Wireless Power Transfer Using Oscillating Magnets - IEEE Xplore ...

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 65, NO. 8, AUGUST 2018

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Wireless Power Transfer Using Oscillating Magnets Sijun Du

, Member, IEEE, Edward K. Chan , Member, IEEE, Bing Wen, John Hong, Hanspeter Widmer , and Charles E. Wheatley , Life Fellow, IEEE

Abstract—The advancement of portable and implantable electronic devices has driven active research into wireless power transfer (WPT) methods. Current commercial WPT techniques using coil coupling have limitations, such as range, frequency, safety, line of sight, orientation, and concurrency. This paper presents a system utilizing oscillating magnets to extract power from a low-level ac magnetic field. This system enables the simultaneous powering of multiple low-power devices in an environment that meets human safety requirements. Compared with conventional WPT techniques, such as RF or acoustic power transfer, the proposed system transfers stable power to electronic nodes placed anywhere within the magnetic field under low driving frequency. Theoretical power calculations and experimental demonstrations are presented to prove the concepts and quantify performance. A custom three-dimensional printed millimeter-size oscillator loaded with two 8 mm3 cubic permanent magnets is employed as the energy receiver to experimentally validate the proposed system. Measurements show that the raw ac electrical power and rectified dc power received by the oscillator achieve 269 and 135 µW, respectively, under 400 µT excitation magnetic field amplitude at 355 Hz. Index Terms—Boost converter, energy harvesting, Helmholtz coil, inductive transducer, magnetic fields, wireless power transfer (WPT).

I. INTRODUCTION HILE the Internet of Things (IoT) continues to expand in both number and variety of deployed devices, keeping them powered unobtrusively is largely an unmet need. Practical systems have concentrated on device efficiency, often reducing function and/or performance so that the battery or device replacement cycle can be lengthened. There is a critical need for a system that can transfer electrical power efficiently without vastly increasing the device size or imposing other require-

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Manuscript received June 20, 2017; revised September 12, 2017, October 10, 2017, and November 18, 2017; accepted December 7, 2017. Date of publication December 22, 2017; date of current version April 2, 2018. (Corresponding author: Edward K. Chan.) S. Du is with Qualcomm Technologies, Inc., San Diego, CA 92121 USA, and also with the Department of Engineering, University of Cambridge, Cambridge CB2 1PZ, U.K. (e-mail: [email protected]). E. K. Chan, B. Wen, J. Hong, H. Widmer, and C. E. Wheatley are with Qualcomm Technologies, Inc., San Diego, CA 92121 USA (e-mail: [email protected]; [email protected]; [email protected]; [email protected]; wheatley@qti. qualcomm.com). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TIE.2017.2786289

ments that would add to the cost and complexity of deployment [1], [2]. Even as the cost of sensors, cameras, transmitters, and other useful widgets decreases, the cost of installing them, most notably with a reliable power supply, decreases little if at all. Currently, these devices are hardwired, use batteries, harvest energy from the environment or use some type of wireless power transfer (WPT). Hardwiring can be expensive and difficult, likely restricting the type of systems that can be built. Batteries have very limited lifetimes and can be very difficult and tedious to replace. Energy harvesting from thermal, vibration, radio-frequency (RF) or solar sources has still not proven to be reliable enough for many critical applications [3]–[7]. On the other hand, coil-to-coil WPT has proven to be reliable in many consumer electronics [8]–[13]. Conventional WPT utilizes focused beams of electromagnetic or acoustic waves to transfer energy between two or more coupled coils [14]–[21] providing the possibility of connectorfree electronics. Wireless charging for smartphones [22], implantable biomedical devices [23], and other portable or mobile electronics [24], [25] is growing fast in popularity and is suitable for transferring relatively large amounts of power (several Watts) but requires close proximity and controlled orientation to be effective [26]–[30]. Some sophisticated beam steering methods to deliver energy at a distance (Energous, Ossia) or through human tissue (Stanford) [31] have also been demonstrated. A WPT method using electromechanical receivers has also been proposed in [32]. Acoustic waves, especially ultrasound, are well suited for delivering power through compatible materials, such as water and human tissue. Commercial success of these methods proves the need and desire for wireless charging with opportunities for techniques that can expand the use conditions [33]. This paper describes a system to transmit power wirelessly by utilizing permanent magnets driven at resonance by a low-level low-frequency magnetic field. Although magneto-mechanical oscillators have been presented in [34] and [35], this paper employs a Helmholtz coil to provide uniform WPT across a potentially large space volume and theoretically and experimentally validates the feasibility and output power to power a light-emitting diode (LED). The system has a high power-tovolume ratio and can be scaled from 10 mm3 to over 1000 mm3 to provide power in the range of 10’s of microwatts to many milliwatts. Depending on the application, the driving magnetic field can be limited to satisfy international commission on nonionizing radiation protection (ICNIRP) [36] field safety limits

This work is licensed under a Creative Commons Attribution 3.0 License. For more information, see http://creativecommons.org/licenses/by/3.0/

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TABLE I COMPARISONS OF DIFFERENT WPT TECHNIQUES Technique Frequency Transfer range

This work

RF

Acoustic

10’s−100’s Hz Anywhere in Helmholtz coil

10’s MHz 10’s cm

10’s kHz−1’s MHz 1’s cm

around people, or raised higher when human safety is not critical. The driving frequency can be between 10’s Hz and 100’s Hz, which is much lower than most of reported WPT techniques, such as RF WPT in 10’s MHz frequency range [17], [21], [37] or ultrasonic WPT between 10’s kHz and 1’s MHz [38]–[40]. The low-frequency driving magnetic field in the proposed method easily penetrates many solids such as concrete and wood, including nonmagnetic metals such as aluminum, copper, and stainless steel, and liquids such as sea water. The magnetic field source can be a planar spiral coil for short range charging over a surface, or a Helmholtz coil pair to power a volume. In addition, power receivers can be placed anywhere in the field even at a substantial distance from the transmitters and mounted on a self-aligning mechanical gimbal to provide a consistent and optimum orientation to the magnetic field. A summary comparing the proposed WPT technique with the existing RF and acoustic WPT systems is shown in Table I. In this paper, Section II presents the architecture and working principle of the proposed WPT system. The mechanical and electrical modeling is presented in Section III. The experimental setup and measurements are provided in Section IV and a conclusion is given in Section V. II. PROPOSED WPT SYSTEM The proposed system contains a source to generate an excitation ac magnetic field, an oscillator with permanent magnets, magnetic coils to extract energy from the oscillating magnets and an interface circuit to rectify the energy induced in the magnetic coils. Fig. 1(a) shows the principle of the proposed system. The excitation magnetic source can be a planar spiral coil for short range charging over a surface, or a Helmholtz coil pair to power a volume, which is chosen to be implemented in this paper. The amplitude of the excitation field B E can range from 10’s to 100’s μT, which is comparable to the Earth’s magnetic field (25–65 μT). As the system aims to provide wireless power in a large space, the Helmholtz coil pair can be made large to cover a room. Contrasted with conventional WPT techniques such as RF, optical or acoustic power transfer which all involve directed energy, the proposed system achieves wireless power transfer uniformly within a potentially large volume of space. Fig. 1(b) shows the simulated results with COMSOL Multiphysics, which shows the uniform magnetic field generated by a Helmholtz coil [41]. Hence, the harvested power by the nodes remains constant throughout the Helmholtz coil coverage volume. As the input frequency is also very low to match the natural frequency of the oscillator, the weak and low-frequency magnetic field is compatible with the exposure recommendations of

the ICNIRP [36], [42]. The oscillator consists of a plate with two cube permanent magnets on each side suspended by two torsional hinges. The size of each magnet can be as small as a few millimeters. A magnetic coil is placed around the oscillator to generate induced power according to Faraday’s Law and the raw electrical power from the coil is further rectified before powering a load. The block architecture of the system and the energy flow are shown in Fig. 2, which describes the operation principle of the WPT mechanism. A function generator generates a sine wave that matches the natural frequency of the magnet oscillator. This sine signal drives a Helmholtz coil to generate an ac magnetic field. The magnet oscillator is excited at resonance in this magnetic field to turn the magnetic energy into mechanical energy. The mechanical energy stored in the oscillator is then turned into electrical form by employing magnetic coils using Faraday’s Law. A dual-path boost-converter interface circuit, which utilizes the coil’s own inductance as the power inductor, is employed to convert the unstable ac electrical power into stable dc power for loads. III. MODELING This section provides the theoretical calculations of the resulting electrical power induced in the magnetic coil of the proposed WPT system. Given the system illustrated in Fig. 1(a), where an external excitation magnetic field with flux density B E resonantly couples energy into the magnetic oscillator, it is important to know how much magnetic power can be extracted and converted into electrical power. The external excitation magnetic field can be described as B E (t) = B0 cos(ωt).

(1)

The harmonic oscillation equation while the oscillator is excited at its natural frequency is given by I θ¨ = −K θ − β θ˙ + m B0 cos(ωt)

(2)

where I is the moment of inertia of the oscillator, θ is the rotational angle of the oscillator, K is the spring constant due to the two torsional hinges, β is the damping coefficient, m is the magnetic moment of the magnets, and B0 is the amplitude of the external excitation magnetic field flux density. Hence, the natural frequency of the torsional oscillator is given as K . (3) ω0 = I The rotational angle of the oscillator is written as θ = θ0 cos(ωt − φ)

(4)

where θ0 is the rotational oscillation amplitude and φ is the phase of the oscillation, which can be expressed as θ0 =

1 m B0 I (ω2 − ω2 )2 + ( β ω)2

(5)

ω β I ω02 − ω2

(6)

0

tan φ =

I

DU et al.: WIRELESS POWER TRANSFER USING OSCILLATING MAGNETS

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Fig. 1.

(a) Working principle of the proposed WPT system and (b) simulated uniform magnetic field generated from the Helmholtz coil.

Fig. 2.

System level block architecture and energy flow.

where ω is the excitation frequency of the excitation magnetic field. When the oscillator is excited at resonance such that ω = ω0 , the above two equations can be rewritten as m B0 βω0 π φ= . 2

In the previous equation, the magnetic moment m can be calculated from the magnetic field strength of the magnet, which is expressed as m=

θ0 =

(7)

Hence, we have the rotational angle of the oscillator (4) written as m B0 π m B0 I θ= = sin(ωt) (8) cos ωt − βω0 2 β K

BM HWT μ0

(9)

where μ0 is the free space permeability and B M is the magnetization of the magnet. H , W, and T are the geometrical length, width, and thickness of the oscillator, respectively, as shown in Fig. 3. The moment of inertia I of the oscillator can be written as HWT 2 H + T2 (10) I =ρ 12

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Fig. 3.


Modeling parameters for x–y–z view vibration mode shape (left) and x–z plane intersection view with magnetic coil.

where ρ is the effective density of the oscillator. As the Q value of the oscillator can be written as Q = I βω0 , the damping coefficient can be expressed as β=

I ω0 . Q

(11)

Hence, replacing parameters in (8) with the previous equations, (8) can be rewritten as m B0 π B0 B M Q θ= = cos ωt − H W T sin(ωt) = θ0 sin(ωt) βω0 2 K μ0 (12) where θ0 = B0KBμM0Q H W T . The equation above shows the torsional angle of the oscillator while it is vibrating at its natural frequency ω = ω0 . The rotational angle is also illustrated in Fig. 3 (right). Next, the induced electrical power in the coil will be calculated to determine how much raw electrical power can be obtained by this wireless energy receiver. Assuming that the clearance between the oscillator and the coil is minimized while making sure that they do not touch during oscillation, the magnetic flux dropoff from the magnet surface to the coil is assumed to be negligible. If the electrical force on the oscillator due to the induced current in the coil is also ignored, the magnetic flux in the magnetic coil can be written as = B M W T sin θ.

(13)

Assuming the angle θ is small, the previous equation can be approximated as ≈ B M W T θ = B M W T θ0 sin(ωt).

(14)

Hence, the induced electromotive force (emf) voltage in the coil is ∂ = −ωN B M W T θ0 cos(ωt) (15) Vcoil0 = −N ∂t where N is the number of turns of the magnetic coil and θ0 = B0 B M Q H W T , as shown in (12). Assuming the inductance and K μ0

the dc resistance of the coil are L and R, the impedance of the coil is (16) Z coil = |R + jωL| = R 2 + ω2 L 2 . The phase between the voltage and current in the coil is ). For the coil used in this imexpressed as ϕc = arctan(− ωL R plementation, L = 0.85 mH, R = 67 and the frequency is 355 Hz. Hence, ϕc ≈ 0 and it is neglected here and in future calculations. Hence, the magnitude of the current in the coil while its two electrodes are shorted is Vcoil ωN B M W T θ0 = −√ cos(ωt). (17) Icoil0 = Z coil R 2 + ω2 L 2 Associated with (17) and (16), the generated power in the coil can be expressed as 2 Pcoil0 = Icoil0 Z coil .

(18)

This is the maximum raw power that needs to be rectified with an interface circuit before it can be used to power load electronics. The real electrical power will be smaller than (18) because the induced current in the coil will generate a magnetic field opposite to the excitation field from the Helmholtz coil. Therefore, the effective excitation magnetic field applied on the oscillator is weaker than that in (1) and the resulting vibration amplitude should be smaller than θ0 calculated in (12). This can easily be observed from the difference in oscillation amplitude between having the coil terminals shorted or open circuited and this will be presented in Section IV. As a result, the real electrical power in the coil is lower than that in (18). In order to make a more accurate estimation of the electrical output power, a few iterative calculations are needed by applying the induced magnetic field from the coil to the system. As the short-circuit current Icoil in the coil is calculated in (17), the induced magnetic field in the coil can be written as ωμ0 N 2 B M W T θ0 cos(ωt). Bcoil = μ0 N Icoil = − √ R 2 + ω2 L 2

(19)


Fig. 4.

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Experimental setup.

The polarization of the induced magnetic field is opposite to the external excitation magnetic field B E . Hence, the effective excitation magnetic field is decreased due to the induced current in the coil and, after the first iteration of the magnetic field, the effective excitation magnetic field on the oscillator can be rewritten as

2 Q H W 2T 2 ωN 2 B0 B M cos(ωt). (20) B E (t) = B0 − √ K R 2 + ω2 L 2 Compared to (1), this magnetic field is weaker than the applied field from the Helmholtz coil. With this updated excitation magnetic field, the new vibration amplitude θ0 can be calculated as

2 Q H W 2T 2 B0 B M Q ωN 2 B M sin(ωt). θ1 = HWT 1 − √ K μ0 K R 2 + ω2 L 2 (21) ωN 2 B 2 Q H W 2 T 2 Defining A = K √ MR 2 +ω2 L 2 , the previous equation can be rewritten as θ1 = θ0 − Aθ0 = θ0 (1 − A).

(22)

The short-circuit current in the coil can then be calculated as ωN B M W T (θ0 − Aθ0 ) cos(ωt) = Icoil0 (1 − A). Icoil1 = − √ R 2 + ω2 L 2 (23) Hence, the resulting electrical power in the coil after the first iteration is Pcoil1 = Pcoil0 (1 − A)2

(24)

As the feedback term is A, the transfer function of a system with negative feedback term A is given as 1/(1 + A). Hence, according to (22) and (24), the final oscillation amplitude and output power can be expressed as θcoil =

θ0 (1 + A)

(25)

Pcoil =

Pcoil0 (1 + A)2

(26)

where θ0 and Pcoil0 are the calculated oscillation amplitude of the oscillator and electrical power in coil without considering

the feedback emf due to the induced current in the coil, which are given in (12) and (18), respectively. This is the raw electrical power consumed inside the coil while the two electrodes are shorted. As the voltage induced in the coil is an unstable ac voltage, an interface circuit is needed to extract the power from the coil and store it in a storage capacitor to power loads. Essential to designing an efficient receiver circuit in a small volume is exploiting the large inductance of the receiver coil itself as part of the ac–dc rectifier. IV. MEASUREMENT RESULTS The proposed WPT system was experimentally evaluated using a three-dimensional (3-D) printed oscillator sandwiched by two 2 mm × 2 mm × 2 mm neodymium magnets. Fig. 4(a) shows the experimental setup and Fig. 4(b) illustrates the working principle of the measurements. In order to generate an ac magnetic field, a Helmholtz coil is driven by a sine wave from a function generator amplified by a power amplifier. A resistor is connected in series with the Helmholtz coil to limit and adjust the electric current in the Helmholtz coil. The driving frequency is tuned to match the resonant frequency of the oscillator placed in the Helmholtz coil. In order to view the oscillation and to find the mechanical specifications of the oscillator, a laser beam is reflected off an optical mirror onto the surface of the oscillator. The beam is then projected onto a projection wall in order to view and measure the oscillation amplitude. Assuming D is the distance between the wall and oscillator and d is the width of the beam projection, the vibration amplitude of the oscillator d . can be calculated as θ = arctan 2D Fig. 5 shows the zoomed-in view of the Helmholtz coil and the energy receiver inside. The Helmholtz coil is commercially available (3B Scientific U21901) with 100 turns and measured 0.78 dc resistance for each coil. The Helmholtz coil is driven by a sine wave signal and the amplitude of the magnetic field in the center of the Helmholtz coil is adjustable between 25 and 400 μT, which is just several times stronger than the earth magnetic field. The oscillator is 3-D printed with a plastic material with the modulus of elasticity around 3000 MPa. A zoomed-in view on the oscillator itself is also provided in the figure. It

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Fig. 5.


Zoomed-in view on the oscillator and permanent magnets.

can be seen that there are two 2 mm cube neodymium magnets placed on both sides of the centered plate, which is suspended by two thin hinges with size of 300 μm × 300 μm × 3000 μm. A model was built in COMSOL for simulations and the dimensions of the simulation model are the same as the device in experiment. The permanent magnets are simulated with NdFeB material with density of 7500 kg/m3 and the hinges are simulated with acrylic plastic with Young’s modulus of 3200 MPa. The COMSOL simulated natural frequency of this oscillator is 322 Hz and the measured natural frequency is around 355 Hz, which approximately matches the simulated result. The Helmholtz coil is driven by a sine wave signal of 355 Hz to excite the oscillator at its resonance. A magnetic coil is made by looping AWG 42 wire (63 μm diameter) with around 600 turns around a plastic core. The coil is fixed at the tip of a micromanipulator and the position of the coil is precisely adjusted to maximize the induced voltage in the coil during the oscillation of the magnets. Fig. 6 shows the frequency response of the oscillator with the amplitude of the excitation magnetic field at 200 μT. During the measurements, a laser beam is used to measure the oscillation amplitude, as explained in Fig. 4(b). The amplitude values in degrees shown in Fig. 6 are calculated from the projection of the laser beam on the wall. The figure shows that the natural frequency of the oscillator is around 355 Hz and the Q value is calculated at around 17. This Q factor is enough to achieve expected results as the proposed wireless power system does not require very high Q to be effective. Fig. 7(a) shows the measured vibration amplitude of the oscillator and the induced voltage in the coil while the excitation magnetic field is varied from 25 to 400 μT. The excitation frequency is set to 355 Hz, which is the resonant frequency of the oscillator. It can be seen from the figure that the vibration amplitude and emf vary approximately linearly with the excitation amplitude. The slight nonlinearity is due to the spring hardening effect at large oscillation amplitude. The spring hardening effect is mainly contributed by the decreased torque at high vibration

Fig. 6. Measured frequency response of the oscillator with the excitation magnetic field amplitude of 200 μT.

amplitudes. While the oscillator is in the stationary position where θ = 0◦ (see the right figure in Fig. 3), the vibration is in the same direction of the excitation magnetic field and the torque applied to the oscillator is the peak, noted as τ0 . When θ goes larger at a higher vibration amplitude, the torque applied to the oscillator by the electromagnetic field is decreased to τ0 cos θ . Therefore, the decreased torque at higher vibration amplitudes introduces the nonlinearity in Fig. 7(a), which is called spring-hardening effect. The inductance and dc resistance of the coil is measured as 0.85 mH and 67 , respectively. Hence, the raw electrical power in the coil can be calculated and the results are shown as solid line in Fig. 7(b). It can be observed that the output power, which is supposed to be quadratic to the excitation field amplitude in linear vibration systems, is now affected by the spring-hardening effect, especially at higher excitation levels. With the excitation field amplitude set at 200 μT, it can be seen that the electrical power can be around 118 μW, which is able to power most low-power wireless sensor nodes. Higher electrical power can also be obtained by increasing the excitation magnetic field amplitude or utilizing larger magnets. It is useful to quantify the effect of the short-circuit current in the coil on the vibration amplitude and finding the factor A in (25) and (26). According to (25), the factor A can be excoil . The vibration amplitudes without coil pressed as A = θ0θ−θ coil (θ0 ) and with short-circuited coil (θcoil ) are measured and shown in Fig. 8. The factor A is calculated for different excitation field amplitudes varying from 50 to 400 μT. According to the theoretically calculated A value just above (22), A is independent of the excitation field amplitude B0 . The decrease of A shown in this figure is due to the spring-hardening effect of the oscillator and this effect can also be observed from the bending trend of the curves in Fig. 7(a). As the electrical power shown in Fig. 7(b) is the raw ac power induced in the coil, this power cannot be directly used to power electronic devices before it is rectified and buffered. Therefore, an interface circuit is needed to extract the power


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Fig. 7. Measured performance results: (a) Vibration amplitude and induced voltage in the coil (emf) and (b) raw electrical output power in the coil and rectified electrical output power storage in capacitors while the excitation magnetic field is varied from 25 to 400 μT.

Fig. 9.

Interface circuit employed in the experiments.

Fig. 8. Vibration amplitude decrease due to the short-circuit current in the coil and the measured factor A.

and store the energy into storage capacitors. The rectified energy is then used to power a load, which is an LED indicator in this implementation. In order to extract energy induced in magnetic coils, different passive or active circuits have been reported [43]–[46]. Fig. 9 shows the circuit diagram of the power extraction circuit employed in this implementation, which includes a dual-path boost converter, two storage capacitors, C S P and C S N , and an LED indicator. The magnetic coil is modeled as a voltage source Vcoil in series with an inductor L coil and a resistor Rcoil . The switch S1 is driven by a switching signal pulse width modulation (PWM) at 21 kHz with 95% duty cycle. In this implementation, the PWM signal is generated from a signal generator to experimentally evaluate the feasibility of the proposed WPT system. In real-world implementations, an on-chip ultralow power oscillator can be designed to provide the PWM signal. Some on-chip low-power oscillators have been presented in [43], [47], and [48], which generate 10’s kHz PWM signals with sub-μ W power consumption. The inherent inductance of the coil serves as the inductor of the boost converter; hence, no separate inductor is needed and the volume and cost of the

circuit can be greatly reduced. Furthermore, the circuit is able to boost very low emf values, especially suitable for the proposed WPT system, since it operates in current-mode. While the switch S1 is ON, the two electrodes of the coil are shorted and the energy is stored in the inductor L coil in electromagnetic form. After the switch is turned OFF, the energy stored in the inductor is then transferred into C S P or C S N according to the polarization of the current in the coil. As the duty cycle of the PWM signal is 95%, the coil can be regarded as being shorted and the effects of using a such boost converter on the vibration amplitude and output power are shown by the measured factor A in Fig. 8. Hence, according to (25) and (26), the decreases in vibration amplitude and output power due to the boost converter can be calculated. In the circuit, Schottky diodes with low 330 mV forward voltage drop are used to minimize power loss in the diodes. The two storage capacitors are 4.7 μF ceramic capacitors. The voltage drop of the LED indicator is 1.6 V and the typical current rate is 1 mA. During the energy extraction, the switch S2 is kept OFF for most of time in order to let the two storage capacitors, C S P and C S N , to be charged to high voltage levels. After a period of time, the switch S2 is pushed and the LED lights up for a short

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Fig. 10.


Printed circuit board of the interface circuit. (a) Pushbutton not pushed. (b) Instant while the button is pushed.

TABLE II PERFORMANCE COMPARISON WITH STATE-OF-THE-ART Reference

Technique

Operating frequency

Transmission range

Peak power (P)

[38] [39] [31] [18] This work

Ultrasonic Ultrasonic Electromagnetic Electromagnetic Oscillating magnets

17 kHz 48.25 kHz 1.6 GHz 4.63 MHz 355 Hz

0.75 cm 88 cm 5 cm 45 cm Anywhere in Helmholtz coil

37 μW 50 μW 200 μW 3.1 W 269 μW

instant until the charge stored in the two capacitors is exhausted. With an excitation field amplitude of 300 μT, the switch can be manually pushed four times per second with the 0.25 s charging time being sufficient to charge the capacitors. Fig. 10 shows the printed circuit board for the measurement and Fig. 10(b) shows the instant while the switch S2 is pushed and the LED is alight. Rectified electrical power stored in the two storage capacitors for different excitation amplitude levels is shown as the dashed line in Fig. 7(b). The energy stored in capacitors is rectified and it is ready to be used as a stable supply with a voltage regulator. Compared to raw electrical power in the coil shown in Fig. 7(b), it can be seen that the power efficiency of the dual-path boost converter is around 50%. The peak power with 400 μT excitation amplitude reaches 135 μW, which is able to power most of wireless electronic sensors. Table II shows the performance comparisons among the stateof-the-art WPT techniques and the proposed technique using oscillating magnets. The second column shows the employed techniques. The following two columns are the operating frequency of driving signals and specified wireless power transmission ranges. Compared to the cited publications, the proposed WPT technique operates at a much lower driving frequency and the transmission range is not limited by a specific value. The energy receivers placed anywhere in the Helmholtz coil can be powered due to the uniform electromagnetic field generated by the Helmholtz coil. The following two columns show the peak electrical power received by the power receiving nodes and

Power receiver size Received power density = (V ) 0.53 cm3 12.6 mm3 14 mm3 701.8 cm3 16 mm3

P V

0.07 mW/cm3 3.97 mW/cm3 14.3 mW/cm3 4.42 mW/cm3 16.8 mW/cm3

estimated volumes of power receivers. The volume values for state-of-the-art publications are calculated from the dimensions given in the papers. The volume value for this work is calculated from the two stacked neodymium magnets of dimension 2 mm × 2 mm × 2 mm. The last column shows the received power density, which is obtained by dividing the electrical power by the receiver volume. From the table, it can be seen that the proposed WPT technique using oscillating magnets achieves the highest power density while it operates at a much lower driving frequency and the receiver can be placed anywhere in the Helmholtz coil. V. CONCLUSION This paper proposed a new WPT system that utilizes resonating magnets to receive power from an ac magnetic field source, which can be a Helmholtz Coil. The proposed system first transforms the input magnetic energy into mechanical energy stored in the resonating magnets. The energy is then extracted from magnetic coils in close proximity to the magnets to generate usable electrical energy. Compared to conventional WPT methods using direct coil-to-coil techniques, the proposed system has three main advantages in frequency, transfer range, and orientation. The input frequency for conventional coil-to-coil technique is usually high, which can be 1 MHz or higher; however, the proposed system can work at 100’s Hz or even lower because the frequency should match the natural frequency of the mechanical


oscillator. The oscillator can also be designed to have natural frequency between 50 Hz and 60 Hz; hence, the power source (Helmholtz coil) can be directly powered by household or mains electricity, which facilitates circuit design and real-world implementations. In terms of the power transfer range, most of conventional WPT systems only allow near-field power transfer, especially for small receivers. For some long-distance WPT systems, precise orientation is required. Hence, conventional WPT systems suffer from a compromise between these two issues. However, the proposed energy receiver (oscillating magnets) can be placed anywhere within the driving Helmholtz coils and, with a mechanically self-aligning gimbal mount, the requirement for orientation is completely eliminated. In order to power energy receiving nodes in an entire room, the Helmholtz coil can also be designed to have the same scale as the room or the building. Hence, the excitation magnetic field exists anywhere in the building and any energy receiving nodes covered by the field can be powered. As the excitation magnetic field amplitude is just a few times higher than the earth magnetic field and the frequency is extremely low, the field strength is compatible with the exposure recommendations of the ICNIRP. In order to prove the proposed system and evaluate the energy transfer performance, a 3-D printed oscillator with two neodymium magnets is fabricated to oscillate in the magnetic field driven by a Helmholtz coil. A laser is employed to view the oscillation and a dual-path boost converter circuit is built to extract power from the magnetic coils. The experimental results show that while the oscillator is excited in a magnetic field of 200 μT amplitude, the raw electrical power obtained in the coil is around 118 μW and the usable rectified electrical power is as high as 55 μW, which is enough to power most of low-power sensors. When the excitation field amplitude is increased to 400 μT, the raw electrical power and rectified electrical power go up to 269 μW and 135 μW, respectively. The electrical power is then used to light an LED indicator to demonstrate useful work. This design focuses on extracting energy from a driven oscillating magnetic field to obtain significant power in a small volume. This system is well suited to power multiple low-powered devices concurrently that consume a relatively small total power. The overall system efficiency is generally lower than other more targeted techniques since the driving magnetic field is not focused on particular devices. It operates in the uncoupled regime where the energy consumer does little to perturb the driving field. This technique enables applications in volumes that were previously inaccessible due to space or material constraints, or significantly impaired by shielding or noise. Power is supplied by magnetic transmitters that can be inherently safe, conforming to the ICNIRP standard when necessary, or going to higher power in more tolerant environments. These transmitters are generally flat making them suitable for installation in or on walls, floors or furniture. Future work will focus on increasing the system size by using a large Helmholtz coil to simultaneously power multiple energy receivers to study the interaction between the nodes. Another task is redesigning the oscillator and the magnets to match the frequency of the mains electricity.

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REFERENCES [1] A. S. M. Z. Kausar, A. W. Reza, M. U. Saleh, and H. Ramiah, “Energizing wireless sensor networks by energy harvesting systems: Scopes, challenges and approaches,” Renewable Sustain. Energy Rev., vol. 38, pp. 973–989, 2014. [Online]. Available: http://www.sciencedirect.com/ science/article/pii/S1364032114004870 [2] T. Rault, A. Bouabdallah, and Y. Challal, “Energy efficiency in wireless sensor networks: A top-down survey,” Comput. Netw., vol. 67, pp. 104–122, 2014. [Online]. Available: http://www.sciencedirect.com/ science/article/pii/S1389128614001418 [3] F. K. Shaikh and S. Zeadally, “Energy harvesting in wireless sensor networks: A comprehensive review,” Renewable Sustain. Energy Rev., vol. 55, pp. 1041–1054, 2016. [Online]. Available: http://www.sciencedirect. com/science/article/pii/S1364032115012629 [4] M. H. Elsheikh et al., “A review on thermoelectric renewable energy: Principle parameters that affect their performance,” Renewable Sustain. Energy Rev., vol. 30, pp. 337–355, 2014. [5] P. Nintanavongsa, “A survey on RF energy harvesting: Circuits and protocols,” Energy Procedia, vol. 56, pp. 414–422, 2014. [Online]. Available: http://www.sciencedirect.com/science/article/pii/S1876610214010364 [6] R. M. Ferdous, A. W. Reza, and M. F. Siddiqui, “Renewable energy harvesting for wireless sensors using passive RFID tag technology: A review,” Renewable Sustain. Energy Rev., vol. 58, pp. 1114–1128, 2016. [Online]. Available: http://www.sciencedirect.com/science/article/ pii/S1364032115017153 [7] A. Khaligh, Z. Peng, and Z. Cong, “Kinetic energy harvesting using piezoelectric and electromagnetic technologies—State of the art,” IEEE Trans. Ind. Electron., vol. 57, no. 3, pp. 850–860, Mar. 2010. [8] A. Kurs, A. Karalis, R. Moffatt, J. D. Joannopoulos, P. Fisher, and M. Soljacic, “Wireless power transfer via strongly coupled magnetic resonances,” Science, vol. 317, no. 5834, pp. 83–86, 2007. [9] S. Y. R. Hui, W. Zhong, and C. K. Lee, “A critical review of recent progress in mid-range wireless power transfer,” IEEE Trans. Power Electron., vol. 29, no. 9, pp. 4500–4511, Sep. 2014. [10] J. I. Agbinya, Wireless Power Transfer, vol. 45. Aalborg, Denmark: River Publishers, 2015. [11] W. Zhong and S. Y. R. Hui, “Charging time control of wireless power transfer systems without using mutual coupling information and wireless communication system,” IEEE Trans. Ind. Electron., vol. 64, no. 1, pp. 228–235, Jan. 2017. [12] L. Xie, Y. Shi, Y. T. Hou, and A. Lou, “Wireless power transfer and applications to sensor networks,” IEEE Wireless Commun., vol. 20, no. 4, pp. 140–145, Aug. 2013. [13] C. C. Huang, C. L. Lin, and Y. K. Wu, “Simultaneous wireless power/data transfer for electric vehicle charging,” IEEE Trans. Ind. Electron., vol. 64, no. 1, pp. 682–690, Jan. 2017. [14] W. C. Brown, “The history of power transmission by radio waves,” IEEE Trans. Microw. Theory Techn., vol. MTT-32, no. 9, pp. 1230–1242, Sep. 1984. [15] J. Zhang, X. Yuan, C. Wang, and Y. He, “Comparative analysis of two-coil and three-coil structures for wireless power transfer,” IEEE Trans. Power Electron., vol. 32, no. 1, pp. 341–352, Jan. 2017. [16] S. Assawaworrarit, X. Yu, and S. Fan, “Robust wireless power transfer using a nonlinear parity–time-symmetric circuit,” Nature, vol. 546, no. 7658, pp. 387–390, 2017. [17] A. P. Sample, D. T. Meyer, and J. R. Smith, “Analysis, experimental results, and range adaptation of magnetically coupled resonators for wireless power transfer,” IEEE Trans. Ind. Electron., vol. 58, no. 2, pp. 544–554, Feb. 2011. [18] X. Liu and G. Wang, “A novel wireless power transfer system with double intermediate resonant coils,” IEEE Trans. Ind. Electron., vol. 63, no. 4, pp. 2174–2180, Feb. 2016. [19] D. Ahn and S. Hong, “A study on magnetic field repeater in wireless power transfer,” IEEE Trans. Ind. Electron., vol. 60, no. 1, pp. 360–371, Apr. 2013. [20] B. L. Cannon, J. F. Hoburg, D. D. Stancil, and S. C. Goldstein, “Magnetic resonant coupling as a potential means for wireless power transfer to multiple small receivers,” IEEE Trans. Power Electron., vol. 24, no. 7, pp. 1819–1825, Jul. 2009. [21] Y. J. Kim, D. Ha, W. J. Chappell, and P. P. Irazoqui, “Selective wireless power transfer for smart power distribution in a miniature-sized multiplereceiver system,” IEEE Trans. Ind. Electron., vol. 63, no. 3, pp. 1853–1862, Mar. 2016.

6268


[22] E. Waffenschmidt, “Wireless power for mobile devices,” in Proc. IEEE 33rd Int. Telecommun. Energy Conf., pp. 1–9. [23] R. F. Xue, K. W. Cheng, and M. Je, “High-efficiency wireless power transfer for biomedical implants by optimal resonant load transformation,” IEEE Trans. Circuits Syst. I, Reg. Papers, vol. 60, no. 4, pp. 867–874, Apr. 2013. [24] Z. Bi, T. Kan, C. C. Mi, Y. Zhang, Z. Zhao, and G. A. Keoleian, “A review of wireless power transfer for electric vehicles: Prospects to enhance sustainable mobility,” Appl. Energy, vol. 179, pp. 413–425, 2016. [25] C. C. Mi, G. Buja, S. Y. Choi, and C. T. Rim, “Modern advances in wireless power transfer systems for roadway powered electric vehicles,” IEEE Trans. Ind. Electron., vol. 63, no. 10, pp. 6533–6545, Oct. 2016. [26] J. P. K. Sampath, A. Alphones, and D. M. Vilathgamuwa, “Figure of merit for the optimization of wireless power transfer system against misalignment tolerance,” IEEE Trans. Power Electron., vol. 32, no. 6, pp. 4359–4369, Jun. 2017. [27] T. Imura and Y. Hori, “Maximizing air gap and efficiency of magnetic resonant coupling for wireless power transfer using equivalent circuit and Neumann formula,” IEEE Trans. Ind. Electron., vol. 58, no. 10, pp. 4746–4752, Oct. 2011. [28] S. H. Lee and R. D. Lorenz, “Development and validation of model for 95%-efficiency 220-W wireless power transfer over a 30-cm air gap,” IEEE Trans. Ind. Appl., vol. 47, no. 6, pp. 2495–2504, Nov./Dec. 2011. [29] D. Liu, H. Hu, and S. V. Georgakopoulos, “Misalignment sensitivity of strongly coupled wireless power transfer systems,” IEEE Trans. Power Electron., vol. 32, no. 7, pp. 5509–5519, Jul. 2017. [30] A. Sample and J. R. Smith, “Experimental results with two wireless power transfer systems,” in Proc. IEEE Radio Wireless Symp., pp. 16–18. [31] J. S. Ho et al., “Wireless power transfer to deep-tissue microimplants,” Proc. Nat. Acad. Sci. USA, vol. 111, no. 22, pp. 7974–7979, 2014. [32] R. C. Vinod, M.-M. Jose Oscar, and P. A. David, “Wireless power transmission to an electromechanical receiver using low-frequency magnetic fields,” Smart Mater. Struct., vol. 21, no. 11, 2012, Art. no. 115017. [Online]. Available: http://stacks.iop.org/0964-1726/21/i=11/a=115017 [33] X. Lu, P. Wang, D. Niyato, D. I. Kim, and Z. Han, “Wireless charging technologies: Fundamentals, standards, and network applications,” IEEE Commun. Surveys Tut., vol. 18, no. 2, pp. 1413–1452, Apr./Jun. 2016. [34] N. P. Cook, S. Dominiak, and H. Widmer, “Wireless power transfer using magneto mechanical systems,” U.S. Patent 8 373 514, Feb. 12, 2013. [35] N. P. Cook, “Wireless delivery of power to a fixed-geometry power part,” U.S. Patent App. 12/115 478, May 5, 2008. [36] J. Lin et al., “ICNIRP guidelines for limiting exposure to time-varying electric and magnetic fields (1 Hz to 100 kHz),” Health Phys., vol. 99, pp. 818–836, 2010. [37] K. Agarwal, R. Jegadeesan, Y. X. Guo, and N. V. Thakor, “Wireless power transfer strategies for implantable bioelectronics: Methodological review,” IEEE Rev. Biomed. Eng., vol. 10, pp. 136–161, 2017. [38] M. G. L. Roes, M. A. M. Hendrix, and J. L. Duarte, “Contactless energy transfer through air by means of ultrasound,” in Proc. 37th Annu. Conf. IEEE Ind. Electron. Soc., 2011, pp. 1238–1243. [39] A. S. Rekhi, B. T. Khuri-Yakub, and A. Arbabian, “Wireless power transfer to millimeter-sized nodes using airborne ultrasound,” IEEE Trans. Ultrason., Ferroelect., Freq. Control, vol. 64, no. 10, pp. 1526–1541, Oct. 2017. [40] J. Charthad, M. J. Weber, T. C. Chang, and A. Arbabian, “A mm-sized implantable medical device (IMD) with ultrasonic power transfer and a hybrid bi-directional data link,” IEEE J. Solid-State Circuits, vol. 50, no. 8, pp. 1741–1753, Aug. 2015. [41] R. Cacak and J. Craig, “Magnetic field uniformity around near-Helmholtz coil configurations,” Rev. Sci. Instrum., vol. 40, no. 11, pp. 1468–1470, 1969. [42] J. C. Lin, “Wireless power transfer for mobile applications, and health effects,” IEEE Antennas Propag. Mag., vol. 55, no. 2, pp. 250–253, Apr. 2013. [43] H. Ulusan, O. Zorlu, A. Muhtaroglu, and H. Kulah, “Highly integrated 3 V supply electronics for electromagnetic energy harvesters with minimum 0.4 Vpeak input,” IEEE Trans. Ind. Electron., vol. 64, no. 7, pp. 5460–5467, Jul. 2017. [44] G. D. Szarka, S. G. Burrow, P. P. Proynov, and B. H. Stark, “Maximum power transfer tracking for ultralow-power electromagnetic energy harvesters,” IEEE Trans. Power Electron., vol. 29, no. 1, pp. 201–212, Jan. 2014. [45] R. Dayal, S. Dwari, and L. Parsa, “Design and implementation of a direct ac-dc boost converter for low-voltage energy harvesting,” IEEE Trans. Ind. Electron., vol. 58, no. 6, pp. 2387–2396, Jun. 2011.

[46] X. Wang, X. Liang, and H. Wei, “A study of electromagnetic vibration energy harvesters with different interface circuits,” Mech. Syst. Signal Process., vol. 58, pp. 376–398, 2015. [Online]. Available: http://www.sciencedirect.com/science/article/pii/S0888327014003987 [47] S. Du, Y. Jia, C. D. Do, and A. A. Seshia, “An efficient SSHI interface with increased input range for piezoelectric energy harvesting under variable conditions,” IEEE J. Solid-State Circuits, vol. 51, no. 11, pp. 2729–2742, Nov. 2016. [48] G. D. Szarka, S. G. Burrow, and B. H. Stark, “Ultralow power, fully autonomous boost rectifier for electromagnetic energy harvesters,” IEEE Trans. Power Electron., vol. 28, no. 7, pp. 3353–3362, Jul. 2013.

Sijun Du (S’14–M’17) received the B.Eng. degree in electrical engineering from University Pierre and Marie Curie, Paris, France, in 2011, the M.Sc. degree in electrical and electronics engineering from Imperial College, London, U.K., in 2012, and the Ph.D. degree in electrical engineering from the University of Cambridge, Cambridge, U.K., in 2017. He was with the Laboratory LIP6 of University Pierre Marie Curie, Paris, and then was a digital IC engineer in Shanghai between 2012 and 2014. He was with the Cambridge Nanoscience Centre, University of Cambridge, for his Ph.D. research between 2014 and 2017. He was an Engineer Intern with Qualcomm Technology, Inc., San Diego, CA, USA, between August 2016 and November 2016. His current research interests include energy harvesters and associated interfaces, power electronics, power management circuits, and rectification circuits.

Edward K. Chan (M’04) received the B.S., M.S., and Ph.D. degrees from Stanford University, Stanford, CA, USA, in 1994, 1996, and 1999, respectively, all in electrical engineering. Upon his graduation, he developed optical and wireless components with Bell Labs, Murray Hill, NJ, USA, in 2000. At Digilens, from 2001 to 2003, he engineered tunable liquid crystal components. Subsequently, he optimized the signal integrity and packaging of Nvidia (2003–2005) and Intel (2005–2010) microprocessors. He is currently the Director of Engineering with Qualcomm Research, San Diego, CA, USA, where he has been designing displays, imagers, sensors, and wireless power systems since 2010, focusing on the multiphysics integration and optimization of a wide range of technologies. He is an inventor of several patents for wireless, optical, and display devices. Dr. Chan was the recipient of the 1994 Henry Ford II Scholar Award at Stanford University.

Bing Wen received the Ph.D. degree in physics from Case Western Reserve University, Cleveland, OH, USA, in 2003. Between 2003 and 2011, he was a Research Scientist with Rockwell Science Center, developing a wide range of novel devices, including beam steering, wave front shaping, motion tracking, and special purpose cameras. In 2011, he joined Qualcomm, San Diego, CA, USA, to develop an microelectromechanical systems (MEMS)-on-glass based reflective display. He has authored or coauthored about 20 publications and holds 20 patents in related areas.


John Hong received the B.S. degree in electrical engineering from the Massachusetts Institute of Technology, Cambridge, MA, USA, in 1982, and the M.S. and Ph.D. degrees in electrical engineering from the California Institute of Technology, Pasadena, CA, USA, in 1983 and 1987, respectively. His dissertation work explored the applications of optics for information processing. He joined Jet Propulsion Laboratory to serve as the Chief Technologist for the Astronomy and Physics Directorate in 2002. In 2008, he joined Qualcomm, San Diego, CA, USA, where he initially worked on computer vision and signal processing efforts. He also led the effort to produce a second-generation microelectromechanical systems (MEMS) display technology involving the integration of MEMS and indium gallium zinc oxide thin-film transistor (IGZO TFT) on glass. Dr. Hong was awarded the SPIE Rudolf Kingslake Award in 1989 and the Rockwell Engineer of the Year Award in 1997 for his work with Rockwell International Science Center, intitially to develop photorefractive nonlinear optics for information storage and processing.

Hanspeter Widmer received the M.Sc. and Ph.D. degrees in electrical engineering from the Federal Institute of Technology (ETH), Zurich, Switzerland, in 1979 and 1991, respectively. He was an Assistant with the Microwave Laboratory, ETH, teaching electrical engineering, transmission line theory, and microwave techniques. In 1984, he started research activities in the area of robust and energy-efficient communication techniques and information theory. In 1988, he started as a Research Engineer with Ascom, Switzerland, where he was leading several research projects in mobile terrestrial and satellite communications as well as in broadband powerline communications. In 2007, he started new research activities in the area of wireless energy transmission. Since 2011, he has been with Qualcomm, Inc., San Diego, CA, USA, as a Principal Engineer working for the Qualcomm Halo wireless electric vehicle charging R&D program.

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Charles (Chuck) E. Wheatley (LF’13) received the B.S. degree in physics from the California Institute of Technology, Pasadena, CA, USA, in 1956, the M.S. degree in electrical engineering from the University of Southern California, Los Angeles, CA, USA, in 1958, and the Ph.D. degree in electrical engineering from the University of California, Los Angeles, CA, USA, in 1972. He has been involved with spread spectrum concepts, including both frequency hopping and direct sequence techniques, for more than 60 years. At Rockwell International, from 1973 to 1981, he contributed to the global positioning system (GPS) on various terminal designs and was also responsible for the design and performance of the GPS spaceborne Rubidium Frequency Standard. From 1982 to 1989, he was with Linkabit Corporation, where he contributed to RF and system design aspects of several military communication systems. He joined Qualcomm, Inc., in 1989, where he is a Senior Vice-President-Technology and Qualcomm Fellow. At Qualcomm, he was primarily involved with various aspects of spread spectrum technology as applied to cellular/personal communications. Starting with its inception in 1989, he contributed to the successful commercial deployment in USA, China, and India. Since then, the technology has evolved through ever increasingly capable versions known as 2G, 3G, 4G, and most recently as LTE. He retired from Qualcomm in 2007, returning in 2008 to work on wireless charging for a variety of devices ranging from portable electronic devices to electric vehicles. Most recently, he consults on the development of large-scale high-speed satellite communication system where Qualcomm acts in a system engineering role. He holds more than 100 patents on various techniques and devices, mostly related to the fields of communications and navigation. He has authored and coauthored more than a dozen articles in various journals and symposiums. Dr. Wheatley is a Caltech Distinguished Alumni, being recognized for his contributions to GPS and cellular communications. His paper “On the Capacity of a Cellular CDMA system” (with coauthors Klein Gilhousen, Irwin Jacobs, Roberto Padovani, Andrew Viterbi, and Lindsay Weaver) in the IEEE TRANSACTIONS ON VEHICULAR TECHNOLOGY, May 1991, was voted 1991 Best Paper Award of the year by the Vehicular Technology Society.