Experimental Study on Boiling Vaporization of Liquid Hydrogen in Nonspreading Pool

Abstract

Research on the boiling vaporization process of liquid hydrogen spilled on solid ground is very important for the safety risk assessment of liquid hydrogen. Since the main source of the heat flux in the vaporization process comes from the ground, the heat flux from the ground into the liquid pool should be studied in-depth. In this paper, the boiling vaporization process of liquid hydrogen on the surface of concrete is studied. The analysis of the boiling process of a liquid pool is conducted by utilizing the boiling curve and historical temperature data collected in close proximity to the surface of the concrete. It was found that the boiling regime of a liquid hydrogen pool on the concrete surface presents non-uniformity, and the film boiling of liquid hydrogen on the concrete surface ended earlier than the results calculated by boiling regime correlations. When the measured temperature in the experiment indicates a transition from film boiling to the transition boiling, the temperature difference between the thermocouple temperature measured at a depth of 2 mm and the boiling point of liquid hydrogen is 130 K higher than the predicted superheat of the minimum heat flux (MHF). In the later stage of the experiment, the average relative error between the experimental value of the vaporization rate and the predicted value of the model is 7.48%. This research advances the understanding of heat transfer between concrete ground and a liquid hydrogen pool. In addition, the experimental data obtained in this study contributes to improving the source term model for safety analysis of liquid hydrogen spills.

1. Introduction

With the industrial revolution and the development of hydrogen energy economy, the utilization of liquid hydrogen has gradually expanded from the field of aerospace to the field of civil industry, transportation, and energy. Various storage methods are available for hydrogen, some of the primary ones being liquid hydrogen, high-pressure gas, and metal hydride storage [1]. As the density of hydrogen gas is only 0.083 kg/m³ under normal temperature and pressure, it is necessary to adopt measures to increase its energy density. Under normal atmosphere, the unit volume energy density of hydrogen gas, liquid hydrogen, and 30 MPa hydrogen gas is 11 MJ/m³, 8640 MJ/m³ and 2160 MJ/m³, respectively. Thus, low-temperature liquid hydrogen storage is one of the effective ways to improve energy density. Once a huge amount of liquid hydrogen is accidentally released, it is easy for a pool of liquid hydrogen to form, and a pool fire and combustion explosion can occur, causing great damage to human beings, the environment, and property. Therefore, research on the safety of liquid hydrogen is the basic premise of its use. The current research on accidental leakage vaporization of cryogenic liquids shows that the ground is the main heat source for liquid vaporization, in which heat convection and radiation can be ignored. Therefore, it is very necessary to further study the heat flux from the ground into the liquid pool.

The research on spill and dispersion of liquid hydrogen is mainly carried out from two directions: experimental and numerical simulation. The Health and Safety Executive (HSE) [2] and Health and Safety Laboratory (HSL) [3] carried out different liquid hydrogen spill experiments respectively. HSE conducted a series of leakage of liquid hydrogen experiments under different diameter nozzles and pressures. The hydrogen concentration sensor and temperature sensor were used to measure the properties of the mixed cloud. The experiment conducted by HSL carried out the equidistant arrangement of thermocouples on the ground, measured the radius of the liquid pool, and conducted the ignition experiment to measure the thermal radiation of the hydrogen cloud after ignition and the speed and range of the flame. Shirai et al. [4] studied the heat transfer between the pool of liquid hydrogen and a horizontal plate heater. The flat plate heater was 10 mm in width, 100 mm in length, and 0.1 mm in thickness, and the critical heat flux (CHF) were investigated from atmospheric to 1.1 MPa. Verfondern et al. [5,6] explored the radius expansion law of liquid hydrogen pools through a series of leakage experiments on water and aluminum surfaces. The research results showed that under the leakage rates of 0.005 m³/s and 0.006 m³/s, the maximum radius of liquid hydrogen pool on water surface and aluminum surface was 0.8 m and 0.5 m, respectively. LAuV prediction model based on shallow water equation is used to predict the radius of the liquid pool of various cryogenic liquids. Nguyen et al. [7] carried out a small experiment on nonspreading boiling vaporization of cryogenic liquid, conducted an experimental study on the evaporation process of cryogenic liquid on concrete ground, and investigated the influence of the depth of liquid pool on evaporation. Takeno et al. [8] carried out experiments in order to calculate the heat flux of liquid hydrogen and liquid oxygen that flowed into the nonspreading pool through mass loss rate. Additionally, the researchers conducted evaporation experiments on substrates made of both dry and wet sand and concluded that the mechanisms of evaporation for wet and dry sand were identical due to extremely low temperatures experienced with liquid hydrogen.

Boiling and vaporization of cryogenic liquid are very complex gas–liquid two-phase flow processes due to the mechanism of bubble dynamics, flow, phase transformation heat, etc. Many scholars carried out numerical simulation studies on this process. Olewski et al. [9] established a computational fluid dynamics (CFD) model based on Rayleigh–Taylor instability and the interface between gas and liquid phases was tracked using the Volume of Fluid (VOF) method to simulate the film boiling of liquid nitrogen, so as to estimate the rate of vapor generation in case of accidental leakage of cryogenic liquid. Liu et al. [10] used CFD to study the ground thermal conductivity, one of the important factors affecting the evaporation rate of a liquid pool. The results show that thermal conductivity has an important influence on the heat transfer and heat flow of a liquid pool. Increasing the thermal conductivity will increase the evaporation rate of a liquid pool on the ground, and the diffusion process of the evaporating gas cloud will be accelerated, but it has little influence on the diffusion range of gas cloud downstream. Dixit [11] numerically simulated the film boiling process of liquid nitrogen (LN2) on the surface of water. On the interface capture method based on VOF, the mass transfer model based on Hertz–Knudsen–Schrage was used to study the influence of single mode and multiple mode on film boiling, and the temperature and velocity fields of three phases were obtained. The results show that the vaporization rate of LN2 increases with the increase of water temperature.

The boiling vaporization of a liquid hydrogen pool is the initial stage of an accidental leakage of liquid hydrogen that has a significant impact on the subsequent diffusion of gas clouds. As described above, current experiments on liquid hydrogen leakage mainly focus on researching the characteristics of gas cloud diffusion after liquid pool evaporation. There are few experimental studies on the vaporization behavior of liquid hydrogen pools, and there is a lack of related experimental data on leakage source terms. This has resulted in current numerical simulations being based mostly on other cryogenic liquids such as LN₂, LNG, and LO₂. The simulation of pool boiling of liquid hydrogen is very scarce. Therefore, this paper conducts experimental research on the boiling vaporization of a liquid hydrogen pool. The present results can be independently compared to those based on the perfect thermal contact model (PTCM) and boiling regime correlations (BRCs). The vaporization rate of the liquid pool over the solid surface and the history of the internal temperature of the solid surface were studied, and the accuracy of predicting of the heat flux of the liquid pool was evaluated.

2. Experimental Setup

The experimental equipment mainly includes mass measurement system, temperature measurement system, liquid hydrogen charging system, and data acquisition system. The schematic diagram of liquid hydrogen leakage experimental equipment is shown in Figure 1. The mass measurement system is used to measure the change of liquid pool mass over time, and the temperature acquisition system is mainly used to measure the temperature of different positions inside the solid surface. At the same time, temperature sensors are placed at different heights on the wall surface to monitor the variation of liquid hydrogen level.

The liquid hydrogen was poured onto a concrete pad of area size 400 by 400 mm and depth of 150 mm. The whole experimental apparatus is wrapped with 100 mm thick polystyrene foam around and on the bottom, which plays the role of heat insulation between the stainless-steel wall and the concrete wall inside the box and the surrounding environment.

The T-type thermocouples with a measuring rage of 73 K–323 K and an accuracy of 0.5 K (TT-T-24-SLE from Omega) were installed at varying depths in the concrete to capture the transient temperature changes in the ground. The depth of the thermocouples inside the concrete is 2 mm, 20 mm, 50 mm, 90 mm, and 150 mm respectively and the position of the thermocouples is shown in Figure 2. Two thermocouples were mounted at the same depth. The experimental data obtained from the Agilent data acquisition instrument is acquired by the Agilent BenchLink Data Logger 3 software, which can be used to select thermocouple types, set acquisition frequencies, and perform local file storage. The digital balance is WT1503LP with a maximum load of 200 kg and resolution of 1 g, which is used to measure the mass change of the pool of liquid hydrogen and its supporting weight measurement software collects mass data every 0.5 s.

An empty container which did not include the concrete ground was used to analyze the heat flux into the liquid pool from the container. The thickness of the wall and bottom of the container is the same, both 2 mm and the container.

The photographs of the vaporization platform and experimental process are shown in Figure 3 and the liquid hydrogen was discharged into the containment at a flow rate that was not measured, so the rate of liquid hydrogen evaporation was analyzed only after the spill was stopped (lasting about 4 min in concrete substrate experiment and 2 min in empty container experiment).

A total of 4 groups of experiments were conducted, including 3 groups with concrete substrates and 1 group without substrates.

3. Results and Discussion

3.1. Two Methods to Estimate the Heat Flux

When discussing the numerical simulation of a coupled set of differential equations for the evaporation of cryogenic liquid, the boundary conditions for heat transfer from the liquid hydrogen pool are usually considered. Two boundary conditions are typically used to determine the heat flux from the ground to the liquid pool: (i) the specified heat flux and (ii) the specified temperature. The boundary condition is developed from predictive correlation boiling heat transfer regimes for a specified heat flux. The boiling regime depends on the temperature difference between liquid hydrogen and the surface. The assumption for the specified temperature boundary condition is perfect thermal contact between the surface and the liquid [12]. The boiling curve calculated by the boiling regime correlations is shown in Figure 4, which is based on the Kutateladze correlation for nucleate boiling, the Breen and Westwater correlation for film boiling, and the interpolation method is used for calculating the transition boiling. The correlations and critical values are shown in

Table 1. Boiling correlations.

Kutateladze correlation [13]	$Nu=3.25\times {10}^{-4}{\left(\frac{{\rho }_{\mathrm{l}}q{C}_{\mathrm{p},\mathrm{l}}{l}_{\mathrm{c}}}{{\rho }_{\mathrm{v}}{h}_{\mathrm{fg}}{k}_{\mathrm{l}}}\right)}^{3/5}\cdot G{r}^{1/8}{\left(\frac{P}{\sqrt{\sigma g{\rho }_{\mathrm{l}}}}\right)}^{7/10}$
Breen and Westwater correlation [14]	$q_{\mathrm{f}}=0.37{\left(\frac{\sigma }{g\left({\rho }_{\mathrm{l}}-{\rho }_{\mathrm{v}}\right)}\right)}^{-0.125}{\left(\frac{{\mu }_{\mathrm{v}}}{k_{\mathrm{v}}^3{\rho }_{\mathrm{v}}\left({\rho }_{\mathrm{l}}-{\rho }_{\mathrm{v}}\right)g{h}_{\mathrm{fg}}{\left(1+0.34C_{\mathrm{p},\mathrm{v}}\Delta T/h_{\mathrm{fg}}\right)}^2}\right)}^{-0.25}\Delta T^{0.75}$
The Critical heat flux [13]	$q_{\mathrm{CHF}}=0.16h_{\mathrm{fg}}{\rho }_{\mathrm{v}}^{0.5}{\left[\sigma g\left({\rho }_{\mathrm{l}}-{\rho }_{\mathrm{v}}\right)\right]}^{0.25}$
The Minimum heat flux [15]	$q_{\mathrm{MHF}}=0.16h_{\mathrm{fg}}{\rho }_{\mathrm{v}}{\left[\frac{\sigma g\left({\rho }_{\mathrm{l}}-{\rho }_{\mathrm{v}}\right)}{{\left({\rho }_{\mathrm{l}}+{\rho }_{\mathrm{v}}\right)}^2}\right]}^{0.25}$
Transition boiling [16]	$q_{\mathrm{t}}=q_{\mathrm{CHF}}{\left(1-\frac{\Delta T-\Delta T_{\mathrm{CHF}}}{\Delta T_{\mathrm{MHF}}-\Delta T_{\mathrm{CHF}}}\right)}^7+q_{\mathrm{MHF}}\left[1-{\left(1-\frac{\Delta T-\Delta T_{\mathrm{CHF}}}{\Delta T_{\mathrm{MHF}}-\Delta T_{\mathrm{CHF}}}\right)}^7\right]$

. The relevant physical parameters of hydrogen at atmospheric pressure were obtained by Reference Fluid Thermodynamic and Transport Properties Database (REFPROP) Version 9.1 computer program of NIST, as shown in

Table 2. Thermal-physical properties of hydrogen at atmospheric pressure.

${\mathit{T}}_{\mathrm{b}},K$	${\rho }_{\mathrm{l}},kg/m^3$	${\rho }_{\mathrm{v}},kg/m^3$	${\mathit{C}}_{\mathrm{p},\mathrm{l}},J/\left(kg\cdot K\right)$	${\mathit{C}}_{\mathrm{p},\mathrm{v}},J/\left(kg\cdot K\right)$	${\mathit{h}}_{\mathrm{fg}},J/kg$	${\mathit{k}}_{\mathrm{l}},W/\left(m\cdot K\right)$	${\mathit{k}}_{\mathrm{v}},W/\left(m\cdot K\right)$	${\nu }_{\mathrm{l}},m^2/s$	${\nu }_{\mathrm{v}},m^2/s$	$\sigma , N/m$
20.324	70.899	1.3164	9747.6	12,018	448,910	0.103	0.017	1.91 × 10−7	7.55 × 10−7	1.91 × 10−3

.

The boiling curve of liquid hydrogen, as shown in Figure 3, represents the change of heat flux with the difference between the saturation temperature of liquid hydrogen and the solid surface temperature. In the experiments of this work, the temperature difference between the saturation temperature of liquid hydrogen and surface temperature of concrete is about 260 K. According to the boiling curve, the liquid pool is in the state of film boiling at the beginning of the pool boiling, and as the temperature of the concrete surface decreases, the boiling regime will change to the transition boiling and then the nucleate boiling. The heat fluxes at critical heat flux and minimum heat flux are in the order of $7.2\times {10}^4 {\mathrm{W}/\mathrm{m}}^2$ and $1.3\times {10}^3 {\mathrm{W}/\mathrm{m}}^2$.

The PTC model assumes that the cryogenic liquid is in direct contact with the ground and that the temperature of the wall is instantaneously equal to the boiling point temperature of the cryogenic liquid. Wall heat flux can then be calculated by the following equation [17]:

$$q{|}_{\mathrm{x}=0}=\frac{k}{{\left(\pi \alpha \right)}^{0.5}}\left(T_{\mathrm{i}}-T_{\mathrm{sat}}\right)t^{-0.5}=A\cdot t^{-0.5}$$

(1)

In this equation, $q{|}_{\mathrm{x}=0}$ is the wall heat flux, $k$ and $\alpha$ are the thermal conductivity and thermal diffusivity of the concrete, respectively, $T_{\mathrm{i}}$ is the initial temperature of the ground, $T_{\mathrm{sat}}$ is the boiling point of the liquid hydrogen, $A$ is the proportionality coefficient, and $t$ is the time.

3.2. Calculation of Thermal Properties of the Concrete

The thermal-physical parameters associated with concrete in the case of liquid hydrogen spills are determined by applying the theory of unsteady one-dimensional heat transfer in a semi-infinite ground. The thermal diffusivity of concrete is calculated using the following equation:

$$\begin{array}{c}\hfill \frac{T-T_{\mathrm{sat}}}{T_{\mathrm{i}}-T_{\mathrm{sat}}}=erf\left(\frac{x}{2\sqrt{\alpha t}}\right)=erf\left(M\right)\hfill \\ \hfill \alpha =\frac{{\left(x/2M\right)}^2}{t}\hfill \end{array}$$

(2)

where $T$ is ground temperature and $x$ is the distance from the ground surface.

The concrete thermal diffusivity as a function of temperature is described by the Equation (2), as depicted in Figure 5. The result indicates an increase in thermal diffusivity as the temperature decreases.

The thermal diffusivity of concrete at a specific temperature is determined using Equation (2). Subsequently, the thermal conductivity of concrete at that temperature is estimated by utilizing the heat flux at the corresponding instant. The thermal properties of the concrete substrate are reported in

Table 3. Thermal-physical properties of concrete.

$\mathit{k}/\left(W\cdot m^{-1}\cdot K^{-1}\right)$	$\alpha /\left(m^2\cdot s^{-1}\right)$	$\rho /\left(kg\cdot m^{-3}\right)$
0.88	1.5775 × 10−7	2096

.

3.3. Measured Temperature and the Predictions from the PTC

The experimentally measured temperature profiles are shown in Figure 6. Temperature data acquired from two thermocouples positioned adjacent to the concrete surface, namely TC1-2 mm and TC2-2 mm, were analyzed and compared. The temperature curve was segmented into three parts, and the boiling regime transition process was observed from the data acquired by these two thermocouples. The temperature profiles of these thermocouples were similar, and both demonstrated a trend indicating the boiling process. The temperature recorded by TC2-2 mm indicated that the liquid pool was in film boiling regime until 269 s. In this stage, the heat transfer between the concrete surface and liquid hydrogen was impeded due to the presence of a vapor film. As the surface temperature decreased, the vapor film collapsed, and transition boiling occurred between 269 s and 326 s. During the transition boiling regime, considerable cooling of the surface was observed. Subsequently, in the nucleate boiling regime, the temperature began leveling off. While the two thermocouples (TC1-2 mm and TC2-2 mm) displayed similar temperature trends, their boiling regime transition times differed. Specifically, TC1-2 mm demonstrated a transition from film to transition boiling earlier than TC2-2 mm, approximately 8 s in advance. The thermocouples positioned close to the concrete surface were extremely responsive to the local surface heat flux, leading to variations in the temperature data measured. Despite their identical depth, the boiling regimes experienced by the concrete surface were not uniform, resulting in differences in the readings obtained by distinct thermocouples. Localized regions of the concrete surface exhibited a collapse in the vapor film, resulting in the generation of vapor core and corresponding bubbles. These bubbles may have contributed to variations in measurement data. The thermocouple TC2-20 mm located at the apparatus’ center at a depth of 20 mm recorded a rapid temperature drop due to the freezing and cracking phenomenon towards the very low temperature above the right side of the concrete surface. Consequently, liquid hydrogen flowed directly into the concrete interior. The temperature measured by the thermocouple located at a depth of 150 mm remained constant throughout the experiment, indicating that the concrete could be analyzed as a semi-infinite solid during the experiment.

By solving the heat transfer equation with the boundary conditions of the PTC model, the prediction of the PTC for the temperature at different depths inside the solid can be obtained as shown in Figure 7.

The results demonstrate that the PTC curve deviates significantly from the experimental values within the first 260 s as the local vapor film begins to collapse. Notably, the PTC model underestimates the temperature at depths of 20 mm and 50 mm during the early stages, but in the later stages, overestimates are observed. Nevertheless, the PTC model agrees with experimental data at greater depths, and near the surface, it tends to underestimate the temperature compared to the actual situation.

An additional experiment was conducted at a concrete surface temperature of 176 K. The temperature distribution inside the experimental concrete is illustrated in Figure 8. The temperature data at a depth of 2 mm indicate that at this initial temperature, the boiling behavior of liquid hydrogen experienced a shorter film boiling process compared to the surface temperature of concrete of Run1, and after 15 s of pool formation, the liquid pool starts to enter transition boiling and enters into a gentle nucleate boiling regime at 43 s.

3.4. Calculation of Container Heat Flux and Concrete Heat Flux

To verify the thermal insulation effect of the experimental apparatus and to quantitatively determine the heat flux from the container, a set of experiments with an empty container that lacked any substrate were conducted first, followed by a set of experiments with a concrete substrate. The mass of the experimental liquid pool versus time is illustrated in Figure 9 and the rate of mass reduction of the liquid pool without a substrate is much smaller than the rate of mass reduction of the liquid pool with concrete. It was assumed that the heat flux from the container wall was equal to the heat flux from the bottom of the pool, based on the presumption that their thickness was the same. The heat flux into the pool from the container was calculated by the following equation:

$$q_{\mathrm{c}}=\frac{m_{\mathrm{evap}}{h}_{\mathrm{fg}}}{A}=\frac{m_{\mathrm{evap}}{h}_{\mathrm{fg}}}{A_{\mathrm{wall}}+L^2}=\frac{m_{\mathrm{evap}}{h}_{\mathrm{fg}}}{10m/\rho +L^2}$$

(3)

where $q_{\mathrm{c}}$ is the heat flux into the liquid pool from the container body, $m_{\mathrm{evap}}$ is the mass loss rate of the liquid pool, $h_{\mathrm{fg}}$ is the latent heat of vaporization, A is the contact area between the liquid and container, $A_{\mathrm{wall}}$ is the contact area between the liquid and container wall, $A_{\mathrm{wall}}=10m/\rho$, $m$ is the pool mass, $\rho$ is the density of liquid hydrogen, and $L$ is the length of container.

By calculating the heat flux from the container to the liquid pool, the heat flux from the concrete into the liquid pool can be calculated by Equation (4):

$$q=\frac{m_{\mathrm{evap}}{h}_{\mathrm{fg}}-A_{\mathrm{wall}}{q}_{\mathrm{c}}}{L^2}.$$

(4)

The heat flux from the concrete substrate and the heat flux from the container are calculated in Figure 10 through Equations (3) and (4). When boiling liquid hydrogen on the concrete, the heat flux from the container wall was small compared to the heat flux from the concrete substrate. The heat flux from the container can be negligible and the vaporization velocity is based on the heat flux from the concrete ground only.

To exclude the influence of liquid hydrogen injection in the early stage, the predicted vaporization rate of PTC was compared with the experimental results when the mass of the liquid pool started to decrease, and the comparison between the predicted and experimental values is shown in Figure 11. The maximum relative error was 11.2% and the average relative error was 7.48%. The main reason for this phenomenon is that in the middle and late stages of vaporization, the temperature of the concrete wall has been cooled, and the superheat is smaller than the initial leakage moment. The operating conditions at this point are close to the PTC assumptions, so its prediction of the vaporization rate is in better agreement with the experimental values. The results are shown in Equation (5):

$$E=\frac{q{|}_{\mathrm{x}=0}}{\rho h_{\mathrm{fg}}}=\frac{1}{\rho h_{\mathrm{fg}}}\frac{k\left(T_{\mathrm{i}}-T_{\mathrm{sat}}\right)}{\sqrt{\pi \alpha t}}.$$

(5)

To verify the boiling correlation equations against experimental results, as shown in Figure 12, the experimental curve was obtained based on the mass loss of the liquid pool. It can be seen from the figure that the overall experimental curve shifts to the right relative to the boiling curve calculated by the correlation equation. This indicates that, on the concrete substrate, the liquid pool transforms from film boiling to transition boiling and nucleate boiling earlier than predicted by the boiling correlation equation. There are two reasons for this phenomenon: on the one hand, surfaces with poor thermal conductivity experience a more rapid temperature decrease during the cooling process, causing local zones to reach the MHF point first, thereby hastening the termination of film boiling. On the other hand, the wettability and porousness of the boiling surface enhance the possibility of solid–liquid contact and nucleation capacity of bubbles once contact is established, greatly deteriorating the stability of the vapor layer and further accelerating the film boiling to the transition boiling. Additionally, the heat flux predicted by the correlation curve at the same degree of superheat is higher than the experimental value. There exists a considerable difference between the boiling curve obtained from the correlation equation and the experimental results, mainly due to the relatively small superheat of the MHF and CHF points calculated by the correlation equation. The superheat of MHF and CHF points calculated theoretically are 5.27 K and 2.85 K, respectively. Whereas, the surface superheat of the liquid hydrogen and concrete in the experimental process is at least 60.81 K and the superheat of MHF is 136 K. However, from the trend of the thermocouple temperature data before reaching the minimum superheat, it can be observed that the boiling pattern of the liquid pool has already transformed from film boiling to transition boiling and nucleate boiling. The temperature difference between the thermocouple temperature measured at a depth of 2mm and boiling point of liquid hydrogen in the experiment is approximately 130 K higher than the BRC’s prediction. Li et al. [18] conducted a series of boiling experiments using liquid nitrogen on the surfaces of rocks and copper and the results demonstrated that the roughness of different solid surfaces can introduce additional vaporization cores, causing an early transition of the boiling regime.

4. Conclusions

In this paper, the vaporization process of a nonspreading pool of liquid hydrogen on a concrete surface is investigated and the impact of concrete surface temperature on the boiling regime of liquid hydrogen is analyzed. Based on the temperature data near the concrete surface, it can be inferred that the liquid pool undergoes the process of film boiling formation, transition boiling, and nucleate boiling. However, the boiling behavior of the liquid hydrogen during the experiment is non-uniform, which prevents the formation of a stable vapor film. Simultaneous occurrences of film boiling, transition boiling, and nucleate boiling are observed. The experiments revealed that the surface of concrete structures affects the boiling heat transfer characteristics of liquid hydrogen and hastens the termination of film boiling which means that the temperature difference measured by experiment when film boiling changes to the transition boiling on the concrete surface has increased by 130 K compared to the predicted value. As for the validation of the PTC model, due to the low thermal conductivity of concrete, the PTC predicts the temperature distribution at deeper depths more accurately, but there is a large error in the temperature prediction at the concrete surface due to its model characteristics. Future studies will mainly analyze the impact of other substrates on the boiling and vaporization rates of liquid hydrogen as well as the relationship between vaporization rate and diffusion behavior of gas cloud.