IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 36, NO.

3, MARCH 2021 3159

Development of a High Power Density Drive System

for Unmanned Aerial Vehicles
Martin Schiestl , Federico Marcolini , Student Member, IEEE, Maurizio Incurvati , Member, IEEE,
Fabio Giulii Capponi , Member, IEEE, Ronald Stärz , Federico Caricchi, Member, IEEE,
Alejandro Secades Rodríguez, and Lukas Wild

Abstract—Unmanned aerial vehicles are characterized by a set bushfire monitoring, agriculture, or surveillance drones [4]–[7].
of requirements, like high efficiency, resiliency, and reliability that Regardless of the mission type, the main challenges in multi-
conflict with the other main requirement of high power density
rotor UAVs design are dictated by the need to maximize payload,
aimed at minimizing the overall weight and size. This article pro-
poses a novel, modular multiphase drive for a quadrotor drone, maneuverability, and flight endurance. Payload and maneu-
realized through the integration of an axial flux permanent magnet verability are related to the size and weight of its propulsion
machine and a GaN-based power electronic converter. After an components and therefore to their power density. On the other
overview of the design process, starting from the propeller choice, side, the flight endurance depends on the stored energy and on
a brief description of the system components is presented. Focus-
the efficiency of the propulsion system, from propellers to the
ing specifically on the power electronic converter, the article then
presents a full analysis of its electrical and thermal performance. energy storage [8]. It is well known that efficiency and power
Extensive experimental tests allows to validate the predictions of density are conflicting requirements.
the design and simulation stages and demonstrated the expected As in all aeronautic applications, even if they do not carry
high power density levels. people or dangerous payload onboard, standards in terms of
Index Terms—Aerospace safety, fault tolerance, gallium nitride, safety are particularly stringent [9], [10]. So far, several studies
integrated design, motor drives, thermal analysis, thermal are reported in the literature, aiming to increase the resiliency
modeling, unmanned aerial vehicles (UAVs), variable speed drives. of these systems through appropriate control in case of failure.
For example, Vey and Lunze [11] analyze the condition of a
complete rotor loss for both a quadrotor and a hexrotor. In the
I. INTRODUCTION latter case, thanks to its redundancy, through a reconfiguration
RONES are unmanned aerial vehicles (UAVs) that can fly, of the control system, just a soft failure is experienced with a
D autonomously or remotely piloted, in open or in confined
spaces, for thousand kilometers or just for few minutes. UAVs are
decrease of maneuverability. However, for a quadrotor UAV, the
complete rotor loss can result in moderate or even catastrophic
therefore designed with very different structures and sizes, ac- failure, especially when the weight increases. In [12] and [13],
cording to the expected performances. One of the most common different approaches have been investigated to avoid crashes in
structure is the so-called multirotor UAV, commonly named also such conditions. Nevertheless, if the thrust is lost in one rotor, the
vertical take-off and landing [1], [2]. It has hoovering capabilities mission capabilities are surely compromised and the more the
and can fly in every direction, horizontally and vertically, with weight the more crucial safety aspects are. While redundancy
high maneuverability [3]. Because of their versatility, multirotor helps to prevent catastrophic failures, it is also clear that it comes
UAVs are gaining more and more interest in a variety of com- at the expense of power density.
mercial and military applications such as: delivery, ambulance, Adoption of a multiphase configuration is another approach
that can be used to increase the resiliency of electric drives since
Manuscript received December 6, 2019; revised March 18, 2020, May 24,
they are still able to operate after a phase winding or inverter leg
2020, and July 13, 2020; accepted July 25, 2020. Date of publication August 4, fault [14], [15]. While this option is given high consideration for
2020; date of current version October 30, 2020. This work was supported part more electric aircraft or fully electric aircraft applications [16],
by Infineon. Recommended for publication by Associate Editor Prof. Dian Guo
Xu. (Corresponding author: Federico Marcolini.)
the scenario is completely different in the field of multirotor
Martin Schiestl, Maurizio Incurvati, Ronald Stärz, Alejandro Secades UAVs, where the most common approach is to use a commercial
Rodríguez, and Lukas Wild are with the Department of Mechatronics, MCI power converter and machine [17]. This consideration leads to
Management Center Innsbruck, 6020 Innsbruck, Austria (e-mail: martin.
schiestl@mci.edu; maurizio.incurvati@mci.edu; ronald.staerz@mci.edu;
suggesting that the application of multiphase configurations to
alejandro.secades@mci.edu; lukas.wild@mci.edu). multirotor UAVs is still an open research field. Also in this case,
Federico Marcolini, Fabio Giulii Capponi, and Federico Caricchi are however, attention should be paid on the effects of this design
with the Department of Astronautical, Electrical and Energy Engineer-
ing, University of Rome “La Sapienza,” 00185 Roma, Italy (e-mail:
choice on the overall drive weight and size.
federico.marcolini@uniroma1.it; fabio.giuliicapponi@uniroma1.it; federico. Finally, thermal design is another critical aspect to consider,
caricchi@uniroma1.it). since not only it allows to extend the flight endurance, but it
Color versions of one or more of the figures in this article are available online
at https://ieeexplore.ieee.org.
also strongly impacts lifetime and therefore reliability of com-
Digital Object Identifier 10.1109/TPEL.2020.3013899 ponents. However, increasing reliability through proper thermal

0885-8993 © 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission.
See https://www.ieee.org/publications/rights/index.html for more information.

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3160 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 36, NO. 3, MARCH 2021

design is another requirement that conflicts with the power

density one.
Following the previous considerations, this article presents a
propulsion system that aims to achieve the high power density
targets for UAVs, while at the same time demonstrating high
efficiency and addressing the resiliency and reliability require-
ments. The considered application is a heavy quadrotor UAVs
that operate inside buildings and therefore the overall size is
limited to ensure maneuverability.
Keywords to reach the desired goal are integration and modu-
larity. Integrated motor drives, and particularly integrated mod-
ular motor drives (IMMD), have been successfully proposed in
traction applications for electric vehicles, which share similar Fig. 1. Schematic depiction of the electric system.
requirements with UAV propulsion [18]–[20], but nothing has
been reported until now related to drones.
This article focuses its attention mostly on the power elec- independent three-phase systems, each with its own neutral
tronic converter, leaving the detailed discussion of the electric point. The power electronic converter can also be split into
machine and the overall control to future publications. It is four three-phase inverters: in this way, smaller and faster power
structured as follows. Section II gives an overview of the entire components can be used, with several benefits in terms of passive
propulsion system, discussing and justifying the main design components size, EMC, and bandwidth. This solution allows
choices. Section III presents the modeling of the inverter and us to implement both traditional control algorithms and fault
its analysis through simulations. Section IV reports the results tolerant control strategies. In fact, in case of failure in one phase,
of the experimental tests. Finally, the conclusion is drawn in the entire three-phase subsystem can be easily taken out of
Section V. service. The drive will however continue to operate with the
remaining healthy windings at rated power for a finite amount
II. SYSTEM OVERVIEW of time or at 75% of rated power indefinitely.
The design of the powertrain features a high degree of in-
The propeller design choice strongly impacts the design of the tegration: in fact, the mechanical frame carries out different
electrical machine and the related power electronic converter. tasks at the same time. It works as a mechanical support for the
Increasing the propeller diameter allows us to obtain higher inverters and the electrical machine; it behaves as a mechanical
thrust for a given mechanical power at the shaft [21], but on interface between the drive and the drone’s framework, and,
the other side, it also limits the maximum operating angular finally, it is the heat sink through which the heat is removed
speed, due to the constraints on the tip speed. On the contrary, from within both the machine and converter. This arrangement
the propeller loading, PL , equal to the thrust per unit propeller allows us to significantly increase the power density, to simplify
area, can be maximized when reducing the diameter. the connections by the elimination of power cables and also to
Therefore, a tradeoff between dimensions, weight, and mitigate possible EMC problems, [18], [20], actually increasing
thrust needs to be achieved [22]. The analysis presented in the reliability of the drive.
Section II-A shows that the optimal sizing is found at a relatively Fig. 1 shows a schematic view of the powertrain, where the
small propeller diameter and high rotational speed. In fact, this propeller is directly coupled to the multiphase electric drive,
solution allows us to maximize the specific propeller loading, while the electrical machine, the power electronic converter and
PLspec , defined as the ratio of the propeller loading with respect the control board are all enclosed in the same housing.
to the mechanical power. As a consequence, also the powertrain Moreover, the massive airflow generated by the propeller
power density can be maximized: in fact, not only the propeller on its axis is used as a forced air cooling mechanism for the
will be smaller and lighter, due to the reduced dimensions, but frame. Therefore, as shown in Fig. 2, since the motor and the
also the electric machine power density will benefit from the power electronic converter are integrated into the same housing,
increase in speed. mechanical and thermal functions can be merged together, to
Following this design choice, it results that the maximum op- increase the power density.
erating fundamental frequency also increases, exceeding the kHz
threshold. At these levels, two challenges arise. On the electrical
A. Propeller Selection
machine side, the design shall take appropriate measures to limit
frequency-dependent losses; on the converter side, an increase in The propeller is required to exert a thrust of at least 40 N,
switching frequency becomes necessary to ensure good quality to carry the payload and to achieve a feasible flight ability.
of the waveforms and accurate current control, with a consequent The study of the suitable propeller geometry is realized in a
impact in terms of losses [23]. range of low flight speeds on the vicinity of such the UAV will
The resiliency of the drive is ensured through a multiphase foreseeably perform (from 1 to 9 m/s, without any inclination
arrangement. In particular, a twelve-phase system is chosen, angle of the flight velocity vector). In particular, as already
since this allows arranging the winding of the machine in four discussed, the propeller diameter needs to be optimized to find

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SCHIESTL et al.: DEVELOPMENT OF A HIGH POWER DENSITY DRIVE SYSTEM FOR UNMANNED AERIAL VEHICLES 3161

Fig. 3. Axial velocity at propeller vicinity, straight up- and downwind.

thrust-to-weight ratio, jeopardizing the specific power figure. In

addition, the lower design speed at the same power reduces also
Fig. 2. Schematic view of the mechanical system.
the power density of the electrical machine.
From the above-mentioned analysis, it can then be concluded
that propeller A is the optimal design in terms of power density:
TABLE I it develops about 45 N of thrust at 15 000 r/min and 1 kW
COMPARISON OF DIFFERENT PROPELLERS
mechanical input power. The noticeable observation from this
analysis is that optimizing at the same time for thrust-to-weight
ratio and specific propeller load leads to high-speed designs, in
excess of 10 000 r/min and in contrast to what is commonly
found in the literature. All that has a significant impact on the
overall drive design.
Given the highly integrated design, it is important to study
the axial air flow, since it runs over the outer surface of the
housing and determines cooling capabilities. When the UAV is
in flight mode, the resulting air axial speed is shown in Fig. 3 as
a function of the propeller radial position and for flight speeds of
the best compromise between thrust-to-power ratio and propeller 1 and 9 m/s. A very flat behavior is obtained, with a minimum of
loading, keeping in mind the limitations that constrain its size 15 m/s. In hovering mode, instead, air axial speed is proportional
[24]. For this design, a maximum value for the tip speed of to the square root of the thrust [24]: when this last one is equal
0.8 Mach number has been considered as a limitation for the to 45 N, an axial air flow of 14 m/s is still found. This analysis
propeller diameter. allows us to calculate the heat transfer coefficient for the thermal
An iterative software in MATLAB has been developed, based design of the drive and confirms that it is going to remain pretty
on the combined Momentum & Blade Element theories for much constant in every operating condition.
minimum loss condition without neglecting wake contraction
as developed and described in [25]. The software develops the
B. Electric Motor
optimal propeller geometry and determines its performances in
parametric form as a function of the number of blades, airfoil Coming to the electrical machine design, an axial flux perma-
data type (or types), hub diameter, and atmospheric conditions. nent magnet (AFPM) machine is considered, since it has already
Table I shows a comparison of three propellers designed been proven that this machine allows greater torque density than
according to different optimization objectives. In propeller B, its radial counterpart [26]. The chosen topology is the torus-type,
the thrust-to-power ratio is maximized, resulting in the design having two rotors and one stator in between [27].
with the highest diameter, which limits the maximum speed Due to space constraints, manufacturing problems cannot be
to 6500 r/min. Propeller C, instead, is designed by optimizing overlooked, thus the number of coils has been fixed to 12, i.e.,
for specific propeller load PLspec , while propeller A represents the minimum number that allows us to obtain four different
the concurrent optimization of specific propeller load and total three-phase subsystems. Concentrated windings [28], [29] are
weight, including its structure. All the designs need to comply also adopted both for the ease of manufacturing and because
with a minimum thrust of 45 N. a significant reduction in the mutual coupling between phases
A comparative analysis shows that propeller A is smaller, is obtained, thus increasing the resiliency of the drive in case
lighter and has much greater PL and PLspec than propeller of fault. Furthermore, when concentrated windings are adopted,
B, at a maximum speed of 15 000 r/min. It is interesting to the end-windings do not overlap: as a consequence, in case of
note that propeller C only slightly outperforms design A in fault, the local over-temperature of the faulty phase does not
terms of specific propeller load PLspec . However, propeller C affect the healthy ones.
results in a much heavier structure with much wider chords From the fractional slot theory [30], it is found that the
and greater thickness of the blades and therefore shows a low best tradeoff in terms of flux linkage and harmonic behavior

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3162 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 36, NO. 3, MARCH 2021

TABLE II
PERFORMANCES OF THE ELECTRIC MACHINE

Fig. 4. Electric motor prototype. (a) Stator twelve-phase concentrated wind-

ing. (b) One rotor disk with magnets.

Fig. 5. One three-phase subsystem back emf measurements (100–10 ms/div).

Fig. 6. Overview of the electronic system including control MCU, power

electronics, and sensors.
is obtained with a number of rotor poles equal to 10. This
sets the fundamental frequency at 1250 Hz, when the speed is
15 000 r/min. At such frequencies, losses mitigation becomes
critical because otherwise the powertrain efficiency (and UAV’s [19]. In particular, for this design, Infineon CoolGaN early
flight endurance) would be significantly affected. Moreover, as engineering samples (CoolGaN-e.e.s.) are chosen.
the size reduces with the increase in power density, it becomes Those are discrete components, with a single GaN switch per
more and more difficult to extract the heat within the desired integrated circuit, rated 100 V and 40 A. One particular feature
temperature limits. For this reason, a coreless AFPM structure of this device is related to its packaging: in fact, the silicon
has been chosen, so that iron losses are eliminated, together with substrate is exposed at the GaN topside, without any additional
a Litz wire winding to limit ac losses in the copper. It is worth plastic cover, thus allowing a very low-thermal resistance path
noting that the absence of the iron core further decreases the for losses extraction. At the present time, however, these devices
mutual coupling between the phases. Details on the design of are not available for a voltage rating that matches more precisely
a coreless AFPM machine with nonoverlapping concentrated the dc-link voltage.
windings can be found, for example, in [31] and [32]. A block diagram of the converter is shown in Fig. 6. It is
Fig. 4 shows the prototype of the twelve-phase coreless Torus- composed of one control board, equipped with an Infineon Aurix
type AFPM machine with nonoverlapped windings that have TC233 automotive grade micro-controller, and one main dc-link
been built. Fig. 5, instead plots, a diagram of the machine’s board plus twelve PCBs each of them realizing an half-bridge
no-load voltages taken during preliminary experimental testing. module. Every half-bridge module mounts.
As it is expected, since each coil spans 150 electrical degrees, 1) Infineon GaN bootstrap driver 2EDL7125G.
the resulting waveforms are very close to sinusoidal. 2) Infineon low voltage CoolGaN-e.e.s.;
Finally, the main machine parameters are listed in Table II. 3) Allegro Hall-Effect current sensor ACS730, with 1 MHz
bandwidth.
Such a modular structure allows full integration between the
C. Inverter Description converter and the electrical machine. In fact, each module is
Consistently with the machine structure, the power electronic located in direct correspondence with each coil, as schematically
converter is composed of four three-phase inverters. Rated input shown in Fig. 7, eliminating interconnection cables. As a result,
voltage is 24 V so that the system can be supplied by six series no connectors are needed (one of the major sources of faults),
connected Li-Po cells; rated output current is 15 Arms at a rated losses and additional voltage drops are eliminated, EMI and
frequency of 1250 Hz. possible overvoltage problems are strongly limited. Moreover,
At such high fundamental frequencies, wide-bandgap devices in case of fault, it is possible to completely switch OFF one half-
have already shown to allow for higher power densities and bridge module or one entire inverter, taking full advantage of
efficiency than their silicon counterparts, especially in IMMD the multiphase arrangement.

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SCHIESTL et al.: DEVELOPMENT OF A HIGH POWER DENSITY DRIVE SYSTEM FOR UNMANNED AERIAL VEHICLES 3163

Fig. 9. Overall control structure of the UAV with redundant central controllers
Fig. 7. Distribution of the half-bridges of the four inverter systems. and independent local controllers.

TABLE III
POWER ELECTRONIC CONVERTER RATED PERFORMANCES

Fig. 10. Three-dimensional cross section of the full assembly.

position sensors for safe a reliable start-up. A space vector

modulation scheme (SVPWM) is adopted to command the GaN
switches.

D. Powertrain Integration
Fig. 10 shows a cross section of the 3-D drawing with all
Fig. 8. (a) Bottom and (b) top views of half-bridge board module. (c) Control components assembled. Starting from the right, it is possible to
electronics board mounted on dc-link PCB. notice the shaft, where the propeller is going to be mounted;
the AFPM motor, composed of two rotors and one stator in
between; the twelve half-bridge PCB modules, mounted radially
The main characteristics of the converter are summarized in in correspondence with each coil; the dc link and control boards.
Table III, while Fig. 8 shows the prototypes of one half-bridge Mounting holes are also visible, that are used to anchor the drive
module PCB and the control board directly mounted on the dc- to the drone framework, and also the bearings that are placed on
link PCB. the front shield and the housing structure.
The whole UAV is composed of four propellers with their Correct radial positioning of the half-bridge PCB modules
integrated propulsion systems and a main body equipped with is ensured through screw terminals on the dc-link board. Each
redundant central controllers (see Fig. 9). Each local Aurix module is then pressed through a metallic clip against one of the
TC233 micro-controller implements a sensorless vector control twelve aluminum spokes, which act both as a mechanical support
scheme for each of the four inverters. The sensorless algorithm and as a thermal extraction path. In fact, the clip directly presses
is based on a back-emf estimation algorithm and used hall-effect the top side of each GaN against the aluminum (see Fig. 11).

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3164 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 36, NO. 3, MARCH 2021

2) The connection among the 12 half bridges and the stator

is flexible.
3) TIM provides a soft surface between electronics and alu-
minum housing.
4) The overall structure is stiff with resonant modes located
at the higher frequencies.
Furthermore, standard inverters not integrated into the motor
would experience on UAV similar level of vibrations and ac-
celerations having the disadvantage of being not modular and
therefore providing a lower reliability/availability.

III. INVERTER MODELING

Fig. 11. Detail of half-bridge PCB positioning. The top side of the GaN is
pressed to the aluminum frame through a thermal interface pad (TIM). This section describes the modeling of the inverter that has
been carried out to confirm the efficiency specification and to
predict both its transient and thermal behavior. First of all, a
TABLE IV loss analysis calculation is described and afterward an equiv-
POWER DENSITIES OF THE ELECTRIC DRIVE
alent circuit that takes into account all parasitic phenomena is
developed. Finally, a thermal analysis is also performed through
the software Ansys Icepack.

A. Losses Estimation
A first estimation of the GaN losses, and the efficiency of the
whole power electronic converter, can be obtained using (1), (2),
[34], and (3)–(5)
 
Electrical insulation between the source terminal of the device 1 ma
Pc,pos_half = Rdson Ipk
2
+ cos φ (1)
and the aluminum frame is guaranteed by a thermal interface 8 3π
pad. This solution is intended to minimize the thermal resistance  
1 ma
from the top side of the GaN to the ambient. The clip, therefore, Pc,neg_half = Rdson Ipk
2
− cos φ (2)
allows the mechanical frame to carry out both its mechanical 8 3π
and thermal functions. where ma is the modulation index and cosφ is the displacement
Finally, the external frame is designed to provide a totally power factor
enclosed manufacturing and is equipped with fins along the
axial direction thus taking advantage of the air flow caused by Rdson Ipk
2
Pc,tot = (3)
the propeller and therefore decreasing the thermal resistance 4
between ambient and frame. Table IV lists the size and weights 1 1
of the different components and calculates the expected power Psw = Vdc Ipk (tr + tf ) fsw + Coss Vdc2
fsw (4)
π 2
densities. The power density of the overall drive, considering   
the cylindrical bulk volume, is approximately 1.43 kW/l, which 2VSD Ipk I2pk
Pdead = fsw tdead + Rdsoff + ΔI 2
is a very good result if it is compared to [33], where a value π 2
of 0.71 kW/l was reached. In addition, the breakdown of the (5)
weights shows that frame and covers are still a significant part
of the overall system. It is therefore possible to further optimize where Pc,pos_half and Pc,neg_half are the conduction losses of
the geometry of the mechanical frame to reduce weight and size the switch, respectively, during the positive and the negative half
and to increase the overall power density. cycles of the sinusoidal current, Pc,tot are the average losses of
the GaN over a fundamental cycle, Psw are the switching losses
including the term related to the output capacitance, and Pdead
E. Considerations on Vibrations are the losses occurring during the dead-time. In (5), Rdsoff is the
Due to its intrinsic modularity, the proposed concept has the differential resistance and VSD is the reverse operation voltage
potential to increase the availability and safety of the overall of the channel both at VGS = 0 V: taken together they model
system. It is however worth to point out that vibrations have to the reverse drain-source characteristic and allow to calculate the
be carefully measured in real operating conditions to evaluate conduction losses during dead-time. All parameters are defined
their effect on the overall reliability. Nevertheless, the system in Table V. The current ripple ΔI can be calculated from the
has been designed to limit the effect of vibrations basing on the dc-link voltage, Vdc , the amplitude modulation ratio, ma , and
following assumptions. the machine inductance. For the selected switching frequency
1) The motor is a slot-less, coreless, type. Therefore, cogging of 200 kHz, it is estimated to be approximately 4 A. Once the
torque and consequent vibrations are not present. current ripple is known, it is also possible to estimate the losses

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SCHIESTL et al.: DEVELOPMENT OF A HIGH POWER DENSITY DRIVE SYSTEM FOR UNMANNED AERIAL VEHICLES 3165

TABLE V
CHARACTERISTICS OF COOLGAN-E.E.S. SWITCHES

TABLE VI
MOSFET VERSUS GaN COMPARISON

for a single GaN according to (3)–(5). The lower part of Table V

shows the estimated losses at nominal operating conditions.
The evolution of the losses over an electric period is shown in
Fig. 12(a) for the case of a top switch. Instead, Fig. 12(b) plots
the losses of one three-phase inverter as a function of the phase
current and the modulation index. The same plot shows also the
constant output power curves for the electrical machine so that Fig. 12. (a) Instantaneous and fundamental cycle average top switch losses.
inverter losses can be identified in each operating condition. (b) Total inverter losses of one three-phase inverter and motor output power.
(c) Relative losses and current ripple versus switching frequency.
Finally, Fig. 12(c) shows the relative behavior of losses and
current ripple for different choices of the switching frequency.
In the analyzed range, converter losses decrease almost linearly
with frequency, but the current ripple increases more than lin-
early in the range between 150 and 200 kHz. Therefore, the B. Electromagnetic Design and Transient Simulations
value of 200 kHz was selected as an overall compromise to limit The design of the dc-link and half-bridge PCB is critical to
the current ripple, resulting in lower harmonic losses inside the ensure system functionality and robustness. Particular care was
machine and better controllability. taken in designing the layout of the latter following well-known
It is also interesting to perform a brief comparison between best practices to minimize parasitics, such as to position the
the selected device and a similarly rated MOSFET (i.e., OptiMOS driver as close as possible to the switch. Rivera-Ramos and
BSC028N06NS). The main results of the comparison are sum- Jimenez [35] provide useful equations to obtain an immediate
marized in Table VI; the gate resistance is optimized to obtain evaluation of parasitic inductances and capacitances, to check if
ideally the same switching transients under the assumption that changes in the layout were needed.
the parasitic inductances remain the same (not true in practice Once the layout was finalized, PCB parasitic capacitances
since the MOSFET has a much greater internal parasitic value). were determined through electromagnetic simulation with the
Under these hypotheses, the two devices show the same losses, software Q3D. The only capacitances of interest are the ones
but the MOSFET exhibits much higher junction-to-case thermal between phase traces and power lines: results led to an estimation
resistance (because of the plastic covering of the top). Further- of Cpar,high = 3.18 pF between the phase trace and the positive
more, the surface area of the GaN is also half with respect to the dc-link bus and of Cpar,low = 2.51 pF between the phase trace
MOSFET, which leads to double power density at the device level. and the negative dc-link bus, (see Fig. 13).
All those considerations confirm the need to adopt wide-bandgap In terms of parasitic inductances, instead, several traces are
semiconductors. of interest: the gate lines to the switches as well as the power

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3166 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 36, NO. 3, MARCH 2021

Fig. 13. Arrangement of HB and main dc-link capacitors and parasitics for
LTSpice simulation of the capacitors’ current.
Fig. 14. Turn-OFF transient of a bottom switch. (a) LTSpice simulation.
(b) Measured waveform (10 ns/div 10 V/div).

TABLE VII
PARASITIC INDUCTANCES AND RESISTANCES FOR THE INVERTER BOARD
vibrations, optimized current/volume ratio, possible EMI issues,
and battery ripple current.
The amount of parallel capacitors is mainly depending on
their rms current Ic,rms which can be initially estimated through
the following equation [41]:
 √ √ 
3 3 9
IC,rms = Iph 2ma + cos φ
2 − ma . (6)
4π π 16
loop inductance (Lloop,HB in Fig. 13). The gate inductances
(LGateHOH, LGateHOL for the high-side switch and LGateLOH, A more detailed calculation of the currents present at Cdc,HB
LGateLOL for the low-side switch) have been both calculated an- can be performed through an LTspice simulation of one half
alytically according to [36] and evaluated through simulations. bridge, including Cpar,high and Cpar,low as well as Lloop,HB , as
Table VII reports the results of the calculations and the values shown in the right end of Fig. 13. These simulation shows, in
extracted from the software. The discrepancies at the paths a worst-case operating scenario, a dc-link capacitor current of
of the low-side gate can be explained by the smaller loop, the half-bridge PCB equal to ICdc,HB,RMS = 6 A which is in
which reduces the self-inductance and is not accounted for in agreement with the initial analytical estimation of (6).
the calculations. In general, however, calculated and simulated In [19], the capacitor volume as a function of switching
valued show a good agreement, ensuring their correct estimation. frequency is shown for three different capacitor technologies:
The above-mentioned values of the parasitics were used for the comparison clearly demonstrates that ceramic capacitors
transient simulations with LTspice, to more accurately recreate allow us to minimize the required volume. Therefore MLCC type
the real switching behavior of the devices. A comparison be- devices from Murata model GCM32DC72A475ME02 were se-
tween simulation and measurements of the transient behavior of lected [42].
the switches has been carried out. An example of the turn-ON According to the device datasheet, this would result in a
transient is shown in Fig. 14. As it can be observed, the simu- temperature increase of over 35 °C if only one capacitor is used,
lation, Fig. 14(a), and the measurement, Fig. 14(b), match quite which reduces to 3 °C when using five parallel capacitors, hence
well. The oscillation frequency as well as the overshoot exhibit decreasing the individual current to IC,cap,RMS = 1.2 A (see
a good agreement, which confirms the correct evaluation of the Fig. 15). As an additional benefit, having multiple capacitors in
parasitics. A small discrepancy of 5 ns can be still observed in parallel allows us to decrease the power loop parasitic inductance
the rise time, which can be explained as a nonperfect modeling Lloop,HB , resulting in smaller overshoot and less stress on the
of the GaN or of the gate drive circuit. switches as well as in a decrease of the overall losses.
Once the parasitics are determined, the optimal sizing of the
dc-link capacitors can be addressed [37]–[39]. The number and
value of the capacitors have been selected according to [40] C. Heat Management and Thermal Simulations
taking especially into account: the nominal current of the drive, The thermal energy generated by the losses of the system
the temperature increases and therefore lifetime, the need for low needs to be dissipated by the outer structure via forced air-
inductance path in the switch layout, mechanical constraints and cooling. Fig. 16 shows the thermal circuit of one inverter board

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SCHIESTL et al.: DEVELOPMENT OF A HIGH POWER DENSITY DRIVE SYSTEM FOR UNMANNED AERIAL VEHICLES 3167

TABLE VIII
THERMAL SIMULATION PARAMETERS AND RESULTS

Fig. 15. MLCC capacitors: temperature rise versus current [43].

Fig. 17. Thermal maps obtained with ICEPACK.

power, even at reduced air flow speed conditions (8.1 m/s instead
Fig. 16. (a) Thermal connection on system level. (b) Equivalent thermal circuit of the nominal value of 15 m/s, to match the experimental test
from the power sources of one inverter to the medium. conditions), both the high-side and low-side switch junction
temperatures, TJHS and TJLS , are below 60 °C.
From the simulations, the equivalent thermal resistance from
and the thermal connections at the system level as well as the
junction to ambient, RJamb , can be estimated as
location of the measurement points. The main path followed by
the heat flow goes from the switch junction over its top (RJT ), RJamb = 6.8 ◦ K/W. (7)
through the thermal interface material (RTIM )and the aluminum
structure (RALU ), till the ambient air (RALUA ). The higher The analysis also allows us to evaluate the effectiveness of the
constructive solution with no plastic cover on the GaN topside.
the airflow velocity at the outer frame, the lower the thermal
resistance between the frame and the ambient air, RALUA , will In fact, from the comparison of the values of RTIM and RJT
a ratio greater of 15 is found, clearly showing that the adopted
become. Any unbalance in losses between the two switches of
GaN packaging allows for a very efficient heat extraction.
the leg will activate also a secondary path from the junction to
its bottom (RJB ), through the board copper (RCu ) to the bottom The thermal analysis, however, needs to include also the
effects of the losses in the motor, since the aluminum frame
of the other switch. The power losses of the driver, PDriver , and
is shared with the power electronics converter. The motor power
the LDO circuitry, PLDO , are two additional power loss sources
connected from the junction to their case (RJC ). losses are transferred to the aluminum frame through RALU
and then to the ambient through the same frame-to-air thermal
All inverter sections are thermally connected to each other via
resistance of converter, RALUA . Assuming 50 W of losses in
the aluminum frame both at the inside and at the outside of the
structure. the machine (from the efficiency at rated conditions, Table II)
and considering that RALU is much smaller than RALUA , it
A finite element thermal simulation of the complete structure
can be estimated that the frame temperature will increase by
is not practicable, since it is very difficult to obtain a converging
mesh due to the small airgaps and the round surfaces. Therefore, 13 °C with respect to the value indicated in Table VIII. All other
temperatures will also increase by the same amount.
exploiting axial symmetry, only a 30° section is simulated,
Finally, a transient thermal simulation has also been per-
stretching it to yield a full orthogonal structure. The half-bridge
module PCB with the complete trace layout is exported into formed, to evaluate the time constant for the junction and yield-
ing a value of τ = 100 s.
Ansys Icepack software, to properly simulate the heat transfer
on the board.
IV. EXPERIMENTAL TESTS
Table VIII lists all important simulation parameters, together
with the temperatures of relevant spots resulting from the sim- The measurement of the converter power losses and so its
ulation. Additionally, Fig. 17 shows the 3-D temperature map efficiency and maximum output power evaluation has several
seen from two different angles. Results show that at the nominal challenges. First, the actual cooling performance of the fins on

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3168 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 36, NO. 3, MARCH 2021

the housing is directly proportional to the airflow speed produced

by the propeller, which is at this stage not available. Additionally,
to measure just the losses of the inverter, an external three-phase
load is required in place of the motor. Finally, the high number
of phases and the high switching frequency and low losses make
it very hard to measure the instantaneous output power.
The conventional method, i.e., a direct measurement of input
and output power, would require a high precision and very fast
instrumentation, which results in high cost, especially for a high
number of phases. Therefore, a nonadiabatic semi-calorimetric
approach [43], [44] has been adopted, which allows accurate
evaluation of total converter losses through multipoint temper-
ature measurements. Fig. 18. Measurement setup of the drive. On the left, the power supplies,
The method requires a calibration phase that is usually carried oscilloscope and current measurement clamps are depicted. On the right, the
out in dc. It consists in applying a known amount of power wind tunnel is visible, together with the UAV drive, the RL loads, and the thermal
data logger. The airflow is in the direction into the tunnel.
losses on the loss carriers and measuring the temperature rise
of specific points, equipped with thermocouples. From these
measurements, the thermal resistance matrix R can be obtained
as

R = Ploss −1 ΔT (8)

where Ploss is the power loss vector containing each loss com-
ponent and ΔT the temperature difference with respect to the
ambient temperature.
Once the dc-calibration procedure is completed and the ther-
mal resistance matrix is obtained, (8) is inverted to calculate
losses from temperature measurements in the desired operation
conditions.
To quantify the error introduced by this method on the evalua-
tion of efficiency and power, both type A and type B uncertainties
are accounted for. This results in an overall uncertainty on the
temperature δT, which can be calculated through the following
equation:

δT (T ) = σlin
2 + σ2
rep (T − Tamb ) + σT C−08 (T )
2 (9)

whereas σ rep and σlin are the errors due to reproducibility and Fig. 19. (a) Phase current and (b) voltage difference between phase and
negative terminal of dc-link waveforms on RL load at rated frequency. The
linearity, respectively, while σ TC-08 is the error introduced by voltage waveform is digitally filtered.
the thermocouples and the data logger.
The relative error on losses calculation, δP/P, depends on the While the electric motor is mounted in the integrated drive, it
relative uncertainty of temperature measurement at the nominal cannot be loaded due to the absence of the propeller. Therefore,
conditions, δT/T, and depends on the relative error of the thermal tests on the converter have been carried out by using four
resistance matrix, δR/R. Moreover, since the matrix R needs now independent three-phase RL loads with isolated star points. The
to be inverted, there is an error propagation phenomena, which heatsink mounted resistors and inductors are clearly visible in
is taken into account with the condition number of the thermal Fig. 18.
resistance matrix, κ(R), yielding to The inductance value was chosen to be 4.2 μH, to match
  the machine inductance. The selected resistance, instead, was
δP δT δR
≤ κ (R) + . (10) 0.375 Ω. In this way, at the rated current, a voltage drop is pro-
P T R
duced that is close to the machine’s back-emf at rated speed plus
the resistive voltage drop on the machine’s resistance. As a con-
A. Measurement Setup sequence, the required phase voltage and the modulation index
Since at the present stage, the propeller prototype is not ma are very similar to the expected ones at rated conditions when
available yet, the drive is installed into a wind tunnel (see Fig. 18) the converter is feeding the motor. Fig. 19(a) shows the phase
to recreate the airflow which would be present if the propeller current waveform at 1250 Hz and 15 Arms. For the same condi-
was mounted on the device. Due to setup limitations, the airflow tions, Fig. 19(b) shows the corresponding voltage difference be-
velocity is set to 8.1 m/s, which is lower than the nominal axial tween a phase and the negative terminal of the dc-link; the wave-
velocity (15 m/s in flight mode). form is digitally lowpass filtered (fpass = 40 kHz, fstop = 80 kHz,

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SCHIESTL et al.: DEVELOPMENT OF A HIGH POWER DENSITY DRIVE SYSTEM FOR UNMANNED AERIAL VEHICLES 3169

TABLE IX
THERMAL MEASUREMENTS

Fig. 20. Efficiency and output power for the total system versus load current
rms.
Astop = −80 dB) to highlight the fundamental. SVPWM is
adopted, consistently to what explained in Section II-B; the
control loop is executed every 50 μs (20 kHz). The third colum in Table IX shows the temperature measure-
ments taken during the experimental tests at nominal operating
B. Power Loss Measurement Procedure conditions. Comparing the simulated results in Table VIII to the
measured temperatures, it can be seen that Tframe and TAlu show
Tests have been performed by placing six thermocouples a good matching, while TCuGnd is off by about 5 °K. However,
around the system: two on the ground plane of phase 1 and this result is reasonable, in consideration of the simplifications
9 inverter legs, two on the aluminum frame near to phase 5 that were made on the geometry.
and 9, one inside the system on the aluminum ribs of phase 9 Fig. 20 illustrates the efficiency and output power of the
and the last one on the dc-link. Figs. 16 and 17 show the positions whole system as a function of the rms of the phase output
of the above thermocouples. current, Irms . The error bars in the efficiency plot are the results
The dc thermal calibration is performed by imposing a posi- of the uncertainties in power calculations, (10), deriving from
tive current between the negative and the positive dc-link termi- the uncertainties introduced by the nonadiabatic semicalori-
nals to place all the devices in reverse conduction mode. In this metric approach. At the nominal value of 15 A, the output
way, it is possible to apply losses homogeneously over all the power Pout of the system is 1222 W and the total efficiency is
GaNs (without the need for any gate signal). Moreover, since 95.1% ± 0.77%. The latter results in a total system power loss of
the reverse conduction equivalent resistance Rdsoff is five times 63.1 W ± 9.83 W.
higher than Rdson , a much lower current is required. If the total losses of the drivers and LDOs are subtracted
From the temperatures and the applied power, the equivalent (the values for a single module are reported in Table VIII),
thermal resistance and, most notably, the relationship between then the measured value of the total system power losses is in
temperature and power can be calculated. Measurements at dif- good agreement with the estimated value of the GaN losses that
ferent power levels showed an almost completely linear relation was shown in Table V. Moreover, the nominal power stated in
between power losses and temperature increase, thus confirming Table III has been reached, leaving headroom for temperature
the validity of the nonadiabatic calorimetric approach, even if the and size as well as power density improvements.
calibration takes place at different power levels than the actual Finally, in terms of transient behavior, the measurements
operating conditions. Any remaining small deviation is, in any resulted in a system thermal time constant τ equal to 100 s,
case, still accounted for in (9) by the standard deviation of the thus confirming the prediction from the simulations. This result
linearity error σ lin . ensures that short overloads will have a low impact on the
junction temperature.
C. Results Summarizing, despite the complexity of the structure both
For the dc thermal calibration, an amount of power higher from a mechanical and an electrical point of view, the adopted
than the rated one is fed into the system. This results in higher analysis and simulation methods have shown to be in very good
temperatures, which decrease the relative error made onto the agreement with measurement results. The developed design
thermal resistance matrix. The reverse input current into the strategy is scalable and can be therefore used for a reliable design
switches is therefore set to 36 A and the reverse voltage of each of bigger systems.
switch is measured, resulting in a total power loss of 100.05 W
for the whole power electronic converter. The resulting calibra-
tion temperatures can be seen in the second column of Table IX. V. CONCLUSION
By comparing TCuGnd-Ph1, TCuGnd-Ph9, and TAlu-Ph10, it In this article, a high-power density integrated modular mul-
can be observed that the temperature rise of Phase 1 and Phase tiphase drive for the propulsion of a UAV was proposed and de-
9 with respect to the aluminum frame is approximately 25 °C. scribed. Detailed electrical and thermal analysis for the converter
Since the main focus of this analysis is the calculation of the was carried out. Finally, experimental tests were performed to
converter losses and since the temperature of the frame shows verify the predictions from the model. The key conclusions can
an homogenous behavior even at nominal conditions, it is easier be summarized as follows.
to take as a reference the temperature of the frame. Hence, the 1) It is possible to design an electric drive that combines at
thermal resistance matrix reduces to the scalar thermal resistance the same time high power density, high efficiency, and a
from the switch junction to the frame, R_JALU. resilient structure to ensure reliability.

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3170 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 36, NO. 3, MARCH 2021

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SCHIESTL et al.: DEVELOPMENT OF A HIGH POWER DENSITY DRIVE SYSTEM FOR UNMANNED AERIAL VEHICLES 3171

Martin Schiestl received the bachelor’s degree in Ronald Stärz received the master’s degree in exper-
mechatronics electrical engineering and the master’s imental physics from the University of Innsbruck,
degree in mechatronics smart technologies from the Innsbruck, Austria.
MCI, Innsbruck, Austria, in 2015 and 2017, respec- He has been the Head of the Hydraulic Engineering
tively. Laboratory, the University of Innsbruck, until 2008.
In April 2016, he began working with the Emerg- Since then, he is working at MCI, where he built up
ing Applications Laboratory, focusing on the topic the study programs in Mechatronics and the Emerging
of resonant wireless charging, amplifier design, and Applications Laboratory, which he leads since 2016.
drive systems. His work focuses on high-frequency power electron-
ics and on probes for fusion experiments.

Federico Marcolini (Student Member, IEEE) re-

ceived the B.S. and M.S. degrees in electrical en-
gineering, in 2015 and 2018, respectively, from Federico Caricchi (Member, IEEE) received the
Sapienza-University of Rome, Rome, Italy, where M.S. and Ph.D. degrees in electrical engineering from
he is currently working toward the Ph.D. degree in the University of Rome “La Sapienza,” Rome, Italy,
electrical engineering. in 1988 and 1994, respectively.
His main research interests include analysis and From 1991 to 2010, he was with the Department
design of permanent magnet machines and wide- of Electrical Engineering, University of Rome “La
bandgap-semiconductor-based power converters. Sapienza,” serving as a Head of the Department from
Mr. Marcolini is a Student Member of IEEE In- 2007 to 2010. Since 2010, he has been with the
dustry Applications Society, IEEE Power Electronics Department of Astronautical, Electrical and Energy
Society, and IEEE Industrial Electronics Society. Engineering as a Full Professor. He has authored or
coauthored more than 100 technical published papers
and more than ten national and international patents. His research interests in-
clude analysis and design of unconventional electric machines, power electronic
equipment, and permanent-magnet motor drives.
Dr. Caricchi is a Registered Professional Engineer in Italy and is a member
Maurizio Incurvati (Member, IEEE) received the of the Italian Association of Electrical and Electronic Engineers, the Italian
M.Sc. degree in electrical engineering with focus on Association for Naval Techniques, and the IEEE Industry Applications Society.
power electronics and the Ph.D. degree from Univer- He is a member of the IEEE IAS Industrial Drives Committee and Electric
sity of Rome “La Sapienza,” Rome, Italy, in 1999 and Machines Committee. He was the recipient of the 2005 First Prize Paper Award
2005, respectively. from the IAS Electric Machines Committee.
He worked with the National Laboratories for nu-
clear physics. The subject of the Ph.D. thesis was
power converters for particle accelerators. For eleven
years, he focused on the development of power con-
verters and control electronics for high-end applica-
tions both for science and industry. He later moved
to renewable industry in the field of wind turbines where he contributed to the Alejandro Secades Rodríguez was born in Llangréu,
development of a highly modular system of three-phase back to back inverters. Asturies, Spain, in 1990. He received the graduate
Since 2017, he has been with the Department of Mechatronics, MCI Manage- degree in aerospace engineering, from the University
ment Center Innsbruck,where he serves as a Lecturer in power electronics and Alfonso X el Sabio, Madrid, Spain, in 2014, and the
sensors. His present research interests include the field of high-speed drives, master’s degree in mechatronics & smart technologies
power electronics applied to robotics, and GaN/SiC applications. from the Management Center Innsbruck, Innsbruck,
Austria, in 2017.
Since 2016, he is currently a Project Developer at
the MCI Mechatronics Department. He specializes in
a distinct variety of topics with special focus on CFD,
aerodynamics, advanced CAD design, and control
Fabio Giulii Capponi (Member, IEEE) received the
engineering.
M.S. and Ph.D. degrees in electrical engineering from
In 2018, Mr. Secades Rodríguez was the recipient of the First Place for the
the University of Rome “La Sapienza,” Rome, Italy,
Best Master Thesis in the Austrian Conference of Mechatronics for his thesis
in 1994 and 1998, respectively.
“Simulation of Active Vibration Damping of a Wind Turbine,” taken place in
Since 1996, he has been with Sapienza-University
Graz, Austria.
of Rome, Rome, Italy, where he is a Full Professor
of power converters, electrical machines and drives
at the Department of Astronautical, Electrical and
Energy Engineering. From 2003 to 2004, he was a
Visiting Scholar with the Wisconsin Electrical Ma-
chines and Power Electronics Consortium (WEM-
PEC), University of Wisconsin, Madison, WI, USA. He has authored or coau- Lukas Wild received the bachelor’s degree in mecha-
thored more than 90 published technical papers. His current research interests tronics/electrical engineering in 2019 from the MCI,
include permanent magnet motor drives and multiphysics design of electrical Innsbruck, Austria, where he is currently working
machines. toward the master’s degree in mechatronics/electrical
Dr. Giulii Capponi was the recipient of the 2014 First Prize Paper Award engineering.
and the 2016 Third Prize Paper Award, both from the IAS Industrial Drives Since October 2018, he has been a Project Collab-
Committee. He is a Registered Professional Engineer in Italy and is member of orator with the Emerging Applications Laboratory,
the IEEE INDUSTRY APPLICATIONS, the IEEE INDUSTRIAL ELECTRONICS, and the working in the field of power electronics and motor
IEEE Power Electronics Societies. He is a member of the IEEE IAS Industrial control.
Drives Committee, the Electric Machines Committee, and the Transportation
Systems Committee.

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