Design, Simulation, and Characterization of a Low-Cost ... - IEEE Xplore

3240

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 60, NO. 8, AUGUST 2013

Design, Simulation, and Characterization of a Low-Cost In-Plane Spark Gap Microswitch With Dual-Trigger Electrode for Pulsed Power Applications Chang Liu, Zhuoqing Yang, Member, IEEE, Guifu Ding, Zhenwei Zhou, Qifa Liu, Yi Huang, and Yu Zheng

Abstract—An in-plane spark gap switch with dual-trigger electrode for pulsed power application has been designed and fabricated based on surface micromachining technology. The switch consists of two main electrodes and a dual-trigger electrode to achieve high peak current and smaller trigger voltage. The simulation result shows that 100-V trigger voltage could break down 30-μm distance and then start up the switch. The main discharge channel length has been set as 800 μm to get the optimized output of rise time combined with peak current. The prototype microswitch has been fabricated and characterized. The peak and rising edge of the discharge current are 3450.75 A and 164 ns, respectively, which well agree with simulation values. Index Terms—Microswitch, particle-in-cell (PIC) simulation, pulsed power, RLC model, spark gap, surface micromachining.

I. I NTRODUCTION

P

ULSED POWER is the scheme by which the accumulated electrical energy can be rapidly released during a short time [1]. Pulsed power systems can generate high power pulses with an ultrafast rise time [2], [3]. It has a variety of applications in material science and plasma physics [4]. Usually, a capacitive pulsed power system consists of four parts: charging unit, energy storage device, closing switch, and load. Among them, the switch is the most important part which can operate at high speed and determines the shape of the output waveform directly. Several kinds of closing switch are designed and fabricated for pulsed power applications, such as pseudospark switch, insulated gate bipolar transistor (IGBT), thyristor, and spark gap switch [5]–[9]. Of these, the spark gap switch is the most commonly used one for its robust, compact, and fast response. It can be used in functioning electroexplosive devices, in the construction of a pulse igniter, and so on [10]–[13]. The spark gap characteristics have been studied extensively both theoretically and experimentally. Guenther and Bettis gave a detailed Manuscript received November 9, 2011; revised February 17, 2012 and April 6, 2012; accepted May 23, 2012. Date of publication June 4, 2012; date of current version April 11, 2013. The authors are with the National Key Laboratory of Science and Technology on Micro/Nano Fabrication, Research Institute of Micro/Nano Science and Technology, Shanghai Jiao Tong University, Shanghai 200240, China (e-mail: [email protected]; [email protected]; [email protected]; [email protected]; [email protected]; [email protected]; [email protected]). 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.2012.2202356

review of the research on spark gap up to 1978 [14]. Several models have been developed to describe spark gap characters: Kushner et al. modeled the arc resistance theoretically [15]. Persephonis et al. gave an induction model for the spark gap discharge [16]. As to experiments, many designs of the spark gap configurations have been raised. For example, Tang and Scuka have built a kind of spark gap with a midplane trigatron prototype [17]. The standoff voltage could reach several kilovolts, and its discharge time is less than a microsecond. However, as a switch based on batch fabrication, piece-part assembly is a problem which could result in imprecision and high cost. In this paper, a spark gap switch that comprises an anode, a cathode, and a trigger electrode which can be monolithically integrated on the same substrate was proposed. In particular, a dual-trigger electrode configuration was used to realize low trigger voltage and in-plane discharge. The test result of the prototype shows satisfactory electrical behavior. II. D EVICE D ESIGN The design guideline of pulsed power microswitch based on microelectromechanical systems (MEMS) technology consists of the conventional spark gap. For conventional electric-triggered spark gap configuration, three electrodes are necessary: where one of them serves as the ground electrode, one of the other is the trigger electrode, and the third is input. Referring to the requirements of the small-scale pulsed power system, the performance parameters should be as follows [18]. 1) The standoff voltage of the switch should be no less than 1000 V. 2) The switch should be triggered at a relatively low voltage signal (e.g., 100 V). 3) The switch should satisfy fast discharge with nanosecond rise time (e.g., 100 ns). 4) The peak output current should be more than 1500 A. The switch designed in this paper was with regard to the MEMS surface micromachining technology. It can overcome the shortcomings of the conventional ones, like high processing cost and large volume which would impede integration in small systems, and by utilizing microfabrication technology, the discharge length can be controlled within the micrometer scale, which can effectively reduce the trigger voltage. Fig. 1 shows the sketch and parameters of the designed spark gap switch. We

0278-0046/$31.00 © 2012 IEEE

LIU et al.: DESIGN, SIMULATION, AND CHARACTERIZATION OF IN-PLANE SPARK GAP MICROSWITCH

Fig. 1. (a) Schematic representation of the spark gap switch and (b) its main parameters.

3241

Fig. 2. Electrostatic field distributions of the discharge space (a) before actuation (trigger electrode connects ground) and (b) after actuation (100-V trigger voltage applied).

TABLE I S TRUCTURAL PARAMETERS OF THE D EVICE (U NIT: M ICROMETER )

are aiming that the main discharge channel can get 1000 V bearable and that 100-V trigger voltage could realize switch closing. The structure parameters are shown in Table I. The operation voltage of the trielectrode spark gap should satisfy [1] 1.5Ug ≤ Ub ≤ 3Ug

(1)

where Ub is the gap breakdown bias and Ug is the designed standoff voltage. The breakdown voltage of the air gap in quasiuniform electric field (E-field) obeys [1] √ (6.72 pd + 24.36pd) . (2) Ub = δ where Ub is in kilovolts, p is the air pressure in bars, d is the gap length in centimeters, and δ is the field enhancement factor. Combining (1) and (2), we could get the designed main gap range 572 μm ≤ d ≤ 1421 μm.

(3)

Such spacing is precisely in the scope of what the traditional process is difficult to achieve but micromachining can easily realize. In our design configuration, we set the gap distance as 800 μm. We set the two main electrodes’ diameters (R) and dual-trigger separation (b) as 4000 μm and 1000 μm, respectively, ensuring that the E-field is nearly uniformly distributed to minimize E-field enhancement and extend device life. We set the trigger electrode in an offset position “a” relative to the center of the left main spark gap, which is of 30 μm [see Fig. 1(b)]. This design aims to minimize the triggered arc length (limited by the processing accuracy, which is 30 μm) in order to maintain the lifetime of the trigger electrode when satisfying the breakdown requirement. The trigger electrode is set as a dual electrode which is placed symmetrically and perpendicularly to the main discharge channel. Therefore, the E-field distribution in the main spark gap is also symmetric. When the switch is triggered, the movement direction of the generated plasma is along the main discharge channel. As a result, the high current requirement could be satisfied. Fig. 2(a) shows the electrostatic

Fig. 3. Operation circuit of the spark gap switch.

field distribution before triggering, when the voltage difference between the two main electrodes is 1000 V, and it can be seen that the electrostatic field distribution of the main gap is nearly uniform (≈1.43 kV/mm) and the maximum field strength is in the vicinity of the trigger electrode, which is about 2.1 kV/mm smaller than the common breakdown electrofield threshold (≈3 kV/mm). When the triggering voltage is applied onto the triggering gap, it can be seen from Fig. 2(b) that the electrostatic field distribution between the left main electrode and the trigger electrode is symmetric. The maximum E-field strength is 5.76 kV/mm, which exceeds the breakdown bias; therefore, the discharge between the trigger electrode and left main electrode will take place, and the generated plasma will move along the main discharge channel and finally actuate the switch. A diagram of the operation circuit is shown in Fig. 3. The trigger electrode connects the positive terminal of the 2.2-mF low-voltage capacitor through the switch K and serves as the trigger terminal of the device. The main electrode (on the right side) makes connection with the positive terminal of the 0.47-μF high-voltage capacitor. The negative terminal of capacitors C1 and C2 and the left main electrode are all connected to ground. The operation of the designed switch is described as follows: Firstly, the high-voltage capacitor C1 and low-voltage capacitor C2 are charged to their rated voltages, respectively. The device will not be actuated until switch K closes. When K closes, the voltage across the trigger electrode and left main electrode exceeds the breakdown voltage bias of the air gap between them, so the device is triggered. The plasma generated in the triggering gap would spread into the main spark gap and cause further gas ionization, which would finally cause the main gap breakdown, and the switch gives out a strong current pulse.

3242

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 60, NO. 8, AUGUST 2013

Fig. 4. Phase space of electrons and ions in main discharge channel [i) is for electrons, and ii) is for nitrogen ions]. (a) t = 0ns, (b) t = 1 ns, (c) t = 20 ns, and (d) t = 25 ns.

III. S IMULATION AND S YSTEM M ODELING The designed switch operates at the arc discharge model [19], for which electrons and ions get exponential increase and the secondary electrons generated by the impact between the charged particle and metal electrode may play an important role. Therefore, for a detailed interpretation, the discharge development was simulated by using a particle-in-cell (PIC) model that runs on the 2-D code OOPIC Pro. OOPIC Pro is a PIC physics simulation code for 2-D (x, y)/(r, z) geometries with 3-D electrostatic and electromagnetic field solvers and Monte Carlo collision and ionization models [20]. To study the breakdown process in a simple way, we set the simulation space as a scale-down version compared to the actual designed switch: a rectangular shape with a height of 4000 μm and a width of 800 μm. As can be seen from Fig. 4, the two trigger electrodes were modeled in the simulation region and were set as conductor with an applied triggering voltage of 100 V. The top and bottom boundaries of the simulation space were modeled as dielectrics with a relative permittivity of one. Particles which strike the two boundaries are absorbed, and the charge would not be accumulated. The right boundary was modeled as an equipotent boundary with a dc bias of 1000 V, and the left one was modeled as a grounded conductor. Secondary electron emission due to ion impact was added on the left boundary and the trigger electrodes with emission + = 0.1. Nitrogen at room temperature filled the coefficient γN2 simulation space at 1 atm. The elastic, excitation, and ionization cross section of N2 are considered. The simulation was run with a time step of 1 ps and grid dimensions of 512 × 1028 to guarantee that particles do not cross one grid during one time step. The simulation conditions were as follows: The initial electrons were set near the trigger electrodes with a number density of 1019 /m3 to present seed electrons after triggering. Monte Carlo collision method was used for the electron-neutral and ion-neutral collisions. The discharge includes electronneutral elastic, excitation, and ionization collisions, as well as ion-neutral charge-exchange collisions. Fig. 4 shows the process of breakdown in the main discharge channel. i) is the phase space of electron, and ii) is that of nitrogen ions. First, initial seed electrons produced by triggering are distributed near the trigger electrode [see Fig. 4(a)].

Fig. 5. Particle number evolution of electrons and ions in the initial discharge process (blue line is for ions, and orange line is for electrons).

Second, electrons were swept away by the background E-field, leaving a positively charged cloud of relatively slow moving ions [see Fig. 4(b)]. The ionization of nitrogen generates N+ 2 and electron. Then, N+ 2 ions move toward the left boundary. Secondary electrons are released from the left main electrode surface by positive nitrogen ions falling on to the latter, which results in a nearly exponential increase of the numbers of positive and negative carriers (see Fig. 5). From Fig. 4(c), we could find that, at a time of 20 ns, a mass of plasma appears near the right boundary, with the number density of electron and nitrogen raised to 1019 /m3 . The size and density of plasma get higher and higher. At 25 ns [see Fig. 4(d)], the plasma spot extends to the entire discharge space, causing the main discharge to take place. Fig. 6 shows the E-field during the whole discharge process. At t = 0, the E-field of the whole discharge area nearly uniformly distributes before the formation of plasma, and electric potential distribution is mainly determined by the structure of the switch. When numbers of particles get exponentially increased, the E-field in the discharge domain is affected by the space charge E-field. The E-field along the main discharge channel gets quickly enhanced until the switch closed. The peak value and rise time of the pulse current rely on structural and environmental parameters. As simulation indicated, the time to form plasma (electron density near the

LIU et al.: DESIGN, SIMULATION, AND CHARACTERIZATION OF IN-PLANE SPARK GAP MICROSWITCH

3243

Fig. 8. System modeling of the test circuit.

Fig. 6. Time evolution of E-field. (a) t = 0 ns, (b) t = 10 ns, (c) t = 20 ns, and (d) t = 25 ns.

Fig. 7. Variation in the peak electron current and time to form plasma with (a) discharge channel length d, where I0 is the peak current for d = 400 nm, and (b) gas pressure P , where I0 is the peak current for P = 0.4 atm.

anode gets to 1019 /m3 ) and the peak current vary with the discharge channel length and air pressure. From Fig. 7(a), we could find that plasma formation time gets shorter as discharge distance increases. This is mainly because electrons can get more collisions with background gas to generate more electrons and ions, which will accelerate the procedure to form plasma. From Fig. 7(a), one can also estimate that electron peak current varies with discharge channel length, first increasing and then reducing. This is because, although electrons would collide

more times to rise the peak current when the channel is longer, they would also move slower for the E-field to become weak when the voltage difference is unchanged. In addition, some environmental factors may also affect the actuation of the microswitch, such as air pressure. In our simulation, we have changed the air pressure from 0.4 to 1.6 atm (considered device packaging and testing needs). Fig. 7(b) shows that peak current drops slightly and plasma formation gets slower as air pressure increased. This is because, when applied E-field is fixed, with the increase in pressure, the number of free electrons generated by the electronic collisions with neutral gas per unit distance (namely, the ionization coefficient α, which is the function of E/P [21]) reduced, which leads plasma discharge channel formation to slow down and peak current to weaken. We should note that other factors may also affect the device performance, such as humidity and temperature (which can affect the initial seed electron density), but through packaging process, these environmental factors can be controlled in small fluctuations to diminish their impact on device performance. When particle numbers get higher and higher, PIC simulation would become computationally resource consuming and may encounter numerical divergence. Therefore, a circuit model was built to compute output waveform theoretically. As seen in Fig. 8, the complete test circuit with the switch can be modeled as a simple RLC circuit. The capacitor in the diagram represents the high-voltage capacitor used in the test circuit. The resistance in circuit R = Rs + Rv , Rs , is the sample resistor (0.01 Ω), and the current is measured through voltage detection across Rs . The spark gap switch is modeled as a time-varied resistance Rv in series of an inductance L. Circuit equations are as follows: dU0 (t) (4) i(t) = − C dt di Uc (t) = L + [Rs + Rv (t)] i. (5) dt A lot of empirical formulas about spark gap resistance and inductance have been developed [22], [23]. We choose the Rompe–Wiezel formula as the variable resistance’s description for our testing conditions is similar (1 atm, d < 0.035 m), and the equivalent inductance is the function of the spark channel length. The two expressions are as follows: d 1 2 0 i (τ )dτ b L = 2d ln a

Rv = 

(6)

2α p

(7)

3244

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 60, NO. 8, AUGUST 2013

where d is the main discharge length, α is the gas constant, p is the pressure in atmospheres, b is the radius of the main electrode, and a is the radius of the spark channel. The differential of (2) is α dRs = − 2 Rs3 (t)i2 (t). dt pd

(8)

Solve (4), (5), and (8) by using numerical methods, and the constants are as follows: C = 0.47 μF, p = 0.1 MPa, α = 0.09 MPa. cm2. V−2. s−1 (the range of the gas constant is 0.08 ∼ 0.1 MPa. cm2. V−2. s−1 ), d = 800μm, and L = 11.2 nH [from (7)]. The initial condition is U c(0) = 1000, i(0) = 0, and r(0) → ∞. The plot in Fig. 9 shows the U c − t and the I − t curve. From Fig. 9, we could find that the I − t curve is an oscillation decay curve, and the peak current and rise time are 4126.2 A and 151.2 ns, respectively, which should be confirmed and compared with the test results. IV. M ICROFABRICATION The designed switch was fabricated on a glass wafer by convenient and low-cost metal electroplating technology. The compositions of nickel electrolyte are Ni[NH2 SO3 ]2 (600 g/L), H3 BO3 (25 g/L), and NiCl2 • 6H2 O (10 g/L). The electroplating conditions are pH 4.0, a temperature at 40 ◦ C, and a current density of 2.0 A/dm2 . The compositions of gold electrolyte are KAu(CN)2 (5 g/L), C8 H4 K2 O12 Sb2 · 3 H2 O (0.1 g/L), and C6 H14 N2 O7 (110 g/L). The electroplating conditions are pH 5.5, a temperature at 40 ◦ C, and a current density of 0.1 A/dm2 . The height of the electrodes is controlled by the electroplating time and monitored by a stylus profiler (Dektak 6M, VECCO, USA). The main fabrication process steps of the spark gap switch are described and sketched in Fig. 10 as follows. (a) A thin Kapton film was spin coated on the glass wafer as insulation substrate. (b) Chromium/copper (Cr/Cu) was sputtered as the seed layer for the plating of the device structure. (c) Then, photoresist was spin coated, and photolithography was completed. (d) The electrodes were electroplated in nickel electrolyte. (e) A thin Au film was electroplated then. (f) The photoresist was dissolved in 15% NaOH solution, and the seed layer was dissolved in ammonia and potassium ferricyanide solution. The released undiced switch was dried on a hot plate in air and then cut into a single device. The optical photograph of the fabricated switch is shown in Fig. 10(g). The dimension of a single switch is 9 mm × 7 mm. The fabricated switch can be used as a trigger switch of an intense electron-beam accelerator [24]. The short closure time and peak current can help to rapidly release energy stored in a primary storage capacitor and provide the initial energy of the accelerated electrons. Its planar scheme and minimized size could be easier to integrate. It can also be utilized in other pulsed power applications such as high intensity discharge (HID) lamps, where the switch can be connected into the starting circuit that applies the starting pulses when the lamp is OFF [10], [11], [13]. Monolithic integration results in compact size,

Fig. 9. Equivalent circuit simulation result. (a) U c − t curve. (b) I − t curve. (c) Peak current and rise time.

Fig. 10. (a)–(f) Process sketch for the fabrication of the microswitch and (g) optical photograph of the switch.

LIU et al.: DESIGN, SIMULATION, AND CHARACTERIZATION OF IN-PLANE SPARK GAP MICROSWITCH

3245

Fig. 11. Sketch of the test setup for the fabricated microspark gap switch.

as well as more precisely high voltage bearable and fast turning on; these will make its applications to be more dominant. V. C HARACTERIZATION The switch was characterized as shown in Fig. 11. The resistance in the circuit should be as small as possible to ensure the shortest rise time, so the sample resistor is chosen as 0.01 Ω, and all interconnects in this test circuit are made with widened copper strip line to reduce parasitic impedance. To measure the current through the switch, the voltage across the sample resistor was measured by using a multichannel oscilloscope (MSO6304A, Agilent), and the current can be calculated by I = V /R. Before the test, the high-voltage capacitor is charged to 1000 V, and the initial voltage is entirely dropped across the main channel, but discharge will not take place as it is lower than the breakdown threshold. Upon triggering, the plasma generated between the trigger and left main electrode can be induced to flow through the main discharge path and release an ultrashort current pulse. As can be seen from Fig. 12, the measured peak current and rise time are labeled in the plot. The peak current through the switch is 3450.75 A, and the rise time is about 164 ns. The test result of our design is compared with the conventional spark gap scheme [8]. The present design has much lower trigger voltage but can get the same short rise time. The test results generally agree with the aforementioned simulated values. Both of the simulated waveform and test results are damped oscillation curves, and the measured peak current is negative because, in the actual test, we have used negative voltage power supply. It can be found that the test peak current is a little lower than theoretical results. It is mainly due to the impact of parasitic resistance in the testing circuit. Similarly, the test rise time is longer than the calculated results because of the circuit inductance and signal delay of transmission line. Moreover, we may also find that the test curve has a relatively weakened damped oscillation compared with the simulated one. This situation attributes to that only resistance varying with time is considered in the system model. However, the resistance factually plays an essential role in the rising process (air

Fig. 12. Test I − t curve of the switch. (a) I − t curve. (b) Peak current and rise time.

gap turn from insulator to conductor). In addition, during the discharge process, the arc resistor gets lower when the contact resistance and parasitic inductance may impact the decay curve to a large extent. Fortunately, the rising time and peak current in the present switch are our focus. In this point, the simulated results have a good agreement with the test ones. In the next work, we will continue to optimize the configuration of the switch and its testing circuit in order to decrease the effect from inductance and other related parameters. VI. C ONCLUSION An in-plane spark gap switch with dual-trigger electrode for pulsed power application is proposed and simulated. The simulation results show that the standoff voltage of the designed switch can reach 1000 V, while it can be triggered at a lower voltage of ∼100 V. The 800 μm is considered as the main discharge channel length based on the optimization. The spark microswitch is successfully fabricated on a glass wafer by surface micromachining, and the obtained prototype is 9 mm × 7 mm. The fabricated microswitch has been tested and characterized utilizing the designed printed circuit board. The test results show that the peak current and the rise time are about 3450.75 A and 164 ns, respectively, which are generally in agreement with the simulated ones. The developed spark microswitch is promising in several pulsed power applications.

3246

IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 60, NO. 8, AUGUST 2013

R EFERENCES [1] H. Bluhm, Pulsed Power Systems: Principles and Applications. Berlin, Germany: Springer-Verlag, 2006. [2] H. Bai and Z. M. Zhao, “Framework and research methodology of short timescale plused power phenomena in high-voltage and high-power converters,” IEEE Trans. Ind. Electron., vol. 56, no. 3, pp. 805–816, Mar. 2009. [3] R. M. Nelms, “A capacitor charging power supply utilizing a Ward converter,” IEEE Trans. Ind. Electron., vol. 39, no. 5, pp. 421–428, Oct. 1992. [4] T. A. Baginski, “A high voltage single shot switch implemented with a MOSFET current source and an avalanche diode,” IEEE Trans. Ind. Electron., vol. 44, no. 2, pp. 167–172, Apr. 1997. [5] K. Frank, C. Bickes, U. Ernst, M. Iberler, J. Meier, U. Prucker, M. Schlaug, J. Schwab, J. Urban, and D. H. H. Hoffmann, “Low-pressure pseudospark switches for ICF pulsed power,” Nucl. Instrum. Methods Phys. Res. A, Accel., Spectr., Detect. Assoc. Equip., vol. 415, no. 1/2, pp. 327–331, Sep. 1998. [6] M. Vellveh, D. Flores, X. Jord, S. Hidalgo, R. Rebolloa, L. Coulbeckb, and P. Waindb, “Design considerations for 6.5 kV IGBT devices,” Microelectron. J., vol. 35, no. 3, pp. 269–275, Mar. 2004. [7] L. K. Tully, E. S. Fulkerson, D. A. Goerz, and R. D. Speer, “Evaluation of light-triggered thyristors for pulsed power applications,” in Proc. IEEE Int. Power Modulators High Voltage Conf., Las Vegas, NV, 2008, pp. 1–4. [8] G. J. J. Winands, Z. Liu, A. J. M. Pemen, E. J. M. van Heesch, and K. Yan, “Long lifetime, triggered, spark-gap switch for repetitive pulsed power application,” Rev. Sci. Instrum., vol. 76, no. 8, pp. 085107-1–085107-6, Aug. 2005. [9] V. Esteve, E. Sanchis-Kilders, J. Jordan, and A. Ferreres, “Improving the efficiency of IGBT series-resonant inverters using pulse density modulation,” IEEE Trans. Ind. Electron., vol. 58, no. 3, pp. 979–987, Mar. 2011. [10] F. L. Tomm, A. Raniere Seidel, A. Campos, M. A. Dalla Costa, and R. N. do Prado, “HID lamp electronic ballast based on chopper converters,” IEEE Trans. Ind. Electron., vol. 59, no. 4, pp. 1799–1807, Apr. 2012. [11] H. J. Chiu, C. H. Li, and Y. K. Lo, “Design and implementation of a single-stage high-frequency HID lamp electronic ballast,” IEEE Trans. Ind. Electron., vol. 55, no. 2, pp. 674–683, Feb. 2008. [12] S. Yu, T. Tsai, F. Wu, and M. W. Wu, “Bipolar narrow-pulse generator with energy-recovery feature for liquid-food sterilization,” IEEE Trans. Ind. Electron., vol. 55, no. 1, pp. 123–132, Jan. 2008. [13] M. Costa and D. G. Vaquero, “Microcontroller-based high power factor electronic ballast to supply metal halide lamps,” IEEE Trans. Ind. Electron., vol. 59, no. 4, pp. 1779–1788, Apr. 2012. [14] A. H. Guenther and J. R. Bettis, “The laser triggering of high-voltage switches,” J. Phys. D, Appl. Phys., vol. 11, no. 12, pp. 1577–1580, Jun. 1978. [15] M. J. Kushner, W. D. Kimura, and S. R. Byron, “Arc resistance of lasertriggered spark gaps,” J. Appl. Phys., vol. 58, no. 5, pp. 1744–1751, Sep. 2009. [16] P. Persephonis, K. Vlachos, and C. Georgiades and Parthenios, “The inductance of the discharge in a spark gap,” J. Appl. Phys., vol. 71, no. 10, pp. 4755–4762, May 1992. [17] H. Tang and V. Dcuka, “The breakdown mechanism of a mid-plane triggered spark gap trigatron,” IEEE Trans. Dielect. Elect. Insul., vol. 3, no. 6, pp. 843–848, Dec. 1996. [18] T. A. Baginski and K. A. Thomas, “A robust one-shot switch for highpower pulse applications,” IEEE Trans. Power Electron., vol. 24, no. 1, pp. 253–259, Jan. 2009. [19] Y. D. Korolev and N. M. Bykov, “High-voltage spark gap in a regime of subnanosecond switching,” IEEE Trans. Plasma Sci., Dec. 20, 2011, to be published. [20] J. P. Verboncoeur, A. B. Langden, and N. T. Gladd, “An object-oriented electromagnetic PIC code,” Comput. Phys. Commun., vol. 87, no. 1/2, pp. 199–211, May 1995. [21] A. J. Davies, C. J. Evans, P. Townsend, and P. M. Woodison, “Computation of axial and radial development of discharges between plane parallel electrodes,” Proc. Inst. Elect. Eng., vol. 124, no. 2, pp. 179–182, Feb. 1977. [22] V. R. Rompe and W. Weizel, “Ueber das toeplersche funkengesetz,” Eur. Phys. J. A, vol. 122, no. 9–12, pp. 636–637, Sep. 1944. [23] T. G. Engel, A. L. Donaldson, and M. Kristiansen, “The pulsed discharge arc resistance and its functional behavior,” IEEE Trans. Plasma Sci., vol. 17, no. 2, pp. 323–329, Apr. 1989. [24] X. B. Cheng, J. L. Liu, B. L. Qian, Z. Chen, and J. H. Feng, “Research of a high-current repetitive triggered spark-gap switch and its applications,” IEEE Trans. Plasma Sci., vol. 38, no. 3, pp. 516–522, Mar. 2010. [25] G. A. Mesyats, S. D. Korovin, V. V. Rostov, V. G. Shpak, and M. I. Yalandin, “The RADAN series of compact pulsed power Generators and their applications,” Proc. IEEE, vol. 92, no. 7, pp. 1166–1179, Jul. 2004.

Chang Liu received the B.Sc. degree in applied physics from Shanghai Jiao Tong University, Shanghai, China, in 2010, where he is currently working toward the M.Sc. degree in the Research Institute of Micro/Nano Science and Technology. He is also currently with the National Key Laboratory of Science and Technology on Micro/Nano Fabrication, Shanghai Jiao Tong University. His research interests include the design, simulation, and fabrication of the MEMS devices.

Zhuoqing Yang (M’12) received the B.Sc. and M.Sc. degrees in electromechanical engineering from Harbin Engineering University, Harbin, China, in 2003 and 2005, respectively, and the Ph.D. degree in microelectronics and solid-state electronics from Shanghai Jiao Tong University, Shanghai, China, in 2010. He is currently an Assistant Professor with the Research Institute of Micro/Nano Science and Technology, Shanghai Jiao Tong University. He is also working as a Japan Society for the Promotion of Science Postdoctoral Fellow in the National Institute of Advanced Industrial Science and Technology in Japan. He is also currently with the National Key Laboratory of Science and Technology on Micro/Nano Fabrication, the Key Laboratory for Thin Film and Microfabrication of Ministry of Education, Shanghai Jiao Tong University. His research interests include the design, simulation, and fabrications of the MEMS/nano-electro-mechanical systems (NEMS) devices.

Guifu Ding received the B.Sc. and M.Sc. degrees from Fudan University, Shanghai, China, in 1984 and 1987, respectively. He is currently a Professor and the Vice-Director of the National Key Laboratory of Science and Technology on Micro/Nano Fabrication, Research Institute of Micro/Nano Science and Technology, Shanghai Jiao Tong University. He is also currently with the Key Laboratory for Thin Film and Microfabrication Technology of Ministry of Education, the National Key Laboratory of Science and Technology on Micro/Nano Fabrication, Shanghai Jiao Tong University. His main research interests include the nanomaterials and the design, simulation, and fabrications of the MEMS/NEMS devices.

Zhenwei Zhou received the B.Sc. and M.Sc. degrees in applied physics and micro and solid electronic engineering from Shanghai Jiao Tong University, Shanghai, China, in 2007 and 2010, respectively. He is currently with the National Key Laboratory of Science and Technology on Micro/Nano Fabrication, Research Institute of Micro/Nano Science and Technology, Shanghai Jiao Tong University. His research interests include the design and simulation of pulsed power devices and the 3-D microfabrication of nonsilicon materials.

Qifa Liu received the B.Sc. degree in chemical engineering and technology from Lu Dong University, Yantai, China, in 2005 and the M.Sc. degree in chemical engineering and technology from Chang Zhou University, Changzhou, China, in 2008. He is currently working toward the Ph.D. degree in the Research Institute of Micro/Nano Science and Technology, Shanghai Jiao Tong University, Shanghai, China. He is also currently with the Science and Technology on Micro/Nano Fabrication Laboratory, Shanghai Jiao Tong University. His research interests include the design, simulation, and fabrication of microelectronic devices.

LIU et al.: DESIGN, SIMULATION, AND CHARACTERIZATION OF IN-PLANE SPARK GAP MICROSWITCH

Yi Huang received the B.Sc. degree in electronics engineering from Shanghai Jiao Tong University, Shanghai, China, in 2010, where he is currently working toward the M.Sc. degree in the Research Institute of Micro/Nano Science and Technology. He is also currently with the Science and Technology on Micro/Nano Fabrication Laboratory, Shanghai Jiao Tong University. His research interests include the design, simulation, and fabrication of microrelay devices.

3247

Yu Zheng received the B.Sc. degree in microelectronic engineering from Soochow University, Suzhou, China, in 2010. She is currently working toward the M.Sc. degree in the Research Institute of Micro/Nano Science and Technology, Shanghai Jiao Tong University, Shanghai, China. She is also currently with the Science and Technology on Micro/Nano Fabrication Laboratory, Shanghai Jiao Tong University. Her research interests include MEMS devices and nonsilicon fabrication technology.