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Conference Paper

NREL/CP-5D00-74785

July 2020

High-Frequency Signature-Based Fault 

Detection for Future MV Distribution Grids

Preprint

Yaswanth Nag Velaga,

1 Kumaraguru Prabakar,1

Akanksha Singh,

1 and Pankaj K. Sen2

1 National Renewable Energy Laboratory

2 Colorado School of Mines

Presented at the 56th IEEE Industrial and Commercial Power Systems 

Technical Conference (I&CPS)

June 29–July 28, 2020
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Conference Paper

NREL/CP-5D00-74785

July 2020

High-Frequency Signature-Based Fault

Detection for Future MV Distribution Grids

Preprint

Yaswanth Nag Velaga,1 Kumaraguru Prabakar,1

Akanksha Singh,1 and Pankaj K. Sen2

1 National Renewable Energy Laboratory

2 Colorado School of Mines

Suggested Citation 

Velaga, Yaswanth Nag, Kumaraguru Prabakar, Akanksha, Singh, and Pankaj K. Sen. 

2020. High-Frequency Signature-Based Fault Detection for Future MV Distribution Grids:

Preprint. Golden, CO: National Renewable Energy Laboratory. NREL/CP-5D00-74785. 

https://www.nrel.gov/docs/fy20osti/74785.pdf. 

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High-Frequency Signature-Based Fault Detection for 

Future MV Distribution Grids 

Yaswanth Nag Velaga∗, Kumaraguru Prabakar∗, Akanksha Singh∗, and Pankaj K. Sen†

∗Power Systems Engineering Center, National Renewable Energy Laboratory, Golden, Colorado 

†Department of Electrical Engineering, Colorado School of Mines, Golden, Colorado 

Abstract—Increasing penetration levels of inverter based distributed energy resources (DERs) impact the legacy distribution 

system protection. Inverter based DERs provide approximately 

1.2-2 pu fault current. In systems with high penetration of 

inverter based DERs, it is diffcult for over-current based protection schemes to differentiate between normal loading conditions and a fault. Directional, distance, and adaptive forms 

of protection schemes are also affected by low fault currents. 

This paper analyzes fault generated traveling wave (TW) based 

high-frequency signatures in the distribution system. In order 

to simulate such signatures, frequency-dependent distributed 

parameter line modeling approach is used in this research 

work to represent distribution lines and underground cables. 

Modifed IEEE 13-bus medium voltage test system is modeled in 

electromagnetic transient simulation tool and multiple transient 

scenarios are simulated in this test system. The results are 

analyzed to understand the high-frequency signatures that can be 

used to detect and locate faults under high penetration of DERs. 

Keywords—Distribution line modeling, distribution system protection, electromagnetic transient program, EMTP, fault analysis,

fault detection, high-frequency signature, traveling wave.

I. INTRODUCTION 

Increasing penetration levels of distributed energy resources (DERs) in distribution network may cause bidirectional power fow and raise several challenges in system 

operations such as protection, voltage rise, frequency control, 

and stability [1]. Out of these, protection requires special 

attention because it can restrict and limit the expansion and 

penetration level of DERs. Traditional synchronous generators 

generally contribute a signifcant amount of fault current: 

approximately 6-10 times the rated generator current. Existing 

overcurrent (OC) schemes utilize this large fault current to 

distinguish between normal and fault condition. 

Inverter based DERs, unlike rotating machines, lack the 

inertia and contribute about 1-2 pu fault current [2]. This 

fault current contribution also varies depending on system 

This work was authored by the National Renewable Energy Laboratory, 

operated by Alliance for Sustainable Energy, LLC, for the U.S. Department 

of Energy (DOE) under Contract No. DE-AC36-08GO28308. This material 

is based upon work supported by the U.S. Department of Energy’s Offce 

of Energy Effciency and Renewable Energy (EERE) under Solar Energy 

Technologies Offce (SETO) Agreement Number 34237. The views expressed 

in the article do not necessarily represent the views of the DOE or the U.S. 

Government. The U.S. Government retains and the publisher, by accepting 

the article for publication, acknowledges that the U.S. Government retains a 

nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce 

the published form of this work, or allow others to do so, for U.S. Government 

purposes. 

confguration and operating conditions. As DER penetration 

increases, the existing OC fault protection schemes may fail 

due to bi-directional power fow. These protection issues are 

discussed in [3]. Additionally, fault ride-through capabilities 

of DERs, that are defned in the interconnection standard [4], 

create additional challenges for existing protection schemes. 

Recent studies have proposed multiple approaches to address these issues. A directional overcurrent protection scheme 

with communications links to all protective devices based 

on the IEC 61850 substation automation system is proposed 

in [5]. But, directional feature is not reliable for close-in faults 

because of voltage collapse on faulted phases [6]. Furthermore, 

the directionality of single-line-to-ground (SLG) faults is diffcult to detect because of the existence of multiple ground 

sources in a distribution system. Also, installation of potential 

transformers, current transformers, and communications at 

every node will be expensive. 

Distance based protection schemes provide inherent directionality with fxed reach and are independent of changing 

system conditions [7]. They are affected by outfeed of tap 

load currents and infeed from generation sources between 

the fault and relay, which makes these schemes over-reach 

and under-reach. Distance protection is also not reliable for 

detecting high-resistance faults and reverse close-in faults. 

Additionally, changing grid conditions make it diffcult to 

fnd the optimal relay settings. And applying such distance 

based protection schemes to a distribution system with DERs 

is challenging. To account for changes in system topology, 

such as grid-connected and islanded operation, the adaptive 

protection scheme has gained traction in recent years [8]–[13]. 

Typically, centralized control detects the network status and 

updates the device settings over a communications link. But, 

they are expensive and practical diffculties are associated with 

their implementation [14]. 

The future distribution system with high penetration of 

inverter based DERs requires innovative protection schemes 

that can handle variable fault current, bi-directional power 

fow, and changing system conditions, etc. This paper proposes 

the traveling wave (TW) based high-frequency signatures for 

medium voltage (MV) distribution system protection. When a 

fault occurs, a very high-frequency voltage and current waves 

are generated at the fault point. These waves are independent 

of the type of fault, topology of the system, and location of the 

fault. In transmission system, transients caused by the faults 

contain energy over a broad frequency spectrum, between 

20 kHz and 2 MHz [15]. There are many challenges to design 

a protection scheme based on TWs because of their behavior 

1

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across different components in the system. 

A study in [16], proposed a single line-to-ground (SLG) 

fault travelling wave based protection for less effective 

grounded systems. A simplifed 35 kV radial distribution system is used. Authors did not analyze the effect of transformers 

and capacitors which are predominantly found in distribution. 

Similarly, the authors in [17] discussed using the protection 

scheme on medium voltages to detect the direction and location using only current measurements. The authors assumed 

that the lines were homogeneous with no discontinuities and 

refections. Because of the presence of a wide-frequency band 

in the generated waves, TWs in the distribution system are 

referred to as high-frequency waves (HFW). 

The rest of this paper is organized as follows. Section II 

briefy discusses the challenges to using TWs in the distribution system. Section III discusses the importance of selecting 

the appropriate line models that can be used in the system 

transient model and provides a comparison among the available 

models. Section IV discusses the modeling of a modifed IEEE 

13-bus distribution system in EMTP-RV without transformers 

to understand the TW behavior in distribution systems and 

presents different simulation cases and results by varying the 

location of the faults and fault type. 

II. HIGH-FREQUENCY TRANSIENTS IN DISTRIBUTION 

SYSTEMS 

Faults on a distribution line generate surges that propagate 

in the form of HFWs at fnite velocity along the line from 

the fault location. The propagation velocity, characteristic 

impedance and the wave shape of these HFWs depend on 

the line parameters: resistance (R), inductance (L), capacitance 

(C), and conductance (G). Generally, line parameters are 

distributed and depend on the physical aspects of the line 

design, such as tower geometry, type of conductor, conductor 

spacing, shield wire, and insulator type. Fig. 1 shows the 

equivalent circuit of a single conductor line of length Δx. 

Consider the simple system shown in Fig. 2 with a overhead line segment of characteristic impedance Zc between the 

buses A and B. The line left of bus A has a characteristic 

impedance Zd, and the line right of the bus B has a characteristic impedance Ze. A fault on the line between terminals 

A-B (as shown in Fig. 2) initiates the wave transients Ai & Bi 

(Voltage and Current) traveling at approximately speed of light. 

As the fault generated transients approach the terminals A-B, 

part of the wave is refected as Ar & Br and rest is transmitted 

as At & Bt to the adjacent sections of the terminal. This 

refection and transmission of the HFW at the junction (or taps) 

are caused by the changes in characteristic impedance. These 

waves also attenuate as they propagate along the line because 

of losses caused by the line resistance (R) and conductance 

(G). Attenuation of these waves are greater in distribution 

systems because of the higher resistance in the lines. 

Use of such TWs for fault detection and fault location in 

transmission system is relatively straightforward by using the 

characteristics of the line, such as length and impedance. But, 

compared to a transmission system, there are challenges in 

using the TW based approach in a distribution network. This 

is due to the presence of unbalanced short lines, underground 

cables, transformers, voltage regulators, capacitor banks and 

RΔx LΔx

GΔx CΔx

Δx

i(x,t) i(x+Δx,t)

v(x,t) v(x+Δx,t)

Fig. 1. Equivalent circuit of single-conductor line 

At A Ai Bi B

Ar Br

Bt

Zd Zc Ze

Ai, Bi - Incident Ar, Br - Reflected

At

, Bt - Transmitted

Fig. 2. Illustration of traveling wave at junctions 

Line Models

Lumped Parameters Distributed Parameters

PI Model Constant

Parameters

(CP)

Frequency - Dependent Parameters (FD)

CP Model Modal Domain Phase Domain

FD Model Universal Line

 Model (ULM)

Fig. 3. General classifcation of line/cable models 

frequent taps (point of discontinuities) which can impact TWs 

in distribution system. 

III. FREQUENCY DEPENDENT LINE MODELING FOR 

TRANSIENT ANALYSIS 

Accurate representation of the component model is crucial 

for all transient analyses. Power systems component modeling 

for electromagnetic transient (EMT) analyses is developed 

by considering the frequency band of the transients to be 

analyzed and the frequency dependence of the parameters. 

Fig. 3 shows the general classifcation of the line models 

applicable to overhead lines and underground cables. Line 

models are divided into two groups: lumped and distributed 

parameter models. Selecting the model group depends on the 

highest frequency involved in the study and, to a lesser extent, 

on the line length. 

A. Lumped Parameter Models

Lumped parameter models, also known as pi models, 

represent the lines and cables by using lumped resistance 

(R), inductance (L), conductance (G), and capacitance (C) 

elements. These values are calculated at a single frequency of 

interest. These models are suffcient for steady-state analysis, 

2

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but they can also be used for transient studies provided the 

parameters are evaluated at the frequency of interest. In some 

cases, these short pi sections are cascaded to approximate the 

distributed nature of the line. The number of such sections 

depends on the expected frequency of the transient [18]. 

Cascaded pi sections are computationally expensive because of 

the increase in matrix size. Because pi models do not represent 

a propagation delay and attenuation of waves, they are best 

suited for steady-state analysis where high-frequency events 

are ignored. 

B. Distributed Parameter Models

Distributed parameter line models are divided into two 

sub-groups: constant parameter (CP) and frequency dependent 

(FD). 

1) Constant Parameter Model: A constant parameter or 

Bergeron’s model calculates the line parameters at a single 

frequency. Then, the calculated line parameters (inductance 

and capacitance) are distributed. The losses in the line are 

represented through lumped resistances at three discrete points 

along the line (1/4 at the both ends, and 1/2 at the middle) [19]. 

Since the line impedances are constant for the whole range 

of frequencies in the study, this approach can give inaccurate 

results if the frequency of the event under study is farther 

from the frequency used to calculate the parameters. Unless 

the frequency of the desired transient is known beforehand, 

the constant parameter model might yield inaccurate results at 

certain frequencies. 

2) Frequency-Dependent Model: The frequency dependent 

(FD) model has the potential to represent the true nature of 

the line. Frequency-dependent line/cable models can further 

be divided into two groups based on the domain in which the 

frequency-dependent line parameters are evaluated: modal and 

phase domain. 

FD model is a modal domain model that utilizes the transformation matrix to transform from phase to mode and mode 

to phase quantities. Frequency dependence of line parameters 

are considered in modal domain. FD model assumes the 

frequency independent transformation matrix which is constant 

and real [20]. Accuracy of the FD model is limited to symmetrical or balanced overhead lines because the transformation 

matrices are frequency dependent for unbalanced lines and 

cables. 

Universal Line Model (ULM) is a phase domain model that 

works directly in the phase domain by taking the full frequency 

dependence of line parameters in the form of Characteristic 

admittance (YC) and the propagation constant (H) matrix 

transfer functions [21]. ULM models avoids the transformation 

matrix and simplifying assumptions in the form of modal to 

phase transformations. This model is highly accurate for any 

overhead line and underground cable confgurations. The ULM 

model or wide-band model is the best choice for most EMT 

studies and used in this study. 

IV. TEST SYSTEM FOR HIGH-FREQUENCY WAVE 

VALIDATION 

A simple test system that represents the common features 

of a typical distribution system is necessary to understand the 

HFWs generated during the transient events. Fig. 4(a) is a modifed version of the IEEE 13-bus test distribution system [22]. 

This modifed test system has one feeder operating at 13.8 kV. 

This feeder runs from the substation for a total length of 4 

miles. There are four taps taken off the main feeder to represent 

real world feeder taps. 

Confgurations for tower structures, overhead lines, and 

underground cables used in the test system are based on [23]. 

The details are presented in Fig. 5. Table I shows the various 

conductors used in the overhead line confgurations. Fig. 5(a), 

(b), & (c) shows the geometric structure of the overhead lines 

and conductor spacing. Concentric neutral cable (250 AWG) 

data taken from [22] is used for the underground cable modeling (shown in Fig. 5(d)). The over head lines, and underground 

cables in the test networks are modeled using the frequency 

dependent modeling approach known as Universal Line Model 

(ULM). 

To compare the simulation results with theoretical time 

taken for the waves to propagate in the lines, alpha mode 

velocities generated by the ULM model is used to create wave 

travel time diagram of the test system. This propagation time 

for HFWs in the test system is shown in Fig. 4(b). Table II 

shows the single and three phase loads modeled in the test 

system. Load modelling ensures that the pre-fault current is 

established in the network steady-state. The source in the 

test network is modeled as voltage source behind equivalent 

Thevenin impedance. The impedance values are  ́ calculated 

from fault MVA at power frequency. 

V. SIMULATION OF TRANSIENT SCENARIOS FOR 

HIGH-FREQUENCY STUDY 

Multiple fault transient scenarios are simulated to characterize the behavior of HFWs in distribution system. The 

simulations are run at a time step of 100 ns. Faults are applied 

at 51.83 ms (Phase A voltage peak) after the initialization. 

Current and voltage probes are installed on every bus and 

TABLE I. OVERHEAD LINE DATA 

Node A Node B Length(mi) Phase Neutral Spacing ID AWG ACSR 

650 632 1 4/0 1 500 

632 633 0.5 1/0 2 505 

633 634 0.3 4/0 1 510 

632 645 0.5 1/0 2 505 

645 646 0.3 4/0 1 510 

646 647 0.3 4/0 1 510 

632 671A 1 4/0 1 500 

671A 671B 1 4/0 1 500 

671B 680 1 4/0 1 500 

671B 675 0.5 1/0 2 505 

671B 684 1 4/0 1 500 

684 611 0.3 4/0 1 510 

684 692 0.3 4/0 1 510 

TABLE II. TEST SYSTEM LOAD DATA 

Bus Type Real Power (MW) Reactive Power (MVAR) 

647 1-ph 0.1 0.05 

634 1-ph 0.1 0.05 

692 1-ph 0.03 0.0005 

611 1-ph 0.1 0.05 

645 3-ph 0.5 0.25 

633 3-ph 5 2.5 

675 3-ph 0.5 0.1 

680 3-ph 2 1 

3

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Tap 1

632

650

671A

671B

680

684

692

611

647 646 645 633 634

A

B

C

N

675

1mi

1mi

1mi

1mi

0.3mi 0.3mi 0.5mi 0.5mi 0.3mi

0.3mi 1mi 0.5mi

0.3mi

13.8kV

Tap 2

Tap 3 Tap 4

OH OH OH OH

OH OH OH OH OH/UG

OH OH OH/UG OH 5.39

μ

s 5.39

μ

s 5.39

μ

s 5.39

μ

s

2.69μs/4.50μs

2.69μs/4.50μs

1.64μs 1.64μs 2.69μs 1.64μs

1.64μs 5.39μs 1.64

μ

s

OH - Overhead Line

UG - Underground Cable

647 646 645 633 634

671B

611 684

671A

680

675

692

632

650

13.8kV

(a) Distribution 13-bus test system 

Fig. 4. Modifed IEEE test system used in electromagnetic transients simulation 

sampled at 10 MHz to record the high-frequency signatures. 

The signals measured from these probes are fltered through a 

high-pass flter with a cut-off frequency of 10 kHz. Two fault 

types, single-line-to-ground (SLG) fault and line-to-line (LL) 

fault are simulated in two locations. First, on the main feeder. 

Second, on the tap 1. 

(b) Distribution 13-bus wave travel time 

A. Overhead Lines with SLG Fault on Bus 680

An SLG fault is applied on Phase A at bus 680. Filtered 

and non-fltered instantaneous values of current and voltage are 

recorded every 10 μs. All the travel time between buses can 

be estimated from Fig. 4(b). Filtered signals of all the buses 

in taps 1 and 3 are plotted to show the arrival times of the 

waves and compare them with the times shown in Fig. 4(b). 

The faulted bus current (680) is included in all the plots as a 

3.6' 3.6'

0.5'

24'

4'

1.5'

0.5'

24'

4'

4'

4'

3.6'

0.5'

24'

4'

4'

0.5'

0.5' (a) ID-500 (b) ID-505 (c) ID-510 (d) ID-515 

Fig. 5. Overhead line tower geometry, and underground cable layout 

4

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51.82 51.83 51.84 51.85 51.86 51.87 51.88 51.89 51.90

Time (ms)

20

15

10

5

0

5

Current (A)

5.3 s

10.7 s

12.3 s

680 (Faulted Bus)

671B

684

611

51.82 51.83 51.84 51.85 51.86 51.87 51.88 51.89 51.90

Time (ms)

8

6

4

2

0

2

Voltage (kV)

5.3 s

10.7 s

12.3 s

680 (Faulted Bus)

671B

684

611

Fig. 6. Filtered currents in tap 3 buses for a SLG Fault 

reference to compare the wave arrival estimates. 

High-frequency current wave magnitudes are small due to 

lower distribution voltage. Current waves are proportional to 

the voltage wave over characteristic impedance. Voltage and 

current waves are initiated at the fault point on Phase A at 

51.83 ms of simulation time. The current waves propagate 

and arrive at bus 671B, part of the energy is refected back 

and the rest gets split between tap 3, tap 4 and main feeder. 

Initial current wave times estimated from Fig. 4(b) in tap 3 

(671B, 684, 611) are 5.39, 10.78, & 12.42 μs respectively. 

From the simulations, they arrive at 5.3, 10.7, & 12.3 μs as 

shown in Fig. 6. Faulted bus 680 records considerably longer 

duration waves with higher magnitude in Fig. 6, compared to 

the waves in tap 3. It is because the wave energy attenuates as 

it travels from fault point and transients recorded at the faulted 

terminal are always double the magnitude due to same polarity 

refections from fault. 

Phase to ground instantaneous values of the bus voltages 

in the tap 3 are shown in Fig. 7. Estimated times of arrival at 

each bus are estimated the same way as the currents. Voltage 

and current have the same wave propagation characteristics, 

so the wave arrival times mentioned in Fig. 6 are applicable 

to Fig. 7. 

Current transients recorded in tap 1 buses are shown in 

Fig. 8. Signals in tap 2 bus timings are very close to the 

estimated ones shown in the timing diagram. Comparing the 

magnitude of the wave that arrived at tap 1 and tap 3 in Fig. 6 

and 8, the former is less because of the attenuation in the 

2 mi of the main feeder. More taps results in the wave energy 

getting split at every tap. As the wave attenuates, it also gets 

distorted in shape, losing its sharp rise characteristic. Initial 

waves arriving at the bus terminals are always distinct from the 

system under normal conditions because these wave transients 

exist only due faults and switching. For faults on the main 

feeder, any taps that are far away from the fault might not see 

the measurable refection from the fault because the energy 

dissipates in every line between the fault and tap. The closest 

taps might see the second wave from the fault. 

Voltage signals in the tap 1 buses are shown in Fig. 9. 

Fig. 7. Filtered voltages in tap 3 buses for a SLG Fault 

51.82 51.83 51.84 51.85 51.86 51.87 51.88 51.89 51.90

Time (ms)

20

15

10

5

0

5

Current (A)

16.1 s

18.7 s

20.4 s

21.9 s

680 (Faulted Bus)

632

645

646

647

Fig. 8. Filtered currents in tap 1 buses for a SLG Fault 

The peak voltage of the traveling wave at bus 632 is about 

1.7 kV compared to 4.1 kV at bus 671B. Because the change 

in characteristic impedance is not great between the bus 632 

three-phase and bus 645 single-phase line, most of the wave 

energy is transmitted to the adjacent line. In distribution systems, different conductors might be used at the same voltage 

level in a line. Different conductors change the line distributed 

parameters, but the change in impedance between the adjacent 

lines is small. Fig. 9 shows that signals in adjacent buses are 

only displaced in time. A greater difference in characteristic 

impedance is seen with underground cables in the network. 

B. Overhead Lines and Underground Cables with a SLG Fault

on Bus 680

Underground cables are mostly used in the distribution 

systems to serve the densely populated areas. The characteristic 

impedance of a cable is almost 10 times lower than that of the 

overhead lines. Wave propagation velocity in a cable is 50-

60% of the speed of light due to the insulation permittivity. 

These crucial differences impact the speed and the attenuation 

5

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51.82 51.83 51.84 51.85 51.86 51.87 51.88 51.89 51.90

Time (ms)

8

6

4

2

0

Voltage (kV)

16.1 s

18.7 s

20.4 s

21.9 s

680 (Faulted Bus)

632

645

646

647

51.830 51.835 51.840 51.845 51.850 51.855 51.860

Time (ms)

4

3

2

1

0

1

Voltage (kV)

2.6 s

4.5 s

671B(OH)

671B(UG)

675(OH)

675(UG)

Fig. 9. Filtered voltages in tap 1 for a SLG Fault 

51.82 51.83 51.84 51.85 51.86 51.87 51.88 51.89 51.90

Time (ms)

8

6

4

2

0

Voltage (kV)

5.3 s

10.7 s

12.3 s

680 (Faulted Bus)

671B

684

611

Fig. 10. Filtered voltages in tap 3 buses for a SLG Fault 

inside the cables. Understanding and studying these differences 

between overhead lines and underground cables are critical 

for creating HFW based protection schemes for systems with 

overhead lines and underground cables. 

In this scenario, two overhead lines between buses 632 

and 645, buses 671B and 675 are replaced by a 3-single core 

cable. Each cable has a central conductor to carry load current 

and fault current, and an outer sheath to carry ground fault 

current. Characteristic impedance varies greatly with different 

cable confgurations. Since the characteristics of current and 

voltage transients are shown to have a same behavior, only 

voltage waves are used for discussion. First, the voltages in 

the tap 3 are shown and later comparison between signals at 

the buses where lines are replaced by cables is made. As seen 

in Fig. 10. underground cables did not impact the travel time 

in the overhead lines. Traveling waves are superimposed by 

an oscillatory signal which frequency is found to be around 

8 kHz. High pass flter tuned at 10 kHz attenuated only part 

of the signal. 

The underground cables cause the travel time to increase 

Fig. 11. Voltage wave travel between overhead and underground in tap 4 for 

a SLG Fault 

compared to the overhead line. The travel time through the 

cable is 4.5 μs, which is same as the simulated time. Voltage 

signals at bus 671B & bus 675 contain smaller magnitudes of 

refections compared to the overhead line as shown in Fig. 11. 

C. Overhead Lines with Line-to-Line (LL) Fault on Bus 680

A LL fault is the second most commonly occurring fault 

in the power system. An SLG fault on bus 680 is replaced 

by a LL fault between phases A and B. Wave transients that 

are propagated in phases A and B are approximately equal in 

magnitude and opposite in polarity. This phenomenon is also 

observed in the current during a LL fault event. When the 

waves propagate down the line in both phases, they attenuate 

and disperse based on the characteristic impedance of two 

phases. For horizontal line and vertical line confguration, 

distributed parameters such as inductance and capacitance are 

the same for all the phases except for the mutual impedance, 

resulting in small differences in magnitude and dispersion of 

the waves in both phases. 

A fault is applied at 49.14 ms, when the voltage difference 

between phases is greater. The arrival times of the waves match 

the estimated values shown in Fig. 4(b). Voltage waves of the 

selected buses in the taps 1 and 3 are shown in Figs. 12 and 13. 

A maximum voltage difference of 17.8 kV is observed between 

the phases at 49.14 ms. An overhead line between the bus 680 

and bus 671B is horizontally confgured with phase B slightly 

above phases A and C. Transients in both the phases look very 

similar in terms of attenuation and dispersion. There is a slight 

difference in the magnitude observed in Fig. 12 because of 

the difference in characteristic impedance. As the waves travel 

away from the fault, magnitude, attenuation and dispersion 

between phases become distinct, as shown in Fig. 13. The line 

between bus 632 and bus 645 is vertically confgured, resulting 

in the wave experiencing different distributed characteristic 

impedances. 

D. Overhead Lines with SLG Fault on Bus 645

All the cases discussed so far have the fault on the main 

feeder. In this case, the fault is assumed in the laterals closer 

6

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49.14 49.15 49.16 49.17 49.18 49.19

Time (ms)

7.5

5.0

2.5

0.0

2.5

5.0

7.5

Voltage (kV)

5.3 s

10.6 s

680(A-ph)

680(B-ph)

671B(A-ph)

671B(B-ph)

684(A-ph)

684(B-ph)

Fig. 12. Voltages in tap 3 buses for a LL Fault 

49.14 49.15 49.16 49.17 49.18 49.19

Time (ms)

7.5

5.0

2.5

0.0

2.5

5.0

7.5

Voltage (kV)

16 s

18.6 s

680(A-ph)

680(B-ph)

632(A-ph)

632(B-ph)

645(A-ph)

645(B-ph)

Fig. 13. Voltages in tap 1 for a LL Fault 

to the substation at bus 645 in tap 1. When the waves move 

away from the fault in both directions, one travels down the 

lateral until it gets refected, and the other reaches the junction 

and gets refected back toward the fault. Voltages in the tap 

are plotted starting with the faulted bus (see Fig. 14 and 15). 

The frst wave reaches bus 646 after 1.6 μs from the fault. The 

characteristic impedance of the single phase lines between the 

buses 645, 646, and 647 is the same. So, the wave launched at 

645 attenuates as it travels through two lines unrefected until 

it reaches bus 647. The waves in the other direction of the 

fault reach the junction at bus 632 after 2.6 μs. 

Voltage signals are used to compare and measure at tap 3 

in the test system. Filtered voltages at buses in tap 3 are plotted 

in Fig. 15. Compared to the signal peak at bus 632, 43% of the 

wave is recorded at 671B. Waves recorded at junction (632, 

671B) stay for a longer duration than buses inside the tap. So, 

the relays located at the taps can analyze the waves for a fault 

on main feeder and tap provided that the fault inside the tap 

is not located far away from the junction. 

51.82 51.83 51.84 51.85 51.86 51.87 51.88 51.89 51.90

Time (ms)

8

6

4

2

0

2

Voltage (kV)

645 (Faulted Bus)

646

632

647

Fig. 14. Voltages in tap 1 for a SLG Fault on bus 645 

51.82 51.83 51.84 51.85 51.86 51.87 51.88 51.89 51.90

Time (ms)

8

6

4

2

0

Voltage (kV)

645 (Faulted Bus)

671B

684

611

Fig. 15. Voltages in tap 3 for a SLG Fault on bus 645 

E. Summary of Results

The simulation results for faults on the main feeder show 

that two taps beyond faulted bus may not see a measurable 

second wave refection from the fault. It is observed from 

the results that line-to-line fault transients attenuation is less 

compared to the faults that involve ground. HFWs gets attenuated faster in the lateral due to high resistance of conductors 

combined with skin effect. This can be observed from the 

results when the fault location is changed from main feeder to 

the tap (see subsections (A) and (D)). In the scenario simulated 

in subsection (B), it is observed that waves travel slower 

through underground cables compared to the overhead lines. 

VI. CONCLUSION 

Limitations of overcurrent, directional, distance and adaptive relaying in the presence of inverter based DER require 

a novel protection technique that does not depend on fault 

current magnitudes and direction of power fow. This paper analyzed the application of traveling wave based high-frequency 

signature for fault detection and location in MV distribution 

systems. A modifed medium voltage IEEE 13-bus test system 

was presented in this paper to analyze fault generated HFWs. 

The need for frequency-dependent line modeling to study 

HFWs was presented and a frequency-dependent line modeling 

7

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approach was used to model the modifed IEEE 13-bus system. 

Different fault scenarios of the test system were simulated and 

results of the EMT simulations were presented. Simulation 

results indicate that the HFWs can be used as unique signatures 

of faults and can be used to detect and locate the faults under 

high DER penetration levels. 

ACKNOWLEDGMENT 

The authors gratefully acknowledge the support of Guohui 

Yuan and David Walter of the U.S. Department of Energy 

Solar Energy Technologies Program, Systems Integration Subprogram, for their valuable advice and funding for this work. 

REFERENCES 

[1] IEEE Working Group D3, “Impact of Distributed Resources on Distribution Relay Protection,” report to the Line Protection Subcommittee 

of the Power System Relay Committee of The IEEE PES, Aug. 2004. 

[2] J. Keller, B. Kroposki, R. Bravo, and S. Robles, “Fault current contribution from single-phase PV inverters,” in 2011 37th IEEE Photovoltaic

Specialists Conference, (Seattle, WA, USA), pp. 001822–001826, IEEE, 

June 2011. 

[3] K. Malmedal, B. Kroposki, and P. K. Sen, “Distributed Energy Resources and Renewable Energy in Distribution Systems: Protection Considerations and Penetration Levels,” in 2008 IEEE Industry Applications

Society Annual Meeting, (Edmonton, Alberta, Canada), pp. 1–8, IEEE, 

Oct. 2008. 

[4] “IEEE Standard for Interconnection and Interoperability of Distributed 

Energy Resources with Associated Electric Power Systems Interfaces,” 

in IEEE Std 1547-2018 (Revision of IEEE Std 1547-2003), pp. 1–138, 

Apr. 2018. 

[5] B. Hadzi-Kostova and Z. Styczynski, “Network Protection in Distribution Systems with Dispersed Generation,” in 2005/2006 PES TD, 

(Dallas, TX, USA), pp. 321–326, IEEE, 2006. 

[6] J. Horak and B. Electric, “Directional Overcurrent Relaying (67) 

Concepts,” in 59th Annual Conference for Protective Relay Engineers, 

(College Station, TX), p. 13, 2006. 

[7] D. Martin, P. Sharma, A. Sinclair, and D. Finney, “Distance protection 

in distribution systems: How it assists with integrating distributed 

resources,” in 2012 65th Annual Conference for Protective Relay

Engineers, (College Station, TX, USA), pp. 166–177, IEEE, Apr. 2012. 

[8] P. Mahat, Z. Chen, and C. L. Bak, “A Simple Adaptive Overcurrent 

Protection of Distribution Systems With Distributed Generation,” IEEE

TRANSACTIONS ON SMART GRID, vol. 2, no. 3, p. 10, 2011. 

[9] S. Ritter, H. Ruttinger, S. Ritter, P. Bretschneider, and D. Westermann, 

“New approaches for smart grid requirements: Grid protection and 

optimization of distribution grid operation,” in 2011 IEEE Power and

Energy Society General Meeting, (Detroit, MI, USA), pp. 1–7, IEEE, 

July 2011. 

[10] J. Yang and Y. Wang, “Review on protection issues of low-voltage 

distribution network with multiple powerelectronic-converter-interfaced 

distribution energy resources,” in International Conference on Renewable Power Generation (RPG 2015), (Beijing, China), pp. 6 .–6 ., 

Institution of Engineering and Technology, 2015. 

[11] H. Muda and P. Jena, “Sequence currents based adaptive protection 

approach for DNs with distributed energy resources,” IET Generation,

Transmission & Distribution, vol. 11, pp. 154–165, Jan. 2017. 

[12] J. Gers and C. Viggiano, “Protective relay setting criteria considering 

DERs and distributed automation,” in 13th International Conference on

Development in Power System Protection 2016 (DPSP), (Edinburgh, 

UK), pp. 9 .–9 ., Institution of Engineering and Technology, 2016. 

[13] B. P. Bhattarai, B. Bak-Jensen, S. Chaudhary, and J. R. Pillai, “An 

adaptive overcurrent protection in smart distribution grid,” in 2015 IEEE

Eindhoven PowerTech, (Eindhoven, Netherlands), pp. 1–6, IEEE, June 

2015. 

[14] M. Norshahrani, H. Mokhlis, A. H. A. Bakar, J. J. Jamian, and 

S. Sukumar, “Progress on Protection Strategies to Mitigate the Impact of 

Renewable Distributed Generation on Distribution Systems,” Energies

2017, no. 1864, 2017. 

[15] E. O. Schweitzer, A. Guzman, M. V. Mynam, V. Skendzic,  ́

B. Kasztenny, and S. Marx, “Locating faults by the traveling waves 

they launch,” in 2014 67th Annual Conference for Protective Relay

Engineers, (College Station, TX), pp. 95–110, IEEE, Apr. 2014. 

[16] X. Dong, B. Wang, B. Dominik, and M. Redefern, “Traveling Wave 

Based Single-Phase-to-Ground Protection Method for Power Distribution System,” CSEE JOURNAL OF POWER AND ENERGY SYSTEMS, 

vol. 1, no. 2, p. 8, 2015. 

[17] N. Davydova and G. Hug, “Traveling wave based protection for medium 

voltage grids with distributed generation,” in 2017 IEEE Manchester

PowerTech, (Manchester, UK), pp. 1–6, IEEE, 2017. 

[18] “Guidelines for Representation of Network Elements When Calculating 

Transients,” 1990. 

[19] Neville Watson and Jos Arrillaga, Power Systems Electromagnetic

Transients Simulation. IET, frst ed. 

[20] J. R. Marti, “Accurate Modelling of Frequency-Dependent Transmission 

Lines in Electromagnetic Transient Simulations,” IEEE Transactions on

Power Apparatus and Systems, vol. PAS-101, pp. 147–157, Jan. 1982. 

[21] A. Morched, B. Gustavsen, and M. Tartibi, “A universal model for 

accurate calculation of electromagnetic transients on overhead lines and 

underground cables,” IEEE Transactions on Power Delivery, vol. 14, 

pp. 1032–1038, July 1999. 

[22] W. H. Kersting, “Radial distribution test feeders,” in 2001 IEEE Power

Engineering Society Winter Meeting. Conference Proceedings (Cat.

No.01CH37194), vol. 2, (Columbus, OH, USA), pp. 908–912, IEEE, 

2001. 

[23] “Distribution Conductor Clearances and Span Limitations,” July 2003. 

8

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