Paper number

PREDICTION OF THE BEHAVIOR OF A CNG TANK EXPOSED TO FLAMES

F. D’Introno, M. Torresi, S. M. Camporeale, B. Fortunato Dipartimento di Ingegneria Meccanica e Gestionale – Politecnico di Bari

C. Amorese, S. De Matthaeis Centro Ricerche FIAT – Valenzano (BA)

2005 SAE_NA section

ABSTRACT

The increasing request of vehicles fuelled by Compressed Natural Gas (CNG) requires that the safety issues related to this kind of fuel should be analyzed more and more thoroughly. So far, the minimum requirement for a CNG tank is to set up a Pressure Relief Device (PRD) triggered by temperature.

This work aims to develop a procedure able to perform thoroughly tank behavior analyses, in presence of fire, in order to reduce the number of expensive experimental tests and to improve the design criteria of tank safety devices. The procedure consists in the solution of the unsteady solid-fluid conjugate heat transfer problem, by means of numerical simulations, in order to evaluate the transient behavior of both the temperature field inside the solid walls of the tank and the thermo-fluid-dynamic field of the compressed gas inside it, when it is exposed to a heating source. Moreover, the melting process of the fusible plug, inside the PRD, is simulated. In order to validate this procedure, experimental tests have been carried out on a suitable test rig.

Several heating conditions, even unlikely, but able to jeopardize safety, have been considered. The condition characterized by a flame located on the opposite side of the PRD has been recognized to be the most critical with respect to explosion risk. In this case, the relationship between time and tank pressure has been numerically evaluated. The fusible plug, inside the PRD, has been checked to be completely melted of the fusible plug before the pressure reaches the safety limit.

INTRODUCTION

Compressed natural gas is the most promising alternative, in the near future, to traditional fuels in order to reduce pollutant emissions from passenger-cars and light and heavy-duty vehicles [1], [2], [3]. Actually, natural gas is characterized by optimum air/fuel mixing, no wall wetting phenomena, accurate fuel metering, no soot and a significant reduction of gaseous pollutant emissions

(CO, NOx and Non Methane Organic Compounds – NMHC). Furthermore, natural gas, being a high-octane fuel, has a good knocking resistance and doesn’t need anti-knocking organic compounds [4], [5], [6]. Moreover, due to its high H/C ratio, a reduction of CO2 emissions, currently considered the main responsible of the “green house” effect [7], can be obtained.

Unfortunately the autonomy of CNG fuelled cars is lower than those fuelled either by petrol or by diesel oil, even using high pressure tanks. Natural gas stored at 200 bar (typical value for this kind of applications) and ambient temperature has a density of 140 kg/m

3, much lower

than that of petrol (750-755 kg/m3) or diesel oil (815-

855 kg/m3); therefore, for the same mass of fuel, the

CNG tank volume is about 5-6 times larger than that of petrol or diesel oil tanks.

Excellent environmental features but limited autonomy make CNG attractive for public transportation (a large fraction of new buses are fed by CNG) or urban passenger car fleets; moreover, due to the increasing number of filling stations and the lower cost of this fuel, the number of private passenger cars, fuelled by CNG, is constantly increasing.

The growing number of vehicles fuelled by natural gas requires to analyze more and more thoroughly the safety issues of this technology in order to raise continuously the safety level.

Current law specifications impose that the tanks must be equipped with a PRD triggered by temperature; this device contains a fusible plug, which melts in case of anomalous temperature increase (up to 110±10°C according to law specifications [8]) allowing the CNG to be discharged in order to avoid explosions.

The PRD could be heated up either by the conductive heat flux through the solid walls of the tank or by the convective heat flux through the compressed gas inside the tank. In order to understand which is the main mechanism, causing the fusible plug melting, both the

transient temperature field in the walls and the thermo-fluid-dynamic field of the CNG tank being exposed to fire have been numerically simulated while. Preliminary test cases have been carried out [9], in order to verify the adopted numerical procedure.

At the same time, the fusible plug has been verified to be completely melted before pressure inside the tank reaches the safety limit (450 bar according to law specifications [8]), by means of a solidification and melting model.

Several heating conditions have been simulated; among them, the particularly unfavorable case with a flame located on the opposite side of the PRD has been considered. In this case, the heat must flow through the whole tank before it can reach the PRD; consequently, the time needed to actuate the PRD could be too high, implying an excessive pressure increase inside the tank. This configuration is taken as reference for the designing criteria of the fusible plug, allowing its safe operation under any other fire condition.

The experimental validation of the model has been carried out by setting up a suitable test rig. For safety reasons, the experiments have been carried out on a tank filled with air at atmospheric pressure. The validated model has been used to simulate the same phenomenon in “real” condition. Finally the fusible plug melting process has been simulated imposing, at the bottom of the tank, the temperature as computed in the previous unsteady simulations; based on the results, a fusible plug design has been proposed allowing its complete melting before reaching the pressure safety limit.

EXPERIMENTAL SET UP

In Fig. 1 the experimental rig is sketched. The tank (diameter D = 250 mm, height H = 1600 mm), with the PRD on the top, is kept in vertical position by means of a metallic frame and is placed on a propane burner; between them a metallic shield is interposed in order to avoid direct flame impingement and to protect the thermocouple (TC1) placed at the bottom of the tank; heat flows from the bottom to the top of the tank, in order to have the heating source as far as possible from the PRD.

The tank is equipped with eight thermocouples (K-type) and a pressure transducer (piezoresistive sensor with integrated amplifier and signal conditioner) inside the tank.

Thermocouples are opportunely placed in order to appreciate temperature gradients along the tank sidewalls (TC2 - TC5), through the wall (TC5 - TC6) and inside the tank (TC7 - TC8).

On the tank flange, two cable passages with gaskets have been placed allowing the instrumentation cabling and, at the same time, the hermetic closure; this condition is essential to evaluate the increase of the

pressure inside the tank during the experiment. Moreover the flange is equipped with a one-direction valve to fill the tank and a PRD for safety.

Fig. 1 Experimental setup

Two internal thermocouples (TC7 - TC8) are suspended by cables at assigned heights; while the third one (TC6) is placed on the internal wall, in a position corresponding to the thermocouple TC5. The external thermocouples are fixed to the tank sidewalls and opportunely shielded from the direct flame impingement.

NUMERICAL MODELS

The commercial code, FLUENT 6.1.22, has been used to carry out the simulations. This code allows to treat contemporarily both fluid and solid regions allowing to solve the conjugate heat transfer problem.

The unsteady compressible three-dimensional RANS equations are discretised by means of a finite volume approach. The pressure velocity coupling is achieved by means of the SIMPLE algorithm (Semi-Implicit Method for Pressure-Linked Equations). The convection terms are discretised using a second order accurate upwind scheme, pressure and viscous terms are discretised by means of a second order accurate centred scheme. A second order implicit time formulation is also used. The Boussinesq approach is applied to relate the Reynolds stresses to the mean velocity gradient. Turbulence has been modeled by means of the standard k-ω model, which solves transport equations for the turbulent kinetic energy (k) and for the specific dissipation rate (ω).

Inside the solid regions the code solves the following energy transport equation:

( ) ( )Tkhρt

c∇⋅∇=

∂

∂ (1)

where ρ is the density, h is the sensible enthalpy, kc is the thermal conductivity and T is the temperature.

In order to verify the numerical procedure dealing with the conjugate heat transfer problem, preliminary test cases have been carried out. The laminar flow over both thin and thick flat plates have been considered (Fig. 2), comparing the numerical results in terms of local Nusselt number Nux either with analytical solution or with numerical results obtained with a different code [10].

Fig. 2 Temperature distribution for the laminar flow over both thin and thick plates

Then, the low turbulent natural convection problem inside an air filled square cavity [11] has been carried out. By means of the standard k-ω model, a good agreement between numerical and experimental data has been obtained (Fig. 3). Further details on these tests are available in [9].

Fig. 3 Nusselt Number along the hot wall

Fig. 4 Computational domain reproducing the test rig

For the numerical simulations, the tank geometry has been simplified neglecting the flange. Actually, a rectangular computational domain has been considered to carry out a 2D axial-symmetric simulation (Fig. 4).

A 5000-cell structured grid has been meshed. The 4mm iron tank walls have been considered carrying out the conjugate heat transfer problem.

Unsteady simulations have been carried out by means of a second order accurate scheme with a 0.5s time-step; the time period considered is 1200s, which means 72000 sub-iterations (a dual time-stepping algorithm has been used with 30 sub-iterations). The selected time period is a compromise between the need for a good comparison with experimental results and to reduce the computational time effort for every test.

In Fig. 4, the model boundary condition are reported. The temperature time law (T = f(t)), imposed at the tank bottom, is obtained from the data recorded by the TC1 thermocouple, during the experimental tests (Fig. 5). To take into account the heat flux at sidewalls from the shield, a fictitious heat source with a linear temperature profile (T∞) has been considered.

In this case, since the metallic shield temperature is variable in time, a suited mean value (T∞ max = 100 °C) has been imposed.

Tab. 1 summarizes all the code options being applied in the numerical simulation.

As far as concerns to the fusible plug melting process, this phenomenon has been carried out separately by means of the melting model available in FLUENT.

This model considers a unique transition zone treated as a porous medium whose porosity is equal to the liquid fraction β. The liquid fraction is defined as follow:

liquidus

liquidussolidus

solidusliquidus

solidus

solidus

if1

if

if0

TTβ

TTTTT

TTβ

TTβ

>=

<<−

−=

<=

(2)

and indicates the cell volume fraction in a liquid state. Actually, the transition region is the domain portion where the liquid fraction ranges from 0 to 1. When the material is completely solidified, porosity becomes zero as well as velocities. Momentum source terms are added to the momentum transport equation accounting for the pressure drops due to solid fraction. Source terms are also added to turbulence transport equations accounting for the reduced porosity of solid regions.

The material enthalpy is computed as summation of the sensible enthalpy and the melting latent heat, equal to the product of the material latent heat time the liquid fraction. The temperature computation follows an iterative procedure between the energy transport equation and the liquid fraction definition, according to Voller and Swaminathan [12].

RESULTS

Fig. 5 and Fig. 6 show respectively the tank bottom temperature and the pressure inside the tank recorded during the three experimental tests.

Tab. 1 FLUENT set up for the simulations

Solver segregated

Fluid compressible

Discretisation 2nd order upwind

Turbulence model

standard k-ω

P = 1 bar

gravity (gy = -9.81m/s2)

Operating condition

floating operating pressure

iron (tank)

Material

air (inside the tank)

unsteady (∆t = 0.5 s)

2nd order implicit Iterative procedure

30 sub iterations every physical time step

Since it was difficult to determine the heat flux from the flame, the temperature time histories were used as boundary conditions at the tank bottom for the three corresponding numerical simulations. Every time the heat flux has been changed, so that different temperature and pressure rises have been obtained.

Fig. 5 Temperature rise at the tank bottom

Fig. 6 Pressure rise inside the tank

The pressure rise was very low (the maximum pressure rise ∆p registered in the third test was only 4500 Pa), due to the high tank volume (78.5 liter), the reduced heat flux and the low time period of the tests. Anyway, the numerical model results effective to compute the pressure variation ∆p for all the three tests, as shown in Fig. 7.

Fig. 8, Fig. 9 show the comparison of the temperature gradients numerically computed and experimentally recorded, respectively along sidewalls (thermocouple TC2 - TC5) and inside (thermocouple TC6 - TC8).

Generally, a good agreement can be seen between numerical and experimental results, even at the tank bottom (mean difference about 10% at thermocouple TC2 and 5% at thermocouple TC8) where the highest difference can be registered. This is due to the radiant heat flux from the shield, being probably underestimated by the numerical model.

TANK BEHAVIOR UNDER FIRE EXPOSURE

The numerical model, which has been experimentally validated, has been used to simulate the behavior of the CNG tank at 200 bar under fire exposure.

The simulation of the tank has been carried out taking into account the presence of the safety valve (Fig. 12). The PRD is integrated inside it, as shown in Fig. 10 and Fig. 11.

The temperature law, which acts on the PRD during the fire exposure, has been obtained and used to simulate the melting process of the fusible plug.

Fig. 7 Comparison of pressure rise predicted by simulation and experiments

Fig. 8 Temperature along sidewalls (t = 1200 s)

Fig. 9 Temperature inside the tank (t = 1200 s)

Fig. 10 Safety valve with integrated PRD

Fig. 11 Longitudinal section of the safety valve

In the computational domain, some details have been neglected such that the axial symmetry hypothesis could be satisfied, reducing the computational effort.

A representative sketch of the computational domain with its boundary conditions is shown in Fig. 12. A 1000 K (727° C) flame temperature, at the bottom of the tank, has been assumed. It has been supposed the heat source has a linear temperature profile along sidewalls.

The numerical method is the same as in the previous simulations (Tab. 1). The simulation time period is equal to the time the inside pressure needs to reach the safety limit (450 bar).

Fig. 12 Boundary conditions for the fire test simulation

Fig. 13 shows the behavior of the pressure inside the tank (being computed as volume integral averaged) and the corresponding temperature of the PRD during the simulation. Considering the CNG at an initial pressure of 200 bar, equal to typical automotive operating pressures, a 450 bar pressure is reached in 1200s while the PRD temperature reaches almost 200°C (Fig. 13).

Fig. 13 CNG pressure and safety valve temperature under fire exposure

Fig. 14 Polynomial fitting of the safety valve temperature

Fig. 14 reports the temperature polynomial curve fit (T = 1.0808e

-4 t

2 + 27, with T in [°C] and t in [s]), used as

boundary condition for the melting process simulation.

The temperature field and the streamlines inside the tank at the end of the simulation are shown in Fig. 15.

Finally, the melting process of the fusible plug inside the PRD has been simulated. The simulation time period has been assumed to be equal to the time needed to the CNG to reach the 450 bar pressure, in order to evaluate the melting level of the fusible plug.

Even if the melting process has been simulated by means of an extremely simplified model (the liquid front moves forward as the fusion temperature isotherm does), the numerical results can be considered reliable enough, for the goal of this work, which doesn’t aim to give a detailed description of the melting process but to estimate the time needed to let it melt.

Tab. 2 Low melting point alloy properties

Density = 10.44 [g/cm3]

Specific heat = 0.140 [J /(g °C)]

Thermal conductivity = 4 [W/(m K)]

Melting point = 110 [°C]

Melting latent heat = 20.9 [J/g]

Fig. 15 Temperature field and streamlines inside the tank after 1200 s

Fig. 16, Fig. 17 show the brass bolt containing the fusible plug and the boundary conditions for the melting process simulation. The bolt is composed by a cylindrical screwed body (to be fixed to the PRD) and a hexagonal head with six holes to allow the extrusion of the melted plug. The temperature time law (T = f(t)), which reproduce the PRD heating (Fig. 14), is imposed as boundary condition at the bottom surface of the bolt; the bolt head, being outside, is considered at 27°C (i.e. 300K). The fusible plug is realized with a low melting point alloy (Bi = 51%, Sn = 25%, Cd = 20%, Pb = 4%; Tab. 2).

Fig. 16 Brass bolt containing the fusible plug

This alloy has been designed on purpose to melt at 110°C, as required by the current law specifications.

Fig. 17 Sketch of the fusible plug inside the brass bolt

Fig. 18 shows the evolution of the liquid front inside the bolt at two different instants: at the end of the simulation (t = 1200 s), the fusible plug isn’t yet completely melted, since the liquid front has only reached the bolt head.

From the numerical results it seems that, under the fire exposure, after a time period of 1200 s, equal to the time the CNG needed to reach a 450 bar pressure, the fusible plug isn’t sufficiently melted to grant its extrusion and the CNG discharge. Nevertheless, it must be considered that

the extrusion of the alloy starts before its complete melting, due to plasticization of the fusible plug under the CNG pressure load that, in this simulation, has been neglected. For instance, considering the high pressure level inside the tank, it can be reasonably supposed that the passage hole inside the bolt head would be opened even if the alloy isn’t yet completely melt.

Fig. 18 Liquid front evolution inside the fusible plug

Anyway to allow the complete opening of the passage holes under any fire exposure without the reaching of the 450 bar limit inside the tank, a new fusible plug, having a volume equal to the liquid portion (Vliquid) computed in the numerical simulation, has been designed.

Fig. 19 Liquid fraction evolution during the simulation

Since at the end of the simulation the liquid fraction is 65% of the whole volume, as being shown in Fig. 19, supposing to leave unchanged the dimension of the bolt

head (Fig. 17) and modifying only the screw body length, the new length is x = 9.5 mm.

Fig. 20 summarizes the final dimensions of the brass bolt; anyway this design, being only based on numerical results, needs to be verified. In fact, the integrity of the fusible plug needs to be granted also under regular operating conditions.

Fig. 20 Fusible plug design

CONCLUSION

The presented study concerns the thermo fluid dynamic analysis inside an automotive CNG tank under a fire exposure, being considered particularly dangerous in terms of safety. The aim of this work is to describe the system behavior under this fire condition and, particularly, verify the right operation of the PRD as ruled by the current law to prevent from high pressure hazard inside the tank.

Suited experimental tests have been set up, trying to reproduce the same phenomena to be carried out by means of numerical simulations and to validate the proposed model. After the model accuracy has been verified through the numerical and experimental result comparison, the tank operating condition under fire exposure has been simulated since no experimental tests could have been carried out for safety reasons.

By means of the numerical model previously validated, the tank behavior under fire exposure has been analyzed, computing the time needed to reach the pressure safety limit inside the tank; based on the obtained results, a design of the fusible plug has been proposed aiming to grant the right operation of the PRD under anomalous increase of the temperature (and pressure) inside the tank.

In consideration of the simplifications, which characterize the model, and the need of a suitable validation of the experimental results (the law imposes a bonfire test to verify the safety system of the CNG tank), this work is a starting point for a deeper analysis on the current safety systems or the design of innovative devices.

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