2.1 Hydrogen storage

Despite their limited gravimetric hydrogen storage densities, MHs offer multiple advantageous properties like compactness, low infrastructure requirements, low energy consumption, loss-free storage and inherent safety due to the endothermic nature of the hydrogen release [Reference Züttel, Remhof, Borgschulte and Friedrichs3, Reference Dornheim and Moreno-piraján5, Reference Lototskyy, Tolj, Pickering, Sita, Barbir and Yartys11–Reference Kazula, De graaf and Enghardt16]. Therefore, MHs may be ideal as a secondary hydrogen source, which serves as an emergency supply. As a drawback, MH storage systems require a thermal management system for heat supply and heat rejection [Reference Baroutaji, Wilberforce, Ramadan and Olabi4]. This can lead to a delay between the heat activation and the release of hydrogen [Reference Colozza and Kohout17].

Hybrid storage systems, which combine solid storage in MHs with compressed gaseous hydrogen (CgH2) storage at high pressures, could be a suitable compromise. These systems achieve higher volumetric storage densities compared to CgH2, while causing only moderate weight penalty [Reference Lototskyy, Tolj, Pickering, Sita, Barbir and Yartys11, Reference Puszkiel18–Reference Zhang, Li, Yuan, Qi, Guo and Zhao25]. Hybrid systems also offer a fast response behaviour, which is mandatory for emergency power supply systems [Reference Schulz, Liang, Hout and Larochelle24]. Disadvantages of hybrid systems include the loss of inherent leakage safety and the need for an additional containment device to avoid MH spillage in the event of a damaged vessel [Reference Takeichi22, Reference Yokota23]. The use of such hybrid storage systems to store reserve fuel was analysed in Ref. (Reference Bahrs, Franke, Kazula and De graaf26). This study presents the achievable mission ranges for a commuter aircraft, which is equipped with a dedicated tank for reserve fuel to allow a diversion to an alternative airport. The results show that an application of a hybrid storage system for this reserve is only feasible when the payload of the aircraft is reduced. Despite the improved gravimetric hydrogen storage densities compared to pure MH storage systems, the energy density of a hybrid system remains too low compared to other energy storage technologies of hydrogen.

2.2 Heat storage

According to Equation (1), the characteristic of the hydride formation allows the use of MHs in thermal applications. MHs possess high thermal storage capacities of up to 2,000kJ/kg_MH and are therefore ideally suited as heat sinks for cooling applications or as heat sources for pre-heating processes [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Bhuiya, Kumar and Kim7, Reference Yang, Wang, Zhang, Meng and Rudolph27, Reference Kölbig, Bürgers and Linder28]. Both use cases appear promising for future electric propulsion systems. The high quantity of waste heat produced at the relatively low operating temperature of the components, especially in polymer electrolyte membrane fuel cells (PEMFC), poses challenges to the thermal management system (TMS) [Reference Webber, Llambrich and Davoudi29–Reference Haran, Madavan and O’connell32]. In these powertrains, an MH reactor (MHR) is suitable for pre-heating a fuel cell (FC) prior to a cold start to reduce aging. Additionally, it can operate during periods of high power demand, like take-off or climb, to absorb some share of the heat generated by the PEMFC [Reference Bürger, Sourmelis terzopoulos, Kretschmer, Kölbig, Brack and Linder33–Reference Boll35]. Both of these operation modes are illustrated in Fig. 3. While a CgH2 tank serves as the main hydrogen source for the FC, the MHR stores the hydrogen only intermediately to generate a heating or cooling effect. No hydrogen is consumed by the operation of the MHR.

Figure 3.

Operation modes of an exemplary FC-powertrain coupled to an MHR: (a) pre-heating of the FC by the heat of absorption and (b) cooling of the FC by the desorption’s heat uptake.

The operating temperatures of PEMFCs restrict the use MH alloys to low-temperature types with lower thermal storage capacities in the range of 184kJ/kg_MH for the example of LaNi_4.85Al_0.15 [Reference Kölbig, Bürgers and Linder28, Reference Kölbig, Bürger and Linder36, Reference Franke, Link and Kazula37]. Another limiting factor occurs when the FC hydrogen intake is directly coupled to the MHR hydrogen outlet. In that case, the FC hydrogen consumption limits the achievable thermal power of the MH, as the hydrogen mass flow dictates the magnitude of the absorption and desorption heat flow [Reference Kölbig, Weckerle, Linder and Bürger34]. In an example of direct coupling, the MH can theoretically absorb up to 35% of the FC’s waste heat when the FC operates at an efficiency of 60% [Reference Franke, Link and Kazula37]. This proportion becomes even larger at higher FC efficiencies.

However, the duration of the cooling effect is limited by the hydrogen capacity of the MHR, since the MH material only absorbs heat during the endothermic desorption of hydrogen. The longer the desired operating time, the more MH material is required, which results in an increased system mass. This mass increase causes the specific cooling power of the MHR to decrease because the thermal power remains constant, as it is linked to the hydrogen mass flow of the FC. Therefore, the use of MHs as a heat sink for component cooling does not seem reasonable for operating times above one minute, as the specific cooling power of the total MHR drops below 1.5kW/kg [Reference Franke, Link and Kazula37]. Nevertheless, the utilisation of MHs as a heat source for de-icing, for pre-heating of the cabin or for warm-up of components like fuel cells, sensors or lubricants may still be reasonable use cases in aviation due to the high thermal storage capacities [Reference Kölbig, Bürgers and Linder28, Reference Bürger, Sourmelis terzopoulos, Kretschmer, Kölbig, Brack and Linder33, Reference Kölbig, Weckerle, Linder and Bürger34, Reference Kölbig, Bürger and Linder36, Reference Moir and Seabridge38].

2.3 Heat pumps

In addition to thermal storage, MHRs can also operate as heat pumps [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Bhuiya, Kumar and Kim7, Reference Yang, Wang, Zhang, Meng and Rudolph27]. A heat pump transfers heat from a low temperature source to a higher temperature sink [Reference Haran, Madavan and O’connell32]. MH-based heat pumps can be driven by the pressure differential between a CgH2 storage system and a consumer, such as an FC, as shown in Fig. 4. Thus, MH-based heat pumps can convert the potential energy of the high-pressure gas into thermal energy for both heating and cooling [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Kölbig, Bürgers and Linder28, Reference Hegner and Weckerle39, Reference Weckerle, Bürger and Linder40]. Until now, this potential energy is throttled to the supply pressure of the consumer and thus lost [Reference Weckerle, Bürger and Linder41]. When MH-based thermal devices utilise the potential energy of the tank pressure, they require only minimal additional amounts of electric energy for their operation [Reference Hegner and Weckerle39].

Figure 4.

Working principle of metal hydride heat pump integrated into an FC-powertrain with a compressed gaseous hydrogen (CgH2) storage according to Ref. (Reference Kölbig, Weckerle, Linder and Bürger34).

When used for cabin air-conditioning or galley cooling, the rather low energy consumption of MH heat pumps could result in a weight advantage compared to conventional state-of-the-art aircraft systems, as less energy has to be carried during the flight. Furthermore, MH-based thermal devices are environmentally friendly, as they can operate without harmful refrigerants [Reference Hegner and Weckerle39]. During cruise flight, the cabin of the aircraft usually requires cooling, even though ambient temperatures are low at high altitude [Reference Scholz42, Reference Herzog43]. Presently, air-conditioning systems of aircrafts comprise air cycle machines (ACM), vapor cycle machines (VCM) or a combination of both [Reference Scholz42, Reference Burschyk44–Reference Moir and Seabridge48]. As VCMs are more efficient, this type serves as a benchmark for the proposed MH-based air-conditioning system in Section 4.1 [Reference Haran, Madavan and O’connell32, Reference Scholz42, Reference Burschyk44, 47, Reference Moir and Seabridge48].

In addition to air-conditioning, MH-based heat pumps have the potential to improve the heat rejection of the propulsion system’s TMS. The limited temperature difference between PEMFCs and ambient air presents a significant challenge for the TMS. The implementation of an MH-based heat pump can increase the temperature difference between the heat transfer fluid (HTF) and the ambient air. Thus, the necessary heat exchanger size (HX) decreases as heat rejection temperature difference increases [Reference Webber, Llambrich and Davoudi29–Reference Hadnum, Pacey and Milne31]. A pair of two MHRs enables the quasi-continuous operation of a heat pump through alternating hydrogen absorption and desorption, as shown in Fig. 5 [Reference Franke, Link and Kazula37]. Subsequent to exiting from the FC, the hot HTF flows through the first MHR, which increases the HTF temperature even further to elevate the heat rejection temperature in the HX. Prior to re-entering the FC, the second MHR removes the quantity of heat that was added by the first MHR, from the HTF. The increase in heat rejection temperature can potentially reduce the size of the HX, although the amount of heat rejected to the ambient air does not change significantly. This is particularly advantageous in hot ambient conditions, where the temperature difference between the PEMFC and the surrounding air is low. At an ambient temperature of 38°C and an FC efficiency of 60%, the integration of an MH-based heat pump could theoretically result in a 7% relative decrease in HX size. However, the additional weight caused by the pair of MHRs functioning as a heat pump exceeds the mass savings of the HX by a factor of 10 [Reference Franke, Link and Kazula37].

Figure 5.

Schematic of the alternating heat pump operation of two MHRs to enhance the FC’s TMS by increasing the temperature difference to the ambient air in the HX.

2.4 Heat switches

Heat switches are utilised in the thermal management of cryogenic applications because they can provide a thermal connection or separation [Reference Shu, Demko and Fesmire49, Reference Jahromi and Sullivan50]. Gas gap heat switches (GGHS) in particular permit variable thermal conductance by adapting the gas pressure in a narrow gap between two components [Reference Shu, Demko and Fesmire49, Reference Prina, Bhandari, Bowman, Paine and Wade51, Reference Prina, Bowman and Kulleck52]. Hydrogen is the optimal operating medium for GGHSs because it has the highest thermal conductivity among all gases [Reference Bhuiya, Kumar and Kim7, Reference Manjavacas and Nieto53]. Despite this, its conductivity remains 100–1,000 times lower than that of solid metals such as copper or titanium [Reference Chan and Ross54]. As a result, GGHS require large surface areas and small gaps, as illustrated in Fig. 6 [Reference Shu, Demko and Fesmire49, Reference Franco, Martins, Catarino and Bonfait63]. In design (A), the hot component inside of a vacuum tubular vessel resembles a heat source that requires temporary cooling. A feedthrough enables the in-/outlet of operating fluids or electric wiring, depending on the interfaces required by the specific device. In design (B) and (C), the hot and cold component resemble the interfaces to an arbitrary heat source and heat sink, respectively. Although both designs possess fins to increase the area for heat transfer, they differ in their outer shape. While design (B) is framed by a tubular barrier, design (C) forms a flat panel. For all three designs, MHs can act as regulating units for the provision and removal of the hydrogen gas [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Bhuiya, Kumar and Kim7, Reference Prina, Bowman and Kulleck52, Reference Chan and Ross54, Reference Vanapalli, Keijzer, Buitelaar and Ter brake62, Reference Bowman64].

Figure 6.

Gas gap heat switch (GGHS) designs: (a) concentric vessel design acc. to Refs (Reference Bowman55–Reference Wade57), (b) overlapping fin design according to Refs (Reference Chan and Ross54, Reference Kimball58–Reference Bywaters and Griffin61) and (c) flat panel design acc. to Refs (Reference Vanapalli, Keijzer, Buitelaar and Ter brake62, Reference Krielaart, Vermeer and Vanapalli68).

The working principle of GGHSs may be suitable for facilitating an efficient thermal management of liquid hydrogen (LH2) tanks with vacuum insulation. Although LH2 is the most promising storage option for aircraft, its need for a TMS to supply the hydrogen and to maintain the tank pressure is a disadvantage [Reference Baroutaji, Wilberforce, Ramadan and Olabi4, Reference Gong and Verstraete65–Reference Colpan and Kovač67]. An MH regulating unit with the capacity to vary the vacuum insulation pressure of an LH2 tank by desorption or absorption of hydrogen can thereby affect the heat transfer through the insulation [Reference Krielaart, Vermeer and Vanapalli68]. In this particular use case, the MH regulating unit acts as a heat amplifier. A small heat supply to the MH for hydrogen desorption results in a larger heat input into the LH2 tank due to the increased heat transfer in its insulation. With such an approach, 95% of the necessary evaporation energy can be drawn from the surroundings [Reference Michel, Fieseler, Meyer and Theißen69]. This reduces the power requirements for the LH2 tank’s TMS. The described approach is similar to the GGHS design (a) from Fig. 6 but with the hot and cold components reversed. Section 4.4 will adapt all the designs from Fig. 6 to the geometry of an LH2 tank and estimate the size of an MH regulating unit to allow the evaluation of the potential of such MH-actuated GGHS for an aircraft’s LH2 tank.

2.5 Hydrogen compressors and actuators

MHs enable the design of thermally driven hydrogen compressors [Reference Weckerle, Bürger and Linder41, Reference Von colbe70–Reference Lototskyy, Yartys, Pollet and Bowman72]. This allows the use of waste heat, resulting in an environmentally friendly and cost-effective hydrogen compression [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Yartys73]. With this ability, MHs may be used to efficiently capture, store and compress the boil-off hydrogen gas released from LH2 tanks for further utilisation, as illustrated in Fig. 7 [74–Reference Rosso and Golben76]. While absorbing boil-off gas from a low pressure LH2 tank, the MH has to be cooled by rejecting the heat of the absorption reaction. An increase in the MH’s temperature would raise its equilibrium pressure, as indicated by the black line, and thereby impede the hydrogen absorption. In turn, this characteristic behaviour is utilised after the hydrogen uptake is completed. By increasing the MH’s temperature, e.g. by using the waste heat from the FC, the hydrogen is released at a higher pressure level to be directly supplied to the FC. The boil-off gas, which occurs periodically during refueling and also continuously during periods when the aircraft is not in service, could be a cost-effective fuel source for ground-based power services [Reference Brewer77]. Recovering this hydrogen decreases the financial loss resulting from the venting of hydrogen boil-off and thereby enhances the efficiency of the LH2 supply chain [Reference Brewer77].

Figure 7.

Schematic Van’t Hoff plot illustrating the working principle of hydrogen compression by metal hydride from low pressure liquid hydrogen (LH2) storage to the supply pressure of a fuel cell (FC).

Compared to the sequential process of hydrogen uptake at low pressure followed by provision at higher pressure for the boil-off recovery, a continuous compressed hydrogen supply can be achieved by utilising at least two MHRs operating in parallel [Reference Von colbe70, Reference Lototskyy, Yartys, Pollet and Bowman72], as shown in Fig. 8. An alternating operation of this parallel process has the potential to enable an onboard compression of hydrogen, which is extracted in gaseous form from a low-pressure LH2 storage, to the supply pressure of the consumer such as an FC or a combustion engine. Therefore, the waste heat of the consumer, which is an FC in Fig. 8, is provided to one of the MHRs to drive the desorption of hydrogen on a pressure level that is desirable for the consumer. In parallel, the other MHR is refilled with hydrogen from the LH2 tank. To enable such low-pressure absorption, this MHR has to be cooled by an additional HX.

Figure 8.

Topology of a metal hydride onboard hydrogen compression system in an FC-powertrain with low pressure liquid hydrogen (LH2) storage.

Furthermore, the compressed hydrogen could be utilised in pneumatic actuators to convert waste heat into mechanical work [Reference Bhuiya, Kumar and Kim7, Reference Lototskyy, Yartys, Pollet and Bowman72, Reference Ino, Sato, Hosono and Izumi78–Reference Vanderhoff and Kim82]. While hydraulic and electrical systems perform most actuation tasks in aircrafts, pneumatic devices are applied as well in minor applications [Reference Moir and Seabridge48]. Exemplary pneumatic aircraft systems, which could potentially be powered by compressed hydrogen instead of compressed air, are engine services on the one hand, like actuating variable guide vanes or the thrust reverser, and aircraft services on the other hand, like pressurising water tanks and hydraulic reservoirs or inflating flexible deicing systems [Reference Moir and Seabridge38]. Nevertheless, replacing air as an operational medium by hydrogen gas leads to additional hazards, such as material embrittlement or the accumulation of explosive gas mixtures [14, Reference Hübert, Boon-brett and Buttner66]. Due to these potential hazards and due to the minor use of pneumatic systems in aviation, only the two use cases ‘boil-off recovery’ and ‘onboard hydrogen compression’ will be further investigated in Section 4.

2.6 Hydrogen purification and separation

Although there are many methods for purifying hydrogen, MHs have received special attention in that field due to their safe, simple and cost-effective operation. MHs achieve a high-level purification of hydrogen, regardless of the input hydrogen quality [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Chen, Wei, Deng, Yang and Zhang9]. Figure 9 shows the procedure of hydrogen purification by MHs. When the impure gas enters the MHR, only the hydrogen is absorbed, while the other gas admixtures occupy the void space between the MH particles. After flushing out these admixtures, the MHR can provide pure hydrogen by desorption. This use of MHs may be promising in FC-powered aircrafts, since hydrogen impurities can cause rapid degradation of PEMFCs [14]. However, onboard purification is not reasonable due to the limited hydrogen exploitation rates of 75–95% and the high weight of such MH systems [Reference Bhuiya, Kumar and Kim7, Reference Chen, Wei, Deng, Yang and Zhang9, Reference Borzenko, Blinov, Dunikov and Malyshenko83].

Figure 9.

Working principle of hydrogen purification by MHs according to Ref. (Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6): (a) filling with impure hydrogen and absorption, (b) flushing of admixtures and (c) desorption of pure hydrogen.

Besides purifying hydrogen gas, MHs can also separate hydrogen from gas mixtures. However, this separation method is only applicable for hydrogen contents exceeding 15% [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Chen, Wei, Deng, Yang and Zhang9]. Additionally, admixtures such as oxygen or water can gradually reduce the capacity and durability of the MH material [Reference Dunikov, Romanov and Mitrokhin84]. Hence, use cases such as removing hydrogen from ambient air in the event of a hydrogen leak or extracting hydrogen from an exhaust gas are unfeasible. Although hydrogen purification and separation by MHs show promise in industrial applications, they are not further considered as use cases in this study due to their expected low potential for a direct use in aviation.

2.7 Hydrogen sensing

Some metals and their hydrides change their electrical, mechanical or optical properties, when being exposed to hydrogen. These characteristics can be used for hydrogen sensing purposes [Reference Bhuiya, Kumar and Kim7, Reference Hübert, Boon-brett and Buttner66, Reference Hübert, Boon-brett, Black and Banach85, Reference Yoshimura86]. MH-based sensors offer selective sensing behaviour and operate without requiring oxygen. The use of MHs as hydrogen detection devices in aviation safety applications was evaluated by Ref. (Reference Franke, Kazula and Enghardt87). The six investigated design principles of MH-based hydrogen sensors enable diverse sensing applications ranging from wide area monitoring to local leakage detection. While sensors based on some principles are already commercially available, other principles are still a field of research. The sensor properties depend on their design principle and can be generalised as follows:

• Electrical resistance sensors: wide detection range and good accuracy, but large size, high weight and high-power consumption

• Cantilever sensors: wide detection range, small size and inherently safe, but slow response and low maturity

• Nanogap sensors: fast response and high sensitivity, but poor upper detection limit and complex fabrication

• Fiber optic sensors: no magnetic interference, inherently safe and suitable for wide area monitoring, but low accuracy, slow response and high cost

• Chemochromic sensors: no infrastructure needed and low cost, but limited hydrogen quantification ability and need for periodic inspections

• Acoustic sensors: wide detection range, high sensitivity and fast response, but challenging manufacturing and low maturity

Despite their various advantages and disadvantages, none of the six design principles meets all the requirements to be used as a safety sensor in aviation [Reference Franke, Kazula and Enghardt87]. However, with continuous growth in hydrogen-based aviation and research progress in MH technology in parallel, MH-based hydrogen sensors may evolve further. As hydrogen leakage detection is expected to be a key technology for hydrogen-fueled aircraft, MH-based sensors should be investigated due to their promising properties [Reference Bray, Burrows and Hadnum88].

2.8 Use case selection for evaluation

Based on the review of possible MH applications, this study selects the following aviation-specific use cases for the evaluation:

1. 1. Hydrogen storage systems as emergency supply

2. 2. Cabin air-conditioning

3. 3. TMS system enhancement for FCs

4. 4. GGHS for thermal management of LH2 tanks

5. 5. Hydrogen boil-off recovery

6. 6. Onboard hydrogen compression

7. 7. Hydrogen safety sensors

3.1 Evaluation criteria

The evaluation criteria should be derived from the requirements list of a technical product, while essential minimum requirements should not be considered [Reference Feldhusen and Grote90, Reference Pahl, Beitz, Blessing, Feldhusen, Grote and Wallace91]. Recommendations and checklists with general requirements from [Reference Feldhusen and Grote8, Reference Kazula, De graaf and Enghardt30, Reference Feldhusen and Grote90, Reference Lindemann93, Reference Sadraey94] serve as guidance material in this study, resulting in the following criteria for the evaluation:

1. 1. Performance: The system’s power and operational behaviour are covered by this criterion. Potential synergies, when multiple tasks can be performed by the same device, increase this performance rating. Additionally, the mass specific powers and energies are compared to appraise also the weight of the system. Possible weight savings in the necessary amount of energy carrier are also taken into account.

2. 2. Ease of integration: This criterion contains installation space requirements and interface requirements like mountings, TMS, wiring and other necessary infrastructure. Moreover, the resilience against ambient conditions and interactions with components in the vicinity is included.

3. 3. Safety: As no hazardous events should occur due to the use of novel MH systems, this criterion considers failure types and their severity and probability. The potential to store hydrogen for emergencies or the ability to perform multiple tasks and thereby increase redundancy improves the overall aircraft safety.

4. 4. Cost: This criterion includes manufacturing and operational costs on the one hand and development costs on the other hand. Besides material costs, the manufacturing costs consider the producibility. Operational costs are influenced by consumables, lifetime or maintenance. The development costs are related to the technology’s maturity, as a low technology readiness level (TRL) implies increased development efforts and risks.

5. 5. Sustainability and Efficiency: As the mitigation of climate change is the main goal for hydrogen-powered aviation, the potential to reduce greenhouse gas emissions is considered by this criterion. This can be achieved by a high system efficiency and by improving the overall efficiency. The latter can be realised by the use of waste heat or by the use of the potential energy of compressed hydrogen. A high material utilisation in combination with recyclability is also valued by this criterion.

3.2 Criteria weighting

The significance of each criterion j is rated by weighting factors w _j . While these factors w _j are real, positive numbers in the range from 0 to 1, their sum must be equal to 1 [Reference Feldhusen and Grote8, Reference Feldhusen and Grote90, Reference Pahl, Beitz, Blessing, Feldhusen, Grote and Wallace91]. The weighting factors w _j are derived in Table 2 by a pairwise comparison. In this table, the criteria are benchmarked against each other according to their importance. The importance values are allocated in the bottom line [Reference Lindemann93]. Subsequently, the weighting factors w _j are obtained by summing the importance values for each criterion and dividing by the total sum of all criteria. The pairwise comparison is suited for this study, where the majority of criteria have to be appraised qualitatively and subjectively [Reference Feldhusen and Grote8, Reference Lindemann93]. As this comparison is conducted by a group of aerospace engineers, the ‘safety’ criterion dominates the other criteria because the safety assessment is a fundamental part in the system development process in aviation [Reference Moir and Seabridge48, 95, 96]. Besides safety, a lightweight design is considered to be a substantial requirement for aircraft systems. As the mass of the components and also weight savings in the amount of energy carrier are covered by the ‘performance’, this criterion also achieves high values in the benchmark.

Table 2.

Criteria weighting via pairwise comparison

3.3 Value assessment

The grade of fulfilment of the criteria is appraised with numerical measures m _i,j . When characteristics and performance parameters are inadequately known, a narrow range of m _i,j from 0 to 4 according to VDI 2225 is reasonable and therefore used in this study as follows [Reference Feldhusen and Grote90, Reference Pahl, Beitz, Blessing, Feldhusen, Grote and Wallace91]:

• very good fulfilment: m _i,j = 4 points = + +

• good fulfilment: m _i,j = 3 points = +

• moderate fulfilment: m _i,j = 2 points = o

• bad fulfilment: m _i,j = 1 point = –

• very bad fulfilment: m _i,j = 0 points = – –

3.4 Overall rating

By multiplying the grade of fulfilment with the weighting factors w _j and dividing by the scale size m _max, the relative rating r _i,j is calculated for each criterion j and each MH use case i, as shown in Equation (2) [Reference Feldhusen and Grote8, Reference Kazula, De graaf and Enghardt30, Reference Feldhusen and Grote90, Reference Pahl, Beitz, Blessing, Feldhusen, Grote and Wallace91]. The sum of all relative ratings r _i,j associated with a use case i results in the use case’s overall rating r _i . The variable k represents to the total number of criteria and n is the total number of possible solutions, which are the MH use cases in this evaluation.

(2) \begin{align}{r_i} = \mathop \sum \limits_{j = 1}^k {r_{i,j}} = \mathop \sum \limits_{j = 1}^k {w_j} \cdot {m_{i,j}}/{m_{{\rm{max}}}}, for\ i = 1 \ldots n\end{align}

3.5 Decision making

The most promising use cases for aviation will be determined with the help of the overall ratings r _i coupled with subjective appraisement to account for uncertainties or the influence of the choice and weighting of the criteria [Reference Ehrlenspiel, Kiewert, Lindemann and Mörtl89].

4.1 Cabin air-conditioning

To evaluate the potential of MHs for air-conditioning (AC), this section sizes a MH-based AC system and its energy consumption and benchmarks this system against a conventional vapor cycle machine (VCM). Figure 11 shows the topology of the MH-based system. A CgH2 storage tank is considered as energy carrier, since this enables the MH-based system to beneficially convert the potential energy of the compressed hydrogen to a cooling effect, which leads to low additional power demands. In contrast to the application of hydrogen compression in Fig. 8, the refilling of the MHRs is realised in Fig. 11 by CgH2 tanks. This high absorption pressure allows the use of MH materials, e.g. alloy 3 from Fig. 2, enabling heat uptake by desorption at temperatures below 0°C to generate a cooling effect. While the desorbed hydrogen is supplied to the FC, the cooling effect is used in the cabin heat exchanger (CHX) for air-conditioning. The ambient heat exchanger (AHX) rejects the heat of absorption of the MHR, which is being refilled.

Figure 11.

Topology of an air-conditioning (AC) system with 2 MH reactors (MHR), ambient heat exchanger (AHX) for waste heat rejection and cabin heat exchanger (CHX) for cooling.

With respect to Ref. (Reference Jiang, Dong and Zhang46), a power P _cool of 2kW is estimated to meet the heat load demands during cruise flight of the reference commuter aircraft. Even though the cooling requirements may be greater on ground [Reference Burschyk44, Reference Martínez101], the comparison solely pertains to cruise conditions, as the MH-based AC system necessitates a hydrogen mass flow to operate. To accomplish initial cooling on the ground, a stationary cooling system could be utilised instead. The comparison between the VCM and the MH-based AC takes into account the weight of the systems and their respective energy demand as well. The system weight of the VCM is calculated on the basis of a specific power SP _VCM of 0.20kW/kg, which is derived from literature sources [102–106]. Table 4 compiles the weight of the MH-based system, while Table 5 lists the energy demands of both systems’ components.

Table 4.

Component weights of an MH-based AC system with a cooling power of 2kW

Table 5.

Power demands of the components of both AC systems

While the blowers of the heat exchangers AHX and CHX generate energy demands for both AC systems, the energy consumption of the coolant pumps only needs to be considered for the MH-based system. On the other hand, the energy demand of the refrigerant compressor is only relevant for the VCM, which results in an overall higher energy consumption compared to the MH system. The total energy demand impacts the respective weight of the energy carrier and its storage system as well as means for energy conversion. According to Fig. 11, the CgH2 tank and the FC are considered to provide the energy not only for the aircraft propulsion but also for the AC system. Thus, the respective weight of the hydrogen fuel as energy carrier and the corresponding weight of the CgH2 storage tank are calculated for both AC systems. Therefore, a gravimetric storage density of 5.5 wt.% is applied for the CgH2 tank, which results in a mass-specific energy SE _CgH2 of 1.83kWh/kg [Reference Durbin and Malardier-jugroot109]. Additionally, the corresponding weight of an FC, which is needed to convert the chemical energy into electrical energy, is considered by a mass-specific power SP _FC of 0.15kW/kg [Reference Apeland110]. The weight results W of the comparison are displayed in Fig. 12, which is based on the Equations (3) and (4).

(3) \begin{align}{W_{{\rm{AC}},{\rm{MH}}}} = {\rm{\;}}\frac{{{P_{{\rm{cool}}}}}}{{S{P_{{\rm{MH}}}}}} + \frac{{{P_{{\rm{pump}}}} + {P_{{\rm{blower}}}}}}{{S{P_{{\rm{FC}}}}}} + \frac{{({P_{{\rm{pump}}}} + {P_{{\rm{blower}}}}) \cdot t}}{{S{E_{{\rm{CgH}}2}} \cdot {\eta _{{\rm{FC}}}}}}\end{align}

(4) \begin{align}{W_{{\rm{AC}},{\rm{VCM}}}} = {\rm{\;}}\frac{{{P_{{\rm{cool}}}}}}{{S{P_{{\rm{VCM}}}}}} + \frac{{{P_{{\rm{comp}}}} + {P_{{\rm{blower}}}}}}{{S{P_{{\rm{FC}}}}}} + \frac{{({P_{comp}} + {P_{{\rm{blower}}}}) \cdot t}}{{S{E_{{\rm{CgH}}2}} \cdot {\eta _{{\rm{FC}}}}}}\end{align}

Figure 12.

Overall weight comparison of a conventional vapor cycle machine (VCM) and an MH-based cabin air-conditioning system both sized for 2kW of cooling power.

The increased power demand of the compressor results in a higher corresponding FC weight for driving the VCM, which offsets the VCM’s higher specific cooling power to some extent. However, due to its lower specific cooling power, the MH-based AC system still displays greater weights for shorter cooling durations. For durations exceeding 2.5 hours, the MH-based AC system achieves an overall weight advantage over the conventional VCM owing to its lower energy requirements.

4.2 Boil-off recovery

For cryogenic LH2 tanks, heat transfer from the ambient environment causes continuous boil-off losses ranging from 0.2% to 3% per day, which have to be vented to avoid overpressure. These losses depend on the size of the tank and the quality of its insulation [Reference Züttel, Remhof, Borgschulte and Friedrichs3, Reference Baroutaji, Wilberforce, Ramadan and Olabi4, Reference Colpan and Kovač67, Reference Ahluwalia, Hua and Peng111, Reference Zhang, Zhao, Niu and Maddy112]. Larger tanks experience smaller losses through heat transfer due to their lower surface-to-volume ratio [Reference Schlapbach and Züttel113]. In addition to continuous boil-off, up to 15% of the transferred LH2 vaporises when re-fueling the storage tank [Reference Fujitani, Nishimura, Nishio, Sato and Yonezu75, Reference Petitpas114, Reference Ghaffari-tabrizi, Haemisch and Lindner115]. These transfer losses strongly depend on the filling method, the amount of delivered hydrogen, the size of the tank, the geometry of the tank, the insulation thickness and the hydrogen mass flow [Reference Petitpas114, Reference Ghaffari-tabrizi, Haemisch and Lindner115]. Therefore, the design of an LH2 aircraft tank involves balancing acceptable boil-off losses on the one hand with tank weight and insulation thickness on the other hand [Reference Brewer77]. For LH2 tanks in aircraft, a boil-off rate in the order of 0.1% per hour is acceptable [Reference Mital, Gyekenyesi, Arnold, Sullivan, Manderscheid and Murthy116]. As a baseline for the sizing, the reference commuter aircraft of Table 3 is considered to be equipped with an LH2 tank to store its hydrogen fuel. An exemplary loss of 2% of the total fuel mass is applied to calculate the necessary amount of MH material for boil-off recovery. Such a share is assumed to be a reasonable value for both the continuous boil-off loss per day as well as the additional loss per re-fueling. For the given tank size, this share of 2% corresponds to a loss 5.4kg of hydrogen.

Besides the quantity of boil-off losses, also the properties of the MH material and of the corresponding MHR affect the sizing results. For capturing low-pressure boil-off gas, the material LaNi₅ and its alloys are suited as they absorb hydrogen rapidly, reversibly and selectively even at relatively low temperatures [Reference Kölbig, Bürger and Linder36, 74]. For example, the LaNi_4.85Al_0.15 alloy has the capability of absorbing hydrogen gas at a pressure of only 2 bar when its temperature does not exceed 30°C [Reference Kölbig, Bürger and Linder36]. This ability enables the capture of boil-off from an LH2 tank, while the ambient air can carry out the heat removal of the exothermic absorption reaction. By transferring waste heat from an FC, which typically operates at a temperature level of 80°C, to the MH for the endothermic desorption, the release of hydrogen takes place up to a pressure of 4 bar. Thus, the MH can directly supply the FC with hydrogen [Reference Kölbig, Bürger and Linder36]. This enables the recovery of boil-off losses from the LH2 supply chain by capturing in an MH storage and re-using the hydrogen in FC-powered vehicles, as illustrated in Fig. 13. Moreover, MH storage systems offer the potential to swiftly replace depleted MH cartridges and are therefore a viable option for emission-free mobile applications such as utility vehicles [Reference Lototskyy, Tolj, Pickering, Sita, Barbir and Yartys11].

Figure 13.

MH cartridge system for boil-off recovery by capturing and re-using.

As the usable hydrogen capacity of the LaNi_4.85Al_0.15 alloy is approximately 1.1 wt.%, it requires 491kg of this material to capture the 2% loss of 5.4kg hydrogen. This amount of hydrogen is capable of supplying multiple utility vehicles such as forklifts, pushback tugs or towing tractors, as 2kg of hydrogen is sufficient for several hours of operation. For example, an FC-powered forklift could operate more than three hours with an average power of 10kW or a light duty industrial vehicle could cover a distance of 80km within 7 hours of operation [Reference Von colbe70, Reference Heung, Motyka and Summers117]. Therefore, it seems reasonable to locate a set of multiple cartridges near the LH2 tanks. A single cartridge with a hydrogen capacity of 2kg requires approximately 180kg of LaNi_4.85Al_0.15 alloy. To determine the total weight of the cartridge, a 1:1 ratio of MH material weight to empty cartridge weight is assumed for a lightweight cartridge design [Reference Bürger, Sourmelis terzopoulos, Kretschmer, Kölbig, Brack and Linder33, Reference Franke, Link and Kazula37]. This results in a total MH cartridge weight of approximately 360kg, which is an order of magnitude heavier compared to a CgH2 tank with a gravimetric storage density of 5.5 wt.% [Reference Durbin and Malardier-jugroot109]. Since utility vehicles require a minimum weight to ensure traction or to avoid overturning, the high weight of the cartridge is advantageous [Reference Lototskyy, Tolj, Klochko, Davids, Swanepoel and Linkov118].

4.3 Onboard hydrogen compression

Like in the previous section, the sizing of the onboard compression system bases upon a topology with an LH2 tank to store the hydrogen fuel of the FC-powered reference aircraft from Table 3. To achieve a low weight, LH2 tanks in aviation should operate at low pressures of 1.2–3.5 bar [Reference Brewer77, Reference Gomez and Smith119–Reference Khandelwal, Karakurt, Sekaran, Sethi and Singh122]. Typical hydrogen supply pressures for FCs range from 1.1 to 10 bar [Reference Lototskyy, Tolj, Pickering, Sita, Barbir and Yartys11, Reference Weckerle, Nasri, Hegner, Bürger and Linder12]. These pressure levels allow the use of a single-stage MH compressor as an onboard hydrogen compression system, since pressure ratios of a factor of 5 are feasible per stage [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Von colbe70–Reference Lototskyy, Yartys, Pollet and Bowman72]. The required size of this MH compressor depends on the achievable half-cycle time and on the maximum hydrogen mass flow demand. In this study, a half-cycle time of 60s is assumed for the sizing, which is considered to be optimistic [Reference Lototskyy, Yartys, Pollet and Bowman72]. The hydrogen mass flow $\dot{m}$ _H2 demand for an FC can be calculated with Equation (5). The variable ΔH _H2 corresponds to the lower heating value of hydrogen [Reference Weckerle, Nasri, Hegner, Bürger and Linder12].

(5) \begin{align}{\dot m_{{\rm{H}}2}} = {P_{{\rm{FC}}}}/\left( {{\eta _{{\rm{FC}}}} \cdot \Delta {H_{{\rm{H}}2}}} \right)\end{align}

The take-off power P _FC of the reference aircraft and its fuel cell efficiency η _FC from Table 3 result in a maximum hydrogen mass flow demand of 25g/s. Based on a usable hydrogen capacity of the MH of 1.1 wt.% in respect to the previous section, an amount of 136kg MH material will be required for the half-cycle of a single MHR. Since an MH compressor requires two MHRs for continuous operation, as shown in Fig. 8, the total amount of MH material for both reactors adds up to 272kg. Considering the same mass ratio of 1:1 to estimate the empty weight of a high-power, lightweight MHR like in the previous section, the total weight for both MHRs constitutes to 544kg. Although this appears competitive with conventional piston compressors for gaseous hydrogen, the weight far exceeds that of LH2 fuel pumps [Reference Brewer77, 123, 124]. For instance, a centrifugal LH2 fuel pump, which achieves a hydrogen mass flow of even 350g/s at similar pressure levels, weighs less than 2kg [Reference Brewer77]. Even when considering a potential reduction of the onboard power demand for compression, the MH compressor seems not competitive to an LH2 fuel pump due to this massive system weight penalty.

4.4 GGHS for thermal management of LH2 tanks

A common form of LH2 tanks is a shape with a cylindrical mid-section and spherical end caps [Reference Colpan and Kovač67, Reference Durbin and Malardier-jugroot109, Reference Ghaffari-tabrizi, Haemisch and Lindner115]. Figure 14 proposes potential layouts for the integration of the GGHS designs from Fig. 6 into the insulation of such an LH2 tank shape. The fuel mass of the reference aircraft from Table 3 results in a volume requirement of 4.1m³, when the hydrogen is stored as a cryogenic liquid [Reference Züttel, Remhof, Borgschulte and Friedrichs3, Reference Colozza and Kohout17]. Exemplary dimensions of an inner vessel to realise this volume are a diameter of 1.60m and a length of the cylindrical part of 0.97m. These dimensions result in an overall surface area of the inner vessel of 12.9m². This resembles the possible heat transfer area A, which is used in the following to estimate the weight and the thermal power of GGHS design (a). Before sizing this MH-actuated GGHS, the thermal power demand for vaporisation of the LH2 fuel is calculated. The heat requirement to vaporise LH2 is 446kJ/kg_H2 [Reference Colozza and Kohout17, Reference Arat and Sürer120]. Multiplying this value with the hydrogen mass flow yields the thermal power which is needed for vaporisation. The take-off mass flow of the reference aircraft leads to a maximum thermal power demand of 11kW to vaporise the LH2 fuel.

Figure 14.

Reference LH2 tank layout and proposals for GGHS integration: (a) concentric vessel design, (b) overlapping fin design in the support and (c) flat panel design in the insulation.

The temperature difference across the GGHS determines the heat flux and thereby the thermal power. While the temperature in the inner vessel of the LH2 tank remains constant at –253°C, at least for a full tank, the ambient temperature on the outside varies depending on weather conditions and altitude [Reference Züttel, Remhof, Borgschulte and Friedrichs3, Reference Arat and Sürer120]. For the purpose of a conservative estimate, an outside temperature of –34°C is selected, which corresponds to the air temperature of the international standard atmosphere at the aircraft’s service ceiling defined in Table 3 [Reference Matthews125]. These boundary conditions lead to a temperature difference ΔT of 219K and an average temperature of –143.5°C. The calculation is based on the assumption of a multilayer insulation, that effectively reduces radiation and convection. Therefore, only the conduction is considered for heat transfer in the ‘on’ state, when the insulation is filled with hydrogen. The conductive heat flow $\dot{Q}$ across a wall with multiple layers x can be determined using Equation (6). The variable g denotes the thickness of the layer, which can be a solid material or a gap in this case, while the variable k represents the layer’s conductivity [Reference Wagner126].

(6) \begin{align}\dot Q = A \cdot \Delta T/\left( {\frac{{{g_1}}}{{{k_1}}} + \frac{{{g_2}}}{{{k_2}}} \ldots \frac{{{g_{\rm{x}}}}}{{{k_{\rm{x}}}}}} \right)\end{align}

Since the solid layers of a multilayer insulation are very thin and their conductivity is high compared to the gas gap, their impact on the achievable heat transfer in the ‘on’ state is assumed to be neglectable. Accordingly, only the gas gaps are considered in this state in Equation (6). The mean thermal conductivity of hydrogen at the average temperature of –143.5°C and at a selected gas pressure of 1 bar is 0.088W/(m·K) [127]. Based on an assumed total gap width of the insulation of 5cm, the achievable heat transfer reaches nearly 5kW. For comparison, a continuous boil-off loss of 2% per day as stated in Section 4.2 corresponds to a heat flow of 0.028kW. With this parasitic heat flow in the ‘off’ state, when no hydrogen is present in the insulation, the ‘on/off’ conductance ratio calculates to approximately 180. As this ratio correlates to other studies, the heat flow estimations are considered to be reasonable [Reference Shu, Demko and Fesmire49, Reference Prina, Bowman and Kulleck52, Reference Martins60, Reference Vanapalli, Keijzer, Buitelaar and Ter brake62]. Based on the gaseous density of hydrogen of 0.187kg/m³ at –143.5°C and 1 bar, a hydrogen amount of 127g is required to fill the total gap volume [128]. To provide this amount of hydrogen by an MH alloy with a usable capacity of 1.1 wt.%, a minimum of 11.5kg of the alloy will be required.

Although the calculated heat transfer of GGHS design (a) of 5kW is not sufficient to cover the maximum hydrogen demand of the FC, this tank design may still be a viable option for generating the lower hydrogen mass flow during cruise flight [Reference Vietze and Weiland129]. Additionally, the achievable thermal power could be improved by decreasing the gap width or by adopting the GGHS designs (b) or (c) to increase the available heat transfer area. As a drawback, those actions result in a compromise as they may deteriorate the insulation quality and, consequently, lead to increased boil-off rates due to higher parasitic losses. An amplification in the applied hydrogen gas pressure to improve the thermal power is not reasonable, as the impact on the thermal conductivity would be neglectable [127]. Instead, the higher pressure would lead to an increased tank weight. However, risks and challenges induced by the cycling of the vacuum pressure and the component temperatures are considered to be major drawbacks of the GGHS integration into the LH2 tank insulation.

5.1 Outlook for boil-off recovery

Since the capture and re-use of boil-off losses with MHs is generally plausible, as shown in Section 4.2, this section investigates the losses along the LH2 pathway and projects an initial economic estimation of such MH-based hydrogen recycling. The illustration of Fig. 13 is a simplified schematic, as the occurrence of boil-off losses depend on the specific means of hydrogen transfer and distribution. Two different hydrogen distribution methods, a distribution by bowser trucks and a pipeline dispenser system, are illustrated in Fig. 16.

Figure 16.

Hydrogen distribution scenarios at airports: (a) bowser distribution and (b) pipeline dispenser system.

In the upcoming initial years of hydrogen usage at airports, the bowser distribution from Fig. 16(a) is likely the preferred scenario due to its lower capital costs [Reference Postma-kurlanc, Leadbetter and Pickard130]. For LH2 transfer from the bowser to a receiving vessel, the bowser is generally pressurised by an onboard vaporiser [Reference Michel, Fieseler, Meyer and Theißen69, Reference Petitpas131]. During transfer, the displaced gaseous hydrogen in the receiving vessel is typically vented and not recovered. Neither is the gaseous hydrogen from the bowser after transfer, when the bowser is de-pressurised again. Those venting losses increase for higher tank pressures and lower tank filling levels, as more hydrogen vapor is present [Reference Petitpas131]. The venting losses from the receiving vessel are likely the most significant along the LH2 pathway [Reference Petitpas114]. The venting losses from the bowser are not significant, when the applicable regulations allow to postpone the de-pressurisation to the time of arrival at the liquefaction plant [Reference Petitpas114, Reference Petitpas131].

With increased use of LH2 at the airport, a cryogenic pipeline system according to Fig. 16(b) may become beneficial [Reference Postma-kurlanc, Leadbetter and Pickard130]. Besides the pipe network with hydrants, such scenario will require refueling equipment to connect and disconnect the hose to the aircraft, to purge the aircraft tank and the hoses and to transfer LH2 by pressurisation or by pumping [Reference Brewer77, Reference Postma-kurlanc, Leadbetter and Pickard130]. The implementation of a hydrogen liquefaction plant and of a cold gas return system allows to recover the vented boil-off gas and thereby limits the losses to the re-condensation energy of the cold gas [Reference Colpan and Kovač67, Reference Brewer77, Reference Postma-kurlanc, Leadbetter and Pickard130].

Table 7 provides an overview of the boil-off gas sources of both scenarios [Reference Brewer77, Reference Petitpas114, Reference Ghaffari-tabrizi, Haemisch and Lindner115, Reference Mangold121, Reference Petitpas131]. Although the bowser scenario is disadvantageous in terms of boil-off recovery and mitigation, it has the potential to allow re-fueling without generating significant hydrogen losses [Reference Mangold121]. Therefore, the bowsers require additional devices like an extra gas storage with a compressor or like the proposed MH cartridge system from Fig. 13.

Table 7.

Sources of boil-off gas for both hydrogen distribution scenarios

^* If liquefaction plant is implemented at airport.

^** If LH2 transfer is realised by LH2 pump instead of tank pressurisation.

Depending on the size of the airport, a cryogenic hydrogen pipeline infrastructure with cold boil-off gas return may not be reasonable [Reference Postma-kurlanc, Leadbetter and Pickard130]. Especially in the near future up to 2040, bowsers will presumably distribute the LH2 to the aircrafts at all airports. While a pipeline infrastructure is expected to be beneficial for medium and large airports after 2045, smaller airports of 10 million passengers per year and below will still rely on bowsers for distribution of LH2 at the airport and even for the delivery of LH2 to the airport, if no liquefaction plant is implemented. In addition, the way of hydrogen delivery to the aircraft may also depend on its parking position, as it may not be reasonable to install cryogenic pipelines to remote positions of the airport apron. Thus, an infrastructure-independent solution for boil-off recycling remains reasonable in the future.

The release of hydrogen to the ambient should not only be prevented because of economic reasons but also because of ecologic reasons and safety issues. Hydrogen indirectly influences global warming and its emission would partially offset the climate benefit of savings in CO₂ emissions [Reference Warwick, Griffiths, Keeble, Archibald, Pyle and Shine132]. As released hydrogen is also a potential hazard, the gas is typically burnt by flare stacks, which implies siting issues on an airfield [Reference Brewer77]. To evaluate the economic motivation for hydrogen recycling, the prospect of hydrogen costs as well as the composition of hydrogen recycling system costs are assessed in the following.

The current cost of green hydrogen, which is produced with renewable energies by electrolysis, is in the range of 5 $\$ $ /kg_H2 [133, Reference Amica, Arneodo larochette and Gennari134]. Studies estimate green hydrogen prices of 3–4 $\$ $ /kg_H2 in 2030 and of 2 to 3 $\$ $ /kg_H2 in 2050 [133, Reference Menn135]. Liquefaction consumes 30–40% of the hydrogen’s chemical energy and adds 1 $\$ $ /kg_H2 [Reference Baroutaji, Wilberforce, Ramadan and Olabi4, Reference Gong and Verstraete65, Reference Durbin and Malardier-jugroot109, Reference Zhang, Zhao, Niu and Maddy112, 133]. Although this cost of liquefaction of 1 $\$ $ /kg_H2 may be applied for benchmarking a boil-off recycling system against a cold gas recovery pipeline according to Fig. 16(b), an overall cost estimation of such scenario is difficult, as the cost of the cryogenic infrastructure is interlaced with the airport infrastructure and operation procedures. Therefore, this study focuses on the comparison of the two infrastructure-independent approaches, the MH-based hydrogen recovery and hydrogen recycling by compression, and compares these two approaches to external venting of hydrogen by burning. As the latter results in a complete loss of the vented boil-off gas, the hydrogen cost of 5 $\$ $ /kg_H2 is applied for the estimation of today’s economic break-even-point of the boil-off recycling. A price of 2 $\$ $ /kg_H2 is applied as an outlook for the future.

As a baseline for benchmarking the economic reasonability of MH-based hydrogen recycling, the costs for hydrogen compression are assessed first. Therefore, the cost for the high-pressure hydrogen storage tank is considered with 500 $\$ $ /kg_H2 [Reference Baroutaji, Wilberforce, Ramadan and Olabi4, Reference Durbin and Malardier-jugroot109, Reference Zhang, Zhao, Niu and Maddy112, Reference Amica, Arneodo larochette and Gennari134, Reference Houchins, James, Huya-kouadio, Desantis and Habibzadeh136]. While the process of compressing hydrogen to a high pressure level is already energy intensive, the capital costs for the required components contribute to a significant share to the costs of the compression process [Reference Amica, Arneodo larochette and Gennari134, Reference Parks, Boyd, Cornish and Remick137]. This study estimates the costs for the high-pressure compression process with 1.50 $\$ $ /kg_H2 [Reference Parks, Boyd, Cornish and Remick137].

The cost estimation for MH hydrogen storage systems is difficult because of the lack of commercial, large-scale applications of such systems [Reference Grégoire Padró and Putsche138]. The cost of the MH material strongly depends on the composition of the alloy and affect the overall capital cost of the system significantly [Reference Von colbe70, Reference Amica, Arneodo larochette and Gennari134, Reference Ni139]. Earlier studies reported a wide range of MH hydrogen storage system costs from 500to 22,000 $\$ $ /kg_H2 [Reference Grégoire Padró and Putsche138, Reference Ni139]. To estimate the cost of an MH material, which is able to desorb the hydrogen by using the waste heat of a low-temperature PEMFC, this study scales the data from Heung et al. [Reference Heung, Motyka and Summers117] to an equivalent of US dollar in 2023. This results in MH material cost of 1,500 $\$ $ /kg_H2 as a lower estimate and of 4,900 $\$ $ /kg_H2 as a higher estimate. The cost for the empty MHR are assessed with 1,500 $\$ $ /kg_H2 [Reference Amica, Arneodo larochette and Gennari134]. Although the capital costs of the MH recycling system are higher compared to the recycling of the boil-off gas by compression, the use of waste heat saves operational costs and can lead to an overall economic benefit [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Lototskyy, Yartys, Pollet and Bowman72, Reference Yartys73]. Figure 17 compares the overall costs of both hydrogen recycling systems in relation to their usage in years y. The graphs are derived from the Equations (7)–(11), which assume a recycling system size with a capacity of 2kg_H2 and a charging and discharging frequency of two cycles per day or 730 cycles per year, respectively. The costs include the capital costs as well as an estimation for the operational costs. As the MH system is driven by waste heat, only minor amounts of electrical energy are required to control the system [Reference Von colbe70]. The chosen MH material should exhibit a high-cyclic stability to allow lifetime operation without the need to replace or regenerate the MH material to keep the maintenance costs at a minimum. Hence, the operational costs of the MH system are assumed with 0.10 $\$ $ /kg_H2.

(7) \begin{align}Cos{t_{{\rm{compression}}}}{\rm{\;}} = 2{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}{\rm{\;}} \cdot \left( {500\,{\rm{\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}} + 730{\rm{\;}} \cdot y \cdot 1.5\,{\rm{\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}} \right)\end{align}

(8) \begin{align}Cos{t_{{\rm{MH}},{\rm{high}}}}{\rm{\;}} = 2{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}{\rm{\;}} \cdot \left( {\left( {4,900 + 1,500} \right){\rm{\;\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}} + 730 \cdot y \cdot 0.1\,{\rm{\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}} \right)\end{align}

(9) \begin{align}Cos{t_{{\rm{MH}},{\rm{low}}}}{\rm{\;}} = 2{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}{\rm{\;}} \cdot \left( {\left( {1,500 + 1,500} \right){\rm{\;\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}} + 730 \cdot y \cdot 0.1\,{\rm{\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}} \right)\end{align}

(10) \begin{align}Cos{t_{{\rm{venting}},{\rm{today}}}}{\rm{\;}} = 2{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}{\rm{\;}} \cdot \left( {730{\rm{\;}} \cdot y \cdot 5.0\,{\rm{\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}} \right)\end{align}

(11) \begin{align}Cos{t_{{\rm{venting}},2050}}{\rm{\;}} = 2{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}{\rm{\;}} \cdot \left( {730{\rm{\;}} \cdot y \cdot 2.0\,{\rm{\$ }}/{\rm{k}}{{\rm{g}}_{{\rm{H}}2}}} \right)\end{align}

Figure 17.

Overall costs of boil-off gas recycling systems with 2kg of hydrogen capacity based on 2 charging and discharging cycles per day in comparison to the costs of external hydrogen venting.

Figure 17 shows a general economic advantage of boil-off recycling compared to today’s cost of hydrogen losses by external venting. For the lower hydrogen cost of the future, the payback of the hydrogen recycling relies on system lifetimes of 2 years and above. Furthermore, the graphs predict an economic benefit of an MH-based hydrogen recycling system compared to recycling by compression, when the usage time exceeds a duration of 2.5 years for the low estimate and 6 years for the high estimate. This implies that the use of MHs for hydrogen recycling is still plausible. Nevertheless, since the cost for the MH material strongly impacts the overall MH system cost, further studies should prospect for a suited MH material with low cost, flat plateau slope, low hysteresis, high cyclic stability, fast kinetics of hydrogen exchange and appropriate low temperature behaviour [Reference Muthukumar, Kumar, Raju, Malleswararao and Rahman6, Reference Lototskyy, Yartys, Pollet and Bowman72]. In addition, a detailed balance-of-plant investigation of the MH system should be performed to precise the forecast of the capital costs and the operational costs.

5.2 Outlook for hydrogen safety sensor installation in aviation

To mitigate the risks of hydrogen exposure in future aviation, a rigorous surveillance should be applied for aircraft and at airports. The use of various sensing principles in parallel is not only reasonable to enhance reliability through redundancy but also to exploit their specific properties [Reference Hübert, Boon-brett and Buttner66, Reference Hübert, Boon-brett, Palmisano and Bader140]. Besides MH-based hydrogen sensors, all other kinds of sensing principles should be considered during sensor selection. An exemplary hydrogen safety sensor installation is illustrated in Fig. 18. Fiber optic sensors can be advantageous for aircrafts because of their inherent safety by optical sensing and their installation flexibility [Reference Franke, Kazula and Enghardt87]. Difficult design spaces and hardly accessible components can be monitored for hydrogen leakage with a single evaluation unit and various sensing elements along an optical wire. After triggering an alarm threshold by an arbitrary sensor, additional chemochromic sensors can help to localise the leakage during repair and maintenance operations. Those sensing elements can be beneficially applied to hydrogen lines, fittings, flanges and connectors to deliver visual awareness of any hydrogen leak [Reference Roberson, Williams, Captain, Mohajeri and Raissi141]. For monitoring the vast spaces of the airport’s aprons, optical open path sensors or ultrasonic gas leak detectors are suited, as they enable wide area monitoring [Reference Hübert, Boon-brett and Buttner66]. Hydrogen leak localisation methods based on sensor networks and machine learning can assist first responders in identifying small leak zones in large confined spaces like hangars [Reference Zhao142].

Figure 18.

Exemplary hydrogen safety sensor installation on an airport.

5.3 Further thermal use cases of metal hydride heat pumps

Although a MH-based cabin AC system can already be competitive to a conventional VCM system as shown in Section 4.1, potential synergies and multiple use cases should be considered to further improve the overall aircraft weight specific performance by replacing other dedicated systems. For example, the MHRs of an AC system can also serve as a heat source for pre-heating during a cold start scenario as well as for de- and anti-icing of components like fuel cells, sensors, valves, lubricants, intake ducts or the leading edges of the wings [Reference Moir and Seabridge48]. The latter is especially beneficial, if the wing surfaces are used as the AHX of the AC anyway. Current state-of-the-art anti-icing devices for wings either deposit a freezing point depressant fluid on the wing surface or heat the leading edge by the engine bleed air or by electric power [Reference Moir and Seabridge38, Reference Moir and Seabridge48, Reference Bray, Burrows and Hadnum88]. As bleed air will not be available in various electrified propulsion system topologies and as electrical heating is energy demanding, MH-based pre-heating can be an appealing alternative [Reference Moir and Seabridge48, Reference Bray, Burrows and Hadnum88].

To provide the desired heating effect by hydrogen absorption, the MH material needs to be charged with hydrogen from one of the aircraft’s hydrogen storage tanks. Therefore, at least some share of the MH material has to offer free capacity for hydrogen absorption. This is guaranteed, when the MHRs were operating as an AC system at the end of the previous flight mission. During this continuous alternating heat pump operation, only half of the MH hydrogen capacity is occupied with hydrogen. When the MH material is fully charged with hydrogen after a pre-heating application, it needs to be regenerated to its half-charged state before being able to operate as a heat pump again. This regeneration can be performed by desorbing hydrogen to the FC and by providing its waste heat to the MH. When the operator already expects an upcoming cold start scenario for the next scheduled flight of an aircraft, it is reasonable to fully discharge the hydrogen capacity of the MHR at the end of the previous flight mission. Thereby, the maximum amount of heat is stored in the MH material to be provided as thermal energy for pre-heating.

The application of MH heat pumps to enhance the TMS of FCs achieved a low rating as a stand-alone use case mainly because of its high weight penalty. However, an MH-based AC system can also serve as a heat sink. Although the heat storage capacity would be much lower compared to a dedicated system for the ‘TMS of FCs’, it can still be reasonable to use the MH-based AC system as a short time cooling booster in an emergency situation. In addition, the MH-based AC system can even serve as an emergency hydrogen supply, as half of the MH hydrogen capacity is always occupied by hydrogen during its continuous alternating operation.

Besides air-conditioning of the cabin, also the cooling of the galley by an MH heat pump should be investigated in further studies. However, as the potential energy of the compressed hydrogen is necessary to drive the heat pump operation, the application of such systems is limited to aircraft with CgH2 storage systems. This narrows the applicability in the aerospace industry, as LH2 is the most promising storage type for hydrogen in aviation [Reference Baroutaji, Wilberforce, Ramadan and Olabi4, Reference Colpan and Kovač67].