Comparative measurement of surge arrester residual voltages by D-dot probes and dividers

Electric Power Systems Research 81 (2011) 1274–1282

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Comparative measurement of surge arrester residual voltages by D-dot probes and dividers I.A. Metwally ∗ Department of Electrical & Computer Engineering, College of Engineering, Sultan Qaboos University, P.O. 33, Al-Khod, Muscat 123, Oman

a r t i c l e

i n f o

Article history: Received 10 November 2010 Received in revised form 16 January 2011 Accepted 24 January 2011 Available online 5 March 2011 Keywords: Arresters Charge coupled devices Impulse testing Pulse generation Voltage dividers Voltage measurement

a b s t r a c t This paper presents the design and testing of coaxial D-dot probes with three different designs, namely, totally closed, totally open and semi-open designs. The latter design gives a good response and has an easy-access assembly for fixing and replacing the surge arresters (SA) under test. Simultaneous measurements of residual voltage for 9-kV SA are conducted by the designed probes, and resistive and universal dividers (RD and MRCD). The divider compensation principle is also introduced and implemented for both dividers. Both perfectly compensated dividers and the designed D-dot probes have an error of about 2.5% in the SA residual voltage due to the unavoidable resistance of connecting leads. SA residual voltages measured by uncompensated short MCRD divider give good agreement with that measured by the D-dot probe. The former has small inductive overshoot which can be neglected as the residual voltage crests before the discharge current by a few microseconds. © 2011 Elsevier B.V. All rights reserved.

1. Introduction Surge protection has been significantly improved since the introduction of zinc oxide (ZnO) material in 1967. However, lightning continues to be the principal cause of inadvertent outages and equipment damage in power systems [1]. ZnO SA has a highly nonlinear current–voltage (I–V) characteristic, which obviates the need for series spark gaps. Consequently, the electrical characteristics are solely determined by the properties of the ZnO blocks. Higher voltage and energy ratings are achieved by adding ZnO blocks in series and by using larger diameter discs or parallel columns of discs, respectively. ZnO material has a temperature dependence that is evident only at low-current densities. The temperature-dependent I–V characteristic is important only for the evaluation of energy absorbed by the SA and should not influence the insulation protective margins [2]. The I–V characteristic depends upon the discharge current waveforms, where faster current rise times result in higher peak voltages [2]. For currents with times to crest in the range of 0.5–4 ␮s or less, any stray inductance in the measurement circuit can result in a higher SA residual voltage than it would be actually produced by the SA. This type of measurement error tends to make the dynamic effects more pronounced.

∗ Tel.: +968 99777512; fax: +968 24413454. E-mail address: [email protected] 0378-7796/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.epsr.2011.01.015

Most data collected by [3] indicates that the SA residual voltage for a given discharge current magnitude is increased by approximately 6–12% as the time to crest of the current is reduced from 8 ␮s to 1.3 ␮s [4]. Indeed, the voltage across the SA is not only a function of the discharge current, but also of the rate of its rise [5]. In addition, the rising rate of residual voltage depends on the current front time and peak value, and kinds of ZnO blocks [6]. Inductive voltage overshoot is another phenomenon which occurs when the time-to-crest of SA discharge currents becomes shorter than approximately 4 ␮s. These voltage overshoots are superimposed on the front of the SA residual voltage waveform which can sometimes exceed the subsequent residual voltage level for the SA. The magnitudes of these overshoots are increased by any unavoidable stray inductance in the measuring loops. The inductive overshoot can be minimized by placing the voltage divider used in the measurement circuit in a coaxial arrangement within the metal-oxide block [6], or by using the D-dot probe or short dividers as it will be demonstrated in this paper. There is no consistent agreement in some technical literature [6–9] concerning the realistic magnitude of these overshoots. There is some evidence, however, that when careful measurements are made, the overshoot does not exceed the subsequent residual voltage for currents cresting in 0.5 ␮s or more [6]. The objectives of this paper are as follows:

• Check the response of a totally closed D-dot probe design [10–12] against uncompensated RD.

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Impulse-voltage measurement has been traditionally performed by capacitive or resistive dividers or with a combination of these components as in the case of universal dividers [13]. The function of the voltage divider is to reduce the high voltage (HV) of some million volts to a voltage level of some 100 V. This low voltage (LV) should be an image of the HV and only reduced by a constant factor “divider ratio”. This ratio should be independent of the applied voltage level, the temperature, the frequency and the surrounding. These dividers are physically large and have to be connected to the test object by a HV long lead. The rise time of fast-front voltages, e.g. SA residual voltage, is much shorter than that of the standard lightning impulse voltages (1.2 ␮s), and even of front chopped impulses [14]. Resistive voltage dividers have a theoretical limitation in the bandwidth which is given by the resistance and the earth capacitance (determined by the height of the divider). The measuring frequency range can be increased by reducing of resistance value but this may create problems due to the increase in the energy dissipated and the influence on the waveform distortion [14]. Damped capacitive voltage dividers are usually used as the standard measuring device for impulse voltages. These dividers have a limited measuring frequency range below that of the fast-front voltages of interest. Depending on the height and choice of the capacitors, i.e. the internal inductance and the necessary damping resistance, these dividers can be adopted to measure fast impulses. Damped capacitive dividers for lower voltages, i.e. smaller physical size, may be used because their response time becomes shorter [14–17]. Universal resistive/capacitive reference voltage dividers (MCRD) are designed for the precise measurement of alternating, direct, lightning and switching impulse voltages [18]. These dividers are based on special measuring capacitors, which guarantee a high stability of the capacitance at both alternating and impulse voltages. There are damping resistors arranged between the internal single HV capacitor packages. Further carefully adjusted damping resistors are located inside the low voltage part and at the beginning of the HV lead. For the measurement of direct voltage, the divider has an additional resistive parallel path. The HV, high-ohmic resistors are arranged in parallel with the separate capacitor stack or in one tube with the capacitor stack. The voltage is applied to the divider over an external damping resistor. It is worth mentioning that a significantly high proportion of the work published about SA residual voltage measurement shows the presence of an initial spike or an inductive overshoot on the front of the residual voltage waveform as shown in Fig. 1 [19]. In addition, there are slow response and fast decay on the rise and fall times, respectively. This especially dominates when measuring the SA residual voltage by uncompensated large impulse voltage dividers at high discharge currents more than 1 kA. Tall dividers do need compensation due to their physical length and the long connecting leads which create a large loop with inductance in the order of about 1 ␮H/m. In practical applications of SA, the presence of a significant overshoot in the transient voltage across the protected device by the SA could be damaging. It is important to identify the nature of this overshoot. That part of the overshoot which may be inherent in the SA, with its nonlinear ZnO elements, housing and terminations, is unavoidable. However, in test procedures, other sources of over-

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• Compare the response of other D-dot probe designs, namely, totally open and semi-open designs. • Demonstrate and apply the divider compensation principle. • Investigate divider response of long and short MCRD.

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Time, µs Fig. 1. The measured residual voltage trace of a 33 kV, 10 kA distribution-class ZnO SA: a 3.5 m high, uncompensated Haefely damped capacitive voltage divider type CS 1000-670 was used, reproduced from [19].

shoot may be more significant as it is illustrated in Fig. 1. The SA residual voltage can inaccurately be measured due to the following reasons: (a) The inductance of the test voltage source and its connections to the SA may, in combination with the relatively large capacitance associated with the high-permittivity ZnO elements, cause an initial voltage overshoot at the SA terminals. (b) High-voltage measurements are usually made with voltage dividers connected in parallel. The transfer characteristics of such dividers are determined by the inherent inductive and capacitive effects associated with both the high- and the lowvoltage arms of the divider, which can limit the response time and cause a voltage overshoot at the high-voltage terminal and/or in the signal output to the recorder. (c) The test circuit arrangement can also be important. If the voltage-divider connection to the SA, then inductive effects are clearly greater. In the application of SA to cable and transformer protection in power systems, this source of voltage overshoot (the distance effect) is unavoidable but can be minimized by installing the SA as close as possible to the protected equipment. (d) Reflection along the diver axial length, and sometimes ringing and/or swinging of the voltage signal when using tall capacitive voltage dividers. (e) Signal cable problems can significantly affect the measured voltage. Therefore, low-losses double-shielded signal cables must be used with proper matching. This paper describes a coaxial D-dot and measurement techniques that minimize the effects discussed in points (a)-(d) and provide residual voltages that are more representative of the actual SA response. This allows more accurate characterization and provides a ‘fingerprint’ of a particular SA transient signal. 3. Principle and design of D-dot probes Electromagnetic pulse (EMP) can be generally detected by four types of probes, namely, (1) antennas, (2) current monitor transformers (coils), (3) B-dot probes (where B is the magnetic flux density), and (4) D-dot probes (where D is the electric flux density). Both B-dot and D-dot probes measure time rate of change of surface current and surface charge density, respectively [20–23]. The operating principle of the D-dot “electric field-coupled” probe is comprehensively discussed in [11,12].

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L3 Charging resistor

Sphere gap R1

L2

L1

Universal Divider (MCRD)

Guard toroid

Resistive Divider (RD)

D-dot probe HVDC

C = 0.56 µF Discharging resistor

δ δ

Attenuator

9-kV Surge arrester Rogowski coil (RC) Impulse current generator (ICG)

VRC

VD-dot

VMCRD

VRD

To digital storage oscilloscopes (DSO) Fig. 2. Schematic representation of the test setup.

Figs. 2 and 3(a) show that the D-dot probe consists of three identical copper toroids of 0.4 m major mean diameter and 0.022 m minor diameter. The three toroids are coaxially placed around the HV electrode “pipe” (1.5 m high and 0.05 m diameter) and each one was mounted by three equally spaced Teflon (PTFE) holders. The LV electrode “aluminum earthed cylinder” (1 m high and 0.44 m diameter) surrounds the test object and the three toroids, and provides a current-return path to earth. In addition, an aluminum disc base is used to provide an additional mechanical support of the complete system and a good uniform current path. In Fig. 3(b) and (c), the aluminum earthed cylinder is replaced by six identical aluminum rods having 1 m long and 0.01 m diameter. The middle “signal” toroid is isolated from the earthed cylinder and fixed at a height of 0.5 m from the base of the earthed cylinder; see Figs. 1 and 2, to allow a separation of about 0.4 m between the SA top terminal and the bottom earthed toroid. The top and bottom toroids are movable and equally spaced above and below the middle “signal” toroid by a separation (ı). These outer toroids are earthed to serve as a shield for the signal toroid from any external electromagnetic interference (EMI) and provide a certain degree of field modification. The effect of the separations of toroids (ı) was numerically investigated [12]. In this work ı is kept at 0.04 m. A 13-m, 50- coaxial cable having a capacitance Cc = 1.34 nF was connected to the signal toroid via a series matching resistance Rm = 50 , and the attenuator shunt capacitance Ca = 1 nF was directly connected between the signal toroid and earth, i.e. at Rm -toroid junction. Corona anywhere in the test circuit can influence the output of the D-dot probe. It should be prevented by optimizing the probe assembly via finite element simulation to calculate the electric field inside [11]. A simple capacitive divider is formed in the HV side by C1 and in the LV side by C2 , the attenuator, the signal cable and the scope capacitances (Cc , Ca and CDSO ). For time greater than twice the cable travel time, the total LV capacitance C2 = C2 + Cc + Ca + CDSO , as all the aforementioned capacitances are connected in parallel. It is worth mentioning that the termination mode of the probe used in the paper (high-impedance termination mode) corresponds

to the use of the probe as a charge meter and not as a D-dot meter [11].

4. Experimental setup Response of the constructed D-dot probe is investigated by using the test circuit shown in Fig. 1 to generate high-impulse currents. In Fig. 1, RD is a Haefely resistive divider R 500, which is used in parallel to the D-dot probe. For the sake of comparison, another type of divider, universal type “MCRD” (HIGHVOLT MCR 0.5/200-100/40), is used. In high-voltage measuring techniques, the use of unit step responses is widely accepted means to characterize high-voltage dividers [24]. The unit step response of the constructed D-dot probe, the Haefely resistive divider “R 500” [25], and the HIGHVOLT universal type “MCRD” [26] has shown that they have experimental response times of 10 ns, 25 ns and 15 ns, respectively. The measuring uncertainty of voltage is <0.7% [25,26]. A third homemade short universal divider, with an experimental response time of 15 ns, is used to show the effect of divider self-inductance on the measured voltage waveform. A distribution-class ZnO siliconerubber-housed SA rated 9 kV was tested inside the D-dot assembly. Its maximum continuous operating voltage and nominal classification (discharge) current are 7.2 kVrms and 10 kApeak , respectively. The impulse current generator (ICG) consists of 5 capacitors rated 100 kV and 2.8 ␮F each (Marx generator principle). These capacitors are charged in parallel via high-ohmic resistances and discharged via 5 sphere gaps in series through the SA. The overall ICG capacitance is C = 0.56 ␮F. Connecting lead inductances between the ICG and the voltage measuring devices, i.e. the D-dot probe, the universal divider (MCRD) and the RD are L1 , L2 and L3 , respectively. The latter one has the highest value because of the large physical size of the RD (2.4 m high). PEM Rogowski current waveform transducer (Rogowski coil “RC”) is used to measure the discharge current. It has a sensitivity of 0.2 mV/A and a maximum impulse current of 30 kApeak . For impulse current waveforms of 12/20 ␮s, 4/6.6 ␮s and 1.5/11 ␮s, the

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Fig. 3. Photos of the D-dot probe with different designs: (a) totally closed, (b) totally open and (c) semi-open.

values of R1 and L1 are respectively about 0.5  and 73 ␮H, 0 and 4 ␮H, and 26  and 4 ␮H. The value of L1 is the sum of the self- and the external inductances, where the former is estimated from the loop physical dimensions and found to be about 4 ␮H. Two-channel LeCroy Wave Surfer 452, 500 MHz, 2 GS/s scopes (DSO) are used to record the SA residual voltage and discharge current. When measuring fast-front voltages, extra care must also be taken in order to avoid errors due to enhanced EMI [14,18]. An aluminum attenuator box was directly connected to the LV electrode “aluminum earthed cylinder”. The LV signals from the RD, the D-dot probe and the Rogowski coil are transferred to the recording equipment via coaxial cables laid inside earthed copper flexible pipes. Fig. 3(a), (b) and (c) respectively illustrate photos of the D-dot probe with different designs, namely, totally closed, totally open and semi-open designs. The latter one has a 0.25-m high aluminum cylinder. The latter two designs are introduced to have an easyaccess assembly for fixing and replacing the SA under test. The scale factor (ratio of the input HV to the output LV at the DSO end) of these three designs are measured by a HV linear resistance and checked by the SA residual voltage measured by the RD. The corresponding scale factors for the three designs shown in Fig. 2 are ∼0.752, ∼1 and ∼0.77 V/kV, respectively. 5. Analysis of divider compensation principle This section presents theoretical analysis of the inductive coupling in the divider loop and the principle of inductive compensation. All conductors with any appreciable length have inherent inductance and the same has SA. Their inductance can significantly contribute to the residual voltage measurement, especially when the arrester is long. This inductive effect is part of the reason that

the residual voltage crests before the discharge current. In addition to the inherent inductance in ZnO SA, the residual voltage is influenced by what is known as the temporal transition behavior of a semiconductor [27–29]. This behavior also appears as inductive effect, which is a function of the ZnO n-type semiconductor transitioning from a non-conductive state to a conductive state. This effect cannot be only seen at fast rising surges but also at all impulse frequencies. Uncompensated long impulse dividers are not suitable to measure the SA fast-front residual voltage because of the superimposed inductive overshoot on the main impulse. The total voltage measured by these dividers (Vtotal (t)) is the instantaneous sum of the actual voltage across the SA and that induced due to the inductive coupling in the divider loop (VSA (t) + Vind (t)). To minimize this inductive coupling, a compensation voltage (Vcomp (t)) must be injected in that loop in the opposite direction, which yields Vtotal (t) = VSA (t) + Vind (t) − Vcomp (t)

(1)

Fig. 4 depicts a part of the test setup to study the inductive coupling in the divider loop, which it will be used to conduct the analytical analysis of the induced voltage in this loop. Finally, the induced emf in the divider loop due to the definition of the current in Eq. (A4) of Appendix A is given by Vind (t) =

o aVch ln(b/o ) (ω cos ωt − ˛ sin ωt) exp(−˛t) 2ωL

(2)

To compensate this emf in the divider loop, an air-cored coil with a certain inductance can be inserted as shown in Figs. 3 and 4 in the divider HV side for a given impulse current waveform. This was primarily invented by HIGHVOLT Company, Germany, and imple-

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Fig. 5(a)–(c) illustrate oscillograms of impulse voltage (measured by uncompensated RD and the totally closed D-dot probe) for front and tail times (tf /tt ) ≈ 11/19 ␮s discharge current with peaks of Ipeak = 1, 5 and 10.5 kApeak , respectively. It can be seen that all discharge currents follow overdamped current waveform, i.e. as the first half cycle in Eq. (A4). From the comparison between Fig. 5(a) and (c), it is obvious that the current tail time decreases from ∼19 ␮s to ∼16.5 ␮s with a little decrease in the front time by ∼1 ␮s. This is attributed to the decrease in angular frequency of the overdamped oscillatory current wave because when the SA conducts at lower currents, it is equivalent to a higher resistance [11,30], see Eqs. (A5)–(A8). In addition, the measured residual voltages by the D-dot probe and the RD show that they have a common actual peak instant which slightly leads that of the current peak as also observed in [31] due to SA self- and connecting lead inductances. It was noticed from the results over a wide discharge current range that the residual voltage rise time decreases with the increase in the discharge current as the SA responds faster. Similar trend was also observed in [6]. In other words, the higher the discharge current, the worse is the measured residual voltage by the RD as it can be seen in Fig. 5, and later in Fig. 6. Fig. 6(a)–(c) show oscillograms of impulse voltage (measured by uncompensated RD and the totally closed D-dot probe) at Ipeak ≈ 7 kA for different discharge current waveforms with tf /tt ≈ 12/20, 4/6.6 and 1.5/11 ␮s, respectively. It is clear that as the current front time becomes shorter, the inductive overshoot dominates on the voltage measured by RD due to the increase in ∂i(t)/∂ t as given by Eq. (A3). Therefore, compensa-

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mented in [30]. The compensation voltage is given by

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Fig. 4. Inductive compensation in the divider loop. (a) Ipeak = 1 kApeak ; (b) Ipeak = 5 kApeak ; (c) Ipeak = 10.5 kApeak .

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Ipeak = 10.5 kApeak.

Fig. 5. Oscillograms of impulse voltage (measured by uncompensated RD and the totally closed D-dot probe) at different levels of ∼11/19 ␮s currents.

tion of the RD is a crucial issue in measuring such SA residual voltages. 6.2. Other D-dot probe designs As it was mentioned earlier three different designs of the Ddot probe were investigated, namely, totally closed, totally open

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(Fig. 3(b)), the LV earthed cylinder is replaced by six aluminum rods arranged hexagonally. Therefore, the capacitance C2 decreases which leads to the increase the scale factor to be ∼1 V/kV. In addition, the value of C2 = C2 + Cc + Ca + CDSO is in the order of few nano Farads and hence this improper matching causes the fast decay of the D-dot probe signal as shown in Fig. 7 in comparison to that for the totally closed design in Fig. 5(c). For the semi-open design shown in Fig. 3(c), the D-dot probe signal is roughly identical to that of the totally closed design, where their scale factors are ∼0.752 and ∼0.77 V/kV, respectively. Therefore, the semi-open design is preferred over the others as it gives good response and has an easy-access to fix and replace the SA. 6.3. Divider compensation

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To demonstrate the divider compensation principle, which was formerly discussed in Section 5, a HIGHVOLT ICG and MCRD were used to test a large ZnO block (4.5 cm high and 6 cm diameter) together with a metallic brass block having the same dimensions, as requested in [28]. The ZnO and the metallic blocks are mounted on top of the other. For the uncompensated MCRD divider case when the ZnO on the bottom, a significant inductive effect can be observed on curve 1 as shown in Fig. 8. When swapping the two blocks, curve 2 shows voltage across the metallic block and the 40

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Fig. 6. Oscillograms of impulse voltage (measured by uncompensated RD and the totally closed D-dot probe) at Ipeak ≈ 7 kApeak with different discharge current waveforms.

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and semi-open designs, see Fig. 3. Fig. 7 illustrates oscillograms of impulse voltage (measured by uncompensated RD and the totally open D-dot probe) at ∼10.8 kApeak (∼10/16.5 ␮s). It can be noticed that the D-dot probe signal decays much faster than that of the RD. This is attributed to the fact that for the totally open D-dot probe

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Time, µs Fig. 8. Oscillograms of impulse voltage (measured by uncompensated and compensated (Lc = 33 ␮H) MCRD) for ∼4.4/9.5 ␮s discharge current waveforms and Ipeak ≈ 121 kApeak .

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Fig. 10. Oscillogram of impulse voltage (measured by the totally closed D-dot probe when the SA is connected to the pipe top end) for ∼4.1/7.1 ␮s discharge current waveforms and Ipeak ≈ 4.5 kApeak .

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connecting leads as well. It can be seen that this voltage waveform follows a di/dt dependence, where the positive peak is higher than the negative one and it has a zero-crossing at the instant of the current peak. The former effect will be explained later in Fig. 9 by introducing the resistive component of this voltage due to the skin effect phenomenon. The second scenario is achieved by using an air-cored compensation coil with inductance Lc = 33 ␮H. Curve 3 in Fig. 8 shows the SA actual residual voltage which can be obtained by perfect compensation (i.e., Vind (t) = −Vcomp (t)) where its peak instant occurs at the zero crossing of curve 2 because the ZnO block behaves as a nonlinear resistance. Alternatively, curve 3 can be obtained by instantaneous subtraction of curves 1 and 2 providing that they are measured at the same discharge current. Fig. 9 illustrates an oscillogram of impulse voltage across the metallic block (measured by perfectly compensated MCRD with Lc = 33 ␮H and when the ZnO block is at the top) for ∼4.4/9.5 ␮s discharge current waveform and Ipeak ≈ 121 kApeak . It is obvious that this voltage is mainly due to the metallic block, coil and connecting lead resistances as it roughly follows the discharge current waveform shown in Fig. 8. The peak value of this voltage represents about 2.4% of the actual residual voltage across the ZnO block (curve 3 in Fig. 8). From the comparison between curve 2 in Figs. 8 and 9, it can be concluded that the inductive component of the voltage dominates over the resistive one as the di(t)/dt is very high, i.e. about 27 kApeak /␮s. Although voltage dividers can perfectly be compensated but there is unavoidable resistive effect. Similar trend of curve 2 in Fig. 8 was observed when using the test circuit shown in Fig. 2. Fig. 10 shows oscillogram of impulse voltage (measured by the totally closed D-dot probe when the SA is connected to the pipe top end) for ∼4.1/7.1 ␮s discharge current waveforms and Ipeak ≈ 4.5 kApeak . The peak value of this voltage represents about 2.6% of the SA residual voltage measured by the D-dot probe. In addition, the D-dot probe signal shows a little inductive effect as the D-dot test module has a very low inductance [10–12,19].

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4

5

6

7

8

9

10

Time, µs

(a) Tall MCRD with and without compensation at ∼10.5kApeak (∼3.8/6.4 µs). 25

5

Short MCRD 20

4

D-dot Probe 15

Voltage, kV

Fig. 9. Oscillogram of impulse voltage across the metallic block (measured by perfectly compensated MCRD with Lc = 33 ␮H) for ∼4.4/9.5 ␮s discharge current waveform and Ipeak ≈ 121 kApeak .

0

Current, kA

-2

3

10

2

Current 5

1

0

0

Current, kA

Voltage, kV

0.4

6.4. Divider size -5 -5

Fig. 11(a) and (b) depict a comparison between the response of long and short MCRD dividers, respectively. In Fig. 11(a), the residual voltage is shown in the case without compensation (Lc = 0), overcompensation (Lc = 176 ␮H) and perfect compensation (Lc = 116 ␮H). It can be seen that the rising part of the residual volt-

0

5

10

15

20

25

-1 30

Time, µs

(b) Uncompensated short MCRD at ∼3 kApeak (∼11/19 µs). Fig. 11. Oscillograms of the residual voltage measured by long and short MCRD dividers and the semi-open D-dot probe, and the discharge current.

I.A. Metwally / Electric Power Systems Research 81 (2011) 1274–1282

age is slowed down even with perfect compensation. On the other hand, Fig. 11(b) illustrates that uncompensated short MCRD divider gives good response because of the low self-inductance and the decrease in the successive reflection along the divider axial length. Although the short MCRD divider signal has a small overshoot and a little distortion in the rising part its overall performance is much better than that of the long one, and its signal coincides on that of the D-dot probe.

  ) law, the magnetic flux density (B(t) in the azimuthal direction a in free space due to the impulse current (i(t)) and at any distance from the current carrying conductor () is given by   B(t) = o H(t) =

Acknowledgments The author would like to thank Prof. V. Hinrichsen, Mr. M. Tuczek and Dr. T. Wietoska, Technische Universität Darmstadt, Germany, for their technical and moral supports, and the kind hospitability and the technical support received at the High-Voltage Laboratories at the same university, as well as the financial support by Alexander von Humboldt Foundation, Bonn, Germany. Appendix A. As it was illustrated in Fig. 4 that part of the test setup was used to study the inductive coupling in the divider loop. From Ampere’s

  o 2p

 i(t)a

(A1)

Hence, the magnetic flux linked to the divider loop ( m (t)) can be calculated as a function of the loop dimensions (a and b in Fig. 4) and the current (i(t)) as



7. Conclusion • D-dot probe operating principle is introduced for a coaxial assembly. The probe consists of three identical copper toroids placed around a high-voltage electrode. Three different designs are investigated, namely, totally closed, totally open and semi-open designs. The latter design gives a good response, has an easyaccess assembly for fixing and replacing the surge arresters (SA) under test and can thus practically be applied in the lab. The totally closed type gives even slightly better response but is not practical in use, as the SA can be only installed with difficulties. • The totally open design, with the best access to the test samples, gives an unacceptable output signal. The introduced D-dot probe acts as a capacitive voltage divider, i.e. high-impedance termination mode. Its output voltage is picked up from an attenuator connected to the middle toroid. The D-dot probe scale factors of these three designs are measured and found to be ∼0.752, ∼1 and ∼0.77 V/kV, respectively. • Simultaneous impulse voltage measurements are done by using the designed D-dot probe, RD, and two universal dividers (long and short MCRD). • Inductive divider compensation principle is introduced and the test results have shown its importance when testing SA. Aircored coils with inductance in the range of few hundred micro Henries can give perfect compensation when they are properly adjusted. On the other hand, the rising part of the residual voltage is distorted even with perfect compensation. Otherwise, good agreement is found between the residual voltages of the properly compensated divider and the D-dot probe. For both devices, an error of about 2.5% is expected due to the resistance of the SA connecting leads. • Very short dividers with short response times, optimized in size and response characteristic for the special purpose, can give the actual residual voltage waveform without compensation similar to the D-dot probe, where a small and unimportant inductive overshoot can be expected from such dividers. • Generally, conventional voltage dividers do need compensation when applied to SA residual voltage measurements as it is demonstrated in this paper. The D-dot probe gives also very good results and can be used as an alternative to conventional voltage dividers. However, matching of the low-voltage side of the former needs careful consideration.

1281

 = → B(t) · ds

m (t) = s

=

a b

o i(t)  · dp dz a  a 2p 

z=0 o

o a ln(b/o ) i(t) 2

(A2)

where o is the radius of the SA current-carrying conductor. According to Faraday’s law, the induced emf in the divider loop yields Vind (t) = −

∂

m (t)

∂t

=−

o a ln(b/o ) ∂i(t) 2 ∂t

(A3)

The ICG capacitor banks are charged to a voltage Vch and discharged when the spark gap is triggered, the current i(t) of an underdamped RLC series circuit is given by [32] i(t) =

V  ch

ωL

exp(−˛t) sin ωt

(A4)

where R, L and C are the total series circuit resistance, inductance and capacitance including the ICG, connecting leads and test object, and ω is the angular frequency. The latter is given by [32] ω=



1 LC



− ˛2

(A5)

where ˛ is the attenuation constant and defined as [32] ˛=

R 2L

(A6)

The current front and tail times (tf and tt ) are respectively given by [32] √ tf ≈ 1.25 LC (A7) √ tt ≈ 2.5 LC (A8) References [1] P.P. Barker, R.T. Mancao, D.J. Kvaltine, D.E. Parrish, Characteristics of lightning surges measured at metal oxide distribution arresters, IEEE Trans. Power Deliv. 8 (January (1)) (1993) 301–310. [2] J.A. Martinez, D.W. Durbak, Parameter determination for modeling systems transients. Part V: Surge arresters, IEEE Trans. Power Deliv. 20 (3) (2005) 2073–2078. [3] IEEE Working Group of Surge Protective Devices Committee, Modeling of current-limiting surge arresters, IEEE Trans. Power Apparatus Syst. PAS-100 (August) (1981) 4033–4040. [4] IEEE Working Group 3.4.11: Application of Surge Protective Devices Subcommittee Surge Protective Devices Committee, Modeling of metal oxide surge arresters, IEEE Trans. Power Deliv. 7 (January (1)) (1992) 302–309. [5] A. Bayadi, N. Harid, K. Zehar, S. Belkhiat, Simulation of metal oxide surge arrester dynamic behavior under fast transients, in: The International Conference on Power Systems Transients – IPST 2003, Paper 14b-1, New Orleans, USA, 2003. [6] W. Schmidt, J. Meppelink, B. Richter, K. Feser, L. Kehl, D. Qiu, Behavior of MOsurge arrester blocks to fast transients, IEEE Trans. Power Deliv. 4 (1) (1989) 292–300. [7] W. Breilmann, Protective characteristics of complete zinc-oxide arresters and of single elements for fast surges, in: Fifth Int. Symp. on High Voltage Eng., Paper 82.04, Braunschweig, Germany, August, 1987. [8] A. Bargigia, M. deNigris, A. Pigini, The response of metal oxide resistors for surge arresters to steep front current impulses, in: Fifth Int. Symp. on High Voltage Eng., Paper 82.01, Braunschweig, Germany, August, 1987. [9] C. Dang, T. Parnell, P. Price, The response of metal oxide surge arresters to steep front current impulses, IEEE Trans. Power Deliv. PWRD-1 (January (1)) (1986) 157–163.

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[10] A. Haddad, P. Naylor, D.M. German, R.T. Waters, A fast transient test module for ZnO surge arresters, IET Proc. Meas. Sci. Technol. 6 (1995) 560–570. [11] I.A. Metwally, D-dot probe for fast-front high-voltage measurement, IEEE Trans. Instrum Meas. 59 (8) (2010) 2211–2219. [12] I.A. Metwally, Coaxial D-dot probe: design and testing, in: Proceedings of Conference on Electrical Insulation and Dielectric Phenomena, vol. 1, Paper 4B-3, Virginia, USA, October 22–25, 1995, pp. 298–301. [13] K. Feser, W. Pfaff, E. Gockenbach, Distortion-free measurement of high impulse voltages, IEEE Trans. Power Deliv. 3 (1988) 857–866. [14] E. Gockenbach, M. Aro, F. Chagas, K. Feser, J. Kuffel, R. Malewski, T. McComb, P. Munoz Rojas, G.G. Rizzi, Measurements of very fast front transients, Electra, CE/SC 33 GA/TF 03 181 (December) (1998) 71–91. [15] G.J. FitzPatrick, E.F. Kelley, Comparative high voltage impulse measurement, J. Res. Natl. Inst. Stand. Technol. 101 (September–October (5)) (1996) 639–658. [16] J. Meppelink, P. Hofer, Design and calibration of a high voltage divider for measurement of a very fast transients in gas insulated switchgear, in: Fifth Int. Symp. on High Voltage Eng., Paper 71.08, Braunschweig, Germany, 1987. [17] F.C. Creed, The Generation and Measurement of High Voltage Impulses, Center Book Publishers, Inc., New Jersey, USA, 1989. [18] IEC 60060-2, High Voltage Test Techniques – Part 2: Measuring Systems, Edition 2.0, 1994–11, 1994. [19] I.A. Metwally, Impulse-voltage measurement of distribution-class surge arresters by D-dot probes, J. Electr. Power Components Syst. (JEPCS), 39, (2011) in press. [20] B. Kelly, EMP from chemical explosions, IS-4 Report Section, Group P-14 Los Alamos National Laboratory, March 1993. [21] L. Ruijin, S. Caixan, Z. Xiaoguang, Development of a transient voltage measuring system, in: Eighth Int. Symp. on High Voltage Eng., Yokohama, Japan, 1993, pp. 141–144. [22] A.P. van Deursen, P.F.M. Guiliclx, van der Laan, A current and voltage sensor combined in one unit, in: Eighth Int. Symp. on High Voltage Eng., Yokohama, Japan, 1993, pp. 463–466. [23] E.I.M. van Heesch, A.P.I. van Deursen, C.A.P. Jacobs, W.F.J. Kersten, P.C.T. van der Laan, Field tests and response of the D/I H.V. measuring system, in: Sixth Int. Symp. on High Voltage Engineering, Paper 42.23, New Orleans, USA, 1989. [24] IEC 60060-2, TC/SC 42, Ed. 3.0 (2010–11), High-Voltage Test Techniques – Part 2: Measuring Systems, 2010. [25] Resistive Impulse Divider Type R. Haefely Test AG, Basel, Switzerland, Dec. 2003, available online: www.haefely.com/pdf/LL R.pdf. [26] Universal Resistive/Capacitive Reference Voltage Divider Type MCR, HIGHVOLT Prüftechnik Dresden GmbH, Germany, Data Sheet No. 5.24/6, available online: www.highvolt.de/datasheets/5-24-6.pdf. [27] IEEE Standard C62.11-2005, Metal-Oxide Surge Arresters for AC Power Circuits (>1 kV), 2005. [28] IEC Standard 60099-4, Ed. 2.2, 2009–05, Surge Arresters – Part 4: Metal-Oxide Surge Arresters without Gaps for A.C. Systems, 2004. [29] J. Woodworth, Arrester Facts 013 Understanding Arrester Discharge Voltage, Arrester Works, Rev 3 03-25-09, pp. 1–8, December 2008.

[30] M. Reinhard, Experimental research on single-impulse energy handling capability of MO-varistors used in high-voltage systems in consideration of a complex failure criterion, Ph.D. Dissertation, Darmstadt University of Technology, Germany, 2009 (in German). [31] C.A. Christodoulou, F.A. Assimakopoulou, I.F. Gonos, I.A. Stathopulos, Simulation of metal oxide surge arresters behavior, in: IEEE Power Electronics Specialists Conference (PESC 2008), Rhodes, Greece, June 15–19, 2008, pp. 1862–1866. [32] M. Abdel-Salam, H. Anis, A. El-Morshedy, R. Radwan, High-Voltage Engineering: Theory and Practice, 2nd ed., Marcel Dekker Inc., New York, USA, 2000. Ibrahim A. Metwally (IEEE M’93–SM’04) was born in 1963. He received the B.Eng. degree in electrical engineering (Hons.), the M.Eng. degree in high-voltage engineering, and the Ph.D. degree in high-voltage engineering from Mansoura University, Mansoura, Egypt, in 1986, 1990, and 1994, respectively. The Ph.D. degree was received in collaboration with Cardiff University, Cardiff, UK. Currently, he is a Professor with the Department of Electrical Engineering, Mansoura University, Mansoura, Egypt. He worked as a Visiting Professor with the University of the Federal Armed Forces, Munich, and Darmstadt University of Technology, Germany from 2000 to 2002 and in the summers of 2003–2007, and in summer 2009, respectively. He has also been on leave as a Professor with the Department of Electrical and Computer Engineering, College of Engineering, Sultan Qaboos University, Muscat, Oman since August 2002. His areas of research include oil- and gas-flow electrification in both electric power apparatus and pipelines, measurements of fast impulse voltages and currents, line insulators and zinc-oxide surge arresters, coronas on overhead transmission lines, impulse voltage characterization and modeling of electrical machines, particle-initiated breakdown in gas-insulated switchgear (GIS) and gas-insulated transmission lines (GITL), power quality, stray-current corrosion in the oil industry, and hazards of lightning strikes to buildings, overhead power lines and aircrafts. He is a Fellow of the Alexander von Humboldt (AvH) Foundation, Bonn, Germany, a Senior Member in the Institute of Electrical and Electronics Engineers (IEEE) and a Member of the International Electrotechnical Commission (IEC). He has published 124 papers, of which more than half have appeared in highly reputed international journals. He has completed 17 industrial projects in UK, Egypt, Germany and Oman, and 16 support services (consultancies) in Oman. He was awarded both the First Rank of the National Prize in Engineering Sciences in 1998 and 2004, and the Late Prof. Dr.-Ing. M. Khalifa’s Prize in Electrical Engineering in 1999 and 2005 from the Egyptian Academy of Scientific Research and Technology. He has been a regular peer reviewer for IEEE Transactions, IET Proceedings, Journal of Electric Power Systems Research, European Transactions on Electrical Power, the International Journal of Emerging Electric Power Systems and Journal of Electrostatics. His biographical profile was published in Who’s Who in Science and Engineering in 2001.