21, rue d'artois, F-75008 Paris http://www.cigre.org A3-301 Session 2004 CIGRÉ CURRENT INTERRUPTION WITH HIGH VOLTAGE AIR-BREAK DISCONNECTORS D.F. PEELO* DF PEELO & ASSOCIATES R.P.P. SMEETS KEMA HIGH-POWER LAB J.H. SAWADA BC HYDRO (Canada) J.G. KRONE HAPAM B.V. (The Netherlands) B.R. SUNGA BC TRANSMISSION CORP. L. VAN DER SLUIS S. KUIVENHOVEN DELFT UNIVERSITY 1. INTRODUCTION The function of high voltage air-break disconnectors is to provide electrical isolation of one part of the power system. This isolation can be normal day-to-day operations or for purposes of maintenance. In the latter regard, disconnectors have a major role in providing safe working conditions and are therefore required by standards to have a higher withstand capability across the open contact gap than to earth. In fulfilling this function, disconnectors are operated under energized conditions and will interrupt current, the type being dependent on the circumstances ranging from transformer magnetizing and capacitive currents of a few amperes at most to loop currents in the tens and hundreds of amperes. Disconnectors do not have interrupting current ratings but, given that they have a contact breaking arrangement, they have a certain current interrupting capability. Standards recognize this fact and the IEC disconnector standard defines a negligible current interrupting capability at 0.5 A and a further bus-transfer (loop) switching capability of up to 1600 A for some disconnectors [1]. Present IEEE standards have a similar negligible current definition but do not recognize bus-transfer as a switching duty [2]. In an earlier version of the IEEE standard, transformer magnetizing and capacitive currents, and small load currents were identified in this context [3]. Deregulation and opposition to the addition of new plant, particularly transmission lines, has resulted in increased demand to use disconnectors to interrupt current in non-traditional ways. The present Cigre WGA3.13 (Changing Network Conditions and System Requirements) has actually identified abnormal disconnector operation as an item of potential interest. Reality is however that this is a complex issue given that an arc is being drawn in air for a variety of disconnector types most commonly vertical break, centre break, double break and pantograph installed in a multitude of physical busbar arrangements and clearances. The intent of this paper is to provide a perspective on this subject based on utility practices and observations and on tests performed at BC Hydro, Eindhoven University of Technology and at KEMA. 2. HISTORICAL BACKGROUND Using air-break disconnectors to interrupt current has been a long-time practice in North American power systems. The IEEE conducted surveys of utility practices in this regard as far back as 1949 and again in 1962 [4, 5]. Likewise, the majority of the literature on the subject is based on the above-noted * dfpeelo@ieee.org
practice. The principal or most-often cited reference is that of Andrews, Janes and Anderson in 1950 [6]. Andrews et al introduced the notion of arc reach and derived empirical equations from a series of field tests run in 1944 to 1948 to determine how far an arc will reach toward adjacent phases or structures when breaking transformer magnetizing and loop currents. The tests were based on a fundamental assumption that the arc would extinguish when it reached a so-called critical length. Accordingly the tests were performed on 33 kv disconnectors. Treating the results as having a general applicability to other voltage levels was a giant leap of faith but some (limited) support for this was provided by later tests at higher voltages [7]. Interestingly, in the IEEE 1962 survey, only two of seventy-one responding utilities stated that they used the Andrews et al approach in spite of endorsement by the IEEE Switchgear Committee [5]. In 1990, the IEEE published a guide for the interruption of transformer magnetizing and capacitive currents based on the tests of Andrews et al [8]. However, scepticism with respect to the basis and manner of the tests has remained over the years, not the least because actual field observations do not reflect those described in the reference. Actual practice has therefore evolved user by user based on need and using a trial-and-error approach. 3. TRANSFORMER MAGNETIZING CURRENTS For modern low loss power transformers, the high side magnetizing current at 100% excitation will be typically less than 2 A, but most often less than 1 A. The current is non-sinusoidal with a high third harmonic content and is stated as an RMS equivalent value as derived in factory core loss measurements. The transformer side transient recovery voltage (TRV) associated with the duty is a critically damped or overdamped oscillation. Because of the non-linearity of the transformer core, the characteristics of the TRV frequency and degree of damping are dependent on the level of excitation. Both quantities actually increase with increasing excitation with the former generally less than 300 Hz. This is based on unpublished studies which further show that the worst case TRV peak across the switching device will not exceed 1.3 pu; in fact, such a value has a very low probability of occurrence and the TRV peak is dominated by the source voltage peak and is therefore close to 1 pu. When the switching device is an air-break disconnector, current interruption is dependent on at least achieving an open contact gap sufficient to withstand the TRV peak. The minimum required contact gap for vertical break disconnectors in the range 72.5 to 550 kv are shown in Fig. 1. It can be readily seen that the interruption of steady state magnetizing current should be achievable before the blade reaches the 45 angle position. This is the simplest case and is applicable provided that thermal effects are absent in the arc (which would certainly be true at less than 1 A) and no inrush currents occur. In the latter regard, inrush current is a common occurrence due to the repeated break-make restriking nature of the switching event. However, because the restrikes will occur close to the TRV peak and thus close to the source voltage peak the inrush currents will tend to have low values. The net effect of the low inrush currents is to prolong the arcing period by their duration [10]. In North America, it is common to use so-called quick-break whip type auxiliary devices on disconnectors in the range 72.5 to 245 kv range in order to achieve a near restrike-free current interruption [9]. These devices are applied principally on vertical break disconnectors and function as follows: as the blade opens the whip is restrained by a captive attachment on the jaw assembly and when the blade reaches a predetermined position the whip releases achieving a high contact tip velocity in the order of 0.5 to 0.6 m/cycle at 60 Hz. The whip has a Fig. 1. Minimum contact gaps for vertical break disconnectors for magnetizing current interruption 2
damping mechanism which ensures that once it passes the blade position, it does not swing back into the open gap. To function correctly, the whip device must meet two basic rules: firstly, at whip release the contact gap should at least equal or preferably exceed those shown in Fig. 1; and secondly, the current must be interrupted before the whip reaches the blade position, otherwise its effect is negated and the arc will transfer to the blade. Field observation of unloaded transformer switching shows that arc reach is insignificant. This is as expected because the event is usually one of dielectric recovery. Andrews et al, however, recorded significant arc reaches in their magnetizing current interruption tests. The explanation for this difference lies in the manner in which the tests were performed. The tests were run at 12 kv to 49 kv with currents up to 35 A giving arc lengths up to 13 m long. The high current values were achieved by overexciting a number of medium voltage transformers in parallel and the set-up was thus more that of interrupting a load current than a transformer magnetizing current. 4. CAPACITIVE CURRENTS Disconnectors switch capacitive currents in those cases where unloaded busbars, (short) lines and (short) cables are taken out of service. Typical values of switching current are given in Table I [10]. Laboratory tests were performed at KEMA in order to gather insight into the relevant phenomena of capacitive current switching. To this aim, capacitive currents in a range from 0.23 2.2 A were interrupted by a horizontal-break 300 kv disconnector in tests with a source voltage of 300/ 3 = 173 kv at 50 Hz. There were no special devices attached to the disconnector for interrupting the current. Of special interest in this test was: (1) the generation of overvoltages (as elaborated earlier in [11]) and (2) the influence of the source side capacitance on the disconnector arc duration. A diagram of the test set-up is shown in Fig. 2. Table I: Typical values of capacitive current Capacitive current (A) at Equipment 72.5 kv 145 kv 245 kv 550 kv type 50 Hz 60 Hz 50 Hz 60 Hz 50 Hz 60 Hz 50 Hz 60 Hz CT* 0.025 0.03 0.033 0.04 0.033 0.04 0.1 0.12 CVT* - - 0.133 0.16 0.22 0.27 0.49 0.59 Busbars*/m 1.7 10-4 2 10-4 0.32 10-3 0.39 10-3 0.54 10-3 0.65 10-3 1.1 10-3 1.3 10-3 Lines/km 0.15 0.18 0.24 0.29 0.42 0.5 1.33 1.6 * For outdoor substations 4.1 Overvoltage generation Fig. 2. Basic test circuit (TRV shaping elements, stray inductances and measuring devices are not shown here) During the test, overvoltages were observed, having peak values up to 2.4 pu (577 kv). The magnitude of these overvoltages is very much dependent on the value of the source side capacitance Cs (see Fig. 2). This dependence is expressed in Fig. 3, where the load side voltage (voltage across Cl) is given for various values of capacitive current and Cs. For smaller values of Cs (for example, a high voltage transformer on the source side), much higher overvoltages are observed than at higher values of Cs. Neglecting the (slight) dependence of overvoltage on arc current, a clear dependence of overvoltage as a function of Cs/Cl is evident, as expressed in Fig. 4. It was observed that the overvoltages occur just before arc extinction. Further, an interesting observation was that the arc existed in two distinct modes (also observed by Knobloch [12]), roughly described as: 3
Erratic mode, showing an erratically moving, bulging arc having a length several times that of the blade tip spacing. This arc appearance mode is observed normally at intermediate and low values of source capacitance (< 60 nf) and is associated with the higher overvoltage values (Fig. 5, upper image). Stiff mode, showing a contraction of the arc from the erratic mode to a form of more or less a straight path between the contacts. No significant overvoltages are observed during this mode. It is observed at lower current and values of Cs 60 nf (Fig. 5, lower image). In Fig. 3, the dotted line roughly indicates the boundary of the existence regions of the two arcing modes. load side voltage peak (pu) 2.5 2.0 1.5 erratic mode stiff mode capacitive current (A) 1.0 0.0 0.5 1.0 1.5 2.0 4.2 Analysis Fig. 4. Peak voltage across load vs. ratio Cs/Cl High-frequency measurements of arc current and load and source side voltages permit to explain the phenomena. In Fig. 6, oscillograms are shown for 30 ms of current (upper trace), load and source side voltages (lower traces) for two test cases. A capacitive current switching arc is a succession of interruptions and restrikes (in contrast to the loop current switching arc, which is burning continuously, see Section 5). Upon restrike of the arc, first the voltages across Cs and Cl (u Cs and u Cl respectively) will equalize through a highfrequency discharge in the loop formed by Cs, Cl and the disconnector. In the upper case of Fig. 6 (Cs/Cl = 2.5), this makes a 28 khz discharge (see inset at left side in Fig. 6). The final level of u cs and u cl after voltage equalization is determined by the ratio Cs/Cl: At high Cs/Cl (Fig. 6 upper) the major excursion is in the load voltage whereas at small value of Cs/Cl (Fig. 6 lower) the major excursion is in the source voltage. Cs 1 nf 6 nf 20 nf 60 nf 100 nf Fig. 3. Measured overvoltages against arc current for various values of source side capacitance Cs load side voltage peak (pu) 2.5 2.0 1.5 1.0 0.01 0.10 1.00 10.00 Cs/Cl The latter situation creates the highest overvoltages since u Cl (together with u Cs ) has to swing back to the 50 Hz steadystate voltage from a level of maximum Fig. 5. Capacitive current arc just before extinction Upper: "erratic" arc, 2 A, Cs/Cl = 0.04 (2.43 m tip spacing) Lower: "stiff" arc, 1 A, Cs/Cl = 3.1 (1.22 m tip spacing) 4
+1 pu trapped load voltage to maximum -1 pu 50 Hz source, and can only do so with an overshoot of maximum 2 pu on top of the -1 pu 50 Hz source. Thus, in theory, 3 pu may be reached but, due to damping, in practice this level is never reached as is evident from Fig. 4. After equalization of voltages on Cs and Cl, a low frequency oscillation arises in the loop consisting of Cs and Cl in parallel, the source inductance and the source, having a frequency between 500 3000 Hz. The resulting transient current becomes much higher than the few amperes in the steady state case. 5
At small Cs/Cl the transient currents in excess of 100 A cannot be interrupted easily, thus creating conditions for continuous arcing during half a 50 Hz period and one restrike per full per 50 Hz period. At large Cs/Cl the transient arcing current is much smaller and of lower frequency (Fig. 6 upper trace) creating easier conditions for restrike at lower voltages and a much higher repetition rate of restriking follows. Thereby the mode of the arc is stiff. 4.3 Arc duration Arc duration was found to be strongly dependent on both arc current and source capacitance Cs. This is shown in Fig. 7. For low values of capacitance, the arc duration rises almost linearly with current. The angular (turning) velocity of each of the disconnector blades is 40 degrees/s, and the angle where the arcing starts is 8 degrees. The 45 degree position is already reached 930 ms after arc initiation. This implies that all the arcs extinguish at blade angles > 45 degrees but always before 90 degrees when the maximum arc duration (ms) 2000 1800 1600 1400 1200 1000 blade tip spacing of 2.8 m has been reached. Nevertheless, the arc reach (the maximum lateral expansion) is always relatively small (even in the test depicted in Fig. 5a with arc duration 2040 ms the arc reach is approximately half the blade tip spacing). This implies that capacitive current switching is less severe than loop switching in terms of risk that the arc reaches adjacent phases or structures. From Fig. 7 it is clear that additional source capacitance (cables, long busbars, instrument transformers) is an effective means to shorten the arc duration just as it is to reduce overvoltages. 5. LOOP SWITCHING Loop switching is the transfer or commutation of current from one circuit to a parallel circuit. Examples of loop switching are: In double busbar schemes, the transfer of current from one busbar to the other. The current can be up to 1600 A and busbar transfer or open circuit voltage (voltage across the open disconnector after current transfer) up to 300 V [1]. For series capacitor banks, the opening of the bypass disconnector to transfer the current from the transmission line to the bank bypass breaker and damping reactor circuit. The current can be as high as 1800 A with an open circuit voltage up to 1000 V or even more in some cases. Cs 1 nf 6 nf 20 nf 60 nf 100 nf 0.0 0.5 1.0 1.5 2.0 2.5 capacitive current (A) Fig. 7. Arc duration versus current and source capacitance For loop switching between transmission lines, most often at subtransmission voltages of 72.5 and 145 kv, the current is limited to about 100 A and an open circuit voltage in the kilovolt range. Loop switching between circuits involving either transformation or cables is usually not practiced nor is it recommended. Fig. 8. Loop switching case schematic Schematically, the loop switching case is as shown in Fig. 8. X S and X L are the series and parallel circuit impedances, I S and I L the initial currents in the disconnector and the parallel circuit, respectively, and I L the total current. As the disconnector opens, it draws an arc which at any one instant has an arc voltage of U a. On a time varying basis, the arc voltage can be expressed as: 6
u a = r a i s (1) where r a is the arc resistance and i s the momentary current. As the disconnector continues to open r a increases due to the increasing arc length and decreasing i s because of commutation to the parallel circuit. The arc will continue to propagate provided that its power input (P) continues to increase: dp > 0 dt (2) However, the arc grows to such a length that eqn. (2) can no longer be satisfied and arc instability followed by total collapse results. This is apparent in Fig. 9 from an actual loop switching test where the peaks of the arc voltage are plotted against the corresponding current values. For eqn (2) to be satisfied, the arc 10 Powertech May 2000, Xt = 60 Ω PT 6_20 PT 7_13 voltage must increase faster than the current is decreasing. 8 Fig. 9 shows that this is initially PT 8_26 PT 8_27 the case but, as the arc PT 7_22 lengthens, the slope of the u-i 6 characteristic becomes less and PT 6_30 less and the arc collapses. Video still images for test no. PT 6_20 of Fig. 9 are shown in Fig. 10 for the last 24 cycles of the arcing period. The images (the view is along the axis of the vertical break disconnector from the jaw end) show the erratic and convoluted nature of the arc and further that any determination of the arc length is essentially impossible. Note the partial arc collapse in the Voltage at peak current [kv] 0 0 50 100 150 200 250 Peak current [A] seventh image: the effect of partial arc collapse is to cause transfer of the current back from the parallel circuit to the disconnector circuit thereby prolonging the arc duration. The luminosity of the arc is sustained if eqn (2) is satisfied but can be seen to change in colour in the second last image signalling a loss of the power input. Note the hot gas remnants in the final image. 6. ACKNOWLEDGMENT The support and advice of Dr.ir. P.H. Schavemaker of Delft University of Technology with the studies partly described in this paper is gratefully acknowledged. 7. REFERENCES 1. IEC 62271-102 Alternating current disconnectors and earthing switches First Edition 2001-12. 2. IEEE Std. C37.30-1997 IEEE Standards for High-Voltage Switches. 4 2 PT 6_20 Fig. 9. U a -i s characteristics for various initial disconnector currents and a loop impedance of 60 ohms 3. ANSI Std. C37.30/IEEE Std. 324-1971 Definitions and Requirements High-Voltage Air Switches, Insulators and Bus Supports. Revised in 1992 and 1997 to become reference 2 above. 4. AIEE Committee Report, Report on Transformer Magnetizing Current and Its Effect on Relaying and Air Switch Operation. AIEE Transactions, Vol. 70, 1951. 5. IEEE Committee Report, Results of Survey on Interrupting Ability of Air Break Switches. IEEE Transactions on Power Apparatus and Systems, Vol. PAS-85, No. 9, September 1966. 7
Fig.10. Vertical break disconnector interrupting 82 A with a loop impedance of 60 ohms (PT 6_20) 6. F.E. Andrews, L.R. Janes and M.A. Anderson, Interrupting Ability of Horn-Gap Switches. AIEE Transactions, Vol. 69, 1950. 7. CEA Project 069T102 Report, The Interrupting Capability of High Voltage Disconnect Switches. July 1982. 8. IEEE C37.36b-1990 IEEE Guide to Current Interruption with Horn-Gap Switches. 9. D.F. Peelo, Unloaded transformers can be switched reliably with disconnect switches. Transmission & Distribution, July 1987. 10. D.F. Peelo, Current Interruption Using High Voltage Air-Break Disconnectors, Ph.D. Thesis, Eindhoven University of Technology, 2004. 11. C. Neumann, Nichtstandardisierte betriebliche Beanspruchungen beim Schalten von Trenn- und Erdungsschaltern im Hochspannungsnetz, Ph.D. Thesis, Darmstadt University, 1992. 12. H. Knobloch, Switching of Capacitive Currents by Outdoor Disconnectors. Fifth International Symposium on High Voltage Engineering, Braunschweig, August 1987. 8