Ground-directional solution to improve selectivity in underground mining power systems protection

Computers and Electrical Engineering 80 (2019) 106491

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Ground-directional solution to improve selectivity in underground mining power systems protectionR,RR German Gutierrez a, David Celeita a,b,∗, Gustavo Ramos a a b

Universidad de los Andes, Department of Electrical and Electronic Engineering, Bogota, Colombia CentraleSup - lec, GeePs - Gnie lectrique et lectronique de Paris, Gif-sur-Yvette, France

a r t i c l e

i n f o

Article history: Received 26 March 2019 Revised 3 October 2019 Accepted 4 October 2019

Keywords: Mining power systems Grounding Protection selectivity Directional overcurrent protection Real-time simulation Hardware-in-the-loop

a b s t r a c t The trend of expanding to larger underground mining systems has motivated increases in the voltage levels used for distribution and utilisation. Such increases have caused the industry to face complexities and challenges that are not common in lower voltage systems. For high voltage levels, the mining regulations require the use of shielded cables, with a resistive grounding method for the neutral. The occurrence of a phase-to-ground fault in a power system with this arrangement causes capacitive current flows from different locations to the point of failure, which reduces the selectivity of the ground overcurrent relays. The solution proposed here for this loss of selectivity is based on the ground-directional over-current algorithm. The methodology was detailed, validated, tested, and assessed using a software test model and subsequently in a real-time hardware-in-the-loop implementation in a playback testbed. The results showed an improvement in the protection performance in the context of different failure scenarios. © 2019 Published by Elsevier Ltd.

1. Introduction Underground mining environments are considered high-risk areas because of the proven presence of explosive gases and dust. Therefore, the regulations for the use of machines and equipment are stricter than those in other industries. The Code of Federal Regulations (CFR) in Title 30 [1], corresponding to those for the exploitation of mineral resources in the United States, provides a set of rules and regulations as an international guide for the safe exploitation of mineral resources in underground mining. The power requirements of the high-capacity machines utilised have significantly increased, which has led to ever-increasing voltage distribution and utilisation levels [2]. operation of these high-capacity systems is inadequate with low voltages. Therefore, it has been necessary to use high voltages (up to 15 kV), with usage levels greater than 1 kV [3]. In [4] a historical analysis of the evolution of the distribution voltage requirements in underground mining was developed. To ensure that high-voltage systems operate at acceptable safety levels in classified zones, the Federal Code of Regulations specifies the need to limit the ground fault current to restrict the dissipated energy, which may produce fires. Traditionally, this requirement has been fulfilled by inserting a resistance in the system neutral, which limits the current flowing through R This project was partially funded by the Administrative Department of Science, Technology and Innovation of Colombia COLCIENCIAS under grant 733 ’Formaci’on de capital humano de alto nivel para el Departamento de Boyac’a2015’. The authors also thank to PTI SA, Potencia y Tecnolog a Incorporadas, Colombia for the tests with the power amplifier. RR This paper is for regular issues of CAEE. Reviews processed and recommended for publication to the Editor-in-Chief by Associate Editor Dr. Shadi A. Aljawarneh. ∗ Corresponding author. E-mail addresses: [email protected], [email protected] (D. Celeita).

https://doi.org/10.1016/j.compeleceng.2019.106491 0045-7906/© 2019 Published by Elsevier Ltd.

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Fig. 1. Power system with resistive grounding and phasor diagram. Based on (a) [7] and (b) [8].

it to 25 A, and adjusting the activation level of the power centre relay of 40% of this current [5]. Because modern power systems can be very extensive, with more than 15 miles (24.1 km) of shielded high-voltage distribution cable, the distributed capacitance of these systems, combined with the limit on the current flowing through the grounding resistance, can create situations where the selectivity of the relays is lost. A phase-to-ground fault in one branch of the circuit causes a capacitive current flow from the other branches to the point of failure, causing the ground-overcurrent relay to trip erroneously [6]. The proposed methodology aims to improve the selectivity in large mining power systems, through the strategic inclusion of a directional protection scheme. Such an implementation could be performed with the existing technology in the power centres of a mine. The residual voltage polarisation magnitude would be determined for ease of implementation. The proposed solution is consistent with safety standards and existing technology. Sections 2 and 3 present the theoretical framework of the current flows,as well as the characteristics of a specific underground mining power system. In Sections 4, an underground mining case study is modelled and validated, which is used in Sections 5 and 6 to study the loss of selectivity. Section 7 shows how the directional relay algorithm is applied to the model, and Section 8 shows how this was validated in hardware. Finally, the conclusions and future work are presented. 2. Current flow in system with resistive grounding method Conventionally in solidly grounded systems the current flowing through the neutral is considered equal to the fault current. However, in systems with a resistive grounding method, the concept extends with the inclusion of the capacitive current. As consequence, it contributes a component to the fault current due to the coupling of the systems to ground through the distributed capacitance of the system [9]. Fig. 1( a) is a reproduction of a figure from [7], where the author represented the capacitive and resistive current flows from the source to the ground. The flow goes through the ground until finding the point of failure, where the currents produce the sum of these flows. The ground fault current (IG) in the phasor representation is given by the sum of the capacitive and resistive currents. The direction of the resistive current is the opposite to that of the voltage of the fault line. Fig. 1(b) is a reproduction of a figure from [8], where the author follows an industrial convention with respect to the flow of the capacitive current, which proposes that the direction of this current is the inverse to the load current of the system. One consideration is that the direction of the fault current (IGF) is the opposite to that of the capacitive current. Thus, the vector sum is found using -IC. 3. Characteristic of power system in underground mining 3.1. Distributed capacitance The cables used in an underground mining power distribution system are subject to extreme operating conditions. In addition to the electrical conditions, they must be designed to withstand bending, abrasion, and crushing factors. Therefore, the manufacturing characteristics of these cables are very robust to withstand severe operating environments [2]. A mining drag wire consists of power and ground conductors, a grounding verification conductor, shield, insulation, tape, and a jacket,

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Fig. 2. Cross section of cable type SHD-GC. Based on [10].

Fig. 2. The type and sturdiness of each cable component depend on the level of voltage to which it is going to be subjected. Generally, low-, medium-, and high-voltage cables are available. SHD-GC high-tension cables use a 0.4 mm wire coating that is fabricated from a semiconductor compound surrounding each phase, along with ethylene-propylene insulation over the 2.8 mm wire coating. A semiconductor ribbon of nylon-butyl is applied helically and superimposed around the insulation of each phase and the outer and inner 5.6 mm jackets [2]. The physical layout of the drag cables results in the presence of a line-to- ground capacitance, distributed along the entire length of the cable. This capacitance is caused by the presence of parallel conductors separated by a dielectric, as shown in Fig. 2. The capacitance can be obtained from the manufacturer’s data, or determined from Eq. (1) [11].

C=



7.35ε



log10 1 +

2t d



pf ft



(1)

where ε is the dielectric constant of the insulation, t is the thickness of the conductor’s insulation and d is the diameter under thex insulation. 3.2. Neutral grounding method Limiting the current flowing through the neutral to 25 A, as is done in some systems, may be inadequate when considering that mines have different specifications in terms of the length and characteristics of the power system. When recommending a design, it is necessary to approximate the capacitive reactance values using the value of the grounding resistance. A high-resistance grounding system is defined as a system in which R0 <= XC , where R0 is the zero sequence resistance per phase of the system and XC is the distributed capacitive reactance of the system. In practical mining energy systems, the zero sequence resistance is dominated by the grounding resistance. Therefore, RN <= 13 XC , where RN is the value of the grounding resistance. In cases where the definition of high resistance grounding is not met, i.e., RN > 13 XC more current flows through the distributed capacitance of the system than through the grounding resistance and the system starts to acquire the characteristics of a system without a ground connection [11]. The recommended value for the resistance must be selected in such a way that Eq. (2) is satisfied:

RN =

|XC0 | 3 (S )

(2)

where |XC0 | is the magnitude of the capacitive reactance distributed per phase of the system and S is a safety factor to allow the expansion of the shielded cables of the system. The selection of the security value is discussed in [12]. 3.3. Effects of distributed capacitance Systems with a resistive grounding method are prone to perceive unwanted characteristics in the power system as a result of inadequate design practices. Oversizing the grounding resistance of the neutral results in more current flowing

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Fig. 3. (a) Capacitive current flows and (b)graphical representation of detection using directional current algorithm.

through the distributed capacitance than through the grounding method, generating dangerous overvoltages that hinder the safe operation of the system. On the other hand, undersizing the grounding resistance results in high fault currents that can cause flies or sparks, which are dangerous for the safe operation of the system [13]. The combination present in a modern operating structure that includes shielded cables, high voltage, and long cable lengths causes problems with the selectivity of the ground fault relay. Even with an adequately designed grounding method, as shown in section VI, it is not possible to eliminate these selectivity problems because the distributed capacitance of the system causes currents to appear in the non-faulty branches that exceed the level of adjustment of the overcurrent relay from phase to ground, resulting in the false triggering of the relays. Fig. 3 (a) recreates an earth fault event in a three-branch mining system. The system shows a phase-to-ground fault in branch 2, which results in a capacitive current flow through the non-faulted branches to the fault point. The magnitude of this current is determined by the capacitance of the branch, which is directly related to the length of the cable. When the value of the capacitive current measured by the current transformer (CT) exceeds the adjustment level of the relay, it decides to disconnect the branch, causing an erroneous trip. 3.4. Directional protection to ground The basic detection principle for the ground directional current determines the phase difference between the residual current and a reference quantity, which may be the residual voltage or the current flowing through the grounding method. The protection measures the phase difference between the current and polarisation magnitude. The characteristic angle of the directional protection must be determined in such a way that the residual current in the selected detection direction causes a current located in the tripping zone. A graphical representation of the directional current detection algorithm is shown in Fig. 3(b). The axis of symmetry is set to 0°, where the polarisation magnitude is located. Based on the characteristic angle selected, a trace is made of the axis on which a perpendicular is projected that determines the detection zone. The rotation of the detection zone will depend on the characteristic angle selected. 4. Case study A three-branch model was used to represent the power system of an underground mine. This system is a simplification of the real models. Nevertheless, it contains sufficient details of the main elements of a real system [3] to allow a precise evaluation of the behaviour before the failure of a phase to earth. The proposed model is shown in Fig. 4. It has a distribution voltage of 12.47 kV, and a utilisation level of 1040 V. Two wire gauges are used, 500 kcmil for the main feeder and 4/0 for the branches feeding the loads. Charges grouped by branch are used to represent the mining equipment. 4.1. Parameters The model of the cable, together with its capacitance and the design of the grounding method, are presented below. Cable sizing: The cables are represented as parameters grouped in a Pi configuration. The cable used is AmerCable reference TB2-604 type MP-GC [7]. The main feeder has a size of 500 kcmil, while 4/0 AWG is used for the other branches. The characteristics of these cables are listed in Table 1.

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Fig. 4. Case study based on [3]. Table 1 Cable 1 parameters. Size

500 kcmil 4/0 AWG

Phase conductor

Ground conductor

R/10 0 0ft

X_L/10 0 0 ft

Size

R/10 0 0ft

0.027 0.063

0.031 0.035

4/0 1

0.063 0.16

Table 2 Cable 2 parameters. Branch

Main 1 2 3

Phase conductor

Ground conductor

Length (ft)

R / 

L (mH/)

R / 

L (mH/)

1000 500 10,000 15,000

0.027 0.0315 0.63 0.945

0.0822 0.0464 0.9284 1.3926

0.0315 0.04 0.08 1.2

0.0822 0.0464 0.9284 1.3926

Table 3 Cable parameters for capacitance calculation. Size

ε

t (in)

d (in)

500 kcmil 4/0 AWG

3.2 3.2

0.175 0.175

0.81722 0.54576

By multiplying the length of each cable by the parameters listed in Table 1, the values for the simulation model listed in Table 2 were obtained. As seen in Fig. 2, there are two ground conductors, modelled as parallel equivalents, and the inductance of the ground conductor is assumed to be equal to that of the phase conductor. The distributed capacitance of the system is obtained from Eq. (1). The parameters for this calculation are listed in Table 3. Neutral resistance grounding: The capacitance of each branch can be determined using the length and type of cable, as represented in the Pi model. Capacitance values per arm of 0.0760 μF, 0.0273 μF, 0.5468 μF and 0.8202 μF are obtained for the main feeder and branches 1, 2, and 3, respectively. Finally, a total system capacitance of 2.94 μF is estimated. According to the methodology shown in section 2.2, for the proper design of the system, a total distributed capacitance of 2.94 μF is obtained. This is considered safety factor 1, assuming that the system analysed does not increase in length. In this way, we have a grounding resistance given by Eq. (2)), obtaining a value of 301 . This value guarantees a current flow through

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Fig. 5. ATPDraw model.

the resistance that is approximately equal to the capacitive current of the system in the event of a phase-to-ground fault condition. 4.2. Validation of model: Case study The model previously described was implemented with ATPDraw. Each of the elements was placed individually by phase, thus achieving greater flexibility in the analysis. Fig. 5 shows the representation of a branch of the circuit, which is composed of the impedashows a representation of a branch of the circuit, which is composed of the impedance of the load, the transformer, and the impedance of the line, divided into two sections, together with the ground cable. Sectioning the line makes it possible to reproduce faults at different distances, which allows a better analysis of the system to be achieved. The ATP data are generated at a time step = 1X10-6. After that, subsampling is applied to reproduce the COMTRADE file and amplify the electrical signals. Once the implemented model shown in Fig. 5 was validated in ATPDraw, the simulation results were compared with theoretical calculations of the fault current, grounding resistance current, and branch current under phase-to-ground fault conditions. A fault was made in phase A to ground at the end of branch 2. Estimates of the current flowing through the method of grounding (INGR ), fault current (IF ) and capacitive current in the third branch (3|I0(3) |) were determined using Eqs. (3)–(5) respectively:

V

|INGR | = √ LL

(3)

3RN

IF ≈

√

3·



VLL 1 RN

+

1

−1

(4)

jXCO 3

 √     3 · VLL   3I0(3 )  ∼ =  XCO(branch ) 

(5)

Therefore, the estimation of the neutral current follows Eq. (3): 12470 IF ≈

−1 = 33.84A √

1 + 1 3· 301 j902.15 3

In a similar way, the capacitive current of the third branch is computed, with a phase-to-ground fault in branch 2. This calculation is made with the total capacitive reactance of line 3:    √3·12470  ∼ 3IA0(3 )  =  = 13.35A 1617

Fig. 6 shows the waveforms of the validation currents of the model, which make it possible to validate the implementation of the model in the software. 5. Direction of current flows and phasor diagrams The parameters of the test circuit corresponding to the branch 3 without load (see Fig. 10) were used to study the behaviour of the current flows. The following values were applied by phase: EG = 12.47kV ; C = 1.64μF ; L = 1.64mH. The grounding could be  resistance    computed using Eq. (2):  1   wC 

  1  2·π ·60·1 .64μF 

RN = 3 = = 539.143 3 Applying Kirchhoff law for the failure node with Eq. (6), which represents the sum of the components of the current in the failure node, yields the following:

INGR = IF + IC

(6)

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IA0(3) INGR

Current (A)

40

7

X: 0.1479 Y: 47.74 X: 0.1667 Y: 33.57

IFault

X: 0.1874 Y: 18.75

20

0

-20

-40 0.1

0.12

0.14

0.16

0.18

0.2

Time (s) Fig. 6. Waveforms of fault current, NGR and residual.

Fig. 7. Case study for current flow analysis.

To validate the directions of the current flows in the system of Fig. 7, two methodologies are proposed. The first is an analysis using signal processing, and the second uses symmetric components. The idea is to obtain equal results with these two methods to infer the behaviours of these currents during fault events. 5.1. Signal processing method A DSP module was developed to measure the currents and voltages, including the time vector. Current meters are located in the directions of the arrows, which represent the expected directions of current flow. The signals are exported from ATPDraw to Matlab, where a fast Fourier transform (FFT) is applied, which makes it possible to quantify the fundamental frequency, magnitude, and phase of each of the currents and voltages. Fig. 8(a) shows the capacitive and resistive current signals for a fault in phase A of the circuit of Fig. 7. An analysis of the signals in time shows a phase jump in the capacitive currents, decreasing their phase angle in comparison with each other. A phasor representation of the system currents, along with the phase voltages, is shown in Fig. 8b. The voltages in the non-faulted phases (B and C) increase to line voltages with a fault angle of 60◦ . The capacitive currents in the non-faulted √ phases increase along with the voltages by a factor of 3 of the pre-fault value, with angles of 60◦ to each other and 90◦ in advance of their reference voltage. The phasor representation of Fig. 8(b) validates the observations made on the signals with the time vector. 5.2. Symmetric components method The fundamental principles developed by Fortescue in 1918 are applied to validate the behavior of the power system during ground fault conditions. The method known as symmetric components assumes that the entire circuit is asymmetric in such a way that it can be expressed in three symmetric systems, as shown in Fig. 9. Using symmetric components in the analysis makes it possible to theoretically validate this concept, by finding the magnitudes and phase angles of the phasors of the voltages and currents, which are then drawn in a phasor diagram.

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Fig. 8. (a) Case study: current waveforms including resistive grounding and (b) signal processing method: phasor diagram of case study with resistive grounding.

Fig. 9. Case study: sequence diagram.

The system of Fig. 9 a subscript of 1 for the positive sequence variables, 2 for the negative, and 0 for the zero sequence. Mutual coupling effects between phases are not considered. The capacitance and reactive impedance values are equal in each of the sequences. To find the sequence values, an analysis using the mesh current method is proposed (see Eq. (7)):

Ax = b ⇒ x = A−1 b

(7)



where vector x represents the mesh currents x = I1x I2x I0x Ifx impedances: ⎡ ⎤ Zs1 + Zc1 0 0 −Zc1 Zs2 + Zc2 0 −Zc2 ⎢ 0 ⎥ A=⎣ ⎦ 0 0 3Zg + Zs0 + Zc0 −Zc0 −Zc1 −Zc2 −Zc0 Zc1 + Zc2 + Zc0 + 3Z f

T

and matrix A is constructed from the mesh

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Fig. 10. (a) Case study with resistive grounding: sequence phasor diagram and (b) capacitive current vs. different cable sizes.



Finally, b represents the sources of the system for each loop b = Es 0 (10)) are then defined to find the capacitive sequence voltages and currents:













V1 = I1x − I f x Zc1 = Ic1 Zc1

0

0

T

. The following equations (Eqs. (8)–

(8)

V2 = I2x − I f x Zc2 = Ic2 Zc2

(9)

V0 = I0x − I f x Zc0 = Ic0 Zc0

(10)

The rotation operator is defined as α = − 12 + j



3 2.

A direct transfer matrix is defined with this operator, which allows a transformation between the sequence and phase dimension, both for voltages and currents, as shown in 11



Va Vb Vc



=

1 1 1

1

α α

2

1

α α2



V0 Ica V1 ; Icb V2 Icc



=

1 1 1

1

α α

2

1

α α2



Ic0 Ic1 Ic2

(11)

Fig. 10 (a) shows the phasor representation of the currents and voltages of the power system. The results correspond entirely to the method of signal processing, which shows a capacitive current in the reverse direction to the fault current. It also shows a phasor of the fault current in advance of the current flowing through the method of grounding (IR). The Kirchhoff current analysis shown in Eq. (6) is mapped. Clearing the fault current, Eq. (12) is available:

IF = IR + (−Ic )

(12)

This equation represents a method to consider the signals in the case of the time domain and the vectors in the phasor domain. This operation makes it possible to obtain the fault current from the capacitive and resistive components of a system with a resistive grounding method. 6. Selectivity of phase-to-ground fault relay Fig. 10 (b) shows a representation of the capacitive currents for cables with different sizes given a voltage of 12.47 kV. The graph makes it possible to quantify the capacitive current with an increase in the length of the cables. This section shows how the loss of selectivity of the relay due to phase-to-ground faults in different locations of the power system was validated using a software of the current flowing through the grounding method as an adjustment value for the relays of each branch, taking as a reference the value obtained in Fig. 6: √ .57 = 9.5A. The values of the currents are expressed in bar charts based on the current measurements durIPickup = 0.4∗33 2

ing a fault taken from the simulations in ATPdraw. For each of the scenarios, the current in the grounding resistance was measured, and the residual currents in the three branches were obtained. Fig. 11 illustrates the percentages of selectivity improvement with the proposed solution under different failure events probabilities, for all cable sizes. As the length of the cable increases, the solution presented in this article acquires greater relevance, while maintaining efficient discrimination against introduction of capacitive current in non-failed branches. In

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Fig. 11. Selectivity improvement.

Fig. 12. Residual currents: (a)phase-to-ground fault in branch 1, (b) phase-to-ground fault in branch 2 and (c) phase-to-ground fault in branch 3.

Fig. 12(a), the current measurements corresponding to a phase-to-ground fault in branch 1. The horizontal dotted line refers to the setting for the activation level of the relay. In this scenario, the relays of branch 1 and branch 3 will be activated. Because the fault occurs in branch 1, it operates correctly. However, this behaviour produces an erroneous trip in branch 3 as a result of the capacitive current that flows through the neutral of the system and exceeds the value set for the relay in that branch. In the same way, a phase-to-ground fault was programmed in branch 2, and the results are shown in Fig. 12(b). In this scenario, firing occurs in branches 2 and 3. There is a correct trip in branch 2, while the relay in branch 3 is activated erroneously. For the third scenario, a fault was programmed in branch 3, and the results obtained are shown in Fig. 12(c). In this case, only the relay in the branch with the fault is activated, which shows the correct operation of the system. The results shown for each of the scenarios computationally validates the effect of the loss of selectivity of the phase-to-ground fault relay, due to the circulation of capacitive currents. For this model, we used a 4/0 cable size, a maximum length of 4572 m in branch 3, and an adjustment current of 40% of the current flowing through the neutral of 9.5 A. When comparing them with the graph, we lose selectivity from 3600 m. When applying the model, this occurs in branch 3, as shown in the software simulations.

7. Directional ground relay solution The use of function 67N is proposed for the protection of each branch. The idea is that the residual current of the faulty branch falls in the operating area of the relay, while the capacitive currents of the other branches remain in the reverse operating area of the relay. The measurements are made in the common bar of all the branches, so that the residual voltage is the same for all the branches, while the current changes depending on the branch that fails.

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Fig. 13. Phasor representation of residual currents and voltages - fault in branch 1.

Fig. 14. Directional current algorithm results: (a) fault in branch 1, (b) fault in branch 2 and (c) fault in branch 3.

The selection of a characteristic angle of 45◦ ensures that the residual current of the branch in fault is located in the middle of the detection zone and also shields the system, allowing it to perform the detection in cases where some component of the residual current prevails. This is because the detection zone for these cases would approach the range between 180◦ and 270◦ ffor currents with predominantly resistive or capacitive components. The three scenarios studied in Section 6 can be assessed to illustrate the algorithm. The first scenario recreates a phaseto-ground fault in branch 1. Residual current and phase voltage signals are exported from ATPdraw to Matlab, the FFT is applied, and the magnitude and phase of the fundamental component are extracted. Fig. 13 sshows the phasor representation of the event analysed. The residual voltage is calculated by the sum of the phase voltages, and the system is located on the axis of symmetry, which is equal to rotating all of the phasors 180◦ (to apply the directional current algorithm shown in Section 3). The results are shown in Fig. 14(a). A second scenario reproduces a phase-to-ground fault in branch 2. The phasor representation is performed, the directional current detection algorithm is applied, and the results are shown in Fig. 14(b). The third scenario recreates a phase-to-ground event in branch 3; the same procedure is applied to branches 1 and 2, obtaining the results shown in Fig. 14(c). The residual current phasor of branch 3 generates a successful relay trip, while branches 2 and 3 do not exceed the set values. Therefore, the relays will not be activated. The evaluations of the phase-toground fault scenarios shown in this section correspond in magnitude to the bar diagrams shown in Section 6. However, when discriminating the failed branches based on the directional current, the problem of selectivity is appropriately solved. 8. Real-time hardware-in-the-loop assessment and playback testing Based on the successful results obtained in simulations, a hardware implementation was performed to validate the proposed methodology for the solution of the loss of selectivity of the relay.

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ATPDraw Comtrade PC PC

LabVIEW USB [VA, VB, BC, IA, IB, IC]

Analog Data [VA, VB, BC, IA, IB, IC]

CompactRIO

Relay Remote HMI

Ethernet

Analog Injection [VA, VB, BC, IA, IB, IC]

DTRS 66

Micom P145

Fig. 15. Real-time hardware-in-the-loop testing framework.

Fig. 16. Directional current algorithm results for fault in branch 1 (a) offline simulation, and (b) real-time test.

This application is an extension of a previous real-time hardware-in-the-loop test bed designed for protective relay control and playback testing [14,15], where the successful validation of virtual relays was performed for transmission line protection [16–18], power swing protection [19], feeder protection [20], overcurrent protection coordination [21] and integration of new algorithms such as mathematical morphology [22]. Fig. 15 shows the scheme used in the laboratory to perform the tests. A COMTRADE file containing the voltage and current signals per phase was extracted from ATPDraw, processed by LabVIEW, and sent to a CompactRIO platform, which transferred the signals to the DTRS 66 equipment for advanced relay testing. These signals were injected into a Micom P145 relay, which in turn used Ethernet to allow the signals to be displayed when a fault event occurred. The first test scenario recreated a phase-to-ground fault in branch 1. Fig. 16 shows the results obtained in the simulation (a) and real-time test (b). By comparing the two representations, it can be concluded that the results are equivalent in both magnitude and phase. The equivalence between these results shows that the suggested methodology for the discrimination of the failed branch works in the hardware and therefore makes it possible to solve the selectivity problem for an overcurrent relay to ground. The second test scenario recreates phase-to-ground fault in branch 2. Fig. 17 shows the results. Making a comparison between the two representations, it is concluded that there is a marked equivalence, which confirms that the methodology works in a general way and could be applied to any underground mining system whose specifications lead to the loss of selectivity by a capacitive current. In the third test scenario, a failure occurred in branch 3. This scenario was also tested in

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Fig. 17. Directional current algorithm results with fault in branch 2 (a) offline simulation, and (b) real-time test.

hardware, and successful results were obtained in the detection of the failed branch and the discrimination of the branches that did not fail by the activation adjustment of the relay. In this scenario, there were no selectivity problems, as shown in Fig. 8(a). Therefore, the use of an overcurrent relay to ground in branch 3 made possible the reliable operation of the system. 9. Conclusions and future work The proposed methodology for a directional current to ground developed in this study improves the selectivity performance. The results obtained both from simulations and hardware implementations validated the functionality of the proposed solution, which also prevents the erroneous triggering of relays. Existing safety mining regulations guarantee the adequate operation of power systems. For this reason, this article maintains the safety considerations outlined in the Federal Code of Mining Regulations of the USA. Under these considerations, the methodology for a directional current to ground was proposed as an ideal complement to eliminate the selectivity problem for the relays of larger mining power systems, while being consistent with the conventional settings. This study considered large-scale mining systems, where there are problems with a loss of selectivity for relays. These recommendations do not have an impact on small systems because the use of the overcurrent protection function to earth allows adequate discrimination of the faulty branches. Declaration of Competing Interest I have no conflict of interest to report. References [1] Leyes E. Code of federal regulations, title 30, mineral resources;. revised as of July 1, 1976, in Code of Federal Regulations, title 30, mineral resources; revised as of July 1, 1976. 1976. National archive; 1976. [2] Novak T. Safety analysis of trailing cables used on 2400-v continuous mining machines. IEEE Trans Ind Appl 2012;48(2):567–74. doi:10.1109/TIA.2012. 2183953. [3] Zhang Y. Active current injection method for limiting ground fault current harmonicsin underground coal mines. University of Kentucky; 2014. Ph.D. thesis. Theses and Dissertations–Mining Engineering. 15., https://uknowledge.uky.edu/mng_etds/15. [4] Novak T, Basar J, Sottile J, Kohler JL. The effects of cable capacitance on longwall power systems. IEEE Trans Ind Appl 2004;40(5):1406–12. doi:10.1109/ TIA.2004.834025. [5] Sottile J, Gnapragasam SJ, Novak T, Kohler JL. Detrimental effects of capacitance on high-resistance-grounded mine distribution systems. IEEE Trans Ind Appl 2006;42(5):1333–9. doi:10.1109/TIA.2006.880844. [6] Novak T, Morley LA, Trutt FC. Sensitive ground-fault relaying. IEEE Trans Ind Appl 1988;24(5):853–61. doi:10.1109/28.8990. [7] Dunki-Jacobs JR, Shields FJ, Pierre CS. Industrial power system grounding design handbook. Dexter, MI : Thomson-Shore; 2007. [8] Paul D. Phasor diagram of a single-phase-ground fault current in a high-resistance grounded power system. In: 2017 IEEE/IAS 53rd Industrial and Commercial Power Systems Technical Conference (I CPS); 2017. p. 1–6. doi:10.1109/ICPS.2017.7945108. [9] Roberts J, Altuve HJ, Hou D. Review of ground fault protection methods for grounded, ungrounded, and compensated distribution systems. In: 27th Annual Western Protective Relay Conference. [2] IEEE Power System Relaying Committee, âǣDistribution Line Protection Practices-Industry Survey Results; 2001. [10] Basar JJ. Improvement of ground-fault relaying selectivity through the application of directional relays to high-voltage longwall mining systems. Virginia Tech; 2004. Ph.D. thesis. http://hdl.handle.net/10919/9888.

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[11] Sottile J, Novak T, Tripathi AK. Best practices for implementing high-resistance grounding in mine power systems. IEEE Trans Ind Appl 2015;51(6):5254–60. doi:10.1109/TIA.2015.2420632. [12] Paul D. High-resistance grounded power-system equivalent circuit damage at the lineâground fault locationâpart i. IEEE Trans Ind Appl 2014;50(6):4179–87. doi:10.1109/TIA.2014.2346702. [13] Novak T. The effects of very-high-resistance grounding on the selectivity of ground-fault relaying in high-voltage longwall power systems. IEEE Trans Ind Appl 2001;37(2):398–406. doi:10.1109/28.913702. [14] Celeita D, Hernandez M, Ramos G, Penafiel N, Rangel M, Bernal JD. Implementation of an educational real-time platform for relaying automation on smart grids. Electr Power Syst Res 2016;130:156–66. doi:10.1016/j.epsr.2015.09.003. [15] Montaa DAM, Rodriguez DFC, Rey DIC, Ramos G. Hardware and software integration as a realist scada environment to test protective relaying control. IEEE Trans Ind Appl 2018;54(2):1208–17. doi:10.1109/TIA.2017.2780051. [16] Celeita D, Meliopoulos AS, Ramos G, Romero L. Dynamic state estimation for double-end traveling wave arrival identification in transmission lines. ElectrPower Syst Res 2019;170:138–49. doi:10.1016/j.epsr.2018.12.019. http://www.sciencedirect.com/science/article/pii/S0378779618304164. [17] Celeita D, Ramos G, Meliopoulos AS. Transmission line protective relay based on recursive least-square filters and weights analysis. Int Rev Model Simul 2018;11(5):277–87. doi:10.15866/iremos.v8i5.7057. Unpublished [18] Celeita D, Flores A, Ramos G, Pohl M. Design of virtual distance protection for offline transmission line relay testing. In: 2018 IEEE 38th Central America and Panama Convention (CONCAPAN XXXVIII); 2018. p. 1–6. doi:10.1109/CONCAPAN.2018.8596426. [19] Rodriguez DFC, Gutierrez M, Toro M, Ramos G. Out-of-step protection modeling for virtual playback testing applied to industrial generators. IEEE Trans Ind Appl 2019. doi:10.1109/TIA.2019.2897670. 1–1 [20] Rodriguez DFC, Osorio JDP, Ramos G. Virtual relay design for feeder protection testing with online simulation. IEEE Trans Ind Appl 2018;54(1):143–9. doi:10.1109/TIA.2017.2741918. [21] Pico JD, Celeita D, Ramos G. Protection coordination analysis under a real-time architecture for industrial distribution systems based on the std ieee 242–2001. IEEE Trans Ind Appl 2016;52(4):2826–33. doi:10.1109/TIA.2016.2538739. [22] Celeita D, Perez JD, Ramos G. Assessment of a decaying dc offset detector on cts measurements applying mathematical morphology. IEEE Trans Ind Appl 2019;55(1):248–55. doi:10.1109/TIA.2018.2867530. German Gutierrez received the degree in Electronic Engineering (2014) from the Universidad Pedagógica y Tecnológica de Colombia Sogamoso, Colombia, along with an M.Sc. (2018) in Electrical Engineering from Universidad de los Andes, Bogota, Colombia. His research interest includes harmonics, industrial automation and power quality. David Celeita received the degree in Electronic Engineering (2011) from the Universidad Distrital, along with an M.Sc. (2014) and a Ph.D (2018) in Electrical Engineering from Universidad de los Andes, Bogota, Colombia. He worked as an automation engineer in low and medium voltage applications, and he was a visiting researcher at Georgia Institute of Technology. He is currently a postdoc researcher in the field of protection algorithms for HV lines at CentraleSupélec and an industry partner in France. His research interests include Protective - Relaying Control, Smart Grids, Advanced Distribution Automation, Fault Location, and Real-Time Simulation. Gustavo Ramos received a degree in Electrical Engineering (1997) from the Universidad Nacional, Manizales, Colombialong with an M.Sc. (1999) and PhD (2008) in Electrical Engineering from Universidad de Los Andes, BogotÃ!‘, Colombia. He is currently an Associate Professor with the Department of Electrical Engineering at the School of Engineering, Universidad de Los Andes and his current research interests include power quality, and transients in grounding systems.