TESTING OF A HIGH-POWER INDUCTION MOTOR WITH THERMOSETTING INSULATION

A. N. Zuev, L. Ya. Klenshtein, V. G. Ryazanov, V. G. Syakov, M. Z. Tuller

(Presented by the scientific seminar of the Departments of Electrical Machines
and General Electrical Engineering)

At the Novosibirsk plant “Sibelektrotyazhmash,” the thermosetting insulation Monolit-II has been introduced for the coil windings of large electrical machines, developed by VEI jointly with SibNIEtI. The coils, insulated with dry glass-mica tape, are placed in the core, which is then impregnated under vacuum with an epoxy compound and baked. It is known that thermosetting insulation differs from compounded mica-tape insulation in having higher electrical, mechanical, and thermophysical characteristics.

Comparative tests were carried out on a production squirrel-cage induction motor of type AZ-500/6000 (500 kW, 6 kV, 57 A, 3000 rpm) and on a motor made on its basis with Monolit-II insulation. The motors differ only in the type of insulation, and its thickness is the same. Naturally, the motors have practically identical electromagnetic characteristics and air consumption. Certain differences in the no-load current values, iron losses, etc., do not exceed the usual deviations observed during acceptance tests of production motors.

Electrical Tests of the Insulation

The motors withstood the tests prescribed by GOST 183-66. A comparison of the dielectric loss values in machines heated to 45°C is given in Table 1, from which it can be seen that, for Monolit-II insulation, they are 1.5 times lower in the range of operating voltages.

A direct comparison of the electrical strength of the types of insulation cannot be made, since breakdown of the windings was not performed in order to preserve the motors; moreover, on the basis of tests of only one pair of motors, such a comparison would be unjustified.

Table 1

Tests on specimens made according to the technology adopted for Monolit-II show that its electric strength is 1.8–2.0 times higher.

Thermal Tests of the Motors

The motors are equipped with thermal monitoring that makes it possible to determine the heating of the active steel and of the air flows. Because of technological and design difficulties, thermocouples were not installed on the active copper of the stator.

The results of measurements in several characteristic modes are given in Table 2.

Table 2

Comparison of the data for modes 3 and 4 shows that Monolit-II provides a considerable reduction in the overheating of the stator winding, not only because of its higher coefficient of thermal conductivity, but also because, when the core with the winding is impregnated, the compound fills the gaps between the coils and the slot walls and impregnates the spacers in the slots and the banding ties in the end-winding parts.

Analysis of the tests (see the Appendix) shows that the equivalent thermal conductance between the copper and the heat-removing medium of the winding (the core steel or the air in the basket of the end-winding parts) increased in motor M by approximately a factor of 1.5, both in the slot and in the end-winding part. The average overheating of the stator winding of motor M relative to the incoming air (\(t = 40^\circ\text{C}\)) in the rated mode is lower by 16.5°C (46.5°C versus 63.0°C).

Processing of the results also shows that, at the permissible average winding overheating of 80°C, motor K is capable of carrying a load of 560 kW, while motor M can carry 670 kW, i.e., 20% higher.

The results of the processing presented should for the time being be regarded as qualitative, since only the testing of a series of machines makes it possible to exclude the possible influence of random factors (for example, the quality of fabrication of the end-region bracing, the density with which the insulation is applied) and to obtain quantitative relationships. However, a significant effect should nevertheless be expected from the use of Monolit-II, especially since its high electric strength makes it possible to reduce the thickness of the ground insulation by approximately 30%.

Insulation reliability tests

It is generally recognized that insulation is subjected to the maximum thermal and mechanical effects in the starting mode at its maximum permissible duration. On motor \(M\), 1000 starts were performed at rated voltage, with a flywheel mass at the maximum permissible under the Specifications; moreover, the operating mode between starts was selected so that the average overheating of the stator winding before the next start was close to the overheating in rated operation (to simulate the so-called “hot-state start”).

Then 50 regimes were carried out to determine the ability of the insulation to withstand the maximum theoretically possible fluctuations of the winding temperature. The motor, cooled on average to \(-20^\circ\text{C}\), was started with a flywheel mass 30% greater than that permitted by the Specifications, then disconnected from the network and supplied with direct current \(1.2\cdot I_{\text{nom}}\) (and) with the average winding temperature brought up to \(+150^\circ\text{C}\).

After the regimes described, the motor withstood insulation breakdown tests at \(U = 15\ \text{kV}\) (GOST 183–66 requires \(13\ \text{kV}\)). During the tests, external inspections of the condition of the winding and measurements of dielectric losses were carried out. The condition of the winding did not differ from the initial one; the value of \(\operatorname{tg}\delta\) underwent no noticeable changes.

Studies of the motor with Monolit-II insulation confirm its high electrical, mechanical, and thermophysical properties and, on the whole, show the high effectiveness of its use in high-power high-voltage electrical machines.

Appendix

The processing of the thermal tests was carried out in the following scope and sequence.

1. Using the method set forth in [1], the regimes with \(I > 0.7 I_{\text{n}}\) (in which measurement errors have less effect) of motor K were processed. The average thermal conductivity of the end regions \(\lambda_1\) and the values of the heat flows from the slot portion of the winding into the end region \(q\) were determined in each regime.

2. From (8) of [1], the value of \(q\) was determined in the no-load regime of motor K.

3. In all regimes, the specific heat flow through the insulation in the slot portion was determined,

\[ p_i = p_2 - \frac{q}{l_2}, \]

and the dependences (Fig. 1) of the average overheating of the core \(\vartheta_{c2}\) on the value \(p_i\) were plotted for the load regimes with \(U = 6\ \text{kV}\) (straight line \(a\)) and for the direct-current supply regime (straight line \(b\)).

1. If, as a first approximation, it is assumed that the cores of both motors are identical in thermal respects in everything not connected with the stator winding (equality of geometry, of the function \(I_2(I_1)\), of the air flows, etc.), then for motor \(M\) the function \(\vartheta_{c2}(p_i)\) should be parallel to straight line \(a\) and located lower by the difference in the values of \(\vartheta_{c2}\) at no load of the motors (straight line \(v\)), while the function expressed by straight line \(b\) should also be valid for motor \(M\). From straight lines \(b\) and \(v\) the values of \(p_i\) (and consequently also \(q\)) are determined in all operating conditions of motor \(M\).

Fig. 1. 1 — no load, \(U=1.0\); 2 — partial load, \(U=1.0\); \(I=0.8\); 3, 4 — full load, \(U=1.0\); \(I=1.0\); 5 — ventilation condition, \(U=0\); \(I=0\); 6 — winding supplied with direct current, \(I=1.0\)

1. For each operating condition, the system of equations (8) and (10) from [1] is solved with respect to the parameters \(\lambda_1\) and \(\lambda_2\) for known \(q\).

The mean values of \(\lambda_1\) and \(\lambda_2\) for motor \(M\) are then determined.

For motor \(K\), the value \(\lambda_2=6.55\ \mathrm{W/(m^2\,^\circ C)}\) was adopted, which corresponds to \(\lambda_1=0.16\ \mathrm{W/(m\,^\circ C)}\). As a result, the following mean values were obtained:

\[ \lambda_2=6.55 \quad \text{and} \quad \lambda_1=3.89 \]
—for motor \(K\),

\[ \lambda_2=10.0 \quad \text{and} \quad \lambda_1=5.76 \]
—for motor \(M\).

The relative deviations of the sought values \(\lambda\) from the means do not exceed \(\pm 6\%\) in all operating conditions (with the exception of the direct-current condition of motor \(M\), where, possibly, the role of measurement errors increases owing to the relatively low \(\vartheta_{c2}\) and \(\vartheta_{\mathrm{av.\,w.}}\)). The values of \(\lambda\) for motor \(M\) are approximately 1.5 times higher.

Fig. 2. Stator-winding temperature rise along the length of the coils in the rated operating mode. 1—motor K, 2—motor M

Figure 2 shows the distribution curves of the stator-winding temperature rises of both motors in the rated operating mode, obtained by calculation using the values (*) and the functions \(\vartheta_{c2}(p_i)\) in Fig. 1.

REFERENCES

1. V. G. Ryazanov. Estimation of the temperature distribution in a winding under nonuniform cooling along its length (see in the present collection).