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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 32, NO. 1, FEBRUARY 2017
A Non-unit Protection Scheme for DC Microgrid
Based on Local Measurements
A. Meghwani, S. C. Srivastava, Senior Member, IEEE, and S. Chakrabarti, Senior Member, IEEE
Abstract—This paper proposes a nonunit protection scheme for
DC microgrids (DC MGs) utilizing only local measurements. This
scheme is developed based on the natural characteristics of DC
current and its first and second derivatives under fault transients.
Since it is based on local measurements, the problems associated
with the communication delay are avoided. The selected protection scheme detects and discriminates the faults within a few
microseconds of its inception. In this paper, a method to calculate the thresholds, used for the protection scheme, is also discussed. The proposed scheme is validated on a ring-type DC MG
architecture under different fault scenarios and tested through
MATLAB/Simulink simulations.
Index Terms—Current derivatives, DC microgrid, local measurements, nonunit protection.
I. INTRODUCTION
N the recent past, significant research and development
efforts have been made to integrate renewable energy resources, such as wind turbines, and photovoltaic systems into
the power distribution networks [1], [2] forming an important
part of smart grid. The energy policy in many countries across
the world envisages increased penetration of the renewable energy resources and Distributed Generations (DGs). For instance,
India is trying to increase the usage of renewable generation up
to 30% by 2030 [3].
A low or medium voltage electrical network, consisting of
distributed resources, especially renewable sources of energy,
storage devices, and loads, is known as a Microgrid [4]. The
electrical network can be AC, DC, or a mix of the two, and may
or may not be connected to the main AC grid. During normal
operating conditions, the microgrid is connected to the AC grid
at the point of common coupling. The loads are supplied from
the local sources and, if necessary, also from the AC grid. If the
power consumed by the loads is less than the power produced
by the local sources, the excess power can be exported to the AC
grid. A DC MG is more suitable where the majority of the loads
contains sensitive electronic equipments. The advantage of the
DC MG compared to an AC microgrid is that loads, sources,
I
Manuscript received December 21, 2015; revised March 10, 2016; accepted
April 5, 2016. Date of publication April 21, 2016; date of current version
January 20, 2017. This work was supported by the Department of Science and
Technology, New Delhi, India, under project no. DST/EE/20100258. Paper no.
TPWRD-01843-2015.
The authors are with the Department of Electrical Engineering, Indian
Institute of Technology, Kanpur 208016, India (e-mail: anjum@iitk.ac.in;
scs@iitk.ac.in; saikatc@iitk.ac.in).
Color versions of one or more of the figures in this paper are available online
at http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TPWRD.2016.2555844
and energy storage can be connected through simpler and more
efficient power-electronic interfaces [5]. Other advantages are,
high efficiency [6], ease of paralleling of sources on the DC bus
[7], and more power transfer capacity [8].
One of the main challenges in adopting the DC distribution
system is the lack of effective protection to the fault in DC
network. The faults in the DC MG can be categorized as:
1) AC side faults: Voltage Source Converters (VSCs) are
widely used as AC-DC or DC-AC converters for electrical power conversion. The protection against AC faults
on grid connected converters forms a part of AC system
protection design [9].
2) Internal faults: VSC internal faults include failure of
power electronics device such as Insulated-Gate Bipolar Transistor (IGBT). In terms of fault tolerant VSCs, the
research efforts have been to protect the system from possible faults by providing a backup converter or redundant
devices [10].
3) DC network faults: A DC network fault, with parallel connected VSCs, is the most severe fault in DC MG. It causes
over current and under voltage due to the presence of
a large filter capacitor of the VSC and low impedance
offered by the cable. The freewheel diodes of VSCs
are subject to over currents and are unable to provide
protection against DC side faults, such as DC link short
circuit, DC cable short circuit, and DC cable ground faults.
These conditions need to be analysed in detail in order to
design an effective protection system.
A line to ground fault is the most common type of fault
in the DC network [11], [12]. The causes can be the physical
damage, environmental stresses, electrical stresses, and cable
ageing. A scheme based on differential current protection was
proposed in [13], [14] to detect the cable fault in DC network.
It requires a reliable communication channel for instantaneous
data transfer between the terminals of the protected element.
Because of the chances of possible communication failure or
data loss, differential protection will require a separate backup
protection scheme. This may increase the total cost and size of
the protection system, and limits its application in microgrids.
Along with poor reliability, usage of communication in protection, also introduces delays, which increases the response
time of the protection scheme. Researchers in [15] proposed a
nonunit protection scheme, which is highly reliable and fast in
operation. In this scheme, as the fault is detected, all the VSCs
are disconnected from the AC side. The capacitors on the DC
side support the load for short duration. The only limitation of
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MEGHWANI et al.: NONUNIT PROTECTION SCHEME FOR DC MICROGRID BASED ON LOCAL MEASUREMENTS
173
equivalent circuit shown in Fig. 1(b). The RLC circuit response
in the frequency domain during this period can be expressed as
[19],
i (s) =
Fig. 1. (a) Faulted network, (b) Its RLC equivalent circuit, and (c) Equivalent
circuit for stage 2.
this method is that it needs a complete shutdown of the system
following the fault. Authors in [16], [17] investigated a range of
unit and nonunit protection solutions for DC distribution systems. It is found that differential protection is the most suitable
for such networks due to its high selectivity and speed of operation. The most stringent requirement for a DC protection system
is its speed of operation. This is due to the presence of the high
rate of rise of fault current and its large steady state value. The
fault must be cleared in a timely manner to avoid damage to
the sensitive power electronic devices and also to keep the fault
current below the interruptible current limit of the breakers.
The goal of the proposed scheme is to detect the fault in a
DC cable and quickly isolate the faulted section. The scheme
is based on locally measured system parameters that do not require any communication channel or system shutdown to clear
the fault. This cause almost no time delay, as the measurements
are local, except for the processing time. The proposed protection scheme is simple and fast to protect the DC MG from
the faults, without interrupting the load. The proposed fault
detection algorithm utilizes fault current, and its first and second
order derivative to discriminate the fault. Based on these parameters, a protection system design framework is developed. The
proposed scheme is tested on a typical DC MG architecture.
II. DC-CABLE GROUND FAULT ANALYSIS
A DC short-circuit fault is the most severe condition for VSCs
[18]. The IGBTs can be blocked for self-protection during faults,
leaving reverse diodes exposed to over currents. Irrespective of
the location of a short circuit on the DC cable, it can be expressed
by an equivalent circuit shown in Fig. 1(a). A short circuit in the
DC cable may result into high fault current due to the presence
of charged capacitors and low impedance offered by the cable.
To find the complete response of the nonlinear circuit, the fault
is analysed in two different stages. Stage one is the natural
response of the RLC circuit and stage two starts when fault
current reaches to peak and capacitor voltage becomes less than
the peak of the input voltage. Further analyses of both the stages
are discussed in the following sections.
A. Stage 1: Capacitor Discharge
Immediately after a line to ground fault, the capacitor starts
discharging through a cable impedance, as depicted in the
vC (0) /L + iL (0) s
1
s2 + R
L s + LC
(1)
where, iL (0) and vC (0) are the current through the inductor
and voltage across the capacitor, respectively, just before the
occurrence of the fault. L and r are the equivalent series inductance and resistance of the cable up to the fault point. R is
the sum of fault resistance RF and r. In this expression, the
converter current contribution to the fault is assumed to be negligible because of the slow response of the converter controllers
[20]. In the time domain, the fault current i (t) can be written
as,
i (t) =
vC (0) −p 1 t
e
− e−p 2 t
L (p2 − p1 )
+
iL (0) −p1 e−p 1 t + p2 e−p 2 t
p2 − p1
where, p1 and p2 are the poles of (1), and are given by,
2
R
R
1
±
−
p1 , p2 =
2L
2L
LC
p1 , p2 = α ± α2 − ω02
(2)
(3)
(4)
Depending upon whether 1/LC is greater than, less than,
or equal to (R/2L)2 , the values of p1 and p2 will be real or
complex. As a result, the fault current can be under damped, over
damped or critically damped. The time taken for the current to
reach peak magnitude for under damped (tu dp ) and over damped
(todp ) circuit conditions can be written as [13],
ω
1
0
tu dp =
tan−1
(5)
ω0
α
todp =
ln (p2 /p1 )
p1 − p2
(6)
By substituting typical values of the cable parameters and
filter capacitor (r, L, and C; refer to Table A1) the time to reach
the peak current, during the fault has been calculated for the two
extreme cases of faults, i.e., for direct short circuit (RF = 0 Ω)
and for high impedance fault (RF = 0.6 Ω). It is found that
the processing time allowed to detect the fault current before
the load voltage collapses in the DC system is much shorter as
compared to the AC system.
B. Stage 2: Diode Freewheeling
As the fault occurs, the capacitor starts discharging with
high fault transients and its rate of rise is limited by the cable impedance. The second stage appears when the voltage of
the DC bus drops to zero and becomes negative, i.e., after todp or
tu dp . This has the effect on reversing the voltage at the converter
terminals and allows the freewheel diodes to conduct, as shown
in Fig. 1(c). In this figure, D represents the equivalent of the,
diodes conducting in any converter leg [21]. This provides an alternative current path irrespective of the IGBT conducting state
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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 32, NO. 1, FEBRUARY 2017
and changes the current response. In this phase, the freewheel
diodes experience almost ten times the nominal current value
which rises rapidly. This work tries to detect the fault on the DC
network before the system enters into the second stage, so that
the sensitive power electronic converters and its components
can be protected from the high fault current.
III. DC FAULT CURRENT CHARACTERISTICS
In (2), the fault current i (t) depends upon vC (0) which is
fixed for a given DC MG and the line current (iL ) just before
the fault. The time derivative of (2) can be written as:
di (t)
vC (0) =
−p1 e−p 1 t + p2 e−p 2 t
dt
L (p2 − p1 )
+
iL (0) 2 −p 1 t
p1 e
− p22 e−p 2 t
p2 − p1
(7)
Immediately after the fault, at t = 0+ , the magnitude of the
current derivative is given by,
vC (0)
di (t)
=
− iL (0) (p1 + p2 )
(8)
dt
L
Substituting the values of p1 and p2 from (3), (8) can be
written as,
vC (0)
di (t)
R
=
− iL (0)
dt
L
L
(9)
For RF = 0, R is equal to the equivalent cable resistance up
to the fault point, denoted by r. For direct short circuit, the
contribution of the iL (0) component, compared to vC (0) to the
di/dt magnitude will be very small and is neglected. Hence, only
vC (0) component may be considered in deciding the primary
threshold for the protection.
The second order derivative of the line current can be obtained
by differentiating (7) with respect to time, and can be written
as,
d2 i (t)
vC (0) 2 −p 1 t
p1 e
=
− p22 e−p 2 t
2
dt
L (p2 − p1 )
+
iL (0) 3 −p 1 t
−p1 e
+ p32 e−p 2 t
p2 − p1
(10)
Immediately after the fault at t = 0+ , the magnitude of the
second derivative can be given by the following equation.
d2 i (t)
vC (0)
(p1 + p2 ) + iL (0) p21 + p22 + p1 p2
=−
2
dt
L
(11)
By substituting the values of p1 and p2 in terms of circuit
parameters, (11) can be written as,
2
R
d2 i (t)
vC (0) R
1
=
−
+
i
(0)
−
(12)
L
dt2
L2
L2
LC
The second derivative magnitude depends upon the equivalent
resistance R = RF + r. The initial line current
contribution to
d2 i/dt2 will be zero for R = Rc , where Rc = L/C. For small
range of RF , i.e., 0 < RF < Rc , the equivalent circuit of the
network during the fault will be under damped. This results in
Fig. 2. Fault current, and its first and second derivative profile for different
values of R F .
high magnitude of the first order current derivative, which is
much larger than its second order derivative.
The test system in Fig. 1(a) is simulated for different values
of RF to observe the current i, di/dt, and d2 i/dt2 profile for
under and over damped circuit conditions. In this work, the
fault is created at t = 0.2 s, and di/dt, d2 i/dt2 are calculated
at t = 0.2001 s, as shown in Fig. 2. It is observed that both the
derivatives are equal in magnitude because of the insufficient
number of data samples to calculate the second order derivative
at the time of fault. This is further explained in Section IV.B.
The d2 i/dt2 is calculated at time t = 0.2002 s, and found that,
as RF increases, d2 i/dt2 magnitude approaches to di/dt value.
The initial line current may have more impact on di/dt and
d2 i/dt2 in the presence of RF .
During normal operation, transients such as step load change
and operating mode change, the calculated first and second order
derivatives should not cross the threshold. Line current i12 and
their derivatives are monitored to see the effect of step load
transient and mode switching transient, as shown in Fig. 3. It
is observed that the calculated derivatives are much smaller
than their respective thresholds, and hence, unwanted tripping
is avoided.
IV. PROTECTION SYSTEM COMPONENTS
The protection scheme for DC MG is analytically developed
in previous sections. The protection system is designed and
developed for DC MG of Fig. A1 (refer to the Appendix) using
various components. Detailed description of these components
and their functionality are provided in the following sections.
A. Protective Device (PD)
A PD, considered in this paper, consists of a digital relay, a
solid state circuit breaker, and a current transducer, as shown
MEGHWANI et al.: NONUNIT PROTECTION SCHEME FOR DC MICROGRID BASED ON LOCAL MEASUREMENTS
175
TABLE I
TIME TO REACH PEAK CURRENT (tu d p ) SEEN BY PDS FOR DIFFERENT FAULTS
IN ms
Fault
F1
F2
F3
F4
F5
tP D 1 . 2
tP D 2 . 2
tP D 3 . 2
tP D 4 . 2
tP D 5 . 2
1.7
2.8
3.1
4.4
4.0
2.2
1.1
2.2
3.3
4.2
3
2.3
1.3
3.4
3.5
4.1
3.0
3.3
3.5
4.0
5
4.5
4.8
3.3
2.4
B. Parameter Determination
To protect the system from line to ground fault, the protection
algorithm is implemented in all P Ds. The current derivative and
its rate of change are numerically calculated as discussed in the
following sections.
1) di/dt Calculation: di/dt is calculated using backward
finite difference approximation [25], and can be written as,
Δi
i (t0 + Δt) − i (t0 )
di
= lim
=
Δ
t→0
dt
Δt
Δt
Fig. 3. Line current i1 2 , and its first and second derivative profile under normal
transients.
Fig. 4.
Protective Device and its components.
in Fig. 4. DC current can be measured by a current transducer
such as Hall effect devices. The sensor output is a voltage signal which facilitates an easier integration with the digital signal
processing devices. Digital relays are equipped with microcontroller for setting the thresholds, and analog to digital converters
to transform the measurements into digital form. The current is
sampled and its rate of change is calculated and compared with
the threshold setting to generate a trip signal. The DC cables
are protected by PDs connected on both the ends as shown in
Fig. A1.
The faulted network in the DC system should be isolated, as
fast as possible, to avoid shutdown. Reference [19] compared
different DC CB technologies to find their suitability in the
DC shipboard. The authors compared solid state CB (SSCB),
Hybrid CB, and electromagnetic CB [22]–[24] on the basis of
their operating time for unmanned aerial vehicle application.
Comparison between the circuit breaker operating times suggests that SSCBs are the most suited for use within DC networks. Therefore, in this paper 50 μs operating time (typical
for SSCB) of circuit breaker is considered. There is no additional delay in the operation because of utilization of the local
measurements.
(13)
and Δi can be written as,
Δi = ik − ik −1
(14)
where, k is the sampling instant, and ik , and ik −1 are the present
and the the previous sampled line current. Since the sampling
time is same for all the measurements, henceforth only current difference is measured at consecutive sampling instants to
calculate the time derivative.
2) d2 i/dt2 Calculation: d2 i/dt2 is calculated for each line
using the present and the previous calculated value of di/dt at
time interval dt, and can be written as:
d2 i
d di
d Δi
Δ2 i
=
(15)
=
=
2
dt
dt dt
dt Δt
(Δt)2
where, Δ2 i can be written as,
Δ2 i = Δik − Δik −1
2
Δ i = ik − 2ik −1 + ik −2
(16)
(17)
The fault occurs at k = 0 and k is increased by one at every
instant of sampling. At k = 1, Δi = Δ2 i because i0 = i−1 =
iL (0). As a result, the actual value of Δ2 i is calculated at
k = 2 and utilized for tripping decision. The calculated Δ2 i
is compared with the respective threshold setting. Once the
calculated value exceeds the threshold level, PD will issue a
trip signal.
3) Sampling Frequency: The first and second order current
differences are likely to be maximum for under damped circuit
condition. To capture the current profile accurately under this
condition, the time to reach the peak current is calculated for
different fault locations. Based on the calculated time the sampling frequency is decided. The current transient during fault
depends upon the network parameter and fault location. The
time required to reach the current at peak value is analytically
calculated for faults at different locations and listed in Table I.
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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 32, NO. 1, FEBRUARY 2017
From Table I, the minimum time to reach the peak of current
is observed as 1.1 ms for fault F2 . To capture the high di/dt
current profile accurately, approximately 1/10th of the time is
taken as sampling period [26].
V. THRESHOLD SETTING CALCULATIONS
The line current is sampled using high bandwidth current
transducers, and converts it into digital form using analog to
digital converters embedded into the digital relay. Other mathematical processing, such as Δi and Δ2 i pickup threshold settings, their comparison with the pre-set values, are implemented
in the digital relay. Δi and Δ2 i threshold setting calculations
are, further, discussed in the following sections.
A. Threshold Settings for Δi
The Δi settings are, further, classified as fixed threshold and
adaptive threshold. Fixed pickup threshold, denoted by Δim ax ,
is calculated once for a given grid configuration. The adaptive
threshold (Δim in ) depends upon the operating condition of the
microgrid.
1) Fixed Threshold: The magnitude of Δi depends upon the
fault location and its impedance. The threshold of Δi, will be
calculated for the direct short circuit fault that occurs at the end
of the line. The equivalent circuit considered in this case will be
under damped for 0 ≤ RF ≤ Rc . During a fault, the expression
for Δi for the time interval Δt can be derived from (9) and
written as,
Δim ax =
vC (0)
Δt
Leq
(18)
where, Leq is the equivalent inductance up to the fault F1 (refer
Fig. A1 in the Appendix). As a fault occurs, the filter capacitor
connected to the VSC starts discharging with di/dt limited by
Leq , and is given by,
Leq = Lc1 +
(Lc5 + Lx ) Lx
Lc5 + 2Lx
(19)
where, Lc5 and Lx are the inductance of the cable from Bus 1 to
Bus 5 and VSC to Bus 1, respectively, as shown in Fig. 5(a). The
fixed threshold settings are calculated based on the equivalent
circuit considered from the P D1.2 up to the fault point. In order
to validate the equivalent model and its parameters in terms of
the frequency response and rise time, the Bode plot of impedance
is drawn for the complete system and its equivalent model as
shown in Fig. 5(b).
The fixed pickup threshold, Δim ax , for P D1.2 is determined by considering the fault F1 , close to P D2.1 , i.e., at
100% of the line inductance (Lc1 ). If fault occurs close to the
P D1.2 , Δi magnitude will be maximum. On the similar lines,
the fixed thresholds are calculated for P Dx.2 and P Dx.1 , where
x {1, 2, ..5} and tabulated in Table II. These settings depend
upon the cable parameter and the DC MG voltage, and are independent of the operating condition of the microgrid [27].
2) Adaptive Threshold: For high impedance fault, the line
loading also plays an important role in deciding the Δi magnitude. To calculate the threshold Δim in for a fault, the worst
Fig. 5.
plots.
(a) Faulted network, its RLC equivalent circuit and, (b) their Bode
TABLE II
FIXED THRESHOLD (Δim a x ) SETTINGS FOR ALL PDS
Bus
No.
x
1
2
3
4
5
Cable Length
P D x . 2 and
P D ( x + 1 ) . 1 (m)
Δ i m a x Fixed
Threshold (Amp)
for P D x . 1
Δ i m a x Fixed
Threshold (Amp)
for P D x . 2
1000
500
200
300
1000
1100
2000
4000
3000
1000
-2000
-4000
-3000
-1100
-1000
case fault impedance of 0.6 Ω is considered. The value of Δi
is calculated for different line loading using (9). Since, the fault
resistance, RF r, the value of Δim in can be calculated as,
Δim in =
vC (0)
0.6
− iL (0)
Δt
Leq
Leq
(20)
The equivalent inductance up to the fault point is calculated
using (20). The value of Δim in depends upon the line current
iL (0) at the time of the fault. The threshold setting for Δi keeps
on varying with the operating condition of the grid. Hence,
the setting for Δim in is self adjusted with different modes of
operation of the microgrid.
B. Issues With the Δi Protection Scheme
For a given DC MG architecture and rating, the change in
fault current Δi depends upon the following parameters:
1) Cable Length: For a direct line to ground fault, the change
in fault current, i.e., Δi, immediately after the fault, depends upon the cable inductance. High current transients
are observed if the fault occurs near the source, i.e., with
small cable length, as shown in Fig. 6(a).
MEGHWANI et al.: NONUNIT PROTECTION SCHEME FOR DC MICROGRID BASED ON LOCAL MEASUREMENTS
177
Fig. 7. Δi vs Δ 2 i for fault F1 at different locations and resistances 0 <
R F < 0.6 Ω.
Fig. 6. Δi1 2 for fault F1 (a) at distance d from P D 1 . 2 (b) for different line
current flows and (c) for different fault resistance R F .
2) Line Loading: Change in the line current also depends
upon the initial value of the line current before the occurrence of a fault. For heavily loaded lines, change in the
fault current will be less as compared to the lightly loaded
lines, as shown in Fig. 6(b).
3) Fault Impedance: Fault impedance introduces a damping
in the current transients, which reduces its rate of rise and
peak value. As a result, Δi reduces as the fault impedance
increases, as shown in Fig. 6(c).
The threshold of Δi setting is calculated for each cable, considering 100% length and 0.6 Ω fault impedance. For small cable
with high impedance fault, i.e., at F1 , the relay P D1.2 issues a
trip signal. The same relay may operate for the fault in longer
cable with small fault impedance, i.e., at F2 . P D1.2 operate for
fault F2 causes an unwanted line outage, which results into poor
power quality. This is a selectivity issue and it becomes more
pronounced as the cable length reduces. To solve this problem,
threshold based on Δ2 i is applied.
C. Adaptive Threshold Settings for Δ2 im in
Similar to Δi calculation, Δ2 i is also derived analytically. The
value of Δ2 im in is calculated for the worst case RF value. The
target is to detect and isolate the fault in the grid before initiation
of the second stage as explained in Section III.B. Any protection
system must locate the fault in DC MG in a timely manner so
that the circuit breaker operates before the current reaches its
peak value. By substituting R = 0.6 Ω in (12), Δ2 im in can be
calculated using:
Δ2 im in = −
vC (0) 0.6
0.36
+ iL (0) 2 Δt2
2
Leq
Leq
(21)
Since, RF Rc , the initial line current effect cannot be
neglected. In this mode, the current is over damped because
(R/Leq )2 1/Leq C and, hence, iL (0) will also contribute
to Δ2 im in . Since the line loading depends upon the operating
mode of the DC MG, the calculation of Δ2 im in should be self
Fig. 8.
Flow chart of the proposed protection scheme.
adjustable. For a given operating condition, Δ2 im in has been
calculated and shown in Fig. 7. For the under damped circuit
conditions, where r < RF < Rc , Δ2 i will be less than Δ2 im in .
VI. FAULT DETECTION ALGORITHM
The currents are continuously sampled and monitored by their
respective PDs connected on both sides of the cable. Under
steady state, the calculated value of Δi and Δ2 i will be zero.
During fault transients, the calculated current difference will
exceed the threshold settings. Fig. 8 shows the flow diagram of
the algorithm used in this paper.
According to the flowchart, for faults with 0 ≤ RF ≤ Rc ,
the calculated Δi will be greater than Δim ax . This causes
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Fig. 9.
IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 32, NO. 1, FEBRUARY 2017
Impact of the short circuit at F1 in Fig. A1.
Fig. 10.
Short circuit at F1 cleared by P D 1 . 2 and P D 2 . 1 .
Fig. 11.
High impedance fault F1 cleared by P D 1 . 2 and P D 2 . 1 .
initiation of trip signal to the breaker, which will isolate the
faulted section of the network. In case of line to ground faults
with Rc ≤ RF ≤ 0.6 Ω, the calculated Δi will lie within the
range of Δim in < Δi < Δim ax . In this case, the decision for
tripping will be transferred to Δ2 i based threshold setting. Calculated Δ2 i will be greater than Δ2 im in for high impedance
faults, where RF lies between Rc and 0.6 Ω.
VII. SIMULATION RESULTS AND OBSERVATIONS
To demonstrate the proposed protection scheme, fault F1
shown in Fig. A1 is simulated at t = 1.5 s. High transient currents as shown in Fig. 9, cause a voltage collapse within 5 ms
of the fault occurrence. This implies that the protection required
for the DC system should be much faster as compared to the AC
system [7].
A. Case I: Fault in Grid Connected Mode
1) Direct Short Circuit Fault: The magnitude of Δi12 is
continuously calculated from sampled current, and compared
with the threshold given in Table II. As the fault F1 occurs,
calculated Δi12 exceeds the pre-set value and generates a trip
signal for their respective PDs. P D1.2 and P D2.1 to isolate
the faulted line, thereby, restoring the system, as shown in
Fig. 10.
The direct line to ground fault is detected within 100 μs,
which ensures that the system does not enter into the second
stage of fault. In this short time span, the load bus voltage
reduces to 0.95 kV and, hence, VSC will not be affected by
the fault. Large filter capacitor discharge through a small RF
causes the line current i12 to reach to 5 pu before the activation of the PD. As a result, the converters and other system
components are protected. The DC bus voltage momentarily
drops to 0.78 pu because of high di/dt during the fault, and
restores back to 1 pu as the fault is cleared.
To further test the reliability and consistency of the scheme,
faults at different locations are simulated, viz., at F3 and F4 , as
shown in Figs. 12 and 13, respectively.
MEGHWANI et al.: NONUNIT PROTECTION SCHEME FOR DC MICROGRID BASED ON LOCAL MEASUREMENTS
Fig. 12.
Fault F3 in grid connected mode.
Fig. 14.
179
High impedance R F = 10 Ω fault F1 .
detect fault is inversely proportional to the fault current. But in
the proposed algorithm, only two operating time are possible:
when only Δi is required to detect the fault and, when both Δi
and Δ2 i are utilized in detecting the fault. In Fig. 11, the fault
F1 is cleared by P D1.2 and P D2.1 . The line current i12 reaches
0.1 kA before the fault clearance due to the low value of the
di/dt.
In this paper, a typical value of RF = 0.6 Ω is considered and
the protection system is designed. The scheme is also simulated
for high ground fault resistance RF = 10 Ω, and the result is
shown in Fig. 14.
B. Case II: Fault in Islanded Mode
Fig. 13.
Fault F4 in grid connected mode.
2) Fault With Rf : The system is tested for the line to ground
fault F1 via fault resistance RF = 0.6 Ω. Other operating conditions and fault location remain same as in the previous case. In
this case, Δi12 < Δim ax and, therefore, Δ2 i is calculated and
compared with its threshold. The fault is detected within 200 μs
because of the requirement of the second calculation and its
comparison. In this scheme, the time to detect the fault does
not depend on the fault impedance, which is typically present
in the traditional AC protection. In AC protection, the time to
To further test the robustness of the proposed algorithm, the
fault F1 is simulated in islanded mode of operation. In this
mode, AC grid is disconnected from the rest of the network and
is not able to supply the DC MG load. As a result, the load
voltage drops to 1.15 kV, or 0.95 pu. Since the AC grid is not
supplying power to the system, current i12 will be same as i23
as shown in Fig. 15. As the fault F1 is cleared, both the line
currents i12 and i23 become zero with no support from the grid.
A direct short circuit fault causes a high di/dt, which results in
the load voltage drop to 1 kV, with the same operating conditions. The impact of the fault F1 with RF = 0.6 Ω, is shown in
Fig. 16.
C. Case III: Measurement Noise
The proposed protection scheme is further tested in the noisy
environment. The noise can be introduced in the signal due to
measuring device or due to the usage of communication channel.
Since, the measurements and their derivatives are processed
180
IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 32, NO. 1, FEBRUARY 2017
Fig. 17.
Fig. 15.
Fault F1 detected under noisy environment.
Direct short circuit fault F1 in islanded mode.
derivatives are much below their pick-up threshold setting and,
hence, cause no unwanted tripping. For more noise in the signal,
the de-noising methods can be utilized before evaluating their
derivatives [29].
VIII. CONCLUSION
An effective protection system scheme for a DC MG has
been proposed in this paper. The natural characteristics of
the fault current have been divided into under damped and
over damped categories, and used to develop the protection
algorithm. There is practically no time delay involved, as the
measurements utilized are local. In this scheme the first and
the second order derivatives of the current have been utilized to detect the low and high impedance faults in the network. The derivatives are applied to discriminate various faults
such as faults on the adjacent lines, and faults with different
impedance values. The thresholds for the derivatives are analytically calculated and utilized in the algorithm. The proposed
protection scheme is simple, fast and suitable for protecting
DC MG from line faults, without interrupting the normal operation of the grid. A protection solution for loop type DC
MG system is demonstrated, which is easy to extend to other
configurations.
Fig. 16.
High impedance fault F1 in islanded mode.
locally, the noise introduced due to the communication link
can be neglected. Therefore, only the measurement noise is
considered. The power distribution factor of the noise in the
measurement can be taken as Gaussian [28]. The system is
simulated for fault F1 with the signal to noise ratio of 20 dB,
as shown in Fig. 17. It is found that the noisy signal and their
APPENDIX
SYSTEM CONFIGURATION
A loop type DC microgrid system is considered for designing
the protection system, as shown in Fig. A1. The DC bus voltage
is controlled by grid-VSC in grid connected mode or by battery
converter in islanded mode [30]. The component ratings of all
the modules are given in Table A1 [31].
MEGHWANI et al.: NONUNIT PROTECTION SCHEME FOR DC MICROGRID BASED ON LOCAL MEASUREMENTS
[10]
[11]
[12]
[13]
[14]
[15]
Fig. A1.
DC MG Architecture considered in this work.
TABLE A1
RATING OF DC MG COMPONENTS
1200 V
DC Grid Voltage
Base Power
Battery DC-DC Converter
Battery
PV Converter
Solar Panel
Grid VSC
Wind Turbine
Cable Resistance
Cable Inductance
Filter Capacitor, C
Load
2 MW
0.5 M W
300 V , 1.3 kAh, Nickel Cadmium
0.5 M W
V m p = 54.7 V , I m p = 5.58 A at STC
1 MW
2 M W , PMSG
10 m Ω per Km
100 μH per Km
25 m F
Constant impedance load 2 M W
[16]
[17]
[18]
[19]
[20]
[21]
ACKNOWLEDGMENT
The authors would like to thank the Department of Science and Technology, New Delhi, India for providing financial support to carry out this research work under project
no.DST/EE/20100258.
[22]
[23]
[24]
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Authors’ photographs and biographies not available at the time of publication.