DOT/FRA/ORD-24-01 Final Report | January 2024 

Improving Thermal Protection of Cryogenic Tank Cars 

Through Testing, Analysis, and Evaluation of Pressure 

Relief Valve System Performance

i 

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1. REPORT DATE (DD-MM-YYYY) 

27-09-2023 

2. REPORT TYPE 

Technical Report 

3. DATES COVERED (From - To) 

8/2022-8/2023 

4. TITLE AND SUBTITLE 

Improving Thermal Protection of Cryogenic Tank Cars Through Testing, Analysis, and 

Evaluation of Pressure Relief Valve System Performance 

5a. CONTRACT NUMBER 

693JJ633C000012 

5b. GRANT NUMBER 

5c. PROGRAM ELEMENT NUMBER 

6. AUTHOR(S) 

Garrett Mattos: ORCID 0000-0003-4682-7110 

Keith Friedman ORCID 0000-0002-3133-4113

Matthew Gacek: ORCID 0009-0009-9882-5791

Jason Huczek: ORCID 0000-0002-2656-853X

5d. PROJECT NUMBER 

5e. TASK NUMBER 

5f. WORK UNIT NUMBER 

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

Friedman Research Corporation Southwest Research Institute (SwRI) 

6500 River Place Blvd, Bldg 7, Ste 250 6220 Culebra Road, Bldg. 143 

Austin, TX 78730 San Antonio, TX 78238

8. PERFORMING ORGANIZATION 

REPORT NUMBER 

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-24-01 

12. DISTRIBUTION/AVAILABILITY STATEMENT 

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

13. SUPPLEMENTARY NOTES 

COR: Francisco Gonzalez, III 

14. ABSTRACT 

The Federal Railroad Administration (FRA) sponsored a research team to conduct high fidelity finite element (FE) modeling to evaluate the effects of postderailment motion on PRV system damage and the performance of structural countermeasures. The team conducted physical tests to measure the 

response of a typical PRV under cryogenic liquid flow as well as the thermal properties of an ignited PRV exhaust plume. Researchers then conducted 

additional FE modeling to apply the experimentally determined PRV exhaust (i.e., jet) fires to an ISO UN-T75 tank and quantify the tank’s response in 

multiple fire-exposed scenarios.

15. SUBJECT TERMS 

Cryogenic fluid, dual phase flow, simulation, fire, LNG 

16. SECURITY CLASSIFICATION OF: 17. LIMITATION OF 

ABSTRACT 

18. NUMBER 

OF 

PAGES

69 

19a. NAME OF RESPONSIBLE PERSON 

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

19b. TELEPHONE NUMBER (Include area code) 

202-689-4316

Standard Form 298 (Rev. 8/98) 

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ii

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

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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

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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) =

=

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1.1 short tons

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

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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

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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 Updated 6/17/98
iii

Acknowledgements

The authors thank Taylor Wharton for providing insight and feedback related to PRV system 

design and performance during this project.
iv

Contents

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

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

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

1.2 Objectives .................................................................................................................... 4

1.3 Overall Approach ........................................................................................................ 4

1.4 Scope ........................................................................................................................... 4

1.5 Organization of the Report .......................................................................................... 4

2. Literature Review of PRV Characteristics .................................................................. 6

2.1 Key PRV Requirements .............................................................................................. 6

2.2 Comprehensive PRV Requirements ............................................................................ 7

3. FE Simulation of Crash Conditions ........................................................................... 10

3.1 FE Model Description ............................................................................................... 10

3.2 Methods ..................................................................................................................... 11

3.3 FE Results .................................................................................................................. 13

3.4 FE Crash Discussion ................................................................................................. 17

4. PRV System Test ....................................................................................................... 18

4.1 Test Methods ............................................................................................................. 18

4.2 Test Results ............................................................................................................... 22

5. PRV Exhaust Flame Test........................................................................................... 30

5.1 Test Methods ............................................................................................................. 30

6. Structural Countermeasures....................................................................................... 42

6.1 Methodology ............................................................................................................. 42

6.2 Results ....................................................................................................................... 43

7. PRV Exhaust Flame Effects ...................................................................................... 46

7.1 Methodology ............................................................................................................. 46

7.2 Results ....................................................................................................................... 48

7.3 Discussion ................................................................................................................. 50

8. Suggested Future Work ............................................................................................. 52

9. Conclusion ................................................................................................................. 53

10. References ................................................................................................................. 55

Appendix A ................................................................................................................................... 56

Appendix B ................................................................................................................................... 57

Abbreviations and Acronyms ....................................................................................................... 59
v 

Illustrations

Figure 1. Jet Fire of PRV exhaust during fire test of LNG filled UN-T75 tank (FRA, 2023) ....... 3

Figure 2. Modified FE model of UN-T75 tank with added PRV piping assembly shown with 

surrounding environment (left) and with isolated piping assembly (right) [LS-DYNA] ..... 10

Figure 3. Physical and FE model of UN-T75 tank positioned on flatbed railcar ......................... 10

Figure 4. Initial position of vehicle combination including rails, ballast, and surrounding terrain; 

FE Model (top) and engineering specifications (bottom) ..................................................... 11

Figure 5. View of model showing LNG within the tank (shown graphically as ‘dots’) at initiation 

of rollover.............................................................................................................................. 11

Figure 6 Pipe damage in rear-end flatcar impacts at 15 mph, 20 mph, 30 mph, and 49 mph (from 

left to right) ........................................................................................................................... 14

Figure 7. Damage to inner tank and internal PRV piping in the Top-quarter – Pole impact at 15 

mph; damage at higher speeds was similar in shape but greater in magnitude .................... 14

Figure 8. Original (left) and reduced pipe cross-section (right) after 25 mph Top-quarter – Pole 

impact .................................................................................................................................... 15

Figure 9. Damage to inner tank and internal PRV piping in the Top-end – Pole impact at 10 mph; 

damage at higher speeds was similar in shape but greater in magnitude.............................. 15

Figure 10. Velocity vs Time for 3 pipe damage severities ........................................................... 15

Figure 11. Methane gas flow through nominal pipe at 120 psi inlet pressure .............................. 16

Figure 12. Methane gas flow through ~50% closed pipe at 120 psi inlet pressure ...................... 16

Figure 13. Methane gas flow through ~80% closed pipe at 120 psi inlet pressure ...................... 16

Figure 14. Methane gas flow through >90% closed pipe at 120 psi inlet pressure ...................... 17

Figure 15. PRV test section arrangement ..................................................................................... 19

Figure 16. PRV test system ........................................................................................................... 19

Figure 17. PRV test system longer discharge ............................................................................... 20

Figure 18. Machined elliptical orifices used for PRV discharge constriction .............................. 21

Figure 19. Baseline test at 130 psi – Run 1 (PRV temperature [°F], PRV pressure [psi], mass 

flow rate [lb/min], and density [kg/m3]) ............................................................................... 22

Figure 20. Baseline test at 130 psi – Run 4 (PRV temperature [°F], PRV pressure [psi], mass 

flow rate [lb/min], and density [kg/m3

]) ............................................................................... 23

Figure 21. Summary of baseline tests – Flow rate vs. tank pressure ............................................ 23

Figure 22. Summary of baseline tests – Fluid density vs. tank pressure ...................................... 24

Figure 23. Comparison of unrestricted vs restricted downstream conditions; note that the 130- 

and 175-psi discharges were conducted with the supply tank half full, and the 190-psi 

discharges were conducted immediately after the supply tank was filled ............................ 24
vi 

Figure 24. 89% Constriction at 130 psi – Run 4 (PRV temperature [°F], PRV pressure [psi], mass 

flow rate [lb/min], and density [kg/m3

]) ............................................................................... 25

Figure 25. Hot test at 240 psi – Run 1 (PRV temperature [°F], PRV pressure [psi], mass flow rate 

[lb/min], and density [kg/m3])............................................................................................... 26

Figure 26. Hot test at PTFE melt temp (PRV temperature [°F], PRV pressure [psi], mass flow 

rate [lb/min], and density [kg/m3])........................................................................................ 26

Figure 27. PRVs after hot testing; temp < melt (left), temp = melt (middle), temp >> melt (right)

............................................................................................................................................... 27

Figure 28. Humid test at 150 psi – Run 4 (PRV temperature [°F], PRV pressure [psi], mass flow 

rate [lb/min], and density [kg/m3])........................................................................................ 27

Figure 29. Summary of mass flow rates with extended exhaust added ........................................ 28

Figure 30. Summary of mass flow rates with extended constricted exhaust added ..................... 29

Figure 31. Large-scale fire engineering and research facility layout ............................................ 30

Figure 32. Schematic of natural gas delivery piping setup ........................................................... 31

Figure 33. Schematic of instrumentation stations for horizontal jet fire tests .............................. 31

Figure 34. Schematic of instrumentation stations for vertical jet fire tests .................................. 32

Figure 35. Natural gas supply into test facility, main cutoff valve, check valve and pipeline flame 

arrestor (left); Transition to 1.5-in diameter pipe and pressure flow rate measurement to data 

acquisition system (right) ...................................................................................................... 32

Figure 36. Pre-test conditions after sand burner is ignited (left); Test in progress with jet flame 

and north measurement state at mid-flame position (right) .................................................. 33

Figure 37. Heat flux time history for Test 15 (3000 SLPM, 4-ft standoff distance and flame 

middle location) .................................................................................................................... 36

Figure 38. Temperature time history for Test 15 (3000 SLPM, 4-ft standoff distance and flame 

middle location) .................................................................................................................... 37

Figure 39. Heat flux time history for Test 22 (3000 SLPM, 8 -ft standoff distance and flame 

middle location) .................................................................................................................... 37

Figure 40. Temperature time history for Test 22 (3000 SLPM, 8 -ft standoff distance and flame 

middle location) .................................................................................................................... 38

Figure 41. Heat flux trend for 4-ft standoff distance and flame tip location ................................ 39

Figure 42. Heat flux trend for 4-ft standoff distance and flame middle location ......................... 39

Figure 43. Heat flux trend for 8-ft standoff distance and flame tip location ................................ 40

Figure 44. Heat flux trend for 8-ft standoff distance and flame middle location ......................... 40

Figure 45. Peak heat flux comparative chart ................................................................................ 41

Figure 46. “Crash Cushion” Countermeasure ............................................................................... 42
vii 

Figure 47. Maximum damage in 15 mph rear impact for baseline tank (top) and countermeasure 

(bottom)................................................................................................................................. 43

Figure 48. Maximum damage in 20 mph rear impact for baseline tank (top) and countermeasure 

(bottom)................................................................................................................................. 44

Figure 49. External PRV Piping Assembly after 30 mph impact with empty flatcar: Baseline Sch 

40 pipe (left); 0.25” thick pipe (center); 0.5” thick pipe (right) ........................................... 45

Figure 50. Maximum deformation of 0.25” (left) and 0.5” (right) thick external PRV pipes ...... 45

Figure 51. Tank and Flat Car Model within the FDS environment .............................................. 46

Figure 52. Example visualizations of the FAR jet fire position in the SIDE (left) and END (right) 

configurations without a pool fire ......................................................................................... 47

Figure 53. Summary of heating rates for the simulated fire conditions ........................................ 48

Figure 54. Incident heat flux on tank exterior in the HIGH pool fire only configuration ............ 49

Figure 55. Incident heat flux on tank exterior in the HIGH pool fire + HIGH jet fire END 

configuration ......................................................................................................................... 49

Figure 56. Molar fraction of oxygen in a LOW, END, CLOSE, Jet configuration (no pool fire) 50

Figure 57. Molar fraction of oxygen in a LOW, END, CLOSE, Jet configuration (LOW pool 

fire) ........................................................................................................................................ 50
viii

Tables

Table 1. Potential effects of a derailment and the associated regulation ........................................ 7

Table 2. FE Crash Simulation Matrix ........................................................................................... 12

Table 3. FE Crash Results ............................................................................................................. 13

Table 4. PRV Flow Test Matrix .................................................................................................... 21

Table 5. Summary of jet fire test matrix ....................................................................................... 33

Table 6. Summary of average test results for north measurement station .................................... 34

Table 7. Summary of average test results for west measurement station ..................................... 35

Table 8. Fire Impingement Simulation Matrix ............................................................................. 47
1 

Executive Summary

In a train derailment, tank cars have a high probability of full or partial inversion of tanks and 

cars. An inversion of a cryogenic tank, such as an ISO UN-T75portable tank, allows liquid or 

dual-phase fluid to be present at the inlet of the pressure relief valve(s) (PRV). However, PRVs 

installed on many cryogenic tanks are not specified for use in liquid or dual-phase flow 

conditions. 

The Federal Railroad Administration (FRA) sponsored a research team to conduct high fidelity 

finite element (FE) modeling to evaluate the effects of post-derailment motion on PRV system 

damage and the performance of structural countermeasures. The team conducted physical tests to 

measure the response of a typical PRV under cryogenic liquid flow as well as the thermal 

properties of an ignited PRV exhaust plume. Researchers then conducted additional FE modeling 

to apply the experimentally determined PRV exhaust (i.e., jet) fires to an ISO UN-T75 tank and 

quantify the tank’s response in multiple fire-exposed scenarios. 

This project builds on previous work, including two full-scale fire tests of ISO UN-T75 tanks 

filled with cryogenic fluid (i.e., Liquefied Nitrogen and Liquefied Natural Gas) and the 

development of validated multi-physics models capable of simulating the structural and thermal 

response of the tank subjected to a pool fire (FRA, 2023). 

Key results from the project include: 

• A discussion of the derailment mechanisms that are likely to result in the PRV assembly 

failing to maintain its performance requirements as defined in the regulations 

• Identification of crash modes and crash severities that are likely to damage the PRV 

piping assembly in a way that would decrease its performance 

• Characterization of PRV response across a range of tank pressures, ambient conditions, 

downstream constrictions, and localized heating

• Characterization of low flow rate jet fire thermal response across a range of mass flow 

rates, orifice sizes, and jet orientations

• Repeatable testing methods

• Demonstration of countermeasures that significantly improve the crashworthiness of the 

PRV piping assembly 

The PRV flow characteristics demonstrated in this project did not negatively affect the 

fireworthiness of cryogenic tanks in the manner researchers anticipated (i.e., due to degraded 

flow), but they did present potential challenges related to the effects on neighboring tanks, 

potentially hindering emergency response, and increasing the rate of content loss. Since the PRV 

exhaust rates under liquid flow are far greater than under vapor/gas flow, a tank would 

completely empty its contents within 60 to 125 minutes of PRV activation in a pool fire scenario. 

While this may reduce the likelihood of a Boiling Liquid Evaporating Vapor Explosion 

(BLEVE) from occurring, it increases the risk of a vapor cloud and the environmental risk due to 

content loss. Additionally, the jet fire that would likely be produced from the discharged 

flammable fluid could impinge on neighboring tanks up to 100 ft away and apply sufficient 

heating to endanger their contents. Emergency responders need to be aware of the higher 
2 

potential for vapor cloud generation due to the greater mass flow rates as well as the danger of 

cryogenic dual-phase PRV exhaust and substantial jet fire potential. 
3 

1. Introduction

The Federal Railroad Administration (FRA) sponsored a research team to conduct high fidelity 

finite element (FE) modeling to evaluate the effects of post-derailment motion on pressure relief 

valve (PRV) system damage and the performance of structural countermeasures. The team 

conducted physical tests to measure the response of a typical PRV under cryogenic liquid flow as 

well as the thermal properties of an ignited PRV exhaust plume. Researchers then conducted 

additional FE modeling to apply the experimentally determined PRV exhaust (i.e., jet) fires to an 

ISO UN-T75 tank and quantify the tank’s response in multiple fire-exposed scenarios. 

The Pipeline and Hazardous Materials Safety Administration (PHMSA) issued a final ruling on 

July 24, 2020 (effective on August 24, 2020) that allowed liquefied natural gas (LNG) to be 

transported by rail in approved tank cars (i.e., enhanced DOT-113 tank cars). In prior research, 

the team worked for several years with FRA and Transport Canada to conduct physical tests and 

create computer models of fire-involved UN-T75 portable tanks to quantify their performance 

and risk of failure. Previous findings related to cryogenic tank response during derailment, tank 

inversion, and fire exposure prompted this follow-on work to investigate the performance of the 

PRV system. This report describes the results of physical testing and physics-based numerical 

modeling used to quantify the performance of the PRV system under derailment conditions and 

how that performance might lead to potential cascading effects. 

1.1 Background

PRVs in cryogenic intermodal tanks and tank cars are necessary to prevent the buildup of interior 

tank pressure when exposed to fire conditions that can lead to catastrophic failure such as a 

Boiling Liquid Expanding Vapor Explosion (BLEVE). The likelihood of tank failure under fire 

conditions is directly related to the ability of the PRV system to perform properly. Crash 

conditions, final tank position (i.e., rollover angle), and PRV system damage have the potential 

to alter PRV performance. Preliminary component-level testing conducted by the project team 

for a Transport Canada project demonstrated that cryogenic fluid flow through PRVs used on 

UN-T75 tanks (i.e., spring-loaded, PTFE sealed, single-phase flow devices) can significantly 

reduce performance under certain conditions. In the same project it was shown that the jet flame 

generated from exhausting high velocity natural gas through the PRV can impinge on a 

neighboring tank and induce a localized heat flux greater than that generated from a pool fire. 

The results of FRA’s most recent UN-T75 fire test supported this finding (FRA, 2023). An 

example of a PRV-exhaust jet fire during a fire test of an LNG-filled UN-T75 tank is shown in 

Figure 1. 

Figure 1. Jet Fire of PRV exhaust during fire test of LNG filled UN-T75 tank (FRA, 2023) 
4 

In this study, the research team characterized PRV performance under inverted and elevated 

pressure conditions and related PRV performance to the probability of tank failure under fire 

conditions. The researchers then present countermeasures to prevent negative outcomes. 

1.2 Objectives

The objectives of this project were to: 

• Identify PRV performance and design requirements from the regulatory and scientific 

literature

• Use validated FE models to estimate the damage caused to the PRV system during select 

derailment scenarios

• Measure PRV performance under cryogenic liquid flow conditions through physical 

testing

• Conduct physical tests to measure the thermal output of an ignited gas jet

• Quantify the effect of measured PRV performance and jet flame exposure on the risk of 

tank failure

1.3 Overall Approach

The research team used a combination of multi-physics modeling and physical testing to measure 

and predict the performance of individual components for a full-scale tank in derailment and fireexposed conditions. FE modeling of a UN-T75 tank was used to simulate the damage to the PRV 

piping assembly during derailment scenarios. Eight PRV devices, typical of the style installed on 

multi-modal cryogenic LNG tanks, were tested under cryogenic liquid flow conditions 

representative of an inverted tank using liquefied nitrogen (LN2) in place of LNG. PRV tests 

were conducted at multiple tank pressures, PRV temperatures, downstream constriction levels, 

environmental conditions, and exhaust pipe lengths. Separate tests were conducted to 

characterize the heat flux generated by ignited natural gas jets at different flow rates, 

orientations, and orifice diameters. The data generated from the physical tests were used to 

inform the inputs of previously validated multi-physics models of fire exposure on a UN-T75 

tank. Previous test data and modeling approaches generated through projects with FRA and 

Transport Canada served as a foundation to this work. 

1.4 Scope

This twelve-month project used virtual and physical testing to evaluate the performance of a UNT75 tank under derailment conditions, PRVs under cryogenic liquid flow conditions, and the 

resulting effect on tank response when exposed to fire. 

1.5 Organization of the Report

The results of a literature and regulation review are presented in Section 2, followed by the 

results of crashworthiness simulations in Section 3. Physical testing, which relied on the 

information obtained from the literature review and the simulations, is presented in Section 4 and

Section 5. The demonstration of countermeasures through FE analysis and the characterization of 

tank thermal response to jet fires are provided in Section 6 and Section 7. Suggested future work 
5 

is discussed in Section 8, and conclusions are presented in Section 9. Fabrication and Test Setup 

schematics are presented in Appendix A and Appendix B. 
6 

2. Literature Review of PRV Characteristics

The research team reviewed selected literature, specification sheets, and American Society of 

Mechanical Engineers (ASME) regulations to identify the performance ranges, design types, and 

most commonly used PRVs for cryogenic tanks. Fire exposure to PRVs has been shown to 

reduce the effective set pressure due to thermal weakening of the internal spring (Prabhakaran et 

al, 2022), although these tests artificially protected the PRV seals during initial heating. During 

the fire testing of an ISO UN-T75 tank with LNG (FRA, 2023) the seals on every external valve 

were observed to fail, leak gas, and produce a flame. The discharge of flammable material 

through PRVs has been observed to increase the severity of the impinging fire (Prabhakaran et 

al, 2022). Most manufacturers provide some warning about using gas/vapor-only PRVs in liquid 

or dual-phase flow conditions (e.g., “Using a standard gas certified trim to flow liquids creates 

chatter and damages valve components. This is not recommended.” (Baker Hughes, n.d.)).

2.1 Key PRV Requirements

The review findings specific to PRV design on multi-modal tanks are summarized below.

• The device type must be spring-loaded1. 

o Seals for such devices are predominantly Polytetrafluoroethylene (PTFE) with a 

maximum service temperature of 500°F.

• The device must fully open at 110 percent maximum allowable working pressure 

(MAWP) and fully close at 10 percent below the discharge pressure2

. 

• The valve capacity is determined with the assumption of gas/vapor flow and must be 

sufficient to prevent the tank from exceeding the test pressure when under complete fire 

engulfment. 

• The discharge capacities for these types of valves typically ranges from 36 to 3162 CFM 

for set pressures between 10 and 400 psi. (Rockwood Swendeman, n.d.). 

o There were no significant differences noted regarding performance or design of 

spring-loaded valves between manufacturers.

• PRVs were most commonly situated external to the tank and often within a metal cabinet.

The PRV selected for testing (Rockwood Swendeman 710NEGH-A114; see Section 4 and

Section 5) meets all the above criteria for an ISO UN-T75 tank and was used in two full scale 

fire tests.

Key design features of the PRV piping assembly include: 

• Backpressure must be minimized, and obstructions prevented within the PRV piping 

assembly3

1 49 CFR § 178.274.f.1.v

2 49 CFR § 178.277.e.1

3 49 CFR § 178.274.f.2
7 

• All PRV inlets must be situated in the vapor space

• Escaping vapor must be directed away from the shell so that it cannot impinge on the 

shell

Table 1 summarizes the potential consequences of a derailment and correlates them with the 

performance requirements outlined in the regulations. This study shows that derailment scenarios 

can result in a tank no longer meeting the design or performance requirement outlined in 49 CFR 

178. Specifically, damage during a derailment could constrict flow within the PRV piping 

assembly (inlet and outlet). A tank inversion would cause the PRV inlet to move from the vapor 

space into the liquid space. Even without a derailment or tank inversion, the escaping vapor from 

one tank, ignited or not, could impinge on a neighboring tank shell. 

Table 1. Potential effects of a derailment and the associated regulation 

Derailment Effect Secondary Effect Affected Requirement

Damage to PRD upstream 

piping

Obstruction leading to PRD which might restrict 

flow 178.274.f.2

Damage to PRD outlet 

piping

Increased back-pressure on relieving device 178.274.f.2

Deformed outlet piping becomes directed toward the 

tank shell 178.274.f.3.i

None As-designed PRD outlet may be nominally directed 

at a neighboring tank 178.274.f.3.i

Rotated or inverted tank PRD inlet no longer in the vapor space 178.274.f.3.i

Damage to general tank 

structure

Protect the devices from damage caused by the 

portable tank overturning 178.274.f.3.ii

Pool fire impingement 

upon PRD

Damage to seals and/or closing mechanism (i.e., 

spring) prevents closing at 10% below discharge 

pressure or fully closing at all

178.277.e.1

2.2 Comprehensive PRV Requirements

A more extensive summary of the regulatory requirement review is provided below. 

Performance-Related Requirements of Pressure Relief Devices for UN Portable Tanks4

• Types5

o Spring Loaded 

o Frangible Disc (not acceptable for liquified gases) 

o Fusible Elements (not acceptable for liquified gases)

• Rated flow capacity shall be marked on the device [using Standard Cubic Meter Per 

Second] (SCMS) 6

4 49 CFR § 178.274 – Specifications for UN portable tanks

5 49 CFR § 178.274.f.1.v

6 49 CFR § 178.274.f.1.v
8 

o For Spring-Loaded PRD the rated flow capacity must be determined according to 

ISO 4126-1 (including Technical Corrigendum 1) and ISO 4126-7 (IBR, see § 

171.7 of subchapter)

o Cross sectional flow areas marked on the device (mm2) 

• Connections to PRDs7

o There must be no obstruction in an opening leading to a vent or PRD which might 

restrict or cut-off the flow from the shell to the device. 

o Vents or pipes from the pressure relief device outlets, when used, must deliver the 

relieved vapor or liquid to the atmosphere in conditions of minimum backpressure

on the relieving devices.

• Location of PRDs8

o All PRD inlets must be situated in the vapor space of the shell.

o The escaping vapor must be directed away from the shell in such a manner that it 

cannot impinge upon the shell. 

o Provisions must be implemented to protect the devices from damage caused by 

the portable tank overturning. 

Performance-Related Requirements of Pressure Relief Devices for UN Portable Tanks 

transporting refrigerated or liquefied gases9

• Opening and closing of PRDs10

o Must open automatically at a pressure not less than the MAWP and be fully open 

at a pressure equal to 110% of the MAWP.

o Must close at a pressure not lower than 10 percent below the pressure at which 

discharge starts and must remain closed at all lower pressures. 

• Capacity and setting of PRDs11

o In the case of a loss of vacuum or loss of 20 percent of the insulation insulated 

with solid materials the combined capacity of all PRDs must be sufficient so that 

the pressure (including accumulation) inside the shell does not exceed 120 percent

of the MAWP. 

o Under the circumstances described in paragraphs (e)(4)(i) and (e)(4)(ii) of this 

section, together with complete fire engulfment, the combined capacity of all 

7 49 CFR § 178.274.f.2

8 49 CFR § 178.274.f.3

9 49 CFR § 178.277 - Requirements for the design, construction, inspection and testing of portable tanks intended 

for the transportation of refrigerated liquefied gases.

10 49 CFR § 178.277.e.1

11 49 CFR § 178.277.e.4
9 

pressure relief devices installed must be sufficient to limit the pressure in the shell 

to the test pressure.

o The required capacity of the relief devices must be calculated in accordance with 

CGA Pamphlet S-1.2 (IBR, see § 171.7 of this subchapter). 

Definitions

• Portable Tank12: Portable tank means a bulk packaging (except a cylinder having a water 

capacity of 1000 pounds or less) designed primarily to be loaded onto, or on, or 

temporarily attached to a transport vehicle or ship and equipped with skids, mountings, or 

accessories to facilitate handling of the tank by mechanical means. It does not include a 

cargo tank, tank car, multi-unit tank car tank, or trailer carrying 3AX, 3AAX, or 3T 

cylinders. 

• UN Portable Tank13: an intermodal tank having a capacity of more than 450 liters (118.9 

gallons). It includes a shell fitted with service equipment and structural equipment, 

including stabilizing members external to the shell and skids, and mountings or 

accessories to facilitate mechanical handling. A UN portable tank must be capable of 

being filled and discharged without the removal of its structural equipment and must be 

capable of being lifted when full. Cargo tanks, rail tank car tanks, non-metallic tanks, 

non-specification tanks, bulk bins, and IBCs and packagings made to cylinder 

specifications are not UN portable tanks. 

• MAWP14: Maximum allowable working pressure (MAWP) means the maximum 

effective gauge pressure permissible at the top of the shell of a loaded portable tank in its 

operating position including the highest effective pressure during filling and discharge.

• Shell15: Shell means the part of the portable tank which retains the refrigerated liquefied 

gas intended for transport, including openings and their closures, but does not include 

service equipment or external structural equipment. 

• Test Pressure16: Test pressure means the pressure to which a tank is subjected to 

determine structural integrity. 

12 49 CFR § 171.8

13 49 CFR § 171.8

14 49 CFR § 178.277

15 49 CFR § 178.320

16 49 CFR § 178.320
10 

3. FE Simulation of Crash Conditions

Researchers generated and validated thermal and structural FE models of an ISO UN-T75 tank in 

previous efforts (Friedman & Mattos, 2018; Mattos & Friedman, 2023). The structural model has 

been used to simulate rollovers, impacts with fixed objects, and impacts with other railcars and 

tanks to investigate the risk of internal tank failure and loss of contents. This model has been 

modified in this project to include pertinent internal and external PRV piping assembly features. 

The model was then subjected to simulated impacts to determine the threshold impact velocity 

producing PRV piping assembly damage. 

3.1 FE Model Description

The existing FE model of the UN-T75 tank was modified using PRV piping assembly details 

provided by Taylor Wharton via drawings and images (Figure 2). These were converted to CAD, 

meshed within LS-DYNA, and added to the existing full model. The materials were defined 

using temperature-dependent thermal and structural properties based on data from the literature 

and previously validated material models. 

Figure 2. Modified FE model of UN-T75 tank with added PRV piping assembly shown with 

surrounding environment (left) and with isolated piping assembly (right) [LS-DYNA]

For all simulations, the UN-T75 tank model was placed on top of a flatbed railcar representative 

of Bethlehem Steel Series car and attached at each corner with constraints representing 150 kN 

ultimate strength twist-locks as per ISO 1496 (Figure 3). The entire vehicle was placed on 

standard freight rails on a standard cross section of mainline track (Figure 4). This placed the 

initial center of gravity of the tank at about 10 feet above the adjacent sub-ballast section. The 

tank was modeled as full of LNG, which was simulated using smooth particle hydrodynamics 

(Figure 5). For all simulations the ground and fixed objects (e.g., pole impact) were assumed to 

be rigid. Other vehicles were modelled with elastic-plastic material properties.

Figure 3. Physical and FE model of UN-T75 tank positioned on flatbed railcar
11 

Figure 4. Initial position of vehicle combination including rails, ballast, and surrounding 

terrain; FE Model (top) and engineering specifications (bottom)

Figure 5. View of model showing LNG within the tank (shown graphically as ‘dots’) at 

initiation of rollover 

3.2 Methods

Two test series were simulated in this task: a series of crashes and a series of flow tests. 

The simulated crash conditions are summarized in Table 2. Each scenario was repeated until the 

team identified the minimum velocity that resulted in significant damage to the PRV piping 

assembly to potentially degrade its performance. “Significant” damage was qualitatively defined 

as deformation of a PRV system component that could result in degradation of the system 

performance.
12 

Table 2. FE Crash Simulation Matrix 

Impact Location Impact Partner Description

Piping cabinet Flatcar Stationary UN-T75 tank & flatcar, rolled 95 

deg, is impacted by an empty flatcar 

Top Center 24-inch rigid pole UN-T75 tank and flat car roll into a 

stationary rigid pole and impact at ~90 deg 

roll angle at the tank center

Top Third 24-inch rigid pole UN-T75 tank and flat car roll into a 

stationary rigid pole and impact at ~90 deg 

roll angle 13 ft from tank end

Top End 24-inch rigid pole UN-T75 tank and flat car roll into a 

stationary rigid pole and impact at ~90 deg 

roll angle 10 ft from the tank end

Rollover Rollover UN-T75 tank and flat car roll off rails onto 

the surrounding terrain
13 

In flow test simulations, the team modeled the flow of methane gas through a straight length of 

2-inch schedule 40 pipe with varying degrees of deformation based on the crash test results. A 

compressible flow solver was used to account for compression of the gas and its effects on the 

state of the fluid as it flowed through the pipe. The tests included a baseline case (no pipe 

damage) and cases with 50, 80, and 95 percent closure. 

3.3 FE Results

FE Crash Results are documented in Table 3. 

Table 3. FE Crash Results

Mode Impact Delta-V Damage

Rear Impact –

Flatcar

15 mph • Exhaust pipe pinched ~50% closed x 2-inch length

• Internal tank undamaged

20 mph • Exhaust pipe pinched ~80% closed x 19-inch length

• Internal tank minor dent

30 mph • Exhaust pipe pinched 90% closed x 20-inch length.

• Annular space pipe pinched 80% closed x 7-inch length

• Internal tank dented

49 mph • Exhaust pipe pinched > 95% closed

• Exhaust pipe torn

• Annular space pipe pinched >95% closed

• Annular space pipe torn

• Internal tank dented

Top-center –

Pole

25 mph • No pipe damage

• Internal tank likely breached

35 mph • Internal pipe minimally damaged

• Internal tank likely breached

49 mph • Internal pipe minimally damaged

• Internal tank breached

Top-quarter –

Pole

15 mph • Internal pipe pinched ~50% closed

• Internal tank crushed ~25% of diameter

25 mph • Internal pipe pinched ~75% closed

• Internal tank crushed ~30% of diameter

35 mph • Internal pipe pinched > 95% closed with potential tearing

• Internal tank crushed ~45% of diameter

49 mph • Internal pipe pinched > 95% closed with potential tearing

• Internal tank breached

Top-end – Pole 10 mph • The internal pipe pinched 60% closed

• Internal tank crushed ~

15 mph • Internal pipe pinched 75% closed with potential tearing

• Internal tank crushed ~10% of diameter

25 mph • Internal pipe pinched 75% closed with potential tearing

• Internal tank crushed ~20% of diameter

35 mph • Internal pipe pinched 75% closed with potential tearing

• Internal tank crushed ~35% of diameter

49 mph • Internal pipe pinched > 95% closed with tearing

• Internal tank crushed ~50% of diameter

Rollover 60 deg/s to 180 deg/s • No notable pipe damage for any of the evaluated rollover 

conditions
14 

The team observed considerable PRV system data, mostly in the form of pipe damage, in

impacts with railcars and rigid poles at speeds as low at 15 mph. No damage to the PRV system 

directly was noted for any of the pure (i.e., without impacts against other objects) rollover cases 

up to 180 deg/s roll velocity. PRV system damage was obviously sensitive to the impact 

location. Longitudinal impacts to the piping cabinet end of the tank caused damage to external, 

post-PRV piping whereas lateral impacts to the tank resulted in damage to the interior piping 

assemblies. At 10-15 mph the 2-inch schedule 40 stainless steel pipes were often pinched 

approximately 50-60 percent closed. Examples of this damage is provided in Figure 6, Figure 7, 

Figure 8, and Figure 9. 

Figure 6 Pipe damage in rear-end flatcar impacts at 15 mph, 20 mph, 30 mph, and 49 mph 

(from left to right) 

Figure 7. Damage to inner tank and internal PRV piping in the Top-quarter – Pole impact 

at 15 mph; damage at higher speeds was similar in shape but greater in magnitude 
15 

Figure 8. Original (left) and reduced pipe cross-section (right) after 25 mph Top-quarter – 

Pole impact

Figure 9. Damage to inner tank and internal PRV piping in the Top-end – Pole impact at 

10 mph; damage at higher speeds was similar in shape but greater in magnitude 

The flow simulations demonstrated that the damage to the pipes significantly degraded the flow 

rate (Figure 10). The amount of deformation for each pipe model is shown in Figure 11, Figure 

12, Figure 13, and Figure 14. The average velocity through the baseline pipe was 49.5 m/s. 

Figure 10. Velocity vs Time for 3 pipe damage severities
16 

Figure 11. Methane gas flow through nominal pipe at 120 psi inlet pressure

Figure 12. Methane gas flow through ~50% closed pipe at 120 psi inlet pressure 

Figure 13. Methane gas flow through ~80% closed pipe at 120 psi inlet pressure
17 

Figure 14. Methane gas flow through >90% closed pipe at 120 psi inlet pressure 

3.4 FE Crash Discussion

Deformation of the internal piping assembly (upstream of the PRVs) or external piping assembly 

(downstream of the PRVs) is possible at impact velocities around 10-15 mph. The extent of 

piping deformation shown in the FE simulations is likely to degrade the capacity of the assembly 

to move vapor and/or liquid through the PRV system and is specifically in contradiction to the 

requirements in 49 CFR 178.274.f.2. The testing described in Section 4 investigated the effect of 

downstream constrictions on flow rates for cryogenic liquid flow through the PRV and found 

that flow was substantially reduced. 

This study only considered derailment crashworthiness of a specific intermodal tank traveling on 

top of a flatcar. This represents a particular scenario that exposes the tank to different impact 

configurations than would be experienced by a tank car or even an intermodal tank double 

stacked on a well car. These other configurations could be investigated using similar modeling 

strategies.
18 

4. PRV System Test

Based on the results of the literature review, FE simulations, and previous tests of the UN-T75 

tank, the team developed a physical test matrix to characterize the flow rate of cryogenic liquid 

nitrogen through a representative PRV assembly. The testing series included varying the inlet 

pressure, downstream constriction, PRV temperature, and environmental conditions (e.g., 

temperature and humidity). Measured output responses included mass flow rate, fluid and PRV 

temperature, and fluid density. 

4.1 Test Methods

The research team, working with Southwest Research Institute (SwRI) with input from FRA, 

designed and constructed an experimental PRV testing system using the 6,000-gallon LN2 vessel 

that provides nitrogen supply to SwRI’s Flow Component Testing Facilities. The PRV testing 

system was developed by 1) constructing the plumbing and controls for the LN2 (cold side) 

supply from an existing valve on the bottom of the LN2 supply vessel, and 2) creating a nitrogen 

gas charge system (gas side) to inject gaseous nitrogen into the head space of the LN2 supply 

vessel to provide gaseous nitrogen to the PRV inlet. This system allowed the team to deliver 

liquid nitrogen to the PRV from the cold side at controlled headspace pressure supplied from the 

gas side. In addition, gaseous nitrogen could be supplied to the isolated upstream side of the PRV 

after an LN2 discharge to check for any gaseous leakage around the seat. A drawing of the PRV 

testing system is provided in Appendix A. The size and pressure rating of the LN2 vessel 

provided sufficient volume to meet and exceed the estimated PRV discharge capacity. The large 

LN2 storage volume allowed more than sufficient flow time for flow discharges beyond one 

minute in duration. 

4.1.1 Test Articles (PRVs) 

The team used Rockwood Swendeman 710NEGH-A114 PRVs in this experiment, specifically 

RXSO bronze ASME coded bodies with a 2-inch orifice, 1.5-inch inlet diameter, and 2-inch 

outlet diameter. These are the same model numbers used on previous UN-T75 tests and represent 

the most common design used on similar tanks. The A114 indicates that it is intended for air or 

gas service with a set pressure of 114 psi. The PRV was sealed with a metal seat and PTFE 

plunger. 

4.1.2 Test Section 

The test section piping arranged for PRV testing was fabricated per the test section and setup 

fabrication drawing shown in Appendix B; the test setup CAD model is shown with additional 

content in Figure 15. The test section included 2-inch schedule 40 304L stainless-steel piping, a 

1-inch full bore automated cryogenic ball valve on the upstream side (upstream supply isolation 

valve), a 2-inch Coriolis flow meter, and a tee with a 1-inch full bore automated cryogenic ball 

valve used as a purge to atmosphere (i.e., purge valve) located just upstream of the PRV. Ports to 

measure pressure and temperature were located upstream of the flow meter and upstream of the 

PRV. The official upstream pressure measurement was taken at three nominal pipe diameters 

(3D) upstream of the PRV connection. A 1.5-inch diameter stainless-steel hose connected the 

LN2 supply vessel to the upstream supply isolation valve. 
19 

Figure 15. PRV test section arrangement

Testing was performed using the system as shown in Figure 16. 

Figure 16. PRV test system
20 

The figures show the LN2 supply plumbing and testing instrumentation. The pressure 

transmitters were mounted on vertical static stand-off tubes used to isolate the pressure 

transmitter diaphragm from the cryogenic temperature. High-accuracy Rosemount 3051S smart 

transmitters were used to measure pressure, and a Micro Motion ELITE CMF200 Coriolis flow 

meter was used to measure the mass flow rate of the nitrogen and the nitrogen fluid density. 

Type K thermocouples were used to measure the PRV temperature, and the fluid temperature 

measured upstream of the flow meter. All test data were measured and recorded electronically 

using an existing facility data acquisition system. This data acquisition and control system was 

also used to control the upstream supply valve and purge valve remotely from the facility control 

room. 

4.1.3 Nitrogen Gas Charge System 

The nitrogen gas charge system is also shown in the test system piping and instrumentation 

diagram (P&ID) in Appendix B. Gaseous nitrogen supplied from SwRI’s nitrogen source 

(pressure-stored at 2,000 - 2,900 psig) was used by reducing pressure through a pressure 

regulator to inject the nitrogen and increase the pressure in the 6,000-gallon LN2 supply vessel. 

A pressure transmitter was used on the gas injection line to measure the LN2 supply vessel 

pressure. Gas was injected until the desired test run start pressure was reached. A small diameter 

(1/4-inch tubing) supply line containing an isolation valve branched from the LN2 injection line 

and connected to the upstream PRV pressure port location on the test section. This line was used 

to apply ambient temperature gaseous LN2 into the test section for PRV seat testing. The 

pressure-reducing regulator was used to control the pressure. 

4.1.4 PRV Discharge Piping Arrangement 

A 2-inch schedule 40 pipe measuring 24-inches-long was used as the PRV discharge plumbing 

for the initial series of tests. A second series of tests added two 90 deg elbows and 12 ft of 2-inch 

piping downstream of the original discharge assembly (Figure 17). This was done to more 

closely replicate the total length of piping downstream of the PRV assembly on a UN-T75 tank. 

Figure 17. PRV test system longer discharge
21 

A PRV discharge piping constriction, used for certain tests, was created by machining ellipticalshaped orifices in a 2-inch pipe cap to replicate a crushed (i.e., oval-shape) pipe. The two PRV 

discharge-constricted orifice sizes with which the PRV was tested are shown in Figure 18. The 

orifice area on the more severe constriction is 0.785 in2 which represents an 89 percent closure. 

The orifice area on the less severe constriction (1.57 in2) represents a 78 percent closure.

Figure 18. Machined elliptical orifices used for PRV discharge constriction

4.1.5 Simulated Fire Test 

A 500,000-BTU propane torch was used to heat the PRV to simulate rapid heating of the PRV 

during a crash scenario. The PRV body temperature was monitored with a Type K thermocouple 

strapped to the body with a ring clamp. The propane torch was modified for safe operation and 

remote control by placing a control valve on the propane feed and equipping the torch with an 

electric igniter.

4.1.6 Test Series Description

All tests were conducted with pressurized LN2 applied at the inlet of the PRV test section. The 

test matrix (Table 4) was developed based on the cause-and-effect analysis that is summarized in 

Table 1 to quantify the effects of potential PRV degradation scenarios. 

Table 4. PRV Flow Test Matrix

Pressure (psi) Constriction Relative Humidity PRV Temperature (°F)

Baseline 130, 150, 190 None Ambient Ambient

Constriction 130, 175, 190 78%, 89% Ambient Ambient

Heated 130, 240 None Ambient 721, 744, 1000

Humid 130, 150 None 80-100% Ambient

Extended exhaust 130, 180 None Ambient Ambient

Extended exhaust + 

constriction 130, 180 89% Ambient Ambient
22 

The primary degradation scenario of interest was the effect of cryogenic liquid flow through the 

PRV assembly. The key test parameters included cryogenic fluid flow to represent the PRV inlet 

no longer in the vapor space, elevated PRV temperature to simulate damage to seals, and 

constricted outlet piping to simulate increased backpressure. Tests at increased ambient humidity 

were performed to replicate the conditions that led to PRV ice-up in previous tests. Test 

conditions were repeated multiple times and data was collected in each test for at least 60 

seconds or until equilibrium was achieved. Testing revealed some sensitivities such as the effect 

of the LN2 supply tank fill level that affected test outcomes in an unexpected manner. 

4.2 Test Results

For each test, the temperature was recorded both upstream of the Coriolis flow meter and at the 

PRV. Pressure was monitored and recorded upstream of the Coriolis flow meter on the gaseous 

nitrogen feed into the headspace and just upstream of the PRV. Mass flow rate and fluid density 

were monitored by the Coriolis flow meter and recorded. For ease of viewing the data in this 

section, the PRV temperature and pressure, as well as the mass flow rate and the fluid density, 

are shown as a function of time. 

4.2.1 Baseline Tests 

The baseline tests consisted of repeatedly discharging LN2 through the PRV at a headspace 

pressure ranging from 130 to 190 psi. During the baseline tests, two feet of 2-inch pipe were 

connected to the downstream discharge from the PRV with no constrictions added. Figure 19

shows a typical discharge experiment. This run happens to be at 130-psi drive pressure, and it is 

the first discharge in a sequence. So, it can be observed that the PRV temperature starts close to 

100°F before rapidly dropping to cryogenic temperatures. 

Figure 19. Baseline test at 130 psi – Run 1 (PRV temperature [°F], PRV pressure [psi], 

mass flow rate [lb/min], and density [kg/m3])

As can be seen, it took over one minute for the PRV to cool to equilibrium, and subsequently, it 

took as long for the flow rate and fluid density to equilibrate. Conversely, on a repeat discharge 

experiment, the PRV was pre-cooled and the system reached equilibrium much more quickly 

(approximately 15 seconds), as shown in Figure 20. 
23 

Figure 20. Baseline test at 130 psi – Run 4 (PRV temperature [°F], PRV pressure [psi], 

mass flow rate [lb/min], and density [kg/m3])

This process was repeated at several discharge pressures. In all cases, it was observed that upon 

closing the upstream control valve to terminate LN2 flow, the PRV reseated and held pressure. 

This can be seen in the orange PRV pressure traces in Figure 19 and Figure 20. 

A summary of equilibrium mass flow rates as a function of discharge pressure for the baseline 

tests is plotted in Figure 21. 

Figure 21. Summary of baseline tests – Flow rate vs. tank pressure

The mass flow rates recorded in these tests were similar to the mass flow rate generated from 

high pressure (~870 psi) natural gas leaking through a 20 mm diameter orifice (Lowesmith & 

Hankinson, 2012). The most surprising feature of these data is that the flow rate was not dictated 

solely by the upstream pressure. It was only discovered after plotting the data that the mass flow 

rate increased significantly after a regularly scheduled filling of the LN2 supply tank was 

completed, even at similar upstream pressures. The cause of this increase became clear once the 

liquid density was plotted in a similar manner (see Figure 22). It was discovered that the density 

of the fluid discharged through the Coriolis flow meter had significantly increased. In the tests 
24 

prior to the LN2 resupply, the fluid density indicated a two-phase flow was entering the Coriolis 

meter. After resupply the flow was at a density indicating it was nearly pure liquid. Not only 

does an increase in fluid density help drive higher mass flow rate, but it also mitigates the choked 

flow issue that is likely associated with the two-phase flow condition. 

Figure 22. Summary of baseline tests – Fluid density vs. tank pressure

4.2.2 Constriction Tests 

The research team also studied the effect of a constriction on the downstream discharge tubing to 

simulate a bent or partially pinched pipe downstream of the PRV. Two different sizes of orifices 

were cut into pipe end caps to repeatedly simulate this type of constriction. Once again, the PRV 

was able to safely discharge and reseat in all of the tests. The summary of the results of these 

tests is shown in Figure 23. 

Figure 23. Comparison of unrestricted vs restricted downstream conditions; note that the 

130- and 175-psi discharges were conducted with the supply tank half full, and the 190-psi 

discharges were conducted immediately after the supply tank was filled
25 

The high-pressure data for the 78 and 89 percent constrictions cannot be directly compared 

because the supply tank was filled while conducting this testing. The impact on the results was 

not fully understood until after the data were processed. Even with this limitation, a consistent 

trend was observed in that the 78 percent downstream constriction resulted in very little flow rate 

decrease (approximately 3.5 to 5.5 percent). The 89 percent downstream constriction resulted in 

a much more significant decrease in mass flow rate, for both the two-phase fluid flow at 130 psi 

and the nearly single-phase liquid flow at 190 psi. A decrease of approximately 30 to 38 percent

was observed for the smaller constriction. In addition to the reduction in equilibrium flow rate, 

the flow rate oscillated considerably upon initial opening of the PRV. This can be observed in 

the example trace shown in Figure 24. 

Figure 24. 89% Constriction at 130 psi – Run 4 (PRV temperature [°F], PRV pressure 

[psi], mass flow rate [lb/min], and density [kg/m3])

This oscillation in the flow rate occurs because upon initial release, the downstream orifice 

creates backpressure in the discharge piping. The result is that the differential pressure across the 

PRV decreases, and it begins to shut. As the PRV chokes off, the flow rate decreases and, 

consequently, so does the backpressure in the downstream pipe. The result is that the PRV 

reopens fully. 

4.2.3 Heated Tests

This section summarizes the results of preheating the PRV to high temperatures to simulate a 

nearby fire or other heat source. This was accomplished by using a propane torch to preheat the 

valve and then discharging LN2 through the PRV. Whenever the PRV was heated above 

approximately 700°F for a significant time, the PTFE plunger in the PRV melted or thermally 

decomposed causing a failure in the ability of the PRV to reseat and hold pressure. The PRV still 

opened successfully, but it no longer functioned as a pressure relief valve with a set pressure. 

An example plot of a hot test is shown in Figure 25 where the mass flow rate reached 800 kg/m3

for a tank pressure of 240 psi. The PRV failed its seat leakage test after any hot test above 700°F. 

The loss of the PTFE plunger due to thermal effects resulted in a slight increase in mass flow rate 

(approximately 14 percent higher) as compared to experiments where the PTFE was undamaged. 
26 

Figure 25. Hot test at 240 psi – Run 1 (PRV temperature [°F], PRV pressure [psi], mass 

flow rate [lb/min], and density [kg/m3])

Tests were conducted just below 721°F, just at 744°F, and above 1050°F, the melt temperature of 

the PTFE seal. In the case of heating to below the melting temperature of the PTFE, the seat 

swelled slightly but it did not cause the PRV to fail to open or affect its performance. In the case 

of heating to just at the melting temperature of PTFE, it was observed that while the PRV had no 

issues opening, the flow rate was much more unstable (see Figure 26).

Figure 26. Hot test at PTFE melt temp (PRV temperature [°F], PRV pressure [psi], mass 

flow rate [lb/min], and density [kg/m3])

The pressure at the PRV decreased slightly over time and the PRV failed the subsequent seat 

leak test. It is unclear if this valve would have functioned properly if another LN2 discharge was 

attempted after sustaining this damage. The springs on the PRVs that reached more than 1,000°F 

were permanently compacted slightly, and it is unclear if the PRV would have been able to reseat 

correctly even if the PTFE had not been completely removed. It is suspected that a softening of 

the spring occurred while being compressed at high pressure, and it remained compressed while 
27 

the LN2 quenched the spring in this compressed position. The thermal damage to each of the 

PRVs can be seen in Figure 27. 

Figure 27. PRVs after hot testing; temp < melt (left), temp = melt (middle), temp >> melt 

(right)

4.2.4 Humid Tests

The relative humidity at the test facility on multiple occasions during testing reached 80 to 95 

percent. Several tests were conducted at high humidity to determine if the PRV would ice over or 

fail to reseat due to ice buildup from ambient moisture in the air. In these tests the PRVs were 

precooled for 2 to 10 minutes down to a temperature approaching -220°F and with discharges of 

several minutes. One such test is shown in Figure 28. 

Figure 28. Humid test at 150 psi – Run 4 (PRV temperature [°F], PRV pressure [psi], mass 

flow rate [lb/min], and density [kg/m3])

Significant ice buildup was seen on the external PRV body, as well as on all the piping. 

However, due to the downstream discharge piping, the moisture and ice could not build up on the 

internal components of the PRV. Even at near-saturated humidity, the volume of air in and 

around the PRV internals did not hold sufficient moisture to build up and restrict the PRV 

mechanism. In all the humid tests the PRV reseated, sealed, and held pressure.
28 

4.2.5 Extended exhaust

The baseline tests were repeated with the addition of an extended exhaust pipe assembly that 

included two 90-degree angles and 12 ft of pipe to more closely replicate the total length and 

quantity of fittings on the fire-tested UN-T75 tank. 

Three tests of multiple discharges were executed: 130 psi (tank full), 130 psi (tank half full), and 

175 psi (tank half full). The first of the baseline tests was performed with a tank head pressure of 

130 psi, but only two of the five discharges were completed. The LN2 supply vessel was almost 

full of liquid, meaning that the built-up head pressure was quickly spent, and significant time 

was needed to allow it to rebuild. It should be noted that LN2 had been freshly supplied to the 

vessel, so it had no time to equilibrate its temperature. In the second baseline test the head 

pressure was once again 130 psi, but the supply vessel was approximately half full. A large 

quantity of fluid had been pumped into separate storage tanks a few hours before the test and

time was allowed for the fluid to equilibrate in temperature. A full five discharges were 

conducted for the baseline test, and a further five were performed with the same head pressure 

and tank conditions, but with the constricted orifice installed. 

The results were comparable to the original tests with shorter exhaust length, as shown in Figure 

29. 

Figure 29. Summary of mass flow rates with extended exhaust added

The 130 psi and 175 psi tests were repeated with the addition of the more sever constriction 

which represents an 89 percent closure of the pipe. Again, the results were similar to the baseline 

conditions as shown in Figure 30. 
29 

Figure 30. Summary of mass flow rates with extended constricted exhaust added
30 

5. PRV Exhaust Flame Test

The thermal characteristics of natural gas flares/jet flames were quantified through a series of 

physical tests at SwRI’s Fire Technology Department. The objective of the testing was to 

generate thermal data for different configurations, flow rates, and pressures that can be used to 

investigate the effects on the thermal response of rail tanks and to use in the validation of future 

simulations (e.g., Fire Dynamics Simulator).

5.1 Test Methods

5.1.1 Facility Overview

Testing was conducted in a large-scale steel-lined facility nominally 65 x 65 x 60 ft. An 

adjustable 55 x 55-ft false ceiling allows simulating structures of different heights; for this 

project it was raised to its full height. Combustion air enters the building through 12 

symmetrically placed louvers around the perimeter of the building. The twin exhaust fans (each 

capable of moving up to 60,000 scfm of air) are operated during test runs to allow for quick and 

efficient smoke extraction. 

Tests can be viewed from an exterior door and an elevated observation room adjacent to the 

building or directly from the test floor, depending on the project scope. The facility also houses 

the data acquisition system, control system for exhaust cans, and fire suppressant system. Figure 

31 shows a basic schematic of the test facility site. 

Figure 31. Large-scale fire engineering and research facility layout
31 

5.1.2 Test Setup

Figure 32 shows a general schematic of the test setup for the NG delivery to the ignition source 

(i.e., propane gas sand burner). All piping used in the setup was schedule 40. The natural gas 

supply was 96.3 percent methane, 1.7 percent nitrogen, 1.6 percent ethane, 0.2 percent carbon 

dioxide, 0.1 percent propane, and 0.1 percent additional trace compounds (i.e., light 

hydrocarbons). Based on the measured constituents, the weighted-average heat of combustion 

was calculated to be 48.99 kJ/g. 

Figure 32. Schematic of natural gas delivery piping setup 

Two measurement stations were set up at specified distances from the visually observed flame 

top or flame middle position of the jet fire and provided room temperature and heat flux at three 

elevations (1.5, 4.5, and 7.5 ft above the floor; the NG pipe was 2 ft above the floor). 

Figure 33 provides schematics showing the relative position of the instrumentation stations for 

the horizontally oriented jet fire tests. Figure 34 shows a schematic of the relative position of the 

instrumentation for the vertically oriented jet fire tests. 

Figure 33. Schematic of instrumentation stations for horizontal jet fire tests

Temperature was measured from 20-ga Type K thermocouple wire (Geocrop model K20-2-G). 

Heat flux was measured from Schmidt-Boelter heat flux transducers with a range of 0-50 kW/m2

(Medtherm model 64-5-19 or 64-5-20). The 2-in diameter line pressure was measured with a 

pressure transducer with a range of 0-200 psig (Honeywell model THE) and the downstream 

discharge pressure was measured with a 0-100 psig range transducer (Tecsis Model 99-6702-

100G). The discharge flow rate was measured with a flow controller with a range of 0-3000 
32 

SLPM (Omega model FMA-2323-V2). All instrument outputs were logged on a data acquisition 

SYSTEM (Yokogawa model DU-100-11) at 1 Hz. 

Figure 34. Schematic of instrumentation stations for vertical jet fire tests

The following nominal test procedure was used for each run. 

1. Ignite the sand burner (pilot) 

2. Start DAQ

3. Open main valve 

4. Adjust line flow rate and/or pressure.

5. Flow for two minutes and collect data.

6. Close main valve and reset for next test

The summary data that is provided in this report is average data over the last 30-120 seconds of 

each run at a given configuration. Additional discussion of post-processing is provided in the test 

results section of this report. Figure 35 shows two photographs of the setup for the natural gas 

supply line and instrumentation to measure flow rate and line pressure. Figure 36 shows two 

photographs of a representative test before and after ignition of a jet fire flame. 

Figure 35. Natural gas supply into test facility, main cutoff valve, check valve and pipeline 

flame arrestor (left); Transition to 1.5-in diameter pipe and pressure flow rate 

measurement to data acquisition system (right)
33 

Figure 36. Pre-test conditions after sand burner is ignited (left); Test in progress with jet 

flame and north measurement state at mid-flame position (right)

5.1.3 Test Plan

The general plan included running tests across a range of natural gas flow rates and recording 

heat flux and temperature at two measuring stations (North and West). Table 5 provides the full 

test matrix. 

Table 5. Summary of jet fire test matrix

Test 

No

Flow 

Rate 

(SLPM)

Standoff 

Distance 

(ft)

North 

Station 

Flame 

Position

North 

Distance to 

Discharge 

Pipe (ft)

West 

Distance to 

Discharge 

Pipe (ft)

General Description /

Additional Comments

1 3000 ~4 ~Tip ND ND Scoping test to find position

2* 3000 4 Tip 7.5 11 First official test

3 3000 4 Tip 7.5 11 Repeat last test

4 2500 4 Tip 6.5 10 Changed flow rate

5 2000 4 Tip 5.5 9 Changed flow rate

6 1500 4 Tip 4.5 8 Changed flow rate

7 1000 4 Tip 3.5 7 Changed flow rate

8 500 4 Tip 1.5 5.5 Changed flow rate

9 3000 8 Tip 7.5 15 Changed standoff distance

10 2500 8 Tip 6.5 14 Changed flow rate

11 2000 8 Tip 5.5 13 Changed flow rate

12 1500 8 Tip 4.5 12 Changed flow rate

13 1000 8 Tip 3.5 11 Changed flow rate

14 500 8 Tip 1.5 9.5 Changed flow rate

15 3000 4 Mid 3.75 11 Changed station position

16 2500 4 Mid 3.75 10 VOID, forgot to move station

17 2500 4 Mid 3.25 10 Repeat test

18 2000 4 Mid 2.75 9 Changed flow rate

19 1500 4 Mid 2.25 8 Changed flow rate

20 1000 4 Mid 1.75 7 Changed flow rate

21 500 4 Mid 0.75 5.5 Changed flow rate

22 3000 8 Mid 3.75 15 Changed standoff distance

23 2500 8 Mid 3.25 14 Changed flow rate

24 2000 8 Mid 2.75 13 Changed flow rate

25 1500 8 Mid 2.25 12 Changed flow rate
34 

Test 

No

Flow 

Rate 

(SLPM)

Standoff 

Distance 

(ft)

North 

Station 

Flame 

Position

North 

Distance to 

Discharge 

Pipe (ft)

West 

Distance to 

Discharge 

Pipe (ft)

General Description /

Additional Comments

26 1000 8 Mid 1.75 11 Changed flow rate

27 500 8 Mid 0.75 9.5 Changed flow rate

28 3000 ~6.5 Scoping 7.5 11 Increased outlet from 1.5” to 3” dia.

29 3000 4 Tip 5 8.5 Adjusted to flame tip position

30 3000 ~0 Scoping 7.5 11 Changed outlet size to 1⁄2-in dia.

31 3000 4 Tip 11.5 15 Adjusted to flame tip position

32 3000 4 Vertical 4 4 1.5-in diameter outlet

33 3000 4 Vertical 4 4 3-in. diameter outlet

34 Range** 4 Vertical 4 4 0.5-in diameter outlet

ND: Not Determined

*Flame was unstable at higher flow rates for this diameter discharge outlet. Testing started at 500 SLPM and ramped 

up to a maximum of 2000 SLPM.

Initial scoping tests were conducted to optimize the ignition procedure and standoff distances for 

heat flux and temperature measurement points. The position of the flame tip was determined 

visually, and this length was divided in half of the tests conducted with the instrumentation 

stations located at the midpoint of the flame length. The following specific parameters were 

evaluated: 

• 500-3000 SLPM natural gas flow rate range 

• 4-ft and 8-ft standoff distances from flame tip and flame middle

• Additional test runs with different diameter outlets

• Additional test runs with vertical discharge orientations

5.1.4 Test Results

Table 6 and Table 7 provide average summary results for the North and West measurement 

stations, respectively. Heat Flux (HF) data are shown in units of kW/m2 and Thermocouple (TC) 

data are shown in units of °C. 

Table 6. Summary of average test results for north measurement station

Test No North Measurement Station (kW/m2

) & (°C)

1.5’ HF 1.5’ TC 4.5’ HF 4.5’ TC 7.5’ HF 7.5’ TC

2 10.1 174.5 15.7 182.4 12.4 177.7

3 9.5 191.3 12.9 192.1 10.5 192.9

4 7.4 168.1 10.9 157.7 9.8 161.1

5 7.9 174.2 12.2 181.8 9.9 169.4

6 6.4 134.1 9.6 147.5 7.5 135.4

7 5 131.7 7.5 153.6 5.6 128.3

8 3.9 98.8 4.6 102.9 2.6 76.7

9 3.8 92.8 5.2 94.5 5.5 96.2

10 4.2 112.8 5.6 110 5.5 108.4

11 3.5 99.1 4.4 103.6 4.6 101.9

12 2.8 86.4 3.3 89.9 3.5 89.9

13 1.8 77.1 2 76.7 2.2 76.8

14 0.9 51.1 0.8 53.7 1 51.8
35 

Test No North Measurement Station (kW/m2

) & (°C)

1.5’ HF 1.5’ TC 4.5’ HF 4.5’ TC 7.5’ HF 7.5’ TC

15 11.9 164.7 16.2 197.3 9.8 155.6

17 11.7 156.7 14.2 199.1 6.5 139

18 9.5 139.6 12.2 183.5 6.9 141.4

19 7 141.5 9.9 165.1 6.5 135.1

20 4.4 119.1 6.1 127 4.6 105.5

21 3.6 110.5 4.1 116.2 2.4 83.7

22 5 98.6 5.9 110.3 4.9 97.3

23 4.8 85.7 5.7 112.5 4.9 106.8

24 3.7 74.4 4.1 94.2 3.7 93.3

25 2.9 75.2 3.3 92 3.1 87.6

26 1.9 68.4 2.1 80.5 2 74.2

27 1 58.2 0.9 60.9 1 57.4

28* 5.1 141.7 8.8 156.7 9.3 163

29 11 214.8 20.7 286.8 18.9 266.6

30* 8.3 112.3 6.5 105.2 3.1 77.2

31 4 135.3 4.7 114.3 2.8 71.6

32 1.9 74.2 4.7 125.3 7.7 141.4

33 2 79.4 5.1 137.7 9.2 163.6

34** 1.2 65.2 3.1 110.2 5.8 120.5

HF: Heat Flux

TC: Thermocouple

*Scoping runs to find correct standoff distance for different outlet diameters

**Values in table are for 2000 SLPM flow rate

Table 7. Summary of average test results for west measurement station

Test No West Measurement Station

1.5’ HF 1.5’ TC 4.5’ HF 4.5’ TC 7.5’ HF 7.5’ TC

2 7.1 116.4 11 151.5 10.6 128.2

3 8.1 135.6 12 185.2 11.5 138.3

4 1.1 36.5 11.1 136.8 9.8 121.3

5 6.7 124.9 10.9 177.2 10.8 116.9

6 2.1 49.9 9.2 90.6 7.1 114.4

7 0.9 36.3 7.2 85 5.1 87.1

8 3.5 87.8 4.6 59.9 2 62.2

9 2.8 71 3.5 72.1 3.8 60.5

10 2.6 72.3 3.3 83.1 3.3 65.9

11 2.2 80.9 2.9 86 2.8 64.3

12 1.9 73.5 2.2 74.4 2.1 60.2

13 1.4 63.5 1.8 55.8 1.3 50.9

14 0.6 50.4 0.7 51.2 0.5 38.8

15 5.7 95.5 10.1 148.7 9.6 103.2

17 6.8 133.9 12 184.9 11.6 128.1

18 5.6 106.2 11 172.7 10.3 122.3

19 4.3 109.1 9 167.8 6.6 101.4

20 3.2 104.5 5.6 143.5 3.7 81.7

21 3.4 96 3.8 124.7 1.5 61.3

22 2.8 74.1 3.5 80.7 3.7 69.3

23 2.5 71.7 3.1 88.8 3.2 67.6

24 2.2 72.2 2.8 82.4 2.7 61.3
36 

Test No West Measurement Station

1.5’ HF 1.5’ TC 4.5’ HF 4.5’ TC 7.5’ HF 7.5’ TC

25 1.9 74.3 2.2 76.7 2 59.9

26 1.3 64.9 1.6 67.8 1.3 52.6

27 0.8 51.7 0.9 58.7 0.6 45.2

28* 4.1 91.2 6.1 112.3 7.9 94.1

29 6.8 137 15.9 149.6 17.9 173

30* 16.8 186.2 43.5 514.2 27 436.9

31 3 71.3 3.5 100.7 8.2 225.8

32 1.7 50.5 3.8 83.4 7.2 97.8

33 1.8 54.3 4 87.4 7.6 101.5

34** 1.1 45.1 2.7 73.4 5.1 86

HF: Heat Flux

TC: Thermocouple

*Scoping runs to find correct standoff distance for different outlet diameters

**Values in table are for 2000 SLPM flow rate

Figure 37 and Figure 38 provide plots of the time history of heat flux and thermocouple (i.e., 

temperature) measurements, respectively, from Test 15 (3000 SLPM, 4-ft Standoff Distance, 

Flame Middle). Figure 39 and Figure 40 provide similar charts but for Test 22 (3000 SLPM, 8 -ft 

Standoff Distance and Flame Middle Location). The temperature data provided in Table 6 and 

Table 7 were averaged over the last 30 seconds of the test (2-2.5 min), but for average heat flux 

data, these values were averaged between the final 60 to 120 s, since these values came to steady 

state much faster than the temperatures. This general approach was taken for all averaging 

calculations.

Figure 37. Heat flux time history for Test 15 (3000 SLPM, 4-ft standoff distance and flame 

middle location) 
37 

Figure 38. Temperature time history for Test 15 (3000 SLPM, 4-ft standoff distance and 

flame middle location) 

Figure 39. Heat flux time history for Test 22 (3000 SLPM, 8 -ft standoff distance and flame 

middle location) 
38 

Figure 40. Temperature time history for Test 22 (3000 SLPM, 8 -ft standoff distance and 

flame middle location) 

5.1.5 Discussion

Figure 41 provides a summary of the average heat flux data collected at a 4-ft standoff distance 

from the flame tip as a function of natural gas flow rate. Figure 42 provides a similar chart but at 

the flame middle position. 

The overall trend of increasing heat flux with increasing flow rate was observed for both series 

of tests. It can also be seen that the 1.5-ft height measurement points seem to have the most 

variation followed by the 7.5-ft height location. The most consistent trend was observed for the 

4.5-ft height which is likely due to the measurement point seeing the highest consistent 

percentage of flame during testing. Peak heat fluxes are slightly higher for the Flame Middle 

measurement position compared to the flame tip position which is likely due to the more 

intermittent exposure close to the flame tip as compared to the middle. 

Figure 43 and Figure 44 show similar comparative graphs for the 8-ft standoff distance. As 

expected, lower heat fluxes were observed compared to the 4-ft standoff distance. The heat 

fluxes measured at the 1.5-ft elevation were again the lowest as compared to the other heights; 

however, the increasing trend with increasing flow rate was more consistent for the 8-ft standoff 

distance compared to the 4-ft standoff distance. Again, this is likely due to a more even 

distribution of heat and less turbulence effect at the 8-ft distance compared to the 4-ft distance. 
39 

Figure 41. Heat flux trend for 4-ft standoff distance and flame tip location 

Figure 42. Heat flux trend for 4-ft standoff distance and flame middle location 
40 

Figure 43. Heat flux trend for 8-ft standoff distance and flame tip location

Figure 44. Heat flux trend for 8-ft standoff distance and flame middle location 
41 

Figure 45 shows a comparison between peak heat fluxes measured for each flow rate and each 

configuration. For the 4-ft standoff distance, the peak heat flux varies from 4-16 kW/m2 across 

the flow rate range. For the 8-ft standoff distance, the heat flux range is 1-6 kW/m2. 

Figure 45. Peak heat flux comparative chart 

At a 4-ft standoff distance from the flame all flow rates (500+ SLPM) generated incident heat 

flux that exceeded the capabilities (~1 kW/m2) of standard personal protective equipment (PPE) 

(Heus & Denhartog, 2017). For the 4-ft standoff distance incident heat flux at flow rates of 

1000+ SLPM exceeded the 3-min working time for firefighter PPE (4.6 kW/m2) and at flow rates 

of 1500+ SLMP exceeded the 5-min working time for aluminized PPE (6.3 kW/m2).

At 8-ft standoff, flow rates of 1000+ SLPM created heat flux that exceeded the ability of industry 

standard PPE to protect personnel for an extended period. Firefighter PPE could withstand the 

heat flux generated at 8 ft for flow rates of 2000 SLPM and below for at least 3 min. The 

capabilities of aluminized PPE are expected to provide sufficient protection at an 8-ft standoff 

distance for flow rates up to 3000 SLPM. 

For reference, the predicted total leak rate through closed fire damaged PRVs and other fire 

damaged valves is 6000 – 12,000 SLPM (FRA, 2023). Note that there are up to 20 valves in 

addition to the PRV(s) located in the piping cabinet, depending on the tank design, which are 

almost certain to fail after 10-20 minutes of fire exposure. Thus, the heat flux generated just from 

ignited natural gas leaking from valves could easily exceed the protective capabilities of 

emergency personnel PPE within an 8-ft radius.
42 

6. Structural Countermeasures

The structural FE model was used to evaluate select design features that could improve the 

crashworthiness of a UN-T75 tank by protecting the PRV piping assembly. Damage to piping 

that can restrict PRV flow was shown to occur at impact speeds as low at 10 mph (see Section 3). 

The countermeasures evaluated were limited to structural features that would protect the PRV 

system in its existing position. Changes in the layout of the PRV piping could provide additional 

benefits. 

6.1 Methodology

The UN-T75 tank FE model was modified to include additional structural members that would 

protect the PRV system from damage in derailment crash conditions. Ad hoc simulations were 

conducted to evaluate custom structural improvements that, based on engineering judgement, 

were likely to reduce the amount of damage sustained by the PRV assembly in two crash 

configurations: end-on impacts to the piping cabinet with a flat car and perpendicular impacts to 

the top end with a rigid pole. The rollover cases were not investigated due to the low risk of 

generating direct damage to the PRV piping assembly. 

Structural components were created in FE and added to the existing model. The crash conditions 

remained the same as described in Section 3. Two countermeasures were considered that 

represented modifications that could reasonably be added without considerably affecting the cost 

or weight of the tank system. A “crash cushion” style countermeasure (Figure 46) was created 

out of 4 x 6 x 1/4 inch square tubing using the same material properties as the tank’s existing 

frame structure. This assembly added 426 kg (939 lb) to the tank, which represents about 3.5 

percent of the tank’s empty weight. The second countermeasure involved increasing the wall 

thickness of the PRV piping assembly from 0.154 inch to 0.25 or 0.5 inch. 

Figure 46. “Crash Cushion” Countermeasure 
43 

6.2 Results

The “crash cushion” countermeasure was able to completely prevent pipe damage in a 15-mph 

end-on impact with a flat car. At 20 mph (1.78 times more energy) the pipe damage with the 

countermeasure was similar to the baseline case at 15 mph.

Figure 47. Maximum damage in 15 mph rear impact for baseline tank (top) and 

countermeasure (bottom)
44 

Figure 48. Maximum damage in 20 mph rear impact for baseline tank (top) and 

countermeasure (bottom)
45 

The thickened pipe countermeasures proved to be very effective in mitigating crush damage to 

the pipe (Figure 49). At 30 mph the baseline pipe was completely pinched closed, the 0.25-inch

thick pipes were >95 percent closed, and the 0.5-inch-thick pipes were 25 percent closed (Figure 

50). Similar results were observed for the internal pipes.

Figure 49. External PRV Piping Assembly after 30 mph impact with empty flatcar: 

Baseline Sch 40 pipe (left); 0.25” thick pipe (center); 0.5” thick pipe (right)

Figure 50. Maximum deformation of 0.25” (left) and 0.5” (right) thick external PRV pipes
46 

7. PRV Exhaust Flame Effects

Ignition of the PRV exhaust is highly likely for flammable fluids when exposed to nearby fire 

such as a pool fire. Due to the moderate mass flow rate through the PRV system under nominal 

conditions and the high mass flow rate through the PRV system under overturn/dual-phase 

conditions (see Section 4), the resulting jet flame can provide a significant heat load to 

neighboring tanks. 

7.1 Methodology

Fire Dynamics Simulation software (FDS, NIST) was used to simulate the subject UN-T75 tank 

and fire conditions. The 3-dimensional tank model included the external tank geometry and a 

representation of a flat car on which it was positioned (Figure 51). The steel materials were 

modeled with appropriate thermal properties. 

Figure 51. Tank and Flat Car Model within the FDS environment

Combinations of two fire loads were applied to the tank/flat car model: a pool fire and a jet 

flame. The pool fire was meant to represent the same size and heat release rate as the fire used in 

the FRA fire tests (Friedman & Mattos, 2018; FRA, 2023). The output data of interest for each of 

the models was the total net steady-state heating to the tank (discussed in the Results section). 

The pool fire was 40 ft x 13 ft (12.19 m x 3.96 m) and provided a heat release rate per unit area 

(HRRPUA) of either 4564 kW/m2 (High) or 2282 kW/m2 (Low). The pool fire was modeled as a 

propane-fed area with a mass flow rate defined to achieve the desired HRRPUA. For reference, 

the target HRRPUA of the FRA fire tests was 4564 kW/m2. 

The jet fire model was meant to replicate the highest (500 lb/min) and lowest (150 lb/min) mass 

flow conditions determined through physical testing in this project (see Section 4). The jet fire 

was either directed at the end of the tank and parallel with its axis (i.e., END) or perpendicular to 

the center of the tank (i.e., SIDE). The jet fire was positioned either 1.1 m (i.e., CLOSE; 

minimum nominal distance between neighboring flat cars), 6.1 m (i.e., MID), or 11.1 m (i.e., 

FAR) from the tank. In each case the jet was positioned 0.1 m above the ground to mimic its 

position on an overturned tank (see Figure 52 for a view of the model setup). The exhaust jet was 

composed of “cryogenic” liquid methane droplets defined with a set of parameters (i.e., spray 

angle, evaporation rate, particle size, and initial velocity) that were calibrated against large-scale, 

high-pressure jet fire test data (Lowesmith & Hankinson, 2012). The thermal properties of the 

droplets and resulting gas were defined by the NIST-provided library within the FDS solver. 
47 

Figure 52. Example visualizations of the FAR jet fire position in the SIDE (left) and END 

(right) configurations without a pool fire 

The matrix of 28 simulation conditions is summarized in Table 8. 

Table 8. Fire Impingement Simulation Matrix

Pool Fire HRRPUA Jet Fire Mass Flow Rate Jet Direction Jet Distance

LOW HIGH END CLOSE

LOW LOW END CLOSE

- HIGH END CLOSE

- LOW END CLOSE

LOW HIGH SIDE CLOSE

LOW LOW SIDE CLOSE

HIGH HIGH SIDE CLOSE

HIGH LOW SIDE CLOSE

No HIGH SIDE CLOSE

No LOW SIDE CLOSE

LOW HIGH END FAR

LOW LOW END FAR

HIGH HIGH END FAR

HIGH LOW END FAR

No HIGH END FAR

No LOW END FAR

LOW HIGH SIDE FAR

LOW LOW SIDE FAR

- HIGH SIDE FAR

- LOW SIDE FAR

LOW HIGH END MID

- HIGH END MID

- LOW END MID

LOW HIGH SIDE MID

- HIGH SIDE MID

- LOW SIDE MID

BASELINE

LOW - - -

HIGH - - -

Pool fire HRRPUA: LOW = 2282 kW/m2; HIGH = 4654 kW/m2 

Jet Fire Mass Flow Rate: LOW = 1.134 kg/s; HIGH = 3.74 kg/s
48 

7.2 Results

In the baseline (pool fire only) simulations the steady-state heating rate across the entire tank was 

4120 kW for the LOW case and 5100 kW for the HIGH case. The addition of a jet fire had 

mixed results in terms of altering the total heating rate of the tank, which can be seen in Figure 

53. 

Up arrow represents the “HIGH” condition and down arrow represents the “LOW” condition 

Figure 53. Summary of heating rates for the simulated fire conditions 

In the LOW pool fire conditions, the addition of the HIGH jet flame marginally increased the 

overall heating rate for some conditions (END_CLOSE, SIDE_CLOSE, SIDE_FAR) while it 

decreased the heating rate in others (END_MID, SIDE_MID, END_FAR). For the combined 

LOW pool fire and LOW jet fire condition all heating rates were equal to or lower than the LOW 

pool fire only condition. 

For the HIGH pool fire condition, all the jet fire configurations resulted in a lower heating rate 

than the baseline case. 

In the jet fire only conditions, the heating rate for most of the HIGH jet conditions was roughly 

half that of the pool fire conditions. The LOW jet fire only cases resulted in heating rates 

approximately 25 percent of the pool fire only cases. The HIGH jet configurations always 

produced greater heating than the LOW jet configuration. In the HIGH jet fire only 

configurations the 1 m standoff distance produced less heating than the 6.1 and 11.1 m distances. 

The jet flame increased the heating on the tank where it directly impinged on it. An example of 

this is demonstrated by comparing Figure 54 and Figure 55. The pool fire alone did not 

appreciably apply direct heat to the tank ends; however, the jet fire directed at the end of the tank 

resulted in an average incident heat flux of approximately 60-100 kW/m2

 and peaking at ~125 

kW/ m2

. Some of the peak heating rates appeared to be offset by the cooling effect of the 

cryogenic liquid methane drops that moved along with the momentum-driven flow. 
49 

Figure 54. Incident heat flux on tank exterior in the HIGH pool fire only configuration 

Figure 55. Incident heat flux on tank exterior in the HIGH pool fire + HIGH jet fire END 

configuration 
50 

7.3 Discussion

It is not surprising that the magnitude of total tank heating rate was relatively unchanged in the 

combined pool fire + jet fire cases. The size of the pool fire itself is sufficient to considerably 

affect the amount of available oxygen around the tank. The additional fuel provided by the 

natural gas jet resulted in a fuel-rich condition with more fuel available than could be burned. 

This is also exemplified in the marginal heating rate increase between the two LOW and HIGH 

Baseline conditions in which the fuel rate was doubled but the heating rate only increased by 24 

percent. This effect can be visualized by the oxygen molar fraction shown along the centerline of 

the model in Figure 56 and Figure 57. Additionally, the strong flow of liquid and gas in the 

exhaust jet resulted in an effective ‘wind’ on the order of 3 to 5 m/s around the tank and up to 30 

m/s below the railcar. The combination of the added jet flame, the richness in the fuel mixture, 

and the added ‘wind’ effectively cancelled out any added heating by the jet itself. 

Figure 56. Molar fraction of oxygen in a LOW, END, CLOSE, Jet configuration (no pool 

fire)

Figure 57. Molar fraction of oxygen in a LOW, END, CLOSE, Jet configuration (LOW 

pool fire)
51 

The localized heating from the jet flames in this study (peak = ~125 kW/m2) was substantially 

less than what was observed from similar simulations of jet flames formed from single (gas) 

phase flow of methane out of the PRVs at the nominal flow rates (peak = 178-292 kW/m2

) 

(Mattos & Friedman, 2023). These differences in peak incident heat flux can be explained by the 

differences in the test conditions. This study simulated the flow of cryogenic liquid methane out 

of a 50 mm pipe orifice at ground level between 1.1 and 11.1 meters from the target tank. The 

previous study simulated the flow of pure methane gas out of a 50 mm pipe at 0.5 m above the 

tank centerline between 1 and 5 m from the target tank. Both the location of the jet flame relative 

to the target tank and the use of cryogenic liquid, which was used to represent the dual phase 

flow observed in the test, likely resulted in the relatively lower incident heating rates. 

Nonetheless, the heat flux provided by the jet flame would be sufficient to thermally fail the 

seals in all the valves of the piping cabinet, included any PRVs, within 10 minutes of exposure or 

less. 
52 

8. Suggested Future Work

The results of the current and previous studies have highlighted knowledge gaps and areas for 

future work. Suggestions for future work are summarized below.

• Use multiphysics modeling to evaluate the crashworthiness and fireworthiness of tanks 

transported within well-cars. Well-cars typically have an open bottom, which may 

increase the thermal load applied from a pool fire compared to tanks transported on flat 

cars. Additionally, the raised walls of a well-car pose a potential impact surface that may 

directly damage external piping or the tank itself. 

• Use the jet fire test data in this study to create validated FDS models. 

• Extend the physical jet fire testing demonstrated in this report to higher flow rates, 

including dual-phase flows.

• Characterize the jet flame impingement on a cryogenic tank and/or external 

valves/piping. Localized heating from a jet fire is likely to degrade valves, the external 

tank, and annular space insulation much more quickly and severely than a pool fire. 

• Evaluate the effects of non-methane components (e.g., ethane, toluene, benzene, etc.) 

within the LNG mixture during heating and PRV exhaust. 

• Conduct full-scale or component-level physical testing to demonstrate countermeasures 

for protecting the PRV piping assembly from damage in a derailment condition. 
53 

9. Conclusion

The literature review, physical testing, and physics-based simulations conducted in this project 

provided novel data and insight into the thermal performance of cryogenic tanks subjected to fire 

conditions. The findings are summarized below. 

• PRVs that meet the regulatory requirements for UN portable tanks and are typical of 

devices found on UN-T75 style tanks, while not specifically designed to exhaust dualphase cryogenic fluid, were subjected to dual-phase flow of cryogenic liquid nitrogen at 

pressures ranging from 130 to 180 psi. Time histories of PRV temperature, mass flow 

rate, fluid density, and fluid pressure were recorded. Testing was very repeatable.

• At tank pressures in the range of those expected to exist in fire-exposed conditions, the 

dual phase mass flow rate through the PRVs was similar to the measured flow rates of 

high pressure (870 psi) natural gas through a 20 mm diameter orifice in the literature. 

• The tested PRVs were found to provide flow rates that exceeded those required to 

maintain a safe working pressure under fire exposure when tested with liquid nitrogen in 

short duration tests. The dual-phase mass flow rate of liquid nitrogen through the PRVs 

was measured to be between 1.134 kg/s and 3.74 kg/s between tank pressures of 130 and 

180 psi. This is 2.8 to 9.4 times greater than the single (gas) phase flow rate for the same 

PRVs. The dual-phase mass loss rate measured in this study is similar to the mass loss 

rate of natural gas through a 20 mm orifice in an 856 psi natural gas pipeline (~2.9 kg/s). 

• A constriction representing 90 percent closure of the discharge pipe reduced the dualphase mass flow rate out of the PRV by approximately 38 percent. The constriction also 

produced an oscillation in the flow which, if allowed to continue, could damage the PRV 

seal, springs, or other components through fluttering. 

• The dual-phase PRV exhaust plume extended from 60 to 100 ft in length, and ignition of 

this plume could provide considerable heating to nearby tanks. 

• All jet fire conditions evaluated in the study present a significant threat to emergency 

responders. 

• A significant increase in flow rate was observed for the liquid nitrogen in the subcooled 

condition. This means that a freshly filled LNG tank would likely empty faster than a 

tank that has had time to heat up during its travel. 

• The risk of a PRV “icing up” resulting from ambient moisture in humid conditions 

encountering the cryogenic piping assembly is predicted to be negligible when a 

sufficient length of pipe is attached to the PRV outlet. 

• Natural gas jet fires were characterized across a range of low mass flow rates, orifice 

sizes, and orientations. These data can be used in the future to validate simulations of fire 

behavior and may be applied to tanks. 

• Two simple structural countermeasures were demonstrated that could reduce the 

likelihood of PRV assembly damage and the associated risk of reduced mass loss rate. In 

some scenarios these countermeasures increased the damage-inducing speed from 10 

mph to over 30 mph. 
54 

• Derailment and/or overturning of a cryogenic tank can result in the PRV assembly no 

longer meeting the performance requirements described in the regulations (49 CFR 

178.274). These include restricted downstream flow, PRV inlet in the liquid space rather 

than the vapor space, and escaping vapor directed at the tank shell or neighboring tank 

shell. The impact of fire will likely prevent the PRV from closing after the pressure 

returns to below the set pressure. 
55 

10. References

Baker Hughes (n.d.). Consolidated 1900/1900 DM Series. 

Federal Railroad Administration (2023). Fire Test of An UN-T75 Portable Tank on A Flat CarPhase II (Report No. RR 23-01).

Friedman, K., & Mattos, G. (2018). Analysis of Liquid Nitrogen Test Data, Transport Canada, 

T8080-170208. 

Heus, R. & Denhartog, E. (2017). Maximum allowable exposure to different heat radiation levels 

in three types of heat protective clothing. Indus. Health, 55, 529-536. 

Lowesmith, B., & Hankinson, G. (2012). Large scale high pressure jet fires involving natural gas 

and natural gas/hydrogen mixtures. Process Safety and Environmental Protection, 90, 

108-120. 

Mattos, G. & Friedman, K. (2023). Final Report: Modelling of ISO [UN-T75] Portable Tank Fire 

Test, Transportation of Dangerous Goods. Transport Canada. 

Prabhakaran, A., et al. (2022). Performance of Tank Car Pressure Relief Devices Under 

Derailment Fire Conditions (Report No. DTFR53-17-C-00015).

Rockwood Swendeman (n.d.). Bronze Safety Relief Valves, Type RXSO. Circle Cryogenics.
56 

Appendix A

PRV Test System Setup
57 

Appendix B

Test Section Fabrication and Setup Drawings
58 
59 

Abbreviations and Acronyms

ACRONYM DEFINITION

ASME American Society of Mechanical Engineers 

BLEVE Boiling Liquid Expanding Vapor Explosion

FDS Fire Dynamics Simulator

FE Finite Element

FRA Federal Railroad Administration

HF Heat Flux

HRR Heat Release Rate

HRRPUA Heat Release Rate Per Unit Area

ISO International Organization for Standardization

LN2 Liquefied Nitrogen

LNG Liquefied Natural Gas

MAWP maximum allowable working pressure

NIST National Institute of Standards and Technology

P&ID piping and instrumentation diagram

PHMSA Pipeline and Hazardous Materials Safety Administration

PPE Personal Protective Equipment

PRV Pressure Relief Valve

PTFE Polytetrafluoroethylene

SCMS Standard Cubic Meter Per Second

SwRI Southwest Research Institute

TC Thermocouple