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Vol.48 No.1
JournalJournal of the Magnetics Society of Japan
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2024

Measurement Technique, High-frequency Devices, Magnetic Imaging
Induced Pulse Voltage and Hysteresis Loss During Magnetization Reversal of Wiegand Wire
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F. Xue, C. Yang, Z. Song, and Y. Takemura…1Power Magnetics

Development of Cylindrical Linear Actuator for Vertical Transfer: Fundamental Consideration of Eﬀect of Shape on
Thrust Characteristics
H. Nosaka, D. Uchino, K. Ogawa, T. Kato, K. Ikeda, A. Endo, T. Narita, and H. Kato

…5

Linear Actuators for High-Speed Reciprocating Motion with Dual Halbach Arrays(Fundamental Consideration of the Eﬀect of Magnet Arrangement on Thrust Characteristics)
M. Tanaka, K. Kimura, J. Kuroda, Y. Majima, D. Uchino, K. Ogawa,T. Kato, K. Ikeda, A. Endo, T. Narita and H. Kato

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Journal of the Magnetics Society of Japan
Vol. 48, No. 1

Electronic Journal URL: https://www.jstage.jst.go.jp/browse/msjmag

CONTENTS
Measurement Technique, High-frequency Devices, Magnetic Imaging
Induced Pulse Voltage and Hysteresis Loss During Magnetization Reversal of Wiegand Wire………………………………………………………… F. Xue, C. Yang, Z. Song, and Y. Takemura

1Power Magnetics
Development of Cylindrical Linear Actuator for Vertical Transfer: Fundamental Considerationof Effect of Shape on Thrust Characteristics…………………………………… H. Nosaka, D. Uchino, K. Ogawa, T. Kato, K. Ikeda, A. Endo,		T. Narita, and H. Kato

5

Linear Actuators for High-Speed Reciprocating Motion with Dual Halbach Arrays(Fundamental Consideration of the Effect of Magnet Arrangement on Thrust Characteristics)………………… M. Tanaka, K. Kimura, J. Kuroda, Y. Majima, D. Uchino, K. Ogawa,T. Kato,		

K. Ikeda, A. Endo, T. Narita and H. Kato

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Copyright ©2024 by the Magnetics Society of Japan.
This article is licensed under the Creative Commons Attribution International License (CC BY 4.0)
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J. Magn. Soc. Jpn., 48, 1-4 (2024)<Paper>Induced pulse voltage and hysteresis loss during magnetization reversal of
Wiegand wire
Fei Xue, Chao Yang*, Zenglu Song*, and Yasushi Takemura

Division of Electrical and Computer Engineering, Yokohama National University, Yokohama 240-8501, Japan*

School of Electrical Engineering, Nanjing Vocational University of Industry Technology, Nanjing 210023, China

When an alternating magnetic field is applied to a Wiegand wire, a series of pulse voltages are induced in a pickup
coil due to fast magnetization reversal known as the large Barkhausen jump. This study aims to analyze thecharacteristics of the pulse voltage, including its height and area, which are determined from the intensity andfrequency of the applied alternating magnetic field. We quantified the energy generated by magnetization reversal ina Wiegand wire. To achieve this, it was necessary to measure the energy consumed by a load resistor connected to the
pickup coil. Additionally, we clarified the energy loss in the Wiegand wire.
Key words: Wiegand wire, Wiegand effect, large Barkhausen jump, magnetization reversal, electricity generator.
1 . Intr oduc tion

2. Methods and Materials

Recent innovations in electronics have enabled thedevelopment of low-energy-consuming and battery-lessdevices. Wiegand sensors1-2), which generate a pulse
voltage during magnetization reversal of the magnetic
wire installed in the sensors, are expected to findapplication in various self-powered modules3). Thesesensors are commercially available and have recently
been used as a power source in battery-free rotary
encoders4). This in turn has necessitated the evaluationof Wiegand sensors, particularly in terms of circuit
applications and energy losses in the circuit.
The fast magnetization reversal in a Wiegand wire
induces a pulse voltage in the pickup coil wound aroundthe wire5-6). This reversal is accompanied by a large
Barkhausen jump7-9). The amplitude and width of theinduced pulse voltage, known as the Wiegand pulse, areindependent of the frequency of the applied magnetic
field10). The energy associated with the output pulse
voltage generated by the Wiegand sensor can be used asa power supply for equipment without batteries11-14). The
application of Wiegand sensors as receiving coils in
wireless power transfer has also been reported15).
Previous research has investigated the energy
produced by the Wiegand sensor as an electricity
generator and the power transferred to the Wiegand
sensor when it is used as a receiver coil in wireless powertransfer10,15). However, the relationship between theenergy consumption and energy loss remains unclear.
This study seeks to evaluate the energies of the induced
pulse voltage, the loss in the Wiegand sensor, and thepower consumption of the load resistance. We quantified
the energy and loss, as well as the efficiency of the energy
consumed in the load resistance.

The Wiegand sensor used in this study comprised a
Wiegand wire (11 mm long and 0.25 mm in diameter) anda pickup coil of 3000 turns wound around the wire. Twoferrite beads (3 mm long, 1.3 mm in inner diameter, and
3.5 mm in outer diameter) were attached at both ends ofthe wire to concentrate the applied magnetic field to thewire (Fig. 1). The atomic composition of the Wiegand wirewas
Fe0.4Co0.5V0.1.
Other
specifications
andcharacterizations of the Wiegand sensor have been
documented in prior research16). As shown in Fig. 1, amagnetic field generated by an excitation coil was
externally applied to the sensors. An alternating current(AC) magnetic field of μ0H = 3–13 mT at 10–1000 Hz wasapplied in this experiment.
A series of pulses of alternating positive and negative
voltages were induced in the pickup coil when an AC
magnetic field was applied to the Wiegand sensor. Figure
2 shows a typical waveform of a positive pulse voltage.
An AC magnetic field of μ0H = 5 mT was applied at 1 kHz.
The output voltage Voutput comprised two components,
VWiegand and Vinduced, as shown in the figure. VWiegand is theCorresponding author: Yasushi Takemura(e-mail: takemura-yasushi-nx@ynu.ac.jp).

Fig. 1 Structure of Wiegand sensor. AC magnetic field
generated by exciting coil is applied to
Wiegand sensor.

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Fig. 2 Typically measured output voltage (black, Voutput)
induced in pickup coil. AC magnetic field of μ0H= 5 mT at 1 kHz was applied. Output voltage
consists of 2 components: Wiegand pulse
voltage (red, VWiegand) and induced voltage fromapplied AC magnetic field (blue, Vinduced).
Wiegand pulse voltage, while Vinduced is the voltage
induced by the applied magnetic field. The pulse width
was approximately 10 μs and independent of thefrequency of the applied magnetic field, which was aunique feature in the magnetization reversal of the
Wiegand wire.
Figure 3 shows the amplitude of the output voltage
Voutput. An AC magnetic field of μ0H = 3–13 mT wasapplied at 10–1000 Hz. The amplitudes of the positive
and negative pulses are asymmetrical. There remains a
region of a higher coercive force in the wire, which does
not contribute to the large Barkhausen jump. The
magnetization of this region is not reversed. The
magnetization of the region exhibiting a large
Barkhausen jump is alternately reversed. The
magnetizations of these two regions can be either parallelor antiparallel5), and the magnetostatic coupling between
the regions is asymmetric, resulting in asymmetricpositive and negative voltages. The output voltage
increases with the intensity of the applied AC magnetic
field. This dependence is due to an increase in the volume
of magnetization reversal in the Wiegand wire. The
amplitude of the pulse voltage increases slightly with theFig. 3 Amplitudes of positive and negative output
voltages induced in Wiegand sensor. AC
magnetic field of μ0H = 3–13 mT at 10–1000
Hz was applied.

2

Fig. 4 Applied field frequency dependence of Wiegand
pulse area. AC magnetic field of μ0H = 3–13 mT
at 10–1000 Hz was applied. Amplitudes of
positive and negative output voltages inducedin Wiegand sensor. AC magnetic field of μ0H =3–13 mT at 10–1000 Hz was applied.
applied field frequency, which is attributed to the voltage
induced by the applied AC field, Vinduced. The amplitudeof the Wiegand pulse voltage, VWiegand, is independent of
the frequency.
The area of the voltage pulse waveform in units of[V•s] corresponds to the energy. Figure 4 shows the areaof the Wiegand pulse voltage, VWiegand. The induced
voltage Vinduced was not included in the calculation of thepulse area. The area of the Wiegand pulse was notdependent on the applied field frequency. The area wasalmost constant for an applied field intensity above μ0H≥ 5 mT, which agrees with the result shown in Fig. 3.
3. Experimental results and discussion
3.1 Energy consumption by load resistanceThis section discusses the energy consumed by the load
resistance connected to the Wiegand sensor. We prepareda circuit comprising a series connection of the pickup coilof the Wiegand sensor and the load resistance R (Fig. 5).
The output voltage across the load resistance wasmeasured, and the energy consumption was calculated as
described later. The DC resistance of the Wiegand sensor

Fig. 5 Experimental circuit to study energy
consumption at load resistance connected
to Wiegand sensor.

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

INDEX(dominated by the pickup coil resistance) was r = 273 Ω.
This is the internal resistance of the Wiegand sensor
used as a power source. The load resistance was variedin the range of R = 30–1000 Ω.
The electric energy consumption by the load resistanceR = 100 Ω during one cycle of the applied AC magnetic
field, Eoutput, was calculated using Eq. (1).
� �������� �����

𝐸𝐸������ � ��

�

𝑑𝑑𝑑𝑑

(1)

The output voltage Voutput in Fig. 2 is equivalent to thevoltage across the load resistance. T is the period of theapplied AC magnetic field. Two positive and negative
pulses were generated during this period. Figure 6(a)shows the energy consumption by the load resistance
Eoutput depending on the applied field frequency. It
increases with increasing the applied field frequency, andis of the order of 0.1 µJ at 10 Hz and 1 µJ at 1 kHz. Eoutput
is equivalent to the summation of the generated energy
associated with the Wiegand pulse EWiegand and theinduced voltage Einduced. EWiegand was calculated using Eq.

(2).
� �������� ������������ �����

𝐸𝐸������� � ��

�

𝑑𝑑𝑑𝑑

(2)

As shown in Fig. 6(b), EWiegand is essentiallyindependent of the applied field frequency, which is a
specific feature of the large Barkhausen jump. However,
the reason for the slight increase in EWiegand at 1 kHz is
not clear. It was also confirmed that Einduced was
proportional to the frequency, based on Faraday’s law of
induction (Fig. 6(c)).
3.2 Relationship of energy consumption with resistancesand hysteresis loss
Figure 7 shows the minor hysteresis loop of the
Wiegand wire derived from the output voltage induced in
the pickup coil. An AC magnetic field of μ0H = 7 mT in
amplitude at 1 kHz was applied. A Barkhausen jump wasobserved at an applied field intensity of μ0H = ±2−3 mT.
The closed area of the hysteresis loop corresponding to
the hysteresis loss was calculated. Figure 8 shows theenergy associated with the hysteresis loss during one
cycle of the applied alternating magnetic field at 1 kHz.
The skin depth at 1 kHz was calculated as 4.9 mm, which
was larger than the radius of the Wiegand wire (0.125mm). Therefore, the eddy current loss in the wire wasnegligible. The hysteresis loss increased with theintensity of the applied AC magnetic field in the range of
μ0H = 0–13 mT.

(a) Eoutput

Fig. 7 Hysteresis loop of Wiegand wire derived frominduced output voltage in pickup coil.
(b) EWiegand(c) Einduced
Fig. 6 Electric energy consumption: (a) Eoutput, (b)EWiegand, and (c) Einduced.

Fig. 8 Energy associated with hysteresis loss in
Wiegand wire during one cycle of appliedalternating magnetic field depending on field
intensity.

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Finally, we analyzed the energy consumption of theentire circuit shown in Fig. 5. Figure 9 shows the energy
consumption by the load resistance Eload and internal
resistance Einternal during one cycle of the applied AC
magnetic field of μ0H = 3–13 mT at 1 kHz. The connectedload resistances were R = 30, 100, 300, and 1000 Ω.
Because the load and internal resistances were seriesconnected, the ratio Eload/Einternal equaled that of these
resistances. The hysteresis loss, Ehysteresis, is also shownin Fig. 9. It was calculated from the area of a minor
hysteresis loop (shown in Fig. 7), which was derived fromthe open circuit voltage induced in the pickup coil. In theexperimental setup shown in Fig. 5, the energy supplied
from the excitation coil to the circuit including the pickup
coil and Wiegand wire is not considered as the input
energy. As the intrinsic value of the hysteresis loss
should not depend on the load resistance, Ehysteresis
increases monotonically with increasing the applied field
intensity.
As shown in Fig. 9, the Eload for R = 300 Ω is larger thanthose for the other resistances. This is due to theimpedance matching between the load resistance and theinternal resistance of 273 Ω. Eload, Ehysteresis, and Einternal
increase with the applied field intensity. The energy
efficiency of the load resistance against the total energy
consumption is a significant parameter for practicalapplications. Figure 9 also shows the energy efficiency
calculated using Eload / (Eload + Ehysteresis + Einternal). The
efficiencies are approximately 8, 24, 40, and 47% for R =30, 100, 300, and 1000 Ω, respectively. They remain

almost constant regardless of the applied field intensity
for μ0H ≥ 5 mT. They increase with increasing R, whichis due to less energy consumption at the internal
resistance with less current through the circuit.
4. Conclusion
We measured the output voltage induced in a pickup
coil during fast magnetization reversal in a Wiegand wire
and analyzed its energy, which was available in a load
resistance connected to the pickup coil. The output
voltage, Voutput, was composed of the voltage originatedfrom the Wiegand pulse, VWiegand, and the voltage inducedby the applied alternating magnetic field, Vinduced. VWiegand
and its pulse area were independent of the frequency of
the applied field, whereas Vinduced was proportional to thefrequency. We determined the energy consumption in theload resistance, Eload to be approximately 0.07 µJ. We also
quantified the energy losses in the internal resistance ofthe Wiegand sensor and energy consumed as themagnetization hysteresis loss of the Wiegand wire. Theenergy efficiency calculated by the ratio of the energy
consumed by the load resistance and energy generated in
the Wiegand sensor was also determined to beapproximately 8–47% depending on the load resistance.
Acknowledgements This study was partially supported
by JSPS KAKENHI JP22K18288.
R e f er e n c e s
1)
2)
3)
4)
5)
6)
7)
8)
9)
10)
11)
12)

Fig. 9 Left vertical axis; Energy consumption by load
resistance (Eload, orange), hysteresis loss
(Ehysteresis, blue), and energy consumption byinternal resistance (Einternal, green) under applied
AC magnetic field of μ0H = 3–13 mT at 1 kHz.
Connected load resistance was R= 30, 100, 300,
and 1000 Ω. Right vertical axis; energy efficiency
consumed by load resistance (red bold circle).

4

13)
14)
15)
16)

J. R. Wiegand and M. Velinsky: U.S. Patent #3,820,090,(1974).
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and F.B. Humphrey: IEEE Trans. Magn., 32, 4532 (1996).
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Kawamura, J. Yamasaki, and I. Ogasawara: IEEE Trans.
Magn., 23, 3242 (1987).
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Chen: Micromachi., 13, 623 (2022).
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Received June 25, 2023; Revised Sep. 14, 2023; Accepted Sep. 27,
20 23

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

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Copyright ©2024 by the Magnetics Society of Japan.
This article is licensed under the Creative Commons Attribution International License (CC BY 4.0)
http://creativecommons.org/licenses/by/4.0/

J. Magn. Soc. Jpn., 48, 5-10 (2024)<Paper>Development of Cylindrical Linear Actuator for Vertical Transfer:
Fundamental Consideration of Effect of Shape on Thrust Characteristics
H. Nosaka, D. Uchino, K. Ogawa*, T. Kato**, K. Ikeda***, A. Endo****, T. Narita, and H. Kato
Tokai Univ., 4-1-1 Kitakaname, Hiratsuka-shi, Kanagawa 259-1292, Japan50-2 Manori, Nishihasama-cho, Gamagori-shi, Aichi 443-0047, Japan
1404-1 Katakuramachi, Hachioji-shi, Tokyo 192-0982, Japan*** Hokkaido University of Science, 7-15-4-1 Maeda, Teine-ku, Sapporo-shi, Hokkaido 006-8585, Japan**** Fukuoka Institute of Technology, 3-30-1 Wajiro-higashi, Higashi-ku, Fukuoka-shi, Fukuoka 811-0295, Japan* Aichi University of Technology,** Tokyo University of Technology,

In conventional elevators, vibration and twisting of the cable become a serious problem when the cable is long.
Therefore, an elevator that omits the counterweight and is moved by an actuator on a single cable with both ends
fixed has been considered. However, contact decreases the efficiency and damages the cable owing to the inclusion offoreign matter. Therefore, a vertical transportation system is proposed that moves above a long, uniform conductor
cable without contact by using a cylindrical linear induction motor (LIM) as the actuator. A cylindrical LIM above
the reaction plate, which is a cylindrical shell-shaped conductor, has the advantage that the magnetic force acts on
the reaction plate uniformly. This suppresses the vibration in the gap direction. However, a conventional LIM drives
the mover in the horizontal direction. Therefore, a model was constructed for analyzing the cylindrical LIM forvertical transfer, and electromagnetic field analyses were performed using the finite-element method. The thicknessof the reaction plate was varied, and the effect on the generated thrust was investigated. The optimal thickness ofthe cable was determined from the results by obtaining the range of the magnetic flux density acting on the cable.
Key words: magnetic recording, soft magnetic material, superconductor, magneto-optical recording, thin film
1. Introduction
In buildings and structures with multiple levels,
elevators are essential to transport people andmaterials vertically through these levels. A schematic
diagram of a conventional elevator is shown in Fig. 1(a).
A conventional elevator consists of a cargo carrying
people and materials, a counterweight of approximately
the same weight as the cargo, a tether connecting thecargo and counterweight, and two pulleys supporting
the tether. A motor is attached to either the upper orthe lower pulley to provide the power necessary forlifting and lowering. In recent years, many large
buildings with heights of several hundred meters havebeen constructed. However, the cables in large buildings
are extremely long, and vibration and twisting of thecables become serious problems1),2). In addition, because
conventional elevators attach one cargo to one tether,
the area occupied by the elevator at a certain level in ahigh-rise building becomes large. Therefore, we areconsidering an elevator that does not use acounterweight and uses an actuator to raise and lower
the cargos on a single tether with fixed ends, as shownin Fig. 1(b). This method allows multiple cargos to beinstalled on a single tether and is expected to increase
the transport volume3).
On the other hand, space development has been
actively carried out, and satellites or materials for spacestations are transported by rockets. The problem withrockets is that they are extremely expensive to
Corresponding author: T. Narita (e-mail: narita@tsc.utokai.ac.jp).

transport, and the safety of the surrounding area is aconcern if a launch failure occurs. A space elevator thatconnects the ground to a geostationary orbit of 36,000
km with a cable and transports materials by ascending
and descending on the cable has been proposed. This
research can be positioned as basic research on theactuator of the climber for the space elevator.
Structures that cannot realistically be equipped withcounterweights will become taller and taller, and spaceelevators will start to become a reality, making it
difficult to install counterweights themselves in somecases.
As shown in Fig. 24)–6), the climber for such anelevator uses rollers or tires driven by a rotating motor
to generate a frictional force between the rollers and thecable. This frictional force causes the elevator to lift andlower. However, the elevator requires a frictional force
that exceeds the weight of the object being lifted. There
are several ways to generate a high frictional force, oneof which is to increase the coefficient of friction bysurface treatment of the cables or tires. Another is toincrease the pressing force. However, these methods areless efficient when foreign objects are entrapped andcause damage to the cable surface. Furthermore,because the friction coefficient of a carbon nanotube,
which is expected to apply to the tether of the spaceelevator, would be very few, a climber with a friction
force would need a larger pressing force.
For vertical devices with long cables, a major
challenge is to reduce maintenance and replacement. To
extend the life of such devices, devices that can beraised and lowered without mechanical contact are

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counterweightFig. 1 Schematic diagram of elevator.

circumferential direction has also been proposed14）.
However, there has been no detailed study of theeffect of the shape of the actuator that raises and lowers
the long cable-like tether on the thrust characteristics.
Therefore, a cylindrical linear induction motor for
vertical movement, in which a cable-shaped conductor
with uniform geometry is moved vertically, wasinvestigated in this study. The optimal tether shape isconsidered to vary depending on the structure of thebuilding, assuming operation in a skyscraper.
In this study, to establish design guidelines for theparameters of the tether for vertical movement, theeffect of varying the diameter of the tether, which
serves as the reaction plate of the cylindrical LIM, on
the thrust characteristics was examined.
2. Fundamental Study on Thrust Characteristics ofLinear Actuators

Fig. 2 Schematic diagram of friction force for verticaltransfer in a conventional climber.
required. Linear synchronous motors and linear
switched-reactance reactance motors are being
investigated to realize such devices7). Linear induction
motors (LIM), which can be driven without contact overa uniformly shaped reaction plate, have been applied ashorizontal propulsion in railroads, maglev, sky rail,
etc.8),9). This study focuses on the characteristics of this
LIM and examines its application to vertical movement,such as in elevators.
A typical LIM installed on a flat reaction plate ispressed against the reaction plate, and the gap between
the LIM and the reaction plate changes. This causes the
LIM to vibrate and to contact the reaction plate.
Furthermore, a flat LIM has difficulty maintaining the
gap against vibrations caused by various disturbances
and torsion.
Furthermore, these LIM has a disadvantage as end
effect and edge effect 10),11). On the other hand, a linear
induction motor with a cylindrical primary has also
been proposed12）.Compared to the flat plate type, thiscylindrical LIM has no coil ends, which increases themain magnetic flux and improves the thrust and power
factor. In addition, the axisymmetric structure cancels
magnetic attractive forces and facilitates support13）.
In addition, a Z-θ drive actuator that can move not
only in the axial direction but also in the

6

2.1 Analysis modelA schematic diagram of the proposed cylindrical LIM
is shown in Fig. 3. This paper did not examine the shapeof the tether for a LIM of a certain shape, but focused on
how the thrust characteristics change when the shape of
the LIM itself is changed. While a hollow tether wasused for the study in the previous paper, the study inthis paper was conducted from a model with a verysmall tether diameter. Since the magnetic field from the
opposing coil may affect the thrust characteristics, ananalytical model with a solid medium tether isconstructed as a basic study. The climber is operated
from an external power supply and has a system suchthat in the event of a loss of power, this mechanicalbrake is activated to stop the climber on the cable. It isnecessary to consider how to feed power to the coils fromthe outside while the climber is elevated, and how
efficient the proposed linear motor is in terms of electric
power. An overview of the LIM analytical model isshown in Table 1. The cylindrical linear induction motor
consists of a core, bobbin, and coil, and it has an outer
diameter of 80 mm, an inner diameter of 32 mm, and aheight of 130 mm. The material used for the core is
S45C, the bobbin is plastic, and the coil is Cu. The coils
are installed at six locations inside the core of theclimber, and each coil is connected to a three-phase AC
power supply, as shown in Fig. 4. The tether material iscopper, and the frequency output from the power supply
is changed to calculate the thrust characteristics acting
on the climber by analysis. In this study, an analyticalmodel of the LIM was created, and a magnetic field
analysis was conducted to understand the lifting andlowering performance and thrust characteristics of the
non-contacting climber. The LIM model is shown in Fig.
5. The radial direction of the LIM is defined as the Xaxis and Y-axis, and the axial direction of the cable isdefined as the Z-axis.

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

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Table 1 Specifications of analysis model.
Core
S45C
Material

LIM

TetherFig. 3 Schematic diagram of proposed climber using
LIM.

Bobbin

Plastic

Coil

Cu

TetherCu

Inner diameter d i

32 mm

Outer diameter d o

80 mm

Height

130 mmMass

4.3 kgDiameter d t

30 mm

Number of coils

6

Air gap

1 mm

Table 2 Analysis conditions.

Parameters

ValuesCurrent Amplitude

7.85 A0-p

Frequency

1-100 Hz

Fig. 4 Circuit diagram of the LIM analysis model.

Cable
U

LIM

U’
V
V’
W
W’

Coil

Z
Y
X

Fig. 5 Analysis model of cylinder-type noncontact
climber.
2.2 Analysis conditions
To evaluate the lifting and lowering performance of an
LIM as a climber, the thrust that can be generated wascompared with the mass of the climber itself. The circuit
diagram of the LIM to be analyzed is shown in Fig. 4.
The winding diameter of the coil in the slot of the model
was set to 1 mm, and the number of turns was set to
182 to achieve a fill factor of 50%. The winding

Fig. 6 Top view of model.
is 0.62 Ω. The model self-weight under each analysis
condition was calculated based on the specifications
shown in Table 1 for the bias of the analytical model.
The analysis conditions are shown in Table 2. 7.85 A
was used for current amplitude of the three-phase AC
power supply to set the coil current density to 10 A/mm²
and the frequency to 1–100 Hz, and an electromagnetic
field transient response analysis was performed. The
analysis time was approximately 0.05 s, with timeincrements of 0.005 s for each calculation step. The
electromagnetic force generated in the Z-axis direction,
which is the axial direction of the tether, was assumed
to be the thrust force.
2.3 Analysis results
The average thrust force at each frequency obtained
from the electromagnetic field analysis are shown in Fig.
7. The figure shows that the vertical thrust was
obtained at all frequencies, that the thrust increased as

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

7

INDEX

according to the tether diameter. Table 3 shows thecharacteristics of each LIM analytical model. An
overview of the analytical model described in Section 2,
which corresponds to Model C, is shown in Fig. 8.

Fig. 7 Analyzed average thrust force (dt = 30 mm).
Table 3 Model specifications.
Model

Material

A

B

Core

S45C

Bobbin

Plastic

Coil

Cu

Tether

LIM

C

D

Cu

Tether diameter d t

10 mm

20 mm

30 mm

40 mm

Inner diameter d i

12 mm

22 mm

32 mm

42 mm

Outer diameter d o

60 mm

70 mm

80 mm

90 mm

Height

130 mm

130 mm

130 mm

130 mmMass

2.9 kg

3.6 kg

4.3 kg

5.0 kgWeight

29 N

36 N

43 N

50 N

Number of coils

6

Air gap

1 mm

Table 4 Analysis conditions.
Model

A

B

C

D

Tether diameter d t [mm]

10

20

30

40

Frequency f [Hz]10-50010-20010-10010-100

Step time [s]0.00010.000250.00050.0005

Number of steps

501

201

101

101

the frequency increased, and the effect of transients also
increased. The same figure shows that the averagevalue of the thrust also increases as the frequency
increases, but the effect of transients also increases.
Furthermore, the average value of the thrust increases
with increasing frequency, but the thrust decreases
after 70 Hz.
3. Effect of Tether Outside Diameter on Thrust
Characteristics of Linear Actuators
3.1 Analytical model
Because of the structure of the elevator underconsideration, it is possible that multiple cargos may beraised and lowered on a single tether. The required
tether strength may vary depending on the height of thestructure. In other words, the tether diameter isconsidered to vary depending on the operating
conditions of the elevator. Therefore, the thrust
characteristics of an LIM moving up and down on atether were studied when the tether diameter waschanged. The tether diameters dt were 10, 20, 30, and
40 mm. Four LIM models, A–D, were designed

8

3.2 Analysis conditions
Four models were analyzed, with the tether diameter
varied in 10-mm increments. Table 4 shows the analysis
conditions for each model. All models were analyzed
with a current of 7.85A and 182 turns per coil. However,preliminary results showed that the frequency at which
the maximum thrust was achieved varied depending on
the tether diameter, so the frequency and number ofsteps were changed for each model as shown in Table 4.
3.3 Analysis results
Vector plots of the magnetic flux density, vector plotsof the current density, and contour plots of the Lorentz
force density for each model are shown in Figs. 9–11.
The magnetic flux and current results indicate that alarge Lorentz force is generated. Furthermore, as thetether diameter increased, the area inside the tether
where the Lorentz force was not generated increased.
The average thrust of each model after 0.02 s isshown in Fig. 12. The horizontal axis of the graph shows
the frequency, and the horizontal axis shows the thrust.
Based on these results, the relationship between tether
diameter and Thrust-to-weight ratio is shown in Fig. 13.
This figure shows that, as the tether diameter increases,
and the thrust-to-weight ratio increases accordingly,
and the frequency at which the highest thrust is
obtained decreases.
4. Conclusion
A vertical transport system using a cylindrical LIM asan actuator was studied. It can move over a uniform
conductor cable without contact. The thrust
characteristics of the system were also studied when thetether diameter was changed. The analysis results show
that, as the tether diameter increases, the thrustgenerated in the climber tends to increase, and thefrequency at which the maximum thrust is obtained
tends to decrease. Also, increasing the diameter tended
to increase the thrust more than the increase in thesurface area of the gap, and the thrust was obtained
more efficiently.
Future studies are planned in which the area of theslot where the coil is wound, the aspect ratio of the coils,
and the distance between the coils are changed to obtaina higher thrust force. It is also necessary to study thepossibility of obtaining higher thrust by changing thematerials of the core and tether to materials with ahigh-saturation magnetic flux density or different
electrical resistivity.
In addition, it is necessary to conduct a detailed
study from the viewpoint of method of power supply andenergy efficiency.

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

INDEX

 
 

  





  

(a) Model A

 (b) Model B

 (c) Model C  (d) Model D(dt = 10 mm)(dt = 20 mm)(dt = 30 mm)(dt = 40 mm)

Fig. 8 Analysis model in each cable thickness.


 



 

(a) Model A  (b) Model B   (c) Model C(dt = 10 mm)(dt = 20 mm)(dt = 30 mm)

(d) Model D  (dt = 40 mm)

Fig. 9 Analyzed current density in each tether diameter.

 

  

 



(a) Model A   (b) Model B   (c) Model C(dt = 10 mm)(dt = 20 mm)(dt = 30 mm)

 
 

(d) Model D(dt = 40 mm)

Fig. 10 Analysed magnetic flux density.

 (a) Model A(dt = 10 mm)

 (b) Model B(dt = 20 mm)

 (c) Model C(dt = 30 mm)

 (d) Model D(dt = 40 mm)

Fig. 11 Contour plot of analysed Lorentz force.

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

9

INDEX

References
1)
2)
3)
4)
5)

6)
7)

Fig. 12 Average thrust.

8)
9)
10)
11)
12)
13)
14)

T. Fujitani and H. Kimura: Trans. JSME, 89, 18-00048(2018).
S. Jia: Powder Technol., 320, 462 (2017).
T. Matsuno: J. Soc. Control Instrum. Eng., 21, 717 (1982).
Y. Ishikawa, K. Otsuka, Y. Yamagiwa, and H. Doi: Acta
Astronaut., 145, 165 (2018).
K. Ishizaki, K. Soonsu, F. Inoue, E. Omoto, and Y.
Ishikawa: JSMERMD, The Proceedings JSME annual
Conference on Robotics and Mechatronics, Yokohama,Japan, 2016, 2A2-17b2 (JSME, Tokyo, 2017).
Y. Aoki: Aeronautical and Space Sci. Jpn., 62, 325 (2014).
R. Coa, E. Su, and M. Lu: IEEE Trans. Appl. Supercond.,
30, 3601205 (2020).
S. Watanabe, T. Koseki, Y. Noda, and M. Miyatake: IEEJ
Trans. Ind. Appl., 137, 44 (2017).
T. Mizuno: IEEJ Trans. Ind. Appl., 120, 479 (2000).
S. Yamamura, H. Ito and Y. Ishikawa: IEEJ J. Inst. Electr.
Eng. Jpn., 91, 309 (1971).
T. Onuki, H. Numata, and T. Yamamura: IEEJ J. Ind.
Appl., 110, 807 (1990).
T. Morizane, T. Koseki, E. Masuda: IEEJ J. Ind. Appl.,
112, 1149 (1992).
A. Hirano, N. Maki, and Y. Tanaka: IEEJ J. Ind. Appl.,
122, 871 (2002).
E. Amiri: IECON 2014 - 40th Annual Conference of theIEEE Industrial Electronics Society, Dallas, TX, USA,
2014, pp. 431-436 (IEEE, Piscataway, New Jersey, 2015).

Received Nov. 23, 2022; Revised Aug. 04, 2023; Accepted Nov.
06, 2023

Fig. 13 Thrust-to-weight ratio of the maximum thrustof each model to its own weight.

10

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

INDEX

Copyright ©2024 by the Magnetics Society of Japan.
This article is licensed under the Creative Commons Attribution International License (CC BY 4.0)
http://creativecommons.org/licenses/by/4.0/

J. Magn. Soc. Jpn., 48, 11-16 (2024)<Paper>Linear Actuators for High-Speed Reciprocating Motion with Dual Halbach
Arrays(Fundamental Consideration of the Effect of Magnet Arrangement on
Thrust Characteristics)
M. Tanaka, K. Kimura, J. Kuroda, Y. Majima, D. Uchino, K. Ogawa,T. Kato*, K. Ikeda**, A. Endo***, T. Narita and H. Kato
Tokai Univ., 4-1-1 Kitakaname, Hiratsuka-shi, Kanagawa, 259-1292, Japan, Hachioji-shi, Tokyo 192-0982, Japan
7-15-4-1 Maeda, Teine-ku, Sapporo-shi, Hokkaido, 006-8585, Japan*** Fukuoka Institute of Technology, 3-30-1 Wajiro-higashi, Higashi-ku, Fukioka-shi, Fukuoka, 811-0295, Japan* Tokyo University of Technology, 1404-1 Katakuramachi** Hokkaido University of Science,

To achieve higher performance in spark-ignition engines, the intake and exhaust valves must be controlled in a
stepless manner according to the engine speed. Our research group has been continuously studying theelectromagnetic drive valve system using a linear actuator. We designed a linear actuator with a dual Halbach arrayof permanent magnets in the stator to increase the thrust of the linear actuator. The thrust characteristics of the
proposed model were investigated using electromagnetic field analysis based on the finite element method. The
analysis confirmed that the Halbach array concentrates the magnetic flux on the coil side, which generates Lorentz
force. From the analysis results, the proposed liner actuator can generate thrust for driving the engine valve, andoptimal permanent magnets in stator can be obtained to generate higher thrust.
Keywords: liner motor, valve drive system, reciprocating engine, dual Halbach array
1. Introduction
Reciprocating motion is used at various points in theoperationofmechanicalproducts.
Generally,
reciprocating motion is converted from the rotationalmotion of a motor or engine by a mechanism, such as a
cam or a piston/crank mechanism. However, direct drives
with linear actuators must utilize high-precision andhigh-speed reciprocating motion. 1)–5)
For example, the cylinder head of an internal
combustion engine (ICE) has intake and exhaust engine
valves. These valves are opened and closed using linear
motion converted by the camshafts that are rotated by
the engine. The cam geometry determines thecommanded motion of the valves, i.e., when and how wide
to open and when to close them. The cam shape is aparameter that significantly affects the engine's
combustion and thus directly affects the performance of
the ICE, such as power output and fuel consumption. In
a typical engine, the engine is operated with a single cam
profile, so there are areas where the cam profiles do notmatch depending on the rpm of operation, resulting inpoor engine performance. A mechanism that switches
cam profiles according to the operating speed has been
proposed to solve this problem. 6) 7) This system has been
implemented in sports cars to improve driving
performance and in some production cars to improve fueleconomy. However, these systems rely on multiple cam                     Corresponding author: T. Narita(e-mail: narita@tsc.u-tokai.ac.jp).

shapes. The engine valves can always be operated with
optimal parameters according to the continuously
changing speed of the ICE. In addition, these systems'
cam shapes, or engine valve operation patterns, aredesigned with restrictions to avoid mechanical problems
such as surging and valve jumps.
Therefore, a method has been proposed to operate theengine valve with a linear actuator without involving the
cam mechanism. This system can operate the engine
valve optimally according to the engine speed without
mechanical limitations. In addition, the output can bevaried in response to accelerator pedal operation without
needing a throttle valve. This improves the combustionof the ICE engine and can be expected to increase power
output and fuel consumption. Several studies haveproposed linear actuators to drive engine valves, focusing
on this advantage. 8)–12) However, in these studies,
although sufficient yokes were installed on the mover, it
is difficult to operate the actuator precisely at the speed
required for ICE engines because of its high inductanceand heavy mover. Therefore, there is a need to develop a
compact linear actuator that can provide high thrust.
The gap magnetic flux must be increased to realize ahigh thrust actuator in a limited space, such as inside acylinder head. Therefore, this research focused on acylindrical moving coil linear actuator with a lightweightmover and small ink dance. In addition, a Halbach array
was used for the stator's permanent magnet array toincrease the gap magnetic flux. A dual Halbach array
was used to further increase the flux density of thepermanent magnets, with Halbach magnet rows on the
inner and outer sides of the moving coil. However, there

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

11

INDEX

Fig. 3 Input three-phase alternating current.
Fig. 1 Schematic of dual Halbach array.

⦿ ⊗⊗⦿ ⊗⊗
Fig. 4 Current directions in each coil.
and outer magnets are different in the cylindrical type,
and the volume of the magnets is considered to bedifferent, resulting in unequal magnetic flux. Therefore,
this paper presents an electromagnetic field analysis of
the changes in thrust characteristics of a cylindricallinear actuator by changing the size of the outerpermanent magnet array.
2. Analysis model of the actuator using dual Halbach
array(a) Bird's view(b) Dual Halbach model (c) Magnetization direction
Fig. 2 Analysis model.
are few examples of this method being used forreciprocating actuators, and no detailed studies havebeen conducted. In addition, the diameters of the inner12

2.1 Outline of dual Halbach array
The linear actuator uses Fleming's left-hand rule togenerate Lorentz force by orthogonalizing the magnetic
flux in the coil to obtain thrust. Therefore, the thrustforce can be increased by increasing the orthogonal
magnetic flux through the coil by changing the directionof magnetization for each permanent magnet.
The dual Halbach array, which concentrates themagnetic flux in the coil by changing the direction ofmagnetization for each permanent magnet, comprises
two types of magnets, as shown in Fig. 1. They include aradially oriented magnet that makes the magnetic flux
orthogonal to the coil and an axially oriented magnet thatguides the magnetic flux from one radially oriented
magnet to another to prevent the flux from going outside
the motor. The radially oriented magnets are used in the
motor. In previous studies, the radial and axial magnets
had the same thickness. Therefore, we considered thatthe thrust could be increased by increasing the thicknessof the radial magnet, which directs the magnetic flux
orthogonally to the coil.
2.2 Linear actuator using Halbach array
To investigate changes in the thrust characteristics bychanging the aspect ratio of the permanent magnet, ananalytical model was created using the 3D CAD software
SOLIDWORKS, as shown in Fig. 2. Fig. 2 (a) shows abirds view of our proposed linear actuator. Fig. 2 (b)Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

INDEX

Table 1 Specifications of permanent magnets
Model

1

2

3

4

5

Thickness of permanent magnet magnetized in radial direction hr [mm]

6.0

7.0

8.0

9.0

9.6

Thickness of permanent magnet magnetized in axial direction ha [mm]

6.0

5.0

4.0

3.0

2.4

(a) Model 1(hr = 6.0 mm,
ha = 6.0 mm)(b) Model 2(c) Model 3(d) Model 4(e) Model 5(hr = 7.0 mm,(hr = 8.0 mm,(hr = 9.0 mm,(hr = 9.6 mm,
ha = 5.0 mm)
ha = 4.0 mm)
ha = 3.0 mm)
ha = 2.4 mm)
Fig. 5 Analysis model of liner actuator applying dual Halbach array.

Table 2 Specifications of analysis model.
Parts
Case

Outer magnet
Inner Magnet

Coil bobbin

Coil

Table 3 Materials in each part of analysis model.
Parts

Materials

49.0 mm

Coil bobbin

YEP-B (Permalloy)

Height70.0 mmValue

YEP-B (Permalloy)

Thickness

2.0 mmInside diameter

34.8 mm

Outer diameter

40.8 mm

Inside diameter

11.2 mm

Outer diameter

22.5 mm

was applied to each coil. Coils 1 and 4 were set to U phase,
Coils 2 and 5 to V phase, and Coils 3 and 6 to W phase as
shown in Fig. 4. The current directions in each coil wereshown in Fig. 4. A case is placed outside the valve, bobbin,and outer permanent magnet. In this analytical model,
the coils, valves, and bobbins act as moving parts.

Inside diameter

6.4 mm

3. Effect of magnet shape on thrust characteristics

Outer diameter10.4 mm

Height72.0 mmThickness

2.0 mmInside diameter

26.3 mm

Outer diameter

32.0 mm

Height

7.5 mm

3.1 Analysis conditions
The area of radial permanent magnets located outside
the coil was increased in steps, and the area of axial
permanent magnets was decreased accordingly toincrease the radial magnets' thickness without changing
the linear actuator's size or the number of permanent
magnets. The permanent magnets' thicknesses werevaried in steps up to 9.6 mm in the radial direction and2.4 mm in the axial direction, with the basic model
having a radial thickness of hr and an axial thickness of
ha of 6.0 mm. The dimensions of each model are shown in
Table 1. The model specifications are shown in Table 2;materials were set as shown in Table 3, and Fig. 5 shows
the analytical model. The number of turns elements inthis analysis model was almost 700,000. The number ofturns of each coil was set to 33 turns. A three-phase

Dimension

ValuesOuter diametershows a cross-sectional view of the analytical model. Six
cylindrical coils and 12 cylindrical permanent magnets
were placed outside and inside the coils, respectively. The
permanent magnets and the outer case constitute thestator. Fig. 2 (c) shows the magnetization direction of thepermanent magnet. The coils are numbered from 1 to 6
in order from the top. Three-phase alternating current

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

13

INDEX



61

58

59

56

T h r us t [ N ]57
Average thrust [N]

55
53

51
49
47

0.0

0.4

0.8
1.2
1.6
Lift amount [mm]

2.0

2.4

(a) Model (hr = 6.0 mm, ha = 6.0 mm)

54
52
50
48
46

61
59

T h r us t [ N ]

53

51
49
0.0

0.4

0 .8
1.2
1.6
Lift amount [mm]

2.0

2.4

Thrust [N](b) Model 2 (hr = 7.0 mm, ha = 5.0 mm)
61
59
57
55
53
51
49
47

Model 3

8.0
7.0
6.0
5.0

4.0
3.0
2.0

1.0
0.0

Model 1Model 2Model 3Model 4Model 5(b) Fluctuation of thrustFig. 7 Analyzed results in each model.

0 .0

0.4

0.8
1.2
1.6
Lift amount [mm]

2.0

2.4

61
59

Thrust [N]Model 5Model(c) Model 3 (hr = 8.0 mm, ha = 4.0 mm)

57
55
53
51
49
47

0.0

0.4

0.8
1.2
1.6
Lift amount [mm]

2.0

2.4

(d) Model 4 (hr = 9.0 mm, ha = 3.0 mm)

Thrust [N]Fluctuation of thrust [N]

55

61
59
57
55
53
51
49
47

0.0

0.4

0.8
1.2
1.6
2.0
Lift amount [mm]2.4(e) Model 5 (hr = 9.6 mm, ha = 2.4 mm)
Fig. 6 Relationship between lift amount and thrust ineach analysis model.

14

Model 2Model 4(a) Average thrust

57

47

Model 1

alternating current was applied to each coil. Coils 1 and4 were set to U phase, Coils 2 and 5 to V phase, and Coils
3 and 6 to W phase. The amplitude of the current was set
to 20 A. In this actuator, 360° of electric angle means 2.4
mm displacement in the axial direction.
Electromagnetic analysis software (JMAG) was used
for analysis. The number of steps in the analysis was 25,
and the displacement of each step was 0.1 mm to obtainthe thrust generated at each step.
3.2 Analysis Results
Fig. 6 shows the relationship between thrust and lift
amount for each model, with the vertical axis showing
the thrust and the horizontal axis showing the lift
amount. From the results, the thrust in each modelshows the same tendency, and Model 2 with hr = 7.0 mm
and ha =5.0 mm generates a higher thrust. The
summarized analysis result is shown in Fig. 7. Fig. 7 (a)shows the average thrust. This figure shows that theincrease in the radial thickness hr more than 7.0 mm in
the model, which was expected to increase flux density inthe air gap, contrarily decreased the average thrust. Fig.
7 (b) shows the fluctuation of thrust obtained by different
thrust forces between maximum and minimum values. In
the view point of thrust fluctuation, there is no differencein each model. Fig. 7 shows a vector plot of the magnetic
flux density at a lift of 1.9 mm, and Fig. 8 shows Lorentz
force density. The vector plots of the flux density andLorentz density for each model at 20 steps are shown in

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

INDEX

Flux density [T]2 .0(a) Model 1(hr = 6.0 mm,
ha = 6.0 mm)(b) Model 2(c) Model 3(d) Model 4(hr = 7.0 mm,(hr = 8.0 mm,(hr = 9.0 mm,
ha = 5.0 mm)
ha = 4.0 mm)
ha = 3.0 mm)
Fig. 8 Analyzed flux density in each analysis model(e) Model 5(hr = 9.6 mm,
ha = 2.4 mm)

Flux density [T]2 .0(a)Model 1(hr = 6.0 mm,
ha = 6.0 mm)(b)Model 2(c)Model 3(d)Model 4(hr = 7.0 mm,(hr = 8.0 mm,(hr = 9.0 mm,
ha = 5.0 mm)
ha = 4.0 mm)
ha = 3.0 mm)
Fig. 9 Analyzed Lorentz force density in each analysis modelFig. 8 and Fig. 9, respectively. The figures show that
excessively increasing the radial permanent magnet area
causes the magnetic flux to exit the case without passing
through the axial permanent magnets. Therefore, in an
actuator with a Halbach array, increasing the radial
permanent magnet area, which makes the magnetic flux
orthogonal to the coil, improves the thrust withoutchanging the size of the actuator. However, the thrust
was lower in Models 4 and 5, in which the radial
permanent magnet the area was greater than in Model 1
with 6.0 mm axial and radial thickness.
4. Conclusion
We investigated the effect of a dual Halbach array on
the thrust force of a linear actuator by altering the areaof permanent magnets arranged in the actuator. In the

1 .5

7 .5 ×

1 .0

5 .0 ×0 .5

2 .5 ×0 .0

0 .0

Lorentz density[ N/m 3 ]
1.0×1
1 07

1 .5

7.5×1
1 06

1 .0

5.0×1
106

0 .5

2.5×1
1 06

0 .0

0 .0

(e)Model 5
hr = 9.6 mm,
ha = 2.4 mm)

Halbach array, the radial permanent magnet area thatorthogonalizes the magnetic flux to the coil was increased,
and the axial permanent magnet area that guides themagnetic flux out of the actuator was decreased.
Consequently, a higher thrust force was generated thanin the model with the same radial and axial permanent
magnet regions. However, if the radial permanent
magnet area is excessively increased, the axial
permanent magnet cannot induce the magnetic flux, andthe flux escapes to the outside of the case. In the future,various studies will be conducted to develop an actuatorwith an even higher thrust.
References
1) I. Martins, J. Esteves, G.D. Marques, and F.P. Da Silva: IEEE
Trans. Veh. Technol. 55, 86 (2006).

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

Lorentz densit[ N/m 3 ]1 .0 ×15

INDEX

2) R. Kornbluh, R. Pelrine, J. Eckerle and J. Joseph: 1998 IEEE
International Conference on Robotics and Automation, R.
Kornbluh, Leuven, Belgium, 1998, 3, pp.2147-2154 (IEEE,
1998)
3) F. Carpi, C. Salaris, and D. De Rossi: Smart Mater. Struct. 16,
S300 (2007).
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Mech. 26, 28 (2018).
5) Y. Keijiro, I. Masayuki, and M. Takasi: Research Reports ofKanagawa Inst. Technol. B, Sci. and Technol., 21, 33 (1997).
6) K. Hatano, K. Iida, H. Higashi, and S. Murata: SAE Trans.
1291 (1993).
7) T.H. Ma: SAE trans., 665 (1988).
8) Y. Takao, M. Yoshifumi, and U. Masaki: DENKI-SEIKO, 74,
147 (2003).

169) S. Akimoto, and A. Masaru: Conference of Tohoku Branch,
The Japan Society of Mechanical Engineers, Japan, 2009, 44,
172 (The Japan Society of Mechanical Engineers, 2009.)
10) O. Masaru, S. Akitomo, and K. Takahide: Conference of
Tohoku Branch, The Japan Society of Mechanical Engineers,
Japan, 2009, 44, 174 (The Japan Society of MechanicalEngineers, 2009.)
11) M. A. Theobald, L. Bruno, and H. Rassem: SAE Trans. 1323
(1994).
12) S. Braune, L. Steven, and M. Paolo: 2006 IEEE Conference
on Computer Aided Control System Design, 2006 IEEE
International Conference on Control Applications, 2006 IEEE
International Symposium on Intelligent Control., S. Braune,
Munich, Germany, 2006, pp.1657-1662 (IEEE, 2006)
Received Dec. 06, 2022; Revised Sep. 27, 2023; Accepted Nov. 06,
2023

Journal of the Magnetics Society of Japan Vol.48, No.1, 2024

INDEX

Editorial Committee Members・Paper Committee Members
S. Yabukami and T. Taniyama (Chairperson), N. H. Pham, D. Oyama and M. Ohtake (Secretary)
H. Aoki
M. GotoT. GotoK. Hioki
S. Inui
M. Iwai
Y. Kamihara
H. Kikuchi
T. Kojima
H. KuraK. Masuda
Y. Nakamura
K. Nishijima
T. Nozaki
T. SatoT. Suetsuna
K. Suzuki
Y. Takamura
K. ThamT. Tanaka
N. Wakiya
S. Yakata
A. YaoS. Yamada
M. Yoshida
N. Adachi
S. HakuT. Maki
M. Sekino
T. Takura

K. Bessho
S. Honda
S. Muroga
Y. Shiota
S. Tamaru

M. DoiS. Isogami
M. NaoeS. Sugahara
S. Yoshimura

T. DoiT. Kawaguchi
T. Narita
I. Tagawa

T. Hasegawa
N. Kikuchi
Y. SatoK. Tajima

K. ItoA. Kuwahata
E. Shikoh
M. Toko
R. Hashimoto
K. Kobayashi
S. Seino
M. Takezawa

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Journal of the Magnetics Society of Japan
Vol. 48 No. 1 (通巻第 331号) 2024年1月1日発行

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