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Problem of synthetic circuit-breaker testing withvoltage injectionL. Blahous, Ph.D.

Indexing terms: Circuit breakers, Electron device testingAbstractWhen testing high voltage circuit breakers synthetically the question arises as to how the test results can becompared with test results in direct tests. In the paper the thermal equivalence between direct testing and themethod of voltage injection is investigated. The circuit breaker is modelled by the Urbanek equation. It is shownthat the prospective voltage alone cannot define the test circuit to ascertain thermal equivalence. To lay out thesynthetic test circuit to obtain results equivalent to direct testing requires the exact knowledge of all circuitelements including the auxiliary breaker and the test breaker. There is no problem, however, if this method isapplied for tests in which the breaker is dielectrically stressed.

1 IntroductionWith the increase of the system voltage of electric networks,not only the rated voltage for circuit breakers increases, but also theshort-circuit power to be interrupted. This development made itnecessary to increase the specific breaking capacity of arc chambers;thus, today, even one single breaking unit of a circuit breaker canhardly be tested directly. Therefore synthetic testing has gained moreand more importance as it is the only way to keep expenses for

circuit-breaker testing within reasonable limits.The question that arises when using synthetic test circuits is towhat extent they can be used, i.e. to what extent the stress on thecircuit-breaker in a synthetic test circuit is equivalent to one in adirect test circuit. It seems relatively easy to define conditions thatascertain equivalence for test circuits with current injection. Thoseconditions are generally accepted and there seems to be agreement

Fig. 1Direct reference circuit

Fig. 2Synthetic test circuitPaper 8399 P, first received 20th April and in revised form 2nd August 1979Dr. B lahous is with the Technical University of Vienna, Gusshausstrasse 25 ,A-1040 Vienna, AustriaPROC. IEE, Vol. 126, No. 12, DECEMBER 1979

that, provided these conditions are fulfilled, the circuit breaker isequally stressed in the synthetic circuit as in the equivalent direct testcircuit.For the voltage-injection synthetic-test method the question of thethermal equivalence still seems to be unsatisfactorily answered. It isthe goal of this work to contribute to the understanding of the inter-action period for that test method.2 Problem of circuit layout

Fig. 1 shows the direct reference circuit and Fig. 2 thatcircuit used for this theoretical investigation on the thermal equiv-alence.In Fig. 2, the damping resistor of the high-voltage circuit is in serieswith the inductance. There are also test circuits in use with theresistor in parallel to the inductance which can mean certainadvantages for shaping the prospective voltage curve. This does not,however, change the results of this investigation.The essential feature of voltage injection is that the prospectivevoltage is not produced by one circuit only but that both the high-current and the high-voltage circuits contribute to it. The initial partof the prospective voltage comes from the high-current circuit,whereas the high-voltage circuit is connected, via the spark gap, aftercurrent zero, contributing the rest of the prospective voltage to thepeak. This means that the test circuit for synthetic testing withvoltage injection is basically different from an equivalent direct circu it.This method proves the claim of the IEC test philosophy that minorchanges in the prospective voltage to the first peak also mean onlyminor changes in the test circuit, to be wrong. That the layout of sucha synthetic test circuit really is a problem is shown by the differentsuggestions on how to do it.Slamecka1 restricts his layout to the approximation of theprospective voltage curve of the direct reference circuit (which fulfilsthe IEC requirements). The question of when this approximation issufficient is not answered . Also, the prospec tive voltage curve of thesynthetic test circuit can follow the prospective voltage curve of thedirect reference circuit in close approximation only to the first peak,then the two curves generally differ considerably. This shows that thepresent IEC requirements are not at all unequivocal. There are twobasically d ifferent circuits yielding two basically different voltagecurves and only because these curves are very similar to the first peakcan both circuits be used to test whether or not the breaker can inter-rupt a certain short circuit power as far as the IEC requirements areconcerned.Kopplin and Pflaum,2 on the other hand, tried to show how to layout the high-current circuit in order to ensure that the breaker isthermally equally stressed as in the direct circuit. In their considera-tions, the prospective voltage is not even mentioned. The onlyproblem they are concerned with is how to lay out the circuit in orderto ensure equal current distortion before current z ero. Their proof,however, provokes a number of objections:(a ) The Mayr model was used to describe the circuit-breaker arc. Itwas shown, however, that before current zero the arc is of the Cassietype (Urbanek,3Swanson, Roidt and T.E. Browne).4(b) The thermal equivalence was proven only for the test breakerbeing identical with the auxiliary breaker. This cannot be generallyascertained in practice. There are test stations that use a welltested breaker as the auxiliary breaker no matter which breaker isbeing tested. Even if both the test breaker and the auxiliary breakerare two different poles of a 3-pole unit, identity between them is not

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ascertained. Their parameters may scatter considerably during aninterruption.(c) Kopplin and Pflaum2 connect the capacitances Ce, Cp and C zto one single capacitance parallel to both breakers. It is questionablewhether this is permissible. It would mean that there is equal voltagedistribution across each breaker which is exactly contrary to therecommendation by Slamecka.1 Both breakers must interrupt thecurrent at exactly the. same time or th e post-arc current of the testbreaker will be interrupted by the auxiliary breaker, which certainlydoes not happen in a direct test circuit. It would also mean that theonly effect of the capacitances Cp and C z is a contribution to thecapacitive current coming mainly from Ce and flowing through bothcircuit-breakers. This approach does not take into account the factthat in parallel to both the auxiliary breaker and the test breaker thereare two different capacitances Cp and C z. Each one forms a closedloop together with the breaker to which it is parallel. Thereforedifferent currents flow in each breaker.(c/) Finally it. was shown that equal cu rrent distortion beforecurrent zero does not necessarily mean equal thermal stress. Whathappens after current zero is at least as important for the thermalbehaviour of the circuit-breaker arc.s

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25 50 75 100Fig. 3Com parison of prospective vo ltage curves of two syn thetic circuitswith direct reference circuit

direct reference circuitRp = 0- Rn = 700 n

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3 Therma l equivalenceThe above questions can only be answered by either testingwhat actually happens when interrupting a short circuit current in asynthetic test circuit with voltage injection or by simulating the whole

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Fig. 5Current in synthetic test circuit w ith Rp = 700 1 around current zero(test breaker identical with auxiliary breaker)current in test breaker current in auxiliary b reaker genera tor current - current through C p

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Fig. 4Currents in synthetic test circuit with Rp = 0 around current zero(test breaker identical with auxiliary breaker)current in test breaker- - curre nt in auxiliary breaker_ _ _ generator current - current through Cp

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Fig. 6Current in synthetic test circuit with auxiliary breaker being muchbetter than test breaker (for circuit details see Table I)current in test breakercurrent in auxiliary breakergenerator currentcurrent through Cp

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Table 1EQUIVALENT CIRCUITS(1) (circuit of Fig. 4)L = \\m\\;Re = 120f t ; C e = 4 0 0 n FCz = 200pF ; K 2 = 0Cp = 90nF;Rp = 0Ua = 1 l-9SkV,(C/ cq = 47-8 kV)

/ = 34-6 kALh = 5mH;Rh = 400f t ; C h = 1 0 / J F/Ch = 57-5 kVTest breaker:Po = 25kW;e = 8-4kV;T = 3 juswd = 400kV/I uxiliary breaker:Po = 25k W;e = 8-4 kV; 71 = 3 jusud = 400kV

(5 )

(7 )

(2) (circuit of Fig. 5)L = l - l mH; / ? e = 120f t ;C c = 400nFC2 = 200pF;7? 2 = 0C p = 90nF ; / ? p = 700 2E7g = 15-63kV,(E7cq = 62-5 kV)/ = 45-25 kALh = 3OmH;7?,, = 0;C h = lOyuFUch = 75-1 kVTei/ breaker:Po = 25kW;

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the breaker is much more severely stressed thermally than in circuit(a) (which is the relevant direct circuit), while circuit (c) is mucheasier than circuit (a), even though the total prospective voltage wouldsuggest circuit (c) to be the most severe one. This example showsagain how little information really can be gained from the prospectivevoltage.The capacitance Cp parallel to the auxiliary breaker has practicallyno influence on the prospective voltage because it is in series with thevery small capacitance C z which completely dominates. So theprospective voltage curve/ that is determined by the high-currentcircuit is defined by Ce. Yet Figs. 4 and 5 show that Cp has consider-able influence on the behaviour of the circuit breakers. In Fig. 3 thecapacitive current across Cp delays current zero in the auxiliarybreaker considerably. This results in a very small thermal stress for theauxiliary breaker which interrupts without any post-arc current. Fig. 4also shows that not all of the capacitive current from Cp flows acrossthe auxiliary breaker. Part of the capacitive energy of Cp, which isloaded to the full arc voltage before current zero, is also fed into thetest breaker just at its most critical period of post arc current, thuscausing a thermal reignition of the arc in the test breaker.

In the case of circuit (c), the capacitance Cp is not loaded to thefull arc voltage as some of its drops on the resistor Rp in series.Because the energy in a capacitance.varies with the voltage squared,there is much less energy stored in Cp and some of this is alsoabsorbed by the resistance Rp. The capacitive current from Cp ismuch smaller and current zero of the auxiliary breaker appears at thesame time as in the test breaker. Only little additional energy is fedinto the test breaker.Fig. 6 shows a case where there is a completely different reason forthe larger thermal stress on the test breaker in the synthetic testcircuit than the one in Fig. 4. The auxiliary breaker shows a muchbetter arc-quenching performance than the test breaker. It interruptsthe current before the test breaker. Therefore the capacitance Cp isput in series with the capacitance Ce which considerably reduces theeffective capacitance parallel to t he test breaker. Once again theinsufficiency of the IEC specifications is shown. Because theprospective voltage only shows the response of the circuit to aninterruption with an ideal breaker, the very important role of theauxiliary breaker principally cannot be taken into account by it.The calculations made for the cases shown in Fig. 4 and 5 suggestthat for a certain value of Rp the thermal stress in the syntheticcircuit is equal to the stress in the direct reference circuit. For thespecial example considered this is the case for Rp = 4 5 0 2. For this

value of Rp the test breaker can interrupt exactly the same shortcircuit power in the synthetic test circuit as in the direct referencecircuit.As, in practice, the identity between auxiliary and test breakercannot be assumed, the question remains whether thermal equivalencecan also be obtained if the breakers are different.Table 1 shows some examples where the auxiliary breaker isdifferent from the test breaker. The auxiliary breaker constants werearbitrarily chosen and the other elements have the values for which

thermal equivalence was obtained. To obtain those values required alot of calculation because the method of trial and error had to beused. No general or empirical law was found on how to lay out thetest circuit to obtain thermal equivalence once the constants for testand auxiliary breaker are known. There was, however, no hint thatthere might be breaker combinations for which thermal equivalencewere principally impossible.4 Conclusion

The method of synthetic testing with voltage injection showsclearly the limits of the IEC specifications imposing conditions onlyon the prospective voltage. The prospective voltage can neither showthe influence of the capacitance Cp parallel to the auxiliary breakernor the influence of the auxiliary breaker itself. Yet both elementshave great influence on the thermal behaviour of the test breaker.Therefore, in a test circuit laid out to fulfil the IEC requiremen ts,only the thermal behaviour of the test breaker is completelyundertermined. A prediction of its thermal behaviour from theprospective voltage curve alone is impossible. If the complete circuittopology is taken into account and both the test breaker and theauxiliary breaker are known, it is possible to lay out the circuit tostress the test breaker thermally equally as a direct circuit does. Theprospective voltage is then only of minor importance.These objections, however, only concern early dielectric or thermalreignition, but by no means late dielectric reignitions occurring manyarc time constants after current zero. There is no doubt that especiallycircuits of low natural frequency, which cannot be simulated bycurrent-injection schemes equivalently and w hich usually do not causethermal b ut late d ielectric reignition, can only be effectively simulatedby voltage-injection schemes. Therefore the question is not which ofthese two type s of circuits are more useful or more reliable but merelyunder which condition they must be used.5 Acknowledgment

The present work was financed by the 'Fo ndszu r Foerderungwissenschaftlicher Forschung' of the Austrian Academy of Sciences.

ReferencesSLAMECKA E.: 'Priifung von Hochspannungs-Leistungsschaltern' (Springer,Berlin, Heidelberg, New York 1966)KOPPLIN H., PFLAUM E.: 'Zur Aequivalenz der synthetischen und direktenPruefungen von Hochleistungsschaltern' ETZ-A, 1963, 84, pp. 14 9-153URBANEK J.: 'The time constant of high voltage circuit-breaker arcs beforecurrent zero' Proc. IEEE, 1971,59, pp. 502-508SWANSON B.W., RO1DT R.M., BROWNE T.E. Jr.: 'A thermal arc model forshort line fault interruption' ETZ-A 1972, 93, pp. 375-38 0BLAHOUS L.: 'Anwendung der dynamischen Lichtbogentheorie auf dasVerhalten von Hochspannungsschaltern in Priifkreisen' Doctorial thesis at theTechnische Universitat of Vienna, 1978URBANEK J.: 'Zur Berechnung des Schaltverhaltens von L.eistungsschaltern,eine erweiterte Mayr-Gleichung', ETZ-A 1972, 93 , pp . 381-385

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