Smart Materials and Structures
Smart Mater. Struct. 26 (2017) 115022 (11pp)

https://doi.org/10.1088/1361-665X/aa8f71

Structural parameter study on polymerbased ultrasonic motor
Jiang Wu , Yosuke Mizuno

and Kentaro Nakamura

Laboratory for Future Interdisciplinary Research of Science and Technology, Tokyo Institute ofTechnology, Yokohama 226-8503, Japan
E-mail: wujiang@sonic.pi.titech.ac.jp
Received 21 June 2017, revised 21 August 2017
Accepted for publication 27 September 2017
Published 20 October 2017
Abstract

Our previous study has shown that traveling-wave rotary ultrasonic motors using polymer-based
vibrators can work in the same way as conventional motors with metal-based vibrators. It is
feasible to enhance the performance, particularly output torques, of polymer-based motors byadjusting several key dimensions of their vibrators. In this study, poly phenylene sulﬁde, afunctional polymer exhibiting low attenuation at ultrasonic frequency, is selected as the vibrating
body, which is activated with a piezoelectric ceramic element bonded on its back surface. The
optimal thicknesses of the polymer-based motors are higher than those of metal-based motors.
When the same voltages were applied, the maximum torques and output powers available withthe polymer-based motors were lower than the values of the metal-based motors with the samestructures. The reasons for the lower torque were explained on the basis of vibration modes.
First, the force factors of the polymer-based vibrators are lower than those of metal-based
vibrators owing to the great difference in the mechanical constants between polymers andpiezoelectric ceramics. Subsequently, though the force factors of polymer-based vibrators can beslightly enhanced by increasing their thicknesses, the unavoidable radial vibrations become
higher and cause undesirable friction loss, which reduces the output torques. Though thepolymer-based motors have rotation speeds comparable to those of metal-based motors, their
output power are lower due to the low electromechanical coupling factors of the polymer-based
vibrators.
Keywords: elliptical motion, force factor, polymer, ultrasonic motor(Some ﬁgures may appear in colour only in the online journal)
light USMs have become increasingly demanded [12]. In ourprevious study, metals were replaced with functional polymers as vibrating bodies of USMs to reduce their weight andenhance the mass production efﬁciencies [12–15]. Poly phenylene sulﬁde (PPS) was selected due to its low damping
coefﬁcient of 0.0025, which is 0.05 and 2 times the values foracrylic resin and stainless steel, respectively, under highamplitude vibration (up to 0.04% strain) [12, 14]. We fabricated a PPS-based ring-shaped vibrator to form a travelingwave rotary USM, and tested its performance [15]. Comparedto the metal-based USMs, the PPS-based USM provided
relatively high rotation speed but low output torque. As PPS
exhibits a much lower elastic modulus and density thanmetals, the optimal design of PPS-based USMs should differ
from those of metal-based ones.

1. Introduction
Ultrasonic motors (USMs), which provide relatively low
speeds, high torques, and quick responses, compared to theelectro-magnetic motors, have been practically applied to apart of cameras and optical instruments for over two decades[1, 2], and also expected as potential key actuators for
robotics [3, 4]. In general, USM consists of a vibrator and arotor, which are pressed to each other to effectively utilize thefriction force. When two channels of voltages with a 90°
phase are applied to a vibrator, a traveling wave or two
degenerated standing waves are excited, and elliptical motion
at vibrator surface drives the rotor [3–6]. Vibrators in theconventional USMs were commonly composed of piezoelectric ceramics and a metal vibrating body [7–11]. Recently,
0964-1726/17/115022+11$33.00

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J Wu et al

Figure 1. (a) Vibrator structure, (b) electrode division of piezoelectric ceramic element, and (c) the 3rd bending mode and elliptical motion at

vibrator surface.

Table 1. Mechanical constants of polymer, metals, and piezoelectric ceramics.

Density (×103 kg m−3)

Material

Elastic modulus (GPa)Poisson’s ratio

Polymer:

PPS

1.35

3.45

0.36

Metals:

Copper alloy
Aluminum alloy
Stainless steel

8.50
2.70
7.90

117
70.3
197

0.31
0.35
0.31

Piezoelectric ceramics

145 (Y11)

7.80

In this paper, ﬁrst, we explore how elliptical shapes andforce factors of the polymer-based vibrators vary as several
key structural parameters are changed. Subsequently, we
attempt to improve the maximum torques (output torque atzero revolution speed) of the polymer-based USMs byadjusting several key dimensions, and discuss the reason for
their relatively low output torques and power compared to themetal-based USMs.

—

one side was evenly divided into 12 parts, while the other side
without division was glued to the back surface of the PPS
vibrating body with epoxy. When four sinusoidal voltages
with phases were applied to the vibrator, a traveling wave inthe 3rd bending mode shown in ﬁgure 1(c) was generated
along the circumference. The vibration velocity components
at the top surface of the vibrator along the z, θ, and r axes are
represented by vz, vθ, and vr, respectively. Ratios vθ/vz andvr/vz show elliptical motion shapes, which are altered by thevibrator dimensions [16, 17].

2. Conﬁguration of USM and shape parameters
Figure 1(a) depicts the conﬁguration of the vibrator, which
consists of a PPS vibrating body and a piezoelectric ceramicelement. The vibrating body had a 30 mm diameter cylindrical part and a 0.5 mm thick bottom disk. Radial slots were
fabricated on the cylindrical part with an interval of 10°. The
vibrator has following shape parameters: the un-slotted partthickness t, the teeth height h, the thickness of piezoelectric
ceramic plate p, the inner diameter of the cylindrical part Din,
and the slot width w. A central hole was made on the bottom
disk to ﬁx the vibrator to a shaft with two bolts. The
mechanical constants of PPS are listed in table 1. A polar
coordinate (z, θ, and r axes) is set on the top surface of thevibrator. The piezoelectric ceramic annular disk (C213, Fuji
Ceramics, Fujinomiya, Japan) had the inner and outer diameters of 20 and 30 mm, respectively, and a thickness of
0.5 mm. Thin silver electrodes were sintered on both sides ofthe piezoelectric ceramic element. As shown in ﬁgure 1(b),3. Elliptical motion shape and force factor versus
structural parameters3.1. Un-slotted part thickness

The elliptical motion shape (ratios vθ/vz and vr/vz) and theforce factor of the vibrator were investigated because they are
essential to USMs [1, 3]. The force factor represents the ratio of
the current to the vibration velocity or equivalently the ratio of
the generated force to the voltage [2, 6]. These indicators reﬂect
the vibrator characteristics, and affect the USM performance[2, 16]. In this work, we discuss the force factor deﬁned for thetangential vibration velocity vθ because it is parallel to therotation direction. The vibration velocity ratios and the force
factors of the PPS-based vibrators with varying un-slotted partthicknesses were obtained through both ﬁnite element analysis(FEA) and experiments. The three orthogonal vibration
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thicker vibrator generates higher force factors and higher
radial vibrations.
3.2. Other shape parameters

Then, we investigated how other parameters affect the elliptical motion shapes and the force factors through FEA.
Figures 4(a) and (b) show how the vibration velocity ratios
and the force factors vary as the piezoelectric ceramic thicknesses p is changed, respectively. Both the ratios vθ/vz andvr/vz have little changes. When the t10 vibrators have 0.05
and 1.50 mm thick piezoelectric ceramic elements, the force
factors are 0.25 and 0.28 N V−1, higher than the values when
the piezoelectric ceramic elements have moderate thicknesses.
Figures 5(a) and (b) demonstrate how the vibration velocity
ratios and the force factors depend on the teeth height h,respectively. The ratios vθ/vz change from 0.5 to 0.9 when theteeth heights ranged from 1 to 6 mm. With decreasing teeth
height, the vibrators provide increases in the force factors andreductions in the ratio vr/vz. The vibration velocity ratios andthe force factors are plotted as a function of the slot width w
in ﬁgures 6(a) and (b), respectively. When their slots arelower than 0.5 mm, both elliptical motion shapes and force
factors exhibit little changes. As shown in ﬁgures 7(a) and(b), the ratios vθ/vz have minimal values when the innerdiameters are approximately 22 mm, where the force factors
reach their peak values. The ratios vr/vz increase sharply as
the radial thicknesses become lower. As ﬁgures 8(a) and (b)show, the optimal inner diameter is determined by the innerdiameter of the piezoelectric ceramic element Dp-in, whereas
the maximal force factor and the minimal ratio vθ/vz are
independent to Dp-in. Since the un-slotted part thickness andteeth height have relatively high inﬂuences on the ellipticalmotion shapes and the force factors among these structuralparameters, we change these parameters to improve the output
torques and powers of the polymer-based USMs.

Figure 2. Simulated (lines) and experimental (dots) results ofresonance frequency.

velocity components were simultaneously measured using an
in-plane laser Doppler velocimeter (IPV100, Polytec, Waldbronn, Germany) and two out-of-plane laser Doppler velocimeters (CLV1000, Polytec, Waldbronn, Germany). The current
was measured with a current probe (P6021, Tektronix, Beaverton, US). The vibration velocities vz and vr were measured at
the antinodes, while the vibration velocity vθ was at the node
under the standing wave excitation. The ratios vθ/vz and vr/vz
in the standing wave are exactly equal to those of the travelingwave [1].
First, we explore how the elliptical motion shape and theforce factor change with varying un-slotted part thickness.
The other parameters are set as p=0.5 mm, h=4 mm,
Din=20 mm, and w=0.5 mm. Figure 2 shows thedependence of the resonance frequency of the 3rd bending
mode on the un-slotted part thickness. It increases from 5 to
11 kHz as the vibrator becomes thicker, and levels off at theun-slotted part thickness of 6 mm. For the cylindrical vibrators, the bending vibrations along the z axis couple with thosealong the r axis. The z-axis bending vibration is dominantly
excited on the thin vibrators, and the resonance frequency
increases because the equivalent stiffness becomes higher
with increasing thicknesses. In contrast, the resonance frequencies on thicker vibrators are mainly determined by theradial thicknesses Dout−Din of their cylindrical parts.
Figures 3(a) and (b) show the vibration velocity ratios and theforce factor as a function of the un-slotted part thickness,
respectively. As shown in ﬁgure 3(a), even when the vibrator
became thicker, the ratio vθ/vz exhibited only small variation.
In contrast, the ratio vr/vz of the vibrator with a 2 mm thickun-slotted part (t2 vibrator) was 0.7, and the value of the t10
vibrator increased to 2.8. With increasing thickness, theequivalent stiffness and mass increase, leading to the reduction in the z-axis vibration velocity [17]. However, theequivalent stiffness and mass in radial direction, determinedby the radial thickness, show almost no change. Thus, theratio vr/vz has an observable increase. As the un-slotted partthickness was varied from 2 to 10 mm, the force factor inﬁgure 3(b) increased from 0.02 to 0.23 N V−1. Observably,

4. Performance of motors with varying dimensions
4.1. Experimental setup

In this section, we experimentally investigated how the motor
performance depended on the structural parameters.
Figure 9(a) illustrates the experimental setup for measuring
the motor performance. The vibrator was clamped to astainless steel shaft, ﬁxed at the center of a stainless steel baseplate. To reduce the friction loss in holding the rotor, as
shown in ﬁgure 9(b), a pivot bearing composed of a conicalhead and a dent, both of which were made of stainless steel,
was used. The conical angles of the bearing head and the dent
were set to 55° and 60°, respectively. The upper shaft of thepivot bearing was inserted into a Teﬂon sleeve, guided along
the central hole in the Teﬂon top plate. The dent was ﬁxed at
the center of the rotor. Preload was applied to the rotor byloading weights on the top of the shaft. As shown inﬁgure 9(c), the rotor consisted of a contacting disk and a ringshaped ﬂywheel, both of which were made of aluminum. The
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Figure 3. Simulated (lines) and experimental (dots) results of (a) elliptical motion shape and (b) force factor versus un-slotted part thicknessof PPS-based vibrator.

Figure 4. (a) Elliptical motion shape and (b) force factor as a function of the thickness of piezoelectric ceramics.

Figure 5. (a) Elliptical motion shape and (b) force factor versus slotted part thickness.

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Figure 6. (a) Elliptical motion shape and (b) force factor as a function of slot width.

Figure 7. (a) Elliptical motion shape and (b) force factor as a function of inner diameter.

Figure 8. (a) Ratios vθ/vz and (b) force factors versus inner diameters of vibrators composed of piezoelectric ceramics with different innerdiameters.

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Figure 9. Testbed for evaluating motor performance. (a) Front view, (b) pivot bearing, and (c) rotor and additional ﬂywheel.

acceleration yielded a reduction. Since the lower and higher
cutoff frequencies of the laser Doppler vibrometer were set to
0.2 Hz and 100 kHz, respectively, the obtained signal included ultrasonic components, which appeared to be the ﬂuctuation in the transient response. By implementing the
exponential ﬁtting curve j(t) on the transient response, themaximum torque Tmax, maximum angular velocity jmax, andmaximum output power Pmax are calculated withd [j (t )]d [j (t )]
,
=J⋅
dt
dt
t=0
jmax = lim [j (t )] ,Tmax = J ⋅ max
Figure 10. Rotation speed measurement with two out-of-plane laser

{

}

t +¥

Doppler velocimeters.

(1 )
(2 )

and
Pmax = J ⋅ max

contacting surface was limited to a portion having outer andinner diameters of 30 and 28 mm, respectively, because thevibrators provided relatively extensive vibrations on the outeredges of the contacting surface (see ﬁgure 1(c)). In this study,
we employed the transient method, repeatedly used in previous studies on USMs [18–20], to evaluate the motor performance. The ﬂywheel has a moment of inertia J of
1.24×10−5 kg m2. As shown in ﬁgure 10, ﬁrst, the transient
response of the rotor speed was measured with two out-ofplan laser Doppler velocimeters: two laser beams illuminatedthe same point on the outer surface of the rotor with the sameangles. The tangential velocity was calculated from the outputs of the two laser Doppler velocimeters on the basis of asimple geometry [21]. Figure 11 gives an example. The
measured transient response of the rotational speed of the
motor with the t5 vibrator is shown in ﬁgure 11(a). The
voltages and the preload are respectively 250 V and 1.04 N.
Initially, the angular velocity increased from 0 to 1.21 rad s−1,
and leveled off at approximately 48 ms. Note that, though theangular velocity became higher in the rise part, the angular{d [j (t )]⋅ j (t ) ,
dt

}

(3 )

respectively. As shown in ﬁgure 11(b), the motor yielded the
maximum torque and angular velocity of 1.39 mNm and1.21 rad s−1, respectively, at the beginning and the end of therise part. The maximum output power reached 0.42 mW whenthe motor exhibited a moderate torque.
4.2. Experimental results

Figure 12 illustrates the motor performance with the t5
vibrator as a function of the driving voltage under different
preloads. As ﬁgure 12(a) shows, there existed a dead region in
the maximum torque when the voltage was lower than
100–150 V. Beyond the dead region, the maximum torqueincreased in proportion to the driving voltage and saturated
according to the preload. At a certain voltage, the maximum
torque became higher with increasing preload [2]. When the
preload exceeded a certain value, the maximum torquedecreased because of friction loss. For example, the maximum torque at 150 V reached 0.4 mNm when the optimal
preload of 0.37 N was applied to the vibrator. The maximum
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Figure 11. (a) Transient response (gray curve) of rotational speed and its exponential ﬁtting curve (black curve), and (b) load characteristics

of PPS-based USM with t5 vibrator.

Figure 12. Motor performance with t5 vibrator as a function of driving voltage under different preloads: (a) maximum torque, (b) maximum

rotation speed at 250 V, and (c) maximum power.

torque under the optimal preload is suitable for representing
the vibrator performance. Figure 12(b) shows that the motor
provided its maximum no-load speed when the minimal
preload and the maximal voltage are applied to the vibrator.
As ﬁgure 12(c) shows, at a certain voltage, the output power

initially became higher with increasing preload, and then,became lower owing to the rapid reduction in the rotation
speed.
Figure 13 shows the motor performance with varying
thickness of the un-slotted part at driving voltages of 250 V.
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Figure 13. Performance of motors with varying un-slotted part thicknesses. (a) Maximum torques at optimal preloads, (b) maximum no-loadrotation speeds, and (c) maximum power when voltages of 250 V were applied.

factor and the r-axis vibration velocity increase, and thenegative effect on the torque caused by the friction loss offsets the positive effect due to the enhancement in the force
factor. Thus, the rapid reduction in the maximum torque inthe thickness range from 8 to 10 mm originates from the large
increase in the friction loss. To quantitatively estimate theoutput torque, a simulation model for the friction on thecontacting surface between the polymer-based vibrator andthe metal rotor should be developed in the future. Since thepolymer-based motor has a softer vibrator and a harder rotor,
it may provide a friction different from the conventional
metal-based motor, of which the vibrator is commonly harderthan the rotor [6, 7, 23, 24]. The previous study shows that theconventional metal-based motor with the same diameteryields the highest torque when its un-slotted part thickness isless than 4 mm [7]. Clearly, a thicker vibrator is needed for
the PPS-based motor. As ﬁgure 13(b) shows, the maximum
no-load speed decreased as the vibrator became thicker
because the θ-axis vibration velocity was lowered. The
maximum power shown in ﬁgure 13(c) decreased from 3.0 to
0.8 mW as the un-slotted part thickness ranged from 2 to
10 mm owing to the reduction in electromechanical coupling
factor. The friction loss also lowers the output power particularly for the thick vibrators.
Figure 14 shows the performance of motors with varying
teeth height. Their un-slotted part thicknesses are 8 mm. The
motors with lower teeth height yielded higher output torques.
When a preload of 6.4 N was applied to the vibrator with

The performance of each motor was evaluated at random
three different positions. Figure 13(a) shows the maximum
torques when optimal preloads were applied to the motors. As
the un-slotted part thickness increased from 2 to 8 mm, themaximum torque exhibited an increase from approximately
0.9–3.0 mNm. However, the maximum torque decreased from
3.0 to 0.8 mNm when the thickness increased from 8 to
10 mm. Under a certain voltage and a sufﬁciently high preload, a thicker vibrator generates a higher driving force to therotor. This is a reason for the increase in the maximum torquewhen the un-slotted part thickness was varied from 2 to 8 mm.
In general, slip exists at contacting surfaces in traveling-wave
USMs because there are velocity differences between thevibrator and the rotor along both the θ and r axes [22]. As
mentioned above, the θ-axis vibration velocity is parallel tothe rotational direction. The points near the crest, which yieldhigher θ-axis vibration velocities than the rotation speed,
accelerates the rotor. Slip is generated on a part of the points
with relatively low vibration velocities [7]. In contrast, since
the r-axis vibration velocities are vertical to the rotation
direction, slip is generated on all contacting points except for
those with radial vibration velocities of zero (theoretically,
six points with the 3rd bending mode). Besides, the r-axis
vibration velocities, particularly in the cases of the thicker
vibrators, are higher than the θ-axis vibration velocities (seeﬁgure 3(a)). Thus, the frictional loss caused by slip should
be dominantly attributed to the r-axis vibration component[23, 24]. As the vibrator becomes thicker, both the force8Smart Mater. Struct. 26 (2017) 115022

J Wu et alvibrator of the PPS-based motor was structurally optimized,
its output torque and power were still lower. To explore thereason for the relatively low torque, we compared the vibration velocity ratios and the force factors between the PPS- andmetal-based vibrators with the same structures. The material
constants of metals are given in table 1. As shown inﬁgure 15(a), there is no observable difference in the ratiosvθ/vz and vr/vz among the vibrators. In contrast, the force
factor in ﬁgure 15(b) shows a great difference among thesematerials. For example, the copper-alloy-based vibrator withan 8 mm-thick un-slotted part exhibits a force factor of
0.99 N V−1, higher than those of the stainless-steel-, aluminum-alloy-, and PPS-based vibrators (0.65, 0.60 and
0.12 N V−1, respectively). To obtain a force factor of
0.15 N V−1, the sufﬁcient thickness is 2.5 mm for the copperalloy-based vibrator and 10 mm for the PPS-based vibrator.
Thus, a thicker vibrator is needed for the PPS-based motor to
generate a driving force comparable to that of the metal-based
motor. Among the tested materials in table 1, the elastic
modulus and the density of copper alloy are relatively close to
those of piezoelectric ceramics, and the copper-alloy-based
vibrator provides a relatively high force factor. In contrast,
there is a signiﬁcant difference in the mechanical constants
between the vibrating body and the piezoelectric ceramicelement of the PPS-based vibrator, and it may result in the
low force factor. Thus, the torque of the PPS-based USM is
difﬁcult to signiﬁcantly enhance by adjusting the dimensions
because it is dominantly restricted by low mechanical constants of PPS. In addition, since the thinner vibrators, whichhave sufﬁciently high force factor, are used in metal-based
motors, their radial vibration components are lower thanpolymer-based ones. The motors made of other commonly
employed functional polymers also face the same problems
because they have almost similar elastic moduli and densities[12, 13]. The lower output powers of PPS-based motors originate from their low electromechanical coupling factors,
which are lower than 0.1 times the value of the stainlesssteel-based vibrator.

Figure 14. Performance of motors with varying teeth height.

t=8 mm and h=1 mm, the maximum torque at 250 V
reached 6.7 mNm, approximately 10 times the value of ourprevious motor prototype [12]. The maximum no-load speed
and the output power became lower and higher, respectively,
as the teeth height decreased.
Figures 7 and 8 indicate that it is difﬁcult to signiﬁcantly
increase the output torque by adjusting the slot width w andthe inner diameter Din since they are close to the optimalvalues. Figure 5 indicates that a larger torque can be obtainedby using the piezoelectric ceramic elements with thicknessesof less than 0.05 mm or over 1.5 mm. In our experiments,
piezoelectric ceramics of less than 0.2 mm in thickness were
unavailable because they were easy to fracture1, 2. Thinpiezoelectric ﬁlms can be obtained by sputtering and hydrothermal synthesis [20, 25]. However, a large-area piezoelectric ﬁlm, e.g., a piezoelectric ceramic element with a
30 mm outer diameter used in this study, has not beenreported. Besides, it is feasible to enhance the output torqueand power by using the piezoelectric ceramic elements withthicknesses of over 1.5 mm, but the USM weight becomeslarger as piezoelectric ceramics have much higher densities
than PPS. In the future, the optimal thickness of piezoelectric
ceramics will be discussed considering the power density (theratio of output torque to weight) of USM.

6. Conclusion
5. Discussions: Performance comparison between
polymer- and metal-based vibrators

To enhance the performance of polymer-based USMs, apreliminary investigation on the shape parameters of theirvibrators was carried out. Through analytical and experimental explorations, the following conclusions wereobtained:

For comparison, we fabricated a stainless-steel-based vibrator
with a 4 mm thick un-slotted part and 4 mm high teeth, andglued a 0.1 mm thick PPS ﬁlm on its top surface to obtain thesame friction characteristics as the PPS-based motors. The
resonance frequency and the electromechanical coupling
factor corresponding to the 3rd bending mode were 25.89 kHz
and 4.9%, respectively. At 250 V, its maximum torque androtation speed were 20.8 mNm and 10.8 rad s−1, respectively,
and its maximum output power reached 31.6 mW. Though the(1) At 250 V, the PPS-based motor with a 30 mm diameteryielded the maximum torque of 6.7 mNm, 10 times thevalue of our previous polymer-based motor.
(2) Relatively low torques of polymer-based motors arecaused by low force factors and undesirable radialvibration components of their vibrators. Relatively low
output power is mainly attributed to the low electromechanical coupling factor.

1

Manual of Fuji Ceramic’s products: http://fujicera.co.jp/product/elements/
Manual of APC International Ltd’s products: https://americanpiezo.com/
product-service/custom-piezoelectric-elements/shapes-sizes.html

29Smart Mater. Struct. 26 (2017) 115022

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Figure 15. (a) Elliptical motion shapes and (b) force factors of PPS-, stainless-steel-, aluminum-, and copper-alloy-based vibrators.

(3) The low force factor of polymer-based vibrator possibly
originates from the difference in the mechanicalconstants between polymer and piezoelectric ceramics.
(4) Higher high force factor and lower radial vibration
component are desirable for not only polymer- but also
metal-based cylindrical vibrators.

Kentaro Nakamura
2899-4484

https://orcid.org/0000-0003-

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Measurement of mechanical quality factors of polymers inThough polymer-based motor cannot generate an output
torque comparable to those of the commercially available
motors, it still has some applications, for example, in thechemical industry because polymer is resistant to acid andalkali [23], or as an actuator in machines where the lightweight is especially required. PPS with a ﬁber reinforced
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Acknowledgments
The authors are indebted to the Daicel Corporation for providing the functional polymers, and the staff members of thePrecision and Manufacturing Center, Technical Department,
Tokyo Institute of Technology, for machining the motor
components used in this study. They also thank Dr Wei Qiu,
who is presently with the Department of Physics, Technical
University of Denmark, and formerly Laboratory for FutureInterdisciplinary Research of Science and Technology, Tokyo
Institute of Technology, for the fruitful discussions with him.
This work was supported by the Adaptable and Seamless
Technology Transfer Program, A-step, Grant Number
AS2711904S, from the Japan Science and Technology
Agency, JST, and the work of J Wu was supported by aGrant-in-Aid JSPS Research Fellows (No. 17J05057).

ORCID iDs
Jiang Wu https://orcid.org/0000-0002-6899-5601
Yosuke Mizuno https://orcid.org/0000-0002-3362-4720
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