-:~~

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Engineering studies for Itaipu convertorstation design

Abs~ract: The paper presents a discussion on some of the engineering studies performed to determine thed~s.lgn param.eters of the Itaipu ~onvertor stations. The following studies are discussed: steady-state con-ditIOns, reactive-power compensatIOn, Insulation co-ordination and arrester protective scheme current stressessystem stability, main characteristics of the master control, AC and DC filter and DC line res;nance. For each'of these st~dy areas, the paper gives a summary of the study methodology used, indicates the main studyresults, and Includes some of the system problems encountered and the solution adopted.

Introduction

Itaipu power plant is a hydroelectric project being constructedby Brazil and Paraguay on the Parana River, with 12600 MWinstalled capacity to be generated by 9 x 823.5 MVA, 50 Hzunits and 9 x 737 MVA, 60 Hz units. In accordance with theagreement signed by the two countries, each one will havethe right to buy one half of the power to be generated byItaipu and, in addition, one country will have the first priorityto buy the excess power that could not be used by the other.Considering the terms of the agreement, that Paraguay wouldonly need the total of its share of the power in the far futureand that Brazilian frequency is 60 Hz, it was decided to con-struct a hybrid AC/DC transmission system with two ± 600 kVbipolar lines [1] of 795 and 8l5km, to transmit the 50Hzgeneration and three 750kV AC transmission lines to transmitthe 60Hz power to Brazil [2,3].

The convertor station specifications were issued inNovember 1978 and the contract was awarded to the ASEA/PROMON Consortium in June 1979. It is the scope of thispaper to report the results of several studies performed, up tonow, comparing them with the requirements of FURNASspecification. These studies were done in Brazil and Sweden bythe Consortium in close co-operation with FURNAS, to finalisethe equipment and control specifications for the Itaipu HVDCconvertor stations.

2 Steady-state conditions: main circuit parameters

The purpose of this study, carried out using the well knownDC formulas, was to establish the main circuit parameters ofthe HYDC transmission as well as its steady-state characteristics,related to the following operating modes:

(a) balanced bipolar(b) monopolar with ground return(c) reduced DC voltage(d) bipole paralleling(e) high reactive absorption(I) reverse power.

The FURNAS specification required transmission of ratedpower of 6300 MW from the rectifier AC 500 kV busbar. Therated power at the inverter substation 345 kV AC busbar wascalculated, based on the minimum line resistance with maxi-mum ambient temperature of 40°C dry and 29°C wet bulb.

Paper 2320 C (P9, Pl0), received 23rd June 1982Mr. Peixoto is with FURNAS Centrais Eletricas SA Rua Real Grandeza219, 13~ andar, Bl.A 22281, Rio de Janeiro, Brazil. Mr. Frontin is alsowith FURNAS Centrais El<'tricas SA, Rua Real Grandeza 219, 16~andar, Bl.C 22281, Rio de Janeiro, Brazil. Mr. Jardini is with THEMAGEngenharia, Praia de Botafogo 5 18, 11~ andar, 22250, Rio de Janeiro,Brazil

The power should be transmitted by two bipoles withtwelve pulse groups per pole. The thyristor valves are air in-sulated, water cooled and designed for indoor use. The valvesare built up as a quadruple valve unit, the single valve beingmade up of eight thyristor modules connected in series, eachcontaining 12 thyristors, type YST-45 (85 mm diameter).Redundancy is 3%.

Rated voltage was chosen as 600kV (at the rectifier) basedon the lowest present worth cost of convertor stations, DC lineand losses when compared with 500 and 550 kV [2]. Theminimum losses were estimated by the Consortium as 18 MWper bipole resulting in a rated current of 2.61 kA. At thespecified low ambient temperature of 30°C dry and 24°C wetbulb, the Consortium guaranteed 3392 MW per bipole withoutadditional cost. An overload capability of 25% for 5 sand 15%for 20 s can be applied to both normal and low ambient con-ditions. The minimum continuous current is 10% of rated. Inthe reverse power mode the Consortium guaranteed, withoutadditional cost, 3007 MW at the 345 kV AC busbar, limited bytransformer tap range.

After a DC line fault, three sequential restarts are permitted,each one with individually adjustable deionisation time from0.05 to 1 s. If they are unsuccessful, one restart is tried with75% of the nominal voltage under the normal operating con-dition. The equipment will be capable of operation in thisreduced voltage condition continuously, that means maximumvalues of 'Y = 36.1°, Q = 38.5° compatible with the valve cir-cuit design, The range of 'Y can also be used for high reactiveabsorption during receiving-system light load.

Taking into acount rated power, rated voltage of ± 600 kV,maximum line resistance (maximum line length and maximumtemperature), chosen nominal values of a = 15° and 'Y = 17°,and contraints related to the maximum short-circuit capabilityof the valves (about 30 kA), the values in Table 1 were deter-mined for each six-pulse bridge.

Load tape changers (LTC) are provided on the AC side ofthe transformers. The LTC on the inverter station transformerscontrols, during steady state, the rectifier DC busbar voltage inthe range 592.5 to 607.5 kV, corresponding to two tap steps toavoid shunting. If 1% measuring tolerance is considered, therange could actually be 586.5 to 613.5 kV. The LTC inthe rectifier controls the firing angle a of the rectifier, a is inthe range 12.5 -17 .50. The LTC range was calculated using the

Trans-former dx, %y Y Li. Ratio (kV)

MVA

8.9 ± 7%500/127.4-127.4940.4/470.2-470.2

8.6 + 8%,-6%345/122-122900.6/450.3-450.3

above data and includes ± 0.75% measuring tolerance in the DCcurrent, ground return circuit of total resitance of 0.7.Q,maximum and minimum DC current, AC voltage variation550-475 kV at Foz do Igua,!u and 362-321 kV at Sao Roque,the four modes of operation (a) to (d) described above,transformer reactance tolerances, resistive voltage drop 0.4%and constant voltage drop of the thyristor 0.21 kV.

A range of - 6 to + 20 steps of 1.25% was determined forthe rectifier and inverter station transformers. This range willcover all conditions except operation at reduced DC voltagewith extreme parameter tolerances and DC voltage higher than525 kV at Foz. A reduced DC voltage of 80% can be obtainedunder such conditions and this deviation from the specificationwas accepted by FURNAS. It should be noted that to preventequipment operation under excessive steady-state voltage twoprotections are provided: one at a level UdioG = (l - k 6.5)UdioL, where k> 1.2 and 6.5 = 1.25% is the tap step, whichblocks the order to change the tap to increase the commutatingvoltage (UdioG equals 1.019 and 1.003 Udion for rectifier andinverter, respectively). The second level is UdioL = 0.99Udiomax (1% is the measuring tolerance) that orders a changeof tap to decrease the commutating voltage (UdioL equals1.034 and 1.022 times Udion for rectifier and inverter,respectively).

3 Reactive-power compensation

The study of the reactive compensation to supply the demandsof the convertors and system comprised the verification of:

(a) reactive balance in steady-state operation(b) voltage variation when switching reactive devices(c) fundamental frequency overvoltages(d) self excitation of synchronous machines(e) low-order harmonic resonance(f) power recovery after system faults.

The reactive compensation in both stations is designed, accord-ing to the specification, to allow transmission of the nominalrating of 700MW per generator with anyone switchableelement out of service, in any stage of development shown inTable 2 and bipolar operation. The reactive compensationshould also permit the transmission of the maximum outputof 727 MW per generator with all elements in service up to amaximum of eight generators.

For Foz do Igua,!u these requirements will be met over thevoltage range 0.95 to 1.05 p.u. at the rectifier 500 kV AC bus-bar, although the equipments are designed to operate up to1.10 p.u. The generators' MVAR capability at the rectifierbusbar was considered part of the reactive compensation.

For the Sao Roque station, these requirements will be metover the voltage range 0.93-1.05 p.U. at the inverter 345 kVAC busbar. With nominal DC infeed and at 0.95 p.u. ACbusbar voltage, there will be zero exchange of reactive powerto the receiving system and at 1.05 p.u. 300 MVAR can besupplied to the receiving AC system. In addition, there will bea further controlled 300 MVAR supply capacity to coveremergency conditions.

Table 2: Stages of development

Bipole voltage

Number ofmachines

kV± 300+ 600, - 300± 600± 600-I- 600= 600

± 300+ 600, - 300± 600

The reactive-power consumption of the convertors wascalculated by the Consortium considering the normal range ofDC voltage and the angles a; (12.5 -17 .5°) and 'Y (17°). Forthis calculation the nominal value of convertor transformerreactance was used based on the FURNAS specification. Thevalues of the compensation are shown in Table 3 and fulfilthe specified requirements unless the average of all transformerreactances is above the nominal value. If the average is 5%above nominal there will be a lack of 75 and 55 MVAR at Fozand S. Roque, respectively. As it is considered unlikely thatsuch deviation will occur, the amount of compensation in-dicated in Table 3 will be retained until a check can be made onthe reactance of the convertor transformers as manufactured.This would permit the determination of the possible need ofextra compensation, allowing this to be installed in due time.

Besides the case of bipolar operation used for dimensioningas per the criteria specified, the other operating modes wereanalysed but with all reactive devices in. The given reactive-power compensation is adequate for the majority of cases.However, in a few situations it will not be possible to operatein the full range of AC voltage in the sending end.

It was required in the specification that the voltage variationwhen switching the largest compensation device should belower than 5% of the preswitching value at Sao Roque. In theproposal, the largest element size was 444 MVAR. As thiswould result in voltage variation higher than 5%, a specialaction in the control of 'Y was proposed to solve the problemduring the scheduled operation. When switching in a filter orcapacitor, 'Y would be rapidly increased and then returned toits initial value slowly. When switching out, 'Y would be in-creased slowly to allow the synchronous condenser to pick theextra reactive power, and then 'Y would be rapidly decreasedafter switching out the device. This control would not coverunscheduled tripping. FURNAS considered this a deviationfrom the specification and the size of the banks was changedto the values of Table 3.

For Foz do Igua,!u, with the filter sizes shown in Table 3,the voltage variation when switching one filter bank will behigher than 5% in the initial stage. To solve this problem,FURNAS is considering the installation of a branch of eachfilter bank on a separate switch bay by advancing the instal-lation of bays that will be later utilised for complete fIlterbanks. This will permit the installation of the required filters,while meeting the voltage-variation criterion.

Another requirement related to reactive compensation wasthat, during complete DC load rejection (blocking of all con·vertor groups), the fundamental frequency overvoltage at theAC convertor busbar should not exceed 1.4 p.u. and should bereduced to 1.05 p.u. of the initial voltage within 5 s, withoutany switching operations. For the rectifier system the voltageat the generator terminals should not exceed 1.3 p.u. Also inthe situation of partial blocking of the convertor groups, atany stage, the fundamental frequency temporary overvoltageswith some convertors remaining in service should not exceedthe design limits of the thyristor valves. In the final instal-lation, when blocking one bipole from an initial operation

Equipment Foz do Igua,>u

Filters 2 X 350 + 3 X 280.3(number X MV AR = 1540.9at 500 and 345 kV)Shunt capacitor(number X MVARat 345 kV)Svnchronous com-pensation (number XMV ARc/MVARi)

3 X 220.8 + 279.8 -296 + 4 X 237 +296.3 = 2482.5

~Gndition of three poles at rated load, the convertor busbar:'undamental frequency voltage should not exceed 1.3 p.u.

For complete blocking it was not possible to keep the ACovervoltage level lower than 1.5 p.u. at the rectifier stationunless contraints of initial voltage and tap position of the step-up transformer were imposed. To solve this problem, the

five equivalent networks were represented on the simulator.The equivalents 1 and 2 were represented by LR circuits wnedto 2.15 or 2 p.u. frequency, respectively. Impedance angle was85° at 60 Hz. Equivalents 3 and 4 matched the same Z (0.:)plot of the actual network, but tuned to 120 and 110Hz. Atthe second harmonic the impedance angle is 75°. Equivalent 5

------'Q I:~",PLC Ii Iters

f2 •...I!J~.•..u 50

P

compensation could be decreased but this would result in lackof reactive power and, consequently, operating restrictions.However, all equipment is able to withstand this overvoltage,and this exception is under consideration. It is important topoint out that, owing to the characteristics of the controlequipment, simultaneous blocking of two bipoles is unlikely tooccur. For internal faults in one convertor only that convertorwill be blocked. Persistent commutation failure during systemAC faults, if initiating blocking, would result in complete loadrejection at the sending end. However, in the case of Itaipucontrol equipment design, for commutation failure due tosystem faults, blocking is not carried out. Blocking wouldoccur only if the commutation failure lasts for longer than onesecond. With reference to self excitation of the Itaipu gener-ators, preliminary investigation by dynamic simulation showedthat, if the number of generators is greater than the number offilter banks at any configuration, no risk of self excitationexists. This will be included in the supervision system. For theinverter station, a synchronous condenser connected with onefilter or capacitor bank must of be permitted and this is~overed by station layout and protection.

With respect to low-order harmonic resonance in the re-:ei\'ing system, in the Consortium proposal, a preliminarylli\'estigation was made regarding the resonant frequency ofthe network in the various stages of the development of the'X system. This was based on an LR equivalent network withi" l:::;Jedance angle of 85° from the fundamental frequency to:",e :::'::d harmonic. The number of synchronous condensersse:e::e~ ,esulted in a resonant frequency of the system plus:11:e, .i::~ synchronous condensers above 2.15 times funda-:-::e:: ::c:. O.·.1ng to changes in system configuration later on, it'.'-.is :":s:.y:e,ed that some combination of system plus compen-satlo" :'J'.::d be in resonance at the second harmonic. Then asim:.:::.::,:, study was conducted to evaluate this condition andto ~J:-'.;::.:,e \\ith the results obtained during the proposalperiod. To compare and analyse the effect of these conditions,

--." "'-J- electrodeto the Iineother pole

represented the actual system with no adjustment. Faultclearing, transformer DC saturation and valve faults (misfireand firing through) were studied and it was concluded thatalthough a second harmonic resonance could be found in theactual system, the angle of the impedance was very low,offering a high damping of the possible oscillation. Conse-quently, the transient performance of the actual system wasbetter than the simple equivalent considered in the proposalwith the same number of synchronous condensers.

It should also be noted that the number of synchronouscondensers chosen was essentially determined by the require-ments of system recovery after faults.

4 Insulation co-ordination and arrester protective scheme

Only zinc-oxide gapless arresters are provided for the ItaipuHYDC system (Fig. 1).

Extensive overvoltages studies were performed to evaluatethe energies and currents for which the arresters would have tobe tested, and to find the maximum current and respectivewaveshape that define the protective level and, consequently,the insulation. The studies covered fundamental frequencyovervoltages, switching-surge overvoltages and lightning surges.In the first group, complete and partial blocking of the con-vertors was examined. In the second group, investigationsof transients were related to load rejection, fault clearing,switching of transformer, filters and lines, valve faults and theconsequences on the AC side, monopolar faults on the DClines, ground faults on the AC phase of the valve side of theconvertor transformers, faulty closing of bypass switches,ground faults on the DC busbar including fast surges, bipoleparalleling, DC mter switching and current extinction in onethree-pulse commutation group. In the third group, the effectof lightning surges injected from AC and DC lines and due toshielding failure were evaluated.

Digital computer programs were used to study fundamentalfrequency and lightning overvoltages; whereas TNA, DCsimulator and computer programs were used to study switching-surge overvoltages. In the latter, as a general rule the TNA and

Table 4: Arrester characteristics and stresses

6Stress Location (F ig. 1) 1A 18 2 3 4A

arrester parameters 48 Foz S. Roque

Number of colums 8 4 6 (3) 2 2 2Energy capability, MWs 8 4.2 6.5 2 3.6 5.3 3.6

Switching Maximum continuoussurge voltage, kVp 206.7 206.7 347.5 10 614 450 296

Maximum energy formstudies, MWs 14.4 4.2 6.5 2 3.6 5.3 3.6Maximum current, kA 6.3 2.1 1.8 60 O.b- 1 1wave shape, J.1.S 1000 1000 1000 (2) 1000 45 X 90 45 X 90Protective level, (1) (1)kV 325/312 316/302 579 110 1057 862 55220% margin(reference) 390/374 379/362 695 132 1269 1035 663

Impulse Maximum current, kA 3 3 3 11 17.5 12.8 10surge Protective level, kV (1) (1)

332/319 332/319 626 101 1380 1050 67725% margin(reference) 415/399 415/399 783 127 1725 1312 847

(1) Lower values refer to Sao Roque station and higher ones to Foz(2) High frequency due to filter discharge wave shape (close to lightning-surge type)(3) Preliminary value(4) AC/DC filter arrester specification not yet finalised (location 5 and 7)

the DC simulator were used for general investigations of worstcases and digital programs used to make the final calculationsof energy and currents in the arrester. This procedure wasadopted because the energy given in TNA and simulatorstudies is considerably lower than that obtained with digitalcalculation owing to the inherent damping of the former tools.In the digital simulation, the arresters were represented by theiraverage characteristic less 2.5% (manufacturing tolerance).The protective level is taken for this same reason as plus 2.5%above the average characteristic. The results presented in termsof arrester stress are shown in Table 4, together with othercharacteristics.

The maximum stress obtained in the studies for the 1Aarrester was 14.4 MWs during an AC phase-to-ground fault onthe valve side of the transfonner in parallel-operation mode ofconvertors and a certain range of fault application instants. Asthis condition has a low probability of occurrence and theenergy specified of 8 MWs covers most of the cases in paralleloperation and all other modes of operation, this risk wasaccepted by FURNAS. In any case, should this high-energy(14.4MWs) condition occur, the protective level will not beexceeded even if the arrester fails. For the other arrester theconditions that gave highest energy were:

arrester lB- current excitation in only one three pulsegroup

arrester 2 - faulty closing of the bypass switcharrester 3 - DC busbar faultarrester 4A/4B - discharge DC line at 1.7 p.u.arrester 6 - fault clearing [4] .

As required by the FURNAS specification, the insulation levelswere selected in the following way:

Non-self-restoring insulation: For the oil insulation the BILis equal to standard IEC value greater than 1.20/0.83 times theswitching surge protective level of the arrester. The BSL isequal to 0.83 BIL. For the values, 15 and 20% margin are usedfor switching surge and lightning/fast surges, respectively. Itshould be noted that minimum values of BIL = 1425 kV andBSL = 1175 kV for the Foz transformer, and BIL = 1050 kVand BSL = 870 kV for the Sao Roque transformer werespecified for AC side.

Self-restoring insulation: For all equipment, except the onesmentioned above, the BSL is equal to standard lEC valuegreater than 1.25 times the switching-surge protective level ofthe arrester, to give an overall risk of failure not greater thanone flashover in 100 years within the substation. The BIL is

Table 5: Self-restoring insulation levels for Foz (1)

Sector Protected by BIL Margin BSL Marginprotected arrester

kV % kV %DC line end 4A 1800 30 1321 25DC switch yard 4B 1675 31 1321 25Valve side ofDC reactor 2+2 1675 34 1448 25Neutral busbar 3 325 221 150 36Transformervalve sidelower D wind 3 + 1B 750 73 533 25lower Y wind 3+1B+1B 1050 37 928 25upper D wind 2+1B 1300 36 1119 25upper Y wind 2+1B+1B 1800 39 1514 25Busbar between 12pulse bridges 2 850 35 724 25Across singlevalve 1A/1 B 399 20 374/363 15-'Ie busbar~:z 8: l;;uEzu 6 1550 48, 1 ~ "75 .-::::=: - ::_:=- 0 ";-:72: 73 ~== -

--~~'"l.~"

Table 6: Nonselfrestoring insulation levels for Foz (1)

Sector Protected by BIL Margin BSL Marginprotected arrester

kV % kV %DC reactorline side 4B 1550 21 1287 21DC reactorvalve side 2+2 1675 33 1390 20Across DCreactor (note 2) 1800 20 not governedTransformervalve sidelower D 3 + 1 B 650 50 540 27lower Y 3+1B+1B 1175 54 975 31upper D 2 + 1 B 1300 36 1079 20upper Y 2+1B+1B 1800 40 1494 23Line sideF5?z Igua<;u 6 1425 36 1175 36Sao Roque 6 1050 55 870 58

(1) For Sao Roque all insulation that depends on arresters 1 A/1 B is slightly lower.(2) There is no arrester across the DC reactor. The margin is applied to maximum overvoltage from the studies.

selected as the best lEC combination from BSL. It should benoted that, for AC equipment, minimum values of ElL =1550kV and BSL = 1175kV, for Foz, and BIL = l175kVand BSL = 900 kV, for Sao Roque were specified. The in-sulation levels for the station were determined by these criteriaand the knowledge of the arrester location involved in theprotection. Tables 5 and 6 show these values.

The purpose of this study was to evaluate the overcurrents dueto short circuit across one valve, phase-to-ground short circuiton the valve side of the convertor transformers, DC line shortcircuits, other DC current transients due to control action orAC busbar short circuits.

Fig. 2 show the condition of short circuit across valve.When the valve 3 starts conduction and there is a fault acrossvalve 1, a phase-to-phase short-circuit occurs. The maximumcurrent value is calculated by

. Id V11- - = - [cos(wt+a+/3)-e-Rt/Lcos(a+mJ

2 2Z

where V is the crest phase-to-phase Thevenin voltage, Z =yR2 + (WL)2 is the Thevenin impedance that includes systemimpedance and convertor transformer impedance, /3 = tan-1

(R/ wL), a is the firing angle and I d is the DC current.The maximum occurs when wt + a + /3 = 1800

, and beforethis maximum occurs a blocking order is given to all convertorsof the pole with the current Id interrupted by the valves in the~1ther3, 6 or l2-pulse groups. This is the reason for the term(- Id /2) in the formula and can be understood if superpositioni, considered (Id injected from the DC line divides into Id/2 in"::C:',e 1 and 3). To find the maximum value based on the;-~:-:~~ulain eqn. 1, many cases need to be investigated including'.:::-:J';S operating conditions in terms of active and the reactive::~.'.~~ and DC/AC voltages that relate the variables V,L' 0: and transformer reactance for each tap position.X,~ 2., Z includes system impedance this should be the mini-:-~.;:~~·.3..l~leforeseen. In the Itaipu design the values obtainedv.e~e ::-.JkA, for the FozdoIgua~ustation,and27.6,fortheSac> RJ::c-...:estation working as rectifier.

All tillS extensive calculation is avoided if the current isestimateci by the phase-to-phase short-circuit calculation usingthe value of V that corresponds to Udiomax, a = amin with Zequal to the minimum transformer reactance Xtmin, then

taking this condition as the study margin. This calculationresults in 33kA for Foz, reducing to 31 kA if the system im-pedance is included, then to 28.3 kA by introducing a and L/R,and finally to 27 kA if Id/2 is subtracted.

A requirement in the specification was that the valvesshould block immediately after this current loop (27 kA), evenwith maximum overvoltage (1.3 p.u.). If this blocking does notoccur, subsequent current loops stress the valve 3 and ceaseonly when the AC breakers open. For this case, the specificationrequired the value capability to withstand two additionalcurrent loops without subsequent blocking. The second andthird current loops are evaluated by the formula in eqn. 1,without the term Id/2 because the DC current was previouslyblocked by the other valves. This results in a maximum currentequal to 28.3 kA for the Foz valves, requiring testing at maxi-mum temperature with the equivalent current of one loop of27 kA and two loops of 28.3 kA.

Related to phase-to-ground faults on the valve side of theconvertor transformer, the calculations were performed con-sidering the voltage and impedance in the circuit path of thefault current. For instance, for faults in phase a of the lower .6..transformer, the voltage is (Va - Vc) = VL and the impedanceis 2Xtmin + Xelect and the circuit path is from the fault pointto A to N to C to E (Fig. 2).

For faults in the phase a of the lower Y, the voltage is thevectorial sum of the phase-to-phase values in the Y and the 1\using the impedance 4Xtmin + Xelect. In all cases, the short-circuit current is below 17 kA.

'EDC reactor

.8Ul

.SOJ

3 5 c.~u

i1 t '30

V/.f3 / :0It'l '0A Ul

NC1JUl0

C u

commutating .t::.u

reactance ~.~

<II

6 2 4

To evaluate the overcurrents on the DC side of thesmoothing reactor and line, the HVDC simulator was used toinvestigate the following conditions:

DC line to ground faultblocking of the inverteruncontrolled rectifier and subsequent blocking of the

invertertransition of the inverter to rectifier mode of operationparallel operation.

The maximum current peak value, 13 kA, was obtained. Forequ~~ment on the valve side of the smoothing reactor thedeclslve current 25 kA was measured at a short-circuit acrossconvertors.

AC busbar short circuits were evaluated with conventionalprocedure and the results are shown in Table 7.

6 Stability studies

During the preparation of the specification FURNAS conductedextensive stability studies in all development stages and with acomplete representation of the system [2,3] .The recovery timeof the DC system after faults was specified as 160 ms, based onthis digital study and a preliminary simulator study. 1985 wasfound to be the most critical year, because it represents thelast stage of the development of the DC system with theweakest receiving system. In 1988 the AC part of the Itaipupower plant is completed and the receiving system is muchstronger.

Owing to the possible difficulties of reproducing thesestudies during the bid preparation, the specification asked fora demonstration of the control behaviour in a reduced ACsystem (total of 14 busbars) that has the same response of theactual network (year 1988) in the first swing, even recognisingt~at the behaviour in the dynamic period could be completelydlfferent. The Consortium did not present the study in theproposal, but guaranteed a performance equivalent of the ref-erence system to be demonstrated afterwards.

After the award of the contract, it was agreed to study theyears 1983, 1985 and 1988 (heavy and light load) with thecomplete system representation. Important conclusions relatedto time to recovery and modulation of the DC system weredetermined. The study concentrated on the 1985 heavy loadcondition [5]; this year, with constant power mode ofcontrol, a reduction in AC voltage will be accompanied by areduction in DC voltage and, consequently, a demand forhigher current. This increase in current aggravates the initialvoltage reduction owing to the increase in reactive-power requi-rement and communication drop. The use of constant currentmode of control would be, in this situation, better for thesystem, or any other measure that avoids high increases incurrent or reductions in voltage (i.e. lower setting of VDCL orincreased synchronous compensation). The studies performedby the Consortium confirmed this principle and showed thatrecovery in 400 ms is better than 160 ms. This will be achievedby introducing a time constant of this order in the constant-power control loop between the DC voltage signal and thecalculation of Iorder = P order/ Ud' During the preparation ofthe specification, power modulation was found to be usefulbecause its action was to avoid increase in current. Studies onmodulation signals made by the Consortium found that amodulation of 'Y derived from AC voltage at Sao Roque (see

Table 7: AC busbar short-circuit current, kArms

Station 30 10 Breaker rating

"4~ 50

= .,' ~ . - - - -

Fig. 3) w?uld give sat.isfactory performance and was preferablefor. cer.tam :easons; mc1uding the fact it does not introduceoscillatlOns m the Paraguayan system. This modulation shouldhave a lower limit so that 'Ydoes not corne below 'Y . = 17°to ensure a sufficient margin for commutation (see Fig,.n3).

VS.Roque

sT,('.sT, )(l.sTZ)

K=Z36, 4° I P u.T) = 2.0

TZ=0.0127l!.y = 13°E

max1m;n ref=17°E

This modulation solved all stability problems for all yearsof the system studies and will be incorporated into the controlsystem at Sao Roque [5] .

7 Main characteristics of the master control

The master control functions are mainly related to the inter-action between the AC systems and the DC link, which areperformed automatically or through action of the systemoperators.

The master control is organised in an hierarchical manner:station control, bipole control and pole control such that thelowest level possible controls the minimum unit or block ofpower with the maximum independence.

The pole control receives the power order from the bipolecontrol and is responsible for the various additional powerorders, such as stabilising of the 50 Hz network and 60 Hzfrequency regulation. The calculation to allow short and longtime overloads, as well as start, stop, paralleling and deparal-leling sequences, are executed at the pole level.

The bipole control is responsible for the transmission of thepower order received from the operator to the pole control. Ir-a loss of transmission capability occurs within one bipole, anadditional power order is transmitted to the other bipole tocompensate for this loss. The selection of the operator controllocation for power control and bipole/pole functions is alsomade at bipole level.

The station control is responsible for the supervision of thenumber of fIlters in operation, due to the related conditions c':'both AC and DC systems, and the reactive-power balan~eRegarding the hardware, microprocessors will be the heart c C

the master control and a redundant microcomputer stmctu:ewill be used by the two highest levels of the control hierarc}-"The functions performed by the microprocessors are differe;'-in each station, mainly in the normal operation mode. T::elead station (Foz do Igua~u) performs almost all the funeti:"related to the power transmission, the trail station (Sao RO':L2receiving and processing the final current order in synchro':-is:-with the lead station.

The communication between stations will utilise 1\\,: :e-dundant and independent channels: power-line carrie: ~:-microwave. The received telegrams will then be C0mDJ:e:. '.achieve a high degree of security. An error-rate l11onit:ri;:: :. -counting feature will be supplied which will allow J :'e~:_ance analysis of the telecommunication._ Regarding the operator, each bipole may ':'e __trom anyone of four different locations. t\C Jt oJ'::.

IS!3.tion control room and bipolc ,>J=-~::-o~:-::~-=~

·1--:~

the master station concept is dependent on the actual controllocation. The station control room at Sao Roque is to be themaster station, in normal conditions, as all control facilitiesare placed there, such as mimic AC and DC panels. When SaoRoque is acting as the master station, there is a possibility ofreceiving the power order and ramp from the system operationcentre (SaC).

The normal control mode will be synchronous powercontrol, 'synchronous' meaning that the pole-current ordergenerated by the lead station is transmitted by telecom to thetrail station. In this case, all control functions summarised asfollows are available:

(a) automatic synchronisation between the current ordersto be applied at both the rectifier and inverter

(b) keeping the earth current to minimum levels regardlessof the voltage levels of the poles

(c) compensation for loss of power when the inverter takesover the current control for any reason (this is performed bythe automatic margin regulator)

(d) automatic power-loss compensation within the bipoledue to the loss of a convertor

(e) possibility to help the speed regulation of the 60 Hzsystem

(I) stabilising the 50 Hz frequency due to loss of Itaipugenerators

(g) possibility of receiving automatic generation controlssignals

(h) utilisation of a continuous overload due to low ambientconditions

(i) utilisation of a short time overload of at least 125%during 5 s and 115% during the next 20 s following systemdisturbances and possibility of paralleling.

In some few situations it is impossible to utilise the synchron-ous power control: loss of telecommunication, loss of polecontrol equipment and complete loss of communicationbetween stations. In these cases, the following operationmodes shall be used:

Asynchronous power control: This mode will be used whenthere is a loss of the telecommunication between the polepower controllers. All the additional inputs (i.e. frequencystabilisation) and facilities are switched off and two operators(one at each station) are necessary to set a new power order.

Current control: This mode is mainly utilised due to loss ofpole control equipment, and its operation procedure is similarto the power asynchronous mode. A device named currentorder medmory allows this operation.

Automatic current margin control (AMC): It is intended tobe used when there is no communication (including voice)between the two stations. In this case, one operator at therectifier station is able to run the system. A follow up controlis utilised at the inverter.

The disturbance indices for harmonic distortion and telephoneinterference required in the specification were:

Foz do Igual.(u Sao Roquelndi\'idual harmonic::s~crtion (Dn)

T·=,~:.~harmonic::s~ = =~ion(D)T~:~~hc:ne-int1uence:'::c~== ( IIF IIT :::',=:UCI (design'J'J~ec~iT,'2 )

These :'::-:-.:~sshould not be exceeded for any foreseeableoperatlr,g condition of the HYDC transmission system and

using the worst point of the envelope of the AC systemimpedance, either with all filters in operation or with a com-plete filter bank out. For the purpose of the performancecalculations, one filter bank should be assumed as comprisingat least one filter branch of each type and at least 15% of thetotal filter fundamental frequency MYAR, independentof the switching scheme.

For the purpose of the calculation each characteristicharmonic current should be assumed as having the maximumvalue possible for the complete range of firing angles, com-mutating reactances and DC currents permitted for theoperating condition considered. The noncharacteristic har-monic currents should also be calculated pessimisticallyby assuming a combination of the most onerous conditionsof variation of firing angle and commutating reactancebetween phases within a bridge, between bridges in 12-pulsepairs, between 12-pulse pairs within a pole, between polesin a bipole, and between bipoles. The individual harmonicsshould be assumed as having their maxima occurring simul-taneously.

The filters should be considered off-tune in the calculationsand their harmonic impedances obtained by assuming acombination of maximum system frequency deviation ofduration exceeding one minute, maxumum temperaturevariation, initial mistuning, capacitor failure to the maxi-mum extent possible prior to the first alarm being generatedand detuning due to component ageing.

The harmonic impedance of the AC system was specifiedby envelopes obtained from extensive digital computercalculations of the system impedance, as a function offrequency for different developments of the system andyears of development. The calculation of Dn, D and TIFshould be done using the worst point within these envelopes.For IT calculations a table of harmonic system impedancesrelated to a specific year was given.

The harmonics have been calculated for one 12-pulsebridge and for Id in the range 10-110% Idn. For the charac-teristic harmonics, the combination of the ranges of 0;, randdx were considered. For the calculation of noncharacteristicharmonics, the same combination of 0; and dx were considered.The asymmetries which are responsible for the generation ofthese harmonics were considered either as statistical (normaldistributions defined by media fJ.- and standard deviation a) oras systematic, as shown in Table 8.

Distribution}l a

Firing angleasymmetry, 0 e1 0 0.02Transformer phasesreactance, % 0 0.33Asymmetry betweenLl-" % 0 0.66Negative sequence, % 0.5

Many runs were done using a Monte Carlo method and adistribution of harmonic magnitude and angles was obtained.A maximum value II was obtained for each harmonic byII = fJ.- + Kp a, where Kp is chosen for 99% confidence level. Toobtain the total harmonic due to n twelve-pulse groups twoapproaches were used: In = nIl (for harmonics of order 6n +1, n odd) and In = nfJ.-+ Kpa Vii (other noncharacteristic har-monics). The linear addition was considered for the 6n ± 1, nodd because they are due to asymmetries between 1 and Ll

windings and could be systematic owing to the design of thetransformer.

For Foz do Igua~u, as a result of these calculations, ascheme with two 350 MVAR and three 280.3 MVAR filterbanks has been proposed. The 350 MVAR banks contain one3rd/5th double tuned branch (69.7 MVAR), one II/13thdouble tuned branch (134 MVR) and one HP branch tunedto the 24th (146.3 MVAR), whereas the 280.3 MVAR bankscontain one II/13th and one HPbrancheach. These differentbranches are to be installed gradually as the number of 50 Hzgenerators at Itaipu power plant increases.

The highest TIF for this scheme with one filter out is36.9. To reduce this figure to the specified values the Con-sortium is studying a change in the HP filters to have 3branches tuned to the 24th harmonics and two branchesdouble tuned to 24th and 36th (or 48th) harmonics.

The total distortion exceeded 4% in many stages, evenwith all filters in (values up to 6.2), caused mainly by the 7thharmonic, for which the individual distortion is up to 1.5%in the final stage. A decision about the need for extra filteringto solve this problem will be taken later on, after a recalculationper bipole of the disturbance indices with the actual measuredvalues of the transformer reactances. This approach has beensuggested due to the fact that the noncharacteristic harmonicsof order 6n ± 1 (n odd), which are considerably influencedby the assumption of a systematic difference between thereactances of the Y and ll.-connected transformer windings,are responsible for a large percentage of the disturbanceindices which exceed the specified limits. Among those non-characteristic harmonics the seventh is particularly influential.It is foreseen that seventh harmonic branch will have to beadded to the present scheme, if additional filters are calledfor by a decision based either on the already calculateddisturbance indices or on the above mentioned recalculatedvalues. It should be noted that an increase in MVAR is notdesirable, because of problems of generator self excitationand dynamic overvoltages.

For Sao Roque a scheme with ten filter banks and twoshunt-capacitor banks has been presented by the Consortium.The total filtering MVAR (2482.5 MVAR) is divided in thefollowing independently switchable units, to be installedgradually as the transmitted power increases:

Bank size, MVAR

220.8 11th,13th

HP

11th, 13th, 3rd/5th

3rd/5th, HP

The impedance of all types of HP arms together reachesminimum around frequencies ranging from the 21 st to the27th harmonic. It is worth noting that the 11th and the 13thharmonic filters are of the damped type, with a configurationsimilar to that of the HP arms. This filter scheme met thespecified criteria. It should be noted that for both stationsthere will be filters tuned to 3rd/5th harmonics. Thesebranches are needed to meet the performance requirementsand are useful to decrease the temporary overvoltages onfault clearing, transformer energisation etc. by minimisingresonant conditions.

With reference to the IT product, the design objectives of35000 and 25000 specified for Foz do Igu~u and Sao Roque,-2spectively, are exceeded considerably for the filtering:~:'.e:·:e proposed. The maximum values of the IT product, , ,,:',~ ,":- :~'e F:z j.J IguJ;;-uand Sao Roque were of the order

, ,.: c:' ~ =:: : :':. :es:-2c:iwly. At Foz the problems are

minimised in comparison with Sao Roque, because the soilresistivity is low (about 500 D-m). At Sao Roque the problemis aggravated because the soil resistivity could be in the range1000 to 5000 D-m and the AC lines entre Sao Paulo city witha very high density of telephone circuits. To evaluate thisproblem FURNAS is conducting an extensive study thatincludes harmonic penetration analysis with three-phaserepresentation of part of the AC network, soil resistivitymeasurement for all frequencies of interest, calculation of theinduced harmonic voltage in the telephone lines as a functionof the separation between the power and telephone lines. Ifhigh interference is confirmed in this new programme ofanalysis, FURNAS is considering the conversion of the twoshunt banks into HP filters.

9 DC filter

During the short time available to perform the, studies rtquiredto determine the data to be included in the convertor speci-fications, it was not possible to do the inductive co-ordinationstudy required to specify the DC filter. Based on theexperience of other projects and very preliminary studies thatconsider different aspects of the inductive co-ordinationsuch as dangerous induced voltages, secondary interferenceand interference on controls, it was decided to ask the biddersto quote two filter alternatives: a design to limit the harmoniccurrent at the terminal of the HVDC line to the values shownin Fig. 4 (level 1) and a design to limit the currents to a mag-nitude equal to three times these values (level 2). Rulesallowing the determination of the cost of a filter design inbetween the values for levelland level 2 were also requested.With this information in hand FURNAS could choose the finalDC filter configuration after contract award. Based on thenoise criteria of 240 mVein an open wire telephone circuit,in bipolar operation and 10 dB higher noise level in mono-polar operation (less than 2% of time), it was decided toinstall the filter level 2 design. The detailed design of thisfilter resulted in a configuration with branches tuned to 100/300 Hz, 1200/2400 Hz, 120/600 Hz at Foz and 100/120 Hz,720 Hz 1440/2160 Hz at Sao Roque. The design was differentfor the two bipoles due to different lengths.

It should be noted that, for the calculation of the harmonicvoltage, the same presumptions related to transformer

I

I6.0

tn 40E~

,or<l:

c~'-:J 1.0u

'=' 08c 060r

'- 0.4aL

rC:J 0.2E><aE 0.1

1000 2000 3000frequency. Hz

Fig. 4 Current for DC filter level 1

-,:~

-~-...=:~~~

~J~~

=~:-~~~

-9

-~~~

="~~':-'"

~~~,~~:-~--~

-=:c,:-~

Table 9: Harmonic currents - bipole 1, amperes

= "equency In Monopolar Bipolarmaximumspecification FI SR FI SR

-,2100 6.0 10.3 9.9120 6.0 8.7 10.0600 1.5 2.0 2.0700 0.9 1.2 2.0

1440 0.91800 0.9 2.02400 0.6 0.9 0.9 1.1 1.2

reactance, asymetries etc. were used as indicated for the ACfilter design in the previous item of this paper.

The filter configuration mentioned in the preceding textwas considered complex, due to the double tuned frequencycharacteristics, and an optimisation of this design was started.This new calculation took into consideration that some ofthe low harmonic current values indicated in Fig. 4 could behigher and still satisfy all aspects of the inductive co-ordination. Therefore, the optimisation should try to obtainbetter results in relation to the telephone interference inSao Roque where the number of telephone circuits is verylarge.

The new filter design has arms tuned to 100/300 Hz,600 Hz 1200 Hz at Foz and 120/360 Hz, 720 Hz, 1440 Hz atSao Roque, and deviates from the specification mainly in thelower harmonics as shown on Table 9.

These higher values were considered acceptable and, inaddition, were compensated by improvements in other fre-quencies, so that the overall interference was reduced asseen in Fig. 5. It should be noted that the curves for Fig. 5

were obtained with the worst consistent set of harmonicvoltages, not with the voltages that gave the maximumindividual harmonic. This design was accepted by FURNASand. as a result, few telephone circuits will need to be changed.

It should be noted that the filters are rated for the con-:itions of one branch out of operation. In this situation, in~i;Jia operation the telephone interference will be higher,'':: o::illess than monopolar operation with all filter branches in.

of the DC lines, around 800 km, indicates a possi-::' [osonance at second harmonic depending on the DC

~e~~::~ sizo, Owing to this a study was carried out in twoJ~~'S: :':~st, ,oJ find the condition where the second harmonic~oS'J~;~~,:o::uld occur and, secondly, to evaluate the systemc'''o~';cltEos j, such a resonant condition.

For the first study a digital calculation was used to evalU3.tothe resonant frequency of the circuit comprising internsiimpedance of the convertors and of the DC reactor at botLends and the DC line. DC filters were not included, as theFURNAS specification required operation without them.If DC mters are included in the calculation, the secondharmonic resonance is avoided as they have branches tunedto that frequency as indicated in the previous item. Variousmodes of operation were considered in the calculation and theresults of the resonant frequencies, as a function of DC reactorsize, are shown on Fig. 6.

130[ , I

1'0 Ii!!If il~~~I~~!ilii!il!!li!i!ii:iiii:iiiilli:i,

~ , ~-l "'. ,I~ , ~I...... b' I90 monopolar·· " Ipo ar

(,300kV "~" ",,600kV

monOPol~r~OOkV ",~ ')1",-I ....•...•.

800,1 0.2 0,270.3

DC reactor inductance04

H

Fig. 6 DC line resonance---- 815 km....... 795 km

To analyse the results it is important to have in mindthat the temporary frequency variation specified was 0.95 ~1.15 p.u. for Foz and 0.985 - 1.06 p.u. for Sao Roque,corresponding to second harmonic frequencies of 95 to 115Hz and 118 to 127 Hz. If all modes of operation are con-sidered together with this frequency range, it can be seen(Fig. 6) that a resonant condition always exists for DCreactor sized from 0.2 to 0.6 H. If higher values are used thesecond harmonic is avoided, but first harmonic resonancecan appear leading to a worse condition.

It should be mentioned that, for the calculation of thecurves shown in Fig 6, a value of 7.33 times the transformerimpedance was considered for the internal impedance of theconvertor. This value was chosen based on some DC simulator,and theoretical investigation was undertaken for the 'char-acteristic harmonics. These conclusions, however, were extra-polated for the low harmonics. On the other hand, it is realisedthat a resonant condition could lead to high currents andovervoltages. However, it is expected that some action of thecontrol of the HVDC system will introduce damping to theseoscillations, reducing the possible overvoltages andovercurrents.

One reactor of size 0.27 H was initially considered becauseof transportation limitation, and also because more than onereactor would not solve all resonance conditions (Fig. 6).With this reactor size the following investigation wasundertaken:

Bipolar operation with two twelve-pulse groups per polewas chosen, because, in this situation, the resonant frequency

4':1

is close to 100 Hz (see Fig. 6). With the system set upcontaining loss compensation devices in transformers, reactorsand valves, a phase-to-ground fault on the valve side of theconvertor transformer was simulated. The rectifier sourcefrequency was adjusted to around 50 Hz to get the maximumtotal direct voltage plus second harmonic on the DC side.Comparative cases were evaluated with different gains in therectifier current control loop.

It was verified that, for normal gain, the overvoltage waslower than 1.4 p.u. With gain reduced to 1/9 of the normalone the overvoltage reached values above 1.7 p.u. These casesindicate the positive damping effect of the controls for thisphenomenon. The concept of control action and the DCreactor size was, in principle, accepted by FURNAS, butsome doubts remained due to possible influence of the simu-lator losses on the phenomenon. It was agreed with theConsortium to accept, in principle, a DC reactor size of 0.27H; but to make final investigation with the actual ItaipuHVDC control system on the new simulator under con-struction that will have much lower losses.

1 PEIXOTO, C.A.O., PORTELA, C.M., LOBLEY, D.J., and JARDINI,J .A.: 'Design optimization and environmental impact for the Itaipu± 600 kV HVDC lines and electrode lines'. Proceedings of CIGREsymposium on transmission lines and environment, Stockholm,Sweden, 1981

2 PEIXOTO, C.A.O.: 'Itaipu 6300 MW HVDC transmission systemfeasibility and planning aspects'. Proceedings of the symposiumon incorporating HVDC power transmission into system planning,Phoenix, Arizona, USA, 1980, pp. 211-236

3 PEIXOTO, C.A.O., CLARKE, C.D., JARDINI, J.A., and MOUNT·FORD, J.D.: 'Simulator and digital studies on Var compensation forthe Itaipu HVDC transmission system'. Proceedings of the IEEE1980 international conference on overvoltage and compensationon integrated AC-DC systems, Winnipeg, Canada, 1980, pp. 27-42

4 BUENO, E.G., JARDINI, J.A., JOHANSON, A., PAAJARVI, B.,and EKSTRON, A.: 'Insulation co-ordination of the Itaipu HVDCstations application of ZnO arresters'. CIGRE SC 33 meeting,Rio de Janeiro, Brazil, 1981

5 PEIXOTO, C.A.O., TAAM, M., FRONTIN, S.O., FIGUEIREDO,E.F., PORANGABA, H., and SVENSSON, S.: 'Itaipu HVDC trans-mission-stability studies'. CIGRE SC 14 meeting, Rio de Janeiro,Brazil, 1981