 in. Swingout at the center of a 

car on a 13-degree curve results in a maximum overall car length of 57 ft-2 in [34]. A baseline 

DOT-105 tank car has a steel shell outer diameter of 8 ft-6 3/10 in (102.3 in). This provides a 

lateral spacing from the outer surface of the shell to the maximum width of approximately 1 ft7/8 in. This is essentially the minimum gap in which external mitigation measures could be 

implemented assuming the foam insulation (4-inch-thick) and jacket are either removed as part 

of the mitigation solution or implemented within it. 

Figure 46. Schematic Cross-section Through DOT-105 Tank Car Within Outer Envelope 

In terms of overall car length, the DOT-105 is a somewhat shorter car with truck centers in the 

region of 35 ft. The possibility therefore exists to decrease the diameter and increase the overall 

car length. This would provide more space around the tank in which to install mitigation 

solutions. For example, Tyrell [34] shows that reducing the tank car internal diameter to 7 ft-11 
37 

3/4 in can be accommodated within the overall length constraints. The new tank would be 

approximately 4 ft-4 in longer, but the lateral space available would increase by approximately 3 

1/4 in. The lateral spacing available for different configurations of DOT-105 tank cars are 

summarized in Table 11. The table shows that the minimum lateral space available is ~8.79

inches (retrofit condition) increasing to a maximum of 16.125 inches (new reduced diameter 

tank). This is important, as more space provides greater opportunity for energy absorption. 

 Table 11. DOT-105 Potential Configurations and Space Available for Side Impact 

Mitigation

# Type Description Outer Diameter

(in)

Lateral Space 

Available (in)

1 Retro-fit Install new mitigation on top of 

existing foam & jacket 110.42 8.79

2 Retro-fit Replace foam & jacket with new 

mitigation 102.30 12.85

3 New – reduced 

diameter

Install new mitigation on top of 

existing foam & jacket 103.87 12.065

4 New – reduced 

diameter

Replace foam & jacket with new 

mitigation 95.75 16.125

4.1.2 CFR § 179.13 – Tank Car Capacity and Gross Weight Limitation

The maximum gross weight of a tank car is 286,000 lbf (129.7 te). The weight of a standard 

DOT-105 chlorine-filled tank car is 247,000 lbf (112.0 te), leaving approximately 39,000 lbf 

(17.7 te) available for new mitigation structure. This equates to approximately 29.4 lbf/ft2 (~144 

kg/m2) based on the calculated surface area of the jacket (see Case #1 in Table 11). The 

allowable weight per area would clearly increase if the foam and jacket were removed or

replaced as part of the new mitigation solution. 

4.1.3 CFR § 179.18 – Thermal Protection Systems

Tank cars requiring thermal protection must have sufficient thermal resistance so there will be no 

release of any lading within the tank car, except through the pressure relief device, when 

subjected to: 

(1) A pool fire for 100 minutes 

(2) A torch fire for 30 minutes 

For the DOT-105 specification, thermal protection is provided by way of the 4-inch-thick foam 

insulation, enclosed by the outer A1011 steel jacket. 

4.2 Mitigation – Enhanced TC128-B 

In the first mitigation solution the team considered the mechanical properties of tank car steel. 

In 2008 the Southwest Research Institute (SwRI) undertook a study to investigate the properties 

of tank car steels retired from the fleet [35]. The study found that most samples tested (97 

percent) met the current AAR TC128-B specification for elemental composition whereas 82

percent of the samples met the tensile property requirements. While tank car specifications allow 
38 

for several different grades of steel, most of the pressure car fleet are manufactured from TC128-

B (>93 percent in 2006). 

In response to a mandate by AAR to develop tougher steels to improve puncture resistance, 

ArcelorMittal (AM) undertook research and development of a new low-carbon tank car steel

[36]. The new steel showed improved toughness properties in normalized and post weld heat 

treated condition while meeting the required tensile properties. 

AM conducted a series of round and notched bar tensile tests to characterize the stress-strain 

relationship and develop a B-W fracture model for both the current steel and new steel. A FE

model of a tank car was developed and subjected to side-impact loading by a spherical object. 

The material in the puncture zone was modelled using both current and new steel, with AM 

reporting an increase in energy absorbed by the new steel to be in the region of 12 – 14 percent. 

The experimental engineering stress-strain curves obtained by AM are shown in Figure 47

below. The tensile strength of the new steel is lower than the current steel, but meets the

minimum tensile strength specified for TC128-B material (81 ksi, or 558.5 MPa). However, the 

failure strain for the new steel is higher, indicating greater ductility. 

Figure 47. Experimental Engineering Stress-Strain Curves for Round and Notched Tensile 

Specimens: a) Current Steel, b) New Steel [36]

AM published the true stress-strain curves developed for use in FE analysis (Figure 48), along 

with the damage initiation criteria for both current and new steel (Figure 49). The damage model 

assumes that the failure initiates at the maximum load carrying capacity of the material (i.e., 

damage initiation) and then shows a subsequent degradation of stiffness progressing to complete 

failure (i.e., damage evolution). Figure 49 shows that damage initiation, which is defined in 

terms of equivalent plastic strain, is dependent on stress triaxiality, and that this is higher for the 

new steel than the current steel. Figure 49 also shows that the round bar and the notched tensile 

tests enabled AM to characterize the damage initiation criteria for tensile failure modes, but not 

for shear or mixed failure modes (i.e., stress triaxialities less than 0.33). Damage evolution, 

which defines the material behavior after damage initiation, is represented by its fracture energy 

dissipated per unit area during the damage process. AM calculated the fracture energy dissipated

for both steels and found those for the new steel to be higher (Table 12).
39 

Figure 48. ArcelorMittal True Stress-strain Curves Derived for Use in FE Analyses [36]

Figure 49. Damage Initiation Criteria for Current and New Steel Grades: a) Current Steel, 

b) New Steel [36]

Table 12. Fracture Energies for Current and New Steel Grades

Specimen 

Current Steel

(J/mm2)

New Steel

(J/mm2) % Difference

Round 3.15 3.25 +4.1

6.35 mm notch 0.96 1.15 +19.8

2.54 mm notch 0.57 0.86 +50.9

1.27 mm notch 0.46 0.64 +39.1
40 

To assess the potential effectiveness of the new steel to improve puncture resistance, the fullscale model described in Section 2 was exercised using a modified version of the Actual TC128-

B material referred to as ‘Enhanced TC128-B.’ Properties for the new virtual Enhanced TC128-

B material were obtained by comparing AM current steel to new steel and scaling the Actual 

TC128-B by a similar amount: 

• Current Steel x Scaling Factor(s) = New Steel

• Actual TC128-B x Scaling Factor(s) = Enhanced TC128-B 

The AM true stress-strain curves and damage initiation criteria curves were digitized and 

overlaid to obtain scaling factors that could be applied to the Actual TC128-B material. As 

shown in Figure 50, it was calculated that factoring the current steel true-stress strain curve by 

0.94 gave a reasonable approximation of the new steel. Similarly, as shown in Figure 51, 

factoring the current steel damage initiation criteria by 1.23 gave a good approximation of the 

new steel. Finally, the average percentage increase in fracture energy, obtained from Table 12, 

was calculated to be +28.5 percent. These scaling factors are summarized in Table 13. 

Figure 50. True Stress-strain Curves: Current and New Steel from ArcelorMittal, Factored 

Steel is Current Steel x 0.94

Figure 51. Damage Initiation Criteria: Current and New Steel from ArcelorMittal, 

Factored Steel is Current Steel x 1.23 
41 

Table 13. Scaling Factors Obtained by Comparing ArcelorMittal Current and New Steels

Material Parameter Scaling Factor

(Current to New)

True stress-strain 0.94

Damage initiation 1.23

Damage Evolution (Damage progression energy) 1.285

The Enhanced TC128-B material model properties obtained when the scaling factors in Table 13

were applied to the Actual TC128-B are shown in Figure 52 and Figure 53. The damage 

progression energy was initially 1400 in-lbf/in2 and was increased to 1799 in-lbf/in2. 

The FE model developed to replicate the ASTM-E8 round bar tensile test and used to verify the

original Actual TC128-B material model (Appendix B) was modified to include the new material 

properties developed for Enhanced TC128-B. The simulation was repeated and a new 

engineering stress-strain curve for Enhanced TC128-B obtained. This was compared to the 

Actual TC128-B material in Figure 54. The overall characteristics match those of the AM current 

and new steels (i.e., the Enhanced TC128-B material has lower strength and greater ductility). 

To assess how these changes translate to improved puncture resistance, the full-scale model was 

used with Enhanced TC128-B in the puncture zone. Test 6 was repeated (indenter speed 15.16 

mph) and the tank found to fail in a similar manner as before. The force-displacement response 

was compared to the baseline response using Actual TC128-B in Figure 55. The peak force 

exhibited by the Enhanced TC128-B version was ~6 percent lower, which aligns with the 0.94 

scaling factor applied to the true stress-strain curve. The displacement to failure was otherwise 

very similar, indicating that the improved ductility of the Enhanced TC128-B did not offer 

significant benefit in this case.

Figure 52. True Stress-strain Curves: Actual TC128-B and Enhanced TC128-B 

True Plastic Strain

True Stress (ksi)

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0.0E+0

2.0E+4

4.0E+4

6.0E+4

8.0E+4

1.0E+5

1.2E+5

1.4E+5

1.6E+5

1.8E+5

Actual TC128-B

Enhanced TC128-B
42 

Figure 53. Damage Initiation Criteria: Actual TC128-B and Enhanced TC128-B 

Figure 54. Engineering Stress-strain Curves: Actual TC128-B and Enhanced TC128-B 

Stress Triaxiality - T

Equivalent Plastic Strain - PEEQ

-0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

Actual TC128-B

Enhanced TC128-B

Engineering Strain

Engineering Stress (ksi)

0 0.05 0.1 0.15 0.2 0.25 0.3

0

20

40

60

80

100

Actual TC128-B

Enhanced TC128-B
43 

Figure 55. Force-displacement Response, Full-Scale Model 15.16 mph; Actual and 

Enhanced TC128-B Materials

4.2.1 Summary 

AM developed a new low-carbon version of TC128-B that was shown to have lower strength but 

greater ductility, with a projected improvement in puncture energy in the range of 12 – 14 

percent. Since the damage models developed for FEA are mesh dependent it was not feasible to 

simply implement the AM material model for new steel into the validated full-scale model. 

However, it was feasible to compare the differences observed between the properties of the AM 

current steel and new steel and apply these to the full-scale model. This was achieved by scaling 

the properties of the validated Actual TC128-B material by the same scaling factors observed 

between the current and new steel, to create a new virtual Enhanced TC128-B material. 

Enhanced TC128-B was not shown to significantly improve puncture resistance. The full-scale 

model punctured at 15.16 mph with the new material model. The survival speed of the full-scale 

model using Actual TC128-B was calculated to be 13 mph. As shown in Table 14, the KE of the 

indenter increased by 16 percent from 13 to 14 mph, and by 36 percent from 13 to 15.16 mph. 

Therefore, consistent with AM findings, the change in material properties may result in an 

increase in puncture speed in the region of 1 mph. Therefore, while beneficial, the improved 

ductility of the steel is not sufficient to result in a step-change improvement in puncture 

resistance. It is potentially part of the solution, but not the solution itself. 

Table 14. Summary of Kinetic Energies

Speed 

(mph)

KE (million-ftlb)% KE Increase 

from 13 mph Comment

13.00 1.679 - Baseline survival speed (Actual TC128-B)

14.00 1.947 16 Potential survival speed (Enhanced TC128-B)

15.16 2.283 36 Test Speed & Failure of Enhanced TC128-B 

Impactor Travel (inches)

Impactor Force (kips)

0 5 10 15 20 25 30 35 40 45

0

200

400

600

800

1000

1200

1400

Actual TC128-B

Enhanced TC128-B
44 

4.3 Mitigation – Shell Thickness

One clear solution for improving puncture resistance is to increase the thickness of the tank shell. 

There is already variation in thickness across different tank car specifications, with shell 

thicknesses as low as 0.562 inch (DOT 105J300W) and as high as 0.954 inch (DOT 105J600I) 

within the TIH tank car category [37]. To investigate the relationship between thickness and 

puncture resistance, the full-scale model was modified to consider shell thicknesses of 1.25 and 

1.50 inches. This was achieved by keeping the mid-plane diameter of the tank shell constant and 

increasing the thickness of the mid-plane shells and solid element region. This meant that the 

diameter of the membrane part, representing the fluid boundary, reduced slightly as did the 

thickness of the foam insulation (i.e., by 0.2375 inches in the 1.25-inch configuration, and 

0.3625 inches in the 1.50-inch configuration). The diameter of the jacket remained the same.

The resulting overall weight increase is shown in Table 15. Based on a standard DOT 105 

chlorine-filled tank car having a weight of approximately 247,000 lbf, there is an approximate 

39,000 lbf weight increase available to remain within the 286,000 lbf limit. The percentage 

weight increases shown in Table 15 were compared against the notional 39,000 lbf available. It 

shows that increasing the shell thickness to 1.50 in consumes approximately 90 percent of the 

available weight and therefore represents a reasonable, practical upper limit for shell thickness. 

Table 15. Tank Shell Weights (Approximate)

Shell Thickness (in) Shell Weight, Full 

(lbf) Weight Increase (lbf) % of Available Weight Increase

0.775 38,900 - -

1.250 61,700 22,800 58

1.500 73,900 35,000 90

The team conducted a series of puncture analyses at varying speeds to establish the maximum 

survival speeds using a 12 x 12-inch indenter for 1.25- and 1.50-inch-thick shells. The results are 

summarized in Table 16. They show that increasing the shell thickness to 1.25 inches increases 

the puncture resistance by 6.0 mph and approximately doubles the amount of kinetic energy 

dissipated. Increasing the shell thickness to 1.50 inches gives a 9.0 mph increase in survival 

speed, which equates to approximately 3 times the kinetic energy. 

Table 16. Results – Tank Shell Thickness Study

Shell Thickness 

(in)

Indenter Speed 

(mph)

Indenter KE

(million-ft-lb)

Response of 

Tank Shell

Speed Increase 

(mph)

KE Increase 

(%)

0.775 13.0 1.679 Survives - -

0.775 14.0 1.947 Punctures - - 

1.250 19.0 3.586 Survives 6.0 113.6

1.250 20.0 3.973 Punctures - -

1.500 22.0 4.808 Survives 9.0 186.4

1.500 23.0 5.255 Punctures - - 

The force-displacement curves for the three survival speeds are shown in Figure 56. They show 

that the thicker shells can accommodate both higher peak forces and greater tank deformations. 

Figure 57 shows cross-sections through both cases at the time of peak force application. 
45 

Figure 56. Force-displacement Response, Full-Scale Model: 0.775, 1.250- and 1.500-inchthick shells – Survival Speeds

Figure 57. Cross-section through Tanks at Peak Force (left: 1.25 inch, right: 1.50 inch) 

Increasing the shell thickness shows significant benefit. It increases the bending strength and 

stiffness of the tank, requiring more energy to deform it, and it improves the local puncture 

resistance. 

4.3.1 Increased Shell Thickness with Enhanced TC128-B Material

Increasing the shell thickness to 1.5 inches increased the puncture resistance of the tank from 

13.0 to 22.0 mph when using the Actual TC128-B material. To ascertain if using the more 

ductile material (i.e., Enhanced TC128-B) would improve the response further, the team repeated 

the full-scale simulations with 1.5-inch-thick shell. Researchers determined that the survival 

speed would increase to 23.5 mph using Enhanced TC128-B. This is a +1.5-mph improvement, 

and an increase in kinetic energy of 14 percent, which is consistent with the 12 – 14 percent 

predicted energy absorption calculated by AM [36]. The results are summarized in Table 17 and 

illustrated in the puncture chart in Figure 58. In this figure, the speeds shown are the calculated 

Impactor Travel (inches)

Impactor Force (kips)

0 5 10 15 20 25 30 35 40 45

0

500

1000

1500

2000

2500

3000

13 mph, 0.775 inch

19 mph, 1.25 inch

22 mph, 1.5 inch
46 

survival speeds, where an increase in speed of +1 mph beyond these values would result in 

failure. 

Table 17 and Figure 58 also show the results of an additional set of sensitivities undertaken in 

which the theoretical maximum shell thickness was used with Enhanced TC128-B material. The 

maximum shell thickness to stay within the 286,000 lbf weight limit was calculated to be 1.575 

inches. This utilized approximately 99 percent of the available 39,000 lbf and resulted in a 

calculated survival speed of 24.5 mph. A sensitivity at double the baseline speed (i.e., 26 mph) 

resulted in failure of the tank. The general behavior of the tank at the 24.5 mph survival speed is 

shown in Figure 59, where (a) shows the tank before impact, (b) shows tank deformation before 

the entire tank assembly mobilizes, (c) shows tank compressed between indenter and wall, and 

(d) shows both the impactor rebounding away from the tank shell and the tank rebounding away 

from the wall. Figure 60 shows a plastic strain contour plot of the inside surface of the tank shell 

after the impact event. It shows that the material remains intact, with locally high plastic strains 

at the impactor corners. 

Table 17. Enhanced TC128-B, 1.50 Inch Shell

Shell Thickness 

(in)

Indenter Speed 

(mph)

Indenter KE

(million-ft-lb)

Response of 

Tank Shell

Speed Increase 

(mph)

KE Increase 

(%)

0.775* 13.0 1.679 Survives - -

0.775* 14.0 1.947 Punctures - - 

1.500 23.5 5.255 Survives 10.5 213.0

1.500 24.5 5.962 Punctures - -

1.575 24.5 5.962 Survives 11.5 255.1

1.575 26.0 6.715 Punctures - - 

*Indicates tank shell material is Actual TC128-B.

Figure 58. DOT-105 Puncture Chart Showing Calculated Survival Speeds (12-inch 

Indenter, Side Impact Scenario)
47 

Figure 59. Overview of Response: 1.575-inch-thick Shell & Enhanced TC128-B at 24.5 mph 

Figure 60. Plastic Strain Contour Plot (Upper Limit set to 0.4) Showing Inside Face of 

Tank Shell
48 

4.3.2 Summary

In summary, the more ductile but slightly lower strength material further improved puncture 

resistance when considering thicker shells. At 1.5-inch-thick, the indenter displacement-tofailure increased from ~39.2 to 42.2 inches when using the Enhanced TC128-B material and in 

doing so facilitated greater energy dissipation. The calculated survival speed of 23.5 mph is 1.8 

times greater than the baseline of 13.0 mph. 

At the calculated theoretical maximum shell thickness of 1.575 inches, the survival speed 

increased to 24.5 mph when implemented with the Enhanced TC128-B material. This indicates 

that using thicker, ‘better’ steel could almost double survival speed and provide a resistance that 

is close to the average mainline derailment speed of 26 mph [37]. 

These results also indicated that improving ductility was of more benefit than increasing 

strength. This has been explored further in subsequent sensitivities. 

Finally, the thickness sensitivities discussed in this section were undertaken using the full-scale 

FE model. Increasing the shell thickness increased the simulation times, primarily due to the 

large number of solid elements required in the thicker puncture zones. For example, in the case 

of the 1.5-inch-thick shell, the number of solid elements approximately doubled to ~1.32 million, 

which significantly increased the computational demand. This partly motivated the need to 

develop the more efficient quarter-symmetry sub-model described previously in Section 2 and 

used in subsequent sensitivities. 

4.4 Mitigation – Material Properties

The thicker shells and the enhanced steel showed varying degrees of benefit in improving 

puncture resistance. To investigate some of the key factors for these improvements, different 

steel properties were considered separately in a parametric study. The features considered were

mass, strength, and ductility. 

4.4.1 Mass

Increasing the plate thickness of the tank shell from 0.775 to 1.5 inches resulted in a weight 

increase of approximately 35,000 lbf. This equates to approximately 12.4 percent of the overall 

tank car weight when using a 1.5-inch-thick shell. To determine the influence of the additional 

mass, a sensitivity was undertaken using the baseline full-scale model with 0.775-inch-thick shell

and Actual TC128-B material and artificially increasing the additional mass gained in the 

thickness increase (35,000 lbf added). Accounting for the half-symmetric modelling approach, 

17,500 lbf of non-structural mass2 was assigned to the tank shell and the baseline simulation at a 

speed of 15.16 mph repeated. The results showed no significant change, with the tank again

puncturing at an impactor travel distance of approximately 38 inches (Figure 61).

2 Non-structural mass is akin to scaling the density of the target elements. The mass-inertia is increased, but the 

element otherwise performs the same. 
49 

Figure 61. Force-displacement Response, Test 6 Baseline Results and Sensitivity with 

35,000 lbf Added Mass

The results indicate that added mass of the magnitudes considered here has little consequence on 

the puncture resistance and therefore, improved performance of the 1.5-inch-thick tank car must 

be achieved elsewhere. 

4.4.2 Material Strength

To investigate the relationship between steel strength and puncture resistance and provide insight 

into the key factors for the improvement offered by thicker steels, the team conducted a series of 

analyses using the quarter-symmetry tank car sub-model. 

First, however, a high strength steel material model was developed and tested. Simple flat bar 

tensile test models were developed in accordance with ASTM E8 (8-inch gauge). The test, 

illustrated in Figure 62, was set-up to compare the characteristics of 0.775-inch-thick plate 

against 1.5-inch-thick plate, and therefore facilitate development of a new virtual material model 

that would replicate the strength of the thicker plate on the thinner specimen. 

Impactor Travel (inches)

Impactor Force (kips)

0 5 10 15 20 25 30 35 40 45

0

200

400

600

800

1000

1200

1400

1600

15.16 mph, 0.775 inch scaled density

15.16 mph, 0.775 inch
50 

Figure 62. Flat Bar Tensile Test Setup: left: 0.775-inch Model, right: 1.5-inch Model

The force-displacement curves extracted from the analyses are shown in Figure 63. For the 1.5-

inch-thick specimen, the peak force increased by ~94 percent (the same as the thickness ratio) 

and the energy dissipated at failure increased by 112 percent when compared to the 0.775-inchthick specimen. 

Figure 63. Force-displacement Response, 0.775- & 1.5-inch-thick Specimens Using Actual 

TC128-B Material Model

A new material model was then developed in which the strength of the Actual TC128-B steel 

was factored up by 1.935 (i.e., the thickness ratio of the 1.5-inch shell vs. 0.775-inch shell). This 

was achieved by factoring up the true stress-plastic strain curve as shown in Figure 64 with all 

other material parameters remaining the same. The new High Strength Actual TC128-B material 

model was assigned to the 0.775-inch-thick flat bar specimen and the tensile test was repeated. 

The engineering stress-strain curves and force-displacement curves obtained are shown in Figure 

65. The results show that the yield stress and ultimate tensile stress were factored up by ~94 
51 

percent, and that the peak force obtained using the High Strength Actual TC128-B matches that 

obtained from the 1.5-inch-thick specimen. The overall energy absorbed at failure is somewhat 

less for the 0.775-inch high strength specimen. This difference is attributed to failure occurring at

different stress triaxialities, manifesting because of the different geometries. Nevertheless, the 

High Strength TC128-B material model developed was considered sufficient to investigate the 

role of material strength in relation to puncture resistance. 

Figure 64. True Plastic Stress-strain Curves for Actual TC128-B & High Strength TC128-

B 

Figure 65. Left: Engineering Stress-strain Curves, Actual vs High Strength TC128-B, and 

right: Force-displacement response, 0.775-inch specimen with High Strength Actual

TC128-B vs 1.5-inch Specimen with Actual TC128-B 

The High Strength TC128-B material model was implemented into the tank car quartersymmetry sub-model and the puncture speed varied in 1 mph increments until the survival speed 

was determined. This has been compared against the survival speed obtained from both the fullscale model and quarter-symmetry sub-model when a shell thickness of 1.5 inches was
52 

implemented (Table 18). Note that the survival speed for the quarter-symmetry model at 1.5 

inches is 1 mph higher than the equivalent full-scale model. This is considered a reasonable 

agreement, with the two models comparing well at the validation speed of 15.16 mph (Section

4.2). Furthermore, the quarter-symmetry model at 1.5-inches-thick failed at 23.5 mph (i.e., a 

small increase in speed). 

Table 18. Survival Speed Summary, High Strength vs. Steel Thickness

Model

Shell 

Thickness 

(in)

Material Indenter Speed 

(mph)

Response of 

Tank Shell

Speed 

Increase 

(mph)

Full-Scale 0.775 Actual TC128-B 13.0 Survives - 

Full-Scale 1.5 Actual TC128-B 22.0 Survives 9.0

Q. Sym. 1.5 Actual TC128-B 23.0 Survives 10.0 

Q. Sym. 0.775 High Strength

Actual TC128-B 20.0 Survives 7.0

The results show that the high strength steel improved the puncture resistance by 7.0 mph, which 

is approximately double the puncture energy (+137 percent). This is broadly consistent with the 

findings of Anderson et al. [38], who previously investigated puncture resistance in tank cars and 

concluded that ‘roughly a 2-fold increase in puncture energy is possible with improved steels.’

Anderson et al. considered three different high strength-low alloy (HSLA) steels in comparison 

to grade TC128-B material, with the highest strength steel grade (HPS 100) having measured 

yield strengths in the region of 115 – 125 ksi, depending on strain rate. The yield strength of the 

Actual TC128-B considered in this study is 71.5 ksi, with the High Strength Actual TC128-B 

having a yield strength of 138 ksi, as summarized in Table 19. 

Table 19. Comparison of Steel Strengths 

Material Yield

(ksi)

Strength

(MPa)

TC128-B (minimum) 50 345

Actual TC128-B 71.5 493

HPS 100 [38] 115 793

High Strength Actual TC128-B 138 951

The yield strength of the High Strength Actual TC128-B steel is therefore very much at the upper 

end of the scale and lies outside of the current TC128-B specification, which requires that the 

UTS is in the range of 81 – 101 ksi. The calculations also assume that ductility remains the same, 

whereas it is normal for higher strength steels to have reduced ductility. Nevertheless, the results 

show significant benefit in using higher strength steel grades and indicate that a significant 

proportion of the energy required to cause puncture of thicker steel can be translated to using 

steel with higher strength, provided ductility can be retained. 

4.4.2.1 Material Strength vs. Structural Strength

The strength component considered above is the material strength (i.e., the resistance of the 

material to yielding and plastic deformation). However, there is also a structural strength and a 
53 

global resistance to bending provided by the 1.5-inch-thick shell that is not captured by simply 

increasing the strength of the material. This is illustrated in Figure 66, which shows a relatively 

flat response of the tank shell in front of the indenter in the 0.775-inch-thick high strength case. 

The 1.5-inch-thick shell by contrast resists the flattening behavior, which increases the local 

punching effect. This behavior is further apparent when comparing the indenter forcedisplacement curves (Figure 67) which shows that the global stiffness of the 1.5-inch-thick shell 

is significantly greater than the 0.775-inch-thick shell with high strength material. Therefore, 

more energy is absorbed by the thicker shell over shorter deformation. 

Figure 66. Comparison of Global Deformation at 22 mph; 0.775-inch-thick High Strength

vs. 1.5-inch-thick Actual TC128-B 

Figure 67. Force-displacement Response at 22 mph, Actual TC128-B 1.5-inch vs High 

Strength TC128-B 0.775-inch

Impactor Travel (inches)

Impactor Force (kips)

0 10 20 30 40 50 60

0

500

1000

1500

2000

2500

3000

Actual TC128-B 1.5 in

High Strength Actual TC128-B 0.775 in
54 

4.4.3 Ductility

The results obtained from simulations using the Enhanced TC128-B material (Section 4.2 and 

Section 4.3.1) indicated that improved puncture resistance could be obtained by increasing 

ductility while reducing the strength of the steel. An approximate 6 percent reduction in steel 

strength, coupled with a 23 percent increase in elongation to failure resulted in approximately 12 

– 14 percent improvement in puncture energy. This suggests that puncture resistance is strongly 

dependent on ductility. 

To further investigate this, a new virtual High Ductility version of the Actual TC128-B material 

model was developed. Recall the Actual TC128-B material model was developed using a B-W 

fracture locus in which the plastic strain at failure is dependent on stress triaxiality. In the High 

Ductility version, the single change made to the material model was to factor the B-W curve up 

by 4, as shown in Figure 68. 

The Round Bar model previously developed to verify the Actual TC128-B material model 

(Appendix B) was utilized with the new High Ductility Actual TC128-B material. The resulting 

engineering stress-strain curves are shown in Figure 69, where the High Ductility results are 

compared against the baseline Actual TC128-B results. The increase in ductility is clear, 

although it should be noted that the baseline Actual TC128-B only just meets the minimum 

ductility requirements (22 percent elongation for 2-inch gauge). It is also noted that while the 

elongation of the High Ductility model is more than 40 percent, it is not significantly higher than 

some other tank car steels previously tested. For example, Kirkpatrick [39] compared the 

engineering stress-strain curves of three tank car steels, with ‘Test 2’ material showing a failure 

strain in the region of 30 – 35 percent. 

Figure 68. B-W Damage Model for Actual TC128-B & High Ductility Actual TC128-B 

Stress Triaxiality - T

Equivalent Plastic Strain - PEEQ

-0.5 0 0.5 1 1.5

0

0.5

1

1.5

2

2.5

3 Actual TC128-B

High Ductility TC128-B
55 

Figure 69. Engineering Stress-strain Curves Produced from Round Bar (2-inch Gauge)

Model Using Actual TC128-B & High Ductility Actual TC128-B 

The High Ductility material model was then implemented in the quarter-symmetry tank car 

model and a series of simulations undertaken in which the indenter speed was increased in 1 mph 

increments until failure was observed. The results are summarized in Table 20, where in each 

case a 1 mph increase in speed resulted in puncture of the tank shell. 

Table 20. Survival Speed Summary, High Ductility vs. Steel Thickness

Model

Shell 

Thickness 

(in)

Material Indenter Speed 

(mph)

Response of 

Tank Shell

Speed 

Increase 

(mph)

Full-Scale 0.775 Actual TC128-B 13.0 Survives - 

Q. Sym. 0.775 High Ductility

Actual TC128-B 19.0 Survives 6.0

The results show that the survival speed increased by 6.0 mph, which is a 114 percent increase in 

kinetic energy, again supporting the view that improved steels could approximately double the 

puncture energy of tank cars. Figure 70 compares the force-displacement response obtained from 

the quarter-symmetry model when the baseline tank car using Actual TC128-B material fails at 

an indenter speed of 14 mph, versus that of the High Ductility model where failure occurs at a 

speed of 20 mph. The impactor travel increases from ~36 to 48 inches, and the peak forces 

increase by 40 percent, from 1309 to 1833 kips. An overview of the High Ductility model at the 

time of peak displacement in the 19-mph survival case is shown in Figure 71. Local deformation

of the shell behind the indenter is notably higher than in previous cases due to the increased 

ductility (Figure 72).

Engineering Strain

Engineering Stress (ksi)

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0.5

0

20

40

60

80

100

Actual TC128-B

High Ductility TC128-B
56 

Figure 70. Force-displacement Response, Actual TC128-B vs High Ductility TC128-B at 14 

& 20 mph 

Figure 71. Section through High Ductility Model, 19 mph, at Time of Peak Indenter 

Displacement

Impactor Travel (inches)

Impactor Force (kips)

0 10 20 30 40 50 60

0

200

400

600

800

1000

1200

1400

1600

1800

2000

Actual TC128-B 14 mph

High Ductility 20 mph
57 

Figure 72. Local Section View through High Ductility Model, 19 mph, at Time of Peak 

Indenter Displacement

4.4.4 Summary

The parameters of mass, strength, and ductility were investigated in isolation. The influence of 

mass was investigated by artificially scaling up the density of the tank shell steel. The mass 

increase was equivalent to increasing the shell thickness from 0.775 to 1.5 inches. It was found 

to have little influence on the overall puncture resistance primarily because the relative overall 

increase in the tank car mass was small (12.4 percent). 

The material parameters strength and ductility were investigated by developing virtual high 

strength and high ductility material models, and in both cases significant improvements in 

puncture resistance were observed. There is a clear benefit to improving steel ductility that has 

been recognized by the tank car industry for some time. The Actual TC128-B material used as 

the baseline in these analyses and validated against physical tests of tank cars performed in the 

late 1980s lies very much at the lower end of the ductility scale, only just meeting the minimum 

requirement of 22 percent elongation for a 2-inch gauge. AAR have petitioned for the timely 

removal of such legacy TIH/PIH tank cars from service [40], specifically the non-normalized3

cars. The benefit of increasing steel ductility was also demonstrated in the assessment of AM’s 

low-carbon Enhanced TC128-B steel, where puncture resistance improved when ductility was 

increased and strength reduced. 

Increasing steel strength was shown to provide similar benefit, but higher strength steels would 

typically be characterized by having reduced ductility. This feature of increasing strength while 

reducing ductility was not captured in this assessment. Furthermore, fabricating the tank shell 

from high strength steel would add complexity to the manufacturing process, particularly where 

heat treatments are used, since this could affect the strength of the steel. One method of 

implementing a high strength steel as part of a mitigation solution would be to use it in a doubletank arrangement, where the high strength steel would be used as an external protection 

mechanism and therefore not subject to the same manufacturing requirements as the tank shell. 

This is discussed further in the next section of the report. 

3 Normalizing is a heat treatment process used to improve metal ductility and toughness after it has been subjected to 

thermal or mechanical hardening processes. 
58 

4.5 Mitigation – Multi-Plate Solutions

4.5.1 Background

In the ATCCRP (see Section 1.1) researchers investigated various protection strategies over the 

course of a decade and arrived at the following key conclusions [41]: 

1. ‘No new high technology design or material that produced significant new protection 

levels (e.g., composites, crushable foams, advanced engineered energy absorbing 

sandwich panels) has been identified. Traditional tank car designs with monolithic layers 

of good quality steel are relatively efficient structures for resisting the impact threats in 

the railroad safety environment.’

2. ‘The only option identified for possible improvements in puncture protection over the 

HM-246 interim specification car designs are potential optimized sandwich designs, 

requiring alternative steels in the jacket (outer tank) for enhanced puncture protection. 

This conclusion is based on puncture analyses performed for various tank design 

configurations. Functional tank car designs for these sandwich cars were not developed in 

the research and the protection levels have not been proven by testing.’ 

The researchers found that adopting a double tank design, or “tank-within-a-tank” concept, 

yielded somewhere in the region 45 – 100 percent improvement in puncture energy over the 

HM-246 interim designs depending on the original tank car specification. Further improvement 

was demonstrated by changing the jacket material to a high strength (HY-130) or high ductility 

(304L stainless) steel. 

4.5.2 Layered Plate Solution

The double tank concept assessed by ATCCRP involved increasing the plate thickness of the 

outer jacket to create a structural barrier offset from the tank shell. In some cases, material was 

moved from the tank shell to the jacket, although this was shown to be of little benefit. In the 

current research, investigation of the relationship between shell thickness and puncture resistance 

indicated similar, if not slightly improved, puncture resistance, albeit with a single impact 

scenario considered and an overall much smaller data set. However, previous tank car puncture 

research had not considered if there was any merit in splitting a thick plate into multiple layers of 

thinner plates. This differs from the double tank concept in that the tank shell itself would be 

split into multiple layers and layered on top of one another, as opposed to being offset like the 

jacket. This layered plate concept was therefore evaluated using a series of FE models. 

4.5.2.1 Punch Plate Model

During research conducted by ARA [39] a series of experimental punch tests were conducted in 

which 0.488-inch-thick TC128-B plates were tested to failure. The tests were used by ARA to 

validate a TC128-B material model, characterized using a B-W failure surface, for use in 

subsequent full-scale FE simulations. The material characterized was ‘Test 2’ TC128-B, details 

of which can be found in Appendix B. Punch tests were carried out on three configurations, as 

summarized in Table 21, with force-displacement curves extracted each time. Numerical 

simulations of the punch tests were performed by ARA using LS-Dyna, with generally good 

agreement achieved (shown in Figure 73). The experimental punch tests were performed at a rate 

of 1-inch/minute and the punch and manhole radii were set to 0.3 and 1.0 inches, respectively. 
59 

Table 21. ARA Punch Plate Tests

Test Punch Diameter 

(in)

Manhole 

Diameter (in)

Test-50-1 1.0 2.5

Test-50-2 1.5 3.0

Test-50-3 2.0 3.5

Figure 73. Punch Test and Comparison of Experimental and Numerical Results

In this research effort the team first replicated the three ARA punch tests by numerical 

simulation before investigating the effects of varying plate thickness and plate layering. A 

quarter-symmetry FE model was developed in Abaqus with the Explicit solver used to simulate 

the tests. Material properties for ‘Test 2’ TC128-B were developed (see Appendix B) and 

assigned to the plate. The punch and manhole were modelled as rigid bodies with the plate 

modelled using deformable solid elements with an element size of 0.085 inches and a coefficient 

of friction of magnitude 0.3 was assigned globally4

. A cross-section through the model is shown 

in Figure 74; a mirror plane has been used to give the impression of the full model. 

Figure 75 shows an overview of Test-50-3 at the point just prior to the onset of failure (left 

image) as well as the progressive failure of the plate (right image), where failure initiates on the 

underside of the plate. A plug approximately the same diameter as the punch was ejected in each 

case, which is consistent with the experimental results. Force-displacement curves for each of the 

tests were extracted and compared to the experimentally obtained values. The results (Figure 76) 

showed that generally good agreement was achieved. The calculated displacement at failure in 2 

out of 3 cases was within 1 percent of the experimental values, while the peak force in each case 

was within 10 percent. On average the simulated peak force was 8.6 percent lower than the 

experimental force. The greatest error observed was the displacement to failure in Test-50-1, 

although notably the ARA simulation showed a similar response and magnitude of error. 

4 Sensitivities were conducted on CoF using 0.3, 0.4, and 0.5, with little observable difference in the results. 
60 

Figure 74. Punch Test Setup, Abaqus

Figure 75. Punch Test-50-3: Left, Mesh Shown; Right, Failure Progression

Figure 76. Punch Test: Force-displacement Response, Experimental vs. TT FEA 

To initially compare the puncture resistance of a single monolithic steel plate against multiple 

layered plates of equivalent overall thickness, the team conducted a series of simulations in 

which a single 1-inch-thick plate was compared against two 0.5-inch-thick plates. The 

simulations considered are summarized in Table 22. First, the original Test-50-3 setup was 

adopted, but with the revised plate configurations (Figure 77) before further sensitivities were 

carried out in which manhole diameter was increased to 6.0 and 7.0 inches. The diameter of the 

manhole was increased to impose more bending and less shear. 
61 

Table 22. New Punch Plate Simulations

Test Plate Configuration Punch Diameter (in) Manhole Diameter 

(in)

Test-50-3 Original (0.488 in) 2.0 3.5

Test-50-3 (single) Single 1 in 2.0 3.5

Test-50-3 (double) Two 1⁄2 in 2.0 3.5

New 6-in single Single 1 in 2.0 6.0

New 6-in double Two 1⁄2 in 2.0 6.0

New 7-in single Single 1 in 2.0 7.0

New 7-in double Two 1⁄2 in 2.0 7.0

Results from the initial set of simulations are shown in Figure 77, where in the two-plate 

configuration plate separation occurs after failure. The force-displacement plot shows that the 

single plate outperforms the double plate across the three key metrics: peak force, displacement 

to failure, and energy absorbed. When the manhole diameter was increased this trend continued, 

albeit the difference between the single and double plate response was less pronounced. The 

results from the 6.0- and 7.0-inch manhole studies are shown in Figure 78. In both cases the peak 

force was approximately equal between the single and double plates, but the puncture energy 

was on average 16.5 percent less for the double plate configuration owing to the reduced 

stiffness. 

Figure 77. Test-50-3 Comparing Results with New (Double) Plate Configurations

Figure 78. Results from 6.0- and 7.0-inch Plate Tests
62 

While the results from the Punch Plate Model indicated that it was detrimental to split the plates, 

the puncture mode was strongly influenced by local shear and did not consider the wider 

influence of boundary conditions and constraints. An alternative flat plate test was therefore 

devised in which some of those factors could be considered. 

4.5.2.2 Flat Plate Model

The Flat Plate Model is shown in Figure 79. It is a quarter-symmetric model featuring a solid 

element puncture region at the center, connected to a shell element region using tie constraints. 

The plate was assigned the material properties developed for Actual TC128-B and the element 

size in the solid element region was set to 0.085 inches (i.e., that for which the B-W failure 

model had been calibrated). The outer edges of the plate were fixed and a rigid 12 x 12-inch 

indenter was assigned a constant velocity of 15 mph to penetrate the plate. In an initial set of 

sensitivities three different plate thicknesses were considered (0.775, 1.25 & 1.50 inch) and the 

force-displacement response of the indenter was recorded (Figure 80). The results, summarized

in Table 23, show that the percentage increases in peak force and puncture energy are broadly 

consistent with the magnitude of increased plate thickness. 

Figure 79. Overview of the Flat Plate Model

Figure 80. Force-displacement Curves, Single Flat Plates of Different Thickness

Impactor Travel (inches)

Impactor Force (kips)

0 2 4 6 8 10 12

0

500

1000

1500

2000

2500

3000

0.775 inch

1.25 inch

1.5 inch
63 

Table 23. Single Flat Plat Results Summary

Thickness Displacement at 

Failure (in)

Energy Increase 

(%)

Force Increase 

(%)

Thickness 

Increase (%)

0.775 9.5 - - -

1.25 9.0 56 50 61

1.50 8.7 86 81 94

In a subsequent parametric study, the influence of splitting a monolithic steel plate into different 

layers was investigated by splitting the 1.5-inch-thick plate into the following three different 

configurations: 

1. Two 0.75-inch-thick plates

2. One 1-inch-thick plate + one 0.5-inch-thick plate

3. Three 0.5-inch-thick plates

Researchers consider two different boundary condition configurations. In the first, the outer 

edges of all the plate layers were assigned fixed edge restraint, while in the second only the 

lowest plate was assigned fixed edge restraint, with the layered plates simply set loosely on top. 

An illustration of the three 0.5-inch-thick plate configuration with all fixed edge restraint is 

shown in Figure 81. For comparison, Figure 82 shows a cross-section through each configuration 

in the cases where only the bottom plate is fixed. In Figure 82, the indenter displacement is the 

same in all cases, with three of the four cases having punctured. The free movement of the upper 

layers is also visible. 

Figure 81. Plastic Strain Contour Plot of 3 0.5-inch-thick Plates with All Fixed Outer Edges

Figure 82. Overview of Free Edge Configurations vs. Monolithic 1.5-inch-thick Plate
64 

The results are summarized in Table 24, where the displacement at peak force, peak force, and 

puncture energies have been compared against the baseline 1.5-inch-thick plate with fixed edge 

restraint. There is a clear distinction between the two boundary condition configurations, with 

the loose top plates improving the puncture resistance and the fixed layered plate configurations 

reducing it. The improvement in puncture energy resulting from splitting the plate in two is 

relatively modest (26 – 30 percent), while there appears to be significant benefit obtained by 

splitting the plate into three separate layers (231 percent improvement). The force-displacement 

curves for the loose upper plate configurations are shown in Figure 83, where the energy 

absorbed is a function of the area under the curves. There is relatively little change in peak force 

for the different configurations, but the three 0.5-inch-thick plate solution can accommodate a 

much greater displacement to failure, more than double that of the single 1.5-inch-thick plate. 

Further investigation of multiple thinner layers using the Flat Plate Model was not undertaken for 

several reasons. First, the 0.5-inch-thick plate represents a reasonable minimum thickness for the 

tank shell. The calibrated element size of 0.085 inches also shows 6 elements through the 

thickness of a 0.5-inch-thick layer, which is sufficient to capture bending behavior correctly, 

although reducing the number of elements through-thickness further would begin to compromise 

accuracy. Finally, the performance gain of multiple layers was unproven on a more realistic tank 

car model. 

Table 24. Layered Plate Results Summary

Configuration Boundary 

Conditions

Displacement at 

Peak Force (in)

Peak Force 

(kips)

Force 

Change (%)

Energy 

Change (%)

Single 1.5 in Fixed 8.44 2435 - -

Two 3⁄4 in Loose upper plates 11.33 2444 + 0.4 +30

1 & 1⁄2 in Loose upper plates 10.28 2452 + 0.7 +26

Three 1⁄2 in Loose upper plates 17.40 2634 + 8.2 +231

Two 3⁄4 in All plates fixed edges 7.91 2353 -3.3 -27

1 & 1⁄2 in All plates fixed edges 7.65 2224 -8.7 -32

Three 1⁄2 in All plates fixed edges 7.91 2270 -6.8 -29

Figure 83. Force-displacement Curves, Loose Upper Plate Configurations vs. Single 1.5-

inch-thick Plate

Impactor Travel (inches)

Impactor Force (kips)

0 2 4 6 8 10 12 14 16 18 20

0

500

1000

1500

2000

2500

3000

Single 1.5 in plate

Two 0.775 in plates

One 1.0 in & 0.5 in plates

Three 0.5 in plates
65 

4.5.2.3 Layered Plate Quarter-Symmetry Tank Car Model

The results from the Punch Plate Model indicated a reduction in puncture performance by 

splitting plates into multiple layers, whereas results from the Flat Plate Model indicated a 

potentially significant benefit depending on the plate constraint conditions. Given this 

uncertainty, the layered plate solution was further investigated using the previously developed 

and validated quarter-symmetry tank car model. In this investigation, the 1.5-inch-thick shell 

model was used as a baseline against which various layered plate solutions were compared. In 

the first sensitivity, the most promising layered solution from the flat plate model (3 x 0.5-inchthick plates) was assessed assuming all three plate layers were continuous (i.e., a-tank-within-atank-within-a-tank, Figure 84). This is most analogous to the flat plate model with fixed end 

restraints applied to all three layers and therefore the puncture performance of the tank was 

expected to reduce in this configuration. 

Figure 84. Illustration of Tank Shell Separated into Three Continuous but Separate Layers

0.5-inch-thick

The model was subjected to impact by a 12 x 12-inch indenter at a speed of 23.5 mph, the speed 

at which the monolithic 1.5-inch-thick shell was calculated to just fail. A comparison of the 

force-displacement response between the two models is shown in Figure 85. The overall 

response was largely as expected, with the layered shell exhibiting lower stiffness and 

undergoing a larger displacement to failure (48.4 vs 42.5 inches). The peak force of the layered 

shell reduced by 19.3 percent, as did the overall energy absorbed (i.e., 6 percent lower). These 

results supported the findings from the Flat Plate Model, where layering the plates and

constraining the outer edges reduced the overall puncture resistance. 

The team then conducted a parametric study in which the outer layers were split

circumferentially to provide free edge restraint to assess if the benefit of providing free edge 

restraint observed in the Flat Plate Model would be carried through to the tank car model. The 

study considered three different plate thickness configurations, as shown in Table 25. In each 

case circumferential splits were implemented in the outer layers at different intervals (i.e., the 

outer layers we considered to be continuous rings with varying plate widths). In the FE model 

these rings were created by imposing physical breaks at distances 25, 50 and 75 inches from the 

centerline (Figure 86). Accounting for the quarter-symmetry modelling technique, this resulted 

in central rings with widths of 50, 100, and 150 inches. The 12 x 12-inch indenter was used in all 

cases and the speed adjusted in 1 mph increments until the survival speed of the tank was 

determined. The results are summarized in Table 25. 
66 

Figure 85. Force-displacement Curves 23.5mph, 3 x 0.5-inch Continuous Shells vs. Single 

1.5-inch-thick Shell

Figure 86. Overview of Implementation of Vertical Splits

Table 25. Layered-Split Plate Results Summary

# Shell Thickness (in) Additional Steel Layers Outer Plate(s) Width

(in) Survival Speed (mph)

1 1.50 - - 23

2 0.50 2 x 0.5 in 50 23

3 0.50 2 x 0.5 in 100 21

4 0.50 2 x 0.5 in 150 21

5 0.775 0.725 in 50 23

6 0.775 0.725 in 100 < 23

7 0.775 0.725 in 150 < 23

8 1.0 0.5 in 50 21

9 1.0 0.5 in 100 21

10 1.0 0.5 in 150 21

Impactor Travel (inches)

Impactor Force (kips)

0 10 20 30 40 50 60

0

500

1000

1500

2000

2500

3000

1.5 inch shell

3 x 0.5 inch shell
67 

The results in Table 25 show that the calculated survival speed of two configurations (Cases 2 

and 5) matched the survival speed of the single 1.5-inch-thick shell but none of the 

configurations resulted in improved puncture resistance. 

The results also show a clear correlation between vertical split plate width and puncture 

resistance, with wider plates having reduced resistance. Figure 87 shows an overview of the 3-

plate solution (Cases 2, 3, and 4) at time 200ms, approximately the time of peak indenter travel

when subjected to a 21-mph impact. In Figure 87, the outer jacket and foam layers have been 

removed from the images to show the response of the outer plate layers when different split 

widths have been implemented. The difference in response of the 50-inch-wide plate between the 

others is apparent. At 50-inches-wide, the longitudinal constraint is low and because of this the 

outer plate can accommodate greater deformation and withstand puncture for longer. This in turn 

delays puncture of the tank shell, enabling more energy to be absorbed over the course of the 

impact. 

In each of the layered-split plate cases examined, failure of the outer plate preceded failure of the 

tank shell. This is illustrated in Figure 88, where failure of the outer 0.5-inch-thick plate precedes 

failure of the 1.0-inch-thick shell. Therefore, implementing measures that would improve the 

resistance of the outer layer to puncture, such as using stronger or more ductile steel, would 

result in improved puncture resistance. 

In summary, splitting the tank shell into multiple layers of thinner plates of equivalent overall 

thickness did not show an improved puncture resistance as predicted by the Flat Plate Model. 

Nevertheless, the results showed that similar or equal puncture resistance could be obtained with 

multiple different plate combinations. Finally, these findings assume that ductility and strength 

remain constant across the different plate thicknesses. This assumption has been investigated in 

detail in Section 5 of this report. 

Figure 87. Response of Layered-split Plate Case #2 at Time 200ms (Jacket and Foam 

Removed for Clarity) 
68 

Figure 88. Failure of 0.5-inch-thick Outer Layer (top) Preceding Failure of 1.0-inch-thick 

Shell (bottom)

4.5.3 Strengthened Jacket Solution

The tank-within-a-tank solution, in which the outer jacket of the tank is strengthened by 

increasing its thickness and/or using stronger or more ductile metals, was considered by 

ATCCRP to provide the most promising solution for enhancing the puncture resistance of tank 

cars. This conclusion had been reached following a comprehensive review of mitigation 

solutions but had not been proven by testing. 

The quarter-symmetry models developed in support of this research had demonstrated very good 

agreement with physical tests. These models were therefore used to assess the strengthened 

jacket solution using validated material models for both the shell and jacket. The different 

configurations investigated are shown in Table 26, with both the jacket and tank shell modelled 

using Actual TC128-B material. Both continuous and split outer jackets were considered as part 

of the study. The thicker jackets were modelled using the same centerline as the original A1011 

jackets, with the thickness of the foam insulation reduced slightly to account for the slightly 

thicker jacket. This was not considered to have any significant bearing on the results given the 

limited role the foam layer plays in determining puncture resistance. 
69 

Table 26. Strengthened Jacket Solution Configurations

# Shell Thickness (in) Jacket Thickness (in) Jacket Plate Width (in) Impact Location

1 0.775 0.725 Continuous - 

2a 0.775 0.725 50 Middle of plate

2b 0.775 0.725 50 At plate boundary

3 0.775 0.725 100 Middle of plate

4 0.775 0.5 50 Middle of plate

5 1.0 0.5 50 Middle of plate

An overview of the model is shown in Figure 89 showing Configuration #2a. The arrow in 

Figure 89 identifies the location of the circumferential break, in this case 25 inches from the 

centerline. In Configuration #2b, the impact has been considered at the location where two plates 

are split, assuming the plates are free and there is no physical restraint between them. Figure 90

and Figure 91 compare these two configurations. These sensitivities were undertaken to 

determine if the location of the impact affects the performance, with impact centered at a plate 

boundary considered likely to represent the worst-case location for the side impact condition due 

to the perceived potential for the indenter to separate and therefore bypass the jacket plates. 

Figure 90 and Figure 91 show full representations of the tank to illustrate the different impact 

locations only. All analyses were conducted using quarter-symmetry techniques. 

Figure 89. Overview of Quarter Symmetry Tank Car Model with Strengthened Jacket 

(Configuration #2a); Arrow Shows Location of the Circumferential Split

Figure 90. Configuration #2a – Indenter Striking Center of 50-inch-wide Split Plate Jacket
70 

Figure 91. Configuration #2b – Indenter Striking at Center of Plate Boundary (No 

Restraint Between Adjacent Plates) 

Another important feature of the models is that within the vicinity of the indenter, the jacket has 

been modelled using solid elements using the same element length and formulation as the tank 

shell. This is shown in Figure 92, in which the solid element region of the jacket measures 15 x 

15-inches. The solid element region is connected to the shell element jacket using the same local 

mesh refinement and tie constraint technique as adopted in the previously validated tank shell 

models. As per previous parametric studies, the indenter speed was adjusted in 1 mph increments 

to determine the survival speed. The results are summarized in Table 27. 

Figure 92. Overview of Puncture Zone Featuring Solid Element Representation of Both 

Jacket and Tank Shell

Table 27. Strengthened Jacket Configurations - Results

# Shell Thickness 

(in)

Jacket Thickness 

(in)

Jacket Plate 

Width (in) Impact Location Survival Speed (mph)

1 0.775 0.725 Continuous - 22

2a 0.775 0.725 50 Middle of plate 24 

2b 0.775 0.725 50 At plate boundary 23

3 0.775 0.725 100 Middle of plate 23

4 0.775 0.5 50 Middle of plate 22

5 1.0 0.5 50 Middle of plate 23

The results showed that implementing a continuous 0.725-inch-thick jacket using Actual TC128-

B material improved the side-impact puncture resistance by 9 mph, which equates to a 186 

percent increase in puncture energy. An overview of this case is shown in Figure 93. In the cases 
71 

where the 0.725-inch-thick jacket was split at 50-inch intervals, the puncture resistance improved 

further, to 24 mph if the indenter strikes the center of the plate and 23 mph if it strikes a plate 

boundary. When the 0.725-inch jacket plate width was increased to 100 inches the effectiveness 

of splitting the jacket was reduced. This is shown in Figure 94 where Configuration #2a survives 

and #3 fails at 24 mph. 

Figure 93. Overview of Configuration #1 – Continuous Outer Jacket

Figure 94. Response of Configurations #2a and #3 at 24 mph 
72 

Configurations #4 and #5 considered a 0.5-inch-thick jacket protecting 0.775- and 1.0-inch-thick

shells. The puncture resistance was not significantly different in either case (22 and 23 mph), 

which is notable in that increasing the shell thickness from 0.775 to 1.0 inches resulted in only 1 

mph increase in resistance. 

4.5.4 Summary

The sub models used to assess the puncture resistance of multiple layered plates indicated 

conflicting results. The local punch model showed that puncture resistance is reduced if a single 

thick plate is split into multiple layers of thinner plate, whereas the flat plate model indicated that 

significant improvement could be obtained provided the right boundary conditions are adopted. 

Analyses of this solution using a validated quarter-symmetry tank car model indicated that the 

puncture resistance of multiple layers of plate was approximately equal to that of a single thick 

plate, with the boundary conditions of the outer plate layers again influencing the response. 

Finally, the strengthened jacket solution was investigated and again showed puncture resistance 

that was comparable to using a tank shell of equivalent thickness (1.5 inches). 

One of the key findings from these analyses was that if using a multiple plate solution, the 

constraint applied to the outer layers of plate (i.e., jacket) influences the puncture resistance. It 

was found from the quarter-symmetry tank car analyses that implementing circumferential 

breaks in the outer plates can slightly improve puncture resistance versus using monolithic 

plates. However, the benefit was found to be relatively small (i.e., 1 – 2 mph improvement) and 

was tested on a relatively small set of side impact scenarios. 

The comparison of puncture resistance between a single thick plate and multiple plates of 

equivalent overall thickness was undertaken on the assumption that ductility remained the same 

across the different plate thicknesses. The influence of ductility on puncture resistance was 

examined in Section 4.4.3 and was a feature of the Enhanced TC128-B material where it was 

shown that improving ductility, even at the cost of reducing strength, was found to improve 

puncture resistance. The assumption of constant ductility at increased shell thickness is therefore 

a key feature in the response of the thicker shells. For example, if ductility were to reduce with 

increasing shell thickness, then adopting a multi-plate solution could have greater benefit, 

particularly as multi-plate solutions offer the potential to use different higher strength or higher 

ductility materials in the outer layers. An investigation of the relationship between ductility and 

plate thickness was therefore undertaken and is discussed in Section 5. 

4.6 Mitigation – Elastomeric Liners

4.6.1 Internal Containment Liner

An alternative mitigation concept not considered previously is the possible use of internal liners 

to prevent hazmat release in the event of puncture. Such liners have the potential to provide an 

additional ‘line of defense’ against hazmat release in cases where puncture occurred but with 

limited extent. The coupled action of the structural mitigation strategies discussed previously 

with internal liners could provide an additional benefit resulting in reduction of the conditional 

probability of release. 

Research partner TT has been involved in previous projects in which an elastomer (i.e., 

polyurea) has been implemented to improve the blast resistance of concrete structures. In one 

such project, constitutive material models of polyurea were developed for use in FE simulations 
73 

and validated against experimental data. The engineering stress-strain curves for both are shown 

in Figure 95, where ‘Data fit’ represents the test data and ‘PLIH’ a von Mises5 plasticity with 

piecewise linear hardening model. The results show that the polyurea has a high elongation with 

an engineering strain-to-failure of approximately 5. 

Figure 95. Experimental and Simulated Engineering Stress-strain Curves for Polyurea

A material model for polyurea (shown in Figure 95) was developed for use in Abaqus, with the 

process described in Appendix D. A single element test using the Abaqus material model was 

undertaken for verification purposes, with the resulting stress-strain curve comparing well 

against the experimental data (Figure 146).

To assess the potential effectiveness of the internal liner solution, a 0.25-inch-thick layer of 

polyurea was implemented in the validated Test 6 baseline quarter-symmetry tank car model 

(i.e., 0.775-inch-thick shell with 12 x 12-inch indenter and no other structural modifications). 

The liner was implemented as a discrete part, positioned between the tank shell and the fluid 

membrane, as illustrated in Figure 96. The liner was modelled using shell elements with a typical 

element length of 2 inches resulting in 11,111 elements within the quarter-symmetric part 

(Figure 97). The general contact algorithm available within Abaqus was used with hard contact 

and a tangential coefficient of friction of 0.3 assigned. This means that the liner can move 

relative to the tank shell and the fluid membrane, but with a frictional penalty. This is somewhat 

different to a spray-applied liner where there would be some adhesion between the liner and the 

tank shell. The only other modification made to the model was to reduce the diameter of the fluid 

membrane by 0.5 inch to facilitate implementation of the liner. 

5 A von Mises stress is a value used to determine if a given material will yield or fracture. It is mostly used for 

ductile materials, such as metals.
74 

Figure 96. Local Cross-section through FE Model Showing Location of Liner

Figure 97. Internal Liner with Mesh Shown

The team conducted simulations at indenter speeds of 15 and 16 mph (i.e., speeds just greater 

than the puncture speed of the baseline model). In both cases the tank shell punctured, but the 

integrity of the liner remained intact showing that it was effective at preventing hazmat release. 

In the 15-mph case there was little change to the overall performance of the tank, with the 

indenter travelling to a depth of 38.7 inches and stopping. This is the same peak depth reached in 

the baseline 15-mph case with no liner. An overview of the response of the model is shown in 

Figure 98, where the left image shows the overall behavior, and the right image shows a plastic

strain contour plot of the liner at the end of the simulation. The peak plastic strain reached in the 

liner is approximately 0.77, which coincidentally is approximately 77 percent of the failure

strain. This behavior is further illustrated in Figure 99. 

Figure 98. Response of Liner Model to 15-mph Impact (Right Image Shows Plastic Strain) 
75 

Figure 99. Plastic Strain Contour Plot Showing Response of Liner to 15-mph Impact

When the indenter speed was increased to 16 mph the indenter travel increased to 42 inches

(+3.3 inches). This in turn increased the demand on the liner, with peak plastic strains found to 

reach 0.924, which is close to the limiting plastic strain. Given the high plastic strains observed it 

was concluded that any subsequent increase in indenter speed would likely result in failure of the 

liner, particularly as the mesh used for the liner was relatively coarse (2-inch element size) and 

mesh refinement may lead to locally higher plastic strains. 

To consider potential variation in liner material properties, the plastic strain-to-failure of the liner 

material was reduced by approximately 50 percent and set to 0.5. The 16-mph simulation was 

repeated using the less ductile liner material with failure of the liner observed to occur at an 

indenter travel of 40 inches. Figure 100 compares the response of the liner in the 16-mph case 

where the liner has the original and reduced ductility material properties. In the reduced ductility 

case, failure of the liner is evident with element deletion occurring. For this type of liner solution 

to be effective a material with high elongation to failure is required. 

Figure 100. Plastic Strain Response of Liner in 16-mph Scenario: left – Polyurea Material, 

right – Reduced Ductility Polyurea

4.6.2 Metal-Elastomer Bilayers

In the analyses described above, the elastomeric liner was considered separate to the tank shell

(i.e., no adhesion between the two components). The liner in that scenario was a ‘bag-type’ liner 

that would not be expected to, nor was it found to, increase the puncture resistance of the shell. 

77 

modulus elastomer would yield an improved response, although this has not been investigated 

and there would therefore likely be a reduction in elongation. 

Figure 101. Force-displacement Curves, Steel-elastomer Bilayer Single Element Tensile 

Tests

In mass terms, a 1.33-inch-thick layer of polyurea has an equivalent metal plate thickness of 

approximately 0.21 inches. As seen in previous investigation of shell thickness, increasing the 

shell thickness from 0.775 inches to 0.985 inches would result in a significant improvement in 

puncture resistance.

4.6.3 Elastomeric Liner Summary

Researchers investigated two types of elastomeric liner: one aimed at providing a secondary 

containment mechanism in cases where the shell just punctures, and another aimed at improving 

the puncture resistance of the tank shell. The two types have very different desired material 

properties, with the containment-type liner requiring large elongation and the bilayer-type liner 

requiring a combination of stiffness and elongation. The polyurea layer considered in this study 

was more suited to the former application, with tank car simulations demonstrating that in the 

side impact case the liner could be effective at containing hazmat at impact speeds in the region 

of 1 to 2 mph greater than the puncture speed. The same polyurea material used as the 

elastomeric component of a bilayer with TC128-B indicated little benefit, particularly when 

compared against an all-metal plate of equivalent mass/area. In these calculations the elastomer 

component has been assessed based on its ability to delay necking in the steel. 

4.7 Mitigation – Blunting

The indenter shape and size are known from various previous studies to strongly influence the 

puncture resistance of tank cars. The standard 6 x 6- and 12 x 12-inch indenters used with side 

and head impact tests are notionally considered to represent a broken rail and a car coupler, 

respectively. Puncture speeds for both indenter sizes have been determined in this research effort 

Displacement (inches)

Force (lbf)

0 0.1 0.2 0.3 0.4 0.5 0.6

0.0E+0

2.0E+4

4.0E+4

6.0E+4

8.0E+4

1.0E+5

1.2E+5

1.4E+5

1.6E+5

1.8E+5

Steel only

Stee + 0.25 inch Polyurea

Steel + 1.33 inch Polyurea
78 

making use of the validated FE models. To investigate the potential effectiveness of blunting 

mitigation solutions, an additional set of simulations were undertaken to determine the survival 

speed of the baseline tank car when subjected to impact by a 24 x 24-inch indenter. The indenter 

was modelled with a 2-inch corner radius and the quarter-symmetry model was used for these 

simulations. Figure 102 illustrates the response of the model to a 22-mph impact. 

Figure 102. Illustration of Impact by 24 x 24-inch Indenter

The calculated survival speeds and associated impact energies for each indenter size are 

summarized in Table 29. The results illustrate the effectiveness of blunting, with a 24-inch 

indenter showing approximately the same puncture resistance as increasing the tank shell 

thickness to 1.5 inches. This assumes that the indenter shape is square, which imposes high local 

stresses at the indenter corners and ultimately results in failure of the tank shell in shear. A 

blunting solution that creates a more uniform (e.g., circular) load on the shell would likely be 

more effective as it would reduce the local shear demand. 

Table 29. Survival Speeds for Different Indenter Sizes 

Indenter Size (in) Survival Speed (mph) Impact Energy 

(million-ft-lb)

6 x 6 8.0 0.636

12 x 12 13.0 1.679

24 x 24 22.0 5.255

To achieve a blunting area equivalent to that imposed by a 24-inch indenter when impacted by a 

12-inch indenter, the 4-inch-wide void space currently occupied by foam insulation would need 

to be increased and the foam replaced by a much more robust and stiffer material. This is 

illustrated in Figure 103, which shows both the 12- and 24-inch indenters positioned adjacent to 

the tank. The required increase in space between the tank shell and jacket would depend on the 

chosen material/solution, but any solution implemented within that space would almost certainly 

have a much higher mass per area than the current, lightweight foam material. The new material 

would also have to provide an equivalent thermal resistance to the foam. This being the case, it is 

unlikely that any solution implemented with the specific role of blunting impact loads would be 

any more efficient than simply increasing the shell or jacket thickness. 
79 

Figure 103. Overview of Side Impact Arrangement with 12 x 12-inch (red) and 24 x 24-inch 

(blue) Indenters
80 

5. Investigation of the Relationship between Plate Thickness and 

Ductility

Investigation of the relationship between tank shell thickness and puncture resistance showed a 

strong correlation between the two. Subsequent investigation of multiple layered and offset plate 

solutions, in which the total plate thickness was the same as a thick (i.e., 1.5-inch-thick) shell, 

showed that similar magnitudes of puncture resistance could be attained in various 

configurations. However, these findings assumed that material ductility remains constant across 

different plate thicknesses. If, as suspected, ductility is adversely affected by increasing plate 

thickness, then the puncture resistance of a single thick shell may be overestimated, and similarly 

the resistance provided by multiple layers of thinner plate underestimated. The research team 

investigated this topic as part of the research effort, building on recent work published by 

Anderson et al. [43] which investigated the fundamental differences between plane strain 

bending and far-field plane strain tension in ductile plate failure. 

Anderson et al. [43] employed the micro-mechanics-based Gurson-Tvergaard-Needleman (GTN)

model to investigate the underlying mechanics behind differences observed in fracture strain of 

thin metal plates; it was previously experimentally observed to be significantly higher in plane 

strain bending than in far-field plane strain tension. In that research, a 1.4-mm-thick plate was 

considered in the bending and tension cases; the absence of thinning in the plane strain bending 

condition was identified as one of the key factors behind lower void growth rate and therefore 

greater strain to failure in that case. The results published for the 1.4-mm-thick plate were used 

in the current research as baseline data against which FE models were developed and verified. 

Thicker plates, such as those used in tank car construction were then investigated by modifying

the FE models to determine how the relationship changes with increasing plate thickness. The 

methodology and results from this study are described in the following sections. 

5.1 Gurson-Tvergaard-Needleman (GTN) Model

The GTN model is a micro-mechanics-based material model that is well suited for capturing the 

complexity of the plate fracture process, accounting for crack initiation where voids grow and 

coalesce to micro-crack formation, as well as stress triaxiality dependence. The model operates 

on the basis that individual elements are of a similar size to the distance between microvoids 

within the metal, with the damage parameters calibrated for a specific element size. A change in 

element size would therefore require redefining the damage parameters. Anderson et al. [43]

used an element size Le = t/128, where t is the plate thickness (i.e., 1.4 mm). This gives an 

element size of 0.01094 mm, resulting in 33,603 elements within a simple 1.4 x 10.0 mm 2-

dimensional plate. Therefore, this model, while well-suited for investigation of highly localized 

failures in small samples, would not be well-suited for use in large tank car simulations due to 

the significant scale increase. The number of elements required within the puncture zone would 

simply be too large. 

5.2 Verification of Plane Strain Tension and Bending Models

In the first step in this investigation, the team reproduced the plane strain tension and plane strain 

bending simulations published in [43]. Following this initial verification, the models were 

modified to investigate the response of thicker plates. 
81 

5.2.1 Far Field Plane Strain Tension Case

An overview of the double-edge notch tension (DENT) test used to investigate far-field plane 

strain tension is shown in Figure 104. The ligament at the center of the specimen approximates a 

plane strain condition (i.e., strain in the x-direction = 0). This can be represented using a 2D 

model with plane strain conditions, where plate thickness, t, is 1.4 mm and plate length, L, is 10 

mm. This gives a length to thickness ratio L/t = 7.1, which is sufficiently large to capture 

through-thickness necking, shear localization, and fracture related to ductile tearing. 

Figure 104. Far-field Plane Strain Test Setup – Only the Highlighted 2D Region with 

Length, L, and Thickness, t, Considered in Numerical Simulations [43]

An overview of the FE model researchers developed is shown in Figure 105, with the material 

parameters summarized in Table 30. Across the central portion of the model, with gauge length 

2t = 2.8 mm, the element length was fixed at 0.01094 mm using an enforced partitioning 

strategy. Outside of the gauge region the mesh was coarsened off for computational efficiency. 

At the top and bottom of the specimen an adjustment was made to the material model to reduce 

the initial void volume fraction, f0, from 0.02 to 0.018 to avoid potential localization at the ends 

of the specimen where the boundary conditions were applied. Similarly, a single element located 

at the edge of the specimen at a 45-degree angle from center was assigned a material model with 

the same parameters as those shown in Table 30, but with a 10 percent reduction in the elastic 

modulus to promote localization at the center of the specimen. This element location is denoted 

by the arrowed ‘X’ in Figure 105. 

The upper and lower edges of the specimen were simultaneously pulled to load the specimen in 

tension. A node at the center of the upper and lower edges was also fixed in the x-direction. 

Four-node bilinear plane strain elements with reduced integration (CPE4R) were assigned 

throughout and the Abaqus Explicit dynamic solver was used (with double precision on), to 

quasi-statically load the specimen. Plastic strain contour plots of the specimen undergoing 

through-thickness necking and failure are shown in Figure 106 and Figure 107. The forcedisplacement response of the specimen was extracted and normalized with respect to both the 

yield stress and the original length. This curve is compared to that produced by Anderson et al. 

[43] in Figure 108 with good agreement observed. 
82 

Figure 105. Overview of 1.4 mm Plane Strain Tension FE Model

Table 30. GTN Material Model Parameters

Parameter Notation Value

Density ρ 7,850 kg/m3

Young’s Modulus E 210 GPa

Poisson’s Ratio v 0.3

Yield Stress σ0 630 MPa

Strain hardening exponent N 0.05

Gurson fitting parameters q1, q2, q3 1.5, 1.0, 2.25

Initial void volume fraction f0 0.02

Critical void volume fraction fC 0.15

Final void volume fraction fF 0.25

Figure 106. Through-thickness Necking and Failure of 1.4 mm Plane Strain Tension 

Sample
83 

Figure 107. Initiation and Failure of 1.4 mm Plane Strain Tension Sample

Figure 106 and Figure 107 illustrate the response of the sample at the various stages of 

deformation. Initially, deformation of the plate was homogeneous, followed by a localization 

setting in and a through-thickness neck developing. Development of the neck increased the stress 

triaxiality at the center of the neck to a level significantly above that of plane strain tension 

Further loading caused localization of plastic straining into a narrow shear 

band at the center of the plate, followed by final fracture. This is consistent with the results 

shown by Anderson et al. [43] and in conjunction with the matching force-displacement response 

verifies the plane strain tension model developed here. 

Figure 108. Normalized Force-displacement Curves for the Far-field Plane Strain Tension 

Tests for N = 0.05 and f0 = 0.02 

5.2.2 Plane Strain Bending Case

An overview of the plane strain bending test is shown in Figure 109. It comprises a deformable 

plate of thickness, t, and length, L, positioned above two rigid rollers with radius Rr. A rigid 

punch with a tip radius Rp was centered above the plate and translated vertically downward to 

simulate the test. The half distance between the rollers is denoted lr and is determined from the 

plate thickness (2lr = 2Rp + 3t) such that the plate and punch cannot jam between the rollers. The 

∆y/L

Fy/(

σ 0A0)

0 0.025 0.05 0.075 0.1

0.0E+0

5.0E-1

1.0E+0

1.5E+0

Anderson et al.

TT Simulation
84 

rollers were constrained against translation, but free to rotate, with a tangential coefficient of 

friction μfr = 0.02. 

Figure 109. The Plane Strain Bending Test Setup [43]

The GTN material parameters used were the same as those adopted for the plane strain tension 

case (Table 30) with the remaining model parameters summarized in Table 31. An overview of 

the FE model developed is shown in Figure 110. 

Table 31. Model Parameters for Plane Strain Bending

Parameter Notation Value

Length of plate L 60 mm

Plate thickness t 1.4 mm

Punch radius Rp 0.43 mm

Roller radius Rr 20 mm

Distance between rollers 2lr 5.06 mm

Figure 110. Overview of FE Model of 1.4 mm Plane Strain Bending Case

Figure 111 shows the response of the simulation at the point where instabilities and shear bands 

form on the tension side of the plate followed by crack initiation and propagation. The results 

show good correlation between those produced by Anderson et al. [43] as well as those observed 
85 

by Dykeman [44] during physical experiments. This includes the apparent ‘flattening’ at the midsection and the observation of surface undulations (i.e., wrinkles) inside the region where the 

curvature has stopped evolving. A normalized bending moment-bending angle curve was also 

produced from the simulated test and compared to that produce by Anderson et al. This is shown 

in Figure 112, where again good agreement was achieved. 

Figure 111. (a) Plastic Strain Contour Plots at Late Stage of Plane Strain Bending, Showing 

Surface Instabilities and Shear Band Formation on the Tension Side of Plate Compared 

with That Photographed by Dykeman Investigating Plane Strain Bending of 22mnb5 

Boron Steel Plate with an Aluminum-Silicon Coating; (B) Crack Propagation and Failure 

of Plate
86 

Figure 112. Normalized Bending Moment-bending Angle Curves for the Plane Strain 

Bending Test

5.3 Modified Plane Strain Tension and Bending Models

Following verification of the plane strain tension and plane strain bending models (and 

associated material models), the team conducted a series of sensitivities in which the plate 

thickness was increased. The plate thicknesses considered in the far-field plane strain tension and 

plane strain bending cases are summarized in Table 32 and Table 33. In each case the ‘notional’

and exact plate thicknesses are given since the exact plate thickness was required to be a multiple 

of the fixed element size (0.0109375 mm). In the plane strain tension cases the aspect ratio of 

length-to-thickness ~7.1:1.0 was maintained in all cases, as illustrated in Figure 113. In all cases 

the gauge (length 2t) was assigned a uniform fine mesh, with the mesh coarsening off outside 

that region. 

The approximate number of elements is also provided in each case, and as the plate thickness 

increases the number of elements increases significantly. This has a corresponding effect on the 

simulation times, with the thicker plates taking longer to run (over 1 week for the 25.4 mm case

on a high-powered cluster with 23 processers). It was also necessary to implement mass scaling 

to reduce the simulation times to reasonable timeframes. 

Table 32. Model Parameters for Plane Strain Tension, Increased Thickness Study

Plate 

Thickness, t

(mm)

Exact Plate 

Thickness, t

(mm)

Plate Length, L (mm) Gauge Length, g (mm) Number Elements

1.4 1.4000 10.0000 2.800 ~34E3

3.0 2.9750 21.1225 5.950 ~152E3

9.0 8.9250 63.3675 17.850 ~1.35E6

25.4 25.3750 180.16250 50.750 ~55E6

α[°]

F punch(R r+lr)/(2

σ

0A0t)

0 25 50 75 100 125 150

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

TT Simulation

Anderson et al.
87 

Table 33. Model Parameters for Plane Strain Bending, Increased Thickness Study

Plate Thickness, t

(mm)

Exact Plate 

Thickness, t

(mm)

Plate Length, L

(mm)

Punch Radius, 

Rp (mm)

Spacing Between 

Rollers, 2lr (mm)

1.4 1.4000 60 0.43 45.06

3.0 2.9750 60 0.91 50.39

9.0 8.9250 80 2.74 72.26

12.7 12.6875 90 3.90 85.86

25.4 25.3750 180 7.79 131.71

For the plane strain tension simulations, the sample size was scaled up as the plate thickness 

increased. However, for the plane strain bending cases, some additional parameters needed to be 

fixed. The roller radius was therefore fixed across cases at Rr = 20 mm. The relationship between 

punch radius Rp and plate thickness t was also assumed to remain constant across all cases, as in 

the baseline case, such that t / Rp = 1.4 / 0.43 = 3.26. The equation used by Anderson et al. for 

determining the gap between rollers (2lr) (shown in Figure 113) was then used to complete the 

arrangement. The length of the sample L was just sufficient to bridge between rollers. This gave 

a length-to-thickness aspect ratio for the plates that was greater than that used in the plane strain 

tension simulations. 

 

Figure 113. Illustration of Plane Strain Tension Samples and Plane Strain Bending 

Arrangement

5.3.1 Far-Field Plane Strain Tension Results

The far-field plane strain tension simulations showed a relatively weak relationship between 

plate thickness and ductility, with ductility increasing slightly with increasing plate thickness 

(see Figure 114 and Table 34). Figure 114 shows the normalized forced-displacement curves for 

each sample while Table 34 provides the calculated strain-to-failure. Some results from the 9.0 

mm tension sample are shown in Figure 115 and Figure 116, where the sample necks, triaxiality 

increases, and then fracture initiates. In Table 34, the fracture strain has been calculated across 

the gauge length (2t) in each case. For conciseness not all results are illustrated here. 
88 

Figure 114. Normalized Force-displacement Curves for the Far-field Plane Strain Tension 

Tests for t = 1.4, 3.0, 9.0, and 25.4 mm 

Figure 115. Plot showing Stress Triaxiality in 9.0 mm Plane Strain Tension Sample at 

Onset of Failure

Figure 116. Plot Showing Plastic Strain in 9.0 mm Plane Strain Tension Sample When 

Failure Initiates and Starts Propagating through the Thickness

∆y/L

Fy/(

σ 0A0)

0 0.025 0.05 0.075 0.1

0.0

0.5

1.0

1.5

1.4 mm

2.975 mm

8.925 mm

25.375 mm
89 

Table 34. Results from Plane Strain Tension, Increased Thickness Study

Plate Thickness, t (mm) Fracture Strain (%)

1.4 7.36

3.0 7.87

9.0 8.62

25.4 8.81

5.3.2 Plane Strain Bending Results

The plane strain bending simulations showed a stronger correlation between ductility and plate 

thickness, with ductility decreasing with increasing plate thickness. Figure 117 shows the 

normalized bending-moment bending angle curves for 4 of the 5 samples investigated. It shows 

that as the plate thickness increases, the bend angle at which failure initiates decreases. 

Moreover, the area under the curve in each case represents the energy absorbed before failure, 

which decreases with increasing plate thickness. 

The output required to plot this curve for the 25.4 mm sample was not requested, therefore it has 

not been included in Figure 117. Nevertheless, the output obtained from the 25.4-mm-thick 

sample enabled calculation of the fracture strain, which has been compared to that of the 

remaining samples in Table 35. 

Figure 117. Comparison of the Normalized Bending Moment-bending Angle Curves for the 

Plane Strain Bending Tests at Increased Thickness

In Table 35, the fracture strain calculated was obtained from the bottom (i.e., tensile) surface 

using the gauge length, 2t, at the point at which fracture first initiates. The procedure for 

extracting the fracture strains is shown in Figure 118. The deformed length of the bottom surface 

was obtained by creating a path along the length of the gauge and plotting the true distance of the 

path. In Figure 118 (c) the true distance has been plotted against plastic strain. It shows that in 

this case (i.e., the 25.4-mm sample) plasticity at the extremities of the gauge is close to zero, and 

then gradually increases toward the center where the average plastic strain is approximately 40 

percent. Local spikes in plastic strain are observed, and these are associated with the formation 

of localizations and instabilities on the bottom surface just prior to fracture initiation. This is 

α[°]

F punch(R r+lr)/(2

σ

0A0t)

0 25 50 75 100 125 150

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

1.4 mm

2.975 mm

8.925 mm

12.6875 mm
90 

illustrated by way of example in Figure 119, which shows the response of the 9.0-mm-thick 

sample when fracture initiates. Note that a similar surface undulation or wrinkling pattern is 

observed in the tension face of the bending sample prior to fracture propagation. 

Table 35. Results from Plane Strain Bending, Increased Thickness Study

Plate Thickness, t (mm) Bend Angle

(deg) Fracture Strain (%)

1.4 106 42.9

3.0 73 32.6

9.0 48 23.5

12.7 43 20.6

25.4 35 17.4

Figure 118. Calculation of Fracture Strain: (A) Initial Undeformed Shape of Gauge 

Region; (B) Deformed Shape of Gauge Region at Initiation of Fracture; (C) Plot Showing 

Plastic Strain vs. Distance Along Deformed Gauge Length (bottom surface); (D) Fracture 

Strain Calculation 
91 

Figure 119. Plastic Strain Contour Plot of 9.0 mm Plane Strain Bending Sample When 

Fracture Initiates

5.4 Plate Thickness vs Ductility Results Summary

The results of the plate thickness versus ductility investigation are summarized in Figure 120, 

which shows that there are fundamental differences between plane strain tension and plane strain 

bending. The first fundamental difference is that the fracture stain in plane strain bending is 

higher than in plane strain tension. This was one of the key findings of Anderson et al. [43] and 

is explained by the absence of necking in the plane strain bending case. When necking occurs in 

plane strain tension the stress triaxiality at the center of the neck increases to levels significantly 

above that of plane strain tension (i.e., 0.577). By contrast, the absence of necking in plane strain 

bending maintains the stress triaxiality on the lower tensile surface equal to that of plane strain 

tension. Figure 121 shows stress triaxiality in the 9 mm plane strain bending simulation with the 

lower contour limit set to 0.4. 

The second fundamental difference observed, and which is unique to this study, is the apparent 

reduction in ductility in the plane strain bending condition that occurs with increasing plate 

thickness, and conversely the relatively little change in ductility that occurs over the same range 

of plate thicknesses (i.e., 1.4 – 25.4 mm) in the plane strain tension condition. The reason for this 

behavior is currently not well understood and further research is necessary to understand the 

mechanism behind the observed differences based on thickness. 
92 

Figure 120. Fracture Strain Versus Plate Thickness for Far-field Plane Strain Tension and 

Plane Strain Bending Conditions

Figure 121. Plot Showing Stress Triaxiality in 9-mm-thick Plane Strain Bending Simulation

These findings are consequential for many industries including the tank car industry since the 

results show that increased plate thickness is associated with reduced ductility in bending. In 

previous sections of this report, it was shown that similar levels of puncture resistance could be 

achieved by increasing shell thickness or implementing a multi-layer solution in which the total 

plate thickness is the same as the equivalent thick shell solution. For reference, side impact 

survival speeds in the region of 23 mph were achievable with both solutions when using an 

equivalent shell thickness of 1.5 inches and a 12 x 12- inch indenter. Those findings assumed 

constant ductility across the different plate thicknesses, which varied from 0.5 to 1.5 inches. 

However, the results of this study show that a penalty is incurred by increasing thickness, and 

therefore the survival speed of the monolithic 1.5-inch-thick shell may be lower than that which 

was previously calculated. In the ductility versus plate thickness study, it was not feasible to run 

the 1.5-inch-thick bend test within a reasonable timeframe since the model size simply became 

too large. However, it was possible to extrapolate out using the data gathered from the 1.4- to
93 

25.4-mm-thick samples. The plane strain bending fracture strains relevant to the thicknesses 

considered within the tank car puncture analyses are summarized in Table 36. In the final column 

of Table 36, the relative change in fracture strain has been calculated in comparison to the 

baseline thickness of 0.775 inches for which the BW curve used in the various puncture analyses 

was validated. The table shows that reducing the plate thickness to 0.5 inches gives a 12 percent

increase in ductility, while increasing the plate thickness to 1.5 inches is projected to reduce the 

ductility by 19 percent. While it was not possible within this research effort to quantify how this 

translates to puncture speed, the results indicate that there is potential benefit to using multiple 

thinner plates and consequences for increasing plate thickness. 

Table 36. Summary of Plane Strain Bending Fracture Strains for Tank Car Applications

Plate Thickness, t (in) Plate Thickness, t (mm) Fracture Strain (%) Relative Change to 

0.775-inch (%)

0.500 12.7 20.6 +12.0

0.775* 19.7* 18.4 0 

1.000 25.4 17.4 -5.4

1.500** 38.1** 14.9 -19.0

*Interpolated value, ** extrapolated value
94 

6. Feasibility, Scalability and Cost

This section discusses the feasibility, scalability, and cost associated with the mitigation 

measures considered in Section 4. The focus is shifted from development to design and where 

possible the following aspects have been considered:

• Feasibility – volumetric and mass constraints, impact on operations and performance, 

inspection and maintenance 

• Scalability – manufacturing methods, joining methods, availability of materials,

repeatability

• Long term performance – fatigue, corrosion, impact on rails

• Cost 

6.1 Enhanced TC128-B 

AM has developed a new low-carbon steel version of TC128-B that they estimate would improve 

puncture energy by approximately 12 – 14 percent over the current version. The new low-carbon 

steel features improved ductility and reduced strength. When similar modifications were made to 

the Actual TC128-B material model developed by Carolan and implemented within validated FE 

models of side impact, puncture energies were predicted to improve by the amount estimated by 

AM. This would show that from a crashworthiness perspective, there is merit to adopting the 

new low-carbon steel in future tank car construction. However, the improvement in puncture 

energy corresponds to a relatively small increase in puncture speed (~1 to 2 mph), and alone it is 

unlikely to significantly change the conditional probability of release. 

From a feasibility, scalability, and cost perspective there is little difference in adopting the new 

steel as it has been designed to meet the requirements of grade TC128-B. 

6.2 High Strength Steel

In Section 4.4.2 a high strength steel was calculated to more than double the puncture energy of 

the tank car if it were to replace the Actual TC128-B material. The high strength steel was 

considered to have a yield stress of 138 ksi (951 MPa) and the same failure strain as the Actual 

TC128-B material (~22 percent for a 2-inch gauge). These results were consistent with similar 

findings by the ATCCRP [41] where similar magnitudes of improvement were calculated across 

a wider range of puncture scenarios when using steels with yield strengths up to 135 ksi (HY130). 

High strength steels, or ultra-high strength steels (UHSS) as they are often called, are available in 

standard sheet thicknesses (3/16 – 4 inches), widths (48 – 120 inches) and lengths (up to 480 

inches) and could be used in tank car manufacture. There is however a wide range of UHSS’s,

and manufacturing methods differ quite significantly (e.g., in terms of welding and heat 

treatment) depending on the type used. Therefore, it is difficult to generically assess the impact 

on cost by transitioning to a UHSS, but it is expected to be somewhat higher than current 

manufacturing costs. 

There would be no direct impact on the rail by adopting this mitigation measure as the mass of 

the tank car would remain broadly the same. 
95 

6.3 High Ductility Steel

In Section 4.4.3 a high ductility steel was investigated and calculated to approximately double 

the puncture energy. The Actual TC128-B material model used in the baseline validated 

simulations had a relatively low ductility, only just meeting the minimum requirement of 22

percent elongation over a 2-inch gauge. The higher ductility version maintained the same 

strength but approximately doubled the elongation to failure. Implementing a high ductility 

material, where the objective is to approximately double the elongation to failure while 

maintaining strength, would require the transition from carbon steel to stainless steel. Types 304 

and 316 stainless steel, for example, have elongations at break in the region of 60 to 70 percent, 

however the yield stress of both materials is only approximately half that of TC128-B. Duplex 

stainless steels have yield strengths closer to that of the Actual TC128-B material considered 

here, but the minimum elongations at break are correspondingly lower as a result (~25 percent).

From a feasibility and scalability perspective, transitioning to stainless steel for the tank shell 

would require significant changes to the manufacturing process. Stainless steel tanks would also 

be more expensive to manufacture and procure, but as evidenced in the oil and gas industry, they 

can provide good long term corrosion resistance in addition to good thermal resistance in the 

event of fire. 

6.4 Thicker Tank Shell

The effect of increasing the tank shell thickness was investigated in Section 4.3, where plate 

thicknesses of 1.25 and 1.50 inches were considered. The shell thickness was shown to strongly 

influence puncture resistance, with the 1.5-inch-thick shell giving an approximate threefold 

increase in puncture energy. However, the findings from the plate thickness versus ductility 

study showed that increasing the plate thickness would reduce the strain to failure in plane strain 

bending. The corresponding influence on puncture resistance has not been quantified, other than 

to note that the findings from the 1.5-inch-thick plate tank car simulations may be overestimated. 

Nevertheless, increasing the shell thickness represents a robust and, in some ways, simple 

solution to improving puncture resistance. From a volumetric and weight perspective, the 

solution satisfies the AAR plate B requirements, with minimal change to the overall dimensions 

of the tank and the overall mass remaining within the 286,000 lb limit. However, there are 

several challenges associated with this solution. The manufacturing process at many facilities 

may need to be adjusted, since rolling 1.5-inch-thick plate to a diameter of 102 inches appears to 

be feasible, specialist plate bending machinery may be required to do so. There would also 

clearly be an increased cost associated with welding. Currently the tank is constructed with two 

elliptically shaped heads and multiple barrel shaped shell sections (i.e., rings) all joined by 

submerged arc welding. If the current welding process could be used with the thicker plate, then 

as a minimum multiple more passes would be required to build up the welded joints. 

There would also be repercussions for rail integrity and overall efficiency of transporting the 

commodity because of the weight increase. The addition of ~29,500 lb to a single tank (i.e., the 

weight increase from moving from a 0.775- to 1.5-inch-thick shell) would increase individual 

wheel loads by ~3.7 ton. This would increase wear on rail and affect useful fatigue life, with 

wheel loading being one of the most influential factors. 
96 

6.5 Multi-Plate Solution

The mitigation solutions discussed in the previous sections require modification to the tank shell 

and are therefore applicable to the design of new tank cars only. However, various multi-plate 

solutions may be applicable to both new and retrofit designs. The solutions investigated in this 

research demonstrated that, for this type of solution, the most important factor to affect puncture 

resistance was the total plate thickness. This is illustrated in Figure 122, where three different 

solutions provide approximately the same level of puncture resistance. 

Figure 122. Different Mitigation Solutions with Total Plate Thickness of 1.5 inches

The multi-plate solutions discussed here refer to the split plate and thicker jacket solutions, with 

several benefits to adopting one of these solutions over the thicker shell solution. One is that the 

mitigation solution can be retrofitted to existing tanks, and another is that the existing tank 

construction methods can be retained with little or no change to the tank shell manufacture 

processes. Further improvement in puncture energies could be obtained by using different 

materials for the outer plate layers (e.g., high strength or high ductility steels). This is a clear 

advantage over the single thick shell solution, but one that was not quantified in this research 

effort. However, the reduction in ductility of thicker steel in bending coupled with the option of 

using high performance steel in the multi-plate solutions indicate that for an equivalent total plate 

thickness the multi-plate solution would provide superior puncture resistance. 

One of the unique findings from this research is that splitting the outer plates into a series of 

unconnected vertical rings in some cases improved puncture resistance. However, the 

improvement was not significantly better than a continuous tank-within-a-tank solution. 

Nevertheless, the results indicate that the level of constraint (i.e., the strength of the connections) 

between adjacent outer plate layers was not consequential for puncture resistance. This provides 

a greater degree of freedom in regard to the manufacturing and joining process, with less 

emphasis placed on ensuring the strength and integrity of joints (i.e., welded connections). 

Another option available with the multi-plate solution is to target the thicker material at the 

locations where there is the highest likelihood of impact (i.e., below the horizontal centerline and 

near the ends of the tank). This is illustrated graphically in Figure 123, which originated from a 

review of the relative frequency of damage to various tank car configurations [45], [46]. 

Targeting material would improve the overall weight efficiency of the tank car, providing further 

cost advantages over the single thicker shell solution. 
97 

Figure 123. Percentage of Damage by Location on Insulated Tank Cars in Mainline 

Accidents [46]

6.6 Liners

Two types of liners were considered in Section 4.6: a liner aimed at providing secondary 

containment in the event the tank shell punctured, and a metal-elastomer bilayer aimed at 

improving the puncture resistance of the tank. The former was shown to be potentially effective 

in situations where the tank shell just punctured (i.e., at impact speeds in the region of 1 – 2 mph 

above the survival speed of the tank). In this scenario, the liner is required to have sufficient 

elongation to accommodate the indenter penetrating some way into the tank. Many tank cars are 

already lined, primarily with the objective of preventing internal corrosion or solvent action 

(specification provided in 49 CFR § 179.201-3 - Lined tanks). It is unlikely that any of these 

lined tanks would have been designed to provide a secondary containment mechanism in the 

event of shell puncture, but they could conceivably be designed to do so. This would require a 

much more specific research effort in collaboration with lining providers to determine the 

potential merits and feasibility of this solution. However, it is a solution that can be applied to 

new or existing tanks, is relatively lightweight, and would not significantly affect the basic 

manufacturing process of tank cars. 

The metal-elastomer bilayer solution aimed at improving puncture resistance of the shell was not 

found to be effective, with an equivalent mass of steel offering better resistance. 

6.7 Blunting Solutions

Blunting was shown in Section 4.7 to provide an effective means of improving puncture 

resistance. However, mitigation solutions developed with the specific aim of blunting the 

indenter are unlikely to provide any improved performance versus either the thicker shell or 

multi-plate solutions. In the side impact case, a 24-inch square indenter provided a similar level 

of puncture resistance (22 mph) to a total plate thickness of 1.5 inches exposed to a 12-inch 

square indenter (23 mph). However, to effectively blunt a 12-inch indenter to act over a 24-inch 

square area would likely require the 4-inch space between the tank shell and jacket to be 

increased and the void filled with a composite material or other novel structure. Additionally, the 

new material or solution would have to provide the same level of thermal performance as the 

current foam insulation. Taking these factors into consideration, it is not recommended to pursue 

solutions that focus on blunting as the primary means of mitigation. 
98 

7. Conclusion

The research team investigated various mitigation solutions aimed at improving the puncture 

resistance of hazmat tank cars by exercising validated FE models subject to side (i.e., shell) 

impact. 

First, the team assessed a new low-carbon version of TC128-B steel developed by AM. The new 

steel featured higher ductility and reduced strength and had been predicted by AM to provide in 

the region of 12 – 14 percent improvement in puncture energy. Using information published by 

AM, researchers developed and evaluated a new pseudo-material model representing the new 

steel. The calculated improvement in puncture energy was very similar to that estimated by AM,

representing an incremental improvement in the puncture resistance of tank cars. 

Within the available tank car space and weight constraints, one of the simplest theoretical 

solutions is to increase the thickness of the tank shell. This applies to the design of new tank cars 

only and is a solution that has been adopted by the tank car industry in generational 

improvements where the head and shell thickness have been increased. The team investigated 

this solution by increasing the tank shell thickness to 1.5 inches and calculating the tank car 

response to side impact. A survival speed of ~22 – 23 mph (an approximate threefold increase in 

energy) was achieved. This utilized ~90 percent of the available weight. 

Recognizing the challenges related to manufacturing a tank car with a 1.5 inch shell, researchers 

conducted a parametric study to investigate the effect of increased ductility and strength. In both 

cases, the puncture energy approximately doubled, with survival speeds increasing to 19 and 20 

mph for the high ductility and high strength materials, respectively. These findings are consistent 

with those in previous research. Implementing a multi-layer (i.e., tank-within-a-tank) solution 

was also considered. The results indicated that in various configurations the survival speed could 

approximately match that of the 1.5-inch-thick shell provided that the total steel thickness was 

the same. The team found that splitting the outer plates into a series of rings did not significantly 

change the puncture resistance, despite separate sub-model analyses indicating otherwise. One 

significant benefit the multi-layer solution offers is the use of increased strength or ductility 

materials for the outer layers. This was not investigated in detail in this research but would likely 

improve the puncture resistance beyond that of a single plate of equivalent thickness. 

Researchers investigated a secondary containment solution in the form of an internal liner, which 

was shown to be potentially effective at impact speeds just above the puncture speed. However, 

it was not effective at improving puncture resistance as a metal-elastomer bilayer. Mitigation 

solutions aimed at blunting the indenter were also considered, with the effectiveness similar to 

that of increasing the shell thickness, but with significant added complexity and cost. Therefore, 

pursuing mitigation solutions aimed primarily at blunting is not recommended. 

Separately, the research team conducted a numerical investigation of the relationship between 

steel ductility and thickness. It was found that increasing steel thickness reduced ductility in the 

plane strain bending condition. This may have consequences for the puncture resistance of 

thicker steel tank cars (i.e., the predicted improvement may not be as high as expected). 

In summary, the most promising mitigation approach to improve tank car puncture resistance is 

likely to be the multi-layer (i.e., tank-within-a-tank) approach, in which different steel grades can 

be used for the outer layers with the specific purpose of enhancing the overall puncture 
99 

resistance of the tank car. These findings match one of the key findings of the Advanced Tank 

Car Collaborative Research Program (ATCCRP), although it is not yet clear what level of 

puncture resistance may be achieved using this approach. The team recommends further research 

to investigate the level of puncture resistance possible using the multi-layer approach. 
100 

8. References

[1] NTSB (2004). Derailment of Canadian Pacific Railway Freight Train 292-16 and 

Subsequent Release of Anhydrous Ammonia Near Minot, North Dakota, January 18,2002

(Railroad Accident Report NTSB/RAR-04/01). 

[2] NTSB (2008). Derailment of Norfolk Southern Railway Company Train 68QB119 with 

Release of Hazardous Materials and Fire, New Brighton, Pennsylvania, October 20, 2006

(Railroad Accident Report NTSB/RAR-08-02). 

[3] NTSB (2009). Derailment of CSX Transportation (CSX) freight train Q380-09, Painesville, 

Ohio, October 10, 2007 (Railroad Accident Report NTSB/RAB-09-02). 

[4] NTSB (2012). Derailment of CN Freight Train U70691-18 With Subsequent Hazardous 

Materials Release and Fire, Cherry Valley, Illinois, June 19, 2009 (Railroad Accident 

Report NTSB/RAR-12-01). 

[5] NTSB (2013). Iowa Interstate Railroad Train Derailment and Subsequent Hazardous 

Materials Release and Fire, Tiskilwa, Iowa, October 7, 2011 (Railroad Accident Report 

NTSB/RAB-13-02). 

[6] NTSB (2014). Norfolk Southern Railway Company Train Derailment and Hazardous 

Materials Release, Columbus, Ohio, July 11, 2012 (Railroad Accident Report NTSB/RAB1408). 

[7] NTSB (2017). Railroad Accident Brief: BNSF Railway Train Derailment and Subsequent 

Train Collision, Release of Hazardous Materials, and Fire, Casselton, North Dakota, 

December 30, 2013 (Railroad Accident Report NTSB/RAB1701). 

[8] TSBC (2014). Runaway and main-track derailment, Lac-Mégantic, Quebec, 06 July 2013

(Railway Investigation Report R13D0054). 

[9] NTSB (2006). Collision of Union Pacific Railroad Train MHOTU-23 with BNSF Railway 

Company Train MEAP-TUL-126-D with Subsequent Derailment and Hazardous Materials 

Release, Macdona, Texas, June 28, 2004 (Railroad Accident Report NTSB/RAR-06/03). 

[10] NTSB (2005). Collision of Norfolk Southern Freight Train 192 with Standing Norfolk 

Southern Local Train P22 with Subsequent Hazardous Materials Release at Graniteville, 

South Carolina, January 6, 2005 (Railroad Accident Report NTSB/RAR-05/04). 

[11] Pipeline and Hazardous Materials Safety Administration (2006). Hazardous Materials: 

Improving the Safety of Railroad Tank Car Transportation of Hazardous Materials. Notice 

of Public Meeting, Federal Register, Vol. 71, No. 222. 

[12] US Department of Transportation (2015). Rule Summary: Enhanced Tank Car Standards 

and Operational Controls for High-Hazard Flammable Trains. 
101 

[13] Code of Federal Regulations (2015). Title 49 – Transportation, Appendix A to Part 179 –

Procedures for Tank-head Puncture-Resistance Test, Pipeline and Hazardous Materials 

Safety Administration. 

[14] Jeong, D. Y. (2009). Probabilistic Approach to Conditional Probability of Release of 

Hazardous Materials from Railroad Tank Cars During Accidents (Report No. IMECE2009-

10872). FRA. 

[15] Jeong, D. Y. (2016). Overview of Full-scale Tank Car Shell Impact Testing and Analysis. 

(Volpe National Transportation Systems Center Report). 

[16] Tyrell, D. C., Jeong, D. Y., Jacobsen, K., & Martinez, E. (2007). Improved Tank Car 

Safety Research (Report No. RTDF2007-46013). FRA.

[17] Tyrell, D. C., Jacobsen, K,. Talamini, B., & Carolan, M. (2007). Developing Strategies for 

Maintaining Tank Car Integrity During Train Accidents (Report No. RTDF2007-46015). 

FRA. 

[18] Carolan, M., Talamini, B., & Tyrell, D. (2008). Update on Ongoing Tank Car 

Crashworthiness Research: Predicted Performance and Fabrication Approach (Report No. 

JRC2008-63052). FRA. 

[19] Carolan, M. E., Jeong, D. Y., Perlman, A. B., Tang, Y. H., & Gonzalez, F. (2011).

Deformation Behavior of Welded Steel Sandwich Panels under Quasi-static Loading

(Report No. JRC2011-56054). FRA. 

[20] Jeong, D. Y., Tyrell, D. C., Carolan, M. E., & Perlman, A. B. (2009). Improved Tank Car 

Development: Ongoing Studies on Sandwich Structures (Report No. JRC2009-63025). 

FRA. 

[21] Jeong, D. Y., Carolan, M. E., Perlman, A. B., & Tang, Y. H. (2011). Deformation Behavior 

of Welded Steel Sandwich Panels under Quasi-static Loading (Report No. 

DOT/FRA/ORD-11/06). FRA. 

[22] Carolan, M. E., Jeong, D. Y., Perlman, A. B., Murty, Y. V., Namboordri, S., Kurtz, R., 

Elzey, R. K., Ananikitpaiboon, S., Tunna, L., & Fries, R. (2013). Application of Welded 

Steel Sandwich Panels for Tank Car Shell Impact Protection (Report No. DOT/FRA/ORD13/19). FRA.

[23] Tang, Y. H., Yu, H., Gordon, J. E., Jeong, D. Y., & Perlman, A. B. (2008). Analysis of 

Railroad Tank Car Shell Impacts Using Finite Element Method (Report No. JRC2008-

63014). FRA.

[24] Yu, H., Jeong, D. Y., Gordon, J. E., & Tang, Y. H. (2007). Analysis of Impact Energy to 

Fracture Un-Notched Charpy Specimens Made from Railroad Tank Car Steel (Report No. 

RTDF2007-46038). FRA. 

[25] Bao, Y. & Wierzbicki, T. (2005). On the cut-off value of negative triaxiality for fracture.

Engineering Fracture Mechanics, 72(7), 1049-1069. 
102 

[26] Nahshon, K. & Hutchinson, J. W. (2008). Modification of the Gurson Model for shear 

failure. European Journal of Mechanics A/Solids, 27(1), 1-17. 

[27] Gonzalez, F. (2021). FRA Hazmat Tank Car Impact Test Research. Railway Age.

[28] Roman, J. (2020). Learning from Crude. NFPA Journal. 

[29] PHMSA (2020). Federal Register / Vol. 85, No. 143/ Friday, July 24, 2020 / Rules and 

Regulations Hazardous Materials: Liquefied Natural Gas by Rail. 

[30] Carolan M., & Przemyslaw, R. (2019). Side Impact Test and Analyses of a DOT-105 Tank 

Car (Report No. DOT/FRA/ORD-19/12). FRA.

[31] Wilson, N., Eshraghi, S., Trevithick, S., Carolan, M., & Rakoczy, P. (2020). Side Impact 

Test and Analyses of a DOT-105 Tank Car (Report No. DOT/FRA/ORD-20/38). FRA.

[32] Elkins, A. (2017). Field Guide to Tank Cars. Third Edition. Association of American 

Railroads. 

[33] Carolan, M., Eshraghi, S., Wilson, N., Trevithick S., & Przemyslaw, R. (2020). Side 

Impact Test and Analyses of a DOT-105 Tank Car – 6 X 6 Inch Indenter (Report No. 

DOT/FRA/ORD-20/38). FRA. 

[34] Tyrell, D. (2007). Developing Strategies for Maintaining Tank Car Integrity During Train 

Accidents (Report No. RTDF2007-46015). FRA. 

[35] McKeighan, P. C., Jeong, D. Y., Cardinal, J. W. (2008). Mechanical Properties Of Tank 

Car Steels Retired From The Fleet (Report No. JRC2009-63060). FRA.

[36] Abotula, S., Konda, B., Walsh, W. J., Challa, V., Ros, T., & Murali, M. (2018). Analysis of 

Puncture Resistant Steel for Tank Cars Steel Manufacturing Simulation and Visualization, 

90(4).

[37] Railway Supply Institute (n.d.). Tank Car Standards in North America. 

[38] Anderson, T. L., Rose, B., McKeighan, P. C., & Kirkpatrick, S. W. (2007). Quantifying 

and Enhancing Puncture Resistance in Railroad Tank Cars Carrying Hazardous Materials, 

Phase II: Development and Validation of a Puncture Resistance Evaluation Methodology. 

Quest Reliability.

[39] Kirkpatrick, S. W. (2010). Detailed Puncture Analyses of Various Tank Car Designs, Final 

Report - Revision 1. Applied Research Associates.

[40] Pipeline and Hazardous Materials Safety Administration (2017). Petition for Rulemaking 

to Phaseout the Use of Tank Cars Not Meeting the HM-246 Specification to Transport 

Toxic-by Inhalation Materials (Docket No. PHMSA-2016-0165). 

[41] Advanced Tank Car Collaborative Research Program Executive Committee (2016). 

ATCCRP Executive Summary and Conclusions. 
103 

[42] Xue, Z., & Hutchinson, J. W. (2005). Neck retardation and enhanced energy absorption in 

metal-elastomer bilayers. Mechanics of Materials, 39(5), 473-487. 

[43] Anderson, R. G., Londono, G. J., Woelke, B. P., & Nielson, K. L. (2020). Fundamental 

differences between plane strain bending and far-field plane strain tension in ductile plate 

failure. Journal of the Mechanics and Physics of Solids, 141. 

[44] Dykeman, J. (2019). Evolving test methods for fracture characterization and prediction.

EWI Workshop 2019. 

[45] Jeong, D. Y., Carolan, M. E., Perlman, A. B., & Tang, Y. H. (2011). Deformation Behavior 

of Welded Steel Sandwich Panels Under Quasi-Static Loading (Report No. JRC2011-

56054). FRA. 

[46] Treichal, T. T. (2007). Specific Locations of Damage and Lading Losses on Tank Heads 

and Shells (Report No. RA 06-04). RSI-AAR Railroad Tank Car Safety Research and Test 

Project

[47] Kirkpatrick, S.W. Ph.D (2010). Detailed Puncture Analyses of Various Tank Car Designs.

Applied Research Associates. 

[48] ASTM International (2008). Standard Test Methods for Tension Testing of Metallic 

Materials (E8/E8M - 16a). 
104 

Appendix A.

Material Properties

This section contains additional information about specific material properties or parameters that 

were used in development of the FE models. 

A1 – Pneumatic Cavity / Air Properties

Table 37 and Table 38 contain information on the material properties used to define the air cavity 

in the full-scale and quarter-symmetry FE models. These properties were obtained from Carolan 

& Przemyslaw [30]. 

Table 37. Air Properties

Parameter Value Units

Universal Gas Constant (R) 73.583 in*lbf / mol*K

Molecular Weight (MW) 1.654 x 10-4 lbf*s2 / mol

Table 38. Molar Specific Heat for Air

Temperature

(K)

Cpmolar 

(in*lbf / mol*K)

250 257.2

300 257.2

350 258.5

400 259.7

A2 – A1011 Jacket Material Properties 

Carolan [30] developed and provided material property data to define the plasticity and damage 

behavior of A1011 based on available tensile test data. The true stress/true plastic strain data 

provided and required by Abaqus is tabulated in Table 39 and shown in Figure 124. 

Table 39. True Stress/True Plastic Stain Inputs for A1011 [30]

True Stress

(psi)

Plastic Strain

(in/in)

47,000 0 

46,000 8.22E-04

46,200 1.20E-02

53,000 3.00E-02

58,000 6.00E-02

62,000 1.10E-01

68,000 1.95E-01

125,000 1.15
105 

Figure 124. A1011 True Stress True Plastic Strain Curve [30]

Carolan developed a B-W failure envelope for A1011 using fully integrated shell elements (S4)

with a fine mesh size of 0.04 inches and demonstrated good agreement between a simulated 

tensile test and the experimental data. In this research effort, the initial simulation was first 

replicated, and then new simulations were undertaken to develop a new B-W curve that could be 

used with larger and therefore more computationally efficient elements. 

The original B-W curve developed by Carolan is shown in Figure 125. When used by the 

research team with S4 elements and a 0.04 inch mesh it resulted in the engineering stress-strain 

curve shown in Figure 126. This compared well to that originally produce by Carolan (Figure 

127). 

In a subsequent iterative process, researchers ran a series of new simulations in which the mesh 

size of the ‘dog bone’ tensile test piece was increased to 0.1196 inches (to provide an aspect ratio 

of 1:1) and the plasticity data points of the B-W curve were scaled by different amounts. The 

element type was also changed from full-integration (S4) to reduced integration (S4R) elements, 

again for computational efficiency. The damage progression energy was not changed and 

remained 1500 in-lbf/in2

. It was determined from the analyses that scaling the B-W curve by a 

factor of 0.9 gave good agreement with the results obtained using the parameters provided by 

Carolan. The resulting engineering stress-strain curves are compared in Figure 128. 

The computational efficiency achieved by using larger elements and switching element type was 

approximately 9-fold. This increase in efficiency was important in the development of an 

efficient sub-model that could be used to assess multiple different mitigation solutions. 

 
106 

Figure 125. B-W Failure Envelope [30]

Figure 126. Engineering Stress Strain Curve Produced by Researchers
107 

Figure 127. Engineering Stress Strain Curve Produced by Carolan [30]

Figure 128. Comparison of Engineering Stress Strain Curves Produced using Carolan 

Parameters (0.04 inch & S4 elements) vs Research Parameters (0.1196 inch, S4R and scaled 

B-W curve)
108 

Figure 129. Plastic Strain Plot of Tensile Tests using Calibrate Material Properties for 

A1011 Material

Figure 129 above shows an overview of the ‘dog bone’ specimen used in the tensile test 

simulations. The plot shows plastic strain as the specimen undergoes necking and failure. 
109 

Appendix B.

TC128-B Material Properties

This section describes the development and verification of the Actual TC128-B material model 

used within the solid element regions in both the full-scale models and applicable sub-models. 

In Appendix F4 of [30], Carolan describes the characterization procedure of three different 

variations of TC128-B material. Two of the variations, referred to as ‘New Normalized’ and ‘Test 

2,’ were developed prior to the full-scale impact test (i.e., Test 6). These two materials were used 

in pre-test FE models to predict the puncture speed and therefore calculate the target ram car 

speed for Test 6. Following the completion of Test 6, a third material referred to as ‘Actual Test 

Material’ was characterized using tensile test data obtained from material cut directly from the 

tank shell. A summary is provided in Table 40 below. 

Table 40. TC128-B Material Models

TC128B Model Performed (Original) Source

Test 2 Pre-Test Kirkpatrick [47]

New Normalized Pre-Test Kirkpatrick [47]

Actual Test Material Post Test 6 Carolan [30]

Carolan provided hardening curves, fracture loci (i.e., B-W Damage Models), and damage 

evolution parameters for each of the three materials, along with information on the element type 

and size for which the material models were calibrated. Note that the B-W damage models and 

associated damage evolution parameters are mesh dependent. The hardening curves and B-W 

damage models for each material are shown in Figure 130 and Figure 131, respectively. 

Figure 130. TC128-B True Stress/True Plastic Strain Curves [30]
110 

Figure 131. TC128B Fracture Loci [30]

A summary of the element formulations and damage evolution parameters used for each material 

model is provided in Table 41. 

Table 41. TC128-B Tensile Coupon Model Characteristics

TC128B Model Mesh & Elements Mesh Size in Neck Damage Progression

Test 2 Bricks – C3D8I 0.085 in 700 in-lbf/in2

Exponential

New Normalized Bricks – C3D8I 0.085 in 700 in-lbf/in2

Exponential

Actual Test Material Bricks – C3D8I 0.085 in 1,400 in-lbf/in2

Linear

To verify the Actual material model for use in the full-scale tank car analyses (and subsequent 

sub-model analyses), the research team conducted a series of Round Bar Tensile test simulations 

using geometry developed in accordance with ASTM E8 [48] (i.e., 2-inch gauge). Sample 

images from the analyses are shown in Figure 132 and Figure 133. 

Figure 132. Necking of Round Bar During Tensile Test
111 

Figure 133. Round Bar Tensile Tests Showing Plastic Strain Contours

The resulting engineering stress-strain curves computed from the simulated tests are shown in

Figure 134. They compared very well to the experimentally obtained curves, as shown in Figure 

135 and Figure 136, in which the calculated engineering stress-strain curves have been overlaid 

onto the experimentally obtained versions. 

Figure 134. Engineering Stress Strain Curves from Round Bar Simulations
112 

Figure 135. Engineering Stress Strain Curves from Researched Round Bar Simulations

Overlaid onto Original Curves (Test 2 and New Normalized, dotted lines) 

Figure 136. Engineering Stress Strain Curve from Researched Round Bar Simulation (solid 

black line) Overlaid onto Original Curve (Actual)

In addition to validating the use of the material model parameters developed by Carolan, the 

results illustrate the variability in material properties of TC128-B. The material specifications for 

TC128-B are summarized in Table 42. Each of the three versions characterized meet these

standards, although the Actual material from Test 6 only just met the required minimum ductility 

limit of 22 percent elongation. It is also notable that there is an upper limit specification for the 

Ultimate Tensile Strength (UTS). 

Table 42. Requirements of TC128-B Material

Property Value

Min. Yield Strength 50 ksi (345 MPa)

UTS Range 81 – 101 ksi (558 – 696 MPa)

Elongation at Failure (2-in gauge) 22%

Elongation at Failure (8-in gauge) 16%
113 

Appendix C.

Full-Scale Model, Potentiometer Results

This appendix contains additional output from the full-scale simulation described in Section 2.3

to validate the FE model against test data. 

Figure 137 compares the change in length in the FE model versus Test 6 for the string 

potentiometer located at the center of the tank (i.e., directly behind the impact zone) and oriented 

in the direction of travel of the indenter. 

Figure 137. Comparison of String Potentiometer Length vs Time, Centre of Tank

Figure 138 compares the change in length in the FE model versus Test 6 for the string 

potentiometer located 24 inches from the center of the tank. Figure 139 compares the same for 

the string potentiometer located 48 inches from the center of the tank and Figure 140 for the skid. 

Figure 141 compares the vertical response and Figure 142 the response at the head of the tank. 

In general, the response of the FE model compares well against the physical test data, 

particularly until puncture occurs at approximately 220 ms. 
114 

Figure 138. Comparison of String Potentiometer Length vs Time, 24-inch Offset from 

Center of Tank

Figure 139. Comparison of String Potentiometer Length vs Time, 48 inch Offset from 

Center of Tank
115 

Figure 140. Comparison of String Potentiometer Length vs Time, Skid

Figure 141. Comparison of String Potentiometer Length vs Time, Vertical
116 

Figure 142. Comparison of String Potentiometer Length vs Time, Head
117 

Appendix D.

Polyurea Material Model

The engineering stress-strain data used to develop the Abaqus material model for the polyurea 

used in the internal liner assessment is provided in Table 43 and shown in Figure 143. A 

simplified nonlinear elastic-plastic material model was used to represent the polyurea, where it is 

assumed that the plastic strain component is non-recoverable. This was considered a reasonable 

simplification since the primary concern with respect to the integrity of the liner under impact 

loading is during the initial loading phase and not during recovery. To include plasticity within 

Abaqus, the stress-strain points past yield must be input in the form of true stress and logarithmic 

plastic strain. The logarithmic plastic strain required by Abaqus can be calculated with the 

equation below: 

The first data point must always correspond to the yield stress (in this case 0.2 ksi, with 

logarithmic plastic strain = 0) with subsequent stress-strains calculated using the relationships 

above. Implementing this approach rendered the true stress – logarithmic plastic strain data 

shown in Figure 144. This is the data used within the Abaqus material model. Other relevant 

material data is provided in Table 44. Ductile damage behavior was defined to facilitate element 

deletion. For simplicity, this was assumed to be independent of stress-triaxiality, and a very 

small displacement to failure upon reaching the fracture strain was assigned to enforce a brittle 

response. 

A single element tensile test was performed on a 0.125-inch-thick by 0.5-inch-wide by 2-inchlong, 4-node fully integrated shell element (S4) to verify the performance of the material model. 

The initial and deformed shape of the element just prior to failure is shown in Figure 145, with 

the engineering stress-strain curve extracted from the analysis shown in Figure 146. Very good 

agreement was obtained between the Abaqus single element test and the experimental stressstrain curve. 

Table 43. Polyurea Input Engineering Stress-strain Data

Engineering Strain (-) Engineering Stress (ksi)

0.000 0.000

0.007 0.200

0.658 2.742

1.400 2.743

1.600 2.800

1.950 2.900

2.250 3.000

2.650 3.150

2.950 3.280

3.250 3.400

3.630 3.520

4.200 3.750

4.900 4.000
118 

5.160 3.880

Figure 143. Polyurea, Input Engineering Stress-strain Curve

Figure 144. Polyurea, Abaqus True Stress – Logarithmic Plastic Strain Curve

Table 44. Polyurea Material Properties Used in FE Model

Property Value

Modulus of Elasticity 30,000 psi

Yield Stress 200 psi

Poisson’s Ratio 0.49

Density 1.25 lbf*s2/in

Logarithmic plastic strain to failure 1.024

Damage evolution (linear displacement) 1E-05 in
119 

Figure 145. Displacement Response of Polyurea Single Element Test

Figure 146. Engineering Stress-Strain Curve Obtained from Abaqus Polyurea Single 

Element Test (black line) Compared to Experimental Data and NLFlex Single Element 

Test Results
120 

Abbreviations and Acronyms

ACRONYM DEFINITION

AAR Association of American Railroads

AM ArcelorMittal

ANSI American National Standards Institute

ARA Applied Research Associates, Inc.

ASTM American Society for Testing and Materials

ATCCRP Advanced Tank Car Collaborative Research Program

B-W Bai-Wierzbicki

DENT Double-Edge Notch Tension

DOT Department of Transportation

FE Finite Element

FEA Finite Element Analysis

FRA Federal Railroad Administration

GTN Gurson Tvergaard Neddleman

HSLA High Strength-Low Alloy

KE Kinetic Energy

LNG Liquified Natural Gas

MOC Memorandum of Cooperation

MPC Multi Point Constraint

MPC multi-point constraint

NGRTCP Next Generation Rail Tank Car Program

NPRM Notice of Proposed Rulemaking

NTSB National Transportation Safety Board

PHMSA Pipeline and Hazardous Materials Safety Administration

PIH Poison Inhalation Hazardous

SPH Smooth Particle Hydrodynamics

SWRI Southwest Research Institute

TC Transport Canada

TIH Toxic Inhalation Hazmat

TT Thornton Tomasetti, Inc.
121 

ACRONYM DEFINITION

TTCI Transportation Technology Center, Inc.

UHSS Ultra-High Strength Steels

UP Union Pacific

USDOT U.S. Department of Transportation

UTLX Union Tank Car Company