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Current interruption using high voltage air-break
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Peelo, D.F.
Published: 01/01/2004
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Peelo, D. F. (2004). Current interruption using high voltage air-break disconnectors Eindhoven: Technische
Universiteit Eindhoven
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CURRENT INTERRUPTION USING
HIGH VOLTAGE AIR-BREAK DISCONNECTORS
PROEFSCHRIFT
ter verkrijging van de graad van doctor aan de
Technische Universiteit Eindhoven, op gezag van de
Rector Magnificus, prof.dr. R.A. van Santen, voor een
commissie aangewezen door het College voor
Promoties in het openbaar te verdedigen
op dinsdag 16 maart 2004 om 16.00 uur
door
David Francis Peelo
geboren te Dublin, Ierland
Dit proefschrift is goedgekeurd door de promotoren:
prof.dr.ir. R.P.P. Smeets
en
prof.ir. L. van der Sluis
CIP-DATA LIBRARY TECHNISCHE UNIVERSITEIT EINDHOVEN
Peelo, David F.
Current interruption using high voltage air-break disconnectors / by David F.
Peelo. – Eindhoven : Technische Universiteit Eindhoven, 2004.
Proefschrift. – ISBN 90-386-1533-7
NUR 959
Trefw.: hoogspanningsschakelaars / boogontladingen / elektrische doorslag.
Subject headings: switchgear / switching / circuit-breaking arcs / electric
breakdown.
A great flame follows a tiny spark.
—Dante Alighieri (1265–1321)
To my wife Anna-Lena, my
children Anna-Maria and Nicholas and
my granddaughters Anna and Emelie
i
SUMMARY
High voltage air-break disconnectors are intended for use as isolators and as such are operated under energized conditions. The disconnectors will therefore be required to interrupt
unloaded transformer magnetizing, capacitive or loop currents in air dependent on the circumstances and the practices of individual utilities. Each of these switching duties is unique
in terms of the arc-circuit interaction, arc sustainability and arc extinction. This research
investigates this arc behaviour with particular emphasis on the loop switching case.
The interruption of unloaded transformer magnetizing currents of 1 A or less is mainly
dependent on achieving a sufficient disconnector contact gap spacing to withstand the transient recovery voltage. For currents greater than 1 A, thermal effects come into play and will
promote longer arcing times. Inrush current may occur, also having the effect of prolonging
the arcing time but not the arc length. The interruption process is one of repetitive breakrestrike with associated restriking overvoltages. The impact of the overvoltages on the transformer insulation structures is a matter for consideration.
For capacitive currents of 1 A or less, successful interruption is dependent on achieving the
minimum disconnector contact gap spacing to withstand the recovery voltage and on the ratio
of the source and load side capacitances CS/CL. For currents greater than 1 A, thermal effects
add to the complexity of the interruption process. The longest arcing times and highest
restriking overvoltages occur when CS/CL < 1. The explanation for this lies in the magnitude
of the equalization voltage immediately after restriking relative to the source voltage and the
associated restrike current magnitudes. A number of arcing modes can be identified dependent on the current magnitude and CS/CL.
The loop switching case is more complex with current interruption having the obvious
dependency on the current magnitude and the loop impedance. The switching duty is one of
current commutation from one circuit to a parallel circuit and arc extinction follows an initial
arc instability. The research shows that the condition for arc instability is similar to that for an
arc in a DC circuit. Potential for arc modelling is examined with a view to enabling simulation of this duty.
The research is principally based on tests and observations on vertical break and centre-break
type disconnectors. The extension of the research results and conclusions to double-break and
pantograph type disconnectors is discussed as is suggestions for further research into the
subject.
iii
SAMENVATTING
Titel: “Het schakelen van stromen met hoogspanningsscheiders in open-lucht opstelling”
Hoogspanningsscheiders hebben als functie netdelen te isoleren en worden geschakeld onder
spanning.
Scheiders zullen derhalve kleine stromen moeten onderbreken die oftewel capacitief van karakter zijn,
oftewel inductief (als magnetiseringsstroom van onbelaste transformatoren) oftewel het gevolg zijn
van een commutatie schakeling, afhankelijk van de praktijken in de diverse energiebedrijven.
Elk van deze schakelhandelingen is uniek in termen van wisselwerking tussen boog en circuit, de
mogelijkheden tot in stand blijven van de boog en uiteindelijk de onderbreking.
Deze studie onderzoekt dit booggedrag in scheiders in open-lucht opstelling, met vooral aandacht
voor het commutatief schakelen: het forceren van bedrijfsstroom uit een netdeel naar een parallel
geschakelde tak.
Het onderbreking van magnetiseringsstromen in onbelaste transformatoren van 1 A of minder wordt
vooral bepaald door het bereiken van een afstand tussen de scheider contacten, voldoende groot om de
transiënte wederkerende spanning te kunnen weerstaan. Bij stromen groter dan 1 A gaan thermische
processen een rol spelen, die langere boogtijden zullen veroorzaken. Inrush stromen kunnen optreden;
deze zullen de boogtijd verlengen, maar vergroten niet de lengte van de boog. Het onderbrekingsproces wordt hier gekenmerkt door een opeenvolging van onderbrekingen en herontstekingen, met de
daarbij behorende overspanningen. De gevolgen van deze overspanningen op de isolatie van
transformator wikkelingen moeten in acht genomen worden.
In het geval van capacitieve stromen van 1 A of minder, wordt een geslaagde onderbreking ten eerste
bepaald door het bereiken van een minimale afstand tussen de scheider contacten om de wederkerende
spanning te kunnen weerstaan en ten tweede door de verhouding van bron- en lastzijde capaciteit
Cs/Cl.
Voor stromen groter dan 1 A, maken thermische processen het onderbreken complexer. De langste
boogtijden en de hoogste overspanningen als gevolg van herontstekingen treden op wanneer
Cs/Cl < 1. De verklaring hiervoor moet gezocht worden in de grootte van de vereffeningsspanning
meteen na de herontsteking ten opzichte van de bronspanning enerzijds, en in de grootte van de
bijbehorende stromen anderzijds. Een aantal verschijningsvormen van de boog kan worden
vastgesteld, afhankelijk van de grootte van de stroom en de verhouding Cs/Cl.
Het commutatief schakelen is ingewikkelder, waarbij de stroom onderbreking wordt bepaald door de
grootte van de stroom en de impedantie van de lus waarin de commutatie plaats vindt. De
schakelhandeling bestaat uit commutatie van stroom uit een circuit naar een parallel circuit waarbij
het doven van de boog het gevolg is van een aanvankelijke instabiliteit.
Dit onderzoek toont aan dat de voorwaarde voor het optreden van een degelijke instabiliteit analoog is
aan die van een gelijkstroom boog.
De mogelijkheden van boog modellering zijn onderzocht met het oog op simulatie van deze
schakelhandeling.
Dit onderzoek richt zich met name op beproevingen en waarnemingen van scheiders met vertikaal
bewegende armen, en scheiders met een centrale scheiding. De extrapolatie van de onderzoeks
resultaten en -conclusies naar scheiders met dubbele onderbreking en pantograaf scheiders wordt
behandeld, en is als aanbeveling voor verder onderzoek op dit gebied neergelegd.
iv
CONTENTS
Summary............................................................................................................ iii
Samenvatting ......................................................................................................iv
1.
High voltage air-break disconnectors .......................................................1
1.1
1.2
1.3
1.4
1.5
2.
Literature review.......................................................................................13
2.1
2.2
2.3
2.4
2.5
2.6
3.
Introduction............................................................................................................33
Analysis..................................................................................................................34
Restriking and its consequences ............................................................................36
Inrush currents .......................................................................................................36
Auxiliary interrupting devices ...............................................................................40
Conclusions............................................................................................................45
Interrupting capacitive currents..............................................................47
4.1
4.2
4.3
4.4
4.5
4.6
4.7
5.
Introduction............................................................................................................13
Transformer magnetizing currents.........................................................................16
Capacitive currents.................................................................................................19
Loop currents .........................................................................................................20
Free burning arcs in air ..........................................................................................22
Conclusions............................................................................................................30
Interrupting transformer magnetizing current .....................................33
3.1
3.2
3.3
3.4
3.5
3.6
4.
Introduction..............................................................................................................1
Standards..................................................................................................................6
Bus and station arrangements ..................................................................................8
Perspective .............................................................................................................10
Objective of the research .......................................................................................10
Introduction............................................................................................................47
Analysis..................................................................................................................47
Auxiliary interrupting devices ...............................................................................49
Field experience .....................................................................................................49
Video record review...............................................................................................50
Capacitive current switching tests 2003 ................................................................63
Conclusions............................................................................................................70
Loop switching...........................................................................................71
5.1 Introduction............................................................................................................71
5.2 Loop switching tests 1999–2000 ...........................................................................71
5.2.1 Initial current, loop impedance and arcing time ........................................72
5.2.2 Arc video record analysis...........................................................................77
5.2.3 Application perspective .............................................................................84
5.3 Electrical characteristics of the arc ........................................................................93
5.4 Conclusions..........................................................................................................110
v
6.
Discussion on use of other disconnector types......................................113
7.
General conclusions and suggestions for further research .................117
8.
References ................................................................................................119
Annex A
Transformer transient recovery voltages.................................125
Annex B
EHV transformer switching case study....................................129
Annex C
Video still images: Figures C1 to C60 ......................................137
Annex D
Auxiliary interrupting devices and capacitive currents .........147
Annex E
Comparative analysis of loop switching tests by Andrews,
Janes and Anderson ...................................................................157
Annex F
Influence of weather...................................................................163
Acknowledgments ...........................................................................................167
Curriculum Vitae ............................................................................................168
vi
Section 1
High voltage air-break disconnectors
1.1
Introduction
The function of air-break disconnectors in high voltage power systems is to provide electrical
and visible isolation of one part of the system. The isolation generally takes two forms:
1. Isolation related to normal day-to-day operation of the power system. For example, shunt
reactors required only during light load periods are switched out using circuit breakers
and then isolated by disconnectors during peak load periods.
2. Isolation related to repair or maintenance on transmission lines or station equipment such
as transformers, circuit breakers and so on.
In the latter regard, the disconnectors are a major contributer to personnel safety. In North
America and no doubt similarly elsewhere, power system safety practices require a so-called
guaranteed point of isolation with a visible break and a disconnector mechanically locked in
the open position meets this requirement. If the disconnector is motor-operated, then the
electrical circuit of the operator is also visibly isolated by means of a knife switch or a
removable fuse link.
To serve the purpose of isolation, disconnectors are required to have a greater voltage withstand capability across the open gap than to ground. The purpose of this is to ensure that
surge voltages originating in the power system or due to lightning activity will more likely
cause flashover to ground than across the open gap. At system voltages of 245 kV and below,
this requirement adds at least 10% above the line to ground voltage withstand capability.1 At
system voltages of 300 kV and above, the requirement is stated as a bias voltage test, i.e. an
AC voltage applied to one side of the disconnector and a switching or lightning surge applied
to the other.
To achieve isolation disconnectors are operated under energized conditions and will thereby
interrupt current, the type of current being dependent on the circumstances. This interruption
of current in air by disconnectors is the subject of this thesis.
To establish the background for the subject, this section further provides an overview of the
following:
•
•
•
•
the different types of high-voltage air break disconnectors in use;
the type and ranges of current to be interrupted by disconnectors;
how the subject is treated in standards;
typical bus and station arrangements.
The section concludes with a perspective on the need for this work and the objectives of this
research effort.
1
Section 1
High voltage air-break disconnectors come in a variety of types and mounting arrangements.
The four most commonly used types are:
•
•
•
•
Vertical break type
Centre side break type
Double side break type
Pantograph type
Of these types, the vertical break type is the most used and is the type primarily considered in
this thesis.
The vertical break disconnector is shown in Fig. 1.1. The active parts of the disconnector are
the hinge end assembly, the blade and the jaw end assembly. The left-most insulator rotates
to open or close the disconnector. The blade is shown open to about a 60° angle from the
closed position, the hinge-end being to the left and the jaw end to the right. This disconnector
type is usually horizontally mounted (base frame horizontal as shown in Fig. 1.1) but can also
be vertically mounted (base frame and
active parts vertical) or, at medium
voltage, inverted mounted (base frame
horizontal with active parts also horizontal but underneath). Standard phase
spacings are used and overhead clearances must be such as to accommodate
the fully open disconnector blade.
Fig. 1.1 Vertical break disconnector horizontally
mounted
Courtesy of HAPAM B.V.
2
The centre break disconnector is
shown in Fig. 1.2. The active parts
consist of two blades which make and
break at the centre and both insulators
rotate to open or close the disconnector. This disconnector type is used
mainly where overhead clearances are
restricted but also where low substation profiles are desired. Because the
blades are reaching horizontally, the
phase spacing must be increased above
standard values and the disconnectors
thus require a larger area than the vertical break type.
High voltage air-break disconnectors
The double break disconnector
is a variation on the centre
break type and is shown in Fig.
1.3. The active parts consist of
two jaw assemblies, one at each
end, and a rotating blade. The
centre insulator rotates to open
or close the disconnector. The
disconnector requires an area
somewhat greater than that for
a centre break disconnector.
The pantograph switch, shown
in Fig. 1.4, is used quite widely
outside of North America and
only to a very limited degree
Fig. 1.2 Centre break disconnector
within North America. The
Courtesy of HAPAM B.V.
active parts consist of a fixed
stirrup arrangement attached to busbar at the top, a scissor type blade and a hinge assembly at
the bottom. The smaller of the two insulators rotates to open or close the disconnector. This
disconnector type clearly requires the least area and, in addition to providing isolation, also
provides transitions from high to low busbars.
Because disconnectors are expected to interrupt certain current levels as discussed later in
this section, it is desirable to avoid arcing on the main contacts or on corona shields (refer to
Fig. 1.1). For this reason disconnectors,
usually by customer request, are
equipped with arcing horns as shown in
Fig. 1.5. The blade is provided with an
arcing tip and on opening the current
commutates from the main contacts to
the arcing contacts thus achieving their
purpose. On closing, making (prestriking) occurs on the arcing horns. In
North American terminology, a disconnector equipped with arcing horns is
referred to as a horn gap disconnector
and requiring of a larger phase spacing
than a disconnector without arcing
horns. The arcing horns do not contribute to current interruption but rather
provide only a location for the arc to
Fig. 1.3 Double break disconnector
root itself.
Courtesy of HAPAM B.V.
3
Section 1
The interrupting capability of disconnectors can be increased by the addition of auxiliary interrupting devices.
These devices include gas-blast devices
(no longer in use), vacuum switches,
commutating devices and, most relevant to this work, whip type devices as
shown in Figs. 1.6 and 1.7. These devices achieve a fast moving contact
effect when the whip releases and are
widely used in North America to interrupt transformer magnetizing current
and small capacitive currents. The
devices are discussed further in
sections 3 and 4 and in Annex D.
High voltage air-break disconnectors
do not have current interrupting ratings.
Courtesy of HAPAM B.V.
However, by virtue of the fact that the
disconnectors have a fixed and a moving contact, they have a certain current interrupting
capability. In brief here but discussed in detail later, the currents to be interrupted are as
follows:
Fig. 1.4 Pantograph disconnectors
4
•
Transformer magnetizing currents:
The current at high voltage is usually less than 2 A equivalent rms
for modern low loss transformers
and is non-sinusoidal with a high
3rd harmonic content. The recovery voltage that appears across the
disconnector after current interruption is the difference between
the source side power frequency
voltage and the transformer side
damped low frequency (less than
300 Hz) oscillation.
•
Capacitive currents: For busbars
with connected instrument transformers, the current is less than
1 A but with some exceptions in
the range 1 to 2 A (EHV series
capacitor bank platforms). For
short lines, currents up to 20 A
Fig. 1.5 Vertical break disconnector jaw assembly
showing the main contacts and the arcing horn
Courtesy of HAPAM B.V.
High voltage air-break disconnectors
Fig. 1.6 Quick-break whip type device on 115 kV vertical break disconnector
(closed position)
Courtesy of Pacific Air Switches Corporation
Fig. 1.7 Quick-break whip type device operation (position just prior to
release of whip)
Courtesy of Pacific Air Switches Corporation
may be applicable. The recovery voltage that appears across the disconnector after
current interruption is the difference between the source side power frequency voltage
and the trapped DC charge related voltage on the busbar or line.
•
Loop currents: Loop currents can be 100 A or more dependent on individual utility
practice. The switching duty is a commutation or transfer of current from one circuit,
such as a busbar or transmission line, to a parallel circuit. In the case of disconnectors,
this is a natural commutation driven by arc voltage. As the arc voltage builds up, the
current in the disconnector is gradually reduced to zero by transfer to the parallel circuit. The power frequency voltage that appears across the disconnector after total
current transfer is referred to as the open circuit voltage or, for the case of current
transfer between busbars, as the bus-transfer voltage. For loop switching between
transmission lines, the open circuit voltage can be as high as 6 or 7 kV, but in most
cases is in the order of a few kV. For loop switching between busbars, the bus-transfer
voltage is less than 1000 V.
5
Section 1
1.2
Standards
Disconnector standards recognize the existence of current interrupting capability. The International Electrotechnical Commission (IEC) defines a disconnector as:1
“A mechanical switching device which provides, in the open position, an isolating distance in accordance with specific requirements.
NOTE: A disconnector is capable of opening and closing a circuit when either negligible
current is broken or made, or when no significant change in the voltage across the terminals of each of the poles of the disconnector occurs. It is also capable of carrying currents
under normal circuit conditions and carrying for a specified time currents under abnormal
conditions such as short-circuit.”
Two additional notes are applicable:
“NOTE 1: “Negligible current” implies currents such as the capacitive currents of bushings, busbars, connections, very short lengths of cable, currents of permanently connected
impedances of circuit-breakers and currents of voltage transformers and dividers. For
rated voltages of 420 kV and below, a current not exceeding 0.5 A is a negligible current
for the purpose of this definition; for rated voltages above 420 kV and currents exceeding
0.5 A, the manufacturer should be consulted.
“No significant change in voltage” refers to such applications as the bypassing of induction voltage regulators or circuit-breakers.
NOTE 2: For a disconnector having a rated voltage of 52 kV and above, a rated ability of
bus-transfer current switching may be assigned.”
With respect to Note 2 above, the applicable rated bus-transfer voltages are given in Annex B
of reference 1 and for convenience shown below:
6
High voltage air-break disconnectors
Bus-transfer is loop switching between busbars within a substation. Gas insulated disconnectors are those associated with gas insulated switchgear or GIS as it is commonly known.
The loops formed by such switchgear are short compared to those found in air insulated
switchgear arrangements and hence the lower recovery or bus-transfer voltages.
In North America, the term disconnecting or disconnect switch is used instead of disconnector. Such a device is defined by the Institute of Electrical and Electronic Engineers (IEEE)
as:2
“A mechanical switching device used for changing the connections in a circuit, or for
isolating a circuit or equipment from the source of power.
NOTE: It is required to carry normal load current continuously and also abnormal or
short-circuit currents for intervals as specified. It is also required to open or close circuits
when negligible current is broken or made, or when no significant change in the voltage
across the terminals of each of the switch poles occurs.”
The definitions are very similar both recognizing an ability to break negligible current. Only
IEC, however, states specific values: up to 0.5 A of capacitive charging current and, for specific disconnectors, a bus-transfer ability of 1600 A against open-circuit voltages of 100 V to
300 V (Annex B of reference 1).
An earlier version of an IEEE standard included the following note:3
“A disconnecting switch and a horn-gap switch have no interrupting rating. However, it is
recognized that they may be required to interrupt the charging current of adjacent buses,
supports and bushings. Under certain conditions, they may interrupt other relatively low
currents, such as:
1. Transformer magnetizing current.
2. Charging currents of lines depending on length, voltage, insulation and other local
conditions.
3. Small load currents.”
Horn-gap disconnectors generally have wider phase spacings. The implication is that such are
used to break currents and that some accommodation should be made for the reach of the arc
towards other phases or grounded structures. In fact, this note recognizes that disconnectors
are commonly used in North America to break small capacitive currents, transformer magnetizing currents and loop currents. The standard was originally an American National Standards Institute (ANSI) standard and was revised in 1992 to become an IEEE standard. At that
time, the above-discussed note was removed. The reason for this was that a guide on current
interruption had been developed4. The guide was based on the work of Andrews et al5 and
Peelo.6 For reasons discussed later in sections 3 and 5 and Annex E, the guide should be
viewed as questionable.
7
Section 1
1.3
Bus and station arrangements
It is evident that, essentially from conception, disconnectors were assumed to have a current
interrupting capability. Bus arrangements in turn exploited this capability thus avoiding the
use of more expensive load-break switches or circuit breakers. Examples of such bus
arrangements are shown in Figs. 1.8 to 1.11. The bus arrangement and required disconnector
current interrupting capability are as follows:
•
Fig. 1.8 shows a single bus arrangement
common at generating stations and the disconnectors are expected to have unloaded
transformer switching capability.
•
Fig. 1.9 shows a double bus arrangement
common outside of North America and the
disconnectors shown are expected to have a
bus-transfer loop switching capability.
•
Fig. 1.10 shows an H-bus arrangement and
the disconnects shown are expected to have
an unloaded transformer switching capability.
•
Fig. 1.11 shows a so-called Jones-type subFig. 1.8 Single bus arrangement
transmission network arrangement common
in North America. Circuit breakers CB2 and
CB5 are normally open and to take line section LS1 out of service the sequence would
be: close CB2, open disconnector B to loop switch the load current and then open disconnector A to switch out the short line length. Disconnectors at the transformers are
used to switch the unloaded units.
Fig. 1.9 Double bus arrangement
8
Fig. 1.10 H-bus arrangement
Fig. 1.11 Jones-type sub-transmission network arrangement
High voltage air-break disconnectors
9
Section 1
1.4
Perspective
To put the foregoing discussion in perspective, we can state:
1. There is added-value in utilizing the (inherent) current interrupting capability of
disconnectors; in fact, without that capability power systems would be difficult, if not impossible, to operate.
2. Surveys conducted by the IEEE in 1949 and 1962 indicate the range of currents involved
in the past which is not to say that they are applicable according to today’s practices.7, 8
The noted magnetizing currents in particular reflect the high loss transformers that were
once common and now being replaced by larger low loss units with quite different
transient recovery voltage characteristics. The surveys are discussed further in subsection
2.1.
3. Despite the often-cited work of Andrews et al,5 practice appears to be one of trial-anderror, i.e. if it worked once then it can be done again under the same circumstances.
Establishing rules is not a trivial task, particularly given that the interrupting medium is
atmospheric air and the arc is free-burning.
4. Looking to the future, deregulation is a major driver to operate power systems more
effectively with fewer planned outages, even short-term outages. In addition, this is
expected to be achieved with existing equipment and puts more onus on breaking circuits
using disconnectors.
5. An overriding concern is personnel safety. Many disconnectors are manually operated
with varying types of mechanical operators and the switching is subject to ongoing complaint and discussion.
1.5
Objective of the research
The use of air-break disconnectors to interrupt current has a mainly trial-and-error basis. The
reason for this is that little or no research effort has been devoted to current interruption in
atmospheric air and to the behaviour of the associated free burning arcs. Unlimited propagation of the arc is obviously not permissible and its representation is not only as a time varying
electrical element but also as a time varying physical element in space. The cases of interest
are those of interrupting transformer magnetizing current, capacitive currents and loop currents and are addressed in this work.
As shown in the literature review in section 2, no consideration has been given up to now to
the conditions that must be satisfied in order for the current to be interrupted and such consideration is the purpose of this thesis. Consequently, the goal of this research is:
•
10
to investigate and interpret free burning arc behaviour from an engineering
perspective;
High voltage air-break disconnectors
•
to advance the engineering basis for the use of air-break disconnectors for the abovenoted switching duties;
•
to determine arc model parameters for the loop switching case to enable simulation of
this switching duty.
This research work is based on experiences and practices provided by BC Hydro, Bonneville
Power Administration, Kinectrics Inc. (formerly Ontario Hydro Research), Manitoba Hydro
and Puget Sound Energy. Testing associated with the work was performed during the period
1999–2003 at Powertech Laboratories Inc. in British Columbia, Canada and at the high
power laboratories of Eindhoven University of Technology and KEMA in The Netherlands.
Disconnector information was provided by HAPAM B.V. of Bunschoten, The Netherlands,
who also made two disconnectors available for testing purposes at KEMA, and by Pacific Air
Switches Corporation of Forest Grove, Oregon.
11
Section 2
Literature review
2.1
Introduction
Literature relating to current interruption using air-break disconnectors is quite sparse and
comes almost exclusively from North American sources. The reason for this is probably the
historically long practice of using disconnectors for this purpose as compared to the other
parts of the world. Further literature of relevance is that relating to free-burning arcs in air
and sourced from North America, Europe and Japan. The review is divided into four topics:
transformer magnetizing currents, capacitive currents, loop switching and finally free-burning
arcs in air. First, however, surveys of North American utility practices are considered.
Surveys of actual practices were conducted by the AIEE (American Institute of Electrical
Engineers later renamed to IEEE) in 1949,7 the IEEE in 19628 and by IREQ in 1979.9 In the
1949 survey, fifty-nine utilities of the time provided responses on successful interruption of
magnetizing current (Fig. 2.1) and successful or unsuccessful interruption of line charging
current (Fig. 2.2).
Magnetizing current (A)
16
14
12
10
8
6
4
2
0
0
50
100
150
200
250
Transformer voltage (kV)
Fig. 2.1 1949 Survey: magnetizing current interruption using horn-gap
disconnectors7
© 1951 AIEE now IEEE
Respondents were asked to define successful operation and fifty replied as follows:
13 37 -
arc is interrupted without operation of the protective relays or system short circuit
but perhaps after disconnector is fully open.
arc is interrupted before disconnector is fully opened.
Further comment or advice included recommending wider phase spacing and overhead clearances; using only vertical break disconnectors; and using operating mechanisms that allow
the blades to open quickly, i.e. not the gear reduction type of mechanism.
13
Section 2
Charging current (A)
25
20
100% successful
15
90 to 99% successful
10
75 to 80% successful
5
0
0
50
100
150
System voltage (kV)
Fig. 2.2 1949 Survey: interruption of line charging current using horn-gap
disconnectors7
© 1951 AIEE now IEEE
The IEEE 1962 survey was more comprehensive including breaking loop currents and recognizing the use of auxiliary arc quenching devices developed in the 1950s. These devices
ranged from vacuum switches and quick-break (whip-type) devices to a blast device that
actually blasted air, N2 or SF6 gas at the arc. The results of the survey based on responses
from seventy-one utilities are shown in Figs. 2.3, 2.4 and 2.5.F1
Magnetizing current (A)
16
14
12
100% successful
10
8
90 to 99%
successful
600 kVA rule
6
4
Power (600 kVA
rule)
2
0
0
100
200
300
400
500
Transformer voltage class (kV)
Fig. 2.3 1962 Survey: magnetizing current interruption using air break disconnectors8
© 1966 IEEE
F1
14
The survey notes the existence of the “600 kVA rule” where the interrupting limit is given by:
3 × current × voltage across the open switch.
The source of the rule is unknown and its use will be reviewed later.
Literature review
Line charging current (A)
30
25
100% successful
20
90 to 99%
successful
With quick-break
devices
600 kVA rule
15
10
Power (600 kVA
rule)
5
0
0
100
200
300
400
500
Voltage class (kV)
Fig. 2.4 1962 Survey: interruption of line charging currents using air
break disconnectors8
© 1966 IEEE
Switched loop current (A)
450
400
350
100% successful
300
250
90 to 99% successful
200
Less than 90%
successful
150
100
50
0
0
50
100
150
Voltage class (kV)
Fig. 2.5 1962 Survey: opening loop circuits using air disconnectors8
© 1966 IEEE
Respondents were also asked if they had standard guides for use in interrupting current with
air break disconnectors. Fifty-nine utilities stated no guides in use; six utilities used “established rule of thumb operations based on past experience;” and three utilities based their decisions on manufacturer’s data. Only two utilities stated that they used the Andrews et al paper
as a guide.5 This is interesting in that this paper is often cited as the classic, if not definitive,
work in this field.
The 1979 survey was conducted by IREQ (Institute de recherche d’Hydro-Quebec) as part of
a Canadian Electrical Association (CEA) sponsored project.9 The survey was addressed to
Canadian utilities only with a total of twelve providing responses. The survey confirmed the
use of disconnectors by all respondents to break currents and showed an increase in the use of
auxiliary interrupting devices.
15
Section 2
2.2
Transformer magnetizing currents
The first experimental study of interrupting current using disconnectors was that of Andrews,
Janes and Anderson of the Public Service Company of Northern Illinois in the 1940s.5, 10, 11
The study covered interrupting transformer magnetizing and line charging currents and loop
switching. The study was based on two assumptions:
1. Arc length is proportional to voltage.
2. An unconfined arc should always extinguish itself if clearance permits the required
growth to a necessary critical length and the gap is wide enough to prevent restriking.
The assumptions have their basis in the work of Ackermann on transmission line power
arcs.12 Ackermann explained the mechanics of arc extinction: the increasing length of the arc
results in an increasing arc resistance and decreasing current until the arc attains a critical
length at which it extinguishes itself. This explanation and its validity is discussed further in
subsection 2.5.
Andrews et al thus viewed the evolving arc as the controlling element in current interruption
using disconnectors. The study set out then to establish a relationship between current, voltage, arc length and arc reach, where arc reach is “defined as the distance from a point midway
between the arc extremities to the most remote point of the arc at the time of its maximum
length” and is illustrated in Fig. 2.6. The magnetizing current interrupting tests were run in a
Fig. 2.6 Arc reach according to Andrews et al5
laboratory set-up on disconnectors in the range 12 kV to 49 kV and currents up to 35 A. The
immediate observation that can be made is that the voltages are low (all being in the medium
voltage range by today’s standards) and the current high (even for the transformer sizes of the
day and those of later years, refer to Figs. 2.1 and 2.3). Arc lengths of up to 13 m were
recorded and the overall results were presented in the form shown in Fig. 2.7. The quantities
used in deriving Fig. 2.7 were the current at the beginning of the arcing period, the opencircuit voltage across the switch just after arc extinction and the arc length when it was at its
longest point.
16
Literature review
Fig. 2.7 Arc reach per kilovolt as a function of initial arc current5
© 1950 AIEE now IEEE
No statistical analysis was done but a limiting envelope, i.e. limit of probable reach, was
drawn. For the range up to 100 A, the limit of probable reach (LPR) is given by:
LPR = 5.03 UocI
…(2.1)
where LPR is in mm, Uoc is the open circuit voltage across the disconnector in kV rms and I
is the initial current interrupted in A rms. This equation has been widely cited even to the
extent of being used in disconnector manufacturer catalogs to determine magnetizing current
interruption ratings.13, 14 More detailed examination of the manner in which the tests were
conducted leads to doubts as to its applicability. Interrupting magnetizing current, as will be
discussed in section 3, is a repetitive break-and-restrike event with the possibility of inrush
current on restriking and yet no mention is made of this in the reference. Successful magnetizing current interruption relies on at least achieving a contact gap able to withstand the
transient recovery voltage across the disconnector. This suggests that the test set-up used by
Andrews et al resembled interrupting load current rather than magnetizing current; in fact, the
one test trace reproduced in the reference supports this notion (Fig. 8 – quality is too poor for
reproduction here): the initial current of 18 A rms is seen to decay to zero as the arc voltage
rises to the value of the applied voltage over a period of 25 cycles. It is noteworthy that both
Abetti10 and Gerngross11 expressed reservations about the manner in which these tests were
conducted and the results obtained. Extrapolation using Eqn. (2.1) is likely to lead to misleadingly long arc reaches but conservative application, all notwithstanding the influence of
inrush current.
Later reported tests by others on actual transformers confirm that unloaded transformer
switching is other than as represented by Andrews et al. Anderson (not the author of the same
name in reference 5) used a vertical break disconnector to switch an unloaded autotransformer 90/120/150 MVA, 230/115.5 kV with a 13.2 kV tertiary.15 The high-side was wyeconnected, solidly earthed neutral and the tertiary delta-connected. The magnitude of the
magnetizing current was 1.73 A at 220 kV and the disconnector interrupted the current at
17
Section 2
contact gaps as short as 0.6 m. The arc was observed as being “uniformly thin and blue with
little thermal energy” and of lengths in the range 0.6 m to 1.5 m depending on the source
arrangement. The switching was further observed to be a repetitive break-restrike event and
the occurrence of inrush current, albeit of very low magnitude, on restriking was noted. The
influence of the delta-connected tertiary is interesting: with two phases conducting the third
phase is energized by circulating current in the tertiary thus preventing prestriking or restriking in that phase16 (refer also to Annex B). Sample oscillographic traces showed that the
transient recovery voltages at the transformer were of a highly damped nature. Luehring and
Fitzgerald also conducted unloaded transformer switching tests, this time on a 345 kV,
200 MVA autotransformer using a double-break disconnector.17 The disconnector interrupted
6.6 A with a maximum arcing time of 0.75 s at a combined contact gap spacing of about
0.9 m. The arcing time is about one-half that reported for a 330 kV vertical break disconnector in similar circumstances.18 This is understandable because, for the same blade tip speed,
the double-break disconnector will achieve the same total contact gap spacing in half the
time. The result supports the notion that successful current interruption is related to establishing a minimum contact gap spacing rather than a critical arc length.
Foti and Lakas reported results in agreement with Eqn. (2.1) but no details are provided.21
As part of the CEA sponsored study, IREQ performed a number of magnetizing current interruption tests at 230 kV, 315 kV and 735 kV.9 Details of the transformers used are not
provided but the current magnitudes ranged from 0.5 A to 9 A. The arc lengths and reaches
were determined and while some of the normalized reach values (cm/kV) were found to
approximately fit within the range reported by Andrews et al, many points were well below
the trend line (Fig. 2.7) as shown in Fig. 2.8. However, any possible agreement would be
coincidental rather than supportive because the two test series were run under totally different
conditions, the only commonality being in the initial magnitudes of the currents. No mention
is made in the CEA report of blade position at current interruption or inrush current. Overvoltages due to restriking were recorded at up to 1.5 pu.
In the early 1950s auxiliary interrupting devices began to be added to disconnectors. For
unloaded transformer switching these tended to be whip-type spring-loaded devices17, 19 and
gas-blast devices.20, 21, 23, 28 The former devices are still used and will be discussed further in
this context and for capacitive currents. Gas-blast devices functioned by blasting the arc with
compressed gas – usually air or nitrogen but also SF6 gas – to assist in extinguishing the arc.
These devices are no longer in use. An alternative approach is to add insertion resistors.21, 22
These resistors are used at EHV levels and consist on a number of porcelain-clad resistors in
series (each about 500 ohms) with contact rings at the top. As the blade opens, the arc makes
contact with successive rings thus inserting an increasing resistance value. The resistor not
only limits the overvoltage magnitudes due to restriking but also limits the associated inrush
currents to low values. Likewise the resistor limits the inrush current due to pre-striking on
closing the disconnector.
18
Normalized Arc Reach (cm/kV)
Literature review
Fig. 2.8 Comparison between IREQ results (≥ 230 kV) and
those of Andrews et al (≤ 138 kV) for magnetizing current
© CEA 1982
2.3
Capacitive currents
Andrews et al conducted a number of line charging current interruption tests (this is usually
called line dropping) at 132 kV.5, 10 Tests were run in the current range 2.2 to 22 A corresponding to line lengths of 8 to 50 km. The arc lengths and reaches were much longer than
for the magnetizing current case due to the trapped charge effect. For example, at 20 A, an
arc reach of 6.1 m was measured. The normalized arc reaches – reach divided by the applied
voltage – were about four times the trend line shown in Fig. 2.7. The paper recommends that
no more than 7 A of line charging current (about 27 km of line) be interrupted using 132 kV
disconnectors.
The CEA study also reports on capacitive current switching tests.9 The test current levels
were to be limited to 1.2 to 2.3 A and the normalized arc reaches were observed to be about
four times those for magnetizing currents of similar magnitude.
Auxiliary interrupting devices are also used to enhance the interruption of capacitive currents.
Whip-type devices were used by Rankin,23 Toomer24 and Patel et al.25 Rankin successfully
dropped 32 km of 138 kV line (8 A) with such a device. Toomer applied the devices on
19
Section 2
115 kV disconnectors to drop up to 17 km of line. Patel et al ran a number of 115 kV line
dropping tests with devices from different suppliers to determine the influence of parallel
lines and line/configuration on line charging currents. The devices exhibited varying degrees
of arcing for current in the range of 4 to 15 A.
Apart from the influence of weather conditions, there is evidence that the performance of
whip-type devices in this application are subject to system conditions and line configurations26 and even phase spacing because of the coupling between phases.F2
Gas blast devices have also been used for this purpose.20, 21, 23, 27 None of these devices
remain in use today. Lastly, insertion resistors have been at EHV levels mainly for overvoltage limitation on dropping long bus runs21, 28 and for the isolating disconnectors on series
capacitor banks.
2.4
Loop currents
The study of loop current interruption is perhaps the most notable part of the work done by
Andrews et al5 and assisted by Abetti10 and Gerngross11 but with reservation with respect to
test conditions and conclusions drawn. The current range tested was 56 to 312 A with open
circuit voltages across the disconnectors of 1 to 7.6 kV. Arcs as long as 12 m were recorded,
the arcs being allowed to propagate freely to ultimate extinction. For the only instance in the
work, the initial assumption as discussed in subsection 2.2 may have some (but very limited)
validity. Loop current interruption is the commutation or transfer of the current from the one
circuit incorporating the disconnector to a parallel circuit. The rate at which this will occur
depends on the impedances in the loop and the rate at which the arc voltage builds up. The
arc is thus being elongated by the disconnector blade motion and thermal and electromagnetic
effects, all while the current is decreasing due to transfer to the parallel path. The easiest case
of this is current transfer between buses within a station and the more difficult case is that of
current transfer between transmission lines.
For loop currents up to 100 A, Andrews et al5 modified Eqn. (2.1) to read:
LPRl = 5.03 ZIS2 × 10-3 mm
…(2.2)
by setting Uoc = ZIS × 10-3 (kV) where Z is the total loop impedance in ohms and IS is the
initial current in the disconnector in amperes rms. Above 100 A, the limit of probable reach
was set at a constant 5.03 mm/kV (Fig. 2.7) giving:
LPRl = 0.503 ZIS mm
…(2.3)
As for the case of magnetizing current, the manner in which the loop current tests were
performed is subject to question and this is discussed in Annex E.
F2
20
Internal BC Hydro report and private communications with various utilities.
Literature review
Permissible Current (A)
McNulty used Eqns. (2.2) and (2.3) to produce a guide for loop switching.29 The calculation
assumes minimum clearances of 1 m in the system voltage range 23 kV to 69 kV and 1.83 m
at 138 kV. No indication of the allowable reach is given and Fig. 2.9 shows permissible loop
current switching at 138 kV based on McNulty’s calculation.
200
180
160
140
120
100
80
60
40
20
0
0
50
100
150
200
250
300
Loop Impedance (Ohms)
Fig. 2.9 Permissible loop current switching at 138 kV according to McNulty29
Shah and Ward describe the case of a failed loop switching attempt at 34.5 kV between
distribution feeders.30 The authors offer a mathematical analysis of the event that has little or
no merit as shown by Harner.31 The loop involved was complex including two transformers.
The CEA study9 also considered loop switching and found normalized arc reaches well below
the limit of probable reach for this case as proposed by Andrews et al.5 This is shown in Fig.
2.10 below. Little detail is provided on how the tests were carried out but it would appear that
a real loop circuit was simulated rather than opening the test disconnectors against a circuit
recovery voltage. The test disconnectors showed contact melting at current above 200 A and
damage due to contact burning (loss of material) at lower currents. In conclusion, the study
recommends that loop switching be based on the results of Andrews et al but limited to
maximum 100 A and further proposes that the permissible reach be up to 80% of the phaseto-phase clearance. This gives a permissible loop switching current (Ip) for a clearance C in
mm as follows:
LPR l = 5.03 × 10 −3 ZI 2p
= 0.8C
and
I p = 12.6
C
Z
…(2.4)
21
Section 2
Fig. 2.10 Comparison of loop switching test results between IREQ9 and
Andrews et al5
© CEA 1982
Experience with loop switching at 765 kV is presented by Keane, Andrei and Halley.32 The
requirement was to interrupt 600 A at a recovery voltage of at least 200 V which is well
within the IEC requirement of 1600 A and 300 V for that system voltage.1 Tests were
conducted on pantograph and vertical break type disconnectors using a synthetic circuit that
produced the desired recovery voltage. Each disconnector was tested at 600 A and 220 V ten
times with a maximum arcing time of 0.83 s and contact damage in the form of pitting was
found on the contacts of both types (in the IREQ tests loss of contact material was found to
start at around 200 A)9 and contact replacement was recommended after such a number of
loop switching operations. The operating guideline for so-called local loop switching states in
part that the operator is to reclose the switch if current interruption does not successfully
occur by the time the disconnector contacts have parted approximately 0.3 m; and that he
must be aware that small particles of molten metal may fall from the disconnector as a result
of contact burning during the arcing period.
2.5
Free burning arcs in air
The basics of electric arcs has been described by Edels.45 The arc can be initiated by various
means with that relevant to this study being the separation of current carrying contacts. The
conducting gas between the contacts has a very high temperature (> 4000°K) and a high
luminosity. The arc has three distinct regions as shown in Fig. 2.11 together with the voltage
22
Literature review
Fig. 2.11 Arc regions and voltage profile
profile along the arc. Note that the voltage profile shown is not to scale and the relative
values of cathode-fall and anode-fall voltages (also known as the electrode voltage drops)
versus the positive column voltage is dependent on the length of the arc. For very short arcs –
and most arc studies have been related to such – the former voltages will dominate. Edels
states that the electrode voltage drop is in the order of 10 V but Browne gives a range of 20 to
40 V for the electrode drop for arcs in air.46 For long arcs the voltage drop in the positive
column dominates and per unit values given by various sources tend to be close: 12 V/cm for
current greater than 50A;46 13 V/cm for fault currents;12 10 V/cm for currents greater than
100 A;43, 53 and 13.4 V/cm for currents greater than 68 A.44
Edels continues:45 “The conducting arc column acts as a normal electrical conductor in the
presence of a magnetic field, although anomalous electromagnetic effects have been observed
at the cathode. Because of its gaseous nature, the arc is easily influenced by gas flow. However, the spatial stability of arcs is greatly dependent on the nature of the cathode material.
Thus with typically refractory cathodes, e.g. carbon, molybdenum and tungsten, the cathode
temperature is high and the arc is relatively stable. With low-melting-point cathodes, e.g.
copper and mercury, the cathode termination is a highly mobile and concentrated spot which
moves constantly over the cathode surface in an irregular fashion. These visual differences
are the result of different cathode mechanisms and lead to the classification of arcs into two
main types – refractory and non-refractory or cold-cathode arcs. Both arc types, however,
have essentially the same column properties.” The consideration here is without doubt short
arcs. This cathode material effect may well have value for the loop switching cases where the
arcs are expected to be short, i.e. low loop impedances. Actual field observations tend to
support this notion in that tungsten arcing horns performed better (arc duration shorter) than
aluminum arcing horns at currents of several hundred amperes and a loop impedance of
0.5 ohm. The exact mechanism responsible for this performance is believed to be related to
the electrode voltage drop.
23
Section 2
Static electric arcs were first studied in the late 1800s and early 1900s by Ayrton,47
Steinmetz48 and others. Both Ayrton and Steinmetz developed arc equations but that of
Ayrton is by far the better known:
V = A + BL +
C + DL
I
…(2.5)
where V is the arc voltage, I is the arc current, L is the arc length and A, B, C and D are constants. The term A represents the sum of the cathode and anode voltage drops, BL the voltage
drop in the positive column and (C + DL)/I the inverse characteristic of the arc. Nottingham49
later showed the Ayrton equation to be a limited approximation for arcs of constant length up
to 15 mm and rewrote it to the form:
V = A + B/In
…(2.6)
where A and B are constants dependent on arc length and electrode material and n dependent
only on electrode material (n was shown to be directly proportional to the absolute temperature of the boiling point of the anode material). However, for longer arc lengths, Nottingham
showed that a variation of Eqn. (2.5) is applicable:
V = A + BL +
C + DL
In
…(2.7)
As proposed by Browne,46 if L is sufficiently large, then DL >> C and A is negligible
compared to BL, and Eqn. (2.7) can be re-written as:
V = L(B +D/In)
…(2.8)
For high currents, the term containing I becomes small and we can write:
V = LB
…(2.9)
Eqn. (2.9) essentially states that the arc voltage per unit length is constant for high currents.
As already noted above, experimental evidence supports this conclusion.
For sufficiently small currents, the term B in Eqn. (2.8) may be neglected giving:
VI n
=D
L
…(2.10)
Eqn. (2.10) is of a form cited for fault and even lower current arcs in early literature and these
equations are listed in Table 2.1. The equations, where appropriate, are discussed in the following and their relevance or otherwise to this work in subsection 2.6.
24
Literature review
Table 2.1 Summary of published arc equations
(V in volts, I in amperes and L in cm)
Source
Year
Arc equation
1906
VI 0.5
= 51
L
Nottingham49 1923
VI 0.67
=K
L
Ackermann12
1928
VI 0.33
= 98
L
Eaton et al53
1931
Tretjak et al43
VI 0.33
= 56
L
Tests performed by Eaton et al and analyzed by
Tretjak et al
Warrington50
1931
VI 0.4
= 286
L
L in this case is described as the distance between
the electrodes and it is assumed that this is along
the arc path (which the text appears to indicate)
Monseth and
Robinson51
1935
VI 0.5
= 104
L
Cited by Andrews et al5
Abetti10
1948
VI 0.4
= 43
L
Loop switching case where V is the open circuit
voltage and I the initial current
Gerngross11
1949
VI 0.36
= 35
L
Loop switching case where V is the open circuit
voltage and I is the initial current
Steinmetz48
46
K dependent on contact material
Browne cited Nottingham’s paper49 but appears to
VI
= constant prefer the Ayrton equation current exponent for
L
the longer arc case
Browne
1955
Maikopar59
1960
Unknown
-
60
V
= 12 +
IP
L
1967
VI0.55
= 45.5
L
Rieder60
Comment
VI p
L
0.4
= 75
Ip is the peak current
Provided by Dr. K. Suzuki of Toshiba
A number of investigators have studied long arcs in air principally in connection with fault
current arcs or secondary arcs. Ackermann, who studied fault current arcs in range 120 A to
11,000 A, derived the equation shown in Table 2.1 by considering a balance between power
input to the arc and power loss from the arc.12 Ackermann’s theory is that as long as the
equation is satisfied, the arc is stable; however, if the arc voltage drops below the value satisfying the equation – as at a critical length – then it becomes unstable and rapidly decays. This
theory does not fit evolving arcs and really contradicts the fact that the arc voltage will
increase as the current decays. The arc evolves due to increasing power input and then
25
Section 2
collapses when the power input is removed.42, 52 Apart from this discussion, Ackermann notes
that the evolution of the arc is limited by partial arc collapses along its length.
Eaton, Peck and Dunham53 studied power arcs in the range 8 to 800 A peak, and the results in
the range of 400 A peak below are of interest with respect to loop switching. Arc voltages
and lengths were measured and the determined arc voltage gradient versus arc current points
are plotted in Fig. 2.12.
Voltage gradient (V/cm)
40
35
30
25
20
15
C
10
5
0
0
100
200
300
400
500
Arc current (A peak)
Fig. 2.12 Arc voltage gradient versus arc current according to Eaton et al53 and
Tretjak et al43 (Lower curve: mean gradient; upper curve: critical gradient)
© 1931 AIEE now IEEE and CIGRE
Tretjak et al investigated AC arcs in air in a high voltage laboratory and then combined the
results with those of Eaton et al to derive the mean equation given in Table 2.1 and shown in
Fig. 2.12. The equation derived for the upper arc gradient limit shown in Fig. 2.12, which
represents the condition of the critical arc length (LC), is:
VI 0.33
= 84
LC
…(2.11)
Warrington50 conducted fault test with currents in the same range as Eaton et al.53 The
purpose of the tests was to examine the influence of arc resistance on distance relays as was
part of the reason for both Ackermann and Eaton et al performing their tests. The constant in
Warrington’s equation (Table 2.1) in particular is five times higher than that of Tretjak et al.
Warrington’s test points and equation are plotted in Fig. 2.13. Both Tretjak et al and
Warrington (erroneously) treat critical length in the same manner as Ackermann, i.e. if the
gradient drops below that given by equation, the arc decays and extinction follows.
26
Voltage gradient (V/cm)
Literature review
100
90
80
70
60
50
40
30
20
10
0
0
200
400
600
800
1000
1200
Arc current (A)
Fig. 2.13 Arc voltage gradient versus arc current according to Warrington50
© 1931 Electrical World
Gross studied free burning arcs in the range 200 A to 400 A.52 Arc lengths were not determined but recordings of the current and arc voltage were made. Fig. 2.14 shows a plot of
these two quantities over the arcing time of 72 cycles at 50 Hz. The plot shows a gradual
250
70
Arc voltage (kV)
Arc current (A)
200
50
40
150
30
100
20
Arc current (A)
Arc voltage (kV)
60
50
10
0
0
1
5
9
13
17 21
25
29
33
37 41
45
49
53 57
61
65
69
73
Time (cycles at 50 Hz)
Fig. 2.14 Arc current and voltage according to Gross52
© 1941 Schweizer Archiv
decrease in current as the arc grows in length and the arc voltage increases. After about
50 cycles, the decrease becomes more rapid as does the increase in arc voltage until extinction occurs. Partial arc collapse is evident at 22, 40, 52 and 65 cycles. The arc resistance and
arc power are shown in Fig. 2.15. The arc power shows a steady increase with the increasing
arc resistance to a point where rapid decay occurs due to a sudden collapse of the power
input. The arc thus continues to evolve as long as the power input is increasing and extinguishes when the input is suddenly removed. This behaviour was also observed by Anjo as
discussed later and in the loop switching tests described in section 5. The U-I characteristic is
shown in Fig. 2.16. The time progression along the characteristic is from right to left. The
loops at 80, 140 and 175 A are partial arc collapse incidents.
27
Section 2
4000
1800
Arc resistance
1600
Arc power
3500
3000
1400
1200
2500
1000
2000
800
1500
600
1000
400
Arc power (kW)
Arc resistance (ohms)
2000
500
200
0
0
1
5
9
13 17 21 25 29 33 37 41 45 49 53 57 61 65 69 73
Time (cycles at 50 Hz)
Fig. 2.15 Arc resistance and arc power according to Gross52
© 1941 Schweizer Archiv
70
Arc voltage (kV)
60
50
40
30
20
10
0
0
50
100
150
200
250
Arc current (A)
Fig. 2.16 Arc U-I characteristic according to Gross52
© 1941 Schweizer Archiv
Strom studied arcs in the range of 68 A to 21.75 kA and lengths of about 3 mm to 1.2 m.44 In
brief summary the study found average arc voltage gradients of 12.2 V/cm below 5000 A and
14.9 V/cm above 5000 A.
Secondary arcs – those following the primary fault current arcs in single phase tripping and
reclosing schemes – are of interest particularly in terms of their extinction mechanism. The
supply voltage that drives the secondary arc current is by electromagnetic and electrostatic
induction from the sound phases and from parallel lines in the case of multi-circuit towers.
This voltage cannot be considered a hard source but tends to produce a constant current even
as the arc evolves in length.51, 59 Numerous system tests have shown that secondary arc
currents are low in magnitude, generally less than 100 A,54-57 and can be symmetrical or
asymmetrical.58 The extinction times of secondary arcs are usually less than one second,
more often low enough to permit reclosing times of 0.33 seconds. This is demonstrated by
Anjo who studied secondary arcs following primary fault currents of 8 kA.42 A plot of the arc
28
Literature review
voltage and current is shown in Fig. 2.17. Once the influence of the primary current has disappeared, the current is approximately 30 A up to the point of extinction, which takes place
over a period of one cycle. Extinction occurs when the arc voltage reaches such a value that
70
Arc voltage
Arc voltage (kV)
60
Arc current
60
50
50
40
40
30
30
20
20
10
10
0
Arc current (A)
70
0
1 2 3 4 5 6 7 8 9 10 1112 13 14 15 16 17 1819 20 21
Time (Cycles at 50 Hz)
Fig. 2.17 Secondary arc voltage and current42
© 1968 Electrical Engineering in Japan
the supply can no longer provide the increasing power necessary to sustain the arc. The arc
power associated with the arc voltage and current of Fig. 2.17 is shown in Fig. 2.18. The
power increases at a steady rate and takes a sudden jump just prior to extinction. The U-I
2500
Arc power (kW)
2000
1500
1000
500
0
1 2 3 4 5 6
7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22
Time (cycles at 50 Hz)
Fig. 2.18 Arc power as derived from arc voltage and current in Fig. 2.1742
© 1968 Electrical Engineering in Japan
characteristic is shown in Fig. 2.19 and it is obvious that it differs greatly from the U-I characteristic of the free burning arcs of Gross (Fig. 2.16). Anjo measured arc lengths up to 9.3 m
with the gradient rising to 6.8 kV/m at extinction.
Abetti10 and Gerngross11 studied the arc films taken by Andrews et al5 and determined the
equations given in Table 2.1. Strangely neither work is cited by Andrews et al.
29
Section 2
70
Arc voltage (kV)
60
50
40
30
20
10
0
0
20
40
60
80
Arc current (A)
Fig. 2.19 Secondary arc U-I characteristic42
© 1968 Electrical Engineering in Japan
2.6
Conclusions
In general, the study of disconnector’s current interrupting capability has not attracted a
research interest appropriate to the lack of available knowledge on the subject. Apart from
Neumann’s62 comprehensive study of air-break and GIS disconnector recovery voltages for
the capacitive current and bus-transfer cases, no major research effort is evident.
The literature provides only a limited insight into the mechanism of current interruption in
air. The principal work of the past is clearly that of Andrews et al in the 1940s and those who
studied the subject later tended to adopt the arc reach approach of that early work. No consideration is thus given to the conditions that must be satisfied in order for the current to be
interrupted and such consideration is the purpose of this thesis.
Past work on free-burning arcs in air is of interest for the case of loop switching. It is fundamental that the arc will propagate as long as its power input is increasing. For loop switching
Arc voltage gradient (V/cm)
100
90
Steinmetz
80
Ackermann
70
Tretjak
60
Warrington
Monseth & Robinson
50
Abetti
40
Gerngross
30
Maikopar
20
Suzuki
10
Rieder
0
0
100
200
300
400
500
Arc current (A)
Fig. 2.20 Plot of arc equations from Table 2.1 for range 0 to 400 A
30
Literature review
Arc voltage gradient (V/cm)
60
Steinmetz
Ackermann
50
Tretjak
40
Monseth & Robinson
Abetti
30
Gerngross
20
Maikopar
Suzuki
10
Rieder
0
0
20
40
60
80
100
Arc current (A)
Fig. 2.21 Plot of arc equations from Table 2.1 for range 0 to 100 A
this means that the arc voltage must increase faster than the current decreases due to commutation to the loop circuit. However, a point is inevitably reached where the rate of change of
the power input goes to zero and the arc will collapse. The arc voltage is a function of the arc
current and the arc length as described by the equations of Table 2.1. These equations are
plotted in Fig. 2.20 and show a reasonable consistency except for that of Warrington. The
plots show an almost constant voltage gradient for currents greater than 100 A. The range is
anywhere from about 3 V/cm to 15 V/cm and it is perhaps no coincidence that power system
study engineers often cite a fault arc voltage gradient of 10 V/cm as a “rule-of-thumb.” For
currents below 100 A, the gradient increases with decrease in current and this range is
expanded in Fig. 2.21 with Warrington’s equation excluded. The conclusion therefore is that
the range below 100 A is more favourable for loop switching because current commutation
promotes a higher arc voltage and in turn further current commutation and ultimate arc
collapse.
31
Section 3
Interrupting transformer magnetizing current
3.1
Introduction
In the past transformer magnetizing currents were in a range up to 15 A (refer to Figs. 2.1 and
2.3). The value of energy has changed this and the low loss transformers of today have magnetizing currents of less than 2 A, often less than 1 A at 100% excitation voltage. The current
is non-sinusoidal with a high 3rd harmonic content. As such current zeros tend to occur prior
to the crest of the applied voltage. The current is usually expressed in terms of an equivalent
RMS value derived during the core loss measurement at the manufacturer’s plant.
Particularly in North America, this switching duty has been treated using the arc reach
approach of Andrews et al.4, 5, 9 Reservation has already been expressed with respect to this
approach and it is important to provide a broader perspective in this regard. Taking the arc
reach approach at face value, it is possible to calculate limiting magnetizing current values
based on arc reach. One such approach gives the following results (Table 3.1):6
Table 3.1 Allowable magnetizing current interrupting levels
System voltage
kV rms
15
27.5
72.5
145
253
550
Magnetizing current
A
4.1
2.9
2.7
2.3
2.1
1.4
The calculation had a conservative basis, the results appeared in practice to be reasonable and
it was used as a basis for an IEEE guide.4 However, from a pragmatic engineering application
perspective, the approach does not take into account blade position at the time of interruption.
Observation at many actual unloaded transformer switching operations has shown that the arc
reach is usually insignificant, in fact the arc showed very little tendency to rise often running
down the opening blade. Furthermore, the current interruption is a repetitive break-restrike
event (refer to Fig. 3.7(a)) that produces restriking overvoltages imposed on the transformer
insulation and also results in power system inrush currents. To describe the switching duty in
engineering terms is to state: to interrupt transformer magnetizing currents up to 2 A with a
recovery voltage equal to the difference between the system applied voltage and the transformer side recovery voltage before the disconnector blade reaches a limiting position. In
dealing with this particular duty it is also necessary to consider the influence of restriking on
the opening operation and on the transformer.
33
Section 3
3.2
Analysis
In Annex A it is shown that, after the interruption of magnetizing current, the transient recovery voltage takes the form of a highly damped oscillation. The degree of damping and the
frequency of the oscillation is dependent on the level of excitation. However, the frequencies
are low – generally less than 300 Hz – and it is reasonable to consider the TRV as having a
(1-cosine) waveshape. A worst case scenario would thus be when the peak of the transformer
side underswing coincides with the peak of the system applied voltage of opposite polarity.
As shown in Annex A, this differential voltage across the disconnector can be taken as
1.3 pu. Vertical break disconnectors are considered in the analysis that follows, which analysis can be extended to other switch types by considering contact gap spacing.
The condition to be satisfied for the current to be interrupted is:
Ugap ≥ 1.3
2
UL
3
…(3.1)
where Ugap is the withstand voltage in kV peak and UL the system voltage in kV rms.
The minimum contact gap required to interrupt magnetizing current can be calculated with
the following assumptions being made:
1. The thermal energy of the arc is not significant. This means that the arc shows little or no
tendency to expand and restrikes do not necessarily occur in the previous arc channel.
2. The opening contact gap is viewed as a rod-rod type gap. This assumption will give a
conservative result because the electrodes in question are of a round nature, i.e. a sphere
on the blade end and corona rings on the jaw assembly.
Rod-rod gap power frequency sparkover values can be found in high-voltage testing standards.33 The 60 Hz values (equally applicable at 50 Hz) are shown in Fig. 3.1 and are mean
values with an error ±8%. Using Fig. 3.1 and known disconnector blade lengths, the required
contact gap spacings are derived (Table 3.2).
Table 3.2 Minimum contact gap spacings for magnetizing current interruption
System
voltage
kV
72.5
145
245
362
420
550
34
Disconnector
blade length
mm
950
1700
2600
3500
4000
4800
Recovery
voltage
kV peak
77
154
260
384
445
583
Mean
150
300
500
740
860
1130
Contact gap
mm
Mean + 8% Mean - 8%
162
138
324
276
540
460
800
680
930
790
1220
1040
Interrupting transformer magnetizing current
60 Hz sparkover voltage
(kV peak)
Rod-rod 60 Hz sparkover voltages
1200
1000
800
600
400
200
0
0
50
100
150
200
250
Gap spacing (cm)
Fig. 3.1 Rod-rod 60 Hz sparkover peak voltages34
The results in Table 3.2 are shown schematically in Fig. 3.2. With some adjustment for
disconnector geometry, it can be seen that the current should be interrupted at about a 15°
blade angle. If a rule is set that the disconnector must interrupt the current before the blade
reaches a 45° angle, then there would appear to be ample margin to achieve this. However, at
higher current values (but not considering inrush currents for the moment) the thermal energy
of the arc will become significant and thermal effects of the arc will also influence the point
of current interruption. We suggest, therefore, that the interrupting capability of the disconnector is the maximum current at which the length of the interrupting path as determined by
geometry of the disconnector is the major controlling element in the interruption process. The
same can be suggested for capacitive currents of the same order, but not for loop currents
where the process is one of current commutation and not current interruption. This notion has
equal relevance for capacitive currents and will be discussed in detail in that context.
Fig. 3.2
35
Section 3
3.3
Restriking and its consequences
Because magnetizing current interruption is a repetitive break-and-restrike process, it will
result in the generation of restriking overvoltages and the occurrence of inrush current. A
case study of using an air break disconnector to switch an unloaded EHV transformer bank is
described in Annex B and well illustrates this point. The issue to be addressed is the impact
of the switching event on the transformer and the remedial measures that can be taken to
eliminate or mitigate any negative consequences of restriking.
While transformers are commonly protected by metal oxide surge arresters, dielectric failure
of some transformer types during unloaded switching have been reported.34, 35 The transformers were delta-connected on the high side and the failures were attributed to a ferroresonance
effect. The failures are probably due to the combination of the transformer type and restriking
in the disconnector. In fact, industry guidelines recommend against using disconnectors to
switch unloaded delta-connected transformers.4 Such failures are the exception and this type
of switching is widely practiced in North America often with the addition of a quick-break
whip type device. With proper design and application (as discussed in the next subsection),
these devices provide an essentially restrike-free magnetizing current interruption.19 Mitigation of prestriking transients can only be achieved at the expense of a faster closing operation
which may be mechanically undesirable.
3.4
Inrush currents
Apart from descriptions in textbooks,36 inrush current is generally viewed in terms of its
influence on other systems such as protection rather than in itself. A literature search revealed
only two references dating back to the early 1950s.37, 38 Inrush current is the initial rush of
magnetizing current on energizing a transformer. This energization can be either a single
event such as the closing or opening of a circuit breaker or a train of events with multiple
prestriking or restriking during closing or opening of a transformer disconnector. When a
transformer is switched out a residual flux is left in the core. On subsequent energizing of the
transformer, the inrush current is determined by the magnitude and direction of the residual
flux and the closing angle on the applied voltage. The inrush current is greatest when the
transformer is energized at voltage zero following which the polarity of the voltage is such
that the flux increases in the direction of the residual flux. The core is driven into saturation
and the transformer draws a high current from the supply network. As the closing angle on
the applied voltage moves away from the zero crossing, the magnitude of the inrush current
becomes less and less. To avoid inrush current altogether, it would be necessary to know the
status of the residual flux and to close at the voltage crest on appropriate polarity.41
Visual observation of arcs during unloaded transformer switching show the thin blue arc
associated with the steady state magnetizing current and following some prestrikes or
restrikes, short-lived bursts of inrush current. Fig. B4 in Annex B shows a typical trace of
inrush current events during the described test. The inrush current is limited in both magnitude and duration, a fact that has been observed by others15, 19, 22 and in other field tests (Fig.
3.3). The highest recorded value of inrush current was 1000 A as compared to a theoretically
36
Interrupting transformer magnetizing current
200
kV
000
-200
0.1
0.2
SECONDS
0.3
0.4
0.1
0.2
SECONDS
0.3
0.4
2
KAMPS
0
-2
Fig. 3.3 Switching unloaded 230 kV transformer with an air break disconnector
possible value of 3000 A. The reason for the limitation is that restriking will tend to occur
close to the peak of the source voltage. A circuit breaker, with its fast closing speed, has the
possibility of closing or prestriking close to a voltage zero crossing and producing a high
inrush current (Fig. 3.4). Magnitude and duration of inrush current are evidently related and
the question now is to consider the influence on the arc and the arcing time.
There are no direct measurements of arc voltage for this case and interpolation from other
sources is necessary. The disconnector will not attempt to interrupt inrush current and the
first conclusion is that the inrush current will prolong the arcing time by its duration. The
inrush current arc and associated power input will leave a hot arc channel with a high ion
density and the subsequent steady state current will follow this channel. The question to be
addressed is: what is the thermal time constant of the hot channel? There is evidence to
suggest that this time is very short indeed and is discussed below.
The duration of power arcs in air has been studied mainly in the context of high-speed threepole and single-pole reclosing following fault current interruption. Early studies of highspeed three-pole reclosing by Sporn and Prince showed that de-energized times of 12 cycles
were adequate to prevent arc reignition on reclosing following fault currents of less than
1 kA.40 Boisseau et al conducted high power laboratory tests on faults initiated across 69, 138
and 230 kV suspension insulator strings and found that for fault currents up to 10 kA, successful three-pole reclosing could be achieved within 6 to 10 cycles.41 This time was found to
37
38
Courtesy of Bonneville Power Administration
Fig. 3.4 Energizing a 230 kV transformer using an SF6 type circuit breaker
Section 3
Interrupting transformer magnetizing current
be independent of a fault current duration in the range 1 to 14 cycles. In the case of single
pole reclosing, the primary power arc is followed by a secondary arc thus prolonging the arc
duration. Anjo et al studied this case for a primary power arc of 8 kA rms followed by secondary arcs in the range of 12 A to 30 A.42 The secondary arcs exhibit a minimum extinction
time of about 12 cycles at 50 Hz. The contributors to this time are the past power input history due to the power arc and the continuing and increasing power input due to the secondary
current. Therefore it can be concluded that the former’s contribution timewise must be less
than 12 cycles (0.24 s). In fact, the secondary arc current as derived from Figs. 3 and 10 of
reference 42 indicates that the high current arc influence time is as short as 5 cycles (100 ms).
This is shown in Fig. 3.5 where the secondary arc current of about 30 A is reached in the
noted time.
70
Arc current (A)
60
50
40
30
20
10
0
0
5
10
15
20
25
Time (cycles at 50 Hz)
Fig. 3.5 Arc current for experiment no. 512 in reference 39
© 1968 Electrical Engineering in Japan
The conclusion that can be drawn from these considerations is that the thermal constants
associated with inrush currents in air – the magnitudes of which are certainly less than 8 kA
and power inputs much less than for sinusoidal waveshapes of the same duration – is in the
order of 100 ms or less.
Applying this logic to inrush current followed by steady state magnetizing current as
described in Annex B, we can argue as follows. The arc length at current interruption was
about 1 m. Based on the work of Tretjak43 and Strom,44 the arc voltage can be taken as
10 V/cm for the current levels in question (arc characteristics are discussed in detail in
connection with loop switching). To simplify the calculation, we assume conservatively that
the inrush current is sinusoidal at 1000 A peak and duration 0.3 seconds. The energy injected
into the arc would then be 212 kJ which is considerably less than the energy of 3.2 MJ
injected into the primary current in Anjo et al’s experiments. The conclusion is that the time
constant of the inrush current arc is less than one cycle at 60 Hz and therefore the inrush
current only contributes to the arcing time by its actual duration.
39
Section 3
Analysis of the videos taken during the BC Hydro test described in Annex B provides compelling support for the inrush current influence scenario discussed above. With reference to
Fig. B4, the sequence of events between the times 3 and 4 seconds is as follows:
1.
2.
3.
4.
5.
6.
7.
8.
Steady state magnetizing current
First major inrush current
Steady state magnetizing current
Minor inrush current
Steady state magnetizing current
Second major inrush current
Steady state magnetizing current
Current interruption
Images from the video records illustrating the sequence of events are shown in Fig. 3.6. The
consistency of the steady state magnetizing current is clear evidence of the lack of influence
of the inrush current on arc propagation beyond extending its duration.
3.5
Auxiliary interrupting devices
The application of auxiliary interrupting devices on disconnectors is common in North
America. To interrupt magnetizing current, the only device in use is the quick-break whip
type device available only at system voltages up to 245 kV. Such a device is shown in Fig.
1.6 and functions as follows. As the blade opens, the whip is restrained by the whip capture
attachment which also functions as the fixed arcing horn (Fig. 1.7). When the blade reaches a
predetermined position, the whip releases achieving a high tip velocity in the order of 0.5 to
0.6 m/cycle at 60 Hz. The whip has a damping mechanism to ensure that, once it passes the
blade position, it does not swing back into the open gap. Two rules apply:
1. At the time that the whip releases the blade position must be such that the contact gaps
shown in Fig. 3.2 are equalled or (preferably) exceeded.
2. The current must be interrupted before the whip reaches the blade position, otherwise its
effect is negated and the arc will transfer to the blade.
By way of example, for a system voltage of 245 kV, the recovery voltage is 260 kV peak
(Table 3.2). From Fig. 3.1 the slope of the rod-gap withstand characteristic is 0.5 kV
peak/mm and a tip velocity of 0.5 m/cycle corresponds to 30 mm/ms, which together gives a
rate of increase of withstand voltage of 15 kV peak/ms. The whip will achieve the required
gap spacing in 17.3 ms. Because the current is interrupted at zero crossings, it is desirable to
have a number of zero crossings while the whip is in motion. This can be arranged by setting
the whip release point to where the blade position well exceeds the minimum contact gap
position.
40
(b) Initiation of first major inrush current event
(c) First major inrush current event
(d) End of first major inrush event
Fig. 3.6 Video images: steady state magnetizing and inrush current sequence (continued)
Courtesy of BC Hydro
Interrupting transformer magnetizing current
41
(a) Steady state magnetizing current
Section 3
42
(e) Steady state magnetizing current
(f) Minor inrush current event
(g) End of minor inrush current event
(h) Steady state magnetizing current
Fig. 3.6 Video images: steady state magnetizing and inrush current sequence (continued)
Courtesy of BC Hydro
(j) End of second major inrush current event
(k) Steady state magnetizing current
(l) Current interruption
43
Fig. 3.6 Video images: steady state magnetizing and inrush current sequence
Courtesy of BC Hydro
Interrupting transformer magnetizing current
(i) Second major inrush current event
Section 3
Observation of a well-designed whip type device in operation shows that it is possible to
achieve a virtually restrike-free magnetizing current interruption.19 At most only an arc of a
few mm in length is visible just as the whip tip releases from the capture attachment. Fig.
3.7(a) and (b) shows the interruption of magnetizing current for a 230 kV using a vertical
break disconnector before and after the addition of a whip type device, respectively. Repeated
restriking is evident in Fig. 3.7(a), but only very minor restriking is seen in Fig. 3.7(b) at the
instant of current interruption, i.e. whip release.
kV
(a) Without whip type device
kV
(b) With whip type device
Fig. 3.7 Switching unloaded 230 kV transformer with vertical break disconnector
without and with whip type device. Primary voltage is measured at the transformer
230 kV bushing.
44
Interrupting transformer magnetizing current
3.6
Conclusions
Conclusions with respect to the interruption of transformer magnetizing current with highvoltage air break disconnectors may be summarized as follows:
•
Steady state magnetizing current interruption is dependent on achieving a minimum
contact gap spacing provided that the current is low enough to avoid thermal effects.
•
The influence of inrush current resulting from restriking is only in terms of prolonging the arc time by its duration plus a few cycles.
•
Arc reach is generally not an issue because the condition for arc propagation – an
increasing power input – is never met. Even inrush current and its power input contribution does not promote arc propagation.
•
To minimize the effect of switching on the transformer, effort should be made to limit
restriking and prestriking transients. For the former in the voltage range up to 230 kV,
auxiliary devices can provide an essentially restrike-free current interruption. Prestriking can only be limited by increasing the closing speed or by the addition of
closing resistors.
•
The results of Andrews et al have little (if any) relevance to unloaded high voltage
transformer switching.
45
Section 4
Interrupting capacitive currents
4.1
Introduction
The capacitive currents in question are those associated with bus lengths with or without connected instrument transformers and short lines. The typical range of these currents is shown
in Table 4.1. In practice, the allowable values of the capacitive currents to be interrupted do
not exceed 5 A at any system voltage (Fig. 4.1).
Table 4.1 Capacitive current range
Equipment
type
CT*
CVT*
Busbars*/m
Lines/km
Capacitive current (A) at
145 kV
245 kV
50 Hz
60 Hz
50 Hz
60 Hz
72.5 kV
50 Hz
60 Hz
0.025
1.7 × 10-4
0.15
0.03
2 × 10-4
0.18
0.033
0.133
0.32 × 10-3
0.24
0.04
0.16
0.39 × 10-3
0.29
0.033
0.22
0.54 × 10-3
0.42
0.04
0.27
0.65 × 10-3
0.5
550 kV
50 Hz
60 Hz
0.1
0.49
1.1 × 10-3
1.33
0.12
0.59
1.3 × 10-3
1.6
Current (A)
* For outdoor substations.
6
Utility 1
5
Utility 2
4
Utility 3
Utility 4
3
Manufacturer 1
2
IEEE Std (ref.4)
1
600 kVA rule (ref. 9)
0
Manufacturer 2
0
200
400
600
System Voltage (kV)
Fig. 4.1 Allowable capacitive current interruption from various sources
As for magnetizing current, the interruption process is a repetitive break-restrike event but
often with a longer arcing duration due to the higher recovery voltage. Overvoltages
produced can be a concern particularly for oil-filled paper insulated condenser type instrument transformers.
4.2
Analysis
This case can be analyzed in the same manner as for magnetizing currents and using the
equivalent assumptions. To allow for some margin, the system voltage is taken to be at 1.1 pu
and thus a recovery voltage peak of 2.2 pu is applicable. The condition to be satisfied for the
current to be interrupted is:
47
Section 4
Ugap ≥ 2.2
2
UL
3
…(4.1)
where Ugap is the gap withstand voltage in kV peak and UL the system voltage in kV rms.
The results of the calculation are shown in Table 4.2 and schematically in Fig. 4.2. Current
interruption by a 45° blade angle would appear to be certain.
Table 4.2 Minimum contract gap spacings for capacitive current interruption
System
voltage
kV
72.5
145
245
362
420
550
Disconnector
blade length
mm
Recovery
voltage
kV peak
950
1700
2600
3500
4000
4800
130
260
440
650
755
988
Mean
252
504
854
1260
1465
1917
Contact gap
mm
Mean + 8% Mean - 8%
272
232
544
464
922
786
1360
1159
1582
1348
2070
1765
Fig. 4.2 Minimum blade angles for capacitive current interruption for vertical
break disconnector
The issues to be considered in this section are:
•
•
•
48
the level of capacitive current that can be interrupted preferably before the disconnector blade reaches the 45° angle position;
the process of current interruption;
external dependencies.
Interrupting capacitive currents
4.3
Auxiliary interrupting devices
Whip-type devices applied to interrupt capacitive currents are discussed in Annex D.
4.4
Field experience
Records of vertical break disconnector successful and failed current interruption attempts are
plotted in Fig. 4.3. In this context, successful means that the current was interrupted before
the blade reached the fully open position and failed that the arc persisted after the blade
Utility 1
7
Utility 2
Utility 3
6
Utility 4
Current (A)
5
Manufacturer 1
IEEE Std (ref.4)
4
600 kVA rule (ref. 9)
3
Manufacturer 2
Disconnector
successes
2
Disconnector failures
1
Linear (Disconnector
successes)
Linear (Disconnector
failures)
0
0
100
200
300
400
500
600
System Voltage (kV)
Fig. 4.3 Vertical break disconnector capacitive current interruption successes and
failures
reached the fully open position. In Fig. 4.3, the points from Fig. 4.1 are also shown but are
now masked (appear in grey).
A linear regression of the success points falls approximately along the upper limit of Fig. 4.1
points. The regression also shows a reasonable agreement with a linear regression of the
600 kVA rule points (Fig. 4.4). The equations for both are:
Isuccess = -0.0064U + 3.7
…(4.2)
I600 kVA = -0.008U + 4.1
…(4.3)
and
where I is the current in A and U the system voltage in kV.
49
Section 4
600 kVA rule
7
6
Disconnector
successes
Current (A)
5
Disconnector
failures
4
3
Linear (600 kVA
rule)
2
Linear
(Disconnector
successes)
1
0
0
200
400
System voltage (kV)
600
Linear
(Disconnector
failures)
Fig. 4.4 Linear regressions of vertical break disconnector current
interruption successes and failures and the 600 kVA rule
A visual perspective of capacitive current interruption is shown in Fig. 4.5 − a 115 kV
disconnector switching out a 10 km long line charging current 2.4 A. The arc is irregular,
exhibits a certain thermal effect rising upwards and is prolonged by energy injection due to
restriking first in the far phase, then the centre phase and finally the near phase. Interruption
occurs just prior to the fully open position. Reach is not significant and is better described by
the approach of Barrett and GreenF3 than that of Andrews et al. Barrett and Green viewed the
arc as being contained within a cylinder whose axis is a line drawn between arc end-points
and whose radius is a line perpendicular to the axis to the extreme point of the arc. Fig. D2 in
Annex D shows a much more intense arc with greater thermal effect for an increase in current
of only about 1 A to 3.3 A. Clearly at some current level below 2.4 A, the thermal effect
starts to be significant and this notion and the role of restrikes is explored in the following
subsections.
4.5
Video record review
Observation of capacitive current arcs in air shows that they can be classified into two types
as follows:
•
F3
50
The first type occurs for current magnitudes below a certain level. At this level, the
arc does not thermally support a sustainable arc channel and each restrike establishes
a new path between the electrodes. No exact study has been made of applicable
current levels but experience indicates that this arc type occurs at currents below 1 A.
J.S. Barrett and M.A. Green, “230 kV Grounding Devices − Inductive and Capacitive Arcs.” Ontario Hydro
Technologies (now Kinectrics Inc.), Report No. A-G-94-19-H, March 1994.
Courtesy of Bonneville Power Administration
51
Interrupting capacitive currents
Fig. 4.5 115 kV vertical break disconnector interrupting line charging current of 2.4 A (continued)
Section 4
52
Restrike far phase
Restrike centre phase
Restrike near phase: note arc colour change on
centre phase indicative of arc collapse
Current interruption: note hot gas remnants near
blade tip on centre phase
Fig. 4.5 115 kV vertical break disconnector interrupting line charging current of 2.4 A
Courtesy of Bonneville Power Administration
Interrupting capacitive currents
•
The second type occurs for current magnitudes of about 1 A or more. In this case, the
arc thermally supports a sustainable arc channel and each restrike follows the established arc channel. The current continues to break-and-restrike until such time as the
channel becomes too long and/or cold for a restrike to occur or that the energy
injected by a restrike is not sufficient to force another loop of power frequency
current.
For the first type of arc, current interruption will occur around the minimum blade angles
shown in Fig. 4.2. For the second type the effect of restriking is to delay current interruption
beyond the above-noted blade angles. Additionally, there is evidence to suggest that the
arcing time is determined, not only by the status of the previous arc channel and its capability
to support restriking, but also by the relative source (CS) and load side capacitances (CL).F4 In
the test at 2 A described in the KEMA report, the longest arcing times were found to occur
when CS/CL < 1 and the disconnector failed to interrupt the current when CS/CL = 0.1.
Knobloch66, 77 and Neumann62 noted a similar dependency with respect to restriking overvoltages. The influence of CS/CL is discussed in detail in subsection 4.6 below.
Fig. 4.6 shows an example of the first type of arc. The current level is approximately 0.5 A
and the arc can be seen to literally dance around the electrodes with no fixed pattern whatsoever.
Fig. 4.7 illustrates the second type of arc. This example is of a 115 kV centre break disconnector interrupting a line charging current of 1.8 A. The disconnector is equipped with a
whip-type device but the whip is too short and releases early transferring the current to the
main contacts. When a longer whip was substituted, the device cleared without a visible arc.
The arcing time was 1.6 seconds. The arcs appear to evolve to a certain length and are
sustained by a balance − perhaps tenuous − between the power input and the power loss and
further injections of energy due to restrikes. The restriking energy does not promote propagation but, similar to inrush current for the magnetizing current case, prolongs the arcing time
by at least its duration of injection plus a number of cycles thereafter.
In another line dropping test at 115 kV and 4.6 A, a similar type of disconnector cleared in
2.3 seconds but only after reaching the fully open position. This disconnector was also
equipped with a whip-type device. However, the adjustment of the device was not correct and
in this case the arc gradually ran down the whip as can be seen in Fig. 4.8. The weather at the
time was inclement with rain and wind. The arc propagates horizontally and whether this is
related to the moving arc on the whip or the wind is not obvious. The arc again exhibits a
limiting length and in fact propagates several times to this length in a series of restrikes and
intermittent partial arc collapses. This length is exceeded only in the final image Fig. 4.8(p),
at which length the power loss is greater than the power input and total collapse is inevitable
provided that no partial arc collapse or restrike occurs in the meantime to sustain it.
F4
KEMA Report of Performance No. 237-86: Development of a test circuit for small capacitive current
interruption by disconnectors. August 1987.
53
Section 4
54
Fig. 4.6 500 kV vertical break disconnector interrupting approximately 0.5 A of busbar and series capacitor bank platform charging current
Courtesy of Bonneville Power Administration
Interrupting capacitive currents
(a) Whip breaks contact on near phase
(b) Arc transfers to the main contacts
(c) Initial steady state arcs
(d) Arcs migrate upwards and become more
convoluted; note similarity between arcs
in centre and far phases
Fig. 4.7 115 kV centre break disconnector interrupting 1.8 A of line charging current (continued)
Courtesy of Puget Sound Energy
55
Section 4
(e) Restrikes in centre phase
(f) Partial arc collapse in near phase; restrike in
centre phase clearly evident
(g) Partial arc collapse in centre phase; effect
of restrike has waned
(h) Arcs appear to be similar in length; restrike
in far phase
Fig. 4.7 115 kV centre break disconnector interrupting 1.8 A of line charging current (continued)
Courtesy of Puget Sound Energy
56
Interrupting capacitive currents
(i) Restrikes in centre and far phases
(j) Partial arc collapse in near phase;
approaches total arc collapse in far phase
(k) Partial arc collapse in middle phase; far
phase has cleared
(l) Restrikes near and middle phases
Fig. 4.7 115 kV centre break disconnector interrupting 1.8 A of line charging current (continued)
Courtesy of Puget Sound Energy
57
Section 4
(m) Restrike in near phase
(n) Restrike wanes in near phase
(o) Centre phase clears; near phase approaches
total arc collapse
(p) Near phase clears; note hot gas remnants
(see (o) also) at former arc high and
presumably hottest points
Fig. 4.7 115 kV centre break disconnector interrupting 1.8 A of line charging current
Courtesy of Puget Sound Energy
58
(b) Steady state current; note that whip movement
results in the double arc image
(c) Arc still on whip propagating horizontally
(d) Partial arc collapse; note hot gas remnants
Fig. 4.8 115 kV centre break disconnector interrupting a line charging current of 4.6 A (continued)
59
Courtesy of Puget Sound Energy
Interrupting capacitive currents
(a) Contact parting
Section 4
60
(e) First restrike
(f) Second restrike
(g) Third restrike
(h) On verge of total arc collapse but see following
image (i)
Fig. 4.8 115 kV centre break disconnector interrupting a line charging current of 4.6 A (continued)
Courtesy of Puget Sound Energy
(j) Fourth restrike
(k) Fifth restrike
(l) Steady state current and partial arc collapse
Fig. 4.8 115 kV centre break disconnector interrupting a line charging current of 4.6 A (continued)
Courtesy of Puget Sound Energy
61
Interrupting capacitive currents
(i) Partial arc collapse and recovery
Section 4
62
(m) Sixth restrike
(n) Seventh restrike; arc has now left the whip and is
on the main contacts
(o) Eighth restrike
(p) Final total arc collapse and current interruption
Fig. 4.8 115 kV centre break disconnector interrupting a line charging current of 4.6 A
Courtesy of Puget Sound Energy
Interrupting capacitive currents
4.6
Capacitive current switching tests 2003
The apparent dependence of successful capacitive current switching on the source and load
side capacitances has been noted earlier. Knobloch66, 67 and Neumann62 made a similar observation with respect to the overvoltages generated during capacitive current switching with a
pantograph type disconnector. To further investigate this matter, a series of tests were conducted at the KEMA High Power Laboratory in November 2003.72
The tests were performed on a 300 kV centre-break type disconnector over a current range of
0.23 A to 2.3 A. The values of CS and CL were varied as shown in Table 4.3 using the test
circuit of Fig. 4.9
Table 4.3 300 kV disconnector test CS and CL combinations
CS (nF)
1.5
6
20
60
100
4.3
X
10.7
X
X
CL (nF)
19.3
X
X
X
X
38.6
X
X
40
X
X
X
X
X
A total number of 57 test shots were carried out at 171.5 kV to 173 kV source voltage and
various combinations of current, CS and CL. The results are discussed in terms of the generated load side overvoltages and the arc duration.
The dependence of the load side overvoltages on CS is evident in Fig. 4.10. There is approximately a 1 pu difference in magnitude between CS values of 1.5 nF and 60 nF. Fig. 4.11
shows the dependence of the overvoltages on the ratio CS/CL. The highest overvoltage values
occur for CS/CL < 1. These results support the conclusions of Knobloch66 and Neumann62 but
noting that higher overvoltages were obtained in the KEMA tests. The overvoltages have no
dependence on the current magnitude (refer to Eqn. (4.7)).
Fig. 4.9 Basic test circuit (TRV elements and measuring devices not
shown)
63
Section 4
Load side overvoltage (pu)
2.6
2.4
1.5 nF
2.2
6 nF
2
20 nF
1.8
60 nF
1.6
100 nF
1.4
Linear (1.5 nF)
Linear (60 nF)
1.2
1
0.0
0.5
1.0
1.5
2.0
2.5
Capacitive current (A)
Fig. 4.10 Load side overvoltages versus switched capacitive current
with CS as parameter
Load side overvoltage (pu)
2.6
2.4
0.23 to 0.6A
2.2
1 to 1.4A
2
2.1 to 2.3A
1.8
1.6
1.4
1.2
1
0
1
2
3
4
5
6
CS/CL
Fig. 4.11 Load side overvoltages versus CS/CL with switched
capacitive current as parameter
2200
Arc duration (ms)
2000
1.5 nF
6 nF
1800
20 nF
60 nF
100 nF
1600
1400
Linear (1.5 nF)
Linear (60 nF)
1200
1000
0
0.5
1
1.5
2
2.5
Capacitive current (A)
Fig. 4.12 Arc duration versus switched capacitive current with CS
as parameter
64
Interrupting capacitive currents
2200
Arc duration (ms)
2000
0.23 to 0.6A
1800
1 to 1.4A
2.1 to 2.3A
1600
1400
1200
1000
0
1
2
3
4
5
6
CS/CL
Fig. 4.13 Arc duration versus CS/CL with switch capacitive current as
parameter
Arc duration shows a similar dependency on CS and on CS/CL as shown in Figs. 4.12 and
4.13, respectively. The worst cases again occur for the lowest CS values and CS/CL < 1.
An interesting observation (also by Knobloch66) during the tests was that the arc exhibited
two distinct modes as shown in Fig. 4.14. In the first erratic mode, the arc evolves erratically
and has a length several times the blade tip spacing (Fig. 4.14 upper image). This mode
occurs for CS values less than 60 nF and CS/CL < 1 and is associated with higher overvoltages
and longer arc durations. The stiff mode is a contraction from the erratic mode to a form of
more or less a straight path between the contacts (Fig. 4.14 lower image). This mode occurs
Fig. 4.14 Capacitive current arc just before extinction
Upper image: Erratic arc 2 A, CS/CL = 0.04 (2.43 m tip spacing)
Lower image: Stiff arc 1 A, CS/CL = 3.1 (1.22 m tip spacing)
65
Section 4
for CS values of 60 nF or greater, CS/CL > 1 and lower values of capacitive current. This
mode results in the lowest overvoltages and shortest arc durations as shown in Figs. 4.15 and
4.16.
The phenomena associated with this switching duty, as described above, can be understood
by considering high frequency measurements of the arc current and the source and load side
voltages. Fig. 4.17 shows oscillograms for two significantly different values of CS/CL. A
capacitive current switching arc is a succession of interruptions and restrike. On restriking the
voltages on CS and CL become equalized through a high frequency discharge in the loop
formed by CS, CL and the disconnector (refer to Fig. 4.9). For the upper oscillogram of Fig.
4.17 with CS/CL = 2.5, this discharge is shown as an inset at 28 kHz. The equalization voltage
is dependent on the value of CS/CL and the overvoltage results from the subsequent transition
between this voltage and the source voltage. If the difference between the two voltages is
great, then the overvoltage will be high (lower oscillogram of Fig. 4.17) and vice versa (upper
oscillogram of Fig. 4.17).
Load side overvaltage (pu)
2.6
2.4
0.23 to 0.6A
2.2
1 to 1.4A
2
2.1 to 2.3A
1.8
Stiff arc
1.6
1.4
1.2
1
0
1
2
3
4
5
6
CS/CL
Fig. 4.15 Load side overvoltages versus CS/CL for stiff arc mode
(shaded points relate to the erratic arc mode)
2200
Arc duration (ms)
2000
0.23 to 0.6A
1 to 1.4A
1800
2.1 to 2.3A
1600
Stiff arc
1400
1200
1000
0
1
2
3
4
5
6
CS/CL
Fig. 4.16 Arc duration versus CS/CL for stiff arc mode (shaded points
relate to the erratic arc mode)
66
Interrupting capacitive currents
3
CS/CL = 2.5
2 . 5
current
2
1 . 5
1
source voltage
0 . 5
550 Hz
0
- 0 . 5
load voltage
voltages: 0.2 pu/div
time: 50 us/div
0. 8
0. 6
- 1
- 1 . 5
restrikes
arcing
0. 4
- 2
0. 2
- 2 . 5
0
6 7 0
6 7 5
6 8 0
6 8 5
6 9 0
6 9 5
7 0 0
voltages: 0.5 pu/div (= 122 kV/div); current: 50 A/div; time: 5 ms/div
- 0.2
3
- 0.4
28 kHz
CS/CL = 0.04
2 . 5
- 0.6
6 7 2. 4
6 7 2. 4 5
67 2 . 5
67 2 . 55
6 72 . 6
current
2
1 . 5
load voltage
1
0 . 5
0
- 0 . 5
source voltage
- 1
- 1 . 5
1 kHz
- 2
2.3 pu
- 2 . 5
1 3 9 0
1 3 9 5
arcing
1 4 0 0
1 4 0 5
1 4 1 0
1 4 1 5
1 4 2 0
Fig. 4.17 Test oscillograms for 2 A capacitive current and
CS/CL values of 0.04 and 2.5
67
Section 4
The equalization voltage can be calculated by considering the circuit in Fig. 4.9. Taking the
voltages on the source and load side capacitances as US and UL, the corresponding charges
are:
QS = USCS
QL = -ULCL
and
Qtotal = USCS + (-ULCL)
After restriking and charge redistribution, the voltages equalize at UE:
or
UE =
Q total
CS + C L
UE =
U SCS − U L C L
CS + C L
…(4.4)
Prior to restriking the voltage across the disconnector UD is:
UD = U S + U L
and substituting in Eqn. (4.4) for UL
UE =
or
U S C S −C L (U D − U S )
CS + C L
U E = US −
UD
1 + CS / C L
…(4.5)
The peak overvoltage value to ground UOV is given by:
U OV = U S + β(U S − U E )
…(4.6)
where β is the damping factor.
Substituting from Eqn. (4.5):

UD
U OV = U S + β
 1 + CS / C L
68



…(4.7)
Interrupting capacitive currents
The dependence of the overvoltage values on CS and CS/CL can now be explained:
•
CS > CL: the term (1 + CS/CL) is large and UOV will consequently be low (Fig. 4.17,
upper oscillogram).
•
CS < CL: the term (1 + CS/CL) is low and UOV will consequently be high (Fig. 4.17,
lower oscillogram).
If CS >> CL, then UOV can theoretically reach 3 pu (no damping with β = 1). However, some
damping is always present as is evident in Fig. 4.11.
CS and CS/CL also influence the arc duration (Figs. 4.12 and 4.13). For large CS/CL the transient arcing currents are low thus creating conditions for a high restriking repetition rate and
ultimate current interruption (Fig. 4.17, upper oscillogram). The arcing mode tends to be of
the stiff variety (Figs. 4.14, lower image and 4.16). For small CS/CL as in Fig. 4.17 lower
oscillogram, the transient currents are higher (as is the energy injected into the arc) and cannot be easily interrupted thus creating conditions for one restrike per cycle and for continuous
arcing of longer duration.
Based on the above analysis and the discussion in subsection 4.5, we can conclude that the
longest arc durations will occur when CS/CL < 1 and the capacitive current is of a magnitude
≥ 1 A such as to produce arc channel remnants thermally conducive to restriking. With
reference to Fig. 4.13, Fig. 4.18 illustrates this conclusion. The arc durations increase
progressively as the current increases from ≤ 0.57 A, through 1 A (where thermal effects are
starting to exert influence) to greater than 2 A (where thermal effects are definitely exerting
an influence).
The frequency of the overvoltage transient following voltage equalization is determined by
the source and CS and CL in parallel. For the upper and lower case of Fig. 4.17, the above
frequencies are 550 Hz and 1 kHz, respectively.
2200
Arc duration (ms)
2000
0.23 to 0.57A
1800
1A
2.1 to 2.25A
1600
1400
1200
1000
0
0.2
0.4
0.6
0.8
1
1.2
CS/CL
Fig. 4.18 Arc duration versus CS/CL (shaded points relate to currents
of 1 A or less)
69
Section 4
Finally, there is also the matter of arc reach. The angular turning velocity of the disconnector
blades is 40 degrees/second with arcing being initiated at 8 degrees. The 45 degree position
(half-open) is reached 930 ms after arc initiation. The arcs for all test cases extinguished for
blade angles greater than 45 degrees but always before the 90 degree position where the
maximum blade tip spacing of 2.8 is reached. Despite this the arc reach in a horizontal or
vertical direction was small. For the test case in the upper image of Fig. 4.14 with an arc
duration of 2040 ms, the arc reach is approximately half the blade tip spacing. This implies
that arc reach for capacitive current switching is much less an issue in terms of risk for
contact with adjacent phases or structures than it is for loop switching.
4.7
Conclusions
The conclusions with respect to the interruption of capacitive currents is based on field
observations and laboratory tests:
70
•
The arc will exhibit a number of arcing modes dependent on the magnitude of the
current and the source (CS) and load (CL) side capacitances.
•
For currents of 1 A or less, thermal effects are not significant and the arc duration is
dependent mainly on achieving the minimum disconnector gap spacing to withstand
the recovery voltage and the source and load side capacitances.
•
For currents greater than 1 A, thermal effects become significant and the arc duration
is dependent on the current magnitude in addition to achieving a minimum contact
gap spacing and the source and load side capacitances.
•
The longest arc durations at any current magnitude occur when CS/CL < 1.
•
The overvoltages associated with capacitive current switching are independent of the
current magnitude and the highest values occur when CS/CL < 1.
•
The 600 kVA rule relates only to current magnitude and any merit that it may have is
coincidental rather than factual.
•
Whip type auxiliary interrupting devices may provide benefit provided that CS/CL > 1.
Caution should be exercised in applications where CS/CL < 1.
Section 5
Loop switching
5.1
Introduction
Loop switching is the one type of current interruption under consideration where the arc
always plays a dominant role. The interruption process is one of commutating the current
from the disconnector path to a parallel path. The range of application is from up to 1600 A
between busbars within a substation down to tens to hundreds of ampere between
transmission lines and between distribution feeders. The industry practice is to only attempt
loop switching with disconnectors as described above. For loops involving cable circuits or
transformation, circuit breakers are used.
The equivalent circuit for loop switching circuit analysis is shown in Fig. 5.1. As will be
shown in this section, the initial current IS in the disconnector will go to zero when the arc
voltage approaches IS(XL + XS) and the open circuit voltage will be XL(IS + IL). Actual loop
switching traces are shown in Figs. 5.19 and 5.20.
iL
iT
iS
ua
Fig. 5.1 Equivalent circuit for loop switching
The condition for successful loop switching is therefore that the above arc voltage in the disconnector must be reached by a certain limiting point in the blade travel. For purposes of this
study, the limit is taken as a 45° angle on a vertical break disconnector.
5.2
Loop switching tests 1999–2000
A series of loop switching tests were run on 230 kV and 25 kV disconnectors at Powertech
Laboratories Inc. in British Columbia, Canada during the period November 1999 to
September 2000. The test circuit was essentially as shown in Fig. 5.1 with the voltage source
arranged to be constant current, i.e. the total current IT is unchanged after the test disconnector interrupts the initial current IS. A total of 205 test shots were conducted, 144 shots on the
230 kV disconnector and 61 shots on the 25 kV disconnector.
Based on the circuit shown in Fig. 5.1 the parameters to be considered are:
•
•
•
•
the disconnector initial current IS
the arc voltage ua
the series and parallel impedances XS and XL; (XS + XL) = loop impedance
and the recovery voltage across the disconnector XLIT
71
Section 5
We can write:
LL
di
di L
= LS S + u a
dt
dt
…(5.1)
and
∧
iT = I T sin ωt
= iS + i L
where LL = XL/ω and LS = XS/ω.
Eqn (5.1) then becomes:
LL
d(i T − i S )
di
= LS S + u a
dt
dt
and differentiating the left term:
ωL L Î T cos ωt = (L S + L L )
di S
+ ua
dt
…(5.2)
and we can further write:
ua = f(iS)
…(5.3)
The quantities recorded in the tests covered the above parameters and in addition the arcing
time ta. This enables some simplification of the analysis of the test results because ta is the
time it takes (for a given value of IS) to build up an arc voltage approaching to IS(XL + XS),
which time is dependent on the values of XL and XS. While Andrews et al hypothesized a
relationship only between IS and the loop impedance (XS + XL), it is more appropriate to
investigate a possible relationship between IS, (XS + XL) and ta if a limit on interrupting time
is to be set. The arc is also examined from the video records in order to provide a perspective
on its reach behaviour in space and time and its visual characteristics. The electrical characteristics of the arcs are examined in subsection 5.3.
5.2.1
Initial current, loop impedance and arcing time
In the experiments conducted by Andrews et al arcing time was not a consideration. Persistence of the arc was viewed as acceptable – even after the test disconnector was fully open –
provided that its reach was not excessive. This approach defies the prudent notion that the
current should be interrupted while the moving contact (the disconnector blade) is still in
motion (successful current interruption is defined in this way). In addition, such a scenario
would be unacceptable on manually operated disconnectors. The test data for the 230 kV and
25 kV disconnectors was therefore analyzed to determine the relationship between the initial
current, the loop impedance and the arcing time.
72
Loop switching
Fig. 5.2 shows the test points giving successful current interruption for the 230 kV disconnector for all tests except those of November 1999. The latter results are treated separately in
Annex F because that test was conducted under very windy conditions in contrast to the fair
Initial current (A)
200
150
100
50
0
0
50
100
150
200
Loop impedance (ohms)
Fig. 5.2 Successful current interruption test points for 230 kV
disconnector
weather conditions during the other tests. While the scatter is wide, there is an apparent
negative sloping trend between the two quantities. This trend becomes more obvious when
the results are segregated by arcing time as shown in Fig. 5.3.
200
<1000ms
1000-1500ms
Initial current (A)
>1500ms
150
100
50
0
0
50
100
150
200
Loop impedance (ohms)
Fig. 5.3 230 kV disconnector successful current interruption test
points with arcing time as parameter
73
Section 5
Similar plots for the 25 kV disconnector tests are shown in Figs. 5.4 and 5.5. It can be seen
that the pattern is similar to that for the 230 kV disconnector. The results of the tests on both
disconnectors are compared in Fig. 5.6. Assuming current interruption by the same blade
angle, the 25 kV disconnector will have a lower capability than the 230 kV disconnector.
Initial current (A)
200
150
100
50
0
0
50
100
150
200
Loop impedance (ohms)
Fig. 5.4 Successful current interruption test points for 25 kV
disconnector
Examination of Figs. 5.3 and 5.5 shows that it should be possible to derive operating limits to
some rule using statistical analysis. The rule proposed for this analysis is that the switch shall
interrupt the current within a maximum arcing time of 1 second which corresponds to about a
30° blade angle. The justification for this rule is as follows:
200
<1000ms
1000-1500ms
Initial current (A)
>1500ms
150
100
50
0
0
50
100
150
200
Loop impedance (ohms)
Fig. 5.5 25 kV disconnector successful current interruption test points
with arcing time as parameter
74
Loop switching
25kV <1000ms
25kV 1000-1500ms
25kV >1500ms
230kV <1000ms
230kV 1000-1500ms
230kV >1500ms
Initial current (A)
200
150
100
50
0
0
50
100
150
200
Loop impedance (ohms)
Fig. 5.6 25 kV and 230 kV disconnector successful current
interruption test points with arcing time as parameter
•
It assures current interruption while the blade is in motion.
•
It provides a degree of arc control in that the arc is extinguished before it has an
opportunity to propagate.
•
It limits any possible phase-to-phase interaction which is important given that the testing was done on a single phase basis.
•
It should avoid any negative influence by the weather given that arc lengths are short.
As noted earlier a major criticism of the work of Andrews et al was the failure to relate
current interruption to blade position or arcing time. The results in the foregoing figures suggest a power relationship between the initial current in the disconnector and the total loop
impedance. We can therefore write:
IS = AX t
−B
where A and B are constants and X t = XS + X L .
Linearizing the above equation gives:
ln IS = −B ln Xt + ln A
which is of the form y = mx + c and a linear regression can be performed to derive the values
of A and B. The results of this regression analysis for the 230 kV disconnector by arcing time
are shown in Fig. 5.7 giving the following equations:
75
Section 5
500-1000ms
1000-1500ms
>1500ms
Power (1000-1500ms)
Power (>1500ms)
Power (500-1000ms)
Initial current (A)
200
150
100
50
0
0
50
100
150
200
Loop impedance (ohms)
Fig. 5.7 Regression analysis for 230 kV disconnector test results with
arcing time as parameter
IS 1500+
= 620 X −t 0.38
…(5.4)
IS 1000−1500
= 340 X −t 0.35
…(5.5)
IS 500−1000
= 150 X −t 0.32
…(5.6)
The next step is to calculate the standard deviation of the regression (also known as the standard error of the estimate) and conservatively to set the interrupting limits at two (2) standard
deviations below the mean. The results of this calculation are shown in Fig. 5.8.
Fig. 5.8 can be read as follows. Below the middle line there is a 2% probability only that the
arcing time will be as high as 1 to 1.5 seconds. Between the middle and upper lines there is a
<1000ms
1000-1500ms
>1500ms
<1000ms, -2SD
1000-1500ms, -2SD
>1500ms, -2SD
200
Initial current (A)
150
100
50
0
0
50
100
150
200
Loop impedance (ohms)
Fig. 5.8 Mean minus two (2) standard deviation regression analysis for
230 kV disconnector results with arcing time as parameter
76
Loop switching
2% probability only that the arcing time will exceed 1.5 seconds. Above the upper line there
is a 98% probability that the arcing time will exceed 1.5 seconds. The equations for the
middle and upper lines are:
IS 2% 1000-1500+ = 210 X −t 0.35
…(5.7)
= 340 X −t 0.38
…(5.8)
IS 98% 1500+
This analysis does not stand alone because arc reach, as described in the next subsection, is a
major consideration.
5.2.2
Arc video record analysis
A. Video camera arrangement
Two digital video cameras were used to record the progression of the arc and its movement
relative to the test 230 kV disconnector. The cameras were placed at jaw assembly height, the
first facing along the longitudinal axis of the closed disconnector and the second facing the
jaw assembly with a line of sight at right angles to that of the first camera. This is in contrast
to the camera arrangement used by Andrews et al who also used two cameras with lines of
sight at right angles to one another but at oblique angles to the longitudinal axis of the test
disconnector. In addition, both cameras were placed on the same side of the test disconnector.
B. General visual observations
The arcs can be said to exhibit an expected randomness but with certain patterns related to the
current level and the arcing time. However, at times the arc exhibits aberrant behaviour for
reasons discussed later. In the following, still picture images from the videos are used to
illustrate arc behaviour.
At low current the arc has the appearance of a reasonably well defined conduction path. As
the current increases, the appearance tends towards that of a poorly defined flame. This is
illustrated in Figs. C1 to C8 in Annex C. Note also that, as the current increases, the upward
motion of the arc changes to motion to the horizontal (refer to subsection 5.2.2C).
The progression of the arc for the case of an initial current of 62 A is shown in Figs. C9 to
C20 in Annex C. In this case the arc exhibited a predominant upward movement and a number of partial arc collapses as it partially short-circuits itself (see Figs. C14 to C17).
As the arc starts to decay beyond recovery, it first exhibits a change of colour at the electrode
followed by a decrease in the diameter (or diameters of several parallel paths) and defragmentation. This is illustrated in Figs. C21 to C31 in Annex C for the case of an initial current
of 82 A. Note that after current interruption (Fig. C29), the remnants of the arc persist for
about 4 cycles (Figs. C30 and C31). At higher currents the effect is even more pronounced:
the case of arc decay at 165 A initial current is shown in Figs. C32 to C37 in Annex C. It is
noteworthy that prior to decay the arc maintains its luminosity despite the decreasing current.
This suggests that power input to the arc is constant or increasing and that decay occurs when
77
Section 5
the power input is suddenly removed resulting in an equally sudden temperature drop and the
change of colour noted above.
The horizontal movement of the arc, particularly at the higher current values, is judged to be
due to the electromagnetic force within the arc loop itself and to be independent of the circuit
arrangement. In the Powertech images in Annex C, the viewing direction is such that the
arcing horn is to the right of the blade. In the majority of the tests, the arc formed a loop
between the blade and the horn and then moved to the right. In a number of cases, refer for
example to Figs. C21 to C31, Fig. C55 and Fig. C58, the arc formed a loop to the left and
then moved to the left. The upward movement of the arc is due in part to blade motion and
then thermal effects.
The arc also exhibited aberrant behaviour. This occurred at the higher currents of 87 A, 90 A
and 103 A in which cases the arc “ran” down the opening blade. The 103 A initial current
case is shown in Figs. C38 to C45 in both front and side images. Clearly at the point of transfer of the current to the arcing contacts, the arc rooted itself on the blade rather than the
moving arcing contact as shown in Fig. C46 in Annex C. A solution to prevent this would
possibly be to offset the arcing contacts to the outside of the jaw assembly.
C. Arc geometry
The intent of the video records is to determine arc length and arc reach. However, it was not
possible to measure arc length with any degree of certainty. The convoluted nature of the arc
makes it impossible to define a clear arc path from one electrode to the other even with two
camera angles. On a limited number of tests shots, it was possible to measure arc reach as
defined by Andrews et al and these test points were plotted on Fig. 5.9 (refer to Fig. 2.5 in
subsection 2.4). The test points are in reasonable agreement with those of IREQ. This sheds
further doubt, i.e. in addition to that already expressed with regard to transformer magnetizing current, on the work of Andrews et al however well intentioned it may be.F5
As the disconnector opens it is expected that the arc will evolve vertically and horizontally. In
the vertical direction, the arc is drawn upwards by the blade motion and in addition by thermal effects and the magnetic force, i.e. the force on the arc due to its self magnetic field. The
absolute acceptable limit for vertical rise of the arc is that it shall never rise above the elevation of the blade tip when the disconnector is in the fully open position. This is simple logic
but reality obviously demands that the disconnector interrupt the current well before this
position is reached. In the horizontal direction, the arc is driven primarily by the magnetic
force and some limitation on reach is required to avoid the possibility of phase-to-phase
flashover. Given this perspective it is only necessary to consider the arc as viewed in the disconnector longitudinal direction.
F5
78
Overall the data and results presented by Andrews et al do not stand up well to scrutiny. The transformer
magnetizing current tests were discussed in subsection 2.2 and a comparative analysis of the loop switching
tests is given in Annex E. On this basis, the IEEE guide must be viewed as questionable.4
Loop switching
Fig. 5.9 Comparison of loop switching test results between
IREQ,9 Andrews et al5 and Peelo
Still images were made of thirty test
shots representative of the range of
current and loop impedance used in
the test series. Arc X and Y coordinates were measured at two cycle
intervals for the entire arcing
period. The X and Y coordinates in
this instance do not denote a single
point but rather two points that correspond to one another (Fig. 5.10).
For each measured image, X is the
maximum horizontal point on the
arc from a line through the arc roots
and Y is the maximum vertical
point on the arc from a reference
point, which in this case is the jaw
end terminal pad. Surprising or not,
the arc – whether it moves to the
left or to the right – exhibits a distinctive pattern of behaviour.
XY plots for initial current ranges of 15 to 50 A, 51 to 90 A and 91 to 165 A are shown in
Figs. 5.11 to 5.13 and a plot for all initial current values in Fig. 5.14. The XY coordinates are
all plotted in one direction only regardless of whether the arc was to the left or right of the
disconnector. These figures show the following:
•
In the range 15 to 50 A (Fig. 5.11), the arc tends to predominantly follow the blade
motion rising due to thermal effects and the X-coordinate values are mostly low.
Fig. 5.10 Measurement of X and Y coordinates
79
Section 5
Y-coordinates (mm)
1200
1000
800
600
15 to 30 A
400
31 to 50 A
200
0
0
200
400
600
800
1000
X-coordinates (mm)
Fig. 5.11 XY plot for initial currents 15 to 50 A
•
In the range 51 to 90 A (Fig. 5.12), the Y-coordinate increases by 0.8 m over and
above that for the 15 to 50 A range and the X-coordinate increases due to electromagnetic forces.
Y-coordinates (mm)
2000
1800
1600
1400
1200
1000
800
600
400
51 to 70 A
71 to 90 A
200
0
0
200
400
600
800
1000
X-coordinates (mm)
Fig. 5.12 XY plot for initial currents 51 to 90 A
•
In the range 91 to 165 A (Fig. 5.13), the Y-coordinate remains relatively unchanged
compared to the 51 to 90 A range but the X-coordinate has doubled. The reason for
this would appear to be that arc collapse (discussed later) provides a natural limiting
mechanism on upward motion of the arc.
•
Viewing all of the results from 15 to 165 A (Fig. 5.14), the XY coordinates show a
triangular pattern, the triangle being bounded by the Y-axis, a horizontal line at Y =
1600 mm and line drawn through the lowest most points. The results are treated
statistically in subsection 5.2.3.
The spread in the XY plots is due to the wide range of current and loop impedances. Clearly
the higher the current or the impedance the longer the arcing period. In Figs. 5.15 to 5.18, XY
plots are shown for four approximately constant current levels and varying impedance.
80
Loop switching
Y-coordinates (mm)
2500
2000
1500
1000
91 to 110 A
500
111 to 130 A
130 to 165 A
0
0
500
1000
1500
2000
X-coordinates (mm)
Fig. 5.13 XY plot for initial currents 91 to 165 A
Y-coordinates (mm)
2500
15 to 30 A
2000
31 to 50 A
51 to 70 A
1500
71 to 90 A
1000
91 to 110 A
111 to 130 A
500
131 to 165 A
0
0
500
1000
1500
2000
X-coordinates (mm)
5.14 XY plot for initial currents 15 to 165 A
To provide a real perspective on Figs. 5.15 to Fig. 5.18, still images of the arc for each of the
cases plotted are shown in Figs. C47 to C60 in Annex C.
Y-coordinates (mm)
1200
1000
24 A and 40 ohms
800
23 A and 100 ohms
600
21 A and 200 ohms
400
29 A and 150 ohms
200
0
0
100
200
300
400
X-coordinates (mm)
Fig. 5.15 XY plots at initial currents 21 to 29 A and varying
loop impedance
81
Y-coordinates (mm)
Section 5
2000
1800
1600
1400
1200
1000
800
600
400
200
0
75 A and 20 ohms
73 A and 40 ohms
69 A and 70 ohms
70 A and 100 ohms
0
200
400
600
800
X-coordinates (mm)
Fig. 5.16 XY plots at initial currents 69 to 75 A and varying loop
impedance
Y-coordinates (mm)
2500
2000
98 A and 40 ohms
1500
96 A and 100 ohms
1000
90 A and 150 ohms
500
0
0
200
400
600
X-coordinates (mm)
Y-coordinates (mm)
Fig. 5.17 XY plots at initial currents 90 to 98 A and varying loop
impedance
2000
1800
1600
1400
1200
1000
800
600
400
200
0
121 A and 60 ohms
124 A and 150 ohms
122 A and 200 ohms
0
500
1000
1500
X-coordinates (mm)
Fig. 5.18 XY plots at initial currents 121 to 124 A and varying
loop impedance
82
Loop switching
D. Arc collapse
Arc collapse occurs when two points along the arc come into contact with one another. An
example of this is as shown in Figs. C14 to C17. Electrically this represents a partial short
circuit of the arc and a decrease in its resistance. The net effect is that the current now transfers back from the parallel circuit into the arc path. This is illustrated in Fig. 5.19 which
shows the original test trace for the test shot on Figs. C14 to C17. The trace shows two major
arc collapses followed by a number of minor collapses and current interruption. At the high
currents, the arc takes on a wriggling appearance with multiple minor arc collapses which
produces a ripple effect on the disconnector current as shown in Fig. 5.20.
Arc collapse is opportunistic but also inevitable due to the convoluted path of the arc. The
longer the arc or – put another way – the longer the arcing time, the more the arc is likely to
experience collapses and thus limit its growth. This is demonstrated in Fig. 5.21 for the tests
on the 25 kV and 230 kV disconnectors.
Legend (top to bottom): 1.
2.
3.
4.
Arc voltage
Total current (iS + iL)
Voltage across loop (Uoc = 8 kV)
Disconnector current (IS = 62 A)
Fig. 5.19 Original test trace: interrupting 62 A with loop
impedance 150 ohms
83
Section 5
Legend (top to bottom): 1.
2.
3.
4.
Arc voltage
Total current (iS + iL)
Voltage across loop (Uoc = 8.9 kV)
Disconnector current (IS = 145 A)
Fig. 5.20 Original test trace: interrupting 145 A with loop impedance
70 ohms
Number of arc collapses
9
8
7
25 kV disconnector
6
5
230 kV disconnector
4
Linear (230 kV
disconnector)
3
2
1
0
0
1000
2000
3000
Arcing time (ms)
Fig. 5.21 Number of arc collapses versus arcing time for 25 kV and
230 kV disconnectors
5.2.3
Application perspective
The next step is to provide a perspective on how the test results described in the foregoing
can be related to actual applications. This will be treated by considering arcing time and arc
reach.
84
Loop switching
A. Arcing time
The arcing times for all the Year 2000 230 kV disconnector tests at Powertech are plotted
against the initial current in Fig. 5.22 with the loop impedance as parameter. There is an
obvious increasing tendency both with initial current and constant loop impedance and with
3000
20 ohms
30 ohms
40 ohms
60 ohms
70 ohms
100 ohms
110/120 ohms
150 ohms
200 ohms
Arcing time (ms)
2500
2000
1500
1000
500
0
0
50
100
150
200
Initial current (A)
Fig. 5.22 Arcing time versus initial current for Powertech Year 2000
tests on 230 kV disconnector
loop impedance and constant initial current. However, loop impedances greater than 60 or
70 ohms are not realistic in practice and we can limit the loop impedance range to maximum
100 ohms to provide some margin. This is shown in Fig. 5.23 and a linear regression applied
Arcing time (ms)
2500
2000
1500
100 ohms or less loop
impedance
1000
Linear (100 ohms or less
loop impedance)
500
y = 8.9679x + 253.32
0
0
50
100
150
200
Initial current (A)
Fig. 5.23 Arcing time versus initial current for loop impedances of
100 ohms or less
to the test points. The standard deviation of the regression (usually termed the standard error
of the estimate) is calculated and shown in Fig. 5.24. Most points are within one standard
deviation particularly at the lower initial current values and we can examine the influence of
loop impedance as shown in Fig. 5.25 to Fig. 5.31.
85
Section 5
100 ohms or less loop impedance
3000
Plus 1 stddev
Arcing time (ms)
2500
Plus 2 stddev
Minus 1 stddev
2000
Minus 2 stddev
1500
Linear (100 ohms or less loop
impedance)
Linear (Plus 1 stddev)
1000
Linear (Plus 2 stddev)
500
Linear (Minus 1 stddev)
Linear (Minus 2 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.24 Linear regression and standard deviations
2500
20 ohms loop impedance
30 ohms loop impedance
Arcing time (ms)
2000
40 ohms loop impedance
60 ohms loop impedance
1500
70 ohms loop impedance
100 ohms loop impedance
Linear regression
1000
Plus 1 stddev
Minus 1 stddev
Linear (Linear regression)
500
Linear (Plus 1 stddev)
Linear (Minus 1 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.25 Linear regression and one standard deviation for various impedances
2500
20 ohms loop impedance
30 ohms loop impedance
Arcing time (ms)
2000
40 ohms loop impedance
60 ohms loop impedance
1500
70 ohms loop impedance
100 ohms loop impedance
Linear regression
1000
Plus 1 stddev
Minus 1 stddev
Linear (Linear regression)
500
Linear (Plus 1 stddev)
Linear (Minus 1 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.26 Linear regression and one standard deviation for 20 ohms loop
impedance
86
Loop switching
2500
20 ohms loop impedance
30 ohms loop impedance
Arcing time (ms)
2000
40 ohms loop impedance
60 ohms loop impedance
70 ohms loop impedance
1500
100 ohms loop impedance
Linear regression
1000
Plus 1 stddev
Minus 1 stddev
Linear (Linear regression)
500
Linear (Plus 1 stddev)
Linear (Minus 1 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.27 Linear regression and one standard deviation for 30 ohms loop
impedance
2500
20 ohms loop impedance
30 ohms loop impedance
Arcing time (ms)
2000
40 ohms loop impedance
60 ohms loop impedance
70 ohms loop impedance
1500
100 ohms loop impedance
Linear regression
1000
Plus 1 stddev
Minus 1 stddev
Linear (Linear regression)
500
Linear (Plus 1 stddev)
Linear (Minus 1 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.28 Linear regression and one standard deviation for 40 ohms loop
impedance
2500
20 ohms loop impedance
30 ohms loop impedance
Arcing time (ms)
2000
40 ohms loop impedance
60 ohms loop impedance
70 ohms loop impedance
1500
100 ohms loop impedance
Linear regression
1000
Plus 1 stddev
Minus 1 stddev
Linear (Linear regression)
500
Linear (Plus 1 stddev)
Linear (Minus 1 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.29 Linear regression and one standard deviation for 60 ohms loop
impedance
87
Section 5
2500
20 ohms loop impedance
30 ohms loop impedance
Arcing time (ms)
2000
40 ohms loop impedance
60 ohms loop impedance
1500
70 ohms loop impedance
100 ohms loop impedance
Linear regression
1000
Plus 1 stddev
Minus 1 stddev
Linear (Linear regression)
500
Linear (Plus 1 stddev)
Linear (Minus 1 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.30 Linear regression and one standard deviation for 70 ohms loop
impedance
2500
20 ohms loop impedance
30 ohms loop impedance
Arcing time (ms)
2000
40 ohms loop impedance
60 ohms loop impedance
1500
70 ohms loop impedance
100 ohms loop impedance
Linear regression
1000
Plus 1 stddev
Minus 1 stddev
Linear (Linear regression)
500
Linear (Plus 1 stddev)
Linear (Minus 1 stddev)
0
0
50
100
150
200
Initial current (A)
Fig. 5.31 Linear regression and one standard deviation for 100 ohms loop
impedance
The maximum arcing time tamax for an initial current of IS can be written as:
tamax = 9 IS + 623
…(5.9)
and is applicable to horizontally mounted vertical break disconnectors having the same
opening speed as the test disconnector. The blade length of the test disconnector was 2.84 m
and the opening time was approximately 4 seconds.
88
Loop switching
B. Arc reach
Arc reach is considered relative to the electrode points, i.e. the blade tip and the arcing horn
tap. Horizontal (Hreach) and vertical (Vreach) are defined as:
Hreach = maximum X-coordinate
Vreach = (maximum Y-coordinate) - 0.38
both quantities being in metres. With reference to Figs. 5.11 to 5.14, the test points for the
various current ranges fall within boundaries as is shown in Fig. 5.32. This diagram assumes
that arc will propagate either to the left or to the right and for example is read as follows: all
of the test points in Fig. 5.12 fall within the boundary designated as C, which boundary also
captures the boundaries A and B for the lower currents. All test points – Fig. 5.14 – thus fall
within boundary D. To provide a three-phase perspective, Fig. 5.33 shows in addition the
superimposed minimum metal-to-metal phase-to-phase clearances as recommended in reference 61 and the fully open position blade tip elevations for 72.5 kV, 145 kV and 230 kV
disconnectors. This figure illustrates the challenge of dealing with the reality of three-phase
loop switching.
The measured horizontal and vertical reach data can be statistically analyzed in the manner as
done above for arcing time. Only the cases of loop impedances of 100 ohms or less are considered. The results for horizontal and vertical reaches are shown in Fig. 5.34 and 5.35,
respectively. The number of points are limited – a total of 21 – and a prudent statistical (as
well as practical) approach is to base application on the plus 2 standard deviation limit. This
gives the following equations:
Hreach = 0.0055 IS + 0.36
…(5.10)
Vreach = 0.007 IS + 0.86
…(5.11)
and
C. Practical considerations
Even without as yet examining the electrical characteristics of the arc, loop switching is a
complex issue. Empirical relationships have been established between initial current, loop
impedance, arcing time and arc reach. These results are presented at face value but noting a
reasonable consistency between tests run at different times. No specific guidance or rules are
stated because this falls within realm of the individual utility. The reason for this is that each
disconnector application is unique, if only in terms of the layout arrangement and associated
clearances. However, certain practical considerations can be observed:
1. In applying the results, personnel safety should be maximized and risk minimized.
2. Extreme limits should be avoided: there should always be an expectation that the
switching operation will be successful.
89
Fig. 5.32 Arc reach boundaries
Section 5
90
Fig. 5.33 Arc reach boundaries and minimum phase-to-phase clearances at 72.5 kV, 145 kV and
Loop switching
91
Section 5
1.4
0.8
Loop impedance 20 to
30 ohms
Loop impedance 40 to
60 ohms
Loop impedance 70 to
100 ohms
Linear regression
0.6
Plus 1 stddev
Horizontal reach (m)
1.2
1
Plus 2 stddev
0.4
Linear (Linear
regression)
Linear (Plus 1 stddev)
0.2
0
0
50
100
150
200
Linear (Plus 2 stddev)
Initial current (A)
Fig. 5.34 Horizontal arc reach for loop impedances of 100 ohms or less
Vertical reach (m)
2
1.8
Loop impedance 20 to
30 ohms
Loop impedance 40 to
60 ohms
Loop impedance 70 to
100 ohms
Linear regression
1.6
1.4
1.2
1
0.8
0.6
0.4
Plus 1 stddev
Plus 2 stddev
Linear (Linear
regression)
Linear (Plus 1 stddev)
0.2
0
0
50
100
150
200
Linear (Plus 2 stddev)
Initial current (A)
Fig. 5.35 Vertical arc reach for loop impedances of 100 ohms or less
3. There is no one encompassing relationship between the initial current, loop impedance,
arcing time and arc reach. Limits should be set with respect to one or two quantities and
then consider the impact on the other quantities. This may be an iterative process.
4. The tests were run on one phase of a three-phase disconnector. For the current range in
question, no electromagnetic influence between phases is expected. However, there is a
risk that arcs on adjacent phases may form towards each other and result in a phase-tophase fault. This situation should be mitigated to the greatest degree possible: there is
some evidence to suggest that placement of the arcing horns may provide some control
over the location of arc formation.
5. The tests were run on a vertical break type disconnector and due caution should be
exercised in applying the results to other disconnector types. This subject is discussed in
section 6.
92
Loop switching
6. The behaviour of the arc dominates this switching case. The electrical characteristics of
the arc and the condition for arc collapse (extinction) are examined in the next subsection.
5.3
Electrical characteristics of the arc
The electrical characteristics of the arc are examined by considering the arc voltage u and the
arc current i. In order, the arc U-I characteristic, the arc resistance U/I, the arc power UI and
the rate of change of arc power d(ui)/dt will be examined and discussed. The intent is to
confirm that the condition for arc stability is
d(ui )
>0
dt
…(5.12)
and to determine related arc-circuit interactions.
MATLAB is used to analyze the data from the Powertech (PT) tests of February and May
2000 and from Eindhoven University of Technology (TUe) and KEMA in 2003. The PT and
KEMA tests were performed on a vertical break type disconnector and the TUe tests on a
centre-break type disconnector. The PT tests used a current source circuit, while the TUe and
KEMA tests used a voltage source circuit.
A. Arc U-I characteristics
The arc U-I characteristics are plotted by taking the peak values of current and the arc voltage
corresponding to that current value. Typical U-I plots for the PT tests are shown in Figs. 5.36
and 5.37 and similar plots for TUe in Fig. 5.38. A U-I plot for the KEMA tests is shown later
in Fig. 5.39.
Review of Figs. 5.36 to 5.38 shows the following:
•
The plots all exhibit the same parabolic shape independent of the disconnector type
and the current or voltage source.
•
The plots all exhibit good consistency and reproducibility between the tests on the
same occasion and on different occasions.
•
The plots generally end at a voltage approaching the initial current times the loop
impedance. This is the point at which the arc becomes unstable and collapses beyond
possible recovery occurs. The arc collapse is discussed further in the following parts
of this subsection.
Direct comparison between tests at the three locations can be made for the cases where the
loop impedances are equal or reasonably close to one another. Comparison between the PT,
TUe and KEMA tests are shown in Fig. 5.39. The comparisons are remarkable for their visual
similarity and further by regression as shown in Table 5.1. The regression is parabolic of the
form:
93
Section 5
Powertech Februari 2000, Xt = 30 Ω
Powertech May 2000, Xt = 30 Ω
5
5
PT 6_19
PT 5_02
4.5
PT 6_21
4.5
PT 5_13
4
PT 7_14
3
PT 7_02
2.5
PT 8_02
2
1.5
PT 8_25
3.5
PT 7_23
3
PT 6_29
2.5
2
1.5
1
1
0.5
0.5
0
PT 8_24
PT 6_02
3.5
Voltage at peak current [kV] →
Voltage at peak current [kV] →
4
0
20
40
60
80
100
120
Peak current [A] →
140
160
180
0
200
0
Powertech Februari 2000, Xt = 60 Ω
PT 5_03
8
PT 5_14
7
PT 6_03
6
PT 7_03
100
150
Peak current [A] →
200
Powertech May 2000, Xt = 60 Ω
PT 6_20
10
PT 7_13
PT 8_26
8
Voltage at peak current [kV] →
Voltage at peak current [kV] →
9
50
PT 8_03
5
4
3
2
PT 8_27
PT 7_22
6
PT 6_30
4
2
1
0
0
50
100
150
Peak current [A] →
200
250
0
0
50
100
150
Peak current [A] →
Fig. 5.36 PT: U-I characteristics for loop impedances 30 and 60 ohms
94
200
250
Loop switching
Powertech Februari 2000, Xt = 100 Ω
Powertech May 2000, Xt = 100 Ω
10
10
PT 5_28
PT 5_08
9
PT 6_16
9
PT 7_17
PT 5_19
8
PT 7_08
6
PT 8_08
5
4
3
PT 8_19
7
PT 7_26
6
PT 6_26
5
4
3
2
2
1
0
PT 8_18
PT 6_08
7
Voltage at peak current [kV] →
Voltage at peak current [kV] →
8
1
0
20
40
60
80
Peak current [A] →
100
120
0
140
0
50
100
150
Peak current [A] →
Powertech Februari 2000, Xt = 150 Ω
Powertech May 2000, Xt = 150 Ω
18
PT 5_09
PT 5_29
PT 5_31
PT 6_13
PT 6_15
PT 7_18
PT 7_20
PT 8_12
PT x 8_16
PT x 8_17
PT x 7_27
PT x 7_29
PT x 6_23
PT x 6_25
PT 5_11
16
16
PT 5_20
x
14
x
x
PT 5_22
PT 6_09
14
Voltage at peak current [kV] →
Voltage at peak current [kV] →
PT 6_11
12
PT 7_09
PT x 7_11
PT x 8_09
10
PT x 8_11
x
8
x
6
x
12
10
8
x
6
4
4
2
2
x
x
0
0
20
40
60
80
100
Peak current [A] →
120
140
160
180
0
0
20
40
60
80
100
120
Peak current [A] →
140
160
180
Fig. 5.37 PT: U-I characteristics for loop impedances 100 and 150 ohms
95
Section 5
Eindhoven University of Technology,
Xs = 35 Ω
TU/e D073
TU/e D074
TU/e D075
TU/e D076
TU/e D077
TU/e D078
TU/e D079
TU/e x D080
TU/e x D081
TU/e x D082
TU/e x D083
TU/e x D084
TU/e x D085
TU/e x D086
TU/e o D087
TU/e o D088
3.5
x x
o
x
x x x
3
o
x
Voltage at peak current [kV] →
Eindhoven University of Technology,
Xs = 56 Ω
2.5
2
1.5
1
0.5
0
0
20
40
60
80
Peak current [A] →
100
120
Eindhoven University of Technology,
Xs = 170 Ω
Eindhoven University of Technology,
Xs = 100 Ω
3.5
TU/e D023
TU/e D017
TU/e D024
x
Voltage at peak current [kV] →
x
TU/e D149
TU/e x D150
2
TU/e x D151
TU/e x D152
TU/e x D153
1.5
1
TU/e D154
x
TU/e D155
2
TU/e x D156
TU/e x D157
TU/e x D158
1.5
TU/e x D159
1
0.5
0.5
0
TU/e D021
x
TU/e D148
x
TU/e D020
2.5
TU/e D028
x
TU/e D019
x
TU/e D027
x
2.5
TU/e D018
TU/e D026
Voltage at peak current [kV] →
3
3
0
5
10
15
20
25
Peak current [A] →
30
35
40
0
0
5
10
15
Peak current [A] →
Fig. 5.38 TUe: U-I characteristics for loop impedances 35, 56, 100 and 170 ohms
96
20
Loop switching
Comparison TUe [X = 75 Ω]
and Powertech [X = 70 Ω] cases
Comparison KEMA and Powertech
cases for X = 20 Ω
1.2
TU/e D050
TU/e D050 fit
TU/e D047
TU/e D047 fit
PT 5_07
PT 5_07 fit
PT 5_18
PT 5_18 fit
PT 5_27
PT 5_27 fit
2.5
2
Voltage at peak current [V] →
1
Voltage at peak current [kV] →
3
KEMA 6025
KEMA 6025 fit
KEMA 6028
KEMA 6028 fit
PT 6_01
PT 6_01 fit
PT 6_22
PT 6_22 fit
PT 6_31
PT 6_31 fit
0.8
0.6
0.4
1.5
1
0.5
0.2
0
0
-0.2
10
20
30
40
50
60
Peak current [A] →
70
80
-0.5
10
90
3
25
30
35
Peak current [A] →
40
45
50
55
20
22
2.5
TU/e D023
TU/e D023 fit
TU/e D027
TU/e D027 fit
PT 5_08
PT 5_08 fit
PT 5_19
PT 5_19 fit
PT 5_28
PT 5_28 fit
2
2
Voltage at peak current [V] →
2.5
Voltage at peak current [V] →
20
Comparison TUe [X = 170 Ω]
and Powertech [X = 150 Ω] cases
Comparison TUe and Powertech
cases for X = 100 Ω
1.5
1
1.5
1
TU/e D018
TU/e D018 fit
TU/e D021
TU/e D021 fit
PT 5_11
PT 5_11 fit
PT 5_22
PT 5_22 fit
PT 5_31
PT 5_31 fit
0.5
0.5
0
0
-0.5
15
5
10
15
20
25
Peak current [A] →
30
35
40
-0.5
4
6
8
10
12
14
Peak current [A] →
16
18
Fig. 5.39 Comparison between PT, TUe and KEMA U-I characteristics
97
Section 5
Table 5.1 Regression analysis for plots in Fig. 5.39
Loop impedance
(ohms)
Test shot
A
(kV/A²)
B
(kV/A)
20
20
20
20
20
70
70
70
70
70
100
100
100
100
100
150
150
150
150
150
KEMA 6025
KEMA 6028
PT 6_01
PT 6_22
PT 6_31
TUe D050
TUe D047
PT 5_07
PT 5_18
PT 5_27
TUe D023
TUe D027
PT 5_08
PT 5_19
PT 5_28
TUe D018
TUe D021
PT 5_11
PT 5_22
PT 5_31
-0.000104
-0.000175
-0.00001115
-0.00007065
-0.000061875
-0.0011292
-0.0015965
-0.00089545
-0.0014191
-0.0011638
-0.0035574
-0.0023317
-0.0017985
-0.003947
-0.0020869
-0.0072819
-0.0067408
-0.0057214
-0.0070318
-0.0082058
-0.0117
-0.0042
-0.02
-0.013865
-0.01409
-0.0043981
0.031036
-0.014521
0.025659
0.0085697
0.081565
0.0011489
-0.022021
0.090545
-0.0074231
0.063179
0.03726
0.025106
0.068585
0.11239
C
(kV)
1.5488
1.3742
1.4764
1.3896
1.4816
2.8568
2.2749
2.6048
2.018
2.294
1.6103
2.9155
2.5186
1.2519
2.5112
2.0734
2.384
1.9601
1.7751
1.3889
U = AI2 + BI + C
Arcing time
(ms)
Initial current
(A)
1340
957
405
264
507
1203
1164
807
708
291
917
890
708
775
500
880
850
675
425
601
53
54
50
50
52
34
34
34
35
33
25
25
23
24
23
15
15
16
16
15
…(5.13)
where U is in kV, I in amperes and A, B and C are constants.
Despite the differences in the disconnector types and circuit sources, the arc as a circuit
element appears to behave in a similar manner in electrically similar circuits (refer also to
subsection 5.4).
B. Arc resistance U/I
The arc resistance U/I for comparable PT, KEMA and TUe tests are shown in Fig. 5.40. For
any individual test, the plots show a similar pattern by way of a gradual build-up of resistance
and a very rapid increase in resistance as the arc goes towards extinction. Partial arc collapses
are clearly evident.
The TUe tests, in particular, exhibit on the whole a slower rate of increase of arc resistance
and thus a longer arcing time. This is judged to be related to the disconnector type (centre
break) and is discussed in section 6.
C. Arc power UI
The arc power UI for comparable PT, KEMA and TUe tests are shown in Fig. 5.41. The plots
show that the arc power increases at an increasing rate to a point where the arc stalls
(d(ui)/dt = 0) and instability and collapse follow. The condition for arc instability is discussed
in the subsection 5.3D.
98
Loop switching
Powertech Februari 2000 and May 2000,
Xt = 20 Ω
KEMA disconnector Juli 2003,
Xs = 20 Ω
KEMA 6025
80
120
54 A
KEMA 6026
KEMA 6027
103 A
Februari 2000
70
97 A
KEMA 6028
100
53 A
60
52 A
103 A
Arc resistance [ Ω ] →
Arc resistance [Ω ] →
May 2000
80
75 A
47 A
71 A
74 A
60
50 A
50
40
30
23 A
40
54 A
54 A
24 A
20
20
10
0
0
100
200
300
400
500
600
Arcing time [ms] →
700
800
900
0
1000
0
200
400
600
800
Arcing time [ms] →
1000
350
300
46 A
Wilczek 2.5 kV
Vis & De Kleijn 2.5 kV
46 A
300
250
46 A
46 A
165 A
78 A
Arc resistance [ Ω ] →
Arc resistance [Ω ] →
46 A
250
122 A
39 A
May 2000
116 A
165 A
200
122 A
40 A
86 A
150
1400
TUe, Xs = 56 Ω
Powertech Februari 2000 and May 2000,
Xt = 60 Ω
Februari 2000
1200
159 A
82 A
47 A
200
45 A
46 46
A A
46 A
150
46 A
46 A
100
45 A
46 A
47 A 46 A
46 A
100
50
50
0
0
500
1000
1500
Arcing time [ms] →
2000
2500
0
0
500
1000
1500
Arcing time [ms] →
Fig. 5.40 Arc resistance for comparable PT, KEMA and TUe tests
99
Section 5
Powertech Februari 2000 and May 2000,
Xt = 20 Ω
KEMA disconnector Juli 2003,
Xs = 20 Ω
140
KEMA 6025
Februari 2000
35
KEMA 6026
120
KEMA 6027
May 2000
30
KEMA 6028
100
54 A
53 A
Arc Power [kW] →
Arc Power [kW] →
25
103 A
80
103 A
97 A
60
74 A
75 A
54 A
54 A
20
15
71 A
40
10
50 A
20
47 A
52 A
5
0
24 A
23 A
0
100
200
300
400
500
600
Arcing time [ms] →
700
800
900
1000
0
0
200
600
800
Arcing time [ms] →
1000
1200
Wilczek 2.5 kV
70
Februari 2000
May 2000
Vis & De Kleijn 2.5 kV
800
46 A
60
46 A
46 A
700
Arc Power [kW] →
Arc Power [kW] →
600
165 A
500
159 A
165 A
400
300
46 A
46AA
47 A
46
46 A
50
46 A
45 A
45 A
46 A47 A
46 A
46 A
40
46 A
46 A
30
122 A
20
116 A
200
122 A
86 A
82 A
10
78 A
100
40 A
0
1400
TUe, Xs = 56 Ω
Powertech Februari 2000 and May 2000,
Xt = 60 Ω
900
400
0
500
39 A
1000
1500
Arcing time [ms] →
2000
2500
0
0
500
1000
Arcing time [ms] →
Fig. 5.41 Arc power for comparable PT, KEMA and TUe tests
100
1500
Loop switching
Powertech May 2000, Xt = 100 Ω
Powertech May 2000, Xt = 100 Ω
1000
500
0
23 A
44 A
-1000
400
d/dt Arc Power [W/s] →
-2000
Arc Power [kW] →
96 A
300
95 A
200
-3000
49 A
70 A
-4000
96 A
-5000
70 A
-6000
70 A
70 A
-7000
44 A
100
49 A
-8000
95 A
-9000
23 A
0
0
200
400
600
800
1000
Arcing time [ms] →
1200
1400
1600
0
200
400
600
800
1000
Arcing time [ms] →
1200
1400
1600
Fig. 5.42 Arc power and rate of change of arc power d(ui)/dt
Fig. 5.42 shows the arc power and rate of change of arc power for the PT tests at 100 ohms
loop impedance. The latter quantity has a waterfall pattern and is initially greater than zero as
it has to be for the arc to propagate. Arc propagation ceases when the rate of change equals
zero and instability and collapse eventually follow.
Another variation is to consider the maximum arc power and the initial current with constant
loop impedance as is done in Fig. 5.43 for a number of PT tests. There is a well fitting quadratic relationship in evidence and the question now is whether an equally good fit exists
between comparable tests at PT and TUe. The answer to this is affirmative (Fig. 5.44) and the
conclusion is that the maximum arc power is determined by the initial current and the loop
impedance and is apparently independent of the disconnector type and source type (voltage or
current).
Powertech Februari 2000 and May 2000
1100
1000
Xt = 60 Ω
900
Maximum Arc Power [kW] →
800
Xt = 70 Ω
700
600
Xt = 100 Ω
500
400
Xt = 30 Ω
300
Xt = 40 Ω
200
Xt = 20 Ω
100
0
0
20
40
60
80
100
Initial Current [A] →
120
140
160
180
Fig. 5.43 Maximum arc power versus initial current with
loop impedance as parameter
101
Section 5
Powertech & Eindhoven
Powertech & Eindhoven
350
350
TU/e Xs = 35 Ω
300
Powertech Xt = 30 Ω
Maximum Arc Power [kW] →
Maximum Arc Power [kW] →
400
250
200
150
300
TU/e Xs = 46 Ω
250
Powertech Xt = 40 Ω
200
150
100
100
50
50
0
0
50
100
0
150
0
20
40
60
Initial Current [A] →
Initial Current [A] →
Powertech & Eindhoven
100
120
Powertech & Eindhoven
1200
900
800
TU/e Xs = 75 Ω
700
Powertech Xt = 70 Ω
TU/e Xs = 56 Ω
1000
Powertech Xt = 60 Ω
800
Maximum Arc Power [kW] →
Maximum Arc Power [kW] →
80
600
400
600
500
400
300
200
200
100
0
0
20
40
60
80
100
Initial Current [A] →
120
140
160
180
0
0
20
40
60
80
100
Initial Current [A] →
120
140
Fig. 5.44 Maximum arc power versus initial current with loop impedance as parameter for
comparable PT and TUe tests
102
160
Loop switching
D. Arc d(ui)/dt characteristics
In subsection 5.2.2B it is noted that the luminosity of the arc is sustained until just prior to arc
extinction. Ignoring the effect of partial arc collapse, this requires an increasing rate of
change of power input particularly since the arc is also continuing to grow in length. This
condition can be expressed as (Eqn. (5.12)):
d(ui)/dt > 0
which in turn can be expanded to
u
di
du
+i
>0
dt
dt
or
u du / dt
+
>0
i di / dt
…(5.14)
But u/i is the arc resistance r and for the same time increment, we can also write
r + du/di > 0
…(5.15)
Eqn. (5.15) is of the same form as describes the condition for arc stability in a DC circuit as
(apparently) first calculated by Kaufmann.63 Using the original notation, Kaufmann’s equation reads:
W+
∂E
>0
∂J
…(5.16)
where W is the resistance of the circuit external to the arc and J is the arc current.
Eqns. (5.14) and (5.15) relate only to the arc and we must now consider their implications.
Rewriting Eqn. (5.14) in a new form:
r/du/dt/di/dt > -1
…(5.17)
we now have the condition or index for arc stability.
With reference to U-I characteristics of Figs. 5.36 and 5.37, r is the instantaneous arc resistance and du/dt/di/dt (or du/di) is the rate of change of arc resistance. Initially, in the period
immediately after contact parting, r will have a low value while du/di has a high value and the
index will be close to zero. As the arc grows in length, r will increase and du/di will decrease
and the index starts to go negative. When the index goes below -1, the arc becomes unstable
and arc collapse may follow unless there is an intervention to restore its stability. A partial
arc collapse would be such an intervention.
103
Section 5
The above reasoning is supported and illustrated by considering the Powertech test cases
shown in Figs. 5.19 (test no. 6_15) and 5.20 (test no. 8_20). Test no. 6_15 is for an initial
current of 62 A and a loop impedance of 150 ohms and is shown in Fig. 5.45. Test no. 8_20 is
for an initial current of 143 A and a loop impedance of 70 ohms and is shown in Fig. 5.46.
The plots from top to bottom in Figs. 5.45 and 5.46 show the arc power, the rate of change
arc power and lastly the stability index, all versus time. Video images from test no. 6_15 are
shown in Figs. C9 to C20 (note that time zero for the figures occurs after time zero on the
traces).
For test no. 6_15, the arc becomes unstable beyond recovery at 1100 ms which time corresponds to the image in Fig. C19. Prior to this, the rate of change of arc power becomes
negative between 600 and 700 ms and between 800 and 900 ms, both indicating a certain arc
instability. The stability index also indicates a tendency towards instability by dipping below
zero. Recovery after the first period is due to a partial arc collapse at about 700 ms (Fig.
C15). This is followed by a rare event at 800 ms: a re-establishment of the previous longer
arc path (Fig. C16) giving an increase in arc resistance. A partial arc collapse follows immediately but is not apparent on the video images.
The test no. 8_20 plots reflect the ripple effect evident on the disconnector current trace (Fig.
5.20). The stability index plot shows a series of potential arc instabilities until the ultimate
instability and collapse occurs at about 1700 ms.
The plots in Figs. 5.45 and 5.46 cannot be used to make any quantitative assessment of when
total arc collapse will follow arc instability. Depending on the degree of smoothing utilized,
the plot appearances will vary. This is illustrated in Fig. 5.47. A seemingly minor event in
one plot can appear as a major event in the other.
E. Discussion on arc modelling
The intent of developing an arc model is to provide a means of calculating the arcing time
given an initial current and the loop impedance. In its simplest form, the model could be a
time varying resistance replacing the disconnector in Fig. 5.1. The typical variations of arc
resistance are as shown in Fig. 5.40: there is an initial gradually rising part followed by an
exponential increase at arc instability and collapse. For shorter arcing times, the resistance r
can be expressed as:
r=
1
a + b ln t
…(5.18)
were t is time in seconds and a and b are constants. For the longer arcing times, Fig. 5.40
shows that gradually rising period can be greatly prolonged and suggests that a time shift
should be introduced into Eqn. (5.18) in some form or other.
104
Loop switching
Powertech May 2000, test: 6_15, Xt = 150 Ω, Initial current = 62 A
No smoothing
3 x Mean over 5 points
Arc power [kW] →
300
250
200
150
100
50
0
200
400
600
Arcing time [ms] →
800
1000
d/dt Arc power [W/ms] →
4000
400
d/dt Arc Power
Arc resistance
3000
2000
300
1000
0
200
-1000
-2000
100
Arc resistance [Ω ] →
0
-3000
-4000
0
200
400
600
Arcing time [ms] →
800
0
1000
Powertech May 2000, test: 6_15, Xt = 150 Ω, Initial current = 62 A
8
No smoothing
3 x Mean over 3 points
6
End of arcing time
Intersection with y = -1
4
rarc ⋅(dîarc /dt)/(dûarc /dt) →
2
0
Line at y = -1
-2
-4
-6
-8
Arcing time: 1157 ms
-10
-12
0
200
∆ t: -63 ms
400
600
800
Arcing time [ms] →
1000
1200
Fig. 5.45 PT test no. 6_15
105
Section 5
No smoothing
3 x Mean over 5 points
600
400
200
0
0
200
400
600
800
1000
Arcing time [ms] →
1200
1400
1600
d/dt Arc power [W/ms] →
4000
200
d/dt Arc Power
Arc resistance
3000
2000
150
1000
0
100
-1000
-2000
50
-3000
-4000
0
200
400
600
800
1000
Arcing time [ms] →
1200
1400
1600
Powertech May 2000, test: 8_20, Xt = 70 Ω, Initial current = 143 A
8
No smoothing
3 x Mean over 3 points
6
End of arcing time
Intersection with y = -1
4
rarc ⋅(dîarc /dt)/(dûarc /dt) →
2
0
Line at y = -1
-2
-4
-6
-8
Arcing time: 1774 ms
-10
-12
0
200
400
∆ t: -46 ms
600
800
1000
1200
Arcing time [ms] →
Fig. 5.46 PT test no. 8_20
106
1400
1600
1800
0
Arc resistance [Ω ] →
Arc power [kW] →
Powertech May 2000, test: 8_20, Xt = 70 Ω, Initial current = 143 A
Loop switching
Powertech May 2000, test: 8_26, Xt = 60 Ω, Initial current = 168 A
8
No smoothing
3 x Mean over 3 points
6
End of arcing time
Intersection with y = -1
4
rarc ⋅(dîarc /dt)/(dûarc /dt) →
2
0
Line at y = -1
-2
-4
-6
-8
Arcing time: 1550 ms
-10
-12
0
200
400
∆ t: -121 ms
600
800
1000
Arcing time [ms] →
1200
1400
1600
Powertech May 2000, test: 8_26, Xt = 60 Ω, Initial current = 168 A
8
No smoothing
3 x Mean over 3 points
6
End of arcing time
Intersection with y = -1
4
rarc ⋅(dîarc /dt)/(dûarc /dt) →
2
0
Line at y = -1
-2
-4
-6
-8
Arcing time: 1550 ms
-10
-12
0
200
400
∆t: -296 ms
600
800
1000
Arcing time [ms] →
1200
1400
1600
Fig. 5.47 PT test no. 8_26 with different smoothing factors:
5 for the upper plot and 7 for the lower plot
107
Section 5
Using Eqn. (5.18) and assuming the applicable constants are known, it is possible to perform
an iterative calculation giving the arcing time when the disconnector current goes to zero.
However, generalization of this approach is difficult because the values of the constants will
vary for different initial currents and loop impedances and even for the same current and loop
impedance (for example no one equation will cover all of the cases shown in Fig. 5.40).
In the literature, the modelling of long arcs in air has been reported in connection with secondary arcs.64, 65 Both studies used a basic arc equation of the form:
dg 1
= (G − g )
dt τ
…(5.19)
where g is the time varying arc conductance, G is the stationary arc conductance and τ is a
time constant. G is defined differently in each study with the evolving arc length being a
common element. A cursory test of the validity of this approach to the loop switching case
shows that it does give a reasonable commutation process representation. However, the
derived arc voltages and currents do not resemble those of the loop switching case.
A difficulty with the modelling overall is the need to know how the arc length and resistance
evolve with time. The convoluted nature of the loop switching arc, particularly at the higher
current values, is evident in the video images in Annex C. Determination of arc length with
any degree of accuracy from the images is judged to be close to impossible. The arc resistances, however, exhibit a repetitive shape and this can most probably be exploited. The
challenge is to find an arc resistance relationship for varying current with constant loop
impedance and then between the various loop impedances. A data mining approach may offer
the best opportunity to achieve this.
To have a totally realistic arc model, the partial arc collapse events also need to be incorporated. The dynamic U-I characteristics for Powertech test nos. 6_16 and 8_20 are shown in
Figs. 5.48 and 5.49, respectively (for clarity the initial first quarter of arcing has been omitted). Despite having a different basis, there is a certain orderliness to them, not the least
because they have the laws of electricity in common. However, the loop switching arcs would
appear to meet the definition of a chaotic system in that they are (1) deterministic through
description by mathematical rules and (2) have mathematical descriptors that are non-linear
in some way. Treating the arc as a chaotic system has in fact been used with success for arcs
in electrical arc furnaces.68–71 The issue in that case is a determination of the fluctuations in
the arcs and consequent voltage flicker. The U-I characteristics of such arcs vary over a
certain range due to the fluctuations, while the loop switching U-I characteristics vary over a
wider range due to commutation of the current and the fluctuations caused by partial arc
collapse. The conclusion is that this approach can be extended to the loop switching case.
108
Loop switching
Powertech May 2000, test: 6_15, Xt = 150 Ω
15
10
Voltage [kV] →
5
0
-5
-10
-15
-100
-80
-60
-40
-20
0
Current [A] →
20
40
60
80
100
Fig. 5.48 Dynamic U-I characteristic for Powertech test no. 6_15
Powertech May 2000, test: 8_20, Xt = 70 Ω
15
10
Voltage [kV] →
5
0
-5
-10
-15
-250
-200
-150
-100
-50
0
Current [A] →
50
100
150
200
250
Fig. 5.49 Dynamic U-I characteristic for Powertech test no. 8_20
109
Section 5
5.4
Conclusions
Loop switching relies on a natural commutation of current from one circuit to a parallel
circuit. Despite the apparent randomness of the arc certain patterns of behaviour and relationships can be found between the initial current in the disconnector, the loop impedance, arcing
time and arc reach. This suggests that the arc can better be described as one of ordered chaos.
Repeat tests on different occasions months apart and as part of the same test series showed a
reasonable consistency enabling a statistical evaluation of the above-noted quantities.
In order to determine and understand the condition that must be satisfied for the arc to extinguish, the arc voltage and current data from the various tests at three locations was examined.
Arc collapse follows a period of arc instability caused by a change in the rate of change of
power input from positive to negative. The actual condition for arc instability has the form as
that for an arc in a DC circuit. This is not unexpected because the arc is sustained by its
power input and the source of the power, whether DC or AC, is not an issue. To put this in
perspective, the arc can be viewed as operating as a series of static arcs of increasing length
as shown in Fig. 5.50. The actual, or dynamic, U-I characteristic is superimposed (refer to
Figs. 5.36 to 5.38) and the arc will thus operate at the successive connection points a, b, … l.
If the arc were to stall at connection point g for example, then the arc would be stable and
sustained. The arc however continues to evolve and eventually reaches a length such that the
two characteristics become tangential (connection point l) and ceases to evolve further
because the power input is now constant or decreasing. Arc instability and collapse follow.
Fig. 5.50 Static and dynamic arc characteristics
110
Loop switching
Further conclusions can be drawn with respect to the U-I characteristics of Figs. 5.36 to 5.39.
Considering Eqn. (5.13), at I = 0 (arc extinction) the voltage U = IS (XS + XL) giving the
value for C; at U = 0 and I = IS (arc initiation), we can write:
AIS2 + BIS + C = 0
…(5.20)
Since C is already known, A and B can be related by solving Eqn. (5.20) resulting in only one
unknown.
Modelling of the arc for this switching event presents a significant challenge. The recommended approach is to represent the arc as a time-varying resistance. This requires further
analysis of the test data and the data mining technique is suggested as the tool for the
purpose. A further approach worth consideration is to extend the arc furnace chaotic system
development to the loop switching case.
Lastly, this research work has provided ample evidence that the assumptions and test methods
used by Andrews et al5 do not stand up to scrutiny. That work should now be viewed as
highly questionable.
111
Section 6
Discussion on use of other disconnector types
Four types of disconnectors are described in section 1 each with its own contact design features and blade movement. Viewed in terms of current interruption, the vertical break and
pantograph types both have a blade movement in a vertical plane and centre break and double
break in a horizontal plane. Therefore, certain similarities in current interruption performance
and arc propagation can be drawn.
One immediate distinction can be made between the vertical break type disconnector and the
other three disconnector types. To make this distinction only the influence of blade movement on the arc is considered:
•
For the vertical break type (Fig. 1.1), the blade movement draws the arc upwards in a
plane along the longitudinal axis of the disconnector. No excursion of the arc into the
space between phases can be attributed to blade movement.
•
For the centre break and double break type disconnectors (Figs. 1.2 and 1.3, respectively), the blade movement draws the arc outwards in a horizontal plane. For the
centre-break disconnector, the blade movement is in the same direction on all three
phases and the arcs are drawn towards the adjacent phase on two of the three poles but
in no case towards one another. For the double break disconnector, the blade movement is in both directions and there is again excursion of the arc horizontally between
phases and, more significantly towards one another, i.e. from the outer phases towards
the centre phase.
•
For the pantograph type (Fig. 1.4), the blade movement draws the arc downward in a
vertical plane at right angles to the axis of the busbars. Due to the scissors effect, the
arc may be drawn on the side towards an adjacent phase and the arcs of adjacent
phases may be drawn towards one another.
For unloaded transformer magnetizing currents of 1 A or less, interruption is dependent on
achieving the required contact gap spacing as discussed in sections 3 and 4, respectively.F6
All four disconnector types can be viewed as being essentially equal in terms of interrupting
capability at these current levels and arc reach is not considered to be an issue. For currents
greater than 1 A, thermal effects will cause the arc to rise vertically and reaching of the arc
will be evident. The vertical break and pantograph type disconnectors are judged to have
similar performance and capability at these current levels. The same can be said for centre
break and double break type disconnectors. Prolonged arcing can be expected and caution
should be exercised in this regard. Additionally, the influence of inrush current should be
recognized.
F6
Refer also to the case study in Annex B.
113
Section 6
For capacitive currents in this regard, the conclusions are similar to those for the above-noted
transformer magnetizing current case. Generally for the longer arc durations, particularly for
currents greater than 1 A, there will be an upward motion of the arc – refer to Figs. 4.5, 4.7,
4.8 and 4.14 (upper image).
The Powertech and KEMA loop switching tests were performed on a vertical break type disconnector, while those at Eindhoven used a centre break type disconnector. The similarities
and differences in the electrical characteristics of the arcs have been discussed in subsection
5.3. The consideration here will deal with arcing time, flame type and the conclusions that
can be drawn.
For the same range current and loop impedance, the centre break disconnector exhibited a
longer arcing time as shown in Fig. 6.1. The Eindhoven results show an obvious decreasing
tendency at the higher currents and a polynomial type regression was used. For the Powertech
1800
Arcing time (ms)
1600
TUe 32 to 170 ohms
1400
1200
Powertech 30 to 150 ohms
1000
800
Poly. (TUe 32 to 170
ohms)
600
Poly. (Powertech 30 to
150 ohms)
400
200
0
0.00
20.00
40.00
60.00
80.00
100.00
Initial current (A)
Fig. 6.1 Comparison of arcing times between Powertech and Eindhoven loop switching tests
results, the same type of regression was used for comparison purposes even though there is
no similar tendency; in fact, a linear regression shows a steady increase in arcing time with
increasing current. This difference in performance can be attributed to the differences in
build-up of arc resistance as noted in subsection 5.3B and as discussed below.
The flame associated with the Eindhoven test arcs as shown in Fig. 6.2 is quite different from
those of the Powertech tests (refer to Annex C). The flame is quite smooth even at the higher
current as compared to the turbulent flames shown in Annex C. The reasons for this difference most probably lie in the realm of arc physics but some relative performance related
conclusions can nonetheless be drawn.
114
Discussion on use of other disconnector types
Fig. 6.2 Typical arc images from the Eindhoven tests
For the vertical break and pantograph type disconnectors, the blade motion is such that it is
drawing out the arc in the direction of its natural propagation. The same cannot be said for the
centre and double break type disconnectors. This reasoning suggests that the influence of the
blade motion is more than might have been thought at first glance. Thus, for the horizontally
opening disconnectors, more onus is placed on the arc itself to achieve current interruption.
The conclusion is that vertical break and pantograph type disconnectors offer superior performance for loop switching than the centre break and double break types.
115
Section 7
General conclusions and suggestions for further research
High voltage air-break disconnectors have a certain but limited capability to interrupt
unloaded transformer magnetizing currents, capacitive currents and loop currents. Each
switching duty is shown to be clearly unique in its own way and thus requiring of individual
review and analysis.
Before dealing with the above-noted switching duties, it is appropriate to address the principal past work in this subject. That work was done by Andrews et al5 in the 1940s and has
largely been viewed as the definitive work on disconnector interrupting capability. It is
shown that the work was based on an invalid assumption which further influenced the test
circuits used. The conclusion is that this work should now be viewed as having questionable
value.
For switching unloaded transformer magnetizing currents of 1 A or less, interruption is
dependent on achieving a minimum contact gap spacing. At higher currents, usually not
exceeding 2 A for modern low-loss transformers, thermal effects become influential and the
required contact gap spacing for interruption becomes greater. The time to reach the interrupting contact gap spacing is a key factor and a higher blade opening speed or the addition
of whip type auxiliary interrupting devices should be considered. It is mandatory to also consider the effect that overvoltages due to restriking may have on the transformer and suitable
elimination or mitigation measures should be implemented. Inrush current, which can follow
as a result of a restrike, does not contribute to the propagation of the arc but rather prolongs
the arcing times by its duration.
For capacitive current switching, successful interruption is more complex being dependent on
a minimum contact gap spacing, the magnitude of the current and the ratio of the source (CS)
to load (CL) side capacitances. The worst cases for arc duration and overvoltage generation
occur when CS/CS < 1. The prudent approach is to limit the current to be interrupted to those
suggested by the standards, i.e. 0.5 A or less.1, 62 Higher currents up to 2 A can be interrupted
(based principally on field experience) provided that it is understood and accepted by the user
that the disconnector may be close to the fully open position at current interruption. Whip
type auxiliary interrupting devices offer advantage and are judged to have an interrupting
capability not exceeding 7 A. The so-called 600 kVA rule is often cited in North America for
this switching case: any merit that this rule may have is judged to be coincidental rather than
factual.
The major part of this work addresses the loop switching case. This case, in contrast to the
two cases of low currents and high recovery voltages discussed above, is one of high currents
up to several hundred amperes and low recovery voltages. Empirical relationships can be
found between the initial current in the disconnector, the loop impedance, the arcing time and
the arc reach. On a theoretical basis, it is shown that the condition to be satisfied for arc stability in this AC case is similar to that for an arc in a DC circuit. For purposes of arc model-
117
Section 7
ling and given the difficulty in defining arc length, an arc resistance versus time is the
suggested approach. Further research, however, is required in this regard.
Four different types of disconnector – vertical break, centre break, double break and pantograph – are commonly used in high voltage substations. By virtue of its blade motion, the
vertical break type is judged to be the most suitable for current interruption.
As to further research, this should be directed towards the loop switching case as follows:
•
To investigate the influence of the electrode material for the case of very high currents
up to 2000 A with loop impedances of 1 ohm or less. There is evidence to suggest that
refractory material such as tungsten will offer advantage over cold cathode material
such as copper and aluminum.
•
To investigate the initial point of arc formation and its influence on the further direction of propagation of the arc. The Powertech tests showed evidence the arc tended to
propagate in a direction away from the side of the disconnector where the arc was first
formed. This suggests that control over the direction of arc propagation may be possible by judicious placement of the arcing horns.
•
To investigate further the development of an arc model. The suggested approaches are
that of treating the arc as a time-varying resistance and/or viewing the arc as an
ordered chaotic system. Some limitation on the range of current and loop impedance
to be addressed would be beneficial and utilities should be consulted in this regard.
118
Section 8
References
1.
IEC 62271-102 Alternating current disconnectors and earthing switches – First edition
2001-12.
2.
IEEE Std C37.100-1992 IEEE Standard Definition for Power Switchgear.
3.
ANSI Std C37.30-1970/IEEE Std 324-1971 Definitions and Requirements for HighVoltage Air Switches, Insulators and Bus Supports. Revised in 1992 and 1997 to
become IEEE Std C37.30-1997 IEEE Standard Requirements for High-Voltage
Switches.
4.
IEEE C37.36b-1990 IEEE Guide to Current Interruption with Horn-Gap Air Switches.
5.
F.E. Andrews, L.R. Janes and M.A. Anderson, “Interrupting Ability of Horn-Gap
Switches,” AIEE Transactions, Vol. 69, 1950.
6.
D.F. Peelo, “Current Interrupting Capability of Air Break Disconnect Switches,” IEEE
Transactions on Power Delivery, Vol. PWRD-1, No. 1, January 1986 and Correction,
IEEE Transactions on Power Delivery, Vol. PWRD-2, No. 4, October 1987.
7.
AIEE Committee Report, “Report on Transformer Magnetizing Current and Its Effect
on Relaying and Air Switch Operation,” AIEE Transactions, Vol. 70, 1951.
8.
IEEE Committee Report, “Results of Survey on Interrupting Ability of Air Break
Switches,” IEEE Transactions on Power Apparatus and Systems, Vol. PAS-85, No. 9,
September 1996.
9.
CEA Project 069 T 102 Report, The Interrupting Capability of High Voltage Disconnect Switches, July 1982.
10.
P.A. Abetti, “Arc Interruption with Disconnecting Switches,” Master Thesis, Illinois
Institute of Technology, January 1948.
11.
J.E. Gerngross, “Arcs on Horn-Gap Switches,” Master Thesis, Illinois Institute of Technology, January 1949.
12.
P. Ackermann, “A Study of Transmission Line Power-Arcs,” The Engineering Journal,
Vol. XI, No. 5, May 1928.
13.
C.S. Beattie and S.C. Killian, Discussion of reference 5, AIEE Transactions, Vol. 69,
1950.
14.
Bulletin 2.3.1-1B, I-T-E Outdoor Air Switches Selection and Application, 8.25 kV
three 800 kV Maximum Voltage. (Date unknown.)
15.
M.W. Anderson, “Power Autotransformer Current Interruption with an Air-Break
Disconnect Switch,” AIEE Transactions, Vol. 74, 1955.
119
Section 8
16.
G.F. Farmer, “A Phenomenon of Air Break Switching of Magnetizing Current,”
Canadian Electrical Association Fall Meeting, Halifax, November 1975.
17.
E.L. Luehring and J.P. Fitzgerald,” Switching the Magnetizing Current of Large
345 kV Transformers with Double-Break Air Switches,” IEEE Transactions on Power
Apparatus and Systems, Vol. PAS-84, No. 10, October 1965.
18.
A.H. Powell, “Dielectric and other problems in the design of a new 330 kV outdoor air
switch,” AIEE Transactions, Vol. 73, February 1954.
19.
D.F. Peelo, “Unloaded transformers can be switched reliably with disconnect
switches,” Transmission & Distribution, July 1987.
20.
P.E. Richardson and A. Foti, “Gas-Blast Switch Tests on 230-kV System,” Electrical
Engineering, August 1956.
21.
A. Foti and J.M. Lakas, “EHV Switch Tests and Switching Surges,” IEEE Transactions
on Power Apparatus and Systems, Vol. 83, No. 3, March 1964.
22.
D.L. Lott, “Surge-Suppressing Resistors Applied to EHV Disconnect Switches,” AllisChalmers Engineering Review. (Reprint date unknown.)
23.
E.C. Rankin, “Experience with Methods of Extending the Capability of High-Voltage
Air Break Switches,” AIEE Transactions, Vol. 79, February 1960.
24.
F.S. Toomer, “Switches of 115 kV (Lines) Updated as Capacity,” Electrical World,
March 1965.
25.
S.G. Patel, W.F. Holcombe and D.E. Parr, “Application of Air-Break Switches for
De-energizing Transmission Lines,” IEEE Transactions on Power Apparatus and
Systems, Vol. 4, No. 1, January 1987.
26.
G.E. Buchanan, Discussion of reference 23, AIEE Transactions, Vol. 79, February
1960.
27.
I.W. Gross, S.C. Killian and J.M. Sheadel, “A 330-kV Air Switch,” AIEE Transactions,
Vol. 73, April 1954.
28.
H.L. Rorden, J.M. Dils, S.B. Griscom, J.W. Skooglund and E. Beck, “Investigations of
Switching Surges Caused by 345-kV Disconnecting-Switch Operation,” AIEE Transactions, Vol. 77, October 1958.
29.
M.B. McNulty, “Guide for Interrupting Loop Circuit,” Transmission & Distribution,
February 1968.
30.
K.R. Shah, “Switching Severity of Horn-Gap Switches,” IEEE Transactions on Power
Apparatus and Systems, Vol. 91, July/August 1972.
120
References
31.
R.H. Harner, Discussion of reference 31, IEEE Transaction on Power Apparatus and
Systems, Vol. 91, July/August 1972.
32.
J.J. Keane, R.G. Andrei and B.R. Halley, “American Electric Power Experience with
Loop Switching in EHV Stations,” IEEE Transactions on Power Delivery, Vol. 6,
No. 4, October 1991.
33.
IEEE Std-4, 1995, IEEE Standard Techniques for High-Voltage Testing.
34.
D.W. Durbak, S.R. Lambert and R.J. Ringlee, “Ferroresonance Can Damage
Transformers During Air Break Switch Operation,” Power Technology, Issue No. 63,
October 1990.
35.
R.J. Carpentier, D.W. Durbak, S.R. Lambert and R.J. Ringlee, “FerroresonanceInduced Failures in High Voltage Transformers by Air Break Switch Operations,” 1990
Annual International Conference of Doble Clients.
36.
L. van der Sluis, “Transients in Power Systems,” John Wiley & Sons Ltd., 2001.
37.
L.F. Blume, G. Camilli, S.B. Farnham and H.A. Peterson, “Transformer Magnetizing
Inrush Currents and Influence on System Operation,” AIEE Transactions, Vol. 63,
1944.
38.
T.R. Specht, “Transformer Magnetizing Inrush Current,” AIEE Transactions, Vol. 70,
1951.
39.
U. Krusi, K. Froehlich and J.H. Brunke, “Controlled Transformer Energization
Considering Residual Flux – Implementation and Experimental Results.” Proceedings
of IASTED International Conference of Power and Energy Systems, Tampa, Florida,
Paper No. 328-143, November 2001.
40.
P. Sporn and D.C. Prince, “Ultrahigh-Speed Reclosing of High-Voltage Transmission
Lines,” AIEE Transactions, Vol. 56, January 1937.
41.
A.C. Boisseau, B.W. Wyman and W.F. Skeats, “Insulator Flashover Deionization
Times as a Factor in Applying High-Speed Reclosing Circuit Breakers,” AIEE Transactions, Vol. 68, 1949.
42.
K. Anjo, H. Terase and Y. Kawaguchi, “Self-Extinction of Arcs Created in Long Air
Gaps,” Electrical Engineering in Japan, Vol. 88, No. 4, 1968.
43.
G.T. Tretjak, V.V. Kaplan and E.I. Kender, “Free Burning Long Power Arcs in Air,”
CIGRE Report No. 324, 1935.
44.
A.P. Strom, “Long 60-Cycle Arcs in Air,” AIEE Transactions, Vol. 65, March 1946.
121
Section 8
45.
H. Edels, “Properties and Theory of the Electric Arc.” Proceeding IEEE, February
1961.
46.
T.E. Browne, “The Electric Arc as a Circuit Element.” Journal of the Electrochemical
Society, January 1955.
47.
H. Ayrton, “The Electric Arc,” The Electrician, London 1902.
48.
C.P. Steinmetz, “Transformation of Electric Power into Light,” Transactions AIEE,
Vol. 25, 1907.
49.
W.B. Nottingham, “A New Equation for the Static Characteristic of the Normal Electric Arc,” Transaction AIEE, Vol. 42, February 1923.
50.
A.R. van C. Warrington, “Reactance Relays Negligibly Affected by Arc Impedance,”
Electrical World, September 19, 1931.
51.
Relay Systems (textbook by Monseth and Robinson), McGraw-Hill Book Company,
New York, 1935.
52.
E.T.B. Gross, “Free Burning Long Power Arcs at High Voltage,” Schweizer Arkiv fűr
Angewandte Wissenschaft und Technik, Vol. 7, 1941.
53.
J.R. Eaton, J.K. Peck and J.M. Dunham, “Experimental Studies of Arcing Faults on a
75 kV Transmission System,” Transaction AIEE, December 1931.
54.
L. Edwards, J.W. Chadwick, H.A. Reich and L.E. Smith, “Single-Pole Switching on
TVA’s Paradise-Davidson 500 kV Line: Design Concepts and Staged Fault Tests,”
IEEE Transactions, PAS-90, No. 6, 1971.
55.
R.M. Hasibar, A.C. Legate, J. Brunke and W.G. Peterson, “The Application of HighSpeed Grounding Switching for Single-Pole Reclosing on 500 kV Power Systems,”
IEEE Transactions, PAS, April 1981.
56.
J.G. Kappenman, G.A. Sweezy, V. Koschik and K.K. Mustaphi, “Staged Fault Tests
with Single Phase Reclosing on the Winnipeg-Twin Cities 500 kV Interconnection,”
IEEE Transactions, PAS-101, No. 3, 1982.
57.
A.J. Fakheri, T.C. Shuter, J.M. Schneider and C.H. Shih, “Single Phase Switching
Tests on the AEP 765 kV System – Extinction Time for Large Secondary Arc
Currents,” IEEE Transactions, PAS-102, No. 8, 1983.
58.
R.M. Hasibar and C.W. Taylor, Discussion of reference 57, IEEE Transactions,
PAS-102, No. 8, 1983.
59.
A.S. Maikopar, “Extinction of an open electric arc,” Elektrichestvo 1960.
122
References
60.
W. Rieder, “Plasma and Lichtbogen.” (Book) Friedr. Vieweg & Sohn, Braunschweig
1967.
61.
ANSI Std C37.32 – 1996 High Voltage Air Disconnect Switches, Interrupter Switches,
Fault Initiating Switches, Grounding Switches, Bus Supports and Accessories, Control
Voltage Ranges – Schedule of Preferred Ratings, Construction Guidelines and
Specifications.
62.
C. Neumann, “Nichtstandardisierte betriebliche Beanspruchungen bein Shalten von
Trenn-und Erdungsschaltern in Hochspannangsnetz” (Non-standard operational stresses
when switching disconnectors and earthing switches in high voltage networks). PhD
thesis, Technischen Hochschule Darmstadt 1992.
63.
W. Kaufmann, “Elektrodynamische Eigentümlichkeiten leitender Gase” (Electrodynamic peculiarities of conducting gas). Ann. d. Phys., 2 1900.
64.
M. Kizilcay and T. Pniok, “Digital Simulation of Fault Arcs in Power Systems,” ETEP
Vol. 1, No. 1, January/February 1991.
65.
A.T. Johns, R.K. Aggarwal and Y.H. Song, “Improved techniques for modelling fault
arcs on faulted EHV transmission systems,” IEE Proceedings – Generation, Transmission and Distribution, Vol. 141, No. 2, March 1994.
66.
H. Knobloch, “Switching of Capacitive Currents by Outdoor Disconnectors,” Fifth
International Symposium on High Voltage Engineering, Braunschweig, August 1987.
67.
H. Knobloch: Schaltern kapazitiver Ströme mit Freilufttrennschaltern. (Switching
capacitive currents with air-break disconnectors.) Internal report Siemens Schaltwerk
Berlin, TVH-VB2716, 1986.
68.
E. Acha, A. Semlyen and N. Rajakovic, “A Harmonic Domain Computational Package
for Non-Linear Problems and Its Application to Electric Arcs,” IEEE Transactions on
Power Delivery, Vol. 5, No. 3, July 1990.
69.
H-D. Chiang, C-W. Liu, P.P. Varaiya, F.F. Wu and M.G. Lauby, “Chaos in a Simple
Power System,” IEEE Transactions on Power Systems, Vol. 8, No. 4, November 1993.
70.
T. Zheng and E.B. Makram, “An Adaptive Arc Furnace Model,” IEEE Transactions on
Power Delivery, Vol. 15, No. 3, July 2000.
71.
O. Ozgun and A. Abur, “Flicker Study Using a Novel Arc Furnace Model,” IEEE
Transactions on Power Delivery, Vol. 17, No. 4, October 2002.
72.
D.F. Peelo, R.P.P. Smeets, L. van der Sluis, S. Kuivenhoven, J.G. Krone, J.H. Sawada
and B.R. Sunga, “Current Interruption with High Voltage Air-Break Disconnectors.”
To be published: Cigre 2004 session.
123
Annex A
Transformer transient recovery voltages
A1. Introduction
Oscillographic traces from actual unloaded transformer field tests show that the transformer
transient recovery voltage is a damped low frequency oscillation.15, 19, 22 Two such traces are
shown in Fig. 3.7 and Fig. B5. It is known that an unloaded transformer can be represented
by a capacitance in parallel with a series RL circuit. However, because of the non-linearity of
the core, no one representation will cover all levels of excitation.
A2. Analysis
The unloaded transformer representation can be derived from the factory core loss measurement test. In the test, the quantities recorded are: the applied voltage Uo line to ground, the
magnetizing current Io (in amperes rms) and the core loss Po. On a single-phase basis, we can
write:
Zo =
Uo
(cos θ + jsin θ)
Io
where cos θ =
Po
U o Io
Zo can now be expressed as:
Zo = Ro +jXo
= Ro + jωLo
The surge impedance Zs is given by:
Zs =
Lo
C
where C is chosen based on experience. The damping factor in turn is given by:
λ=
Zs
R0
since the R0, L0 and C form a series resonance circuit.
Lastly, the frequency of the oscillation is given by:
F=
1
2π L o C
125
Annex A
As will be demonstrated later, the values of Io and Po vary significantly with Uo (level of
excitation) and likewise the representation and its characteristics.
Transformer transient recovery voltages cannot be treated theoretically because “standard”
transformers do not exist. An empirical approach is necessary and this is done by considering
transformers on the BC Hydro system. While the number of transformers on the system is in
the many hundreds, the range of ratings at any system voltage level is limited but characteristics vary by vintage and manufacturer. A base sample of 230 kV transformers reflecting
rating, vintage (1950s to the present) and manufacturer was taken and the transient recovery
voltage characteristics were calculated as discussed above. To test the results, smaller samples of 500 kV, 138 kV and 60 kV transformers and the calculations performed. Finally, as a
totally independent source, the calculation was done for six transformers rated at 127 kV,
130 kV, 150 kV, 230 kV, 410 kV and 525 kV whose core loss details are given in a CIGRE
survey report.AF1 The results of the calculations are plotted in Fig. A1 and can be seen to fit
350
Frequency (Hz)
300
500 kV transformers
250
230 kV transformers
200
138 kV transformers
150
60 kV transformers
100
CIGRE survey
50
0
0
5
10
15
20
25
Damping factor (Lamda)
Fig. A1 Transformer transient recovery voltage characteristics at 100%
excitation
into a distinct pattern. The lowest damping factor values were 0.77 and 0.78 at 88 and 91 Hz,
respectively. All other values were 1 or greater. The calculations also show the following:
1. The damping factor and frequency tends to decrease with increasing system voltage as
shown in Figs. A2 and A3.
2. The damping factor and frequency tends to increase with the level of excitation as shown
in Figs. A4 and A5.
AF1
E. Colombo and G. Santagostino, “Results of the enquiries on actual network conditions when switching
magnetizing and small inductive currents and on transformer and shunt reactor saturation characteristics,”
Electra 94, May 1984.
126
Transformer transient recovery voltages
Damping factor (Lamda)
25
BC Hydro
transformer data
20
CIGRE survey data
15
Power (BC Hydro
transformer data)
10
Power (CIGRE
survey data)
5
0
0
200
400
600
System voltage (kV)
Fig. A2 Transformer transient recovery voltage damping factor versus
system voltage
350
Frequency (Hz)
300
BC Hydro
transformer data
250
CIGRE survey data
200
150
Power (BC Hydro
transformer data)
100
Power (CIGRE
survey data)
50
0
0
200
400
600
System voltage (kV)
Fig. A3 Transformer transient recovery voltage frequency versus system
voltage
80
90%
excitation
Damping factor (Lamda)
70
100%
excitation
60
50
110%
excitation
40
Linear (90%
excitation)
30
20
10
0
0
200
400
600
Linear
(100%
excitation)
Power
(110%
excitation)
System voltage (kV)
Fig. A4 Transformer transient recovery voltage damping factor versus
system voltage with level of excitation as parameter
127
Annex A
90%
excitation
450
400
100%
excitation
Frequency (Hz)
350
300
110%
excitation
250
200
Power (90%
excitation)
150
100
50
0
0
200
400
System voltage (kV)
600
Power
(100%
excitation)
Power
(110%
excitation)
Fig. A5 Transformer transient recovery voltage frequency versus system
voltage with level of excitation as parameter
A3. Discussion
Switching unloaded transformers has not attracted a great deal of interest. Field tests are
therefore quite rare and published papers on the subject rarer still. Of these papers, only a
limited few include actual transient recovery voltage measurements. However, some unpublished test data exists and is of value.
During the period 1977 to the present, BC Hydro carried out a number of switching tests on
unloaded 500 kV and 230 kV transformers. The oscillographic records from the tests all show
a damped oscillation virtually identical to that shown in Fig. 3.7(b) and Fig. B5. In no cases
was the underswing greater than 0.3 pu and in no case did the oscillation persist beyond the
underswing loop. Published traces by others show a similar result.15, 22 In addition private
access to unpublished test results from another major utility in North America also supports
the result. The 1.3 pu limit cited in subsection 3.2 is thus based on the best available field test
data.
As at least a first approximation Fig. A1 does demonstrate that the transformer transient
recovery voltage is a damped oscillation. The lowest damping factors are associated with the
lowest oscillation frequencies and comparison to generalized damping curves show that this
is the area where an underswing of the order 0.3 pu is possible. For the disconnector gap to be
actually stressed at 1.3 pu, the peak of the underswing must coincide with a peak of opposite
polarity of the source voltage. For underexcited transformers, conditions may occur where the
0.3 pu value can be exceeded, however the lower applied voltage will provide some compensation for the lower damping factor. Similarly the tendency to demand lower and lower core
losses may result in values exceeding 0.3 pu.
128
Annex B
EHV transformer switching case study
B1. Introduction
The switching of unloaded EHV transformers with disconnectors is a special case for a number of reasons. Firstly, there are no auxiliary interrupting devices suitable for use at this
voltage level and the disconnector must interrupt the magnetizing current through the repetitive break-restrike process. Secondly, restriking results in overvoltages and inrush current and
it is desirable that some form of control or limitation be exercised over these phenomena. In
the latter regard, it is understood that the transformer has a surge arrester connected at its high
voltage terminals.
The transformers to be switched are a bank of three single-phase generator transformers each
rated at 512.5 kV/13.8 kV, 105.3 MVA and connected wye/delta. The conditions to be met to
permit using a 500 kV vertical break disconnector for unloaded switching are:
1. The disconnector should interrupt the magnetizing current before the blade reaches the
45° angle position.
2. Occurrence of restriking should be minimized.
3. Magnitude of restriking overvoltages should not exceed a pre-determined value.
B2. Analysis
As shown already in Annex A, an unloaded transformer is a non-linear circuit. The magnetizing current can more than double as the excitation voltage goes from 90% to 100% or 100%
to 110%. For the low loss 500 kV transformers of today, the magnetizing current is unlikely
to exceed 2 A. For the above-noted transformers, the characteristics of the transformer transient recovery voltage (TRV) are (Table B1):
Table B1 Transformer TRV characteristics
Excitation
%
90
100
110
Damping
factor
0.93
1.07
9.92
Frequency
Hz
76
78
132
Using generalized damping curves,1 it can be shown that the maximum voltage across the
disconnector after current interruption will be less than or equal to 500 kV at power frequency as a first (and reasonable) approximation. With reference to Fig. 3.1, a gap spacing of
at least 1 m required corresponding to a blade angle of about 15° (refer to Fig. 3.2).
129
Annex B
The worst case restriking overvoltage will occur if the TRV peak coincides with the peak of
opposite polarity of the system voltage. This overvoltage (Um) is given by:
Um = Usm + β (voltage across switch)
where Usm is the system peak voltage to ground and β is the damping factor associated with
the restriking current. At 100% excitation, we can write:
Um ≤
512.5 2
3
+ β (500)
≤ 918 kV peak for β = 1 (no damping)
≤ 2.2 pu
Actual tests on similar transformers indicate that β is in the range 0.5 to 0.75. Using these
values, the expected value of Um is in the range 668 kV peak to 793 kV peak or 1.6 pu to
1.9 pu.
From the transformer’s perspective it is desirable that the restriking overvoltages do not
exceed a value given by the following:
SIL × 0.75 × 0.85
minimum insulation margin of 15%
transformer used test value
transformer as-new switching impulse level
giving 0.6375 × 1390 = 886 kV peak (2.12 pu in this case).
Natural damping in the restriking circuit would appear to limit the overvoltages to within the
desired limit. The influence of the (mandatory) surge arrester will provide further benefit. A
396 kV rated metal oxide arrester will limit the overvoltages to around 1.4 to 1.5 pu.
B3. Mitigation
The purpose of mitigation is to limit the incident of restriking and can be done by using the
most suitable disconnector type. The intent is to achieve a contact gap of 1 m in the shortest
time possible. This is shown in Table B2. In order of ranking, the best disconnector types to
use are: double break, pantograph, vertical break and (lastly) centre break.
Another option used by some utilities is to add insertion resistors as discussed in subsection
2.2.
130
EHV transformer switching case study
Table B2 500 kV single-pole disconnector type timing comparison
time* Time to 1 m contact gap
Disconnector type Opening
s
s
Vertical break
5–6
0.64 – 0.76
Double break
4–5
0.53 – 0.67
Centre break
4–5
1.6 – 2
Pantograph
3–4
0.53 – 0.77
* Blade in motion.
B4. Test
The results of an actual test confirmed the conclusions of the analysis and provided further
insight into the switching event. The transformer bank described in the analysis was energized and de-energized using a vertical break disconnector. The disconnector was three-pole
ganged operated and thus had a very slow opening time. In addition, the disconnector has a
long (rod-type) fixed arcing horn such that it takes the blade 3 to 4 seconds to clear the horn.
The net effect of the slow blade speed and the fixed horn is to promote prestriking and
restriking.
Measurements were made of the prestriking and restriking overvoltages at the transformer
terminals using bushing capacitive tap devices. The inrush currents were measured on the
neutral for all three single-phase transformer units. Video records were taken of each switching event. Fig. 3.6 is taken from these records.
The cumulative frequency of the prestriking overvoltages is shown in Fig. B1 and that of the
restriking overvoltages in Fig. B2. The two plots are very similar indicating that closing is as
severe as opening as far as overvoltage magnitude is concerned. Opening, however, resulted
in the greater number of overvoltages, on average over 20% more than for closing. This is the
expected result because the arcing time on opening is always longer than that during closing.
In both cases approximately 10% of the overvoltages are limited by the surge arrester, the
highest recorded overvoltage being 1.44 pu.
Cumulative frequency (%)
120
Positive polarity
100
Negative polarity
80
60
40
20
0
1.02-1.07
1.08-1.13
1.14-1.19
1.20-1.25
1.26-1.31
1.32-1.37
1.38-1.44
Overvoltage factor ranges (pu)
Fig. B1 Closing prestriking overvoltage factor distribution
131
Annex B
Cumulative frequency (%)
120
100
Positive polarity
80
Negative polarity
60
40
20
0
1.02-1.07 1.08-1.13 1.14-1.19 1.20-1.25 1.26-1.31 1.32-1.37 1.38-1.44
Overvoltage factor ranges (pu)
Fig. B2 Opening restriking overvoltage factor distribution
The distribution of the inrush currents are shown in Fig. B3. While values in excess of
3000 A are possible, the magnitudes are typically less than 1000 A suggesting that the prestrike or restrike is occurring close to the source voltage peak. Some polarity dependence
appears to be in evidence but the number of test shots (eight closes and opens) is probably too
Cumulative frequency (%)
120
100
Close positive
80
Close negative
60
Open positive
40
Open negative
20
0
0-500
501-1000
1001-1500
Current ranges (A)
Fig. B3 Inrush current distribution
low for proper statistical evaluation. The highest inrush currents occurred in Test 7 and the
trace is shown in Fig. B4. Inrush current can be seen to occur at times of 2.3, 2.9, 3.1 and
3.7 s. The other transients on the trace are believed to be crosstalk from restrikes that did not
result in inrush current.
The transformer side TRV is shown in Fig. B5. The TRV is of a critically damped nature as
was anticipated. The TRV is evident at time 3600 ms and more so following a last restrike at
time 3640 ms.
132
EHV transformer switching case study
The remaining parameter of interest is the contact gap spacing at current interruption and at
first prestrike. The blade angles at current interruption were determined from the video
records and are shown in Fig. B6 because of the high fixed arcing horn, current interruption
occurs in the range of blade angle 38° to 50° corresponding to a contact gap range of 800 mm
to 2000 m. The largest blade angle of 50° corresponded to the worst case of inrush current
shown in Fig. B4. On closing, the first prestrikes always occurred at blade angles of 40° or
less (contact gap 1 m or less). Lateral arc reach was insignificant.
kA
3
GO07I
2
1
0
-1
-2
0
1
2
3
4s
Fig. B4 Trace showing the highest magnitude of inrush current
133
Annex B
MV
0.6
0.4
0.2
0.0
-0.2
-0.4
-0.6
3300
MV
0.6
0.4
0.2
0.0
-0.2
-0.4
-0.6
3300
MV
0.8
0.6
0.4
0.2
0.0
-0.2
-0.4
-0.6
3300
GO08VA
3400
3500
3600
3700
3800
4000ms
GO08VB
3400
3500
3600
3700
3800
3900
4000ms
GO08VC
3400
3500
3600
3700
3800
Fig. B5 Transformer side TRVs
Fig. B6 500 kV transformer magnetizing current
interruption with vertical break disconnector
134
3900
3900
4000ms
EHV transformer switching case study
The results of the test support the assertions made in section 3. In summary, these are:
1. The transformer side TRV is a critically damped low frequency oscillation.
2. Comparison of Figs. 3.2 and B6 confirm that the contact gap appears to be the controlling
element.
3. The influence of inrush current is primarily to extend the arcing time by its duration
which tends to be short at less than 0.5 seconds. This is evident in the higher upper range
of interruption angle, Fig. B6 as compared to Fig. 3.2.
4. Prestriking and restriking occur close to the source voltage peak and the associated
overvoltages are therefore limited in magnitude.
B5. References
1. A. Greenwood and T.H. Lee, “Generalized Damping Curves and Their Use in Solving
Power-Switching Transients,” AIEE Transactions, August 1963.
135
Annex C
Video still images: Figures C1 to C60
Fig. C1
Initial current 24 A
Fig. C2
Initial current 44 A
Fig. C3
Initial current 62 A
Fig. C4
Initial current 82 A
Fig. C5
Initial current 98 A
Fig. C6
Initial current 121 A
137
Annex C
Fig. C7
Initial current 145 A
Fig. C8
Initial current 165 A
Fig. C9
Arc progression from 62 A initial current:
time zero
Fig. C10
Arc progression from 62 A initial current:
time zero + 100 ms
Fig. C11
Arc progression from 62 A initial current:
time zero + 200 ms
Fig. C12
Arc progression from 62 A initial current:
time zero + 300 ms
138
Video still images: Figures C1 to C60
Fig. C13
Arc progression from 62 A initial current:
time zero + 400 ms
Fig. C14
Arc progression from 62 A initial current:
time zero + 500 ms
Fig. C15
Arc progression from 62 A initial current:
time zero + 600 ms
Fig. C16
Arc progression from 62 A initial current:
time zero + 700 ms
Fig. C17
Arc progression from 62 A initial current:
time zero + 800 ms
Fig. C18
Arc progression from 62 A initial current:
time zero + 900 ms
139
Annex C
Fig. C19
Arc progression from 62 A initial current:
time zero + 1000 ms
Fig. C20
Arc progression from 62 A initial current:
time zero + 1100 ms
Fig. C21
Arc decay at 82 A initial current:
time zero
Fig. C22
Arc decay at 82 A initial current:
time zero + 2 cycles
Fig. C23
Arc decay at 82 A initial current:
time zero + 4 cycles
Fig. C24
Arc decay at 82 A initial current:
time zero + 6 cycles
140
Video still images: Figures C1 to C60
Fig. C25
Arc decay at 82 A initial current:
time zero + 8 cycles
Fig. C26
Arc decay at 82 A initial current:
time zero + 10 cycles
Fig. C27
Arc decay at 82 A initial current:
time zero + 12 cycles
Fig. C28
Arc decay at 82 A initial current:
time zero + 14 cycles
Fig. C29
Arc decay at 82 A initial current:
time zero + 16 cycles
Fig. C30
Arc decay at 82 A initial current:
time zero + 18 cycles
141
Annex C
142
Fig. C31
Arc decay at 82 A initial current:
time zero + 20 cycles
Fig. C32
Arc decay at 165 A initial current:
time zero
Fig. C33
Arc decay at 165 A initial current:
time zero + 2 cycles
Fig. C34
Arc decay at 165 A initial current:
time zero + 4 cycles
Fig. C35
Arc decay at 165 A initial current:
time zero + 6 cycles
Fig. C36
Arc decay at 165 A initial current:
time zero + 8 cycles
Video still images: Figures C1 to C60
Fig. C37
Arc decay at 165 A initial current:
time zero + 10 cycles
Fig. C38
Aberrant arc behaviour at 103 A initial current:
time zero front image
Fig. C39
Aberrant arc behaviour at 103 A initial current:
time zero + 2 cycles front image
Fig. C40
Aberrant arc behaviour at 103 A initial current:
time zero + 4 cycles front image
Fig. C41
Aberrant arc behaviour at 103 A initial current:
time zero + 6 cycles front image
Fig. C42
Aberrant arc behaviour at 103 A initial current:
time zero side image
143
Annex C
Fig. C43
Aberrant arc behaviour at 103 A initial current:
time zero + 2 cycles side image
Fig. C44
Aberrant arc behaviour at 103 A initial current:
time zero + 4 cycles side image
Fig. C45
Aberrant arc behaviour at 103 A initial current:
time zero + 6 cycles side image
Fig. C46
Aberrant arc behaviour at 103 A initial current:
arc transfer to disconnector blade
Fig. C47
Interrupting 24 A with loop impedance
40 ohms
Fig. C48
Interrupting 23 A with loop impedance
100 ohms
144
Video still images: Figures C1 to C60
Fig. C49
Interrupting 21 A with loop impedance
200 ohms
Fig. C50
Interrupting 29 A with loop impedance
150 ohms
Fig. C51
Interrupting 75 A with loop impedance
20 ohms
Fig. C52
Interrupting 73 A with loop impedance
40 ohms
Fig. C53
Interrupting 69 A with loop impedance
70 ohms
Fig. C54
Interrupting 70 A with loop impedance
100 ohms
145
Annex C
Fig. C55
Interrupting 98 A with loop impedance
40 ohms
Fig. C56
Interrupting 96 A with loop impedance
100 ohms
Fig. C57
Interrupting 90 A with loop impedance
150 ohms
Fig. C58
Interrupting 121 A with loop impedance
60 ohms
Fig. C59
Interrupting 124 A with loop impedance
150 ohms
Fig. C60
Interrupting 122 A with loop impedance
200 ohms
146
Annex D
Auxiliary interrupting devices and capacitive currents
D1. Introduction
Whip-type auxiliary interrupting devices are applied on disconnectors to provide an enhanced
line dropping capability. A less common so-called rigid arm device has also been used for
this purpose (Fig. D1). This device consists on a hinged actuating arm attached to the disconnector jaw assembly. As the blade opens, a latch attached to the blade lifts the actuating arm
and charges a torsion spring mechanism. When the blade reaches a predetermined open
contact gap, the latch releases the actuating arm and the arm and the blade now move in
opposite directions just achieving current interruption. These devices tend to have lower tip
velocities than the whip type devices.
Fig. D1 Rigid arm device
Note: this device is not one of the devices of Tables D3 and D4
Courtesy of Pacific Air Switches Corporation
In Fig. D2 still images are shown of a 115 kV pole-mounted vertical break disconnector
switching out a line length of about 14 km. The charging current is 3.3 A. The arc can be seen
to be solid but quite convoluted and the current is interrupted after the blade passes the 45°
angle position. Note also that the restrike follows along the previous arc channel and that hot
gas persists momentarily after current interruption. The total arcing time is about 1.5 seconds.
In Fig. D3 a whip-type device is applied to the same disconnector and the same line length
switched out. A momentary arc is drawn on all three phases with that on the centre phase
being the longest. No restrikes occurred and the blade at current interruption is about 45°.
The total arcing time is less than 0.25 seconds.
147
Annex D
Contact parting
Fig. D2 115 kV vertical break disconnector interrupting line charging current of 3.3 A (continued)
Courtesy of Bonneville Power Administration
148
Auxiliary interrupting devices and capacitive currents
Restrikes near and far phases
Restrikes all phases
Current interruption: note hot gas remnants
Fig. D2 115 kV vertical break disconnector interrupting line charging current of 3.3 A
Courtesy of Bonneville Power Administration
149
Annex D
Just prior to whip release
Arc follows whip tip trajectory
Current interruption
Fig. D3 115 kV vertical break disconnector (as in Fig. D2) equipped with whip-type device
interrupting line charging current of 3.3 A
Courtesy of Bonneville Power Administration
These devices cannot be described as precision current interrupting devices being dependent
on the geometry of the set-up among other variables. Tests run on the same device and the
same current levels in different laboratories will not necessarily yield the same results. Laboratory tests performed by two utilities demonstrates this varying performance as does actual
experience in the field.
D2. Laboratory tests
Allegheny Power Service Corporation ran a series of tests on two whip-type and two rigid
arm type devices.DF1 The devices were intended for use on 115 kV disconnectors. The results
of the tests are given in Tables D1 to D4.
DF1
T.J. Jackson, “138 kV Air Switch Interrupting Device Tests.” Presented at PEA Electrical Equipment
Meeting, Hershey, Pennsylvania, May 1977.
150
Auxiliary interrupting devices and capacitive currents
Table D1 Whip-type device 1
Voltage Current
Average arc length
Failures/Shots
kV
A
m
64*
4.7
0/20
< 0.3
6.2
0/20
< 0.3
9.2
0/20
< 0.3
11.0
0/20
< 0.3
12.2
0/19
< 0.3
13.5
0/20
< 0.3
15.7
0/20
< 0.3
23.7
0/20
< 0.3
27.1
0/20
< 0.3
32.7
0/20
< 0.6
39.9
0/20
< 1.5
105*
7.9
0/20
< 0.3
10.6
1/20
< 0.6
15.0
0/20
< 0.9
20.8
3/20
< 0.9
24.3
0/20
< 1.5
31.0
5/7
< 1.8
78.7‡
3.6
0/5
< 0.9
7.2
0/5
< 0.9
80‡
6.2
0/10
< 0.3
6.2
0/3
< 0.3
17.6
0/2
< 0.6
* Single-phase laboratory tests.
‡ Three-phase field tests.
151
Annex D
Table D2 Whip-type device 2
Voltage Current
Average arc length
Failures/Shots
kV
A
m
64*
4.7
0/20
< 0.3
6.2
0/20
< 0.3
9.2
0/20
< 0.3
11.0
0/20
< 0.3
12.2
0/20
< 0.3
13.5
0/20
< 0.3
15.7
0/20
< 0.3
23.7
0/20
< 0.3
27.1
0/20
< 0.6
32.7
0/20
< 0.6
39.9
0/23
< 0.6
105*
7.9
1/20
< 0.6
10.6
2/24
< 0.6
15.0
4/20
< 0.6
20.8
10/20
< 0.6
78.1‡
3.6
0/10
< 0.6
78.3
7.2
0/10
< 0.6
80‡
6.2
0/10
< 0.3
17.6
1/2
< 0.6
6.2
0/3
< 0.3
* Single-phase laboratory tests.
‡ Three-phase field tests.
Table D3 Rigid arm type device 1
Voltage Current
Average arc length
Failures/Shots
kV
A
m
64*
4.7
0/20
< 0.3
6.2
0/20
< 0.6
9.2
0/20
< 0.6
11.0
0/20
< 0.6
12.2
1/20
< 0.9
13.5
3/20
< 0.9
15.7
1/20
< 0.9
18.3
4/20
< 1.2
105*
7.9
6/7
5.0
2/2
* Single-phase laboratory tests.
152
Auxiliary interrupting devices and capacitive currents
Table D4 Rigid arm type device 2
Voltage Current
Average arc length
Failures/Shots
kV
A
m
64*
9.2
0/20
< 0.3
13.5
0/20
< 0.9
15.7
0/20
< 1.2
18.3
0/20
< 1.2
20.4
0/20
< 1.5
23.7
0/20
< 1.2
27.1
0/20
< 1.5
32.7
0/20
< 1.5
39.9
0/20
< 1.5
105*
15.0
9/20
< 1.2
10.6
9/15
< 1.2
7.9
9/20
< 1.2
* Single-phase laboratory tests.
Failure in the test means that the auxiliary interrupting device has not interrupted the current
by the time it reaches its neutral position. For the whip-type device, this is parallel to the
blade and, for the rigid arm device, its rest position. Whip-type device 1 was judged to offer
the best potential performance and suitable for dropping 138 kV lines up to 64 km in length
(approximately 20 A) without exceeding a limiting arc length of 0.9 m. A disconnector with
arcing horns only was considered suitable for dropping 138 kV lines up to about 6 km in
length (approximately 2 A – refer to Fig. 4.1).
BPA conducted a series of laboratory tests on a single-phase 115 kV vertical break disconnector with various auxiliary interrupting devices (termed as quick break devices).DF2 Failure
was defined as discussed above and the results are summarized in Table D5. BPA ranked performance based on the current level at which the device first showed marginal performance
and it is interesting to note that these are generally much lower than those of the Allegheny
Power laboratory tests (the BPA test voltage was 84 kV). Whip #5 was actually the same
whip-type device as tested in 1972 and exhibited approximately one-half the earlier test result
at about the same voltage. This result tends to support actual field experience with the device
as noted in subsection 2.3 and discussed below.
DF2
BPA Division of Laboratories: Laboratory Report ELE-89-39, “Tests to Determine the Interrupting
Capabilities of 115 kV Quick Break Devices.” March 1989.
153
Annex D
Table D5 BPA test results for various auxiliary interrupting devices
Suitable to switch Drop 115 kV line
currents less than lengths less than
A
km
Whip #1
3
12
Rigid arm #1
5.4
21
Whip #2
5.5
22
Rigid arm #2
7.3
29
Whip #4
8
32
Whip #5
9
36
Whip #6
11
44
Device type
The difference in performance of the same device in two separate laboratory tests cannot be
definitively explained. Both tests were run on a single-phase basis using variable shunt
capacitor banks as the load circuit. Some variation can be expected in the mounting of the
device but this is not considered sufficient to explain the difference.
The field tests on the whip-type device 1 in Table D3 are more indicative of performance. As
installed, the device protrudes beyond the jaw assembly by about 70 cm. This can be viewed
as increasing the effective blade length and, with reference to Fig. 4.2 for a disconnector at
145 kV, gives an equivalent minimum blade angle for current interruption at less than 15°.
This in turn gives a minimum gap spacing in the expected range. However, the average arc
lengths in one case were recorded as being less than 0.9 m for currents of 3.2 A and 7.2 A.
This is equivalent to about the minimum gap spacing for a 245 kV disconnector and to a
neutral blade position of about 30° on a 145 kV disconnector. This suggests therefore that, to
interrupt the current by a neutral blade position of 45°, the allowable current is in the range
close to 7.2 A.
D3. Field experience
Experience with whip-type devices has shown a clear dependency on phase spacing. The
devices are most commonly used for line dropping at 72.5 kV and 145 kV. Field data from
various utility sources for a particular device applied at 115 kV and 138 kV is given in Table
D6.
154
Auxiliary interrupting devices and capacitive currents
Table D6 Field data for whip device at 138 and 115 kV
Source
Supplier
Utility 1
Utility 2
Utility 3
Utility 4
Utility 5
Voltage Phase spacing Current (A) to be interrupted
kV
m
Successful
Failed
132
Unknown
10.25
138
24
138
11−17.6
138
3
13
3
15.3
2.13
9.6
138
3.66
1.92
138
Unknown
7.2
138
2.13
9.6
115
3
16
A further complication is the weather: devices which perform well in dry weather fail in wet
or very humid conditions. This is believed to be due, at least in part, to increased current
caused by leakage across line insulators and corona losses.
155
Annex E
Comparative analysis of loop switching tests by Andrews, Janes
and Anderson
E1. Introduction
Andrews et al5 conducted their loop switching tests in July 1945 and November 1946 and the
combined results are shown in Fig. 2.7. The test data points – initial currents, loop impedances, arcing times, arc lengths and open circuit voltages – are given by Gerngross11 thus
enabling a detailed analysis and comparison to the Powertech loop switching test results from
February and May 2000.
E2. Comparison
The scope of the two test series is shown in Fig. E1. Andrews et al conducted their tests over
a narrow range of loop impedance, necessarily so because transmission lines were used to
achieve the impedance values, and a current range of 60 A to 312 A. The Powertech tests
350
All Powertech tests
Initial current (A)
300
Andrews et al tests
250
200
150
100
50
0
0
50
100
150
200
250
Loop impedance (ohms)
Fig. E1 Scope of Andrews et al and Powertech test series
were run in a high power laboratory and greater flexibility with respect to impedance and
current was possible. The common area is for impedances less than 50 ohms and currents
greater than 60 A.
The test series are first compared for consistency. The Powertech test results are shown in
Fig. E2 and indicate good agreement between arcing times and therefore consistent results.
Fig. E3 shows the Andrews et al test points and only three points are common to both test
times. The results associated with these are given in Table E1. The points show a reasonable
consistency. However, the arcing times are remarkably short and the arc lengths remarkably
long and is discussed further later.
157
Annex E
160
February: arcing time <=1s
140
May: arcing time <=1s
Initial current (A)
120
100
February: arcing time 1 to
2.5s
80
May: arcing time 1 to 2.5s
60
40
Pow er (February: arcing time
<=1s)
20
Pow er (May: arcing time
<=1s)
0
0
50
100
150
200
250
Pow er (February: arcing time
1 to 2.5s)
Pow er (May: arcing time 1 to
2.5s)
Loop impedance (ohms)
Fig. E2 Comparison between Powertech February and May 2000 test results
350
Initial current (A)
300
250
Test points: July 1945
200
Test points: November
1946
150
100
50
0
0
10
20
30
40
50
Loop impedance (ohms)
Fig. E3 Comparison between Andrews et al July 1945 and November 1946 test
points
Table E1 Andrews et al common test points
Quantity
158
Test time
July 1945
November 1946
Initial current (A)
242
238
249
Loop impedance (ohms)
24.5
24.9
23.8
Arcing time (s)
0.38
0.45
0.29
Arc length (m)
7.3
8.9
10
Comparative analysis of loop switching tests by Andrews, Janes and Anderson
The next comparison is to check for expected variations between the parameters of initial
current, loop impedance and arcing time. Both test series exhibit the expected variation of
increasing arcing with increasing loop impedance for constant current as shown in Figs. E4
and E5. Note, however, the large difference in arcing time.
3000
6 ohms loop
impedance
Arcing time (ms)
2500
10 to 16 ohms loop
impedance
2000
1500
20 to 40 ohms loop
impedance
1000
Linear (6 ohms loop
impedance)
500
Linear (10 to 16
ohms loop
impedance)
0
0
100
200
300
Initial current (A)
Linear (20 to 40
ohms loop
impedance)
Fig. E4 Powertech test results
11 to 18 ohms loop
impedance
1000
900
22 to 30 ohms loop
impedance
Arcing time (ms)
800
700
600
31 to 43 ohms loop
impedance
500
400
Linear (11 to 18
ohms loop
impedance)
300
200
Linear (22 to 30
ohms loop
impedance)
100
0
0
100
200
300
400
Initial current (A)
Linear (31 to 43
ohms loop
impedance)
Fig. E5 Andrews et al test results
Arc length and arc reach are compared in Fig. E6. The arc lengths are those given by
Gerngross and the arc reaches are based on the statement in the reference that the arc reaches
never exceeded half the arc length.5 Arc reach as defined by Andrews et al can be in any
direction and can be directly compared to the horizontal reaches from the Powertech data.
159
Annex E
Andrews et al: arc
length
14
Arc length or arc reach (m)
12
Andrews et al:
maximum arc reach
10
Powertech:
horizontal arc reach
8
6
Linear (Powertech:
horizontal arc
reach)
4
Linear (Andrews et
al: maximum arc
reach)
2
0
0
100
200
300
Linear (Andrews et
al: arc length)
400
Initial current (A)
Fig. E6 Comparison of arc reach or length for Andrews et al tests versus
Powertech tests
The difference between the reaches is quite extraordinary, all the more so given that Andrews
et al tests were run on a 33 kV disconnector with an open gap of probably no more than
0.5 m. Observation of the Powertech test arcs, particularly those at the higher currents, shows
that the arc tends to be limited by an apparent balance between the magnetic force driving the
arc outward and partial arc collapses, and of course the diminishing rate of increase of power
input – refer to Fig. 5.14.
E3. Conclusions
The difference in the results between the two test series, performed some fifty years apart, is
significant. In Fig. E7 measured arc reaches for three cases at 70 ohms loop impedance from
the Powertech tests are compared to values calculated using Eqns. (2.2) and (2.3) as follows.
4
33 A Powertech
reach
Vertical reach (m)
3.5
3
69 A Powertech
reach
2.5
103 A Powertech
reach
2
1.5
33 A Andrews
reach
1
0.5
69 A Andrews
reach
0
0
1
2
3
Horizontal reach (m)
4
103 A Andrews
reach
Fig. E7 Measured Powertech arc reaches versus calculated Andrews et al arc
reaches for a loop impedance of 70 ohms
160
Comparative analysis of loop switching tests by Andrews, Janes and Anderson
33 A and 70 ohms:
LPRl = 5.03 × 70 × 332 × 10-3
= 383 mm
69 A and 70 ohms:
LPRl = 5.03 × 70 × 692 × 10-3
= 1676 mm
103 A and 70 ohms:
LPRl = 0.503 × 70 × 103
= 3626 mm
The plot compares the actual arc reach direction for the Powertech tests against the locus of
the possible reach directions for the calculated points. At 33 A, the calculated reach is lower
than the actual reach and at 69 A and 103 A the calculated reach is greater to much greater
than the actual reaches. The question is: What is the reason for the difference in the arc
reaches between the two test series?
To answer that question one possible theory is that the differing circuit arrangements causes
the difference. The Powertech circuits used reactors, while the Andrews et al tests used
transmission line loops. The latter loops were not pure impedances but incorporated transformation: from 33 kV to 12 kV in the series circuit and 12 kV to 132 kV and then 132 kV to
33 kV in the parallel circuit. It is difficult to judge the exact influence of transformation
beyond adding a lumped impedance but it is implausible it could impact on arc development
and propagation.EF1
A more probable theory is that the arc was somehow drawn out rapidly in a horizontal direction. Gerngross11 in his thesis shows a photograph (unfortunately an unclear photograph) of a
device described as “arc measuring device used for drawing arc out horizontally.” This theory
is credible when we recall that the tests were based on the premise of a critical arc length to
achieve current interruption. The arc length then became the issue and artificially drawing it
out introduces an element of forced commutation, rather than natural commutation of current
to the parallel circuit. This would explain the long arcs and the short arcing times but absolute
proof, however, is not forthcoming.
EF1
McNulty29 advocated that transformation could be treated simply as a series impedance. However, many
utilities do not permit loop switching where transformation is involved no doubt based on bad experience
with the practice.
161
Annex F
Influence of weather
F1.
Introduction
No studies have been conducted of weather as it relates to current interruption using air break
disconnectors. The weather element of most interest is that of wind and that will be the main
focus of this annex. Temperature, humidity and air pressure are viewed as more relevant to
static air gap considerations than to the dynamic conditions associated with arcs.FF1 With
regard to heat transfer, air temperature is not significant compared to the arc temperature and
the high differential promotes heat transfer from the arc to the ambient air.
F2.
Literature
Weather is mentioned in a number of the references discussed in section 2. Warrington50
notes that wind can have the effect of drawing out the arc to even longer lengths and states:
Besides the effect of increasing length the arc resistance is further raised by the
cooling effect of the wind which causes deionization and reduction in cross-section of
the arc.
This statement appears, however, to be speculation rather than a proven fact. Andrews et al5
describe their experience with wind during transformer magnetizing current and loop
switching tests:
It was observed that wind conditions greatly affected the shape of an arc and its
direction but did not appreciably affect its length. With light wind it tended to be very
irregular and circuitous. With strong wind it tended to be blown out horizontally
windward, often like an elongated “V.”
This statement in part contradicts Warrington’s speculation and, in fact, the loop switching
tests performed at Powertech Laboratories tend to support these observations.
Anderson15 conducted his magnetizing current interruption tests under varying weather
conditions:
Weather conditions were adverse, it was cold and windy with showers of mixed rain
and snow.
Anderson noted further that the maximum contact gap at current interruption was about
600 mm. This compares well with the calculated values of Fig. 3.2 for a 230 kV disconnector.
The CEA report9 also discusses weather but mainly in the context of the above references.
FF1
Barrett and Green (see subsection 4.4) in their experiments tracked these quantities but found no correlation
to arc reach.
163
Annex F
F3.
Discussion
Wind can be expected to have an influence on arcs in air because it enhances heat transfer.
The influence, whether the arc is thin as in Fig. 4.8(p) or voluminous as in Fig. C59, is to
increase the power loss and thus promoting arc collapse and current interruption.
Filter and Amm conducted capacitive current interrupting tests on live line connectors at
115 kV and 230 kV.FF2 The range of current was 0.42 to 1.38 A at 115 kV and 0.1 to 0.83 A
at 230 kV. Wind velocity was recorded during the tests and treated on the following basis: no
wind was defined as wind velocity less than 6.5 km/hour and wind as wind velocity greater
than 6.5 km/hour. The wind condition resulted in the shorter arc reaches. At 1 A, the wind
reach was about 60% of that with no wind and at 0.5 A the wind reach was about 50% of that
with no wind. The report concluded that about 0.5 A could be interrupted by the connectors
without excessive reach.
The loop switching tests on the 230 kV disconnector at Powertech Laboratories were run in
November 1999, February 2000 and May 2000. The weather for the latter two test series was
fair with little or no wind, while the weather for the November test series was a storm with
high winds gusting at times. No record was kept of wind velocities, which varied anyway
from test shot to test shot, but it is reasonable that on average the test condition was one of
wind as compared to no wind. The November tests were in the approximate current ranges
20 A to 30 A and 50 A to 60 A and varying loop impedances. The results with arcing time as
parameter are plotted in Fig. F1 together with the results for the same ranges of current from
the later no wind tests. For the same arcing time ranges there is a shifting to the right for the
wind condition particularly in the lower current range. This is demonstrated more explicitly
70
Wind: arcing time <=500ms
Initial current (A)
60
Wind: arcing time 5011000ms
Wind: arcing time 10011500ms
No wind: arcing time
<=500ms
No wind: arcing time 5011000ms
50
40
30
20
No wind: arcing time 10011500ms
10
0
0
200
400
600
Loop impedance (ohms)
Fig. F1 Initial current versus loop impedance for conditions of wind and no wind
FF2
R. Filter and D.E. Amm, “115 kV and 230 kV Arc Reach Tests at Kleinburg Outdoor Test Facility.”
Ontario Hydro Research Division (now Kinectrics Inc.), Report No. 90-146-K, August 1990.
164
Influence of weather
in Fig. F2: only the points for the arcing time of 501 to 1000 ms are plotted and loop impedances are limited to common values. The regression shows that the wind condition is more
favourable to current interruption.
70
Initial current (A)
60
50
Wind: arcing time 5011000ms
40
No w ind: arcing time 5011000ms
30
Linear (Wind: arcing time
501-1000ms)
20
Linear (No w ind: arcing
time 501-1000ms)
10
0
0
100
200
300
Loop impedance (ohms)
Fig. F2 Initial current versus loop impedance for conditions of wind and no wind
and arcing time 501 to 1000 ms
F4.
Conclusions
Wind can have a positive effect on the interruption of current in air. However, it may promote
movement of the arc which increases the risk for contact with adjacent phases or grounded
structures. Operators should exercise discretion before opening disconnectors to break
circuits under wind conditions.
165
ACKNOWLEDGMENTS
I am greatly indebted to Prof. dr. R. Smeets for accepting me as a promotion candidate. I am
further indebted to him and the other members of the Core Promotion Committee, Prof.
L. van der Sluis, Prof. G. Damstra and Prof. dr. H-H. Schramm for their advice and encouragement during the work.
I want to particularly thank J.G. Krone, Managing Director, HAPAM B.V. for providing high
voltage disconnectors for the tests run at KEMA and of course KEMA for providing the test
time. My thanks also to Dr. P. Schavemaker and S. Kuivenhoven, TU Delft for their interest,
ideas and help with the MATLAB work.
I am further grateful to many others who supported the work in one form or other: to
R. Threlkeld, Senior Vice-President, BC Hydro T&D (now retired) and G. Smyrl, VicePresident, BC Hydro for approval to use the Powertech and other test data; to Dr. J. Brunke
and S. Lowder, Bonneville Power Administration, for sharing BPA practices and providing
video clips; to J. Shipek and M. Clifton of Puget Sound Energy for the line dropping video
tapes; to Dr. H. Knobloch, Siemens AG, for his report on capacitive current switching; to
Drs. W. Chisholm and J. Kuffel of Kinectrics Inc. (formerly Ontario Hydro Research) for
reports on past studies for Ontario Hydro; to Dr. J. Chand, Manitoba Hydro, for providing
experiences with auxiliary interrupting devices; to A. Crino, Pacific Air Switches Corporation, for information on disconnectors and auxiliary interrupting devices; to M. Alexander,
P. Gillan, J. Sawada and B. Sunga, BC Hydro, for reviewing various drafts; to P. Crawford,
V. Bright, R. Tennant and A. Wong, BC Hydro Information Centre, for the literature searches
and references; to B. Avent for the Powertech video images; and to B. Sunga and M. Hogg
for drafting various figures.
Lastly, I express my sincere thanks to Sandra Giasson for the word processing of this thesis
and her endless patience through the many changes and drafts.
167
CURRICULUM VITAE
David Peelo was born on 28 February 1943 in Dublin, Ireland. After completing high school,
he studied electrical engineering at University College Dublin and graduated cum laude in
1965. His first employment was at the ASEA High Voltage Laboratory in Ludvika, Sweden.
In 1973 he joined BC Hydro in Vancouver, British Columbia, Canada eventually becoming a
switchgear and switching specialist. He took early retirement in 2001 to pursue a new career
as an independent consultant and to undertake postgraduate work as represented by this
thesis. He is active in Cigre, IEC and IEEE committees and working groups and has authored
or coauthored over 40 technical papers.
168