Field Tests of Generator Differential Protection with Balanced No-Load EMF of Current Transformers

L. V. Baginsky

(Presented by Prof. I. D. Kutyavin, Doctor of Technical Sciences)

Field tests were carried out in order to verify the principal conclusions concerning the properties of generator differential protection with balanced no-load EMF of current transformers, set forth in [1] and [2]. From [1] it is known that the current transformers in the specified circuit operate in the no-load mode. In [2] it is shown that such an operating mode is permissible if the cross section of the steel of the spiral core does not exceed \(q \leq 17\ \mathrm{cm}^2\), and if the number of turns of the secondary winding \(w_2\) and the value of \(q\) are selected from the condition of limiting the maximum possible value of the amplitude of the secondary EMF, \(E_{2i} \leq 1000\ \mathrm{V}\). The possibility of such a limitation lies in the fact that, with a sinusoidal primary current \(I_1\), the dependence \(E_{2m} = f(I_1)\) is nonlinear, and \(E_{2m}\) increases more slowly than \(I_1\).

The first stage of the field tests consisted in checking these propositions under operating conditions. For this purpose, an experimental specimen was made by modifying current transformers of type TPSF 10--4000--3. A hot junction of a carefully calibrated copper-constantan thermocouple was fastened to the inner surface of the core torus. The core was insulated with three layers of micalenta and two layers of asbestos tape. The secondary winding was made with PSD wire and was not impregnated. The current transformer manufactured in this way had the following parameters: length of the mean magnetic line of the core \(l_{\mathrm{av}} = 104\ \mathrm{cm}\), \(q = 5.9\ \mathrm{cm}^2\), \(w_2 = 10\) turns. Under the test conditions, a number of measures were provided to ensure a sufficient reliability margin in the experimental results, namely:

1. The core cross section was increased by a factor of 2--3 in comparison with the requirements of the protection circuit under study.

2. The temperature measured was that of the most heated section of the core surface, not the temperature of the secondary winding.

3. The core was insulated with unimpregnated asbestos tape, which substantially worsened its cooling conditions.

The thermal-stability tests were performed by passing the operating current of a 1000 kVA transformer, \(I_1 = 750\ \mathrm{A}\), for a long time through the primary winding of the tested current transformer, which corresponds to a field strength in the core \(H = 10.2\ \mathrm{A/cm}\). At

In this case the overheating temperature of the steel proved to be \(\tau_c = 8^\circ\mathrm{C}\). The experiment was then repeated on a 50 MW generator at \(I_1 = 3200\) A (\(H = 43.5\) A/cm), and the overheating temperature obtained was \(\tau_c = 39^\circ\mathrm{C}\). On the basis of these measurements one may conclude that the dependence is approximately linear. Therefore the value of \(\tau_c\) can also be estimated approximately for other values of \(I_1\). For example, at the nominal primary current of the same current transformer, \(I_{1n} = 5000\) A, \(H = 68\) A/cm and, consequently, one may expect \(\tau_c = 61^\circ\mathrm{C}\). Assuming that the ambient temperature should not exceed \(\Theta_0 = 35^\circ\mathrm{C}\), we obtain the maximum possible core temperature \(\Theta_c = 96^\circ\mathrm{C}\). Therefore, for manufacturing current transformers at \(I_{1n} \le 5000\) A, PBD wire with impregnation may be used, insulating the core with electrical cardboard. At \(I_{1n} = 6000\) A the temperature of the steel will increase only slightly, since \(l_{\mathrm{av}}\) will increase substantially and the value of the field strength in the core will rise little.

Thus, in order to ensure thermal stability when the core is heated in the no-load regime, the secondary winding should be made of PBD wire with impregnation if \(I_{1n} \le 5000\) A, and of PSD wire if \(I_{1n} = 6000\) A. From the standpoint of thermal stability, there are no obstacles to placing current transformers of differential protection with no-load emf balance in the same casing with transformers of another purpose.

Fig. 1

The described sample of the current transformer was also used to obtain the dependence of the specific amplitude of the secondary emf on the strength of the sinusoidal magnetic field in the core, \(l_{2m} = f(H)\). For this purpose the test sample was mounted on a 1000 kVA transformer and on 25 and 50 MW generators connected in short circuit. The results of this experiment are given in Fig. 1. They indicate the nonlinearity of the dependence \(l_{2m} = f(H)\), repeating the character of the curve in Fig. 3 in [2].

To verify the basic properties of differential protection with no-load emf balance of the current transformers described in [1], and also for a quantitative assessment of these properties, a three-phase set of this protection was manufactured. The parameters of the transformers

of this set were as follows: \(q = 2.34\ \text{cm}^{2}\), \(l_{\mathrm{av}} = 77\ \text{cm}\), \(w_2 = 340\) turns. The cores were not annealed and were insulated with electrical cardboard. The secondary winding was made with PSD wire without impregnation, since, in the time allotted for preparation for the tests, it was impossible in the workshop conditions of the Department of Electrical Stations, Networks, and Systems of the Tomsk Polytechnic Institute to wind it with PBD wire with impregnation. Figure 2 gives the no-load characteristics of the current transformers of the circuit under test, \(E_2 = f(I_1 w_1)\). The greatest difference is observed between the characteristics of transformers Nos. 1, 5, and 6 on the one hand, and transformer No. 2 on the other. This difference occurs in the region of medium inductions and reaches 41.5%, which is probably explained by substantial deformations of the core strips of transformer No. 2, which they experienced when the core was rewound.

The circuit of differential protection with balancing of the no-load emf of the current transformers was installed on a 25 MW, 10.5 kV generator. The current transformers were distributed by phases in such a way as to obtain the largest unbalance currents during the tests. For this purpose, transformers Nos. 1, 5, and 6 were placed on the neutral terminals of the windings of phases \(K\), \(Z\), and \(Zh\), respectively, and transformers Nos. 2, 3, and 4 respectively on the same phases, but on the side of the line terminals. Thus, the greatest discrepancy of characteristics was in the current transformers of phase \(K\) (41.5%).

Fig. 2.

The connection diagram of the secondary and operating circuits during the tests is shown in Fig. 3. Current relays \(T\) used were relays of type ET-521/0.6; the intermediate relays were P-EP-101, and the additional inductive resistances \(D\) were reactors with a small air gap. The latter were not used in all experiments. With the aid of an oscillograph, the currents of the secondary and primary circuits were recorded (trace \(I_1\) was supplied from class 0.5 measuring current transformers \(TT_i\)) and the line of contacts of relay \(P\).

Phase \(K\) of the circuit was in the most severe conditions because of the mismatch of the characteristics of the current transformers. Therefore, the results obtained during the tests of this phase of the circuit are the most representative. In what follows we shall consider primarily these results.

As is known, a sudden short circuit in the zone of differential protection with balancing of the no-load emf, installed on separately operating generators with a large stator-circuit time constant, may be cleared with a delay. The delay may be overcome by increasing the time constant of the secondary circuit \(T_2\) [1] or by increasing the sensitivity coefficient

\(\kappa_4\). The main purpose of arranging sudden short circuits in the protection zone in this experiment is to verify the second method of eliminating delays. In all, 12 sudden short circuits were produced in the protection zone at a generator voltage \(U_2 = 6.3\) kV, which yielded a transient-current value \(I'' = 9\) kA.

Fig. 3.

Phase \(K\) of the circuit was adjusted to a time-constant value \(T_2 = 0.01\) sec. With an active resistance of the circuit \(r_2 = 2.42\) ohm, this required the inductive reactance of the relay circuit to be \(\omega L_2 = 6.2\) ohm. On the basis of laboratory tests of the circuit with current transformers Nos. 1 and 2, for the given values of \(r_2\) and \(\omega L_2\) the relay setting \(I_y = 0.5\) A was adopted, with a reliability factor \(k_n = 2\) (Fig. 6). This setting provided a value \(\kappa_4 \simeq 4.4\) for the conditions of this experiment, when the greatest value of the primary current was \(I_1 = 9\) kA (Fig. 5).

Fig. 4.

The tests showed that in none of the cases were there any delays in the operation of phase \(K\) of the circuit. A typical oscillogram of the pro-

The process in the circuits of phase \(K\) of the protection under a sudden short circuit in the zone is shown in Fig. 4. Because of the limited permitted number of short-circuit switchings, it was not possible to test phase \(K\) of the circuit at other values of \(T_2\) and \(k_4\). We shall only note that, at approximately the same value of \(T_2\) for phase 3 of the circuit (\(T_2 = 0.012\) sec), but at \(k_4 \simeq 2.5\), delays in operation of the protection occurred. Thus, increasing the value of \(k_4\) is an effective means of eliminating delays in protection operation when the generator is operating isolated. Since the inductive additional resistances in the relay circuit do not increase the values of \(k_4\), it is advisable to dispense with them if the selectivity of protection operation is thereby ensured.

To determine the possibility of delays arising in operation when the protected generator is operating in the system, on phase \(Zh\) of the circuit, instead of a choke, a shorting jumper was installed, connecting the phase conductor of the circuit to the neutral conductor. This simulated an internal fault with equal currents flowing into the fault point from the system side and from the generator. It should be noted that the values of \(T_2\) and \(r_2\) in this test were considerably lower than in the simulated condition. In fact, in the event of a short circuit in the protection zone with two-sided feeding, each current transformer is loaded by a load consisting of half the sum of the resistances of the relay \(z_r\) and the connecting phase conductor \(z_{pr}\), which for the tested circuit is \(1.5\) ohms. In this test the load was equal to \(0.5(z_{pr} + z_{onp} + 2z_r)\), where \(z_{onp}\) is the resistance of the neutral conductor, which gave a resulting resistance of about \(3.2\) ohms. The setting of the relay of phase \(Zh\) of the circuit was selected as \(I_y = 0.3\) A at \(k_n > 2\). However, even under these conditions not a single case of delayed operation was observed. This same test indicates sufficient thermal stability of the coils of relay ET-521/0.6 (one-second current \(I_{1s} = 45\) A with the coils connected in series) in the short-circuit condition within the protection zone with two-sided feeding, since the current values in the relay did not exceed \(8\) A.

To determine the primary operating current \(I_{1sr}\), as well as the current in the relay in the event of a fault in the protection zone \(I_{2n}\), the dependence \(I_{2n} = f(I_1 w_1)\) was taken with the above-specified parameters of the relay circuit of phase \(K\) and one-sided feeding of the short-circuit point in the protection zone. The results of this test are shown in Fig. 5 by curve 2. In the same place, curve 1 shows the result of testing the pair of current transformers Nos. 1–5 under laboratory conditions at the same values of \(r_2\) and \(\omega L_2\) as for curve 2. The discrepancy between curves 1 and 2 is explained by the fact that, when curve 1 was taken, the current transformer whose primary winding was carrying current \(I_1\) was loaded more heavily by the inductive resistance of the second transformer than when curve 2 was taken (Fig. 2). Since the difference in the characteristics of current transformers Nos. 1 and 2 is most significant in the region of small and medium inductions, it should be expected that at very large \(I_1\) curves 1 and 2 must coincide. It was not possible to verify this point, since it was not possible to obtain a steady-state short-circuit current greater than \(2800\) A.

As can be seen from Fig. 5, the placement of current transformers with the higher characteristic on the neutral-terminal side gives significant advantages in the value of \(I_{1sr}\) compared with their placement on the line-terminal side. For example, at \(I_y = 0.5\) A, if current transformer No. 1 is placed on the neutral-terminal side, \(I_{1sr} = 220\) A, or \(I_{1sr} = 0.13 I_{n2}\), whereas with coincident characteristics, \(I'_{1sr} = 600\) A, or \(I'_{1sr} = 0.35 I_{n2}\). When current transformer No. 1 is placed on the line-terminal side, one can expect \(I_{1sr} \geq 0.7 I_{n2}\).

Such an arrangement of the current transformers can be implemented without additional time expenditure, since even at present the volt-ampere characteristics of the current transformers of the differential protection are taken by the operating personnel.

Fig. 5.

Thus, in this experiment, with an additional inductive resistance in the relay circuit of \(5.2\ \Omega\) and \(k_n = 2\), phase \(K\) of the protection has a primary pickup current \(I_{\text{sr}} = 0.13 I_k\) and a sensitivity coefficient \(k_4 \simeq 4.4\). In reality, the current of a two-pole short circuit at the terminals of the protected generator will be not \(9\), but \(12\ \text{kA}\), and therefore \(k_4 \simeq 6\).

The selectivity of the protection under test was checked by measuring the unbalance currents \(I_2\) in the steady-state condition and by oscillographing the currents in the protection circuits and in the contact line of relay \(\Pi\) (Fig. 2) during external sudden short circuits.

The greatest value of the unbalance current in the steady-state condition was determined by taking the characteristics \(I_2 = f(I_1 w_1)\), where \(I_2\) was measured by means of relay ET-521/0.6. These dependences were first taken in the laboratory during preparation for the tests, and then during the tests; in both cases identical results were obtained. As expected, the greatest unbalance current was recorded in phase \(K\) of the circuit (Fig. 6). Its maximum is located at \(I_1 \simeq 330\ \text{A}\) and reaches the value

Fig. 6.

0.25 A, and at \(\omega L_2 = 5.89\ \Omega\). It follows from Fig. 6 that increasing \(\omega L_2\) slightly reduces the maximum value of \(I_2\), but, on the other hand, it greatly reduces \(I_2\) at large values of \(I_1\). This is explained by the predominance of the fundamental harmonic in the current \(I_2\) at small values of \(I_1\), caused by a considerable divergence of the characteristics of the current transformers. As \(I_1\) increases, the content of higher harmonics in the current \(I_2\) rises sharply because of saturation of the steel of the transformer cores, as a result of which \(I_2\) is substantially limited by \(\omega L_2\). It is therefore inexpedient to try to limit the current \(I_2\) by additional inductive resistances if its maximum lies in the region of small currents. These considerations made it possible to exclude, from the circuit of the phase \(K\) relay, the choke before the external short-circuit test. The relay coils were connected in series and the current \(I_2\) was measured again. It had the following values: at \(I_1 = 2800\ \mathrm{A}\), \(I_2 = 0.15\ \mathrm{A}\), and at \(I_1 = 350\ \mathrm{A}\), \(I_2 = 0.3\ \mathrm{A}\).

In the two other phases of the circuit it was not possible to measure the currents \(I_2\) during the tests carried out, since, with the relay coils connected in series, they did not operate at \(I_y = 0.15\ \mathrm{A}\). This indicates very high values of the security factors in these phases of the circuit.

Sudden short circuits outside the protection zone were included in the test program in order to check the selectivity of the protection, and also to show that the largest unbalance currents are caused by the periodic primary current. To place the protection under the most severe conditions, the relay coils were connected in parallel to reduce their resistance, and minimum settings \(I_y = 0.3\ \mathrm{A}\) were set on the relays of phases \(Z\) and \(Zh\). On the relay of phase \(K\), the setting could not be made minimum, since with the coils connected in parallel and \(I_1 = 350\ \mathrm{A}\), \(I_2 \simeq 0.45\ \mathrm{A}\); that is, when using an ET-521/0.6 relay, the value \(k_n = 2\) could not be maintained. However, the ET-521/0.6 relay was not replaced by an ET-521/2 relay for the following reasons. The value of the largest unbalance current of phase \(K\) occurs at small \(I_1\) and therefore depends only on the ratio of the current-transformer characteristics in the region of small and medium inductions and does not depend on \(I_1\) and \(T_1\). Since the characteristics of the current transformers are very stable, in this case it is possible to reduce the values of \(k_n\) substantially, down to the values adopted in maximum-current protection. In this test, the relay slider was moved to the right up to the limit, which provided \(I_y \simeq 0.65\ \mathrm{A}\).

With this adjustment of the circuit, the protected generator was switched six times onto a short-circuit link installed outside the protection zone, at \(U_2 = 6.3\ \mathrm{kV}\). Not a single operation of the protection was observed in this test. Analysis of the oscillograms shows that an aperiodic component does not appear in the current \(I_2\), and, when an aperiodic component is present in the primary current, the current \(I_2\) decreases sharply. These observations fully agree with the conclusions made in [1], and make it possible to measure the largest value of the unbalance current without conducting special tests during normal operation of the generator.

Since the introduction of additional resistances into the relay circuit is undesirable for reasons of operating reliability of the circuit and because of their negative influence on the value of \(k_i\), the resistance of the relay coils should be used to limit the higher harmonics of the unbalance current. For this purpose it is necessary to use, in the circuit, a relay with predominantly inductive coil resistance (ET-521/0.6 or ET-521/2 with the coils connected in series). The same relay can be used to measure the unbalance current in the case when the characte...

the current-transformer cores differ considerably. In order to maintain the values \(k_n=(1.5—2)\), it may become necessary to shift the relay slider beyond the setting scale to the right. In this connection, when adjusting the relay at the factory it is desirable to position the return-spring arm so that it permits the relay slider to be shifted as far as possible to the right. Then the selection of settings can be carried out as follows. In the case of good agreement of the current-transformer characteristics, the current \(I_2\) cannot be measured by means of the relay. Then the maximum setting should be adopted with the relay coils connected in series, which provides the value \(k_n>2\). If the current \(I_2\) can be measured by means of the relay, but it does not increase as \(I_1\) is decreased, then the setting must be determined as \(I_y=(2—3)I_{2n}\), where the current \(I_{2n}\) is the current \(I_2\) changed at \(I_1=I_{n2}\). For a protected generator capacity of 25 MW or more, one should take \(k_n=2\). If the maximum value of the current \(I_2\) occurs at small \(I_1\), one should take \(k_n=(1.3—1.5)\).

After the completion of testing the protection, it was left in trial operation on the same generator, acting on a signal. During trial operation, 6 external short circuits were recorded, to which the protection did not respond.

As follows from [1], high values of the amplitude of the secondary e.m.f. of current transformers, determined by the properties of the core steel, and the tendency to limit the value \(E_{2m}\) to too small a magnitude, may constitute an obstacle to obtaining the required sensitivity of differential protection with balancing of the no-load e.m.f. at high capacity of the protected generator. In the course of operating the protection, the values of \(E_{2m}\) were measured at full generator load. The current transformers installed on the neutral leads had the highest values \(E_{2m}=520\) V ampl.; the smallest value \(E_{2m}\) was exhibited by current transformer No. 2, \(E_{2m}=300\) V ampl., which is explained by the gentle course of its characteristic in the region of low inductions. The presence of a significant unbalance current in phase \(K\) of the circuit reduced the voltage between the phase and neutral wires of the circuit to 375 V ampl. at 520 V ampl. in the two other phases.

Consequently, in order to reduce \(E_{2m}\), it is necessary to have gentle characteristics of the current transformers in the region of low inductions. This can be achieved with a simultaneous simplification of the manufacturing technology of current transformers, if annealing of the cores after their manufacture is abandoned. The omission of annealing cannot lower the values of \(K_1\), since the transformers operate at very high inductions, and the magnetization curve in the region of high inductions does not increase upon annealing [3]. On the other hand, the limitation proposed in [2] for the value \(E_{2m}<1000\) V ampl. can hardly be considered justified, since the highest values arise for a short time (during external short circuits) and cannot be dangerous for the insulation. In fact, the secondary circuits of the protection are usually tested with a megohmmeter at 2.5 kV. Therefore it is advisable to limit only the voltage amplitude at \(I_1=I_{n2}\) to the value \(E_{2m}<250\cdot\sqrt{2}=350\) V ampl., for reasons of safety in servicing the circuit. In this case \(E_{2m}\) may reach values of 1500—1800 V ampl. during supercurrents.

The protection was in continuous operation for one year. Since during this time the insulation deteriorated neither in the protection circuits nor in the current transformers, it may be considered that the current transformers are sufficiently stable with respect to heating by losses

into steel in the no-load mode, the protection circuits easily withstand the increased voltages.

Conclusions

1. The no-load operating mode of the current transformers in the circuit under study is not an obstacle to its use on generators of any capacity.

2. The consumption of copper and steel for manufacturing current transformers for such a circuit is 5–10 times less than for class “D” transformers. The transformer cores need not be annealed.

3. The greatest importance is attached to the unbalance current caused by the periodic primary current.

4. If the cores of the current transformers of a differential protection with no-load e.m.f. balancing are made from an equal number of strips of steel of the same grade, it has the following characteristics: \(I_{1\mathrm{cp}} \leqslant 0.3 I_{н2}\), \(\hat{K}_4 \geqslant 7\).

5. In the absence of additional resistances in the relay circuit, the protection under study will not have tripping delays in the event of a short circuit within the zone, even during isolated operation of the protected generator, if \(k_1 \geqslant 5\text{–}6\).

6. In the protection under study it is advisable to use a relay with predominantly inductive resistance (ET-521 0.6 or ET-521/2 with series connection of the coils).

Engineers V. S. Anufriev, M. G. Baryshnikov, B. A. Borkovsky, L. R. Vereshchagin, Ya. B. Vishnevsky, N. A. Dulzon, Yu. Ya. Kirichenko, V. S. Kolesnikov, and Yu. A. Smirnov took an active part in the preparation and conduct of the production tests; the author expresses deep gratitude to them.

References

1. Baginsky L. V. On the use of current transformers with a small volume of steel for differential protection of generators, Proceedings of the Tomsk Order of the Red Banner of Labor Polytechnic Institute named after S. M. Kirov, vol. 94, 1958.

2. Baginsky L. V. On the prolonged operation of current transformers with an open secondary winding, Proceedings of the Tomsk Order of the Red Banner of Labor Polytechnic Institute named after S. M. Kirov, vol. 94, 1958.

3. Zaimovsky A. S. and Chudnovskaya L. A. Magnetic Materials, GEI, 1957.