Synthesis of Fast Onload Multitap-Changing Clamped-Hard-Switching AC Stabilizers

862 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 21, NO. 2, APRIL 2006

Synthesis of Fast Onload Multitap-ChangingClamped-Hard-Switching AC Stabilizers

Joaquín Vaquero López, Juan Carlos Campo Rodríguez, Member, IEEE, Santiago Monteso Fernández,Salvador Martínez García, Senior Member, IEEE, and Miguel Angel Pérez García

Abstract—Power insulated-gate bipolar transistor (IGBT) andother gate turnoff switches have allowed to analyze, in a previouspaper, four new circuits for onload-tap-changing ac stabilizers thattypically permit to make up to 7 or 45 (depending on the particularcircuit solution) changes per half cycle of the mains voltage. Thissecond and last paper considers the synthesis aspects of these sub-cyclic stabilizers giving design rules, and a synthesis calculus sheetactualized in a 100-kVA numeric example. A comparative table forthe four circuit solutions, involving both technical and economicalfigures, is given as the final step of the iterative analysis and syn-thesis cost comparative method performed along the research. Pre-liminary experimental results are obtained from a 1-kVA prototypeand a first evaluation of the possibilities of reducing voltage distor-tions is demonstrated in the fields of voltage low harmonics andflicker by means of a high-accuracy graphic simulator.

Index Terms—AC voltage stabilizer, onload tap-changer, powerconditioning, power quality, voltage control.

I. INTRODUCTION

MODERN power transistors and other power switchesshowing easy turn-on and turn-off operation invite the

search for new topologies in the field of onload tap-changingac voltage stabilizers able to commutate several times everyhalf cycle of the mains voltage (subcyclic commutation).The analysis of four topologies based on the hard switchingtechnique with clamped voltage and their historic backgroundhas been done in a previous paper [1]. These solutions werededicated to multitap stabilizers with and without a compen-sating transformer in the range of kilowatts to megawatts forlow- and medium-voltage ac installations. That circuit analysiscompleted the first step of a systematic analysis and synthesiscost comparative search method [2] on the topologies for thisnew stabilizers family, originally done over nine solutions. Thecircuits selected [1, Sec. III] show the ability to commutate in0.15 ms (solution 1: nonshared load) and in 1 ms (sol. 2: sharedload; sol. 3: nonshared load; sol. 4: mixed load) allowing about7 to 45 tap changes every half cycle of the mains voltage.

Manuscript received October 6, 2004; revised July 6, 2005. This work wassupported in part by the Spanish Ministry of Science and Technology and in partby the IBERDROLA Program of Electric Research. Publication funds from theVicerrectorado de Investigación, Universidad Nacional de Educción a Distancia,Madrid, Spain. Paper no. TPWRD-00473-2004.

J. Vaquero López, S. Monteso Fernández, and S. Martínez García are withthe Department of Electrical, Electronic and Control Engineering, UniversidadNacional de Educación a Distancia (UNED), Madrid E 28040, Spain (e-mail:[email protected]; [email protected]; [email protected]).

J. C. Campo Rodríguez and M. A. Pérez García are with the Departmentof Electrical, Electronic, Computing and Systems Engineering, Univer-sidad de Oviedo, Gijón E 33204, Spain (e-mail: [email protected];[email protected]).

Digital Object Identifier 10.1109/TPWRD.2005.864041

TABLE IEQUIPMENT GENERAL SPECIFICATIONS (VALUES FOR A

TYPICAL SINGLE-PHASE, 230-V, 50-HZ, 100-kVA STABILIZER

ARE INCLUDED IN BRACKETS)

TABLE IIAUXILIARY SPECIFICATIONS FOR SOLUTION 3 (MIXED LOAD), COMPENSATING

STRUCTURE. EVERY VOLTAGE DROP IS DEFINED AS PROJECTED ONTO THE

MAINS VOLTAGE VECTOR. (VALUES FOR A TYPICAL SINGLE-PHASE, 230-V,50-HZ, 100-kVA STABILIZER ARE INCLUDED IN BRACKETS)

II. SYNTHESIS

The synthesis step of the search method obtains, for a par-ticular solution, the mathematical expressions of the electricalrates and working conditions of the components as a functionof the equipment general specifications (Table I) and of theauxiliary specification of the given solution (Table II), in thispaper solution 3—mixed load—with compensating structure.For space reasons, only this solution and structure shall be fullydeveloped. In order to add numerical meaning to this very sym-bolic step of the search, Tables I and II, the synthesis formulassheet (Table III) and the comparative Table IV include, in

0885-8977/$20.00 © 2006 IEEE

VAQUERO LÓPEZ et al.: SYNTHESIS OF FAST ONLOAD MULTITAP-CHANGING AC STABILIZERS 863

TABLE IIIPROTOTYPES SYNTHESIS CALCULUS SHEET FOR SOLUTION 3 WITH COMPENSATING STRUCTURE [SEE FIG. 1a)], GIVING THE DESIGN PARAMETERS AND STRESS

VALUES OF THE COMPONENTS. (SOME EXPRESSIONS ARE NOT SOLVED OR SIMPLIFIED DELIBERATELY TO SHOW THEIR ORIGIN) (IN BRACKETS, NUMERICAL

VALUES RESULTING FOR A TYPICAL SINGLE-PHASE, 230-V 50-Hz 100-kVA STABILIZER EXAMPLE, AS DEFINED IN TABLES I AND II)


TABLE IVCOMPARISON AMONG THE PROPOSED SOLUTIONS IMPLEMENTED AROUND A SIMPLIFIED CALCULUS SHEET. SOLUTION 3 COLUMN REPEATS SOME OF

THE RESULTS GIVEN IN TABLE IV TO MAKE THE COMPARISON EASIER. LETTERS IN PARENTHESIS REFER TO EXPRESSIONS IN TABLE IV,CORRESPONDING TO SOLUTION 3. (IN BRACKETS, NUMERICAL VALUES RESULTING FOR THE TYPICAL SINGLE-PHASE, 230-V, 50-Hz, 100-kVA

STABILIZER EXAMPLE CONSIDERED IN TABLE III AS DEFINED IN TABLES I AND II)

brackets, actual values for a typical example single-phase,230-V 50-Hz 100-kVA stabilizer.

Descriptions and formulas in this part are referred to in Fig. 1(which repeats Fig. 1(b) in [1] adding coil windings and auxil-iary switches). The number of taps selected for that figure (8.24taps) is the round-down integer of the result (eight taps) derivedfrom expression (4) for the synthesis example in Table III.

Solution 3 (mixed load) has been preferred to solutions 1 and2 to show this synthesis step because of its higher complexity.Solution 4, although includes the extra auxiliary short-circuitwinding, does not add significant conceptual meaning to thefamily. Any other solution can be easily synthesized followingthe method here exposed for solution 3.

A. Main Tap Distribution and Transformer Ratios

The approximate number of changes of conducting states(including sharing and no-sharing working modes) needed

for a given input and output voltage deviation, addingan empirical 0.5 term to account for sample and controlerrors, and assuming that the ratio of the main transformerapproximates to 1, is: see (1) at the bottom of the page.is the transforming ratio of the compensating transformer(CT; Fig. 1) (turns 7-3/turns 5-6) and , the ratio of themain transformer (MT; Fig. 1) (turns 7-4/turns 1-2). Ratio

- -gives the position of the common terminal Com.Tap in the secondary of the MT. Ratio

- -gives the turns of the secondary lower winding (Com. Tap toTap 8) related to the complete winding.

In general, the number of conducting states needed to accom-plish a given number of changes is

(2)

(1)


Fig. 1 (a) Single-phase multitap changer with a compensating structure showingthe main transformer (MT) with eight taps in its secondary winding. Dependingon the active tap, a voltage in phase, or counter phase, with respect to themains voltage is applied to the primary of CT (5-6). Its secondary (3-7) addsor subtracts, respectively, to the voltage of the MT secondary. Only the powerfraction needed to compensate for deviations of the mains voltage is controlledby the switching devices. (b) Semiconductor structure of the main switches Sto S , and the auxiliary switches S and S .

and for solution 3 considered, the number of conducting statesderived from a given number of taps is

(3)

Solving relations (1)–(3), the approximate number of tapsneeded is (4), shown at the bottom of the page.

The defined transforming ratios must be derived, taking intoaccount the voltage drop in components in two extreme condi-tions (EC).

EC 1—Minimum input voltage, nominal load: Voltagedrops in components subtract from the output voltage,acting against the stabilizer operation. The tap changeradds its maximal voltage selecting an extreme tap [tap1 in Fig. 1(a)].EC 2—Maximum input voltage, no load: Voltagedrops subtract from the input voltage, now facilitatingthe stabilizer operation (notice the plus sign in thesecond term of (7). The tap changer subtracts themaximum voltage selecting the other extreme tap [tap8 in Fig. 1(b)].

Fig. 2. Graphic method to calculate the intertap voltages: Segment a isthe output voltage variation range, b the maximum output voltage, c is themaximum input voltage, and d is the voltage between Tap 8 and the midpointof Taps 8-7. Using the relationship b=a = c =d , d is obtained. Subtractingd from c , a new value c is the result, iterating to reach the other extremecondition (voltage between the medium point of Taps 2-1 and Tap 1), all of theintertap voltages are obtained.

In the extreme condition 1, the load voltage must show theallowed negative deviation , so

(5)

where is the secondary voltage of MT for this condition.

In the extreme condition 2, the output voltage must show theallowed positive deviation , so

(6)

where is the secondary voltage of MT for this condition.

In order to reduce the equivalent power (see its definitionbelow) of CT to a minimum, the position of Com. Tap, not nec-essarily in the midpoint of the secondary of the MT must guar-antee that the primary voltage module of CT be the same forboth extreme conditions

(7)

Solving conditions (5)–(7), the values of , andare obtained. The one’s complement of the latter gives

.The number of turns between taps is not constant, as the main

criteria to calculate them is to adjust to its maximum allowedvalue ( , denoted by in Fig. 2); the voltage vari-ation in the load for any tap change. In this way the number of

(4)


changes needed (hence, the number or taps) is minimized. Sothe number of turns between upper taps (Tap 1 to Tap 2, etc.) ishigher than that between lower taps (Tap 8 to Tap 7, etc.) as theinput voltage in extreme condition 1 is lower than in extremecondition 2. Then, the turns between taps vary gradually alongthe secondary of MT and can be easily calculated by means ofFig. 2. In the geometrical relationship , the un-known term (the voltage between tap 8 and the medium pointof taps 7 and 8) can be derived arithmetically from the otherterms ( ; ; )and, repeating the calculation for every next pair of conductingstates, all and terms, as well as the corresponding intertapvoltages and turns, are obtained. The term at the bottom ofthat figure accounts for the sum of the voltage drops reduced tothe output terminals in the extreme condition 1, whose value isshown in (8) at the bottom of the page.

In practice, as the calculated turn number for each tap mustbe rounded to either the previous or the next integer, the tapposition in the main transformer can be derived also withenough precision designing Fig. 2 for the actual stabilizer ingraph paper. In both ways of calculation, arithmetic or graphic,it must be kept in mind that the combination of all segments

in Fig. 2 corresponds (for this example with compensatingstructure) to the whole MT secondary turns. A final handmadesmall adjustment of the tap positions can be eventually needed,always in search of an exponential tap distribution all along theMT secondary.

B. Equivalent Power of Transformers

The 50-Hz isolation transformer power (here referred asequivalent power), as defined in [2] (in short, power ratingof the simple isolation transformer for 50 Hz and sinusoidalwaveform that can be made using the same core and copperneeded to make an specific transformer or coil), has to bederived for MT, CT, and for the coil in order to obtain theircosts easily from common 50-Hz transformers price lists.

On deriving the equivalent power of MT, the worst case forthe current in the secondary upper half (Tap 1 to Com. Tap) is theextreme condition 1, as the primary current of CT(where the factor , which is also used for the MT primarycurrent, accounts for the modulus increment due to the magne-tizing current) adds to the nominal load current . The worstcase for the lower half (Com. Tap to Tap 8) is also the extremecondition 1, in which circulates. Note that in any conditionwith a selected tap under Com.Tap (from Tap 6 to Tap 8, or anypair among them in case of shared load), the primary current ofCT detracts to the load current circulating in the secondary ofMT. The positive deviations of the input voltage and the mag-netizing current are not considered in the final formulas for theequivalent power (see the results in Table III) because they areaccounted for in the ordinary design of the standard 50-Hz trans-formers to be consulted.

C. Auxiliary Taps and Multiple-Winding Coil

For the practical design of the multiple-winding coil, thevoltage , seen by any pair of active equalizing windings insharing mode, can be approximated to a medium value

(9)

The inductance of an individual equalizing winding isderived from the condition imposed in [1, Sec. III-C1], whichlimits the value of the short-circuit current between taps to60% of the nominal load (for the compensating structure in thiscase, to 60% of the nominal load reflected to the primary circuitof CT) resulting in

(10)

so that

(11)

The worst case for the magnetizing current, which determinesthe core section and the maximum induction, is the one ex-plained in [1, Sec. III-A2a], for the nonsharing mode, partic-ularized for a tap change during the maximum value of the loadcurrent . Also taking into account the overload factor

, the transforming ratio of CT and its magnetizing current(factor ), the design peak value becomes

(12)

where the addend must rise in case of overlapping tapchanges. Note that in sharing mode, the load current does notimpose any magnetic effect on the core, as its two halves circu-lating through two consecutive windings compensate each other.Both windings act as an autotransformer fed by . Then, themaximum instantaneous value of the magnetizing current (re-duced to a single winding) is two times the peak value of theequalizing current (as circulates along both windings).If is chosen to be bigger than 60% of the load current re-flected to the primary of CT, the value resulting from this con-dition exceeds (12) and must be taken as the worstcase.

The turns ratio of the auxiliary winding depends on theovervoltage allowed in the equalizing windings during the en-ergy-return interval and the voltage of the auxiliary windings

– – (see [1, Sec. III-A2b]), the latter chosen to ob-tain a convenient operating voltage in the auxiliary switches.The higher the overvoltage, the shorter the interval. A reason-able maximum value for the overvoltage, derived from the ex-periments, is two times , so that

(13)

(8)


The peak value of the current through the auxiliary windingand its associated switches is derived from the analysis in [1,Sec. III-A2b and Fig. 4(c)], assuming a tap change during themaximum value of the load current. Following the observationson deriving (12), we obtain:

(14)

As the duration of the recovery pulse depends inversely uponthe instantaneous value of the auxiliary taps voltage, in orderto avoid long tap-changing times, it can be advisable to inhibitchanges when the instantaneous value of the mains voltage is,say, less than one-third of its maximum, that is 9.5 before andafter the zero-crossing. The contribution of this part of the mainsvoltage is low for most practical purposes, so that such an in-hibition adversely affects the stabilizer performance to a lowdegree. To deduce the typical duration of the recovery pulse,the medium value of the mains voltage during the resulting freechanging period (0.6 of the maximum value) can be applied.For an instantaneous value of the load current equal to half ofthe maximum value, and modifying the expression given in [1,Sec. III-A2b] for compensating structure, we obtain

(15)

According to the triangular waveshape of this current (in [1],Fig. 4), and assuming tap changes along one mains cyclelasting at an averaged current of 50% of the value given in(14), the following rms value is obtained:

(16)

The coil equivalent power is derived easily after the windingsrms voltage and current values. See the result in Table III.

D. Switches

The selection of the switches is determined by both the repeti-tive peak of voltage and current. The worst case for an individualof each switch family (main and auxiliary) is applied usuallyto all members. For both families, the maximum stress voltagehappens for the extreme condition 2. For the main switches, thepeak voltage of the MT secondary

is applied to when is on. For the auxiliary switches,the peak voltage is given by the same expression affected by theratio .

The worst repetitive peak current for the main switches is thecurrent in the primary of CT for extreme condition 1 (denotedby on Table III) affected by the peak waveform factorand the overload factor . For the auxiliary switches, thatcurrent is given by (14).

The common component equivalent concept, helpful to es-timate manufacturing costs, can be also extended to capacitors

and semiconductors [2], [3]. It has been applied in Table IV, onlyfor the switches due to space reasons, following the rules givenin [2].

E. Synthesis Sheet

Table III summarizes the synthesis step of the search method.It shows the symbolic formulas for the design parameters andthe working conditions of the components as a function of theequipment specs given in Tables I and II. The formulas are ob-tained from the equations derived in the analysis Section III of[1] and the design relationships already considered in this syn-thesis section. The following general design assumptions arealso considered.

1) As electromagnetic components tolerate substantial over-load during 1 min, the overload factor has been ne-glected except for the current in semiconductors and forthe magnetizing current in the coil. No protection has beenconsidered against the output short circuit. The overcur-rent in the switches for this condition would be canceledby shortcircuiting the primary of CT through an extrathyristor pair in antiparallel, while inhibiting the switches.

2) The efficiency values given for the transformers in Table IIdo not appear explicitly in the synthesis (Table III) be-cause the synthesis formulas are clarified if the losses ef-fects are accounted for by means of the voltage drop pa-rameters, also given in Table II. For simplicity reasons,in the resulting expressions, the latter have been inten-tionally defined as active voltage drops (projected ontothe mains voltage vector). The efficiency parameters re-main in Table II to inform the designer about the qualityassumed in the transformers from which the numericalvalues given for the voltage drops are derived.

For stabilizers of more than 100 kVA with ten main taps ormore, it can be interesting in order to obtain a better profit of theMT secondary winding and the main switches, to connect ter-minal 6 of CT [Fig. 1(a)] to a second array of taps (distributedalso along the MT secondary), instead of to the common tap.This new array needs its own set of main switches, multiplewinding coil and auxiliary switches. The adding or detractingvoltage obtained for the CT primary is twice the voltage ob-tained with a single array for the same voltage stress in the mainswitches, which, in turn, divide by two the current in that pri-mary and in the main switches. The number of conducting statesderived from two equal arrays summing taps for solution3 is two units less than the one derived in (3) for a single array

(17)

III. SOLUTIONS COMPARISON

Solution 1 is faster, as the recovery interval is shorter due tothe low value needed for the equalizing coils inductance . So-lutions 2, 3, and 4 involve high values for and, hence, longercommutating times. This advantage of solution 1 is enhanced inequipment with a high intertap voltage (stabilizers with a high


mains voltage allowed deviations of mains voltage and not manytaps).

Solution 2 is preferable when the resulting current in theswitches approaches the limit of available insulated-gate bipolartransistors (IGBTs), as the current through each tap approacheshalf the load current. As in solution 4, the equalizing coils donot affect the output voltage in steady-state operation. It has thedisadvantage of needing an extra tap when compared to solution1 for the same number of conducting states. It is worthwhileto highlight that the supercyclic (commutation every one ormore half cycles) version of solution 2 (shared load stabilizersimplemented with triacs [4, patent 2], [5]) has revealed, alongtwo decades of field experience, outstanding reliability whenworking at moderate loads, as it maintains a not-too-impairedoperation even with several switches (not consecutive) out ofservice.

Solution 3 is interesting when a very fast response is notmandatory and the resulting voltage drop in the equalizing coilsfor the nonsharing mode is acceptable.

Solution 4 is the most complete if a fast response is notneeded. The only advantage over solution 3 is the ability toreduce the voltage drop in for nonsharing mode. This haspractical meaning only when the intertap voltage is relativelyhigh.

Soft switching tap changers [6], [7] are more suitable thanthose explained here for dual-tap high-frequency chopping sta-bilizers, as the reduction of switching losses is important. Theyare less applicable, in general, to multitap topologies.

Seminatural switching [8] tap changers are in a midpointposition between hard and soft switching regarding losses. Intheir actual developing status, they do not allow shared load normixed load modes (interesting for high-power applications) asthe proposed solutions does.

By means of the iterative analysis and synthesis cost compar-ative search method mentioned in [1, Sec. II], the comparisonamong the studied solutions (Table IV) has been made up fol-lowing the evaluation criteria proposed in [2]. It resumes thenew dedicated synthesis calculations not shown here developedfor solutions 1, 2, and 4 in [11], similar to that of Section IIfor solution 3. The usage coefficient for the transformers andcoils ensemble is defined [2] as the ratio between the equipmentoutput rated power and the sum of the equivalent powersof transformers and coils. The semiconductor usage coefficientis defined [2] as the ratio of the equipment rated output power

and the sum of rms current times rms voltage products re-ferred to all switches.

On deriving Table IV, these following criteria have been con-sidered.

— To help comparison, the column assigned to solution 3in Table IV repeats several formulas from Table III.

— All letters in parenthesis refer to expressions derived forsolution 3 in Table III, although they are used in othersolutions.

— The equivalent power of MT derived for solution 3 andits example figure (120.7 kVA) has been translated to theother solutions, ignoring some negligible differences.

Fig. 3. A 1-kVA single-phase prototype block diagram. The programmable acsource is used as a source and load by adjusting both module and angle of oneof the phases. This allows demanding current with any module and phase fromthe ac stabilizer and helps testing commutation in every condition.

Fig. 4. Practical implementation of the coil windings in the 1-kVAsingle-phase prototype (for 9 or less, main taps) searching a good compromisemagnetic coupling—interlayer capacitances. The auxiliary winding has beeninterlaced in the same degree with every equalizing coil.

— The expression for the equivalent power of CT inTable III is usable in any solution as it depends only ongeneral specifications.

— The equivalent power of the coils for solution 1 uses theexpression obtained for solution 3, changing the voltagesupported by the fictitious value thatrenders an appropriated small value for well below0,1 I.

— In all solutions, the voltage withstood by the auxiliaryswitches is the voltage across the MT secondary.

— In solution 4, it is considered an extra short-circuitswitch.

IV. EXPERIMENTAL RESULTS

A single-phase, 230-V 50-Hz 1-kVA 9 (or less) tap prototypewith a control circuit on the ADSP-2181 processor has beenimplemented according to Fig. 1. It has been tested for solutions1, 2, and 3 to verify the respective switching process and tosimulate basic applied control operations, as canceling a fastmains voltage step (Figs. 3–6).

Special care has been taken when implementing the mul-tiple-winding coil to ensure high magnetic coupling betweenthe auxiliary winding and every equalizing winding and alsoto prevent too high interlayer capacitances which could lead


Fig. 5. Prototype response to an input negative voltage step from 342 V (peak)to 280 V (peak) operating as in solution 1, nonoverlapping mode. Upper signali , vertical scale 500 mA/div.; Lower signal: Voltage at CT primary, verticalscale 50 V/div. Horizontal scale 0.5 ms/div. There is not a new tap change untilthe last one has finished (until i comes down to zero).

Fig. 6. Prototype response (lower trace) to an input voltage step (upper trace)from 264 V (peak) to 358 V (peak) operating as solution 1, overlapping mode.Vertical scale 100 V/div. Horizontal scale 2 ms/div.

to undesirable current peaks during the switching process. Anacceptable compromise between leakage inductances and straycapacitances has been reached by an appropriated windings dis-tribution (Fig. 4).

The main component rates and constructing parameters areshown in Table V. Special low-loss practical design has beenimplemented for transformers and coil to make the results ob-tained from this small power prototype more readily scalable tohigher power equipment.

Fig. 5 shows the internal response to an input voltage negativestep ( to ) when working asin solution 1 in nonoverlapping mode. The whole process (sixchanges) lasts 1.7 ms.

Fig. 6 shows the external response to an input voltage positivestep (from to ) working asin solution 1, overlapping mode. The lasting process has beenreduced to 0.4 ms although more changes than those shown inFig. 5 are involved.

TABLE VCONSTRUCTING PARAMETERS OF THE 220-V, 50-HZ, 1-kVA PROTOTYPE.

SEE THE CIRCUIT IN FIG. 1

A. Graphic Simulator

Based on the exhaustive knowledge obtained through the it-erative analysis and synthesis search referred to in [1, Sec. II], aspecific and very accurate simulator can be made for any solu-tion consisting of the graphic replication of the time evolutionof the input and output stabilizer electric variables. The time-de-pendent signals are constructed upon their particular mathemat-ical laws obtained from the analysis step for each operative in-terval related to a specific circuit status. These laws have been,in turn, checked for every solution and working condition of theprototype. The high accuracy achievable allows this simulatorto substitute the physical stabilizer (or other component basedsimulation, as in PSPICE) when studying the overall stabilizeroperation as part of an electric application problem by meansof a general control system simulation. Since the graphic simu-lator works in the time domain as the stabilizer itself, its primaryconnection to the simulated or real control system does not needany transfer function. Nevertheless, deriving a transfer functionfor the stabilizer can be an important aid for the conceptual un-derstanding and the achievement of the control algorithm for theoverall practical system.

Fig. 7 shows a very basic description of the graphic sim-ulator internal blocks and their functional relation with theinput–output waves and with the control circuit. As this sim-ulator operates at the real-time scale, the control circuit canbe first implemented on a simulation tool and finally on theelectronic microprocessor or similar.

Also, the internal variables, as current and voltage in theswitches, transformers, and coils, can be simulated into the


Fig. 7. Conceptual description of the graphic simulator and its relationshipwith the control circuit and the input (mains) and output (load) electrical waves.

graphic simulator. This allows the tool to also study, in detail,the operation and stress of any specific component undercritical conditions, a very convenient facility in high-powerapplications, since full-scale testing can be difficult.

The mutual dependence among the variables evolution andthe parameters defining the circuit operation (as, for instance,the lasting of the energy-return interval upon the total cur-rent change of ; [1, Sec. III-A2b] canbe easily taken into account in the mathematical network of thegraphic simulator, so that a very high fidelity can be obtainedin the time-domain replication of the variables evolution of thestabilizer in any condition, both steady and transient states. Theerror resulting for the instantaneous replication of the physicalvariables values can be fixed around 0.5% for steady states and3% for transients. Both errors depend on the detail degree of theanalysis step, the completeness of the simulator mathematicalnetwork, and the fidelity of the parameters derived for the com-ponents in the synthesis step. The high fidelity achievable in thissort of tool suggested to propose, in precedent works applied toother converters [9], the name of real structure simulation for thebasis of the method. As every kind of simulation tries to emulatea real structure, here, a more restrictive name (graphic simula-tion) describing the specific procedure has been preferred.

This graphic simulation concept is independent from the sim-ulation environment chosen. Here, the tools Simulink1 (for theblock diagrams) and Stateflow1 (for the state flux) integrated inthe Matlab1 environment have been chosen, profiting on theirmany advantages in solving control problems. This environmentallows also to easily include extra power components, not con-sidered in the nucleus of the graphic simulator, by means of theSimPowerSystem1.

Fig. 8 shows both physical and simulated curves for the pri-mary voltage of CT and the current in the auxiliary coil duringa tap change for solution 1. Snubbers have been removed for amore detailed comparison. Figs. 9 and 10 show the applicationof solution 1 to voltage harmonic and flicker reduction emulatedby the graphic simulator. To highlight its performance, in bothcases, the control algorithm has been simplified to a change oftap when the output voltage goes out of the fixed margin (basicon–off control), ignoring many possible control improvements.The results prove the graphic simulation tool as a good analyzerof plant applications.

1MATLAB, SimPowerSystems, Simulink, and Stateflow, are registered trade-marks of The Matlab Works, Inc., Natick, MA.

Fig. 8. Comparison among prototype and simulator results during a tap changeto a higher voltage tap in solution. 1(a) Upper trace: CT input voltage, 25 V/div.,50 �s=div.; lower trace: i , 0.5 A/div. (b) The same curves obtained in thegraphic simulator. The average error is less than 2%.

Fig. 9. Stabilizer operation in third harmonic voltage reduction with a simpleon–off control. Simulation based on solution 1 (17 main taps; 50 �s for overlapinterval). (a) Input voltage THD = 10% and harmonic content 10% thirdphase, 0.0% 5th and 0.0% 7th. (b) Ideal output voltage 230 V, 50 Hz. (c) Outputvoltage THD = 2:66% and harmonic content 1.92% third, 1.55% fifth, and0.742% seventh.

Fig. 10. Stabilizer operation in flicker reduction with a simple on–off control.Simulation based on solution 3, mixed load (eight main taps, 100 �s for overlapinterval). a: Input voltage: 230 V, 50 Hz, 10% amplitude modulation at 8 Hz. b:Output voltage showing a residual modulation of 0.6% approximately.


V. COMPARISON WITH OTHER LINE CONDITIONERS

Traditional onload multitap changers are devoted to compen-sate slow voltage variations. When no other mains disturbancesmatter, they are by no means the cheapest and sturdiest solution.The fast-acting versions here proposed (seven to 47 changesper half cycle) can compete with series-connected voltage lineconditioners based upon high-frequency active filters [12]–[15]only when the voltage disturbance to be mitigated includes sub-harmonics (Fig. 10) and harmonics up to the fifth order (Fig. 9for mitigation of a third-order harmonic). Nevertheless, multitapchangers are basically different from high-frequency voltage ac-tive filters both in the power circuit and the control strategy. Thesimplicity and advantage of the multitap-changer power circuitcannot be impaired when a power transformer is already in-volved in the line to be stabilized, as the improvement action isthen as simple as to add controlled taps to the existing machine.

From the operation and control point of view, the multitapchanger switches only when needed and remains in the sametap until important voltage disturbances occur. This fact adds ro-bustness to the power circuit, reduces the switching losses, andforces to operate the last control blocks in an on–off strategy,which, in turn, can be commanded by operative controllers ofmore or less complexity (long-term voltage or current com-pensation, compound compensation, fast voltage compensation,etc.).

Voltage line conditioners based on an active filter include apower inverter that constantly switches under some pwm pat-tern, thus increasing switching losses with respect to multitapchangers. The inverter needs, in turn, a power feeder (an extrarectifier or in universal power conditioning [16]–[19], the par-allel current filter). This means that, in general, and unlike tap-changers, active filters perform twice the mitigating power in-volved. Nevertheless, they can attenuate higher order harmonics(typically up to the 19th order if switching at 10 kHz) than thefast-acting changers here proposed.

VI. CONCLUSION

The feasibility of subcyclic multitap-changing ac voltage sta-bilizers with clamped voltage hard switching has been demon-strated for four solutions of the switching circuit. They includethe direct and compensating structure for the main circuit, thesharing and nonsharing load modes for the steady-state oper-ation, and the overlapping and nonoverlapping modes for theswitching operation, covered by patents [20].

The results of this new class of stabilizers confirms the pos-sibility of reaching from 7 to 45 (depending on the solutionand modes selected) tap changes every half cycle in a 50-Hzmains wave, with moderate switching losses, using a standardpower IGBT, fast diodes, and well-coupled inductors. This al-lows the tap changers to extend their operation from the classiccorrection of long-term voltage variations to the attenuation offlicker, voltage harmonics, and instability problems, as envis-aged by the electrical net experts of the 1970s [10] for medium-and low-voltage ac lines.

The iterative study here applied consists of the analysis ofnine solutions, the synthesis of their calculus sheets, and the

comparison among the resulting effective power and other com-ponent working rates for the best four solutions. It demonstratesto be a very effective search method.

Predicted waves by analysis show good agreement with theexperimental results in a low-power prototype.

An accurate specific graphic simulator (based on the math-ematical replication of waves after an exhaustive analysis andphysical test for every working condition) has been developedfor each solution to help on power and control design and test.

First simulated applications to the reduction of voltage har-monics and flicker show quite promising results for both the cir-cuits and their graphic simulator.

REFERENCES

[1] J. Vaquero, J. C. Campo, S. Monteso, S. Martínez, and M. A. Pérez,“Analysis of fast on-load multi-tap-changing clamped-hard-switchingAC stabilizers,” IEEE Trans. Power Del., vol. 21, no. 2, pp. 852–861,Apr. 2006.

[2] F. Barrero, S. Martínez, F. Yeves, and P. M. Martínez, “Active powerfilters for line conditioning: A critical evaluation,” IEEE Trans. PowerDel., vol. 15, no. 1, pp. 319–325, Jan. 2000.

[3] J. C. Campo, J. Vaquero, M. A. Pérez, and S. Martínez, “Dual-tap stabi-lizer with mixed seminatural switching. Analysis and synthesis,” IEEETrans. Power Del., vol. 20, no. 3, pp. 2315–2326, Jul. 2005.

[4] S. Martínez, “Perfeccionamientos en Equipos Electrónicos Para Reg-ulación de Tensión Alterna…,” Spanish Patents 500.523 (Oct. 1981),500.524 (Oct. 1981), 522.497 (Oct. 1984).

[5] , “Sharing load AC tap-changing stabilizers” (in Spanish), MundoElectrón., no. 166, pp. 127–135, Oct. 1986.

[6] G. Villegas, J. Vaquero, R. Echevarría, S. Horta, M. A. Pérez, and S.Martínez, “Quasiresonant fast on-load tap changer stabilizer,” in Proc.IEEE-Cenidet CIEP, Morelia, Mexico, Oct. 1998.

[7] J. C. Campo, “Functional and topological analysis of ultra-fast differen-tial ac stabilizers,” Ph.D. thesis (in Spanish), Dpt. Ingen. Eléct., Electrón.Comput. Sistem., Univ. Oviedo, Gijon, Spain, 2000.

[8] N. Burany, “Safe control of four-quadrant switches,” in Proc. IEEE IASAnnu. Meeting, Oct. 1989, pp. 1190–1194.

[9] S. Lorenzo, J. M. Ruiz, F. Aldana, and M. Shaker, “A new modeling andsimulation CAD package for power converter design,” IEEE Trans. Ind.Electron., vol. 37, no. 5, pp. 387–397, Oct. 1990.

[10] G. Musgrave and D. O’Kelly, “Improvement for power system transmis-sion by solid-state techniques,” in Proc. Inst. Elect. Eng. Conf. Publica-tions, Dec. 1974, pp. 228–233.

[11] J. Vaquero, “AC on-load tap-changing fast voltage stabilizers for powerquality improvement,” Ph.D. dissertation (in Spanish), National Dis-tance Univ., Madrid, Spain, 2000.

[12] W. M. Grady, M. J. Samotyj, and A. H. Noyola, “Survey of active powerline conditioning methodologies,” IEEE Trans. Power Del., vol. 5, no.3, pp. 1536–1542, Jul. 1990.

[13] H. Akagi, “Trends in active power line conditioners,” IEEE Trans. PowerElectron., vol. 9, no. 3, pp. 263–268, May 1994.

[14] F. Yeves, J. Carpio, S. Martínez, V. Feliu, and C. Pumar, “Iterativevoltage and current high frequency line conditioner,” Eur. PowerElectron. J., vol. 5, no. 3/4, pp. 38–43, Jan. 1999.

[15] X. Lei, D. Retzmann, and M. Weinhold, “Improvements of powerquality with advanced power electronic equipment,” in Proc. IEEE Int.Conf. Electric Utility Deregulation Restructuring and Power Technolo-gies, London, U.K., 2000.

[16] T. Tanaka and H. Akagi, “A new combined system of series active andshunt passive filter aiming at harmonic compensation for large capacitythyristor converters,” IEEE IECON, 1991.

[17] L. Gyugyi, “A unified power flow control concept for flexible AC trans-mission systems,” Proc. Inst. Elect. Eng. C, vol. 139, no. 4, pp. 323–331,Jul. 1992.

[18] S. Martínez, V. Feliu, F. Yeves, J. L. Iribarren, and P. M. Martínez, “Dis-positif de Condicionnement de Ligne Pour Reduire ou Eliminer Les Per-turbations,” French Patent no. d’enr. 94 069631, 05, 1994.

[19] F. Barrero, S. Martínez, F. Yeves, F. Mur, and P. M. Martínez, “Universaland reconfigurable to UPS active power filter for line conditioning,”IEEE Trans. Power Del., vol. 18, no. 1, pp. 283–290, Jan. 2003.


[20] J. Vaquero, S. Martínez, M. A. Pérez, and J. C. Campo, “Perfec-cionamientos introducidos en equipos electrónicos de Tomas Múlti-ples…,” Spanish Patents 2 198 208 (16/1/2004), 2 198 209 (16/1/2004),2 198 210 (16/1/2004), 1994.

Joaquín Vaquero López was born in Madrid,Spain. He received the M.Sc. degree in electricalengineering from the Polytechnic University ofMadrid, Madrid, in 1994, and the Ph.D. degree fromthe National Distance University of Spain, Madrid,in 2000.

In 1995, he joined the Department of Electrical,Electronic and Control Engineering, Universidad Na-cional de Educación a Distancia, Madrid, as AssistantProfessor. His research interests are fast tap changersand equipment to improve power quality.

Juan Carlos Campo Rodríguez (M’97) was bornin Gijon, Spain, in 1970. He received the M.Sc. andPh.D. degrees in electrical engineering from the Uni-versity of Oviedo, Gijon, Spain, in 1995 and 2000,respectively.

Currently, he is an Associate Professor withthe Department of Electrical and Electronic Engi-neering, University of Oviedo. His current researchinterests include power quality, line conditioners,and tap changers.

Santiago Monteso Fernández was born in Madrid,Spain. He received the M.Sc. degree in electrical en-gineering from the National Distance University ofSpain, Madrid, in 2002.

Since 2003, he has collaborated with the De-partment of Electrical, Electronic and ControlEngineering with the Department of Electrical,Electronic and Control Engineering, UniversidadNacional de Educación a Distancia, Madrid, as As-sistant Professor. His research interests are fast-tapchangers, equipment to improve power quality, and

graphic simulator for power-electronic circuits.

Salvador Martínez García (SM’90) received theM.Sc. and Ph.D. degrees in electrical engineeringfrom the Polytechnic University of Madrid, Madrid,Spain, in 1966 and 1969, respectively.

Currently, he is a Full Professor with the Univer-sidad Nacional de Educación a Distancia, Madrid,where he was Associate Professor from 1979 to 1982.From 1975 to 1979, he was Associate Professorwith the Polytechnic University of Madrid. He was aDesign Engineer concerned with power-electronicsequipment with several companies. His research

interests are integrated magnetics and power-line conditioners.Dr. Martinez was a member of the IEEE IAS Industrial Static Converter Com-

mittee EWG from 1990 to 1997.

Miguel Angel Pérez García was born in Turon, As-turias, Spain, in 1962. He received the M.Sc. degreein industrial electronics engineering and the Ph.D. de-gree in electrical engineering from Oviedo Univer-sity, Gijon, Spain.

Currently, he is an Associate Professor in the De-partment of Electrical and Electronics Engineering,Oviedo University, where he manages several R&Dprojects in the areas of electronic instrumentation andpower electronics.

Synthesis of Fast Onload Multitap-Changing Clamped-Hard-Switching AC Stabilizers

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