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Actuators
Volume 6
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Actuators, Volume 6, Issue 2 (June 2017) – 9 articles

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2995 KiB  
Article
Measuring the Temperature Increase of an Ultrasonic Motor in a 3-Tesla Magnetic Resonance Imaging System
by Peyman Shokrollahi, James M. Drake and Andrew A. Goldenberg
Actuators 2017, 6(2), 20; https://doi.org/10.3390/act6020020 - 6 Jun 2017
Cited by 3 | Viewed by 7357
Abstract
This paper aims to evaluate the temperature increase caused by a 3.0-T magnetic resonance imaging (MRI) system on an ultrasonic motor (USM) used to actuate surgical robots in the MRI environment. Four fiber-optic temperature sensors were attached to the USM. Temperature was monitored [...] Read more.
This paper aims to evaluate the temperature increase caused by a 3.0-T magnetic resonance imaging (MRI) system on an ultrasonic motor (USM) used to actuate surgical robots in the MRI environment. Four fiber-optic temperature sensors were attached to the USM. Temperature was monitored outside the five-Gauss boundary and then inside the bore for 20 min while the USM was powered on. The USM temperature was tested for two states of the scanner, “off” and “on”, by employing common clinical imaging sequences and echo planar imaging sequences. The USM showed a slight temperature increase while operating in the static field of the MRI. A considerable temperature increase (~10 °C) was observed when the scanner was on. The temperature increased to 60 °C, which is beyond the acceptable safe temperature and can result in thermal burns. Most of the temperature increase (80%) was due to effects of the static field on the motion of the rotating parts of the motor, while the remainder (20%) derived from heat deposited in the conductive components of the USM due to radiofrequency pulses and gradient field changes. To solve the temperature increase, the metal components of the USM’s case can be replaced by silicon carbide. Full article
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Figure 1
<p>Structure of the ultrasonic motor described in [<a href="#B18-actuators-06-00020" class="html-bibr">18</a>].</p> Full article ">Figure 2
<p>PUMR 40E, Piezoelectric Technology Co., Ltd., Seoul, Korea.</p> Full article ">Figure 3
<p>Sensor locations: (<b>a</b>) Motor front view; (<b>b</b>) Motor rear view.</p> Full article ">Figure 4
<p>The motor orientation with respect to the bore.</p> Full article ">Figure 5
<p>Motor temperature increase in static field, when the scanner was off.</p> Full article ">Figure 6
<p>(<b>a</b>) Temperature of the operating ultrasonic motor (USM) outside the magnetic resonance imaging (MRI) system; (<b>b</b>) Temperature of the operating USM inside the bore when the scanner was on.</p> Full article ">Figure 7
<p>Motor temperature increase while applying Echo Planar Imaging (EPI) sequences.</p> Full article ">Figure 8
<p>(<b>a</b>) The T2-weighted turbo spin echo (T2W) transverse image of the phantom in the presence of silicon carbide sheet; (<b>b</b>) the same image in the absence of the sheet; (<b>c</b>) the subtraction of these two images (<b>a</b>,<b>b</b>).</p> Full article ">
3767 KiB  
Article
Design and Characterization of In-Plane Piezoelectric Microactuators
by Javier Toledo, Victor Ruiz-Díez, Alex Diaz-Molina, David Ruiz, Alberto Donoso, José Carlos Bellido, Elisabeth Wistrela, Martin Kucera, Ulrich Schmid, Jorge Hernando-García and José Luis Sánchez-Rojas
Actuators 2017, 6(2), 19; https://doi.org/10.3390/act6020019 - 3 Jun 2017
Cited by 15 | Viewed by 9435
Abstract
In this paper, two different piezoelectric microactuator designs are studied. The corresponding devices were designed for optimal in-plane displacements and different high flexibilities, proven by electrical and optical characterization. Both actuators presented two dominant vibrational modes in the frequency range below 1 MHz: [...] Read more.
In this paper, two different piezoelectric microactuator designs are studied. The corresponding devices were designed for optimal in-plane displacements and different high flexibilities, proven by electrical and optical characterization. Both actuators presented two dominant vibrational modes in the frequency range below 1 MHz: an out-of-plane bending and an in-plane extensional mode. Nevertheless, the latter mode is the only one that allows the use of the device as a modal in-plane actuator. Finite Element Method (FEM) simulations confirmed that the displacement per applied voltage was superior for the low-stiffness actuator, which was also verified through optical measurements in a quasi-static analysis, obtaining a displacement per volt of 0.22 and 0.13 nm/V for the low-stiffness and high-stiffness actuator, respectively. In addition, electrical measurements were performed using an impedance analyzer which, in combination with the optical characterization in resonance, allowed the determination of the electromechanical and stiffness coefficients. The low-stiffness actuator exhibited a stiffness coefficient of 5 × 104 N/m, thus being more suitable as a modal actuator than the high-stiffness actuator with a stiffness of 2.5 × 105 N/m. Full article
(This article belongs to the Special Issue MEMS-based Actuators)
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Figure 1

Figure 1
<p>Design domain and boundary conditions: (<b>a</b>) top view; (<b>b</b>) side view. <math display="inline"> <semantics> <mrow> <msub> <mi>χ</mi> <mi>s</mi> </msub> </mrow> </semantics> </math> and <math display="inline"> <semantics> <mrow> <msub> <mi>χ</mi> <mi>p</mi> </msub> </mrow> </semantics> </math> are the material and polarization profile design variables, respectively. The force <math display="inline"> <semantics> <mrow> <msub> <mi>f</mi> <mrow> <mi>i</mi> <mi>n</mi> </mrow> </msub> </mrow> </semantics> </math> and displacement <math display="inline"> <semantics> <mrow> <msub> <mi>u</mi> <mrow> <mi>i</mi> <mi>n</mi> </mrow> </msub> </mrow> </semantics> </math> at the input port are also depicted.</p> Full article ">Figure 2
<p>Low-stiffness actuator: (<b>a</b>) structure variable; (<b>b</b>) polarization profile for <math display="inline"> <semantics> <mrow> <msub> <mi>f</mi> <mrow> <mi>i</mi> <mi>n</mi> </mrow> </msub> <mo>=</mo> <mn>1</mn> <mo> </mo> <mi mathvariant="normal">N</mi> </mrow> </semantics> </math> and <math display="inline"> <semantics> <mrow> <msubsup> <mi>u</mi> <mrow> <mi>i</mi> <mi>n</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msubsup> <mo>=</mo> <mn>20</mn> <mo> </mo> <mi mathvariant="sans-serif">μ</mi> <mi mathvariant="normal">m</mi> </mrow> </semantics> </math></p> Full article ">Figure 3
<p>High-stiffness actuator: (<b>a</b>) structure variable; (<b>b</b>) polarization profile for <math display="inline"> <semantics> <mrow> <msub> <mi>f</mi> <mrow> <mi>i</mi> <mi>n</mi> </mrow> </msub> <mo>=</mo> <mn>10</mn> <mo> </mo> <mi mathvariant="normal">N</mi> </mrow> </semantics> </math> and <math display="inline"> <semantics> <mrow> <msubsup> <mi>u</mi> <mrow> <mi>i</mi> <mi>n</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msubsup> </mrow> </semantics> </math> = 50 µm.</p> Full article ">Figure 4
<p>Micrograph of the MEMS actuators: (<b>a</b>) low-stiffness actuator; (<b>b</b>) high-stiffness actuator.</p> Full article ">Figure 5
<p>Average out-of-plane displacement (<span class="html-italic">z</span>-axis) from the optical characterization of the low-stiffness actuator. The measured modal shape and estimated quality factor are shown next to each peak.</p> Full article ">Figure 6
<p>Average out-of-plane displacement (<span class="html-italic">z</span>-axis) from the optical characterization of the high-stiffness actuator. The measured modal shape and estimated quality factor are shown next to each peak.</p> Full article ">Figure 7
<p>Average displacement in the <span class="html-italic">y</span>-axis at 200 Hz as a function of the applied voltage and its linear fit. The error bar represents the measurement resolution (1 nm).</p> Full article ">Figure 8
<p>Average displacement at the resonant frequency as a function of the applied voltage and its linear fit.</p> Full article ">
13560 KiB  
Article
Static and Dynamic Studies of Electro-Active Polymer Actuators and Integration in a Demonstrator
by Pauline Poncet, Fabrice Casset, Antoine Latour, Fabrice Domingues Dos Santos, Sébastien Pawlak, Romain Gwoziecki, Arnaud Devos, Patrick Emery and Stéphane Fanget
Actuators 2017, 6(2), 18; https://doi.org/10.3390/act6020018 - 4 May 2017
Cited by 21 | Viewed by 9124
Abstract
Nowadays, the haptic effect is used and developed for many applications—particularly in the automotive industry, where the mechanical feedback induced by a haptic system enables the user to receive information while their attention is kept on the road and on driving. This article [...] Read more.
Nowadays, the haptic effect is used and developed for many applications—particularly in the automotive industry, where the mechanical feedback induced by a haptic system enables the user to receive information while their attention is kept on the road and on driving. This article presents the development of a vibrotactile button based on printed piezoelectric polymer actuation. Firstly, the characterization of the electro-active polymer used as the actuator and the development of a model able to predict the electromechanical behavior of this device are summarized. Then, the design of circular membranes and their dynamic characterization are presented. Finally, this work is concluded with the construction of a fully functional demonstrator, integrating haptic buttons leading to a clear haptic sensation for the user. Full article
(This article belongs to the Special Issue MEMS-based Actuators)
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Figure 1
<p>Schematic cross section of the piezoelectric EAP stack that actuates a movable mechanical layer (beam or circular membrane) using the unimorph effect. The location of the cross section is illustrated on the photograph of the polymer actuators.</p> Full article ">Figure 2
<p>(<b>a</b>) Schematic cross section of the sample studied using the PU technique. (<b>b</b>) Measured transient optical reflectivity signal allowing for the detection of acoustic propagation (oscillation) and reflection at the free interface (reflectivity step near 105 ps).</p> Full article ">Figure 3
<p>Concordance between experimental measurements and models responses in terms of maximal displacement as a function of applied voltage, on a cantilever actuator.</p> Full article ">Figure 4
<p>Optimization of the R<sub>EAP</sub>/R<sub>sub</sub> ratio to obtain maximal displacement for circular actuator. Sketch of the structural membrane and piezoelectric actuator, simulated in the developed model.</p> Full article ">Figure 5
<p>Frequency sweep from 0 to 20 kHz on a 12 mm membrane at 6 V<sub>RMS</sub>. Three main resonance modes are observed, respectively, at 3.45 kHz, 6.57 kHz, and 12.3 kHz. Mode shape is illustrated on each peak.</p> Full article ">Figure 6
<p>Image of the first resonant mode at 6 V<sub>RMS</sub> obtained by measurement with the laser Doppler vibrometer, at the center of a 12 mm membrane, and comparison with the expected theoretical mode shape, obtained by FEM modeling.</p> Full article ">Figure 7
<p>Resonance frequency of the first mode as a function of the membrane diameter. Comparison between the FEM model, the analytical model, and experimental measurements.</p> Full article ">Figure 8
<p>Correspondence between experimental measurements and the FEM model value of a 12 mm membrane maximal displacement as a function of the applied voltage for the first resonance mode.</p> Full article ">Figure 9
<p>Demonstrator integrating vibrotactile buttons actuated by electro-active polymers.</p> Full article ">
2606 KiB  
Article
Environmental Effects on the Polypyrrole Tri-layer Actuator
by Nirul Masurkar, Kawsar Jamil and Leela Mohana Reddy Arava
Actuators 2017, 6(2), 17; https://doi.org/10.3390/act6020017 - 26 Apr 2017
Cited by 7 | Viewed by 10010
Abstract
Electroactive polymer actuators such as polypyrrole (PPy) are exciting candidates to drive autonomous devices that require low weight and low power. A simple PPy tri-layer bending type cantilever which operates in the air has been demonstrated previously, but the environmental effect on this [...] Read more.
Electroactive polymer actuators such as polypyrrole (PPy) are exciting candidates to drive autonomous devices that require low weight and low power. A simple PPy tri-layer bending type cantilever which operates in the air has been demonstrated previously, but the environmental effect on this actuator is still unknown. The major obstacle in the development of the PPy tri-layer actuator is to create proper packaging that reduces oxidation of the electrolyte and maintains constant displacement. Here, we report the variation in the displacement as well as the charge transfer at the different environmental condition. PPy trilayer actuators were fabricated by depositing polypyrrole on gold-coated porous poly(vinylidene fluoride) (PVDF) using the electro-synthesis method. It has been demonstrated that the charge transfer of tri-layer actuators is more in an inert environment than in open air. In addition, tri-layer actuators show constant deflection and enhancement of life due to the negligible oxidation rate of the electrolyte in an inert environment. Full article
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Figure 1
<p>(<b>a</b>) Schematic diagram of the tri-layer actuator during oxidation and reduction; (<b>b</b>) Variation in molecular chain of the polypyrrole (PPy) during applied potential difference.</p> Full article ">Figure 2
<p>Illustration of the process steps for the tri-layer actuator; fabrication steps and images after polymerization. (<b>a</b>) (1) A 5 cm × 5 cm piece of commercial PVDF membrane is cut from the roll; (2) Sputter deposition of 5 nm Au on both sides of the PVDF; (3) Electrochemical synthesis of PPy on gold electrode. (<b>b</b>) The final actuator with the dimensions of 5 cm × 5 cm, optical image of cross section and cut out from the membrane.</p> Full article ">Figure 3
<p>(<b>a</b>) X-ray Diffraction (XRD) spectra of PPy doped with LiTFSi at −16 °C; (<b>b</b>) SEM images of electrodeposited polypyrrole (PPy) from pyrrole monomers; (<b>c</b>) Raman spectra of PPy on Gold and poly(vinylidene fluoride) (PVDF).</p> Full article ">Figure 4
<p>Images of different frame grabs showing actuation of a PPy tri-layer actuator in open air and the glove box. (<b>a</b>–<b>c</b>) show the actuation at ±1.5 V in open air, while (<b>d</b>–<b>f</b>) show the actuation inside the glove box, i.e., inert environment.</p> Full article ">Figure 5
<p>(<b>a</b>) Oxidation and reduction currents of the PPy tri-layer actuator in the glove box and open air at ±1.5 V with the same dimensions; (<b>b</b>) Displacement of the polypyrrole actuator contacted by the Kelvin clip in the glove box and open air; (<b>c</b>) Step input voltage and output current in the inert and open environments.</p> Full article ">Figure 6
<p>(<b>a</b>) Input voltage of ±1.5 V square wave; (<b>b</b>) Charge transfer in as inert environment and open air; (<b>c</b>) Displacement vs. charge plot with respective time of the polypyrrole actuator contacted by Kelvin clip in the glove box and in open air.</p> Full article ">
6155 KiB  
Article
Design and Evaluation of a Semi-Active Magneto-rheological Mount for a Wheel Loader Cabin
by Soon-Yong Yang, Chulhee Han, Sang-Un Shin and Seung-Bok Choi
Actuators 2017, 6(2), 16; https://doi.org/10.3390/act6020016 - 20 Apr 2017
Cited by 14 | Viewed by 8780
Abstract
In this study, a semi-active magneto-rheological (MR) mount is designed and manufactured to minimize unwanted vibrations for the cabin of heavy vehicles. Normally, working conditions in heavy vehicles are extremely rugged. Usually, the heavy vehicles use passive rubber mounts for the reduction of [...] Read more.
In this study, a semi-active magneto-rheological (MR) mount is designed and manufactured to minimize unwanted vibrations for the cabin of heavy vehicles. Normally, working conditions in heavy vehicles are extremely rugged. Usually, the heavy vehicles use passive rubber mounts for the reduction of vibrations from road. However, the passive mount has definite performance limitations because the passive mount has a fixed resonance frequency when the design is finished. An MR application is one of the solutions because the viscosity of MR fluid can be controlled. As a first step, an experimental apparatus was established for performance evaluation of the mounts. The apparatus has hydraulic excitatory, force, and displacement sensors. Performance of two different passive mounts used in industrial fields were evaluated. The passive mount data of force-displacement, force-velocity, and displacement transmissibility were collected and tested. After that, an MR mount was designed and manufactured that provides better performance using the passive mount data. The MR mount uses two different flow paths, annular duct and radial channels, for generating the required damping force. The field-dependent damping forces were then evaluated with respect to the moving stroke and input current. In this work, in order to control the damping force, an on-off controller associated with the fast Fourier transform (FFT) was used. The control results of the MR mount were compared with the results of passive rubber mounts. It was shown that the semi-active MR mount can attenuate vibrations more effectively at all frequency ranges compared with the passive rubber mount. Full article
(This article belongs to the Special Issue Magnetorheological Fluids, Devices, and Integrated Adaptive Systems)
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Figure 1
<p>Commercialized passive rubber mounts for cabin.</p> Full article ">Figure 2
<p>Experimental apparatus for damping force measurement. DAQ = data acquisition; RVDT = rotary variable differential transformer; LVDT = linear variable differential transformer.</p> Full article ">Figure 3
<p>Dynamic characteristics of passive mounts: (<b>a</b>) force-displacement diagram; (<b>b</b>) force-velocity diagram.</p> Full article ">Figure 4
<p>Transmissibility of passive rubber mounts.</p> Full article ">Figure 5
<p>Properties of magneto-rheological (MR) fluid: (<b>a</b>) general properties of MR fluid; (<b>b</b>) B-<span class="html-italic">H</span> curve; (<b>c</b>) <math display="inline"> <semantics> <mrow> <msub> <mi>τ</mi> <mi>y</mi> </msub> </mrow> </semantics> </math>-<span class="html-italic">H</span> curve.</p> Full article ">Figure 6
<p>Modeling of MR mount: (<b>a</b>) schematic structure; (<b>b</b>) design parameters and flow path.</p> Full article ">Figure 7
<p>Photograph of stiffness parts: (<b>a</b>) rubber part; (<b>b</b>) spring part.</p> Full article ">Figure 8
<p>Components of the fluid flow path: (<b>a</b>) annular path; (<b>b</b>) modified annular path; (<b>c</b>) radial path.</p> Full article ">Figure 9
<p>Simulation result of magnetic flux intensity.</p> Full article ">Figure 10
<p>Photograph of the MR mount: (<b>a</b>) parts; (<b>b</b>) assembled.</p> Full article ">Figure 11
<p>Properties of the proposed MR mounts: (<b>a</b>) force-displacement diagram; (<b>b</b>) force-velocity diagram.</p> Full article ">Figure 12
<p>Transmissibility diagram of the MR mounts.</p> Full article ">Figure 13
<p>Comparison of the passive rubber mount and the MR mount.</p> Full article ">
3265 KiB  
Article
Power Split Based Dual Hemispherical Continuously Variable Transmission
by Douwe Dresscher, Mark Naves, Theo J. A. De Vries, Martijn Buijze and Stefano Stramigioli
Actuators 2017, 6(2), 15; https://doi.org/10.3390/act6020015 - 10 Apr 2017
Cited by 6 | Viewed by 9391
Abstract
In this work, we present a new continuously variable transmission concept: the Dual-Hemi Continuously Variable Transmission (CVT). It is designed to have properties we believe are required to apply continuously variable transmissions in robotics to their full potential. These properties are a transformation [...] Read more.
In this work, we present a new continuously variable transmission concept: the Dual-Hemi Continuously Variable Transmission (CVT). It is designed to have properties we believe are required to apply continuously variable transmissions in robotics to their full potential. These properties are a transformation range that includes both positive and negative ratios, back-drivability under all conditions, kinematically decoupled reconfiguration, high efficiency of the transmission, and a reconfiguration mechanism requiring little work for changing the transmission ratio. The design of the Dual-Hemi CVT and a prototype realisation are discussed in detail. We show that the Dual-Hemi CVT has the aforementioned desired properties. Experiments show that the efficiency of the CVT is above 90% for a large part of the range of operation of the CVT. Significant stiction in the transmission, combined with a relatively low bandwidth for changing the transmission ratio, may cause problems when applying the DH-CVT as part of an actuator in a control loop. Full article
(This article belongs to the Special Issue Variable Stiffness and Variable Impedance Actuators)
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Figure 1

Figure 1
<p>A conceptual drawing of the Dual-Hemi CVT. To change the transmission ratio, the blue hemisphere rolls over the red one around the <span class="html-italic">x</span>-axis of the frame <math display="inline"> <semantics> <msub> <mi>ψ</mi> <mn>1</mn> </msub> </semantics> </math> in the point of contact (black dot in the figure). (<b>a</b>) An overview drawing of the conceptual mechanism. (<b>b</b>) A zoom on the contact point, with the addition of frame <math display="inline"> <semantics> <msub> <mi>ψ</mi> <mn>1</mn> </msub> </semantics> </math>. (The initial idea of the Dual-Hemi CVT is by Peter Rutgers, DEMCON, Enschede, The Netherlands.)</p> Full article ">Figure 2
<p>The complete transmission ratio range of the DH-CVT.</p> Full article ">Figure 3
<p>Schematic overview of a PS-CVT and a preview of the mechanical realisation (will be discussed in depth in <a href="#sec5-actuators-06-00015" class="html-sec">Section 5</a>). The colour coding illustrates the relation between parts in the schematic figure the drawing of the mechanical system. The parts connected to the ring gear of the planetary gear are blue, the parts connected to the cage of the planets of the planetary gear are coloured red and the parts connected to the sun gear of the planetary gear are coloured green. (<b>a</b>) Schematic overview of a PS-CVT. (<b>b</b>) Drawing of the mechanical system.</p> Full article ">Figure 4
<p>Transmission ratio of PS-CVT for typical values of <math display="inline"> <semantics> <mi mathvariant="script">X</mi> </semantics> </math>.</p> Full article ">Figure 5
<p>Transmission ratio of PS-CVT and DH-CVT.</p> Full article ">Figure 6
<p>A conceptual drawing of the Dual-Hemi CVT (<b>left</b>) in combination with a planetary gear (<b>right</b>). The fixed ratio transmissions are not shown.</p> Full article ">Figure 7
<p>Maximum allowable input torque <math display="inline"> <semantics> <msub> <mi>τ</mi> <mn>1</mn> </msub> </semantics> </math>.</p> Full article ">Figure 8
<p>Output torque <math display="inline"> <semantics> <msub> <mi>τ</mi> <mn>7</mn> </msub> </semantics> </math> over the selected configuration range.</p> Full article ">Figure 9
<p>An image showing the drive train of the dual-hemi CVT. The use of colours is consistent with <a href="#actuators-06-00015-f006" class="html-fig">Figure 6</a>: the parts connected to the ring gear of the planetary gear are blue, the parts connected to the cage of the planets of the planetary gear are coloured red and the parts connected to the sun gear of the planetary gear are coloured green.</p> Full article ">Figure 10
<p>An image showing the drive train and reconfiguration mechanism of the dual-hemi CVT.</p> Full article ">Figure 11
<p>A photo of the Dual-Hemi CVT. Photo by Wanda Tuerlinckx.</p> Full article ">Figure 12
<p>Simplified IPM of the transmission system.</p> Full article ">Figure 13
<p>Friction model.</p> Full article ">Figure 14
<p>Stiction, measurements and fitted curves.</p> Full article ">Figure 15
<p>Coulomb friction, measurements and fitted curves.</p> Full article ">Figure 16
<p>Efficiency for negative output velocities.</p> Full article ">Figure 17
<p>Efficiency for positive output velocities.</p> Full article ">Figure 18
<p>Schematic illustration of the reconfiguration mechanism.</p> Full article ">Figure 19
<p>Inverse transmission ratio of the PS-CVT.</p> Full article ">Figure 20
<p>Max magnitude of the torque that can be applied to the reconfiguration mechanism.</p> Full article ">Figure 21
<p>Scaling of the friction wheel radius with output torque.</p> Full article ">
14439 KiB  
Article
Development of a Compact Axial Active Magnetic Bearing with a Function of Two-Tilt-Motion Control
by Yuji Ishino, Takeshi Mizuno, Masaya Takasaki, Masayuki Hara and Daisuke Yamaguchi
Actuators 2017, 6(2), 14; https://doi.org/10.3390/act6020014 - 30 Mar 2017
Cited by 9 | Viewed by 9476
Abstract
A compact axial active magnetic bearing with a function of two-tilt-motion control is fabricated which has a new configuration of magnetic poles. They consist of four cylindrical poles with coils and a single common pole whose opposite plane of the rotor has a [...] Read more.
A compact axial active magnetic bearing with a function of two-tilt-motion control is fabricated which has a new configuration of magnetic poles. They consist of four cylindrical poles with coils and a single common pole whose opposite plane of the rotor has a permanent magnet to achieve multi-degree-of-freedom zero power control. Modal control is applied because local zero power control may make the whole system unstable when the number of control channels is larger than the number of freedoms of motion to be controlled. In the developed system, a disk-shape rotor is sandwiched between two axial magnetic bearing stators that are operated differentially. Such a configuration makes it possible to rotate the rotor without disturbing the axial motion. The characteristics of the fabricated magnetic bearing system including rotation are studied experimentally. Full article
(This article belongs to the Special Issue Active Magnetic Bearing Actuators)
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Figure 1

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<p>Schematic view of the proposed axial magnetic bearing.</p> Full article ">Figure 2
<p>Photograph of the fabricated three-degree-of-freedom magnetic suspension system.</p> Full article ">Figure 3
<p>Photograph of the upper electromagnet unit with displacement sensors.</p> Full article ">Figure 4
<p>Photograph of the floator attached permanent magnets and a sensor target.</p> Full article ">Figure 5
<p>Coordinate definition and the location of coils and sensors.</p> Full article ">Figure 6
<p>Block diagram of the control system for the translational motion.</p> Full article ">Figure 7
<p>Magnetic flux density in an analysis model with the proposed pole arrangement.</p> Full article ">Figure 8
<p>Relation between the displacement of the floator and the attractive force when the equivalent current is applied (analytical results).</p> Full article ">Figure 9
<p>Relation between the square of resonance angular frequency to the proportional gain (<b>a</b>) in the translational motion along the <span class="html-italic">z</span> axis; (<b>b</b>) in the rotational motion around the <span class="html-italic">x</span> axis; (<b>c</b>) in the rotational motion around the <span class="html-italic">y</span> axis.</p> Full article ">Figure 9 Cont.
<p>Relation between the square of resonance angular frequency to the proportional gain (<b>a</b>) in the translational motion along the <span class="html-italic">z</span> axis; (<b>b</b>) in the rotational motion around the <span class="html-italic">x</span> axis; (<b>c</b>) in the rotational motion around the <span class="html-italic">y</span> axis.</p> Full article ">Figure 10
<p>Block diagram of zero power feedback controller for the translational motion along the <span class="html-italic">z</span> axis.</p> Full article ">Figure 11
<p>Frequency characteristics of the simulated disturbance to the displacement with the zero power control. (<b>a</b>) Translational motion along the <span class="html-italic">z</span> axis; (<b>b</b>) Rotational motion around the <span class="html-italic">x</span> axis.</p> Full article ">Figure 12
<p>Step characteristics of equivalent current. (<b>a</b>) Translational motion along the <span class="html-italic">z</span> axis; (<b>b</b>) Rotational motion around the <span class="html-italic">x</span> axis; (<b>c</b>) Rotational motion around the <span class="html-italic">y</span> axis.</p> Full article ">Figure 12 Cont.
<p>Step characteristics of equivalent current. (<b>a</b>) Translational motion along the <span class="html-italic">z</span> axis; (<b>b</b>) Rotational motion around the <span class="html-italic">x</span> axis; (<b>c</b>) Rotational motion around the <span class="html-italic">y</span> axis.</p> Full article ">Figure 13
<p>Step characteristics of the displacement and angles. (<b>a</b>) Displacement along the <span class="html-italic">z</span> axis; (<b>b</b>) Angle around the <span class="html-italic">x</span> axis; (<b>c</b>) Angle around the <span class="html-italic">y</span> axis.</p> Full article ">Figure 13 Cont.
<p>Step characteristics of the displacement and angles. (<b>a</b>) Displacement along the <span class="html-italic">z</span> axis; (<b>b</b>) Angle around the <span class="html-italic">x</span> axis; (<b>c</b>) Angle around the <span class="html-italic">y</span> axis.</p> Full article ">Figure 14
<p>Step characteristics of the coil current in the electromagnet 1.</p> Full article ">Figure 15
<p>Time history of rotation speed.</p> Full article ">Figure 16
<p>Output of horizontal sensor. One pulse per rotation is generated which is used to estimate the rotational speed. (<b>a</b>) 20 seconds after the start of rotation; (<b>b</b>) approaching to the critical speeds.</p> Full article ">
2570 KiB  
Article
A System Identification Technique Using Bias Current Perturbation for the Determination of the Magnetic Axes of an Active Magnetic Bearing
by Dewey Spangler, Robert Prins and Mary Kasarda
Actuators 2017, 6(2), 13; https://doi.org/10.3390/act6020013 - 28 Mar 2017
Cited by 4 | Viewed by 7129
Abstract
Inherent in every Active Magnetic Bearing (AMB) are differences between the expected geometric axes and the actual magnetic axes due to a combination of discrepancies, including physical variation from manufacturing tolerances and misalignment from mechanical assembly, fringing and leakage effects, as well as [...] Read more.
Inherent in every Active Magnetic Bearing (AMB) are differences between the expected geometric axes and the actual magnetic axes due to a combination of discrepancies, including physical variation from manufacturing tolerances and misalignment from mechanical assembly, fringing and leakage effects, as well as variations in magnetic material properties within a single AMB. A method is presented here for locating the magnetic axes of an AMB that will facilitate the accurate characterization of the bearing air gaps for potential improvement in field tuning, performance analyses and certain shaft force measurement techniques. This paper presents an extension of the application of the bias current perturbation method for the determination of the magnetic center to the determination of magnetic axes for the further development of accurate current-based force measurement techniques. Full article
(This article belongs to the Special Issue Active Magnetic Bearing Actuators)
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Graphical abstract

Graphical abstract
Full article ">Figure 1
<p>Equilibrium of a magnetically-suspended object.</p> Full article ">Figure 2
<p>Experimental rotor test stand configuration (photo: [<a href="#B2-actuators-06-00013" class="html-bibr">2</a>]).</p> Full article ">Figure 3
<p>Active Magnetic Bearing (AMB) configuration.</p> Full article ">Figure 4
<p>Geometric coordinates for rotor space Quadrant 1.</p> Full article ">Figure 5
<p>Rotation of geometric coordinate axes in Quadrant 1.</p> Full article ">Figure 6
<p>Transformed coordinates with respect to geometric origin.</p> Full article ">Figure 7
<p>Spatial relationship between geometric and effective origins. All values shown are in microns (µm).</p> Full article ">Figure 8
<p>Iteration vs. set point and <span class="html-italic">v</span> axes MPM coordinate.</p> Full article ">Figure 9
<p>Relationship between geometric, transformed and effective origins.</p> Full article ">Figure 10
<p>Geometric vs. effective coordinates for the <span class="html-italic">v</span> axis.</p> Full article ">
883 KiB  
Article
Modeling and Realization of a Bearingless Flux-Switching Slice Motor
by Wolfgang Gruber and Karlo Radman
Actuators 2017, 6(2), 12; https://doi.org/10.3390/act6020012 - 27 Mar 2017
Cited by 14 | Viewed by 9772
Abstract
This work introduces a novel bearingless slice motor design: the bearingless flux-switching slice motor. In contrast to state-of-the-art bearingless slice motors, the rotor in this new design does not include any permanent rotor magnets. This offers advantages for disposable devices, such as those [...] Read more.
This work introduces a novel bearingless slice motor design: the bearingless flux-switching slice motor. In contrast to state-of-the-art bearingless slice motors, the rotor in this new design does not include any permanent rotor magnets. This offers advantages for disposable devices, such as those used in the medical industry, and extends the range of bearingless slice motors toward high-temperature applications. In this study, our focus is on the analytical modeling of the suspension force torque generation of a single coil and the bearingless motor. We assessed motor performance in relation to motor topology by applying performance factors. A prototype motor was optimized, designed, and manufactured. We also presented the state-of-the-art nonlinear feedback control scheme used. The motor was operated, and both static and dynamic measurements were taken on a test bench, thus successfully demonstrating the functionality and applicability of the novel bearingless slice motor concept. Full article
(This article belongs to the Special Issue Active Magnetic Bearing Actuators)
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Figure 1

Figure 1
<p>Cross-section of a flux-switching motor with twelve stator and 10 rotor teeth. The arrows indicate the direction of magnetization.</p> Full article ">Figure 2
<p>Principles of torque and force generation: Torque is created by an energized stator coil with angle-dependent permanent magnetic flux linkage (illustrated at the top for two different rotor angles), while force is created by field strengthening and weakening (illustrated at the bottom).</p> Full article ">Figure 3
<p>Unrolled cross-sectional part of the flux-switching motor, with the field lines created by the stator coil current (<b>left</b>) and one permanent magnet placed between the stator teeth (<b>right</b>).</p> Full article ">Figure 4
<p>Normalized permeances and air gap flux densities of the flux-switching motor for mechanical rotor angles <span class="html-italic">ϕ</span> of 0° (<b>left</b>) and 16° (<b>right</b>).</p> Full article ">Figure 5
<p>Analytically computed and separated suspension force components; all curves were normalized with the maximum of <span class="html-italic">F<sub>x,l</sub></span>(<span class="html-italic">ϕ</span>)<span class="html-italic">i<sub>s</sub></span>.</p> Full article ">Figure 6
<p>Analytically computed flux linkages (<b>left</b>) and separated torque components (<b>right</b>); all curves were normalized with the maximum of Ψ<span class="html-italic"><sub>PM,s</sub></span> and <span class="html-italic">T<sub>z,l</sub></span>.</p> Full article ">Figure 7
<p>Connection of two opposing coils in series to compensate for the quadratic force component, and to decouple force and torque.</p> Full article ">Figure 8
<p>Single-phase torque and force characteristics of the linearized and decoupled bearingless flux-switching slice motor.</p> Full article ">Figure 9
<p>Bearingless flux-switching motor compositions with 12 stator teeth and double three-phase winding system.</p> Full article ">Figure 10
<p>Block diagram of the nonlinear feedback control scheme for controlling the double three-phase bearingless flux-switching slice motor.</p> Full article ">Figure 11
<p>Geometric parameters of the bearingless flux-switching motor.</p> Full article ">Figure 12
<p>Photograph of the bearingless flux-switching motor prototype.</p> Full article ">Figure 13
<p>Comparison of the finite-element simulated and measured single-phase characteristics of the suspension forces (<b>left</b>), and the torque (<b>right</b>).</p> Full article ">Figure 14
<p>Reactions of the <span class="html-italic">x</span>- and <span class="html-italic">y</span>-positions following a step in the set-point value (<b>left</b>), and a step change in the suspension force (<b>right</b>).</p> Full article ">Figure 15
<p>Efficiency and load factor of the manufactured bearingless flux-switching motor.</p> Full article ">
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