NREL is a national laboratory of the U.S. Department of Energy, Office of Energy 

Efficiency & Renewable Energy, operated by the Alliance for Sustainable Energy, LLC.

Contract No. DE-AC36-08GO28308

Wind Power Plant Enhancement 

with a Fault-Current Limiter 

Preprint

E. Muljadi and V. Gevorgian

National Renewable Energy Laboratory

F. DeLaRosa

Zenergy Power Inc.

To be presented at the IEEE PES Annual Meeting

Detroit, Michigan

July 24-29, 2011

Conference Paper

NREL/CP-5500-49781

March 2011
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1 

Wind Power Plant Enhancement with a Fault-Current Limiter

Abstract— An important aspect of wind power plant (WPP) 

impact studies is to evaluate the short-circuit current (SCC) 

contribution of the plant into the transmission network under 

different fault conditions. This task can be challenging to protection engineers due to the topology differences between different 

types of wind turbine generators (WTGs) and the conventional 

generating units.

While it is important to size the circuit breaker to accommodate the available SCC, the available rating of a circuit breaker 

is limited. A fault-current limiter (FCL) has the capability to 

limit the SCC, and in some cases, may be necessary to limit the 

SCC from a WPP to defer the replacement of a circuit breaker 

currently installed in the transmission line.

This paper investigates the capability of a saturable core FCL 

to limit the SCC of different types of WTGs. Different faults are 

simulated to investigate the effectiveness of the FCL to limit the 

SCC and to reduce transient torque during faults. Several cases 

will be considered to demonstrate the benefits of using FCLs in 

unique situations. 

Index Terms — Fault contribution, saturable core faultcurrent limiter, induction generator, protection, short circuit, 

wind power plant, wind turbine. 

I. INTRODUCTION 

nergy and environmental issues have become one of the 

biggest challenges facing the world. In response to energy 

needs and environmental concerns, renewable energy technologies are considered the future technologies of choice [1], 

[2]. Renewable energy is harvested from nature, and it is 

clean and free. However, it is widely accepted that renewable 

energy is not a panacea that comes without challenges. With 

the federal government’s aggressive goal of achieving 20% 

wind energy penetration by 2030, it is necessary to understand the challenges that must be overcome when using renewable energy.

In the years to come, there will be more and more wind 

power plants (WPPs) connected to the grid. With the goal of 

20% wind penetration by 2030, the WPP’s operation should 

be well planned. The power system switchgear and power 

system protection for WPPs should be carefully designed to 

be compatible with the operation of conventional synchronous generators connected to the same grid. This paper illustrates the behavior of short-circuit current (SCC) contributions for different types of WTGs.

A. A Typical Wind Power Plant

As shown in Figure 1, a WPP consists of many (hundreds) 

of wind turbine generators (WTGs). Currently, available 

WTG sizes are between 1 MW and 5 MW. The WTG is connected at a low voltage level (e.g. 600 Volt), and a padmounted transformer is connected to step up the voltage to 

34.5 kV.

Several turbines are strung together and connected in a 

daisy chain fashion. The collector system is connected to the 

substation transformer where the voltage is stepped up to 

higher voltage (e.g., 230 kV) and the power is transmitted 

over long distance. Thus, a 300-MW wind power plant may

consist of 300 turbines connected to the grid.

In this paper, the fault-current limiter (FCL) is proposed to 

be installed at the collector side (34.5 kV) of the substation 

transformer. Thus, the FCL will be able to protect the entire 

WPP. 

B. Organization of the Paper

The organization of this paper is as follows; in section II, 

the FLC description and method of operation will be covered. 

In section III, the SCC characteristics of different WTG types 

will be presented. In section IV, the characteristics of different faults will be discussed. While in section V, the effectiveness of FCLs will be presented. Finally, in section VI, the 

conclusion will summarize the paper findings. 

E

E. Muljadi

Fellow, IEEE

NREL

1617 Cole Blvd.

Golden, CO 80401,

USA

eduard.muljadi@nrel.gov

V. Gevorgian

Member, IEEE

NREL

1617 Cole Blvd.

Golden, CO 80401,

USA

vahan.gevorgian@nrel.gov

F. DeLaRosa

Senior Member, IEEE

Zenergy Power Inc.

1616 Rollins Road

Burlinggame, CA 94010

USA

francisco.delarosa@zenergypower.com

Figure 1. A typical WPP layout.

POI or 

connection 

to the grid Collector System 

Station

Feeders and Laterals (overhead 

and/or underground)

Individual WTGs

Interconnection 

Transmission Line
2 

II. FAULT CURRENT LIMITER

A. Principle of Operation

The saturable core FCLs [3] consist of a set of coils wound 

around one or more ferromagnetic cores. A high-temperature 

superconducting (HTS) magnet is coupled to the core region 

in such a way that the DC magnetization force can saturate 

the magnetic material. As shown in Figure 2, each AC phase 

consists of two coils connected in series and wound around 

two core regions. Under normal operating conditions, the 

operating point on the Φ-I curve is always in the saturated 

region due to the DC bias, as depicted in Figure 3(a). AC positive flux shifts the operation point further into the saturated 

region due to the magnetic field of one of the AC coils that 

boosts the existing DC bias. The other AC coil has an opposite effect and its magnetic field and bucks the existing DC 

bias. All this happens in the region shaded in green in Figure 

3(a), where the slope of the curve and therefore the resultant 

inductance of the FCL is very small. Under a fault, the operating point on the Φ-I plane is transitioned through the high 

slope, and therefore high inductance regions, as illustrated in 

red in Figure 3(a). 

The corresponding flux changes in each of the coils are 

shown in Figure 3b under normal conditions and in Figure 3c 

under short-curcuit conditions. 

The total voltage E is shown as the thick line in Figures 3d 

and 3e, where the unsymmetrical portions cancel each other. 

Note that there is a 90o phase shift between the generated voltage and the flux changes.

total buck boost += eee

where: 

t

e

∆

∆Φ =

The corresponding voltages generated across the coils are

shown in Figure 3d. As shown in Figure 3d, the generated 

Fig.2. Dual iron cores saturated by an HTS DC-coil in a single-phase FCL, 

adapted from [3]. 

Ε 

O’ 

O” 

i

B”

A” 

O’ 

B’ O” 

A’ 

i

Ε 

2”

O” 

O’ O’ 

O” 

1” 

1’ 2’

d) Generated voltage e) Generated voltage

normal condition SC condition

1” 2’ 

∆φ 

O’ 

O” 

i

B”

A” 

O’ 

O” 

B’ 

A’ 

i

∆φ 

2”

O’ O’ 

O” O” 

1’ 

O’ 

O” 

b) Flux changes c) Flux changes

normal condition SC condition

O’ 

O” 

Figure 3. Operating region for normal and short-circuit conditions.

a) φ-I characteristic of the FCL for buck coil (dotted line) 

and boost coil (dashed line).

 
3 

voltage in the boost coil and buck coil are symmetrical, and

from the series connection of the coils and the way they were 

connected, the total generated voltage (E) across the series 

coils is the sum of both voltages generated across each coil. 

The total voltage E is shown as the thick line. Note that there 

is a 90o phase shift between the generated voltage and the flux 

changes.

In the short-circuit condition, the SCC is very large, thus it 

drives the operating points to operate beyond the linear region 

of the φ−i curves. As shown in Figure 3a, the operating point 

moves along O”1”O”2”O” in the φ−i, for the buck coil and 

similarly, the operating point moves along O’1’O’2’O’ for the 

boost coil where the operation reaches outside the linear region. The corresponding flux changes in each of the coils is

shown in Figure 3c. As shown in Figure 3c, the flux changes 

will be very large for the same amount of current change and 

the voltage drop across the FCL will also be very large.

As shown in Figure 3a, a very small change in the flux 

(∆φnormal) is produced when the system operates under normal 

current because the operating points are in the saturation region of the magnetic circuit. The corresponding voltages 

generated across the coils are shown in Figure 3d. As shown 

in Figure 3d, the generated voltage in the boost coil and the 

buck coil are non symmetrical, however, from the series connection of the coils and the way they were connected, the total 

generated voltage (E) across the series coils is the sum of the 

voltages generated across each coil.

On the other hand, the change in flux (∆φfault) is very large

when a fault current is flowing in the coil. During the fault, 

the operating points are driven out of the saturation region 

into the linear region of the magnetic circuit, producing a 

large back EMF across the AC coils as depicted in Figure 

3(e). 

B. Basic Equations in an FCL

The nonlinear model which describes the behavior of the 

device is based on the physical principle described above. 

The core material has B-H characteristics that can be approximated by an inverse tangent function [3]. We begin by estimating the linkage magnetic field in the core section seen by 

one AC coil as

ac sat

sat

ac iI B

I

K

K

B iB 2] 2 1[ tan )((

)

2

1 tan (

2 )( max

max

1

1

+ +−−

+ −

− = −

−

π π

π π

where iac is the instantaneous AC line current, Imax is the line 

current it takes to fully saturate the cores, at which point the 

average magnetic field is Bsat, and K determines the range of 

line currents where the magnetic state of the cores are actively 

changing from saturated to unsaturated. The equation is 

scaled in such a way that B = Bdc_bias at iac = 0 (bias point, 

corresponding to point O’ or O” in Figure 3a), and the majority of the change in field occurs just before iac = Imax. 

For a value of Bsat = 2.2 Tesla, and Imax=30kA, and two 

different values of K, the differential flux density ∆B(iac)

caused by AC currents is shown in Figure 4.

The effective inductance of an individual AC coil in the 

FCL can be computed as follows: 

ac

ac

ac core i

iB AnL

∂

∂ = )( ~

Where 

~

L is the differential inductance, nac is the number 

of turns in the AC coil that carries the line current, and Acore is 

the cross-sectional area of the ferromagnetic core material. 

Figure 5 shows the plot of the differential inductance for different operating currents for K=2 and K=5. In some cases, 

this model is improved by imposing two additional conditions. First, L must be greater than or equal to an additional 

parameter called Lair when the differential current δiac is 

small. This value represents the insertion impedance as an 

equivalent air-core inductance Lair corresponding to line 

“AO”B” or A’O’B’ in Figure 3a. Furthermore, if δiac is 

greater than Imax, then L moves beyond the saturation region 

into linear magnetization region. This accounts for the fact 

that when the line current is very large, the FCL’s magnetic 

core is reverse saturated and the impedance is once again approximately equal to the insertion impedance of an equivalent 

air-core inductor. Note that the inductance changes significantly before the value of defined Imax. 

Fig. 4. The differential B(AC) for different operating currents.

Fig. 5. The differential inductance for different operating currents.
4 

The induced voltage or back EMF across this section of 

the FCL is 

t

i LV

∂

∂ =

As shown in Figure 6, the voltage drop across the FCL varies depending on the operating point of the FCL. 

The equations above provide a general framework for describing the behavior of an FCL in an electric circuit. This 

framework requires four input parameters: Imax, Bsat, K, and 

Lair. As shown in Figure 4 through Figure 6, the value of K

influences the shape of the curve. The value of K is designed 

based on the dimensions, the characteristics of ferromagnetic 

material used (i.e., non linear relative permeability, μr), and 

the choice of the number of turns used for the AC coil. The 

characteristics can be shaped to satisfy the design criteria.

As described in the following sections, we use finite element method (FEM) simulations to calculate the average 

magnetic flux for various static values of superconducting 

currents and line currents. We then take these flux values and 

use a least-squares fitting procedure to determine the above 

parameters. Finally, we use some form of electrical simulation software (such as PSCAD®) to implement the nonlinear 

inductance model and compare it to experimental results from 

the extensive tests performed on the 15-kV FCL device [3]. 

III. THE EFFECT OF FCLS ON SCC CONTRIBUTION FOR WTG

TYPE 1 – INDUCTION GENERATOR

To investigate the impact of FCLs on SCC contribution, an 

FCL is installed in series with the line at the low voltage side 

of the substation transformer. In Figure 7, the single-line diagram of a typical WPP is shown [4]. With the insertion of the 

FCL, the SCC from all WTGs within the wind plant will be 

subjected to the change in FCL inductance due to the changes 

in current.

First, let us consider the WTG Type 1 (induction generator) used in the WPP investigated. The block diagram of the 

wind turbine Type 1 is shown in Figure 8. 

Fig. 7. FCL inserted at the low voltage side of a 34.5-kV substation transformer. 

Fig.8. Type 1 WTG. 

A. Three Phase Faults

In this section, the aggregated contribution of all WTGs in 

a WPP is considered. The main difference between the characteristic of the three phase fault in a synchronous generator 

and an induction generator is the nature of field excitation. In 

a synchronous generator, the field winding is controlled by 

the field excitation and the field current continues to be 

present during the fault, generating the EMF voltage E driving the fault currents [5-6]. In an induction generator, when 

there is a three phase fault at the terminals, the voltage at the 

stator winding terminals drops to zero and there is no grid 

voltage available to maintain the magnetic flux. The magnetic flux in the air gap eventually dissipates as the fault 

progresses [7-9]. 

Fig.9. Voltage at the substation transformer for a three phase fault. 

Fig. 6. Approximate voltage drop for different operating currents.
5 

Figure 9 shows the voltage at the low voltage side of the 

substation transformer when the three phase fault occurs. As 

shown, the voltage quickly dies out when the three phase fault 

sustains for many cycles, however, when the FCL is inserted 

in series with the line, the voltage dies out in a longer time 

constant.

The fault currents of the two systems are compared. It is 

shown in Figure 10 that a significant current reduction is 

achieved when the FCL is installed. Without the FCL, the 

peak of the fault current reaches 10 kA. With the FCL installed, it can reduce the limit the peak of fault current down 

to 4 kA.

Fig. 10. Fault currents at the substation transformer for a three phase fault. 

B. Single line-to-ground (SLG) Fault 

The single line-to-ground (SLG) fault has significant implications on WTG applications, especially for Type 1 and 

Type 2 machines. Under this type of fault, unequal current 

flows in the three phases and the unfaulted phases are still 

available to excite the generator fluxes. Thus, the system 

within the power plant is operated under unbalanced condition at 60 Hz operation until the fault is removed.

 The FCL is represented as variable impedance in the diagram. Under an SLG fault, the FCL behaves differently 

among the three phases. The larger the current flow in any 

phase, the larger the apparent impedance of the FCL at that 

phase. As a result, it has a balancing impact on the overall 

system. 

Figure 11 shows the system without the FCL and produces 

a larger fault current in the faulted line. The unbalanced series impedance among the three phases with the FCL is such 

that the system tends to have less fault current on the faulted 

line.

From Figure 11, the inductance in the faulted line (phase a) 

appears to be the largest among the three lines. It is also 

shown that due to the variation of the instantaneous inductance, the nonlinear variation due to the magnetic saturation 

creates a less than perfect sinusoidal current during the fault.

The comparison between the two simulations of an SLG

fault with and without FCL in the circuit shows that the presence of the FCL in the circuit will create more balanced currents in the system during a fault event.

Fig.11. Fault currents at the substation transformer for a single line–toground fault.

Figure 12 shows the three sequence currents (positive, negative, and zero sequence currents). The response of the system with and without FCL is shown on the same graphs.

It is shown that for the positive, negative, and zero sequence currents, there is a significant reduction of transient 

currents during the first few cycles during the faults. This 

reduction can significantly impact the transient torque production.

Fig.12. The sequence currents at the substation transformer 

Fig. 13. The transient torque reduction with a FCL installed in the circuit. 
6 

The concept of minimizing unbalanced voltage on induction machines with unsymmetrical series components has 

been proven to be effective [10-11]. The same concept can 

be extended to a WPP by using a FCL at the substation transformer.

Figure 13 shows the impact of a FCL on the transient torque reduction in an induction generator during an SLG fault. 

As shown in the figure, there is a significant reduction in the 

fault current, as well as a significant reduction in the torque 

spike in the induction generator after the FCL is installed in 

series with the circuit.

In this operation, it is important to note that if the operation under unbalanced condition is left unnoticed, it may 

cause serious damage to the induction generator due to the 

unequal heating in the stator windings and the torque pulsations that may excite mechanical modes of the WTGs [12]. 

Although an unbalanced fault occurs for a very short duration, 

nevertheless, the resulting torque spike may damage the gear 

box from frequent faults that may occur in the lines.

IV. THE EFFECT OF FCL FOR DIFFERENT WTG TYPES 

In this section, the SCC contributions of different WTGs 

for faults at the terminal of the generator are discussed. The 

three phase faults for different types of WTGs are short in 

duration because the air gap flux collapses without the support of sufficient terminal voltages. The duration of the fault 

is the longest for a Type 1 WTG.

Type 2 WTG has a lower SCC and a shorter time constant. 

The insertion of external rotor resistance into the rotor 

winding in the circuit reduces the SCC and the duration of the 

SCC. 

Similarly, for a Type 3 WTG, the SCC and the duration of 

the three-phase-to-ground fault is shorter than for a Type 1 

WTG. In a Type 3 WTG, there is a crowbar circuit that is 

used to protect the DC bus and the power switches in the 

power converter. The crowbar, in effect, adds adjustable effective rotor resistance and can even short circuit the rotor 

winding.

For a Type 4 wind turbine, the generator and grid is separated by a power converter. The power converter is limited 

by the current carrying capability of the power semiconductor 

used. Thus, the SCC contribution from a Type 4 wind turbine 

is limited to 110% of rated current (assuming temporary overload current is designed to be 110% rated). The SCCs from 

the WTG can be maintained in balance by controlling the 

power converter.

The SLG fault is the most likely to occur in the power grid. 

The terminal voltage and currents are sustained longer because the line voltages, except from one phase, are able to 

sustain air gap flux. The SC continues to flow until the fault 

is cleared.

The FCL is also installed in Type 2 and Type 3 WTGs and 

the reduction in SCCs and the torque transient are compared 

between the system without FCL and the system with the 

FCL. The comparison in SCC reductions are tabulated in Table I, and the comparison in the transient torque reductions 

are tabulated in Table II. The values listed are the ratio of the 

values for systems with FCLs with respect to the values for 

systems without FCLs. The baseline used is the system without FCL, thus, values below 1.0 per unit indicate the reduction for the system with an FCL.

TABLE I 

THE RATIO OF THE SCC WITH AND W/O FCL

FOR DIFFERENT TYPES OF WTGS

TYPE 1 TYPE 2 TYPE 3

3LG 0.44 0.5 0.99

SLG 0.33 0.34 0.75

LL 0.47 0.38 0.71

LLG 0.33 0.36 0.70

TABLE II

THE RATIO OF THE TRANSIENT TORQUES WITH AND W/O FCL 

FOR DIFFERENT TYPES OF WTGS

TYPE 1 TYPE 2 TYPE 3

3LG 0.63 0.62 0.99

SLG 0.55 0.55 0.71

LL 0.66 0.63 0.77

LLG 0.66 0.65 0.67

V. CONCLUSIONS

In this paper, the role of FCLs is described in detail for 

Type 1 WTG. From section III, it is shown that the FCL can 

effectively reduce the SCC for the Type I wind turbine generator. It is also shown that for single line-to-ground and other types of unsymmetrical faults, the FCL behaves like an 

equalizer such that not only the reduction in the positive, negative, and zero sequence currents are achieved, but also the 

timing of the reduction is very beneficial. The most significant reduction of those sequence currents occurs during the 

transient (at the few cycles in the beginning of the faults), 

thus, as a result, the transient torque is reduced very significantly.

In section IV, the SCC contribution for different types of 

WTGs is tabulated as the ratio of SCC for a WPP with FCLs 

to the SCC for a WPP without FCLs. A table showing the 

reduction in transient torque for different faults and different 

WTGs is also tabulated in the same fashion. As shown in

both tables, there is significant improvement in the SCC and 

the transient torque reductions for a WPP with FCLs. 

The designs of the FCL for different types of WTGs must 

be designed specific to the WTG technology because the nature of the SCC contribution is unique for each type of WTG. 
7 

VI. ACKNOWLEDGMENT

This work is supported by the U.S. Department of Energy 

and California Energy Commission.

VII. REFERENCES

[1] U.S. Department of Energy – Energy Efficiency and Renewable Energy, “20% Wind Energy by 2030 – Increasing Wind Energy’s Contribution to U.S. Electricity Supply,” May, 2008. 

[2] J. Charles Smith, Michael R. Milligan, Edgar A. DeMeo and Brian 

Parsons, "Utility wind Integration and operating impact state of the art," 

IEEE Trans. Power Systems, vol. 22, pp. 900-908, Aug. 2007.

[3] Moriconi, F., Koshnick, N., De La Rosa, F. and Singh, A., “Modeling 

and Test Valiadation of a 15 kV 24 MVA Superconducting Fault Current Limiter”, in Proc. of IEEE/PES Transmission and Distribution

Conference, New Orleans, USA, April 2010.

[4] WECC Modeling and Validation Work Group, “WECC Wind Power 

Plant Power Flow Modeling Guide Prepared.” May 2008.

[5] Short-circuit currents in three-phase a.c. systems–Part 0: Calculation of 

currents, IEC Standard 60909-0, July 2001.

[6] IEEE Application Guide for AC High-Voltage Circuit Breakers Rated 

on a Symmetrical Current Basis, IEEE Standard C37.10-2000, Sept. 

2000.

[7] J. Moren, S.W.H. de Haan, “Short-Circuit Current of Wind Turbines 

with Doubly Fed Induction Generator,” IEEE Transactions on Energy 

Conversion, Vol. 22, No. 1, March 2007.

[8] IEEE PES Wind Plant Collector System Design Working Group, ”Power Transformer Application for Wind Plant Substation,” Proceedings of 

the 2009 IEEE Power and Energy Society General Meeting.

[9] V. Gevorgian, E. Muljadi, " Wind Power Plant Short-Circuit Current 

Contribution for Different Fault and Wind Turbine Topologies", Presented at the 9th International Workshop on Large Scale of Wind Power 

into Power Systems, Quebec City, Quebec, Canada, July 18-19, 2010 

[10] E. Muljadi, R. Schiferl, and T.A. Lipo, "Unsymmetrical Triggering of 

Phase Back Controller for Unbalanced Operation of Induction Machine", IEEE Transactions on Industry Applications, Vol. IA-21, No. 4, 

May/June 1985, pp. 669- 677 

[11] J. Oyama, F. Profumo, E. Muljadi and T.A. Lipo, " Design and Performance of a Digitally Based Voltage Controller for Correcting Phase 

Unbalance in Induction Machines." IEEE Transactions on Industry Applications Vol. 26, No. 3, May/June 1990, pp. 425-433. 

[12] E. Muljadi, T.Batan, D.Yildirim, C.P. Butterfield, "Understanding the 

Unbalanced-Voltage Problem in Wind Turbine Generation," IEEEIndustry Applications Society Annual Conference, October 3-7, 1999 in 

Phoenix, Arizona.

VIII. BIOGRAPHIES

Eduard Muljadi (M’82-SM’94-F’10) received his Ph. D. (in Electrical 

Engineering) from the University of Wisconsin, Madison. From 1988 to 

1992, he taught at California State University, Fresno, CA. In June 1992, he 

joined the National Renewable Energy Laboratory in Golden, Colorado. His 

current research interests are in the fields of electric machines, power electronics, and power systems in general with emphasis on renewable energy 

applications. He is member of Eta Kappa Nu, Sigma Xi and a Fellow of the 

IEEE. He is involved in the activities of the IEEE Industry Application Society (IAS), Power Electronics Society, and Power and Energy Society 

(PES).

He is currently a member of various committees of the IAS, and a member 

of the Working Group on Renewable Technologies and Dynamic Performance Wind Generation Task Force of the PES. He holds two patents in 

power conversion for renewable energy.

Vahan Gevorgian (M’97) graduated from the Yerevan Polytechnic Institute (Armenia) in 1986. During his studies, he concentrated on electrical 

machines. His thesis research dealt with doubly-fed induction generators for 

stand-alone power Systems. He obtained his Ph.D. degree in electrical engineering from the State Engineering University of Armenia in 1993. His dissertation was devoted to a modeling of electrical transients in large wind 

turbine generators.

Dr. Gevorgian is currently working at the National Wind Technology

Center (NWTC) of National Renewable Energy Laboratory (NREL) in Golden, Colorado, USA, as a research engineer. His current interests include 

modeling and testing of various applications of small wind turbine based 

power systems.

Francisco DeLaRosa (M'91, SM'00) joined Zenergy Power Inc. in April 

2008 as Director of Electrical Engineering. Before joining Zenergy Power, 

Inc., Francisco had held various positions in R&D, consultancy and training 

in the electric power industry for around 30 years. Francisco holds a PhD 

degree in Electrical Engineering from Uppsala University, Sweden. He is a 

Collective Member of CIGRE and a Senior Member of IEEE PES where he 

contributes in several WG’s including the TF on FCL Testing. His main 

interests include the assessment and integration of new technologies in the 

electric power system.
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Wind Power Plant Enhancement with a Fault-Current Limiter: 

Preprint

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DE-AC36-08GO28308

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V. Gevorgian, E. Muljadi, and F. DeLaRosa

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NREL

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14. ABSTRACT (Maximum 200 Words)

This paper investigates the capability of a saturable core fault-current limiter to limit the short circuit current of 

different types of wind turbine generators. Different faults are simulated to investigate the effectiveness of the FCL to 

limit the SCC and to reduce transient torque during faults. Several cases will be considered to demonstrate the 

benefits of using FCLs in unique situations. 

15. SUBJECT TERMS

Wind power plant; short circuit current; fault current; saturable core; voltage fault

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