05 overvoltages and insulation coordination

390

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Industrial electrical network design guide T & D 6 883 427/AE

5. OVERVOLTAGES AND INSULATION CO-ORDINATION

Different types of overvoltage may occur in industrial networks. Devices must therefore be

installed to reduce their magnitude and the insulation level of equipment must be chosen so

that fault risks are reduced to an acceptable level.

5.1. Overvoltages

An overvoltage is any voltage between one phase conductor and earth, or between phase

conductors having a peak value exceeding the corresponding peak of the highest voltage for

equipment, defined in standard IEC 71-1.

An overvoltage is said to be of differential mode if it occurs between phase conductors or

between different circuits. It is said to be of common mode if it occurs between one phase

conductor and the frame or earth.

n origin of overvoltages

Overvoltages can be of internal or external origin.

o internal origin

These overvoltages are caused by a given network element and only depend on the

characteristics and structure of the network itself.

For example, the overvoltage that occurs when a transformer's magnetizing current is

interrupted.

o external origin

These overvoltages are caused or transmitted by elements outside the network, for example:

- overvoltage caused by lightning

- spread of HV overvoltage through a transformer to the internal network of a factory.

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n classification of overvoltages

Standard IEC 71-1 gives the classification of overvoltages according to their duration and form.

According to the duration, a distinction is made between temporary overvoltages and transient

overvoltages:

- temporary overvoltage: power frequency overvoltages of relatively long duration (from

several periods to several seconds).

- transient overvoltage: short-duration overvoltage lasting only several milliseconds, which

may be oscillatory and is generally highly damped.

Transient overvoltages are divided into:

. slow-front overvoltage

. fast-front overvoltage

. very-fast-front overvoltage.

n standard voltage forms

Standard IEC 71-1 gives the standardised wave forms used to carry out tests on equipment:

- short-duration power frequency voltage: this is a sinusoidal voltage with a frequency

between 48 Hz and 62 Hz and a duration equal to 60 s.

- switching impulse: this is an impulse voltage having a time to peak of 250 µs and a time to

half-value of 2500 µs.

- lightning impulse: this is an impulse voltage having a front time of 1.2 µs and a time to

half-value of 50 µs.

n consequences of overvoltages

Overvoltages in electrical networks cause equipment degradation, a drop in service continuity

and are a hazard to the safety of persons.

The consequences can be very varied depending on the type of overvoltages, their magnitude

and their duration. They are summed up as follows:

- breakdown in the insulating dielectric of equipment in the case where the overvoltage

exceeds the specified withstand

- degradation of equipment through ageing, caused by non-destructive but repetitive

overvoltages

- loss of power supply caused by the destruction of network elements

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- disturbance of control, monitoring and communication circuits by conduction or

electromagnetic radiation

- electrodynamic stress (destruction or deformation of equipment) and thermal stress

(elements melting, fire, explosion) essentially caused by lightning impulses

- hazard to man and animals following rises in potential and occurrence of step and touch

voltages.

5.1.2. Power frequency overvoltages

Power frequency overvoltages are generally caused by:

- an earth fault

- resonance or ferro-resonance

- neutral conductor breakdown

- a generator voltage regulator or transformer on-load tap changer fault

- overcompensation of reactive energy following a varmeter regulator fault

- load shedding, notably when the supply source is a generator

5.1.2.1. Overvoltage caused by an earth fault

Overvoltages caused by the occurrence of an earth fault greatly depend on the neutral

earthing system of the given network.

n unearthed (MV or LV) or impedance earthed (MV) neutral

Figure 5-1 shows that on occurrence of a solid earth fault, the voltage between the neutral

point and earth becomes equal to the single-phase voltage:

V VNeutral n=

Vn : nominal single-phase voltage

For a fault on phase 1, V VNeutral = − 1 .

The phase-earth voltage of healthy phases thus becomes equal to the phase-to-phase

voltage:

V V V V VE Neutral2 2 2 1= + = −

V V V V VE Neutral3 3 3 1= + = −

whence V V VE E n2 3 3= =

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2

3

1

V3

V2

V1

earth

fault V E2 V E3V E1VNeutral ZNeutral

V2 V3

V1

Neutral

V E2 V E3

V E1 0

V1 , V2 , V3 : phase-neutral voltages

V E1 , V E2 , V E3 : phase-earth voltages

ZNeutral : earthing impedance ( ZNeutral = ∞ for an unearthed neutral)

Figure 5-1: overvoltage on an unearthed or impedance earthed network

on occurrence of a phase-to-earth fault

Note 1 : for an impedance earthed neutral, the value of ZNeutral is much greater than the value of the

transformer and cable impedances and the fault resistance, which is why V VNeutral = − 1 .

Note 2 : in overhead public distribution networks, there are highly resistive faults (several kΩ), having avalue close to or higher than the earthing impedance. In this case, a highly resistive fault will

cause an overvoltage lower than 3Vn .

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n solidly earthed neutral (HV or MV)

On occurrence of an earth fault on one network phase, a high current is generated which

circulates in the circuit formed by the fault phase, earth and neutral earth electrode

(see fig. 5-2).

At the fault point, the three-phase voltage system is disturbed. The fault phase voltage in

relation to earth is almost zero if we neglect the fault resistance. The voltages of the other two

phases in relation to earth are higher than the single-phase voltage, while remaining lower

than the phase-to-phase voltage.

V E3

V3ZT

ZT

ZT

ZC

ZC

ZC

Rf

V2

V1

fault

Re

V E1 V E2

V1 , V2 , V3 : single-phase voltages

ZT : transformer impedance

ZC : cable impedance

Re : neutral earth electrode resistance

R f : fault resistance

Figure 5-2: equivalent diagram of a phase-earth fault when the neutral is solidly earthed

Thus, we can define an earth fault factor k characterising the phase-earth overvoltage

occurring on the healthy phases:

V V k VE E n2 3= =


The symmetrical component calculation method (see § 4.2.2. of the Protection guide) can be

used to determine the value of k in relation to the positive, negative and zero-sequence

impedances:

kZ a Z a Z

Z Z Z R f= −

+ ++ + +

13

2

2 0

2 0

(1) ( ) ( )

(1) ( ) ( )

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In most networks, generators are sufficiently far away to take the approximation Z Z(1) ( )= 2 ; we

thus have:

( )k

a Z Z

Z Z R f= +

−

+ +12 3

1 0

1 0

( ) ( )

( ) ( )

Nomographs can be used to determine factor k for a zero fault resistance ( R f = 0 ) in relation

to the ratios R

X

( )

( )

0

1

andX

X

( )

( )

0

1

for R( )1 0= and R X( ) ( ).1 10 5= (see fig. 5.3. et 5.4.).

where:

R( )1 : positive-sequence resistance seen from the fault point

X( )1 : positive-sequence reactance seen from the fault point

R( )0 : zero-sequence resistance seen from the fault point

X( )0 : zero-sequence reactance seen from the fault point

When the fault resistance is not zero, we can see in the formula expressing k that the

overvoltage is weaker. The calculation of the overvoltage with a zero fault resistance thus

provides an excess value.

If we again use the diagram in figure 5-2, we can determine these impedances for a practical

case:

by taking:

Z R j X

Z R j X

Z R j X

Z R j X

T T T

C C C

T T T

C C C

= +

= +

= +

= +

positive - sequence impedances

zero - sequence impedances( ) ( )

( ) ( )

0 0

0 0

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we can determine:R R RT C( )1 = +

X X XT C( )1 = +

R R R Re T C( )0 3= + +

X X XT C( ) ( ) ( )0 0 0= +

Note: A factor 3 appears before Re . The reason for this is explained in figure 4-11 of the Industrial

network protection guide.

8

7

6

5

4

3

2

1

1 2 3 4 5 6 7

8

k = 1.7

k = 1.6

k = 1.5

k = 1.4

k = 1.3k = 1.2

R

X (1)

(0)

X

X(1)

(0)

Figure 5-3: earth fault factor in relation to ratiosX

X

( )

( )

0

1

andR

X

( )

( )

0

1

for R( )1 0= and R f = 0

8

7

6

5

4

3

2

1

1 2 3 4 5 6 7

8

k = 1.7

k = 1.6k = 1.5

k = 1.4

k = 1.3

k = 1.2

X

k = 1.5

R

X (1)

(0)

X (1)

(0)

Figure 5-4: earth fault factor in relation to ratiosX

X

( )

( )

0

1

andR

X

( )

( )

0

1

for R X( ) ( ).1 10 5= and R f = 0

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o example

Let us consider a YNyn, 33 kV/11 kV transformer with a power rating of S MVAn = 24 (see

IEC 909-2 table 3 A) supplying a network with 240 mm² aluminium cables the longest outgoing

feeder of which is 5 km. The neutral earth electrode resistance is 0.5 Ω.

- transformer characteristics:

Usc = 24 2. %

R

X

T

T

= 0 046.

X

X

T

T

( ).

00 7=

we can deduce( )

X UU

ST sc

n

n

= × = ××

×=

2 3

60 242

11 10

24 10122. . Ω

RT = 0 056. Ω

X T( ) .0 0 85= Ω

Note: the value of Usc is extremely high in relation to the transformers feeding a network with a

limiting resistor earthed neutral. The transformer here is a United Kingdom transformer adaptedto the solidly earthed neutral system.

The short-circuit voltage has been chosen high on purpose so as to minimise the short-circuit

current. Indeed, if Usc is high, the valueR

X

( )

( )

0

1

is minimised ( )since X X XT C( )1 = + , which

decreases the overvoltage factor (see fig. 5-3 and 5-4).

- cable characteristics:

RL

SkmC = = × =ρ 0 036 1000

240015

.. /Ω

X kmC = 01. /Ω

We assume that X X kmC C( ) . /0 3 0 3= = Ω .

Note: the value of X C( )0 is highly variable (from 0.2 to 4 X( )1 ) depending on what the cable is made

of and the return via the earth (remote earth, screen or earthing conductor).

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For a solid fault ( R f = 0 ) at the transformer terminals:

R RT( ) .1 0 056= = Ω

R R Re T( ) . . .0 3 3 0 5 0 056 156= + = × + = Ω

X XT( ) .1 122= = Ω

X X T( ) ( ) .0 0 0 85= = Ω

whence R X( ) ( ).1 10 05 0= ≅

R

X

( )

( )

.0

1

128=

X

X

( )

( )

.0

1

0 70=

Figure 5-3 shows that k is between 1.4 and 1.5.

For a solid fault ( R f = 0 ) 5 km away from the transformer:

R R RT C( ) . . .1 0 056 015 5 0 81= + = + × = Ω

R R R Re T C( ) . . . .0 3 3 0 5 0 056 015 5 2 31= + + = × + + × = Ω

X X XT C( ) . . .1 122 01 5 1 72= + = + × = Ω

X X XT C( ) ( ) ( ) . . .0 0 0 0 85 0 3 5 2 35= + = + × = Ω

whence R X( ) ( ).471 10=

R

X

( )

( )

.0

1

134=

X

X

( )

( )

.0

1

137=

Figure 5-4 shows that k is between 1.2 and 1.3.

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n TN earthing system

The current of an earth fault circulates in the protective conductor. The neutral earth electrode

resistance is thus not used to determine the zero-sequence impedance

(see fig. 5-5).

V3

ZT

ZT

ZT

ZC

ZC

ZC

V2

V1V VM2

V VM3

VMZPERe

V1 , V2 , V3 : single-phase voltages

ZT : transformer impedance

ZC : cable impedance

ZPE : protective conductor impedance

VM : potential of exposed conductive parts (masses) in relation to earth

Re : neutral earth electrode resistance

Figure 5-5: equivalent diagram of an earth fault in a TN earthing system

We are interested in the overvoltage of the healthy phases in relation to the exposed

conductive part, which determines whether or not an insulation fault may occur on the other

load: kV V

V

V V

VM

M

n

M

n

= − = −2 3 .

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For a transformer or a cable in low voltage, we can take the zero-sequence impedance to be

approximately equal to the positive-sequence impedance: Z ZT T( )0 = and Z ZC C( )0 = .

We thus have Z Z Z ZT C PE( )0 3= + +

Z Z ZT C( )1 = +

whence ( )ka Z

Z Z Z

a Z

Z Z ZM

PE

T C PE

PE

PE T C

= −+ +

= −+ +

13

31 for a solid fault ( R f = 0 )

a ej

=2

3

π : rotation operator of 120°

The overvoltage will be maximum when ZT is negligible compared with Z ZPE C+ , which is the

case for a long length cable.

Thus ka Z

Z ZM

PE

PE C

≤ −+

1

kM will be maximum when the protective conductor cross-sectional area is as small as

possible, i.e. equal to half the phase conductor cross-sectional area; thus R RPE C= 2 .

For an aluminium cable cross-sectional area smaller than 120 mm², the reactance can be

neglected compared with the resistance, which thus gives us:

Z

Z Z

R

R R

PE

PE C

PE

PE C+≅

+= 23

since R RPE C= 2

whence k aM ≤ −1 2

3

k jM ≤ − − +

1

2

3

1

2

3

2

kM ≤1.45

We can show that for a cable with a large cross-sectional area (> 120 mm²), the overvoltage

will be lower than in the case of a small cross-sectional area.

401




n TT earthing system (see fig. 5-6)

V3

ZT

ZT

ZT

ZC

ZC

ZC

V2

V1I f

I f

VN

load 1 load 2

RM1VM1 RM2Re

Re : substation earth electrode resistance

RM1 : load 1 and fault load earth electrode resistance

RM2 : load 2 earth electrode resistance

VM1 : load 1 and fault load phase-to-earth voltage

Figure 5-6: equivalent diagram of an earth fault in a TT earthing system

We want to know the overvoltage of the healthy phases in relation to the exposed conductive

part, which determines whether or not an insulation fault may occur on the other load:

kV V

V

V V

VM

M

n

M

n

=−

=−2 3

In low voltage, the neutral and load earth electrode resistances are very high in relation to the

transformer and cable impedance ( ZT and ZC are roughly several tens of mΩ).

We can thus write that the fault current is:

IV

R Rf

e M

=+1

1

and ( )Z Z IT C f+ ≅ 0

The exposed conductive part of load 1 is connected to phase 1 by the fault (zero impedance).

The voltage of one healthy phase of this load in relation to the frame is V V2 1− or V V3 1−(since ( )Z Z IT C f+ ≅ 0 ) , whence kM = =3 1 73. .

402




The exposed conductive part of load 2 is at the same potential as the remote earth.

The voltage of one healthy phase of this load in relation to the exposed conductive part is

therefore V VNeutral2 − or V VNeutral3 − :

V V V R I VR V

R RV a V

R

R RV

a R

R RNeutral e f

e

e M

e

e M

e

e M2 2 2

12 2 2 1− = − = −

+= −

+= −

+

whence ka R

R RM

e

e M

= −+

1

for R R kM e M= =, 132.

for R R kM e M> <, 132.

The earth electrode resistance of a group of loads is in general higher than the substation

earth electrode resistance. The overvoltage coefficient will thus be lower than 1.32 on load 2.

The overvoltage factor is maximum in the TT earthing system for a load having an exposed

conductive part connected to the same earth electrode as the fault load, we thus have

kM = 3

n recapitulative table of maximum earth fault overvoltages in relation to the neutral

earthing system

Medium and high voltage (1) Low voltage (2)

solidly earthed neutral

(HV or MV)

unearthed or

impedance

earthed neutral

(MV)

TN system TT system IT system

< 1.73 *

(generally 1.2 to 1.4)

1.73 1.45 1.73 1.73

(1) : phase-earth overvoltage

(2) : phase-exposed-conductive-part overvoltage

(*) : a network with a solidly earthed neutral is generally made up so as to limit overvoltages to values close to 1.2

to 1.4.

Table 5-1: maximum overvoltage factor in relation to neutral earthing system

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n consequences on equipment selection

The overvoltage factor and fault duration influence the choice of equipment insulation voltage

level.

o solidly or limiting impedance earthed neutral in MV, or TT and TN earthing system

in LV

Rapid clearance of the fault, and thus a short overvoltage time, means that the switchgear

phase-earth insulation level does not have to be higher than the nominal single-phase voltage.

o unearthed neutral in MV or IT earthing system in LV

Since the power supply does not have to be interrupted on occurrence of a first fault, the

overvoltage is likely to occur for a long period of time (several hours). It is therefore advisable

to choose switchgear with a phase-earth insulation level that is suitable for the nominal phase-

to-phase voltage.

Note: some manufacturers give a phase-earth insulation withstand equal to the single-phase voltage,but stipulate that their switchgear can be implemented in an unearthed neutral network. Thereare also switchgear standards that specify an insulation level compatible with use in anunearthed neutral network.

404




5.1.2.2. Resonance and ferro-resonance

resonance

The presence of inductive L , capacitive C and resistive R elements, connected, either in

series or in parallel, causes spreading of current and voltage having values which may be

dangerous for equipment.

series resonance

Figure 5-7 shows a series R L C, , circuit at the terminals of which a voltage U is applied.

L CRI

U

Figure 5-7: series R L C, , circuit fed by a voltage U

The voltage U is the vectorial sum of the voltages at the terminals of each element:

U U U U

R I j L IjC

R L C= + +

= + +ωω1

The vectorial diagram in figure 5-8 shows that for certain values of L and C , the voltages at

the terminals of the inductance and capacitance may be higher than the network voltage U :

U

1

jCI

j L I

R I

Figure 5-8: vectorial diagram of a series R L C, , circuit fed by a voltage U

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The resonance phenomenon occurs when U UL C= − :

j L IjC

ωω

= −1

LCω2 1=

We thus have U R I= ; the series inductance and capacitance behave like a short circuit.

For given values of L and C , the angular frequency ω r such that LC rω2 1= is said to be

a resonant angular frequency.

An overvoltage factor f is thus defined which is the ratio of the voltage UL (or UC ) to the

supply voltage U :

fU

U

L I

R I

L r= = ω

fL

R RC

r

r

= =ωω1

parallel resonance

Figure 5-9 shows a parallel R L C, , circuit at the terminals of which a current source J is

applied.

RUL

CJ

IR IL IC

Figure 5-9: parallel R L C, , circuit fed by a current source J

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The voltage U is common to the three elements.

We have the following relation:

JR j L

jC U= + +

1 1

ωω

The resonance phenomenon occurs when I IL C= − :

U

j LjC U

ωω= −

LCω 2 1=

We thus have U R J= ; the inductance and capacitance behave like an open circuit.

For given values of L and C , the angular frequency ω r such that LC rω2 1= is said to be

a resonant angular frequency.

An overvoltage factor is thus defined which is the ratio:

- between the voltage that is produced at the terminals of the parallel R L C, , circuit when

the resonance occurs

- and the voltage that would be produced on occurrence of the resonance if the inductance

(or capacitance) were the only circuit element

fR J

L Jr=

ω

fR

LR C

rr= =

ωω

The most current example of parallel resonance is the case of a network having harmonic

currents (patterned by current sources) and reactive energy compensation capacitors

(see § 8.1.5).

407




example: resonance in a Petersen coil earthed HV/MV substation

Figure 5-10 shows the diagram of a Petersen coil earthed HV/MV substation when an HV

earth fault flows through the common earth electrode.

CI f

Ve

HV / MV transformerHV

Re1

Re

L Rc c,

ZMV

I f : HV earth fault current

L Rc c, : Petersen coil inductance and resistance

R Re e, 1: earth electrode resistances

C : MV cable phase-earth capacitance

Ve : rise in substation earth potential

ZMV : sum of MV cable and transformer impedances

Figure 5-10: HV earth fault in an HV/MV substation with a Petersen coil earthed neutral

The symmetrical component method gives us the fault current value as (see § 4.2.2 of the

Network protection guide):

IV

Z Z Zf

n=+ +3

1 2 0( ) ( ) ( )

where Z Z ZT( )1 = + l

Z Z ZT( )2 = + l

Z Z RT e e( ( ) )0 0 1+ +l

Z ZT T, ( )0 :HV transformer positive-sequence (or negative-sequence) and zero-sequence impedances

Z Zl l, ( )0 : HV line positive-sequence (or negative-sequence) and zero-sequence impedances

408




In high voltage, the substation earth electrode value ( Re ) is very low compared with the

transformer and line impedances. The fault current is thus independent of Re ; it is thus

considered to be a source of current with a value of I f .

The equivalent Thevenin’s diagram of the current source I f with an internal impedance of

Re is shown in figure 5-11.

I f

equivalent

V R Ie feRe

Re

Figure 5-11: equivalent Thevenin’s diagram of the current source I f with an internal impedance of Re

The equivalent MV network diagram is thus that shown in figure 5-12.

V R Ie fe

C C C

Re Rc Lc

ZMV ZMV ZMV

Figure 5-12: equivalent MV network diagram on occurrence of an earth fault on the substation HV side

The transformer and cable impedances are negligible compared with the cable phase-earth

capacitance: ZC

MV <<1

ω .

409




The simplified MV network diagram is thus that shown in figure 5-13.

V R Ie fe 3C VC

VL

Re Rc Lc

Figure 5-13: simplified diagram

Let VL be the voltage at the inductance terminals.

We have

( )VL

R R j LVL

c

e c c C

e=+ + −

ω

ω ω1

3

In the case of a Petersen coil earthed neutral, (resonance) tuning between the inductance and

the MV cable capacitance is aimed at as far as possible. We thus have : LC

cωω

≈1

3 and

V VC L≈ whence VL

R RVL

c

e ce=

+ω .

To minimise the rise in substation earth potential (Ve ), the resistance earth electrode must be

as weak as possible (of the order of 0.5 Ω).

We can thus neglect Re compared with Rc , which thus gives us:

V VL

RV QVL C

c

ce e= = =

ω

V Q R IC e f=

Q : coil quality factor

VC : is equal to the MV cable phase-earth overvoltage in this case

The coil quality factor must not therefore be too high in order to avoid the risk of a very high

overvoltage.

This is why, in some cases, a resistor must be connected in parallel with the coil, in order to

reduce the quality factor.

410




Numerical application:

Let us take a 63/5.5 kV substation where:

V kVn = =5 5

33175

..

( )I kV kAf 63 3=

Re = 0 5. Ω

QL

R

c

c

= =ω

4

The rise in potential is: V R I Ve e f= × =1500 .

The phase-earth voltage in the cables is: V Q VC e= × .

V VC = × =1500 4 6 000

V VC n=189.

The overvoltage in the cables is roughly twice the nominal phase-earth voltage.

It can be dangerous if the substation earth electrode is of poor quality. Indeed, for Re = 3Ωwe will have V VC n=113. .

It is thus essential to limit the value of Re .

411




n ferro-resonance

o parallel ferro-resonance (see fig. 5-14)

Let us take a circuit made up of a parallel-connected capacitance, a coil with a saturable iron

core and a resistance. Let R be the resistance, C the capacitance and L the inductance

which varies with the current flowing through the coil and the voltage at the circuit terminals.

C V

IR

R L

ICIL

IT

Figure 5-14: parallel ferro-resonance

The total current IT flowing through the circuit is then given by the relation (1):

( )IV

Rj C V IT L= + −ω (1)

We cannot express IL as a function of V , owing to the saturation.

The rms values are given by the relation (2):

( )IV

RC V IT L

22

2

2= + −ω (2)

We can thus write relation (3) as follows:

IV

RC V IT L

22

2− = −ω

(3)

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This equation can be graphically resolved by plotting, as a function of V, the curves

representing functions (see fig. 5-15):

I IV

RT= −2

2

2(a)

I C V IL= −ω (b)

For any value of IT , the intersection of curves (a) and (b) gives the V solutions of equation

(3); figure 5-15 shows the graphic resolution of this equation.

Curve (a) is an ellipse having the equation:

V

RI IT

2

22 2+ =

and having one half axis which is equal to IT and the other to R IT . An ellipse corresponds

to each total current value IT .

Curve ( )I VL presents a very steep slope when V increases owing to the saturation of the

coil's iron core: ( ) ( )I VV

L VL =

ω .

On saturation, ( )L V becomes very weak and the current then highly increases (see fig. 5-15).

Curve I C VC = ω is a linear function of V (see fig. 5-15).

Curve (b) shows the development of ( )I I C V IC L L− = −ω as a function of the voltage.

The OSA portion of curve (b) corresponds to a lead current in relation to the voltage owing to

the preponderance of the capacitive current. On the other hand, the AB part corresponds to a

lag current, since the inductive current is preponderant. The intersection of ellipse (a) and

curve (b) can give:

- an operating point Q if ellipse (a) is inside ellipse (a") passing through point A

- three operating points M N P, , if ellipse (a) is between ellipses (a') and (a")

- two points S T, if ellipse (a) is equal to ellipse (a')

- a single point X if ellipse (a) is outside ellipse (a').

The ferro-resonant mechanism is described below.

413




With the circuit being initially unused, the total current IT is zero, as well as the voltage V , and

ellipse (a) is reduced to point O . If the current increases, the length of the axes of ellipse (a)

increases and the voltage rises, the operating point M moves along branch OS of curve (b).

When the total current exceeds the value IT' for which ellipse (a') cuts curve (b) at S , the

operating point suddenly jumps from point M to point T located on branch AB of curve (b), it

then moves along this branch. The voltage thus suddenly increases, going from VS to VT , and

then it continues to increase if the current IT increases.

If the total current now decreases, the operating point moves along branch AB and stays

there, even if the current drops below the value IT' corresponding to ellipse (a'). When the

current reaches the value IT , the operating point is P instead of M . It only returns to

branch OS if the current drops below the value IT'' corresponding to ellipse (a") passing

through point A . When this occurs, the operating point suddenly jumps from A to Q , and

the voltage from VA to VQ .

We can thus see that two stable operating conditions, for which the voltage at the circuit

terminals takes very different values, for example VM and VP , can correspond to the same

rms current value IT .

Finally, if the initial operating conditions correspond to a weak voltage (branch OS ), with a

resulting capacitive current, it is possible that, following a sudden change in operating

conditions leading to a transient phenomenon (overcurrent or overvoltage), the resulting

current becomes inductive and the voltage maintains a high value, even once the disturbance

has disappeared.

Ferro-resonance can be avoided if the resistance R is sufficiently weak for ellipse (a) to

remain within zone OSA , even when there is a high overcurrent.

resonance

capacitive operating

conditions

inductive operating conditions

IT'''

I

IT'

IT''

IT

O

IL

I C VC

(b)

B

X (a''')

(a')

(a)

T

PNM

Q

S

(a'')

VVQ VM VS VN VP VTVA

A

C V IL

Figure 5-15: parallel ferro-resonance - graphic resolution

414




o series ferro-resonance (see fig. 5-16)

Let us take a series circuit made up of a resistance, a coil with a saturable iron core and a

capacitance. We have:

V R I j VI

CL= + −

ω

(1)

We cannot express VL as a function of I , owing to the saturation.

If we move to rms values, we can write:

V R I VI

CL

2 2 22

= + −

ω

(2)

or: V R I VI

CL

2 2 22

− = −

ω

(3)

V R I L II

C

2 2 2− = −ωω

(4)

CLR

V

VR VL VC

I

Figure 5-16: series ferro-resonance

As for the parallel circuit, this equation can be graphically resolved as a function of I ,

by plotting curves (see fig. 5-17):

v V R I= −2 2 2

and v VI

CL= −

ω

Curve ( )V IL presents a very small slope when I increases owing to the saturation of the

coil's iron core ( ) ( )V I L I VL = ω .

415




On saturation, ( )L I becomes very weak and the voltage almost stops increasing when I

rises.

The network operating point is located at the intersection of curve (b) having the equation:

v VI

CL= −

ω

and ellipse (a) having the equation:

v V R I= −2 2 2

There are three possible operating points: M N P, , . M and P are stable, N is unstable.

A voltage disturbance can make the circuit move from point M to point P . This results in a

high current and high overvoltages at the inductance and capacitance terminals. Ferro-

resonance can be avoided if the resistance R is sufficiently high for ellipse (a) to stay within

zone OSA , even when there is a high overvoltage.

resonance

O

A

X(a''')

(a')

(a)

T

PN

M

Q

S

(a'')

V

I

V'''

V'

V''

V

(b)

VC

VL

IQ IM IS IN IA IP IT IX

Figure 5-17: series ferro-resonance - graphic resolution

416




o example of parallel ferro-resonance - unearthed neutral three-phase network

(see fig. 5-18)

Let us consider a three-phase network with unearthed neutral having a capacitance C

between each phase and the earth. Furthermore, a voltage transformer, with a similar

magnetizing inductance to a saturable core reactor, is connected between each phase and

earth. A parallel inductance-capacitance circuit thus appears between each phase and earth.

Parallel ferro-resonance can then be sparked between the capacitance and voltage

transformer of the same phase.

This ferro-resonance may occur following a transient overcurrent or overvoltage caused by a

switching operation and notably when the network is energized. Owing to the existing phase

displacements between the voltages of the three network conductors, the overcurrents and

switching overvoltages do not have the same magnitude in the three phases. Ferro-resonance

can thus very easily occur on only two phases, phases 2 and 3 for example. The voltages of

these two phases in relation to earth correspond to points located on portion AB of curve (b)

(see fig. 5-15). The voltage of phase 1 corresponds to a point located on the OS part of this

curve.

For phases 2 and 3, the capacitance-inductance assembly behaves like an inductance, and for

phase 1, like a capacitance. If we plot the voltage vector diagram, we can see:

- that the phase 1 voltage in relation to earth is weak

- that the voltages in relation to earth of the other two phases are very high

- that there is a very high potential difference between the neutral point and earth

(see fig. 5-18).

417




These overvoltages will cause a breakdown in equipment insulation if provisions to limit them

are not taken.

V3

V2

V1

CL

CL

Ph 3

Ph 2

Ph 1

v T3v T2

v T1

vN

N

V1

V2V3

Ferro-resonance occuring between two phases

N

CL

Figure 5-18: parallel ferro-resonance in an unearthed neutral network

protection against the risks of parallel ferro-resonance

A voltage transformer ( )VT charged by a resistor r behaves like a saturable (magnetizing)

inductor in parallel with this resistor.

Thus, in an unearthed network, if a charging resistor is connected to the secondary of the

voltage transformers, the L-C parallel circuits, made up of these transformers and network

cable capacitances, are transformed into R-L-C parallel circuits, such that if the resistors are

correctly sized, the risk of ferro-resonance outlined previously can be avoided (ellipse (a)

remains inside the zone 0SA - see fig. 5-15):

- the resistors must be sufficiently weak to be efficient

- they must not be too weak, so that the VT are not overcharged and their accuracy is

maintained.

418




In the case of VT with a single secondary, a charging resistor is installed on each phase

(see fig. 5-19).

A resistance value equal to 68 Ω is recommended for a secondary voltage of100

3V .

VT VT VT

r

r

r

measurements

Figure 5-19: protection against risks of ferro-resonance using resistors with single secondary VT

419




In the case of VT with two secondaries, a resistor is installed in the open delta of one of the

two (see fig. 5-20).

It is recommended that a power above 50 W be dissipated in the resistor on occurrence of a

phase-earth fault.

For a secondary voltage of100

3V , on occurrence of a solid earth fault, the voltage at the

resistor terminals is equal to 100 V; the resistance value is then determined:

( )R ≤

100

50

2

R ≤ 200Ω

VT VT VT

measurements r

Figure 5-20: protection against the risks of ferro-resonance via a resistor with two-secondary VT

420




o example of series ferro-resonance (see fig. 5-21)

Figure 5-21 shows a solidly earthed network feeding a three-phase transformer having a delta-

connected primary. This can also apply to a star-connected transformer with an unearthed

neutral. If, when the switch is closed, one of the poles remains accidentally open or closes

late, for example the pole of phase 1, series ferro-resonance may occur in the circuit including:

- the magnetizing inductance of transformer windings AC or BC

- the capacitance of phase 1 in relation to earth.

Very high overvoltages can occur at the transformer terminals and between phase 1 and the

earth.

This type of ferro-resonance has frequently been encountered on HV networks with solidly

earthed neutral. It may also occur when a switch is opened. The means of protecting against

this type of ferro-resonance consists in inserting a resistor in the supply transformer neutral

point earthing. This solution does not however provide total protection since ferro-resonance

can, for example, occur in the circuit including the transformer AC winding and the

capacitances of phases 1 and 3 in relation to earth.

V3

V2

V1

CCC

Ph 3

Ph 2

Ph 1

switchA

LB

C

L

L

I f

Figure 5-21: series ferro-resonance

421




5.1.2.3. Neutral conductor breakdown

Let us consider the diagram in figure 5-22 , where Z1 , Z2 and Z3 represent the equivalent

impedances per phase of all the loads downstream of the neutral breakdown point.

If the phases are perfectly balanced, the voltage system is not disturbed.

In the event of load unbalance, the neutral point is displaced and the phase-neutral voltages

move close to the phase-to-phase voltage for the least loaded phases, while for the loaded

phases (weak impedance), they drop below the single-phase voltage.

V3

V2

V1

Z3

Z2

Z1N

neutral breakdown

Figure 5-22: equivalent diagram of an LV network during neutral breakdown

Using the superposition theorem, we can show that:

VZ Z

Z Z ZV

Z Z

Z Z ZV

Z Z

Z Z ZVN =

+

+

+

+

+

1 2

3 1 23

1 3

2 1 32

2 3

1 2 31

//

//

//

//

//

//(1)

The voltage applied to the terminals of a single-phase load on phase 3, for example, will be:

V V VN N3 3= −

If we know that V a V221= and V a V a j3 1

1

2

3

2= = − +

,

then we can calculate V N3 , for example, for the following impedances:

Z R1 =Z R2 2=Z R3 10=

(We have taken resistive loads to simplify the calculations.)

422




By applying formula (1), we find:

Va

VN = +

4 9

16

2

1

we then have V V V a V VN N N3 3 1= − = −

Vj

VN3 115 10 3

16= − +

whence V VN n3 1= .43

Similarly, we can determine: V VN n2 114= . and V VN n1 0 6= .


We can see that once the most sensitive single-phase loads have broken down, there are

successive breakdowns, following the development of the phenomenon which worsens the

unbalance ( Z3 increases after the breakdowns and consequently V N3 increases); this is an

avalanche phenomenon.

This risk thus underlines that it is preferable to well balance the loads on the three phases.

423




5.1.3. Switching overvoltages

When switching to energize or de-energize loads transient overvoltages occur on the network.

These overvoltages are all the more dangerous if the current interrupted is inductive or

capacitive. The magnitude, frequency and damping duration of these transient overvoltages

depend on the given network characteristics and the mechanical and dielectric characteristics

of the switching device.

5.1.3.1. Interrupting principle

Interrupting an electric current using an ideal device involves the resistance of the device going

from zero before interruption to an infinite value just after interruption. The interruption occurs

the instant the current crosses zero.

It is impossible to make such an ideal device, but with the interrupting techniques being based

on the behaviour of the electric arc in different dielectric media we can come close to it.

n circuit-breaker interruption

The instant the current is interrupted, an electric arc is created between the terminals of the

switching device. The conductive electric arc tends to be held by the ionizing phenomenon of

the dielectric caused by the energy dissipated.

Around current zero crossing, the dissipated energy decreases dropping below the thermal

energy supplied to the medium, the arc cools down and its resistance increases.

When the current crosses zero, the arc resistance becomes infinite and the arc is interrupted.

Between the start and end of interruption, the voltage between the poles of the switching

device goes from zero to the network voltage. This change gives rise to a high frequency

transient phenomenon called the transient recovery voltage (see fig. 5-23).

424




I V

t

VA

L

C

R

I

VAV

L R, : inductance and resistance equivalent to the network upstream of the circuit-breaker

C : upstream network capacitance

Figure 5-23: transient recovery voltage during circuit-breaker interruption

n fuse interruption

On occurrence of a short circuit, the value of the current flowing through the fuse is higher than

its nominal fusing value.

Interruption can thus occur at any instant and not necessarily the moment the current crosses

zero.

Figure 5-24 gives an example of a transient overvoltage which occurs on the network after a

wire fuse has fused.

t

Volts

225

1000

~ 1 ms

Figure 5-24: transient overvoltage on fusion of a wire fuse

425




5.1.3.2. Load switching

n de-energizing loads

o inductive load

single-phase circuit

Let us consider the equivalent single-phase diagram in figure 5-25 with an ideal circuit-breaker

CB which has a zero arc resistance the instant the contacts separate and which carries out

interruption when the current crosses zero. Before operation of the circuit-breaker, between

points A and B, there is a voltage drop due to the load current flowing through Ls .

At the instant of interruption, the voltage at B suddenly reaches the voltage at A and the

capacitance Cs is charged through Ls . The energy exchanges between Cs and Ls make

voltage oscillations at frequencies of 5 to 10 kHz occur.

The voltage at C suddenly decreases to zero and the capacitance Cp is then discharged

through L . The energy exchanges between Cp and L create voltage oscillations at

frequencies going from 1 to 100 KHz.

BA C

L

Ls

Is IL

CpCs

ID

I0

CB

Lp

VA

Ls : network inductance upstream of the circuit-breaker

Cs : network capacitance upstream of the circuit-breaker

L : load inductance

Lp : stray inductance

Cp : network capacitance downstream of the circuit-breaker

CB : circuit-breaker

Figure 5-25: interruption in an inductive load network

426




The phenomena observed are illustrated by curves in figure 5-26.

t

VB

t

t

VC

I s

t

ID

t

IL

t

t0 t1

t

VA

V V VD B C= −

t0 : separation of contacts

t1 : zero current

Figure 5-26: interruption cycle of an ideal device

427




three-phase circuit

When the three-phase circuit in figure 5-27 is interrupted, the first phase which sees the

current crossing zero interrupts this current. There follows a transient current circulating in the

two uninterrupted phases. Thus, if phase 1 interrupts the current first a transient voltage is

obtained between points C1 , C2 and earth which is capable of reaching a value of 2 $Vn for

an ideal circuit-breaker. For an actual circuit-breaker, the overvoltage coefficient is higher than

or equal to 2.

$Vn : peak value of the phase-neutral nominal voltage

Note: the current crosses zero on the following phase after 1/3 of a period (7 ms at 50 Hz), while theperiod of oscillations is roughly 1 ms.

Ls

C pCs Lp

A2 V2B2 C2 L2

Ls

C pCs Lp

A3 V3B3 C3 L3

Ls

C pCs

V1A1B1 C1 L1

Lp

N

Figure 5-27: equivalent diagram of a three-phase circuit during interruption

restrike phenomenon

The instant a circuit is interrupted, the voltage at the terminals of the circuit-breaker quickly

increases (roughly from 0.1 to 0.5 kV/µs). If the circuit-breaker poles separate shortly before

the current reaches zero (for an inductive circuit, this corresponds to the maximum voltage),

regeneration of the dielectric medium may not be sufficient to withstand the stress-voltage.

Indeed, in this case, the voltage is maximum and the poles are closer together.

Renewed breakdown then occurs accompanied by overvoltages with a peak to peak

magnitude of 2 $Vn . This phenomenon is called restrike.

428




multiple restrike

If we consider the single-phase diagram in figure 5-25, we can see that in the case of restrike,

the voltage at point C almost instantaneously reaches the voltage at point B .

The capacitance Cp is charged by a high frequency current (roughly 1 MHz) circulating in the

L C CBp s, , and Cp circuit.

This high frequency current very quickly crosses zero (1 µs).

If the circuit-breaker manages to interrupt the current at that moment, the restrike phenomenon

is repeated as the distance between the circuit-breaker contacts is still very small.

Furthermore, the peak-to-peak magnitude of the oscillation is then equal to 4 $Vn .

The overvoltage increase makes the occurrence of a second breakdown highly probable.

Indeed, the increase in dielectric withstand through the increase in the distance between the

circuit-breaker contacts may be lower than the increase in overvoltage.

This is why a multiple restrike phenomenon occurs with overvoltages of increasing magnitude

(see fig. 5-28).

In theory, such a phenomenon may generate overvoltages having a peak value equal to the

dielectric withstand limit of the open device, without a definite interruption of the current being

obtained. In practice, this case remains exceptional as it is enough for one of the restrikes to

allow the power frequency current to be restored; a new current half wave then flows through

the circuit-breaker. The circuit-breaker interrupts this half-wave the moment it crosses zero

when the distance between the contacts is sufficient. Thus the types of circuit-breakers

undergoing multiple restrike usually manage to interrupt the current without causing

overvoltages of very high magnitude.

t

VC

Figure 5-28: voltage VC in case of interruption with multiple restrike

429




chopping current (weak inductive currents)

When weak currents, notably lower than the nominal current of the circuit-breaker, are

interrupted, the arc which occurs occupies a small volume. It consequently undergoes

considerable cooling linked to the circuit-breaker's capacity to interrupt much higher currents.

Owing to this fact, the arc becomes unstable and its voltage may present relatively large

variations, while its absolute value remains lower than the network voltage (case of SF6 or

vacuum). These voltage variations may generate high frequency oscillating currents, with a

magnitude that may reach 10% of the current at 50 Hz, in the nearby capacitances (C L Cs p p, ,

circuit in figure 5-25). Superposing these high frequency currents on the current at 50 Hz

results in multiple crossings of the current through zero around zero of the fundamental wave

(see fig. 5-29).

The circuit-breaker interrupts the current the first time it crosses zero while the load current

(only the current at 50Hz) is not zero. The value of this current represents what we call the

chopping current ( )Ichop .

extinctionpossible

"chopping"current

current in thecircuit-breaker

50 Hz wave

Ichop

Figure 5-29: superposition of a high frequency oscillating current

on a power frequency current

430




The current is then interrupted as in the case in figure 5-25 except for the peak-to-peak

magnitude of the oscillations, due to the presence of energy stored in L1

2

2L Ia

which is

added to that in the capacitance Cp1

2

2C Vp n$

.

If $maxVc is half the peak-to-peak maximum value of the oscillation at point C, we can write:

1

2

1

2

1

2

2 2 2C V C V L Ip c p n a$ $max = +

$ $maxV V

L

CIc n

pa= +2 2 in single-phase.

$Vn : phase-neutral nominal voltage peak value

For a three-phase circuit $Vn must be added in order to take into account the transient

operating conditions linked to the non-simultaneous interruption of the phases, whence:

$ $ $maxV V V

L

CIc n n

pa= + +2 2

This phenomenon is notably problematic in the case of an arc furnace transformer power

supply.

Indeed, the transformer is generally connected not very far away from the busbar. Thus the

value of Cp is very weak and therefore the value of $maxVc high.

We can determine $maxVc by taking:

L : transformer leakage inductance

Cp : capacitance of the cable linking the circuit-breaker to the transformer

Ia : transformer magnetizing current

Schneider carried out an analysis for a single-phase arc furnace transformer where:

VV

n = 150003

; L H= 8 26. ; C nFp =14 75. ; I Aa = 4 36.

we find $ . $maxV Vc n= 8 5

Installing an R C, circuit in parallel with the circuit-breaker allowed the overvoltage to be

reduced to 2 $Vn .

431




virtual chopping current - simultaneous interruption of the three phases

The transients generated by the first phase that creates overvoltages may cause, owing to the

capacitive coupling between the phases, oscillating currents inside circuits L C Cp p s, , of the

other phases.

It is thus possible to obtain zero current in these phases, immediately (several hundreds of a

microsecond) after interruption of the first phase.

If the circuit-breaker interrupts such currents, a chopping current phenomenon is then created

with very high chopping current and overvoltage values.

chopping current and multiple restrike

Current chopping and multiple restrike are frequently linked.

Overvoltages caused by current chopping can themselves lead to restrike. They are almost

systematic in the case of the virtual chopping current.

o capacitive loads (see fig. 5-30)

Interruption of capacitive circuits, such as a capacitor bank or off-load cable, raises less

difficulties than the interruption of inductive circuits.

Indeed, the capacitances remain charged at the peak value of the 50 Hz wave after extinction

of the arc when the current reaches zero and the recurrence of voltage at the switchgear

terminals is accompanied by a 50 Hz wave.

Nevertheless, one half period after interruption, the device is subjected to a voltage equal to

twice the 50 Hz peak voltage ( )2 $Vn .

If the speed and dielectric withstand of the device are not sufficient to withstand this stress,

restrike may occur. It is followed by a voltage reversal at the terminals of the capacitances,

raising them to a phase-neutral voltage equal to 3 $Vn maximum (if damping is neglected).

432




When the supply voltage reverses back, a half period later, the potential difference at the

device terminals then reaches 4 $Vn . Such an overvoltage can obviously cause renewed

restrike between the device contacts, and the previously described oscillation mechanism is

renewed with increased magnitude, leading to a new rise in the phase-neutral voltage of the

capacitances ( )5 $Vn .

The cumulative effect of multiple restrike is obviously highly dangerous for the network

components as for the device itself.

This rise in overvoltages can be avoided by choosing the appropriate equipment, i.e. which

does not allow restrike.

20 msV

t

interruption

$Vn 2 $Vn

4 $Vn

V VC n3$

V VC n5$

V VC n$

Figure 5-30: voltage rise on separation of a capacitor bank from

the network by a slow operating device

433




n energizing a load

o inductive circuit

When a device closes, on an inductive circuit (no-load transformer, motor on starting), there is

a moment when the dielectric withstand between its contacts drops below the applied voltage.

A breakdown occurs causing sudden zero voltage at the device terminals.

This is accompanied by oscillations with stray capacitances which cause high frequency

currents to circulate in the circuit-breaker.

Depending on the speed of the device, prestrikes may or may not occur up to complete closing

of the poles.

Multiple prestrike is accompanied by successive overvoltages which decrease until the device

is completely closed.

The phenomenon is highly complex and involves several parameters:

- the characteristics of the switching device

- the characteristic impedance of the connections

- the natural frequencies of the load circuit

which means that a mathematical simulation model is required to pre-determine the

overvoltage values.

o capacitive circuit (capacitor bank)

When a capacitor bank is energized via a slow operating device, prestrike occurs between the

contacts close to the wave peak of 50 Hz.

A damped oscillation in the LC system in figure 5-31 then occurs at a frequency above

50 Hz concentrated around the peak. In this case the maximum overvoltage is 2 $Vn . It

corresponds to the maximum overvoltage admissible by the capacitors (see IEC 831-1 for LV

and 871-1 for MV or HV).

With a faster device, prestrike does not necessarily occur around the 50 Hz peak and

consequently the overvoltage is smaller.

When put out of service, the bank remains charged at a voltage going from 0 to the peak

voltage of the network.

434




If the bank is energized shortly afterwards, a breakdown due to the application of a voltage of

opposite polarity may give rise to an overvoltage of 3 $Vn .

C

CB

U

L

Figure 5-31: closing operation of a capacitive circuit

To ensure the safety of persons, the capacitor banks are fitted with a discharging resistor

having a time constant allowing 75 V to be reached after 3 minutes in LV and 10 minutes in

HV.

n Means of protecting loads

The phenomena created by de-energizing (or energizing) loads, which we have studied, lead

to transient overvoltages which may be dangerous for both loads and other network elements.

Table 5-2 gives the level of overvoltages and their characteristics for each phenomenon

studied.

Occurrence of

phenomenon

Number of

overvoltage

peaks

Overvoltage

value

dU/dt order of

magnitude

Remark

Chopping

current

at every

interruption

1 2 to 4 $Vn 0.1 kV/µs favours restrike

Multiple

restrike

interruption with

separation

close to zero

current

0 to 20 2 to 7 $Vn 10 kV/µs

Prestrike at every closing 1 to 50 2.5 $Vn 10 kV/µs

$Vn : phase-neutral voltage peak value

Table 5-2: different types of overvoltage

435




The loads affected by these phenomena are off-load transformers, neutral point coils (neutral

reactance earthing) and motors during the starting period for inductive circuits as well as

capacitor banks for capacitive circuits.

Transformers undergo impulse wave dielectric tests; because of this, they are better built than

motors to be able to withstand the transients caused by restrike (see IEC 76-3).

The case of motors is different. At each start, they must withstand the transients caused by

prestrike. Moreover, even if interruption during the starting period does not occur very often, it

is nevertheless a possibility and they are then subjected to multiple restrike.

Motors are thus especially sensitive to multiple prestrike, because of its high rate of

occurrence, as well as to multiple restrike, due to the magnitude of the overvoltages produced.

These overvoltages cause deterioration of the insulation of the first turns.

In order to limit overvoltages, Zn0 type surge arresters can be connected in parallel with the

load.

But the best method consists in using switching devices suitable for the type of application.

Table 5-3 gives the behaviour of medium voltage switchgear with respect to the phenomena

relating to the switching overvoltages studied.

Switchgear Multiple

prestrike on

closing

Current

chopping

Multiple

restrike

Overall behaviour

Puffer-type SF6 circuit-

breaker

no weak no No problem. Below 300

kW, use a rotating arc

SF6 circuit-breaker.

Rotating arc SF6 circuit-

breaker and contactor

no no no No problem.

Vacuum circuit-breaker yes yes yes Use surge arresters

Vacuum contactor yes weak yes Use surge arresters

Magnetic blast circuit-

breaker and contactor

no no no No problem.

Minimum oil circuit-breaker no yes yes Use surge arresters

Table 5-3: behaviour of medium voltage switchgear

436




5.1.3.3. Circuit-breaker clearance of phase-earth faults

Let us consider the three-phase network shown in figure 5-32 in which phase 1 is affected by

an earth fault.

In this case, the network is equivalent to the diagram in figure 5-33 which corresponds to the

case examined in paragraph 5.1.2.1.

At the start of contact separation, the arc voltage is weak and remains constant.

On the other hand, just before interruption, this voltage, called the extinction voltage, increases

to a more or less high value which may exceed $Vn . This voltage depends on the type of

circuit-breaker (air, oil, SF6 , vacuum) as well as the arc extinction technique (cooling,

lengthening, rotating arc).

When the current crosses zero, the arc is extinguished and the recovery voltage magnitude

will depend on the extinction voltage as follows:

- for the case of neutral earthing via resistance (the fault current is in phase in relation to the

voltage), the extinction voltage limits the magnitude of recovery voltage oscillations

- for the case of neutral earthing via reactance (the fault current is phase shifted byπ2

in

relation to the voltage), the extinction voltage increases the magnitude of oscillations.

After interruption, restrike may take place if re-generation of the dielectric medium is not fast

enough in relation to the rise in recovery voltage. In this case, the magnitude of oscillations

may reach double the size of the first recovery voltage.

If we neglect the transformer and line impedances, the voltage at the terminals of the neutral

earthing impedance ( )VN is equal to the difference between the supply voltage and the

voltage at the circuit-breaker terminals. The voltage VN is vectorially added to the voltage of

the healthy phases and may lead to the latter reaching higher overvoltages than the

overvoltages observed on the fault phase.

The curves in figure 5-34 give the overvoltage levels recorded on occurrence of an earth fault

in relation to the network characteristics and the earthing impedance.

437




We can see that reactance earthing of the neutral (case with restrike) clearly increases the

magnitude of the overvoltages. Resistance earthing is thus preferable. In the latter case, we

see that the overvoltages do not exceed 240 % when the ratio of the current in the earthing

resistor to the network capacitive current is equal to 2 (see fig. 5-34). In networks with

resistance earthing, the following relation should therefore always be respected if possible:

I IrN C> 2

IrN : current in the neutral earthing resistor during the fault

IC : currents in the network phase-earth capacitances (see § 4.3 of Protection guide)

V3

V2

V1

C C C

Ph 3

Ph 2

Ph 1

I f

CB

ZNrN

or

ZN : neutral earthing impedance (or rN )

C : phase-earth capacitance

I f : fault current

CB : circuit-breaker

V V V1 2 3, , : single-phase voltages

Figure 5-32: phase-earth fault clearance

C

I f

CB

ZNrN

or

Xnet

~

IrN

IC

Xnet : network reactance

C : fault phase earth capacitance

ZN or rN : neutral earthing impedance (or resistance rN )

I f : fault current

Figure 5-33: fault circuit on occurrence of a phase-earth fault

438




High resistance earthing with restrike in the

fault or circuit-breaker, case of industrial

networks for which I to ArN < 20 30

(see Protection guide - § 10.1.1.).

The overvoltage depends on the ratioI

I

rn

C

Limiting resistor earthing with restrike in the

fault or circuit-breaker, case of public

distribution networks for which IrN is equal

to several hundred to 1 000 A. The overvoltage

depends on the ratior

X

N

d

0.5 1 1.5 2 2.5

460

300

260240

200

100

healthy

phases

neutral

fault

phaseI

I

rN

C

%

transient voltage as

% of the nominal

single-phase voltage

peak value

3 6

250

150

100

50

healthy

phases

neutral

faultphase

200

4 5 8 10 20 30 50 70 90

%


% of the nominal


peak value

r

X

N

(1)

I V rrN n N= / : current in the neutral resistor during

the fault

I C VC n= 3 ω : vectorial sum of current in the phase-

earth capacitances

If I IrN C≥ 2 , the overvoltage does not exceed 240 %

rN : neutral point earthing resistance

X( )1 : network positive-sequence reactance

Reactance earthing, case of public distribution networks for which IXN is equal to 1 000 to

several thousand amps

Case without restrike in the circuit-breaker Case with restrike in the circuit-breaker

0

100

theoretical limits

without

damping200

300

400

2 4 6 8 10 12 14 16

C

N

A

B

%


% of the nominal


peak value

X

X

N

(1) 0

100

theoretical limits

withoutdamping

200

300

400

2 4 6 8 10 12 14 16

C

N

A

B

500

%


% of the nominal


peak value

X

X

N

(1)

X( )1 : network positive-sequence reactance

X N : neutral point earthing reactance

A : earth fault phase

B C, : healthy phases

N : voltage at reactance terminals

Figure 5-34: transient overvoltages depending on the type of neutral earthing

during a phase-earth fault

439


The publication, translation and reproduction, either wholly orpartly, of this document are not allowed without our written consent.


Reactance earthing: (network reactance)

Voltage at circuit-breaker terminals Voltage at the terminals of the reactance

Resistance earthing: (network reactance)

Voltage at circuit-breaker terminals Voltage at the terminals of the resistor

: arc extinction voltage

Figure 5-35: transient voltage on circuit-breaker opening during a permanent phase-earth fault

440




5.1.4. Atmospheric overvoltages

n general

The earth and the electrosphere, the conductive area in the atmosphere (about 50 to 100 km

thick), constitute a natural spherical capacitor which is charged by ionization, thus producing

an electric field directed towards the ground of roughly several hundred volts/metre

Since air is not very conductive, there is thus an associated permanent conduction current, of

roughly 1 500 A for the entire earth's globe. Electrical balance is ensured during discharges by

rain and strokes of lightning.

The formation of storm clouds, masses of water in the form of aerosols, is accompanied by

charge separation electrostatic phenomena: the positively charged light particles are driven by

the rising air currents, and the negatively charged heavy particles fall because of their weight.

At the base of the cloud, there may also be islets of positive charges where heavy rains are

located.

On an overall macroscopic scale, a dipole is created.

When the breakdown withstand limit gradient is reached, a discharge is produced inside the

cloud or between clouds or between the cloud and the ground. In the latter case, it is referred

to as lightning.

The cloud-ground electric field can reach 15 to 20 kV/metre on flat ground. But the presence of

obstacles deforms and locally increases this field by a factor of 10 to 100 or even 1 000

depending on the form of the obstacles (also called the "peak effect"). The atmospheric air

ionizing threshold is thus reached, i.e. roughly 30 kV/cm, and corona effect discharges are

produced. When these discharges are located on fairly high objects (tower, chimney, pylon)

they may divert lightning to this objects.

o classification and characteristics of strokes of lightning

Strokes of lightning are classed according to the origin of the discharge (or leader) and their

polarity.

Depending on the leader origin, the stroke of lightning may be:

- either descending from the clouds to the ground in the case of fairly flat land

- or ascending from the ground to the clouds in the case of mountainous land.

Depending on the polarity the following distinctions between lightning strokes are made:

- negative when the negative part of the cloud is discharged, which represents 80 % of cases

in temperate countries

- positive when the positive part of the cloud is discharged.

441




form and magnitude of the lightning wave

The physical phenomenon of lightning corresponds to a source of impulse current the actual

form of which is highly variable: it consists of a front rising up to the maximum magnitude of

several miscroseconds to 20 µs followed by a decreasing tail of several tens of µs (see

figure 5-36).

Figure 5-36: oscillogram of a lightning current

The magnitude of strokes of lightning varies according to a log-normal distribution law. We can

thus determine the probability of a given magnitude being exceeded (see figure 5-37). We can

see, for example, that for the average curve (IEEE), the probability of exceeding a magnitude

of 100 kA is 5 %. This means that 95 % of lightning strokes have a magnitude less than

100 kA.

Figure 5-37: probability of exceeding positive and negative lightning stroke magnitudes,

according to IEEE (experimental statistic)

442




Similarly, the steepness of the wave front varies according to a log-normal distribution law. Let us

determine the probability of exceeding a given front steepness (see fig. 5-38). We can see that the

probability of exceeding a front steepness of 50 kA/µs of a negative stroke of lightning is 20 %.

Figure 5-38: probability of exceeding the front steepnesses of positive and negative

lightning currents according to IEEE (experimental statistic)

standard wave form

The lightning impulse wave form given by IEC 71-1 is a 1.2/50 µs wave (see fig. 5-39):

- rise time to the maximum value of 1.2 µs

- time to half-value of 50 µs.

Figure 5-39: standard lightning impulse voltage wave form (IEC 71-1)

443




lightning density

On a world-wide scale, 63 billion discharges are recorded on average each year which

corresponds to 100 discharges per second. In France, this figure varies from 1.5 to 2 million

lightning strokes per year.

We then define the lightning density as being the number of days per year on which

thunder has been heard in a place.

In France, the average value of is 20 with a variation range going from 10 in channel

coastal regions up to over 30 in mountainous regions.

The value of may be much higher and reach 180 in tropical Africa or Indonesia.

lightning strike density

The lightning strike density represents the number of lightning strikes per km2 per year,

whatever their current value levels.

In France, varies between 2 and 6 lightning strikes/km2/year depending on the region.

o lightning impact mechanism

The lightning impact mechanism begins with a leader from a cloud which approaches the

ground at a low speed. When the electric field is sufficient, sudden conduction is established

giving rise to the lightning discharge.

An experimental practical approach has enabled the relation linking the current of the

lightning strike to the distance between the starting point (leader position) and discharge point

(point of impact connected to the earth) to be found:

or according to the authors.

: striking distance, in m

: lightning current, in kA

444




By applying an electro-geometrical model to a vertical rod with a height (see fig. 5-40-a),

we can show that there are two distinct zones:

- zone 1 : this is located between the ground and the parabola which is the locus of the

equidistant points of and the ground; the instant the flash occurs, any leader

located in this zone will touch the ground since it is nearer to this than to

- zone 2 : this is located above the parabola; the instant the flash occurs, any leader located

in this zone will be picked up by point on the vertical rod as soon as the

distance between and the leader is less than the striking distance .

Figure 5-40-a: diagram of different protection zones offered by a vertical rod

For a lightning current with a value of , and thus a given striking distance, the distance x between

the point of impact on the ground and the point where the rod is fixed to the ground (called the rod

pick-up radius) is:

if

if

The rod pick-up radius is thus all the greater the more intensive the lightning stroke.

For very weak currents, the pick-up radius becomes less than the height of the rod which is then

able to pick up the current along its length. This has been experimentally proved.

445


The publication, translation and reproduction , either partly of this document are not allowed without our written consent.


application to equipment protection using a lightning conductor

The lightning conductor diverts lightning to itself in order to protect equipment. Its principle is

based on the striking distance; tapered rods are placed at the top of equipment to be

protected, they are connected to the earth by the most direct path (the lightning conductors

surrounding the structure to be protected and interconnected with the earthing network).

The electrogeometric model allows the zone to be protected to be determined using the fictive

sphere method.

The point of impact of the lightning is determined by the object on the ground the closest to the

leader starting distance d . Everything happens as if the leader was surrounded by a fictive

sphere with a radius d moving with it. For good protection, the fictive sphere rolling on the

ground reaches the lightning conductor without touching the objects to be protected

(see fig. 5-40-b).

Protection against direct lightning strikes is approximately good in a cone the top of which is

the top of the lightning conductor and the half-angle at the top is 45 °.

leader

d = critical striking distance

fictive

sphere

protected

zone

(cone)

lightning

conductor

45°

Figure 5-40-b: determining the zone protected by a lightning conductor

using the "fictive" sphere method

446




n direct lightning strike (on phase conductors)

When lightning strikes the phase conductor of a line, the current ( )i t is shared out in equal

quantities on either side of the point of impact and is spread along the conductors. These have

a wave impedance Z the value of which is between 300 and 500 Ω. This impedance is that

seen by the wave front, is independent of the length of the line and of a different type from the

impedance at 50 Hz.

This results in a voltage wave of:

( ) ( )U t Z

i t= .

2

which spreads along the line (see fig. 5-41).

U i

t

i

U Zi

=2

Figure 5-41: lightning strike on a phase conductor

447




Depending on the magnitude of the lightning current, two cases may occur:

o full impulse propagation

If the maximum voltage U ZI

maxmax=

2

is below the sparkover voltage Ua of the insulator

string, the entire (full) wave spreads along the line.

o chopped impulse propagation

In the case where U Uamax ≥ , as a first approximation, insulator sparkover occurs at the value

of Ua , and a phase-earth fault occurs at 50 Hz due to the arc being maintained. The lightning

that is propagated is thus broken at the maximum value corresponding to Ua .

The lightning current causing this flashover, for a given line, is called the critical current IC

such that:

IU

ZC

a= 2

For lines, the order of magnitude of IC is:

- 5.5 kA at 225 kV, which corresponds to a probability of exceeding the magnitude according

to the IEEE method of 95 % (see figure 5-37)

- 8.5 kA at 400 kV, which corresponds to a probability of exceeding the magnitude according

to the IEEE method of 92 % (see figure 5-37).

In medium voltage, flashover is systematic in the case of a stroke of lightning occurring due to

the small distances in the air of the insulator string. This flashover of the insulator gives rise to

a phase-earth fault current, called a follow current, which is held at the power frequency of 50

Hz until it is cleared by the protections.

448




n indirect lightning strikes (lightning protection rope or pylons)

When lightning strikes the line protection rope, part of the current flows through the pylon since

the protection rope is connected to it (see fig. 5-42).

This results in a potential rise at the top of the pylon the value of which depends on the self

inductance L of the pylon and the resistance R of the earth electrode:

( ) ( ) ( )U t k R i t L

di t

dt= +

k : ratio of the current shunted into the pylon by the incident current

i

Uk . i

L

R

k . i

lightning strike

protection rope

U k R I Ldi

dt= × +

Figure 5-42: lightning strike on a protection rope

The voltage U may reach the impulse sparkover voltage of the insulators and cause a

breakdown. This is "back-flashover". Part of the current is then propagated along the affected

phase(s) towards the users. This current is in general greater than that of a direct lightning

strike.

449




In extra high voltage (> 220 kV), back-flashover is unlikely (the flashover level of the insulators

is high), which is why it is useful to install protection ropes thus limiting the number of service

interruptions. But below 90 kV back-flashover occurs even if the value of the earth electrode

resistance is low (< 15 Ω); the usefulness of protection ropes is thus limited (more frequent

service interruptions).

o induced impulse

A stroke of lightning that falls anywhere on the ground behaves like an electromagnetic field

radiation source.

The steeper the rising front of the lightning current the greater the radiation.

For front steepnesses of 50 to 100 kA/µs, the effects of this field will be felt several hundreds

of metres, if not kilometres, away.

The magnetic field H at a point located at a distance of r from a circuit through which a

current I flows, is given in the relation:

HI

r=2π

This field creates induced voltages in the neighbouring circuits which are able to reach

dangerous values both for equipment and persons.

case of a loop

Let us consider the loop formed by the supply cable and the telecommunication link in figure

5-43, with a surface S and located 100 m from the lightning impact which has a current rising

front steepness of 80 kA/µs.

450




The induced voltage is given in the relation:

ed

dtSdB

dtSdH

dt= − = − = −φ µ0

µ = × −0

74 10π : magnetic permeability of the vacuum

nowdH

dt r

dI

dtA m s= =

×× × = ×−

1

2

1

2 10080

10

10127 10

3

66

π π/ /

whence e kV= × × × =−4 10 120 127 10 197 6π

A phase-earth overvoltage of 19 kV thus occurs on the loop. This has a very short duration

( ≈ µ1 s ) but can cause insulation breakdown.

To avoid this risk, the surfaces of the loops must be reduced.

telecommunication link computer

supply cable

printer

circuit

loop

surface = 120 m²

earth

lightning impulse

phase-earth insulation

subjected to 19 kV ( 1 µs)

100 m

front steepness = 80 kA/µs

magnetic field

Figure 5-43: circuit loop

451




n impulse wave transference in a transformer (see IEC 71-2 - appendix A)

In lightning impulse conditions, the transformer behaves like a capacitive divider with a ratio of

s ≤ 0.4 . It is equivalent to a capacitance Ct (see figure 5-44-a).

lightning wave

U1

U sU0 1

Ctequivalent

U sU0 1

U1 : impulse voltage on the high voltage terminal

U0 : no-load voltage transferred

Figure 5-44-a: impulse wave transference in a transformer

U0 represents the no-load overvoltage, i.e. when the secondary outgoing terminals are not

connected to any cables or lines. This overvoltage is generally not acceptable by the

transformer.

In reality, the transformer is connected to a network with a capacitiance Cs . This plays the

role of a voltage divider with the transformer capacitance Ct (see fig. 5-44-b).

U sU0 1=

Ct

Cs U2

U2 : voltage transferred to the secondary with a network

Figure 5-44-b: transformer with its equivalent network

The voltage transferred to the secondary is thus:

UC

C CsUt

t s2 1=

+

452




The values of Ct are generally between 1 and 10 nF. The capacitance of a cable is close to

0.4 nF/m. Thus, several tens of metres of cable will greatly reduce the overvoltage transferred

to the secondary.

In general, the network is sufficiently widespread for the overvoltage transferred not to raise

any difficulties.

However, in the case of a short cable, e.g. between a specific transformer and a load (arc

furnace, etc.), the overvoltage transferred may be unacceptable for the equipment on the low

voltage side.

To reduce the magnitude of the impulse transferred, it is possible to:

- use a surge arrester with a lower sparkover voltage on the high voltage side

- install a surge arrester on the low voltage side between each phase and earth

- increase the capacitance between each phase and earth on the low voltage side.

5.1.5. Propagation of overvoltages

Overhead lines and cables constitute a propagation media for any overvoltage wave likely to

occur on a network.

For high frequencies (case of switching and lightning overvoltages), the line is characterised by

its so-called "characteristic" or "wave" impedance:

ZL

Cc ≈

L : line inductance

C : line capacitance

We can see that this impedance is independent of the length of the line.

The speed of the wave propagation on an overhead line is close to the speed of light:

c = ×3 108 m/s (300 m/µs)

for cables, it is equal to vc

r

=ε

ε r : relative permittivity of the cable insulating material

453




The value of v is close to 150 m/µs.

This gives us an idea of the way a lightning wave spreads along a conductor. Figure 5-45

shows how a lightning wave spreads along an overhead line in relation to time and space.

t (to x constant)

V front : 200 kV / µs

2 µs

400 kV

developmentin time

x (to t constant)

V front : 0.66 kV / m

600 m = 300 x 2 µs

400 kV

spread inspace

Figure 5-45: diagram showing how a lightning wave spreads along an overhead line

in relation to time and space

Let us closely examine the phenomenon that is produced at a point M , where a change of

impedance exists, separating two circuits with characterstic impedances of Z1 and Z2

(see fig. 5-46).

i1'

MZ1 Z2

i1

v1 v2

i2

v1'

Z Z1 2, : upstream and downstream characteristic impedances

v i1 1, : incident wave upstream of M

v i2 2, : wave transferred downstream of M

v i1 1' ', : wave reflected upstream of M

Figure 5-46: propagation of a wave at a change of impedance point M

454




Upstream of M , we have:

v Z i1 1 1= and v Z i1 1 1' '= − (1)

immediately downstream of M :

v Z i2 2 2= (2)

at point M :

v v v2 1 1= + ' and i i i2 1 1= + ' (3)

We can thus deduce:

( )v v v v Z i v Z i i2 1 1 1 1 1 1 1 2 1= + = − = − −' '

whencev

Z

Z Zv2

2

2 112=

+×

In particular:

- for a line short-circuited to earth, Z2 0= ; we can deduce from this that v2 0= and v v1 1' = −

- for a conductor without a change of impedance, Z Z2 1= ; we can deduce from this that

v v2 1= and v1 0' =

- for an open line, Z2 = ∞ ; we can deduce from this that v v2 12= and v v1 1' = .

To conclude, at the point of change of impedance, the maximum voltage value may reach

double the incident wave. This is the case of a line feeding a transformer as its impedance in

relation to the lightning wave is very high in relation to the characteristic impedance of the line.

455




5.2. Overvoltage protection devices

5.2.1. Principle of protection

The protection of installations and persons against overvoltages is greatly improved when

disturbances flow to earth, and this is done as close as possible to the sources of disturbance.

This requires low impedance earth electrodes to be implemented.

Thus, three overvoltage protection levels can be distinguished:

n 1st protection level

The objective is to avoid a direct impact on structures by catching the lightning and directing it

towards designated flow points, via:

- lightning conductors, whose principle is based on the striking distance; a rod placed at the

top of a structure to be protected captures the lightning and evacuates it through the

earthing network (see fig. 5-40-b)

- meshed or Faraday cages

- lightning protection ropes (see fig. 5-42).

n 2nd protection level

Its aim is to ensure that the basic impulse level (BIL) of the substation components has not

been exceeded.

In HV, this type of protection is established using elements ensuring that the lightning wave

flows to earth, such as:

- spark-gaps

- HV surge arresters.

n 3rd protection level

Used in LV as an extra protection for sensitive equipment (computers, telecommunication

devices, etc.).

It uses:

- series filters

- overvoltage limiters

- LV surge arresters.

456




5.2.2. Spark-gaps

n operation

The spark-gap is a simple device made up of two electrodes, the first connected to the

conductor to be protected and the second connected to earth.

At the place where it is installed in the network, the spark-gap constitutes a weak point where

overvoltages can flow to earth and thus protects the equipment.

The sparkover voltage of the spark-gap is set by adjusting the distance in the air between the

electrodes so as to obtain a margin between the impulse withstand of the equipment to be

protected and the impulse sparkover voltage of the spark-gap (see fig. 5-47). For example,

B = 40 mm on French public EDF 20 kV networks.

bird proof rod

earth electrode phase electrode

45°

electrode

holder

B

rigid

anchoring chain

device for adjusting B

and locking the electrode

45°

Figure 5-47: MV spark-gap with birdproof rod

457




n advantages

The main advantages of spark-gaps:

- their low price

- their simplicity

- the possibility of setting the sparkover voltage.

n drawbacks

- The sparkover characteristics of the spark-gap are highly variable (up to 40 %) depending

on the atmospheric conditions (temperature, humidity, pressure) which modify the ionization

of the dielectric medium (air) between the electrodes.

- the sparkover level depends on the overvoltage.

- spark-gap sparkover causes a power frequency phase-to-earth short circuit owing to the arc

being maintained. The short circuit lasts until it is cleared by the switching devices (this

short circuit is called a follow current). This means that it is necessary to install shunt circuit-

breakers or rapid reclosing system on the circuit-breaker located upstream. Because of this,

the spark-gaps are unsuitable for the protection of an installation against switching

overvoltages.

- the sparkover caused by a steep front overvoltage is not instantaneous. Due to this delay,

the voltage actually reached in the network is higher than the chosen protection level. To

take this phenomenon into account, it is necessary to study the voltage-time curves of the

spark-gap.

- sparkover causes the appearance of a steep front broken wave which could damage the

windings of the transformers or motors located nearby.

Although still used in certain public networks, spark-gaps are currently being replaced by surge

arresters.

458




5.2.3. Surge arresters

To overcome the drawbacks of spark-gaps, different models of surge arresters have been

designed with the aim of ensuring better protection of installations and good continuity of

service.

Non-linear resistor type gapped surge arresters are especially found in HV and MV

installations which have been in operation for several years. The current tendency is to use

zinc oxide surge arresters which provide better performance.

n definitions

Surge arrester discharge current

The surge or impulse current which flows through the arrester after a sparkover of the series

gaps.

Surge arrester follow current

The current from the connected power source which flows through an arrester following the

passage of discharge current.

Surge arrester residual voltage

The voltage that appears between the terminals of an arrester during the passage of discharge

current.

5.2.3.1. Non-linear resistor type gapped surge arresters (see IEC 99-1)

n operating principle

In this type of surge arrester, a variable resistor (varistor), which limits the current after the

passage of the impulse wave, is associated with a spark gap.

After evacuation of the impulse wave to earth, the surge arrester is only subjected to the

network voltage and the follow current is limited by the varistor.

The arc is systematically extinguished after the 50 Hz wave of the single-phase-to-earth fault

current has reached zero.

Owing to the variation of the resistance, the residual voltage is maintained close to the

sparkover level. Indeed, this resistance decreases with the increase in current.

Various techniques have been used to make non-linear resistor type gapped arresters. The

most conventional method uses a silicon carbide (SiC) resistor.

Some surge arresters also have voltage grading systems (resistive or capacitive dividers) and

arc blowing systems (magnets or blow-out coils).

459




n characteristics

Variable resistor type surge arresters are characterised by:

- the rated voltage, which is the maximum specified value of the power frequency rms voltage

permitted between its terminals for which the surge arrester is designed to function

correctly. This voltage can be continuously applied to the surge arrester without this

modifying its operating characteristics.

- the sparkover voltages for the different wave forms (power frequency, switching impulse,

lightning impulse, etc.).

- the impulse current evacuation capacity.

5.2.3.2. Zinc oxide ( ZnO ) surge arresters

n operating principle

Figure 5-48 shows that, unlike the non-linear resistor type gapped surge arrester, the zinc

oxide surge arrester is only made up of a highly non-linear variable resistor.

The resistance goes from 1.5 MΩ at the duty voltage (which corresponds to a leakage current

below 10 mA) to 15 Ω during discharge.

Following the passage of the discharge current, the voltage at the terminals of the surge

arrester become equal to the network voltage. The current which flows through the surge

arrester is very weak and is stabilised around the value of the earth leakage current.

Because of the high non-linearity of the ZnO surge arrester a high current variation causes a

low voltage variation (see fig. 5-49).

For example, when the current is multiplied by 107, the voltage is only multiplied by 1.8.

460




connecting spindle

flange(aluminium alloy)

exhaust pipe andoverpressure device

in the upper and

lower flanges

fault indication

plate

exhaust pipe

flange

ring clamping

device

overpressure device

prestressed tightnessdevice

rubber seal

compression spring

porcelain enclosure

thermal shield

spacer

washer

rivet

elastic stirrup

blocksZnO

Figure 5-48: example of the structure of a ZnO surge arrester in a porcelain enclosure

for 20 kV networks

600500

400

300

200

100

.001 .01 .1 1 10 100 10000

peak kV

I

U

1000

SiC

linear

Z On

SiC : non-linear resistor type gapped surge arrester made up of a silicon carbide resistor

ZnO : zinc oxide surge arrester

linear : U curve proportional to I

Figure 5-49: characteristics of two surge arresters having the same 550 kV/10 kA protection level

461




n characteristics

ZnO surge arresters are characterised by:

- the steady-state voltage which is the permitted specified value of the power frequency rms

voltage that can be continuously applied between the terminals of the surge arrester

- the rated voltage which is the maximum power frequency rms voltage permitted between its

terminals for which the surge arrester is designed to operate correctly in the temporary

overvoltage conditions defined in the operating tests (a power frequency overvoltage of 10

seconds is applied to the surge arrester - see IEC 99-4)

- the protection level defined at random as being the residual voltage of the surge arrester

when it is subjected to a given current impulse (5,10 or 20 kA according to the class), with a

wave form of 8/20 µs

- steep front current impulse (1 µs), lightning impulse (8/20 µs), long duration impulse, and

switching impulse withstand

- nominal discharge current.

Table 5-4 gives an example of the characteristics of a phase-to-earth ZnO surge arrester for

a 20 kV public distribution network (with tripping on occurrence of the first fault).

Maximum steady-state voltage (phase-earth) 12.7 kV

Rated voltage 24 kV

Residual voltage for nominal discharge current < 75 kV

Nominal discharge current (8/20 µs wave) 5 kA

Impulse current withstand (4/10 µs wave) 65 kA

Table 5-4: example of the characteristics of a ZnO surge arrester for a 20 kV network

462




n choice of zinc oxide surge arresters in HV

The general method for choosing a zinc oxide surge arrester in HV consists in determining its

characteristic parameters using the network data, at the place where it will be installed.

The parameters characterising the surge arrester are:

- UC , steady-state voltage

- Ur , rated voltage

- Ind , nominal discharge current

- discharge class and energy capacity

- mechanical characteristics.

The data relative to the network are:

- Um , highest phase-to-phase voltage applied to equipment

- TOV temporary overvoltages (appearing on occurrence of an earth fault or load shedding

on the public distribution network).

The choice of the surge arrester involves making a compromise between the equipment

protection levels and the energy capacity of the surge arrester.

The protection level must be as low as possible for the equipment withstand. This involves the

lowest voltage rating possible and thus greater difficulty withstanding temporary overvoltages.

o determining UC and Ur

simplified method using equipment characteristics

The voltages UC and Ur may be directly determined using the highest voltage for the

equipment Um :

UU

Cm≥3

U Ur C= ×125.

463




more accurate method using temporary overvoltages

The simplified method has a drawback as it does not take into account the real requirements

of the network which are generally lower thanUm

3.

The temporary overvoltages likely to occur in a network are of two types:

- overvoltages due to a phase-earth fault the clearance time of which depends on the

protection system (see table 5.1 - the earth overvoltage factor is equal to 1.73 for unearthed

or impedance earthed networks)

- overvoltage due to load-shedding on the public distribution network, of the order of 15 %

but able to reach 35 % in some networks.

The temporary overvoltage value to be taken into account is the product of the earth fault

overvoltage and load shedding factors.

- specific case

If one of the temporary overvoltages lasts over 2 hours, it is considered to be a steady-state

condition for the surge arrester and thus UC is chosen to be equal to this overvoltage and

U Ur C= ×125.

- general case

A surge arrester's capacity to withstand temporary overvoltages is given in relation to an

equivalent voltage lasting 10 seconds ( )U s10 expressed in the following equation:

U TOVT

s1010

=

ηwhere η ≅ 0 02.

T : overvoltage duration

TOV : overvoltage value

This formula allows the 10 second overvoltage which would cause the same stress on the

surge arrester to be calculated for each temporary overvoltage.

The duration of the temporary overvoltage must be between several seconds and two to three

hours (U TOVs10 0 97= ×. for T s= 2 and U TOVs10 114= ×. for T hours= 2 ).

The rated voltage of the surge arrester will be chosen to be above or equal to the maximum

value of the equivalent 10 second voltages: ( )U Ur s≥ max 10 .

We will take UU

Cm≥3

464




o nominal discharge current Ind

In practice, for the voltage range 1 52kV U kVm≤ ≤ , two values of Ind are available: 5 kA and

10 kA.

The value I kAnd =10 is chosen for areas with a high lightning density.

o discharge class and energy capacity

These are determined by testing or comparison with identical projects.

o mechanical characteristics

The IEC 99-4 and 99-5 standards fix the allowable pressure limit (expressed in "kA") which

must be met for the three-phase short circuit at the surge arrester terminals.

The surge arrester characteristics will also be checked in relation to:

- the ambient temperature

- the altitude

- the level of pollution

- the mechanical resistance to the wind, seismic stress, frost.

o surge arrester protection level

The protection level of the surge arrester at the installation point corresponds to the residual

voltage ( )Ursd at its terminals when its nominal discharge current flows through it.

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5.2.3.3. Installation of HV and MV surge arresters

In HV and MV electrical networks, surge arresters are installed at the entrance to the

substation to ensure protection of the substation transformer and equipment. This protection

only works if the protection distance and the installation rules are respected.

n protection distance

The wave propagation phenomenon studied in § 5.1.5. shows that at the point of reflection

(e.g. MV/LV transformer), the overvoltage reaches double the value of the incident wave.

The surge arrester peaks at a sparkover voltage Ursd (equal to the residual voltage for ZnO

surge arresters).

If it is located a considerable distance away, the maximum voltage at the location of the

equipment to be protected will thus be 2Ursd . Now, the equipment impulse withstand is

generally lower than 2Ursd .

To overcome this drawback, the surge arrester is installed at a shorter distance away than the

"protection" distance D . The surge arrester then undergoes the sum of the incident wave and

the reflected wave. It is thus sparked for an incident wave below Ursd .

Assuming that at the equipment termination point, the wave is totally reflected, we can show

that the overvoltage in relation to the equipment is limited to U U rD

vrsd= + 2

rdV

dt= : rise front steepness of the voltage wave, kV/µs

v : wave propagation speed, m/µs

For a lightning impulse withstand voltage Ul , the surge arrester must therefore be located at

a distance D such that:

U rD

vUrsd l+ ≤2

whence DU U

rvl rsd≤

−⋅

2

466




Numerical application:

Let us consider the example illustrated in figure 5-50:

U kVl =125 , case of an MV/LV transformer complying with IEC 76.3

U kVrsd = 75 , residual voltage of the surge arrester

r kV s= 300 / µ , voltage wave rise front steepness

v m s= 300 / µ , for an overhead line (speed of light)

we then have D ≤ −×

×125 75

2 300300

D m≤ 25

The surge arrester must therefore be installed less than 25 m away from the transformer for

the overvoltage not to exceed the lightning impulse withstand value.

lightning impulse

overhead line

surge arrester

transformer

A BD

ZCZC

Figure 5-50: protection distance of a surge arrester protecting a transformer fed

by an overhead line

467




5.2.4. protection of LV installations

n general

LV installations are protected against overvoltages by installing devices in parallel; 3 types of

devices are used:

- overvoltage limiters located on the secondary of MV/LV transformers (only in an IT

earthing system); they only provide protection against power frequency overvoltages

- low voltage surge arresters installed in LV switchboards or incorporated in loads

- surge diverters designed to protect telephone networks, LV terminal boxes and loads.

The main technologies used are:

- zener diodes

- gas discharge tubes

- zinc oxide varistors.

Zener diodes have the drawback of only ensuring the protection of a precise point in the

network. The gas discharge tube requires the addition of a varistor to prevent follow current.

Variable resistor-type surge arresters are currently the most cost-effective solution owing to

their simplicity and reliability

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n LV surge arrester installation rules

The equipment is only protected properly if certain installation rules are followed:

- rule 1

The length of the connection between the surge arrester and its disconnecting circuit-breaker

must be below 0.5 m.

disconnectingcircuit-breaker

L < 50 cm load to be

protected

Figure 5-51: diagram of connections

- rule 2

The outgoing feeders of the protected conductors must be connected to the terminals of the

surge arrester and its disconnecting circuit-breaker.

- rule 3

The loop surfaces must be reduced by tightly grouping together the incoming, phase, neutral

and PE wires.

- rule 4

the incoming wires of the surge arrester (polluted) must be moved away from the protected

outgoing wires (healthy) in order to avoid any possible electromagnetic coupling.

- rule 5

The cables must be flattened against the metal structures of the box in order to reduce frame

loops and thus benefit from a reduction in disturbances.

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n connection layout according to the earthing system

In figures 5-52-a and 5-52-b the connection layouts of the LV surge arrester are shown for

different earthing systems.

main earthterminal

(entrenchedloop)

load earthelectrode

electrical switchboard

disconnecting

circuit-breaker

equipment

to be protected

PE

LV neutral

earth electrode

PE

surge arrester

RCD

Ph1

Ph2

Ph3

N

TT earthing system


disconnecting

circuit-breaker

equipment

to be protected

PE

LV neutralearth electrode

main earthterminal

(entrenchedloop)

loadearth

electrode

PE

overvoltage

limiterPIM

surge

arrester

Ph1

Ph2

Ph3

N

IT earthing system

Figure 5-52-a: connection layout of an LV surge arrester for TT and IT earthing systems

470




main earthterminal

(entrenched

loop)

load earthelectrode


equipment to

be protected

surge arrester

LV neutralearth electrode

PEN

disconnecting

circuit-breaker

Ph1

Ph2

Ph3

PEN

TNC earthing system

main earth

terminal(entrenchedloop)

load earth

electrode


disconnecting

circuit-breakerequipment to

be protected

PE

LV neutral

earth electrode

PE

surge arrester

PE

Ph1

Ph2

Ph3

N

TNS earthing system

Figure 5-52-b: connection layout of an LV surge arrester

for TNC and TNS earthing systems

471




5.3. Insulation co-ordination in an industrial electrical network

5.3.1. General

Co-ordinating the insulation of an installation consists in determining the insulation

characteristics necessary for the various network elements, in view to obtaining a withstand

level that matches the normal voltages, as well as the different overvoltages.

Its ultimate purpose is to provide dependable and optimised energy distribution.

Optimal insulation co-ordination gives the best cost-effective ratio between the different

parameters depending on it:

- cost of equipment insulation

- cost of overvoltage protections

- cost of failures (loss of operation and destruction of equipment), taking into account their

probability of occurrence.

With the cost of overinsulating equipment being very high, the insulation cannot be rated to

withstand the stress of all the overvoltages studied in paragraph 5.1.

Overcoming the damaging effects of overvoltages supposes an initial approach which consists

in dealing with the phenomena that generate them, which is not always very easy. Indeed, if

using the appropriate arc interruption techniques the switchgear switching overvoltages can be

limited, it is impossible to prevent lightning strikes.

n clearance (see fig. 5-53)

This term covers two notions:

- gas clearance (air, SF6, etc.), which is the shortest path between two conductive parts.

- creepage distance: this is also the shortest path between two conductors, but following the

outer surface of a solid insulating material (e.g. insulator).

472




The clearance is directly related to the withstand of the equipment to different overvoltages.

creepage

distancedistancein air

distancein air

Figure 5-53: air clearance and creepage distance

n overvoltage withstand

The overvoltage withstand depends on the type of overvoltage applied (magnitude, wave form,

frequency and duration, etc.).

It is also influenced by external factors such as:

- ageing

- environmental conditions (humidity, pollution)

- variation in air or insulating gas pressure.

n withstand voltage

Electrical equipment is characterised by its withstand voltage to different types of overvoltages.

We can thus distinguish:

- the power frequency withstand voltage

- the switching impulse withstand voltage

- the lightning impulse withstand voltage.

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o power frequency withstand voltage

This corresponds to the equipment withstand to power frequency overvoltages likely to occur

on the network and the duration of which depends on the network operating and protection

mode.

The equipment withstand is tested by applying a sinusoidal voltage with a frequency of

between 48 Hz and 62 Hz for one minute. The test is valid for nominal network frequencies of

50 Hz and 60 Hz (see IEC 71-1).

o switching impulse withstand voltage

This characterises the equipment withstand to switching impulses (only for equipment with a

standard voltage above or equal to 300 kV).

The equipment test (see IEC 60-1) is performed by applying a wave with a front time of 250 µs

and a time to half-value of 2500 µs.

o lightning impulse withstand voltage

This characterises the equipment withstand to the 1.2 µs / 50 µs lightning voltage wave.

This withstand voltage concerns all voltage ranges, including low voltage.

o examples of equipment withstand (see table 5-5)

Highest voltage for the

equipment

Um (kV) (1)

(r.m.s. value)

Standard short-duration

power frequency withstand

voltage (kV)

(r.m.s.)

Standard lightning impulse

withstand voltage (kV)

(peak value)

3.6 10 2040

7.2 20 4060

12 28 607595

17.5 38 7595

24 50 95125145

36 70 145170

52 95 250

72.5 140 325

(1) Um is the highest rms value of the phase-to-phase voltage for which the equipment is specified.

474




Table 5-5: standard withstand voltages for 3.6 kV < Um < 72.5 kV

475




5.3.2. Reduction in risks and overvoltage levels

The risks of overvoltages, and consequently the danger they represent for persons and

equipment, can be greatly reduced if certain measures of protection are taken:

- limiting substation earth electrode resistances in order to reduce power frequency

overvoltages

- reducing switching overvoltages by choosing suitable switchgear (interruption in SF6)

- making lightning impulses flow to earth by a first clipping operation (surge arrester or spark-

gap at the entrance to the substation) with limitation of the earth electrode resistances and

pylon impedances

- limiting the residual voltage from the first clipping operation by HV surge arrester which is

transferred to the downstream network by providing a second protection level on the

transformer secondary

- protection of sensitive equipment in LV (computers, telecommunications, automatic

systems, etc.) by connecting series filters and/or overvoltage limiters to it.

5.3.2.1. Rise in potential of LV exposed conductive parts following an MV fault in the

transformer substation

In this paragraph, we propose to study overvoltages in LV caused by an earth fault on the MV

side in an MV/LV substation, and the measures to be taken in order to protect equipment and

persons, in compliance with IEC 364-4-442.

The values of rises in potential of the substation or LV installation exposed conductive parts

depend on the values of the earth electrode resistances, the fault current values and the

earthing system.

n earthing in transformer substations

A single earth electrode must be installed in a transformer substation, to which must be

connected:

- the transformer tank

- the metallic coverings of high voltage cables

- the earth conductors of high voltage installations

- the exposed conductive parts of high voltage and low voltage equipment

- the extraneous conductive parts.

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n symbols

In the following paragraphs, the symbols used have the following signification:

Im : part of the earth fault current in the high voltage installation which flows through the earth electrode of the

transformer substation exposed conductive parts

Re : transformer substation earth electrode resistance

V : low voltage installation phase-to-neutral voltage

U : low voltage installation phase-to-phase voltage

U f : fault voltage in the low voltage installation, between the exposed conductive parts and earth

U1 : stress-voltage in the transformer substation low voltage equipment

U2 : stress-voltage in the installation low voltage equipment

n TN a− and IT a− earthing systems (see fig. 5-54)

In these two systems, the substation, neutral and installation earth electrodes are the same.

Inside the equipotential area, the ground and exposed conductive part potentials increase

simultaneously. The touch voltage U f is then zero.

On the other hand, outside this area, the ground potential remains equal to that of the remote

earth, while the potential of the exposed conductive parts increases to U R If e m= .

Thus, when there are exposed conductive parts outside of the equipotential area and the

touch voltage U R If e m= cannot be cleared in the time defined in tables 2-3-a and 2-3-b, the

TN a− and IT a− earthing systems are not acceptable in relation to the protection of

persons.

To overcome this drawback, the following provisions must be taken:

- TN a− earthing system: the neutral of the LV installation must be connected to a separate

earth electrode, which is the case in the TN b− earthing system (see fig. 5-55)

- IT a− earthing system: the exposed conductive parts of the LV installation must be

connected to a separate earth electrode from that of the substation, which is the case in the

IT b− earthing system (see fig. 5-56).

TN b− and IT b− earthing systems allow dangerous touch voltages to be cleared but make

overvoltages occur:

- in the installation LV equipment for the IT b− earthing system

- in the substation LV equipment for the TN b− earthing system.

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MV ph 1

ph 2

ph 3

PEN

LV

U1 U2

Im

U f 0

U V1

equipotential zone outside

zone

substation LV installation

U U V2 1

U R If e m

Re

TN a−

MV LV

U1 U2

Im

U f 0

equipotential zone outsite zone

Z

U V1 3*

U U V2 1 3*

(*) a first LV fault is present

Re

U R If e m

IT a−

Figure 5-54: rise in potential of TN-a and IT-a earthing systems

478




n TN b− , TT b− and IT c− earthing systems (see fig. 5-55)

In these three systems, we can see a rise in potential of the exposed conductive parts of the

substation U1 such that:

U R I Ve m1 = + for TN b− and TT b− earthing systems

U R I Ve m1 3= + . for IT c− earthing systems with the presence

of a first fault on the LV side

Depending on the maximum current value Im , the values of Re must be limited so that U1

remains below the power frequency withstand voltage Utp of the substation equipment.

U U tp1 ≤

Table 5-6 gives the maximum values of Re for different values of Im and Utp .

Values at Re not to be exceeded

Fault current Im

(A)

Utp = 2 000 V

Class I

Utp = 4 000 V

Class II

Utp = 10 000 V

Special class

TN b− ; TT b− IT c− TN b− ; TT b− ; IT c− TN b− ; TT b− ; IT c−

300 A 5.9 Ω 5.3 Ω 12 Ω 30 Ω

1 000 A 1.8 Ω 1.6 Ω 3.6 Ω 10 Ω

5 000 A 0.35 Ω 0.32 Ω 0.72 Ω 2 Ω

Table 5-6: maximum values of Re in TN b− , TT b− and IT c− earthing systems

479




MV ph 1

ph 2

ph 3

PEN

LV

U1 U2

Im RB

U R I V

U V

U

m

f

1

2

0

e

Re

U f

TN b−

MV ph 1

ph 2

ph 3

N

LV

U1 U 2

Im RB RAU f

U R I V

U V

U

e m

f

1

2

0

Re

TT b−

MV LV

U1 U 2

Im Z RA U fIf

(*) a first LV fault is present

U R I V

U V

U R I U

m

f A f L

1

2

3

3

*e

Re

IT c−

Figure 5-55: rise in potential in TN b− , TT b− and IT c− earthing systems

480




nTT a− and IT b− earthing systems

In these two cases the substation earth electrode and that of the neutral are common.

The LV installation earth electrode is separate.

The earth fault current flows through the common earth electrode (neutral/substation).

As shown in figure 5-56, we can see that there is a risk of breakdown for the LV equipment

whose earth electrode is separate from that of the substation.

The following conditions must be met:

U R I VtM e m> + for the TT a− earthing system

and U R I VtM e m> + 3 for the IT b− earthing system

whenceR

U V

I

RU V

I

etM

m

etM

m

<−

<−

3

for the TT a− earthing system

for the IT b− earthing system

where:

U tM : power frequency withstand voltage of the installation LV equipment equal to 2 1000V + for V = 220 to

250 V, i.e. 1500 V

Table 5-7 gives the values of Re for different values of Im .

TT a− IT b−

Im = 300 A 4 Ω 3.5 Ω

Im = 1000 A 1.2 Ω 1 Ω

Im = 5000 A 0.24 Ω 0.2 Ω

Table 5-7: maximum values of Re in TT a− and IT b− earthing systems

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MV L 1

L 2

L 3

N

LV

U1 U2

Im

substation LV installation

RA

U V1 U R I V

U

m

f

2

0

eU f

Re

TT a−

MV LV

U1 U2

Im RA

U f

If

Z

(*) a firs t LV fault is present

U V

U R I V

U R I U

m

f A f L

1

2

3

3

*

*e

Re

IT b−

Figure 5-56: Rise in potential in TT a− and IT b− earthing systems

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n recapitulative table of touch voltages and overvoltages which occur for each earthing

system

TN a− IT a− TT a− IT b− TN b− TT b− IT c−

Touch voltage Y Y N N N N N

Overvoltage of LV

installation exposed

conductive parts

N N Y Y N N N

Overvoltage of

substation exposed

conductive parts

N N N N Y Y Y

Y : yes

N : no

Table 5-8: touch voltages and overvoltages which occur for

each earthing system

5.3.2.2. Rise in potential of the LV exposed conductive parts on occurrence of a

lightning impulse

When a lightning overvoltage from the distribution network flows to earth in an MV/LV

substation through a protection device (surge arrester or MV spark-gap), there follows a rise in

potential of the substation LV exposed conductive parts and/or those of the installation which

depends on the earthing system.

The level of overvoltages transferred in LV depends on the clipped value Ursd and the earth

electrode values.

To ensure protection of the LV switchgear against these overvoltages, LV surge arresters must

be installed and the resistance of the substation earth electrode limited so that the equipment

lightning impulse withstand voltage is not exceeded.

n limiting earth electrode impedances

As for the case of the MV earth fault, the limit values of the earth electrode impedances are

calculated for each earthing system.

483




The overvoltage at a point on the network where the impedance changes is given in the

relation:

vZ

Z Zv2

2

1 212=

+(see § 5.1.4)

v U rsd1 = : corresponds in this case to the clipped overvoltage

v2 : overvoltage of the substation exposed conductive parts

Z Zc1 = : characteristic impedance of the medium voltage line

Z Ze2 = : substation earth electrode impedance

We thus have:

vZ

Z ZUe

c ersd2 2=

+.

The equipment lightning impulse voltage Utc must be above the overvoltage v2 , whence:

UZ

Z ZUtc

e

c ersd≥

+. 2

( )ZZ

ec

UUrsd

tc

≤−21

For U kVrsd = 120 and Zc = 330Ω , the impulse impedance Ze is equal to 1.5 times the

resistance Re measured in low frequency: RZ

ee=15.

.

The condition on the value of the substation earth electrode impedance is thus:

( )RZ

ecUUrsd

tc

≤× −15 1.

The maximum values of Re for the different earthing systems are given in table 5-9.

Earthing system TN b− , TT b− , IT c− TT a− , IT b−

Utc (kV) 4 8 20 3

Re 3.8 7.7 20.2 2.7

Table 5-9: maximum values of the MV/LV substation earth electrode resistances

recommended for limiting MV atmospheric overvoltages transferred in LV

05 overvoltages and insulation coordination

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