Liquid hydrogen releases show dense gas behavior

international journal of hydrogen energy xxx (xxxx) xxx

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Liquid hydrogen releases show dense gas behavior Olav Roald Hansen Lloyd’s Register - Energy, Bergen, Norway

highlights
The paper describes under what conditions LH2 releases show dense gas behavior.
Hazard distances for LH2 bunkering extends further than for LNG.
For releases indoor the potential for strong explosions is much higher for hydrogen than LNG.
Hydrogen fueled vessels must be designed to higher safety standard than LNG vessels.

article info

abstract

Article history:

The number of maritime initiatives with hydrogen as alternative fuel is increasing. While

Received 15 November 2018

most of the early projects aim at using compressed hydrogen the use of liquid hydrogen

Received in revised form

(LH2) is more practical and is expected to become more attractive for implementation on

31 August 2019

larger vessels due to more efficient storage, bunkering and handling of the fuel. In the

Accepted 6 September 2019

industry there seems to be some confusion regarding the behavior of LH2 releases, in

Available online xxx

particular whether a release into air will behave like a dense gas or a buoyant gas. The understanding of this aspect is critical to optimize design with regard to safety. This article

Keywords:

will explain the expected behavior of LH2 releases and discuss expected hazard distances

Liquid hydrogen

from LH2 releases relative to gaseous hydrogen releases and LNG. Some other safety

dispersion

concerns of LH2, like indoor releases, releases from vent masts, potential BLEVEs and RPTs

explosion safety

are also discussed. The article explains why a higher safety standard may be required

dispersion explosion safety

when designing hydrogen fueled vessels than for existing LNG fueled vessels.

bunkering

© 2019 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved.

IGF-code

Introduction In April 2018 the International Maritime Organization (IMO) agreed ambitions to reduce greenhouse gas (GHG) emissions from shipping by 50% in 2050 relative to 2008 levels, see Ref. [20]. This adds to previously agreed ambitions to significantly cap sulphur emissions by 2020. Independently of the IMO decisions there has been a drive to cut GHG emissions from domestic transportation in many countries. In 2018 in Norway 60% of all new cars registered were chargeable, more than half of these (31%) were fully electric [23]. In the main

cities Bergen and Oslo the electric car fraction is highest among cities in the world [17]. Strong economic incentives are main explanation to this. Across Norway there are more than 100 car ferry connections in operation, and an assessment [4] concluded that 84 ferries on 60 connections could be run fully electric. By 2018 electric car ferries are introduced partly on two connections, by 2021/22 the road authorities expect 70 electric ferries in operation. Battery technology has a range limitation due to weight and time to charge. Hydrogen and fuel cell technology may be used to extend the range of electric ferries. In Norway the road authorities have decided to

E-mail address: [email protected]. https://doi.org/10.1016/j.ijhydene.2019.09.060 0360-3199/© 2019 Hydrogen Energy Publications LLC. Published by Elsevier Ltd. All rights reserved. Please cite this article as: Hansen OR, Liquid hydrogen releases show dense gas behavior, International Journal of Hydrogen Energy, https://doi.org/10.1016/j.ijhydene.2019.09.060

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invest in a hybrid hydrogen-electric car ferry (80 cars). The contract to build and operate the ferry for 10 years was awarded end of 2018, and the ferry will be in operation 2021. There are also numerous fast passenger ferries in Norway for which zero-emission will be requirement in new 10 year contracts. For many of these the most viable zero-emission concepts seem to be based on using hydrogen as an energy carrier. Greenhouse gas emission targets are far from the only incentives for emission free shipping. Local air pollution from cruise ships and ferries in narrow fjords is another. In 2018 the Norwegian parliament adopted a resolution to ensure emission free maritime transport by cruise ships and ferries in its world heritage fjords as soon as technically possible, and no later than by 2026 [22]. Zero emission technologies which generally have electric propulsion also tend to emit significantly less noise than conventional fuel alternatives. As a consequence, hydrogen based zero emission vessels are being developed for the commercial maritime transportation infrastructure in Norway. Hydrogen can be stored as compressed gas (e.g. 200e350 bar), cryogenic liquid (20e30 K), in metal hydrides or chemically bound in liquids (e.g. methanol, ammonia, toluene/LOHC or other substances). All of these have a significantly lower energy density per volume than conventional fuels (e.g. diesel), see estimates in Table 1, and there may be further logistics challenges to be handled. For current zero emission fast passenger and car ferry initiatives compressed hydrogen or LH2 seem to be the most relevant alternatives for storage, partly due to requirements on how the hydrogen shall be produced. The use of compressed hydrogen gas is less technically demanding and will be the preferred solution when practical. Using LH2 storage may give a more efficient bunkering process and a higher energy density, and may be the preferred solution for vessels with high power consumption. Currently there are only a few countries in Europe in which there is LH2 production, Norway is not among these. Initiatives to build production facilities for LH2 in Norway exist and local production will likely be established once there is a demand from the maritime industry. Being the first element of the periodic system, hydrogen has extreme properties in many ways. Some of these properties are important for the safety. In Table 2 selected properties are compared to those of methane.

Table 2 e Comparison of selected hydrogen and methane properties (various sources). Properties

Hydrogen

Molecular weight Density at 0  C Speed of sound Boiling point Liquid density Flammable range in air Stoichiometric concentration in air Minimum ignition energy (MIE) Maximum laminar burning velocity

Methane

2 g/mol 16 g/mol 0.090 kg/m3 0.72 kg/m3 1270 m/s 450 m/s 20.4 K (253  C) 111 K (162  C) 420 kg/m3 70.8 kg/m3 4%e75% 5%e14% 29.6% 9.5% 0.017 mJ 2.7 m/s

0.29 mJ 0.4 m/s

For explosion safety assessments the low minimum ignition energy and the high laminar burning velocity can be important. Still, these apply for hydrogen concentrations around and above stoichiometry (29.6% in air), for concentrations below 10% in air hydrogen is less reactive than methane. For unconfined gas release scenarios two other parameters from Table 2 are therefore very important. The high speed of sound ensures a quick and efficient mixing and dilution with air from a sonic release, while the low gas density at ambient temperatures (7% of air) ensures that any hydrogen released from a compressed gas system will quickly disperse upwards and disappear after losing its initial release momentum. To minimize hazard distances barrier walls may be installed to ensure early deflection of hydrogen releases, as the buoyancy of hydrogen gas, even when diluted in air, is strong at reactive concentrations. A safe design of a system handling compressed hydrogen gas outdoor is therefore not too challenging in most cases. Does the same apply for liquid hydrogen releases?

Liquid hydrogen release dynamics The dynamics of liquid hydrogen releases depend on storage pressure, temperature and whether the release scenario is impinging, confined or into the open. Liquid hydrogen for maritime applications may be handled (bunkering, tank storage and transfer) at temperatures between 20 and 30 K corresponding to saturation pressures of 1e10 bar. A liquid hydrogen release may be described as follows:

Table 1 e Hydrogen energy storage density, weight of tanks and systems to be added. Hydrogen storage

H2 density 3

Compressed 700 bar Compressed 350 bar Compressed 250 bar LH2 Ammonia

41.3 kg/m 24.3 kg/m3 18.2 kg/m3 70.8 kg/m3 129 kg/m3

Methanol Toluene (or LOHC e liquid organic H2 carriers)

99 kg/m3 75 kg/m3

Conventional fuels (MGO and diesel)

~300 kg/m3

Comment Higher pressures require thicker walls and will dominate weight of storage

Weight of cryogenic tanks significant Hydrogen 17% of ammonia weight, refrigeration or pressure tanks and systems will add to weight Hydrogen 12.5% of methanol weight, tanks and systems will add to weight Hydrogen 8.7% of toluene weight, tanks and systems, and used fuel will add to weight Energy density expressed in equivalent hydrogen amount, 14% of mass is hydrogen (~120 kg/m3)

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Outflow phase: Due to the low liquid density of LH2 the outflow velocity will be high (~2.5 times higher than for LNG). At 1 bar overpressure outflow velocity of more than 50 m/s is expected, at 4 bar overpressure more than 100 m/s. Flashing: If the hydrogen is at temperatures above its boiling point an immediate evaporation (flash) of ~2% of LH2 can be expected for each degree K superheat, this will lead to a volume increase (expansion) of around 120% per K. With a storage temperature 5 K above the boiling point immediate flashing of 10% of the LH2 is assumed leading to a plume expansion of a factor 7. (Calculation assumes a constant pressure heat capacity Cp; ~1.25% flash per degree K would apply if Cv is used). Condensed air phase: LH2-temperatures are well below boiling and freezing points of main components of air, oxygen has a boiling point of 90 K and a freezing point of 54 K while nitrogen has a boiling point of 77 K and a freezing point of 63 K. In the near-field outside the release multiphase hydrogen spray may co-exist with oxygen and nitrogen solid or liquid particles. “Pool”: If the LH2 spray impinges before this multiphase region has been heated and evaporated by dilution with ambient air, particles of oxygen and nitrogen may deposit possibly together with some LH2 and receive heat from substrate or objects. The dynamics inside such a multi-phase pile of solid air particles and LH2 may be complicated, due to preferential melting and boiling of oxygen and nitrogen, atmospheres rich on oxygen may be seen. Small increases in oxygen concentration in the local atmosphere can lead to substantial increases in gas reactivity. Cold gas plume: Outside the condensed air phase near-field region the mixture can be assumed to return to gas phase (except fog/ice-particles due to humidity). A plume of gaseous hydrogen and air from a pipe/spray release will initially be significantly colder than ambient air and disperse showing dense gas behavior. A plume resulting from pool evaporation, with evaporation heat provided from substrate, will initially have neutral density in ambient air, see Table 3. With dilution, in particular in humid air, the plumes will gradually become buoyant [9,10] and may rise upwards, see estimated plume densities with humidity in Fig. 1. For LH2 releases into vacuum insulated volumes pool evaporation can be expected main source of phase change. For releases into helium inerted spaces more pool formation than in air would be expected, and evaporated hydrogen is expected to show a clear dense gas behavior. In Table 3 gas and spray densities for natural gas (methane), ammonia and hydrogen are compared to the

density of air at 20  C. At room temperature methane and ammonia densities are very similar, both gases are strongly buoyant with a density less than 60% of the density of ambient air. For comparison the hydrogen gas density is only 7% of that of ambient air. In case of a pool spill the liquefied gases will be boiling. The evaporation heat will primarily be provided by heat transfer from the ground before the evaporating gas is mixed with ambient temperature air. From Table 3 the gas densities at the boiling points indicate that for such a situation natural gas will be denser than air while ammonia evaporating from a pool will be buoyant. Hydrogen evaporating from a pool (or solid air particles slush) will be almost neutral compared to ambient air. As suggested above for releases from LH2-piping at a few bar overpressure and a few degrees K superheat the release will likely have a significant release velocity. The plumes will likely expand due to immediate flashing, and in most cases a liquid spray behavior is expected, even for sprays impinging onto nearby surfaces. This way most of the LH2 evaporation heat will be provided from surrounding air, which will be cooled. In the right column of Table 3 an indicative evaporated spray density (relative to ambient air) is shown for the three gases. For LNG and ammonia evaporated spray densities have been calculated using the FLACS “flash” utility program [8], this program calculates evaporation of a horizontal flashing spray and estimates a gas/air source term (including the evaporated spray density) at the location where all of the spray is evaporated. Due to the low LH2-temperature causing the air condensation region, it is difficult to estimate a similar spray density once the LH2 has evaporated. The gas plume density for the LH2-spray is instead estimated at the location where the gas/air plume has reached the boiling point of oxygen and returned to gas phase. Since gas phase hydrogen has a significantly higher density at its boiling point than ammonia gas, it would be expected that the density of an evaporated LH2-spray is higher than the density of the evaporated ammonia spray (~1.3 times density of ambient air). There should however be no doubt that the initial plume from an LH2-spray will be denser than ambient air and show dense gas behavior. With further dilution with ambient air the hydrogen plume gradually becomes less dense, neutral and eventually buoyant relative to ambient air. Air humidity has a significant effect contributing to plume buoyancy as the heat added to the plume due to water vapor condensation has more impact on plume buoyancy than the reduced volume due to phase change from water vapor to fog. This is clearly illustrated in

Table 3 e LH2 plume properties compared to LNG and ammonia. Liquefied gas

Molecular weight

Gas/spray density relative to dry air at 20  C

Boiling point 

Natural gas (LNG) Ammonia Hydrogen (LH2) a

16 g/mole 17 g/mole 2 g/mole



162 C (111 K) 33  C (240 K) 253  C (20 K)

20 C

Boiling point

Evaporated spray density

0.55 0.59 0.07

1.5 0.73 1.01

> 1.5 (~2.3) > 1 (~1.3) > 1 (~1.1a)

The LNG and ammonia spray densities are estimated at the moment of full evaporation, while the LH2 spray density is estimated at end of the condensed air phase. The location of full LH2 evaporation will be inside the condensed air region, and here the spray density relative to ambient air will be significantly higher (closer to 2).

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Fig. 1 e Estimated plume density for hydrogen plumes from LH2-sprays for 20  C (upper plot) and 0  C (lower plot) for different air humidity (dry air, 50% and 100% relative humidity). 100% relative humidity corresponds to an absolute humidity of 2.3 % (20  C) and 0.60% (0  C), respectively.

Fig. 1 where estimated hydrogen plume density in ambient air of 20  C and 0  C, respectively, is compared to the density of ambient air for dry air, 50% and 100% relative humidity. 100% relative humidity corresponds to 2.3% and 0.6% absolute humidity for 20  C and 0  C respectively. From the plots it can be seen that for dry air (zero humidity) the diluted hydrogen plumes are in practice neutral in ambient air, the marginal buoyancy estimated in Fig. 1 for diluted plumes is not expected to lift plumes off the ground. For the situations with more air humidity it is predicted that a stronger buoyancy effect is seen in warm weather and for high relative humidity (RH) than for cooler temperatures and moderate humidity. The strongest buoyancy is predicted for the case with the highest absolute humidity (2.3% for 100% RH @ 20  C) with minimum plume density of 0.92 relative to ambient air, followed by minimum plume density of 0.96 (1.15% absolute humidity, i.e. 50% RH @ 20  C), and 0.98 and 0.99 for 0.6% and 0.3% absolute humidity (100% and 50% RH @ 0  C). The buoyancy effect thus seems to correlate well to the absolute humidity of air. A dispersing plume of cold hydrogen and air will not immediately lift off the ground even if it has become

marginally buoyant. This will depend on the momentum of the plume and on actual wind profiles. Typical wind profiles with increasing velocity with elevation will tend to keep slightly buoyant plumes on the ground.

Research on LH2 spills Experiments with LH2-spills are few. This is likely both due to the limited availability of LH2 in most countries, but also the limited use in society. The very low temperature of LH2 makes it technically challenging to perform and instrument meaningful experiments of good quality. In the following sections two of the more interesting experimental test series are discussed.

NASA experiments In 1980 NASA researchers [27] performed a series of seven LH2 spill tests at the NASA White Sands test site in New Mexico. 5.7 m3 LH2 (~400 kg) was spilled, a helium back-pressure of 6.9 bara ensured flow momentum. Temperatures and

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concentrations were measured downwind of the spill. After the first test, from which results were not reported, a diffuser was mounted on the outlet to ensure horizontal, radial outflow with minimized spill momentum impact and erosion of the ground. Tests 2, 4, 5 and 6 were spilled over 24e40s giving average spill rates of 10e15 kg/s, while tests 3 and 7 had lower average spill rates of around 4.7 kg/s and 0.8 kg/s, respectively. In Fig. 2 visible plumes for tests 2, 4, 5 and 6 are shown. The visible plume is primarily fog due to the air humidity and was estimated to extend to hydrogen concentrations around 8%. This is similar to the lower flammability limit for downwards flame propagation reported to 8.5e9.45% by Ref. [6] and to around 8% by Ref. [25]. This means that only if the clouds get ignited within the visible plume flames can be expected to burn back to the release source. For ignition above and beyond the visible plume flame pockets should be local and primarily propagate upwards. Fig. 2 indicates that the LH2-plumes (fog) tend to remain near the ground for 50e100 m prior to lifting off. In Test 2, the first test with diffuser mounted, the plume rises more sharply. This is explained by the authors to be because much of the spill hit the gravel ground directly at a high momentum and bounced upwards. When analyzing the plume development it is worthwhile to notice that Test 2 had by far the highest absolute humidity among the NASA-tests (An estimated 1.45%, see overview of tests in Table 4), and the plume could become buoyant before hydrogen concentrations are diluted to 50% in air. The combination of possible upwards flow momentum and heat transfer from ground impingement, plume becoming buoyant due to high humidity and low winds (1.6 m/ s) should be sufficient to explain the immediate lift off the ground for the hydrogen plume of Test 2.

5

In Table 4 atmospheric conditions reported in the tests are shown, in addition the absolute humidity is estimated. Due to the pool spill character of the release heat transfer from the ground can be expected to help vaporize the LH2, thus the resulting vapor plume should be slightly less dense than plumes from LH2-sprays (Fig. 1). The cold gas will thereafter dilute in ambient air, heat up and gradually become buoyant and lift off the ground. The higher the absolute humidity and the lower the wind speed, the sooner the buoyant lift-off can be expected. Despite the simplicity of this assessment the observed plume behaviors seem consistent with this conclusion. For the four plumes in Fig. 2 it is seen that the higher absolute humidity, the sooner the plume lifts off the ground. Test 6 may be an exception, despite a marginally lower absolute humidity than Test 5 its plume seems to lift-off earlier. This can be explained by the much lower wind speed for Test 6 which gives more time for the buoyancy forces to work. For the large spill rate scenarios, the authors commented that there was a strong turbulence in the plume which would enhance the heat transfer from the ground. This effect is likely most significant for low wind scenarios. In Test 7 the release rate was one order of magnitude lower than for the other tests, which also led to a much lower release induced turbulence, and thus less initial plume dilution and heat exchange with the ground. In this test it was commented that the visible, dense/neutral plume along the ground (assumed to represent ~8% hydrogen concentration) extended “well beyond the range of the instrumentation towers (120 m)”. The paper also commented observations in tests from late 1950s by A.D. Little Company [2] in which comparable moderate liquid release rates gave visible dense plumes well beyond 100 m from release location.

Fig. 2 e Visible fog plumes from NASA tests with high release rates (Reprinted from Ref. [27] with permission from Elsevier).

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Table 4 e Test conditions NASA tests. Test 1 2 3 4 5 6 7

Average spill rate

Wind

7 kg/s 10 kg/s 4.7 kg/s 12 kg/s 17 kg/s 11 kg/s 0.8 kg/s

2.7e3.1 m/s 1.3e1.8 m/s 4.5 m/s 3.1e3.6 m/s 6.3 m/s 2.2 m/s 3.1 m/s

Temperature

HSL liquid release tests In 2010 and 2011 UK Health and Safety Laboratories performed two series of unignited and ignited liquid hydrogen release experiments [11,26]. For the reported tests the hydrogen tank was first depressurized and cooled to its boiling point and thereafter some LH2 was vaporized to build a vapor pressure in the tank of 1 barg. With this back-pressure LH2 was released from a 1” diameter mostly vacuum insulated line giving an estimated outflow of 1 L/s LH2 (~70 g/s). The different release scenario setups from the HSL test series are described in Table 5. In the horizontal tests onto the ground some LH2-pool and solid air deposits were observed. HSL reported two different plume behaviors in these tests. For the tests with significant wind (around 2 m/s) the visible plume remained near the ground, while for low wind scenarios (<1 m/s) the visible plume would lift off the ground. The explanation is likely that the lower wind gives more time for heat transfer from the ground, and for the plume to lift once vapor concentration dilutes to levels where plumes become buoyant. In Fig. 3 reported hydrogen concentrations are shown for unignited Test 5. Concentrations up to 13e14% are reported at 6 m distance (0.25 m above ground) and 8% at 7.5 m distance. The wind speed in this test was around 2.7 m/s. For the downwards impinging releases the plume buoyancy is not commented in the reports. For the horizontal free jet release 0.86 m above ground dense plume behavior is observed, see Fig. 4. The plume seems to fall to ground around 4 m from the release point. This behavior is expected since the initial plume is denser than ambient air and there is no heat exchange with the ground prior to the plume touch-down. Thus the plume will initially be clearly denser than ambient air. Similar release characteristics can be expected for any significant LH2 release into the air from piping. A storage temperature higher than the LH2 atmospheric boiling point will lead to immediate partial boiling/flashing during the release. With higher storage temperatures the vapor pressure is also increasing which will

Table 5 e HSL LH2 release scenarios. Test

Spill rate

Height

Unignited Test 5 þ seven ignited tests Unignited Test 6 & 10 Unignited Test 7

70 g/s Onto the ground

Direction Horizontal

70 g/s 100 mm above ground Downwards 70 g/s 860 mm above ground Horizontal

30 24 26 15 12 15 17

Relative humidity

Absolute humidity

18% 49% 27% 43% 43% 29% 29%

0.76% 1.45% 0.90% 0.73% 0.60% 0.49% 0.56%



C  C  C  C  C  C  C

Table 6 e Comparison LH2 and LNG characteristics, instrument releases (5 barg and 10 mm). Property Composition Temperature Density Outflow velocity Leak rate (Cd ¼ 0.62) Combustion heata Energy density LFL a

LH2

LNG

Ratio

100% H2 22 K 68.9 kg/m3 120 m/s 404 g/s 57 MJ/s 0.51 MJ/m3

93%CH4, 5%C2H6, 2%C3H8 113 K 445 kg/m3 47 m/s 1030 g/s 57 MJ/s 2.0 MJ/m3

1 : 6.5 2.6 : 1 1 : 2.5 1:1 1:4

Based on higher heating value HHV. LNG value is for methane, actual mixture likely ~3e4% lower.

normally increase the outflow velocity. Both flashing and increased outflow velocity will increase the turbulence level in the plume, unless there is strong interaction with heat sources (ground, confinement or impinged objects) the plume will initially remain denser than air. With further dilution with ambient air (in particular if air is humid) the plume will gradually become buoyant.

Modelling approach for LH2 pipe releases A detailed modelling of LH2 releases is complex, in particular the phase change dynamics (hydrogen and air) and heat transfer from the surroundings are challenging. For flashing releases of liquefied gases like propane, chlorine and ammonia the particle-air plume is typically cooled 30e40 K below the boiling point of the liquefied gases at the moment of full evaporation. For LH2 with a boiling point of 20.4 K it is necessary to model the heat release due to phase change of air (N2 and O2), if the air phase change is ignored when calculating plume properties for the estimated LH2-spray the plume temperature could fall below 0 K. Ground temperatures of 16 K were reported by Ref. [26] below the pool in a horizontal ground release, this subcooling of the ground is likely caused by the LH2 pool evaporation. In practice the plume temperature for a free spray release of hydrogen into air at atmospheric pressure should not drop below the atmospheric boiling point of hydrogen of 20.4 K. For a proper modelling of liquid hydrogen release and dispersion to estimate exclusion zones in realistic scenarios 3D computational fluid dynamics (CFD) tools should be applied. It is unrealistic to expect that the detailed near-field dynamics of liquid hydrogen releases can be properly modelled by a CFD-tool. Even if models were implemented in a

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Fig. 3 e Visible fog plumes from HSL unignited horizontal ground release Test 5 (upper plot) and reported concentrations at sensor array 0.25 m above ground (lower plot). Sensors are placed on the array of 5 vertical poles located from 3 m to 7.5 m from the release location (From Ref. [26]. Plots published by the HSE and reproduced under the Open Government License.

Fig. 4 e Visible fog plumes from HSL unignited horizontal elevated release Test 7 (From Ref. [26]). Plot published by the HSE and reproduced under the Open Government License. Please cite this article as: Hansen OR, Liquid hydrogen releases show dense gas behavior, International Journal of Hydrogen Energy, https://doi.org/10.1016/j.ijhydene.2019.09.060

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CFD-tool to predict phase change of air and solid air particle interaction and deposition, the phenomena are far too complex and proper experiments too few and with large uncertainties, thus the precision of such a tool could not be verified. In the CFD-tool FLACS [8] the spray dynamics of liquefied gases can be modelled in two different ways. The simpler modelling approach uses a utility program ‘flash’ to calculate the free jet conditions at the moment all the liquid droplets are evaporated. Thus a source term for gas dispersion is specified a certain distance away from the release, with a given air entrainment, a plume cross-section, velocity and temperature. This pseudo-source approach is considered sufficiently accurate provided the region of the plume with droplets is limited and ends before the plume reaches the ground or interacts with objects influencing the plume dispersion. For situations where a more accurate near-field modelling is required a homogeneous equilibrium model (HEM) is implemented in FLACS. The HEM-model approach assumes instantaneous phase change according to the vapor pressure curve [15]. With the HEM assumption of immediate phase change and pressure equalization the plume temperature in a free jet LH2 release should not drop below the atmospheric boiling point of hydrogen (20.4 K). The reality with finite times for phase change and pressure equalization is not expected to deviate much from this. A HEM-model for hydrogen in FLACS, with air phase change modelled, is not yet available. A development version of a HEM-model was however previously used to model HSL Test 7 [16]. To model the dynamics of LH2 releases and dispersion for free jets it is therefore proposed to use a pseudo-source approach with similarities to the flash utility program approach in FLACS. The air condensation/freezing region is assumed to be limited and reversible, and a source term of cold gas mixture of evaporated LH2 and air is defined. FLACS is primarily a compressible gas phase solver in which condensation/freezing of air is ignored. The gas source term mixture of cold hydrogen and air can thus be defined near the release location at temperature near the hydrogen boiling point. To define a proper source term the steps below are followed: 1) Estimate the LH2 outflow rate using the Bernoulli equation combined with an adequate orifice discharge coefficient. For long pipes, small pipe diameters and significant hole sizes friction pressure losses and in-pipe flashing due to heat transfer from pipe can reduce the outflow rate significantly. 2) Estimate the immediate flash fraction due to storage temperature above atmospheric. 3) Estimate the heating rate required to vaporize the remaining unflashed LH2 released. 4) Estimate the amount of ambient temperature air to be cooled and entrained to provide the required heating to vaporize the remaining unflashed LH2. 5) Estimate the LH2 release velocity and momentum from LH2 storage pressure, pipe friction and potential flash fraction prior to release. Thereafter estimate a velocity of the pseudo-source plume with the air entrained, this should be lower than the LH2 exit velocity. The momentum

of the cold air-hydrogen jet can be assumed to be higher than the momentum of the LH2 release due to the flow field set up by the developed jet. For the HSL horizontal free jet release (Test 7) the LH2 was stored at its atmospheric boiling point, thus with assumed no immediate flash. A 1 barg vapor pressure was built up in the tank using an evaporator to push LH2 through the pipe. The 100 pipe consists of a 20 m vacuum insulated section and a short uninsulated section (1.6 m) before the release location. Using the Bernoulli-equation assuming storage pressure 1 barg and no friction or phase change would predict an LH2 release velocity of 53 m/s for a 1” pipe. Assuming a discharge coefficient Cd~1.0 this would give a release rate of 1.9 kg/s. In the experiments an LH2 release rate of only 1 L/s (0.07 kg/s) was reported. There are potentially two main reasons for this significant deviation, these are friction pressure losses and phase change due to heat transfer from the pipe. With a 20 m long 1” pipe the pressure losses will be significant at high flow speeds. By iteratively using a pressure loss calculator [7] assuming a 0.1 mm surface roughness for the flexible pipe gives a likely exit flow rate of 5.5 L/s (0.39 kg/s) with flow velocity of 11 m/s rather than 53 m/s and pipe pressure loss around 0.95 bar. Liquid flow friction pressure losses can thus not explain the low flow rate of 1 L/s. In the tests a pressure of 0.2 barg was reported upstream the release location indicating a pressure loss in the 20 m pipe of 0.8 bar. With the reported flow rate 1 L/s (2 m/s) the friction losses should be low, of the order 0.03 bar. The low flow rate is therefore assumed to be due to heat transfer from the pipe leading to partial flashing of LH2 inside the pipe, both in the vacuum insulated pipe (required to explain pressure loss of 0.8 bar) and in the uninsulated end pipe. The uninsulated pipe will be cooled by the LH2. Assuming typical heat transfer coefficients of 50 W/m2K for air-steel and 500 W/m2K for steelLH2 for the 1.6 m uninsulated pipe, a pipe temperature just below 50 K may be seen. This would vaporize an estimated 2% of LH2, enough to reduce liquid hold-up to half. As mentioned, it also seems likely that LH2 has evaporated inside the 20 m vacuum line which would further reduce the liquid fraction of flow and efficient flow rate. In the past there are examples from large-scale cryogenic release experiments where release rates have been lower than expected due to unintended flashing inside supply pipe. One example is the Phoenix LNG Test 1 [5] where the planned LNG release rate was 0.42 m3/s (~180 kg/s) while at most 0.12 m3/s (50 kg/s) was achieved in the actual test. The main explanatory factor was flashing inside supply pipe and less efficient two-phase flow. In the HSL experimental procedures prior to releasing LH2 the piping was cooled to below air condensation temperature using liquid nitrogen and boil-off hydrogen. Still it is expected that the pipes will be further cooled when the LH2 release is started, which could lead to some increase in flow rate with time. This effect can be suspected e.g. in Test 6, but effect is not significant. In Fig. 5 a FLACS CFD simulation of the HSL Test 7 using the pseudo-source approach for LH2-releases is shown. Because of source term uncertainty due to the flashing inside pipes

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Fig. 5 e Simulation of HSL Test 7 (71 g/s release) using proposed approach. Upper plot shows predicted gas volume fraction (FMOLE) in a vertical cut plane, while predicted temperature profile 11 cm below the jet axis is shown in the lower plot.

simulations were performed with exit velocities from 10 to 30 m/s. The plume touch-down to ground is reported to be around 4 m from the release, corresponding well with the simulation with release velocity of 20 m/s (indicating a likely flash fraction well above 10%). It should be noted that the maximum predicted distance to 15% hydrogen concentration is around 15 m. For comparison the distance to 15% concentration for a 71 g/s sonic gas release of hydrogen at ambient temperature will be less than 3 m, for the liquid release the hazard distance is thus 5 times longer. In Fig. 5 the predicted temperature profile just below the axis of the jet is also shown, with a minimum around 120 K about 1 m from the release and increase to 220e260 K 4e6 m from the release, fairly consistent with observations in experiment [16].

Comparison LH2 and LNG hazard distances bunkering For LNG fueled ships a bunkering risk assessment is required according to ISO 20519:2017 [21]. In this assessment the following zones shall be defined:
Hazardous zone IEC 60079-10-1 [18].


Safety Zone (only access for essential people for the bunkering operation)
Monitoring and Security Area (outer area monitored to prevent people from entering Safety Zone) The Safety Zone can be defined in two ways. It can be established as the maximum LFL-distance from defined credible/dimensioning releases which could be:
Constant release from instrument connection
Rupture of filling hose spilling content of hose only (ESDvalves closed) Alternatively, a quantitative risk assessment can be performed, and the Safety Zone is defined as a minimum to cover the inner exclusion zone (individual annual fatality rate per annum IRPA > 105). The Monitoring and Security Area can similarly be based on best judgement (if credible release approach is used) or to cover the middle exclusion zone (IRPA >106) from a risk assessment. For LH2 there is no specific standard for bunkering. The Norwegian Directorate for Civil Protection (DSB), who is the safety authority overseeing that the major hazards legislation is followed by the industry, has instead requested that the

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ISO20519 LNG standard is used for bunkering studies of LH2 vessels. In the following a comparison assessment of LNG and LH2 hazard distances from credible instrument connection releases is shown. The scenario studied is a release at 5 barg through a D ¼ 10 mm instrument connection, LH2 and LNG compared. The pressure level is sufficient for LH2 transfer, while for LNG pressures closer to 10 barg are common. To find maximum flammable distances it will be required to assess a few different scenarios for each release, e.g. variations in release direction, wind direction and speed. Maximum hazard distances are often found for wind speeds around 2 m/s or for releases with momentum along the wind direction or impinging releases with low momentum to limit initial mixing with air. In this example two scenarios are simulated, these are releases from the rear of a road tanker, both downwards release in 2 m/s wind and horizontal release in 2 m/s tailwind. A typical standard LNG composition is assumed (93% methane). Temperatures for both LNG and LH2 are 1e2  C above atmospheric boiling point, it is assumed that vaporizers are used to generate 5 barg filling pressure inside the road tanker. All liquid is assumed to flash with no pool generation at ground. For downwards release of LNG this may seem to be a conservative assumption, however, without bunding and LNG collection the total evaporation rate could quickly approach the simulated release rate. In Fig. 6 simulation results with the FLACS CFD software are compared. For LNG the FLACS homogeneous equilibrium model (HEM) is used [15], which is considered best practice for LNG sprays. From the release temperature an immediate flash fraction is estimated to get necessary input for the HEMmodel. The release momentum is found using the Bernoulli equation. For LH2 releases no HEM-model is available and the modelling approach described above has been applied. No friction losses or phase change have been assumed prior to the release due to the small hole size. For both LNG and LH2 a transient simulation is performed using the compressible version of FLACS, and the simulation has continued until a nearly steady plume has developed.

The computational domain is flat and extended sufficiently far away from the release so that boundary effects will not influence the results. A Cartesian grid is applied. Near the release and towards the ground, where the highest concentration gradients are expected, a 0.25 m grid resolution is used. Across the release a 7 cm grid resolution is used for the LH2release, only marginally larger than area of expanded plume as per FLACS guidelines. For the LNG release a 5 cm grid resolution across the release is used for the same reason. Further away from the source where gradients are smaller, the grid size is gradually increased. The simulations are performed assuming insignificant humidity (cold winter day). The characteristics of the releases are shown in Table 6. For LNG the LFL-distances extend about 36 m (downwards release) and 25 m (horizontal release). LFL-distances for LNG are somewhat longer than the predicted 15% hydrogen contour which has a similar energy density. This is mainly due to the higher release velocity for LH2 giving a better mixing with air. The LFL for hydrogen is however only 4%, and in Fig. 7 the predicted LFL-contours from the hydrogen simulations are shown. The LFL-distance for the horizontal release is 5 times longer for LH2 than for LNG, and twice as long for the downwards release. The predicted flammable cloud is 12e38 times larger. The estimated explosive cloud, expressed as equivalent stoichiometric Q9 cloud [12], is only 1.5 to 2 times larger for the LH2 scenario, see Table 7. The release scenarios and rates modelled are within the expected range for credible releases per ISO 20519 in a bunkering study. For LNG an extensive validation effort has been carried out for FLACS [13], and predictions are expected to be accurate. While the 36 m LFL-distance would not be too challenging to define as safety zone most places, the 122 m LFL-distance for LH2 would be more challenging and likely prohibitive. Hydrogen at concentrations between 4% (LFL) and 8% has a very low reactivity, is challenging to ignite, and if ignited the flames will only propagate upwards. For people present inside a hydrogen plume below 8% concentration the likelihood to become burnt will be very low, this should primarily happen if the cloud would

Fig. 6 e Downwards release (upper) and horizontal release (lower) with LH2 (left) and LNG (right) from 5 barg release through D ¼ 10 mm instrument connection in 2 m/s wind from left. Plumes of similar energy density ~50% of stoichiometric concentration are shown, i.e. 5% (LFL for natural gas) and 15% (hydrogen). Cold winter day with little humidity is assumed. Please cite this article as: Hansen OR, Liquid hydrogen releases show dense gas behavior, International Journal of Hydrogen Energy, https://doi.org/10.1016/j.ijhydene.2019.09.060

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Fig. 7 e Downwards release (upper) and horizontal release (middle) with LH2 from 5 barg tank through D ¼ 10 mm instrument connection in 2 m/s wind from left. Hydrogen LFL-plumes (4%) are shown. In the lower plot the ground level concentrations for the horizontal release is shown (release direction left from X ¼ 0 m, wind 2 m/s from right). Cold winter day with little humidity is assumed.

Table 7 e Predicted LFL-distances and cloud sizes, instrument releases (5 barg and 10 mm hole size), horizontal and downwards releases with 2 m/s wind, cold winter day with insignificant humidity assumed. Results LFL-distancea Flammable cloud Explosive cloud (Q9) a

LH2 Horizontal/Down

LNG Horizontal/Down

Ratio LH2:LNG

122 m (52 m) / 67 m (43 m) 1800 m3 / 1500 m3 29 m3 / 60 m3

25 m / 36 m 47 m3 / 130 m3 15 m3 / 40 m3

2:1 to 5:1 (1.2:1 to 2:1) 12:1 to 38:1 1.5:1 to 1.9:1

Distance to 8% hydrogen concentration (downwards flammability limit) is shown in brackets.

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ignite at their feet. The temperature and heat from flames at concentrations below 8% would also be low, 370e700  C, with limited risk for severe injury for the expected short time exposure. Thus there are good reasons to define the 8% downwards LFL concentration for hydrogen as the relevant LFL-concentration for this type of study. The Safety Zone would then shrink to about 52 m, which is still significant. If higher humidity of air had been assumed this would have reduced the 4% LFL-hazard distance as the plume might lift off the ground, for the 8% contour the reduction would be less. If the plume can fill confined spaces, care should be taken, as hazard potential from an ignited hydrogen cloud at 8% can still be significant. To justify a further reduction of the bunkering Safety Zone according to ISO20519 a quantitative risk assessment should be carried out estimating fatality risk contours. Such a study should consider release frequencies and rates for a larger variation of hole sizes, ignition probabilities, frequency of various wind conditions and directional dependencies to estimate location specific fatality risk from flashfires, jet-fires and explosions. For bunkering sites with limited activity the 105/year fatality risk contour can be significantly shorter than 52 m. With a high activity level larger safety zones might be required. Bunkering studies not recognizing that LH2-releases can be expected to show dense gas behavior predicting buoyant plumes, may severely underestimate the bunkering Safety Zone. A too small bunkering Safety Zone may potentially expose bunkering staff and passengers to higher risks than considered acceptable.

Further safety challenges with LH2 Hydrogen fueled ships must perform risk assessments as required by the IGF-code [19], alternative design option. In this assessment an equivalent safety level to conventionally fueled vessels must be demonstrated. An explosion study is a mandatory part of such an assessment. One example of such a concept risk assessment for a compressed gas hydrogen fueled high speed passenger ferry is presented by Ref. [1]. In the following some safety challenges with LH2, relevant to consider in risk assessments, are discussed.

Indoor LH2 releases If the LH2-tank is located indoor below deck there may be potential for both LH2- and gas phase H2-releases. Normally there will be a tank connection space (TCS) next to the tank with bunkering lines, and LH2 fuel lines to a vaporizer and a pressure build-up unit. The TCS is normally well ventilated, in the IGF-code [19] 30 air changes per hour (ACH) is required as minimum for LNG-vessels. In Table 8 a coarse risk evaluation is performed in which LH2- and gas phase H2-leak rates are estimated for various hole sizes (1e5 mm diameter) at typical tank pressures 3 and 6 barg. Often used discharge coefficients of 0.62 and 0.85 are assumed for liquid and gas phase, respectively. In the assessment forced ventilation of 30 ACH and a 43 m3 net volume of the TCS are assumed. The maximum average concentration in the TCS is estimated using a simple stirred-tank-reactor model (inlet air and leaked gas are injected while the average gas-air mixture is extracted for each time interval). This way an approximate maximum gas concentration inside the TCS is predicted, as well as the time it takes to reach an average concentration of 4% (LFL). The assessment in Table 8 shows that for the same hole sizes the estimated leak rates for LH2 may be 18 and 15 times higher at 3 and 6 barg than for gas phase H2 releases at 20  C, while cryogenic gas phase releases at 30 K, representative for piping connected to tank vapor space, would have 3 times higher release rate than releases at 20  C. A leak rate of 1.3 g/s is required to reach average concentration of 4% (LFL) inside the TCS while 2.8 g/s is required to reach 8%. Even the smallest hole size considered in Table 8 (1 mm diameter) gave leak rates higher than this for LH2. For the gas phase releases at 20  C hole sizes of 2 mm would not give average concentration above LFL, while 3 mm releases could give concentrations between 4 and 8%. For the 30 K gas phase releases 2 mm releases from 3 barg would not give flammable clouds. 2 mm 6 barg releases would give concentrations between 4 and 8%, while releases from 3 mm hole size would give concentrations well above 8% for both pressures. CFD-calculations will give more accurate results, predicting somewhat higher or lower concentrations and spatial concentration variations depending on the inlet and outlet of the ventilation system relative to the relevant leak scenarios.

Table 8 e Release rates from LH2 releases and H2-gas releases and maximum average concentration in 43 m3 TCS. Release rate shown in bold may give concentrations above 8% while release rates in italic gives concentration above LFL but below 8%. Leak diameter @ pressure 1 2 3 5 1 2 3 5

mm mm mm mm mm mm mm mm

@ @ @ @ @ @ @ @

3 3 3 3 6 6 6 6

barg barg barg barg barg barg barg barg

Release rates LH2 3.1 g/s 12.2 g/s 27.5 g/s 76.3 g/s 4.3 g/s 17 g/s 39 g/s 108 g/s

H2 (30 K) H2 (20  C) 0.52 g/s 2.1 g/s 4.7 g/s 13 g/s 0.91 g/s 3.6 g/s 8.2 g/s 23 g/s

0.17 g/s 0.67 g/s 1.5 g/s 4.2 g/s 0.29 g/s 1.2 g/s 2.6 g/s 7.3 g/s

Ratio H2 leak rate LH2: H2 (30 K):H2 (20  C) 18 18 18 18 15 15 15 15

: 3.1 : 3.1 : 3.1 : 3.1 : 3.1 : 3.1 : 3.1 : 3.1

:1 :1 :1 :1 :1 :1 :1 :1

Maximum H2% concentration (time to LFL is reached) LH2

H2 (30 K)

H2 (20  C)

8.7% (57s) 25% (12s) 39% (5s) 59% (1s) 11% (37s) 30% (8s) 46% (3s) 64% (1s)

1.7% () 6.5% (94s) 14% (34s) 30% (11s) 3.0% () 10% (45s) 21% (17s) 43% (6s)

0.56% () 2.2% () 4.6% (190s) 11% (38s) 0.96% () 3.8% () 7.5% (70s) 17% (20s)

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Fig. 8 e Predicted explosion pressures from methane and hydrogen as function of gas concentration in air in a 43 m3 tank connection space with explosion vent duct dimensions 0.38 m £ 0.38 m and 1.0 m £ 1.0 m, respectively.

Due to the wide flammability range and high reactivity of hydrogen it is critical to ensure that maximum release rates of LH2 or gas phase hydrogen into the TCS or other indoor rooms must be kept low enough to prevent flammable atmospheres to develop given the ventilation present. If the TCS is equipped with a properly dimensioned explosion vent duct, slightly higher concentrations could be acceptable. In Fig. 8 explosion pressures predicted with FLACS CFD-tool are shown for methane and hydrogen as function of gas concentration for the 43 m3 TCS with vent ducts of 0.38 m square and 1.0 m square. FLACS is a well validated and recognized tool to predict explosion consequences both for natural gas and hydrogen and with certain settings it is also possible to predict DDT and detonations using FLACS [14]. For methane the large vent duct seems sufficient to limit the worst-case explosion pressure to around 1 barg. This may damage TCS-walls, but likely not lead to any further escalations for a robust ship design. For hydrogen higher pressures are predicted over a much wider concentration range. Severe explosions with escalation may be feared for significant gas concentrations. The effect of vent duct diameter is significant, with the larger vent duct predicted overpressures are below 1 barg for concentrations up to 15%, while for the smaller vent duct pressures exceed 1 barg already at 10% gas concentration. For a risk-based design a properly dimensioned vent duct could improve safety and increase the tolerance with regard to maximum possible leak rates within the TCS. For the explosion pressures predicted in Fig. 8 room temperature (20  C) is assumed inside the TCS at time of ignition. If an LH2 spray release is generating the gas cloud a temperature of 59  C prior to ignition could be expected for a stoichiometric mixture inside the TCS (a mixture at 45% hydrogen concentration would be at 105  C). At these temperatures there would be 37% and 74% more gas-air molecules to burn inside the TCS, respectively. This will

likely lead to significantly higher explosion overpressures for the quite confined TCS, despite that the gas reactivity should be somewhat reduced. The flame speed models for very low temperatures in FLACS predicts lower flame speeds, but the precision is uncertain. The heat release during combustion should however be well estimated. The predicted overpressures for high reactivity mixtures with FLACS at low temperatures inside the poorly ventilated TCS are expected to be representative. For the smaller vent duct (0.38 m square) pressures are predicted to increase from 6.7 barg to 9.2 barg for the stoichiometric mixture if the gas cloud results from a cryogenic release (gas or liquid), while a pressure increase from 6.0 barg to 10 barg is predicted for the colder 45% mixture. The significant increase is despite the reduced reactivity for cold mixtures and is due to the significantly increased energy release during combustion. For LNG-fueled vessels there is normally a potential for severe explosions to happen inside the TCS, but this would require that most of the TCS is filled to methane concentrations between 8 and 12% which ignites. Methane releases in the TCS from LNG (and cold vapor) will be more likely to stratify due to a low release velocity and higher density than air, thus most likely only a limited fraction of the TCS volume will be at reactive gas concentrations simultaneously. Maximum overpressures for methane will then be significantly lower than presented in Fig. 8. The higher reactivity, wider flammability range, higher release velocity and ignition probability of hydrogen relative to methane make it important to design LH2 fueled vessels to higher safety standards than existing LNG vessels. The probability for a severe explosion if a significant release happens inside the TCS may be several orders of magnitude higher for an LH2 fueled vessel than for an LNG fueled vessel. To comply with the IGF-code requirement of equivalent safety level the hydrogen systems must therefore be designed to prevent the possibility for severe explosions.

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Vent mast releases Depending on the vessel and design there may be situations where LH2 can be sent to the gas mast, this could be during bunkering or for major releases into double piping between tank and vaporizer. With the gas mast initially at ambient temperature there will be a significant heat transfer from the gas mast to vaporize the LH2. The cooling depends on the heat capacity, length, diameter and thickness of the gas mast as well as the LH2 flow rate into the mast. Depending on scenario it may take 15s, 30s, 60s or more until the mast is sufficiently cooled and LH2 will reach the exit of the mast. Once this happens the hydrogen plume leaving the gas mast will

become denser than ambient air and may fall towards deck/ sea. For smaller vessels, e.g. a fast passenger ferry, this may be of more concern than for larger vessels, since the gas mast is shorter. This might give a higher plume concentration at touch-down of the plume, and with a higher possibility of hitting vulnerable targets than for releases from a taller gas mast. In most cases the wind and/or the speed of the vessel will prevent dense plumes from becoming a safety concern to people on board. In Fig. 9 FLACS CFD simulations of hydrogen plumes from 500 g/s LH2 releases from a 15 m tall 150 mm diameter gas mast are shown for 0.5 and 2.0 m/s wind for cold winter days (10  C, no humidity) and typical spring days (þ15  C, 1.2%

Fig. 9 e FLACS simulation of 500 g/s LH2 release from a 15 m tall gas mast on cold, dry winter day ¡10  C (left) and typical spring day þ15  C (right). Wind speed 0.5 m/s and 2.0 m/s from fore, plots are shown 30s and 150s after mast has been sufficiently cooled for LH2 to reach top of vent mast. White plume shows hydrogen concentrations above 15% while color legend show flammable concentrations onto deck. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.) Please cite this article as: Hansen OR, Liquid hydrogen releases show dense gas behavior, International Journal of Hydrogen Energy, https://doi.org/10.1016/j.ijhydene.2019.09.060

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absolute humidity). For the low wind scenarios the plume is predicted to fall onto deck at concentrations above 15%. For the cold winter day the plume remains dense after falling to deck and continues falling towards sea, while the spring day plume with higher humidity is predicted to rise upwards after falling to the deck. For the 2 m/s wind the cold winter day plume is predicted to make a touchdown onto the deck initially before establishing a higher trajectory falling clear of the aft of the vessel. For the spring day plume with 2 m/s wind the plume goes clear of the deck before becoming buoyant. In most cases such a release should be detected and stopped well before mast would be cooled sufficiently to release LH2. The main potential for such scenarios may be during bunkering and should be reflected when defining bunkering safety zones.

LH2 spills into outer double tank A release from the inner LH2 tank into the outer vacuum insulated tank will be heated when in contact with the outer tank. Due to evaporation of LH2 pressure will build up and cold hydrogen gas will be vented through the pressure relief valve into the vent mast. The outer tank should be of double containment type and able to collect the released LH2 without failing, if not, a severe leak scenario from the inner tank could be catastrophic. If the release is significant the evaporation and pressure build-up may be high, and it is important to dimension the pressure relief valve and gas mast so that risk for failure of outer tank due to pressure is negligible. Measures could also be taken to limit the initial pool size and evaporation rate inside the outer tank.

BLEVE and trapped LH2 scenarios An LH2 tank with a properly dimensioned gas mast relief vent should generally be safe in an external fire situation. The heat from the fire will increase the evaporation rate of LH2 with vapor sent to the gas mast. If the gas mast venting capacity is sufficient and can handle increased heat transfer also after a potential weakened outer tank and loss of vacuum, this scenario could be managed. The fire scenario may be mostly relevant to LH2-tanks outdoor above deck, below deck problematic fires should be prevented by limiting access of oxygen and/or flammables. If the venting capacity is not sufficiently dimensioned the pressure inside the tank may increase until the tank fails. LH2tank pressure relief vents may open when tank pressure reaches around 8 bar, and tanks may fail catastrophically, maybe at 15e20 bar, if the pressure continues to increase. One important question is whether this will result in a BLEVE (Boiling Liquid Expanding Vapor Explosion). To predict whether a BLEVE may take place the Reid superheat temperature limit criterion is often applied [24]. This simple relation works well for a wide range of compounds. According to this homogeneous boiling and BLEVE can be expected for tank ruptures with significant hole sizes to give rapid pressure loss at temperatures above TSTL ðKÞ ¼ 0:895 x Tc ðKÞ For hydrogen the critical temperature is 33.18 K and critical

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pressure is 13 bar. Using Reid STL criterion LH2 BLEVEs should be feared at temperatures above 29.7 K corresponding to a vapor pressure of 7.7 bar. This condition is very close to or within the possible operational range for LH2. Does this mean that an explosive BLEVE and tank explosion would result if a tank saw a sudden significant pressure loss? For LNG the TSTL is 171 K, which is almost 60 K above the boiling point with a vapor pressure of 24 bar. Above this a BLEVE may be feared if the tank fails. With a sudden pressure loss at this temperature 23% of the LNG in a tank will immediately boil. As LNG will expand 241 times when boiling, a BLEVE at 171 K will thus in milliseconds expand 56 times. For comparison LH2 at 29.7 K is only 9.3 K above its boiling point, and with a sudden pressure loss to atmospheric pressure only 11.8% of LH2 will immediately boil. LH2 at 29.7 K will expand 42 times when boiling thus with an 11.8% immediate boiling a volume increase of about factor 6 is predicted. The expanding LH2-droplets are cooled to their boiling point and will cool, condense and freeze surrounding air in the process. Air will get compressed around 700e800 times, and the net effect when LH2 mixes with ambient temperature air may be limited, possibly a weak implosion, making the actual expansion of LH2 less than factor 6. From these considerations it is difficult to imagine a traditional, violent BLEVE-scenario developing with LH2 around 30 K. Since the air entrained in the expanding LH2 will be frozen, the potential for ignition and significant combustion and pressure effects is unclear. No experiments with LH2 BLEVEs are found in literature. It will be very interesting to see the results from such experiments in the future. For the LH2-piping, however, it is important that no LH2 is trapped inside sections of the pipe closed in both ends by ESD (emergency shutdown) valves. This would result in uncontrolled pressure build-up due to heat transfer from surroundings and failure of the pipe and explosive release of flashing LH2. The amount of hydrogen inside pipes will however be limited, and the main risk foreseen from such trapped LH2 scenarios will be a rupture of a pipe with a potential following explosion or flash-fire if the mixture should ignite after mixing with air.

RPT Rapid phase transition is a phenomenon of some concern for LNG. This is often split between early RPT, potentially seen during LNG release into water, and delayed RPT which can be seen after a significant fraction of the lighter components of the LNG has evaporated. The RPT takes place once the boiling point of the remaining components gets closer to the sea temperature so that insulating film-boiling layer collapses [3]. The delayed RPT mechanism has a higher potential for explosions than the early RPT mechanism. For LH2 only the early RPT mechanism is relevant, as there will be no preferential boiling leaving denser compounds with a boiling point close to that of water. Due to the very low density of LH2 a spill falling onto the sea is unlikely to

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penetrate water significantly. If LH2 is released at significant momentum below the water surface, some RPT effects may be seen, but this is assumed to be much less than for a delayed LNG RPT.

Conclusions There are currently numerous zero emission initiatives within shipping, and political decisions or private initiatives may initiate many more in the years to come. Hydrogen is the chosen energy carrier for many of these vessels. Scaling up from smaller demonstration vessels to larger vessels the use of liquid hydrogen LH2 may be attractive due to a higher storage density and faster bunkering compared to compressed gas, in addition the leak potential and risk from systems with numerous high pressure tanks can become significant. It is important that stakeholders have a proper understanding of the physics related to potential loss of containment incidents with LH2 and can design ships and safety systems in the most optimal way. Some of the initiatives planning to build LH2 fueled vessels do not seem to acknowledge that LH2 released into air will initially show dense gas behavior rather than rising into the sky and disappear like releases of compressed hydrogen gas. The current work aims to clarify the release and dispersion mechanisms of LH2, and to explain that LFL-distances for LH2 releases can be significantly longer than for comparable LNG-releases, not the opposite. Some further hazards of concern with regard to using LH2 for maritime applications are discussed, illustrating why hydrogen fueled vessels may require to be designed to higher safety standard than current LNG fueled vessels. A proper understanding of these aspects is required for a safe and optimized design of LH2 fueled vessels.

references

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Please cite this article as: Hansen OR, Liquid hydrogen releases show dense gas behavior, International Journal of Hydrogen Energy, https://doi.org/10.1016/j.ijhydene.2019.09.060