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New Fault Current Limiters for Utility
Substations – Design, Analysis,
Construction, and Testing
By
Frank Darmann
Tim Beales
Australian Superconductors
A Wholly Owned Division of SC Power Systems, Inc. USA
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New Fault Current Limiters in Utility Substations – Design,
Analysis, Construction, and Testing
Frank Darmann, PhD
Tim Beales, PhD
Australian Superconductors
A Wholly Owned Subsidiary of SC Power Systems, Inc. USA
Abstract
The fault current on the low voltage (LV) side of a sub-station is dependent on the
transformer impedance, the number of transformers connected to the LV bus, and the
incoming fault level on the high voltage (HV) side. Switchgear and other plant on the 22 and
11 kV distribution side typically have a fault rating of 500 and 250 MVA respectively. Due to
the increase of incoming fault levels, or the installation of another transformer, fault levels at
the LV bus can sometimes exceed these ratings. Typical standard solutions to this problem
include upgrading the LV switchgear, installing Neutral earthing resistors (NERs), or
operating the LV bus in a “split bus mode”. A less common solution is to install passive
series limiting reactors or fault current limiters (FCLs), which can reduce the fault current but
at the expense of regulation and not insignificant energy losses.
In this paper, a new generation of passive series fault current limiter will be described which
present none of the disadvantages of conventional FCLs. The improved properties include a
negligible reactance during normal conditions, high impedance during fault conditions, very
low copper loss, removal of the transient and DC component, and limitation of the steady
state fault current. These advantages are achieved by employing three phase saturated core
fault current limiters with the saturating current provided by a single high efficiency, high
current density DC superconducting coil. These superconducting fault current limiters (SCFCL) are still passive devices and do not suffer from switching delays.
Introduction
The need to limit faults in Australia's electricity supply system is a necessity from both
practical and regulatory requirements. A EU study has estimated that about 150 faults per
year per 100 km of line occur in distribution lines in Europe.
There are many reasons for wishing to limit fault currents at the zone substation level.
These include:
1. To meet the regulatory requirements,
2. To assist in the parallel operation of existing transformers to improve reliability
without resorting to use of high impedance transformers,
3. To reduce expenditure on system upgrades allowing existing switchgear to remain in
circuit,
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4. To extend the lifetime of substation equipment, including the circuit breakers and
transformers,
5. To enable flexibility in substation upgrades,
6. To increase safety, reliability, and improve voltage sag levels for upstream customers.
7. To prevent saturation of current transformers during faults
Superconducting FCLs have an automatic fault sensing response, are self-triggering, have a
very short response period of less than 1 ms, and in the normal operating conditions have
negligible reactive voltage drop and energy loss. Superconducting FCLs can be designed to be
sited at several key locations within a network.
There are several designs that can employ superconductors to limit fault currents: (i) the
resistive model that uses a single superconductor resistive element; (ii) the resistive shunt model
that uses a resistive superconductor element with a resistive shunt; (iii) the inductive shunt
model that uses a resistive superconductor element with an inductive shunt; (iv) the screenedmagnetic core inductive model that uses a cylinder made of superconductor surrounding a
magnetic core; (v) the shorted-secondary inductive model that uses a short-circuit
superconductor winding or ring surrounding a magnetic core; (vi) the electronic-switching
inductive model that uses a superconductor coil in association with a power electronic circuit;
and (vii) the saturated magnetic core inductive model that uses a superconductor winding and a
magnetic core.
The superconducting FCL design being developed by AS is a saturable inductor-type. The
operating principle is an iron core that is wound with two coils, one normal AC and the other
superconducting DC. The HTS DC winding is energized sufficiently to saturate the iron core
with magnetic flux. Consequently, when current flows in the AC winding attached in series to
the load AC system, it will not significantly change the magnetic flux, and therefore the
impedance to the AC current is very low. However, should a fault occur in the system
upstream from the FCL, the AC current rises and takes the iron core out of saturation. As the
magnetic flux has now changed, the impedance of the AC winding increases and the fault
current can be limited.
Australian Superconductors first looked at the generic properties of superconducting FCLs in
medium and high voltage networks, and began to focus exclusively on the saturated magnetic
core inductive model. The saturable FCL design being developed by AS has a number of
advantages over other superconducting FCL designs: (i) it is designed to use HTS tape as the
DC winding, so there are miniscule operational losses; (ii) the saturable FCL design uses HTS
tape which is the most robust, practicable and reliable conductor compared to other potential
HTS current limiting elements; (iii) there can be no superconducting quench with this design;
(iv) the superconductor is not connected to the rated AC system current and therefore, the
HTS need not be rated the same as the power system current levels; (v) the required
impedance to limit the fault current can be easily achieved by controlling the number of AC
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winding turns; (vi) the fact that the fault current capability is essentially determined by the
NA value of the DC coil, means that a series of fault current levels may be accommodated by
this design using an off-load tap changer-type arrangement; (vii) the HTS cryogenic system is
very simple, since the high voltages and currents of the power system do not pass through it;
and (viii) in the unlikely event that the HTS FCL should fail for any reason, the normal power
supply will not be cut off.
The operational and design details of the Australian Superconductors SC-FCL working
prototype will be given in a later section.
Conventional Reactor Designs
Conventional series reactors are very similar to transformers, in that they use paper-insulated
windings with copper conductors. The steel core however, is not continuous, but is built with
an air gap. This forms a high reluctance magnetic flux path, which increases the terminal
impedance of the copper windings. There are four basic types of conventional series reactors
used to provide a high impedance: (i) those with disk or helical windings that are cast in
concrete with an air core; (ii) those with their coils immersed in oil with an iron core that
contains an air gap; (iii) those with their coils immersed in oil with an eddy current induced
magnetic shield and no iron core; and (iv) those with their coils immersed in oil with an
electromagnetic shield and no iron core.
Ideally, series current limiting reactors would not contain an iron core, since high fault
currents saturate the iron and this tends to reduce the terminal impedance. This feature
reduces the fault current level that can be limited. Hence, to circumvent this problem, reactor
Designs (iii) and (iv) listed above have no iron core. In the iron-free type of reactor,
mechanical and thermal considerations determine the magnitude of the fault current that can
be limited, rather than any magnetic considerations.
The type of reactor cast in concrete (Design (i)) completely eliminates iron from the device. It
consists of helical or disk type windings of copper conductor supported by symmetrical
concrete posts arranged around a circle. The posts have no steel re-enforcement. The
relatively high heat capacity of concrete provides excellent temperature control. The
disadvantage of this design is that there is a considerable external magnetic field, and this
poses a health and safety hazard to personnel working near to the device. In addition, the
concrete has to be cast without any defects, and the reactor must be housed in a room of nonmagnetic materials and so, for example, a concrete structure with steel re-enforcement is not
acceptable.
The same considerations do not apply to the oil-filled reactors of Designs (ii) to (iv). Owing
to the iron content of these designs, the magnetic field is constrained to move along the
lowest reluctance path, which will be through the iron core. Several types of design exist. One
particular three-phase design, invented by ABB and manufactured under licence by Alstom in
Australia, uses three long cylindrical stacks (one per phase) of alternating transformer steel
and non-magnetic slate/epoxy matting acting as separator between the laminations to provide
the "air” gap. This arrangement can be shielded with aluminium to limit the external field, or
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be clad in low-loss iron laminations to prevent any external field. The air gaps make up no
more than about 1% of the magnetic flux path, but this is sufficient to reduce the flux density
of the device to a level such that, even at fault currents of 10 to 12 times the normal rated
current, the core remains unsaturated, and the reactance only drops by a few percent
compared to the steady state reactance.
This report will show, however, that even for very small reactors, the copper losses associated
with these types of reactors are about two orders of magnitude greater than an HTS FCL is.
Owing to the complexity of the circuit equations (Equations A1-A4 in Appendix 1), and the
need to study the effect of faults at various “points-on-wave”, the only viable option was to
use electromagnetic transient (EMT) analysis using the Power Systems Computer Aided
Design (PSCAD) software package [1].
PSCAD is an industry standard electromagnetic transients simulation program used by
engineers, scientists, utilities, consultants, research, and academic institutions all around the
world. PSCAD is a graphical user interface used to set up the circuit to be analysed, and
EMTDC is the solution engine. It is used in the planning, operation, and design of power
systems, and in the teaching and research into power systems, power electronic systems, and
their control. The capability of the package is evident from many publications that exist in the
public domain in relation to the studies carried out. The package is being continuously refined
and improved by a team of researchers, engineers, and software experts.
The Power Quality Centre at the University of Wollongong has access to the latest PSCAD
software developed in Canada.
Placement of SC-FCLs in T&D Networks
The most appropriate location of SC-FCLs was identified to be on the secondary side of each
transformer circuit, as shown in Figure 1, such that each transformer in a substation is
individually protected. Each FCL is rated at the maximum load current of each transformer
circuit.
Other placement possibilities exist, notably, the high voltage side of the transformers at the
sub-station or on the low voltage side of the transformers at the terminal station. These two
options will effectively limit the incoming fault level, however, the complexity of the SCFCL design would also increase owing to the higher voltage level required. Therefore, in the
short term, we considered a low voltage FCL design (operating at 22 or 11 kV) and the
connection scheme shown in Figure 1 to be the most applicable solution for rapid
implementation in the Australian grid.
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Xl
Xl
Figure 1.
Individual transformer protection using SC-FCLs
Principles of the Australian Superconductors Designed SC- FCL
Previous work in academia has shown that the design of a single-phase DC saturated
superconducting FCL requires two separate closed iron cores as shown in Figure 3. A
separate DC HTS winding that is wound in the form of a solenoidal coil energizes each core
independently. The cores are both saturated to a designated field intensity value, Hdc, in the
opposite sense, according to the standard dot notation used in the industry, as shown in Figure
2. Typically, Hdc will have a value between 9,000 and 20,000 Am-1. The DC current flows out
of the page in the positive cycle saturated core, and into the page in the negative cycle
saturated core. The corresponding points on the DC magnetisation curve of the cores are
denoted by (Bdc, Hdc) and (-Bdc, -Hdc), respectively.
In the above concept, the superconducting DC windings will be maintained at a constant
temperature of 77 K (-196 ºC) as they are immersed in liquid nitrogen, and the whole
assembly is contained in an isothermal vacuum-jacketed cryostat, with its outer walls at room
temperature.
The ampere-turns required for each HTS DC coil to achieve the required value of Hdc in a
wound core (i.e no air gaps) is given approximately by
NI = 2(2w + 2h)Hdc
(1)
where N is the number of total DC turns, I is the DC excitation current, w is the average core
width in the plane of the paper of Figure 2, h is the average core height in the plane of the
paper of Figure 2, and Hdc is the required field intensity of the saturated core.
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dc
AC linkage windings, n turns
DC saturating windings, N turns
Figure 2
Schematic of single phase SC-FCL layout
The AC windings in the above model are then arranged in such a manner that the AC flux
arising from each coil is in the opposite sense to the DC core magnetisation of each. The
instantaneous inductance may be found from the differential or incremental permeability, diff
μdiff = (dB/dH) | Bdc
=
ΔB/ ΔH | Bdc
(2)
where ΔB and ΔH are the maximum extents of the minor hysteresis loop at the DC bias
points, Bdc and Hdc. Appendix 1 details the derivation of an equation approximating diff.
The instantaneous impedance presented to the network may be expressed in phasor notation
as
Z = R + 2πf(n2A/l)μdiffj
(3)
where R is the resistance of the AC coils, f = frequency of operation (50 Hz), l = mean flux
path length of the iron core assuming no air gaps, j = square root of -1 (the imaginary
number), and n = the number of turns of the AC winding. The resistance of the AC coils, R, is
small compared to the imaginary part of the impedance during a fault.
For an effective HTS fault current limiter, the normal operating inductance of the core must
be small, so as not to impose any unnecessary regulation of the line or impedance to the
current flow. This is normally achieved by ensuring that the DC flux density in the iron core,
Bdc is 1.9 – 2.2 T, and thereby ensuring that μdiff is approximately zero.
The DC field is chosen such that an oscillatory fault current of peak value, If, will increase the
differential permeability to a large value during a fault condition. The size of the cores, DC
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current, and DC turns are calculated based on the fault level and the permeability of the iron
so that
nIf(max)/l = Hdc
(4)
and
nIf(min)/l = Hdc - Hdc(sat)
(5)
where n is the number of AC turns, l is the length of the wound core magnetic circuit, Hdc is
the DC field intensity of the DC coil, Hdc(sat) is the field intensity required to saturate the core,
If(max) is the maximum fault current that the HTS FCL is required to limit, and If(min) is the
minimum fault current that the HTS FCL is required to limit.
To limit the fault current, the number of AC linkage windings is chosen such that
n = NI/If
(6)
Owing to the oscillatory nature of the fault current, two separated cores, as shown in Figure 2,
are required to provide different senses of differential permeability to the AC windings and
thereby limit the fault in both positive and negative cycles of the fault current. A three-phase
FCL to the above concept would therefore require six saturated cores and would require six
separate HTS DC windings.
Australian Superconductors three-phase superconducting FCL
Figure 3 shows a schematic representation of a three-phase SC-FCL design patented by
Australian Superconductors that takes the basic design concepts given above, and builds in
added advantages for commercial devices. Figure 3. highlights the placement of the core
cross-sections in relation to the single superconducting winding used and the AC copper
linkage turns. This design maximises the phase-to-phase clearance between the copper
linkage coils. One design advantage is the use of a single superconducting coil to saturate all
the iron cores to the required design value.
The iron core laminations chosen for the SC-FCL are Kawasaki type 35RG155 laminations.
Since there is only a small steady state perturbation in the hysteresis loop about the DC
position, expensive low loss core steels are not required, and thicker, cheaper laminations
than those used in transformers can be employed. The incremental permeability of the
35RG155 laminations was fitted to a closed-form curve to allow the magnetisation properties
to be used in the PSCAD modelling software package. Figure 4 shows a linear depiction of
the magnetisation curve, and a graphical explanation of the operating principals of a
superconducting FCL.
At 50 Hz and under normal operating conditions (i.e., no fault), the magnetisation response of
the iron core will oscillate about the operating point shown in Figure 4. In this condition, μdiff
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w
d
Basic footprint = (3.9.d+2.w)^2
Iron core leg, diameter: d, window: w x h
AC coils
Cryostat
HTS coil and support
Footprint
Figure 3:
Schematic of an SC Power/Australian Superconductor’s 3-phase SC-FCL.
Figure 4.
Magnetisation properties of core steel laminations, 35RG155
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is approximately that of air (the FCL is essentially an air core inductor, μdiff = 4π x 10-7), and
the terminal reactance of the device is essentially zero owing to the very few ac turns.
In a fault condition, the magnetisation of the iron will decrease to a point where μdiff has a
maximum value of μdiff = 0.1275, (five orders of magnitude greater) and the terminal
inductance increases to a value of a few henries, depending on the specific FCL design used.
Essentially, the FCL behaves like a current-controlled reactance. A plot of the terminal
inductance versus terminal current for one such FCL design is shown in Figure 4. Each peak
is produced by each half of the iron core. In this case, the inductance was designed to have a
maximum at a fault current of 2.5 kA to effectively clip the fault current to this value. The
two peaks effectively form a high impedance barrier to the fault current.
Terminal inductance (i)
10.0000
1.0000
0.1000
Fault
conditions
0.0100
0.0010
Normal
operating
conditions
0.0001
0.0000
-5000
-4000
-3000
-2000
-1000
0
1000
2000
3000
4000
5000
Terminal current (i, amps)
Figure 5
Typical plot of the SC-FCL instantaneous terminal inductance versus instantaneous terminal
current of a SC-FCL This particular design would limit the fault current to 2.5 kA.
The steady state inductance of this design is negligible, owing to the fact that the incremental
permeability at the DC operating point is approximately 4x10-7. In addition, the resistance of
the DC linkage coils is also small compared to the power rating. Hence, during normal
operating conditions, the FCL is invisible to the network, and requires no extra voltage
regulation to fluctuations in normal load levels.
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Advantages of DC saturated SC- FCLs over series limiting reactors
The traditional approach to limit fault currents has been to employ higher reactance
transformers, and to install series limiting reactors. Series reactors impose significant losses at
the substation, which can amount to hundreds of kW that are generated 24 hours a day over
the life of the substation. A conventional series reactor, for example with an inductance of 0.4
mH and rated at 11 kV/2.4 kA would have a copper loss component of 100 kW, and would
waste 877 MWh annually in lost energy. At a price of $20/MWh, this represents a cost of
$17.5k per annum, or $0.5 M over 30 years. A SC-FCL can reduce fault current levels
without introducing a high impedance during steady state (non-fault) operation. This study
has calculated that the overall steady state losses of a SC-FCL (i.e including DC current coils)
are below 10% of the losses of a conventional reactor. The advantages of SC FCLs over
conventional reactors are:
1. The steady state reactance of the Australian Superconductors FCL design is
negligible.
2. The copper loss of the Australian Superconductors FCL design is low, approximately
10 kW.
3. The windings of the Australian Superconductors FCL design are significantly smaller
in size owing to the high current density of the superconductor windings (10,000
A/cm2 compared to 350 A/cm2).
4. The superconductor windings of the Australian Superconductors FCL design are
isothermal, so there is no long-term degradation in the insulation.
5. There is no oil in the of the Australian Superconductors FCL design, which reduces
the risk of fire, lowers maintenance, and reduces the need for civil works.
6. The coolant used for superconducting windings of the Australian Superconductors
FCL design is liquid nitrogen, which is environmentally benign.
7. The fault level of the Australian Superconductors FCL design can be adjusted by
tuning the DC current to the If value required.
8. The Australian Superconductors FCL design has flexible dimensions.
9. The Australian Superconductors FCL design is able to completely clip initial
transients.
10. The lifetime of the main portion of the HTS FCL is unlimited, owing to its cryogenic
operating temperature, which essentially preserves the windings. The copper coils can
be replaced or upgraded at a later date, because they are relatively small and cheap.
11. HTS FCLs are future proof. Whereas conventional reactors are rated for a single line
current and a single fault current level, HTS FCLs do not suffer this limitation. Future
increases in the fault current level can be easily accommodated: increases in the rated
line current can be accommodated by changing the AC coils, and the original DC HTS
coils and iron core can remain.
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Performance Study of Australian Superconductors SC-FCL Design at Specific
Australian Utility Sites
Site 1: Substation ST1
Substation ST1 has two Delta: Star, 45 MVA 132/11 kV transformers each with an
impedance of 67 % on a 100 MVA base. The total primary line current is rated at 400 A, and
the secondary line current is rated at 4,800 A. The substation occupies an area of 100 x 82 m2
and is located in a high population growth area. The load is fed from 14 radial feeders off the
site to a mixture of domestic and industrial users, with the property lines of residential houses
abutting the once isolated site. The only area available for a superconducting FCL is on a
blue-metal area of 5.58 x 7.06 m2 adjacent to, and between the existing transformers. The
majority of the free land on site is unsuitable for the installation of power equipment owing to
the significant induction from the overhead 132 kV lines that traverse it.
At this site, a single transformer is used to supply the load, in which case the contribution to
the fault impedance is the full 67%. However, both transformers may be required to operate
in parallel for approximately three months of the year and in this case, the steady state fault
current level is 272 MVA (14.3 kA line current) for a bolted, three-phase to ground short
circuit. To reduce the fault level, a split bus arrangement is employed. A SC-FCL was design
to limit the fault level to 250 MVA (13.1 kA) on a solid bus arrangement.
The electrical layout of the ST1 substation was modelled using PSCAD software. This
confirmed the transient nature of the faults at the ST1 substation, and confirmed the
behaviour of the SC-FCL under a variety of fault conditions. The worst case of a three-phase
to ground bolted short circuit coinciding with a voltage zero on one phase was used to
demonstrate the behaviour of the FCL. However, the behaviour of the FCL under other types
of fault were also examined, such as phase to phase and phase to ground faults. Figure 6
shows the PSCAD network lay out for ST1 with both transformers in service.
Design and Space Considerations of the SC-FCL at ST1
Figure 7 shows the 5.58 x 7.06 m2 available footprint area for a superconducting FCL at the
ST1 substation. This is an area located between the two transformers. In addition, the height
is limited to approximately 3.0 m. A similar space is available for a second superconducting
FCL to the left of the transformer location. A superconducting FCL was designed to fit into
the required volume using Australian Superconductors' design. The FCL dimensions are 2.2 x
2.2 x 2.5 m3. However, if there was a need to limit the height, for example, for transport or
other clearance reasons, then the FCL could be redesigned, while still meeting all the
electrical requirements. As an example, an alternative footprint is shown in Figure 6, which
has dimensions of 4.0 x 4.0 x 1.2 m3. Both designs have an approximate copper winding loss
of 6 kW. Scope is also available for designing a very small footprint SC-FCL (1.4 x 1.4 m)
but which has a correspondingly greater height (4.2 m).
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To cable box
5580
2500
7060
2200
5580
Figure 6.
ST1. Volume options, which will enclose the SC-FCL, and the available space between the
existing transformers. All dimensions are in mm.
Electromagnetic Transients Analysis at ST1 Using PSCAD/EMTDC
The PSCAD software package was used to model the performance of the ST1 substation
under electromagnetic transients (EMT) (i.e., faults). It was confirmed that the most severe
(i.e., highest amplitude) transients in the current occurred when the fault coincided with a
voltage zero crossing on one of the phases. Owing to the symmetry of the system, all faults
were synchronised to a zero crossing of the Phase A voltage waveform.
We proposed to locate a three-phase FCL in each of the 11kV/2.4kA circuits at ST1, as
shown in the PSCAD diagram, Figure 7. With no FCL in service, the fault current in each
circuit is nominally 7.15 kA rms, which gives a fault current of 14.3 kA rms in the bus bar.
To meet the requirements of the Integral Energy engineers, the Australian Superconductors’
superconducting FCL design had to be able to limit the fault current in each circuit to 6.55 kA
rms.
There were two design options for the FCL that were considered, the final design choice
would depend on the specification of the installed switchgear and circuit breakers: (i) FCL
Design 1 was configured to limit the steady state rms current waveform per circuit to 6.55 kA
rms, and substantially reduce the transient peak; and (ii) FCL Design 2 was configured to
completely limit the fault current waveform to below 9.3 kA peak during the transient and
steady state operating conditions.
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PSCAD model of the Rooty Hill sub-station
132 kV
11 kV
14.3 kA rms
7.15 kA rms
13.1 kA rms
A
B
#1
Ia1
B
B
LFCL1
+
Ib1
Ib
#2
C
C
Ic1
LFCL2
+
Ic
0.3
C
C
132.0
11.0
+
A
A
45 [MVA]
0.3
Ia
LFCL3
3 Phase
RMS
Ea(11kVbus)
A
B
A
C
#1
132.0
#2
11.0
Ia2
FAULTS
C
ABC->G
B
Ib2
LFCL4
+
C
Ic2
LFCL5
+
faultcontrol
B
45 [MVA]
+
A
Faultcontrol
A
0.021095 0.3
B
Va
6.55 kA rms
Timed
Breaker
Logic
Closed@t0
LFCL6
7.15 kA rms
1.192
0.00285
1.192
0.00285
1.192
0.00285
6.55 kA rms
Figure 7.
PSCAD circuit diagram for ST1 showing the existing fault current (272 MVA) and the
required fault current (250 MVA).
ST1 SC- FCL Design 1
Figure 8 shows the current waveform in phase A both before and after a fault, and with and
without a SC-FCL Design of type 1. Only one phase is shown for clarity, as the other phase
current waveforms are similar. The peak transient is reduced from 15 kA to 12 kA and the
steady state fault level is limited to 6.55 kA rms (9.3 kA peak) as required.
ST1 SC- FCL Design 2
Figure 9 shows the waveforms of the Phase A current for a SC-FCL of design type 2 at
substation ST1 compared to the normal situation (without SC-FCL) and an appropriately
rated series reactor (0.126 ). Both the transient and steady state waveform are limited by the
SC-FCL, but the series reactor only reduces the steady state response. The advantage of the
SC-FCL is that the transient is clipped as well as the steady state response, and this cannot be
achieved with a conventional series reactor.
To ensure the suitability of the FCL under different types of faults at different point on wave
timings, FCL Design 2 was used in an EMT analysis on faults at a voltage zero of the single
phase to ground type and phase to phase type. In addition, the analysis was also carried out
for a line to ground fault at a voltage maximum, and a three phase to ground fault at a current
maximum was also checked with very similar results.
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Figure 8
Fault current on phase A with and without a SC-FCL of design type 1 at ST1
(Three-phase fault occurs at a voltage zero crossing on phase A)
Figure 9
Clipped fault current - with SC-FCL of design type 2.
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Site 2: Substation ST2
Substation ST2 has three Star: Star, 20 MVA 66/22 kV transformers each with an impedance
of 10 % on a 20 MVA base. The primary line current is 175 A, and the secondary line current
is rated at 525 A per transformer. The substation occupies an area of 86 x 40 m2 and was
nominated because of its high fault level (three transformers are already installed) and its
strategic importance. The 22 kV load is fed from 13 feeders off the site to a mixture of
resistor (NER) has been installed to reduce the current for line to ground faults. The 22 kV
fault currents are summarised as follows:
1. Three phase = 12.3 kA (4.1 kA per phase),
2. Phase to ground = 2.3 kA with an 8 Ω NER installed,
3. Phase to ground = 13.0 kA without an NER installed.
Design of a SC- FCL for ST2
As an exercise to develop an alternative to the installed NER, a SC-FCL was designed to
limit the fault current to 2.5 kA per circuit. In this case, the design, which completely clips the
fault current, was the preferred option.
The PSCAD software was used to set up a network diagram of the ST2 substation site to
determine the transient response to various types of faults, both with and without a SC-FCL.
Figure 10 shows the PSCAD circuit diagram. The assumptions made about the ST2
substation were as follows:
1.
2.
3.
4.
5.
The lagging load power factor = 0.8.
The X/R ratio of the 66 kV generation = 20.
The steady state load current per circuit = 0.42 kA (i.e. 1.25 kA / 3).
All three transformers are in circuit.
All fault analyses were carried out with the fault occurring on the 22 kV bus at the
substation without including any extra line impedance.
6. No NER was included in the circuit.
It was confirmed that the worst-case transients were produced when the fault occurred on a
voltage zero (phase to ground) on one of the three phases.
Several protection regimes were looked at, and the one presented in this report is the one
considered to be the best solution. This was to insert a three-phase superconducting FCL into
each circuit on the secondary side of each 20 MVA transformer, as shown in Figure 10. A
three-phase superconducting FCL was designed to reduce the steady state fault current in each
circuit to 2.5 kA peak. The dimensions of the SC-FCL are approximately 2.0 x 2.0 x 2.4 m2.
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Va
A
A
0.00646 0.3
B
B
#1
A
Ia1
+
B
B
Ib1
LFCL1
+
Ib
C
C
Ic1
LFCL2
+
Ic
A
20 [MVA]
#2
0.3
C
C
66.0
22.0
0.3
Ia
LFCL3
3 Phase
RMS
Ea(22kVbus)
Faultcontrol
A
B
B
20 [MVA]
#1
C
#2
66.0
22.0
A
Ia2
+
B
Ib2
LFCL4
+
C
Ic2
LFCL5
+
FAULTS
C
ABC->G
faultcontrol
A
Timed
Breaker
Logic
Closed@t0
LFCL6
A
B
C
20 [MVA]
#1
#2
66.0
22.0
8.129
0.0194
8.129
0.0194
8.129
0.0194
+
A
Ia3
B
LFCL7
+
Ib3
C
Ic3
LFCL8
+
LFCL9
Figure 10
PSCAD circuit diagram of the Heatherton site
Electromagnetic Transients Analysis at ST2 Using PSCAD
Figure 11 shows the results of the transient analysis of the ST2 substation for a three-phase to
ground fault on the 22 kV bus without a superconducting FCL present. The circuit load
current before the fault was 0.42 kA rms, and the maximum transient fault current in each
transformer circuit was 10.6 kA occurring at t = 0.21 s, half a wave cycle, after the fault
occurred. A 5-kA DC offset in Phase A is clearly visible, and the steady state fault current is
6.2 kA peak or 4.4 kA rms.
Figure 12 shows the results of the transient analysis of ST2 substation with a three-phase
superconducting FCL inserted into each transformer circuit between the secondary and the
bus. The maximum fault current was limited to 2.5 kA both in the transient and in the steady
state.
The ST2 substation FCL design was chosen such that the transient and the steady state fault
were both limited to a maximum peak current of 2.5 kA. However, other FCL configurations
could have been employed. For example, it is possible to limit the transient current to any
given figure (say, 3.5 kA), and also limit the steady state to a given figure (say, 2.5 kA).
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Figure 11.
Transient fault currents per circuit at ST2 for a three phase to ground bolted short circuit.
Figure 12
Steady state three phase fault response of Heatherton with SC-FCL.
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Transient analysis of the same circuit was undertaken for the following types of faults with
very similar results to those presented.
1.
2.
3.
4.
Single phase to ground, voltage zero crossing on that phase.
Single phase to ground, current maximum crossing on that phase.
Phase to phase, voltage zero crossing on one of those phases.
Three phase to ground, current maximum crossing.
It was estimated that to achieve the same reduction in steady state fault current, a fixed
conventional reactance of 2.04  would be required in each phase of each transformer circuit
to reduce the final steady state rms current (in each circuit) to 2.5 kA. The response of the
ST2 substation to a bolted three phase fault was analysed using PSCAD analysis with each of
the FCL’s replaced by fixed 2.04 Ω reactors (one in each phase of each transformer circuit).
Although the steady state rms fault current is indeed reduced to 2.5 kA (Figure 13), the initial
peak remains at 6 kA. The additional advantages of the SC-FCL include the ability to clip the
transient to 2.5 kA, and reduce the DC component by half.
Figure 13
Phase currents per circuit at ST2 for a three phase to ground fault occurring at a voltage zero
with 2.04 conventional series reactors installed.
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The Australian Superconductors Prototype SC-FCL
A working 1 MVA single-phase prototype SC-FCL has been built and tested at the Australian
superconductors plant. Figure 14 shows the essential components.
440 mm
650 mm
Figure 14.
The Australian Superconductors prototype single phase SC-FCL showing 2 separate iron
cores with copper linkage coils (nac), and two cryogenic vessels holding the SC coils (N)
For demonstration purposes and for confirming the mathematical model of operation, the
steady state current is set to 40 A rms through a 2 3.4 kW resistive load. A 6 V battery and
0.2  resistor provides a current of 30 A dc through the superconducting coils. A contactor
and push button operated trip coil provide a user operated short circuit across the 2 
resistive load. A sealed quick-lag type circuit breaker serves as a visual demonstration to the
operation of the FCL. The CB time lag characteristics are such that a fault current above
400A will trip the coil within half a cycle (0.01 s) and a fault current below 400A will take
more than 2 s to trip.
Prototype Simulations and Measurements
Figure 15 shows the results of the PSCAD/EMTDC simulations with and without the
prototype FCL in circuit for a number of different AC, linkage turns. Figure 16 shows the
measured fault current with and without the SC-FCL in circuit. The load is resistive and
hence the steady state fault current is achieved within 2 – 3 cycles. As can be seen, the SCFCL virtually eliminates the transient fault current from 2.4 kA to 0.4 kA. Figure 17 shows
the steady state fault current in comparison to the simulation. The disagreement in the shape
of the current waveform is due to the approximation used for the incremental permeability
(Equation A1).
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Figure 15
Results of PSCAD simulations with various ac linkage turns (nac)
Figure 16
Measured fault currents with and without a SC-FCL in circuit
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Figure 17
A comparison of the steady state fault current with the SC-FCL in circuit.
Simulated and measured waveforms shown.
Terminal Inductance of the SC-FCL
Figure 18 shows the instantaneous terminal inductance (i.e on the AC side) as a function of
time, and superimposed on the fault current waveform. These curves offer insight into the
working principal of the SC-FCL and how the regulation problems imposed by series reactors
are overcome. Firstly, in the steady state, the terminal impedance is negligible which is a
direct consequence of the saturated iron core. When a fault condition occurs, the terminal
impedance increases to 188 Ohms at the set fault current limit (NI/nac), which effectively
clips it and prevents further increases.
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Figure 18
Instantaneous terminal inductance of the SC-FCL in relation to the fault current
Cooling Considerations for the Superconducting Coil
The DC superconducting coils used to energise the magnetic circuit will be immersed in
liquid nitrogen under a slight positive pressure of approximately 0.1-0.6 bar, depending on
the specific design parameters. Immersing the coils in liquid nitrogen simply cools the coils to
a temperature below their superconducting transition temperature and so maintains the coils
in a zero resistance state. In the DC mode, there is zero resistance to the passage of an
electrical current.
Liquid nitrogen is available in tonnage daily tonnage quantities throughout Australia from
several companies, the largest being BOC Gases, Linde, and Air Liquide.
The liquid nitrogen is kept in a large thermos-type flask called a vacuum-insulated cryostat,
because it will tend to evaporate owing to heat ingress from the ambient. The cryostats
proposed for the Australian Superconductors design are to be manufactured of stainless steel,
and the vacuum insulation can keep the liquid nitrogen evaporation rate to a very small, finite
value. The actual evaporation rate will depend on the cryostat design. However, for the two
FCL designs considered in this report, the approximate evaporation rate is 20 litres of liquid
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nitrogen per day. The Australian Superconductors FCL design can employ two techniques
that can address the liquid nitrogen evaporation problem.
Total Loss System
Owing to the fact that gaseous nitrogen is environmentally benign, and liquid nitrogen is
available for as little as 40 cents a litre, it can be economical to allow the nitrogen gas to
simply vent to the atmosphere, and simply replenish the liquid from a large capacity (≥1,000litre) storage vessel. The cost of replacing this liquid nitrogen lost would be in the range of $8
a day. Automatic re-fill systems, which detect the level of the liquid nitrogen to an accuracy
of 1% and can initiate a solenoid valve to re-fill the cryostat, are available for under $4,000.
Owing to the relatively low cost of this type of replenishment system, a large redundancy can
be cheaply built in. In addition, sufficient liquid reserve can be built into the cryostat to cover
for up to two weeks or more, by making it artificially larger than necessary.
Re-liquefaction Using a Cryocooler
An alternative cooling arrangement is a closed cycle system where no nitrogen gas is allowed
to escape. In this arrangement, a “cold head”, which resembles a copper block, is inductioncooled by a cryocooler refrigeration unit, and is used to reliquefy the nitrogen gas directly in
the cryostat. The cold head sits directly above the liquid nitrogen pool in the FCL cryostat.
The cold head and is fed with cold, pressurised helium gas, which acts as the refrigerant (or
working fluid) to maintain the cold head at a temperature between 66 and 70 K (-207 and 203 ºC). The nitrogen gas that has evaporated from the pool of liquid in the FCL cryostat
condenses on the surface of the cold head, where it exchanges its latent heat and recondenses
back into liquid nitrogen. Once the droplets of condensed nitrogen reach a given size, they
simply fall back down into the cryostat under the influence of gravity.
A small cryocooler of only 2 kW capacity would be able to cope with a liquid nitrogen
evaporation rate of 10 litres a day, and a 5.5 kW capacity cryocooler would be able to
reliquefy a nitrogen evaporation rate of 40 litres a day. This is because, in terms of their
cooling power, cryocoolers rapidly increase in efficiency with increasing size of the
cryocooler. This fact almost single-handedly dictates why superconducting devices become
much more economically competitive as the system power requirements increase.
The size of a cryocooler re-liquefaction system has a volume of 1 x 0.8 x 0.7 m3, and weighs
200 kg. The cold head inserted into the cryostat is essentially cylindrical in shape with a
diameter of 100 mm and a length of 500 mm. A length of about 250 mm protrudes outside the
FCL cryostat.
A typical system to reliquefy 40 litres of liquid nitrogen a day would cost approximately
$70,000 to purchase, ship, and install, and would cost an average of $2,000 per year to
maintain. The maintenance would consist of replacement of some gasket and O-ring seals
every 8,000 hours, and the replacement of a cryocooler part (the regenerator) every 36,000
hours. The maintenance time is expected to be approximately 2 hours per year on average.
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TechCon® 2003 Asia-Pacific
Maintenance requires a shut down of the cryocooler, but not of the FCL device, as the liquid
nitrogen buffer within the cryostat will keep the superconducting coils cold.
Conclusions
It was shown in this report that superconducting FCLs can reduce the fault level at
substations, enabling a previously split-bar operation to be made solid, as in the case of ST1,
or displacing the need for a neutral earthing resistor, as in the case of ST2.
It was also shown that unlike conventional reactors, SC- FCLs can completely clip the fault
current transient to a predetermined level. In addition, a relatively small conventional reactor
at ST1 would consume about 100 kW of energy, whereas the AS superconducting FCL would
use only about 6 kW.
Superconducting FCLs are also future proof, in that their rating can be upgraded by
replacement of an item with a small capital value (i.e., the AC linkage coils) compared to cost
of replacing the entire device. It was also shown that SC-FCLs could be made sufficiently
small to fit in the available space at the substations located at ST1 and ST2.
The provision for cooling the DC coils of the superconducting FCL is cheap, of low risk, and
an existing solution exists, which is reliable and proven in the field. In particular, the most
economical cooling technique is to use a simple storage vessel containing liquid nitrogen, and
this can be topped-up using an auto re-filling scheme. Sufficient redundancy can easily be
added to the cooling system to allow for maintenance periods.
Biography
Specialising in high power superconducting equipment, Frank Darmann has carried out
projects for the Ministry of Energy and Utilities, the Electricity Supply Association of
Australia, and has built an efficient transformer from high temperature superconductors. He is
currently testing a prototype saturable fault current limiter, measuring the high voltage
breakdown of materials in liquid nitrogen, designing high voltage bushings for use in liquid
nitrogen, and designing high power superconducting transformers using FEM techniques.
Frank Darmann, PhD graduated from Monash University with honours degrees in Physics (H,
1992) specialising in electro-magnetics, and Electrical Engineering (H, 1995) where he
specialised in power engineering and communications. His Engineering honours thesis was
on the subject of embedded generation in utilities. He completed his doctoral thesis in 2002
from the University of Wollongong in Electrical Engineering while working at Australian
Superconductors. He has presented his findings at over a dozen overseas and local
conferences to experts in the field of superconductivity.
Professor Tim Beales (PhD, BSc) is Manager Director of Australian Superconductors a
wholly owned subsidiary of SC Power systems, Inc. USA.
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Appendix 1. Derivation of a Model for the Terminal Inductance of a DC
Saturated SC-FCL using Laminations of type: Kawasaki 35RG155.
Step 1. Using a graphical method, the derivative of the magnetisation curve (B-H curve) of
the specified laminations was linearised over the range of H from zero to 30,000 A/m with an
adjustable step size to suit the curve – very dense near H = 0 (step size = 1 A/m) and very
sparsely when saturated near H = 1000 (step size = 1000 A/m). The piece wise linear curve
was then fitted to an equation using public domain software. The fitted expression is given by
Equation A1.
dB
 e a . H c . eb / H
dH
(A1)
Where a = 4.6377358, b= -12.751086, c = -2.6760766
Strictly speaking, this is not the incremental permeability, diff, in Equation 2; however, it
does serve as a good approximation and is suitable for predicting the overall behaviour of the
SC-FCL.
Equations A2 and A3 describe the inductance of a coil in an iron circuit which is biased by a
separate source.
L
n 2 A dB
.
l dH
(A2)
Where n = ac coupling turns, A = iron core cross sectional area, l = magnetic length, N = DC
ampere turns, and where:
1
H  ( N I dc  ni)
l
(A3)
where i = ac current, Idc = dc current
As can be appreciated, expressions A2 and A3 will lead to a terminal inductance which is a
strong function of the instantaneous current, I, that is, L = f (i). Equation A4 gives the full
analytical expression for the terminal inductance of the SC-FCL derived in this work.
L(i ) 
n2 A
l
 NI  ni c
 bl 
 bl 
NI  ni c

.
exp

(
)
(
) . exp 
a



. e .
l
l
 ( NI  ni) 
 ( NI  ni) 


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 (A4)




Where:
L is the overall inductance as a function of the instantaneous value of the current in the
copper coils,
i = i (t) is the instantaneous value of the current in the copper coils,
n = Turns in each of the copper linkage coils,
N = Turns in the DC HTS coil,
A is the area of the iron core circuit,
l = length of the wound magnetic circuit,
e = exponential function ~ 2.71828182
References
[1] Report of the SCENET Working Group, “Power Applications of Superconductivity - a
Roadmap for Europe”.
[2] The developer of the PSCAD software package is the Manitoba HVDC Research Centre
(a wholly owned subsidiary of Manitoba Hydro, Canada) of Cree Crescent, Winnipeg,
Manitoba, Canada R3J 3W1, http://www.hvdc.ca). A personal edition of PSCAD/EMTDC
with limited node capabilities can be downloaded from http://www.pscad.com for evaluation.
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