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Numerical analysis of accidental hydrogen release in a laboratory
international journal of hydrogen energy 35 (2010) 4409–4419
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Numerical analysis of accidental hydrogen release in a laboratory M. Heitsch*, D. Baraldi, P. Moretto JRC, Institute for Energy, Pb 2, NL-1755 ZG Petten, The Netherlands
article info
abstract
Article history:
In a hydrogen laboratory hydrogen may accidentally leak from the high pressure systems
Received 22 August 2009
and it may form a flammable mixture after mixing with air. CFD (Computational Fluid
Received in revised form
Dynamics) simulations of accidental hydrogen releases in a laboratory are performed in
14 January 2010
order to estimate the size and the resident time of the flammable cloud inside the labo-
Accepted 14 January 2010
ratory. Moreover the effectiveness of the hydrogen sensor positions in detecting the
Available online 20 March 2010
presence of the hydrogen in the laboratory within few seconds has also been investigated. In a first scenario, the hydrogen coming from the storage is pumped by a compressor into
Keywords:
the laboratory pipelines at constant pressure. To be conservative, it is assumed that the
Hydrogen
pressure can be maintained when the leak in the pipeline occurs. In a second scenario, the
Safety
hydrogen flows directly from the storage into the laboratory pipeline without the
Jet release
compressor, and the pressure in this case is decreasing with time. In both cases, shut-off
Critical flow
valves have been neglected in order to consider a worst case scenario. The simulation
Flammable cloud
results from the two accident scenarios are compared, showing that the configuration with the compressor will potentially produce larger amounts of flammable mixture in the laboratory. ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved.
1.
Introduction
In the last years international and national research programmes have invested large amounts of resources in extensive experimental research aiming at the development of technologies for hydrogen production, storage, transport and conversion. To this purpose laboratories and facilities have been newly constructed or converted with an increased local hydrogen storage and usage quantities. To guarantee safe operations various regulations and guidelines are available to designers and operators, such as the Pressurised Equipment Directive [1] and the ATEX directive [2] in Europe. To model accidental situations and eventually assist in a safety design, various analysis methods are advised by the guidelines, such as the Failure Mode and Effects Analysis (FMEA), the Hazard and Operability Studies (HAZOP), and the
Failure Mode and Effects Criticality Analysis (FMECA). As effective tools to be used to evaluate in details specific accidental scenarios, Computational Fluid Dynamics (CFD) analysis can be performed and the use of CFD is also advised by the guidelines. CFD codes have been extensively used to simulated hydrogen accident scenarios in garages [3–7], in tunnels [8–13] and in re-fuelling stations [14,15]. The authors performed CFD investigations of accidental release scenarios in a relatively small laboratory (approximately 1.2 kg of hydrogen stored) for hydrogen storage technologies. The safety of the laboratory was designed by means of the FMEA method and it respects the above mentioned directives. By performing the CFD studies, various goals were pursued: the evaluation of the laboratory general safety level with a more sophisticated 3-D method, the correctness of the safety design assumption and more specifically the
* Corresponding author. Tel.: þ31(0)2245140; fax: þ31(0)224565623. E-mail address: [email protected] (M. Heitsch). 0360-3199/$ – see front matter ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2010.01.044
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optimisation of the hydrogen detection system used for an early alarm and automatic operative shut down. To this purpose various release scenarios have been adopted and by CFD simulation flow rates and volumes of flammability were calculated and local hydrogen concentration increases were identified. We report here only on the worst case scenario, beyond the maximum credible accident which was considered during the risk assessment and the safety design. This scenario corresponds to the full and instantaneous break of a high pressure hydrogen distribution pipe. The authors believe that this publication could stimulate the adoption of 3dimensional simulation tools for the optimisation of safety design also for less critical cases, such as small hydrogen laboratories and equipments. The simulations were carried out with the CFD code CFX. Different modelling options are investigated and results in terms of flammable mass and volume in the laboratory as a consequence of the hydrogen releases are given. The work presented here is an extension of the simulations presented earlier in [16].
2.
Problem description
The laboratory under study is a typical installation for the investigation of hydrogen storage capacity in materials. It is equipped with a hydrogen supply and a distribution system which delivers hydrogen at 20 MPa maximum to several working places where the experimental equipment is installed. The laboratory safety has been designed by performing an FMEA (Failure Mode and Effects Analysis) and determining the maximal credible accident. In order to validate the assumption, various leakage scenarios at different locations have been simulated by numerical simulations. We report here on the simulation results of a release case which has not been considered during safety design because beyond the maximum credible accident; the instantaneous rupture of a hydrogen distribution pipe at 20 MPa. The numerical simulations were carried out by means of the CFD (Computational Fluid Dynamics) code Ansys CFX [17]. Gas distribution in the laboratory under specific boundary conditions was calculated at first, followed by the calculation of the size and the duration of a flammable hydrogen cloud in the laboratory. Finally, the effectiveness of the positions of existing hydrogen sensors in the laboratory can be checked. The main micro processes in the release jet of the leak were resolved to a certain degree but have not been the main focus of this work. The resolution of the jet structures is necessary for the correct prediction of the hydrogen dispersion in the far-field of the jet.
2.1.
a volume of approximately 285 m3. A photo of a part of the laboratory is given in Fig. 1. Hydrogen is delivered at 20 MPa at maximum from a 50 l gas bottle in combination with a compressor. Bottle and compressor are located outside the laboratory. Two hydrogen distribution operative configurations are available: the use of a compressor allows to maintain the hydrogen delivery to the laboratory at constant pressure, while the direct connection of the hydrogen gas bottle to the distribution line causes a continuous decrease of the pressure due to hydrogen consumption. However, it simplifies the system also from the safety control point of view. In the following the two configurations will be referred to as ‘‘bottle-compressor’’ and as ‘‘bottle-only’’. The maximum amount of hydrogen which would be released in the laboratory in case of failure of all automatic and manual shut down procedures can be calculated by considering a full bottle, a full compressor buffer tank, all the pipes at 20 MPa and all the experimental equipments at full storage capacity. That amount is about 1.2 kg. A distributed ventilation system provides the laboratory with a constant flow of air. As shown in Fig. 2, in the ventilation system there are 8 inlet openings located on the ceiling, and 10 extraction openings, 6 of which are placed on the ceiling and 4 on a wall, at the floor level (in Fig. 2 the inlets are shown in blue while the exhaust openings are marked in red and turquoise). In addition, air extraction fume hoods are located on top of each test stand in order to collect hydrogen possibly released during experiments and discharge it to the roof of the building (names are RIG, IVR, AMC; see Fig. 2). Two independent leak detection systems are available: an overflow valve in the main hydrogen supply line, and safety hydrogen detectors under each fume hood. When the flow rate in the hydrogen supply system or local hydrogen concentrations in the lab exceed certain threshold values, an
The laboratory
The laboratory under study is a typical facility for the investigation of hydrogen storage capacity in materials located at the Institute for Energy of the Joint Research Centre (JRC). It consists of sorption equipments based on gravimetric, volumetric and spectrometric principles. They are performed typically in a pressure range of 0.1 up to a maximum of 20 MPa and with temperature from 77 K up to 273 K. The laboratory is approximately 18 m long, 8.5 m wide and 3.5 m high, with
Fig. 1 – Picture of a part of the laboratory: Vin indicates the positions of the ventilation inlets, Vout the positions (on the ceiling and near the ground) of the ventilation outlets, H two of the fume hoods positioned on top of the test rigs. Number 2 indicates the position of the hydrogen leak that is investigated in the simulations.
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Fig. 2 – Ventilation system of the laboratory (inlets – blue; outlets – red, turquoise). The black arrow indicates the viewing direction of the observer in the photo of Fig. 1. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
automatic shut down procedure starts within less of a second. In that case the hydrogen supply is interrupted and the distribution system is depressurized. Any detector triggers the shut down procedure when the volumetric hydrogen concentration passes 0.4%. This corresponds to one-tenth of the lower flammability limit (LFL) of hydrogen in air which is about 4%. Local air velocities were measured by means of a calibrated anemometer at all openings of the ventilation system and then averaged in order to obtain a single value per opening. The velocities range is between 1 and 2 m/s, and the total airflow entering the laboratory sums up to approximately 15000 m3/h. The same procedure was applied to the air extraction pipes under the fume hoods, where the air velocity range is between 1.5 and 5.5 m/s. These data have been used in the CFD calculations as boundary conditions.
2.2.
Accident scenario
The selected accident scenario consists in the instantaneous rupture of a hydrogen pipe with an internal diameter of 4 mm. In order to study the release of the full hydrogen content in the laboratory, it has been further assumed that all the automatic detection and shut down systems do not operate. The only available operative system is the ventilation system of the laboratory. The hydrogen system is under full pressure (20 MPa) and therefore the total available amount of hydrogen (1.2 kg) in the bottle and in the pipe system is released into the room. A realistic location of a leak is where pipes are connected. There are several pipe connections in the laboratory. However, the pipe joints located underneath the fume hood RIG (Fig. 2) is most often subject to human intervention. It is therefore considered to be the location where the simulated pipe rupture takes place. The full cross section of the pipe with 4 mm diameter serves as release opening. Further, a horizontal blowing direction is assumed towards to the opposite wall of the laboratory (direction to hood AMC). The accident can be subdivided into two phases. The first phase is dominated by the critical release of hydrogen into the laboratory. The ventilation is active but is only able to influence the gas distribution at a distance from the jet. The flow
field created by the jet is dominating. The second phase starts when the pressure in the hydrogen storage equals to the environmental pressure and the hydrogen outflow stops. Then the forced ventilation system becomes able to remove more effectively the accumulated hydrogen from the laboratory. Both phases are covered by the simulations.
3.
Simulation preparation
Prior to the simulation of the hydrogen leak, the flow field in the laboratory due to the ventilation flows is calculated and used as initial state for the leak simulations. Two principal model options concerning the hydrogen leak were applied. The first option assumes a constant sonic flow through the leak until the total amount of hydrogen from the hydrogen storage outside the laboratory has been discharged. It corresponds to the operative case ‘‘bottle-compressor’’ described in 2.1. This operative mode assumes that the hydrogen compressor is able to maintain a constant maximum pressure of 20 MPa in the hydrogen distribution lines. The assumption of a constant pressure during the whole discharge is a simplification of the real situation, because the compressor is not expected to be able to guarantee the requested constant pressure especially towards the end of the transient. However, it simplifies considerably the calculation and it represents a conservative approach from safety point of view, because it assumes a quicker release of the available gas than in the real case. The second option includes the storage by its volume and simulates the hydrogen leak as an expansion process from the storage starting from an initial internal pressure of 20 MPa. This modelling option considers the relation between pressure and mass in the storage and calculates consequently a time dependent leak flow. It corresponds to the ‘‘bottleonly’’ operative case described in 2.1, where the hydrogen compressor is bypassed and a gas bottle at 20 MPa is directly connected to the hydrogen distribution line. The second option could also use a simplifying approach to calculate the leak flow into the laboratory. In literature several algorithms are proposed, for example by Birch et al. [18] and
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by Xiao et al. [19]. These algorithms avoid the full calculation of an under-expanded jet with very small cells close to the leak and short time steps by providing an algorithm which allows starting the calculation downstream of the physical leak. The algorithm preserves the main parameters of the jet flow, simplifies the calculation and saves computing time. However, deficiencies are observed in the turbulence level and the actual temperature at the position where the simplified calculation starts (fictitious diameter). Gas entrainment upstream of the position of the fictitious diameter cannot be dealt with. Further, in the context of the laboratory connected to a relatively small tank (transient behaviour), an extra tool would be needed to calculate the leak flow rate depending on the pressure in the hydrogen tank. In order to avoid the need of an extra code together with additional assumptions on flow parameters at the position of the fictitious diameter it was decided to calculate the leak flow without a simplifying algorithm. This is accomplished at the expense of a higher computational effort and demonstrates at the same time that a full simulation as part of a long transient is possible with average computer resources. The two modelling options concerning the hydrogen leak and the results obtained are discussed in detail in the following sections.
3.1.
Computer model of the laboratory
The numerical model of the laboratory from Fig. 1 contains only the medium and large size components. All small size devices like computers, analytical equipment, cables and pipes were neglected because it is assumed that they cannot influence the global hydrogen distribution. All venting openings, fixed side boards, the fume hoods and two larger devices under the fume hoods RIG and IVR are included in the computational model. Some of the ventilation inlets (blue colour in Fig. 2) release air in two directions some only in one direction. The ventilation outlets in the model were not located directly in the walls but were placed outside of the laboratory to allow the flow to develop in such a way that the pressure across the openings becomes uniform. Fig. 2 indicates the approximate location of the assumed leak under hood RIG. Two principal modelling options for the leak flow are followed in this paper. One option predefines the mass flow through the leak and only a short pipe to introduce the flow is included in the numerical grid. The other option calculates the mass flow from the actual mass, temperature and pressure in the hydrogen storage against the conditions in the laboratory. The size of the hydrogen storage was calculated to contain the total hydrogen mass available in the whole hydrogen system. This is a conservative approach because it neglects the possibility of some sections of the system to be isolated which therefore would not contribute to the hydrogen release. For this option the storage and the leaking pipe are included in the numerical grid. The mesh for the numerical simulations is composed of hexahedral cells in most of the model. Only close to the roof and downstream the leak an unstructured mesh composed of tetrahedral and pyramidal cells was employed. The use of a hybrid mesh with structured and unstructured cells offers
the flexibility needed for inserting very small entities (leak pipe) into a given hexahedral mesh and keeps the number of cells within manageable limits. The total number of nodes in the mesh is 690,400 (the code CFX is node based) which corresponds to 1,027,267 elements (tetrahedra, pyramids and hexahedra). The leak pipe of 4 mm diameter is resolved by 33 hexahedral cells over the cross section.
3.2.
Code set-up
Ansys CFX version 10 [17] was applied for all simulations. The following code options were selected except otherwise stated. The selection of options was based on the experience gained by the authors during the validation work presented in [20–22]. The total energy model of CFX was applied, because it includes the kinetic energy contributions which are extremely important in high speed gas flows. The SST (k–u` shear stress transport, see ref. [17]) turbulence model with standard parameters was selected because of the super-sonic character of the flow. For the transient term in the transport equations the Second Order Backward Euler discretization scheme was used. The advection terms in the discrete conservation equations employed the High Resolution scheme of CFX. This numerical scheme includes a blending factor to the First Order Advection scheme according to the local variable gradients. The multi-component gas mixture in the laboratory was modelled with hydrogen, oxygen, nitrogen individually. The ideal gas law was applied. It is known that a real gas law could improve the accuracy of the release simulations from high pressure systems [23,24]. The larger the pressure ratio is, the larger is the difference for the gas density that is calculated with the ideal gas law compared with the real gas law. For a storage pressure of 700 bar, the hydrogen density is over-estimated by 45% using the ideal gas law [23]. For the initial storage pressure of 200 bar, the over-estimation is less than 10% and for the initial pressure (about 100 bar) at the leakage position the over-estimation is less than 2% [23]. Moreover the high density hydrogen flow predicted by the ideal gas model produces a larger flow rate throughout the whole release process and therefore the ideal gas law represents a conservative approximation [24]. It should be also emphasized that in the simulation the total hydrogen mass that is released into the laboratory was kept the same (1.2 kg). Therefore using the ideal gas law was considered an acceptable approximation. The measured flows at inlets and outlets of the laboratory ventilation system were used as boundary conditions and were kept constant during the simulations. The initial thermodynamic state was assumed to be 293 K (20 C) and 0.1 MPa (1 bar) in the laboratory and 20 MPa in the storage/bottle at the same temperature. Heat transfer to walls at constant wall temperatures was included in all simulations. This choice is very important especially in the regions of the jet expansion where very low gas temperatures exist.
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4.
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Simulation results
The calculations presented require a certain amount of computer resources. The initial state in the laboratory was obtained by calculation of the flow field imposed by the forced ventilation. This run was performed within about 4 h wall clock time on eight processors of an AMD Opteron (1.8 GHz) based Linux cluster machine. The subsequent leak simulation could be finished within about 25 days based on the same hardware. The numerical capacity of recent computers develops quickly. With the latest processors available today computing times might be less than half of the times indicated.
4.1.
Fig. 3 – Examples of the flow field in the lab under normal working conditions.
3.3.
Ventilated flow in the laboratory
The initial state in the laboratory before the hydrogen discharge occurs is determined by the forced ventilation system. Inflow and outflow velocities through the ventilation openings in the laboratory were measured. These data were provided to CFX as boundary condition in the calculation of the spatial flow field. Vertical cuts through this converged flow field at different positions are depicted in Fig. 3. It can be seen that higher velocities are found around the inlet and outlet openings of the ventilation. These openings are distributed along the ceiling and the side wall where the pipe rupture is assumed to occur. In some regions of the laboratory there are almost stagnant flow conditions with flow speeds of 0.1 m/s or below.
Constant flow rate: the ‘‘bottle-compressor’’ case
The assumption of a constant sonic flow through the leak corresponds to the ‘‘compressor-bottle’’ configuration, which makes use of an idealised compressor able to maintain a constant pressure to the hydrogen distribution line. A pressure of 10 MPa was assumed at the leak position together with the critical speed of hydrogen in air. The flow was stopped when the total available hydrogen mass was released. The hydrogen flow history was specified as a boundary condition in the code. Fig. 4 shows the resulting hydrogen mass flow rate in relation to the accumulated mass released. The total release time amounts to about 12 s. A main result expected from the simulations is the calculation of the size of flammable hydrogen clouds in the laboratory. The lowest hydrogen concentration at which ignition is possible (LFL) is about 4% in air. Fig. 5 illustrates the extension of these clouds during two different stages of the accident by an iso-surface of hydrogen concentration equal to 4%. The walls of the laboratory are coloured by local hydrogen concentrations according to the colour key given in the figure. In the early stage of the release hydrogen is found mainly
Fig. 4 – ‘‘Bottle-compressor’’ operative mode: Hydrogen flow through leak and total mass released.
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Fig. 5 – ‘‘Bottle-compressor’’ operative mode: Iso-surfaces of hydrogen concentration at the lower flammability limit (LFL – 4%) during the hydrogen release (left-hand side, approximately 4 s from release start) and later after completion of the release in the distribution phase (right-hand side, approximately 20 s after start of transient).
along the jet (left-hand side in Fig. 5). After the end of the release phase hydrogen is spread over most parts of the laboratory and is then slowly removed by the ventilation system. The maximum cloud size with hydrogen concentrations of 10% (representing the lower limit for substantial flame acceleration) or higher is found to be about 64 m3. Results are discussed quantitatively and in more detail in comparison with the simulations of the following section in paragraph 5.
4.2.
Variable flow rate: the ‘‘bottle-only’’ case
The other operative mode of the laboratory is the ‘‘bottleonly’’ mode, consisting in the direct connection of the
hydrogen gas bottle to the hydrogen distribution line without the hydrogen compressor. Also for this operation mode the simulations start from the converged ventilation run as initial state. The results shown were obtained with the finest mesh available. Investigations on the influence of the mesh resolution are discussed separately (Section 4.3). Due to the limited size of the storage its internal pressure drops quickly and both the density and the mass flow through the leak decrease with increasing time. The calculated mass flow rate from the leak is shown in Fig. 6. For a very limited early time the leak flow is almost constant but reduces later continuously until the pressures in the hydrogen storage and in the laboratory equalise. The release process is finished after about 90 s.
Fig. 6 – ‘‘Bottle-only’’ operative mode: Mass flow rate through the leak (circle points to the very early discharge phase shown in the second graph).
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Fig. 7 – ‘‘Bottle-only’’ operative mode: Iso-surface of hydrogen concentration at the lower flammability limit (LFL, 4%) in the lab during the hydrogen release (left-hand side, approximately 4 s after beginning) and later in the distribution phase (righthand side, approximately 20 s after start of transient).
Considering the size of the flammable volume (4% limit) Fig. 7 shows similar results compared with the first approach illustrated in Fig. 5. However for the current operation mode the flammable cloud remains smaller because the outflow of the same hydrogen mass is distributed over a longer time. The 4%-flammable cloud volume reaches a maximum of about 132 m3 whereas the 10% volume is nearly negligible and is less than 1 m3. With these results illustrated in Fig. 8 it becomes evident that flammable conditions exist only for a limited time of about 100 s. The chance of a fast deflagration is small due to the low hydrogen concentrations in the cloud. In the laboratory hydrogen sensors are installed at selected positions in order to detect hydrogen releases as early as possible. Two sensors are located under the hoods RIG and AMC (Fig. 2). Their approximate positions are shown in Fig. 9. The hydrogen mol fraction at these locations passes the alert level of 0.4% very early in the transient (less than 1 s). This would enable any kind of counter measures to mitigate the transient. However, for the purpose of studying the complete release evolution, as explained earlier, mitigation measures were not taken into account in this work.
The simulation of the hydrogen leak in the laboratory must also include a realistic simulation of the jet flow itself from initially 20 MPa in the storage to atmospheric pressure. Although not the main focus of this work the numerical grid was set-up sufficiently fine that the microstructures of the jet become visible. The pressure ratio between hydrogen storage and laboratory is much larger than the critical pressure ratio of hydrogen (about 1.9); therefore an under-expanded jet develops. Further, the pressure at the pipe exit is calculated to be much higher (with full bottle about 85 bar) than the environmental pressure which leads to a highly under-expanded jet. The maximum Mach number in the jet calculated is about 6.7 and the maximum gas speed reaches a value of approximately 2500 m/s. Due to the strong under-expansion of the jet at the position of the maximum Mach number the temperature drops to a very low value. This value is related to a local pressure of 4100 Pa. The under-expansion region is followed by the Mach disk which was found to be located at about 8 diameters downstream the release pipe with 4 mm diameter. There are many investigations on highly under-expanded jets available in the literature. Two examples are [25,26].
Fig. 8 – Size of the flammable volume with concentrations of 4% (left-hand side) and 10% (right-hand side) (‘‘bottle-only’’ operative mode).
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Fig. 9 – Time dependent mole fraction at two hydrogen sensor locations in the lab (‘‘bottle-only’’ operative mode).
Ref. [25] describes a numerical study on a critical hydrogen jet and ref. [26] discusses jet correlations derived from experimental data. The values calculated in this work are slightly different from the ones reported in [25]. In ref. [25] a maximum Mach number of nearly 9 is predicted. The Mach disk following the peak Mach number is found at 9.2 diameters. A detailed review of experimental investigations and relevant correlations on the position and size of the Mach disk is found in ref. [26]. The correlations of Ashkenas and Sherman [27] and the experimental data of Love [28] are used by Lehnasch [26] to calculate the size of the Mach disk that is in a range of 6–8 pipe diameters for the given pressure ratio. Those values are in good agreement with the own simulation results. It can be stated that the main microstructures of the underexpanded critical jet were calculated with acceptable accuracy in the simulations. This supports the credibility of the predictions presented in this work.
4.3.
Sensitivity studies
The following studies were carried out for the ‘‘bottle-only’’ case which includes the hydrogen storage in the mesh.
4.3.1.
Mesh resolution
A proper spatial resolution of the volume around the sonic jet is important for a realistic simulation of the hydrogen dispersion in the whole laboratory. Therefore, the mesh resolution in the leak area was consecutively increased in order to identify an influence on the predictions of flammable mass and volume. Three different computational meshes have been generated: a coarse, a medium and a fine mesh. Two of the three meshes prepared for the simulations use local grid refinement by an internal conic volume in the jet region which gives good control over the local cell size. The cone starts always at the leak opening and extends far enough downstream to include the full jet. It consists mostly of unstructured cells and is linked to the hexahedral mesh of the
lab. A summary of major grid parameters of the three meshes used is given in Table 1. The refinement of mesh structure in the jet region influences the relevant results in a significant manner only in the area close to the jet. In Fig. 10 Mach number profiles along the jet axis are displayed. With increasing spatial resolution the maximum Mach number at a distance of 10 mm downstream of the pipe exit increases considerably. Directly at the leak exit (before the final expansion downstream of the leak opening occurs) however, no influence is seen (black lines in Fig. 10). If global results such as the flammable mass and flammable volume history are considered no remarkable changes are observed when applying the different meshes. It can be concluded that the numerical predictions of the relevant parameters for hydrogen risk assessment in the laboratory were obtained with sufficient spatial resolution because further refinement of the grid does not change these results. This is also confirmed by the visualisation of the flammable cloud (LFL, 4%) in size and shape for the two finer meshes applied. The differences in the hydrogen distribution patterns are small as shown in Fig. 11.
4.3.2.
Turbulence model
Alternatively to the SST (Shear Stress Transport) turbulence model the RNG (Renormalisation Group) k–3 model was applied to simulate the hydrogen leak. Inside the fluid domain far away from boundaries the SST model behaves like
Table 1 – Main parameters of the three meshes applied. Average equivalent Number Total cell length in leak of elements number of region [mm] in leak cone elements Coarse mesh Medium mesh Fine mesh
38 26.2 14.4
– 9302 72 337
897 778 918 681 1 005 025
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by the RNG k–3 model has an about 20% higher peak value (about 180 m3) compared to the SST model simulation.
5.
Fig. 10 – Influence of local mesh resolution on calculated Mach numbers in the jet.
a standard k–3 model whereas the RNG k–3 model includes an extra term in the dissipation equation and uses different model constants. The RNG k–3 model provided good results for a sub-sonic jet analysed in a recent work [22]. Therefore this model was selected to compare to the SST model simulations. In the current work both turbulence models required similar computing times during the simulations but the SST model showed better convergence behaviour. Results were compared for two selected parameters. For the Mach number as a local parameter no influence of the turbulence model at a monitor point position at the centreline of the jet and 0.01 m downstream of the leak opening was seen. The second parameter to be compared was the flammable volume (4–75% hydrogen). This is an integrated parameter. It is influenced by the choice of the turbulence model. The jet simulated by the RNG k–3 model had a tendency to be more compact with less entrainment of air than the one calculated by the SST model. Therefore the flammable volume calculated
Discussion
The two principal modelling strategies followed in this work end up with quite different results. As shown in before, the mass flows from the leak are very different. In the ‘‘bottlecompressor’’ operative case a constant leak flow (Fig. 4) was assumed. With the second option ‘‘bottle-only’’ the leak flow was calculated as a high pressure discharge process from the hydrogen storage (Fig. 6). The decreasing hydrogen pressure and density in the storage reduces the mass flow through the leak and spreads the release over a much longer time compared with the first modelling option. In Fig. 12 the resulting flammable masses are compared for both cases. The longer outflow time in the case of the hydrogen storage/bottle in the calculation leads to a considerably smaller amount of mass which might be ignited (LFL). Further, the simplified modelling option with constant leak flow can be considered as a conservative approach in terms of the hydrogen risk as it results in a much larger flammable mass. This option is much less numerically demanding but considerably overestimates the amount of ignitable hydrogen mixture. Iso-surfaces of LFL for both modelling options can be compared in Figs. 5 and 7. The flammable clouds look very similar during the release. With both simulation options detached regions from the flammable cloud exist only at a late time after the end of the release. Then some isolated hydrogen–air pockets exist for a short time but these do not contribute to the flammability in the lab. The modelling results obtained are stable. They do not depend on the mesh resolution. The choice of the turbulence model has some influence on the parameters of interest like size of flammable volume. Within the scope of this work assumptions had to be made on the accident scenario and certain modelling options (e.g. real versus ideal gas). The
Fig. 11 – Hydrogen distribution in the lab at the moment of maximum cloud size for medium (left-hand side) and fine mesh (right-hand side) (‘‘bottle-only’’ operative mode).
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Fig. 12 – Flammable mass in the lab for both modelling options in comparison to the total mass released through the leak.
choice of the turbulence model is another contribution to a degree of uncertainty of the results presented.
6.
Conclusions
A conservative accident scenario was defined to evaluate a worst case situation in a hydrogen laboratory. On this basis simulations of the accidental release of hydrogen were carried out. The purpose of the work was to assess the formation of hydrogen clouds inside the laboratory. For this reason the integrated volume of regions which fulfil the ignition criteria of hydrogen were calculated during and after the hydrogen release. Two principal modelling options were followed corresponding to two typical operative modes of the laboratory. In the first approach a compressor is assumed to maintain the pressure in the hydrogen distribution line at a fixed pressure and the predefined critical mass flow across the leak is constant. In the second approach, the compressor is not included and the discharge occurs through pipes from a high pressure bottle/storage. Care was taken to resolve also the main microstructures of the leak jet by local mesh refinements. As a major result of ‘‘bottle-only’’ simulations it was calculated that flammability at the lower limit of 4% exists over approximately 100 s in a maximum volume of 130 m3. This volume was calculated with the SST turbulence model. The 4% flammability limit enables ignitability but flame propagation is only possible in upright vertical direction. The limit of 10% which represents the lower value of substantial flame acceleration was only passed in less than 1 m3 (corresponding to 9 g hydrogen). Additional simulations were performed with different mesh resolution and with an alternative turbulence model. The mesh investigations did not show a dependency of the major results concerning the hydrogen distribution. The choice of the turbulence model however has consequences in the calculation of the hydrogen release jet. The alternative RNG k–e model predicted a higher value of the
peak flammable volume at 4% hydrogen concentration. Combustion at this lower limit is very weak and therefore the increase in the uncertainty of the results limited. In terms of laboratory safety, the operative solution without the compressor is preferable, as it is also expected from an intuitive point of view, because it releases the total available hydrogen mass in a longer period, allowing the ventilation system to be more effective and reducing the maximum value of the flammable hydrogen mass/volume. To study the full evolution of hydrogen in the laboratory this work has disregarded the automatic safety system, e.g. shut-off valves. Experimental tests have shown that in reality the depressurisation starts approximately 2–3 s after that the hydrogen concentration at a sensor passes the 10% of the LFL value (0.4% of hydrogen in air). This will reduce the fraction of the hydrogen mass released into the laboratory (Fig. 12). In the cases studied it is finally interesting to note that the first sensor experiencing an increase of hydrogen concentration is the sensor positioned near the wall opposite to the leak position and not that positioned precisely above it. This is due to the behaviour of the hydrogen jet during the first release phase. The result demonstrates the importance to locate sensors in an optimal way. In the case of the laboratory under study typical accidental releases consist in ‘‘small’’ leaks, where buoyancy dominates from the very start of release. The sensors have been located ‘‘intuitively’’ on top and in front of each high pressure equipment and have demonstrated their usefulness. But in cases where hydrogen jets could be expected, an intuitive approach will not be enough and CFD tools should be used intensively in conjunction with risk assessment techniques to calculate locations with the most probable ‘‘early hydrogen increase’’. A possible extension of the presented work is the inclusion of real gas behaviour and the simulation of a hydrogen deflagration across the flammable cloud. Property data of the modelled gas components in a wide range of pressure and temperature are needed to carry out the super-critical jet expansion. Improvements on the property data may also lead to a better quality of the simulation results.
international journal of hydrogen energy 35 (2010) 4409–4419
references
[1] Directive 97/23/EC of the European Parliament and of the Council of 29 May 1997 on the approximation of the laws of the Member States concerning pressure equipment OJ L 181, 09/07/1997. [2] Directive 94/9/EC of the European Parliament and the Council of 23 March 1994 on the approximation of the laws of the Member States concerning equipment and protective systems intended for use in potentially explosive atmospheres (OJ L 100, 19.4.1994). [3] Breitung W, Necker G, Kaup B, Veser A. Numerical simulation of hydrogen release in a private garage. In: Proceedings of hypothesis IV, 9–14 September 2001, Strahlsund, Germany. pp. 368. [4] Papanikolaou EA, Venetsanos AG. ‘‘CFD modelling for helium releases in a private garage without forced ventilation’’. In: first International conference on hydrogen safety, Pisa, September; 2005. [5] Barley CD, Gawlik K. Buoyancy-driven ventilation of hydrogen from buildings: laboratory test and model validation. International Journal of Hydrogen Energy 2009;34: 5592–603. [6] Venetsanos AG, Papanikolaou E, Delichatsios M, Garcia J, Hansen OR, Heitsch M, et al. An inter-comparison exercise on the capabilities of CFD models to predict the short and long term distribution and mixing of hydrogen in a garage. International Journal of Hydrogen Energy 2009;34: 5912–23. [7] Gallego E, Migoya E, Martı´n-Valdepen˜as JM, Crespo A, Garcı´a J, Venetsanos A, et al. An intercomparison exercise on the capabilities of CFD models to predict distribution and mixing of H2 in a closed vessel. International Journal of Hydrogen Energy 2007;32:2235–45. [8] Breitung W, Bielert U, Necker G, Veser A, Wetzel FJ, Pehr K. Numerical simulation and safety evaluation of tunnel accidents with a hydrogen powered vehicle. In: 13th World hydrogen energy conference, Beijing, China; 12–15 June; 2000. [9] Mukai S, Suzuki J, Mitsuishi H, Oyakawa K, Watanabe S. CFD simulation on diffusion of leaked hydrogen caused by vehicle accident in tunnels. In: first International conference on hydrogen safety, Pisa, Italy; 8–10 September; 2005. [10] Venetsanos AG, Baraldi D, Adams P, Heggem PS, Wilkening H. CFD modelling of hydrogen release, dispersion and combustion for automotive scenarios. Journal of Loss Prevention in the Process Industries 2008;21:162–84. [11] Molkov V, Verbecke F, Makarov D. LES of hydrogen–air deflagrations in a 78.5 m tunnel. Combustion Science and Technology 2008;180:1–13. [12] Baraldi D, Kotchourko A, Lelyakin A, Yanez J, Middha P, Hansen OR, et al. An inter-comparison exercise on CFD model capabilities to simulate hydrogen deflagrations in a tunnel. International Journal of Hydrogen Energy 2009;34:7862–72.
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[13] Middha P, Hansen OR. CFD simulation study to investigate the risk from hydrogen vehicles in tunnels. International Journal of Hydrogen Energy 2009;34:5875–86. [14] Baraldi D, Venetsanos AG, Papanikolaou E, Heitsch M, Dallas V. Numerical analysis of release, dispersion and combustion of liquid hydrogen in a mock-up hydrogen refuelling station. Journal of Loss Prevention in the Process Industries 2009;22:303–15. [15] Makarov D, Verbecke F, Molkov V, Roe O, Skotenne M, Kotchourko A, et al. An inter-comparison exercise on CFD model capabilities to predict a hydrogen explosion in a simulated vehicle refuelling environment. International Journal of Hydrogen Energy 2009;34:2800–14. [16] Heitsch M, Baraldi D, Moretto P, Wilkening H. Safety of laboratories for new hydrogen techniques. In: International conference on hydrogen safety, San Sebastian; 2007. [17] CFX-10, Documentation. Southpointe, 275 Technology Drive, Canonsburg, PA 15317, USA: ANSYS Inc., http://www.ansys. com; 2006. [18] Birch AD, Brown DR, Dodson MG, Swaffield F. The structure and concentration decay of high pressure jets of natural gas. Combustion Science and Technology 1984;36:249–61. [19] Xiao J, Travis JR, Breitung W. Hydrogen release from a highpressureGH2 reservoir in case of a small leak. 3. In: International conference on hydrogen safety, Ajaccio; 2009. [20] Heitsch M, Wilkening H, Baraldi D. CFD simulations of light gas release and mixing in the Battelle model-containment with CFX. Nuclear Engineering and Design 2008;238:618–26. [21] Wilkening H, Baraldi D, Heitsch M. On the importance of validation when using commercial CFD codes in nuclear reactor safety. CFD4NRS OECD/NEA & IAEA Workshop. Garching; 5–7 Sept 2006. [22] Heitsch M, Baraldi D, Wilkening H. Validation of CFD for containment jet flows including condensation. XCFD4NRS OECD/NEA & IAEA Workshop. Grenoble; 10–12 Sept 2008. [23] Tchouvelev AV, Cheng Z, Agranat VM, Zhubrin SV. Effectiveness of small barriers as means to reduce clearance distances. International Journal of Hydrogen Energy 2007;32: 1409–15. [24] Mohamed k, Paraschivoiu M. Real gas simulation of hydrogen release from a high pressure chamber. International Journal of Hydrogen Energy 2005;30:903–12. [25] Xu BP, Zhang JP, Wen JX, Dembele S, Karwatzki J. Numerical study of a highly under-expanded hydrogen jet. In: International conference on hydrogen safety, Pisa, Italy; 8–10 September 2005. [26] Lehnasch G. Contribution a` l’E´tude Nume´rique des Jets Supersoniques Sous-De´tendus. Thesis. E´cole Nationale Supe´rieure de Me´chanique et d’Ae´ronautique; 2005. [27] Ashkenas H, Sherman FS. The structure and utilization of supersonic free jets in low density wind tunnel. In: Advances in applied mechanics – rarefied gas dynamics. Academic, New York; 1966. p. 84–105. [28] Love ES, Grigsby CE, Lee LP. Experimental and theoretical studies of axisymmetric free jets. Technical report technical report R-6. Langley Research Center, Langley Field, Va; 1959.
Available at www.sciencedirect.com
journal homepage: www.elsevier.com/locate/he
Numerical analysis of accidental hydrogen release in a laboratory M. Heitsch*, D. Baraldi, P. Moretto JRC, Institute for Energy, Pb 2, NL-1755 ZG Petten, The Netherlands
article info
abstract
Article history:
In a hydrogen laboratory hydrogen may accidentally leak from the high pressure systems
Received 22 August 2009
and it may form a flammable mixture after mixing with air. CFD (Computational Fluid
Received in revised form
Dynamics) simulations of accidental hydrogen releases in a laboratory are performed in
14 January 2010
order to estimate the size and the resident time of the flammable cloud inside the labo-
Accepted 14 January 2010
ratory. Moreover the effectiveness of the hydrogen sensor positions in detecting the
Available online 20 March 2010
presence of the hydrogen in the laboratory within few seconds has also been investigated. In a first scenario, the hydrogen coming from the storage is pumped by a compressor into
Keywords:
the laboratory pipelines at constant pressure. To be conservative, it is assumed that the
Hydrogen
pressure can be maintained when the leak in the pipeline occurs. In a second scenario, the
Safety
hydrogen flows directly from the storage into the laboratory pipeline without the
Jet release
compressor, and the pressure in this case is decreasing with time. In both cases, shut-off
Critical flow
valves have been neglected in order to consider a worst case scenario. The simulation
Flammable cloud
results from the two accident scenarios are compared, showing that the configuration with the compressor will potentially produce larger amounts of flammable mixture in the laboratory. ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved.
1.
Introduction
In the last years international and national research programmes have invested large amounts of resources in extensive experimental research aiming at the development of technologies for hydrogen production, storage, transport and conversion. To this purpose laboratories and facilities have been newly constructed or converted with an increased local hydrogen storage and usage quantities. To guarantee safe operations various regulations and guidelines are available to designers and operators, such as the Pressurised Equipment Directive [1] and the ATEX directive [2] in Europe. To model accidental situations and eventually assist in a safety design, various analysis methods are advised by the guidelines, such as the Failure Mode and Effects Analysis (FMEA), the Hazard and Operability Studies (HAZOP), and the
Failure Mode and Effects Criticality Analysis (FMECA). As effective tools to be used to evaluate in details specific accidental scenarios, Computational Fluid Dynamics (CFD) analysis can be performed and the use of CFD is also advised by the guidelines. CFD codes have been extensively used to simulated hydrogen accident scenarios in garages [3–7], in tunnels [8–13] and in re-fuelling stations [14,15]. The authors performed CFD investigations of accidental release scenarios in a relatively small laboratory (approximately 1.2 kg of hydrogen stored) for hydrogen storage technologies. The safety of the laboratory was designed by means of the FMEA method and it respects the above mentioned directives. By performing the CFD studies, various goals were pursued: the evaluation of the laboratory general safety level with a more sophisticated 3-D method, the correctness of the safety design assumption and more specifically the
* Corresponding author. Tel.: þ31(0)2245140; fax: þ31(0)224565623. E-mail address: [email protected] (M. Heitsch). 0360-3199/$ – see front matter ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2010.01.044
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optimisation of the hydrogen detection system used for an early alarm and automatic operative shut down. To this purpose various release scenarios have been adopted and by CFD simulation flow rates and volumes of flammability were calculated and local hydrogen concentration increases were identified. We report here only on the worst case scenario, beyond the maximum credible accident which was considered during the risk assessment and the safety design. This scenario corresponds to the full and instantaneous break of a high pressure hydrogen distribution pipe. The authors believe that this publication could stimulate the adoption of 3dimensional simulation tools for the optimisation of safety design also for less critical cases, such as small hydrogen laboratories and equipments. The simulations were carried out with the CFD code CFX. Different modelling options are investigated and results in terms of flammable mass and volume in the laboratory as a consequence of the hydrogen releases are given. The work presented here is an extension of the simulations presented earlier in [16].
2.
Problem description
The laboratory under study is a typical installation for the investigation of hydrogen storage capacity in materials. It is equipped with a hydrogen supply and a distribution system which delivers hydrogen at 20 MPa maximum to several working places where the experimental equipment is installed. The laboratory safety has been designed by performing an FMEA (Failure Mode and Effects Analysis) and determining the maximal credible accident. In order to validate the assumption, various leakage scenarios at different locations have been simulated by numerical simulations. We report here on the simulation results of a release case which has not been considered during safety design because beyond the maximum credible accident; the instantaneous rupture of a hydrogen distribution pipe at 20 MPa. The numerical simulations were carried out by means of the CFD (Computational Fluid Dynamics) code Ansys CFX [17]. Gas distribution in the laboratory under specific boundary conditions was calculated at first, followed by the calculation of the size and the duration of a flammable hydrogen cloud in the laboratory. Finally, the effectiveness of the positions of existing hydrogen sensors in the laboratory can be checked. The main micro processes in the release jet of the leak were resolved to a certain degree but have not been the main focus of this work. The resolution of the jet structures is necessary for the correct prediction of the hydrogen dispersion in the far-field of the jet.
2.1.
a volume of approximately 285 m3. A photo of a part of the laboratory is given in Fig. 1. Hydrogen is delivered at 20 MPa at maximum from a 50 l gas bottle in combination with a compressor. Bottle and compressor are located outside the laboratory. Two hydrogen distribution operative configurations are available: the use of a compressor allows to maintain the hydrogen delivery to the laboratory at constant pressure, while the direct connection of the hydrogen gas bottle to the distribution line causes a continuous decrease of the pressure due to hydrogen consumption. However, it simplifies the system also from the safety control point of view. In the following the two configurations will be referred to as ‘‘bottle-compressor’’ and as ‘‘bottle-only’’. The maximum amount of hydrogen which would be released in the laboratory in case of failure of all automatic and manual shut down procedures can be calculated by considering a full bottle, a full compressor buffer tank, all the pipes at 20 MPa and all the experimental equipments at full storage capacity. That amount is about 1.2 kg. A distributed ventilation system provides the laboratory with a constant flow of air. As shown in Fig. 2, in the ventilation system there are 8 inlet openings located on the ceiling, and 10 extraction openings, 6 of which are placed on the ceiling and 4 on a wall, at the floor level (in Fig. 2 the inlets are shown in blue while the exhaust openings are marked in red and turquoise). In addition, air extraction fume hoods are located on top of each test stand in order to collect hydrogen possibly released during experiments and discharge it to the roof of the building (names are RIG, IVR, AMC; see Fig. 2). Two independent leak detection systems are available: an overflow valve in the main hydrogen supply line, and safety hydrogen detectors under each fume hood. When the flow rate in the hydrogen supply system or local hydrogen concentrations in the lab exceed certain threshold values, an
The laboratory
The laboratory under study is a typical facility for the investigation of hydrogen storage capacity in materials located at the Institute for Energy of the Joint Research Centre (JRC). It consists of sorption equipments based on gravimetric, volumetric and spectrometric principles. They are performed typically in a pressure range of 0.1 up to a maximum of 20 MPa and with temperature from 77 K up to 273 K. The laboratory is approximately 18 m long, 8.5 m wide and 3.5 m high, with
Fig. 1 – Picture of a part of the laboratory: Vin indicates the positions of the ventilation inlets, Vout the positions (on the ceiling and near the ground) of the ventilation outlets, H two of the fume hoods positioned on top of the test rigs. Number 2 indicates the position of the hydrogen leak that is investigated in the simulations.
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Fig. 2 – Ventilation system of the laboratory (inlets – blue; outlets – red, turquoise). The black arrow indicates the viewing direction of the observer in the photo of Fig. 1. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
automatic shut down procedure starts within less of a second. In that case the hydrogen supply is interrupted and the distribution system is depressurized. Any detector triggers the shut down procedure when the volumetric hydrogen concentration passes 0.4%. This corresponds to one-tenth of the lower flammability limit (LFL) of hydrogen in air which is about 4%. Local air velocities were measured by means of a calibrated anemometer at all openings of the ventilation system and then averaged in order to obtain a single value per opening. The velocities range is between 1 and 2 m/s, and the total airflow entering the laboratory sums up to approximately 15000 m3/h. The same procedure was applied to the air extraction pipes under the fume hoods, where the air velocity range is between 1.5 and 5.5 m/s. These data have been used in the CFD calculations as boundary conditions.
2.2.
Accident scenario
The selected accident scenario consists in the instantaneous rupture of a hydrogen pipe with an internal diameter of 4 mm. In order to study the release of the full hydrogen content in the laboratory, it has been further assumed that all the automatic detection and shut down systems do not operate. The only available operative system is the ventilation system of the laboratory. The hydrogen system is under full pressure (20 MPa) and therefore the total available amount of hydrogen (1.2 kg) in the bottle and in the pipe system is released into the room. A realistic location of a leak is where pipes are connected. There are several pipe connections in the laboratory. However, the pipe joints located underneath the fume hood RIG (Fig. 2) is most often subject to human intervention. It is therefore considered to be the location where the simulated pipe rupture takes place. The full cross section of the pipe with 4 mm diameter serves as release opening. Further, a horizontal blowing direction is assumed towards to the opposite wall of the laboratory (direction to hood AMC). The accident can be subdivided into two phases. The first phase is dominated by the critical release of hydrogen into the laboratory. The ventilation is active but is only able to influence the gas distribution at a distance from the jet. The flow
field created by the jet is dominating. The second phase starts when the pressure in the hydrogen storage equals to the environmental pressure and the hydrogen outflow stops. Then the forced ventilation system becomes able to remove more effectively the accumulated hydrogen from the laboratory. Both phases are covered by the simulations.
3.
Simulation preparation
Prior to the simulation of the hydrogen leak, the flow field in the laboratory due to the ventilation flows is calculated and used as initial state for the leak simulations. Two principal model options concerning the hydrogen leak were applied. The first option assumes a constant sonic flow through the leak until the total amount of hydrogen from the hydrogen storage outside the laboratory has been discharged. It corresponds to the operative case ‘‘bottle-compressor’’ described in 2.1. This operative mode assumes that the hydrogen compressor is able to maintain a constant maximum pressure of 20 MPa in the hydrogen distribution lines. The assumption of a constant pressure during the whole discharge is a simplification of the real situation, because the compressor is not expected to be able to guarantee the requested constant pressure especially towards the end of the transient. However, it simplifies considerably the calculation and it represents a conservative approach from safety point of view, because it assumes a quicker release of the available gas than in the real case. The second option includes the storage by its volume and simulates the hydrogen leak as an expansion process from the storage starting from an initial internal pressure of 20 MPa. This modelling option considers the relation between pressure and mass in the storage and calculates consequently a time dependent leak flow. It corresponds to the ‘‘bottleonly’’ operative case described in 2.1, where the hydrogen compressor is bypassed and a gas bottle at 20 MPa is directly connected to the hydrogen distribution line. The second option could also use a simplifying approach to calculate the leak flow into the laboratory. In literature several algorithms are proposed, for example by Birch et al. [18] and
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by Xiao et al. [19]. These algorithms avoid the full calculation of an under-expanded jet with very small cells close to the leak and short time steps by providing an algorithm which allows starting the calculation downstream of the physical leak. The algorithm preserves the main parameters of the jet flow, simplifies the calculation and saves computing time. However, deficiencies are observed in the turbulence level and the actual temperature at the position where the simplified calculation starts (fictitious diameter). Gas entrainment upstream of the position of the fictitious diameter cannot be dealt with. Further, in the context of the laboratory connected to a relatively small tank (transient behaviour), an extra tool would be needed to calculate the leak flow rate depending on the pressure in the hydrogen tank. In order to avoid the need of an extra code together with additional assumptions on flow parameters at the position of the fictitious diameter it was decided to calculate the leak flow without a simplifying algorithm. This is accomplished at the expense of a higher computational effort and demonstrates at the same time that a full simulation as part of a long transient is possible with average computer resources. The two modelling options concerning the hydrogen leak and the results obtained are discussed in detail in the following sections.
3.1.
Computer model of the laboratory
The numerical model of the laboratory from Fig. 1 contains only the medium and large size components. All small size devices like computers, analytical equipment, cables and pipes were neglected because it is assumed that they cannot influence the global hydrogen distribution. All venting openings, fixed side boards, the fume hoods and two larger devices under the fume hoods RIG and IVR are included in the computational model. Some of the ventilation inlets (blue colour in Fig. 2) release air in two directions some only in one direction. The ventilation outlets in the model were not located directly in the walls but were placed outside of the laboratory to allow the flow to develop in such a way that the pressure across the openings becomes uniform. Fig. 2 indicates the approximate location of the assumed leak under hood RIG. Two principal modelling options for the leak flow are followed in this paper. One option predefines the mass flow through the leak and only a short pipe to introduce the flow is included in the numerical grid. The other option calculates the mass flow from the actual mass, temperature and pressure in the hydrogen storage against the conditions in the laboratory. The size of the hydrogen storage was calculated to contain the total hydrogen mass available in the whole hydrogen system. This is a conservative approach because it neglects the possibility of some sections of the system to be isolated which therefore would not contribute to the hydrogen release. For this option the storage and the leaking pipe are included in the numerical grid. The mesh for the numerical simulations is composed of hexahedral cells in most of the model. Only close to the roof and downstream the leak an unstructured mesh composed of tetrahedral and pyramidal cells was employed. The use of a hybrid mesh with structured and unstructured cells offers
the flexibility needed for inserting very small entities (leak pipe) into a given hexahedral mesh and keeps the number of cells within manageable limits. The total number of nodes in the mesh is 690,400 (the code CFX is node based) which corresponds to 1,027,267 elements (tetrahedra, pyramids and hexahedra). The leak pipe of 4 mm diameter is resolved by 33 hexahedral cells over the cross section.
3.2.
Code set-up
Ansys CFX version 10 [17] was applied for all simulations. The following code options were selected except otherwise stated. The selection of options was based on the experience gained by the authors during the validation work presented in [20–22]. The total energy model of CFX was applied, because it includes the kinetic energy contributions which are extremely important in high speed gas flows. The SST (k–u` shear stress transport, see ref. [17]) turbulence model with standard parameters was selected because of the super-sonic character of the flow. For the transient term in the transport equations the Second Order Backward Euler discretization scheme was used. The advection terms in the discrete conservation equations employed the High Resolution scheme of CFX. This numerical scheme includes a blending factor to the First Order Advection scheme according to the local variable gradients. The multi-component gas mixture in the laboratory was modelled with hydrogen, oxygen, nitrogen individually. The ideal gas law was applied. It is known that a real gas law could improve the accuracy of the release simulations from high pressure systems [23,24]. The larger the pressure ratio is, the larger is the difference for the gas density that is calculated with the ideal gas law compared with the real gas law. For a storage pressure of 700 bar, the hydrogen density is over-estimated by 45% using the ideal gas law [23]. For the initial storage pressure of 200 bar, the over-estimation is less than 10% and for the initial pressure (about 100 bar) at the leakage position the over-estimation is less than 2% [23]. Moreover the high density hydrogen flow predicted by the ideal gas model produces a larger flow rate throughout the whole release process and therefore the ideal gas law represents a conservative approximation [24]. It should be also emphasized that in the simulation the total hydrogen mass that is released into the laboratory was kept the same (1.2 kg). Therefore using the ideal gas law was considered an acceptable approximation. The measured flows at inlets and outlets of the laboratory ventilation system were used as boundary conditions and were kept constant during the simulations. The initial thermodynamic state was assumed to be 293 K (20 C) and 0.1 MPa (1 bar) in the laboratory and 20 MPa in the storage/bottle at the same temperature. Heat transfer to walls at constant wall temperatures was included in all simulations. This choice is very important especially in the regions of the jet expansion where very low gas temperatures exist.
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4.
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Simulation results
The calculations presented require a certain amount of computer resources. The initial state in the laboratory was obtained by calculation of the flow field imposed by the forced ventilation. This run was performed within about 4 h wall clock time on eight processors of an AMD Opteron (1.8 GHz) based Linux cluster machine. The subsequent leak simulation could be finished within about 25 days based on the same hardware. The numerical capacity of recent computers develops quickly. With the latest processors available today computing times might be less than half of the times indicated.
4.1.
Fig. 3 – Examples of the flow field in the lab under normal working conditions.
3.3.
Ventilated flow in the laboratory
The initial state in the laboratory before the hydrogen discharge occurs is determined by the forced ventilation system. Inflow and outflow velocities through the ventilation openings in the laboratory were measured. These data were provided to CFX as boundary condition in the calculation of the spatial flow field. Vertical cuts through this converged flow field at different positions are depicted in Fig. 3. It can be seen that higher velocities are found around the inlet and outlet openings of the ventilation. These openings are distributed along the ceiling and the side wall where the pipe rupture is assumed to occur. In some regions of the laboratory there are almost stagnant flow conditions with flow speeds of 0.1 m/s or below.
Constant flow rate: the ‘‘bottle-compressor’’ case
The assumption of a constant sonic flow through the leak corresponds to the ‘‘compressor-bottle’’ configuration, which makes use of an idealised compressor able to maintain a constant pressure to the hydrogen distribution line. A pressure of 10 MPa was assumed at the leak position together with the critical speed of hydrogen in air. The flow was stopped when the total available hydrogen mass was released. The hydrogen flow history was specified as a boundary condition in the code. Fig. 4 shows the resulting hydrogen mass flow rate in relation to the accumulated mass released. The total release time amounts to about 12 s. A main result expected from the simulations is the calculation of the size of flammable hydrogen clouds in the laboratory. The lowest hydrogen concentration at which ignition is possible (LFL) is about 4% in air. Fig. 5 illustrates the extension of these clouds during two different stages of the accident by an iso-surface of hydrogen concentration equal to 4%. The walls of the laboratory are coloured by local hydrogen concentrations according to the colour key given in the figure. In the early stage of the release hydrogen is found mainly
Fig. 4 – ‘‘Bottle-compressor’’ operative mode: Hydrogen flow through leak and total mass released.
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Fig. 5 – ‘‘Bottle-compressor’’ operative mode: Iso-surfaces of hydrogen concentration at the lower flammability limit (LFL – 4%) during the hydrogen release (left-hand side, approximately 4 s from release start) and later after completion of the release in the distribution phase (right-hand side, approximately 20 s after start of transient).
along the jet (left-hand side in Fig. 5). After the end of the release phase hydrogen is spread over most parts of the laboratory and is then slowly removed by the ventilation system. The maximum cloud size with hydrogen concentrations of 10% (representing the lower limit for substantial flame acceleration) or higher is found to be about 64 m3. Results are discussed quantitatively and in more detail in comparison with the simulations of the following section in paragraph 5.
4.2.
Variable flow rate: the ‘‘bottle-only’’ case
The other operative mode of the laboratory is the ‘‘bottleonly’’ mode, consisting in the direct connection of the
hydrogen gas bottle to the hydrogen distribution line without the hydrogen compressor. Also for this operation mode the simulations start from the converged ventilation run as initial state. The results shown were obtained with the finest mesh available. Investigations on the influence of the mesh resolution are discussed separately (Section 4.3). Due to the limited size of the storage its internal pressure drops quickly and both the density and the mass flow through the leak decrease with increasing time. The calculated mass flow rate from the leak is shown in Fig. 6. For a very limited early time the leak flow is almost constant but reduces later continuously until the pressures in the hydrogen storage and in the laboratory equalise. The release process is finished after about 90 s.
Fig. 6 – ‘‘Bottle-only’’ operative mode: Mass flow rate through the leak (circle points to the very early discharge phase shown in the second graph).
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Fig. 7 – ‘‘Bottle-only’’ operative mode: Iso-surface of hydrogen concentration at the lower flammability limit (LFL, 4%) in the lab during the hydrogen release (left-hand side, approximately 4 s after beginning) and later in the distribution phase (righthand side, approximately 20 s after start of transient).
Considering the size of the flammable volume (4% limit) Fig. 7 shows similar results compared with the first approach illustrated in Fig. 5. However for the current operation mode the flammable cloud remains smaller because the outflow of the same hydrogen mass is distributed over a longer time. The 4%-flammable cloud volume reaches a maximum of about 132 m3 whereas the 10% volume is nearly negligible and is less than 1 m3. With these results illustrated in Fig. 8 it becomes evident that flammable conditions exist only for a limited time of about 100 s. The chance of a fast deflagration is small due to the low hydrogen concentrations in the cloud. In the laboratory hydrogen sensors are installed at selected positions in order to detect hydrogen releases as early as possible. Two sensors are located under the hoods RIG and AMC (Fig. 2). Their approximate positions are shown in Fig. 9. The hydrogen mol fraction at these locations passes the alert level of 0.4% very early in the transient (less than 1 s). This would enable any kind of counter measures to mitigate the transient. However, for the purpose of studying the complete release evolution, as explained earlier, mitigation measures were not taken into account in this work.
The simulation of the hydrogen leak in the laboratory must also include a realistic simulation of the jet flow itself from initially 20 MPa in the storage to atmospheric pressure. Although not the main focus of this work the numerical grid was set-up sufficiently fine that the microstructures of the jet become visible. The pressure ratio between hydrogen storage and laboratory is much larger than the critical pressure ratio of hydrogen (about 1.9); therefore an under-expanded jet develops. Further, the pressure at the pipe exit is calculated to be much higher (with full bottle about 85 bar) than the environmental pressure which leads to a highly under-expanded jet. The maximum Mach number in the jet calculated is about 6.7 and the maximum gas speed reaches a value of approximately 2500 m/s. Due to the strong under-expansion of the jet at the position of the maximum Mach number the temperature drops to a very low value. This value is related to a local pressure of 4100 Pa. The under-expansion region is followed by the Mach disk which was found to be located at about 8 diameters downstream the release pipe with 4 mm diameter. There are many investigations on highly under-expanded jets available in the literature. Two examples are [25,26].
Fig. 8 – Size of the flammable volume with concentrations of 4% (left-hand side) and 10% (right-hand side) (‘‘bottle-only’’ operative mode).
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Fig. 9 – Time dependent mole fraction at two hydrogen sensor locations in the lab (‘‘bottle-only’’ operative mode).
Ref. [25] describes a numerical study on a critical hydrogen jet and ref. [26] discusses jet correlations derived from experimental data. The values calculated in this work are slightly different from the ones reported in [25]. In ref. [25] a maximum Mach number of nearly 9 is predicted. The Mach disk following the peak Mach number is found at 9.2 diameters. A detailed review of experimental investigations and relevant correlations on the position and size of the Mach disk is found in ref. [26]. The correlations of Ashkenas and Sherman [27] and the experimental data of Love [28] are used by Lehnasch [26] to calculate the size of the Mach disk that is in a range of 6–8 pipe diameters for the given pressure ratio. Those values are in good agreement with the own simulation results. It can be stated that the main microstructures of the underexpanded critical jet were calculated with acceptable accuracy in the simulations. This supports the credibility of the predictions presented in this work.
4.3.
Sensitivity studies
The following studies were carried out for the ‘‘bottle-only’’ case which includes the hydrogen storage in the mesh.
4.3.1.
Mesh resolution
A proper spatial resolution of the volume around the sonic jet is important for a realistic simulation of the hydrogen dispersion in the whole laboratory. Therefore, the mesh resolution in the leak area was consecutively increased in order to identify an influence on the predictions of flammable mass and volume. Three different computational meshes have been generated: a coarse, a medium and a fine mesh. Two of the three meshes prepared for the simulations use local grid refinement by an internal conic volume in the jet region which gives good control over the local cell size. The cone starts always at the leak opening and extends far enough downstream to include the full jet. It consists mostly of unstructured cells and is linked to the hexahedral mesh of the
lab. A summary of major grid parameters of the three meshes used is given in Table 1. The refinement of mesh structure in the jet region influences the relevant results in a significant manner only in the area close to the jet. In Fig. 10 Mach number profiles along the jet axis are displayed. With increasing spatial resolution the maximum Mach number at a distance of 10 mm downstream of the pipe exit increases considerably. Directly at the leak exit (before the final expansion downstream of the leak opening occurs) however, no influence is seen (black lines in Fig. 10). If global results such as the flammable mass and flammable volume history are considered no remarkable changes are observed when applying the different meshes. It can be concluded that the numerical predictions of the relevant parameters for hydrogen risk assessment in the laboratory were obtained with sufficient spatial resolution because further refinement of the grid does not change these results. This is also confirmed by the visualisation of the flammable cloud (LFL, 4%) in size and shape for the two finer meshes applied. The differences in the hydrogen distribution patterns are small as shown in Fig. 11.
4.3.2.
Turbulence model
Alternatively to the SST (Shear Stress Transport) turbulence model the RNG (Renormalisation Group) k–3 model was applied to simulate the hydrogen leak. Inside the fluid domain far away from boundaries the SST model behaves like
Table 1 – Main parameters of the three meshes applied. Average equivalent Number Total cell length in leak of elements number of region [mm] in leak cone elements Coarse mesh Medium mesh Fine mesh
38 26.2 14.4
– 9302 72 337
897 778 918 681 1 005 025
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by the RNG k–3 model has an about 20% higher peak value (about 180 m3) compared to the SST model simulation.
5.
Fig. 10 – Influence of local mesh resolution on calculated Mach numbers in the jet.
a standard k–3 model whereas the RNG k–3 model includes an extra term in the dissipation equation and uses different model constants. The RNG k–3 model provided good results for a sub-sonic jet analysed in a recent work [22]. Therefore this model was selected to compare to the SST model simulations. In the current work both turbulence models required similar computing times during the simulations but the SST model showed better convergence behaviour. Results were compared for two selected parameters. For the Mach number as a local parameter no influence of the turbulence model at a monitor point position at the centreline of the jet and 0.01 m downstream of the leak opening was seen. The second parameter to be compared was the flammable volume (4–75% hydrogen). This is an integrated parameter. It is influenced by the choice of the turbulence model. The jet simulated by the RNG k–3 model had a tendency to be more compact with less entrainment of air than the one calculated by the SST model. Therefore the flammable volume calculated
Discussion
The two principal modelling strategies followed in this work end up with quite different results. As shown in before, the mass flows from the leak are very different. In the ‘‘bottlecompressor’’ operative case a constant leak flow (Fig. 4) was assumed. With the second option ‘‘bottle-only’’ the leak flow was calculated as a high pressure discharge process from the hydrogen storage (Fig. 6). The decreasing hydrogen pressure and density in the storage reduces the mass flow through the leak and spreads the release over a much longer time compared with the first modelling option. In Fig. 12 the resulting flammable masses are compared for both cases. The longer outflow time in the case of the hydrogen storage/bottle in the calculation leads to a considerably smaller amount of mass which might be ignited (LFL). Further, the simplified modelling option with constant leak flow can be considered as a conservative approach in terms of the hydrogen risk as it results in a much larger flammable mass. This option is much less numerically demanding but considerably overestimates the amount of ignitable hydrogen mixture. Iso-surfaces of LFL for both modelling options can be compared in Figs. 5 and 7. The flammable clouds look very similar during the release. With both simulation options detached regions from the flammable cloud exist only at a late time after the end of the release. Then some isolated hydrogen–air pockets exist for a short time but these do not contribute to the flammability in the lab. The modelling results obtained are stable. They do not depend on the mesh resolution. The choice of the turbulence model has some influence on the parameters of interest like size of flammable volume. Within the scope of this work assumptions had to be made on the accident scenario and certain modelling options (e.g. real versus ideal gas). The
Fig. 11 – Hydrogen distribution in the lab at the moment of maximum cloud size for medium (left-hand side) and fine mesh (right-hand side) (‘‘bottle-only’’ operative mode).
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Fig. 12 – Flammable mass in the lab for both modelling options in comparison to the total mass released through the leak.
choice of the turbulence model is another contribution to a degree of uncertainty of the results presented.
6.
Conclusions
A conservative accident scenario was defined to evaluate a worst case situation in a hydrogen laboratory. On this basis simulations of the accidental release of hydrogen were carried out. The purpose of the work was to assess the formation of hydrogen clouds inside the laboratory. For this reason the integrated volume of regions which fulfil the ignition criteria of hydrogen were calculated during and after the hydrogen release. Two principal modelling options were followed corresponding to two typical operative modes of the laboratory. In the first approach a compressor is assumed to maintain the pressure in the hydrogen distribution line at a fixed pressure and the predefined critical mass flow across the leak is constant. In the second approach, the compressor is not included and the discharge occurs through pipes from a high pressure bottle/storage. Care was taken to resolve also the main microstructures of the leak jet by local mesh refinements. As a major result of ‘‘bottle-only’’ simulations it was calculated that flammability at the lower limit of 4% exists over approximately 100 s in a maximum volume of 130 m3. This volume was calculated with the SST turbulence model. The 4% flammability limit enables ignitability but flame propagation is only possible in upright vertical direction. The limit of 10% which represents the lower value of substantial flame acceleration was only passed in less than 1 m3 (corresponding to 9 g hydrogen). Additional simulations were performed with different mesh resolution and with an alternative turbulence model. The mesh investigations did not show a dependency of the major results concerning the hydrogen distribution. The choice of the turbulence model however has consequences in the calculation of the hydrogen release jet. The alternative RNG k–e model predicted a higher value of the
peak flammable volume at 4% hydrogen concentration. Combustion at this lower limit is very weak and therefore the increase in the uncertainty of the results limited. In terms of laboratory safety, the operative solution without the compressor is preferable, as it is also expected from an intuitive point of view, because it releases the total available hydrogen mass in a longer period, allowing the ventilation system to be more effective and reducing the maximum value of the flammable hydrogen mass/volume. To study the full evolution of hydrogen in the laboratory this work has disregarded the automatic safety system, e.g. shut-off valves. Experimental tests have shown that in reality the depressurisation starts approximately 2–3 s after that the hydrogen concentration at a sensor passes the 10% of the LFL value (0.4% of hydrogen in air). This will reduce the fraction of the hydrogen mass released into the laboratory (Fig. 12). In the cases studied it is finally interesting to note that the first sensor experiencing an increase of hydrogen concentration is the sensor positioned near the wall opposite to the leak position and not that positioned precisely above it. This is due to the behaviour of the hydrogen jet during the first release phase. The result demonstrates the importance to locate sensors in an optimal way. In the case of the laboratory under study typical accidental releases consist in ‘‘small’’ leaks, where buoyancy dominates from the very start of release. The sensors have been located ‘‘intuitively’’ on top and in front of each high pressure equipment and have demonstrated their usefulness. But in cases where hydrogen jets could be expected, an intuitive approach will not be enough and CFD tools should be used intensively in conjunction with risk assessment techniques to calculate locations with the most probable ‘‘early hydrogen increase’’. A possible extension of the presented work is the inclusion of real gas behaviour and the simulation of a hydrogen deflagration across the flammable cloud. Property data of the modelled gas components in a wide range of pressure and temperature are needed to carry out the super-critical jet expansion. Improvements on the property data may also lead to a better quality of the simulation results.
international journal of hydrogen energy 35 (2010) 4409–4419
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