On Ground-Fault Protection in Mine Section Networks

R. F. Trofimov

The question of protection against ground faults has not, up to the present time, received a proper solution. The proposed methods of protection against ground faults in mine section networks are not technically perfect.

The need to apply ground-fault protection is caused by the fact that, with the grounding devices existing in mines, having a resistance of 2 ohms (practically greater), the permissible touch voltage of 40 volts is ensured at a ground-fault current

\[ J=\frac{U}{r}=\frac{40}{2}=20a. \]

At fault currents greater than \(20a\), safety from electric shock is not ensured. It must be borne in mind that double faults in a low-voltage mine network can produce currents of several thousand amperes. The passage of such currents through the grounding network is dangerous, since it can create a touch voltage exceeding the permissible value. It is known that emergency currents passing through a person for \(0.1\)—\(0.2\) sec. can cause severe shock. To ensure the permissible touch voltage of \(40v\), it would be necessary to reduce the resistance of the protective grounding by increasing the amount of metal, which is economically disadvantageous and technically irrational.

Insulation-condition monitoring schemes increase the danger of electric shock when high-power apparatus is used. Moreover, such signaling is insufficient, since the light signal indicating the occurrence of a fault may either not be noticed or be noticed only with a considerable delay, which from the operating point of view is considered inadmissible.

The circumstances indicated require the use of protection that responds to single-phase ground faults.

Single-Phase Ground Faults in a Network with an Isolated Neutral Point

During normal operation of networks, charging currents and currents consumed by the power and lighting loads occur; these are superimposed on one another. In ground faults there also appear fault currents—zero-sequence currents—which are expressed by the following equations:

currents up to the fault location

\[ J_0=U\cdot j\cdot \omega\cdot C_x, \]

beyond the fault point

\[ J_0''=-U\cdot j\omega \cdot C\cdot (L-x) \]

and the ground-fault currents

\[ J_3=3\cdot U\cdot j\omega \cdot C\cdot L, \]

where \(U\) is the phase voltage,
\(C\) is the capacitance of a phase conductor with respect to ground,
\(L\) is the length of the network,
\(x\) is the distance to the fault point.

The physical picture of this condition is as follows. Ground faults are accompanied by a reduction of the voltage of the defective conductor with respect to ground to zero (a solid fault) and by an increase in the voltage of the healthy phases with respect to ground by a factor of \(\sqrt{3}\). This condition is equivalent to applying, at the fault point, to all phases a negative phase voltage of the damaged conductor, which causes a ground-fault current. The ground-fault current passes through the point of connection of the conductor with the ground, through the capacitances of the conductors with respect to ground, through the intact conductors and the transformer windings, and then returns along the conductor to the fault point. In this case, in the intact conductors

Fig. 1

Fig. 2

there is an increase in the normal current, while in the phase that has become connected to ground, the currents of the healthy phases add together, thereby creating a single-phase overload of the transformer.

The current of the damaged phase, unlike that of the healthy phases, does not decrease starting from the current source but, on the contrary, increases up to the point of the ground fault. The latter will be clear if one refers to the diagram shown in Fig. 1.

In mines where the networks have a radial configuration, currents from other parts of the network of the given section will also flow to the ground-fault point. The complete picture of the distribution of the total fault current for this case is given in Fig. 2.

Ground-fault currents obtained experimentally

In order to determine the actual ground-fault currents in systems with an isolated neutral point, experimental measurements were carried out. The current measurements were made according to the circuit shown in Fig. 3. We shall present data characterizing the networks of a mine with a steep dip (mine No. 1) and a mine with a gentle dip (mine No. 2), Table 1 (only for 2 mines).

At other mines, the ground-fault currents had values close to the lower and middle limits. The data presented in Table 3 prove that ground-fault currents are very small and are not capable of actuating ground-fault protections, except for complex ones possessing high sensitivity; but these currents are sufficient to produce a considerable thermal effect.

Fig. 3

Table 1

Table 2

Table 3

Secondary currents in A for the number of winding turns

| Primary-winding current | \multicolumn{5}{c}{Number of primary-winding turns: 10} | \multicolumn{5}{c}{Number of primary-winding turns: 20} | \multicolumn{5}{c}{Number of primary-winding turns: 30} | \multicolumn{5}{c}{Number of primary-winding turns: 40} |
|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|---:|
| Secondary-winding turns | 10 | 20 | 30 | 40 | 50 | 10 | 20 | 30 | 40 | 50 | 10 | 20 | 30 | 40 | 50 | 10 | 20 | 30 | 40 | 50 |
| 2 | — | — | — | — | — | — | 0.50 | 0.50 | 0.50 | 0.50 | 0.65 | 1.0 | 1.0 | 0.85 | 0.75 | 1.0 | 1.55 | 1.5 | 1.3 | 1.1 |
| 3 | 0.25 | 0.25 | 0.3 | — | — | 0.75 | 1.05 | 1.05 | 1.0 | 0.75 | 1.1 | 1.65 | 1.65 | 1.55 | 1.3 | 1.35 | 2.25 | 2.30 | 2.1 | 1.8 |
| 4 | 0.45 | 0.55 | 0.55 | 0.50 | 0.2 | 1.1 | 1.6 | 1.6 | 1.25 | 1.15 | 1.35 | 2.2 | 2.38 | 2.13 | 1.85 | 1.60 | 2.80 | 3.1 | 2.85 | 2.5 |
| 5 | 0.6 | 0.8 | 0.65 | 0.6 | 0.55 | 1.3 | 1.95 | 1.95 | 1.70 | 1.55 | 1.58 | 2.6 | 2.85 | 2.7 | 2.35 | 1.80 | 3.2 | 3.75 | 3.5 | 3.2 |

Single-phase ground faults in a network with a grounded neutral point

For systems with the neutral point of the transformer grounded through a resistance, the ground-fault current may be determined by the modified Oberdorfer formula

\[ J_z = U \frac{1 - 3 \cdot R_n \cdot j \cdot \omega \cdot C \cdot L}{R_n - r_u \left(1 - 3 \cdot R_n \cdot j \cdot \omega \cdot C \cdot L\right)}, \]

i.e., the ground-fault current depends on the magnitude of the resistance through which the neutral point of the transformer is grounded \((R_n)\), and on the resistance of the arc and of the network insulation \((r_u)\).

The complete picture of the current distribution is shown in Fig. 4. It is evident from it that grounding the neutral point of the transformer through the resistance \(R_n\) creates a large asymmetry of the currents and increases the ground-fault current and the current in the damaged conductor by the amount

\[ I_n = \frac{U}{R_n}. \]

Fig. 4

Existing protection methods

Protection against ground faults can be implemented on the basis of using: 1) zero-sequence currents, 2) zero-sequence voltages, 3) zero-sequence power, 4) replacement of the protective grounding by a protective relay (the circuit of engineer S. V. Sokolov and RWE), and 5) checking the insulation resistance before the closing of the starter contacts (PMV-type starters).

The first method has found application in systems both with an isolated neutral and with a grounded neutral. In systems with an isolated neutral, if the network consists of one or two radial feeders, selective protection by current relays in the event of ground faults cannot be achieved. Moreover, the implementation of protection operating on capacitive current is difficult because of the small capacitive currents. Grounding the transformer neutral makes it possible to provide selective protection against ground faults.

The so-called “shielded” cables used in England, with grounded concentric copper sheaths surrounding each working conductor separately, or all conductors together, reduce the danger of electric shock.

In the event of mechanical damage to a cable, the damaged working conductors are first connected to the grounded braid, causing a short circuit of one or several phases to ground and, consequently, operation of the maximum relay or leakage relay.

The use of zero-sequence voltages does not ensure selectivity. The use of zero-sequence power complicates the protection circuits. The RWE circuits and those of Engineer S. V. Sokolov require complex and sensitive protection, which does not ensure safety when faults occur before they are disconnected.

PMV-type starters have ground-fault protection, but this protection operates only at the moment the starter is switched on (when the “run” button is pressed). In all other cases it does not operate; moreover, it has low sensitivity, since the pickup current of the leakage relay RU is 1.75 a. This current, passing through the fault point after the armature has been pulled in, decreases to 0.9 a and is interrupted by pressing the “stop” button (BR1) of the local control built into the starter housing, a considerable time after the occurrence of the ground fault.

Proposed Ground-Fault Protection Circuits

Mine ground-fault protection must meet the following requirements: 1) suitability for systems with neutral grounding through resistance; 2) safety with respect to fires and explosions; 3) continuity of

Fig. 5

operation; 4) selectivity; 5) simplicity of design and sufficient mechanical strength.

Protection meeting these requirements can be provided by the circuit shown in Fig. 5. According to this circuit, ground-fault protection is implemented on the principle of using zero-sequence currents, with the use of a balancing (differential) transformer as a filter.

current (1) placed in the starter. Considering that zero-sequence currents in systems with an isolated neutral are insufficient for tripping, they must be artificially increased to the current of (smaller) arc formation by applying neutral grounding through a resistance. It is known that interruption of the leakage-current circuit between metallic conductors and the soil of the working can cause a gas explosion only when there is an arc, for the formation of which a current of 3–5 A is required. Ground-fault protection in PMV-type starters permits a ground-fault current of 1.75 A, which after pickup of the core decreases to 0.9 A and is not interrupted for a long time.

PMV-type starters may be used as starters, but there is a need to design a new starter. In PMV-type starters the control circuit should be changed and simplified. Of the existing three relays RU, RB1, and RB2, it is necessary to leave two relays: one for controlling the contacts of the starter and the second as an interlocking relay for ground faults. Ground-fault protection should consist of a zero-sequence current filter and a grounding relay RZ connected into the circuit of this filter. The complete diagram of a starter with ground-fault protection is shown in Fig. 6.

In the absence of a ground fault, as indicated above, the starter is controlled by one interlocking relay RB2. Starting is effected by pressing the “run” button of local or remote control. In this case relay RB2

Fig. 6

operates and closes its contacts in the circuit of the starter closing coil. This is followed by closure of the contacts of the power circuit. For tripping there are the normally closed “stop” buttons and BR1.

In the event of a ground fault, the grounding relay RZ operates and closes its RZ contacts in the circuit of relay RB1. The latter, when it operates, self-locks and opens the circuit of relay RB2. This gives the command to open the contacts of the starter power circuit. Relay RB1 is released by pressing the BR1 button (local control). Reliable operation of the circuit is obtained because, in relay RB1, the upper contacts close before the lower contacts of the RB2 relay circuit open. This is achieved by raising (bending upward) the lower fixed contacts of relay RB1.

This entire circuit can be simplified if relay RB1 is omitted and the circuit of relay RB2 is opened by opening the lower contacts of the grounding relay RZ. But this makes the protection coarser, because it becomes necessary to increase the pressure of the moving contacts on the lower fixed contacts of relay RZ. The circuit for this case is shown in Fig. 7.

When the grounding relay operates, it must self-lock. For this purpose, either a mechanical or an electrical principle may be used.

The first principle requires complex devices for interlocking. The simplest is the second method. In this case the ground-fault relay must be provided with an additional coil connected to the secondary winding of the voltage transformer through the contacts of the ground-fault relay and the additional contacts BRZ of the “stop” pushbutton of local control. Pressing this pushbutton will effect the interlocking.

Fig. 7

Filters of zero-sequence currents

A zero-sequence current filter can be implemented by means of current transformers connected in Holmgreen’s circuit. The presence of unbalance currents and leakage currents lowers the accuracy of Holmgreen’s circuit, not to mention that it is complex, since it requires 3 current transformers.

Simpler in construction and more accurate in operation is a balanced (differential) current transformer having one core on which the primary and secondary windings are arranged.

Under normal operation the sum of the magnetizing ampere-turns is

\[ i_{AH} w_1 + i_{BH} w_1 + i_{CH} w_1 = 0, \]

where \(w_1\) is the number of turns of the primary winding of the filter, and \(i_{AH}\), \(i_{BH}\), \(i_{CH}\) are the magnetizing currents of phases \(A\), \(B\), and \(C\).

Owing to the absence of unbalance currents, the current in the relay is

\[ i_p = \frac{w_1}{w_2} (i_{AH} + i_{BH} + i_{CH}) = 0. \]

During a ground fault, the effect of zero-sequence currents appears, and then

\[ i_p = \frac{w_1}{w_2} [-3 i_0 + (i_{AH} + i_{BH} + i_{CH})]. \]

Taking into account the no-load currents, which cannot be neglected, the current in the relay is

\[ i_p = \frac{w_1}{w_2} [-3 i_0 + (i_{Ax} + i_{Bx} + i_{Cx})], \]

i.e., it is equal to the sum of the zero-sequence currents and the leakage currents going to magnetization and covering iron losses.

The balancing current transformer must operate at small ampere-turns. In order to determine the dependence of the secondary currents on the primary currents, experimental studies were carried out with various types of current transformers. For the core of a coil-type current transformer having a cross section of \(3 \times 4.2 = 12.6\ \text{cm}^2\), an average magnetic-circuit length of \(25.4\ \text{cm}\), numbers of turns 10, 15, 20, 25, 30, and a transformation ratio equal to 1, the data are given in Table 2.

The unbalance currents were determined during motor starting and load changes for various types of mine motors. In all cases these currents were equal to zero.

The experimental specimen subjected to further tests was designed for a rated primary current of \(75\ a\), had a core weight of \(0.718\ \text{kg}\), and a total weight (with conductors) of \(2.2\ \text{kg}\) with very small overall dimensions.

The primary and secondary windings were made with taps with numbers of turns 10, 20, 30, 40, and 50. The secondary currents were determined for different numbers of winding turns and changes in the primary current. The results of the investigations are given in Table 3.

These results prove the possibility of using balancing transformers for large operating primary currents as zero-sequence current filters, with their small overall dimensions and low weight. The transformation ratio should be taken equal to unity, and the number of turns \(w_1 = w_2 = 30\). Such a filter may be located on the reverse side of a PMV-type starter panel, above the disconnectors. The data given in Tables 2 and 3 prove the impossibility of using Ferranti-type transformers in ground-fault protection schemes under mine conditions.

Relay and its parameters

As indicated above, the relay must operate instantaneously in the event of single-phase ground faults. It must be connected to the secondary winding of the zero-sequence current filter, which has a small number of ampere-turns. Taking these circumstances and the specific mine conditions into account, the following requirements must be imposed on the grounding relay:

1) minimum power consumption and minimum pickup current;
2) a sufficient return ratio;
3) continuity of operation;
4) minimum operating time;
5) invariability of the characteristic under all possible operating conditions;
6) simplicity of construction and operation;
7) high mechanical strength;
8) ensuring normal operation at starter inclinations of up to \(15^\circ\).

These conditions can be satisfied by relays of an electromagnetic system, both of the clapper type and of the armored type. A clapper-type electromagnetic relay was adopted as the experimental specimen.

To determine the pickup currents, the winding was provided with taps at 50 and 75 turns. In general, the relay winding had 100 turns.

The values of the relay pickup currents as a function of the number of turns of the relay winding are given in Table 4.

Table 4

What should be regarded as high sensitivity? According to calculations, the operating power was less than 0.1 va.

Sensitivity of ground-fault protection

The currents of the primary winding (ground fault) of the above-mentioned filter at which the relay operated are given in Table 5.

Table 5

This proves that highly sensitive ground-fault protection using a low-turn zero-sequence current filter, in combination with an electromagnetic-system relay, can be implemented. This also refutes English practice, which demonstrates the impossibility of ground-fault protection operation in mines at a leakage current of up to 7.5% of the rated current. The protection considered, as is evident from the experimental data, operates at a current of the rated value

\[ \frac{1.0}{75}\cdot 100 = 1.33\%;\quad (w_1 = w_2 = 30), \]

which is 5.63 times more sensitive than English systems and 1.75 times more sensitive than the RУ leakage relay of PMV-type starters.

The sensitivity of the protection in terms of the number of protected turns of the stator winding is usually taken to be not less than 70% (dead zone 30%).

If the ground-fault current is assumed to be 3 a, which is less than the arc-formation current, then the sensitivity for motors whose stator winding is connected in star will be equal to:

\[ \frac{3 - 1}{3}\cdot 100 = 66.7\%, \]

which is quite permissible.

For motors whose stator winding is connected in delta, 100-percent turn protection will be obtained.

The sensitivity of the protection can be increased by changing the parameters of the filter and relay, and also by manufacturing them more accurately.

If $I_0$ denotes the minimum current at which the ground-fault protection can operate (1 a), and $I_{0\max}$ the maximum permissible ground-fault current (3 a), then the earth-fault completeness coefficient will be equal to:

\[ a = \frac{I_0}{I_{0\max}} = \frac{1}{3} = 0.333, \]

which is permissible, since it lies within the limits recommended by the standards, 0.2–0.5.

According to the oscillograms, the time from the beginning of the ground fault to disconnection of the faulty section of the network is up to 3–5 periods.

Selectivity of Protection

Mine section networks are radial, with one-sided supply. In networks with an isolated neutral, the magnitude of the current depends on the length of the network and, under certain conditions, may decrease so much that the protection ceases to operate, even if the earth-fault current normally exceeds the relay pickup current. To eliminate this, as noted above, it is necessary to ground the transformer neutral through an active resistance, thereby artificially increasing the earth-fault current. Selectivity can be achieved by applying time grading. The operating time delay of the protection for the terminal section (distribution point) may be taken as zero, and at section substations as 0.5 sec. For this purpose, feeder circuit breakers must be equipped with earth-fault protection and time relays, preferably of the thermal type manufactured by KhETZ, as ensuring simplicity of design.

Conclusion

The question of methods of protection against single-phase earth faults is exceptionally urgent, since its correct solution will increase the safety of work in mines, will make it possible to move to a higher voltage, and will significantly simplify power-supply schemes and protective grounding devices by increasing their resistances.

In systems with neutral grounding through a resistance, among all possible methods of protection against single-phase earth faults, the most technically advanced method should be considered the one proposed in the present work, since it ensures simplicity of the scheme, selectivity, continuity of operation, and high sensitivity.

Since this method requires an artificial increase in earth-fault currents to the value necessary for relay operation, concerns may arise with respect to fires and explosions of mine gas. However, these currents will be smaller than arc-formation currents (protection against a spark is impossible and is not needed), and they will last for 3–5 periods, not 7–8 days, as occurs with existing methods of protection against earth faults. In addition, the proposed protection will prevent the transition into double earth faults—two- and three-phase short circuits—and thus it will replace maximum-current protection in starters.