SWITCHING PHENOMENA IN MEDIUM VOLTAGE SYSTEMS GOOD ENGINEERING PRACTICE ON THE APPLICATION OF
VACUUM CIRCUIT-BREAKERS AND CONTACTORS
Ansgar Mueller
SiemensAG
P.O. Box 3240
91050 Erlangen
Germany
Dieter Saemann
SiemensAG
P.O. Box 3240
91050 Erlangen
Germany
Abstract - This paper summarises the transient phenomena
and overvoltages when switching with vacuum circuit­
breakers and vacuum contactors in medium voltage systems
(1 kV < U :s; 52 kV). It describes the interaction between
switching devices and network as well as the most influenc­
ing factors for the shape and magnitude of overvoltages.
Based on studies, laboratory and field tests, engineering
and mitigation guidelines are presented intended to cope with
transient overvoltage phenomena. It is the aim of this paper
to demonstrate that adequate surge protection ensures safe
and reliable system operation with vacuum switching de­
vices.
Index Terms
vacuum circuit-breaker, vacuum contactor,
switching transients, medium voltage, current chopping,
multiple reignition, surge arrester, RC circuit, protection.
-
I.
INTRODUCTION
During the recent decades, starting in the 1970ies, vacuum
replaced almost all earlier arc-quenching media for circuit­
breakers and contactors in medium voltage systems. Today
about 80% of new installations employ vacuum switching
technology, the others use SF6 gas. "Old" quenching princi­
ples, such as e.g. bulk-oil, minimum-oil, air-magnetic or air­
blast, almost disappeared from the market.
Fig. 1 Vacuum interrupter
In the public perception vacuum
circuit-breakers (VCB) and switch­
ing overvoltages sometimes are
seen as inseparable twins, al­
though the physical effects also
occur with all other switching
devices. Many incidents have
been reported where the break­
down of insulation was combined
with
high-frequency
switching
transients, said to be a character­
istic of vacuum breakers. Those
events triggered intensive re­
search all over the world.
Numerous investigations have been carried out to assess
the switching transients associated with vacuum circuit­
breakers and to design adequate surge protection. In fact,
the VCB was more intensively investigated since its market
entrance than any other breaker type before. The merit of
this intensive research is deep scientific knowledge about the
switching behaviour, the interaction between circuit-breaker
and network. On this basis concepts for surge protection
could be developed which resulted in higher safety and reli­
ability than with any other switching principle. This is why
most operators use VCB in their network today.
At the PCIC Europe Conference 2007 a paper [1] was
published which reported failures in oil & gas installations
where vacuum circuit-breakers were installed, it analysed the
causes and proposed surge protection measures. The pre­
sent article is intended as a follow-up. It discusses the basic
transient phenomena when switching of inductive or capaci­
tive currents and presents good engineering practice for
adequate insulation coordination and surge protection based
on theoretical investigations, laboratory and field tests.
II.
SWITCHING IN MEDIUM VOLTAGE SYSTEMS
Every switching operation causes more or less coined
transients. Inductive and capacitive circuits (Table 1) produce
different transients in terms of frequency, wave shape and
magnitude. In adjustable circuits both inductive and capaci­
tive conditions may occur depending on the relevant applica­
tion.
TABLE I
Switching duties in medium voltage systems
Type of circuit
Switching duties (examples)
Inductive circuits
Transformer (no-load)
Motor
Shunt reactor
Arc furnace
Arc suppression coil
Traction power supply
Capacitive circuits
Cable I Overhead line
Capacitor bank
(Harmonic) Filter circuit
Adjustable circuits
Generator
Power converter
In detail the switching transients depend on the
interaction between breaker and network,
influence of network parameters,
characteristics of the HV equipment connected.
Thorough system engineering must take into account three
consecutive steps when configuring surge protection for an
individual installation:
basic surge protection depending on the load
extended measures corresponding with the network
precautionary measures related to the industrial
(production) process.
The following sections describe the switching transients
and mitigation methods in principle.
III.
INDUCTIVE CIRCUITS
Thus the overvoltages on switching off unloaded transform­
ers are far below being critical. Switching tests on magnetiz­
ing currents from less than 1 A up to 20 A at operating volt­
ages between 6 kV and 30 kV yield overvoltage factors (see
Appendix) of maximum k = 3. 1.
When inductive circuits are switched, overvoltages can oc­
cur with all switching principles especially during switching off
[2]. Three possible causes are to be considered.
A.
B.
Current chopping
Multiple re-ignitions and virtual current chopping
When switching small inductive currents (approx. 20 A up
to some hundred amps), higher overvoltages can be gener­
ated if the arc reignites after the first current interruption, and
if the device is then able to interrupt the high-frequency tran­
sient current, which appears after the reignition. This process
always includes a transient reaction between the capacitance
on the system side and that on the load side. If it occurs
repetitively, it is defined as "multiple reignition". The voltage
increases with each reignition (voltage escalation), so that
high overvoltages can result.
Current chopping can occur, when switching off very small
inductive currents, for example up to approx. 20 A with
unloaded transformers. If the current is suddenly interrupted
(chopped) before the natural current zero, magnetic energy
remains in the iron of the disconnected circuit, proportional to
the chopped current, see Fig. 2. This magnetic energy is
discharged through the capacitance, mainly the cable capaci­
tance, see equation (1). Transposing it (2) estimates the
maximum overvoltage (first peak). As L and C are given by
the system, the only item that can be influenced is the magni­
tude of the chopped current ich' So, one objective is to de­
velop breakers with small chopping currents. However, the
chopping effect cannot be reduced at will, because intensive
arc quenching is required for other switching duties, such as
switching of capacitive currents.
U
(lJs]
t [ms]
I
I
I
I
10
Fig. 2 Current chopping during inductive switching
I
-·
2
L·
·2
I
=
where L
C
iCh
I
-·
2
C·
u
2
(1)
(2)
load inductance
load side capacitance
chopping current
UN
UL
Umax
P
P
,-
0.4
IOn
v
�
t
(lJs]
With an increasing voltage, caused by reignition, the cor­
respondingly high frequency transient current rises with each
reignition. If this transient current (iT in Fig. 5 and 6) is cou­
pled inductively / capacitively into the other two phases,
which are still carrying the power-frequency current, high­
frequency current zeros (is, iR) may also appear there. If the
breaker interrupts in one of these current zeros, this is called
virtual (induced) current chopping. The high-frequency cur­
rent only flows in the immediate vicinity of the switchgear,
whereas the load still carries the 50 Hz current. Current inter­
ruption at a HF current zero described above has the same
effect on the 50 Hz current as real current chopping at this
time. So, equation (2) mentioned in subclause A will also
estimate the height of this possible overvoltage.
system voltage
voltage at the load
maximum voltage at load
Probability of
occurrence
Ich Value of
chopping current
0.3
.--
0.2
f--
0.1
0
QI
v
d i electric strength
Fig. 4 Reignitions during inductive switching (principle)
With circuit-breakers, using gas for arc quenching, the de­
sign of the arc quenching unit determines the chopping cur­
rent. In contrast to this the chopping value of vacuum circuit­
breakers only depends on the contact material. Modern
contact material ranges at 3 to 4 A for circuit-breakers; and
for some contactors the value can even be below 1A.
0.5�
Recovering
I
0
2
3
r------,
4
Fig. 5 Coupling to the
phases S and R due to
multiple reignitions in
phase T
Ich
5
A
6
Fig. 3 Chopping current of vacuum circuit-breakers
2
3.
4.
5.
High-frequency
�
current zero in the �
last poles-to-clear,
caused by induction
C.
As the virtual current chopping can occur at high instanta­
neous values of the 50 Hz current, high overvoltages are
possible, which must be controlled by surge arresters.
500
f. (kHz]
t
100
so
....�
. ....-.
--c> eN network capacitance
(sum of cables connected to the busbar)
Range In which higher
overvoitages due to
multiple
re-Ignltlons
Fig. 8 Network and load capacitances on closing
The arrangement in Fig. 8 shows an inductive load, which
is connected to the switchgear by a cable. The load capaci­
tance Cc results from the cable, whereas the stray capaci­
tance of the load (e.g. transformer, motor) is some multiples
of ten less, so it can be neglected in this treatment. Many
cables are connected to the busbars, which results in a very
large network capacitance (CN --+ oc). Thus full transient
voltage appears at the load side of the breaker.
can occur.
10
5
0,5
1
2
3
4
o
0,5
5
10
Unloaded transformers
Arc furnace transformers
Neutral earthing transformers
with petersen coil
Motors beeing started
so 102
5
6
7
8
5·102 10' 5·10' 10'
Breaking
current IE<l
5·10'
Shunt reactors
Transformers with
short-circuited secondaries
System short-circuits
Short-circuits with reactors in series
Ll
Fig. 9 Voltages at time t1 and trace in first pole-to-close
In the time scale of the oscillations described in Fig. 9, the
power frequency voltage remains almost constant. At time t1
the line to earth voltage UL1 is at the maximum + 1 p.u. (Fig.
9 left) and the first closing pole is in line L 1. The pre-arcing
over the contact gap L 1 causes a travelling wave which runs
periodically along the cable. In the first pole-to-close, L 1, this
builds an oscillating voltage with an overvoltage factor k = 2
(Fig. 9 right). The travelling wave sees the inductive load as a
quasi open cable end, since the wave resistance of, for ex­
ample, motors and transformers lies in the region of 1000
Ohms, in contrast to 10 Ohms of cables. At the breaker, the
cable connection has a wave resistance approaching zero.
Thus the natural cable frequency fc = v I (4·1); where v =
wave velocity (approx. 150 m/lJs for XLPE cables), I = length
of cable.
Switching off a purely inductive load (e.g. motor dur­
ing run-up, locked rotor or jogging).
Opening of the first pole-to-clear < 0.5 ms prior to
current zero. This corresponds in three-phase sys­
tems to a probability of occurrence of 15 % at 50 Hz
1
(or 18% at 60 Hz) .
1 At 50 Hz: current zero occurs every 3.3 ms in one of the three
phases. Thus the probability is 0.5 ms / 3.3 ms
=
U [p.u.]
o
Multiple reignitions and virtual current chopping do not oc­
cur on every breaking operation or in every circuit. Figure 7
marks the range of higher overvoltage. The current (ab­
scissa) is given by load inductance and system voltage. The
frequency (ordinate) results from the load inductance and the
capacitance of the connecting cables (ranging from only few
meters to some hundred meters). Five prerequisites must be
simultaneously fulfilled for multiple reignitions plus virtual
current chopping:
2.
2
+1 p.u.
Fig. 7 Switching duties with inductive circuits
(example of 12 kV systems)
1.
Closing transients
Closing transients can occur with all switching devices, re­
gardless of the method of arc extinction and arc quenching
medium used. Every closing of an electrical circuit in high
voltage systems will cause pre-arcing with transient oscilla­
tions. The following theoretical treatment describes the clos­
ing conditions in the most unfavourable case, in which the
maximum transients occur. Damping and other effects are
neglected.
Virtual current chopping in the phases S and R
(principle, transients not drawn to scale)
Fig. 6
Build-up of the dielectric strength in the contact gap
lagging rise of transient recovery voltage (TRV).
Combination of supply side and load side capaci­
tances and inductances result in a di/dt value that
enables the breaker to quench at high-frequency
current zero.
The current to be switched off has to be smaller
than the limit current up to which virtual current
chopping may occur. For vacuum circuit-breaker
the value is I < 600 A.
0,15
3
the capacitor. Due to this, the voltage theoretically oscillates
to a value corresponding to the capacitor voltage + the in­
stantaneous system voltage (time t2), but this value is not
fully reached due to the existing system damping. Multiple
recharging (more reignitions) can generate very high switch­
ing overvoltages, so that the switchgear insulation may be
overstressed and flashover may occur.
L1
( ------------ ------------------------- ',I
��
' --,
L2
__
�
:
:
(�:::..: ------
.
/��;� );/7//�;]J;;7/�/7J;///��;;////
uN
Fig. 10 Oscillations in last poles-to-close before contact touch
Simultaneously with the closing of the first pole, an equalis­
ing oscillation of frequency fLoad is set up in the two last clos­
ing poles L2 and L3. This is represented by the circuit in Fig.
10. Through the bridging of the contact gaps in the last poles
to close, the oscillation fLoad moderates into an oscillation with
the natural frequency of the cable fc (Fig. 11) which then
oscillates into the instantaneous value of the operating fre­
quency voltage (half line-to-earth voltage). Overvoltages up
to k = 3 result.
2
UN
us
C
uN
Uc
Us
ic
System voltage
Capacitor voltage
Voltage across breaker
Capacitor current
LN
System inductance
C
Capacitance on load side
u
U [p.u.]
fc
o +---��----�_4_+�_+�-__.
f\
.. t
-1
-2
Fig. 12 Restrikes during capacitive switching
-3
Fig. 11 Voltage trace in last poles to close
For safe operation of capacitive circuits it is essential that
restrikes do not occur.
Related to the peak value of the line-to-earth voltage,
overvoltages with factor k = 2 can occur in the conductor of
the first pole to close and, in the conductors of the last poles
to close, with k = 3 resp. k' = 5 (bipolar). Theoretically, a
value of k = 6 is possible for the overvoltage between the
phases of the last poles to close. In practice, the stresses are
far less than the theoretical maximum values, because of the
statistical distribution of the overvoltages and damping ef­
fects.
IV.
B.
CAPACITIVE CIRCUITS
A distinction has to be made here between switching off
and switching on capacitive loads as different stresses may
arise.
A.
Closing onto and paralleling of capacitors
(back-to-back switching)
10
Making onto a short-circuit on entry into
the vortage maximum with maximum
rate of rise of the current
Switching off
2 Inrush current on back-to-back switching
When switching off at a current zero (time t1), the capacitor
remains charged at the peak value of the source voltage (uc).
The system voltage (UN) continues changing sinusoidally and
reaches its opposite peak value after 10 ms. The recovery
voltage (difference between Uc and UN) now starts rising
slowly. In this case, the stress is not caused by the rate of
rise as with inductive switching, but by the absolute value of
the voltage. If there is a new ignition within 5 ms after arc
quenching, this is called a reignition. This type of new ignition
is not dangerous. If the new ignitions take place after a de­
energized pause of more than 5 ms, they are called restrikes.
However, if they appear after approx. 10 ms, these restrikes
can be the origin of high switching overvoltages for the fol­
lowing reason: The restrike recharges the residual energy of
Fig. 13 Current variation when paralleling a capacitor
When the breaker contacts approach each other, pre­
arcing occurs across the open gap before galvanic contact.
At this moment, a transient process takes place between the
system and the capacitor, in which making currents up to
some 10 kA with frequencies up to several kHz can appear.
As pre-arcing occurs approx. 1 . . . 2 ms prior to galvanic
contact, the full transient current (2 in Fig. 13) flows through
the arc when switching on capacitors. In contrast to this,
when making onto a short-circuit (1) the instantaneous cur­
rent value at this time is much smaller. The dynamic forces of
high currents during pre-arcing may slow down contact
4
movement and hamper latching or promote contact welding.
So back-to-back capacitor switching is harder than making
onto short-circuit with the same values of current amplitude.
When configuring the application the maximum permissible
inrush current of the circuit-breaker must be considered.
V.
1,0
0,9
0,8 -I----Jl�-----I
INFLUENCE OF THE NETWORK
Switching transients do not only depend on the nature of
the load circuit but also on the network on the feeding side.
The network configuration influences the shape and the
magnitude of transients on the load and busbar side of the
breaker. Thus the network as a whole also determines which
kind of surge protection must be installed.
A.
-,;:;-:-----=--.--,---,--,
0,5
�
:::.:..:.
==���
:::.:....=..::=::;..:::....:::
-"
-t-<I-----'-;=::..:
°
Fig. 15
Cables and travelling waves
C.
Where impulse voltages occur, the voltage along a cable is
no longer constant. The cable then acts as a travelling-wave
conductor. This effect is obtained when the impulse front time
is shorter than the travelling time in the cable; i.e. in mathe­
matical terms when the condition T « (II v) is met, where T =
front time of impulse; I = cable length; v = propagation veloc­
ity of the wave in the cable. Depending on the cable type, v =
0.3 to 0.6·c (c = velocity of light 300 m/lJs). Cables with a
length of less than 30 . . . 40 m ("short cables") do not there­
fore act as travelling-wave conductors in normal switchgear
installations as regards the phenomena described here.
UBb
UM
Zc
ZM
busbar voltage
y
refractive index
5
10
15
20
Share of load side voltage as a function of the
number of cables connected to the busbar
Switching status of the busbar
The location of cables connected to the busbar also affects
virtual current chopping. Chopping is more likely if high­
frequency and power-frequency currents share a long com­
mon loop along the busbar between incomer and panel to be
switched off (Fig. 16, left). In contrast, cables connected
outside the power-frequency current flow (Fig. 16 right) re­
duce the coupling along the busbar and thus the probability
of virtual chopping in that the cable capacitances form a
bypath for the high-frequencies. Altogether, the probability of
occurrence of overvoltage may widely vary with one and the
same load circuit, depending on the switching status.
load voltage
Cable impedance
�i50Hz
�150Hz
Loadimpedance
(e.g. motor)
r= UBb = ZM -Zc
1
+
ZM +Zc
UM
Fig. 14 Voltage rise at the load due to travelling waves
Ignition of the contact gap is a matter of a few nanosec­
onds and causes a very steep fronted voltage variation to
take place on both sides of the contact gap within a few
microseconds. Reignition thus gives rise to a travelling wave.
The surge impedance changes at the transition from the
cable to the inductive load, e.g. a motor (ZM » Zc) so that
the travelling wave is reflected and a voltage rise occurs at
the motor terminals. Fig. 14 shows the relationships and the
definition of the refractive index. Theoretically, the latter can
reach a value of y = 2, but in practice (field and test lab
measurements) only values up to y = 1.7 have occurred,
where the rise times of the surge are in the order of 1 IJs.
B.
Incomer
High probability
Fig. 16
Incomer
Low probability
Probability of virtual current chopping dependent on
the panel arrangement
VI.
INFLUENCE OF THE LOAD
In view of the impact on the insulation, the magnitude and
wave shape alone are only one part of the decisive factors.
The load itself plays the other part of the role. As all inductive
loads consist of windings whose design (shape of the con­
ductor, arrangement of the turns, type of main and interturn
insulation) determine the stress on the insulation and which
parts of it may be particularly at risk. As motor insulation is
more susceptible to overvoltage (compared to other inductive
loads), it is taken as example in the following.
Cables connected to the busbar
Fig. 15 shows the equivalent circuit of an installation in
which the cables connected to the busbar are represented by
their surge impedance Ze. On ignition of the contact gap, the
impulse waves occurring on the right and left of the gap
depend on the surge impedance values.
As can be seen from the graph for the load side voltage UL
the impulse voltage stressing of the load side increases with
the number of cables connected to the busbar. Reversely,
the voltage stress on the busbar itself rises with only few
cables connected. There are configurations where surge
protection is also required for the busbar side (incoming
feeder) of the switchgear installation.
A.
Voltage distribution within windings
Within a winding the distribution of steep surge voltages is
not linear and depends on the rate-of-rise of the overvoltage.
Because of the propagation effect the voltage across the coil
5
TABLE II
depends mainly on its conductor length. Fig. 17 shows the
example of a 10 kV motor, where the voltages across the line
end coils dropped to noncritical values after the third coil. As
regards stressing of the motor insulation, a distinction must
be made between the unipolar component of travelling waves
(which stresses the motor winding with respect to earth) and
the bipolar voltage range of the travelling wave (voltage
surge) by which the interturn insulation of the line end coils in
particular is stressed. With rise times in the order of 1.7 IJs,
recorded in numerous tests - see section VII, A - the first coil
is stressed with 65% of the voltage at the terminals. The
same may apply in principle to transformer windings.
Laboratory and field tests on inductive load switching
40+----!--t-:=O_;;o;;::l:--1
3rd ceil
,7
2
!,-S
4
Test object
Subject of
investigation
Berlin
Factory
Test lab
Motors
10 kV
Mannheim
FGH Test lab
(PEHLA)
Motors
6 kV
Fundamental research on overvoltage phenomena, i.e.
multiple re-ignitions
and virtual current
chopping
Heilbronn
Thermal
power station
Isar II
Nuclear
power station
Pleinting
Thermal
power station
V 61klingen
Thermal
power station
Otlenhofen
Substation
Berlin
Factory
Test lab
Meitingen
Steel factory
A.
T1_
Fig. 17 Voltage distribution across the coils of a motor wind­
ing under synthetic impulse voltage
B.
Installation
Research on the
Motors
6 kV and
10 kV
Shunt reactor
30 kV
influence of network
parameters
Investigation of the
dielectric stress on
equipment in real
installations
Verification of various surge protection
methods
Transformers
6 kV to 30 kV
Overvoltages on
no-load switching
Capacitor
banks 20 kV
Inrush currents,
breaking currents
Arc furnace
transformer
Surge protection
methods
The following focuses on inductive switching and summa­
rises some major issues of overvoltage test results and the
corresponding surge protection when investigating the
switching of motors with vacuum circuit-breakers.
20+----!---'--+-----1-----1
o
Site
Occurrence of multiple re-ignitions
Not every breaking operation is associated with multiple
reignitions and severe overvoltages. Fig. 18 shows the result
of 914 breaking operations with a 10 kV motor under worst
case conditions [3]. 148 of the 914 breaking operations were
followed by multiple reignitions, this is 16 %. This percentage
corresponds very well with the prerequisite No. 2 (refer to
section III. D.) predicting a probability of 15 % (at 50 Hz).
Resonances within windings
Windings have characteristic natural resonances because
of the inductive and capacitive coupling of the individual
winding turns with one another and to earth. The frequency
response of this system is determined by the geometry of the
winding. External pulses may excite resonant oscillations.
When the exciting pulse frequency equals a resonance fre­
quency, this may cause impermissible voltage rises at certain
points inside the winding while the voltage is still in the per­
missible range at the external terminals.
If the repetition frequency of multiple reignitions is within
the range of the resonance frequencies, there is the risk of
resonant oscillations in the winding giving rise to high internal
overvoltages. This can only be prevented by e.g. means of
an external RC damping circuit, not by surge arresters at the
outer terminals.
VII. LABORATORY AND FIELD TEST
Besides theoretical investigations numerous practical tests
have been carried out (in addition to the standard type tests)
on inductive and capacitive load switching since the early
development of the vacuum circuit-breaker. In regular inter­
vals some tests are retaken in order to check and ensure
constant characteristics of vacuum circuit-breakers and vac­
uum contactors.
Fig. 18 Test series of 914 breaking operations of a 10 kV I
400 kW motor during starting (every vertical line
stands for an operation with multiple re-ignitions)
6
B.
A.
Probability of occurrence of overvo/tage peaks
Well proven techniques to cope with surges is to install
- surge arresters to limit the magnitude of surges,
- surge capacitors to attenuate their rate-of-rise,
RC circuits to damp high-frequency transients and pre­
vent repetitive impacts.
When the overvoltage phenomena associated with motor
switching were investigated, in total 2500 switching tests
were carried out using different system configurations (cable
types I lengths, motor currents, busbar arrangements) and
surge suppression measures in order assess the dielectric
stress and to verify the most effective protection methods.
The above mentioned 914 cycles - using the worst-case
configuration - were evaluated to back up statistically the
reliability of the system referred to as "vacuum circuit-breaker
fitted with surge arresters". Multiple reignitions occurred in 16
% of the start-stop cycles and virtual current chopping in 14
%. It might be assumed that each multiple reignition or virtual
current chopping went along with unduly high overvoltage.
However, this is not true. The magnitude of overvoltage also
is subject to statistical distribution.
i�9
90
(J
In general there is consensus among professional experts
to use these protection elements, and corresponding rec­
ommendations were published at earlier conferences, refer
for example to [4], [5]. In detail, however, there is a margin of
discretion which measures shall be applied so that recom­
mendations may vary. Insulation coordination must weigh
technical and economical aspects and conclude in a justifi­
able balance between efficient protection for reliable service
and acceptable costs. A guideline to configure surge protec­
tion is given in the following.
I
98
. %
95
B.
Maximum values of unipolar and -
-- - -
bibolar overvoltages recorded at
- -
worst-case configuration
-- 1--- -- -- 1---
TABLE III
Standard surge protection for MV switching duties
50
f--Unipolar
(J
10
5
2
o
Fig. 19
l!I---L--
�
V '\
,
\
r t---Bipolar
38kV
20
60
I
80 kV 100
Switching duty
Recommendation
Distribution and/or
power transforme(
No surge protection for normal distribution
transformer.
¢ For unit
Surge arresters to be fitted in case of:
- connection to overhead line
transformers see
-- t- --
molor, converter
or furnace
Motor
jReference probability
�59k�\
40
--
Basic protection measures - Step 1
Table III lists common switching duties in MV systems and
the corresponding recommendations
914 switching tests using the
-
Means of overvo/tage protection
- insulation level is "List 1", IEC 60071-1
- insulation is aged or its level is unknown
- switching rate in normal service is high
Surge arresters if Istart
<
600 A;
Additionally RC circuits are fitted:
- if insulation is not in line with IEC 60034-15
- if insulation is aged or its level is unknown
- if intended switching rate is high
- at neutral of starting auto-/transformer
D�
Cumulative probability of occurrence of
overvoltage peak values (10 kV motor)
Fig. 19 shows the cumulative probability of occurrence of
the unipolar and bipolar voltages measured at the motor
terminals. Out of 914 tests only 2% (reference value for sta­
tistical insulation coordination to IEC 60071) caused over­
voltages higher than 38 kV (unipolar) or 59 kV (bipolar).
VIII. ENGINEERING AND MITIGATION GUIDELINES
Surge protection should consider three consecutive steps.
- Step 1 is the basic protection required by the nature of
the load circuit.
- Step 2 considers the network configuration as whole
and special modes of operation. These may require
additional measures, e.g. protection on the busbar or
feeding side (incoming feeder).
- Step 3 takes into account further criteria related to the
industrial process. Beyond the switching duty and net­
work conditions, other aspects may predominate and
advise protection as a precautionary measure.
Generator
Surge arresters if short-circuit current
Ik < 600 A; for RC circuits ¢ see "Motor"
Converter
transformer
Surge arresters
Shunt reactor
RC circuits (individual configuration required),
surge arrester if I < 600 A
Short-circuit
limiting reactor
Surge arresters
Petersen coil
Surge arresters
Furnace
transformer
Surge arrester + RC circuits
(individual configuration required)
Capacitive circuits
If breaker corresponds with class 2 to IEC
62271-100, no surge protection is required
Overhead line
or cable
Circuit-breaker of class C1 is suitable; however, class C2 is preferable.
Capacitor bank
(back-to-back)
Circuit-breaker of class C2 only.
Consider inrush current limit and increased
thermal loading by harmonic currents.
Filter circuit
Circuit-breaker of class C2 only.
Consider voltage rise due to harmonics.
2
7
In this table "distribution transformer" means that all downstream
feeders are equipped with a separate breaker (switch); hence in
normal service the transformer is switched under no-load condition.
The inductive switching duties which require protection cor­
respond to those which fall in or touch the marked range of
Fig. 7, i.e. where higher overvoltages caused by reignitions
and virtual current chopping are likely to occur. Capacitive
circuits do not require surge protection due to nature of the
transients if the breaker is restrike-free.
C.
the voltage rise at the motor terminals due to travelling
waves. Suitable surge arresters for this kind of motor protec­
tion are available up to approx. 15 kV operating voltage.
E.
Surge capacitors
They reduce the rate-of-rise of surges by increasing the
line-earth capacitance, and are mainly used to protect
against "non-switching" surges. For example, a frequent
usage is the protection of open transformer secondary termi­
nals against surges transferred from the primary side [6].
For inductive load circuits, however, where multiple reigni­
tions may occur, pure capacitors are considered less favour­
able. Located close to the breaker they may promote virtual
current chopping, because the concentrated capacitances
couple the phases with low-impedance for high-frequency
transients. By contrast, RC circuits avoid the unwished, un­
damped phase coupling by the inbuilt resistance.
Surge arresters
Arresters limit the overvoltage peak value and suffice most
applications, i.e. where the rise time of the overvoltage and
the excitation of resonant oscillations are not critical or no
issue. When configuring surge arresters the nature of the
overvoltage must be taken into account. As switching opera­
tions initiate transient oscillations with different instantaneous
polarity of the voltages in the phases (differential mode over­
voltage), the voltage between them can add up to higher
values than to earth. Arresters must limit the overvoltages
between the phases (L-L) just as phase-to-earth (L-E). The
objective is to achieve this with 3 arresters in phase-earth
connection. If this is not possible, one needs a 6 arrester
format; see Fig. 20.
F.
RC circuits
RC-circuits form a bypass for high frequency currents (and
voltages) in that the transients are diverted to earth instead of
reaching the equipment to be protected. They damp travel­
ling waves and prevent their refraction at the load-side cable
terminals thus reducing repetitive stress of the insulation or
the excitation of resonance in windings in case of multiple
reignitions. However, due to its resistor RC circuits do not
limit the peak value or the rate-of rise of a surge.
Cable (bar)
u
res
-
2
1.15
res
-
n
: ::
�
u < O.83·Up
<�. O.83·Up
LOW-impedanc
dual earthing
1.15
_
Fig. 20 Protection by surge arresters
Fig. 22 RC circuits for susceptible equipment
Ures
Up
The RC installation at the load is most economical since its
surge capacitor then is smallest, whereas installation at the
other cable end requires larger capacitors. As capacitors
"suck" higher frequencies, it is essential to design the RC
circuit also according to the harmonic content prevailing in
the installation, i.e. considering the heat losses.
In any case, feed and earth cable must be low-inductance
connections, with a time constant L / R :5 50 ns (where L is
the total inductance of RC circuit and cable, R is the resis­
tance of the RC circuit).
0.83
y,
D.
residual voltage of arrester ("protection level")
rated lightning impulse withstand voltage of equipment
reduction factor for switching impulse voltage (due to pro­
longed time to half value of wave tail)
since the line-line overvoltage can rise to twice the line-earth
voltage; the arresters must limit the overvoltage to half the
permissible line-line value.
Special arresters for motors
Normally surge arresters are installed at the equipment to
be protected. However, arresters are susceptible against
vibration and high ambient temperatures prevailing in a motor
terminal box. Moreover, arresters would require a large ter­
minal box for which most often there is no space.
G.
The standard surge protection described in the previous
sections ensures trouble-free operation regarding the load
side. For the major part of all applications the basic protec­
tion suffices all requirements. However, there are applica­
tions which require further consideration.
Step 2 considers the network as a whole. Regarding
switching transients, arrangements with low line-earth capaci­
tance on the busbar side (few number of cables connected,
and/or of short length) may require additional surge arresters
on the busbar or at the incoming feeder. In rare cases also
RC circuits may be necessary on the busbar side.
It should also be considered that overvoltages are not only
caused by switching actions. There may be sources like
overhead line connections, networks with frequent fault tran­
sients (earth fault, short-circuit) due to construction activity,
U2",1,7·U1
U2
Additional protection measures - Step 2
at 1 IJS rise time
of overvoltage
Fig. 21 Special surge arresters for motors can be safely
installed in the switchgear panel
So it is recommended that arresters be fitted at the breaker
or contactor (within the switchgear assembly), regardless the
cable length. There is no loss of the protective effect if the
residual voltage of the arrester is tuned low enough, covering
8
triggered a debate on the application of vacuum switching
technology. Mr Tastet encouraged this paper.
mobile cable installations (e.g. open cast mining) etc. Last
but not least there may be transients capacitively or induc­
tively transferred from other systems, so that, overall, surge
protection is required in any case.
H.
XI.
Precautionary protection measures - Step 3
[1]
Step 3 goes beyond the switching duty and network and
takes into account all the boundary conditions relevant for the
individual installation. In fact, other aspects than the switch­
ing duty may predominate and may advise to invest in addi­
tional protection. Table IV summarizes some issues. Not all
conditions and operational actions occurring during the ser­
vice life may be known in advance. During planning only the
foreseeable effects can be considered. There may be a high­
grade industrial (production) process, where the loss of
power supply causes safety problems or unduly consequen­
tial damages. For these applications individually configured
protection ensures safe operation and is quasi an insurance
for the electrical investment.
[2]
[3]
[4]
TABLE IV
Conditions requiring customised solutions
Special conditions
Examples
High frequency of occurrence of
overvoltage
High switching rate, frequent
earth faults, surges transferred
from the infeeding network
Low insulation level of
equipment
Dry-type transformers;
insulation level "List 1"
Adverse ambient conditions
Site altitude, lightning
Degradation of insulation
Aging, environment (salt, dust)
System configuration
Panel arrangement
Operational conditions
Changes during service life,
switching actions during commissioning or maintenance
[5]
[6]
REFERENCES
J. Tastet, P. Angays;
Safe Implementation of HV Vacuum Switches in Oil &
Gas Installations;
PCIC Conference Record, 2007
F. Battiwala, H. Fink, M. Rimmrott, W. Schultz; "3AF
Vacuum Circuit-Breakers and 3TL Vacuum Contactors
in System Operation", Siemens Power Engineering,
Vol. lll (198 1), Special Issue "Vacuum Switchgear for
Voltages up to 36 kV"
A. Luxa, A. Priess; "Switching of Motors During Start­
Up", Siemens, Power Engineering & Automation VII
(1985), Special Issue "Medium Voltage Technology"
(ISBN 3-8009-384-80)
C. Vollet, B. de Metz Noblat; "Protecting HV Motors
against switching overvoltages", PCIC Conference Re­
cord, 2007, PA 15
D. Penkov et al.; "Overvoltage protection study on
vacuum breaker switched MV motors", PCIC Confer­
ence Record, 2008, WE 16
J. C. Das, "Surge transference through transformers",
IEEE Industry Applications Magazine, SepUOct 2003,
pp 24-32.
XII. APPENDIX
U [p.u.]
1
Unipolar overvoltage
Uunipolar
ÛLE
Ubipolar
k
=
t
-1
IX.
CONCLUSION
k'
X.
network
voltage
(line-line)
ULE Nnetwork voltage line-earth U",,"3
Fig. 23
..fi
.J3 ·Um
Bipolar overvoltage
Urn Maximum
The interaction between breaker and network results in
transient phenomena, where every breaker type has different
and distinct characteristics on inductive or capacitive currents
depending on its arc quenching principle. Vacuum circuit­
breakers may cause switching transients in a certain range of
inductive currents that require surge protection. Those tran­
sient phenomena and their boundary conditions are very well
investigated and necessary mitigation techniques are avail­
able. The established surge protection methods suffice all
applications in MV networks: standardised protection can
cover most of the power supply installations. Susceptible
industrial processes under specific network conditions are
served with customised surge protection. Meanwhile three
decades of service experience prove that these protection
methods safely govern all transients and ensure trouble-free
operation.
Switching phenomena are only one facet of vacuum
switching technology. In fact the vacuum circuit-breaker
spread over the market due to its best overall performance in
view of mechanical and electrical requirements, endurance,
reliability and maintainability.
UUJlipolar
Ub;potar
..fi
.J3 ·Um
Definition of overvoltage factors k and k'
XIII. VITAE
The lead author is Ansgar Muller of Siemens. He received
his Diploma in Electrical Power Engineering 1982 at Tech­
nische Hochschule Darmstadt, Germany. Since 1983 he has
been working for Siemens in the energy distribution division,
where he conducted studies on medium voltage switching
transients, insulation coordination, electromagnetic compati­
bility and electromagnetic fields. He is secretary of the stan­
dardisation committee CENELEC TC 17AC High-Voltage
Switchgear and Controlgear, member of the German NC and
IEC working groups related to HV switchgear.
Dr. Dieter Samann is Business Development Manager for
Medium Voltage Components at Siemens AG. After studying
at Technical University Braunschweig he was technical assis­
tant with receipt of Doctor's degree in electrical engineering
in the field of vacuum switches in 1984. At Siemens he is an
expert regarding vacuum switching technology and has pub­
lished a lot of articles about vacuum switches and their be­
haviour.
ACKNOWLEDGEMENTS
The authors would like to thank Jacques Tastet and Phil­
lippe Angays, Technip, whose conference contribution [1]
9