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Jun 24, 2009 - possible to fix them to ground-voltage level. G. Antecedents on Live Line Installation and Maintenance. The stabilizer solution to be proposed in ...
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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 24, NO. 3, JULY 2009

Feasibility of Electronic Tap-Changing Stabilizers for Medium Voltage Lines—Precedents and New Configurations Salvador Martínez García, Senior Member, IEEE, Juan Carlos Campo Rodríguez, Member, IEEE, José Antonio Jardini, Fellow, IEEE, Joaquín Vaquero López, Alfonso Ibarzábal Segura, and Pedro María Martínez Cid

Abstract—Thyristor-based onload tap-changing ac voltage stabilizers are cheap and robust. They have replaced most mechanical tap-changers in low voltage applications from 300 VA to 300 kVA. Nevertheless, this replacement hardily applies to tap-changers associated to transformers feeding medium-voltage lines (typically 69 kV primary, 34.5 kV line, 10 MVA) which need periodical maintenance of contacts and oil. The Electric Power Research Institute (EPRI) has studied the feasibility of this replacement. It detected economical problems derived from the need for series association of thyristors to manage the high voltages involved, and from the current overload developed under line fault. The paper reviews the configurations used in that field and proposes new solutions, using a compensating transformer in the main circuit and multi-winding coils in the commutating circuit, with reduced overload effect and no series association of thyristors, drastically decreasing their number and rating. The stabilizer can be installed at any point of the line and the electronic circuit can be fixed to ground. Subsequent works study and synthesize several commutating circuits in detail. Index Terms—AC voltage stabilizer, onload tap-changer (OLTC), power conditioning, power quality, voltage control.

I. INTRODUCTION C voltage stabilizers based on thyristor-commutated onload tap-changers (OLTC) are reliable and have almost completely replaced mechanical switching equipment in low-voltage applications, single- and three-phase, ranging from

A

Manuscript received May 02, 2008. Current version published June 24, 2009. This work was supported by IBERDROLA S. A., Spain, I+D program, ESTRAP Project: Feasibility of Electronic Medium Voltage Line Stabilizers, 1994–1997. Paper no. TPWRD-00322-2008. S. Martinez is with the Departamento de Ingeniería Eléctrica, Electrónica y de Control (DIEEC) de la Universidad Nacional de Educación a Distancia (UNED), Madrid 28040, Spain (e-mail: [email protected]). J. C. Campo is with the Departamento de Ingeniería Eléctrica, Electrónica, Computación y Sistemas de la Universidad de Oviedo, Gijón 33205, Asturias, Spain (e-mail: [email protected]). J. A. Jardini is with the Departamento de Energia e Automação Electricas da Escola Politécinca da Universidade de São Paulo, São Paulo 05508-900, Brazil (e-mail: [email protected]). J. Vaquero is with the Departamento de Tecnología Electrónica de la Universidad Rey Juan Carlos, 28933 Madrid, Spain (e-mail: [email protected]). A. Ibarzábal is is with the INCOESA, Bo Bidecoeche, Vedia 48390, Vizcaya, Spain (e-mail: [email protected]). P. M. Martínez is with IBERDROLA S.A., Dirección de Innovación, Calidad y Medio Ambiente, Bilbao 48008, Vizcaya, Spain (e-mail: [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/TPWRD.2009.2021032

300 VA to 300 kVA. This replacement does not apply to the mechanical tap-changers associated to feeding transformers of long medium-voltage lines (typically 69 kV primary, 34.5 kV secondary, 10 MVA), usually housed in the transformer tank, needing periodical renewal of switches and isolating oil [1]. The EPRI [2], [3] and other research institutes have evaluated the feasibility of such a replacement, both in voltage stabilizers and phase regulators, concluding (see II.F.1) that, although technically possible, the equipment derived from the simple substitution of mechanical switches by their semiconductor counterparts is too expensive because of the amount of devices in series association needed to obtain medium-voltage switches, and because of the high over-current under line faults. They recommend looking for modified topologies needing less semiconductors. The long-standing positive results [4]–[12] on solid-state tap-changers with commutating current limitation based on multi-winding coils inspired us to explore their abilities for medium-voltage lines. This work studies the feasibility of a direct structure (without compensating transformer) and a seriescompensated structure, both commutated by multi-winding coil. The steady and commutating operations are analysed to derive expressions for the electrical working conditions of switches, transformers and coils, thus enabling their economical evaluation. The circuits exposed summarise the results of an iterative analysis and synthesis cost comparative search method applied in other research on power electronics [13], [14], some of them bringing the tap-changers to the field of fast voltage and phase regulators [15]–[17]. The direct configuration here analyzed in IV.A is similar to those studied by EPRI. The results confirm its high cost and a strong cost reduction for the new series-compensating solutions here proposed, accomplished by the synergic effect of the compensating transformer, the multi-winding commutating coil and a bypass fault switch. II. ANTECEDENTS A. Electromechanical Stabilizers The early generalization of ac voltage in the electric lines was an obvious encouragement to take advantage of tap-changers associated to the transformers involved in energy distribution to accomplish easy voltage stabilization. Taps were first selected manually and then electromechanically. Variants made from iron toroid core and brushed over a one-layered coil are

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Fig. 1. Fac simile of a mechanical tap-changer German patent 474 613, 1926, awarded to B. Jansen. The incoming tap closes before the outcoming tap opens, thus avoiding output voltage interruptions. The internal overcurrent during overlap is limited by resistors, shorcircuited after the operation. By courtesy of Maschinenfabrik Reinhausen (MR), Regensburg, Germany.

very common today in laboratories and industrial plants for one-phase low-voltage regulators in the range of 100 VA to 15 kVA I . For voltage regulation and stabilization, from 30 kVA to 300 kVA III , three-limb vertical core equipments with brushes contolled by servomotor are still in use. Long-lasting graphite materials used in the brushes give high reliability to this classic, although inherently slow (about 1 s correcting time), equipment. B. Early Line-Voltage Correctors Line breakers, load interrupters (both manual and automatic) and other simple devices intended to help power management have been used in medium-voltage lines since the twenties, located at any point from head to end of the lay-out. Remote control versions have been available since the seventies (Fig. 4, (left)). In Europe, most ac tap correctors are derived from patents such us the German 474 613 one for mechanical resistive changers, awarded to B. Jansen in 1926 (Fig. 1). This technology was acquired by the MR company and brought about important improvements in 1973. C. Present Electromechanical Line Voltage Correctors Electromechanical changers are used today in connection with transformers feeding long medium-voltage lines (11.9 kV to 34.5 kV; 3.15 MVA to 40 MVA) from high-voltage lines (30 kV to 138 kV). The need for making the change before break to avoid interruptions necessarily creates internal short circuits requiring current limiting components, resistive or/and inductive. Improvements are made on transformers and commutating circuits [18]–[22]. The switches are normally housed in the transformer tank [Fig. 2, (upper left corner)]. In the U.S., the Cooper Power Systems Co. has combined the inventions of T. A. Edison with the innovations of the McGraw Co. (first tap-changing voltage correctors in 1959) and currently makes more refined variants with inductive short-circuit limitations [Figs. 3 and 4 (right)]. Stabilizers with no connection to the line feeding transformer, as showed in Figs. 3 and 4 on the right, are useful in long lines with loads connected all along, correcting most voltage problems. Mechanical changers have been diversified by other European, American, Australian and, more recently, Asian companies. There are variants with electronically-aided commutation

Fig. 2. Medium-voltage line transformer 132 kV/21.5 kV, 50 MVA with the covering tank removed. It contains an electromechanical onload tap-changer (OLTC, upper left) made by MR. Courtesy of INCOESA, Vedia, Spain.

Fig. 3. Single-phase, 32-step on-load voltage stabilizer with inductive limitation of internal short circuits. It operates as an autotransformer with a compensating winding in series with the load. The polarity of this winding can be reversed by a switch and the turns changed by an 8-positions tap-changer associated to a bridging reactor. Courtesy of Cooper Power Systems [23], Waukesha, WI.

Fig. 4. Outdoor equipment for medium-voltage line management. Left: Line breaker. Right: McGraw-Edison-type voltage stabilizer (see Fig. 3). By courtesy of IBERDROLA, Dirección de Innovación y Calidad, Spain, Madrid area.

[19]–[23] and some pure electronic variants [24]–[33]. The later are not able to replace mechanical solutions in high and medium-voltage lines yet. Although the frequency of operation imposed by the system to tap-changers is very low (typically several commutations per hour) and the minimum time between changes is limited by the control circuit (to no less that several minutes), the wearing of switches and oil make periodical and expensive revisions necessary, which is the main problem with mechanical tap-changers. The aim of the work conducted is to replace mechanical tapchangers by thyristor-based maintenance-free solutions.

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Fig. 5. (a) Direct configuration tap-changing stabilizer using a multi-winding coil to limit commutation over-current with 50 Hz transformer equivalent power of 5% of the nominal power, typically. Optional short-circuit winding L and switch S are explained in III.C.4. (b) Current in the outgoing (up) and the incoming triac (bottom) in a ten-tap, shared load, 1 kVA, 230 V 15% input, 2:2% output, at nominal load. Worst case current peaks. Courtesy of Electrónicas Boar, Madrid, Spain.

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D. Electronic Tap-Changing Low-Voltage Stabilizers. Industry and Office Applications Low-cost plastic case thyristors and triacs up to 40 A, 800 V made it possible to replace most mechanical tap-changers by their electronic counterparts in low-voltage applications under 10 kVA in the eighties. Both zero-crossing current commutation and resistor- or inductor-aided commutation have been used to reduce internal over-current at commutation [4]–[7], [24]–[33]. The multi-winding limiting coil with alternated corresponding terminals [4]–[7] is especially useful in this respect, see Fig. 5. It reduces the commutation over-current to a single impulse (8.3 ms or 10 ms) of about four times (seven times in the worst case) the repetitive peak with nominal load, while pure resistive limiters reduce it to no less than fifteen times, especially in equipments over 10 kVA. Since the nonrepetitive peak current allowed by triacs and alternistors is about thirteen times its repetitive peak (four to nine times in an antiparallel thyristor pair) and assuming a security coefficient of 0.333 in rms current [34], the multi-winding coil in practice eliminates the need to oversize the semiconductors because of commutation over-current. In fact, for the worst case of the aforementioned numerical values (nonrepetitive peak current allowed by the semiconductor equal to four times its repetitive current peak), the switch can be able to withstand a nonrepetitive current peak of %

%

(1)

of the repetitive peak under nominal load, while the multi-winding coil demands no more than 700%. Since the equivalent 50 Hz power of that coil (power of the 50 Hz transformer feasible with the same iron and cooper [13]) is about 3% to 11% of the nominal power of the stabilizer, the global cost is not significantly increased. The coil allows for two main steady-state operation modes: shared load by two adjacent taps and nonshared load. The first mode [4] (2nd and 3rd pat.), [5] yields in practice to very reliable equipments, as it performs rather well even with some switches (not adjacent ones) out of operation. Stabilizers based

Fig. 6. Configurations for electronic low-voltage tap-changing stabilizers. (a) Direct configuration, suitable between 1 and 4 kVA. (b) Compensating series transformer—the single switch-comb configuration is suitable between 5 and 50 kVA. The switch current is typically one third of the value in the direct configuration. (c) Compensating series transformer—double switch-comb configuration, suitable over 50 kVA. Switch current is typically one sixth of direct configuration value. The multi-winding coil shown in Fig. 5 can be used in all configurations. In both compensating transformer configurations the taps can be associated to the primary or to the secondary of the main transformer.

on multi-winding coil made by Electrónicas Boar in the 1980s (later absorbed by Chloride) in the range of 0.3 to 300 kVA, sold in more than 45 countries, drastically lowered failure frequency. E. Advantages of the Series-Compensating Transformer Another reason for the low-voltage electronic tap-changers success is the series-compensating transformer configuration [Fig. 6(b) and (c)], replacing direct configuration [Figs. 5(a) and 6(a)]. In a typical stabilizer with % input and % output deviation, the current in the switches is divided by three [Fig. 6(b)] or six [Fig. 6(c)], compared with the current in the switches of the direct configuration [Fig. 6(a)]. The switches voltage is not increased, thus making it possible to reduce their cost and to increase the power of practical equipment with no parallel thyristor association. The power handed by the compensating transformer is about 15% of the nominal power. The series-compensating transformer and its economical autotransformer variant [23] were used in very early electromechanical voltage stabilizers. The scheme in Fig. 3 shows an example of it. It is also used in stabilizers not based on

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MARTÍNEZ GARCÍA et al.: FEASIBILITY OF ELECTRONIC TAP-CHANGING STABILIZERS FOR MV LINES

tap-changing, such as ferroresonant and phase-controlled reactor stabilizers. An added advantage of the compensating transformer is to reduce high over-currents in the switches as those imposed by line faults. In this abnormal condition the stabilization function of the equipment is not mandatory, the tap switches are opened and the primary of the compensating transformer is short-circuited (see Figs. 6 by means of an extra fault deviation switch and 11). The short-circuit impedance of the compensating transformer appears between the mains and the load. Once the output opens and the tapcurrent comes down to normal values, switches are restored. The series-compensating transformer and the switches can be connected before or after the main transformer of the stabilizer. In the first case [Fig. 6(b) and (c)], the main transformer sees almost constant voltage, which means a very economical core design. In the second case, the switches are less exposed to line over-voltages because of the isolation provided by the main transformer. F. Problems Found in Previous Studies and Solutions 1) EPRI Reports: Although modern monitoring and control techniques improve the performance of both the mechanical and the electronic versions [35]–[42], tap-changing ac voltage stabilizers and phase regulators show some common problems, when applied to medium- and high-voltage lines, on account of their discrete-regulation principle. The 1988 and 1990 EPRI reports [2], [3] on improved on-load-tap changers discussed the viability of electronic versions to reduce the high-cost maintenance of mechanical solutions. The main conclusions were as follows. 1) Electronic stabilizers derived by simple substitution of switches in the mechanical equipment are not possible due to the high number of semiconductors needed and their cost. 2) New topologies minimizing the number of switches needed to obtain a given number of voltage changes would help. 3) Even with less switches, the need to associate the semiconductors in series to support high steady-state and transient voltage raises their number and complicates protection. 4) A no less important problem is the very high over-current in medium-voltage lines developed as a consequence of phase-to-phase and phase-to-ground faults. They reach up to 1 200% the nominal current and last up to 100 ms, the opening time of the ordinary line over-current protectors. These over-currents are well supported by mechanical taps providing that changes are inhibited in the meantime, but in electronic versions they make oversizing the thyristors necessary, as their over-current rate1 for 100 ms is about 200% of the nominal current, far from the 1 200% needed. EPRI concludes that “major breakthroughs are required before LTC equipment using solid-state devices can become economically attractive for ordinary power transformers. Such 1High current thyristors allow during 100 ms about 200% of the nominal current. Even taking into account the 0.333 current security coefficient mentioned in II-D [34] (and loosing it during faults), only 600% of the nominal load could be supported with thyristors not oversized by this reason, thus still under 1 200%.

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breakthroughs could come by developing ways to avoid the impact on device ratings of system short-circuit currents and transient voltages”2. 2) Tap-Changer Association With the Line Transformer: The need to reduce slow voltage variations is common in mediumvoltage (13 kV to 34.5 kV) very long and/or undersized (poor) lines. The scalar line voltage drop reaches up to 10% of the nominal voltage. Assuming an additional 9% scalar drop in a typical 69 kV/34.5 kV, 10 MVA feeding transformer, the drop rises to 19% from the high voltage feeding line to the end of the medium-voltage line. Although the stability in the highvoltage side is usually good and the legal variation allowed in the medium-voltage inlets is about %, the aforementioned transformer-plus-line drop is usually compensated by a tap-changer associated with the line transformer [42], [43]. Since the actual voltage at the end of the line is not known, a compound control criteria (input voltage plus load current) is used to select the tap [18], [42]–[44]. When the customers are connected all along the medium-voltage line, it can be technically impossible to supply a correct voltage to all of them by means of a single OLTC associated to the feeding transformer. In these cases one or more additional OLTC are installed in selected intermediate points of the line (see Fig. 4, (right): two single-phase OLTCs connected in V-configuration serving a three-wire line [23]). The following sections show that using an isolating seriescompensating transformer topology allows for the tap-changer to be installed in any point of the line. If connected at the output of the line-feeding transformer, it can replace the associated OLTC, see Fig. 10. These options are not usually considered by the studies on feasibility of electronic tap-changers, as [2], [3]. Besides the advantages in lowering the number and the current rate of the switches, seen in Section II-E, this topology makes it possible to fix them to ground-voltage level. G. Antecedents on Live Line Installation and Maintenance The stabilizer solution to be proposed in this paper can be connected (and disconnected for maintenance purposes) to a live medium-voltage line. Live working in overhead lines is a mature and reliable activity today regulated by very precise protocols from the electric companies [45] and governments [46]–[48]. III. ON CONFIGURATIONS, OPERATION MODES AND COMMUTATION A. Iterative Analysis and Synthesis Cost Comparative Search The method applied to explore the power circuit configurations has been described in [13] and [14], when applied to the active filters search. Applications to tap changers can be seen in [8], [10], [11], [15]–[17]. It consists of the following steps: analysis of new circuits; synthesis of component rates as a function of equipment specifications; study of circuit variants; derivation of component cost by transformation to standard components; technical and economical comparison of solutions. This paper compares the advantages of the general configurations of stabilizers and introduces the commutation circuits. Other solutions are studied in subsequent works. All of them are based on natural zero-current commutation aided by the multi2EPRI

Perspective in the Technical Portofolio corresponding to [2].

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TABLE I STATES FROM A GIVEN NUMBER OF TAPS. IN BRACKETS TOTAL TAPS FOR THE EXAMPLE

winding coil seen in II-D and their variants [4]–[7] because of the low overcurrent and overvoltage developed in the thyristors during the changes and the good practical results in low-voltage applications. They can be referred to as thyristor-based, slowcommutated tap-changers, acting once in a half-cycle or at a lower rate (supercyclic commutation). B. Configurations Comparison Irrespective of the circuit configuration, the number of voltage state changes needed by a line stabilizer is a function , of the input high-voltage deviation to be compensated ( positive and negative per unit deviation, respectively) and the . The voltage drops allowed output voltage deviations and in the with nominal load in the mains transformer also have to be compensated. As most mechanical line tap-changing line stabilizers, the electronic counterpart must offer the ability to add or detract an optional per-unit output . By neglecting the voltage drops in the internal deviation transformers and coils and in the switches, the number of changes needed is (2)

Fig. 7. Shared load mode. Steady state with taps 1 and 2 activated. Ideal switches and coil with perfect coupling. Load power factor 0.8 inductive.

is added or subtracted under the extreme condition. For the preof the nominal. This vious numerical example means that the maximum power managed by the compensating transformer is 14.5% of the nominal power. As shown in Section II-E, in both structures the taps can be in the primary or in the secondary of the main transformer. In the direct structure with taps on the primary side, the semiconductors can be used to open the circuit in case of a line fault, but they must be sized to support the fault current reflected to the primary and the inrush current when connecting the transformer. Using autotransformers in the compensating structure brings about an important reduction in their size and very economical equipments (Fig. 3). If the isolation transformers are maintained the electronic switch combs [Fig. 6(b) and (c)] can be fixed to ground, which simplifies the isolation components and maintenance, as shown at the end of II-F-2. C. Operation Modes Comparison. Tap Commutation

accounts for errors in voltage-sensing and where control management. As an example, for (legal and , the result maximum 0.07), (eight in practice). Obviously, the number of is required voltage states is (3) The taps needed to obtain a given number of states depends on the operation mode (shared, nonshared and mixed load, see III-C) [5], [6] and the configuration (Fig. 6) resulting in expressions (4)–(9) in Table I. Expressions (7)–(9) also apply to the direct configuration with double-comb (not contemplated in Fig. 6), sometimes used for low power with the second comb working as a voltage vernier. The configurations to be studied here belong to the direct [Fig. 6(a)] and the compensating transformer [Fig. 6(b) and (c)] structures. The secondary of this transformer adds or subtracts to the line the voltage needed to keep the load voltage within limits. In an optimal design, the same scalar value (10)

In all solutions considered here, the multi-winding coil seen in II-D, Fig. 5(a), is used to reduce the internal short-circuit current during commutation. This coil enables four steady state operation modes, each one associated to a tap commutating procedure. They shall be studied here with the minimum extension needed to make a technical-economical comparison in part IV. 1) Shared Load: a) Steady State Operation: The alternated corresponding terminals of the multi-winding coil easily lend themselves to sharing the load by two consecutive taps [4] (2nd pat.), [5]–[7]. Both operating windings make a 1:2 ratio autotransformer that distributes the load in a 1:1 proportion between the selected taps (taps 1 and 2 in Fig. 7). The output voltage must be the mean value of both active taps voltage. A magnetizing current circulates in both equalizing windings , where is the voltage between consecutive taps and the inducmodifies the scalar tance of a single winding. In practice and to about 0.4 and 0.6 times the nominal tap currents current. The power rating of the stabilizer feasible with a given semiconductor is then multiplied by (1/0.6). b) Commutating Process: Fig. 8 shows the equivalent circuits and the relevant variables evolution for the change from taps 1–2 to taps 2–3. • Turning off the outcoming switch. Interval to .

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MARTÍNEZ GARCÍA et al.: FEASIBILITY OF ELECTRONIC TAP-CHANGING STABILIZERS FOR MV LINES

Fig. 8. Commutating process for shared load. (a) Equivalent circuit before t : from t to t , the outcoming switch S is de-excited but still conducting. (b) Equivalent circuit from t on: the incoming switch S is turned on and shares the load with the permanent switch S . (c) Relevant current and voltage waves. The equivalent circuit from t to t , with S on and S ; S off, is not represented. Note that i is I before t and goes towards -I after t .

While the branches and , for instance, are , see Fig. 8(a), the outturned on, each one producing coming switch is de-excited at time . It is effectively turned off after the first zero-crossing of its current at time ’. The load is now fed by branch until is turned on at time . The equivalent circuit for ’ to is not given in Fig. 8. This subinterval lasts less than half a cycle is in practice 3 to 7 times its repetitive and the peak value of peak during steady state operation with nominal load. (defined as The equivalent total magnetizing current flowing through a single winding, entering in dotted terminal) until ’ and from ’ to . The peak value is of must not saturate the core to keep the limitation ability during the next interval. . • Turning on the incoming switch

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is At time , the switch of the incoming branch remains conducting. The turned on while branch is chosen slightly over half a cycle to time duration ensure that extinguishes before . A new equalizing current operates whose magnetizing effect is opposite the effect , due to the change of doted terminals. Nevertheless, of the process does not imply any step in the magnetizing current, thus avoiding over-voltage in the load and switches. reaches zero at time In fact, when the load current is . Then, at time this current flowing through one winding produces the same magflowing through two windings. The netizing effect that magnetization sign is also maintained as deduced after the dot of the involved terminals. is closed at time evolves in a few cycles Once from to the new steady state value . 2) Nonshared Load: a) Steady State Operation: The circuit in Figs. 7, 8 can also operate with a single active branch in each voltage state [6], [7]. One winding and one switch support the nominal load current and the overloads, a less advantageous condition that in shared load for the coil and switches. The number of voltage states equals the number of taps, see (5), one state more than for shared load, see (4). The voltage drop in the coil (90 in advance with respect to ) can be up to four times the drop in the switch, not appreciably disturbing the load voltage. b) Commutating Process: The commutation to a next tap has two steps. First, the switch of the incoming branch is turned on and starts operating in shared load with the outgoing branch, as explained in III.C.1. After a few milliseconds, the outgoing switch is de-excited with no regard to its current. It is effectively turned off in the first current zero crossing, as seen in III-C-1.b for . No steps are forced in the magnetizing current, nor is voltage transient generated. Since this operation in shared load lasts no more than one or two half-cycles, a higher value for the equalizing current than in shared load is permitted and a low coil voltage drop in steady state is easily obtained. The equivalent power of the coil is also reduced. 3) Mixed Load (Shared and Nonshared Load): By combining the above modes, a mixed operation is obtained. The load can be fed in steady state either by one single branch or by two consecutive branches [4] (1st and 3rd pat.), [6]. The number of voltage states, see (6), is the sum of the states in both modes. The switches and the coil windings must be rated must be a compromise drawn from for nonshared load. the recommended values for shared and nonshared load. In a change, the stabilizer passes from a given mode to the other one. Only the following changes are possible: 1) Change from one operating branch, nonshared load, to two branches (incoming the next or the precedent tap) shared load. 2) Change from two branches, shared load, to a single branch (outgoing: either the superior or the inferior tap), nonshared load. In any case the commutation follows one of the processes seen in III-C-1.b or III-C-2.b. 4) Nonshared Load and Mixed Load With Coil Short-Circuit Winding: If the coil voltage drop in nonshared load is too high

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(as in equipment with high values of the inter-tap voltage which, in turn, happens when input and output voltage deviation helps keep that can be high), adding a short-circuit winding drop low. and do not operate during commutation. is shortened by means of an extra switch After a change [4] (1st and 3rd pat), Fig. 5(a). The inductive drop in the tap winding is eliminated leaving the resistive drops and the conducting drop. closes one cycle reflected value of the after a tap change and opens one cycle before a new change. is the same as in a The power managed in winding one-tap winding since both operate assembled as a transformer. is reduced by the turns ratio of with The current in respect to a main winding. If this ratio is chosen over ten, the becomes negligible. cost of

IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 24, NO. 3, JULY 2009

Fig. 9. Solution 1. Three-phase stabilizer, direct configuration, taps in the secondary. Only phase R is detailed. Values for a typical medium voltage line.

D. Coil Magnetic Evolution. Equivalent 50 Hz Transformer Power If a commutation starts with no heed paid to the magnetic state of the coil core, the induction can in practice reach a transient value up to 2.8 times the steady state value, [5], [6]. Then, typically a steady induction of 0.5 T must be chosen to guarantee no more than 1.4 T in transients which, in turn, results in a high core section. The resulting worst half-cycle peak current during commutation found in equipment designed under this criteria, as mentioned in II.D, is less than 4 times the steady repetitive peak current with nominal load for shared load, less than 6 times for nonshared load and less than 7 times for mixed load (when operating in single tap). For stabilizers over 100 kVA it is economically advisable to limit the maximum induction (thus reducing the security factor from 2.8 to 1.3 or less) by choosing a convenient instant to start commutation. This requires a delay of one cycle or less from the tap change order of the control circuit. The delay has no practical effects in slow acting (supercyclic) stabilizers. The equivalent 50 Hz transformer power of the coil is reduced in the same proportion as the maximum induction. The commutation start time can be chosen by monitoring the core induction, by voltage and current phase criteria or by a mixed procedure. IV. INITIAL ANALYSIS AN SYNTHESIS OF CONFIGURATIONS Here the direct and the compensating transformer configurations (see III-B) shall be briefly analyzed and synthesized to compare their pros and cons independently of the operation mode (shared, nonshared and mixed load, see III-C) applied. As this study will show, the direct configuration (Solution 1) is at a clear disadvantage, both technically and economically. It is similar to the solutions studied by EPRI [2], [3]. The stabilizers derived from the compensating configuration here proposed (Solutions 2 to 5) can be installed in any point of the line with no association to the line transformer, thanks to an auxiliary feeding shunt transformer. A. Stabilizers With Direct Configuration (Solution 1) The electronic switches are linked to the secondary of the HV/MV transformer that feeds the medium voltage line. They manage the complete line power and voltage. In many mechanical stabilizers the switches are in the primary to reduce the cur-

Fig. 10. Stabilizer, with series compensating transformer CT, connected at the output of the line feeding transformer. Circuit variant with independent shunt transformer ST feeding the tap changer. To calculate ST power, the single comb is assumed. Only one phase is detailed.

Fig. 11. Single-phase circuit of the stabilizer for solutions with compensating transformer CT and fault deviation switch S . The tap topology chosen in the figure is a single switch-comb of five taps (Fig. 6(b)). The shunt transformer ST feeds CT through de tap comb and the limiting multiple winding coil L L. ST also adds a fixed voltage to the line to accomplish balanced operation of CT. An optional medium voltage disconnecting switch MDS makes live installation and removal possible.

0

rent rate. In electronic equipments this advantage is cancelled by the need for more series-connected semiconductors. Fig. 9 shows the general circuit with only one phase in detail for the sake of simplicity. Three-phase equipment is normally made with three wye-connected single-phase circuits, although other connections are sometimes preferred for specific applications and control strategies [22]. 1) Thyristor Rating: The steady estate current in the switches (nonshared load; mixed load in the nonshared condition) is the nominal line current . The maximum overload typiduring 1 h every 8 cally previewed by line operators is 0.2

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h and, under fault condition, 11.5 during 0.1 s. Normally, all semiconductors are chosen so as to withstand the voltage of the upper tap. The worst case voltage (maximum line voltage condition) is derived taking into account that the transforming ratio between the secondary upper tap and the primary must guarantee the minimum allowed load voltage plus the optional , while compensating for output deviation the voltage drops and , under minimum input voltage . Then (for the numerical limits given in the example in III-B, for the line of 34.5 kV and 10 MVA in Fig. 9, and for the said 11.5 overload rating), each switch withstands

MVA kV A rms during A rms during kV repetitive peak

s

(11) (12)

kV

(13)

kV rms kV

Taking into account the usual security coefficients [34] such as for current; 0.83 supplementary for parallel association; for voltage; 0.7 supplementary for series association; the ratings for the series-parallel ensemble in each switch are A rms during h A A repetitive peak during s kV kV repetitive peak

(14) (15) (16)

An economical and easy option (March 2007), which is not in the limit of the available current and voltage, could be the thyristor Westcode K1947ZC450 (330 /unit) whose characteristics are A RMS C and for an antiparallel pair A A rms A av A C and for an antiparallel pair A A rms A repetitive peak during A nonrepetitive peak V repetitive peak

1497

) are needed for the equipment. Thyristors for the coil shortseen in III-C-4, if any, have been ignored circuit switches because of their low rating. This confirms the EPRI conclusion seen in II-F-1: the electronic upgrade of mechanical stabilizers by simple substitution of switches in the direct configuration is possible but expensive. 2) Mains Transformer Power: The effective power of the mains transformer exceeds the nominal power [10 MVA in the example] due to the extra winding necessary to feed the extreme tap operating with minimum input voltage and maximum load. The extra secondary voltage needed to compensate for the negative input deviation and the line drop , in per unit of , so that the power the nominal voltage, is of the transformer rises in that proportion. In transformers cooled by oil, the typical transient line overload condition (20% during 1 h every 8 h, as shown in IV-A-1) for this equipment makes it necessary to overrate the power following the average quadratic rule. Afterwards, the effective power of the transformer for this configuration and the values of the example is

MVA MVA

The per unit extra secondary voltage has been calculated for nonshared load for the sake of simplicity. For other operation modes similar values are obtained. The average quadratic factor 1.027 also would apply for a line transformer without taps. 3) Equivalent 50 Hz Transformer Power of the Multi-Winding Coil: The mixed load mode is considered here because it gives the best economical results. Every winding withstand the nominal current in nonshared load, although in shared load the current is lower. The voltage handled by any winding is equal to the output voltage change which, in turn, must not exceed the . This value allowed step of the output voltage can be affected in practice by 0.9 due to voltage drop in the conducting thyristors. Pressumably, a 0.8 coefficient acounts for the under rate of the windings section due to nonsimultaneous operation. Thus, the equivalent power becomes

MVA ms

ms

Subsequently, a series association of kV kV steps is needed. Parallel association is not required. Thus, thyristors (59 400 ) for the three phases per tap in one phase are needed. For the best case, mixed load [five taps per phase, see expression (6)], 900 thyristors (297 000

(17)

(18)

where 6 is the number of windings (five main windings and one short-circuit winding, see Sections III-B and III-C.4). 1.2 is the core section overrating if the commutating instant is optimized (see III-D). B. Stabilizers With Series Compensating Transformer and Fault Current Deviation (Solutions 2 to 5) They show the advantages listed in Sections II-E and III-B as follows.

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• The power managed electronically is, typically, 15% of the nominal. • The series association of thyristors is replaced by parallel association, which is more reliable and economical. • Overload in the switches due to line faults is practically eliminated. • The switches achieve galvanic isolation so that the assembly can be separated from the transformers and fixed to ground voltage. Both thyristor isolation and live maintenance become easier. • The stabilizer can be settled at any point of the line, with no connection to the feeding line transformer. The general circuit for these solutions is given in Fig. 10. As shown in Fig. 3, and in Section II-E, Fig. 6(b), (c), the secondary of the series compensating transformer CT adds or subtracts the voltage required by the line at the settling site chosen. The voltage needed to feed the primary is obtained by means of a shunt transformer ST and associated taps. The power managed by CT is proportional to the voltage added-subtracted to the line voltage, typically 15% of the nominal power as advised in Section II.E. This figure also shows the power managed by the switches. The optimum design (minimum managed power) seeks equal voltages added and subtracted by the secondary of CT in the extreme conditions. If the combination of required line input and output voltage limits, MT drop and line drop do not result in an overall symmetrical voltage addition and subtraction, an extra balance winding in the ST primary, in series with CT primary, is needed to attain balanced operation of CT. The extra winding works as an autotransformer with the main primary winding and adds half the difference between the system adding and subtracting voltages. For the example in Section III-B, the extreme conditions are: • Extreme condition 1: Minimum input voltage, nominal load. The maximum phase-neutral voltage to be added by the stabilizer is (numerical values for the example given in III-B)

kV

(19)

• Extreme condition 2: Maximum input voltage, no load. The maximum phase-neutral voltage to be subtracted is

kV

(20)

The phase-neutral voltage added by the balance winding becomes

and the voltage added by the CT secondary in extreme condition 1 (equal to the voltage subtracted in extreme condition 2) is

kV

(22)

This is a very economical configuration if the voltage given . The operation and by expression (22) is lower than 0.3 commutation modes seen in Section III-C also apply to this group of solutions. Besides the stated reduction of the power managed by the in the example, see (22)), CT switches (reduced to 0.145 helps to drastically cut down their overcurrent due to line faults connected by means of an extra fault deviation switch across the primary, see Fig. 11. Once the fault is detected, the conducting main switches are inhibited turning off at the first current zero crossing. One half-cycle later (to avoid internal is turned on absorbing the line over-current short-circuit) reflected into the CT primary until the ordinary line protection switch opens (normally after 100 ms maximum). In this way the maximum line current (typically, 12.5 times the nominal) reflected to CT primary is supported by the main switches during no more than during one or two half cycles, and by are usually open, they do not 100 ms. As the thyristors of need cooling heatsinks. Under fault condition, the stabilizing operation is inhibited and the equipment is seen by the line as a short-circuited CT in series with the balance winding of ST. The rising effect of ST in the line fault current is compensated approximately by the short-circuit impedance of CT. 1) Single Switch-Comb Circuit: The single-comb configuration, Fig. 11, shall be first studied in detail for the sake of conceptual and circuit simplicity. a) Thyristor Rating: Tap thyristors: The line current reflected in the CT primary is the steady state current in any conducting tap switch for the nonshared load, and the maximum value of this current for mixed load. For single-comb, Figs. 6(b) and 11, the voltage seen by the extreme tap switches is the secondary voltage of ST, which must be chosen to take maximum advantage of available medium voltage thyristors without series association. For using thyristors of 4 500 V repetitive peak , for instance, the secondary rms voltage of ST V. Following the example must be 4 500 V in Section III-B, the balanced configuration in Figs. 10, 11, and supposing, for the sake of simplicity, that the common tap is in the middle of ST secondary, the sec./prim. ratio of CT (ignoring voltage drop in switches and ohmic drop in coils), according to (22), is

V V kV

(21)

(23)

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where is the voltage drop in CT for nominal load in per unit of its own voltage. Extending to every tap switch the voltage requirements for the extreme taps, each switch withstands

MVA

kV A rms during

(24)

A nonrepetitive ms peak V repetitive peak

(25) (26)

Introducing the security coefficients seen in Section IV-A-1, the nominal ratings for every tap semiconductor ensemble become A A rms during h A A nonrep ms peak V V repetitive peak as ab initio

(27) (28) (29)

By using the thyristor Westcode K1974ZC450 as in the direct configuration in Section IV-A-1, two units in parallel are needed. After the ratings related in Section IV-A-1 for a single SCR, the ratings for the tap ensemble (four SCR: two paralleled antiparallel pairs) become A

change with respect to a standard nontap transformer of the nominal power (10 MVA in the example). c) Effective Power of the Compensating Transformer CT: To guarantee independent voltage stabilization of the phases, three single-phase compensating transformers are needed. The power managed by one CT in the extreme condition 1 (see IV-B) is the product of the added phase voltage, expressed in (22), by the nominal line current . Taking into account the 1 h overload as in (17), the CT effective power becomes

MVA

(33)

d) Effective Power of the Shunt Transformer ST: In extreme condition 1, the single phase power given by the balance winding is the product of the voltage expressed in (21) by the nominal line current . The single-phase power given by the upper half of the secondary to CT is given by (33) plus CT losses. The lower half manages similar power in the extreme condition 2. A 1.4 factor is applied as the secondary halves do not operate simultaneously. Approaching the sum of unitary losses in CT, the coils and the switches to four times the CT , the equivalent voltage unitary drop under load 50 Hz power of ST becomes

MVA C permanent

(30)

A A nonrepetitive V repetitive peak

1499

ms peak

(31) (32)

that surpasses the requirements given by (27) and (28). The final result is four individual thyristors per tap, which, for three-phase equipment, means 12 thyristors (3 960 ) per tap in one phase. For the best case (mixed load) of the proposed example (5 taps per phase), this gives 60 thyristors (19 800 ) for the equipment. Coil short-circuit switches are ignored as in IV-A-1. withstands the fault Fault deviation thyristors: Switch line current reflected to the primary of CT and, for the singlecomb circuit (Fig. 11), supports half the voltage seen by the tap switches, as one side of the CT primary is connected to the common tap. After the security coefficients seen in IV-A-1, the is expressed in (25) exrepetitive peak current rating for tended to 100 ms. The voltage is half the value expressed in can be made, for unification, with 12 (26). Three switches units K1947ZD450 (two paralleled antiparallel pairs per phase) used for the taps, although underused in voltage. This represents an additional 3 960 to the cost of the equipment. b) Mains Transformer Power: The equipment is independent of the mains feeding transformer which does not need to

(34)

e) Equivalent 50 Hz Transformer Power of the Multi-Winding Coil: As for the direct configuration, the mixed load mode is considered. The value given in (18) is also valid for the compensating configurations as, although the , its current is multiplied winding voltage is now divided by by the same ratio. 2) Double Switch-Comb Circuit: In the double-comb configuration [Fig. 6(c)]—more suitable than the single-comb for high power, as in the example (10 MVA)—we take (as in IV-B-1-a) 1 273 V again for the complete ST secondary. The voltage stress ratio and the current of the switches does not change. The in the taps are divided by two, as the voltage in CT primary for the extreme conditions is twice the value for single-comb. For the example, shared load, 6 taps per comb result, see (9). Following the calculus procedure of IV-B-1 for the single-comb, the results are: a) Thyristor Rating: Tap thyristors: The current demand per tap is given by (27) and (28) divided by two. The repetitive peak voltage is given by (29). By using the same K1974ZC450, a single antiparallel pair per tap (no parallel or series associataps per phase) 36 tion needed) results. For mixed load ( thyristors (11 880 ) for the equipment are needed. In both the single and the double-comb circuit, the cost of the (see III-C-4) to shorcircuit the coil, if any, has small switch not been considered.

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Fault deviation thyristors: Switch withstands twice the voltage (4 500 V repetitive peak) and half the current (expression (25) divided by two: 6 774 A repetitive peak during 100 ms) needed with the single-comb. For the example, six units (one antiparallel pair per phase) of the aforementioned thyristor are needed, adding 1 980 to the cost of the equipment. Comparison of the thyristor rating with direct configuration: The results in IV-B-1.a and IV-B-2.a for the assembly of tap and fault deviation thyristors show that the compensating transformer configuration improves the poor expectations of early economical studies carried out excluding this configuration [2], [3]. Although both the single and the double switch-comb drastically reduce the number of thyristors, the double-comb has a clear advantage. b) Mains Transformer and Compensating Transformer Power: The same considerations than for single-comb apply to the mains transformer. The compensating transformer has the same secondary than in the single comb. The primary has half the current and twice the voltage so that the same effective power (33) applies. The ratio ’ is half the value given in (20). c) Effective Power of the Shunt Transformer ST: The working conditions of the balance winding are the same as that for single-comb. Only one secondary is needed with 1 273 V rms to have the same nominal repetitive peak voltage rate in is the switches than for single-comb (4 500 V). The ratio half the value given in (23). Equation (34) for the equivalent 50 Hz power of ST with a single-comb, becomes the following for the double-comb:

MVA

(35)

d) Equivalent 50 Hz Transformer Power of the Multi-Winding Coils: For the proposed example, mixed load, double comb, two multi-winding coils per phase with three tap coils (plus one short circuit coil) are needed. The voltage winding is twice and the current is half the respective values for single comb. By modifying expression (18), the equivalent power of each coil becomes

MVA

(36)

TABLE II THYRISTORS AND MAGNETIC COMPONENTS IN DIRECT AND COMPENSATING CONFIGURATIONS FOR THE 10 MVA STABILIZER EXAMPLE INTRODUCED IN =/UNIT, MARCH, III-B (CHOOSEN THYRISTOR: WESTCODE K1947ZC450 (330 C 2007) COMPARED TO THE CLASSICAL ELECTROMECHANICAL SOLUTION

magnetic components have to be considered for a more complete comparison. In the first column, the three-phase direct mechanical configuration (as appears in Fig. 6(a) for single-phase) is given. The thyristor’s cost of the compensating transformer, double-comb, configuration is lower than the cost of the electromechanical switch assembly (Fig. 2, upper, left). V. LIVE WORKING INSTALLATION AND REPAIR As a consequence of their independence from the line feeding transformer, the stabilizers derived from the compensating configuration proposed in IV-B can be connected to any point of the line to regulate the downstream voltage or to adjust the voltage of a derived line. The stabilizer (in the single-phase equivalent circuit) is seen by the line as a dipole (see Fig. 11) similar to the medium voltage disconnecting switches used to isolate parts of a complex net [49], [50]. Subsequently, the stabilizer can be connected and removed by using the same live working installation procedure [45]–[47] developed by the electric companies for such switches without service interruption, see Fig. 12 and Table III. Since the power-electronic assembly is at ground potential, some maintenance operations can be performed in live status. VI. AUTOREMOVAL In case of stabilizer internal malfunction, the tap switches and associated circuits can be functionally eliminated by following the removal sequence in Table III up to point 3, keeping in coordination the OLTC control and the MDS control. The line can remain operative during the stabilizer reparation in situ or the removal and reparation in a workshop, in this case following the complete removal sequence.

C. Quick Economical Comparison of Direct and Compensating Transformer Configurations

VII. CONCLUSION

Table II compares the thyristors and the magnetic components needed in the direct and the compensating configurations, mixed load, to implement the 10 MVA stabilizer of the example introduced in III-B. The compensating configuration (both with single and double-comb) adds extra advantages as a free connection to any point of the line (see Section V). The costs of the

The feasibility of voltage stabilizers for medium voltage lines based on electronic tap changers has been reviewed drawing from the advantages of using a multi-winding commutating coil, largely proved in power distribution applications. This coil reduces the commutation stress of the thyristors so that overrating is not necessary.

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TABLE III LIVE WORKING INSTALLATION STEPS PLANNED FOR THE COMPENSATING TRANSFORMER CONFIGURATION STABILIZER PROPOSED IN FIGS. 10–12

The configuration proposed needs additional electromagnetic components (multi-winding commutating coil, shunt transformer and compensating transformer) whose power is keep under reasonable limits (about 15% of the nominal power). The advanatages of the proposed stabilizer are as follows. • Isolation of the electronic ensemble with respect to the line. It can be connected to ground potential. • Series association of the thyristor is not needed. For power of more than 10 MVA, an easy parallel association would be needed. • The fault line current is derived from the tap thyristors to a specific, not expensive, fault deviation thyristor switch. • Connection to any point in the line without any connection to the head of the line feeding the transformer. • The configuration that is chosen facilitates eventual live working installation and repair. REFERENCES

Fig. 12. (a) Single-phase conceptual scheme for the line connection of the compensating configuration stabilizer proposed in Section IV-B. In parallel with the stabilizer there is a classic mechanical disconnecting switch (MDS). (b) Stabilizer inhibited, MDS open, S closed. (c) Equivalent circuit for (b) status. Only the short circuit impedance of CT remains in series with the line. (d) Possible ground installation at a generic line point allowing a live working connection and removal. Single-phase representation after Fig. 1 of [49]. Aerial installation as in Fig. 4, right, is also possible.

The study confirms the high costs of thyristors previewed [2], [3] if the classic direct configuration used in most mechanical tap changers is considered. Nevertheless, it also demonstrates drastic reductions and the possibility of adding technical advantages by using compensating transformer configurations, well-known in low- and medium-voltage low power applications. This paper shows the feasibility of a 69 kV/34.5 kV, 10 MVA stabilizer for long medium voltage lines by using 42 standard thyristors rated 1 847 A, 4 500 V, with no need of series, nor parallel, association.

[1] Transformers Committee of the IEEE Power Engineering Society, , IEEE Standard Requirements for Load Tap Changers. New York, 1995. [2] P. Wood, V. Bapat, and R. P. Putkovich, “Study of improved loadtap-changing for transformers and phase-angle regulators,” EPRI Rep. EL-6079, Nov. 1988, p. 148. [3] “Study of improved load-tap-changing for transformers and phase-angle regulators,” Apr. 1990, p. 108, EPRI Staff, EPRI, Rep. EL-6764. [4] S. Martinez, E. BOAR, and S. A. , “Perfeccionamientos en equipos electronicos para regulación de tension alterna. . .,” Spanish Patents 500.523, (Oct. 1981), 500.524 (Oct. 1981), 522.497 (Oct. 1984). [5] S. Martinez, “Estabilizadores de c.a. por pasos con carga compartida,” Mundo Electron., no. 166, pp. 127–135, Oct. 1986. [6] D. Gahigiro, M. A. Erro, and S. Martinez, “AC tap changing stabilizers with limited commutating current,” in Proc. IEEE 1st Power Electronics Int. Congr., Cuernavaca, México, Aug. 1992, pp. 195–208. [7] S. Martinez, Alimentacion Electronica de Equipos Informaticos y Otras Cargas Criticas, ser. Electrotecnologias. Aravaca, Madrid: McGrawHill Interamericana, IBERDROLA, EVE), 1992. [8] G. Villegas, J. Vaquero, R. Echevarria, S. Horta, M. A. Perez, and S. Martinez, “Quasi-resonant fast on-load tap changer stabilizer,” presented at the IEEE-Cenidet, Morelia, Mexico, Oct. 1998. [9] J. Vaquero, J. C. Campo, S. Martinez, and M. A. Perez, “Fast on-load tap changers possibilities using new power devices,” presented at the Min, 10th Annu. Meeting EWG/IEEE/IAS/IPCC&PEDCC, Aveiro, Portugal, Jun. 1–2, 1999. [10] J. Vaquero, “AC on-load tap-changing fast voltage stabilizers for power quality improvement,” Ph.D. dissertation, Universidad a Distancia, Madrid, Spain, Apr. 2000. [11] J. C. Campo, “Functional and topological analysis of ultra-fast differential ac stabilizers,” Ph.D. dissertation, Universidad de Oviedo, Gijón, Spain, Oct. 2000.

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[12] J. C. Campo, J. Vaquero, M. A. Perez, and S. Martinez, “Fast twotaps changing AC stabilizer with seminatural switching,” Elect. Power Quality Utilization, vol. 6, no. 1, pp. 111–114, Dec. 2000. [13] F. Barrero, S. Martínez, F. Yeves, and P. M. Martinez, “Active power filters for line conditioning: A critical evaluation,” IEEE Trans. Power Del., vol. 15, no. 1, pp. 319–325, Jan. 2000. [14] F. Barrero, S. Martinez, F. Yeves, F. Mur, and P. M. Martinez, “Universal and reconfigurable to UPS active power filter for line conditioning,” IEEE Trans. Power Del., vol. 18, no. 1, pp. 283–290, Jan. 2003. [15] J. C. Campo, J. Vaquero, M. A. Perez, and S. Martinez, “Dual-tap chopping stabilizer with mixed seminatural switching. Analysis and synthesis,” IEEE Trans. Power Del., vol. 20, no. 3, pp. 2315–2326, Jul. 2005. [16] J. Vaquero, J. C. Campo, S. Monteso, S. Martinez, and M. A. Perez, “Analysis of fast onload multitap-changing clamped-hard-switching AC stabilizers,” IEEE Trans. Power Del., vol. 21, no. 2, pp. 852–861, Apr. 2006. [17] J. Vaquero, J. C. Campo, S. Monteso, S. Martinez, and M. A. Perez, “Synthesis of fast onload multitap-changing clamped-hard-switching AC stabilizers,” IEEE Trans. Power Del., vol. 21, no. 2, pp. 862–872, Apr. 2006. [18] G. Musgrave and D. O’Kelly, Improvement for Power System Transmission by Solid-State Techniques. London, U.K.: Inst. Elect. Eng. Conf. Publ., Dec. 1974, pp. 228–233, no. 123. [19] G. H. Cooke and K. T. Williams, “New thyristor assisted diverter switch for on load transformer tap changers,” Proc. Inst. Elect. Eng. B, vol. 139, no. 6, pp. 507–511, Dec. 1974. [20] R. Shuttleworth, X. Tian, C. Fan, and A. Power, “New tap changing scheme,” Proc. Inst. Elect. Eng., Elect. Power Appl., vol. 143, no. 1, pp. 108–112, Jan. 1996. [21] A. Krämer and J. Ruff, “Transformers for phase angle regulation considering the selection of on-load tap-changers,” IEEE Trans. Power Del., vol. 13, no. 2, pp. 518–525, Apr. 1998. [22] M. T. Bishop, J. D. Foster, and D. A. Down, “Single-phase voltage regulators and three-phase systems,” IEEE Ind. Appl. Mag., vol. 2, no. 4, pp. 38–44, Jul./Aug. 1996. [23] CPS staff, Cooper Power Systems, , Voltage regulators Bull. B22597020, Jul. 2005. [24] J. Arrillaga, “A static alternative to the transformer on-load tap-changer,” IEEE Trans. Power App. Syst., vol. PAS-99, no. 1, pp. 86–91, Jan./Feb. 1980. [25] R. M. Mathur and R. S. Basati, “A thyristor controlled static phaseshifter for AC power transmission,” IEEE Trans. Power App. Syst., vol. PAS-100, no. 5, pp. 2650–2655, May 1981. [26] E. C. Servetas and V. Vlachakis, “A new AC voltage regulator using thyristors,” IEEE Trans. Ind. Electr. Control Inst., vol. IECI, no. 2, pp. 140–145, May 1981. [27] G. Güth, G. R. Baker, and P. Eglin, “Static thyristor-controlled regulating transformer for AC transmission,” presented at the Inst. Elect. Eng. Int. Conf. Thyristor and Variable Static Equipment for AC and DC Transmission, London, U.K., Nov./Dec. 1981. [28] F. Q. Yousef-Zai and D. O’Kelly, “Solid-state on-load transformer tap changer,” Proc. Inst. Elect. Eng., Electr. Power Appl., vol. 143, no. 6, pp. 481–491, Nov. 1996. [29] O. Demirci, D. A. Torrey, R. C. Degeneff, F. K. Schaeffer, and R. H. Frazer, “A new approach to solid-state on load tap changing transformers,” IEEE Trans. Power Del., vol. 13, no. 3, pp. 952–961, Jul. 1998. [30] P. Bauer and S. W. H. de Haan, “Electronic tap changer for 500 kVA/10 kVA distribution transformers: Design, experimental results and impact in distribution networks,” presented at the IEEE-IAS Annu. Meeting, St. Louis, MO, Oct. 1998. [31] H. Jiang, R. Shuttleworth, B. A. T. Al Zahawi, X. Tian, and A. Power, “Fast response GTO assisted novel tap changer,” IEEE Trans. Power Del., vol. 16, no. 1, pp. 111–115, Jan. 2001. [32] J. Faiz and B. Siahkolah, “New solid-state onload tap-changers topology for distribution transformers,” IEEE Trans. Power Del., vol. 18, no. 1, pp. 136–141, Jan. 2003. [33] J. Faiz and B. Siahkolah, “Optimal configurations for taps of windings and power electronic switches in electronic tap-changers,” Proc. Inst. Elect. Eng., Gen. Transm. Distrib., vol. 149, no. 5, pp. 517–524, Jul. 2002. [34] J. F. Ruiz-Garrido, Security coefficients for power thyristors in electronic power converters SEPSA, Pinto, Madrid, Spain, Tech. Rep. 14, Jun. 2000.

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[35] K. Imhof, F. Oesch, and I. Nordanly, “Modelling of tap-changer transformers in an energy management system,” IEEE Trans. Power Syst., vol. 11, no. 1, pp. 428–434, Feb. 1996. [36] B. Kasztenny, E. Rosolowski, J. Izykowski, M. M. Saha, and B. Hillstrom, “Fuzzy logic controller for on-load transformer tap changer,” IEEE Trans. Power Del., vol. 13, no. 1, pp. 164–170, Jan. 1997. [37] M. Crape, L. Gertmar, A. Haböck, W. Leonhard, and D. Povh, Eds., “Power electronics and control by micoroelectronics in future energy systems,” in Proc. Lausanne Conf. Rep., EPE J., Apr. 2000, vol. 10, no. 1, pp. 6–10, 40–46. [38] P. Kang and D. Birtwhistle, “Condition assessment of power transformer on-load tap-changers using wavelet anaysis,” IEEE Trans. Power Del., vol. 16, no. 3, pp. 394–400, Jul. 2001. [39] P. Kang and D. Birtwhistle, “Condition monitoring of power transformer on-load tap-changers. Part 1: Automatic condition diagnostics,” Proc. Inst. Elect. Eng., Gen. Transm. Distrib., vol. 148, no. 4, pp. 301–311, Jul. 2001. [40] P. Kang and D. Birtwhistle, “Condition assessment of power transformer onload tapchangers using wavelet anaysis and self-organizing map: Field evaluation,” IEEE Trans. Power Del., vol. 18, no. 1, pp. 78–84, Jan. 2003. [41] M. M. Adibi, R. A. Polyak, I. A. Griva, L. Mili, and S. Ammari, “Optimal transformer tap selection using modified barrier-augmented lagrangian method,” IEEE Trans. Power Syst., vol. 18, no. 1, pp. 251–257, Feb. 2003. [42] G. N. Korres, P. J. Katsikas, and C. Contaxis, “Transormer tap setting observability in state estimation,” IEEE Trans. Power Syst., vol. 19, no. 2, pp. 699–706, May 2004. [43] Consideraciones a la Regulación de Tensión en ST a Los Niveles de MT. Bilbao, Spain, Sept. 1992, Personal DERED/PLANIFICACIÓN, Manual de Organización y Procedimientos, IBERDROLA S.A.. [44] F. S. Roesner and T. J. Sillers, “Understanding voltage regulators,” Elect. Power Int., p. 32, 58, Mar. 1995. [45] Trabajos en Tensión, Método a Distancia, Instalación de Órgano de Corte de Red, Simple Circuito, Aislamiento de Amarre. Bilbao, Spain, Sep. 1996, Personal Distribución y Clientes, Manual Técnico de Distribución y Clientes MTDYC 2.25.70, IBERDROLA-DIDYC. [46] Office of the Queensland Parliamentary Counsel, “High voltage live line work,” in Electric Safety Regulation 2002 (Division 3). [Online]. Available: http://www.deir.qld.gov.au/electricalsafety/business/workers/live/safely/index.htm. [47] “Manual de Construção e Manutenção de Redes de Distribução-Standard Operation Procedures for Live-Line Works in Spacer Cable,” CEMIG, Minas Gerais Brazil, 1998. [48] Live Working, IEC/TC78, , Int. Elect. Committee., 2003. [49] Órgano de Corte en Red4a ed. May 2002, NORMA IBERDROLA, NI 74.53.01. [50] “Fichas técnicas de lineas aereas de media tensión 12/20 kV,” Regulador de Tensión Para Lineas de MT; LAM70 Organo de Corte en Red de Intemperie (OCRI), Feb. 2003, IBERDROLA, DITEC-NOMAN, MT 2.03.96-II, 1a Ed.: LAM68.

Salvador Martínez García (M’88–SM’90) was born in Spain in 1942. He received the M.Sc. and Ph.D. degrees in electrical engineering from the Polytechnic University of Madrid, Madrid, Spain, in 1966 and 1969, respectively. He was Associate Professor at the Polytechnic University of Madrid from 1975 to 1979 and at the National Distance University of Spain, Madrid, from 1979 to 1982, where he has been Full Professor since 1982. He was a Design Engineer of power-electronics equipment in several companies for eight years. His research interests are in integrated magnetics and power-line conditioners. Prof. García was a member of the IEEE Industry Applications Society Industrial Static Converter Committee EWG from 1990 to 1997.

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MARTÍNEZ GARCÍA et al.: FEASIBILITY OF ELECTRONIC TAP-CHANGING STABILIZERS FOR MV LINES

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Juan Carlos Campo Rodríguez (M’97) born in Spain in 1970. He received the M.Sc. degree in electrical engineering from the Oviedo University, Asturias, Spain, in 1995 and the Ph.D. degree in electrical engineering in 2000. Currently, he is an Associate Professor in the Department of Electrical and Electronic Engineering at Oviedo University, where he has been since 1995. He has conducted research on a fast-tap changers for ac voltage control. His current research interests include power quality, line conditioners, and tap changers.

Alfonso Ibarzábal Segura was born in Spain in 1945. He received the M.Sc. degree in electrical engineering from the Polytechnic University of Bilbao, Bilbao, Spain, in 1974. He is the Technical Director of INCOESA, dedicated to the development and manufacture of distribution power transformers, medium-voltage transformers, and onload tap changers. Mr. Segura is a member of the Standards Technical Committee of AENOR CTN207/GT 14, “Power Transformers.”

José Antonio Jardini (M’66–SM’78–F’90) was born in Brazil in 1941. He received the M.Sc. and Ph.D. degrees from the Polytechnic School of São Paulo University, EPUSP, São Paulo, Brazil, in 1971 and 1973, respectively. He was with Themag Eng. Ltd. for 25 years, where he conducted numerous studies on power systems and participated in many important Brazilian projects, including Itaipu. Currently, he is a Professor in the Departamento de Energia e Automação Electricas da Escola Politécinca da Universidade de São Paulo, São Paulo, Brazil, teaching power analysis and digital automation. He has conducted research work in power systems for many Brazilian utilities. Prof. Jardini is a member of the CIGRE Working Group on HVDC transmission.

Pedro María Martínez Cid was born in Spain in 1956. He received the M.Sc. degree in electronic and control engineering from the National Distance University of Spain, Madrid, in 1989. Since 1995, he has been a Lecturer with the National Distance University of Spain at the Associated Centre of Bilbao. In 1979, he joined the electrical company IBERDROLA S.A. where he has conducted research-and-development projects on flexible ac transmission systems (FACTS), including the ESTRAP Project among them. He is with the Division for Innovation, Power Quality, and Ambient Studies. Mr. Cid is a member of several international committees for the promotion of FACTS.

Joaquín Vaquero López was born in Spain in 1968. He received the M.Sc. degree in electrical engineering from the Polytechnic University of Madrid, Madrid, Spain, in 1994 and the Ph.D. degree from the National Distance University of Spain, Madrid, in 2000. In 1995, he joined the Department of Electrical, Electronic, and Control Engineering as Assistant Professor at the Universidad Rey Juan Carlos. He was a Design Engineer of power-electronics equipment with SEPSA from 2000 to 2007. He joined the Electronics Technology Department at Rey Juan Carlos University of Madrid as an Associate Professor with research interests in fast multi-tap changers and equipment to improve power quality.

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