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Experimental and numerical investigation of localized fire testfor high-pressure hydrogen storage tanks

Jinyang Zheng a, Kesheng Ou a, Zhengli Hua a, Yongzhi Zhao a, Ping Xu b,*, Jun Hu c,Bing Han c

a Institute of Process Equipment, Zhejiang University, Hangzhou 310027, Chinab Institute of Applied Mechanics, Zhejiang University, Hangzhou 310027, ChinacDalian Boiler & Pressure Vessel Supervision and Inspection Institute, Dalian 116013, China

a r t i c l e i n f o

Article history:

Received 14 November 2012

Received in revised form

11 February 2013

Accepted 12 February 2013

Available online xxx

Keywords:

Hydrogen storage tank

Localized fire test

Pressure relief device

Global Technical Regulation

* Corresponding author. Tel./fax: þ86 571879E-mail addresses: pec101@zju.edu.cn, pin

Please cite this article in press as: Zhenpressure hydrogen storage tanks,j.ijhydene.2013.02.052

0360-3199/$ e see front matter Copyright ªhttp://dx.doi.org/10.1016/j.ijhydene.2013.02.0

a b s t r a c t

Vehicle fires may cause localized fires on on-board high-pressure hydrogen storage tanks.

To verify the safety performance of such tanks under localized fire exposure, a localized

fire test was proposed in the Global Technical Regulation for Hydrogen Fuel Cell Vehicles.

However, practicality and validity of the proposed test still require further verification. In

this paper, this new fire test was experimentally investigated using the type 3 tanks.

Influences of hydrogen and air as the filling media were studied. A three-dimensional

computational fluid dynamics model was developed to analyze the effects of filling pres-

sure and localized fire exposure time on the activation of thermally-activated pressure

relief device (TPRD). The experimental results showed that temperature distribution on the

tank surface was uneven around the circumference. The rising temperature of internal

hydrogen or air contributed little to TPRD activation. The simulation results indicated that

TPRD activation time was slightly affected by the variations of the filling pressures, but it

increased when the localized fire exposure time was extended.

Copyright ª 2013, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights

reserved.

1. Introduction showed that about 40% of the vehicle fires caused on-board

Hydrogen fuel cell vehicles (HFCVs) are being demonstrated

worldwide and gradually brought to the market [1e3]. The

safety performance of hydrogen storage tanks in HFCVs at the

accidental fire scenario has drawn extensive attention due to

the flammability and explosibility of high-pressure hydrogen

[4]. To prevent explosion in the event of vehicle fires, a specific

thermally-activated pressure relief device (TPRD) is required

to be fitted onto the end boss of the tank in existing standards

[5e8].

Vehicle fire tests, which were conducted by the Japan Auto-

mobile Research Institute and US automobile manufacturers,

53393.gxu@zju.edu.cn (P. Xu).

g J, et al., ExperimentaInternational Journal

2013, Hydrogen Energy P52

storage tanks to experience localized fire [9]. Moreover, in-

service failure of compressed natural gas (CNG) tanks during

the past decade also indicated that the majority of fire-related

vehicle accidents were caused by localized fire [10,11]. Before

turning into an engulfing fire, localized fire can locally degrade

the tank wall and even cause the tank to burst. In such a

circumstance, TPRD cannot be activated in time to release in-

ternal high-pressure gas because of inadequate heat exposure.

However, localized fire exposure has not been addressed in

existing standards [5e8].

Given the potential severity of such an event, a localized

fire test has been proposed in the Global Technical Regulation

l and numerical investigation of localized fire test for high-of Hydrogen Energy (2013), http://dx.doi.org/10.1016/

ublications, LLC. Published by Elsevier Ltd. All rights reserved.

i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n en e r g y x x x ( 2 0 1 3 ) 1e82

(GTR) for HFCVs [12]. The main purpose of the proposed test

is to evaluate the safety performance of on-board storage

tanks subjected to localized fire. The tank is required to

experience a period of localized and then engulfing flame

impingement to accomplish the test purpose. The test con-

ditions are based on information derived from actual vehicle

fire tests [9].

There are few investigations conducted on this new fire

test. Tamura et al. [13] developed a propane burner to imple-

ment localized and engulfing fire, and studied the effects of

flame configurations on test results. Ruban et al. [14] per-

formed a localized fire exposure burst test, and showed that

the tank burst after 5 min and 20 s of fire exposure. So far,

practicality and validity of the proposed fire test still need to

be further verified. For example, installation locations of the

thermocouples are not definitely specified in GTR. Likewise,

influences of hydrogen and air as the filling media on test

results have not been experimentally confirmed.

In this paper, experimental investigation of localized fire

test was conducted on two type 3 tanks. Temperature varia-

tions during the fire experiments were obtained. Differences

in the test results that were caused by the use of hydrogen and

air as the filling media were studied in detail. Based on the

whole flame impingement process of the fire experiment, a

three-dimensional computational fluid dynamics (CFD)model

was developed to analyze the effects of filling pressure and

localized fire exposure time on TPRD activation.

Fig. 1 e Localized fire test set-up.

2. Experimental methods

2.1. Test set-up

A fire source control device was developed for the fire exper-

iments according to the test requirements in GTR. As shown in

Fig. 1, the fire source control device was composed of a kero-

sene burner, a tank bracket and a guide rail. The kerosene

burner with a length of 1650mm and a width of 600mm could

freely move along the guide rail to implement localized and

engulfing flame impingement on the test tank. Each fire

experiment used an equal volume of kerosene (i.e., 60 L).Wind

shields were applied to ensure uniform heating.

The fire experiments were conducted on two identical type

3 tanks that were respectively filled to about 18 MPa with

hydrogen and compressed air. As described in Fig. 1, a TPRD

was fitted onto the end boss of each tank. The activation

temperature of fusible plug of the TPRD was (383 � 5) K. The

tank was horizontally positioned with its bottom approxi-

mately 100 mm above the level of kerosene.

As specified in GTR, the localized fire area with a length of

approximately 250 mm was located on the tank furthest from

the TPRD (Fig. 1). During the experiments, temperatures

around the outside surface of the tank and the TPRD were

monitored by 19 thermocouples (type K). Nine thermocouples

were in the localized fire area. All the thermocouples were not

more than 250 mm apart, and located about 15 mm from the

outside surface of the tank in accordance with the test re-

quirements in GTR. The same arrangement of the thermo-

couples was used for each fire experiment. The pressure

variation of internal gas was monitored by a pressure

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transducer. A camcorder was used to record the whole test

process. The ambient temperature was about 273 K.

2.2. Test procedure

In this experimental investigation, a localized fire test proce-

dure was defined as shown in Fig. 2. During the localized fire,

the kerosene burner was placed on the side furthest from the

TPRD to ensure that only the localized fire area of the tankwas

exposed to fire. The flame directly impinged the tank surface

across its entire diameter. Temperatures of the thermocou-

ples located in the localized fire area should rise to about 873 K

during the first 180 s.

After the first 180 s period of localized flame impingement,

the kerosene burner was remotely pulled along the guide rail

to make the tank at the center of the burner. Thus, the fire

fully engulfed the tank along its entire length and width.

Temperatures of the thermocouples around the tank surface

should ascend to approximately 1073 K before the TPRD

opened to vent internal high-pressure gas. The test would be

terminated when the tank pressure fell to less than 1 MPa.

3. Numerical modeling

3.1. Mathematical model

A three-dimensional CFD model for simulating the whole

flame impingement process of localized fire test was

l and numerical investigation of localized fire test for high-of Hydrogen Energy (2013), http://dx.doi.org/10.1016/

Fig. 2 e Localized fire test procedure.

Fig. 3 e Schematic of the calculation region and partial

view of the grid structure.

i n t e rn a t i o n a l j o u rn a l o f h y d r o g e n en e r g y x x x ( 2 0 1 3 ) 1e8 3

developed using the CFD code FLUENT. The period of gas

discharge after TPRD activation was not included in the

model. To simplify the analysis model, the following as-

sumptions were proposed:

(1) The level and burning rate of kerosene were kept constant

since only a small amount of kerosene was burned before

TPRD activation in the experiment.

(2) The tightly attached layers of the tank wall ensured the

continuous variation of the temperature between adjacent

interfaces in the tank wall.

(3) This numerical analysis focused on the heat transfer

behavior of the tank under specific fire condition. Thus, we

assumed that fire had no effect on the structural integrity

of the tank.

The current model incorporated with various sub-models

such as the conductive and convective heat transfer model,

the renormalization-group (RNG) ke 3model, and the species

transport & finite-rate chemistry model. The mathematical

equations of the sub-models were described in our previous

paper [15], as well as the standard equations of mass, mo-

mentum and energy conservations. Moreover, the burning

rate of kerosene _m00 can be calculated by [16]:

_m00 ¼ _m00N

�1� e�kbD

�(1)

where _m00N is the burning rate of the kerosene pool with an

infinite diameter, _m00N ¼ 0:039 kg$m�2$s�1; kb is a constant,

kb ¼ 3.5 m�1; and D is the equivalent diameter of the fire

source in the experiment.

National Institute of Standards and Technology (NIST)

database [17] includes real-gas equations of state for hydrogen

and air which are separately described in Refs. [18,19] in detail.

Thus, NIST database was used for the analysis of internal

hydrogen and air. In addition, coupled wall was used in the

simulation of the whole localized and engulfing flame

impingement process. The SIMPLE algorithm was applied for

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the coupled solving of pressure and velocity field under the

unsteady governing conditions. The discretization of govern-

ing equations was accomplished using the second-order up-

wind scheme.

3.2. Geometrical model

A schematic of the calculation region is shown in Fig. 3. The

whole calculation region was a hemispheroid with a radius of

10 m. Pressure boundary was applied as the outlet in the

model. The fire source was divided into localized and engulf-

ing fire sources. According to the experiment, the fire source

was 1650mm long and 600mmwide. The tankwas positioned

100mm above the fire source. The length of the tank part over

the localized fire source was 250 mm. The localized fire

exposure time was set as 180 s.

Geometric parameters of the tank were identical to those

of the tank used in the experiment. As shown in Fig. 3, the

tank wall consisted of an aluminum liner and a carbon-fiber/

epoxy laminate. A copper alloy TPRD was located at the end

boss of the tank. TPRD was assumed to be activated when

the temperature of the active point (fusible plug of the

TPRD) reached 383 K. The initial temperature of the whole

l and numerical investigation of localized fire test for high-of Hydrogen Energy (2013), http://dx.doi.org/10.1016/

0 100 200 300 400 500 600 700200

400

600

800

1000

1200

1400

1600

1800

2000

Top surface in localized fire area (#2,9) Middle surface in localized fire area (#1,3,4,7,8) Bottom surface in localized fire area (#5,6) Top surface outside localized fire area (#10,17) Middle surface outside localized fire area (#11,12,15,16) Bottom surface outside localized fire area (#13,14) Region around TPRD (#18,19)

Tem

pera

ture

/K

Time/s

TPRD activation

(a) When the filling medium of the tank was hydrogen

i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n en e r g y x x x ( 2 0 1 3 ) 1e84

calculation region was set to 273 K in accordance with the

ambient temperature of the fire experiment. Thermophysical

properties of the tank and TPRDmaterials are listed in Table 1

[20e22]. Thermal conductivity and specific heat of carbon-

fiber/epoxy laminate were defined as the function of tem-

perature [20e22].

The hexahedral structured grid was adopted to mesh the

tank and its internal region, and unstructured grid was used

for some other parts of the model. The grids around the tank

and fuel combustion region were refined. A partial view of the

grid structure is also plotted in Fig. 3. Three models of 211,420

cells, 336,697 cells and 470,233 cells were developed, respec-

tively. The differences in calculation results between the

model of 336,697 cells and the model of 470,233 cells were

quite small. Since the calculating capacity of our high per-

formance server was enough for the developed models, the

model of 470,233 cells was chosen for the numerical analysis.

0 100 200 300 400 500 600 700 800 900200

400

600

800

1000

1200

1400

1600

1800

2000

Top surface in localized fire area (#2,9) Middle surface in localized fire area (#1,3,4,7,8) Bottom surface in localized fire area (#5,6) Top surface outside localized fire area (#10,17) Middle surface outside localized fire area (#11,12,15,16) Bottom surface outside localized fire area (#13,14) Region around TPRD (#18,19)

Tem

pera

ture

/K

Time/s

TPRD activation

(b) When the filling medium of the tank was compressed air

Fig. 4 e Average temperature variations of the

thermocouples located on the top, middle and bottom

surfaces of the tank and in the region around TPRD. #1e19

is the thermocouple number.

4. Results and discussion

4.1. Temperature variations during the experiments

During the fire experiments, temperature variations were

obtained from the thermocouples around the tank surface

and TPRD. The recorded temperatures of the thermocouples

in the same region can be averaged. The average temperature

variations of the thermocouples in different regions are

shown in Fig. 4. Temperatures differed for the thermocouples

that were located on the top, middle and bottom surfaces of

the tank. Namely, the temperature distribution on the tank

surface was uneven around the circumference. Thus, the

monitored temperatures would be affected by the thermo-

couple locations. However, installation locations of the ther-

mocouples have not been definitely specified in GTR. Random

installation of the thermocouples may produce inconsistent

test procedures and results.

As described in Fig. 4, the thermocouple temperatures of

the middle surface could immediately respond and quickly

rise to reach the target temperatures (873 K within 180 s and

1073 K after 180 s) due to the impingement of outer flame.

However, the thermocouple temperatures of the top and

bottom surfaces increased slowly and failed to achieve the

required temperatures before TPRD activation. The slow

temperature rises of the thermocouples on the top and bottom

surfaces may be caused by the shielding of the tank itself and

kerosene evaporation, respectively. Therefore, thermocouples

should be installed on both sides of the middle surface of the

Table 1 e Thermophysical properties of the tank andTPRD materials [20e22].

Material Densitykg/m3

ThermalconductivityW/(m$K)

Specificheat

J/(kg$K)

Aluminum 2700 218 902

Carbon-fiber/

epoxy laminate

1750 0.304 þ 0.000852T �19 þ 3T

Copper alloy 8978 387.6 381Fig. 5 e Experimental pressure and calculated temperature

variations of internal gas.

Please cite this article in press as: Zheng J, et al., Experimental and numerical investigation of localized fire test for high-pressure hydrogen storage tanks, International Journal of Hydrogen Energy (2013), http://dx.doi.org/10.1016/j.ijhydene.2013.02.052

Fig. 6 e Temperature distribution of internal hydrogen and outside surface of the tank.

i n t e rn a t i o n a l j o u rn a l o f h y d r o g e n en e r g y x x x ( 2 0 1 3 ) 1e8 5

Please cite this article in press as: Zheng J, et al., Experimental and numerical investigation of localized fire test for high-pressure hydrogen storage tanks, International Journal of Hydrogen Energy (2013), http://dx.doi.org/10.1016/j.ijhydene.2013.02.052

0 100 200 300 400 500 600100

150

200

250

300

350

400

450

500

10

15

20

25

30

35

40

45

50

Calculated temperature Simulation temperature

Tem

pera

ture

/K

Time/s

Filling medium: H2

Filling pressure: 18.05 MPa

Pre

ssur

e/M

Pa

Experimental pressure Simulation pressure

Fig. 7 e Comparison of pressure and temperature rises of

internal hydrogen between the simulation and

experimental results.

i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n en e r g y x x x ( 2 0 1 3 ) 1e86

tank along the longitudinal axis in the localized fire test. This

process ensures that the required temperatures can be

reached in time and the requirement of flame covering the

entire tank diameter can be satisfied.

In addition, the rise rates of temperatures around the TPRD

differed between the two fire tests (Fig. 4). This phenomenon

may be caused by the incompletely uniform heating of the

TPRD during the experiments. After the TPRD opened to

release internal gas, different temperature variations around

the TPRD were due to the different flammability of hydrogen

and air.

4.2. Influences of hydrogen and air as filling media

According to the ambient temperature and initial pressure of

the tank, the densities of internal hydrogen and air could be

obtained from the NIST database. Moreover, the densities

were kept constant before TPRD activation based on the iso-

choric properties of hydrogen and air. Therefore, the rise in

temperature could be calculated via the NIST database using

the values of the rise in pressure recorded through the ex-

periments. The experimental pressure and calculated tem-

perature variations of internal gas are presented in Fig. 5.

TPRD was activated during the period of engulfing flame

impingement in the two fire experiments. The pressure and

temperature rise processes of internal hydrogen and air were

similar from the ignition of the fire to the start of venting

through TPRD.

The air temperature was still much lower than the acti-

vation temperature of TPRD when the device was activated

(Fig. 5). The temperature of internal hydrogen was only

slightly higher than the activation temperature. However, the

temperatures around the TPRD were 453 K and 587 K for the

two fire tests at the time of TPRD activation, as observed in

Fig. 4. Therefore, TPRD may be activated mainly by the heat

from the outside of the tank. Moreover, the rising temperature

of internal hydrogen or air contributed little to TPRD

activation.

In addition, the pressure values of internal hydrogen and

air slowly increased before TPRD activation (Fig. 5). Based on

this figure, the increasing rate of hydrogen pressure

(0.0157 MPa/s) was only 6% larger than that of air pressure

(0.0148 MPa/s). So, the difference in the rising pressure of in-

ternal hydrogen and air was small. However, the discharge

rate of hydrogen gas was much higher than that of air. The

significant difference in the discharge rates was likewise due

to the different flammability of hydrogen and air. As shown in

Fig. 5, the discharged hydrogen deflagrated, whereas com-

pressed air only jetted to the environment. Deflagration of the

released hydrogen could greatly speed up the discharge.

Furthermore, the higher temperature and larger flame scale

caused by hydrogen deflagration would accelerate the degra-

dation of the tank. Even so, both tanks did not burst during the

fire experiments.

4.3. Comparison between simulation and experimentalresults

Temperature distribution of internal hydrogen and outside

surface of the tank is shown in Fig. 6. The temperature

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distribution was not uniform inside the tank, especially at the

calculation time of 180 s (Fig. 6a). In addition, the temperature

of outside surface unevenly distributed around the tank

circumference, and the temperature of the middle surface

was higher than that of the top and bottom surfaces (Fig. 6b).

The temperature distribution of outside surface obtained from

the simulation (Fig. 6b) was similar to that observed in the

experiment (Fig. 4). The differences in the temperature values

between the simulation and experiment were due to the fact

that the thermocouples were not in contact with the outside

surface of the tank in the fire tests.

Moreover, a comparison of pressure and temperature rises

of internal hydrogen between the simulation and experi-

mental results is described in Fig. 7. The hydrogen discharge

process was not included in the model, and thus only the

variations of pressure and temperature before TPRD activa-

tionwere compared. Since the temperature distribution inside

the tankwas uneven, volume-average temperature of internal

hydrogen was recorded and compared with the NIST calcu-

lated temperature. The simulation results can be basically in

agreement with the experimental data (Fig. 7).

4.4. Effect of filling pressure

As required in GTR, the tank should be pressurized to its

nominal working pressure (NWP) in the localized fire test. The

NWPs of on-board hydrogen storage tanks are different and

generally set as 35 MPa and 70 MPa. Moreover, the actual

pressure of the tank in HFCV greatly varies during the service.

So, the effects of different filling pressures on TPRD activation

were studied through the simulation.

The TPRD activation time and maximum pressure of in-

ternal gas with different filling pressures are shown in Fig. 8.

The variations of the filling pressures had a weak influence on

the TPRD activation time. This characteristic may be attrib-

uted to the rising temperature of the active point (fusible plug

of the TPRD), which was mainly caused by the heat trans-

mitted from the outside of the tank during the engulfing fire.

Furthermore, TPRD activation time was slightly affected by

hydrogen and air as the filling media with the same filling

pressure. The minimum test pressure of the tank is specified

l and numerical investigation of localized fire test for high-of Hydrogen Energy (2013), http://dx.doi.org/10.1016/

0

10

20

30

40

50

60

70

80

90

100

18.05MPa 35MPa 70MPa0

100

200

300

400

500

600

700

Max

. pre

ssur

e be

fore

TPR

D a

ctiv

atio

n/M

Pa

TPR

D a

ctiv

atio

n tim

e/s

Filling pressure/MPa

Time-H2 Max. pressure-H2

Time-Air Max. pressure-Air

Localized fire exposure time: 180s

Fig. 8 e TPRD activation time and maximum pressure of

internal gas with different filling pressures.

i n t e rn a t i o n a l j o u rn a l o f h y d r o g e n en e r g y x x x ( 2 0 1 3 ) 1e8 7

as 1.5 NWP in existing standards [5e8]. For the different filling

pressures in current simulation, the maximum pressures of

internal gas (Fig. 8) were still lower than the minimum test

pressures due to the slow pressure rise during the flame

impingement process.

4.5. Effect of localized fire exposure time

Vehicle fire test data showed that the period of on-board

storage tanks subjected to localized fire usually lasted for

less than 600 s [9]. The recommended localized fire exposure

time is still under discussion. Consequently, the effects of

different localized fire exposure times on TPRD activation

were analyzed in current simulation.

The TPRD activation time and maximum pressure of in-

ternal gas with different localized fire exposure times are

described in Fig. 9. TPRD activation time increased when the

localized fire exposure time was extended. However, the

maximum pressure of internal gas did not significantly vary

with the localized fire exposure time.

The case that the tank consistently experienced localized

fire was also studied to confirm the TPRD activation time.

0

8

16

24

32

40

180s 480s 600s0

200

400

600

800

1000

Max

. pre

ssur

e be

fore

TPR

D a

ctiv

atio

n/M

Pa

TPR

D a

ctiv

atio

n tim

e/s

Localized fire exposure time/s

Time-H2 Max. pressure-H2

Time-Air Max. pressure-Air

Filling pressure: 18.05MPa

Fig. 9 e TPRD activation time and maximum pressure of

internal gas with different localized fire exposure times.

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Simulation results indicated that TPRD could not be activated

within 1800 s of localized fire exposure. Based on the data

from the localized fire exposure burst test [13,14], the tank

probably burst before TPRD was activated to release internal

high-pressure gas in such a circumstance.

5. Conclusions

The experimental investigation of localized fire test was

conducted on two type 3 tanks. A three-dimensional CFD

model for simulating the whole flame impingement process

was developed. Based on the experimental and numerical

studies, the following conclusions can be obtained:

(1) The temperature distribution on the tank surface was

uneven around the circumference during the flame

impingement process. Thermocouples should be installed

on both sides of the middle surface of the tank along the

longitudinal axis in the localized fire test.

(2) The rising temperature of internal hydrogen or air

contributed little to TPRD activation. TPRD may be acti-

vated mainly by the heat from the outside of the tank in

the fire experiments.

(3) The difference in the rising pressure of internal hydrogen

and air was small. However, hydrogen could discharge

much faster than air. Regardless of these differences, both

tanks did not burst during the fire tests.

(4) The TPRD activation time was slightly affected by the

variations of the filling pressures, but it increased when

the localized fire exposure time was extended.

Acknowledgments

This research is supported by the National High Technology

Research and Development Program of China (863 Program,

Grant No. 2012AA051504), the Research Fund for the Doctoral

Program of Higher Education of China (RFDP, Grant No.

20110101130004), and the National Natural Science Founda-

tion of China (NSFC, Grant No. 51206145), and the Funda-

mental Research Funds for the Central Universities.

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