Design Study for a Superconducting High-Power Fan Drive for a Long-Range Aircraft

Abstract

New aerodynamic aircraft concepts enable the storage of volumetric liquid hydrogen (LH2). Additionally, the low temperatures of LH2 enable technologies such as the superconductivity of electrical fan drives and power distribution components. An increased power density of the onboard wiring harness and the electrical machine can be expected. The highest system efficiency and the smallest fuel and tank weight will be achieved with a highly efficient energy conversion by the fuel cell from LH2 to electrical energy. This publication shows a comprehensive study for cryogenic fan drives based on experimental-driven tape superconductor investigations, mission profile-based considerations, design analyses of superconducting electrical machines, and studies of the cooling concepts. A cryogenic system cannot be considered without a feasible cooling concept. Here, an approach with a safe He-based cooling system is proposed, using the LH2 flow to the fuel cell as a heat sink for the losses in the electrical system.

1. Introduction

Future emission-free long-distance commercial aircraft have to use the synergies between distributed electric drives and aerodynamic improvements to save energy, as proposed in [1,2]. The study [3] is based on the proposals of [1,2] and illustrates the multi-fidelity design optimization of the long-range blended wing body in the sustainable and energy-efficient aviation (Project: Strategy-EXC 2163/1, SE²A consortium; Figure 1). These aerodynamics improvements can enable increased fuel efficiencies compared to designs based on the conventional tube–wing design with two or four propulsors. Here, some new air-frame technologies are used; for example, active flow control, active load alleviation, body layer ingestion (BLI), new materials, and structure concepts. Additionally, the distributed electrical propulsion (DEP) necessary for BLI enables a highly redundant drive system.

As it was shown in [4], for an electrical drive system with conventional copper winding, the optimal current density in the winding with respect to system weight balance including the cooling devices is approximately 30–40 A/mm², yielding a power to weight ratio of approximately 15 kW/kg of the active material. This value is nearly halved if the power electronics and cooling system are taken into account. A more preferable option then, is electrical machines with superconducting windings. As already discussed in [5], cryogenic hydrogen (LH2) is the most suitable approach for the storage of emission-free fuel. For superconducting machines, power densities at least of 20 kW/kg [6] can be expected. In some papers, up to 55 kW/kg [7] is reported. As the power losses are much smaller with superconducting machines, a much lighter and smaller cooling system can be expected, using the hydrogen flow to the fuel cell as a heat sink.

The preferred aircraft design offers space for volumetric cryotanks because of LH2’s poor volumetric energy density [5]. Here, the feasibility of the power electronics is also discussed in detail and an approach for a lightweight highly efficient 20 kHz high-power inverter presented.

In [8], possible solutions for the challenges of the onboard electrical power system are proposed. Wide band gap semiconductors, hybrid circuit breakers, and a DC voltage of 3 kV are the main key words. Future research and development will also focus on a feasible voltage range. Superconductors might be the solution for voltage reduction. As a power source, fuel cells combined with an advanced battery for peak loads are proposed as the most efficient (approximately 55–60%) solution. This requires high-power batteries with a high capacity of more than 2000 Wh/kg and 500 Wh/kg which will be available in the near future. An intermediate solution could be the combination of hydrogen-fueled gas turbines with superconducting generators. This design would have the disadvantage of significantly lower efficiency (approximately 35%) than a fuel cell/battery system. In [9], the control of a fuel cell/battery system is studied and in [10] a fuel cell-powered experimental small aircraft is presented.

As so far numerous technical challenges but no principal obstacles from physics have been recognized, the focus is now centered on the design study of superconducting electrical machines for fan drives.

2. Mission Profile

The operating conditions form the basis of the design of electrical machines. In the aviation sector, mission profiles have to be clarified to define the boundary conditions and design parameters. Here, a mission profile is calculated using an optimization methodology like the one shown in [3]. Figure 2 shows the calculated mission profile of SE²A long-distance aircraft. Four load cases (LCs) with their Mechanical Shaft Power (PShaft) specifications are sufficient to study the system’s feasibility.

The profile displays a peak demand during climb and a long constant power demand during cruise. Additional peak power may be required during the landing phase if the landing procedure is aborted and a go around is necessary. Depending on the response time of the fuel cell, a relatively small battery could cover this peak power demand, but a conventional gas turbine drive also has a considerable response time The combination of battery and fuel cell yields potential here for future optimization. Charging this battery is easily possible during cruising.

3. Electrical Drive System

In this design study, twelve superconducting electrical drives powered by a highly efficient fuel cell supplied with hydrogen from cryotanks are proposed for propulsion. As the energy conversion in a fuel cell system has a better efficiency than in a turbo generator system, one finds the smallest mass flow of hydrogen with a fuel cell what defines the critical situation for the design of the cooling system of the superconducting components. That is why the focus of this study is on fuel cell operation.

Table 1. LCs based on the mission profile analysis.

Load Case		Total Shaft Power	Shaft Power per Propulsor
		MW	MW
LC1	Peak Rating	64.8	5.4
LC2	Climb	54	4.5
LC3	Cruise	37.4	3.12
LC4	Landing	6.24	0.52

concludes the mission profile in four load cases (LC) to define the specifications for the design of the drives.

The designing LC will be the peak power (LC1) because of its losses. The time of the climb phase is too long to operate the powertrain in an overload strategy, as used in automotive drives to increase the power density of electrical components. A further challenging LC is the cruise (LC3) because this has the longest operation time of more than 700 min. Therefore, the electrical powertrain has to show good efficiencies at partial loads. At landing (LC4), the average power consumption is low but the hydrogen flow to the fuel cell is also the smallest. The losses and the temperatures in the cooling system must be balanced thoroughly here.

As discussed in [5,8], a 3 kV intermediate circuit is suitable for the high power demand with respect to the currents and blocking voltage capability of modern SiC MOSFET semiconductors. This leads to a 9 or 12 phase winding in the electrical machine combined with 3 or 4 three-phase systems. The additional advantage of multiphase winding is a fairly high redundancy in case of failures in the inverter or the winding of each machine. Combined with the 12 propulsors of the hybrid wing–body design a double redundant drive system is achieved.

4. Calculation Approaches

In [7,11], a theoretical design approach with a fully superconducting machine with an extremely high power density of 55 kW/kg was calculated. Here, MgB₂ HTSL filaments in a copper matrix were considered at a temperature level between 25 and 30 K with a transition temperature of MgB₂ of 40 K. Using YBCO band conductors promises further weight savings and reduced AC losses if cooled down to the 30–40 K level. Regarding the wide temperature span between the temperature level of LH2 (20–25 K) and the transition temperature of YBCO conductors of 90 K, a very high safety margin against quenching is given and high current densities under AC fields can be expected.

There are multiple challenges when designing a machine with superconducting stator (AC) and rotor (DC) windings:

• The superconducting state depends on the current, the temperature, and the flux density that the superconductor has to face. Assuming the critical state model of type II superconductors with critical temperature T_c, critical field strength H_c, and critical current I_c, the transport current has to remain below the level of the critical current in all load cases. AC losses in the superconductor are mainly hysteresis losses, while in the covering copper or silver layers, eddy current losses occur. To calculate the safe transport currents in winding and excitation depends on the knowledge of the field components inside the wire.
• The design of the winding has to ensure that these losses are safely removed by the cooling system.
• The cooling of the machine with cold hydrogen or helium of temperature 25–50 K has to be designed taking into account the thermal restrictions of the material. The cooling of the rotor winding is also achieved by hydrogen or helium circulating through the rotor structure while rotation supports the transport of the cooling fluid. In any case, ice forming on cold parts of the machines or the supplying pipes has to be strictly avoided.
• Although a large cooling power is available with cryogenic hydrogen or helium circulating through the drives, the thermal insulation technology has to be designed carefully to avoid a too high introduction of heat from the ambient into the cryogenic system. Here, proven technologies can be used. This also reduces the size and weight of the cryo pipes feeding the coolant to the drives.
• The electrical insulation system has to take into account the different dielectric behavior of the cooling gas.
• Superconducting machines exhibit different behavior when subjected to fault situations such as short circuit or dynamic load changes. As the windings have nearly no resistance, no damping effects can be expected, except additional damping is provided for fault situations. Under normal operation, damping is provided by the control in the inverter as with conventional synchronous machines as well. The large magnetic gap of superconducting machines consequently has a small inductance of the stator winding, causing large short circuit currents and rather high oscillating torques in the shaft which could affect gear and fan/propeller construction.
• As the drives are fed by superconducting DC cables connected to the inverter and from there an AC cable connection to the machine is required, it would be advantageous to have the inverter inside the cooling system operating at the 200–300 K temperature level as shown in [5]. Here, the properties of the semiconductors pose a strict limitation to the cooling design of the inverter.
• With the small inductance of the machine, high switching frequencies of the inverter are required to achieve currents in the windings with acceptable time harmonics.

In this paper, the machine design of a synchronous machine with superconducting windings made of tape conductors in the rotor and the stator under the restrictions found in [5] is discussed.

5. System Components

5.1. Electrical Machine

In order to provide the technical solutions for the stated requirements, all 12 drives combined in this aircraft design have to provide, i.e., the total power and thrust. The boundary conditions for a single drive are presented in

Table 2. Design data of the electrical machine.

Max. continuous power	5.4 MW	Start and climb flight
Max. continuous torque	10,320 Nm	No field weakening
Power at travelling speed and height	3.12 MW
Max. speed	5000 1/min, Ω = 523.6 1/s
Fan diameter	Dd = 1.2 m

.

A further feature to reduce the weight is the choice of a Fe-Co lamination (saturation flux density 2.4 T).

5.2. Electromagnetic Design

According to the above requirements, a suitable feasible electromagnetic design has to be found and investigated. Initially, the machine type must be determined as a basis for subsequent steps. A high power factor of the machine is crucial to minimize the mass of the inverter. Therefore, the choice falls on the synchronous machine type. Important boundary conditions for the hts drive in this research are to use presently available tapes and to meet the requirements for a gas cooling system with the flow of hydrogen, i.e., to the fuel cells as a thermal sink. As the temperature range of REBCO tape is larger and the Tc is also higher than in, i.e., MgB2, in this investigation REBCO tape has been used. Choosing the temperature range of the hts conductors is vital for the loss calculation. Preferably, gaseous Helium is being used in this study as a direct coolant (cp ≈ 20.8 J/mol·K, λ ≈ 0.04 W/m·K). An alternative could be Neon (vapor cp ≈ 21 J/mol·K, λ ≈ 0.009 W/m·K, for the liquid state the cp values rise to an approximately double value), but Helium is safer in case of possible cooling errors to the lower temperature range of 20 K of LH2 and it also provides a broader temperature range at similar physical properties. Direct cooling with liquid or gaseous hydrogen would of course be desired from a thermal perspective, as the cooling medium would enable a very low temperature range, thus increasing the lift factor of the superconducting tape and increasing the critical current density. On the other hand, the piping system between heat exchangers and the 12 machines can be of considerable size with a lot of flanges, seals, and extension compensators, which are subject to vibrations and ambient temperature changes during flight. So, to avoid the risk of fatal leakage, an inert cooling gas is the best choice. On the other hand, direct hydrogen exposition might cause also material damage to the hts tape and other motor parts. It also can severely damage permanent magnets, if these should be used on the rotor and are directly exposed. Figure 3 shows the principal arrangement and measures of the drive model for the investigations.

Superconducting drive systems can be designed in various ways. One main subject is the design of the magnetic circuit and the flux path. The magnetic field can be conducted either through an air-volume (vacuum) or through an iron yoke. The maximum power densities of a superconducting drive can be realized by building lightweight ironless motor. On the rotor side, excitation can be realized with permanent magnets (i.e., Halbach array) or superconductors. Superconducting rotor excitation has the advantage of the possibility to control the excitation, but difficulties in cooling realization can arise.

As depicted above, the chosen design utilizes iron yokes and places the coils in between. The central idea here is to place the stator coils in a fiber-reinforced cartridge system inside the stator yoke, only cryo-cooling the coils and leaving the iron at atmospheric temperature, cooled, i.e., by the airflow of the fan. A form fit combined with an elastic adhesive would enable the torque transfer to the stator yoke. Although iron yokes are disadvantageous in terms of power density, there are other very beneficial factors. Avoiding air coils reduces costs, as less tape length is needed. The permeability of iron, or preferably cobalt–iron, even if close to saturation, is substantially larger than that of air. This enhances the magnetic field significantly. An air-coiled system requires a higher magnetomotive force, which only can be provided by increasing the coil turns (leaving aside higher currents) and thus length of superconducting tape. Moreover, an iron yoke also provides mechanical stability and robust support in vibration environments. Providing magnetic shielding is another important property. Fringe fields are reduced, which contributes to safety and reduces interference with electronic devices.

The simulations performed in this work showed that circumferential flux densities are significantly higher in ironless designs, leading to considerably increased losses in the superconductors. The investigated design uses a warm iron yoke at the expense of a lower power density, but on the other hand this leads to the gain of an improved loss situation throughout the operating range, which can make use of the on-board heat sink in the form of LH2. By this strategy, the cryo-cooling system can be designed without contributing larger masses to the total weight. The iron losses in the warm yoke can easily be removed by the airflow of the fan.

Another important design element is the winding. Contrary to a classical copper wire, superconducting tape cannot be deformed in a way that exceeds elasticity and deformation limits; or in other words, it cannot be bent into a free-form shape, as is common in, i.e., automotive hair-pin windings or classical windings, where the end winding often must provide a change in layers to be sufficiently compact. Here, a two-tier winding concept with concentric coils (if q > 1) was chosen to eliminate crossings in the end windings. In this nine-phase case there is only one coil crossing the layers. The two-tier configuration can simplify the winding process, making it easier to automate and reduce the potential for errors. A further advantage is that this winding type is operating at a fundamental frequency and less harmonics of the B-field in the airgap are generated. The proposed principal coil arrangement is shown in Figure 4. The coils are placed in two separate layers. The coil assemblies are placed into a fiber-reinforced fixation structure, which in turn is mounted inside the stator yoke. The green-colored structure enables the gas flow, it insulates the cryogenic parts against the environment and absorbs the high forces of the torque-generating propulsion drive.

Flight safety is generally increased by relying on redundant systems. Electric drive systems can contribute to this by adding more phases to the stator. In this example, a redundant nine-phase current supply is used. In Figure 5 and Figure 6, the arrangement is displayed for a six-pole machine with independent inverter and coil systems, one can notice the large airgap (yellow) between the iron yokes (blue) and the two tiers of coils; here, an example is using 2 mm tape. The red areas are sections for rotor excitation.

The condensed results of the machine design fulfilling the discussed requirements are shown in

Table 3. Technical data of the resulting electrical machine.

Rotor circumferential speed	94 m/s
Number of pole pairs	3
Number of phases	9
Max. feeding frequency	250 Hz	superconducting tape winding
Intermediate circuit voltage	3 kV	reinforced insulation, a lower voltage design is also possible
Phase voltage	1280 V
Specific tangential force	275.5 kN/m2
Power density, continuous operation	32 kW/kg
Efficiency	>99%	electric drive

.

The air gap field components play a significant role in building forces and torque in the electromagnetic system. Circumferential field components are shown in Figure 7 for a fixed airgap position next to the superconducting windings. When performing a Fourier analysis of the two curves of Figure 7, the first and third order are shown to be dominant. The first order varies with the current and increases in this example from around 105 mT at full load to approximately 195 mT when setting the drive current to zero. These variations in the circumferential field depend on the stator current, radial position, the rotor field, winding topology, phase angle, and the presence or absence of an iron yoke. In this example, losses due to tangential field increase with decreasing current. One countermeasure could be to shift the phase angle of the drive current to field weakening or to decrease the rotor current in case of an externally excited system.

5.2.1. Losses in the Stator Winding

In the first step, preceding the calculations, measurements were performed for the 2 mm tape at 77 K to generate a base for the further calculations and simulations. The tape is a narrow 2 mm tape from Faraday Factory Japan. As the central idea is to reduce losses, this tape was chosen because it is not coated with copper to reduce further eddy current losses. HTS-tapes are frequently coated with a further copper layer, i.e., with 40 µm thickness to improve current tolerance and handling, but in motor coils this leads to additional losses when exposing the conductors to an alternating perpendicular field. The placement of the coils in slots can reduce these fields, but this depends on the desired power to weight ratio of the system. Only a 3 µm silver layer is deposited on the hts layer which typically has a thickness below 1 µm. This is necessary for contact and protection of the superconductor. The investigated coil is an air coil, losses in this setting are mainly resulting from the transport current and local perpendicular field components due to the coil self-field. Losses are detected by measuring current, voltage, and phase angle, which is achieved by a power analyzer. The principal arrangement is shown in Figure 8.

An overview of the real setup is given in Figure 9. The coil is placed in the center, submerged in liquid nitrogen. To the right, there are voltage test leads and high current connectors for the tape.

Measurements were performed with varying sinusoidal currents up to 30 A rms, resulting in a 0.42 ratio of the peak current to I_c. The testing frequencies ranged from 50 to 500 Hz, with a short test of 1000 Hz, but only up to 12.5 A rms to reduce the risk of coil failure. As can be seen in Figure 10, the total coil losses rise exponentially with the current, but due to the narrow tape they stay in a low range. A linear dependency of the losses from frequency can be seen up to 300 Hz. Above 300 Hz, the rise of measured losses differs from a linear context. This could be due to the high ratio of $I/I_c=0.42$ and the increased frequencies or local temperature increase. A similar behavior at 77 K is observed in [12]. Further measurements at lower temperatures, i.e., 30 K will be performed in the future. For the investigated machine design, a 250 Hz max. frequency produces losses which are in the range of the cooling capability of the system and lead to the high efficiency of the drive. The low losses can already be viewed at 77 K in Figure 10 and are estimated to decrease at a design temperature of 30 K. The 2 mm, but also the 3 mm tapes, have the advantage of a low height in the airgap and they can be formed into coils easily. A glimpse on potential problems for later constructions of superconducting low loss ac-coils could be noticed during testing. As the thermal capacity of these thin tapes is very low, a coil voltage/current observation has to be implemented combined with a fast shut down for the winding. Furthermore, a temperature scan of the coil has to be performed at ac load to qualify the coils and to detect the slightest errors of the hts layer. There is definitely a need for future online ac coil error detection procedures and their development for electric hts drives.

The losses in the stator winding are composed of hysteresis and eddy current losses, depending on the two-dimensional field situation in the air gap of the machine. As the circumferential speed of a superconducting machine is not extremely high (below 150 m/s) aerodynamic friction losses will be neglectable, and only bearing and iron losses will remain as losses.

The tape losses were calculated with the help of numerical calculations and checked via the analytical equations given by Mawatari [13,14], based on the work of Brandt [15] and Halse [16]. The transport losses for an infinite vertical stack of HTS tapes can be calculated in J/m per cycle by

$$Q_t={\displaystyle \frac{{\mu }_0{Ic}^2{i}_0^2}{\pi }}{\int }_0^1\left(1-2s\right)ln\left[{\displaystyle \frac{{cosh}^2\left(\frac{\pi w}{L_y}\right)}{{cosh}^2\left(\frac{\pi i_{0}sw}{L_y}\right)}}-1\right]ds$$

(1)

With I_c: critical current; i_0: percentage of I_c; w: half tape width; L_y: distance between tapes.

The hysteretic part for an infinite vertical stack of the losses with a maximum magnetic field strength H₀ is calculated by

$$Q_h={\left({\displaystyle \frac{L_y}{\pi w}}\right)}^2{h}_0^2{\int }_0^1\left(1-2s\right)ln\left[{\displaystyle \frac{{sinh}^2\left(\frac{\pi w}{L_y}\right)}{{cosh}^2\left(h_{0}s\right)}}+1\right]ds$$

(2)

With h₀ = $\pi$H₀/j_c d; d: tape thickness; j_c: critical current density.

For verification of the analytic equations, a model for stacks of superconducting tape was generated, as shown in

Table 4. Data for the stator winding.

Tape type	REBCO
Tape width	2 mm
Tape thickness	<70 µm
Total tape length (stator)	2110 m (2 tapes parallel)
Ic 77 K	84 A
Coating	2 µm Silver
Fundamental frequency	250 Hz

. (REBCO tape) (Figure 11). Two important insights could be drawn from this comparison: analytic equations for the stack already yield good results and they only lack additional losses of the end tapes.

As can be seen in Figure 8, the external field can only partially penetrate the stack, unlike, for example, in copper conductors. This reduces losses in the stack. Only the tapes situated at the right and left of the stack show higher values, i.e., factor 3 to center tapes.

The results of the loss calculations for the stator tape are shown in Figure 12. The displayed total losses have been calculated for different current loads, circumferential fields, and parallel tape numbers, in order to gain information on the influence and to determine a suitable compromise regarding tape length and losses. Tape losses in electric drives with constant rotor excitation do not reach zero at low currents because of the alternating rotor-field in relation to the stator. Higher currents lead to exponentially increased losses which can clearly be observed for a single tape carrying the total current. A reduction in transport current in relation to the critical current reduces losses. This can be achieved by employing multiple parallel tapes to divide the total current, which results in a compromise between losses and tape length. In the case of an aircraft propulsion system relying on LH2 as energy source, the aim is to design the loss situation according to a feasible heat flow. For the following investigations, the number of parallel tapes has been set to two, as this forms a good compromise between low losses at low power demands and losses at high power demands. This is also important regarding tape length. Adding one parallel tape to the winding will add approximately 1000 m to the total tape length.

Tape width also contributes to the loss situation. A variation in the loss calculation for 2 and 3 mm HTS tape material was performed. The difference is shown in Figure 13. Negative values show a lower loss for the 3 mm tape, which can clearly be seen in the left side of the diagram for lower numbers of tapes.

A drawback of a broader tape in addition to mass and additional HTS area is the loss situation at lower stator currents, where the rotor excitation is dominant. Therefore, a controllable excitation would be very beneficial.

The temperature also has an impact on losses, as can be observed in Figure 14. The comparison was made for 30 and 40 K. At loads below 60%, at 40 K, lower losses are observed. Although lower temperatures offer a higher lift factor and thus a higher I_c, higher temperatures can be beneficial at the C-axis magnetic field load. This temperature range was chosen to enable safe operation. It could also be possible to further increase the temperature of the tape at a higher load. For this investigation, one can state that the cooling system is able to transport the heat flow, so a temperature range up to 40 K can be used.

The resulting HTS-losses for each load-case are shown in

Table 5. Losses in the stator winding.

Load Case	Losses [W]
1	4011
2	2320
3	1208
4	507.6

. Additionally, heat from the ambience entering the cryogenic part of the machine must be considered. In this case it is estimated to be 100 W in addition.

Compared to gas turbines, electrical machines are not subject to atmospheric conditions such as temperature and ambient pressure with respect to their power output. That is why they can be operated on the natural quadratic torque speed characteristic of a fluid flow machine. For each operating point, torque and speed can be adjusted to the requirements of the driven fan. As these vary with air pressure and temperature, each operation point has its own cubic load parabola for the power of the drive, due to the current flight altitude. The maximum power requirement is found on the ground level. As the fan design is actually not defined, the losses for constant speed are calculated, which puts the loss results on the conservative side.

5.2.2. Losses in the Rotor Winding

In the shielded rotor winding (DC), neglectable losses are expected, which can be transferred to the gas flow in the mechanical air gap.

6. Cooling System Design

Consideration of the cooling system is of vital importance to the implementation of superconductivity. Here, it is assumed that the safest approach is to feed the hydrogen using the shortest possible route from the tank to the fuel cell or the turbo generators. As a cooling gas, helium or neon may be considered. With respect to viscosity and thermal conductivity, helium seems to be the most viable alternative, although its heat capacity is smaller than that of hydrogen. To use the hydrogen flow as heat sink, the helium cooling circuits are then coupled to the hydrogen flow by heat exchangers. As the losses of the superconducting machines and the cables are very small, the heat exchangers will have a moderate size and weight. The number of required heat exchangers depends on the design philosophy and the risk assessment of the propulsion system. If each machine with its inverter is coupled to the heat sink by its own heat exchanger, the most redundant propulsion system is achieved. Gas cooling with hydrogen was previously used successfully in power station generators of the 300 MW class. Compared to a liquid cooling system, gas cooling enables a more simple cooling system, i.e., without space consuming pipe systems in certain cooling areas.

These conceptual considerations also fit the ideas presented in [5], where the power electronic inverters are positioned close to the electrical machines but are fitted with a separate cooling circuit.

In the machine, the heat transfer to the cooling gas contains three components: the heating of the cooling gas by the transferred power, the temperature drop between cooling tube walls and cooling gas, and the temperature drop inside the winding supporting structure and the superconducting tapes.

7. Cooling Design and Simulations

Several proposals were made for hydrogen- or helium-cooled machines; e.g., in [17], a capillary cooling of the superconducting coils with hydrogen in a two-phase (vapor, fluid) state is proposed. As hydrogen is excluded for safety reasons, and helium is always in the gaseous phase in the relevant temperature range, this approach is not applicable here.

In [5], the total mass flow of hydrogen was already calculated for the example considered here, assuming an efficiency of the fuel cell of 55% and 99% each for the machines, inverters, and the DC backbone. This also incorporates the 500 kW demand of the on-board electric system. In between, higher efficiencies of machines and the DC backbone are found and investigated, yielding the modified data shown in

Table 6. Estimated required ${\dot{m}}_H$.

Load Case	Total Power	Total H2	H2 Flow per Engine	He Flow per Engine
	(incl. Subs.)
	MW	g·s−1	g·s−1	g·s−1
LC1	122	1016	84.7	76
LC2	102	847	70.6	45
LC3	70.7	588	49	30
LC4	12.2	102	8.5	20

for the load cases of Table 1. The helium flow given here is the smallest flow that allows temperatures in the winding between 30 and 40 K. It was calculated with the help of numerical CFD simulation.

The cooling method of the tapes is a central design point of hts machines. The design depends on the losses on the one hand, and on the other hand, as hts motor tape coils fulfill multiple functions, it also depends on voltage, forces, and oscillations. Furthermore, thermal expansion must be regarded. A simple approach would be to integrate pipes which follow along the tape and thermal conductors to extract the heat. This can be feasible for weaker heat fluxes. Additional thermal conductors, especially metallic types, increase losses and package sizes. The concept presented here cools the hts components, and the iron parts are left warm. The coil components have to be mounted into a housing, which permits a gas flow and provides mechanical stability, as well as insulation. Experience in thermal insulation has also been gained in the construction and assembly of a high-load superconducting bearing [18].

The general aim of the cooling design is to generate a high heat transfer coefficient at the lowest possible mass flow and lowest pressure drop to decrease pump power, mass, and also heat exchanger mass. The proposed solution to this is to drastically parallelize the flow of the cooling gas by multiple cooling channels parallel to the winding axis of the coil. As helium molecules are very small, these channels can have small diameters of some tenths of a millimeter. Figure 15 shows an overview of the principle for four parallel tapes of a racetrack coil with integrated cooling slots.

Although the flow in these capillary tubes is nearly laminar, analytical calculations show that the difference between the temperatures of cooling gas flow and tube wall is striving towards zero, yielding high heat transfer coefficients of more than 1000 W/(m²·K) at a pressure drop of approximately 10 mbar. This was also confirmed by CFD calculations (Figure 16).

The slotted spacer is part of the coil and is integrated into each winding. As the gas flow is perpendicular to the winding direction, the coil has to be placed in a housing equipped with a small gap for the gas transport. All parts can be manufactured glass or carbon fiber-reinforced, which would also be beneficial to the integration of distributed cooling gas supply channels. A cylindrical vacuum housing with mechanical spacers and thermal reflectors has to be placed around the fiber-reinforced coil fixation housings to provide the thermal separation from the iron yoke. Furthermore, a plate has to be installed below and above the tape to stabilize the tapes and to conduct the high forces to the coil housing. Several CFD simulations were performed to determine gas parameters and temperatures depending on the load. One result is shown below. For this simulation, the SOLIDWORKS^® Flow-Simulation was used with an input of all relevant cryogenic properties data. The objective was to keep the hts temperatures for given loads between 30 and 40 K at low gas flow rates, also at static peak load. These simulations yield, i.e., flow rates which are used for simulations of the complete circuit. For the inlet, a temperature of 30 K and a pressure of 2 bar were assumed. Boundary conditions were set to be symmetric for this case study. Due to low Reynolds numbers, the simulation was set to laminar in the course of the calculations.

Figure 16 displays the temperature results for the maximum load. Two tapes are cooled in parallel. For the whole machine, a mass flow of 76 g/s GHe at 30 K is needed. The tape then reaches a maximum temperature of 40 K. The pressure drop is rather low, reaching max. 10 mBar, decreasing to approximately 3 mBar in low load cases. A heat transfer coefficient of approximately 1500 W/(m²·K) is reached with a Nusselt value of appr. 3.8 and a tube diameter of 100 µm.

The considerable low values in Table 6 are due to H2’s high gravimetric energy density. Moreover, the ${\dot{m}}_H$ is only available if the plane operates in flight mode. At the airport, a small separate supply of cooling gas is necessary to maintain the superconducting state of all cryogenic devices.

The goal for superconducting tape winding is a temperature range between 30 and 50 K, to keep a safe distance to the transition temperature and to utilize high lift factors under AC operation. To maintain a low temperature of, e.g., 25–30 K, of the helium flowing into the machine, a reverse Brayton cycle cooler (RBCC) is a viable option although it adds some extra weight to the system. The system sketch in Figure 17 shows the main components and exemplary data.

$$P_{me}={\displaystyle \frac{P_{Cw}-P_{Tw}}{{\eta }_m}}, P_{th}=P_C-P_T, P_{Cw}={\displaystyle \frac{P_C}{{\eta }_{pcC}}}, P_{Tw}=P_T{\eta }_{pcT}$$

(3)

• P_me: electrical power of the motor; η_m: motor efficiency.
• P_th: thermal power difference.
  P_Cw, P_Tw: shaft power of compressor and turbine.
• P_T, P_C: thermal power of turbine and compressor.
• η_pcT, η_pcC: polytropic efficiencies of turbine and compressor.

Compressor and turbine are fitted on the same shaft, the motor delivers the power difference. The electrical power consumption of the motor is calculated from the efficiency chain of turbine, compressor and motor.

The basic equations for the RBCC process are given in [19]. The thermophysical data for hydrogen (para- and ortho hydrogen) and helium are taken from [20,21]. Table 6 gives the main input data of the RBCC for this example. The weight calculation for the turbo compressor engine set is taken also from [19] for industrial machines. It should be kept in mind that there is an optimization potential if aerospace technology is applied to the construction. The weight will increase significantly if more than one stage is required on the turbine or the compressor.

As a heat exchanger, a plate heat exchanger in cross-counter flow design (Figure 18) was chosen. Preferably, it is built from thin laser-welded aluminum plates. The distance between the plates was within 0.5 mm, adapted to the properties of the cooling gases. By this design, a good heat transfer coefficient is achieved. For the heat exchanger design, in a first step, the average logarithmic temperature difference was calculated from the temperatures of both media flows for all load cases. This is given by

$$\begin{gathered} {∆T}_{max}=T_2-T_{wa} \\ {∆T}_{min}=T_3-T_{we} \end{gathered}$$

(4)

$${∆T}_{mlog}={\displaystyle \frac{{∆T}_{max}-{∆T}_{min}}{ln\left(\frac{{∆T}_{max}}{{∆T}_{min}}\right)}}$$

(5)

The heat transfer in the heat exchanger itself is composed from the temperature drops between He and the heat exchanger wall, and from the temperature drop between the wall and H2. The temperature drop inside the wall was neglected. So, in a second step, the middle velocity and pressure drops were calculated with the equations given in [22] for compressible media. To calculate the Nusselt numbers, the Gnielinski equation [22] was used. With the obtained temperatures, the design parameters of the heat exchanger as number of cells or length and height of the cells were adapted to the average logarithmic temperature difference.

In Figure 17, the calculated exemplary process data for one electrical 5.4 MW fan drive are given for the designing of LC 1. The shaft power of the compressor is dominant. With a polytropic efficiency of 0.8, an additional increase in temperature of 16.6 K has to be dissipated in the heat exchanger. To fully utilize an efficient counter flow heat exchanger, the temperature on the secondary entry T₂ (Point 2) has to be above the primary exit temperature T_wa (Point 6). From the power balance, the heat exchanger has to transfer the machine losses plus the compressor power input, minus the power extracted by the turbine. The results for this cooling circuit are shown in

Table 7. Data of the exemplary cooling set.

RBCC mass engine set	mRBCC	62 kg
Electrical efficiency	ηcc	0.71
He mass flow ${\dot{m}}_{He}$	LC1	0.076 kg/s
	LC2	0.045 kg/s
	LC3	0.03 kg/s
	LC4	0.1 kg/s
Heat exchanger mass without piping	mWT	0.5 kg
Heat exchanger transfer coefficient kWT @	LC1	2600 W/(m2·K)

.

It should be kept in mind that the hydrogen mass flow is prescribed by the consumption of the fuel cell. Increasing the He mass flow on the secondary side will reduce temperature T₂ but also increase the compressor power. The temperature T_wa drops then, with a small gradient, while the negative gradient of T₂ is much higher. Furthermore, the temperature drop, ΔT_WT, in the heat exchanger must be taken into account.

At a certain mass flow on the secondary equal temperatures T₂ = T_wa + ΔT_WT appear as the limiting condition. With the data chosen here, there is a safe distance to this limit.

As a second order condition, the pressure ratio in the turbine was limited to 3.5, and in the compressor to 2.5, which are close to practice values for turbomachinery with radial machines.

The limits detected with the RBC cooler led to considerations to abandon the request for a low entry temperature T₄ = 25 K and to study a simple fan-driven secondary circuit as depicted in Figure 19. This would avoid the largest part of the compressor power and allow a higher He mass flow with a smaller temperature rise in the machine. Additionally, a considerable saving of weight is possible.

In the simplified cooling system of Figure 19, the same basic data as for Figure 17 were used. As there is a higher sensitivity against the total pressure difference, the pressure drops in the piping (1a-1) and (3-3a) were estimated with additional 10 mbar.

By controlling the fan speed and pressure, the mass flow of He can be adopted to the different load cases. The exit temperatures at point 1 at the same mass flow are higher than with the RBCC, but are in the tolerable region for the superconductor. Optimization of the mass flow is possible. The limiting condition for the mass flow is now given by the equality of T₃ and $T_e$ +$∆T_{WT}$, with $∆T_{WT}$ of approximately 5–8 K. A significant saving in weight, cost, and power consumption are the main advantages of the simplified cooling system. The resulting data is presented in

Table 8. Data of the simplified cooling set.

Estimated fan mass	mBl	6 kg
He mass flow ${\dot{m}}_{He}$	LC1	0.076 kg/s
	LC2	0.045 kg/s
	LC3	0.030 kg/s
	LC4	0.045 kg/s
Number of cells per medium	nzm	35
Cell height and length	hz/Lz	50/105 mm
Heat exchanger mass without piping	mWT	0.7 kg
Heat exchanger transfer coefficient	kWT @ LC1	1500 W/(m2·K)

. The simplified cooling system requires a lower power of below 0.8 kW compared to the Brayton cycle with about 4 kW.

A new challenge arises with load case 4: as the heat exchanger is dimensioned for LC1, a difference of the middle temperatures at LC4 (landing phase) arises. The basic reason is that the hydrogen mass flow is too small to remove the losses in LC4 at the given temperature levels. Here, an increase in the hydrogen flow from 8 g/s to 13 g/s can solve this problem. The surplus hydrogen is then recirculated to the tank. One solution to this problem could be to design a distributed helium supply system in the coil fixture to enable a helium flow reduction in the end windings at low power flight states, where losses in the active area (iron) are dominating.

In the case of a controllable rotor excitation, recirculation in this operation point would not be necessary. As a summary of above considerations, a simplified overview of a possible cooling setup of a sc motor and inverter, utilizing a blower and a heat exchanger is shown in Figure 20. The superconducting drive is electrically powered by a fuel cell, a backup battery and an inverter.

The secondary cooling circuit cools the inverter and superconducting drive via helium, using the flow of LH2 to the fuel cell as a heat sink. From the sc machine perspective, a cryo-cooled inverter and superconducting power lines would be very beneficial to the cooling of the coils. The contribution and impact of the inverter and current lead on the cooling system are subject to further investigations, which are starting in late 2024.

8. Mass Balance

The resulting mass balance for the drive, inverter, and cooling system can be estimated due to the above calculations. The masses (

Table 9. Mass estimations.

Machine active mass	166 kg
Inverter	211 kg
Cooling system (Fan), without piping	7 kg
Cooling system (RBCC), without piping	63 kg

.) are the results of the presented assumptions. The goal also was to keep the tape length and frequency in a reasonable framework. The fan-cooling system has its advantages in terms of mass.

9. Discussion and Conclusions

Superconducting drives for emission-free aircraft are subject to global research. Several prototypes have already been built and the development of tape material is still continuing. Numerous requirements have to be taken into account in the multidisciplinary design of hts drives, which opens various new fields of research. As external cryogenic cooling equipment possess large mass, a more preferable strategy is to utilize the available cryogenic flow of LH2 on-board the aircraft as a heat sink. The investigations show that it is also possible to make use of a secondary gas-cooling circuit, recooling the gas with the LH2 flow, which is needed for powering an aircraft. Furthermore, the hts drive provides a lightweight, power-dense solution, which also significantly exceeds the efficiency of other drive systems.

In this paper, the basic design approaches for a superconducting machine for aviation fan drives, the basic design requirements, the losses, and the cooling system have been addressed. Tools for cooling design, which imply the collected cryogenic properties’ data, loss calculation in the tapes, and models for CFD calculation, have been developed to support the design of future hts drives in different configurations. Further steps for aviation have to be taken to enable superconducting technologies apart from cooling issues. REBCO tape has to improve regarding quality, long-term stability and AC-losses. From a mechanical view, the influence of thermal and mechanical cycling will have to be looked at in the future. Focus on in-depth research is necessary for the optimization of the winding and the mechanical construction, as well as for several failure scenarios (e.g., short circuit). However, until now the researchers have found no principal obstacles which could impede the application of superconducting drives in emission-free aviation.