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Parametric analysis on multi-stage high pressure reducing valve for hydrogen decompression
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Parametric analysis on multi-stage high pressure reducing valve for hydrogen decompression Fu-qiang Chen a,b, Xiao-dong Ren a, Bo Hu a, Xue-song Li a,*, Chun-wei Gu a, Zhi-jiang Jin b,** a
Key Laboratory for Thermal Science and Power Engineering of Ministry of Education, Department of Energy and Power Engineering, Tsinghua University, Beijing, 100084, China b Institute of Process Equipment, College of Energy Engineering, Zhejiang University, Hangzhou, 310027, China
highlights Demanding conditions are imposed on the pressure reducing system of hydrogen fuel cell electric vehicle. Multi-stage high pressure reducing valve for hydrogen decompression is proposed. Effects of different structural parameters on the internal flow characteristics are investigated. A better hydrogen decompression process is achieved with the parametric results.
article info
abstract
Article history:
Hydrogen fuel cell electric vehicle (FCEV) can reduce air pollution as well as achieve effi-
Received 29 July 2019
cient use of hydrogen energy. Farther travel distance requires larger hydrogen storage
Received in revised form
pressure, thereby imposing more demanding working conditions on the pressure reducing
29 September 2019
system. In this paper, a multi-stage high pressure reducing valve (MSHPRV) for hydrogen
Accepted 1 October 2019
decompression in FCEV is proposed, and the effects of different structural parameters on
Available online 1 November 2019
its internal flow characteristics are investigated to achieve a better hydrogen decompression process. Results show that compared with perforated plate, multi-stage perforated
Keywords:
sleeves and valve core hold the dominant position in hydrogen throttling process. Larger
Multi-stage high pressure reducing
multi-stage perforated sleeve diameter, perforated plate diameter and pressure ratio relate
valve
to larger hydrogen kinetic energy, turbulence vortex and energy consumption. However,
Structural parameters
with the increase of perforated plate stage and perforated plate radius, the turbulent in-
Fuel cell electric vehicle
tensity and energy consumption inside MSHPRV decreases correspondingly. This study can
Hydrogen decompression
provide some technical supports for achieving hydrogen decompression in FCEV when
Computational fluid dynamics
facing harsh working conditions, or help with dealing energy conversion during decompression process. © 2019 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.
* Corresponding author. ** Corresponding author. E-mail addresses: [email protected] (X.-s. Li), [email protected] (Z.-j. Jin). https://doi.org/10.1016/j.ijhydene.2019.10.004 0360-3199/© 2019 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.
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Introduction Hydrogen FCEV can achieve nearly zero exhaust emission. It can prevent environmental pollution as well as achieve efficient use of hydrogen energy [1]. In order to make FCEV reach a farther travel distance, greater demands are put on its pressure reducing system [2]. Farther travel distance requires larger hydrogen storage pressure, here, FCEV running 200 km requires hydrogen pressure in storage tank to be maintained at 35 MPa [3]. In this harsh condition, the traditional pressure reducing valves for hydrogen FCEV will be difficult to regulate pressure. Therefore, it is necessary to develop better pressure reducing systems to cope with the harsh working condition inside FCEV. Research on control valves for hydrogen energy utilization has been carried out by many researchers. Kim et al. [4] conducted a numerical study on the performance of polymer electrolyte membrane fuel cell vehicle with variable operating pressure to investigate the effect of blower and backpressure control valve. Jin et al. [5] proposed a novel high multi-stage pressure reducing valve for hydrogen stable decompression in hydrogen refueling station. Ferreira et al. [6] investigated the effect of sudden valve opening on hydrogen generation rate in batch sodium borohydride hydrolysis systems. Cao and Kedziora [7] conducted the innovative designs of an in-tank hydrogen valve towards direct metal laser sintering compatibility and fatigue life enhancement. Chen et al. [8] presented a numerical study on reversed hydrogen flows through the multi-stage Tesla valves during hydrogen decompression for hydrogen fuel cell. Menaa et al. [9] focused on optimizing the hydrogen timed manifold injection system through valve lift law and hydrogen injection parameters to prevent backfire phenomena. Wang et al. [10] conducted research on sealing performance and self-acting valve reliability to find out the method to improve the reliability of high-pressure oil-free reciprocating compressors for hydrogen refueling stations. Lee et al. [11] investigated the applications of intake and exhaust valve timing variation and lean boosting to a hydrogen engine with external mixture to identify the possibility of achieving high power and efficiency without backfire generation. Moreover, many researchers conducted study on other valves for hydrogen FCEV or related utilizations [12e17]. Tsujimura et al. [18] and Yu et al. [19] both paid attention to the pressure reducing systems for hydrogen decompression. Hydrogen decompression in fuel cell system has been widely studied by some researchers, which can provide some guidance on the design of the pressure reducing valves. Ye et al. [20] designed a new fuzzy control system to control the hydrogen pressure in fuel cell system. Corgnale et al. [21] built a 2-L adsorbent prototype tank for fuel cell driven vehicles and investigated the hydrogen decompression. Kuroki et al. [22] developed a dynamic simulation approach to propose an optimal hydrogen refueling method. Lin et al. [23] proposed a method for determining the optimal delivered hydrogen pressure for fuel cell electric vehicles. Yin YB et al. [24], Michael et al. [25] and Jung et al. [26] all paid attention to hydrogen decompression in hydrogen FCEV. Hussein et al. [27] performed a numerical study of pressure peaking for ignited hydrogen releases in an enclosure. Tanc¸ et al. [28] presented
an overview of the next quarter century vision of hydrogen fuel cell electric vehicles. Bauer et al. [29] evaluated two different refueling station concepts for fuel cell passenger cars with 70 MPa technology. Talpacci et al. [30], Yue et al. [31], Fernandez et al. [32], Robledo et al. [33] and Kelly et al. [34] all focused attentions on different aspects of the hydrogen FCEV. Liu et al. [35] proposed a computational fluid dynamics model to predict the decompression wave speed of high-pressure hydrogen-natural gas mixtures in pipelines. Except for fuel cell system, hydrogen decompression in vehicle gas cylinder or hydrogen refueling station was also widely studied [36e41]. Ramadhani et al. [42] presented an innovative design and analysis of solid oxide fuel cell based poly-generation system for residential applications. In the previous study [2], a two-step high pressure reducing system for hydrogen FCEV was proposed, and the effect of different valve openings on hydrogen flow was investigated. It was found that the new system works well for hydrogen decompression under complex conditions. However, a better hydrogen flow inside MSHPRV with smaller turbulent intensity and lower energy consumption has not been realized. Therefore, in this paper, the effects of different structural parameters on the internal flow characteristics of MSHPRV are investigated, with the purpose of achieving a better hydrogen decompression process for FCEV.
Model description Mathematical model Hydrogen is hydrogenated from the hydrogen refueling station to the vehicle gas cylinder, and then it is delivered to the fuel cell stack. The hydrogen transfer process can be approximated taken as a fixed volume, variable mass deflation process. It is assumed that the hydrogen in the gas cylinder has sufficient heat exchange with the outside world, therefore, the hydrogen transfer process is considered to be isothermal. Hydrogen mass flow through the orifice at subsonic flow is described as follows [43]. vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 3 u u 2k kþ1 u k 7 Pi u 2k 6 6 P0 P0 7 Qm ¼ fs ðS0 ; Pi ; P0 ; TÞ ¼ CS0 pffiffiffiffiffiffi u 5 Pi RT tk 1 4 Pi
(1)
Hydrogen mass flow through the orifice at supersonic flow is described as follows. Pi Qm ¼ fc ðS0 ; Pi ; TÞ ¼ CS0 pffiffiffiffiffiffi RT
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 k1 2k 2 kþ1 kþ1
(2)
Where, C refers to flow coefficient of the orifice, S0 refers to throttling area, T refers to absolute temperature of hydrogen, k refers to adiabatic ratio coefficient of hydrogen, Pi refers to inlet pressure of throttling component, P0 refers to outlet pressure of throttling component. Hydrogen continuity equation is described as follows. Qm ¼
dM dt
(3)
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hydrogen pressure reducing process totally, which can contribute to more gradual flow and smaller turbulent intensity. Changes in the throttling areas can in turn changes the hydrogen flow state inside MSHPRV. Fig. 1(b) shows the mesh model of MSHPRV. 1/2 model of MSHPRV is chosen as the computational domain since the flow model is symmetrical. Considering the structural complexity of MSHPRV, meshing is performed using hybrid grid technology. Structured grid is used for inlet, middle and outlet pipelines, whose structures are regular. Unstructured grid is used for multi-stage perforated sleeves, valve core and valve chamber. Fluid medium inside MSHPRV is hydrogen with initial pressure 35 MPa. Pressure inlet and outlet are chosen as the boundary conditions respectively. Outlet pressure of MSHPRV and the previously designed multi-stage muffler is 5 MPa and 0.16 MPa respectively. However, as a component which dominants the hydrogen decompression process, only MSHPRV is considered in this study. Center plane in MSHPRV is set as the symmetry. Steady-state solver and the second-order upwind scheme are adopted. RNG k-ε turbulence model are used since it has a better accuracy in the study of flow with turbulent vortex. Grid-independent verification is necessary to ensure the accuracy of numerical method adopted in this study. Table 1 shows that the mass flow rate in MSHPRV changes with different grid numbers at inlet pressure 3.0 MPa. It can be seen from Table 1 that, the relative error of mass flow rate is 14.1%, 5.0% and 3.2% for grid number 1.18 107, 1.2 107 and 1.22 107 respectively. It can be found that grid number 1.2 107 is a suitable chose. Therefore, it can be used as an optimal solution to ensure efficiency and accuracy of numerical calculation.
Fig. 1 e Physical model of MSHPRV. Where, M refers to hydrogen mass flow in the gas cylinder, Qm refers to hydrogen flux. Pressure dynamic equation is described as follows. 2
3
RT V 7 dM dP 6 2a m ¼6 M7 2 2 4 5 dt dt ðmVÞ V mb M
(4)
Where, P refers to cylinder pressure, T refers to cylinder temperature, R refers to hydrogen gas constant, V refers to hydrogen storage volume, m refers to hydrogen molar mass.
Model verification MSHPRV is a newly designed high pressure reducing valve. Nowadays, it is only conceptually designed for hydrogen FCEV. However, it is already used in some process industries with steam medium. Since the numerical model adopted in MSHPRV for hydrogen flow in FCEV is similar to that for steam flow in process industry, the numerical method is verified by experimental method with steam medium. Table 2 shows the experimental and numerical results of flow rate inside MSHPRV with steam medium. It can be seen
Computational model and boundary conditions Fig. 1(a) shows the three-dimensional model of MSHPRV. It can be seen from Fig. 1(a) that, MSHPRV is mainly composed of inlet valve chamber, multi-stage perforated sleeves, valve core, valve steam, perforated plate, valve body, outlet valve chamber and the others. Working principle of MSHPRV was described in the previous work [2]. It can achieve five-stage
Table 2 e Experimental and numerical results of flow rate inside MSHPRV. L/Lmax qm numerical (t/h) qm experimental (t/h) d(t/h)
0.2 110 114.5 4.5
0.4 120 120.5 0.5
0.6 125 126.03 1.03
0.8 133 131.73 1.27
1.0 140 137.43 2.57
Table 1 e Grid-independent verification. N (106) Flux (kg/s) d(%)
11.6 16 15.2
11.8 24 14.1
12 28.5 5.0
12.2 30 3.2
13 31 3.1
14 31.5 3.1
15 31.8 3.1
16 31.7 3.1
17 31.8 3.1
18 31.7 3.1
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Fig. 2 e MSHPRV with different perforated plate stages.
Fig. 3 e Pressure distribution of symmetry plane with different perforated plate stages (MPa).
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regardless of temperature change. Therefore, it can be indirectly concluded that the numerical model adopted in MSHPRV for hydrogen flow in FCEV is reliable.
Results and discussion In order to obtain a better hydrogen decompression process inside MSHPRV under harsh working conditions as high pressure and large pressure ratio, the effects of different structural parameters on its internal flow characteristics are investigated. Here, the structural parameters of main throttling components, such as different perforated plate stages, perforated plate diameters, multi-stage sleeve diameters, perforated plate radii and pressure ratios, are studied.
Effect of perforated plate stage Fig. 4 e Velocity distribution along Z direction with fourstage perforated plates (m/s).
Table 3 e Velocity and turbulence parameter maximum difference with different perforated plate stages. Stages Dvmax (m/s) Dkmax (m2/s2)(106) Dεmax (m2/s3)(1012) 1 2 3 4
108 104 103 90
2.8 2.2 2.1 1.8
6 6 5.8 5.5
from Table 2 that, the relative mass flow rate error d between experimental and numerical results inside MSHPRV with relative valve opening L/Lmax 0.2, 0.4, 0.6, 0.8 and 1.0 is 4.5 t/h, 0.5 t/h, 1.03 t/h, 1.27 t/h and 2.57 t/h respectively. Therefore, the maximum difference of flow rate between numerical and experimental results is located at L/Lmax 0.2, and the error d is only about 4%. It can be found that compared with experimental data, the numerical results is not much different from it. Temperature difference between steam and hydrogen under high pressure is large. However, energy equation is closed in this study, and attention is only paid to the flow
Different perforated plate stages will cause changes in the throttling areas, which in turn changes the hydrogen flow state. Therefore, the effect of perforated plate stage on hydrogen internal flow inside MSHPRV is investigated in this part. Fig. 2 shows the three-dimensional flow model of MSHPRV with different perforated plate stages. Fig. 3 shows the pressure contours of symmetry plane with different perforated plate stages. It can be seen from Fig. 3 that, taking MSHPRV with one-stage perforated plate as an example, hydrogen pressure reduces from 35 MPa to 10 MPa at the multi-stage perforated sleeves and valve core, and from 10 MPa to 5 MPa at the last-stage perforated plate. Therefore, compared with perforated plate, multi-stage perforated sleeves and valve core play a more important role in hydrogen decompression. Moreover, hydrogen pressure changing gradient mainly reflects at those throttling components, while evenly distributed in the inlet chamber and outlet chamber. Different perforated plate stages do not have a great impact on the hydrogen pressure reduction, because perforated sleeves and valve core take the dominant position. Fig. 4 shows the velocity distribution of MSHPRV along Z direction with four-stage perforated plates. It can be seen from Fig. 4 that, corresponding to the pressure change, the hydrogen
Fig. 5 e MSHPRV with different perforated plate diameters.
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Fig. 6 e Velocity distribution of symmetry plane with different perforated plate diameters (m/s).
velocity changes greatly at the throttling elements. There are two peaks in the curve indicated by the red ellipse box, representing the two sides of the multi-stage sleeves and valve core. Moreover, there are four peaks in the red square box, which correspond to the four-stage perforated plates. This is because as the areas of the throttling elements decrease, the
flow rate increases, and the fluid kinetic energy increases correspondingly. According to the law of energy conservation, the potential energy is reduced, resulting in pressure loss. It can be found that, the maximum velocity at the multi-stage sleeves is around 100 m/s, while about 40 m/s at the perforated plates. Therefore, just as mentioned above, compared
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Effect of perforated plate diameter
Fig. 7 e Turbulent kinetic energy along Z direction with different perforated plate diameters.
with perforated plate, multi-stage perforated sleeves and valve core hold the dominant position in hydrogen throttling process. Table 3 shows the maximum difference of velocity, turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different perforated plate stages. It can be seen from Table 3 that, the maximum difference of velocity inside MSHPRV with perforated plate stage 1, 2, 3 and 4 is 108 m/s, 104 m/s, 103 m/s and 90 m/s respectively. Moreover, the correspondingly maximum difference of turbulent kinetic energy is 2.8 106 m2/s2, 2.2 106 m2/s2, 2.1 106 m2/s2 and 1.8 106 m2/s2. It can be found that with the increase of perforated plate stage, the turbulent flow can be gradually reduced.
Different perforated plate diameters will also change the throttling areas. Fig. 5 shows the three-dimensional flow model of MSHPRV with different perforated plate diameters. The original perforated plate diameter for engineering application is 8 mm, therefore, taking the two values before and after 8 mm with an arithmetic progression. Fig. 6 shows the velocity distribution of symmetry plane with different perforated plate diameters. It can be seen from Fig. 6 that, the maximum velocity inside MSHPRV exists at the multi-stage sleeves or the perforated plate, and compared with perforated plate, the maximum velocity mainly reflects at the perforated sleeves and valve core. Meanwhile, the maximum velocity at the multi-stage perforated sleeves is 108 m/s, 110 m/s, 110 m/s, 116 m/s and 120 m/s for MSHPRV with perforated plate diameter 4 mm, 6 mm, 8 mm, 10 mm and 12 mm respectively. Moreover, with the increasing of perforated plate diameter, the velocity gradient in the valve chamber after multi-stage sleeves and valve core increases correspondingly. However, the velocity gradient in the valve chamber after perforated plate decreases. It can be seen from Fig. 6 that, the turbulent vortex in the valve chamber increases with the increase of perforated plate diameter, while decreases at the outlet pipeline. Changing of perforated plate diameter has a great impact on the hydrogen internal flow inside the whole MSHPRV. Just as mentioned above, multistage perforated sleeves and valve core hold the dominant position in hydrogen throttling process. Therefore, it can be concluded that large perforated plate diameter will result in large hydrogen kinetic energy. Fig. 7 shows the turbulent kinetic energy of MSHPRV along Z direction with different perforated plate diameters. It can be seen from Fig. 7 that, in accordance with Fig. 4, there exist three peaks in the curve, which located at the multi-stage
Fig. 8 e MSHPRV with different sleeve diameters.
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Fig. 9 e Velocity and streamlines distribution of symmetry plane with different sleeve diameters (m/s).
sleeves and perforated plate respectively. Therefore, the turbulent vortex inside MSHPRV is mainly reflected at the throttling components like multi-stage perforated sleeves, valve core and perforated plate. Moreover, the maximum turbulent kinetic energy which marked in black circle box is about 1.05 106 m2/s2, 1.80 106 m2/s2, 2.15 106 m2/s2, 2.25 106 m2/s2 and 2.38 106 m2/s2 for MSHPRV with perforated plate diameter 4 mm, 6 mm, 8 mm, 10 mm and 12 mm respectively. Large perforated plate diameter refers to large turbulent kinetic energy. Therefore, it can be concluded that with the increasing of perforated plate diameter, the turbulent intensity and energy consumption will increase correspondingly.
Effect of sleeve diameter Multi-stage perforated sleeves and valve core hold the dominant position in hydrogen throttling process, therefore, a parametric study on multi-stage perforated sleeves is
conducted. Fig. 8 shows the three-dimensional flow model of MSHPRV with different perforated sleeve diameters. The original multi-stage perforated sleeves diameter for engineering application is 5 mm, therefore, taking the two values before and after 5 mm using an arithmetic progression with a tolerance of 1. Fig. 9 shows the velocity and streamlines distribution of symmetry plane with different sleeve diameters. It is known from the above studies that the maximum hydrogen velocity inside MSHPRV is mainly distributed at the multi-stage perforated sleeves, so it is partially enlarged to study the effect of sleeve diameter on hydrogen flow. It can be seen from Fig. 9 that, with the increasing of sleeve diameter, the maximum velocity distribution range which located at the multi-stage perforated sleeves and valve core increases gradually. At the same time, the velocity gradient in the valve chamber after perforated sleeves and valve core increases correspondingly. Moreover, it can be found from the streamlines distribution that, the vortex becomes larger gradually as
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Fig. 10 e Turbulence parameter maximum difference with different sleeve diameters.
Fig. 11 e Pressure and velocity distributions of symmetry plane with different perforated plate radii (MPa) (m/s).
the sleeve diameter increases. Therefore, it can be concluded that larger multi-stage perforated sleeve diameter relates to larger hydrogen kinetic energy and turbulence vortex. Fig. 10 shows the maximum difference of turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different perforated sleeve diameters. It can be seen from Fig. 10 that, the maximum difference of turbulent kinetic
energy inside MSHPRV with perforated sleeve diameter 3 mm, 4 mm, 5 mm, 6 mm and 7 mm is 1.4 106 m2/s2, 1.5 106 m2/ s2, 1.8 106 m2/s2, 1.8 106 m2/s2 and 2.0 106 m2/s2 respectively. Moreover, the correspondingly maximum difference of turbulent dissipation rate is 5.0 1012 m2/s3, 6.0 1012 m2/s3, 5.6 1012 m2/s3, 7.0 1012 m2/s3 and 8.5 1012 m2/s3. Therefore, it can be concluded that with the
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increase of multi-stage perforated sleeve diameter, the turbulent intensity and energy consumption inside MSHPRV becomes larger.
Effect of perforated plate radius The purpose of chamfering is to change the straight transition to a smooth transition. Compared with straight transition, smooth transition can make the hydrogen flow at the corners to be more uniform and the turbulent intensity to be smaller. Therefore, chamfering can help change the hydrogen flow state inside MSHPRV. The first row of Fig. 11 shows the threedimensional flow model of MSHPRV with different perforated plate radii. Fig. 11 shows the pressure and velocity distributions of symmetry plane with different perforated plate radii. It can be seen from Fig. 11 that, with the increasing of perforated plate radii, the light blue area gradually shrinks and the dark blue area gradually increases, which indicates that the pressure distribution in the valve chamber after the multi-stage perforated sleeves and valve core tends to be uniform. Moreover, the velocity gradient in the valve chamber decreases correspondingly, which can be seen from the third row of Fig. 11 that the light blue area shrinks gradually. In the meantime, the turbulent vortex area decreases with the increase of perforated plate radius. Therefore, it can be concluded that larger perforated plate radii contribute to a better hydrogen flow state inside MSHPRV. Fig. 12 shows the maximum difference of turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different perforated sleeve radii. It can be seen from Fig. 12 that, the maximum difference of turbulent kinetic energy inside MSHPRV with perforated plate radius 0 mm, 0.5 mm, 1 mm, 1.5 mm and 2 mm is 2.8 106 m2/s2, 2.2 106 m2/s2, 2.1 106 m2/s2, 1.9 106 m2/s2 and 1.8 106 m2/s2 respectively. Therefore, it can be found that larger perforated plate radius relates to smaller turbulent intensity inside MSHPRV. That is because chamfering can help change the hydrogen flow state inside MSHPRV. Moreover, the correspondingly maximum difference of turbulent dissipation rate is 6.0 1012 m2/s3,
Fig. 12 e Turbulence parameter maximum difference with different perforated plate radii.
Table 4 e Velocity and turbulence parameter maximum difference with different pressure ratios. p1/p2 4 5 6 7 8
Dvmax (m/s)
Dkmax (m2/s2)(106)
Dεmax (m2/s3)(1012)
75 90 90 100 120
1 1.4 1.8 1.8 2.8
1.8 2.8 3.5 6 8
4.0 1012 m2/s3, 3.8 1012 m2/s3, 3.4 1012 m2/s3 and 3.1 1012 m2/s3. It again shows that the perforated plate radius is beneficial for improving the internal turbulent flow field inside MSHPRV. Therefore, it can be concluded that with the increase of perforated plate radius, the turbulent intensity and energy consumption inside MSHPRV decreases correspondingly.
Effect of pressure ratio Pressure ratio is closely related to the boundary conditions of hydrogen flow inside MSHPRV. Therefore, the effect of pressure ratio on the flow characteristics is investigated. Pressure distribution of MSHPRV along Z direction with different pressure ratios is investigated. There exist three peaks in the curves of pressure decompression with different pressure ratios, which located at the two sides of multi-stage sleeves and perforated plate respectively. The change of the flow area at the throttling components changes the flow rate and the kinetic energy of hydrogen, which in turn causes a certain pressure loss, so the pressure gradient is mainly reflected at such throttling components as multi-stage perforated sleeves, valve core and perforated plate. Moreover, the maximum pressure gradient at the multi-stage perforated sleeves and valve core is 9.8 MPa, 14.7 MPa, 19.8 MPa, 24.6 MPa and 29.5 MPa for MSHPRV with pressure ratio 4, 5, 6, 7 and 8 respectively. However, the maximum pressure gradient at the perforated plate is 5.2 MPa, 5.3 MPa, 5.2 MPa, 5.4 MPa and 5.5 MPa correspondingly. Therefore, compared with perforated plate, hydrogen pressure and velocity gradients are mainly reflected at the multi-stage perforated sleeves. Table 4 shows the maximum difference of velocity, turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different pressure ratios. It can be seen from Table 4 that, the maximum difference of velocity inside MSHPRV with pressure ratio 4, 5, 6, 7 and 8 is 75 m/s, 90 m/s, 90 m/s, 100 m/s and 120 m/s respectively. Larger pressure ratio relates to larger velocity difference. Moreover, the correspondingly maximum difference of turbulent kinetic energy is 1 106 m2/s2, 1.4 106 m2/s2, 1.8 106 m2/s2, 1.8 106 m2/s2 and 2.8 106 m2/s2. Therefore, the turbulent kinetic energy increases with the increase of pressure ratio. The maximum difference of turbulent dissipation rate is 1.8 1012 m2/s3, 2.8 1012 m2/s3, 3.5 1012 m2/s3, 6 1012 m2/s3 and 8 1012 m2/s3. It can be found that with the increase of pressure ratio, the maximum difference of velocity and turbulent parameters all increase accordingly. Therefore, it can be concluded that larger pressure ratio relates to larger turbulent flow and energy consumption inside MSHPRV.
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Conclusions A multi-stage high pressure reducing valve (MSHPRV) for hydrogen decompression in FCEV is proposed. The effects of different structural parameters on internal flow characteristics of MSHPRV are investigated, with the purpose of achieving a better hydrogen decompression process for FCEV. Results show that compared with perforated plate, multi-stage perforated sleeves and valve core hold the dominant position in hydrogen throttling process. Different perforated plate stages do not have a great impact on the hydrogen pressure reduction. However, with the increase of perforated plate stage, the turbulent flow can be gradually reduced. Secondly, large perforated plate diameter will result in large hydrogen kinetic energy. With the increasing of perforated plate diameter, the turbulent intensity and energy consumption will increase correspondingly. Thirdly, larger multi-stage perforated sleeve diameter relates to larger hydrogen kinetic energy, turbulence vortex and energy consumption. Meanwhile, larger perforated plate radii contribute to a better hydrogen flow state inside MSHPRV. With the increase of perforated plate radius, the turbulent intensity and energy consumption inside MSHPRV decreases correspondingly. Finally, larger pressure ratio relates to larger turbulent flow and energy consumption inside MSHPRV. In conclusion, better hydrogen flow characteristics inside MSHPRV for hydrogen FCEV are achieved with larger perforated plate stage, smaller perforated plate and sleeve diameter, larger perforated plate radii and smaller pressure ratio. In the meanwhile, the appropriate structural parameters should be chosen by considering the processing difficulty and manufacturing cost. This research work can provide some technical supports for achieving hydrogen decompression in FCEV when facing harsh working conditions, or to whom dealing with energy conversion and better flow state.
Acknowledgements This work is supported by the National Natural Science Foundation of China through Grant No. 51875514, the Key project of Natural Science Foundation of Zhejiang Province, China through Grant No. LZ17E050002, and the Zhejiang Key Research & Development Project, grant number 2019C01025.
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[23] Lin ZH, Ou SQ, Elgowainy A, Reddi K, Veenstra M, Verduzco L. A method for determining the optimal delivered hydrogen pressure for fuel cell electric vehicles. Appl Energy 2018;216:183e94. [24] Yin YB, Zhang L, Li L, Shen L, Fu JY. Study on pressure and velocity fields of pneumatic high-pressure reduction valve for hydrogen vehicles. Chin J Constr Mach 2011;9:1e6 [in Chinese]. [25] Michael MK, Brendan S, Marc CS, Donald D, Samuelsen GS, Jacob B. Air quality impacts of fuel cell electric hydrogen vehicles with high levels of renewable power generation. Int J Hydrogen Energy 2016;41:16592e603. [26] Jung A, Oh JK, Han K, Kim MS. An experimental study on the hydrogen crossover in polymer electrolyte membrane fuel cells for various current densities. Appl Energy 2016;175:212e7. [27] Hussein HG, Brennan S, Shentsov V, Makarov D, Molkov V. Numerical validation of pressure peaking from an ignited hydrogen release in a laboratory-scale enclosure and application to a garage scenario. Int J Hydrogen Energy 2018;43:17954e68. lu E, Aydın K. Overview of the next [28] Tanc¸ B, Arat HT, Baltacıog quarter century vision of hydrogen fuel cell electric vehicles. Int J Hydrogen Energy 2019;44:10120e8. [29] Bauer A, Mayer T, Semmel M, Morales MAG, Wind J. Energetic evaluation of hydrogen refueling stations with liquid or gaseous stored hydrogen. Int J Hydrogen Energy 2019;44:6795e812. [30] Talpacci E, Reub M, Grube T, Cilibrizzi P, Gunnella R, Robinius M, Stolten D. Effect of cascade storage system topology on the cooling energy consumption in fueling stations for hydrogen vehicles. Int J Hydrogen Energy 2018;43:6256e65. [31] Yue ML, Jemei S, Gouriveau R, Zerhouni N. Review on healthconscious energy management strategies for fuel cell hybrid electric vehicles: degradation models and strategies. Int J Hydrogen Energy 2019;44:6844e61. [32] Fernandez RA, Cilleruelo FB, Martinez IV. A new approach to battery powered electric vehicles: a hydrogen fuel-cell-based range extender system. Int J Hydrogen Energy 2016;41:4808e19. [33] Robledo CB, Oldenbroek V, Abbruzzese F, van Wijk AJM. Integrating a hydrogen fuel cell electric vehicle with vehicle-
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Available online at www.sciencedirect.com
ScienceDirect journal homepage: www.elsevier.com/locate/he
Parametric analysis on multi-stage high pressure reducing valve for hydrogen decompression Fu-qiang Chen a,b, Xiao-dong Ren a, Bo Hu a, Xue-song Li a,*, Chun-wei Gu a, Zhi-jiang Jin b,** a
Key Laboratory for Thermal Science and Power Engineering of Ministry of Education, Department of Energy and Power Engineering, Tsinghua University, Beijing, 100084, China b Institute of Process Equipment, College of Energy Engineering, Zhejiang University, Hangzhou, 310027, China
highlights Demanding conditions are imposed on the pressure reducing system of hydrogen fuel cell electric vehicle. Multi-stage high pressure reducing valve for hydrogen decompression is proposed. Effects of different structural parameters on the internal flow characteristics are investigated. A better hydrogen decompression process is achieved with the parametric results.
article info
abstract
Article history:
Hydrogen fuel cell electric vehicle (FCEV) can reduce air pollution as well as achieve effi-
Received 29 July 2019
cient use of hydrogen energy. Farther travel distance requires larger hydrogen storage
Received in revised form
pressure, thereby imposing more demanding working conditions on the pressure reducing
29 September 2019
system. In this paper, a multi-stage high pressure reducing valve (MSHPRV) for hydrogen
Accepted 1 October 2019
decompression in FCEV is proposed, and the effects of different structural parameters on
Available online 1 November 2019
its internal flow characteristics are investigated to achieve a better hydrogen decompression process. Results show that compared with perforated plate, multi-stage perforated
Keywords:
sleeves and valve core hold the dominant position in hydrogen throttling process. Larger
Multi-stage high pressure reducing
multi-stage perforated sleeve diameter, perforated plate diameter and pressure ratio relate
valve
to larger hydrogen kinetic energy, turbulence vortex and energy consumption. However,
Structural parameters
with the increase of perforated plate stage and perforated plate radius, the turbulent in-
Fuel cell electric vehicle
tensity and energy consumption inside MSHPRV decreases correspondingly. This study can
Hydrogen decompression
provide some technical supports for achieving hydrogen decompression in FCEV when
Computational fluid dynamics
facing harsh working conditions, or help with dealing energy conversion during decompression process. © 2019 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.
* Corresponding author. ** Corresponding author. E-mail addresses: [email protected] (X.-s. Li), [email protected] (Z.-j. Jin). https://doi.org/10.1016/j.ijhydene.2019.10.004 0360-3199/© 2019 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.
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Introduction Hydrogen FCEV can achieve nearly zero exhaust emission. It can prevent environmental pollution as well as achieve efficient use of hydrogen energy [1]. In order to make FCEV reach a farther travel distance, greater demands are put on its pressure reducing system [2]. Farther travel distance requires larger hydrogen storage pressure, here, FCEV running 200 km requires hydrogen pressure in storage tank to be maintained at 35 MPa [3]. In this harsh condition, the traditional pressure reducing valves for hydrogen FCEV will be difficult to regulate pressure. Therefore, it is necessary to develop better pressure reducing systems to cope with the harsh working condition inside FCEV. Research on control valves for hydrogen energy utilization has been carried out by many researchers. Kim et al. [4] conducted a numerical study on the performance of polymer electrolyte membrane fuel cell vehicle with variable operating pressure to investigate the effect of blower and backpressure control valve. Jin et al. [5] proposed a novel high multi-stage pressure reducing valve for hydrogen stable decompression in hydrogen refueling station. Ferreira et al. [6] investigated the effect of sudden valve opening on hydrogen generation rate in batch sodium borohydride hydrolysis systems. Cao and Kedziora [7] conducted the innovative designs of an in-tank hydrogen valve towards direct metal laser sintering compatibility and fatigue life enhancement. Chen et al. [8] presented a numerical study on reversed hydrogen flows through the multi-stage Tesla valves during hydrogen decompression for hydrogen fuel cell. Menaa et al. [9] focused on optimizing the hydrogen timed manifold injection system through valve lift law and hydrogen injection parameters to prevent backfire phenomena. Wang et al. [10] conducted research on sealing performance and self-acting valve reliability to find out the method to improve the reliability of high-pressure oil-free reciprocating compressors for hydrogen refueling stations. Lee et al. [11] investigated the applications of intake and exhaust valve timing variation and lean boosting to a hydrogen engine with external mixture to identify the possibility of achieving high power and efficiency without backfire generation. Moreover, many researchers conducted study on other valves for hydrogen FCEV or related utilizations [12e17]. Tsujimura et al. [18] and Yu et al. [19] both paid attention to the pressure reducing systems for hydrogen decompression. Hydrogen decompression in fuel cell system has been widely studied by some researchers, which can provide some guidance on the design of the pressure reducing valves. Ye et al. [20] designed a new fuzzy control system to control the hydrogen pressure in fuel cell system. Corgnale et al. [21] built a 2-L adsorbent prototype tank for fuel cell driven vehicles and investigated the hydrogen decompression. Kuroki et al. [22] developed a dynamic simulation approach to propose an optimal hydrogen refueling method. Lin et al. [23] proposed a method for determining the optimal delivered hydrogen pressure for fuel cell electric vehicles. Yin YB et al. [24], Michael et al. [25] and Jung et al. [26] all paid attention to hydrogen decompression in hydrogen FCEV. Hussein et al. [27] performed a numerical study of pressure peaking for ignited hydrogen releases in an enclosure. Tanc¸ et al. [28] presented
an overview of the next quarter century vision of hydrogen fuel cell electric vehicles. Bauer et al. [29] evaluated two different refueling station concepts for fuel cell passenger cars with 70 MPa technology. Talpacci et al. [30], Yue et al. [31], Fernandez et al. [32], Robledo et al. [33] and Kelly et al. [34] all focused attentions on different aspects of the hydrogen FCEV. Liu et al. [35] proposed a computational fluid dynamics model to predict the decompression wave speed of high-pressure hydrogen-natural gas mixtures in pipelines. Except for fuel cell system, hydrogen decompression in vehicle gas cylinder or hydrogen refueling station was also widely studied [36e41]. Ramadhani et al. [42] presented an innovative design and analysis of solid oxide fuel cell based poly-generation system for residential applications. In the previous study [2], a two-step high pressure reducing system for hydrogen FCEV was proposed, and the effect of different valve openings on hydrogen flow was investigated. It was found that the new system works well for hydrogen decompression under complex conditions. However, a better hydrogen flow inside MSHPRV with smaller turbulent intensity and lower energy consumption has not been realized. Therefore, in this paper, the effects of different structural parameters on the internal flow characteristics of MSHPRV are investigated, with the purpose of achieving a better hydrogen decompression process for FCEV.
Model description Mathematical model Hydrogen is hydrogenated from the hydrogen refueling station to the vehicle gas cylinder, and then it is delivered to the fuel cell stack. The hydrogen transfer process can be approximated taken as a fixed volume, variable mass deflation process. It is assumed that the hydrogen in the gas cylinder has sufficient heat exchange with the outside world, therefore, the hydrogen transfer process is considered to be isothermal. Hydrogen mass flow through the orifice at subsonic flow is described as follows [43]. vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 3 u u 2k kþ1 u k 7 Pi u 2k 6 6 P0 P0 7 Qm ¼ fs ðS0 ; Pi ; P0 ; TÞ ¼ CS0 pffiffiffiffiffiffi u 5 Pi RT tk 1 4 Pi
(1)
Hydrogen mass flow through the orifice at supersonic flow is described as follows. Pi Qm ¼ fc ðS0 ; Pi ; TÞ ¼ CS0 pffiffiffiffiffiffi RT
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 k1 2k 2 kþ1 kþ1
(2)
Where, C refers to flow coefficient of the orifice, S0 refers to throttling area, T refers to absolute temperature of hydrogen, k refers to adiabatic ratio coefficient of hydrogen, Pi refers to inlet pressure of throttling component, P0 refers to outlet pressure of throttling component. Hydrogen continuity equation is described as follows. Qm ¼
dM dt
(3)
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hydrogen pressure reducing process totally, which can contribute to more gradual flow and smaller turbulent intensity. Changes in the throttling areas can in turn changes the hydrogen flow state inside MSHPRV. Fig. 1(b) shows the mesh model of MSHPRV. 1/2 model of MSHPRV is chosen as the computational domain since the flow model is symmetrical. Considering the structural complexity of MSHPRV, meshing is performed using hybrid grid technology. Structured grid is used for inlet, middle and outlet pipelines, whose structures are regular. Unstructured grid is used for multi-stage perforated sleeves, valve core and valve chamber. Fluid medium inside MSHPRV is hydrogen with initial pressure 35 MPa. Pressure inlet and outlet are chosen as the boundary conditions respectively. Outlet pressure of MSHPRV and the previously designed multi-stage muffler is 5 MPa and 0.16 MPa respectively. However, as a component which dominants the hydrogen decompression process, only MSHPRV is considered in this study. Center plane in MSHPRV is set as the symmetry. Steady-state solver and the second-order upwind scheme are adopted. RNG k-ε turbulence model are used since it has a better accuracy in the study of flow with turbulent vortex. Grid-independent verification is necessary to ensure the accuracy of numerical method adopted in this study. Table 1 shows that the mass flow rate in MSHPRV changes with different grid numbers at inlet pressure 3.0 MPa. It can be seen from Table 1 that, the relative error of mass flow rate is 14.1%, 5.0% and 3.2% for grid number 1.18 107, 1.2 107 and 1.22 107 respectively. It can be found that grid number 1.2 107 is a suitable chose. Therefore, it can be used as an optimal solution to ensure efficiency and accuracy of numerical calculation.
Fig. 1 e Physical model of MSHPRV. Where, M refers to hydrogen mass flow in the gas cylinder, Qm refers to hydrogen flux. Pressure dynamic equation is described as follows. 2
3
RT V 7 dM dP 6 2a m ¼6 M7 2 2 4 5 dt dt ðmVÞ V mb M
(4)
Where, P refers to cylinder pressure, T refers to cylinder temperature, R refers to hydrogen gas constant, V refers to hydrogen storage volume, m refers to hydrogen molar mass.
Model verification MSHPRV is a newly designed high pressure reducing valve. Nowadays, it is only conceptually designed for hydrogen FCEV. However, it is already used in some process industries with steam medium. Since the numerical model adopted in MSHPRV for hydrogen flow in FCEV is similar to that for steam flow in process industry, the numerical method is verified by experimental method with steam medium. Table 2 shows the experimental and numerical results of flow rate inside MSHPRV with steam medium. It can be seen
Computational model and boundary conditions Fig. 1(a) shows the three-dimensional model of MSHPRV. It can be seen from Fig. 1(a) that, MSHPRV is mainly composed of inlet valve chamber, multi-stage perforated sleeves, valve core, valve steam, perforated plate, valve body, outlet valve chamber and the others. Working principle of MSHPRV was described in the previous work [2]. It can achieve five-stage
Table 2 e Experimental and numerical results of flow rate inside MSHPRV. L/Lmax qm numerical (t/h) qm experimental (t/h) d(t/h)
0.2 110 114.5 4.5
0.4 120 120.5 0.5
0.6 125 126.03 1.03
0.8 133 131.73 1.27
1.0 140 137.43 2.57
Table 1 e Grid-independent verification. N (106) Flux (kg/s) d(%)
11.6 16 15.2
11.8 24 14.1
12 28.5 5.0
12.2 30 3.2
13 31 3.1
14 31.5 3.1
15 31.8 3.1
16 31.7 3.1
17 31.8 3.1
18 31.7 3.1
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Fig. 2 e MSHPRV with different perforated plate stages.
Fig. 3 e Pressure distribution of symmetry plane with different perforated plate stages (MPa).
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regardless of temperature change. Therefore, it can be indirectly concluded that the numerical model adopted in MSHPRV for hydrogen flow in FCEV is reliable.
Results and discussion In order to obtain a better hydrogen decompression process inside MSHPRV under harsh working conditions as high pressure and large pressure ratio, the effects of different structural parameters on its internal flow characteristics are investigated. Here, the structural parameters of main throttling components, such as different perforated plate stages, perforated plate diameters, multi-stage sleeve diameters, perforated plate radii and pressure ratios, are studied.
Effect of perforated plate stage Fig. 4 e Velocity distribution along Z direction with fourstage perforated plates (m/s).
Table 3 e Velocity and turbulence parameter maximum difference with different perforated plate stages. Stages Dvmax (m/s) Dkmax (m2/s2)(106) Dεmax (m2/s3)(1012) 1 2 3 4
108 104 103 90
2.8 2.2 2.1 1.8
6 6 5.8 5.5
from Table 2 that, the relative mass flow rate error d between experimental and numerical results inside MSHPRV with relative valve opening L/Lmax 0.2, 0.4, 0.6, 0.8 and 1.0 is 4.5 t/h, 0.5 t/h, 1.03 t/h, 1.27 t/h and 2.57 t/h respectively. Therefore, the maximum difference of flow rate between numerical and experimental results is located at L/Lmax 0.2, and the error d is only about 4%. It can be found that compared with experimental data, the numerical results is not much different from it. Temperature difference between steam and hydrogen under high pressure is large. However, energy equation is closed in this study, and attention is only paid to the flow
Different perforated plate stages will cause changes in the throttling areas, which in turn changes the hydrogen flow state. Therefore, the effect of perforated plate stage on hydrogen internal flow inside MSHPRV is investigated in this part. Fig. 2 shows the three-dimensional flow model of MSHPRV with different perforated plate stages. Fig. 3 shows the pressure contours of symmetry plane with different perforated plate stages. It can be seen from Fig. 3 that, taking MSHPRV with one-stage perforated plate as an example, hydrogen pressure reduces from 35 MPa to 10 MPa at the multi-stage perforated sleeves and valve core, and from 10 MPa to 5 MPa at the last-stage perforated plate. Therefore, compared with perforated plate, multi-stage perforated sleeves and valve core play a more important role in hydrogen decompression. Moreover, hydrogen pressure changing gradient mainly reflects at those throttling components, while evenly distributed in the inlet chamber and outlet chamber. Different perforated plate stages do not have a great impact on the hydrogen pressure reduction, because perforated sleeves and valve core take the dominant position. Fig. 4 shows the velocity distribution of MSHPRV along Z direction with four-stage perforated plates. It can be seen from Fig. 4 that, corresponding to the pressure change, the hydrogen
Fig. 5 e MSHPRV with different perforated plate diameters.
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Fig. 6 e Velocity distribution of symmetry plane with different perforated plate diameters (m/s).
velocity changes greatly at the throttling elements. There are two peaks in the curve indicated by the red ellipse box, representing the two sides of the multi-stage sleeves and valve core. Moreover, there are four peaks in the red square box, which correspond to the four-stage perforated plates. This is because as the areas of the throttling elements decrease, the
flow rate increases, and the fluid kinetic energy increases correspondingly. According to the law of energy conservation, the potential energy is reduced, resulting in pressure loss. It can be found that, the maximum velocity at the multi-stage sleeves is around 100 m/s, while about 40 m/s at the perforated plates. Therefore, just as mentioned above, compared
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Effect of perforated plate diameter
Fig. 7 e Turbulent kinetic energy along Z direction with different perforated plate diameters.
with perforated plate, multi-stage perforated sleeves and valve core hold the dominant position in hydrogen throttling process. Table 3 shows the maximum difference of velocity, turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different perforated plate stages. It can be seen from Table 3 that, the maximum difference of velocity inside MSHPRV with perforated plate stage 1, 2, 3 and 4 is 108 m/s, 104 m/s, 103 m/s and 90 m/s respectively. Moreover, the correspondingly maximum difference of turbulent kinetic energy is 2.8 106 m2/s2, 2.2 106 m2/s2, 2.1 106 m2/s2 and 1.8 106 m2/s2. It can be found that with the increase of perforated plate stage, the turbulent flow can be gradually reduced.
Different perforated plate diameters will also change the throttling areas. Fig. 5 shows the three-dimensional flow model of MSHPRV with different perforated plate diameters. The original perforated plate diameter for engineering application is 8 mm, therefore, taking the two values before and after 8 mm with an arithmetic progression. Fig. 6 shows the velocity distribution of symmetry plane with different perforated plate diameters. It can be seen from Fig. 6 that, the maximum velocity inside MSHPRV exists at the multi-stage sleeves or the perforated plate, and compared with perforated plate, the maximum velocity mainly reflects at the perforated sleeves and valve core. Meanwhile, the maximum velocity at the multi-stage perforated sleeves is 108 m/s, 110 m/s, 110 m/s, 116 m/s and 120 m/s for MSHPRV with perforated plate diameter 4 mm, 6 mm, 8 mm, 10 mm and 12 mm respectively. Moreover, with the increasing of perforated plate diameter, the velocity gradient in the valve chamber after multi-stage sleeves and valve core increases correspondingly. However, the velocity gradient in the valve chamber after perforated plate decreases. It can be seen from Fig. 6 that, the turbulent vortex in the valve chamber increases with the increase of perforated plate diameter, while decreases at the outlet pipeline. Changing of perforated plate diameter has a great impact on the hydrogen internal flow inside the whole MSHPRV. Just as mentioned above, multistage perforated sleeves and valve core hold the dominant position in hydrogen throttling process. Therefore, it can be concluded that large perforated plate diameter will result in large hydrogen kinetic energy. Fig. 7 shows the turbulent kinetic energy of MSHPRV along Z direction with different perforated plate diameters. It can be seen from Fig. 7 that, in accordance with Fig. 4, there exist three peaks in the curve, which located at the multi-stage
Fig. 8 e MSHPRV with different sleeve diameters.
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Fig. 9 e Velocity and streamlines distribution of symmetry plane with different sleeve diameters (m/s).
sleeves and perforated plate respectively. Therefore, the turbulent vortex inside MSHPRV is mainly reflected at the throttling components like multi-stage perforated sleeves, valve core and perforated plate. Moreover, the maximum turbulent kinetic energy which marked in black circle box is about 1.05 106 m2/s2, 1.80 106 m2/s2, 2.15 106 m2/s2, 2.25 106 m2/s2 and 2.38 106 m2/s2 for MSHPRV with perforated plate diameter 4 mm, 6 mm, 8 mm, 10 mm and 12 mm respectively. Large perforated plate diameter refers to large turbulent kinetic energy. Therefore, it can be concluded that with the increasing of perforated plate diameter, the turbulent intensity and energy consumption will increase correspondingly.
Effect of sleeve diameter Multi-stage perforated sleeves and valve core hold the dominant position in hydrogen throttling process, therefore, a parametric study on multi-stage perforated sleeves is
conducted. Fig. 8 shows the three-dimensional flow model of MSHPRV with different perforated sleeve diameters. The original multi-stage perforated sleeves diameter for engineering application is 5 mm, therefore, taking the two values before and after 5 mm using an arithmetic progression with a tolerance of 1. Fig. 9 shows the velocity and streamlines distribution of symmetry plane with different sleeve diameters. It is known from the above studies that the maximum hydrogen velocity inside MSHPRV is mainly distributed at the multi-stage perforated sleeves, so it is partially enlarged to study the effect of sleeve diameter on hydrogen flow. It can be seen from Fig. 9 that, with the increasing of sleeve diameter, the maximum velocity distribution range which located at the multi-stage perforated sleeves and valve core increases gradually. At the same time, the velocity gradient in the valve chamber after perforated sleeves and valve core increases correspondingly. Moreover, it can be found from the streamlines distribution that, the vortex becomes larger gradually as
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Fig. 10 e Turbulence parameter maximum difference with different sleeve diameters.
Fig. 11 e Pressure and velocity distributions of symmetry plane with different perforated plate radii (MPa) (m/s).
the sleeve diameter increases. Therefore, it can be concluded that larger multi-stage perforated sleeve diameter relates to larger hydrogen kinetic energy and turbulence vortex. Fig. 10 shows the maximum difference of turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different perforated sleeve diameters. It can be seen from Fig. 10 that, the maximum difference of turbulent kinetic
energy inside MSHPRV with perforated sleeve diameter 3 mm, 4 mm, 5 mm, 6 mm and 7 mm is 1.4 106 m2/s2, 1.5 106 m2/ s2, 1.8 106 m2/s2, 1.8 106 m2/s2 and 2.0 106 m2/s2 respectively. Moreover, the correspondingly maximum difference of turbulent dissipation rate is 5.0 1012 m2/s3, 6.0 1012 m2/s3, 5.6 1012 m2/s3, 7.0 1012 m2/s3 and 8.5 1012 m2/s3. Therefore, it can be concluded that with the
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increase of multi-stage perforated sleeve diameter, the turbulent intensity and energy consumption inside MSHPRV becomes larger.
Effect of perforated plate radius The purpose of chamfering is to change the straight transition to a smooth transition. Compared with straight transition, smooth transition can make the hydrogen flow at the corners to be more uniform and the turbulent intensity to be smaller. Therefore, chamfering can help change the hydrogen flow state inside MSHPRV. The first row of Fig. 11 shows the threedimensional flow model of MSHPRV with different perforated plate radii. Fig. 11 shows the pressure and velocity distributions of symmetry plane with different perforated plate radii. It can be seen from Fig. 11 that, with the increasing of perforated plate radii, the light blue area gradually shrinks and the dark blue area gradually increases, which indicates that the pressure distribution in the valve chamber after the multi-stage perforated sleeves and valve core tends to be uniform. Moreover, the velocity gradient in the valve chamber decreases correspondingly, which can be seen from the third row of Fig. 11 that the light blue area shrinks gradually. In the meantime, the turbulent vortex area decreases with the increase of perforated plate radius. Therefore, it can be concluded that larger perforated plate radii contribute to a better hydrogen flow state inside MSHPRV. Fig. 12 shows the maximum difference of turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different perforated sleeve radii. It can be seen from Fig. 12 that, the maximum difference of turbulent kinetic energy inside MSHPRV with perforated plate radius 0 mm, 0.5 mm, 1 mm, 1.5 mm and 2 mm is 2.8 106 m2/s2, 2.2 106 m2/s2, 2.1 106 m2/s2, 1.9 106 m2/s2 and 1.8 106 m2/s2 respectively. Therefore, it can be found that larger perforated plate radius relates to smaller turbulent intensity inside MSHPRV. That is because chamfering can help change the hydrogen flow state inside MSHPRV. Moreover, the correspondingly maximum difference of turbulent dissipation rate is 6.0 1012 m2/s3,
Fig. 12 e Turbulence parameter maximum difference with different perforated plate radii.
Table 4 e Velocity and turbulence parameter maximum difference with different pressure ratios. p1/p2 4 5 6 7 8
Dvmax (m/s)
Dkmax (m2/s2)(106)
Dεmax (m2/s3)(1012)
75 90 90 100 120
1 1.4 1.8 1.8 2.8
1.8 2.8 3.5 6 8
4.0 1012 m2/s3, 3.8 1012 m2/s3, 3.4 1012 m2/s3 and 3.1 1012 m2/s3. It again shows that the perforated plate radius is beneficial for improving the internal turbulent flow field inside MSHPRV. Therefore, it can be concluded that with the increase of perforated plate radius, the turbulent intensity and energy consumption inside MSHPRV decreases correspondingly.
Effect of pressure ratio Pressure ratio is closely related to the boundary conditions of hydrogen flow inside MSHPRV. Therefore, the effect of pressure ratio on the flow characteristics is investigated. Pressure distribution of MSHPRV along Z direction with different pressure ratios is investigated. There exist three peaks in the curves of pressure decompression with different pressure ratios, which located at the two sides of multi-stage sleeves and perforated plate respectively. The change of the flow area at the throttling components changes the flow rate and the kinetic energy of hydrogen, which in turn causes a certain pressure loss, so the pressure gradient is mainly reflected at such throttling components as multi-stage perforated sleeves, valve core and perforated plate. Moreover, the maximum pressure gradient at the multi-stage perforated sleeves and valve core is 9.8 MPa, 14.7 MPa, 19.8 MPa, 24.6 MPa and 29.5 MPa for MSHPRV with pressure ratio 4, 5, 6, 7 and 8 respectively. However, the maximum pressure gradient at the perforated plate is 5.2 MPa, 5.3 MPa, 5.2 MPa, 5.4 MPa and 5.5 MPa correspondingly. Therefore, compared with perforated plate, hydrogen pressure and velocity gradients are mainly reflected at the multi-stage perforated sleeves. Table 4 shows the maximum difference of velocity, turbulent kinetic energy and turbulent dissipation rate inside MSHPRV with different pressure ratios. It can be seen from Table 4 that, the maximum difference of velocity inside MSHPRV with pressure ratio 4, 5, 6, 7 and 8 is 75 m/s, 90 m/s, 90 m/s, 100 m/s and 120 m/s respectively. Larger pressure ratio relates to larger velocity difference. Moreover, the correspondingly maximum difference of turbulent kinetic energy is 1 106 m2/s2, 1.4 106 m2/s2, 1.8 106 m2/s2, 1.8 106 m2/s2 and 2.8 106 m2/s2. Therefore, the turbulent kinetic energy increases with the increase of pressure ratio. The maximum difference of turbulent dissipation rate is 1.8 1012 m2/s3, 2.8 1012 m2/s3, 3.5 1012 m2/s3, 6 1012 m2/s3 and 8 1012 m2/s3. It can be found that with the increase of pressure ratio, the maximum difference of velocity and turbulent parameters all increase accordingly. Therefore, it can be concluded that larger pressure ratio relates to larger turbulent flow and energy consumption inside MSHPRV.
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 4 4 ( 2 0 1 9 ) 3 1 2 6 3 e3 1 2 7 4
Conclusions A multi-stage high pressure reducing valve (MSHPRV) for hydrogen decompression in FCEV is proposed. The effects of different structural parameters on internal flow characteristics of MSHPRV are investigated, with the purpose of achieving a better hydrogen decompression process for FCEV. Results show that compared with perforated plate, multi-stage perforated sleeves and valve core hold the dominant position in hydrogen throttling process. Different perforated plate stages do not have a great impact on the hydrogen pressure reduction. However, with the increase of perforated plate stage, the turbulent flow can be gradually reduced. Secondly, large perforated plate diameter will result in large hydrogen kinetic energy. With the increasing of perforated plate diameter, the turbulent intensity and energy consumption will increase correspondingly. Thirdly, larger multi-stage perforated sleeve diameter relates to larger hydrogen kinetic energy, turbulence vortex and energy consumption. Meanwhile, larger perforated plate radii contribute to a better hydrogen flow state inside MSHPRV. With the increase of perforated plate radius, the turbulent intensity and energy consumption inside MSHPRV decreases correspondingly. Finally, larger pressure ratio relates to larger turbulent flow and energy consumption inside MSHPRV. In conclusion, better hydrogen flow characteristics inside MSHPRV for hydrogen FCEV are achieved with larger perforated plate stage, smaller perforated plate and sleeve diameter, larger perforated plate radii and smaller pressure ratio. In the meanwhile, the appropriate structural parameters should be chosen by considering the processing difficulty and manufacturing cost. This research work can provide some technical supports for achieving hydrogen decompression in FCEV when facing harsh working conditions, or to whom dealing with energy conversion and better flow state.
Acknowledgements This work is supported by the National Natural Science Foundation of China through Grant No. 51875514, the Key project of Natural Science Foundation of Zhejiang Province, China through Grant No. LZ17E050002, and the Zhejiang Key Research & Development Project, grant number 2019C01025.
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