NHTSA Hydrogen Fuel System Research - 812133 ... Energy... · 2.5 Conclusion ... Four fixed orifices established the leak rates. Nominal orifice sizes were selected to provide the

May

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Technical Report Documentation Page

1. Report No. 2. Government Accession No. 3. Recipient’s Catalog No.

DOT HS 812 133

4. Title and Subtitle 5. Report Date Compressed Hydrogen Container Fueling Options for Crash Testing May 2015

6. Performing Organization Code

7. Author(s) 8. Performing Organization Report No.

Doug Pape, Andrew Cox 9. Performing Organization Name and Address 10. Work Unit No. (TRAIS)

Battelle Memorial Institute 505 King Avenue 11. Contract or Grant No.

Columbus, OH 43201 DTNH22-8-D-00080

12. Sponsoring Agency Name and Address 13. Type of Report and Period

National Highway Traffic Safety Administration Office of Vehicle Safety Research

Final Report

1200 New Jersey Avenue SE. 14. Sponsoring Agency Code

Washington, DC 20590 15. Supplementary Notes Barbara Hennessey was the Contracting Officer’s Technical Representative (COTR) for this project.

16. Abstract This project involved three series of experiments to assess the post-crash fuel system integrity test procedures specified in SAE 2578, Recommended Practice for General Fuel Cell Vehicle Safety. The first test series compared cylinder vulnerability to axial and lateral impact at high and low states of fill to determine the worst case condition for leak or rupture. The second test series assessed the use of helium as a non-flammable surrogate fill gas for crash testing. In the third series of tests, three Honda CNG Civics were retrofitted with hydrogen containers and subjected to front, side and rear crashes to obtain loading, displacement, and intrusion data on fuel system components, and to develop test procedures to remotely defuel high pressure systems after testing, for the safety of laboratory personnel approaching the vehicles post-crash.

17. Key Words 18. Distribution Statement

Hydrogen safety , fuel cell vehicle, fuel system integrity, alternative fuels

Document is available to the public from the National Technical Information Service www.ntis.gov

19. Security Classif. (of this report) 20. Security Classif. (of this page) 21. No. of Pages 22. Price

Unclassified 118

Form DOT F 1700.7 (8-72) Reproduction of completed page authorized

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TABLE OF CONTENTS

PAGE

EXECUTIVE SUMMARY .......................................................................................................... ix Dynamic Impact Crush Tests of Fuel Containers .............................................................. ix Leak Rate Measurements and Characterization ...................................................................x Crash Tests of Mockup Hydrogen Vehicles ..................................................................... xii Related Task Orders ......................................................................................................... xiii

1.0 INTRODUCTION..............................................................................................................1 1.1 Dynamic Impact Crush Tests of Fuel Containers ....................................................2 1.2 Leak Rate Measurements .........................................................................................2 1.3 Crash Tests of Mockup Hydrogen Vehicles ............................................................3

2.0 DYNAMIC IMPACT CRUSH TESTS OF FUEL CONTAINERS ..............................4 2.1 Summary of the Dynamic Impact Tests...................................................................4 2.2 Introduction and Background ..................................................................................5

2.2.1 Test Objective ..............................................................................................5 2.2.2 Construction of Hydrogen Containers .........................................................5

2.3 Test Conditions ........................................................................................................7 2.3.1 Selection of Containers for the Test Program ..............................................7 2.3.2 Selection of Impact Energy ..........................................................................8

2.3.2.1 Analysis of Prior FMVSS 301 Compliance Tests ........................ 8 2.3.2.2 Comparison With Prior Container Impact Research .................. 10 2.3.2.3 Decision for the Drop Energy ..................................................... 12

2.3.3 Test Matrix .................................................................................................12 2.3.4 Test Apparatus ...........................................................................................14 2.3.5 Test Procedure ...........................................................................................17

2.4 Results ....................................................................................................................17 2.4.1 Horizontal Impact ......................................................................................18 2.4.2 Vertical Impact...........................................................................................22

2.5 Conclusion .............................................................................................................25

3.0 LEAK RATE MEASUREMENTS AND CHARACTERIZATION ...........................27 3.1 Summary of the Leak Rate Characterization .........................................................27 3.2 Introduction and Background ................................................................................27

3.2.1 Test Objective ............................................................................................28 3.2.2 Key Assumptions of SAE J2578, Appendix A ..........................................29

3.3 Test Conditions ......................................................................................................32 3.3.1 Technical Approach ...................................................................................32 3.3.2 Test Matrix .................................................................................................35 3.3.3 Instruments .................................................................................................36

3.4 Results ....................................................................................................................37 3.4.1 Raw Data ....................................................................................................37 3.4.2 Initial Flow Rates .......................................................................................42 3.4.3 Mass Estimation from Pressure and Temperature .....................................44

3.4.3.1 Comparison of Mass Flow Meter and Gas Density Change ...... 45

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3.4.3.2 Mass Loss Estimated by a Purely Empirical Formula.................47 3.4.3.3 Mass Loss by Formulas in Appendix A of SAE J2578 ...............49 3.4.3.4 Mass Loss Estimated by Japanese Blue Book Formulas.............52

3.5 Conclusion .............................................................................................................. 54

4.0 CRASH TESTS OF MOCKUP HYDROGEN VEHICLES ........................................ 55 4.1 Summary of the Crash Tests .................................................................................. 55 4.2 Introduction and Background ................................................................................. 56

4.2.1 Test Objective............................................................................................. 56 4.2.2 Context of the Crash Tests ......................................................................... 56

4.3 Test Conditions ....................................................................................................... 57 4.3.1 Technical Approach ................................................................................... 57 4.3.2 Test Matrix ................................................................................................. 57 4.3.3 Test Vehicles .............................................................................................. 61

4.3.3.1 Physical Description ....................................................................61 4.3.3.2 Fuel System ..................................................................................62

4.3.4 Facilities and Equipment ............................................................................ 65 4.3.4.1 Test Site .......................................................................................65 4.3.4.2 Instruments ...................................................................................65 4.3.4.3 Cameras ........................................................................................68

4.4 Results .................................................................................................................... 68 4.4.1 Qualitative Description of Damage to the Fuel Systems ........................... 68

4.4.1.1 Front Crash..................................................................................69 4.4.1.2 Rear Crash ...................................................................................69 4.4.1.3 Side Crash ...................................................................................72

4.4.2 Transient Motion During the Crashes ........................................................ 76 4.4.2.1 Front Crash..................................................................................76 4.4.2.2 Rear Crash ....................................................................................79 4.4.2.3 Side Crash ...................................................................................82

4.4.3 Pressure Integrity ........................................................................................ 85 4.4.4 Damage to the Containers .......................................................................... 86

4.5 Conclusion .............................................................................................................. 86

5.0 OVERALL CONCLUSIONS .......................................................................................... 88 5.1 Crash Test Conditions ............................................................................................ 88 5.2 Fuel Options ........................................................................................................... 89 5.3 Crash Procedure ..................................................................................................... 89 5.4 Further Work .......................................................................................................... 90

REFERENCES ............................................................................................................................. 91

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LIST OF APPENDICES

Appendix A: Photographs of the Crush Test ............................................................ A-1

Appendix B: Load Time Histor ies From the Crush Test .......................................... B-1

Appendix C: Leak Rate Test Plan .............................................................................. C-1

Appendix D: Leak Rate Time Histor ies ..................................................................... D-1

Appendix E: Full-Vehicle Crash Test Plan and Safety Plan ..................................... E-2

Appendix F: TTI Crash Test Repor ts ......................................................................... F-1

Appendix G: Time Histor ies of the Crash Tests ....................................................... G-1

Appendix H: Crash Sensor Specifications ................................................................. H-1

LIST OF TABLES

Table 2-1. Energy dissipated in NHTSA FMVSS 301 compliance tests......................................... 9 Table 2-2. Peak forces in NHTSA FMVSS 301 compliance tests. ................................................ 10 Table 2-3. Comparison of the high-level results of the prior impact tests with the current

test plan. .................................................................................................................... 11 Table 2-4. Test conditions for crush test. ....................................................................................... 13 Table 2-5. Specifications of containers used for the crush and leak tests. ..................................... 16 Table 2-6. Specifications of instrumentation and DAQ. ................................................................ 17 Table 2-7. Results from crush test by test condition. ..................................................................... 18 Table 3-1. Four fixed orifices established the leak rates. Nominal orifice sizes were

selected to provide the flow rates in the light blue cells. Flow rates in white cells were predicted from nominal conditions. Actual orifice diameters were measured with a microscope. Actual initial flow rates are discussed in Section 3.4.2. ............................................................................................................ 34

Table 4-1. Conditions and camera locations. ................................................................................. 59 Table 4-2. Weights of the vehicles in the crash tests. .................................................................... 61 Table 4-3. Fuel container comparison. ........................................................................................... 62 Table 4-4. Planned and actual conditions at impact. ...................................................................... 68 Table 4-5. Distances above the ground where the crush profile was measured. ........................... 73 Table 4-6. Exterior crush profile of the side impact vehicle. ......................................................... 73 Table 4-7. Density ratios from the hydrogen crash tests. ............................................................... 85 Table 4-8. Density ratios from the nitrogen pressurization tests. .................................................. 86

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LIST OF FIGURES

Figure ES-1. Stills from the high-speed video of a crush test. . ....................................................... x Figure ES-2. The amount of hydrogen leakage was predicted via the two formulas for

surrogate conditions in Appendix A of SAE J2578. . . . ........................................... xi Figure ES-3. This is the same data as in the pevious figure, but the formulas in Appendix

A of SAE J2578 have been adjusted for changes in compressibility of the gas. .......................................................................................................................... xii

Figure ES-4. In the rear crash, the moving deformable barrier passed through the crumple zone in the trunk and directly impacted the hydrogen fuel container. . ................. xiii

Figure 2-1. Cross-section of a typical Dynetek Industries Ltd. Type 3 container. .......................... 6 Figure 2-2. Cross-section of a typical Lincoln Composites Inc. Type 4 container. ........................ 6 Figure 2-3. Schematic of the drop test apparatus. .......................................................................... 14 Figure 2-4. Photograph of the drop test apparatus. ........................................................................ 15 Figure 2-5. Photograph of the drop weight. ................................................................................... 15 Figure 2-6. Stills from the high-speed video of Test #5, horizontal impact on a 350-bar

Type 4 container filled to 100% service pressure. ................................................... 19 Figure 2-7. Force time histories for Tests #1, #2, and #3 of the Type 3 containers. . . ................. 20 Figure 2-8. Force time histories for Tests #4, #5, and #6 of the Type 4 containers. . ................... 21 Figure 2-9. Stills from the high-speed video of Test #9, vertical impact on a 700-bar Type

3 container filled to 10% service pressure. ............................................................... 23 Figure 2-10. Time history of the load for Test #8, in which the container ruptured. .................... 25 Figure 3-1. Appendix A of SAE J2578 provides formulas for estimating the mass of

leaked hydrogen in three possible fueling conditions for crash tests. . .................... 28 Figure 3-2. The approximation for the density of high-pressure hydrogen at 15° C in SAE

J2578 agrees well with NIST values. ....................................................................... 30 Figure 3-3. The approximation for the density of high-pressure helium at 15° C in SAE

J2578 agrees well with NIST values. ....................................................................... 30 Figure 3-4. The approximation for the density of low-pressure hydrogen at 15° C in SAE

J2578 agrees well with NIST values. ....................................................................... 31 Figure 3-5. The approximation for the density of helium in the Japanese standard agrees

well with NIST values. ............................................................................................. 31 Figure 3-6. Scanning electron microscope images of the (1) small, (2) medium, (3) large,

and (4) extra large orifices. ....................................................................................... 35 Figure 3-7. Temperature and pressure measuring instruments immediately upstream and

downstream of the orifice. ........................................................................................ 37 Figure 3-8. Mass flow rate and container pressure (hydrogen, Type 3 container, 350 bar,

135 µ L orifice). ........................................................................................................ 38 Figure 3-9. The hydrogen in the container cooled during the leak of the case in Figure 3-

8. ................................................................................................................................ 38 Figure 3-10. Mass flow rate and container pressure (hydrogen, Type 4 container, 70 bar,

22 µ S Orifice). ......................................................................................................... 39 Figure 3-11. Mass flow rate and container pressure (hydrogen, type 3 container, 35 bar,

135 µ L Orifice). ....................................................................................................... 39

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Figure 3-12. Mass flow rate and container pressure (hydrogen, type 3 container, 70 bar, 95 µ M Orifice). Flow was switched from the Coriolis flow meter to the Laminar flow meter about 100 min into the leak. .................................................... 40

Figure 3-13. Mass flow rate and container pressure (hydrogen, Type 4 container, 350 bar, 192 µ XL Orifice). .................................................................................................... 40

Figure 3-14. Mass in containers calculated from temperature and pressure measurements. ......... 42 Figure 3-15. Measured and predicted initial flow rates of the hydrogen experiments. ................. 43 Figure 3-16. Measured and predicted initial flow rates of the helium experiments. ..................... 43 Figure 3-17. Discharge coefficients, calculated using the initial flow rates of the hydrogen

experiments. .............................................................................................................. 44 Figure 3-18. Discharge coefficients, calculated using the initial flow rates of the helium

experiments. .............................................................................................................. 44 Figure 3-19. Comparison of mass leak calculated from flow meter and measurements of

pressure and temperature for high-pressure experiments with M, L, and XL orifices....................................................................................................................... 46

Figure 3-20. Comparison of mass leak calculated from flow meter and measurements of pressure and temperature for low-pressure experiments with M, L, and XL orifices....................................................................................................................... 46

Figure 3-21. Comparison of mass leak calculated from flow meter and measurements of pressure and temperature for experiments with the 22-µ S orifice. ......................... 47

Figure 3-22. Predicted leaked mass of high-pressure hydrogen using Equation 2. ....................... 48 Figure 3-23. Comparison of amount of hydrogen leaked when a surrogate gas predicted

non-conformance to the non-conformance criterion. ............................................... 49 Figure 3-24. The amount of hydrogen leakage was predicted via the two formulas for

surrogate conditions in Appendix A of SAE J2578. . . . .......................................... 50 Figure 3-25. This is the same data as in the pevious figure, but the formulas in Appendix

A of SAE J2578 have been adjusted for changes in compressibility of the gas. . .......................................................................................................................... 51

Figure 3-26. Comparison of predicted and measured hydrogen mass flow rates for high pressure experiments with M, L, and XL orifices. ................................................... 53

Figure 3-27. Comparison of predicted and measured hydrogen mass flow rates for low pressure experiments with M, L, and XL orifices. ................................................... 53

Figure 3-28. Comparison of predicted and measured hydrogen mass flow rates for experiments with S orifices. ..................................................................................... 53

Figure 4-1. The impact area for the side crash was to the rear of the area for a compliance test so the panel adjacent to the fuel container would be directly struck. . .............. 60

Figure 4-2. The rear of the original fuel container, with the trunk panel removed. ...................... 63 Figure 4-3. The hydrogen container installed in the trunk of a test vehicle. ................................. 63 Figure 4-4. The front of the hydrogen container, showing the modification to the

mounting frame. ........................................................................................................ 64 Figure 4-5. The crashes were at the south end of TTI’s Riverside Campus. ................................. 66 Figure 4-6. The mounting configuration for the string potentiometers in the front crash

test vehicle. ............................................................................................................... 67 Figure 4-7. The moving deformable barrier in the rear crash pushed the trunk lid to the

fuel container. .......................................................................................................... 70

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Figure 4-8. In this view through the driver-side rear door, the yellow box shows where the anchor for the container mounting failed during the rear crash. ........................ 71

Figure 4-9. The impact of the rear crash pushed the fuel container into the rear seat area. . ................................................................................................................................... 72

Figure 4-10. The vertical lines drawn on the vehicle following the side crash indicate the longitudinal positions where the deformations were measured. .............................. 74

Figure 4-11. The maximum deformation in the side crash was in the panel of the rear door. .......................................................................................................................... 75

Figure 4-12. This photograph of the vehicle following the side crash shows the penetratation caused by the impact. The outward displacement of the rear door is visible. ........................................................................................................... 75

Figure 4-13. Displacement of the end of the container on the driver’s side during the front crash. ......................................................................................................................... 77

Figure 4-14. Vehicle body acceleration during the front crash. ..................................................... 78 Figure 4-15. Container longitudinal accelerations superimposed over the vehicle body

acceleration during the front crash. .......................................................................... 78 Figure 4-16. Container vertical accelerations superimposed over the vehicle body vertical

acceleration during the front crash. .......................................................................... 79 Figure 4-17. Displacement of the passenger side end of the container during the rear

crash. ......................................................................................................................... 80 Figure 4-18. Vehicle body acceleration due to the rear impact from the MDB. ........................... 81 Figure 4-19. Longitudinal accelerations of the container superimposed over the vehicle

body acceleration during the rear crash. ................................................................... 82 Figure 4-20. Passenger side container displacement relative to the vehicle body during

the side crash. ............................................................................................................ 83 Figure 4-21. Vehicle rear seat X-Y resultant acceleration during the side crash........................... 84 Figure 4-22. Container X-Y resultant accelerations superimposed over vehicle rear seat

resultant acceleration during the side crash. ............................................................. 84

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EXECUTIVE SUMMARY

Hydrogen fueled motor vehicles offer many advantages in terms of pollution and efficiency over gasoline fueled vehicles. At the same time, they pose hazards that are not necessarily more severe but are certainly different from those in vehicles with gasoline fuel. The National Highway Traffic Safety Administration has contracted with a team led by Battelle to generate technical data for NHTSA to consider should the agency write future safety performance standards for hydrogen fuel vehicles.

The tests in this task order were conducted with the goal of enabling NHTSA to resolve some key questions for specifying the conditions and pass/fail criteria for hydrogen vehicle crash tests. This effort had previously been affected by the limited availability and expense of hydrogen vehicles, preventing their use in fuel system integrity crash and leakage tests.

Three separate technical tasks were conducted to assess the post-crash fuel system integrity test procedures specified in SAE J2578, “Recommended Practices for General Fuel Cell Vehicle Safety.” First, drop weight crush tests with static containers assessed their vulnerability to impact damage in a sequence of conditions. A full internal pressure stiffens a container, thereby limiting its deformation on impact. A container struck on its side flexes more if it has a lower pressure, presumably causing more latent damage to the wall. A container struck on its end and constrained on its other end, which is an unlikely crash condition, will leak or rupture if the energy is sufficient. The second series of experiments related a container’s pressure drop over time to its mass leak rate, so that a pass/fail criterion can be specified in terms of an easily measured pressure drop, following the pattern of FMVSS No. 303. The data generally confirmed the procedure in SAE J2578 for experiments run with high-pressure hydrogen. Bias, evident in the formula for the equivalence of hydrogen and helium, could be mostly removed by accounting for more thermodynamic properties than the standard currently does. Measurement errors prevent meaningful mass estimates at leak quantities of a few grams, below those anticipated by the standard but near what another task order found to be potentially harmful. The final technical task was a set of three full-vehicle crash tests with a mockup hydrogen fuel system. These crashes demonstrated how a hydrogen fueled vehicle might behave in FMVSS No. 301 front, side, and rear crash tests. Damage to the host passenger cars was typical of FMVSS No. 301 crashes. The fuel container in one of the crashes sustained superficial damage. Though the stainless steel tubing was deformed substantially by the impact, it held the pressure. Damage to the mockup hydrogen fuel system was documented so that it can serve as a baseline of comparison in future crash or sled tests.

Dynamic Impact Crush Tests of Fuel Containers

A dead weight was dropped onto hydrogen containers. These impacts were quite severe, as though the entire energy dissipated in a vehicle crash test were applied directly to the container.

Containers that were struck on their sides withstood the impacts without leaking. Those that were pressurized to only 10 percent of their service pressure deformed considerably in the process. Containers that were struck on their ends developed severe leaks if they were pressurized to 10 percent of their service pressure, and with one exception, ruptured if they were fully pressurized.

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A crash test of a hydrogen fuel system would be performed to probe the most vulnerable condition of the system. Presumably, if the system passes such a test, then it would be robust in other conditions. The worst pressure condition for crash testing hydrogen fuel systems is not necessarily the maximum pressure. Composite fuel containers are damaged when they are forced to flex and deform by external impact. High pressures tend to stiffen fuel containers and piping, preventing them from deforming, thereby reducing the induced damage. Consequently, engineering experience and previous studies suggest fuel containers and fuel system components are more likely to leak in a crash if they are at low pressures, than at high pressures. Furthermore, 350 bar (5,000 psi) fuel containers have thinner walls and will flex more than 700 bar (10,000 psi) fuel containers under the same external loading, making them more likely to be damaged in a crash. These greater strains would be expected to leave more residual damage. Quantifying the residual damage to the containers in this experiment would require that the containers be sectioned for metallographic inspection or that they be subjected to pressure testing, either burst or cyclical.

The level of damage depends upon the orientation of impact on the container, whether axial or transverse. All of the container failures in this experiment resulted from the vertical (axial) impacts. The more highly pressurized containers failed by rupture, which is a more severe outcome than the leakage of the low-pressure containers. This is in contrast to the horizontal impacts, where the more severe outcome was presumed to be the greater flexing of the less pressurized containers. A tentative conclusion to draw from these observations is that high internal pressure protects a container from moderate impacts but can exacerbate the failure mode if the container does fail. Testing this hypothesis requires further axial-impact experiments to determine if there is an intermediate energy where a container with low pressure sustains more damage than one with high pressure. As most pressurized fuel containers in passenger cars are mounted transversely, axial impact corresponds to a side-impact crash. The failures in this experiment occurred when the container was between rigid surfaces on both ends, which is more severe than realistic crash conditions.

1 2 3 1. Drop weight released and coming down, prior to impact

2. Moment of maximum compression

3. First bounce of drop weight off the container and container off the load cell

Figure ES-1. Stills from the high-speed video of a crush test. The container’s bouncing shows that it was able to recover resiliently after absorbing the energy of impact.

Leak Rate Measurements and Characterization

A full-matrix test of two container constructions (Type 3 and Type 4), two wall thicknesses (service pressure of 350 and 700 bar), two initial pressurizations (100 and 10 percent of service

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pressure), and two gases (hydrogen and helium) was conducted. Each of these combinations was repeated three times, with the gas leaking through a fixed orifice to provide a “small,” “medium,” or “ large” leak rate.

Appendix A of SAE J2578 has three formulas for estimating an amount of hydrogen leakage from a container using pressure and temperature measurements. The first formula, a direct measure for crash tests with high-pressure hydrogen, and well predicted the experimentally measured values. Two other formulas in the standard are for surrogate conditions—low-pressure instead of high-pressure hydrogen and high-pressure helium instead of hydrogen. Both of these formulas were found to have a bias error. The formula for low-pressure hydrogen can be corrected by accounting for an extra effect in the theory. Accounting for this same effect removes some, but not all, of the bias error from the formula for high-pressure helium.

The first figure below shows how well the formulas for the two surrogate conditions predicted the mass of hydrogen that actually leaked in the corresponding high-pressure hydrogen case. The percent errors are not large, but a bias is evident. At the extreme pressures of these experiments, hydrogen and helium do not behave as ideal gases—their properties vary with increasing pressure. The spread of points is much reduced in the second figure, where the formulas in SAE J2578 have been adjusted to account for how the compressibility of hydrogen and helium change as the pressure rises.

Figure ES-2. The amount of hydrogen leakage was predicted via the two formulas for surrogate conditions in Appendix A of SAE J2578. This amount is compared with the loss in the corresponding experiment with high-pressure hydrogen. Predictions from high-pressure helium are within about 10 percent, but they are all above the Y=X line. Predictions from low-pressure hydrogen are within about 20 percent, all below the Y=X line.

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Figure ES-3. This is the same data as in the pevious figure, but the formulas in Appendix A of SAE J2578 have been adjusted for changes in compressibility of the gas. Both predictions have been improved, but evidence of a small bias in the formula for high-pressure helium remains.

Crash Tests of Mockup Hydrogen Vehicles

Three model year 2009 Honda Civic GX natural gas vehicles were purchased for this experiment. The CNG fuel container in this model is mounted transversely in the trunk space, immediately behind the rear seat back. The CNG containers were removed and replaced with similarly sized 350 bar (5,000 psi) Type 4 hydrogen fuel containers. Stainless steel tubing, rated for use with the hydrogen, was installed. Because the vehicles do not need to run for FMVSS No. 301 crash tests, no attempt was made to make the internal combustion engines work with hydrogen. The vehicles were ballasted as they would be for compliance testing.

The modified vehicles were subjected to front, side, and rear crashes as are specified in the test procedures for FMVSS No. 301 compliance tests. The vehicles behaved as they typically do in such crashes. The mockup hydrogen fuel systems in all three crashes were able to maintain their internal pressure following the crash, even though the fuel system in the side crash sustained moderate damage and that in the rear crash sustained substantial damage.

The crashes demonstrated that it is possible for a hydrogen fuel system to contain the hydrogen in a severe crash test, if it is properly protected and installed. At the same time, the deformation of the mounting and the plumbing, particularly in the rear crash, highlights the importance of specifying components that can withstand a crash.

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Figure ES-4. In the rear crash, the moving deformable barrier passed through the crumple zone in the trunk and directly impacted the hydrogen fuel container. The container is visible in the gap between the rear bumper and the trunk lid.

Related Task Orders

NHTSA’s hydrogen research program includes other activities related to the present tests.

• Durability Testing (Task Order 2). In this activity, the integrity of containers will be assessed after they have experienced pressurization and temperature cycles intended to simulate a lifetime of harsh service.

• Post-Crash Hydrogen Leakage Limits and Fire Safety (Task Order 3). This task order provides data to support assessments of acceptable post crash leakage rates. Hydrogen gas was flowed at a controlled rate into a passenger car interior. A spark ignited the hydrogen, and the consequences to the vehicle and to dummies in it were noted. The experiment was repeated for a number of different flow rates. A draft report on this project [5] is under review.

• Electrical Isolation (Task Order 4). Battelle developed and verified an alternative electrical isolation test procedure for hydrogen fuel cell vehicles. The refined final procedure was performed on two hydrogen fueled vehicles, confirming that the detailed steps and instrumentation can accurately test electrical isolation on an inactive fuel cell. This task order is completed [4].

• Published Literature Review (Task Order 5). The final report on this project [3] cited more than 100 references on hydrogen dispersion and ignition, tests of the container, fast

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fueling, and incidents. Findings were discussed in light of NHTSA’s needs. This task order is completed.

• Electrical Protective Barrier (Task Order 6). Battelle is assessing the need for an electrical protective barrier in fuel cell vehicles. Battelle is examining various failure modes and evaluating test procedures.

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1.0 INTRODUCTION

NHTSA promotes the safety of vehicles through several means, including setting and enforcing safety performance standards for motor vehicles and associated equipment through the Federal Motor Vehicle Safety Standards. Recognizing the unique hazards and issues associated with use of hydrogen fuel, which is colorless, odorless, stored at extremely high pressure, and has a flammability range of 4 to 75 percent in air, NHTSA is undertaking risk assessment studies to quantify potentially unsafe conditions; developing performance tests to address these conditions; and evaluating procedures to ensure hydrogen-fueled vehicles exhibit a level of safety equivalent to that of conventionally fueled vehicles.

Hydrogen-fueled vehicles are expected to significantly reduce the amount of pollutants expelled into the environment. Furthermore, hydrogen can be generated from a number of energy sources, thus helping to address energy security. Hydrogen vehicles introduce hazards that differ from those of the conventional fueled vehicles that consumers, mechanics, fire safety personnel, the public, and engineers already know and understand. Nevertheless, the public demands that hydrogen vehicles be no more hazardous to own and operate than conventional gasoline or diesel fueled vehicles. The objective of this Task Order is to develop data for NHTSA to consider in determining the most appropriate fueling conditions for crash tests of vehicles with hydrogen vehicle fuel systems.

To ensure fuel system integrity of passenger vehicles in front, side, and rear impact crashes, NHTSA has promulgated regulations that impose limits on post-crash fuel leakage under representative crash conditions. These conditions are defined in FMVSS Nos. 301, Fuel System Integrity [8], and 303, Fuel System Integrity of Compressed Natural Gas (CNG) Vehicles [9]. FMVSS No. 301 limits liquid fuel leakage in a fixed or moving barrier crash test to 28 g from impact until motion of the vehicle has ceased, and shall not exceed a total of 142 g in the 5-minute period following cessation of motion. For the subsequent 25-minute period, fuel spillage during any 1-minute interval shall not exceed 28 g. FMVSS No. 303, which applies to bi-fuel, dedicated, and dual fuel CNG vehicles, limits the leakage of compressed natural gas in a fixed or moving barrier crash test to an energy equivalent measured by a post-crash pressure drop in the high pressure portion of the fuel system through the 60 minute period following cessation of motion. Similar testing requirements need to be developed for hydrogen-fueled vehicles. Toward this end, NHTSA has tasked a team led by Battelle to evaluate various technical aspects of the safety of hydrogen fueled vehicles. Battelle was the prime contractor leading a team to conduct these experiments. Battelle was responsible for the overall organization of the project and analysis of the data. Powertech had primary responsibility for carrying out the crush tests and the leak tests. Powertech also modified the three CNG vehicles to have mockup hydrogen fuel systems and supported the crash tests. Texas Transportation Institute (TTI) conducted the crash tests. TRC, Inc., built the moving deformable barrier for the crash tests and provided valuable consulting on FMVSS No. 301 compliance testing.

2

1.1 Dynamic Impact (Crush) Tests of Fuel Containers

A unique feature of pressurized fuel containers is that the internal pressure significantly affects the likelihood and even the nature of the possible failure. A container is stiffened by its internal pressure, so that a container filled to its full service pressure is less likely to flex or deform under a given impact than a nearly empty container. Therefore, a container at lower pressure could be more easily deformed and possibly suffer damage from an impact that a fully pressurized container would otherwise withstand. This is especially true for lateral or transversely oriented impacts.

Composite containers are especially vulnerable to axial impacts in the end opening area. Different composite containers designs have end openings that will differ in diameter, sealing mechanisms, and degree of protrusion from the dome. Some designs are more vulnerable than others, but all designs have areas of elevated stress concentrations stemming from both design and manufacturing limitations. When a container is impacted, a crack is typically initiated somewhere in the transition from the dome end to the face of the opening, as the impact force is concentrated on the protruding end boss. The degree and severity of crack propagation from this point depend on factors such as internal pressure, amount of displacement of the end boss, thickness of the liner or laminate, and type of container.

The purpose of the impact tests is to explore these phenomena in both Type 3 and Type 4 hydrogen containers. The series of experiments provided data on the effects of construction, wall thickness (design service pressure), impact orientation, and internal pressure. These experiments will lead to a better understanding of how Type 3 and Type 4 hydrogen containers will behave in a crash, and will assist NHTSA in identifying the set of conditions in which the containers are most vulnerable to damage.

1.2 Leak Rate Measurements

The possibility of a fire resulting from a fuel container leak depends in part on the amount of fuel leaking. The actual mass of leaked hydrogen following a crash test would be difficult to measure directly, but the pressure drop in a container can be readily measured. Inferring leak quantity by a pressure drop is the approach taken by FMVSS No. 303 for CNG vehicles. The relationship between mass flow and pressure drop can be estimated from ideal gas laws, but real gases can depart significantly from ideal behavior, particularly at the high pressures at which hydrogen will be stored onboard vehicles.

Ideally, the limit would be zero—fuel containers would not leak at all following a crash. True zero is difficult to measure, so a small but reasonable and safe amount must be specified. These experiments included reasonable leak rates under controlled scenarios, and the data can show how to measure specific mass leak rates through a container’s pressure drop. A companion project [5] has examined the consequences of hypothetical leaks of various amounts of hydrogen.

The SAE J2578 standard [22], presents a number of calculations and simulations of hydrogen and helium leaks. An important purpose of the experiments was to produce data to verify the calculations and the equivalence of hydrogen and helium under likely test conditions.

3

1.3 Crash Tests of Mockup Hydrogen Vehicles

The purpose of the full-vehicle crash tests was to demonstrate the behavior of a hydrogen container and its mounting in a crash test. Accelerations and deflections were measured to quantify the conditions endured by a container, its mountings, and the associated plumbing. The nature of the damage was qualitatively described, and the results were photographed.

Whereas the impact tests were expected to produce a number of failures and did so, the crash tests were not expected to lead to a catastrophic failure. Though the fuel systems were damaged by the crashes, they held their pressure. Because the behavior of the vehicles in the crashes was typical of that sustained by small passenger cars in FMVSS No. 301-style crashes, the response of the fuel system can be considered representative of what might be sustained by a trunk-mounted hydrogen fuel container in such a crash. Thus, it can guide the interpretation of future sled tests to further assess the vulnerabilities of hydrogen fuel systems.

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2.0 DYNAMIC IMPACT CRUSH TESTS OF FUEL CONTAINERS

These experiments determined the behavior of hydrogen containers themselves when they are subjected to the full energy of a crash test. When the container is mounted in a vehicle in a crash, as in Section 4.0, it is more protected from the impact energy, though not entirely so.

2.1 Summary of the Dynamic Impact Tests

This experiment examined the behavior of hydrogen fuel containers when subjected to worst case constraint scenarios with impact energies simulating those occurring in FMVSS No. 301 or 303 crash tests. Orientation, internal pressure, and installation characteristics were varied. The goal was to determine what combination of conditions led to a ‘worst case’ failure mode and provide a basis of establishing realistic test parameters for future hydrogen vehicle crash tests.

The test program was structured in three groups of three tests on Type 3 and Type 4 fuel containers. Each group examined a possible configuration of the container under practical service installations, ranging from the most robust case (impact in the horizontal orientation with an end plug) to most vulnerable case (impact in the vertical orientation with a valve), while the test conditions within each group examined various characteristics of a container such as wall thickness and pressure contained in the container. The three groups examined were:

• Horizontal orientation (simulating an impact to the front or rear of a vehicle with a transversely mounted container) with an end plug (this is theoretically the most robust case and was shown to be so);

• Vertical orientation (simulating a side impact crash) with an end plug; and

• Vertical orientation with a valve.

Within each test group, the test conditions were

• Thinnest container sidewall (350 bar design service pressure) and near empty fuel system (filled with hydrogen to 10% of service pressure), the theoretical worst case scenario;

• Thinnest container sidewall and full fuel system (filled with hydrogen to 100% service pressure);

• Thickest container sidewall (700 bar design service pressure) and near empty fuel system

None of the containers in the horizontally oriented test group leaked or ruptured. Containers with less pressure were less stiff and deformed more under the impact load than did those that were fully pressurized. Because these tests imparted more energy to the containers than they are likely to experience in FMVSS No. 301 compliance crash conditions, data suggest that the containers struck in this orientation in a compliance crash test are unlikely to fail.

All containers in the two vertically oriented test groups failed by either leakage or rupture. Containers that were nearly empty leaked and those that were fully pressurized ruptured. The leak rates from the leakage failure modes were significant with only one small leak. Whether a valve or end plug was used as the container fitting appeared to have no effect on the failure mode. In the one case where a fully pressurized 700-bar container was struck on its end, a slow

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leak came from plastic deformation of the steel plug but the container itself had no apparent damage. As with the horizontally impacted containers, the test conditions were more severe than the container would experience in a compliance crash test. A container in a vehicle would typically be mounted with straps, so it would not be constrained between the impacting force and a rigid surface as was the case in these tests.

It must be noted that the boundary conditions for these tests were more severe than is likely in a crash. The container was fully constrained in the axial direction and subjected directly to the full impact force. During a crash, a container mounted in a vehicle would likely experience much less deflection as the vehicle and mounting system would absorb a significant portion of this energy.

2.2 Introduction and Background

Containers for pressurized gas fuel in motor vehicles are constructed differently than containers for liquid gasoline, so they are susceptible to different kinds of damage. Containers for hydrogen are built differently than those for compressed natural gas, so they must be tested separately.

2.2.1 Test Objective

The objective was to understand the behavior of hydrogen fuel containers when they are subjected to an impact energy representing those of FMVSS No. 301, Fuel System Integrity. The experiments generated data to enable the determination of a worst-case configuration and state of internal pressurization for fuel systems in hydrogen vehicles subjected to compliance crash tests. The effects of container orientation, design wall thickness, degree of internal pressure, and end fitting configuration were examined by recording pressure drop, impact force, and failure mode.

2.2.2 Construction of Hydrogen Containers

Containers for pressurized gas fuel in motor vehicles are constructed differently than containers for liquid gasoline, so they are susceptible to different kinds of damage. They are typically designed as all metal liner (Type 1), metal liner with composite wrap over the cylindrical sidewall (Type 2), metal liner with full composite overwrap (Type 3), or plastic liner with full composite overwrap construction (Type 4). All major manufacturers of hydrogen vehicles are using Type 3 or Type 4 container designs, so they were the focus of the test program.

The main differing design feature between Type 3 and Type 4 construction is the liner material. This liner acts to contain the gas while the composite overwrap provides the strength to withstand the high storage pressures.

In a typical Type 3 vehicle fuel container (Figure 2-1), the liner is constructed from seamless, extruded tube (usually aluminum for weight/strength ratio, ductility, and corrosion resistance) with a technique of heating and mechanical forming to close the end domes into a ported end or closed end. A ported end would be later threaded to accommodate a container valve or end plug. The liner is then fully wrapped using a composite fiber and resin matrix (usually carbon fiber and epoxy) and cured. The metal liner in this design can take up to 20 percent of the wall stresses from internal pressurization with the remainder borne by the composite laminate.

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Figure 2-1. Cross-section of a typical Dynetek Industries Ltd. Type 3 container.

In a typical Type 4 vehicle fuel container (Figure 2-2), the liner is constructed from extruded plastic tube (usually some form of polyethylene for cost, expansion, permeation, and good temperature extreme resistance) with injection or rotation-molded end domes bonded to the sidewall portion. Each dome contains an integral metal boss end to accommodate a container valve or end plug. The liner assembly is then fully wrapped using a composite fiber and resin matrix (usually carbon fiber and epoxy) and cured. The composite laminate takes 100 percent of the load of the internal pressure.

Figure 2-2. Cross-section of a typical Lincoln Composites Inc. Type 4 container.

A unique feature of pressurized fuel containers is that the internal pressure significantly affects the likelihood and even the nature of the possible failure. A container is stiffened by its internal pressure, so that a container filled to its full design service pressure is less likely to flex or deform under a given impact than a nearly empty container. Therefore, a container at lower

Aluminum Liner Carbon Fiber

Overwrap

Boss and Neck Area

HDPE Liner

Fiberglass outer layer

Aluminum Boss

Plastic/Alunimum Boss/Liner Interlocking Interface

Dome/Sidewall Plastic Weld

Carbon/glass hybrid overwrap

Energy absorbing material

Neck area

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pressure could be more easily deformed and possibly suffer more damage from an impact than a fully pressurized container. Mitsuishi, Oshino, and Watanabe [14] have shown that this is especially true for lateral or transverse oriented impacts [11], and the present experiment confirmed this as well.

The internal pressure of a container determines the level of stress of the liner. Upon axial impact at low internal container pressure, the container is less stiff and the boss can collapse into the dome. In cases where the internal pressure does not exceed the rupture stress of the liner, the outcome will most likely be leakage, as in the experiments of Mitsuishi, Oshino, and Shogo Watanabe [14]. The newly initiated cracks do not significantly propagate and gas is allowed to escape through the damaged areas. In impacts on containers with high internal pressure, the cracks propagate rapidly and typically result in rupture. Mitsuishi and colleagues have shown in Type 3 containers that an analysis of containers fractured during similar crush tests revealed pure tensile fracture surfaces despite the compressive force of impact.

Composite containers are especially vulnerable to axial impacts due to the geometry of the dome transitioning to an end boss and end opening area. While composite container designs have end openings that differ in diameter, sealing mechanisms, and degree of protrusion from the dome, all designs have a dome that contains at least one end boss with an opening. This neck area of the end boss is the most vulnerable portion of the container due to the elevated stress concentrations stemming from its geometry as well as design and manufacturing limitations. When impacted, a crack is typically initiated in this neck area as impact forces are concentrated on the protruding end boss. The degree of crack initiation stems from:

• Boss geometry (locations and degrees of stress concentration),

• Container construction (all metal dome and boss as in Type 3 or plastic dome interfacing with metal end boss as in Type 4), and

• Manufacturing quality (depth of folds in the neck region or strength of the liner).

Severity of crack propagation from this point depends on factors such as:

• Internal pressure, • Amount of displacement of the end boss, • Thickness of the liner laminates, and • Type of container construction.

2.3 Test Conditions

The containers for this test were selected to be representative of those that are being considered by auto manufacturers for use in hydrogen-fueled passenger vehicles. The crash energies were intended to be the most severe that could realistically occur in a FMVSS No. 301-style set of crash tests.

2.3.1 Selection of Containers for the Test Program

The study assumed a typical fuel system configuration in a passenger vehicle. This means the system would consist of one or more containers (each with a valve and pressure relief device)

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connected together and mounted transversely in the rear portion of a vehicle, typically under the rear trunk compartment, rear seat or a combination thereof. In such a system, each individual container can be considered an iterative ‘design element’ , meaning that a system could be produced by repeating each container or ‘design element’ to create a fuel system. The test program focuses on the impact energy or forces affecting this one individual design element or container rather than an entire system.

Only Type 3 and Type 4 containers were studied. A container representative, in volume and dimensions, of a typical vehicle fuel container was selected for the test program. A configuration of a double-ported container (container with openings on either end) was selected as the worst case scenario. Containers are available with closed domes but because the closed end dome has no neck or lacks the geometry to accommodate an opening, the stress concentrations are greatly reduced or even eliminated in some designs.

Major vehicle manufacturers are considering 350 bar as the minimum and 700 bar as the maximum design service pressure. These design service pressures were selected as the thinnest wall container (350-bar design) and thickest wall container (700-bar design) for the test program. Containers for 350 bar have thinner sidewalls and will flex more than 700-bar containers under the same external loading, making them more likely to be damaged in a crash, if all other conditions are the same.

Dynetek Industries Ltd. and Lincoln Composites Inc. were chosen as the manufacturers to provide Type 3 and Type 4 fuel containers, respectively. Their fuel containers are currently being used in natural gas and hydrogen vehicles worldwide in 250 bar, 350 bar, and 700 bar fuel systems. The models (specifications shown in Table 2-5) selected for this test program are actual vehicle fuel containers currently in service around the world.

2.3.2 Selection of Impact Energy

Battelle selected a small number of recent compliance tests from NHTSA’s database [18] to estimate a realistic worst case of the conditions a fuel container might endure in a crash test. The cases were examined from a kinetic energy standpoint and a peak force standpoint. The energies were compared with those used in Mitsuishi’s similar prior study [11].

2.3.2.1 Analysis of Prior FMVSS 301 Compliance Tests

Table 2-1 lists the selected compliance tests and their estimated “high” and “ low” impact energy values. Battelle noted the velocities of the moving barrier and the test vehicle before and after the impact. The “ low” energy in the table is the change in kinetic energy of the (filled) container, assuming it is rigidly mounted to the chassis and experiences the same change in velocity as the test vehicle. It is a lower bound approximating what a well-protected container would experience in a highway crash. The “high” energy is the total energy dissipated at impact—the difference between the total kinetic energy of the vehicle (or moving barrier) before impact and the kinetic energy after. In the extreme case, if the fuel container itself was the bumper and absorbed all of the energy lost in the collision, this is the amount. It is the highest upper bound on the energy a container would possibly dissipate in this kind of crash.

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Table 2-1. Energy dissipated in NHTSA FMVSS 301 compliance tests.

Condition NHTSA

Compliance Test Number

Vehicle Energy, kJ

Low High

Front 30 MPH 6378 2008 Kia Spectra 4.2 127.8

Side 33 MPH 5986 1995 Honda Civic 4 Door 1.1 62.1

5984 1996 Pontiac Grand Am 0.4 65.6

Rear 30 MPH 5553 2005 Volkswagen Golf 4-Door Hatchback 1.6 44.0

5214 2003 Jeep Liberty 1.3 49.2

Estimated Rear 50 MPH

(see note) 2005 Volkswagen Golf 4-Door Hatchback 4.5 122.2

(see note) 2003 Jeep Liberty 3.7 136.6

Front 28.5 MPH (this program) 2009 hydrogen mockup vehicle - 124.1

Side 31.7 MPH (this program) 2009 hydrogen mockup vehicle - 66.4

Rear 51.8 MPH (this program) 2009 hydrogen mockup vehicle - 204.0

Note: There were no 50-mph rear impact tests reports on the NHTSA Web site. These energy values were estimated by scaling the rear 30-mph values by the square of the speed ratio. The crashes from the present program are added for comparison; they were not available at the time the drop tests were planned.

Table 2-2 lists the estimated “ low” and “high” peak forces in the same compliance tests. The “ low” value was estimated by multiplying the peak acceleration of the test vehicle by the mass of the filled container. If the container were well protected inside the vehicle and rigidly mounted to the chassis, this would be the force exerted on it by its mounting. The “high” value is the product of the peak acceleration and the mass of the test vehicle (or of the moving barrier in the case of side and rear impacts). This is the peak force at the interface between the test vehicle and the moving or fixed barrier. If the impact penetrated to the surface of the container such that it bore the full load of the impact, the “high” estimate would be that force. As with the “high” energy values, this is an extreme, yet objective, upper bound.

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Table 2-2. Peak forces in NHTSA FMVSS 301 compliance tests.

Condition NHTSA

Compliance Test Number

Vehicle Peak Force, kN

Low High

Front 30 MPH 6378 2008 Kia Spectra 30.9 940

Side 33 MPH 5986 1995 Honda Civic 4 Door 12.3 259

5984 1996 Pontiac Grand Am 13.2 256

Rear 30 MPH 5553 2005 Volkswagen Golf 4-Door Hatchback 14.1 444

5214 2003 Jeep Liberty 18.5 796

Estimated Rear 50 MPH

(see note) 2005 Volkswagen Golf 4-Door Hatchback 23.4 740

(see note) 2003 Jeep Liberty 30.9 1327

Note: There were no 50-mph rear impact tests reports on the NHTSA Web site. These force values were estimated by scaling the rear 30-mph values by the speed ratio.

The conditions for the drop tests were to approximate a worst case crash, yet be grounded in reality. Speaking in round numbers, the “high” energy values are roughly in the range of 50 to 100 kJ. The “high” values for the peak force are 500 to 1,000 kN.

2.3.2.2 Comparison with Prior Container Impact Research

The Japanese Automotive Research Institute has published research on drop tests similar to those in this study. Mitsuishi et al. [11] reported that only 50 kJ was sufficient to rupture a container if it was applied directly to the container in an axial impact. Containers that are impacted on their sides, rather than their domes, can survive much greater impacts. Mitsuishi also found that the peak force developed by the container occurred in the first 100 mm of deformation. Because the container in the crush test, unlike a crash, is fully constrained, the amount of deformation in a crush test can be limited by a stopper. That 50 kJ value is close to the energy dissipated in the compliance tests in Table 2-1.

An impact can have essentially three outcomes—the container can rupture in a catastrophic failure, it can leak, or it can maintain its pressure integrity, though perhaps with damage. Some of the cases reported by Mitsuishi, Oshino, and Watanabe are similar to those in the NHTSA test plan. Their outcomes of the corresponding tests are listed in Table 2-3. The tests by Mitsuishi et al. resulted in some cases in a leak and in some cases in a rupture. The objective for the NHTSA tests is to determine the worst case impact and pressurization conditions.

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Table 2-3. Comparison of the high-level results of the prior impact tests with the current test plan.

.NHTSA Test Plan Result of the Closest Mitsuishi Case [11]

Container Service

Pressure (bar)

Container Orientation

Pressure

Type 3 Container Type 4 Container

Size Crush mm

Result Crush mm

Result

350 Horizontal 10% SP

350 Horizontal 100% SP 3a 25 No Leak

700 Horizontal 10% SP

350 Vertical 10% SP 3a 3b

138 100

Leak Leak

100 Leak

350 Vertical 100% SP

3a 3a 3a 3a 3b 3b

50 88

138 138 50

100

Rupture Rupture Rupture Rupture

Leak Leak

50 100 100

Leak Rupture Rupture

700 Vertical 10% SP

Notes concerning the tests in Mitsuish,i Oshino, and Shogo Watanabe [14]:

• Two sizes of Type 3 container, designated 3a and 3b, were used. See the table at right. • The “crush” column identifies the permitted crush distance, as shown in Figure 4-1. • Partially filled containers were at 70 bar (20% of the 350-bar service pressure). • Containers were filled with helium, except for the two in the circle, which were filled with hydrogen.

These two cases were filled with hydrogen; others, with helium.

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2.3.2.3 Decision for the Drop Energy

By the energy analysis, the force analysis, and the outcome analysis, the conditions of Mitsuishi’s group are close to those that were desired for the NTHSA tests. The contractor team recommended that these tests be run at exactly the same conditions as the previous tests so the results of these rare tests could be more directly compared.

The weight was 2,500 kg and it was dropped from a height of 2.0 m, providing a kinetic energy at impact of 50 kJ.

Wooden stoppers adjacent to the container (Figure 2-3) were 100 mm below the container’s top surface to limit the deflection. This maximum allowable deformation of 100 mm was selected based on discussions in Section 3.2.1 where Mitsuishi found that the peak force was observed in the first 100 mm of deformation. The allowable deformation was kept constant through all test conditions to provide consistency and allow the comparison of all outcomes.

2.3.3 Test Matrix

A pressurized fuel container must be strong and robust. The weak points of a cylindrical fuel container arise from the need for openings to render it useful for storage. The presence and geometry of these openings, as well as design and manufacturing limitations in producing them, generate stress concentrations and crack initiation points. The containers were oriented in the horizontal (transverse impact onto the sidewall) and vertical (axial impact onto the end boss) directions to verify that contact with the end boss is indeed a worst case scenario.

The worst pressure condition for crash testing hydrogen fuel system is not necessarily maximum internal pressure. Composite fuel containers are damaged when they are forced to flex and deform by external impact. Internal pressure tends to stiffen fuel containers and piping, preventing them from deforming, thereby reducing the induced damage. Consequently, work was planned on the hypothesis that fuel containers and fuel system components are more likely to leak in a crash if they are at lower pressures, than higher pressures.

The key variables being examined for this test program are:

i) Impact or ientation (hor izontal versus vertical). To evaluate the effect of a container mounted transversely in the rear portion of a vehicle being impacted from the front/rear or side. The fuel containers were impacted in the horizontal (front/rear impact condition) and vertical (side impact condition) directions.

ii) End fitting component (end plug versus valve). To evaluate the effect of a blank/ported end plug versus an in-tank valve containing an integral pressure relief device (PRD). The use of the end plug rather than a valve ensures potential container failure will originate in the container and not in an associated component. The valve case was meant to examine whether the inclusion of the valve increases susceptibility of the container design element (fuel container + valve) to leakage.

iii) Container type (Type 3 versus Type 4). To compare the effect of different container constructions on potential failure characteristics.

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iv) Design service pressure (350 bar versus 700 bar ). To compare the effect of container wall thickness on potential damage or failure mode.

v) Internal pressure (10% service pressure versus 100% service pressure). To compare the effect of internal pressure on potential damage or failure mode of the fuel container. The 100% service pressure was selected to represent a full fueling condition at maximum rigidity and 10% service pressure was selected to represent a near-empty fuel condition at near-minimum rigidity. The 10% service pressure condition also reflects the lower or minimum pressure limit during a typical container design qualification pressure cycle test for the HGV2 standard [2].

Table 2-4. Test conditions for crush test.

Container Service

Pressure (bar)

Container Orientation

Pressure Plug/Valve Rationale

350 Horizontal 10% SP Plug Horizontal worst case, thinnest wall and lowest pressure

350 Horizontal 100% SP Plug Confirms 10% SP (Test 1) is worst case for impact

700 Horizontal 10% SP Plug Confirms thinnest wall (Test 1) is worst case for impact

350 Vertical 10% SP Plug Vertical worst case, thinnest wall and lowest pressure

350 Vertical 100% SP Plug Confirms 10% SP (Test 4) is worst case for impact

700 Vertical 10% SP Plug Confirms thinnest wall (Test 4) is worst case for impact

350 Vertical 10% SP Valve/PRD Demonstrates effect of valve, worst case vertical (Test 4)

350 Vertical 100% SP Valve/PRD Confirms 10% SP (Test 7) is worst case for valve impact

700 Vertical 10% SP Valve/PRD Confirms thinnest wall (Test 7) is worst case for valve impact

These nine cases were carried out once for Type 3 containers and once for Type 4 containers.

The test matrix shown in Table 2-4 was structured in three groups of three tests. Each group examined a possible configuration of the container under practical service installations, ranging from the most robust case (impact in the horizontal orientation with end plug) to most vulnerable case (impact in the vertical orientation with a valve), while the test conditions within each group examined various characteristics of a container such as wall thickness and pressure contained in the container. The testing was conducted for Type 3 and Type 4 containers for a total of 18 tests.

One deviation of note from the test plan was that no “with-valve” test was conducted for the 700-bar Type 4 container. Due to the port configuration of the test container and the limited availability of 700-bar valves, no valve was available to perform the test. Instead, Test 18 was a unique condition where the 700 bar Type 4 container was impacted in the vertical orientation at 100% service pressure.

For all test conditions, the container was fully constrained by having the bottom surface of the container resting against a non-deformable steel plate mounted on a load cell. This was to ensure a fixed deformation, decrease test variability, and provide a force measurement during the crush event. This fully constrained condition was necessary for the testing; however, it exceeds the reasonable worst case conditions of an actual vehicle crash.

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2.3.4 Test Apparatus

The testing was conducted at Powertech’s Remote Destructive Testing Range located near Hope, British Columbia.

A drop testing apparatus, constructed for the purpose of subjecting pressure containers to dynamic crush scenarios, was used for this test. A schematic of the apparatus is in Figure 2-3 and a photograph in Figure 2-4. The apparatus consisted of a drop weight constrained by two guide rails for approximately 5 m of vertical travel. The weight was held and released using a magnet suspended from a cross-beam above the apparatus. The area where the container was situated was enclosed with steel with a cut-out to allow cameras to view the test. The enclosed space was intended to capture pieces of the pressure vessel during rupture for safety and post-test documentation.

Figure 2-3. Schematic of the drop test apparatus.

TEST CONTAINER

LOAD CELL

STOPPER

P

PRESSURE SENSOR

MAGNET

DROP WEIGHT

CRUSH DISTANCE (100 mm)

EMERGENCY VENT LINE

ACCELEROMETER

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Figure 2-4. Photograph of the drop test apparatus.

The drop weight, shown in Figure 2-5, was a steel enclosure containing sand with a total mass of 2,500 kg. The impact surface was a 50-mm-thick steel plate welded to the bottom of the drop weight.

Figure 2-5. Photograph of the drop weight.

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The container crush zone was limited to a specific displacement of 100 mm. Wood blocks measuring 255 mm x 255 mm were used to stop the weight at the desired displacement.

The specifications of the containers are in Table 2-5. Containers were selected for the crush test based on availability of 700-bar units both in Type 3 and Type 4 construction from the two manufacturers that were also available at the 350-bar pressure, with the extra condition that they were all approximately the same volume. All these factors converged around a water capacity of 60 L. Note that the aspect ratio of the 700-bar Type 4 container is more round than cylindrical. Containers of these same four specifications were used for the leak rate characterization in Section 3.0. The 350-bar Type 4 container container is the same specification as those installed for the crash tests, as shown in Table 4-3.

The containers were provided by the manufacturer with an end plug installed in each end. The end plugs were custom machined using mild 4000-series steel and were 50 mm in diameter and approximately 50 mm long. End plugs installed into the top end of the container had no ports and the end plugs installed into the bottom end of the container had two quarter-inch FNPT ports machined into opposing sidewalls of the end plug. Swagelok SS-20-VS4 needle valves were installed in each port and connected to two separate vent lines intended as redundant emergency vent lines as well as pressure measurement lines.

Table 2-5. Specifications of containers used for the crush and leak tests.

Container Type Type 3 Type 4

Manufacturer Dynetek Industries Ltd. Lincoln Composites Inc.

Construction Aluminum liner,

full carbon fiber overwrap Plastic liner,

full carbon fiber overwrap

Service Pressure 350 bar 700 bar 350 bar 700 bar

Volume (Water Capacity)

61 L 58 L 65 L 66 L

Mass (nominal, empty)

32 kg 56.5 kg 32 kg 67 kg

Length 1,304 mm 1,230 mm 815 mm 512 mm

Diameter (nominal) 297 mm 329 mm 400 mm 692 mm

For the horizontal crush tests, the containers were oriented horizontally on a 12-mm-thick steel plate that was as long as the container and rested directly on the load cell.

For the vertical crush tests, the containers were kept upright by 12-mm-diameter ready-rods and electrical tape. The configuration was intended to keep the container upright during the test but not secure the container in place, possibly affecting the behavior or failure mode of the container during or after impact. The bottom end of the container rested on a 100 mm thick steel plate that fit directly into the sensing portion of the load cell.

Specifications of the equipment used for instrumentation and data acquisition are given in Table 2-6.

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Table 2-6. Specifications of instrumentation and DAQ.

Measurement Manufacturer Model Specification

Pressure Stellar Technologies Inc. GT160-series 0 – 40,000 psi, 0.1-4.9 VDC output, 0.1% FS accuracy

Data Acquisition (for pressure only)

National Instruments PCI DAQ Card with custom programmed LabView software

SCB-68 Used at 1 kHz frequency

Acceleration Dytran Inc.

Load Powertech Labs Inc. (custom manufactured)

n/a 0 – 4.5 MN load capacity

Data Acquisition (all other high frequency data)

Soltec Corp. TA220-2300MF Used at 100 kHz frequency with 100 Hz low-pass filter

High-Speed Imaging

Photron Ultima APX High-speed imaging was set at 250 fps – 500 fps depending on available lighting conditions

The valve for the 350 bar test cases was chosen as a GFI Control Systems Inc. model XTV-195X2 with an integral metal eutectic type PRD. The valve for the 700 bar test case was chosen as a Dynetek model BV-700-01-0 with an integral glass bulb type PRD.

2.3.5 Test Procedure

The weight was lifted part way, to a height that allowed for safety pins to lock the weight in place. All additional safeties for the weight were put in place to ensure three levels of safety. The appropriate stoppers were positioned. The container was then placed into position in the appropriate orientation for the test conditions. The container was secured, and the vent and pressure measurement lines were connected. The needle valves were then opened and pressure was confirmed by the DAQ system. All cameras were set to record and two levels of safety were removed. The weight was remotely lifted to the test height and the final level of safety remotely removed. The high frequency DAQ system and the high-speed camera were set to event trigger mode. Upon clearance and final checks, the magnet was deactivated allowing the weight to drop and impact the container. If the container remained pressurized, the pressure was remotely released before the impact site could be accessed by Powertech personnel.

2.4 Results

The outcome of the testing resulted in one of three scenarios: no effect (no leakage or rupture), leakage, or rupture. In scenarios where there was no leakage or rupture, visual examination of the container revealed qualitative observations ranging from showing no evidence of external damage to evidence of fiber breakage in impacted areas on the top or bottom surfaces.

The distinction between leakage and rupture is the time duration and nature of the pressure release event. A leakage can be considered “ the release of the container contents through a defect

18

or crack.” A rupture can be considered “a sudden and unstable damage propagation in the structural components of the container” [2].

The results of each test condition are in Table 2-7. Photographs of the containers after the crashes are in Appendix A. A complete set of load time histories is in Appendix B.

Table 2-7. Results from crush test by test condition.

Test Condition Type 3 (Dynetek) Type 4 (Lincoln)

Service Pressure Orientation

Test Pressure

End Test

# Result

Peak Force

Test #

Result Peak Force

350 Horizontal 10% Plug 1 No leak or rupture, no visible damage

636 kN 4 No leak or rupture,

fiber breakage on top surface

645 kN


1,596 kN 5 No leak or rupture,

fiber breakage on top and bottom surface

1,632 kN


1,005 kN 6 No leak or rupture,

minor fiber breakage on top surface

1,347 kN

350 Vertical 10% Plug 7 Leakage

(34.7 to 0.1 bar in 4.25 s)

370 kN 10 Leakage

(35.6 to 0.2 bar in 2.55 s)

286 kN

350 Vertical 100% Plug 8 Rupture 672 kN 11 Rupture,

plastic welds failed No data

700 Vertical 10% Plug 9 Leakage

(73.1 to 0.2 bar in 3.5 s)

699 kN 12 Leakage,

via sheared fitting 557 kN

700 Vertical 100% Plug 18

Leakage, via deformed end plug

(700 bar to 0 bar in 14 hours or ~50,000 s)

1,536 kN

350 Vertical 10% Valve 13 Leakage

(36.3 to 0.2 bar in 1.95 s)

237 kN 16 Leakage

(35.2 to 0.1 bar in 2 s) 237 kN

350 Vertical 100% Valve 14 Rupture No data 17 Small rupture or

big leak 575 kN

700 Vertical 10% Valve 15 Leakage

(72.4 to 0 bar in 1.8 s)

656 kN

2.4.1 Horizontal Impact

All horizontally oriented containers, anticipated in the test plan as the most robust test condition, were able to withstand the impact without leakage or rupture regardless of internal pressure. This confirms that fuel containers are able to withstand very high forces without failure when

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impacted on their side, or during front or rear crash scenarios. The high-speed video footage (example in Figure 2-6) shows that these fuel containers were actually subjected to multiple impacts. The video shows the 2,500-kg drop weight bouncing off the container (not the wooden stoppers), back upwards a distance, then falling and impacting the container, sometimes numerous times.

1 2 3

4 5 6 4. Drop weight released and coming down, prior to impact

5. Moment of drop weight impacting container

6. Moment of maximum compression

7. First bounce of drop weight off the container and container off the load cell

8. Drop weight falling again after first bounce contacting container again

9. Final resting position of drop weight on container

Figure 2-6. Stills from the high-speed video of Test #5, horizontal impact on a 350-bar Type 4 container filled to 100% service pressure.

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0

200

400

600

800

1000

1200

1400

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Test #1 - 10% SP, thin wall Test #2 - 100% SP, thin wall Test #3 - 10% SP, thick wall

Figure 2-7. Force time histories for Tests #1, #2, and #3 of the Type 3 containers. The higher, shorter trace for the fully pressurized Test #2 shows the container was stiffened by its internal pressure. The less pressurized containers allowed more deformation, with the thin wall in Test #1 being the least stiff and spreading the load over the longest time.

The horizontal impact cases confirmed the hypothesis that the internal pressure would stiffen the container. The impact on the Type 3 350-bar container filled to 350 bar, Test 2 or the dotted red line in Figure 2-7, is indicative of a mass landing on a stiff spring—the force is rather high and the duration of the contact is short. Contrast this with the impact on an identical container filled to only 10% service pressure, Test 1 or the solid blue line in Figure 2-8. A softer spring cushions the landing, slowing the mass over a longer duration and lessening the peak force. A softer spring also compresses more under the shock load than does a stiff spring. This effect is difficult to see in the stills in Figure 2-6 (which is only for Test 1), but it is readily apparent to one watching the high-speed videos of Tests 1 and 2. The thicker walled 700-bar container, filled to only 10% service pressure, in the green dashed line, draws out the impact over roughly the same duration as the blue line but generates a higher peak force because of its thicker wall than the 350-bar container.

Figure 2-8 shows the force time histories for the corresponding three impacts on horizontal Type 4 containers. The dotted red line, for the fully pressurized 350-bar container of Test 5, is brief and peaked. The solid blue line, for the partially filled container of Test 4, is of longer duration and lower amplitude, again showing that the nearly empty container is a softer spring than the fully pressurized container. The unpressurized 700-bar container, shown in the green dashed line, is stiffer, more like the fully pressurized 350-bar container. Its effective stiffness is due to its geometry—the 700-bar Type 4 container is shorter than the other containers. The

21

weight struck all of the other containers fully in their cylindrical section, but part of the dome on both ends of the 700-bar Type 4 container was under the weight. The ends of containers are stiffer radially than is the constant-radius cylindrical portion.

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Test #4 - 10% SP, thin wall Test #5 - 100% SP, thin wall Test #6 - 10% SP, thick wall

Figure 2-8. Force time histories for Tests #4, #5, and #6 of the Type 4 containers. The behavior is similar to that of the Type 3 containers in the previous figure, except that the less pressurized thick wall container (Test #6) is almost as stiff as the fully pressurized thin wall container (Test #4).

One reason for the robustness of the fuel container in this orientation is the geometry of the container relative to the location of contact on the container. The container was impacted at the sidewall where the container has no significant stress concentrations and has most of its rigidity. The sidewall is strengthened by the internal pressure, and on impact, is allowed to flex. The ductility of the liner (whether it be aluminum in the case of a Type 3 or plastic for Type 4) and the lack of stress concentrations allows this deflection without initiating cracks. Evidence for this deflection can be seen in the high-speed footage as well as observation of the fiber breakage seen in the Type 4 horizontal crush tests. Because Type 4 fuel containers have a flexible plastic liner rather than a metal liner, the liner does not provide structural support to the composite overwrap when subject to external compressive forces. To increase toughness of their composite laminate, Lincoln Composites Inc. uses a thin, sacrificial, non-load bearing glass fiber layer over their load-bearing carbon fiber wrap. The fiber breakage comes from the compressive forces and corresponding deflection of this outermost layer.

22

2.4.2 Vertical Impact

All containers struck in the axial direction were damaged, resulting in leakage or rupture. This confirms that a side impact crash, specifically those directly impacting the end dome of a transversely mounted container, would be worse than an impact on the container’s side. Even if the side of a vehicle were to strike a nearly rigid object the loading would not be as severe as in these tests, where the container was constrained by a rigid base opposite the impacting weight. Containers in vehicles are typically mounted by circumferential straps and the end opposite the striking object would be free.

All containers of both types at 10% service pressure in the vertical orientation leaked. The degree of leakage was significant: the container fully vented within 4 s. This was usually the result of a sizeable fracture of the liner (Type 3) or disbonding of the liner/boss interface (Type 4) in the neck area. There were no minute or slow leaks observed in these test cases. The dominant failure mechanism of the container of both types at 100% service pressure was rupture.

The most common mechanism for leakage was damage to the neck region (high-speed video captures shown in Figure 2-9) of the container either at the top end (contacting the drop weight) or the bottom end (contacting the load cell). It is difficult to conclusively determine what conditions would have caused the failure of one specific end of the container as opposed to the other end. The force of the impact drove the end boss down into the container (or up into the container for cases where the bottom end failed) causing sufficient deflection of the neck and initiating fractures in the shoulder of the neck where the highest stress concentration exists.

23

1 2 3

4 5 6

7 8

Figure 2-9. Stills from the high-speed video of Test #9, verticalal impact on a 700-bar Type 3 container filled to 10% service pressure.

Because the impact surface contacted the end plug or valve installed in the end boss, the forces of the impact were translated directly to the boss and consequently, the liner itself.

For Type 3 designs, the impact caused a fracture in the neck region. These arise from manufacturing limitations during forming of the neck. Folds inevitably occur in the internal surface of the liner as the end is formed and excess material is built up. Manufacturers carefully control the forming process to manage fold depth and ensure the fold depth does not exceed the maximum defect size allowed in their design. In this way, the folds are not initiation points for fatigue or fracture failure under normal service conditions. This is verified by the manufacturer in design qualification testing by ensuring fatigue failure during the pressure cycle test or fracture failures during the burst test does not originate in the neck area. However a direct impact to the end boss in the axial direction, especially of the magnitude seen in this study, fractures the neck with each fold acting as an initiation point.

For Type 4 designs, the metallic end boss is usually machined and therefore does not have the neck folds of a Type 3 design. However, the design and construction of a Type 4 container presents its own design challenges in resisting damage from a direct axial impact. The most significant of these is the liner/boss interface (refer to Figure 2-2). In a Type 4 container, the

24

impact drives the end boss into the dome while the internal pressure pushes the dome and boss outward, causing yielding in the plastic dome and possible disbonding at the boss/liner interface.

The container failure mode was dependent on the degree of stress present in the liner and composite laminate. It is apparent that when the stress levels were low (10% service pressure cases), the laminate was able to arrest crack propagation in the neck region causing leakage but not rupture. When the stress levels were high (100% service pressure), the cracks in the neck region propagated rapidly resulting in rupture. Assuming the impact dealt at least the same degree of damage to the composite for the 10% and 100% service pressure test cases, the ruptures would have occurred as a function of the inability of the laminate to arrest crack propagation at high stress levels and not due to damage to the composite from the impact.

The absence of a rupture in the case of Test 18, where a 700-bar container was subjected to impact in the vertical orientation at 100% service pressure, demonstrated an interesting case. While the 350-bar cases at 100% service pressure resulted in rupture, the 700-bar test case did not. Recognizing that only one container underwent this test case, it is nonetheless conceivable that the higher pressure, even though possibly subjecting the same neck region to even higher stresses than 350-bar cases, caused sufficient rigidity in the container that the impact energy was not high enough to cause a degree of deflection resulting in damage. Leakage occurred only because the end plug was deformed, not from damage to the container itself. An anecdotal observation from this test was that the impact forces were sufficient to plastically deform a 4000-series steel end plug and not a fully pressurized 700-bar fuel container. This deformation caused an extremely small leak (700 bar to empty over 14 hours) as reported in Table 2-7. The leak was from the plastically deformed steel end plug, not the container itself.

There was little difference in outcome between cases where a valve or end plug was at the point of impact. It can be concluded that the valves tested did not induce any additional failure mode unique to their presence. There was a possibility of the glass bulb type PRD, integral to the valve used in Test 15, fracturing as a result of the shock of impact but the bulb or the PRD actuation mechanism did not leak or fail.

These results can be well correlated with the results from testing conducted by Mitsuishi et al. In that test program (shown in Table 2-3), all test cases resulted in rupture except for three. Two test cases were Type 3 and Type 4 containers filled to 100% service pressure subjected to a vertically oriented impact with a 50 mm displacement resulting in leakage. It can be inferred that, in limiting the displacement to 50 mm, the containers were not subjected to the peak forces seen during this study. One test case of a Type 3 container filled to 100% service pressure subjected to a vertically oriented impact with a 100 mm displacement, corresponding to the test parameters in this study, resulted in leakage. This container was a Type 3 container manufactured by Structural Composite Industries LLC and indicates construction could affect the outcome as well.

It is important to note that the ruptures of the test are in part a result of the test conditions. A container in an actual vehicle crash is unlikely to be constrained between two essentially rigid surfaces, as were the steel plates in the drop weight and the base. Even if a vehicle were to slide into a nearly rigid bridge abutment, the straps holding the container are unlikely to generate a peak force as high as those in the tests.

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Rupture @ approx 55 mm displacement

Figure 2-10. Time history of the load for Test #8, in which the container ruptured.

2.5 Conclusion

This study examined the behavior of hydrogen fuel containers of Type 3 and Type 4 construction when subjected to impact energies simulating those resulting from FMVSS No. 301 or 303 test crashes. Orientation, internal pressure, and installation were varied. The purpose was to determine what could be considered a worst-case set of conditions and provide a basis of establishing realistic test parameters for future hydrogen vehicle crash tests. With only a single impact per test condition and limited range of conditions, the experiment’s conclusions must be carefully interpreted, but trends were clearly present and results were consistent with an understanding of the situation. A second caution in drawing conclusions is that the test conditions of 50-kJ impact energy on a fully constrained container are more severe than a container is likely to experience in a highway crash.

The videos of the horizontal impact tests, with the container bouncing between the anvil and the 2,500-kg weight, show just how violent these tests were. The containers are quite robust when struck on their sides, with all of the containers holding their pressure and only the Type 4 containers showing some visible surface damage. The low-pressure condition appeared to be more severe on the side-impacted containers because these containers flexed more during the impact. These greater strains would be expected to leave more residual damage. Quantifying the residual damage would require that the containers be sectioned for metallographic inspection or that they be subjected to pressure testing, either burst or cyclical.

Of the test conditions in this study, containers were most susceptible to damage when the impact was axially directed at their end. As most pressurized fuel containers in passenger cars are mounted transversely, axial impact corresponds to a side-impact crash. The conditions with the

26

most severe results were thinner wall (350 bar) containers at 100% service pressure. If these test conditions are adopted for future hydrogen vehicle crash testing, care should be taken to accommodate a potential rupture failure mode.

All of the container failures in this experiment resulted from the vertical (axial) impacts. The more highly pressurized containers failed by rupture, which is a more severe outcome than the leakage of the low-pressure containers. This is in contrast to the horizontal impacts, where the more severe outcome was presumed to be the greater flexing of the low-pressure containers. A tentative conclusion to draw from these observations is that high internal pressure protects a container from moderate impacts but can exacerbate the failure mode if the container does fail. Testing this hypothesis requires further axial-impact experiments to determine if there is an intermediate energy where a container with low pressure container sustains more damage than one with high pressure.

It is important to note that the container was fully constrained during this test program. This condition forced buckling of the container, which is unrealistic in the case of an actual vehicle crash. It is unlikely the container mounting assembly or associated tubing connections would withstand the peak forces in Table 2-7, therefore making it unlikely that the container in an actual crash would be subjected to the full forces measured in this experiment.

This study addressed blunt impact, not penetration. Odegard and Thomas [18] impacted containers with cylindrical projectiles. In summary, their results are that an energy of 2 to 8 kJ (in the “ low” range of Table 2-1) is sufficient to penetrate a container if the projectile is 10 to 50 mm diameter.

The containers this project were free of manufacturing defects and prior damage. A container with a defect would be more likely to rupture.

27

3.0 LEAK RATE MEASUREMENTS AND CHARACTERIZATION

The second of the three technical tasks was the measurement and characterization of controlled leaks of hydrogen and helium from fuel containers. Unlike a spilled liquid fuel, a leaked gas does not accumulate on the pavement of the crash test site. The permissible leak from a CNG container in an FMVSS No. 303 crash test is specified in terms of a pressure drop in the container. The purpose of this technical task was to relate the amount of gas flowing from a hydrogen fuel container to a measurable pressure drop in the container, with special consideration on the accuracy of the measurements and the scaling equivalence between hydrogen and helium.

3.1 Summary of the Leak Rate Characterization

Hydrogen, at initially high pressure, was leaked through a known orifice in exactly the same conditions as less hazardous surrogates: low-pressure hydrogen and high-and low-pressure helium. The resulting data allowed direct comparison of the total mass flow and changes in pressure and temperature for the less hazardous surrogate conditions.

A full matrix of controlled leak experiments consisting of 48 separate cases was conducted. Hydrogen and helium, beginning at high and low pressures, was allowed to flow through fixed orifices for a period of at least 60 minutes. Temperatures and pressures were measured throughout, and mass flowmeters measured the gas as it exited.

The procedure in Appendix A of SAE J2578 for calculating the mass of leaked hydrogen from a pressure measurement proved to be accurate. Parallel formulas, for running a case in a surrogate condition and then inferring what the leaked mass of high-pressure hydrogen would have been, showed evidence of a bias error. The bias is due at least in part to an effect neglected by the formulas. The change in compressibility of the gases over the extreme range of pressures is appreciable. The formula for the low-pressure hydrogen case seemed satisfactory when it was modified to account for the pressure-dependent compressibility. The results for the high-pressure helium case were improved when this correction was included, but a remaining bias of a few percent was not fully explained.

The formulas of SAE J2578 were valid at the approximate leak rates anticipated in the standard. Measurement of temperature and pressure was not a viable method of quantifying leak rates two orders of magnitude below those rates.


The objective of these experiments was to experimentally confirm the approach of measuring pressure drop and scaling flow rates in Appendix A of SAE J2578 [22].

One way to assess the integrity of the fuel system following a compliance crash test is to measure the drop in its internal pressure. SAE J2578 provides formulas for determining whether the amount of leaked hydrogen exceeds a certain threshold by monitoring the pressure over a prescribed time period. Recognizing the danger of conducting compliance tests with a container

28

filled to high pressure with hydrogen, the standard provides alternative formulas for conducting the tests with hydrogen at a low pressure or helium at a high pressure. The standard contains three sets of steps for measuring a pressure drop and then estimating the mass of hydrogen that would leak from a high-pressure container through a given orifice (See Figure 3-1).

Formula 1: High-pressure hydrogen (Equations A-3, A-4, A-6, and A-8) Formula 2: High-pressure helium (Equations A-10, A-11, and A-12) Formula 3: Low-pressure hydrogen (Equations A-13, A-14, A-15, and A-16) (J2578 has no formula for low-pressure helium.)

In addition, the Japanese Blue Book [12] provides a pass criterion based on a pressure measurement.

Figure 3-1. Appendix A of SAE J2578 provides formulas for estimating the mass of leaked hydrogen in three possible fueling conditions for crash tests. The purpose of these experiments was to assess the formulas with data from simulated leaks.

Hydrogen hazardous if released

Helium harmless in small quantities

high pressure rupture is hazardous

low pressure less hazardous

high pressure rupture is hazardous

low pressure less hazardous

Leaked mass What matters for fire hazard but hard to quantify

Crashes on the street will be this case

Decreased pressure meaningless by itself but possible to measure


The purpose of these experiments was to generate pressure and flow rate data under a variety of controlled, representative leak conditions and to determine whether the scaling of low-pressure leaks and helium leaks to represent high-pressure hydrogen leaks is appropriate as formulated in Appendix A of SAE J2578.

This experiment characterized leak rates to aid NHTSA in developing acceptance criteria for potential future fuel system integrity tests, where hydrogen fuel vehicles may be subjected to crashes similar to those in FMVSS Nos. 301 and 303. The goal was data and an understanding of a container’s pressure drop in relation to the amount of gas that has flowed through a simulated leak. The leak rates were selected to be representative of those that might be considered in a hydrogen vehicle crash test. As the tests for both FMVSS Nos. 301 and 303 use non-flammable simulants, data were generated both for hydrogen (the fuel itself) and for helium (the likely non-flammable simulant).

Fo

rmu

la 1

?

Formula 3?

Formula 2?

29

3.2.2 Key Assumptions of SAE J2578, Appendix A

SAE J2578 starts with the energy content of the two fuels and calculates a mass of hydrogen, 606 grams, having the same energy as the amount of gasoline permitted to leak in FMVSS No. 301. The standard then goes on to show how pressure transducers can be used to ascertain whether the amount of hydrogen lost from a fuel container exceeds this threshold.

SAE J2578 establishes a maximum amount of hydrogen that can leak from a vehicle during a compliance test. This standard also describes how high pressure helium and low pressure hydrogen can be substituted for hydrogen at pressures near the container service pressure. The document’s assumptions permit no more than 606 g of hydrogen to leak from the vehicle when the container initial temperature is 15° C and the container initial pressure is at its service pressure. The document provides formulas that compensate for experiments that are not performed precisely at 15° C or at the container service pressure. Mass loss is calculated from initial pressure and final temperature and pressure measurements of the container. From these measurements, density is calculated from a quadratic equation that has been fit to the actual density of the gas over the range of expected pressures. These density equations closely match the material property values from the National Institute of Standards and Technology (NIST) [19], with a relative error of around 1 percent or less over the range of interest. Equation A6 in SAE J2578 for high pressure hydrogen is valid from 170 bar (17 MPa) to 870 bar (87 MPa), as shown in Figure 3-2. The standard’s Equation A11 for helium is a sound approximation from 140 bar (14 MPa) to 750 bar (75 MPa), as shown in Figure 3-3. The standard’s equation A15, part of its low-pressure hydrogen formula, is valid from less than 1 bar (0.1 MPa) to 270 bar (27 MPa), as in Figure 3-4. Finally, the Japanese Blue Book standard contains a quadratic fit to the density for helium gas, which is compared with NIST values in Figure 3-5. From these formulas, the allowable mass loss of the test gas can be calculated for a particular condition.

30

Figure 3-2. The approximation for the density of high-pressure hydrogen at 15° C in SAE J2578 agrees well with NIST values.

Figure 3-3. The approximation for the density of high-pressure helium at 15° C in SAE J2578 agrees well with NIST values.

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Figure 3-4. The approximation for the density of low-pressure hydrogen at 15° C in SAE J2578 agrees well with NIST values.

Figure 3-5. The approximation for the density of helium in the Japanese standard agrees well with NIST values.

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A prior report to NHTSA on post-crash hydrogen leakage limits and fire safety research [5] addressed the first of the two topics in Appendix A of SAE J2578—the consequences of leaks of given amounts. The report concluded that the “ leak rate is not the most important metric, but instead, the volume of hydrogen leaked into car compartments.” Undesirable consequences occur if the concentration is allowed to “accumulate locally to ~5%, a level just exceeding the lower flammability limit of hydrogen, ~4%. It appears to be unimportant if this lower flammability limit is reached via a low leak rate after a long duration (up to 60 min) or a high leak rate (up to 118 lpm) over a very short duration.”

The present experiments were intended to produce data to assess the second topic—the appendix’s formulas for calculating leak rates from pressure loss. The question is the suitability of using a pressure transducer to ascertain leak rate or (its integral) total leaked mass.

The SAE J2578 standard explicitly ignores any hydrogen in the low-pressure portion of the system [24, page 35, note a] and focuses on the high-pressure containers.

3.3 Test Conditions

The experiment attempted to establish the conditions envisioned in SAE J2578—hydrogen or helium leaking through variously sized orifices that established anticipated leak rates. The experimentally measured pressure drops and flow rates were compared with predictions from established theory of compressible gas flow using real material properties and with the formulae of SAE J2578. The complete test plan is in Appendix C.

3.3.1 Technical Approach

If a container is damaged in a hypothetical crash, certain leak paths may be created from the interior to the atmosphere. For a given leak path, the mass flow rate of leaking hydrogen will depend on the conditions in the container, most notably the pressure. A higher internal pressure will cause a higher leak flow rate. As the leak progresses, the pressure will decrease and the flow of hydrogen will slow. If the container is filled with a simulant gas (i.e., helium), the flow rate will be different.

A leak path in a damaged container could conceivably be a pinhole with uniform geometry, or it could be a tortuous, irregular path through the fiber windings and epoxy matrix. Engineers studying fluid flow have found that a leak path can ultimately be described by an equivalent diameter and a shape factor (called a “discharge coefficient” ). These two quantities can be combined to form a single parameter that characterizes the leak path’s effect on the flow rate. The leak path in these tests were simulated by fixed orifices that were selected to yield particular gas flow rates. One can imagine the leak path in a composite container changing somewhat during the leak, due to flow wearing away loose pieces in the path or by the geometry changing as the stress in the wall is relieved. Nevertheless, these effects were neglected and the orifices remained fixed throughout the experiments.

When a gas flows from a high-pressure region through an orifice to a region of much lower pressure, the flow rate of the gas molecules will reach the gas’s speed of sound. This condition can be described as “sonic” or “Mach 1” or “choked flow.” The mass flow rate for a particular

33

gas in choked flow is determined by the properties of the orifice and the upstream stagnation conditions, which in these experiment were the pressure and temperature in the container. (This is different from the orifice flow meter, where the volume flow rate can be calculated by measuring the pressure upstream and downstream of an orifice of known properties, but where the flow rate through the orifice is well below the speed of sound.) The mass flow rate of choked flow can be predicted from an equation that is found in textbooks on compressible fluid flow [10],

)1()1(

1

2 −+

��

��

�

+=

γγ

γγ

TRZ

MPACm

γ

wD� (1)

where

= mass flow rate CD = discharge coefficient A = area of the orifice P = upstream (stagnation) pressure γ = ratio of specific heats Mw = molecular weight Z = compressibility Rγ = universal gas constant T = upstream (stagnation) temperature.

The work plan [21] for this task order anticipated that the leak rates in this experiment would range from permeation levels up to leak rates similar to those observed in the crush tests of Chapter 2.0. Some of those test cases had no leak at all; those with the breached container vented the entire contents in less than 5 s (See Table 2-7). Unfortunately, the crush tests did not provide a “representative” leak rate from a damaged container. Appendix A of J2578 [22, on page 35 following Equation A2] does provide guidance on a representative leak rate. Calculating the equivalent energy leak rate from the permissible mass leak rates for liquid fuel in FMVSS No. 301, the corresponding flow rate of hydrogen is approximately 117 standard liters per minute. Battelle decided that the target “medium” flow rate for the leak tests would be this value. The “high” flow rate was to be twice this value or 234 standard liters per minute. The “ low” or “permeation” level was 3 sL/min, the flow rate calculated to bring the volume of a mid-size passenger compartment to the lower flammability limit over an hour [5, page 108].

Four fixed orifices, designated small, medium, large, and extra large, were used to establish the flow rates. The diameters of the orifices and the predicted initial flow rates are listed in Table 3-1. Four flow rate entries in the table are shaded light blue. They are the design flow rates that controlled the experiment. The orifices were sized according to Equation 1 to provide these flow rates for hydrogen at 700 bar (for three cases) and 350 bar (for one case). The other flow rates in the table were then calculated from the respective conditions’ gas and initial pressure. The actual diameters of the orifices, which differed from the nominal sizes, are listed in the table. The hole size through the thickness varied slightly; the choking condition was assumed to occur at the smallest actual diameter, which is stated in the table. Scanning electron micrographs of the orifices are in Figure 3-6. Where possible, the diameter was calculated by picking coordinates

34

around the edge of the orifice using an optical microscope and calculating the best fit circle. For the small orifice, the diameter was measured on the scanning electron microscope since the optical microscope did not have enough resolution to clearly view the edge of the small orifice. The orifices were laser cut in rupture discs. The discs were retained in high-pressure safety heads during the experiments.

Table 3-1. Four fixed orifices established the leak rates. Nominal orifice sizes were selected to provide the flow rates in the light blue cells. Flow rates in white cells were predicted from nominal conditions. Actual orifice diameters were measured with a microscope. Actual initial flow rates are discussed in Section 3.4.2.

Orifice Predicted Initial Leak Rate, standard liters per minute

Designation Diameter (µ)

700 bar 350 bar 70 bar 35 bar

H2 He H2 He H2 He H2 He nominal actual

S 14.5 22 3.00 2.35 1.64 1.26 0.356 0.267 0.179 0.135

M 90.9 95 117 91.7 13.9 10.4

L 128 135 234 183 128 98.1 27.8 20.9 14.0 10.5

XL 174 192 234 179 25.6 19.2

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Figure 3-6. Scanning electron microscope images of the (1) small, (2) medium, (3) large, and (4) extra large orifices.

3.3.2 Test Matrix

This experiment had a full-matrix design:

Test medium: hydrogen and helium gas – To observe the relationship between the two gases. While hydrogen is the fuel, the crash testing will most likely be conducted with helium as a non-flammable and inert alternative with similar flow behavior

Container type: Type 3 and Type 4 containers – To observe the different temperature response of the two construction types

Container service pressure: 350 and 700 bar – To observe the effect of container construction (wall thickness) and varying state of charge

Container initial state of charge: 10% and 100% service pressure – To observe the effect of state of charge on the flow behavior

Flow (Leak) rates - low, medium, and high flow rates – To observe the pressure drop and gas temperature of the container at varying flow rates.

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Table 3-1 has 24 entries of predicted initial flowrate. All cases in the table were executed once with a Type 3 container and again with a Type 4 container, for a total of 48 test cases.

The four containers used for these experiments had the same specifications as those used for the dynamic impact tests, which are in Table 2-5. These containers, at 58 to 66 L nominal water volume, are roughly half the smallest size that are likely to be used in production fuel cell vehicles. A given leaked mass, therefore, will produce about twice the change in container pressure that would be expected in a compliance test on a production vehicle. Thus, detecting and quantifying a certain leak in these experiments is easier than it would be in a compliance test.

3.3.3 Instruments

The pressure in the container was monitored with a pressure transducer. Separate transducers were used for cases beginning with 100% or 10% of service pressure. Internal and external thermocouples measured the gas temperature. The flow rate in most cases was measured with a coriolis type mass flow meter. The advantage of a Coriolis meter is that it directly measures mass flow; there is no conversion factor between a calibration gas and hydrogen. The flow rate in the cases with the smallest orifice was too low to use the coriolis meter, so a laminar flow meter was used instead. This meter measures a volumetric flow and internally converts to a mass flow by measuring the pressure and temperature of the gas.

The instruments that measured temperature and pressure immediately upstream and downstream of the orifice are shown in Figure 3-7. Temperature and pressure measurements immediately upstream of the orifice were not significantly different than the stagnation temperature and pressure. Calculations of the kinetic energy and head loss between the tank and the orifice indicated that these properties would have a negligible affect on the pressure and temperature of the gas entering the orifice.

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Swagelok ¼” Needle Valve (Fully open during test)

Orifice (Fully open): 4.4 mm

All tubing ¼” 316SS, 1.7 mm wall, 3.0 mm ID

Leak Orifice (Table 3-1)

Swagelok ¼” Tee Orifice: 4.8 mm

Swagelok ¼” FNPT Branch Tee Orifice: 4.8 mm

Figure 3-7. Temperature and pressure measuring instruments immediately upstream and downstream of the orifice.

The complete list of instruments for these tests is listed in the test plan in Appendix C.

3.4 Results

All 48 cases in the full matrix were conducted as planned. The flow rates and thermodynamic conditions were close to what was expected, within reasonable measurement errors. Pressure and temperature measurements proved to be adequate to estimate total mass flow in some instances but not in others. In cases where the total mass flow was a small fraction of the mass in the container, the error attributable to the pressure and temperature sensors was significant. Due to the magnitude of this error, the flow rate in these cases cannot be accurately characterized by pressure and temperature measurements alone.

3.4.1 Raw Data

Container stagnation pressure and temperature were measured through the entire duration of each case. Sample plots of the container temperature, container pressure, and measured mass flow rate

38

versus time are shown in Figure 3-8 through Figure 3-13. Plots of all 48 cases are in Appendix D. All cases began with the flow exiting through the Coriolis meter, which was more accurate in most instances. Flow was switched to the laminar flow meter when its accuracy was better than that of the coriolis meter. Which meter was active is indicated by the color of the flow rate curve. A typical plot of the pressure and mass flow rate time history is shown in Figure 3-8. The magnitude of the container pressure and mass flow rate decrease at similar and smooth rates. The internal container temperature for this same case, plotted in Figure 3-9, dropped considerably during the leak, as should be expected.

Figure 3-8. Mass flow rate and container pressure (hydrogen, Type 3 container, 350 bar, 135 µ L orifice).

Figure 3-9. The hydrogen in the container cooled during the leak of the case in Figure 3-8.

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Figure 3-10 depicts the mass flow rate and container pressure during a test with a low flow rate. The container pressure and mass flow rate remained nearly constant during such tests.

Figure 3-10. Mass flow rate and container pressure (hydrogen, Type 4 container, 70 bar, 22 µ S Orifice).

In Figure 3-11, the mass flow rate measured by the Coriolis mass flow meter wavers. This phenomenon, confirmed by the manufacturer, is likely due to the measurement error of the flow meter towards the low end of its measurement range.

Figure 3-11. Mass flow rate and container pressure (hydrogen, type 3 container, 35 bar, 135 µ L Orifice).

In some cases, as shown in Figure 3-12, flow was diverted from the Coriolis mass flow meter to the laminar mass flow meter part way through the test. Changing flow meters decreased the amount of measurement error because the laminar flow meter had better accuracy than the

40

Coriolis meter at low flow rates. The maximum measurable flow rate of the laminar flow meter precluded its use for the entire experiment.

Figure 3-12. Mass flow rate and container pressure (hydrogen, type 3 container, 70 bar, 95 µ M Orifice). Flow was switched from the Coriolis flow meter to the Laminar flow meter about 100 min into the leak.

During the experiments with the greatest mass flow rates, the measured mass flow rate did not decrease smoothly with time at the beginning of each test. An example is shown in Figure 3-13. When this occurred, only the data in the smooth portion of the mass flow rate curve was used to characterize the orifice.

Figure 3-13. Mass flow rate and container pressure (hydrogen, Type 4 container, 350 bar, 192 µ XL Orifice).

Experimental error must be considered. The manufacturer’s datasheets for each sensor were consulted for the amount of measurement error. The pressure sensors were subject to

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±0.5 percent error full-scale. The temperature sensors were assumed to have a total error of ±3.2° C due to the manufacturer’s specifications and a small allowance for non-uniform temperature within the container. The Coriolis flow meter had an accuracy of ±0.083 g/min. The Laminar flow meter accuracy was a function of flow rate. For our experiments, this accuracy was calculated to be less than ±0.010 g/min. The measurement error was significant during experiments with low mass flow rates. The container temperature and pressure changed very little during the cases with low flow rates. Sometimes this change in container pressure and temperature was within the measurement error of the sensor. Figure 3-14 illustrates the calculated initial and final mass of hydrogen in a container for three high pressure hydrogen experiments, conducted with the Type 3 container. The experiment shown on the left had a relatively low flow rate. The experiment shown on the right had a relatively high flow rate. The experiment in the middle had a medium flow rate. Error bars indicate the amount of error attributable to the measurement equipment used during these experiments. For the low flow rate case, the range of measurement error is greater than the measured change in mass. For the other two cases, the measured mass loss was much greater than the experimental error. The measurement error attributable to the mass flow meters was also more significant at low flow rates.

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Figure 3-14. Mass in containers calculated from temperature and pressure measurements.

3.4.2 Initial Flow Rates

As a check of the self-consistency of the data, the initial flow rate of each case was compared with the initial flow rate predicted by Equation 1. The initial flow rates are plotted against the predicted flow rates in Figure 3-15 and Figure 3-16. In nearly all cases, the measured initial flow rate was less than the predicted initial flow rate. A discharge coefficient of 1 in Equation 1 had been assumed for the compoarison, and a discharge coefficient is usually less than unity, so this was to be expected. Applying a discharge coefficient that is less than unity would predict lower flow rates more representative of what was measured.

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Figure 3-15. Measured and predicted initial flow rates of the hydrogen experiments.

Figure 3-16. Measured and predicted initial flow rates of the helium experiments.

An effective discharge coefficient was calculated from the measured initial mass flow rate, measured gas properties, and known orifice size. The initial discharge coefficient for each case is plotted in Figure 3-17 and Figure 3-18. The discharge coefficient is plotted as a function of upstream stagnation pressure. The results presented in Figure 3-17 and Figure 3-18 illustrate that the discharge coefficient of each orifice was not constant for all of the experiments that were performed.

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Figure 3-17. Discharge coefficients, calculated using the initial flow rates of the hydrogen experiments.

Figure 3-18. Discharge coefficients, calculated using the initial flow rates of the helium experiments.

3.4.3 Mass Estimation From Pressure and Temperature

These experiments had a mass flowmeter on the outlet to measure how much hydrogen escaped through the simulated leak. In a compliance test, the location of a leak would not be known in advance, precluding use of a flowmeter to measure the flow directly. The mass of escaping hydrogen would have to be estimated from measurements of temperature and pressure. The analysis that follows presents several approaches to doing so, and in each case the estimated

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mass lost is determined using only pressure and temperature measurements, simulating the data available during a compliance test.

The analysis below estimates the mass of hydrogen that escaped from a container that was initially pressurized with hydrogen to its full service pressure. First, a strictly empirical formula without any reference to theory is derived. Then the two existing standards, Appendix A of SAE J2578 [24] and the Japanese Blue Book [12], are evaluated. Their aim was to establish a criterion as to whether a container passes a test because the mass leaked was below a specified amount. Their formulas can be critiqued in light of the data and analysis of this set of measurements.

Before beginning the assessment of the pressure and temperature formulas, the mass estimates from the pressure and temperature sensors are compared with the output of the mass flow meters.

3.4.3.1 Comparison of Mass Flow Meter and Gas Density Change

The instrumentation of this experiment allowed the amount of mass loss during each case to be quantified by two different methods. The mass loss could be calculated from tank pressure and temperature measurements taken at the beginning and end of the experiment. With these measurements, the density of the gas could be determined using the published NIST values [19]. The actual mass is determined by multiplying the density by the nominal volume of the container. An alternative method of determining the total mass loss is based on the mass flow rate measured by one or both flowmeters. The mass flow rates, integrated over the duration of the experiment yield the total mass loss. The calculated mass loss using these two methods is compared in Figure 3-19 through Figure 3-21. The data is shown in three figures to allow appropriate scaling of the axes for the wide range of measured mass losses.

These figures show that the measurements are self-consistent, within the nominal accuracy of the instruments. The pressure and temperature measurements are more reliable than the flowmeter measurements, which were subject to vagaries including those in Figure 3-11. Therefore, the amount of leaked mass that will serve as the basis for comparison in all following analysis is the mass calculated from the pressure and temperature measurements of the container, at the relevant initial and final moments. Densities were determined from published NIST values [19].

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Figure 3-19. Comparison of mass leak calculated from flow meter and measurements of pressure and temperature for high-pressure experiments with M, L, and XL orifices.

Figure 3-20. Comparison of mass leak calculated from flow meter and measurements of pressure and temperature for low-pressure experiments with M, L, and XL orifices.

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Figure 3-21. Comparison of mass leak calculated from flow meter and measurements of pressure and temperature for experiments with the 22-µ S orifice.

3.4.3.2 Mass Loss Estimated by a Purely Empirical Formula

The next step will be to compare the leaked mass of hydrogen with the predictions in Appendix A of SAE J2579 and the Japanese standard. First, though, as an example for comparison, a set of purely empirical relations between pressure and temperature changes and the corresponding mass leakage was developed. Separate relations were developed for each of the four estimations on page 27. These highly linearized approximations, developed without reference to theory, have the form of Equations (2) and (3).

(2)

(3) where ∆m = amount of mass leaked K = constant derived from the experimental results for each gas P = stagnation pressure V = container volume (water capacity) ρ = density of gas from NIST [19] using P and T T = stagnation temperature ( )H = high pressure hydrogen ( )S = surrogate gas ( )B = beginning of case

( )E = end of case

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The least squares method was used to determine the coefficients that best predicted the mass loss of an analogous high pressure hydrogen experiment. Due to the large measurement error of container pressure and temperature change for low-flow-rate experiments, the small orifice results were omitted from this analysis. The measurement error did not allow the surrogate mass loss to be determined accurately. The KS value for low-presure hydrogen was 0.88. The values for high- and low-presssure helium were near each other, 0.60 and 0.58, respectively. The accuracy of the empirical formula in predicting the leaked mass of high pressure hydrogen from surrogate gas experiments is shown in Figure 3-22.

Figure 3-22. Predicted leaked mass of high-pressure hydrogen using Equation 2.

The mass of leaked high pressure hydrogen predicted by Equation 2 had a maximum error of 9 percent for both high- and low-pressure helium, and 12 percent for low-pressure hydrogen. This error can be attributed to measurement inaccuracy and nonlinearities of the gas. The derived equation assumes a linear relation between initial pressure and leaked mass. The flow rate over the entire length of the experiment is affected by the instantaneous pressure of the container. At higher flow rates, the container pressure decreases more rapidly. Also, properties of the tested gas, such as compressibility and the ratio of specific heats, which appear in Equation 1, are nonlinear with respect to pressure changes. These nonlinear effects are not captured by scaling the total mass flow by initial conditions of the container.

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Compares the leaked mass predictions of the purely empirical formula with the actual mass as determined using the NIST values for density. The following two subsections will compare the predictions from the formulas of the existing safety standards to the actual mass determinations.

3.4.3.3 Mass Loss by Formulas in Appendix A of SAE J2578

As noted in Figure 3-1, Appendix A of SAE J2578 provides three formulas for determining whether a leak is within a passing criterion. The three formulas will be examined in order. The first formula determines the amount of leaked hydrogen from a beginning pressure and temperature measurement and an ending pressure measurement. It is identical in principle, though not in detail, to the method used in this report to determine the actual mass loss. The most notable difference is that SAE J2578 uses a curve fit to a density relation while this report uses the NIST values. As shown in Figure 3-2, the density approximation for high-pressure hydrogen is quite good, so the predictions of SAE J2578 should be expected to be close. Indeed they are, as shown in Figure 3-23. TAML is based on terminolgy in SAE J2578—it is the time into the test where the mass flow reaches the allowable mass leakage. The mass of the failure criterion in the standard is nominally 606 g, but it is subject to adjustments for initial temperature and pressure.

Figure 3-23. Comparison of amount of hydrogen leaked when a surrogate gas predicted non-conformance to the non-conformance criterion.

The other two formulas in Appendix A of SAE J2578 use leaks under surrogate conditions (high-pressure helium and low-pressure hydrogen) to determine whether the allowable leakage mass would have been reached had the case been run with high-pressure hydrogen. The procedures in the standard were followed to determine the time of allowable mass leakage for each of the test cases. The mass loss in the corresponding high-pressure hydrogen case was calculated at this same time. If the SAE formulas had no error, then the high pressure hydrogen mass loss at the determined time would be equal to the allowable mass loss in the high-pressure hydrogen case.

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For example, the case of the 95-µ-diameter orifice for low-pressure hydrogen in a 350-bar Type 3 container was compared with the actual leaked mass in the case of the same orifice and same container but with high-pressure hydrogen. The accuracy of the SAE approach for predicting high-pressure hydrogen leakage from surrogate gases is plotted in Figure 3-24.

Figure 3-24. The amount of hydrogen leakage was predicted via the two formulas for surrogate conditions in Appendix A of SAE J2578. This amount is compared with the loss in the corresponding experiment with high-pressure hydrogen. Predictions from high-pressure helium are within about 10 percent, but they are all above the Y=X line. Predictions from low-pressure hydrogen are with about 20 percent, all but one below the Y=X line.

The predictions of the formulas in Appendix A of SAE J2578 ranged from approximately zero up to 10 percent for high-pressure helium. The predictions for low-pressure hydrogen were from just above zero down to about -20 percent. Considering measurement inaccuracies and necessary approximations, these predictions are reasonable. In fact, when the mass loss from pressure and temperature changes were compared with mass flow meter measurements in Figure 3-19 and Figure 3-20, the discrepancies were as much as 10 percent. The next observation to make from Figure 3-24 is that both formulas have a bias error. Spoken simply, all of the blue markers are above the line and nearly all of the green markers are below the line. The bias appears to be due to the formulas’ neglecting the changes in gas properties that occur over hundreds of bar change. By Equation 1, the mass flow rate depends on the upstream pressure when all else is held constant. The formulas take the flow to vary linearly with the pressure, which is a suitable approximation when the pressure changes are small and the variations in other properties can be neglected. The compressibility of both hydrogen and helium varies appreciably over the range of pressures in these experiments, as indicated by the curvature in the pressure-density relations in Figure 3-2 and Figure 3-3. For hydrogen at 15° C, the compressibility at 70 bar is 30 percent less than at 700 bar. (The mass flow rate also depends on the specific heat ratio, but its contribution to the discrepancy is much smaller. Similarly, though

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the viscosity does not appear explicity in Equation 1, it does affect the discharge coefficient but the contribution is not significant here.) A possible compressibility correction factor is shown in Equation (4). This correction factor was multiplied by the Allowable Mass Leakage (AML) calculated by the SAE formulas to obtain a new AML that accounted for compressibility changes. The correction factor assumes that the ratio of compressibilities is constant during the entire experiment and uses the initial compressibility of the gas in Equation (4). In cases where the pressue falls significantly during the test, even this assumption is an approximation.

(4)

where

KZ = compressibility correction factor Z = compressibility factor at the temperature and pressure from NIST [19] SP = tank service pressure TR = reference temperature (15° C used in Appendix A of SAE J2578) IP = initial pressure TS = initial temperature

The results of Figure 3-24 were adjusted for compressibility changes and replotted as Figure 3-25.

Figure 3-25. This is the same data as in the pevious figure, but the formulas in Appendix A of SAE J2578 have been adjusted for changes in compressibility of the gas. Both have been improved, but evidence of a bias in the formula for high-pressure helium remains.

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Accounting for compressibility changes due to pressure results in a more accurate correlation between low- and high-pressure hydrogen results. A small bias remains in the predictionsn for high-pressure helium. This bias may be to be due to the approximation for converting an amount of hydrogen leaked to a similar amount of helium leaked, Equation A9 in Appendix A of SAE J2578. It may be due to the formula’s not accounting for the change in temperature of the gas in the container during the leak. Its source was not further explored.

3.4.3.4 Mass Loss Estimated by Japanese Blue Book Formulas

The Japanese Blue Book [12] establishes a maximum amount of hydrogen that can be leaked during a compliance test. If the mass leak rate of hydrogen exceeds 131 NL per minute averaged over a 60-minute test, the vehicle fails the compliance test. This amount of mass leaked is less conservative than the SAE standard of 606 g. A 131-NL-per-minute leak would amount to 700 g over the course of a 60-minute test. The Japanese Blue Book standard provides an equation to compensate for stagnation temperatures different from 0° C, where a normal liter is defined. The standard assumes that the initial stagnation pressure is representative of the service pressure of the container, and provides no compensation for container pressure. Mass loss is calculated from temperature and pressure measurements of the container. Similar to the SAE formulas, density is calculated from a quadratic equation that has been fit to the actual density of the gas over the range of expected pressures. Once the amount of mass leaked has been calculated, the Japanese standard provides an equation to convert helium mass loss to hydrogen mass loss if helium was used as the surrogate gas. There are no separate equations for lower pressure surrogate gases. The Japanese formulas do not indicate how to compensate for experiments started at different pressures, initial starting points in the experimental data were chosen where the temperature-compensated-pressure was nearly equal between the helium and analogous hydrogen experiment. Except for the small orifice experiments, where the pressure did not change considerably over the length of the experiment, equal pressures within 0.1 percent were chosen as the starting pressure for the experiment. The mass loss that occurred during the 60 minutes following this pressure measurement was calculated and converted to NL per minute. The accuracy of the Japanese Blue Book formulas is shown in Figure 3-26 through Figure 3-28. The data is shown in three figures to allow appropriate scaling of the chart axes for the wide range of measured and predicted mass losses.

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Figure 3-26. Comparison of predicted and measured hydrogen mass flow rates for high pressure experiments with M, L, and XL orifices.

Figure 3-27. Comparison of predicted and measured hydrogen mass flow rates for low pressure experiments with M, L, and XL orifices.

Figure 3-28. Comparison of predicted and measured hydrogen mass flow rates for experiments with S orifices.

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According to the experiments performed in this study, the accuracy limits of the Japanese Blue Book formulas are within 20 percent near the failure criterion of 131 NL per minute. The accuracy of the predictions remains is 20 percent except where very low flow rates are observed. The increased inaccuracy is likely due to increased measurement error, precluding a precise calculation of mass in the container.

3.5 Conclusion

The leak experiments confirm that the formulas presented in Appendix A of SAE J2578 and the Japanese Blue Book are reasonable approximations. The formula for high-pressure hydrogen, based solely on a well-matched curve fit of the pressure-density relation, gives results as well as can be achieved by this approach. The formulas for the two surrogate conditions, high-pressure helium and low-pressure hydrogen, are close but not as good. By the data of this experiment, a bias error was evident in each case. The bias was removed almost completely from the formula for low-pressure hydrogen by accounting (in part) for the change in compressibility of hydrogen over the 10-fold pressure difference between the low- and high-pressure cases. Making this same adjustment for the high-pressure helium formula improved the situation, but some bias remained. The formulas in SAE J2578 include an adjustment for the starting temperature of a test. The gas in the container cools considerably as the pressure drops, but the formula provides no adjustment for the internal container temperature at the end of a test. The effect of this assumption was not explored. A new formula for based on a test with low-pressure helium could be developed using the same principles as the existing formulas. The experimental matrix was fully completed with both Type 3 and Type 4 containers. No evidence of the effect of container construction was observed in the flow data. A research question at the outset was whether the metal liner of a Type 3 container permit heat conduction to the outside and affect the results. Data from the two container constructions was plotted separately in every figure; but, if any such effect is present, it is smaller than other effects. The repetition did serve as source of extra data for analysis and consistency checks. The leak rates from the smallest orifices, as measured by the mass flowmeter, were close to predictions, so the equipment can be taken to have been functioning properly. However, the minuscule changes in pressure over the duration of the tests were too small to be reliably measured. If a much longer hold period is impractical, then a different approach to detecting these leaks may be needed.

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4.0 CRASH TESTS OF MOCKUP HYDROGEN VEHICLES

The culmination of this project was a series of crashes of a passenger car with a mockup hydrogen fuel system. The pressure-holding ability of the fuel systems were measured following the crashes, and the fuel systems were inspected for damage.

4.1 Summary of the Crash Tests

Three model year 2009 Honda Civic GX natural gas vehicles were purchased for this experiment, and their CNG fuel systems were removed and replaced with similarly sized hydrogen fuel systems. These were tests of the mockup fuel systems, not of the vehicles, which were merely a test bed. The three vehicles were subjected to front, rear, and side crashes, as in FMVSS No. 301.

As the crashes were planned, details were selected to provide the greatest challenge to the hydrogen fuel systems, consistent with the test procedures. The pressures for the containers were chosen to be those that showed the most damage to the container in the crush tests of Section 2.0: The front and rear crashes, where the impact is essentially lateral against the transversely mounted container, used a pressure of 10% service pressure. The container in the side crash, with the potential for an axial impact to the container, was pressurized to 100% service pressure. The impact offset for the rear crash was selected to be to the driver’s side, which is the side where the fuel port and solenoid valve for the container are. In the only significant departure from FMVSS No. 301 testing, the impact point on the side crash was aimed aft of its usual location (See Figure 4-1), so the impact would be against the fuel port and have a greater chance of damaging the plumbing.

The deformations and accelerations observed in this experiment were typical of similarly sized vehicles in these tests. The front crash compressed the engine to the front axle and the peak vehicle acceleration was 28 g. The moving deformable barrier rode up on the rear bumper of the test vehicle and the behavior of the crumple zone was again typical. The penetration of the moving deformable barrier in the side crash was limited by the B pillar and the rear axle, though the door panel was deformed a maximum of 162 mm.

As was expected, the front crash caused little damage to the mockup hydrogen fuel system because the container is in the rear of the vehicle. The low-pressure fuel line leading to the engine was bent by the impact, but there was no evidence of any leaking. Though the side crash deformed the vehicle body in the vicinity of the fuel container, damage was again not severe. The container was displaced in its mounts and nearby fuel lines were deformed, but the fuel port was functional following the crash. After passing through the crumple zone in the trunk, the moving deformable barrier in the rear crash directly struck the fuel container. The rear crash caused extensive damage to the fuel system mounting and deformed much of the plumbing. The container surface was abraded in several locations. The pressure integrity of all three fuel systems was intact following the crashes.

These crashes show that a hydrogen fuel system can be protected from the kinds of crashes specified in FMVSS No. 301. Damage to the fuel system in the rear crash appeared to be extensive by visual observation, but no leaked hydrogen was detected and the system could hold

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pressure following the crash. The impacts of the crush tests showed that the container is robust even if all the the crash energy is directed to the container. The crash test stress the importance of building the plumbing so that it can tolerate deformations on the order of tens of millimeters (several inches).


The behavior of the vehicle itself in FMVSS No. 301 crashes is well documented through the many compliance tests run by NHTSA, whose results are published [18]. These three crashes focused on the behavior of the fuel systems—the container, its mounting, and the associated piping.


The purpose of the full-vehicle crash tests was to demonstrate the behavior of a hydrogen container and its mounting in a crash test. Accelerations and deformations were measured to quantify the conditions endured by the test vehicle and more importantly the container during a crash test.

4.2.2 Context of the Crash Tests

NHTSA’s existing regulations for fuel system safety, FMVSS No. 301 for gasoline and other liquid fuels and FMVSS No. 303 for compressed natual gas, provide for three crash tests—impacts to the front, side, and rear of a vehicle. Three vehicles were outfitted with containers for compressed hydrogen, and crashed, one each on the front, side, and rear, in crashes patterned after NHTSA’s test procedures for FMVSS No. 301. The primary focus in the tests was to determine the behavior of the mockup hydrogen fuel system and its supporting structures; the vehicles in the crashes were ballasted with lead weights and sand bags as a surrogate for the dummies.

Because the purpose of these tests is to understand the consequences to a hydrogen fuel system in a crash, the crash conditions selected were those in which the fuel container is most vulnerable. In NHTSA’s compliance tests the fuel container is filled with an inert material (Stoddard solvent instead of gasoline, or pressurized nitrogen instead of CNG), but these crashes were conducted with pressurized hydrogen in each vehicle’s fuel system. There were no fires or container ruptures during the testing.

These crash tests on mockup hydrogen vehicles provide data and insight on the behavior of the hydrogen fuel system and its supporting structure during a crash. This data may provide guidance on how test “bucks” or sleds can be used to test hydrogen fuel systems without resorting to the expense of full-vehicle testing. Because they eliminate differences caused by the vehicle structure and the subtle variations inherent in crash testing, sled tests are more controlled than full-vehicle crashes. Data from full-vehicle crash tests can lead eventually to specifications for sled tests as well as to better interpretations of the energy and crush distance of impact tests.

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4.3 Test Conditions

The team crashed three small passenger sedans with mockup hydrogen fuel systems, generally following the procedures of the front, rear, and side crashes in FMVSS No. 301. The impacts were recorded by accelerometers and other instruments in the vehicle and by video cameras. Damage to the fuel systems caused by the crashes was documented. The procedures for conducting the crashes and collecting the data are summarized in this section. The complete test plan and the associated safety analysis are in Appendix E.

4.3.1 Technical Approach

The technical approach was to crash vehicles with mockup hydrogen fuel systems, with their fuel containers in conditions that have been shown to be the most vulnerable to damage through impacts.

The team purchased three CNG-fueled sedans. The natural gas systems were removed, and pressurized hydrogen fuel systems were installed, simulating a hydrogen vehicle construction. Each vehicle was fitted with one Type 4 hydrogen container, nominally 65 liters water volume.

The crashes were conducted by the Texas Transportation Institute (TTI) at its Riverside Campus near College Station, Texas. The crashes were patterned after the front, side, and rear crash test conditions of FMVSS No. 301 [8] and NHTSA’s associated test procedures [15 , 16, and 17]. Pressure and temperature in the fuel container were to be monitored for one hour following each crash. (Due to equipment malfunctions, the one-hour hold periods following the crashes were truncated. Separate one-hour nitrogen pressurization tests were subsequently run on each vehicle. Neither the post-crash hold nor the supplemental test showed evidence of a leak.) The fuel systems in the crashed vehicles were inspected and photographed following the crashes. The vehichle body deformation was also documented following the crashes.

4.3.2 Test Matrix

The crash tests consisted of one sequence of front, side, and rear impact tests, as in FMVSS No. 301:

• Vehicle towed forward into a fixed barrier, nominally at 30 mph • Moving barrier towed, nominally at 33 mph, into the side of a stationary vehicle, and • Moving barrier towed, nominally at 50 mph, into the rear of stationary vehicle.

The crashes generally followed NHTSA’s compliance test procedures, as indicated in Table 4-1. The front crash was into a rigid concrete barrier, and the other two crashes used a moving deformable barrier (MDB) [6]. The side crash was to the left side of the vehicle because that is where the fill port and the most internal plumbing were located. The point of impact was closer to the rear of the vehicle than in a standard side impact test [16, page 32] so that the moving deformable barrier was directed at the hydrogen container rather than at the passenger compartment. As shown in Figure 4-1, the face of the moving deformable barrier engaged the B-pillar and extended to about 300 mm aft of the fuel port. The rear impact crash will be offset to the left, as in the figure, again to direct the energy toward the hydrogen plumbing.

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Because these are research tests and not compliance tests, there were departures from the standard procedures. Most notably, elements of NHTSA’s test procedures that apply to components other than the fuel system were excluded; the vehicles were ballasted with lead blocks and sand bags rather than dummies. The fuel containers in these crashes were pressurized with hydrogen.

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Table 4-1. Conditions and camera locations.

Test Procedure

Graphic

Front TP-301-04

[15]

AND TWO HIGH-SPEED CAMERAS

REAL TIME CAMERA AND TWO HIGH-SPEED CAMERAS

Rear TP-301R-02

[17]

Side TP-214D-08

[16]

(Note that this sketch has been modified to show that the MDB struck

the vehicle aft of the location in the NHTSA Test Procedure.)



RT& HS

RT& HS

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MDB location for a 214 testCenter line of the fuel container

Welding rod aimed here

MDB location for this test

Figure 4-1. The impact area for the side crash was to the rear of the area for a compliance test so the panel adjacent to the fuel container would be directly struck. The upper photo was taken after the crash; the lower, from the high-speed movie a moment before impact.

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4.3.3 Test Vehicles

The vehicles that were crashed were three identical Model Year 2009 Honda Civic GX natural gas vehicles. Powertech replaced the existing CNG fuel systems and with mockup hydrogen fuel systems. Modifications to the original vehicle frame and container mounting were kept to a minimum, although some changes were required to accommodate the slightly smaller hydrogen container. The original engine remained in place but was not operable. A fuel cell and associated drive train was not installed. To the extent possible, ballast was added and non-functional weight removed to keep the vehicle’s mass distribution unchanged. Powertech installed pressure release solenoids to safely vent fuel following the crash as a safety precaution. The post-crash pressure release system is described more fully in Appendix E.

The re-fitted vehicles were pressure tested to a nominal 350 bar with hydrogen to check for leaks, and the fuel containers were de-fueled and re-filled with nitrogen to about 1 psi before shipment. Battelle installed all of the sensors in the vehicles and tested their operation before the vehicles were shipped to the test site. The data recording equipment was installed in Texas by TTI.

These vehicles were selected because the existence of an original CNG fuel system would simplify installation of a mockup hydrogen fuel system. Changes to the container mounting and the fuel system were kept to a minimum. Nevertheless, modifications were necessary, so the behavior of the mockup fuel system in these crashes cannot be construed to be representative of how these vehicles’ original fuel systems would behave in a crash.

4.3.3.1 Physical Description

The test bed vehicles were mid-size four-door sedans. They were built with five designated seating positions, two in the front and three in the rear. Following the formulas of the respective NHTSA compliance test procedures, the target test weight for the vehicles was calculated to be 1520 kg, as shown in Table 4 of Appendix E. The actual test weights of the three vehicles are in Table 4-2.

Table 4-2. Weights of the vehicles in the crash tests.

Crash Direction Target Test

Weight Actual Test Weight

Range Actual Test

Weight

Front 1,520 kg 1,511 ≤ ____ ≤ 1,520 1,514 kg

Rear 1,520 kg 1,510 ≤ ____ ≤ 1,515 1,509 kg

Side 1,520 kg 1,511 ≤ ____ ≤ 1,515.5 1,509 kg

The vehicles were merely a test bed for the experimental fuel system. These crashes were not tests of the production vehicles. Because of the extensive modifications to the vehicle and fuel system, the results of these experiments are not necessarily indicative of how the production vehicles would perform in compliance tests.

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4.3.3.2 Fuel System

The vehicles were purchased new, with the original CNG fuel system intact. The major modification was to replace the 250-bar CNG container in each vehicle with a 350-bar hydrogen container. Other components that are wetted by the fuel were replaced with components rated for hydrogen. All high-pressure tubing was 1/4 in., 0.065 in., 316 stainless steel. Table 4-3 compares the containers manufactured in the cars by Honda with those that were installed for the crash tests.

Table 4-3. Fuel container comparison.

As the Vehicles were Manufactured

As the Vehicles were Tested

Construction Type 3 Type 4

Manufacturer Structural

Composites Industries, LLC

Lincoln Composites

Model ALT 847 RG50A16-03306

Service Pressure 250 bar 350 bar

Volume (Water Capacity)

115 L 65 L

Mass (nominal, empty) 50 kg 32 kg

Length 893 mm 815 mm

Diameter (nominal) 460 mm 400 mm

Figure 4-2 is a photograph of a vehicle before modifications. The CNG fuel container was exposed for this photograph by opening the trunk and removing two cover panels. The container is mounted transversely and is located behind the rear seat back cushion. The solenoid valve is on the driver (left-hand) side of the container, and the fuel receptacle (with its door open in the photograph) is on the driver side of the vehicle.

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Figure 4-2. The rear of the original fuel container, with the trunk panel removed.

Figure 4-3. The hydrogen container installed in the trunk of a test vehicle.

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Figure 4-3 is a photograph of the hydrogen container installed in place of the CNG container. As much of the original mounting hardware was retained as possible. The straps around the container are segmented; one segment was removed from each strap to accommodate the smaller diameter of the hydrogen containers. The bolts for the hydrogen straps are longer than were the original bolts. Figure 4-4 is a photograph taken from the driver side rear door, looking backwards. The rear seat cushion has been removed. This photograph shows the only physical modification of the container-mounting frame made by Powertech. The front cross arm was moved approximately 35 mm (1-1/2 in.) toward the rear of the vehicle to account for the smaller diameter container. The previous manner of fitment was a bolt through the frame base arm and cross-arm. This new fitment was replicated using an equivalent Grade 8 through-bolt (shown in the circle). Additional photographs of the mockup hydrogen fuel system are in the Test Plan in Appendix E and the in the pre-crash photographs of Appendix F.

Figure 4-4. The front of the hydrogen container, showing the modification to the mounting frame.

A sheet metal stamping was behind the container across the entire width of the interior, and a triangular stamping was on each side of the vehicle in front of the container. The rear stampings were removed from all three vehicles, and both front stampings were removed from the rear and side crash vehicles and the driver side front stamping from the front crash vehicle, to facilitate mounting instruments. To the extent that these members could carry a load had they been in place, the experimental crashes were harsher on the fuel system than would have been a crash with the original equipment.

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4.3.4 Facilities and Equipment

The crashes were conducted by TTI personnel at their location, using their instruments and cameras, with their normal crash test procedures adapted to the hydrogen tests. Supplemental instruments specific to the hydrogen crashes were installed as well.

4.3.4.1 Test Site

The remote, outdoor location of TTI’s Riverside Campus was an ideal location for these experiments requiring a large exclusion zone for safety. Figure 4-5 is a map of the campus showing the location of the crashes and the safety perimeter.

4.3.4.2 Instruments

The vehicles were outfitted with a series of accelerometers, pressure transducers, temperature sensors, and string potentiometers. The moving deformable barrier had two triaxial accelerometers. A complete list of the transducers and mounting locations is in Appendix E.

In each test 16 Endevco 7264C accelerometers were used to record the dynamics of the vehicle and the hydrogen container during the crash. The accelerometers were mounted such that accelerations due to the yaw, pitch and roll of the container could be collected. The orientations of the accelerometers complied with the conventions in SAE J211. The accelerometers were secured to triaxial mounting blocks. The approximate location of each accelerometer block was

• Triax 1 was located at the x-y center of gravity for the vehicle. • Triax 2 was located near the rear seat cross member. • Triax 3 was mounted axially on the container end on the passenger side. • Triax 4 was mounted axially on the container end on the driver side. • Triax 5 was mounted to the container carrage frame rear mounting bolt. • Uniax 1 was mounted to the top of the container for measuring the pitch during the crash.

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Figure 4-5. The crashes were at the south end of TTI’s Riverside Campus.

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Celesco crash rated string potentiometers were used to measure the linear displacement of the container in the direction of each crash. The body of each string potentiometer was mounted to a series of movable rails which were mounted to the vehicle. Once the string potentiometers were in line with container ends, they were secured to the rails. One string from each potentiometer was then secured to each end of the container in the front and rear crashes. Figure 4-6 shows the string potentiometers and mounting rails for the front crash test vehicle. The string potentiometers are circled in yellow. In the side crash the potentiometers were mounted transversely to measure the side to side movement of the container relative to the vehicle body and relative to the container mounting frame. The string potentiometer data was sampled at 20 kHz.

Accelerations of the container and the vehicle body were measured using Endevco piezoresistive accelerometers. All acceleration data was collected using a DTS TDAS-PRO at a sampling rate of 10 kHz and was filtered post-test in accordance with CFC60. The pre-crash and post-crash measured locations of each accelerometer for each vehicle are in Appendix H. The accelerometers were mounted such that they adheared to the specifications in SAE J211. Where necessary the output polarity was switched to match the sign convention of SAE J211.

The pressure transducers and temperature sensor were used to characterize the state properties of the hydrogen gas inside the container. Since the density of gases are dependent on both temperature and pressure, measurements of both are required to determine the presence of a leak. A drop in pressure alone does not indicate a leak.

Three AST4000 hydrogen compatible pressure transducers were used to measure the hydrogen pressure inside the container and low pressure line. Two of the three pressure transducers were plumbed to measure the container pressure. The third pressure transducer was plumbed into the the mock fuel line to monitor the low pressure beyond the single stage regulator. The pressure transducers and a hydrogen leak sensor were sampled at 200 Hz.

Figure 4-6. The mounting configuration for the string potentiometers in the front crash test vehicle.

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4.3.4.3 Cameras

Two real time and seven high-speed motion picture cameras captured each crash. One high-speed camera was in the crash vehicle between the two front seats looking backward. The fuel container was clearly visible from the onboard high-speed camera because the cushion for the rear seat had been removed. Several stationary ground-level cameras were mounted to the side of the impact point. An overhead high-speed camera captured the impact from a height of roughly 15 m (50 ft) .

4.4 Results

The crashes were carried out according to the plan with only the smallest of deviations. The planned and actual speeds of the crashes are listed in Table 4-4.

This discussion focuses on the effect of the crashes on the fuel systems. Descriptions of the crashes, as they pertain to the whole vehicle, are in TTI’s test reports in Appendix F. A complete set of time histories is in Appendix G. The responses of the fuel systems to the crashes were assessed by qualitative description, dynamic measurement during the crash, by pressure measurement, and by post-crash inspection.

Table 4-4. Planned and actual conditions at impact.

Crash Direction Speed Angle

Planned Actual Planned Actual

Front 48 kph

(30 mph) 45.8 kph

(28.6 mph) 90° 89.3°

Rear 80 kph

(50 mph) 83.4 kph

(52.2 mph) 90° 89.9°

Side 53 kph

(33 mph) 51.0 kph

(31.9 mph) 27° 30.0°

4.4.1 Qualitative Description of Damage to the Fuel Systems

Not surprisingly, the front crash did the least damage to the rear-mounted fuel system. The fuel container rocked in its mounting, but it returned essentially to its original location. Of the three crashes, the side crash was intermediate in its effect on the fuel system. The rear crash caused extensive damage to the container mount and superficial damage to the container wrapping.

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4.4.1.1 Front Crash

The fuel container in the front crash rocked in its mounts but returned to its original position. The fuel container in the rear crash was moved within its mount, and the mount was moved significantly within the vehicle. The exhaust pipe hanger failed. The distance between the engine and the rear bumper was reduced during the crash, and the exhaust pipe shortened by buckling near a bend in the horizontal plane. The fuel line shortened by bending at the point where it bends to turn up into the engine compartment.

4.4.1.2 Rear Crash

The container in the rear crash moved toward the right in its mount, with the two straps slanting as a parallelogram. The container permanently rolled forward in its mount. The container entered the passenger compartment by coming forward through the position of the rear seat cushion. The cushion was not in place but is not structural.

Had the triangular gussets been in place and remained in place, the fuel system would not have come as far forward. The gussets would have prevented the two horizontal members from opening to permit a passage for the fuel container. Had a bolt sheared or (more likely) a gusset torn from the bolt, some energy would have been absorbed, but the opening would again be permitted.

Immediately at impact, the crumple zone in the test vehicle's trunk took damage, with the body of the test vehicle only slowly accelerating until the MDB had passed through the trunk. The MDB rode up on the left rear tire of the test vehicle, lifting the MDB’s front tires off the pavement. The brakes on the MDB were never applied, so there was no force to pull the MDB away from the test vehicle. The test vehicle's right rear tire went flat, so it provided a braking force to slow the combined vehicles and yawing them to the left.

The MDB struck the container itself. Figure 4-7 shows the trunk lid against the container. On either side of the license plate location are the two bolts on the straps of the container mount, which punctured through the trunk lid. This direct load on the container damaged the mounting brackets. On the left side the bolt, in the circle in Figure 4-8, pulled out of the vehicle body. The fore-aft mounting member is bent at nearly a right angle. The failure was different on the right side. The finger points to the hole in the fore-aft member from which a bolt was pulled. The fore-aft member stayed attached to the vehicle body on the right side.

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Figure 4-7. The moving deformable barrier in the rear crash pushed the trunk lid to the fuel container. (The tail pipe and right edge of the bumper are less deformed because the impact was offset to the left.)

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Figure 4-8. In this view through the driver-side rear door, the yellow box shows where the anchor for the container mounting failed during the rear crash.

Figure 4-9 shows the effect of the crash on the fuel system itself. The impact pushed the container past one of the structural bolts. The bolt scratched and broke fiber wrappings on the surface of the container and penetrated into the protective foam underneath. The area of this damage is in the dotted circle in the photograph. Note the gap between the strap and the container at the location where the strap is anchored. The solid green arrow points to a gap on the near side; a similar gap is visible by the far anchor in the photograph. Before the crash, the strap conformed to the circular form of the container. The strap has been straightened in plastic deformation caused by the container pushing it forward. The container itself would have been momentarily deformed while doing so. The permanent translation of the container with respect to the vehicle caused the tubing between the vehicle body and the container to be bent and stretched. The outlined yellow arrow in the photograph indicates one such location.

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Figure 4-9. The impact of the rear crash pushed the fuel container into the rear seat area. The green solid arrow indicates where the strap has been permanently deformed by the container’s pushing it against the mount. The yellow hollow arrow points to distortion in the vent line. When the fuel container came forward, a structural bolt marred its surface in the area inside the dotted circle. This photo looks through the passenger-side rear door.

4.4.1.3 Side Crash

Penetration caused by the side crash in compliance tests is typically less than the distance to where the hydrogen fuel container was mounted, and such was the case in this crash. Even though the moving deformable barrier was aimed aft of its location in a compliance test, damage to the fuel system was minimal. The fuel inlet line was the only part with permanent deformation, and it was functional following the crash. The profile of the side-crash vehicle was measured at longitudinal intervals of 300 mm. The forwardmost measurement was at the front axle, as shown in Figure 4-10. The reference plane for the measurements was 914 mm (36 in.) from the vehicle centerline. Measurements were made at four levels; Table 4-5 lists their heights above the ground. Level 1 is the sill, the same as Level 1 in the test procedure [16, page 32]. Level 4 is the roofline, the same as the Level 5 in the test procedure except that the roof of the test vehicle is more curved than that in the sketch.

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Table 4-5. Distances above the ground where the crush profile was measured.

Pre-Test Post Test

Level 1 279 254

Level 2 635 635

Level 3 940 914

Level 4 1,397 1,372

Units are mm.

The measurements are in Table 4-6 and Figure 4-11. The maximum deformation, which was in the door panel, was 162 mm. The maximum at the 2700-mm location, near the container, was 145 mm, and the deformation close to the fuel port was 84 mm. As often occurs in these tests, the impacted door becomes concave and the upper part of the door moves outward, as indicated by the negative numbers in the final column of the table and is evident in the photo in Figure 4-12.

Table 4-6. Exterior crush profile of the side impact vehicle.

Pre-Test Post-Test Difference

1 2 3 4 1 2 3 4 1 2 3 4

0 75 48 90 58 15 10

300 80 43 142 98 50 137 18 7 -5

600 67 47 37 87 52 58 20 5 21

900 61 46 119 86 55 125 25 9 6

1200 56 42 110 304 88 57 126 295 32 15 16 -9

1500 51 40 100 313 99 95 132 299 48 55 32 -14

1800 51 39 99 324 106 201 204 307 55 162 105 -17

2100 42 38 90 340 117 196 186 324 75 158 96 -16

2400 46 35 95 335 94 185 185 331 48 150 90 -4

2700 45 35 100 101 180 184 56 145 84

3000 45 43 124 122 165 174 77 122 50

3300 85 87 162 141 123 169 56 36 7

Units are mm.

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Figure 4-10. The vertical lines drawn on the vehicle following the side crash indicate the longitudinal positions where the deformations were measured.

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Figure 4-11. The maximum deformation in the side crash was in the panel of the rear door.

Figure 4-12. This photograph of the vehicle following the side crash shows the penetratation caused by the impact. The outward displacement of the rear door is visible.

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4.4.2 Transient Motion During the Crashes

Transient motion of the vehicle and container during the impact will be discussed in order for the front, rear, and side crashes. Highlights of the data sufficient for the discussion are presented here; plots of all string potentiometer and accelerometer data are in Appendix G. Figure 4-13, the plot of the string potentiometer reading during the front crash, exemplifies how the plots are formatted. Time zero is the moment of impact—the time when a tape switch on the front bumper first made contact with the rigid barrier. The legend in the corner of each plot identifies each trace by color and style. The first word in the legend entry is front, rear, or side, indicating the crash direction. The second entry is the string potentiometer name or accelerometer channel number (See Appendix H). The final entry is a brief description of the sensor’s location.

4.4.2.1 Front Crash

Displacements of the left and rights end of the container were nearly identical in the front crash. Figure 4-13 shows measurement by the string potentiometer on the driver side. Time zero on the figure is where the tape switch on the front bumper of the vehicle contacted the concrete barrier. The vehicle was released from the towing cable approximately 200 ms prior to impact. The mechanism for the release is a slip fitting striking a fixed trigger, and this is presumably the source of the small oscillation before the main impact. After time zero, the container pitched forward, the maximum displacement of the container being roughly 25 mm in front of its original location. The container then rebounded and oscillated. These oscillations gradually damped out, and the container came to rest 7 mm in front of its original location.

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Figure 4-13. Displacement of the end of the container on the driver’s side during the front crash.

The vehicle body attained a peak acceleration of 28 g during impact. Figure 4-14 shows the acceleration pulse of the vehicle. The negative acceleration indicates that the vehicle slowed at the impact. The acceleration pulses of the container ends and the container top are superimposed over the vehicle body acceleration in Figure 4-15. Consistent with with the displacement data from the string potentiometer, the container rocks back and forth on its mounting. Note that the acceleration at the container top (the green line in the figure) has a higher amplitude than the acceleration along the container axis (the red line).

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Figure 4-14. Vehicle body acceleration during the front crash.

Figure 4-15. Container longitudinal accelerations superimposed over the vehicle body acceleration during the front crash.

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While the accelerations in the direction of the impact are the most severe, the vertical component to the accelerations is also significant. The container experiences a vertical maximum acceleration, which exceeds the vertical acceleration of the vehicle during the impact. Figure 4-16 shows the vertical accelerations of each end of the container superimposed over the vertical acceleration of the vehicle body. This indicates that the peak vertical acceleration experienced by the container is roughly double that of the peak vertical acceleration experienced by the vehicle.

Figure 4-16. Container vertical accelerations superimposed over the vehicle body vertical acceleration during the front crash.

4.4.2.2 Rear Crash

In the rear crash, the container began to pitch backwards as expected. However, due to the MDB riding up onto the rear wheel, the container was thrust forward into the passenger compartment. Figure 4-17 shows the rear crash string potentiometer output for the passenger side of the container. Because the test vehicle was stationary prior to impact, there is no pre-crash oscillations as in the front crash. Immediately after impact, the container is displaced backwards slightly, then was pushed forward almost 250 mm. The container rebounded slightly. The lack of oscillations is likey due to the container coming in contact with surrounding objects resulting from the large forward displacement. The mounts for the string potentiometers were deformed in the impact, so only in the initial movement is the quantity meaningful. Figure 4-9 shows a close up of the passenger side of the container contacting a black cross-member, which damaged the outer wrappings of the container.

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Figure 4-17. Displacement of the passenger side end of the container during the rear crash.

The rear crash peak acceleration of the vehicle body in the crash direction was 23 g as shown in Figure 4-18. The positive acceleration means that the vehicle was accelerated forward due to the collision. The vehicle’s motion continued well beyond 1 s after the crash. Thus, the bumps observed after 0.2 s are a result of the vehicle’s continued forward motion.

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Figure 4-18. Vehicle body acceleration due to the rear impact from the MDB.

The MDB rode up on the rear driver side tire and directly impacted the container (Figure 4-7). As a result, the container experienced an acceleration peak around 170 g and a subsequent rebound peak around -70 g. Figure 4-19 shows the container accelerations superimposed over the vehicle body acceleration. This sudden advancement and slight retreat are evident in the in-car movie. The acceleration of the vehicle center of gravity (the blue lines in Figure 4-18 and Figure 4-19) has a brief peak at the moment of impact and then rises gradually as MDB works through the rear crumple zone. The very high peak acceleration of the container (the red and green lines in Figure 4-19) are consistent with a direct impact. At this moment, the crumple zone is exhausted, and the vehicle maximum acceleration is when the elastic energy stored in the container and its mounts transfers to the vehicle body during the rebound.

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Figure 4-19. Longitudinal accelerations of the container superimposed over the vehicle body acceleration during the rear crash.

4.4.2.3 Side Crash

The string potentiometers for the side crash were mounted tranversly so that the lateral motion of the container could be measured. One string potentiometer measured the container displacement relative to the container mounting frame, and the second string potentiometer measure the mounting frame relative to the vehicle body. Figure 4-20 shows the combined data from both string potentiomaters, resulting in the container motion relative to the vehicle body. Motion towards the passenger side of the vehicle is positive.

Immediately after impact, the container moves a maximum of 18 mm towards the driver side. There is a slight stutter on the rebound due to the compound motion of the container and the mounting frame moving out of unision. The oscillations after the rebound are smaller and damp out more quickly than in the front crash. The net final displacement of approximately 4 mm towards the driver side is completely due to the container shifting its location relative to its mounting frame.

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Figure 4-20. Passenger side container displacement relative to the vehicle body during the side crash.

The vehicle motion in the side crash was not parallel on one of its axes, so the side crash accelerations are plotted here as resultants in the X-Y (horizontal) plane. The individual X and Y axis data are in Appendix G. Figure 4-21 shows the X-Y resultant acceleration of the rear seat. The peak vehicle acceleration was slightly less than 30 g. The container experienced slightly higher peak acceleration than the vehicle during the side crash test. From Figure 4-22 the peak resultant acceleration of the container end closest to the impact region was approximately 10 g higher than the acceleration of the vehicle. The end of the container of the more protected side of the vehicle experienced a peak resultant acceleration that was roughly 5 g higher than the vehicle. The container also experienced some rocking after the rebound. This agrees with the graph of the string potentiometer data in Figure 4-20.

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Figure 4-21. Vehicle rear seat X-Y resultant acceleration during the side crash.

Figure 4-22. Container X-Y resultant accelerations superimposed over vehicle rear seat resultant acceleration during the side crash.

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4.4.3 Pressure Integrity

Pressure and temperature in the fuel container were to be monitored for one hour following each crash. Equipment malfunctions limited the data collection following the front and rear impacts to 327 s (just under five and a half minutes). A separate equipment malfunction in the side impact released the pressurized hydrogen prematurely, so pressure monitoring was not possible. Therefore supplemental one-hour nitrogen pressurization tests were run on each vehicle. Neither the brief hold periods immediately following the two crashes nor the subsequent pressurizations showed any evidence of a leak.

The leak detection analysis was performed by comparing the densities of the hydrogen before and after the crash. These densities were calculated from the recorded hydrogen pressure and temperature by the NIST thermophysical property calculator [19]. This calculator is based on the equations of state for hydrogen [11] and nitrogen [25].

In the event of a leak, the mass of the hydrogen post-crash would be less than the mass of hydrogen pre-crash. Assuming the volume of the container is the approximately the same throughout the test, the densities of the hydrogen before and after the crash can be compared. If the ratio of the post crash density to the pre crash density is equal to unity, then no leak occurred. The ratio of densities was compared to the propagated uncertianty from the pressure and temperature measurements. The uncertainty in the hydrogen density ratios are around 3.5 percent. A propagated uncertainty of less than 0.7 percent would be required to detect a leak. In the analysis, the errors considered for the pressure transducers included non-linearity, non-repeatability, hysteresis, and temperature compensation span error. The errors considered in the temperature sensors were, span error, non-linearity, and errors due to the conditioning of the RTD signal.

The results from the hydrogen crash tests are in Table 4-7. The nominal density ratios are indistinguishable from unity. The deviations from unity are within the bound of the uncertainty. No discernable hydrogen leak occurred following the front or rear crash tests.

Table 4-7. Density ratios from the hydrogen crash tests.

Crash Direction

Density Ratio Result

Nominal Maximum

Uncertainty

Front 0.993 0.035 No Leak

Rear 0.995 0.035 No Leak

Side - - *

* Malfunction of the post crash vent system released the hydrogen prematurely.

For the one-hour nitrogen pressurization tests a similar analysis was performed. The density ratios were compared over a time span of one hour. Fuel systems were individually pressurized to a nominal 35 bar with nitrogen, and the pressure and temperature in each container were

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monitored for one hour. The instrumentation used in the hydrogen tests were re-used for the nitrogen tests with one exception. The side crash vehicle’s 700-bar pressure transducers were replaced with 70-bar transducers from the front crash test vehicle. The 700-bar transducers were not sufficiently sensitive for the 35-bar supplemental pressurization.

Table 4-8. Density ratios from the nitrogen pressurization tests.

Crash Direction

Density Ratio Result

Nominal Maximum

Uncertainty

Front 0.985 0.038 No Leak

Rear 1.001 0.038 No Leak

Side 1.005 0.038 No Leak

Because the density ratios in Table 4-7 and Table 4-8 are close to unity and their difference is much less than the maximum uncertainty, the conclusion can be drawn that no detectable leak occurred.

4.4.4 Damage to the Containers

Containers were removed from the vehicles following the crashes so their entire surface could be inspected. The containers in the front and side crashes had no visible signs of damage. The container from the rear crash had abrasions in several places; most were superficial. The surface was penetrated near the end, as shown in Figure 4-9, and many fibers were broken. This gouge was into the fiberglass outer layer, which is indicated in the cross section in Figure 2-2. This outer layer is to protect the dome of the pressure vessel in an impact; it is not itself structural. The gouge at its deepest point was in excess of 3 mm. The manufacturer of these containers, Lincoln Composites, Inc., publishes guidelines that abrasion damage limited to the first 0.25 to 0.90 mm (0.010 to 0.035 in.) does not diminish the integrity of the container and can be re-worked in the field [12]. Gouges deeper than 1.27 mm (0.05 in.) require that the container be condemned.

4.5 Conclusion

At the highest level the results of the crashes were as expected—there was no leak or fire. The most significant finding of the crashes was the amount of damage that the fuel system tolerated in the rear crash. The mockup hydrogen fuel system was moved from its original position, its mounts were ripped from their anchors, and the stainless tubing was stretched and twisted, but the system held pressure. With its provision for redundant post-crash hydrogen venting, the mockup fuel system had more and longer lengths of exposed tubing than would a production hydrogen fuel system, so it was more susceptible to damage, but the tubing and fittings held. This outcome indicates that properly specified and installed plumbing can hold pressure during at least some instances of severe crash damage.

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A sled test could be designed to further explore the susceptibility to damage of the container itself in various fueling conditions. The container could be mounted in hardware of various designs being considered and subjected to longitudinal and lateral accelerations that are typical of the crash pulses in the various FMVSS No. 301 compliance tests. A sled test would be inadequate to test the entire fuel system. The plumbing external to the container is subject to significant deformations, depending on its location within the vehicle and its mounting points. Certainly component-level tests, such as bending a pipe or tensioning a fitting, could be devised, but a crash test of an entire vehicle would be more consistent with NHTSA’s usual approach of testing the system in place.

The crashes demonstrated some of the kinds of damage to which a fuel system might be susceptible. The high-pressure fuel lines external to the fuel container were pulled and bent during the crashes. Future test vehicles and production vehicles must be built with high-quality plumbing that can maintain its integrity in such large deformation. The ability of the PRD line to direct fuel to the exterior of the vehicle must be maintained. While the container is quite robust in the case of blunt impact, prior research [18] has shown that sharp objects can penetrate the pressure boundary. The surface damage to the container caused by the structural bolt in the rear crash (Figure 4-9) is a reminder that puncture hazards must be minimized by measures such as not using self-tapping screws in the vicinity of the fuel system. As with all compliance tests, these crashes are of an idealized condition. Cargo in the trunk, such as a set of golf clubs or a large tool box, could diminish the effect of the crumple zone and pose a danger to the fuel container.

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5.0 OVERALL CONCLUSIONS

The three crash tests in the current FMVSS No. 301 were shown to pose a serious challenge to a mockup hydrogen fuel system in a small sedan. Crashes were performed based on the front, rear, and side compliance test crashes, with the particular conditions selected to be the potentially most severe to the hydrogen fuel containers. Despite the serious challenge, the fuel systems maintained their integrity with no leak or fire. No unusual vulnerabilities or anomalies were observed.

The dynamic impact tests, in which a weight was dropped on a container, confirmed that extreme impacting a container on its end with an essentially rigid object can cause catastrophic failure. The aluminum honeycomb in the existing FMVSS Nos. 214 and 301 test is unlikely to cause this failure. The essentially rigid pole of the other FMVSS No. 214 crash test would be a greater threat to the end of a container.

The formulas in Appendix A of SAE J2578 were shown to agree reasonably well with the leak data. They could benefit from adjustments to account effects that have been neglected. The most important such effect is the increase in hydrogen’s and helium’s density between atmospheric and the high test pressures. The more significant finding is that measurements with readily available pressure and temperature transducers are not sufficiently accurate to detect the small leaks that have been shown by another task order to be potentially dangerous. A new approach to detecting leaks may be warranted.

5.1 Crash Test Conditions

Following is what the project team learned concerning the conduct of crash tests for hydrogen fueled vehicles.

The container and the rest of the fuel system were protected by the vehicle as a system. The straps allowed the container to move within its mounting, cushioning the blow of the initial impact. The crumple zones of the vehicle gradually dissipated the energy of the crash and protected the container as it was mounted in the interior of the vehicle in this case.

The expectation before the project was that the containers would be quite robust. That was largely confirmed by both the crush tests and the crash tests. Surface damage to the container in a crash is possible. Long, sharp, vehicle components (or cargo) in the vicinity of the container threaten to impale or abrade it, as exemplified in the rear crash. Rupture, as occurred in longitudinal impacts in the crush tests, would occur only in the most extreme crash conditions where both ends of the container are constrained. The rear crash caused significant displacement of the container within the vehicle and deformation of the tubing. The tubing, fittings, and threads survived the crash because they are ductile and of high quality.

The crush tests showed that the most vulnerable portion of a container is its valve. The results, summarized in Table 2-7, were that the only leaks or ruptures occurred when the container was struck on its end. The container in this particular experiment was sufficiently small that its end was not contacted during the side crash. The procedure for the side crash was modified so that

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the moving deformable barrier was directed more at the fuel container than the occupant compartment, but the fuel system was not severely damaged. Because the deformable aluminum honeycomb is unlikely to cause a rupture, it would be difficult to make an argument that vehicles should be subjected to two such impacts, one at the current location to test occupant protection and again at a location farther back to test the fuel system. A modified FMVSS No. 214 pole test would pose a greater threat to the valve. Striking the container with a rigid object during a crash test would be more severe than striking it with an aluminum honeycomb but less severe than constraining it between two rigid surfaces.

5.2 Fuel Options

The reason for selecting the low container pressurization for the front and rear crashes was borne out by the crash tests. The front and rear crashes were expected to load the container through straps similarly to the horizontal impacts in the crush tests, which had shown the greater wall deformations occur at lower internal pressures. The permanent deformation of the straps following the rear crash (Figure 4-9) clearly shows that the container’s side was deformed during the crash.

All of the container failures in the dynamic impact tests occurred in the vertical orientation, where the weight struck the container on its end. The more highly pressurized containers failed by rupture, which is a more severe outcome than the leakage of the low-pressure containers. The containers in these experiments were between two pieces of concrete with steel plates on their surfaces. Only if a large fraction of the crash energy were directed to a rigidly constrained container would this failure be expected. The pressure for the side crash was selected to be full service pressure because fully pressurized containers the crush tests tended to rupture while the nearly empty containers only leaked (Table 2-7). The valve on the end of the container was not engaged during the side crash in this study, so the reason for the high pressure was not met. If a different crash test article strikes the valve, a full pressure would then be the most vulnerable condition.

If a fuel system were equally vulnerable at low and high pressure, safety would be a reason for testing with low pressure. On the other hand, a leak of a given physical size is easier to detect through a pressure drop if the initial pressure is higher, but the effect is not strong.

5.3 Crash Procedure

A necessary part of this project was developing procedures to safely conduct crash tests of vehicles fueled with hydrogen. There were precautions for fueling the container, for protecting personnel during the crash, and for releasing the fuel following the crash. The mockup fuel system built specially for these experiments did not have some of the safeguards that would presumably be present in a production vehicle. Many of the precautions for the experiments would be prudent during continued development of test procedures. Ultimately, a compliance test would demonstrate the effectiveness of safeguards that mitigate fire and high pressure hazards for rescue workers as well as for occupants.

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Though it was not a focus of this project, the safety procedures developed for the experimental crashes can serve as guidelines for future research. A radio-controlled valve to vent the high pressure hydrogen from a safe standoff distance was essential to the project. As confidence is gained in the ability of a damaged container to hold pressure, at least for a short time, the need for remotely venting the fuel may be diminished.

5.4 Further Work

The crush and crash tests in this program examined two extremes of protection for the container. The dynamic impact tests examined the strength of the container and its valve, without any hardware that might mount it in a vehicle. The full-vehicle crash tests, on the other hand, had an entire mockup fuel system in a representative mounting in a vehicle. Continued research can focus on subsystems. A next step would be to continue the dynamic impact tests with the containers in a vehicle-type mounting. In this way, well controlled impacts can be directed to a more realistic assembly. Similarly, a complete fuel system, including tubing and a PRD and regulator, can be challenged in sled tests that mimic crash tests but are less expensive.

Further work with respect to the dynamic impact tests would include examining the containers that appeared to survive the horizontal impacts to determine the extent of hidden damage. All of the containers struck on their side surfaces held pressure. Conventional wisdom is that the less pressurized containers suffered more damage because of their greater deformation. Further testing of the containers, either metallographic examination or pressure testing, is necessary to confirm the conclusion.

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2. American National Standards Institute. (2007, July). NGV2-2007: American National Standard for Natural Gas Vehicle Containers. New York: Author.

3. Flamberg, S., Rose, S., & Stephens, D. (2010), February). Analysis of published hydrogen vehicle and safety research. (Report No DOT HS 811 267). Washington, DC: National Highway Traffic Safety Administration. Available at www.nhtsa.gov/DOT/NHTSA/NVS/Crashworthiness/Alternative%20Energy%20Vehicle%20Systems%20Safety%20Research/811267.pdf

4. National Highway Traffic Safety Administration. (2012, March). Hydrogen fuel cell vehicle fuel system integrity research: Electrical isolation test procedure development and verification. (Report No. DOT HS 811 553. Washington, DC: Author. Available at www.nhtsa.gov/DOT/NHTSA/NVS/Crashworthiness/Alternative%20Energy%20Vehicle%20Systems%20Safety%20Research/811553.pdf

5. Reuther, J. J., John, J. S., Shawcross, P. E., & Kimmel, G. (2013, October). Post-crash hydrogen leakage limits and fire safety research. (Report No. DOT HS 811 816). Washington, DC: National Highway Traffic Safety Administration. Available at www.nhtsa.gov/DOT/NHTSA/NVS/Crashworthiness/Alternative%20Energy%20Vehicle%20Systems%20Safety%20Research/811816.pdf

6. 49 CFR 587 - Subpart B - Side Impact Moving Deformable Barrier. Available at http://frwebgate.access.gpo.gov/cgi-bin/get-cfr.cgi?TITLE=49&PART=587&SUBPART=b&TYPE=PDF

7. Endress + Hauser. (2009, October) Technical information, Proline Promass 80A, 83A. Coriolis mass flow measuring system. Reinach, Switzerland: Authors. Available online at https://portal.endress.com/wa001/dla/50000000158/000/06/TI054DEN_1009.pdf

8. 49 CFR Part 571, Docket No. NHTSA-00-8248, RIN 2127 AF36 Federal Motor Vehicle Safety Standard No. 301; Fuel system integrity. Available online at www.nhtsa.gov/cars/rules/rulings/301nprm/index.html

9. 49 CFR 571.303, Federal Motor Vehicle Safety Standard No. 303; Fuel system integrity of compressed natural gas vehicles. Available online at www.gpo.gov/fdsys/granule/CFR-2011-title49-vol6/CFR-2011-title49-vol6-sec571-30Kayser, K., & Shambaugh, R. (1991). Discharge coefficients for compressible flow through small-diameter orifices and convergent nozzles. Chemical Enγineerinγ Science, 46; (7); pp. 1697-1711.

10. Leachman, J. W., \Jacobsen, R. T., Penoncello, S. G., & Lemmon, E. W. (2009). Fundamental equations of state for parahydrogen, normal hydrogen, and orthohydrogen.

92

Journal of Physical and Chemical Reference Data, 38(3); pp. 721-748. Available at http://link.aip.org/link/JPCRBU/v38/i3/p721/s1

11. Japan Automobile Standards Internationalization Center. Automobile Type Approval Handbook for Japanese Certification [a.k.a “ the Blue Book” ]. (Annual electronic publication). Tokyo: Author. Available at www.jasic.org/e/08_publication/bb/20_handbook.htm

The following JASIC standards and test procedures are relevant: Technical Standards for Construction and Device of Motor Vehicles

11-1-4-17 Attachment 17, Technical standard for fuel leakage in collisions, etc. 11-1-4-100 Attachment 100, Technical standard for fuel systems of motor vehicles

fueled by compressed hydrogen gas Test Procedures

TRIAS 33-2005 Test procedure for fuel leakage in collisions, etc. TRIAS 66-2005 Test procedure for fuel systems of motor vehicles fueled by

compressed hydrogen gas

12. Lincoln Composites, Inc. (2008, January). Fuel container inspection guidelines: Fuel containers & fuel storage systems. (Product No. 15653 Rev D). Lincoln, NE: Author. Available at www.ngvi.com/Documents/Lincoln_Guidelines_000.pdf

13. Mitsuishi, H., Oshino, K., & Watanabe, S. (2005). Dynamic crush test on hydrogen pressurized cylinder. [In Japanese:] Japanese Automotive Research Institute Research Journal, 27(7). Also in (in English): International Conference on Hydrogen Safety, Sept 8-10, 2005; Pisa, Italy.

14. NHTSA. (2007, January 17). Fuel system inteγrity - frontal fixed barrier impacts, standard reγulation on fuel system inteγrity - frontal fixed barrier impacts. (NHTSA Laboratory Test Procedure TP-301-04). Washington, DC: Author. Available at www.nhtsa.gov/DOT/NHTSA/Vehicle%20Safety/Test%20Procedures/Associated%20Files/TP-301-04.pdf

15. NHTSA. (2006, December 15). Side impact protection – Dynamic, standard reγulation on dynamic side impact protection. (NHTSA Laboratory Test Procedure TP-214D-08). Washington, DC: Author. Available at www.nhtsa.gov/DOT/NHTSA/Vehicle%20Safety/Test%20Procedures/Associated%20Files/TP214D-08Part1.pdf (Appendices A and B refer to the Side Impact Dummy. Appendix C, linked below, defines the aluminum honeycomb face on the moving barrier: www.nhtsa.gov/DOT/NHTSA/Vehicle%20Safety/Test%20Procedures/Associated%20Files/TP214D-08APP_C.pdf )

16. NHTSA. (2007, January 17). Fuel system inteγrity – Rear movinγ barrier impacts, standard reγulation on fuel system inteγrity -- rear movinγ barrier impacts. (NHTSA Laboratory Test Procedure TP-301R-02). Washington, DC: Author. Available at

93

www.nhtsa.gov/DOT/NHTSA/Vehicle%20Safety/Test%20Procedures/Associated%20Files/TP-301R-02.pdf

17. NHTSA. (n.d.) Office of Vehicle Safety Compliance, Compliance Database. (Online database). Washington, DC: Author. Available at www.nhtsa.gov/cars/problems/comply/

18. National Institute of Standards and Technology. (2011). Thermophysical properites for hydrogen. (Webpage). Available at http://webbook.nist.gov/chemistry/fluid/

19. Odegard Jr., B. C., & Thomas, G. J. (2001). Testing of high pressure hydrogen composite tanks. (Report No. NREL/ CP-570-30535). In: Proceedinγs of the 2001 DOE Hydroγen Proγram Review, Baltimore, MD, April 17-19, 2001, pp. 666-678. Golden, CO: National Renewable Energy Laboratory. Available at http://www.nrel.gov/docs/fy01osti/30535.pdf

20. Pape, D., Stephens, D., & Horacek, P. (November 19, 2008) Work plan for compressed hydrogen container fuel options for crash testing. (Unpublished document from Contract No. DTNH22-08-D-00080 Task Order 1). Washington, DC: National Highway Traffic Safety Administration.

21. Rose, S., Flamberg, S., & Stephens, D. (2009, December 24). Engineering assessment of current and future vehicle technologies, FMVSS No. 303; Fuel system integrity of compressed natural gas vehicles, FMVSS No. 304; Compressed natural gas fuel container integrity. (Final Report on Contract No. DTNH22-08-D-00085, Task Order No. 001). Washington, DC: National Highway Traffic Safety Administration. Available at

22. SAE J211, Revised July 2007. Instrumentation for impact test - part 1 - electronic instrumentation. Available at http://standards.sae.org/j211/1_200707.

23. SAE J2578 Revised January 2009. Recommended practice for general fuel cell vehicle safety. Appendix A: Post-crash criteria for compressed hydrogen systems. Available at www.sae.org/technical/standards/J2578_200901.

24. Span, R., Lemmon, E. W., Jacobsen, R. T., Wagner, W., & Yokozeki, A. (2000). A Reference equation of state for the thermodynamic properties of nitrogen for temperatures from 63.151 to 1000 K and pressures to 2200 MPa. Journal of Physical and Chemical Reference Data, 29, pp. 1361. Available at http://link.aip.org/link/JPCRBU/v29/i6/p1361/s1.

A-1

APPENDIX A:

PHOTOGRAPHS OF THE CRUSH TEST

File available on Alternative Energy link www.nhtsa.gov/Research/Crashworthiness

B-1

APPENDIX B:

LOAD TIME HISTORIES FROM THE CRUSH TEST


C-1

APPENDIX C:

LEAK RATE TEST PLAN


D-1

APPENDIX D:

LEAK RATE TIME HISTORIES


E-2

APPENDIX E:

FULL-VEHCILE CRASH TEST PLAN

AND SAFETY PLAN


F-1

APPENDIX F:

TTI CRASH TEST REPORTS


G-1

APPENDIX G:

TIME HISTORIES OF THE CRASH TESTS


H-1

APPENDIX H:

CRASH SENSOR SPECIFICATIONS


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