Cigre_544 -- Metal Oxide (MO) Surge Arresters - Stresses and Test Procedures

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544

MO Surge ArrestersStresses and Test Procedures

Working GroupA3.17

August 2013



MO SURGE ARRESTERS

WG A3.17

Members

B. Richter, Convenor (CH),

J.L. De Franco (BR), R. Göhler (DE), F. Greuter (CH), V. Hinrichsen (DE), M. Holzer (AU), S. Ishibe

(JP), Y. Ishizaki (JP), B. Johnnerfelt (SE), M. Kobayashi (JP), K. Lahti (FI), T.M. Ohnstad (NO), R.S.

Perkins (CN), M. Reinhard (DE), J.H. Sawada (CA), A. Sironi (IT)

Corresponding Members

A. Dellallibera (BR), R. Diaz (AR), S. Vizintin (SL), Y. K. Tong (GB)

Copyright © 2013

“Ownership of a CIGRE publication, whether in paper form or on electronic support only infers rightof use for personal purposes. Are prohibited, except if explicitly agreed by CIGRE, total or partialreproduction of the publication for use other than personal and transfer to a third party; hencecirculation on any intranet or other company network is forbidden”.

Disclaimer notice

“CIGRE gives no warranty or assurance about the contents of this publication, nor does it accept

any responsibility, as to the accuracy or exhaustiveness of the information. All implied warrantiesand conditions are excluded to the maximum extent permitted by law”.

ISBN: 978-2-85873-239-5



MO Surge Arresters-Stresses and Test Procedures

Page 2

MO Surge ArrestersSTRESSES AND TEST PROCEDURES

Table of Contents

EXECUTIVE SUMMARY ............................................................................................................. 4

Foreword ........................................................................................................................................................................... 8

1. Stresses on Surge Arresters ................... ........................... ....................... ........................... ....................... ............... 11 1.1 Introduction ........................................................................................................................................................... 11 1.2 Stresses from three phase systems ........................ ....................... ........................... ....................... ........................ 11

1.2.1 General ........................................................................................................................................................... 11

1.2.2 Temporary Overvoltages ................................................................................................................................. 11 1.2.3 Slow front/switching overvoltages .................................................................................................................. 12 1.2.4 CB and DS TRVs ............................................................................................................................................... 13

1.3 Stresses from HVDC networks ................................................................................................................................ 15 1.3.1 Introduction.................................................................................................................................................... 15 1.3.2 Stresses on surge arresters.............................................................................................................................. 17 1.3.3 Creepage distance and clearance in air............................................................................................................ 19 1.3.4 Overvoltage limiting characteristics of arresters ................................... .......................... ........................ ......... 19 1.3.5 Surge arresters in a converter station.............................................................................................................. 21 1.3.5.1 AC bus arrester (Type A) ........................ ....................... ........................... .......................... ........................ ... 22 1.3.5.2 Valve arrester (Type B) ................................................................................................................................. 22

1.4 Stresses in traction systems .................................................................................................................................... 24

1.4.1 General ........................................................................................................................................................... 24 1.4.2 Voltages in traction systems ........................... ....................... ........................... ....................... ........................ 25 1.4.3 MO surge arresters for d.c. traction systems ................................................................................................... 26 1.4.4 MO surge arresters for a.c. traction systems ................................................................................................... 26

1.5 Stresses from Lightning .......................................................................................................................................... 27 1.5.1 Introduction.................................................................................................................................................... 27 1.5.2 Lightning surges. ............................................................................................................................................. 27 1.5.3 Examples from transient analysis. ................................................................................................................... 28 1.5.4 Lightning Statistics .......................................................................................................................................... 29 1.5.5 Winter lightning. ............................................................................................................................................. 29 1.5.6 Parameters of summer and winter lightning current ......................... ........................ ....................... ............... 32

1.6 Ambient stresses .................................................................................................................................................... 37

1.6.1 Mechanical stresses ........................................................................................................................................ 37 1.6.2 Pollution ......................................................................................................................................................... 46 1.6.3 Humidity ......................................................................................................................................................... 50 1.6.4 Combined humidity and AC stresses ........................ .......................... ........................ .......................... ............ 53 1.6.5 Exposures to low ambient temperatures .................... ....................... ........................... ....................... ............ 55 1.6.6 Biological growth ............................................................................................................................................ 56

1.7 Short circuit currents.............................................................................................................................................. 58

2. Functional parameters and design of MO Surge Arresters ............................. ....................... ........................... ......... 62 2.1 Function and relevant parameters .......................................................................................................................... 62

2.1.1 Introduction.................................................................................................................................................... 62 2.1.2 Currents and voltages ..................................................................................................................................... 64




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2.1.3 Coordination of insulation and selection of arresters ................... ........................... ....................... .................. 67 2.2 MO-Varistors: state of the art and actual trends .................. ........................... ....................... ........................ ......... 69

2.2.1 Electrical properties of the metal-oxide resistor .............................................................................................. 69

2.2.2 Microstructure of Metal-Oxide resistors ....................... ........................... ....................... ........................ ......... 70 2.2.3 The manufacturing process ............................................................................................................................. 72 2.2.4 Electrical testing of Metal-Oxide varistors ....................................................................................................... 73 2.2.5 From grain boundaries to varistor blocks ...................... ........................... ....................... ........................ ......... 73 2.2.6 Failure modes of varistor blocks ...................................................................................................................... 79 2.2.7 Long-term stability of ZnO varistors................................................................................................................. 83 2.2.8 Trends and open issues ................................................................................................................................... 85

2.3 Design of surge arresters ........................................................................................................................................ 87 2.3.1 Foreword ........................................................................................................................................................ 87 2.3.2 Design principles of polymer housed HV arresters ....................... ........................ .......................... .................. 87 2.3.3 The mechanical supporting structure .............................................................................................................. 87 2.3.4 Outer housing and sheds................................................................................................................................. 92 2.3.5 Design principles of polymer housed MV arresters ................... ........................... ....................... ..................... 92 2.3.6 Conclusion ...................................................................................................................................................... 93

2.4 Special designs of surge arresters ........................................................................................................................... 94 2.4.1 Separable and Dead front Arresters ................................................................................................................ 94 2.4.2 Under-oil Arresters ......................................................................................................................................... 96

2.5 SF6 gas insulated MO surge arresters ...................................................................................................................... 98 2.6 Integrated Arrester Systems ................................................................................................................................. 102

3. Energy handling capability of MO surge arresters .......................... ........................ ....................... ......................... 104 3.1 Summery ............................................................................................................................................................. 104 3.2 Introduction ......................................................................................................................................................... 104 3.3 The different aspects of “energy handling capability” ................... ........................... ....................... ...................... 105

3.3.1 Thermal energy handling capability ................... ....................... ........................ ....................... ...................... 105

3.3.2 Impulse energy handling capability ............................................................................................................... 107 3.4 State of knowledge about energy handling of MO arresters ...................... ........................ ....................... ............. 108

3.4.1 A brief review of the relevant literature ........................................................................................................ 108 3.4.2 Results of an experimental investigation initiated by Cigré WG A3.17 ........................... ....................... .......... 112

3.5 Energy handling capability in international arrester standards ........................... ....................... ........................ .... 129 3.5.1 General ......................................................................................................................................................... 129 3.5.2 Energy handling issues in standard IEC 60099-4............................................................................................. 130 3.5.3 Energy handling issues in standard IEEE C62.11 ............... ........................... ....................... ........................ .... 136 3.5.4 Energy handling issues in other national standards ................... ........................... ....................... ................... 138 3.5.5 Conclusion and outlook ................................................................................................................................. 138

4. Summary .................................................................................................................................................................... 139

APPENDIX 1 ................... ....................... ........................... ....................... ........................... ....................... ...................... 140 References ..................................................................................................................................................................... 141




Page 4

EXECUTIVE SUMMARYThe Cigré Technical Brochure TB 60 was published in 1991 describing effects on gapless metal oxide surge

arresters (MO arresters) from various electrical stresses encountered in 3-phase AC systems. Since then,

continued improvements in equipment technologies coupled with the interest of the de-regulated power industry tomaximize utilization of the existing infrastructure has revolutionized the MO arrester applications and the expected

performances in an environment, characterized by higher stress levels.

Today’s proven confidence in the reliability and capability of modern MO arresters offers new possibilities of

overvoltage protection and improved management of power system disturbances.

The Working Group A3.17 of SC A3 took the task to evaluate the stresses on MO arresters and to review the

existing test procedures. Further on, the actual state of MO arrester designs was investigated, as well as the

various applications in different types of electrical networks.

Emphasis was given to the MO resistors as the active part of the MO arresters. A research project was started to

experimentally investigate the energy handling capability of the MO resistors, which is a key design criterion for a

reliable arrester application. The resulting Technical Brochure covers and describes the actual MO resistor and

arrester technology and the results of the first part of the research project on the energy handling capability of MO

resistors.

Electrical stresses on MO arresters can be divided into stresses at power frequency, which can have long time

durations, and transient stresses of short time duration resulting from switching and lightning. IEC 60071-4

proposes some recommendations for the evaluation of overvoltages, based on the use of numerical programs.

The different stress types seen by a MO arrester are:

Temporary Overvoltages

A temporary overvoltage (TOV) is an oscillatory phase-to-ground or phase-to-phase condition that is of relatively

long duration and is undamped or only weakly damped. TOV are one of the most crucial stresses to an MO

arrester and are detrimental for its layout.

The following origins of TOV are typically considered:

- Earth fault temporary overvoltages occur in a large part dependent on the effectiveness of system earthing.

Guidance for the determination of TOV amplitudes is given in IEC 60099-5 and IEC 60071-2.

- Disconnection of a load will cause the voltage to rise at the source side of the operating circuit breaker. The

amplitude of the overvoltage depends on the disconnected load and the short-circuit strength of the feeding

substation. The amplitude of load rejection overvoltages is usually not constant during their duration.

Accurate calculations have to consider many parameters.

- Voltage rise along long unloaded lines (Ferranti effect).

- Harmonic overvoltages, originating from e.g. DC converters or saturated transformers.

- Resonances, in particular Ferro resonances.

- Overvoltages due to flashover between two systems of different system voltages installed on the same

tower.

Slow-front overvoltages

Slow-front overvoltages, in most cases generated by switching or faults, are associated with load switching or fault

clearing. Different switching cases have to be considered: line re-energization, switching of capacitive loads and

inductive loads.




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Fast-front overvoltages

Fast-front overvoltages are in many cases caused by thunderstorms and occur all over the world. The heaviest

thunderstorms with the most intensive lightning will normally be experienced in the equator region. Other sources

are, for instance, current chopping of breakers or back flashovers.

In low voltage (LV) power systems up to 1 kV and medium voltage (MV) power systems (1 kV < Us 52 kV) the

distribution lines are generally of lower height and less exposed to direct flashes than transmission lines. Most of

the occurring overvoltages are then due to induced voltages originating from lightning to surrounding structures.

High voltage (HV) systems in the range of 52 kV < U s 245 kV can be found in transmission and sub-transmission

rural areas. Direct strokes, back flashovers and induced overvoltages will statistically result in a higher stress for

the installed arresters than in other voltage systems.

Transmission lines in extra-high voltage (EHV) with 245 kV < Us 800 kV and ultra-high voltage (UHV) systems

above 800 kV have steel towers with shield wires and are in spite of their height above ground well protected

against direct lightning strokes to the phase wires. Most of the lightning will hit the towers or the shield wires, and

only shielding failures and back flashovers will cause a critical surge in the phase wire.

In general, in 90% of all cases the lightning flashes are negative flashes from cloud to ground. However, some

countries, such as Norway or Japan, experience rather often thunderstorms during winter. Typical weather

conditions to create the winter thunderstorms are strong winds from the west, which transport warm air from the

ocean to the mountains of the mainland. The typical positive lightning flashes of winter thunderstorms transfer

higher charge than negative lightning flashes, which are typical for summer thunderstorms.

HVDC networks

Since the late 1970s overvoltage protection of HVDC converter stations has been based exclusively on MO

arresters. This is due to their superior protection characteristics and their reliable performance when connected in

parallel to the sensitive converters.

The continuous operating voltage stress for HVDC MO arresters differs from that of a normal a.c. arrester in that it

consists of not only the fundamental frequency voltage but also of components of direct voltage, fundamental

frequency voltage and harmonic voltages, and high frequency transients. These waveforms require other

dimensioning rules for the continuous operating voltage and some specific tests of the MO arresters, e.g. the

accelerated ageing procedure, as described in the emerging IEC 60099-9. Furthermore, polarity reversals might be

an issue.

Ambient stresses

Mechanical stresses like seismic loads strongly affect the structure and materials used for the design of the MO

arresters. Vibrations as well as static and dynamic loads have to be considered and appropriate test procedures

have been developed accordingly.

Ambient stresses can be very different in the different regions of the world. Very cold climates with ice and snow

loads have to be considered as well as climates of high temperature and high humidity. Observations of biologicalgrowth on the surface of polymer insulation have been made in various places. Three types of organic growth have

been identified: Algae, Fungi and Lichen. Despite all the reports of biological growth on the insulation in some

areas of the world there are up to now no known failures of MO arresters caused by it. Animal impact may be

another issue in some countries of the world, e.g. Australia, where cockatoos would nibble on specific types of

polymeric material.




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MO Resisto rs

Steady progress has been made over the last decades in MO resistor technology, their application in overvoltage

surge protection devices and the understanding of the basic mechanisms of nonlinear conduction, energy handling

capability etc. A lot of new insights have been gained, new physical phenomena have been observed, improvedand more consistent models have been developed and much progress has been made in simulations related to

materials and components.

The nonlinear conduction mechanism of the material can be traced back to individual grain boundaries in the

ceramics, which show a typical value of the switching or breakdown voltage UB of approximately 3.2 V– 3.4 V each.

Combining many grain boundaries in series and in parallel within an MO element allows to scale the voltage and

current characteristic of an MO resistor. For a sufficient large number of grain boundaries, the field strength E and

current density J then describes the material characteristic more generally.

Design of MO arresters

Different basic design principles are used for high voltage arresters and medium voltage arresters. In the high

voltage field mechanical requirements are much higher than in normal distribution applications. For this reason

porcelain housings are still used besides the growing number of hollow core insulators, so called tube designs,

and direct molded designs. For distribution arresters in medium voltage systems, porcelain housings have rapidly

disappeared and the direct molded design is used almost exclusively today.

Energy handling capability of MO resistors

The energy handling capability is a key property of MO arresters and has many different aspects, which are only

partly or not at all reflected in the actual standards. At least, though this list may be not complete, they have to be

divided into:

- “thermal” energy handling capability

- “impulse” energy handling capability

For the “impulse” energy handling capability single impulse stress, multiple impulse stress (without sufficient

cooling between the impulses), and repeated impulse stress (with sufficient cooling between the stresses) have to

be considered.

Thermal energy handling capability, on the other hand, can only be considered for complete arresters, as it is

mainly affected by the heat dissipation capability of the overall arrester design, besides the electrical properties of

the MO block.

For a deeper understanding of the energy handling capability of MO resistors and the relevant parameters, the

working group A3.17 initiated a research project to evaluate the energy handling capability under different impulse

stresses such as rectangular impulse currents, sine half waves, alternating currents and double exponential high-

current impulses. More than 3000 specimens of commercially available MO resistors from seven well established

American, European and Japanese manufacturers were tested. Two basically different sizes of MO resistors were

considered, one for application in high voltage arresters (“Size 1”: 40...45 mm in height, 60 mm diameter) and

one for application in medium voltage arresters (“Size 2”: 30...40 mm in height, 40 mm diameter).

For the tests with impulse stresses, an extended failure criterion, beyond simple visible damages, was introduced

for the first time to differentiate the various failure modes and to quantify early changes in the electrical material

characteristics. The a.c. tests were performed up to mechanical failure.

It turned out that for the a.c. tests up to failure the statistical evaluation gives better information on very low failure

probabilities compared to the impulse stress tests (characterized by their mean failure probabilities).

Some of the most important conclusions from the research program, as discussed in more detail in the TB, are:

- Energy handling capability has generally been improved over the last decade by the established

manufacturers.




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- Energy handling capability increases with current density.

- Statistical evaluation is easier to perform for a.c. tests and leads to more reliable predictions than for

impulse testing.

- Due to different dominating failure mechanisms, the energy handling capability is somewhat lower for “Size

2” resistors.

- For the lightning current impulse ( 90/200 s), recently introduced in the arrester standard IEC 60099-4,

energy handling capability may be affected by flashover phenomena.

Outlook

The follow-up working group of A3.17 (A3.25: Metal oxide varistors and surge arresters for emerging system

conditions) is working on:

- Further aspects of the energy handling capability such as durability (repeated impulses) or combined

stresses

- UHV arresters- Consequences of increasing the field strength of MO resistors

- Long term ageing of MO resistors

- Consequences of the axial temperature distribution in an MO arrester

The outcome of WG A3.25 will be given in an up-coming Technical Brochure.




Page 8

Foreword

Cigré Technical Brochure 60 (TB 60) METAL OXIDE ARRESTERS IN AC SYSTEMS written by Working Group 06

of Study Committee 33, published April 1991, describes the severity with which system parameters affect arrester

performance and how system performance is affected by the arresters. The main intention was to give detailedinformation on the application of the new type of surge arrester at this time. In addition the IEC standards for testing

and application developed in parallel. Many of the results of TB 60 were incorporated at this times in the new IEC

standards for metal oxide surge arresters (MO arresters).

TB 60 addresses, besides some basics about the characteristics of the MO material, the application in high voltage

3-phase transmission systems with 50/60 Hz and MO arresters with porcelain housings only.

Since the 1990s the application of MO arresters has increased in general and, due to the relatively simple and

robust mechanical design of the MO arresters compared to the conventional gapped arresters with SiC resistors,

new applications have become possible.

Continuous basic research on the MO material as well as the introduction of polymeric materials for the housings of

MO arresters for all system voltages has brought new and deeper knowledge and new application possibilities at

the same time. Today almost 100% of the medium voltage MO arresters have a polymeric housing, porcelain types

are not produced anymore on large scale. In the high voltage field more than 50% of all designs are of the

polymeric type, with increasing share. This development has brought new possibilities and as usual new questions.

The mechanical and pollution performance is of course different for polymeric and porcelain designs.

The MO material itself has been studied continuously over the years, which has brought better understanding of

the overall characteristics and better MO resistors with respect to electrical characteristics, homogeneity, long term

stability and energy withstand capability.

The number of manufacturers of MO resistors and arresters has increased, as well as the application of MO

arresters. Nowadays MO arresters are installed in a.c. and d.c. power systems with very different voltage levels,

from 660 V d.c. in traction systems up to 800 kV d.c. in HVDC systems, up to 1100 kV a.c. in UHV systems, and

they are used in substations, in cable systems, as line arresters etc., to give only some examples. In parallel, the

application of zinc oxide based MO varistors developed into a mass market for low voltage and electronicapplications, but this development is not described here.

The continuous development and the field experience with the MO arresters made it necessary to review the actual

state of the technology as well as the validity of the existing standards for testing MO resistors and arresters. An

example is for instance the classification of MO arresters in line discharge classes. The line discharge classes for

MO arresters are based on the energy that may be stored in transmission lines of different system voltages. This

classification works well as long as only 3-phase transmission systems up to 550 kV system voltage are being

considered. Various new applications in all electrical power systems, including UHV and HVDC, traction systems,

distribution systems etc. makes it necessary to reconsider the classification according to line discharge classes. For

this reason a critical review of the existing international standards was performed with emphasis on the energy

handling capability of MO resistors. To get a clearer and deeper understanding of “energy” related to MO resistors

and arresters the working group initialized a research program on energy handling withstand capability of MO

resistors. For the first time several thousand MO resistors for medium and high voltage application from manydifferent manufacturers were tested up to the limits and relations between the type of current impulse stress and

the failure mode of the MO resistors were evaluated.

Following the title Evaluation of stresses of Surge Arresters and appropriate test procedures and the scope of

working group A3.17 of Cigré SC A3, High Voltage Equipment, the TB is structured in the following sections:

Section 1. Stresses on Surge arresters

describes in general the different types of stresses on MO arresters, which may influence the performance of the

arresters. Naturally the performance of polymer housed and metal clad arresters is different in many aspects to the

performance of the “classical” designs with porcelain housings.




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- Subsection 1.2 (following a general introduction in 1.1) gives an overview about the stresses in 3-phase systems

with specific attention to temporary overvoltages and switching overvoltages. This is of special interest for system

studies.

- The subsection 1.3 addresses the special case of the very different voltage wave shapes and related stresses inHVDC systems. The performance of the MO arresters under voltage stresses different from pure a.c. or pure d.c.

needs careful consideration.

- The specific conditions of d.c. and a.c. traction systems are dealt with in subsection 1.4.

- Stresses from lightning are discussed in subsection 1.5. Lightning parameters are given and the severe and

special cases of winter lightning is addressed. Results from studies and evaluations about the occurrence of

lightning stresses in different systems are given as examples.

- Subsection 1.6 deals with various stresses from the ambient. This can be divided into static and dynamic stresses

and the severe case of seismic stresses, which is especially important for larger equipment with mechanically

sensitive internal design like SF6 gas insulated (GIS) arresters. Further on, long term stresses with pollution and

humidity, as well as very low temperatures and temperature cycles, are of importance if polymeric insulation is

concerned. Biological growth is addressed in brief.

- Finally, subsection 1.7 deals with the electrical and mechanical stress under overload conditions.

Section 2. Func tional parameters and design o f MO Surge Arresters

deals with material and design aspects of MO resistors and arresters, respectively.

Surge arresters constitute an indispensable means for insulation coordination in electrical power supply systems.

A general definition states that a surge protective device is a device that is intended to limit transient overvoltages

and divert currents. Two different principles exist: voltage switching devices based on a spark gap (which are the

old gapped arresters with SiC resistors), and voltage clamping devices based on varistor technology. In the high

voltage community the today’s devices are of the voltage clamping type and are called MO surge arresters, or

shortly arrester. A MO arrester has, simply speaking, to protect important and expensive electrical equipment

against damages resulting from overvoltages.

- Subsection 2.1 gives details about the voltage-current-characteristics of MO surge arresters, shows the current

and voltage wave forms as specified and standardized in international standards.

- Subsection 2.2 provides an overview of the material science of the MO material, the production process, and

leads from the micrometer scale of a single grain boundary up to the complete MO resistor and arrester. Possible

failure modes of the MO resistor and the long term performance of the material are addressed.

- In subsection 2.3 the different design principles of medium and high voltage arresters are shown.

- MO surge arresters with designs adapted to specific applications are dealt with in subsections 2.4, 2.5 and 2.6.

Section 3. Energy handling capability of MO surge arresters

deals with the need of a critical review of the existing standards and gives the details of the research project.

- In subsection 3.1 (summary) and 3.2 (introduction) the motivation of the performed research project on energy

handling capability is given and the general results are summarized.

- Subsection 3.3 explains the different aspects of energy handling capability for MO resistors and complete MO

surge arresters.

- In subsection 3.4 the state of knowledge and the initiated research project on energy handling capability of MO

resistors is described in detail.

- Subsection 3.5 finally gives a critical review of the many different aspects of the energy capability of MO surge

arresters in international standards.




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Section 4. Summary

summarizes the work of WG SC A3.17 and points out the influence on the actual standardization work in IEC TC

37.

APPENDIX 1gives an overview about Cigré Technical Brochures related to MO surge arresters and their application.

Considering the actual development and discussion in the field of MO surge arresters some subjects have been

addressed more in detail than others. For instance the electrical stresses on MO surge arresters for application in

HVDC systems are of increasing interest because of the increasing numbers of HVDC lines. Further, MO surge

arresters (and high voltage equipment in general) for application in UHV systems require special attention with

regard to pollution and seismic stresses and possible test procedures. The development and the variety of possible

arrester designs made it necessary to go into the details of the actual designs available on the market. Questions

regarding the long term stability and the energy handling capability of the MO resistors can only be dealt with when

the material properties are given in detail.

The content of this TB was discussed and agreed by all members of the working group. The sections were written

by one or more authors in charge. Each section starts with a short introduction to the specific subject and ends witha short conclusion. That’s why each single section can be read by itself without necessarily reading the other

sections.




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1. Stresses on Surge Arresters

1.1 Introduction

Surge arresters are widely applied in HV and MV a.c. and d.c. systems. They provide overvoltage protection fromthe generator in the power plant up to the end-user, including protection of substations, overhead-lines, and cables.

They are installed in fixed installations and in traction systems. Due to the world wide applications under very

different and sometimes severe ambient conditions a variety of stresses occur.

According to the title of the working group “Evaluation of stresses of Surge Arresters and appropriate test

procedures” the working group collected all stresses that can occur.

The addressed stresses, like electrical stresses from the system, from lightning and from ambient are described in

the following chapters.

1.2 Stresses from three phase systems

Author in charge: Jack Sawada

1.2.1 GENERALCIGRE WG 06 published Technical Brochure TB60 in 1991 which describes effects on gapless metal oxide surge

arresters (SA) from various electrical stresses encountered in AC systems. Since then, continued improvements in

equipment technologies coupled with de-regulated power industry’s interest in maximizing utilization of existing

infrastructure has revolutionized SA applications and expected performances in a more stressful environment.

Confidence in the reliability and capabilities of modern SAs and power electronic based equipment offer improved

management of power system disturbances. On the other hand, such power systems which allow increasing

number of distributed generations tapping into existing transmission and distribution circuits or circuits of different

voltage classes sharing common towers increase operational complexity and higher occurrences of network stress.

1.2.2 TEMPORARY OVERVOLTAGES

A temporary overvoltage (TOV) is an oscillatory phase-to-ground or phase-to-phase condition that is of relativelylong duration and is undamped or only weakly damped. TOV magnitudes are determinable and the stress on

surge arresters and insulation is considered in steady-state terms. The following causes of temporary overvoltages

are typically considered:

Earth fault overvoltages occur in a large part dependent on the effectiveness of system grounding. Guidance for

the determination of temporary overvoltage amplitudes is given in Annex of IEC 60099-5. The duration of the

overvoltage corresponds to the period of the fault (until fault clearing). Within earthed neutral systems it is generally

less than 1 s. For resonant earthed neutral systems, with fault clearing, it is generally less than 10 s and systems

without earth fault clearing the duration may be several hours.

Load rejection, following disconnection of a load will cause the voltage to rise at the source side of the operating

circuit breaker. The amplitude of the overvoltage depends on the disconnected load and the short-circuit strength of

the feeding substation. The temporary overvoltages have particularly high amplitudes after full load rejection at

generator transformers depending on magnetizing and over speed conditions. The amplitude of load rejectionovervoltages is usually not constant during its duration. Accurate calculations have to consider many parameters,

the following typical values of such overvoltages may be considered

In moderately extended systems, a full load rejection can give rise to phase-to-earth overvoltages with amplitude

usually below 1.2 p.u. The overvoltage duration depends on the operation of voltage-control equipment and may be

up to several minutes.

In extended systems, after a full load rejection, the phase-to-earth overvoltages may reach 1.5 p.u. or even more

when Ferranti or resonance effects occur. Their duration may be in the order of some seconds.




Page 12

For load rejection of generator transformers, the temporary overvoltages may reach amplitudes up to 1.4 p.u. for

turbo generators and up to 1.5 p.u. for hydro generators. The duration is approximately 3 s.

TOVs from following causes may also require consideration depending on the nature of the network:

- Voltage rise along long unloaded lines (Ferranti effect).

- Harmonic overvoltages, e.g. DC converters or saturated transformers.

- Back feed through interconnected transformer windings, e.g. dual transformer station with common secondary

bus during fault clearing or single-phase switched three-phase transformer with an unbalanced secondary load.

Resonance: Linear resonance can be either series or parallel which could involve both large currents and/or

voltages. Ferro-resonance modes may be sub-harmonic or harmonic with the latter producing higher temporary

overvoltages. Temporary overvoltages from resonance should not form the basis for the surge arrester selection.

The use of a surge arrester to damp out resonance is not effective and unproven.

Combination of TOVs: Combinations of TOVs such as earth faults and load rejection may result in higher

temporary overvoltage values than those from a single event. When combinations are considered sufficientlyprobable, overvoltages from each cause have to be compounded taking into account actual system configuration.

Severe TOV Cases: Occurrences of severe or repetitive TOVs are possible when different voltage circuits share

common towers and flashovers occur between high and low voltage conductors and possibly with multiple re-close

operations. While it is practical to provide limited SA protection for moderate TOVs of short duration, more severe

and/or sustained cases could cause multiple equipment and SA failures. To safeguard against such events it might

be worthwhile to install some designated SAs with lower protective level to relieve stress on parallel SAs and

equipment.

1.2.3 SLOW FRONT/SWITCHING OVERVOLTAGESSlow front overvoltages generated by switching have traditionally been associated with lines, load or fault clearing.

With the emergence of f lexible AC transmission system (FACTS) devices such as static var compensators(SVC),

static compensators (STATCOM), and thyristor controlled series capacitors(TCSC) which require power-electronictypes of switching and SA protection for both the electronic and power components.

1.2.3.1 Line ReclosingRandom high-speed reclosing on transmission lines with trapped charges generates travelling waves on the phase

conductors which may cause insulator flashover(s) to the tower(s) along the line if not controlled. Especially critical

is the case at remote end without terminal equipment such as shunt reactors, transformers or SAs which may

cause a doubling of the incident surge.

There are various methods of controlling line switching overvoltages, like traditional closing resistors, line switching

surge arresters and more recently, circuit breaker (CB) controllers. With advancements in SA technology, low

protective level line arresters with high energy capacity have been introduced which when combined with CB

staggered-pole closing is considered adequate for limiting switching overvoltages on short to medium line lengths.

Alternatively, modern CBs have fewer interrupters and more consistent point-on-wave (POW) switching capabilities

so that controlled switching is now a practical and economical option.

Unlike lightning related applications where arresters may be installed at consecutive structures, arresters to control

switching surges are only needed at both ends of the line and possibly one or two other locations along the line

depending on the SIWL of the line insulation, arrester protective level and the length of the line. For one or two

point installations, arresters are applied near the midpoint or approximately one third and two thirds of the line

length, respectively.

In theory, with recent developments of intelligent multi-purpose POW CB controllers [refs: ABB CATCO, AREVA

RPH3, SIEMENS PSD3] and emergence of modern CBs with precision and consistent closing capabilities,

switching overvoltages seem to be virtually eliminated, regardless of length. In practice, however, line SAs are still




Page 13

considered necessary to ensure line switching performance and reliability when controllers are out-of-service or

CBs misoperate.

1.2.3.2 Capacitive LoadsCapacitor bank energization can generate both voltage and current transients. Also, switches used for capacitive

switching have been improved to reduce restrikes but not eliminated. Therefore, some methods of reducing the

severity of transients are often incorporated such as current-limiting resistors or reactors, controlled switching and

SAs to minimize restrike possibilities and also provide overvoltage protection.

Controlled switching is typically reserved for large capacitor banks and/or networks with weak source strengths or

sensitive loads. Since controller and/or CB misoperations and local faults are possible, both current limiting

reactors and SAs may be applied to reduce the severity of inrush and outrush transients.

SAs may be applied in parallel with capacitors and sometimes with current-limiting reactors or more often phase-to-

ground for overvoltage protection. During re-strikes, unprotected capacitor overvoltages can exceed 3 p.u. while

SAs can reduce them well below 3 p.u. SAs applied with low voltage current-limiting reactors can be subjected to

both fast and slow front transient stresses during lightning and normal switching operations but severest stress is

still encountered during restrike conditions.

1.2.3.3 Inductive loadsSwitching small inductive load currents is considered a challenge for CBs designed for interrupting large fault

currents. Very fast reignition and/or restrike transients can also damage wound equipment like shunt reactors and

transformers due to uneven winding voltage distribution.

Simple control methods can be applied with CBs on large reactors for de-energization which minimizes reignition

possibilities by advancing the contact opening or equivalently increasing the switch’s dielectric strength before

interruption takes place naturally near current zero. Special controllers for transformer energization are now

available which monitors remanent flux left on transformers and closes optimally with flux conditions to minimize

energization disturbances. SAs are also applied with reactive equipment to reduce switch restrike possibilities and

also provide basic overvoltage protection.

Large switching transients can appear across phase-phase insulation of terminal equipment connected to

simultaneously switched lines or shunt capacitor banks. When special phase-phase switching transient

requirements have not been specified for such equipment, it might be necessary to apply switching controls and/or

additional SAs directly in parallel across those insulation to attain acceptable protective margins.

1.2.3.4 Flexible AC Transmission System (FACTS) DevicesPower electronics offer fast switching capabilities and unlike mechanical switches can be turned off before natural

current zeros. Therefore with proper design, it is possible for these devices to initiate system control within ½ cycle

of overvoltage or overload detection.

Some SAs are usually applied to protect major components from normal external stresses. FACTS devices are

frequently or continuously switched when in service and for security measures, SA protection is applied internally

as countermeasures against abnormal control and/or equipment misoperations. Besides evaluating arrester

protective and thermal energy requirements, transient interaction possibilities between SAs and power electronic

devices must be carefully examined.

1.2.4 CB AND DS TRVSInterrupting fault and even load currents can generate severe transient recovery voltages (TRV) across the switch.

If the TRVs are too fast or too large relative to the switch thermal or dielectric recovery rate, switches can re-ignite

or restrike, resulting in switch failure and/or damages to unprotected equipment caused by single or multiple

restrike transients.

Both switch TRV performance and equipment protection can be improved by appropriate SA application but

arrester failure from excessive energy absorption during multiple restrikes might be expected. Normally, SAs are

applied phase-to-ground on one or both terminals of the switch to limit respective overvoltages. In some special




Page 14

cases, SAs may be applied directly in parallel with the switch to provide more effective TRV control but continuous

or temporary overvoltages expected during open switch conditions must be carefully examined.




Page 15

1.3 Stresses from HVDC networks

Authors in charge: Bengt Johnnerfelt and Reinhard Göhler

1.3.1 INTRODUCTIONThe most significant differences for arrester applied in HVDC systems compared to normal AC-applications are the

wave shapes of the actual COV and TOV, and that the verification of the energy stresses becomes more complex.

For some applications the most severe energy stresses are sometimes not even followed by any significant service

voltage so that thermal instability cannot occur.

For the arresters indoors, for example in the valve hall, insulation withstand tests are not relevant and should be

skipped.

Continuous op erating voltages

As the HVDC arresters can be applied at a variety of different positions, with many different wave shapes of the

service voltage, it is unpractical to give the voltage values in r.m.s. values. Therefore the continuous operating

voltage always has to be given in crest values together with the wave shape. Hence additional definitions for the

operating voltages are needed.

- CCOV, which is the highest crest value of the continuous operating voltage excluding possible

commutation overshoots.

- Non-significant CCOV, which is a continuous operating voltage of such low amplitude that the power losses

generated can never initiate a thermal run-away after energy injections. Each manufacturer will have to

give its limit for significant power losses of their designs.

- PCOV, which is the highest crest value of the continuous operating voltage including possible commutation

overshoots.

- ECOV, which is the equivalent ac- or dc-voltage, used in operating duty tests, having at least the same

power losses as the actual CCOV at the actual temperature after energy injections.

At the AC-yard there are applications apart from the normal AC-bus, e. g. in the filters. These HVDC-arresters may

be stressed with sinusoidal voltages but with higher frequencies, like for example 3rd

, 5th, 11

th, 13

th harmonics of the

nominal power frequency.

In the converter station there are more complex wave shapes. Typically there are also commutation overshoots.

The wave shapes for valve arresters have a dc-component and a very short voltage peak of the opposite direction.

In the valve hall there may also be so called bridge arresters across 6- or 12-pulse groups. Their operating voltage

has commutation overshoots together with a high dc-component but with no voltage peak of opposite polarity. It is

important to notice that the commutation overshoots are more or less influenced by the arresters themselves

depending on the valve set up, so studies should be performed also with the arresters present to see the true

commutation overshoots, as the arresters may damp them significantly.

In the DC-yard there may also be filter arresters at different frequencies apart from the DC-bus arresters which see

a pure dc-voltage. Other applications may be across the smoothing reactor, which see a non-significant CCOV.

There may also be neutral bus arresters in the valve hall or in the DC-yard. They also have non-significant CCOV.

Accelerated ageing tests should be performed with wave shapes similar or obviously worse than the actual ones.

Also for determination of power losses at the actual wave shapes it is necessary to generate a variety of different

wave shapes in order to get the proper ECOV to use in operating duty tests. One way to solve this is to generate

similar wave shapes on the low voltage side and then amplify this voltage with an amplifier up to appropriate levels

for testing on individual ZnO-discs. For the filter applications a frequency generator can be used.

Worth noticing is also that some HVDC system may reverse the dc-polarity even after a long time with one polarity.

This means that there has to be shown in accelerated ageing tests, after 1000h testing, that the ZnO-discs, which

are exposed to CCOV with a high dc-component, can cope with this without causing excessive power losses or

other ageing effects.




Page 16

Energy stresses

The aim of the type testing for HVDC arresters is to verify both energy capability of the ZnO-discs themselves as

well as to verify their thermal stability after maximum energy stresses followed by CCOV. The operating duty tests

should in principal follow the IEC procedures, but standard test parameters for line discharge classes acc. to IECare typically not usable.

System studies are nearly always performed, resulting in energy stresses and typical transient wave shapes. For

HVDC applications it is proposed that the long-duration current impulse withstand test is substituted by a high

energy impulse withstand test with 6 impulses of the maximum energy requirement from the system studies,

separated by one minute apart. Then it has to be decided which test wave shape that best cover the actual energy

stresses; sinusoidal, half “sine wave” shapes or rectangular, which can typically be generated in arrester test labs

for operating duty tests. Guidance which wave shape to use can be found in Chapter 3, Energy Handling Issues.

The operating duty tests should be performed similar to the switching surge operating duty tests for line discharge

classes 4 and 5 in IEC 60099-4, with test samples preheated to 60ºC, but with the two long-duration current

impulses substituted with one energy impulse having at least the same energy as calculated in the system studies.

In some cases the maximum energy stress may also come from two consecutive impulses and then two energyimpulses one minute apart are used in the testing. After the energy injection, the test samples shall be exposed to

an a.c.- or d.c.-voltage voltage that generates the same or higher power losses as the actual wave shape. If there

are TOV stresses, the calculated energy from these stresses should be generated either by adding it to the energy

impulse or test the test samples with an equivalent ac- or dc-voltage generating the same or higher energy. TOV

stresses of the a.c.- or d.c.-bus arresters can of course be tested with the actual voltage and duration. It is also

recommended to use the same three test samples in both the long-duration current impulse withstand test and the

switching surge operating duty test.

It is not unusual that for the most severe fault scenario the converter station is closed down afterwards. In these

cases there is no need to verify thermal stability after this energy injection. But if there are other fault scenarios

followed by CCOV also this has to be verified. So for some arresters there may be one high energy value verified in

the high energy impulse withstand test and a lower energy verified in the operating duty tests.

In some cases like the neutral bus arrester application there is never any significant CCOV, so all the disc-tests canbe performed on open ZnO-disc sections. For this application, which often consists of several parallel columns,

there may also exists very rare fault scenarios with so extremely high energy requirements that it is more

economical to use the arrester as a sacrificial device with a failure of one or two columns. In this case special

consideration to the short-circuit tests may be necessary, in order to easily facilitate a restart of the converter

station.

Since the late 1970s, overvoltage protection of HVDC converter stations has been based exclusively on metal-

oxide surge arresters. This is due to their superior protection characteristics and their reliable performance when

connected in series or parallel with other arresters.

The basic principles when selecting the arrester arrangement are that:

- Overvoltages generated on the a.c. side should be limited by arresters on the a.c. side

- Overvoltages generated on the d.c. or earth electrode line should be limited by d.c. line arresters and

neutral bus arresters

- For overvoltages within the HVDC converter station, critical components should be directly protected by

arresters connected close to the components, such as valve arresters

Information about selection, application and testing of HVDC surge arresters is given in the “Application guide for

metal oxide surge arresters without gaps for HVDC converter stations” prepared by CIGRE working group 33/14-05

and published in 1986. Further information is given in IEC 60071-5 “Insulation Coordination – Part 5: Procedures

for high voltage direct current (HVDC) converter stations”. Parts of the CIGRE guideline are already included in IEC

60071-5




Page 17

The CIGRE guideline is divided into 7 chapters:

- Chapter 1: Scope

- Chapter 2: Metal oxide arresters characteristics

- Chapter 3: Arrester schemes and stresses on HVDC converter station arresters- Chapter 4: Studies for determination of arrester stresses

- Chapter 5: ZnO arrester to limit temporary overvoltages

- Chapter 6: Rules for determination of arrester capabilities and arrester test requirements

- Chapter 7: Arrester testing

1.3.2 STRESSES ON SURGE ARRESTERS

1.3.2.1Continuous operating voltagesThe continuous operating voltage for HVDC arresters differs from that for normal a.c. arresters in that it consists of

not simply the fundamental frequency voltage but rather of components of direct voltage, fundamental frequency

voltage and harmonic voltages, and high frequency transients (see Figure 1.1).

Special attention must be paid to the commutation overshoots caused by switching action of the valves withrespect to energy absorption in the valve arresters and other arresters on the d.c. side.

The continuous operating voltage waveform for the valve arrester is shown in Figure 1.2. The CCOV is proportional

to the Udim, and is given by:

0dim 23

vU U CCOV

Udio ideal no-load direct voltage (IEC 60633)

Udim maximum value of Udio

Uv0 no-load phase-to-phase voltage on the valve side of converter transformer, r.m.s. value

Operation with large delay angles increases the commutation overshoots.

(equation 1.1)




Page 18

Figure 1 .1: Typical waveforms of conti nuous operat ing volt ages at various locati ons in the convert erstat ion. Al l individual figures show the volt age over t ime.

1.3.2.2 Sources and types of overvoltagesOvervoltages on the a.c. side may originate from switching, faults, load rejection or lightning. Overvoltages on the

d.c. side may originate from either the a.c. system or the d.c. line or from in-station flashovers or other fault events.

1.3.2.2.1 Slow-front and temporary overvoltages

Slow-front and temporary overvoltages occurring on the a.c. side are important to the study of arrester applications.Together with the highest a.c. operating voltages they determine the overvoltage protection levels. Slow-front

overvoltages can be caused by switching of transformers, reactors, static vac compensators, a.c. filters and

capacitor banks and by fault initiation and fault clearing as well as by closing and reclosing of lines. Slow-front

overvoltages caused by events occurring close to the converter a.c. bus are relatively high in comparison to those

which originate at locations in the a.c. network remote from the HVDC converter station.

The d.c. side insulation co-ordination for slow-front overvoltages and temporary overvoltages is mainly determined

by fault on the d.c. side. Events to be considered include d.c. line-to-earth faults, d.c. side switching operations,

events resulting in an open earth electrode line, generation of superimposed a.c. voltages due to faults in the

converter control (e.g. complete loss of control pulses) misfiring, commutation failures, earth faults and short-

circuits within the converter unit.




Page 19

1.3.2.2.2 Fast-front, very fast-front and steep-front overvoltagesTravelling waves such as those caused by lightning strokes on the a.c. side or on the d.c. line are attenuated due

to the presence of a.c. filters, d.c. filters, large shunt capacitor banks, series reactance and shunt capacitance to

earth. Steep-front overvoltages caused by earth faults in the HVDC converter station, including locations inside the

valve hall, are important for insulation co-ordination. These overvoltages typically have a front time of the order

0,5 µs to 1,0 µs and durations up to 10 µs. In the a.c. switchyard section, very fast-front overvoltages with front

times of 5 ns to 150 ns may also be initiated by operation of disconnectors or circuit breakers in gas-insulated

switchgear (GIS).

1.3.3 CREEPAGE DISTANCE AND CLEARANCE IN AIRThe creepage distance on the insulators is one of the factors that dictate the performance of external insulations at

continuous operating voltages (a.c. or d.c.). Contamination on the insulators reduces their ability to support the

operating voltages, particularly during wet conditions. When wet weather conditions concentrate the pollution on

some parts of the surface of the insulators, the non-uniform distribution of pollution and increase in leakage current

creates dry zones resulting in uneven voltage stresses and this can initiate the process of flashover. Rain, snow,

dew or fog are some of the weather conditions that can initiate this process. The withstand capability ofcontaminated insulators is also affected by other factors such as the shed profile, the orientation angle and the

diameter of the insulators.

The base voltage used together with the specific creepage distance is as follows:

- for the insulation on the a.c. side of the converter (a.c. equipment): the highest value of operating voltage

expressed as the r.m.s. voltage phase-to-phase (IEC 60815);the minimum recommended creepage

distances are defined in terms of mm per kV (phase-phase). Typically the range is between 16 mm to 31

mm/kV.

- for the insulation on the d.c. side of the converter (d.c. equipment): the d.c. system voltage for the

insulation to earth, or a corresponding average value of the voltage across the insulation for insulations

between two energized parts.

The trend in the industry for several years has been to use larger specific creepage distances in HVDCapplications. For example, creepage distances as high as 60 mm/kV have been used in HVDC systems. However,

such an increase in the specific creepage distance did not eliminate the external flashovers.

The specific creepage distance of 60 mm/kV in a d.c. system corresponds to about 35 mm/kV in an a.c. system.

The use of composite housings for surge arresters has been successful also with smaller specific creepage

distances.

For an indoor clean environment, a minimum specific creepage distance of about 14 mm/kV has been widely used

and has not experienced any flashover.

For both d.c. and impulse voltages the positive polarity has lower withstand voltage than the negative polarity.

1.3.4 OVERVOLTAGE LIMITING CHARACTERISTICS OF ARRESTERSMetal-oxide surge arresters without gaps are used for the protection of equipment in HVDC converter stations.

These arresters provide superior overvoltage protection for equipment due to their low dynamic impedance and

high energy absorption capability. The ability of the metal-oxide arrester blocks to share arrester discharge energy

when connected in parallel if they are selected to have closely matched characteristics allows any desired

discharge energy capability to be realized. Metal-oxide blocks may be connected in several parallel paths within

one arrester unit and several arrester units may be connected in parallel to achieve the desired energy capability.

Also, parallel connection of metal-oxide blocks may be used to reduce the residual voltage of the arrester, if

required.

The protective characteristics of an arrester are defined by the residual arrester voltages for maximum steep-front,

lightning and switching current impulses that can occur in service.




Page 20

The amplitude of the current for which the protective level is specified, which is referred to as the co-ordinationcurrent, is usually selected differently for different types of current wave shapes and locations of the arresters.These co-ordination currents are determined from detailed studies carried out during the final stages of the design.

The arresters used on the a.c. side are usually specified as for arresters in a normal a.c. system by their ratedvoltage and maximum continuous operating voltage.

For the arresters on the d.c. side of a HVDC converter station, the rated voltage is not defined and continuousoperating voltage is defined differently because the voltage wave shape which continuously appears across thearresters consists, in many cases, of superimposed direct, fundamental and harmonic components and, in somecases, also commutation overshoots. The arresters are specified in terms of:

- PCOV peak continuous operating voltage- CCOV crest value of continuous operating voltage- ECOV equivalent continuous operating voltage

This means that the tests specified for these arresters shall be adjusted for the particular applications, differentfrom standard tests usually applicable for a.c. arresters.

The required energy capability of the arresters shall consider the applicable wave shapes as well as theamplitudes, duration and the number of respective discharges.

Figure 1.2: Opera ting voltage of a valve arrester, recti fier operation




Page 21

1.3.5 SURGE ARRESTERS IN A CONVERTER STATION An HVDC converter station includes a number of different surge arresters for protection of the different pieces of

equipment. There are basically six types of surge arresters, which are commonly denominated by the letters “A”

through “F”.

Figure 1.3 : Different types of surge arresters in a HVDC convert er stat ion

Event Arresters

A B C D E F

Earth fault d.c. pole X X X

Lightning from d.c. line X X X

Slow-front overvoltage from d.c. line X X X

Lightning from earthed electrode line X

Earth fault a.c. phase on valve side X X X

Current extinction X X

Loss of return path, monopolar operation or commutation failure X

Earth faults and switching operation, a.c. side X X X X X

Lightning from a.c. system X X

Station shielding failure (if applicable) X X

Table 1.1: Events stressing the different arresters




Page 22

1.3.5.1 AC BUS ARRESTER (TYPE A)The a.c. side of an HVDC converter station is protected by arresters at the converter transformers and at other

locations. These arresters are designed according to the criteria for a.c. applications.

1.3.5.2 VALVE ARRESTER (TYPE B)The dimensioning of the “B”- and “C”-arresters for protection of the semiconductors of the valve tower is particularly

critical. On the one hand, the protection level must be maintained as low as possible in order to protect the very

sensitive semiconductors and to minimize the number of these very costly components. On the other hand, the

voltage and current wave shape across the arresters is extremely non sinusoidal and dependent on the load

conditions and power flow of the HVDC converter station. As a consequence, the power dissipation of the arrester

is variable with the load conditions and it is difficult to find the right compromise between protection of the valve

tower and safe operation of the arrester. Simulation of the HVDC station including the various possible faults is an

important tool for determination of the arrester voltage and current stress.

The valve arrester continuous operating voltage consists of sine wave sections with commutation overshoots (see

Figure 1.1). The peak continuous operating voltage (PCOV), which includes the commutation overshoot, shall be

considered when the reference voltage of the arrester is determined. The commutation overshoot is dependent on

the firing angle.

The maximum temporary overvoltages are transferred from the a.c. side during fault clearances combined with load

rejections close to the HVDC converter station.

The events producing significant valve arrester currents of switching character are as follows:

- earth fault between the converter transformer and the valve in the commutating group at highest potential;

- clearing of an a.c. fault close to the HVDC converter station;

- current extinction in only one commutating group (if applicable).

Depending on current rating, control system dynamics, inductance of the d.c. reactor, and the protection scheme,

the phase to earth fault will be dimensioning for the energy and current rating of the arresters.

The valve arresters can in general only be subject to fast-front and steep-fronted overvoltages at back-flashoversand earth faults within the converter area. The most critical case for steep-front overvoltages is normally an earth

fault on the valve side of the converter transformer of the bridge with the highest d.c. potential.

1.3.5.3 Converter unit arrester (Type C) A converter unit arrester may be connected between the d.c. terminals of a 12-pulse bridge. The maximum

operating voltage is composed of the maximum direct voltage from one converter unit plus the 12-pulse ripple.

The converter unit arresters are normally not exposed to high discharge currents of switching character. The

arrester may limit overvoltages due to lightning stresses propagating into the valve area, although these stresses

are not decisive for the arrester.

1.3.5.4 DC bus and DC line/cable arrester (Type D; DB and DL)

The maximum operating voltage is almost a pure d.c. voltage. These arresters are mainly subjected to lightningstresses. Critical slow-front overvoltages can often be avoided by suitable selection of the parameters in the main

circuit, thus avoiding critical resonances.

When the HVDC line comprises overhead line sections as well as cable sections, consideration should be given to

the application of surge arresters at the cable-overhead line junction to prevent excessive overvoltages on the

cable due to reflection of travelling waves. At HVDC links with very long cables, the energy rating of the cable

arresters is decided by the discharge of the cable from the highest voltage.




Page 23

1.3.5.5 Neutral bus arrester (Type E)The operating voltage of the neutral bus arrester is normally low. These arresters are provided to protect

equipment from fast-front overvoltages entering the neutral bus and to discharge large energies during the

following contingencies:

- earth fault on the d.c. bus;

- earth fault between the valves and the converter transformer;

- loss of return path during monopolar operation.

An earth fault on the d.c. bus will cause the d.c. filter to discharge through the neutral bus arrester, giving a very

high but short current peak.

1.3.5.6 AC and DC filter arrester (Type F; FA/FD)The ratings of a.c. arresters are normally determined by the transient events. The events to be considered with

respect to filter arrester duties are slow-front plus temporary overvoltages on the a.c. bus and discharge of the filter

capacitors during earth faults on the filter bus.

The normal operating voltage of the d.c. filter reactor arrester is low. Arrester duties are mainly determined by filtercapacitor discharge transients resulting from earth faults on the d.c. pole.




Page 24

1.4 Stresses in traction systems

Author in charge: Bernhard Richter

1.4.1 GENERALThe overvoltage protection of the electrical traction systems has an increasing importance nowadays. This is not

only by the railways which are supplied with a.c. voltage but also increasingly by the d.c. railways.

The long-distance railway system is electrified with 3 kV d.c. voltage on over 70 000 km rails (in year 2000, that

means about 38% of the total length of the rails of the electrical railways) and with 1,5 kV d.c. voltage on more than

20 000 km (about 11%). That means that about half of the world-wide railway length of the long-distance traffic is

operated with direct-current. The length of the electrified rails by the outer suburban service, including local trains,

which operate with a d.c. voltage under 1000 V, is about 25 000 km. These figures show the extent of the d.c.

voltage systems by railways and also the importance of an optimal overvoltage protection which is adjusted to the

specific demands of the d.c. voltage railways.

Increasing use of electronic equipment in and close to the rails and overhead lines (safety and signaling

equipment) need protection against overvoltages. Further on each breakdown of the power supply leads to an

interruption of the train service.

Lightning strokes are the most dangerous threats for railway networks. Overhead lines and trains can be hit by

direct or nearby lightning. For this reason very high charges can be transferred into the overhead lines and the

installed surge protective devices have to withstand high energy stresses. On the other hand the continuous

voltage in d.c. traction systems is naturally a d.c. voltage, which means that all surge protective devices have to be

designed for d.c. application. For this reason MO surge arresters without gaps are used in the d.c. power supply of

the traction systems. It stands for itself that the used MO surge arresters have to be long term stable under d.c.

voltage stress.

The application and dimensioning of metal oxide surge arresters (MO surge arresters) without spark-gaps in

alternating current networks with 50/60 Hz and 16,7 Hz of the railway supply is not very different from the one of

the general energy supply. Requirements and tests for MO surge arresters for application in a.c. traction systems

are similar to the ones for MO surge arresters without gaps for three phase power systems and [IEC 2009] applies. A separate international standard is not available.

As can be seen in Table 1.2 the voltages in traction systems have a strong fluctuation depending on the load in the

system, which is given by the number of accelerating and breaking trains in a power supply section. Due to this fact

the voltage Umax2 should be considered the relevant precondition when choosing the continuous voltage Uc. This

applies for the a.c. as well as for the d.c. system.

In general the electrical and mechanical requirements for MO surge arresters for application in traction systems are

very high. Arresters installed on traction vehicles have to withstand high mechanical stresses, especially vibrations,

mechanical shocks and high wind loads in case of high speed trains have to be considered.

Because rails, trains and train stations are public places the safety of the surge arresters is an important point. The

arresters should have fail-safe performance in case of an overload.

The electrical requirements and tests for MO surge arresters for application in d.c. traction systems are given in the

new European Standard EN 50526-1 [EN 2012].




Page 25

1.4.2 VOLTAGES IN TRACTION SYSTEMSThe values for the operating voltages of the railway facilities with the admissible deviations are defined in the

European Standard EN 50163. The most important values and definitions are given in the following Table 1.2 and

Figure 1.4.

Nominal voltage

Un

V

Highest permanent

voltage

Umax1

V

Highest non-permanent

voltage

Umax2

V

Umax3

V

DC systems (mean values)

600 720 770 1015

750 900 950 1269

1500 1800 1950 2538

3000 3600 3900 5075

AC systems (r.m.s. values)

15 000 (16,7 Hz) 17 250 18 000 24 311

25 000 (50 Hz) 27 500 29 000 38 746

Note 1: Umax3 is a calculated value for an overvoltage at t = 20 ms.

Note 2: The voltage values for Umax2 can become 800 V in the 600 volt net and 1 000 V in the 750 V net, in case ofregenerative breaking.

Table 1.2: Voltages in traction systems

Figure 1.4 : Highest values of volt age occurr ing in the system depending on time durati on




Page 26

The definitions of the voltages are as follows:

- Un : designated voltage for a system.

- Umax1: maximum value of the voltage likely to be present indefinitely.

- Umax2: maximum value of the voltage likely to be present for maximum 5 min.

Figure 1.5 shows as an example the possible voltage fluctuations on the current collector of a metropolitan d.c.

traction system. For the purpose of overvoltage protection only the maximum values are of importance, because

they designate the values for the design of the MO surge arresters used for the protection in traction systems.

Figure 1.5: Vol tage at the current collector of a d.c. traction vehicle for a period of 15 min, urbant ransport at ion system

1.4.3 MO SURGE ARRESTERS FOR D.C. TRACTION SYSTEMSThe surge arresters are classified by their charge transfer capability Qt and their nominal discharge current In.

The classes DC-A, DC-B and DC-C correspond to increasing discharge requirements. The selection of the

appropriate class shall be based on system requirements.

Class DC-A has a nominal current of In = 10 kA and a charge transfer capability of Qt = 1.0 As

Class DC-B has a nominal current of In = 10 kA and a charge transfer capability of Qt = 2.5 As

Class DC-C has a nominal current of In = 20 kA and a charge transfer capability of Q

t = 7.5 As

An optional test is intended to prove the ability of the arrester to withstand direct lightning currents. The

requirements for the direct lightning current Iimp are 2 kA peak value for class DC-A, 5 kA peak value for class DC-B

and 15 kA peak value for the class DC-C.

1.4.4 MO SURGE ARRESTERS FOR A.C. TRACTION SYSTEMS AC surge arresters for traction systems are classified in the same way as MO surge arresters for three phase

power systems according [IEC 2009] and [IEC 2000].




Page 27

1.5 Stresses from Lightning

Authors in charge: Trond Ohnstad and Yoshihiro Ishizaki

1.5.1 INTRODUCTIONLightning and thunderstorms occur all over the world, from far north in Norway to far south in South-Africa. The

heaviest thunderstorms with the most intense lightning will normally be experienced in an area of about ± 2000 km

along the equator.

Lightning has always been a problem to telecommunication and electricity systems and surge arresters have

become an important asset to protect people and equipment against dangerous over-voltages caused by lightning.

As a general rule surge arresters are installed close to the equipment it shall protect.

1.5.2 LIGHTNING SURGES.Surge arresters in electrical systems are due to stress caused by lightning surges in case of:

- Lightning stroke to an incoming power line.

- Direct stroke to the substation.

- Induced voltage from a nearby lightning stroke.

During a thunder storm a power line could probably be hit by several l ightning strokes, either directly to the phase

conductors or most likely to the shielding wires, if any present.

The strokes will cause an earth fault on the line, and initiate a switching sequence of the circuit breakers in the

substations at both line ends.

The following lightning surges and switching surges will eventually stress the surge arresters at the line ends or

elsewhere in the substations.

The degree of stress to the arrester depends on several factors like: distance to the place of l ightning stroke,

striking point on the voltage curve, lightning current amplitude, earth resistance, the total flash charge, tower

earthing impedance and if the power line has shielding wires or not.

Different electricity systems with different voltage levels will see different levels of lightning surges and cause

different level of stress to the arresters.

1.5.2.1 LV- (Us up to 1 kV) and MV- systems (1 kV < Us 52kV)Power distribution lines are generally of lower height and less exposed to direct flashes than transmission lines with

higher voltage. The substations are often in-house and well protected against direct strokes to the bus bars. The

number of overvoltages exceeding the basic insulation level in these systems is dominated by induced

overvoltages caused by lightning strokes to or in the surroundings. Due to a large number of surge arresters used

in these systems the energy is split on several units and failures due to stress caused by lightning are rare. Surge

arresters for these systems are rather small and cheap and are easy to keep as spares and to replace when

necessary.

1.5.2.2 HV-systems (52kV < Us 245kV)Systems within these voltage levels consists of both distribution and transmission lines but often still in a rural area

where the lines and substations to a certain degree are protected against lightning by surrounded houses, towers

and trees. Combinations of surge arrester stress due to direct strokes, back-flashovers and induced voltages will

statistically result in a higher failure rate caused by lightning than in any other systems.

1.5.2.3 EHV-systems (245kV < U s 800kV) and UHV-systems above 800kVTransmission lines with steel towers and shield wires are in spite of the height above ground well protected against

direct lightning strokes to the phase lines. Most of the lightning will hit the towers or the shield wires, and only a

back-flashover will cause a critical surge in the phase-line.




Page 28

Operation experience from different utilities in different places on earth implies surge arrester failure due to

lightning only in cases with very close and nearby strokes to the substation or in an incoming transmission line

without shielding wires. The surge arrester energy capacity is dimensioned according to the much higher energy in

switching surges and normally the much lower stress from lightning is not a problem.

System voltage Us kV

Insulation characteristic Lightning overvoltage Surge arrester stress

0.23 - 52 BIL < induced lightningsurges.Determined by lightningovervoltages

Induced surges can causeflashover.Direct stroke to a line orsubstation is very rare.

No critical stress due tolightning.

52 - 245 BIL determined bylightning overvoltages.

Induced surges can causeflashover.Direct stroke to shieldwires with back flashover,and direct stroke to phaseconductors.

Critical stress fromlightning can occur.High energy ratingsneeded.

245 - 800 and above BIL > induced lightning

surges.BIL > surges caused bydirect stroke to the phase.BIL determined byswitching surges.

Direct stroke to the power

line, to a phase conductoror shield wire.

Critical stress only due to

no shielding or shieldingfailure or high earthresistance combined witha nearby stroke.

Table 1.3: Lightning stresses in different system voltages

1.5.3 EXAMPLES FROM TRANSIENT ANALYSIS.

1.5.3.1 Norwegian 145kV and 420kV system. A study of energy stress on surge arresters due to lightning in a 145kV and a 420kV substation [Tra 1994],

concludes the following about the average accumulated stress during a thunder storm:

- 145 kV without shielding wires 2,20 kJ / kV

- 145 kV with shielding wires 0,02 kJ / kV

- 420 kV with shielding wires 0,03 kJ / kV

According to [Fuk 1997] only lightning strokes hitting the line without shielding wires within 2 km to the substation

will cause critical level of stress to the arresters, about 60 kJ/kV.

In power systems with voltage > 100kV the modern metal oxide surge arrester will be able to withstand the stress

caused by lightning as long as it is not a lightning stroke directly or close to the arrester in a power line without

shielding wires.

1.5.3.2 Norwegian line arresters 300kV

A study of energy stress due to lightning on transmission line surge arresters in a 300kV line in the south part ofNorway gave the following energy levels :

- Direct stroke to the phase wire (10kA) 0.88kJ/kV

- Stroke to the shielding wire (100kA) 0.04kJ/kV

- Stroke to a tower without shielding wires 15.8kJ/kV

1.5.3.3 115kV Transmission Line surge Arresters An EPRI study [Bir 1997] of lightning on an 115kV transmission line with surge arresters calculated the energy

dissipated in the surge arresters with stroke to the shielding wire and direct stroke to a phase wire:

- Direct stroke to a phase wire (50kA) 2.95 kJ/kV




Page 29

- Stroke to a shield wire (50kA) 0.10 kJ/kV

1.5.4 LIGHTNING STATISTICSSince 10-20 years several countries throughout the world have established a lightning location and registration

system which include information about lightning current amplitude, polarity and striking site and time.

Information from these systems are very important establishing databases and to get good lightning statistics.

Flash density per square km and year is an important input in all analysis concerning probability of failures due to

lightning.

Area/Country Maximum flash

density

50% value

positive

50% value

negative

50% value

total

Norway 1 23 17 18

Norway

west coast

0.6 37 22 24

Japan (winter) - - - 24

(51kA 16% value)

Cigré (all world) 0.1 - 200 - - 35

Table 1.4: Lightning parameters in different parts of the world

1.5.5 WINTER LIGHTNING.

1.5.5.1 GeneralNorway as well as Japan experience rather often thunder storms during winter.

In Norway the winter lightning occurs most frequent along the west coast and not very far inland from the sea.

Typical weather conditions to create the thunder storms are strong winds from the west which bring rather warm air

from the ocean in to the Norwegian mainland. The warm air is pressed upwards and collides with the much colder

air coming from the mountains and the Norwegian inland. Normally heavy clouds are building up and eventually

initiate thunderstorms. Observations from the Norwegian lightning location and registration system shows a

proportional higher number of positive lightning strokes during the winter storms than during the summer storms.

The positive strokes have in general higher energy and higher lightning current than the negative strokes.

Along the northern part of the west coast the winter lightning occurs more frequent than summer lightning, but in

the southern part the summer lightning is much more common and more frequent. In Norway there are no

indications of more failures in the electricity systems caused by lightning in the winter than in the summer.

Winter lightning studies from the engineering side started in 1978 in Japan, because the winter lightning occurred

primarily along the coast of the Sea of Japan which brought about peculiar lightning faults on EHV transmission

lines. The characteristics of winter lightning are described below, based on measurement results so far and are

compared with summer lightning.




Page 30

1.5.5.2 Characteristics of Winter Thunderstorm[Mic 2007]

Figure 1.6 shows the schematic evolution of winter thunderclouds. The winter thunderclouds are smaller compared

with the summer thunderclouds.

The characteristic features of the electrical activity of winter thunderclouds have been extensively investigated.

The representative main results are as follows:

- The duration of the lightning activity of an individual storm is short (usually less than 30 min) and the

frequency of lightning discharge is very low.

- About 30 – 40% of all ground flashes lower from cloud positive charge. This percentage is remarkably high

compared with that of a few percent for summer thunderclouds.

- Ground flashes of more than 300 C are occasionally observed [Hac 2008].

Winter thunderclouds in this area are formed by the advection of Siberian air masses, which are dry polar air

masses, over the relatively warm Sea of Japan.

Figure 1.6: Schematic chart illustrating radar echoes associated w ith the cycle of a t hundercloud(Chisholm and Renick, 19 72 ); in the low er part of this figure the temporal var iat ion of each radar echo

reflectiv it y in a convective cloud is illust rated in correspondence wit h the echo lif e stages.

1.5.5.3 Lightning current parameters

Table 1.5 summarizes typical lightning current parameters in Japan. As is widely known around the world, lightningin Japan can be broadly classified into summer lightning, which occurs frequently in the summer months, both in

the mountains and flatlands, and winter lightning, which occurs frequently in the winter months, primarily along the

coast of the Sea of Japan. Consequently, the Lightning Protection Design Guide for Transmission Line must

propose designs that take into account the characteristics of both types of lightning.




Page 31

Table 1.5: Characteristic comparison between summer and winter lightning [AIE 1950, Uma,Uma 1987]

1.5.5.4 Probability distribution of lightning stroke peak currentFor the cumulative probability distribution of lightning stroke peak current, a variety of them have been proposed

mainly targeting at summer l ightning, both in Japan and abroad CRI 1976 . In Figure 1.7, curve 1 is the one

recommended in the old guide [AIE 1950] (logarithmic normal type: Average value=26kA and logI=0.325), while

curve 5 is an AIEE distribution curve [Uma ] (exponential type). In addition, curve 2 is a distribution created based

on both positive and negative polarity data from 103 winter lightning strokes recorded between January 1979 and

July 1986 in Kashiwazaki and Fukui on the coast of the Sea of Japan. Note that since the cumulative frequency

distribution of lightning peak current differs by the geographical region, we must consider distributions that are

appropriate for each one.

Based on a comparison of probability distributions, this guide recommends curve 1 given by equation 1.2, which

agrees with the observational results for both summer and winter lightning, as the probability distribution of

lightning stroke peak current.

Note that it is possible to calculate the probability for any current value with the above equation, but we do not

recommend it for portions exceeding the range in Figure 1.7.

(equation 1.2)




Page 32

Figure 1 .7: Various data of cumulative probabilit y dist ribut ion of lightning stroke peak current[CRI1976]

The winter lightning phenomenon along the coast of the Sea of Japan occurs, is specific in this region on the globe.

The increase of positive ground flashes and frequent occurrence of upward lightning are considered as main

features of the winter lightning in this region.

1.5.6 PARAMETERS OF SUMMER AND WINTER LIGHTNING CURRENT

1.5.6.1 Wave front and wave tail duration of lightning currentsCumulative probability distribution curves of wave front duration of the summer and winter lightning current are

shown in Figure 1.8 [Ike 1981] [Asa 1994]. The 2 s value is usually used to represent the wave front duration of

summer lightning and winter lightning current, but the actual 50% value is longer than 2 s as shown in Figure 1.8.

As for the winter lightning it is furthermore long. The 22 examples of winter lightning discharges, which accompany

strong electromagnetic emission and heavy current, are classified and shown in the figure as type A.

p r o b a

b i l i t y ( % )




Page 33

Figure 1.8: Probability of the wave front duration of lightning current [Ike 1981 ] [A sa 1994]

As for the range of wave tail duration of lightning current, it is approximately 10~100 s with the 50% value of

30~50 s [Ike 1981].

The observation result for the winter lightning current in Japan is shown in Figure 1.9 [CRI 1989]. The figure

suggests the following:

- 50% value of wave tail duration is 50 s.

- 50% value of wave tail duration of negative polarity lightning is only 25 s.- 10% of negative polarity currents have longer wave tail duration than about 100 s.

- 50% value of wave tail duration of positive polarity lightning is 650 s, and 10% value of it exceeds several

thousand s.

Summer lightning Winter lightning

p r o b a b i l i t y

p r o b a b i l i t y ( % )

Wave front duration µs Wave front duration µs

Data of Fukui &

Kashiwazaki Type A




Page 34

Figure 1.9: Probability of li ghtning current wave tail duration of w inter lightning [CRI 1989 ]

As understood from the relationship between wave front duration and wave tail duration in Figure 1.10, the longerthe wave front duration, the wave front duration also tends to be longer. According to data of Figure 1.11, whichwere measured by Berger [Ber 1975], the 10% value of the wave tail duration of negative first stroke currentsexceeds 150 s. When focused on the positive polarity lightning, the 10% value exceeds 1 ms.

Figure 1.10 : Relationship between the wave front and wave tail durat ion of t he lightning current ofw inter lightning [CRI 198 9]

Wave tail duration50%: 50.1 µs16%: 794 µs


Wave tail duration µs

Wave front duration µs

W a v e t a i l d u r a t i o n µ s




Page 35

Figure 1.11 : Probability of wave tail duration of lightning current [ Ber 1975]

Figure 1.12 : Probability of electric charges of w inter lightning current [Miy 1992]

1: negative first lightning stroke

2: negative

3: positive lightning stroke

Wave tail duration µs



Amount of electric charge (C)

: Total

30% :2.9C




Page 36

Figure 1.13: Example of posit ive polarit y current shape of w inter lightning measured at Fukui in Japan inFebruary 1983 [CRI 1995]

1.5.6.2 Electric charge of lightning currents

The electric charge quantity of lightning is calculated by time integration of the lightning current. An example of thecumulative probability distribution of the electric charge quantity of the winter lightning current is shown in Figure

1.12 [Miy 1992]. The figure suggests the following:

- Approximately 10% of the winter lightning currents have the electric charge quantity exceeding 100 C.

- The 50% value of the electric charge quantity of the positive polarity lightning is 20 C, which is 20 times of

the 50% value of negative polarity lightning.

- Lightning currents with large electric charge quantity have very long wave tail duration rather than high

crest values. Figure 1.13 shows an example of the current wave shape of positive polarity with duration

exceeding 4 ms observed in winter.

- According to electrical charge data of summer lightning shown in Figure 1.14 measured by Berger [Ber

1975], the 50% value of the electrical charge of positive polarity lightning is 80 C, which is 10 times or more

of the 50% value of 7.5 C for the negative polarity lightning.

Figure 1.14: Probabili ty of electr ical charges of summer lightning [Ber 1975]

Time

C u r r e n t

e a k

v a l u e

1: negative first l ightning stroke

3: positive lightning stroke


Electrical charge C




Page 37

1.6 Ambient stresses

1.6.1 MECHANICAL STRESSES Author in charge: Shinji Ishibe

Arresters are usually mounted vertically or horizontally, but there are various erection alternatives (suspension,

under-hung, etc.) especially in polymer arresters. Mechanical stresses occurred in arresters are strongly dependent

on the erection configuration and external forces and can be categorized into static load, vibration load and seismic

load.

(1) Static load

The following loads are considered with regard to arresters in the actual fields.

a) Electromagnetic force

Since the continuous current thorough an arrester is of the order of a few mA, electromagnetic forces are usually

not considered. When the arrester is directly connected to the main circuit that may receive electromagnetic forces,

the effect of these forces on arresters should be considered.

b) Thermal effect

This is the load due to a thermal expansion of the line conductor corresponding to ambient temperature change or

main current change. Arresters should be connected by flexible leads so as not to be applied such loads.

c) Load during assembling

This load may be applied when a lead is connected to the arrester top. The installation work shall be carried out

taking the cantilever strength of the arrester into consideration.

d) Line pull

Flexible leads should be used to connect arresters so as to minimize the loads due to the weight of the leads.

e) Wind load

The wind speed of 34 m/s in IEC 60099-4 and 40 m/s in JEC-2371 are specified. But wind loads are generally not a

problem for arresters [100–300 N/m (per unit length of porcelain type arrester) at the wind speed of 40 m/s].

e) Snow load

The effect of snow loads is larger in the horizontally mounted arresters than in the vertically mounted arresters.

According to the installation configuration the effect of the snow on the lead should be considered

The mechanical strength of arresters against the above loads is typically evaluated by the cantilever strength. The

IEC 60099-4 standard covers the test method of the cantilever strength, which is a manufacturer’s declared value.

Porcelain type arresters have basically strong cantilever strength and the above loads may not be a great concern.

But in case of polymer arresters, the cantilever strength of which is usually 1000–3000 Nm may be sensitive to

these loads depending on the installation configuration. Therefore a guide for the cantilever strength of arresters is

required for manufacturers to design the mechanical strength of their arresters and for users to design the

installation and connection of arresters.

(2) Vibration load

This load would be severe if the vibration frequency were close to the natural frequencies of the internal part of the

arrester. A large vibration load may cause shifts or cracks of ZnO elements.




Page 38

a) Transportation vibration

Arresters are transported by trucks, trailers, freighters or airplanes and subjected to their inherent frequencies and

shocks due to bad roads, blocks, sudden braking or landing. These loads are usually taken into consideration in

the construction of arresters: e.g. insulators for supporting ZnO elements or bumpers between ZnO elements andthe porcelain wall. In case of polymer arresters where ZnO elements are directly molded, these loads may not be

large.

Some manufacturers evaluate the capability against these loads by the actual transportation test, where a truck or

trailer runs through normal roads, highways and bad roads. The vibration test with simulated waveforms is also

carried out instead of the actual transportation test. Shocks watching labels or meters are sometimes used for

monitoring loads when a high shock is expected during transportation.

b) Suspension vibration

A suspension type arrester may be subjected to a continuous vibration of its natural frequency due to the wind or

the movement of the line. This load may be small for the arrester body because acceleration is not so high and the

natural frequency of the suspension system is lower than that of the arrester, but the long-term reliability of the lead

connecting and arrester suspending parts should be confirmed.

(3) Seismic load

This load should be considered when the arrester is installed in the area where a large earthquake is expected. As

a seismic event is rare to occur, it is unreasonable to expect all of the above static loads to occur simultaneously.

There are three typical guides, IEC, IEEE and JEAG (Japan), which have been published and revised through the

experience of the large earthquakes as shown in Table 1.6.




Page 39

Year 1965- 1970- 1975- 1980- 1985- 1990- 1995- 2000- 2005-

JEAG

-5003

1st

(1980)

Revision

(1998)

IEEE

-693

1s

(1984)

Revision

(1997)

Revision

(2005)

IEC[Note]

68-2-57

(1989)

68-3-3

(1991)

1166

(1993)

1463

(1996)

62271-2

(2002)

Earthquake

in Japan

Niigata

(1964)

Off-

Tokachi

(1968)

Off-

Miyagi

(1978)

Nihonkai

-Chuubu

(1983)

Hokaido

-Nanseioki

(1993)

Hyogoken

-Nanbu

(1995)

Tottoriken

-Seibu

(2000)

Niigata

-Chuuetu

(2004)

Earthquake

in US

Sanfernand

(1971)

Northridge

(1994)

Other earthquake

Kocaeli,

Shu-shu

(1999)

Indonesia

Off-

Sumatra

(2004)

Pakistan

(2005)

[Note] IEC 68-2-57(IEC 60068-2-57): Time history method

IEC 68-3-3(IEC 60068-3-3): Seismic test methods for equipments

IEC 1166(IEC 62271-300): Seismic qualification of alternating current circuit breakers

IEC 1463(IEC 61463): Bushings – Seismic qualification

IEC 62271-2: High-voltage switchgear and control gear–Seismic qualification for voltages of 72.5 kV and above

Table 1.6: Publication and revision history of the guide for seismic test [JEA 1998]

Table 1.7 shows the seismic test conditions of IEC, IEEE and JEAG for arresters, where the acceleration response

of JEAG is the most severe. The concept of the JEAG with resonant 3 cycles sine wave is different from that of IEC

and ANSI with artificial earthquake waves which comply with the required response spectrum (RRS) as explained

below.

Seismic loads of arresters are also strongly dependent on the connecting leads [Oka 1986]. The guides of IEEE

and JEAG require sufficient flexible lead slack that allows for any relative deflection of the equipment that will occur

during an earthquake.

Static 0.5 GD namic 0.3 G




Page 40

Arrester standard IEC60099-4 IEEE C62.11 JEC-2371

Referred guide for seismic test

IEC62271-300

(Circuit breaker)

IEC60068-3-3

(General)

IEEE-693

(General)

JEAG-5003

(General)

Seismic test

method

Voltage rating Not detailed 90kV Test

(90kV

Analysis)170 kV

Input

acceleration

0.5 G High

0.3 G Moderate

0.2 G Low

0.5 G High

0.25 G Moderate 0.3 G

Frequency

range0.535 Hz 0.333 Hz 0.510 Hz

Waveform

Artificial earthquake

which complies with

RRS

Artificial earthquake

which complies with

RRS

Resonant 3 cycles

sine wave (Note 2)

Max. acceleration response

at 2% damping

(Single-degree-of-freedom-

system)

1.4G (High)

0.85G (Moderate)

0.56G (Low)

1.62G (High)

0.81G (Moderate)2.35G

Acceptance

criteria

For structural

parts

Specified

in the guide

Specified

in the guide

Specified

in the guide

For arrester

performanceNot detailed

(Note 1) Not detailed Not detailed

(Note 1) Only check items (reference voltage, partial discharge, leakage check) are listed in Annex M of IEC

60099-4.

(Note 2) When the inherent natural frequency of the equipment is higher (lower) than 10 Hz (0.5 Hz), the frequency

of 10 Hz (0.5 Hz) shall be used.

Table 1.7: Comparison of seismic test guides for arresters

RRSRRS




Page 41

(3.1) Test cond ition o f Japanese seismic test guide

The arrester standards of IEEE and JEC require having the seismic withstand capability according to the referred

guides. The IEC60099-4 standard requires the seismic test in case of the agreement between the manufacturerand user in Annex M and refers the IEC61166 (IEC62271-300), which is the guide for circuit breakers. As arrestersof higher rating have more sensitive and complex responses against the earthquake, the actual seismic test maybe necessary. Therefore the guide for the seismic test and evaluation method for arresters is required.

(3.1.1) Backgroun d of the test condition

(1) Acceleration

Figure 1.15 shows the map that indicates the value of horizontal acceleration on the ground surface expected inJapan once at the interval of 75 years (return period of 75 years). According to this figure, the horizontalaccelerations are less than 0.3 G in almost all regions in Japan.

Figure 1 .15 : Distr ibution of horizontal accelerat ion on the ground surface expected once at the intervalof 75 years [JEA 1998 ]

Table 1.7 shows the seismic intensity scale used by the Japanese Meteorological Agency. The acceleration of 0.3G (294 Gal) corresponds to the level of the seismic intensity .

Figure 1.16 shows the seismic records in the past 75 years from 1921 to 1995 in Japan. This figure shows that thepast records of the acceleration are below 0.3 G (294 Gal) in almost all regions in Japan. Please note that thementioned scale was revised in 1997.




Page 42

Intensity scale Designation Acceleration (Gal)

0 No feeling < 0.8

A slight earthquake 0.82.5

A weak earthquake 2.58

A rather strong earthquake 825

A strong earthquake 2580

A very strong earthquake 80250

A disastrous earthquake 250400

A very disastrous earthquake > 400

Table 1.8: Seismic intensity scale used in 1949 – 1997 by the Japanese Meteorological Agency[JEA 1998]

Figure 1 .16 : Seismic records in the past 75 years from 1921 to 1995 in Japan [JEA 1998 ]

7

2

Hokkaido

6

5

Tohoku

2 5

4

2

Kanto

93

3

Chubu

7

Hokuriku

95

7

Kansai

3

Chugoku

3

Chugoku

58

Kyushu

Sismic intensity scale

[Acceleration (Gal)][25-80] [80-250][250-400]




Page 43

(2) Waveform

As the waveform of an earthquake is dependent on the ground condition between the epicenter and the equipment,it is not practical to specify an earthquake waveform for the test. Porcelain type apparatuses are destroyed at the

peak acceleration of the response and the destruction is not affected by the duration and the waveform of thevibration, therefore the quasi-resonant method by resonant N cycle sine wave is adopted as the Japanese seismictest.

(2.1) Frequency

The predominant frequencies of earthquakes in Japan are 0.5–10 Hz, which overlaps the natural frequencies ofporcelain type apparatuses, bushings, arresters, etc. So the severest test condition is realized when the test waveis specified as a sine wave with the natural frequency of the apparatus.

(2.2) Cycle

The acceleration response factors of a simplified single-degree-of-freedom model to resonant 1–3 cycles sinewaves are compared with the actual 615 earthquake records on the ground surface in Figure 1.17. The response

factor in the resonant 2 cycles sine wave covers the most responses of the actual earthquake records.

In addition, the amplification effects due to the existence of foundations (1.2 times) and other unknown factors (1.1times) are considered. As the correction factor of 1.3 (1.2x1.1) corresponds to the ratio of the response factor toresonant 3 cycles sine wave to that to resonant 2 cycles sine wave (6.1/4.7 as shown in Figure 1.16), resonant 3cycles sine wave is adopted as the seismic test wave.

Figure 1.17: Comparison of t he response fact ors betw een a simpli fied single-degree model t o resonant1-3 cycles sine waves and the actual 615 eart hquake records in Japan [ JEA 1998 ]




Page 44

(3.1.2) Example of test

Table 1.9 and Figure 1.18 shows the seismic test result of the 500 kV GIS-arrester in accordance with JEAG-5003.

As the measured inherent natural frequency of 17 Hz was higher than 10 Hz, the 3 cycles sine wave of 10 Hz was

applied. The test with the EL Centro earthquake wave was also carried out for reference.

Direction

ItemX Y Z

Natural frequency [Hz] 24 17 > 30

Seismic test ofJEAG-5003

Wave shape 3 cycles sine wave of 10 Hz

Input acceleration [G] 0.3 0.31 0.31

Responsefactor

Tank (A4) 1.4 1.4 1.0

Middle of Internal parts (A3) 2.3 2.9 1.0

Top of Internal parts (A2) 2.4 3.0 1.1

Seismic test with

the EL Centroearthquake wave

Wave shape Actual earthquake wave

Input acceleration [G] 0.35 0.35 ---

Response

factor

Tank (A4) 1.1 1.0 ---

Middle of Internal parts (A3) 1.3 1.0 ---

Top of Internal parts (A2) 1.3 1.1 ---

Table 1.9: Seismic test results of 500 kV GIS arrester in accordance with JEAG-5003 [Shi 1997]




Page 45

a: Example of wave shapes

Figure 1 .18: Seismic test of 500 kV GIS arrester in accordance with JEAG-5003 [Shi1997] . The axis X,Y

and Z give the direct ions of accelerat ion, see Table 1.9.

b: 500 kV GIS shaking table,seismic test for horizontalinstallation

c: Internal view of 500 kVGIS arrester

Insulating rod (S1)

Tank (A4)

Middle of internal parts (A3)

Top of internal parts (A2)

Input acceleration (A1)

Time

A c c e l e r a t i o n

Input acceleration (A1) Tank (A4)

Shaking table

Y

Z

X

Top ofinternal parts (A2)

Middle of

internal parts (A3)

Insulating rod (S1)ZnO element

Shield

1 . 5

m

X

Z

Y




Page 46

1.6.2 POLLUTION Authors in charge: Bernhard Richter and Yoshihiro Ishizaki

One of the unsolved problems, and therefore discussed in Cigré working groups again, is the pollution performance

of polymer housed surge arresters for high voltage applications. Due to the length of the multi-unit designs testprocedures and criteria could not be agreed upon in the past. This was discussed as well in WG A3.21 and WG

A3.22 when dealing with 1000 kV UHV aspects. The principle problems are addressed in the following.

Figure 1.19 illustrates three possible mechanisms that may affect a multi-unit MO HV arrester in a polluted

environment, see also [Ric 2007].

A special problem for HV arresters of Type A (see 2.3 “Design of surge arresters”) with considerable gas in the

inside of the insulator may be the radial field strength as shown in Figure 1.20. Radial field stress, however, is also

a concern for Type B arresters. Here the radial voltage drops across small distance of only few millimeters between

the outer surface and the MO column, and a weak design may lead to puncture of the insulation material. As the

possible radial field stress increases with the distance between two metal fittings, the maximum unit length is

limited.

Figure 1.19 : Possible risks due to pollution

3

2

1

Risk of "internal"partial discharges,degradation of theMO resistors anddeterioration of the

supporting structure

Risk of partialheating of the activeparts (see Annex Fof IEC 60099-4)

Risk of externalflashover (seeIEC 60507)

33

22

11

Risk of "internal"partial discharges,degradation of theMO resistors anddeterioration of the

supporting structure

Risk of partialheating of the activeparts (see Annex Fof IEC 60099-4)

Risk of externalflashover (seeIEC 60507)




Page 47

Figure 1.20 : Possible volt age distr ibut ion of an arrester unit under pollut ed condit ions

Table 1.10 gives an overview about proposed methods for pollution tests of polymer housed MO surgearresters. The main discussed and open point is the way of treating the polymer surface of the insulatinghousing to get realistic results in the test, which can be compared to long term stresses in the system.

Item Solid layer Sprayed pollutant Salt fog

Standard to be

referred IEC 60507JEC 0201

IEEE C62.11IEC 60507

Artificial

pollution

procedure

After the polluted insulator

dried, the pollutant wetted by

clean fog

Pollutant slurry

sprayed on insulator

Salt fog generated

Soluble

componentNaCl NaCl NaCl NaCl

Non-solible

component

Kieselguhr

SiO2

Kaolin

or Tonoko

Tonoko

or Bentonite---

Note

Recovery of hydrophobicity

during testing considered

No special testing

facilities required

Recovery of

hydrophobicity

considered

Table 1.10: Applicable pollution methods for polymer housed arresters

A method to remove the hydrophobicity of a polymeric housing for test purposes, as practiced in Japan, isdescribed for information in the following, see Figure 1.21.

MO column

Conductive

layer

Gas

Solid

U a x i a l , i n t

Uradial

MO column

Conductive

layer

Gas

Solid

U a x i a l , i n t

Uradial




Page 48

Remark 1: Tonoko

Tonoko is fine inorganic powder from particular kinds of natural rock, which has been originally used as fineabrasive or filler for some traditional craft works in Japan. Tonoko is also introduced in IEC 60507 as inert material,

of which characteristics are given in Table 1.11.

Weight composition (%)Granulometry

(cumulative distribution) mmVolume

conductivitys20 (S/m)SiO2 Al2O3 Fe2O3 H2O 16% 50% 84%

60 - 70 10 – 20 4 - 8 - 0.8 - 1.5 3 – 5 8 - 15 0.002-0.01

Table 1.11: Characteristics of Tonoko [IEC 1991]

Figure 1 .21 : Procedure to remove hydrophobicit y and to apply contaminant for solid layer method

[Nai1996][Ish2008]

Remark 2: representative methods to remove hydrophobicity

The following procedure, as shown in Figure 1.21, is suggested to remove hydrophobicity on silicone surface ofMOR temporarily for the testing, without any damage on the surface or any additional chemical agent in thepollutant.

a) Prepare Tonoko slurry, which contains approx. 1 kg of Tonoko in 1 liter of water.

b) Spray the Tonoko slurry as uniformly as possible on the hydrophobic MOR surface.

c) Dry the polluted surface under natural ambient condition.




Page 49

d) Wash off the deposited Tonoko roughly, by running tap water, for example. After this process some amount of

Tonoko will remain on the surface, which suppresses recovery of the hydrophobicity temporarily. It is important

to conduct testing during hydrophobicity is completely lost.

Voltage application method

Figure 1.22 shows the test with only the MCOV to apply to surge arrester under the polluted condition according

with IEC 60099-4, Annex F.

Figure 1.23 shows the voltage application method to superimpose a temporary overvoltage (the sound phase

voltage at the time of single phase ground fault) on the continuous operating voltage (Uc) according to JEC

standard. This test procedure is considered as one of solutions for the risk concerned the external flashover at TOV

of surge arresters.

Figure 1 .22: Continuous volt age U c apply ing method according to IEC standard [ IEC 2006 ]

Figure 1 .23 : Temporary overvo ltage (TOV) applying method according to JEC standard

[JEC 2003]

3 min.

10 min.

UC contaminant

application

first sequence 3 min 1 continuous operating voltage

E 2 temporary over voltage

2 1

× 4 sequences

1 min 1 min 1 min 1 min 1 min 1 min 1 min 1 min

contaminantapplication




Page 50

1.6.3 HUMIDITY Author in charge: Kari Lahti

1.6.3.1 Ambient humidity stressThe humidity stressing of arresters is caused by different forms of ambient humidity; precipitation and content of

moisture in air. Factors like length and recurrence of moist periods also affect the stressing, likewise the properties

of periods with lower air humidity content and precipitation. Especially in the case of polymer housed equipment

atmospheric humidity stress has to be handled as an entity by considering the total weather type over longer

periods of time. Water may permeate through polymeric materials, and dry periods have a significant effect on this

process.

Atmospheric humidity stress varies a lot depending on the location, season etc. Equatorial areas, especially those

with tropical rainforest climates, typically introduce hard stresses of this kind with frequent rain and high absolute air

humidity. In areas like, for example, Europe these stresses are much lower. An overview of the humidity related

climatological conditions can be made based on levels of relative humidity (RH), precipitation and number of

precipitation days per year. Because the humidity stress conducted on electrical equipment is proportional to the

water vapor pressure and not to the relative humidity of the air the air temperature also has to be kept in mind

when considering the stresses in different parts of the world.

A map of World’s climatic zones (by Okolowicz) is given in Figure 1.24 with the most humid areas indicated in

colors. As can be seen in the figure, equatorial areas in South America, Africa and Indonesia are examples of

regions with highest atmospheric humidity stress.

Figure 1.24 : The wor ld’ s climat ic zones. The most humid climates in equatorial and tropical zones areindicated by colors [Mar92 ]




Page 51

1.6.3.2 Number of days with thunderstorms An indication of the variation of the lightning caused stresses over the world can be seen in the isokeraunic map

(Figure 1.25). Even better indication would be a ground flash density map but no such map covering the whole

world is, unfortunately, available. An estimation of the ground flash density can, anyhow, be calculated from the

thunderstorm day data for example by empirical expression by Anderson.

25.104.0 Dg T N (equation 1.3)

Ng = Ground flash density (km-2

yr -1

)

TD= Number of days with thunderstorms

Figure 1 .25 : Annual number of days w ith thunderstorms [Mar 1992]




Page 52

1.6.3.3 Polymer housed surge arresters under humidity stressesTightness of any surge arrester type is an important factor which is for polymer housed arresters tested by the

moisture ingress test (IEC 60099-4). The test is a combination of thermal - mechanical pre stressing and following

immersion test in boiling water during which the moisture are not allowed to penetrate inside an arrester and form

remarkable power losses, partial discharges or deviation in residual voltage.

The conditions during the above test are really abnormal but information of the behavior of polymer housed

arresters under more realistic ambient conditions can be found in the literature [Lah 1999] [Lah 2002] [Lah 2003]

[Kan 1997] [Kan 1998]. A summary of main results of such tests are given in the following. The corresponding tests

were performed for new, commercial medium voltage arresters during 1996 – 2001.

1.6.3.4 Moisture penetration into arrestersFigure 1.26 presents the general behavior of internal leakage currents of some arresters during a test in very humid

air (“humidity chamber test”). In this test the arresters were subjected to very humid ambient conditions (+30°C –

+35°C, RH 95 – 100 %, artificial rain periods). Moisture ingress into the arrester interior can thus be evaluated by

this figure. In Table 1.12 the tested MV arresters are divided into three groups according to their internal structure

and manufacturing technique.

Group Arrester types Description

I A,B,C,H Housing molded directly onto the

arrester body, no end caps

II D,E Housing manufactured separately

and pressed or slipped onto the

arrester body, end caps

III F,G As type II but with considerable

internal gas space

Table 1.12: Tested medium voltage arrester types according to internal structure

Figure 1.26: internal leakage currents of MO surge arresters according to their internal structure duringthe humidit y chamber test

0

10

20

30

40

50

60

70

80

90

100

0 100 200 300 400 500 600 700 800

Testing Time (days)

L e a k a g e C u r r e n t ( A ) Group I

Group II

Group III




Page 53

Based on the results, at least slight moisture penetration into arresters interior is possible in most of the arrester

types studied. However, remarkable levels of internal leakage current was reached by all the arresters of type III,

most of the arresters of type II but none of the arresters of type I. In this context one has to keep in mind that these

tests are performed for certain MV arrester types and, for example, in HV arresters also clearly different structures

are utilized.

In general, same kinds of results/behaviors have been achieved also in hot water immersion tests [Lah 1999] [Lah

2003]. Based on this result hot water immersion tests are applicable in sealing testing of polymer housed arresters.

In some cases some problems may be found in hot water immersion testing of arresters with housings with good

surface properties and high diffusion coefficient (e.g. some silicones). Quite high internal power losses may be

measured immediately after immersion test, which result does not necessarily reflect real service situation where

also good surface properties (e.g. hydrophobicity) affects the total behavior.

1.6.4 COMBINED HUMIDITY AND AC STRESSESContinuous AC stress typically limits the internal moisture content and corresponding power losses of a polymer

housed arrester into a lower level than it would be without the AC stress. If an arrester has a void free structure with

internal interfaces bonded together small amounts of moisture may still collect into interfacial areas under very

humid ambient. Heating caused by the consequent power losses chance the balance of water vapor partialpressures and part of the moisture diffuses out from the interfacial areas. The situation changes if the interfaces

loosen or especially when there are voids inside an arrester where water can permeate in and collect to.

Some results of a test case [Lah 2002] where continuous AC voltage stresses were applied to MV arresters having

internal moisture content are given in the following. The tested arresters were preconditioned in a hot water bath

before the actual test to obtain sets of arresters with some humidity inside. 12 kV AC stress was applied on the

arresters but also two and four week periods without voltage stress were applied. At test week 78 the ambient

temperature was increased from +30C to +50C. The internal power losses of these pre-defected specimens were

recorded and analyzed over the test period. Indication of the structures of the different arrester types are given in

Table 1.12.

1.6.4.1 General results

Only one of the arresters (E4, Figure 1.28) out of 26 failed totally during the first 102 weeks of the test. Thebehavior of internal power losses of some of the MOAs is shown in Figures 1.27 to 1.29.

In the beginning of the test the internal power losses of all the specimens decreased until an equilibrium state in

moisture diffusion was reached. Lowest loss levels were reached by the type I (direct molding) silicone (high

diffusion coefficient) housed arresters (Figure 1.29). During test periods without voltage stress internal humidity

content may increase and evidence of such can be seen as quite high peaks in losses immediately after

reapplication of test voltage. A clearly increasing trend of internal power losses can be seen for arrester F1 (with

internal gas space) from the beginning of the test indicating moisture accumulation inside the arrester.

In general, all rapid changes (increases) in stresses on arresters are the most demanding situations under very

high ambient humidity (e.g. rain forest conditions) since the internal power losses may reach relatively high levels

before the moisture equilibrium is reached.




Page 54

Figure 1 .27: Internal pow er losses of type D (internal structure type II) at 12 kV during the test

Figure 1.28 : Internal power losses of t ype E (internal structure type II) and F (internal structure type III)at 12 kV during the test

Figure 1 .29 : Internal power losses of t ype A and H (both of internal structure type I) at 12 kV duringtest

0

1

2

3

4

5

67

8

9

10

0 10 20 30 40 50 60 70 80 90 100

Test duration (weeks)

P o w e r l o s s e s ( W )

D1

D2

D3D4

D5

D6

0

1

2

3

4

5

6

7

8

9

10

0 10 20 30 40 50 60 70 80 90 100Test duration (weeks)

P o w e r l o s s e s ( W )

E1

E2

E3

E4

E5

E6

E7

F1

0

1

2

3

4

5

0 10 20 30 40 50 60 70 80 90 100

Test duration (weeks)

P o w e r l o s

s e s ( W )

H1

H2

H3

H4

A1A2




Page 55

1.6.5 EXPOSURES TO LOW AMBIENT TEMPERATURESExposures of arresters to very low (even -60°C) but rather short (2...6 days) temperature stresses have been

studied in laboratory conditions [Kan 1997]. It was shown that these stresses did not in general cause any

remarkable changes in the AC or DC leakage current behavior or in the residual voltages of the tested arrester

types. Field experiences gathered over the cold regions of the world support this result.

1.6.5.1 Effects of ice coveringsStudies of the effect of ice coverings on the electrical behavior of surge arresters have not been reported widely.

Laboratory investigations of the behavior of one two unit HV arrester type (Ur = 120kV) have, anyhow, been

reported in [Kan 1998] where the behavior of this arrester type has been studied under AC- and switching impulse

voltage stresses.

In addition to the outages in transmission networks due to flashovers of ice-covered equipment, icing of a metal

oxide surge arrester consisting of two or more units in series may have harmful effects on the performance of the

MOA. With AC voltage stress an unevenly ice-coated MOA may be thermally stressed due to the leakage current

transition from the ice covering of one unit to the interior of another unit not covered with so much ice. Whether the

stressed unit stands up to this situation or not depends on the thermal properties of the unit, the duration of the

active arcing period and the possible external flashovers on the MOA, which may clean the surface of ice and thusdecrease the leakage current stress after the flashover. The porcelain arrester investigated in this research stood

up to the different types of freezing rain conditions without thermal instability.

With switching impulse current surges the residual voltage across an ice-covered unit can cause an external

flashover. Such a high current surge stresses the varistors of the lower unit operating normally.




Page 56

1.6.6 BIOLOGICAL GROWTH Author in charge: Trond Ohnstad

Polymer insulators have become a common choice for surge arresters in transmission and distribution systems,

due to their many advantages over traditional porcelain insulators.

Though there are many types of composite materials used in insulators, silicon rubber is the material used by most

manufacturers today.

In many parts of the world there have been reports of observation of biological growth on the surface of the

polymer arrester insulation. Most of the reports concerns growth of algae and fungi on insulators of silicon rubber

in hot and humid climates and in clean environment.

In Scandinavia biological growth has been observed in typical inland substations surrounded by forest, places

which can be very humid and warm during summertime.

In [Gut 2003] the biological or organic growth is described as micro-organisms that colonize polymeric materials

and attach to the surface by forming a bio film. A bio film consists of micro-organisms embedded in a highly

hydrated matrix of extra cellular substances. The water content is 80-95 % and the cells themselves make up aminor fraction of the bio film.

Three types of organic growth have been identified, Algae, Fungi and Lichen.

Algae can be seen as a simple plant, producing its own food from carbon dioxide and water by utilizing sunlight, i.e.

photosynthesis. They also need some mineral nutrients which are taken from the environment. They can be found

almost everywhere, in sea and fresh water, on rocks, in soil. They spread by wind, water and animal movements

and multiply under certain climate conditions, i.e. under favorable temperature, humidity and sun radiation

Fungi are multi-cellular organisms, and composed of long, thread-like filaments. The body of the organism is called

mycelium and they multiply through sending out spores. Fungi cannot produce their food by photosynthesis;

instead they absorb nutrients from the surrounding environments.

Lichen is a special plant group consisting of fungi and algae which live in symbiosis with each other and adhere torocks, stones, trees, etc. The algae are dispersed in the fungi mycelium. The fungi take nutrients from algae while

algae take nothing but water from fungi.

Despite all the reports of biological growth on insulators there are no known reports of any failures of surge

arresters caused by it.

The biological growth will certainly reduce the hydrophobicity, but this might not be a problem in a clean

environment. The problem is more of a visual character, the insulators change color to green or black and the

insulator do not look very reliable.

Observations indicate a dependency of the polymer formulation to the degree of biological growth on the insulators.

This should be closer investigated.

If more severe problems should occur in the future, there will be a need of effective mitigation techniques.




Page 57

Figure 1 .30: Example of biological grow th.




Page 58

1.7 Short circuit currents


An arrester overload is a very rare event. However, it cannot in principle be ruled out, not even in the case of anover dimensioned arrester. Possible causes are, for instance, direct lightning strikes occurring near the arrester, orpower frequency voltage transfer from a higher to a lower voltage system, for example on a transmission line withseveral voltage levels crossing each other and suffering a conductor failure or galloping. In transmission systemsthis occurs considerably less frequently than in distribution systems. In some special applications an MO arrester isintentionally overloaded and falls into short circuit state, in this way protecting sensitive and expensive equipment.This application is then called “sacrificial arrester”.

In case of an overload of an MO arrester some or all MO resistors in the arrester will flash over and an arc willoccur internally in the housing between the two flanges. The full short circuit current of the net, which appearswhere the arrester is actually installed, flows through this arc. As a result, an abrupt increase of pressure developswithin the housing and stresses mechanically the design. In addition the hot arc gets in contact with the housing,which may lead in the case of porcelain housings to a thermal cracking of the housing. The described scenario isvalid for all arrester designs with considerable enclosed gas volume. To avoid dramatical failures (violentshattering) pressure relief devices, integrated in the flanges, are needed, typically on both sides of each arresterunit. The pressure relief devices have to open within a few milliseconds so that the arc can commutate to theoutside of the arrester preventing shattering of the housing.Based on the design of the housing with pressure relief devices the “pressure relief behaviour” had to be testedwith specific “pressure relief tests”. Figure 1.31 shows the principle steps of a pressure relief of a porcelain orhollow core housed arrester unit.

Figure 1.31: Pressure rel ief of a porcelain housed arrester unit.

left : puncture and fl ashover of indiv idual MO resistorsmiddle: internal arc along the full length of the unitright: opening of pressure relief devices and venting of the unit

As the new polymer insulated arresters with direct molding no longer contain enclosed gas volumes in theirhousing, it makes sense to refer more generally to “short circuit behavior”, and accordingly to “short circuit tests”.No defined pressure builds up in this type of arrester housings and no special pressure relief device is needed.Instead the arc propagating along the MO resistors in the arrester seeks a path through the housing wall to anarbitrary point or points, which have been specially provided in the design.Independent of the design of the arrester, the goal always remains the same: in the case of an arrester overloadingthe housing must either remain intact, or if it breaks, the housing fragments and ejected parts must fall down in anarrow and defined area.




Page 59

Figure 1 .32 : Short circuit behavior of a cage design polymer housed arrester unit (courtesy ABB).1: the arrester has fai led and gas begins to be expelled through the housing

2: the gas streams tr igger an external flashover and the internal arc is commutated to the outside

The maximum short circuit current, defined for a flowing time of 200 ms, is the rated short circuit current Is, given assymmetric current in r.m.s. value with power frequency. The manufacturer has to state the rated short circuitcurrent the arrester is intended to withstand. Table 1.13 gives the rated short circuit currents for short circuit testsas they are defined in IEC 60099-4.

Rated short circuit current200 ms / A

Reduced short circuit currents200 ms / A

Low short circuit current1 s / A

80 000 50 000 25 000

600 ± 200

63 000 25 000 12 000

50 000 25 000 12 000

40 000 25 000 12 000

31 500 12 000 6 000

20 000 12 000 6 000

16 000 6 000 3 000

10 000 6 000 3 000

5 000 3 000 1 500

Table 1.13: Standard short circuit currents for tests purposes of MO arrestersOnce the rated short circuit current is claimed for an MO arrester the type tests have to be performed with all theshort circuit currents given in Table 1.13. The required tests at reduced short circuit currents shall help avoidingthat the short circuit performance is optimized for high short circuit currents only while at lower short circuit currentsthe housing might violently shatter before the pressure relief devices can open.




Page 60

To avoid misunderstandings, it has to be pointed out that the short circuit current with power frequency, driven bythe system voltage, is flowing through the surge arrester, or more likely through the arc, only after the arrester wasoverloaded. Therefore, the MO arrester was destroyed and failed in a controlled way.

Figure 1.33 shows as example porcelain housed MO arrester units after successful short circuit tests. Both theresults are considered to be positive according the pass criteria as defined in the test standard IEC 60099-4.Thermal cracking, as to be seen in Figure 1.33 (right) is acceptable, as long as all parts heavier than 60 g remain ina well-defined enclosure. The MO arrester is of course destroyed in both cases, but it failed in a well-definedmanner.

Figure 1.33 : Examples for successful short ci rcuit tests of porcelain housed arrester units.left : intact porcelain body, r ight: secondary break of porcelain body(courtesy Siemens).

Rated short circuit current t ested I s = 63 kA for 200 ms.

Figure 1.34 shows examples of short circuit tests performed on polymer housed MO arrester units. Both results areto be considered positive. In both cases the arrester units are destroyed, but without cracking or ejecting hard partslike MO resistors or parts of them. In case of the cage design, Figure 1.34 right, only soft polymeric material isripped off.




Page 61

Figure 1.34 : Example of successful short circuit tests on polymer housed MO arrester units.lef t: FRP hollow insulat or (cour tesy Siemens),

right: cage design (courtesy ABB).In both cases the short circui t cur rent w as I s = 63 kA for 200 ms.

It has to be understood clearly that an overload is a normal operation for a surge arrester. Therefore, surgearresters have to be designed for an overload. Important is that the arrester fails in a controlled and safe way. Thisis an important safety issue when arresters are installed close to public places, sensitive infrastructure or forinstance on the roof of a traction vehicle in railway applications.




Page 62

2. Functional parameters and design of MO Surge Arresters

2.1 Function and relevant parameters


2.1.1 INTRODUCTION A metal-oxide surge arrester without gaps (MO arrester) is an arrester having nonlinear metal-oxide resistors

connected in series and/or in parallel without any integrated series or parallel spark gaps. The wording surge

arrester is used in the HV and MV community and describes different designs of MO arresters. In the LV field it is

common to speak in general about Surge Protective Devices (SPDs), which covers different technologies and

design types, e.g. spark gaps, metal oxide varistors and combinations of them including disconnecting devices etc.

The function of a surge arrester with an active part consisting of a series connection of MO resistors is very simple.

In the event of a voltage increase at the arrester’s terminals, the current rises according to the characteristic curve,

see Figure 2.1, continually and without delay, which means that there is no actual spark over, but that the arrester

skips over to the conducting condition. After the overvoltage subsides the current becomes smaller according to the

characteristic curve. A subsequent current, such as those that arise with spark-gaps and spark-gapped arresters,does not exist; it flows only the so-called almost pure capacitive leakage current ic of about 1mA.

Figure 2.1 : Volt age-current characteristic of a MO surge arrester

a - Lower part (capacitive), b - knee point, c - strongly non-linear part, d - upper part (“turn up” area),

A - Operating point (continuous operating voltage Uc),

B - Protective level Up (nominal discharge current In).




Page 63

In Figure 2.2 a more technical diagram is given, indicating the standardised definitions.

Figure 2.2 : Voltage-current character istic of a MO arrester with I n = 10 kA, l ine discharge class 2. Thevolt age is normalized to the residual voltage of the arrester at I n . The values are given as peak values

for the volt age (l inear scale) and the current (logarit hmic scale). Shown are t ypical values.

In Figure 2.3 typical voltage and current wave forms are given for MO surge arrester as defined in IEC 60099-4 and

ANSI/IEEE C62.11. In the lower part at Uc the arrester acts as a capacitor, the current is in the range of 1 mA and

below. At the knee point b (Figure 2.1) at Ur and Uref the arrester starts to conduct, the ohmic content of the current

is increasing rapidly with a slight voltage increase. At U ref the current has a dominantly ohmic component.

Temporary overvoltages have to be considered in the voltage region between point’s b and c.

In the low current region up to point b power frequency currents and voltages have to be considered (range of

continuous operation). In the region above b the protective characteristic of the MO arrester is of importance.

That’s why in this range the voltage-current-characteristic is defined by impulse currents of different wave shapes

and current magnitudes (Figure 2.2).

For simulations of the performance of MO surge arresters normally a voltage-current-characteristic is used that

starts at some 10 A in the region c and goes up to maximal 40 kA in the region d. Normally an impulse current

wave shape of 8/20 s is considered.




Page 64

Uc Ur Uref

Isw In Ist Ihc

Figure 2 .3: Typical voltage and current w aveforms of a MO surge arrester.

- Uc continuous operating voltage

- Ur rated voltage

- Uref reference voltage

- Isw switching current impulse, wave shape 30/60 s

- In nominal, lightning current impulse, wave shape 8/20 s

- Ist steep current impulse, wave shape 1/.. s

- Ihc high current impulse, wave shape 4/10 s

The following paragraph briefly explains typical current and voltage waveforms in different areas of the

characteristic curve.

2.1.2 CURRENTS AND VOLTAGES

Continuous op erating voltage Uc: Designated permissible r.m.s value of power-frequency voltage that may be

applied continuously between the arrester terminals.

Continuous current ic: Current flowing through the arrester when energized at the continuous operating voltage.

The MO arrester behaves in an almost purely capacitive manner in the region of the continuous operating voltage.

The current is around 1 mA and almost 90° electrically shifted compared to the voltage. The power losses in this

region can be neglected.

The continuous current is also known as leakage current.

Rated voltage Ur : Maximum permissible r.m.s. value of power-frequency voltage between the arrester terminals at

which it is designed to operate correctly under temporary overvoltage conditions as established in the operating

duty tests.

Briefly: the rated voltage Ur is the voltage value, which is applied for t = 10 s in the operating duty test in order to

simulate a temporary overvoltage in the system. The relationship between the rated voltage Ur and the continuous

operating voltage Uc is generally Ur /Uc = 1.25. This is understood as a given fact, but it is not defined anywhere.

Other ratios, such as Ur /Uc, can be chosen. The rated voltage has no other importance although it is often used

when choosing an arrester.




Page 65

Reference voltage Uref : Peak value of the power-frequency voltage divided by 2, which is applied to the arrester

to obtain the reference current.

Reference current iref : Peak value (the higher peak value of the two polarities if the current is asymmetrical) of the

resistive component of a power-frequency current used to determine the reference voltage of an arrester.The reference current is chosen by the manufacturer in such a way that it lies above the knee point of the voltage-

current characteristic and has a dominant ohmic component. Therefore, the influences of the stray capacitance of

the arrester at the measurement of the reference voltage are not to be taken into account. The reference voltages,

which are measured at single MO resistors, can be added to give the reference voltage of the entire arrester.

Reference voltage U1mA and reference current with DC voltage:

A reference current and the reference voltage for DC voltage belonging to it are often also demanded instead of a

given reference current for AC voltage. It is now common practice to specify the DC voltage, which is applied with a

direct current of 1 mA to the terminals, no matter what the diameters of the MO resistors are. Both types of

information, the reference current and the reference voltage for AC voltage and for DC voltage, are in principle

equal. Both of these types information describe a point on the voltage-current characteristic of an arrester, where

the influences of the stray capacitance can be ignored. All the tests performed according to IEC are always based

on the reference current and the reference voltage for AC voltage. Reference current and reference voltage withDC voltage are additional information, which can be received from the manufacturer.

Residual voltage Ures : Peak value of voltage that appears between the arrester terminals during the passage of a

discharge current.

The residual voltage of a MO resistor or MO arrester is determined with surges having different wave forms and

current heights and it is given in tables or as a voltage-current characteristic on a curve. The measurements are

generally performed on MO resistors. As the measurement is mostly performed in regions of the characteristic

where the ohmic part of the current is dominant, the capacitive stray influences can be ignored. The residual

voltages measured on single MO resistors can be summed up as the residual voltages of the whole arrester.

Lightning impulse protective level Upl: Maximum permissible peak voltage on the terminals of an arrester

subjected to the nominal discharge current. Corresponds to the guaranteed residual voltage Ures at In.

Switching impulse protective level Ups: Maximum permissible peak value on the terminals of an arrester

subjected to switching impulses.

Lightning current impulse: Current impulse with the wave shape 8/20 s. The virtual front time is 8 s and the

time to half-value on the tail is 20 s. The lightning current impulse reproduces approximately the current impulse

produced by a lightning stroke in a conductor after an insulator flashover. This current impulse travels as a

transient wave along the line.

Nominal di scharge current of an arrester In: The peak value of the lightning current impulse that is used to

classify an arrester. The nominal discharge current and the line discharge class of an arrester are correlated to the

system voltages and prescribe the test parameters, see Table 2.1. Recommendations for the choice of the nominal

discharge currents and the line discharge classes for different system voltages are to be found in IEC 2000 and

IEC 2009.

High current impulse Ihc: Peak value of discharge current having a 4/10 s impulse shape. The high current

impulse should reproduce a lightning stroke close to an arrester and it is used with medium voltage arresters of the

line discharge class 1 as a proof of thermal stability. It represents not only an energetic stress but also a dielectric

one, taking into consideration the high residual voltage that occurs with a high current impulse with a peak value of

100 kA. It is however, necessary to strongly emphasize that a high current impulse with an amplitude of 100 kA is

not the same as a real lightning current of the same amplitude. The real lightning current of this amplitude

measured during a thunderstorm lasts longer than several 100 s. Though such strong lightning currents and

impulse shapes are very rare and appear only under special conditions, such as during winter l ightning in hilly

coastal areas.




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Switching current impulse Isw: Peak value of discharge current with a virtual front time between 30 s and 100

s, and a virtual time to half-value on the tail of roughly twice the virtual front time. The switching current impulses

are used to determine the voltage-current characteristic, and in connection with the line discharge class are also

used to determine the energy which must be absorbed during the operation test. The current amplitudes lie

between 125 A and 2 kA, and roughly reproduce the load of an arrester produced by overvoltages, which were

caused by circuit breaker operation.

Steep current impulse: Current impulse with a virtual front time of 1s and a virtual time to half-value on the tail

not longer than 20 s. The steep current impulses are used to determine the voltage-current characteristic. They

have amplitudes up to 20 kA and roughly reproduce steep current impulses like those which may appear with

disconnector operation, re-striking, back flashes, and vacuum circuit breakers.

All the current impulses described above (except the high current impulse) are used to determine the voltage-

current characteristic of a MO arrester. It must be considered that only the virtual front time and the amplitude of

the current impulses are decisive for the residual voltage and not the virtual time to half-value on the tail. That is the

reason why the tolerance for the virtual front times is very tight, and contrastingly, the tolerances for the virtual

times to half-value on the tail are very broad.

Long-duration current impulse Ild: Also called rectangular wave (Irw) or square wave, a long-duration current

impulse is a rectangular impulse that rises rapidly to its peak value and remains constant for a specified period of

time before it falls rapidly to zero. The length of the current pulse duration is correlated to the line discharge class

of an arrester. Rectangular impulses are used in laboratories during the type tests with long-duration current

impulses, and during the operating duty test of MO arresters having line discharge classes 2 to 5, in order to inject

the energy in the arrester. The current amplitudes are up to 2 kA and reproduce the load of an arrester when a

charged transmission line discharges into the arrester in case of an overvoltage occurrence.

It is now regarded as a matter of course to use a rectangular wave of 2 ms duration to compare different MO

arresters, although there is no norm established for doing so. Specified is either the amplitude of the rectangular

wave for a specific MO arrester or the energy transferred into the arrester during the flow of the rectangular current.

Line discharge class: The line discharge class is the only possible way to specify the energy absorption capability

of an arrester provided in IEC 60099-4. The line discharge classes 1 to 5 are defined with growing demands. They

differ from one another due to the test parameters of the line discharge tests. The energy W is calculated from the

line discharge class in connection with the residual voltage of the switching current impulse. This calculated energy

must be injected with each discharge in a MO resistor during the test with a long-duration current impulse Ild (line

discharge test). Two corresponding line discharges are loaded in the arrester during the operating duty test as a

proof of thermal stability.

W = Ures (UL – Ures) 1/Z T (equation 2.1)

Ures = Residual voltage of the switching current impulse. Here, Ures is the lowest value of the residual voltage

measured at the test sample with the lower value of the switching current impulse.

UL = Charging voltage of the current impulse generator used in test labs for producing the long-duration current

impulse Ild.

Z = Surge impedance current impulse generator.

T = Duration of the long-duration current impulse

The parameter of the line discharge classes are derived from the stored energy of long transmission lines, see

Table 2.1.

That is the reason why the line discharge classes have no direct importance in medium voltage systems. They

serve here only to distinguish the energy handling capability of different arresters.




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In

kA

LD Us

kV

L

km

ZL

T

ms

10 1 245 300 450 2.0

10 2 300 300 400 2.0

10 3 420 360 350 2.4

20 4 525 420 325 2.8

20 5 765 480 300 3.2

L = the approximate length of the transmission line.

ZL = the approximate surge impedance of the transmission line

Table 2.1: Correlation between line discharge classes and parameters of transmission lines. The

duration T of the long-duration current impulse I ld is also given.

Rated short circuit current Is: The r.m.s. value of the highest symmetrical short circuit current, which can flow

after an overload of the arrester through the arc short circuiting the MO resistors without violent shattering of the

housing. The proof of the specified value specified by the manufacturer is conducted in the short circuit test.

2.1.3 COORDINATION OF INSULATION AND SELECTION OF ARRESTERS

The coordination of the insulation is the matching between the dielectrically withstand of the electrical equipment

taking into consideration the ambient conditions and the possible overvoltages in a system.

For economic reasons, it is not possible to insulate electrical equipment against all overvoltages that may occur.

That is why surge arresters are installed to limit the overvoltages up to a value that is not critical for the electricalequipment. Therefore, a MO arrester ensures that the maximum voltage that appears at the electrical equipment

always stays below the guaranteed withstand value of the insulation of an electrical device.

In the following the very basics of insulation co-ordination are given, see also Figure 2.4.

An arrester has to fulfill two fundamental tasks:

- It has to limit the occurring overvoltage to a value that is not critical for the electrical equipment and

- It has to guarantee a safe and reliable service in the system.

The continuous operating voltage Uc is to be chosen in such a way that the arrester can withstand all power

frequency voltages and also temporary overvoltages without being overloaded in any possible situation. This

means that T Uc must be always higher than the maximum possible temporary overvoltages UTOV in the system.

Comment: Ferromagnetic resonances are the exception. The ferromagnetic resonances can become so high and

exist so long that they may not be taken into consideration by the dimensioning of the continuous voltage if the

arrester should still be able to fulfill its protection function in a meaningful way. If ferromagnetic resonances appear,

then this generally means that the arrester is overloaded. The system user should take the necessary measures to

avoid ferromagnetic resonances.

The MO arrester can fulfill its function of protection properly if the lightning impulse protection level Upl lies clearly

below the lightning impulse withstanding voltage (LIWV) of the electrical equipment to be protected, the safety

factor Ks is also to be taken into consideration. The point is to set the voltage-current characteristic of the arrester in

a way that both requirements are met.




Page 68

It makes sense to choose the continuous operating voltage Uc a l ittle bit higher than was calculated (for instance

10%). As a rule, there is enough distance between the maximum admissible voltage at the electrical equipment

and the protection level of the arrester.

Figure 2.4: Comparison of the possible occurr ing volt ages in the system, the wi thstand voltage of theelectr ical equipment and the parameters of the MO arresters.




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2.2 MO-Varistors: state of the art and actual trends

Authors in charge: Felix Greuter, Roger Perkins and Manfred Holzer

The purpose of this chapter is to lay a foundation of understanding of the basic behavior of the metal-oxide resistor

and its consequences for the surge arrester. In the former technical brochure TB 60 the knowledge gained in the

70-ies was summarized. Since then major progress has been made in the technology of metal oxide (MO)

varistors, their application in overvoltage surge protection devices and in the understanding of the basic

mechanisms. A lot of new insights have been gained, new physical phenomena have been observed, improved

and more consistent models have been developed and much progress has been made in simulations related to

materials and components. These topics are briefly addressed in this chapter.

2.2.1 ELECTRICAL PROPERTIES OF THE METAL-OXIDE RESISTOR

Some fundamentals are necessary for an understanding of the electrical behavior of the metal-oxide resistor.

Figure 2.5a shows the DC voltage characteristic of a ZnO varistor. The sharp transition from the insulating to the

conducting state, which takes place at the breakdown voltage UB, is the outstanding feature of this strongly non-

linear and voltage-dependent resistor. The switching is both extremely fast (in the range of pico- to nano-seconds)

and also fully reversible, i.e. the resistor reverts to blocking the current flow as soon as the applied voltage U falls

below UB.

A Pre-breakdown region B Breakdown region C Upturn region

1 DC voltage characteristic 2 AC voltage characteristic 3 Residual voltage characteristic

E Field strength J Current density UG Continuous operating voltage (DC)

UB Breakdown (or switching) voltage

Uv Continuous operating voltage (AC, 50 Hz)

Up Residual voltage, 8/20 s

Resistivity Non-linearity exponent (U)

Figure 2.5a: Linear presentat ion of thecharacteristic of a metal- oxide resistor

for t he high voltage sector

Figure 2.5b: Log-log plot of thenormalized J(E)-characteristi cs of atypical ZnO varistor [ Gre 1989 ]




Page 70

By careful dimensioning of the geometry and controlled manufacture of the metal-oxide ceramic, the breakdown

voltage per element can be set to values within a very wide range (UB 3 V to >104 V). This allows realizing

protection devices for electronic circuits up to ultra-high voltage systems.

The switching mechanism of the material can be traced back to individual grain boundaries in the ceramics, whicheach show a typical value for UB of app.3.2-3.4 V. For a general review see [Lev 1989], [Gre 1990], [Cla 1999],

[Bue 2008]. Combining many grain boundaries in series and in parallel within a MO-element then allows scaling the

voltage and current characteristic of a MO-ceramic block. For a sufficient number of grain boundaries the field

strength E and current density J then describe the material characteristic more generally. In the log-log

representation of the characteristic (Figure 2.5b) there are three distinct regions, i.e. the pre-breakdown region A,

the breakdown regime B, and the upturn region C.

During normal system operation, in which no overvoltages occur, the voltage applied to the arrester is the

continuous operating value (UV or COV for AC or UG for DC), which lies in the upper part of the pre-breakdown

region. The breakdown region is characterized by a very high non-linearity in the current-voltage curve. As Figure

2.5b shows, it is described quantitatively by the non-linearity exponent (E), which is a function of the voltage U or

applied field E, respectively. Its maximum values are typically around 20-70, but values above 100 have been

observed. The rated voltage UR, which also dictates the arrester's range of application, lies in the region of UB. Inthe extreme case, where current densities are very high or very low, approaches unity (ohmic behavior), the two

regions being as much as 12 current decades apart [Gre 1989] [Per 2003]! In the case of a high, transient

overvoltage, the varistor state lies in the upper part of the breakdown or in the upturn region. This protective range

is characterized by the residual voltage UP, which depends upon the wave shape and amplitude of the impulse

current. For an arrester to effectively suppress voltage transients, the difference between the residual voltage and

the continuous operating or rated voltage must be small. However, it is important to recognize that the highly

temperature and voltage dependent power loss Pv(V,T) generated at UV limits the maximum continuous operating

voltage (MCOV) which is possible. The power loss Pv must be low enough to satisfy the conditions for a thermally

stable state under possibly simultaneous conditions of elevated operating voltage, elevated temperature, aging,

pollution and energy absorption. A broad range of thermal stability is ensured by low power losses and the efficient

dissipation of heat in the arrester. This is also dependent on its detailed internal and external design and the

materials used to make them.

The AC-characteristics shown in Figure 2.5b is the result of the superimposed capacitive and resistive currents

flowing through the grain boundaries under a time dependent voltage stress. It is common practice to plot the

current peak vs. the voltage peak under AC. Well below UV the current is predominantly capacitive, while it is

strongly (non-linear) resistive for peak values above UB . To a first approximation, the AC response can be

described by using the DC-curve for the resistive part and the (small signal) dielectric permittivity of the material,

which is rather high (few hundreds). In some cases, for the so called “AC resistive” currents also the current value

I(t) in phase with the voltage maximum is used, which provides an improved (but not full) approximation to the

power losses. As indicated in Figure 2.5b, in the breakdown region the peak AC voltage is a few percent above the

steady state DC voltage due to the dynamic effects of charge trapping at the grain boundaries [Gre 1990].

It is worth mentioning that the phenomenon of electrical ageing, leading to thermal instability and failure, was a

serious one early in the development of the ZnO varistor. Today it has largely been brought under control by most

qualified manufacturers. However ageing is not only a result of long term exposure to operating voltages, it canalso be generated by current impulses (impulse degradation). Equally, not only the power losses at operating

voltage and/or reference voltage are affected, in some severe cases the discharge voltage can also be in- or

decreased [Lev 1989], [Per 2003].

2.2.2 MICROSTRUCTURE OF METAL-OXIDE RESISTORSThe metal-oxide resistor is made of a ceramic based on ZnO, a wide band gap semiconducting material. Its special

electrical properties are the direct result of its microstructure. Viewed through a microscope, the structure is seen to

be made up of minute ZnO crystals or grains, approximately 10-20m in size. The core of the grains is a good

electrical conductor ( 1 cm). At the grain boundaries, however, electrostatic potential barriers are built-up,

which form a highly insulating (electrostatically repulsing) region not more than 100 nm thick on each side of the ca.

1nm thick interface. It are these tiny grain boundary potential barriers, which control the current flow in the pre-




Page 71

breakdown, breakdown and lower part of the upturn region. By adding a few percent of selected doping elements

such as Bi, Sb, Co, Mn etc. to ZnO and using a suitable sintering process, it is possible to influence both the

conductivity of the ZnO grains and the properties of the high-resistance grain boundaries. The microstructure of the

ceramic is dominated by the densely packed ZnO grains resembling irregular polyhedrons. It are the common

interfaces (or grain boundaries) of these ZnO-polyhedrons, which are the electrical active part of the material.

Photographs of fracture surfaces provide a clear picture of this structure (Figure 2.6).

Figure 2.6 : Typical microstr ucture of a ZnO-varistor aft er fract uring preferentially along the grainboundaries (dark: doped ZnO grai ns, white: Bi2O3-phase at t riple points, grey: Spinel secondary phase

Most of the admixed bismuth oxide collects as a separate phase at the triple points at which the adjacent grainedges make contact. Also found at these points is a spinel phase in the form of fine grains, which are most easily

distinguished from the ZnO grains by their smaller size and more regular, octahedral shape. The grain boundaries

themselves, however, are not quite free of bismuth, although this cannot be seen in Figure 2.6. Using highly

sensitive techniques for analysis, it is however possible to detect at these boundaries the presence of fractions of

an atomic monolayer of both bismuth and oxygen atoms, which are essential for the electrical function of the

varistor [Cla 1999], [Chi 1998], [Kob 1998], [Sat 2007], [Stu 1990], [Elf 2002], [Che 1996]. During the sintering

process the bismuth oxide melts to form a liquid phase, which dissolves, at least in part, the other doping

substances and promotes their uniform distribution. The liquid phase also favors grain growth and dense sintering.

The spinel precipitates, on the other hand, inhibit grain growth and generate a uniform distribution of the ZnO grain

size. More recently the frequently observed inversion (or twin) boundaries were also recognized to be crucial for

the grain growth mechanism [Ber 2008].




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2.2.3 THE MANUFACTURING PROCESSThe basic steps of the manufacturing process are shown in Figure 2.7. For a reliable high-performance metal-oxide

varistor to be produced, each of these steps must be well understood and optimally carried out [Lev 1989], [Gre

1989], [Per 2003].

1 2 3 4 5 6 7Figure 2.7 : Manufacturing process for met al-ox ide resistors

1. Production of homogeneous slurry by wet-mixing of oxide powders2. Drying and granulation in a spray dryer3. Compacting the granulate to form resistor blocks4. Sintering to obtain dense ceramic bodies5. Addition of electrical contacts and a protective coating6. Electrical testing7. MO resistors ready for assembly

The basic material used to manufacture metal-oxide resistors is very fine grain ZnO with particle sizes of about1 m, to which as many as ten or more doping elements are added in the form of fine oxide powders. Its actual

composition differs from manufacturer to manufacturer. The proportion by weight of all additives together is typically

10 percent, with the share of the individual components ranging from ppm to percent. The purity and fineness of

the metal-oxide powders and the homogeneity of the mixture are, therefore, of immense importance for the quality

of the end-product. To achieve the required homogeneity the powder is usually treated in several different

processing steps. Sometimes the metal-oxides are wholly or partially heat-treated or calcined with none, part or all

of the ZnO powder to complete some of the solid-state chemical reactions before the sintering process is carried

out. Almost always a grinding operation is necessary to make the overall grain-size distribution smaller and

narrower and thoroughly mix the powders. Additional mixing operations are also used, in particular to mix the

smaller quantity of metal-oxide material with the larger quantity of ZnO. High shear mixing is frequently used to

achieve high homogeneity and various organic processing aids (dispersants, binders etc.) are added.

After these powder processes have been completed, the mixture or slurry has to be spray-dried to remove thewater and obtain a dry granulate that is beneficial for processing. Sometimes the spray-drying operation is also

carried out for the calcining operation mentioned above. The spheroidal granules obtained by spray-drying have

about 100 m in diameter, flow very easily and can be easily compacted under pressure. This takes place in the

next production stage, during which the granulate is compressed into disc-shaped blocks using a dry, uniaxial,

hydraulic press. The “green” blocks have approximately 50 to 60 percent of the theoretical density at this stage. It is

important here to ensure a uniform high density throughout the block, and that there are no defects present.

The blocks are finally sintered at about 1100-1300ºC, which has the effect of fully densifying the compacted powder

into a solid ceramic body with virtually no remaining porosity. In a prior step that occurs at lower temperature, the

organic additives mentioned above are pyrolyzed. This is a critical step, which requires very different conditions of

heat treatment and a continuous flow of fresh air. It can easily introduce flaws and voids into the body. During the




Page 73

sintering process the submicron-sized powder particles are united by means of diffusion and grow into large single

crystalline grains, where at the same time the additive dopants are built into the crystal lattices and the grain

boundaries are formed.

Ready-to-use MO resistors are obtained by adding high-conductivity metal contacts to the flat surfaces andapplying a coating to the resistor's peripheral surface to protect it from the environment. Frequently and beneficially

this is a glass coating, but other organic or inorganic materials have also been used. The coating is often referred

to as a “passivation”, in analogy with those used in solid-state semiconductors. Therefore, whatever material is

used, it is important that the coating not only has high dielectric and thermal withstand capability but also does not

change the properties of the varistor material underneath it.

2.2.4 ELECTRICAL TESTING OF METAL-OXIDE VARISTORSBefore leaving the production line each block has to pass a series of tests to verify its electrical properties, with

additional sampling of the resistors for ageing tests. These tests vary from one manufacturer to another, depending

upon their product quality and company standards. They typically include discharge voltage UP, AC or DC

reference voltage, power loss at continuous operating voltage, long duration and/or high amplitude current

impulses (LCLD, HCSD) and ageing. Some of them are performed on every MO disk whilst others are by nature

almost destructive and therefore sample tests. Some of these values are typically marked on the disk along withproduction information such as the lot or batch number. These almost always include the discharge voltage. Often

the information is coded and not disclosed to third parties.

2.2.5 FROM GRAIN BOUNDARIES TO VARISTOR BLOCKSEarly in the history of ZnO-varistors it was already realized that the varistor action is a grain boundary phenomenon

and a variety of models have been developed [Lev 1989], [Gre 1990], [Cla 1999], [Bue 2008]. These models have

continuously been refined and put on a sounder physical basis. Also a more consistent understanding on the

microstructure down to the atomic level at the grain boundaries has evolved in the 80-ies and 90-ies [Lev 1989],

[Gre 1990], [Chi 1998],[Kob 1998], [Sat 2007], [Stu 1990], [Elf 2002], [Che 1996]. Figure 2.8 is a schematic picture

of the (electrical) microstructure of the varistor ceramics. At the grain boundaries extra electrons are trapped in

interfacial defect states, which lead to electrostatic potential barriers (Double Schottky Barriers, DSB) as illustrated

in the band diagram in Figure 2.9.

Figure 2 .8: Schematic view of t he“electr ical” microstructure of a MO-

varistor [Gre 1989]

Figure 2.9 : Band diagram of a single grainboundary, showing the Double Schottky

Barriers DSB formed by charge trapping ininterface states [Gre 1989 ] (dash-

dotted/ dot ted lines: Fermi or quasi Fermilevel [Bla 1986] )




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At the grain boundaries thin intergranular films have consistently been observed, which seems to be an equilibrium

feature decorating all ZnO interfaces in the microstructure [Cla 1999], [Chi 1998], [Kob 1998], [Sat 2007], [Stu

1990], [Elf 2002], [Che 1996]. These films have a thickness of 1 nm, consist of an amorphous Bi2O3-ZnO solid

solution of reduced density and their Bi-concentration corresponds to an equivalent of 0.5-1 atomic layer. In

addition an excess of oxygen ( 0.5-1 monolayer) has been detected at these electrically active interfaces and has

been shown to be essential for the electrical characteristics [Stu 1990]. The negative charge trapped in this

intergranular film (or at its interface to the ZnO grains) is compensated by the ionized, positive donors in the

adjacent space charge (or depletion) regions to achieve charge neutrality. I t are these tiny electrostatic potential

barriers B , which control the current flow through the material. Note that these barriers are very thin ( 200nm,

compared to the grain size of 10-20 m) and all the voltage drops occur on the positive biased side of the high

resistive grain boundaries, except in the upturn region, where the finite conductivity of the highly doped ZnO-grains

0.1-1 cm) [Lev 1989], [Cab 2004] starts to add to the voltage drop.

For a good varistor characteristics the potential barrier B(U) should stay high with increasing voltage U to prevent

conduction electrons to be thermally activated over the barrier and to generate a leakage current I(U) or power

loss, respectively (note that: I~exp(B(U) + EF)/kT, for eU >>kT, EF : Fermi energy). If the interface states Ni(E)

indicated in Figure 2.9 have a high density and a suitable energy distribution within the band gap of ZnO, then the

barrier height B(U) hardly changes with voltage and the leakage currents stay low. This is referred to as pinning ofthe potential barrier B(U) and eventually even -values <1 can thereby result in the pre-breakdown. If all interface

states Ni(E) however would be filled, the total interface charge Qi ( Qi =Ni(E)dE ) would be constant and B(U)

would decay rapidly for U>0, leading to a poor characteristics. If however filled and unfilled interface states are

available, this decay of B(U) is counteracted by extra electrons being trapped in the unfilled states, Qi(U) will

increase and stabilize B(U) at a high level (note that: B(U) ~ Qi2/N0 , N0: free carrier density in ZnO). Such a

pinning mechanism for interfacial barriers is also well known for other semiconductors [Gre 1990]. From various

studies the picture evolves that these pinning states are characteristic defect levels of the ZnO crystal lattice

(and/or the intergranular film) and hence are always positioned at the same energy within the energy gap of ZnO

[Gre 1990], [Gre 1986], [Gre 1995]. Due to its chemical nature, ZnO never is fully stochiometric and always

contains a small excess (few ppm) of Zn-atoms (e.g. either via Zn-interstitials or O-vacancies in the crystal lattice).

By varying dopants and the sintering and/or heat treatment in the manufacturing process the defect equilibrium of

ZnO is changed and hence the concentration (but not the energy position) of these pinning states is changed. This

affects the voltage response of the barrier B(U) and thereby different J(E) characteristics will result. A high density

of unfilled states (at U=0) is favorable for low leakage currents (and low power losses), whereas a low density

results in a gradual decay of B(U) and in a higher power loss in the pre-breakdown [Gre 1990], Gre 1986].

Quantitatively the defect equilibrium in ZnO-varistors is still badly understood [Mah 1983], [Suk 1988], [Koh 2000],

[Car 2003], but qualitatively this “DSB defect model” [Gre 1990], [Gre 1986], together with today’s knowledge on

the grain boundary films [Chi 1998], explains why so many different material recipes lead to a varistor behavior and

why there are subtle, but technically important differences in the electrical characteristics depending on the recipe

and process used by the manufacturers.

Next we discuss the breakdown region with its high nonlinearities and a well reproducible switching voltage UB of

ca. 3.2-3.4 V per grain boundary, largely independent on the materials formulation. In the past, several models

were able to explain -values up to ca. 20 [Lev 1989], whereas for good varistors -values well above 50 are

realized today. In the l iterature often the term “varistor” is used whenever -values >1 are measured for the J(E)

characteristics, although this is misleading for the present topic. There are various mechanisms in solid statephysics leading to deviations from ohmic behavior, but the real challenge is to explain the high nonlinearities >20.

To achieve such high nonlinearities two ingredients are needed:

i) first a strong pinning of B(U) stabilizing B at typically 0.8-1eV in the pre breakdown

ii) second a mechanism leading to a raid decay of B(U) for U > UB , since for eU >>kT

- (eU/kT)·dB/dU, i.e. U and dB/dU both have to be high for a high nonlinearity!

In the case of a strong pinning of the Fermi level (or B), very high electrical fields build-up on the positively biased

side of the grain boundary. For sufficiently high doping of the bulk of the ZnO ( N0 > ca. 1017

/cm3) the fields at the

interface reach values as high as 0.5-1 MV/cm, high enough to create hot electrons similar to what is known for




Page 75

other semiconductors with high carrier mobility’s (e.g. GaAs). These hot electrons can create holes in the valence

band by impact ionization as soon as their energy near the edge of the space charge region is above the band gap

of ZnO (ca. 3.2 eV). The positively charged holes will diffuse back to the grain boundary within less than 1 nsec.

There they compensate part of the negative interface charge Qi . As a consequence this will lower B and increase

the current across the barrier - and hence also the number of hot electrons created will increase further. This hole

production above a threshold level typical for ZnO will trigger a positive feedback mechanism, which leads to a

rapid decay of B(U) with increasing U and hence to high nonlinearities ! For a sufficiently high hole generation

rate, the energy gain (B+eU) of the electrons must be in the range of 4 eV, which explains the observed switching-

or breakdown-voltage of typically 3.2-3.4 V per grain boundary (B(UB) 0.8 eV, details also depend on N0 and

Ni(E) [Bla 1986]). The stabilizing element in this avalanche-type feedback mechanism is the electron-hole

recombination at the interface.

Figure 2.10: Band diagram scheme of the hole induced breakdow n mechanism[Gre 1989]




Page 76

With this “hole induced breakdown model”, developed by Pike, Greuter and Blatter [Gre 1990], [Bla 1986], [Pik

1984], most of the unusual breakdown phenomena in ZnO-varistors can be (semi)quantitatively understood, like

e.g. the high nonlinearities (>>20), the negative small signal capacitance around UB and the electroluminescence

phenomena observed at the switching point [Gre 1990], [Gre 1998], [Pik 1984], [Pik 1985], [Gre 1984]. The

electroluminescence comes from the fraction of holes, which recombine directly with electrons in the conduction

band, thereby emitting light in the UV-region ( h 3.2 eV ). This direct observation of the band-band

recombination is the most direct evidence for the hole induced breakdown model [Pik 1985], [Gre 1984]. Besides

the valence band states, also defect levels in the band gap of ZnO can be ionized by the hot electrons and their

recombination leads to a strong luminescence in the visible range, like the dominating emission at 700nm (red)

from the doping with Co [Gre 1998], [Pik 1985], [Gre 1984], [Glo 1981], [Cor 1990]. When viewed under the

microscope this strong emission can be observed at every grain boundary along a current filament, which each

light up like a small light emitting diode (see Figure 2.11).

With the model developed by Pike-Greuter-Blatter a good and partly even quantitative understanding is achievedfor the pre-breakdown and breakdown region. This is the most powerful and accepted model today for consistently

explaining the different varistor phenomena. Also the dynamic effects of voltage overshoot under fast pulses can

be understood by this model, although some interface parameters (e.g. interface recombination rates) are not yet

known precisely enough to make the model quantitative for this case [Tua 1988]. No effort has so far been made to

calculate the AC large signal response. Qualitatively the model also explains the slightly higher AC-breakdown

voltage compared to the steady state DC-value UB (see Figure 2.5b) and the observed asymmetries in the resistive

current component, which are due to the charge trapping and de-trapping dynamics [Gre 1990].

Next we briefly address the statistical and microstructural effects, which are caused by the network of grain

boundaries inside a varistor element. In the real microstructure the individual grain boundaries are arranged in a

partly disordered 3-dimensional network and the net current density of a MO-element may well deviate from the

Figure 2.11 : Electroluminescence observed under the light microscope [Gre1990] ,[Gre 19 98] . A surface contact geometry is used with a contact at the top andbottom of the picture. Each grain boundary l ights up by a short l ine roughly

perpendicular to the current fi lament connect ing both elect rodes. Some filamentscan be seen to disappear under the surface and to reappear closer to the other

surf ace electrode. The mater ial i s stressed in the breakdown region U> U B




Page 77

local current density seen by a specific ZnO-ZnO grain boundary. Electroluminescence pictures as in Figure 2.11

[Gre 1990], [Gre 1998] made, for the first time, directly visible that the current flow through a MO-block is of

filamentary nature. Similar insight also can be gained by new imaging techniques with high speed infrared

cameras, as shown by Wang et. al. [Wan 1998], Figure 2.13. Several groups [Bar 1996], [Voj 1996], [Che 2002],

[Zha 2005], [Lee 1999], [Wan 1998], [And 2003], [Bog 2000] now have performed simulations of such random

networks in 2-dimensions (e.g. of Voronoi-type), considering different types of irregularities. They provide

interesting insights into various questions, which are difficult to access by experiments, like the role of disorder, the

influence of the grain size distribution, local variations in nonlinearity and switching voltage, fluctuations in barrier

heights, local hot spots etc.

Figure 2.12: Local distr ibut ion of current i n a random netw ork containing varistor -type grain boundaries

for different positi ons on the U/I-characteri stics: a) ohmic region w ith a homogeneous distr ibuti on of thecurrent density (V= 1.5) , b) in the breakdown wit h clear filamentary conduction (V= 3.5) , c) higher up in

the breakdown region at the tr ansiti on to the upturn (V= 4.3 ) and d) in the upturn-region where thecurrent distr ibution reverts to a more homogeneous sit uation (V= 6) [Bar 1996] . The gray-level

represents the relat ive value of the current through a grain boundary, normalized to the total current.




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Figure 2 .13 : Infrared thermal images taken w ith a high speed camera for t hin slices of low volt agevaristors wit h large grain size [Wan 1998 ]

Depending on the position on the macroscopic U/I-characteristics, the flow pattern through the microstructure can

be quite different and far from homogeneous, as is illustrated by Figure 2.12. In the ohmic regions (U << UB and U

>> UB ) the current flow is rather homogeneous throughout the structure, whereas in the high nonlinear regimestrong current filaments can develop and lead to local overheating, as nicely demonstrated in Figure 2.13. Heat

generation and heat spreading then can become inhomogeneous in the microstructure and may lead to local

thermal runaway and local thermo-mechanical stresses [Wan 1998], [Bog 2000]. For a high irregularity in the grain

size distribution it is possible that two or several filaments join into a single one, which increases the risk of

microscopic hot spots if operated at this condition for a long time (see Figure 2.12). Clusters of several large grains,

as an example, are an obvious condition for creating hot spots [Bar 1996], [Voj 1996], [Che 2002], [Zha 2005], [Lee

1999], [Wan 1998], [And 2003]. For thin samples, as typically used for LV-devices, such statistical disorder effects

are even more critical than for larger volume elements [Wan 1998]. In 3-dimensions some of these reported critical

phenomena are less severe due to the higher number of possible paths available for the current filaments.

However, up to now 3D-simulations have not yet been done for nonlinear random networks, but will certainly come

the more the computational power develops.

In the above network simulations often cases are discussed, where a significant fraction of the boundaries are

assumed to be not varistor-active [Bar 1996], [He 2004], i.e. either ohmic with some assumed conductivity, bad

junctions with poor nonlinearities or even being insulating. The secondary phases like Bi2O3, spinel or pyrochlor

and clusters thereof represent such insulation pockets within the microstructure and this certainly has an influence

on the U/I-characteristics, as experimentally shown by artificially generating such inclusions [Gre 1998]. The reports

of non-active ZnO/ZnO boundaries, however, have to be considered with caution, as experimentally it is extremely

difficult to prepare polished surface structures for local measurements without destroying and short-circuiting the

sensitive grain boundaries. Unfortunately, most authors do not check and comment on such possible artefacts.

Based on the recent understanding that the amorphous grain boundary films are wetting all ZnO-interfaces [Chi

1998], it is rather unlikely that some boundaries should not be varistor-active; local variations in the boundary

properties (barrier heights etc), however are more likely and have to be expected.




Page 79

The simulation studies clearly underline that the optimization and homogenization of the microstructure is

important. This probably is one of the efforts common to all varistor manufacturers. Clear progress has been made

in ceramic process technology over the past decades, as can, for example, be seen in the recent comparative

study on the energy handling capability of major varistor manufacturers [Rin 1997], [Rei 2008] and this report.

Besides the microscopic non-homogeneities discussed above there are also non-homogeneities on the

mesoscopic (ca. 100 m, size of spray granule) and macroscopic ( mm-cm ) scale, which have to be controlled in

varistor manufacturing and which affect the overall performance of a varistor block [Gre 1998], [And 2003], [Ste

2004]. Additional materials challenges to be solved are given by the passivation layer (environmental, dielectric and

thermo-mechanical stresses) and the metallization (adhesion, contacting, rim structure) [Per 2003], [Bog 2000],

[And 2000]. The role of macroscopic (mm-cm) electrical non-homogeneities on the energy handling capability has

been studied by simulations [Eda 1984], [Bar 1996], [Nie 1989] and will be addressed in the following section.

2.2.6 FAILURE MODES OF VARISTOR BLOCKSIn qualification testing and in the field the varistor elements have to cope with a variety of stresses like: long term

stability tests, discharge voltage, temporary overvoltage, long duration impulses (1-4 ms square wave), high current

impulses (90/200 s, 4/10 s etc), short circuit behavior (after forced prefailing; shattering test) etc. Most of these

aspects are discussed at other places in this report or in the literature [Gre 1998], [Bar 1996], [Voj 1996], [Che

2002], [Zha 2005], [Lee 1999], [Wan 1998], [And 2003], [Bog 2000], [Bal 2004], [Rin 1997], [Rei 2008], [He 2004],

[And 2000], [Eda 1984], [Bar 1996], [Nie 1989], [Hag 1997], [Voj 1997], [Len 2000], [Mah 2001], [Miz 1983], [Ste

2004].

Here we only briefly discuss the phenomena directly related to the varistor blocks. The different stresses can lead

to a variety of failure modes, like:

- thermal runaway

- puncture from current concentrations followed by local thermal runaway and melting

- cracking due to localized heating (with or without puncturing)

- cracking due to thermo-elastic stresses during high current impulses (even for a perfectly homogeneous

block)

- flashover from high dielectric stresses at the rim or surface of the blocks.

An ideally homogeneous block can only fail either by thermal runaway (heat input faster than cooling) or by

fracturing under high current stresses, where the stresses exceed the mechanical strength of the material. In

reality, however non-homogeneities are always present, be it in the microstructure or on the macro scale of the

block. These imperfections will give rise to local overheating followed by either puncturing or/and

compressive/tensile stresses leading to fracturing of the ceramics. Various reports can be found in the literature on

the different types and degrees of non-homogeneities observed and how they translate to mechanical stresses

[Lev 1989], [Eda 1984], [Bar 1996], [Nie 1989]. Figure 2.14 illustrates by IR-measurements the electrical non-

homogeneities, which can be present in a non-perfect varistor block [Ste 2004].




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Figure 2.14 :Left : Cross sect ion along the axis of a stat ion arrester block (Ø 63 mm), view ed by a fast IR-cameraupon exposure to a square wave impulse of a few ms duration and a current ampli tude of 100 -200 A

[Ste 2004 ].Right: light intensity prof ile proport ional to the temperature, which in turn is proport ional to the current

flow ing through the block.

For a complete simulation of the failure behavior, the electrical, thermal and mechanical properties of aninhomogeneous varistor block have to be calculated in 3 dimensions in a coupled mode. This is a very demandingtask and has not yet been done with finite element methods. Parts of the problem however have been addressedand are illustrated in Figure 2.15 and Figure 2.16. Figure 2.15 shows the coupled thermal electrical simulation forthe rim of a varistor block, where the metallization is made with an edge margin of 2 mm. The equithermal contour

plots nicely show that this non-metallized rim leads to a current concentration at the metallization edge, where highlocal temperatures can occur and can trigger a puncturing.

In general, puncturing has to be expected, if the local heating is faster than the heat spreading and the ceramiclocally is heated to above the melting point (> 750-850 C), where the melt can be progressively ejected, startingfrom the surface of the hot spot. The local heating simultaneously also creates thermo-mechanical stresses and, ifthese stresses reach the mechanical strength of the material before the hot spot has reached the melting point,then mechanical fracturing can occur prior to puncturing. Alternatively cracking can follow the puncturing, ifsufficient energy is deposited in the channel. Both phenomena, puncturing and cracking under long durationimpulse, hence have the same origin, but the outcome will depend on such parameters as the local energy inputrate, size and geometry of inhomogeneity, local stresses generated, fracture work needed etc.

A deeper insight in these coupled electrical-thermal-mechanical phenomena was first provided by Bartkowiak et. al.by using continuum mechanics and assuming simple failure criteria to study the failure modes for the energytesting of varistor blocks containing hot spots [Bar 1996]. The hot spots were simulated by assuming a small axial,cylindrical area with reduced breakdown voltage (-5%) in an otherwise homogenous block. Radial, axial andtangential stresses were then calculated self consistently on the basis of a theoretical U/I-characteristic for differentcurrent densities and two different block sizes. Although the assumed inhomogeneity is rather high compared totoday’s high quality blocks (and for symmetry reasons was located in the center of the block), the model predictsquite well the regimes for the most likely failure modes to occur: thermal runaway, puncture, cracking under tensionand cracking under compression. The results for a distribution arrester disk are illustrated in Fig. 2.16. Note thatdepending on the range of current density, different failure modes can show up on a block.




Page 81

Figure 2 .15 : Equithermal contour plot for an edge margin of 2 mm of a vari stor block as obtained fromcoupled thermal electrical simulations [And 2000].

Figure 2.16 : a) Time to fai lure t f and b) energy handling capability of a distr ibuti on-class type of MOdisk; the diff erent limit ing failure modes are shown [Bar 1996] .

For the simpler case, where the temperature field across the block is known, the mechanical stresses have beencalculated by Nied [Nie 1989]. Despite the simplifications made the work from Bartkowiak et.al. provides a rather




Page 82

detailed and useful insight into the overall (3D) behavior of distribution and station blocks during the energy

handling capability tests. Note that local hot spots (or channels) can lead to cracking patterns parallel or

perpendicular to the block axis, depending on the impulse amplitude/duration and/or the block size [Bog 2000], [Rin

1997], [Rei 2008], [Eda 1984], [Bar 1996], [Nie 1989].

The above failure modes for energy stresses in the (sub) millisecond to second range have to be clearly separated

from the failures observed under high current impulses, like the 4/10 s pulse. Here a new type of failure mode

appears, which has been identified [Hag 1997] as being a very special case of a thermo elastic stress generation

resulting from the extremely high heating rates produced by this fast impulse: Heating rates up to 107 K/s can be

generated inside a block of a distribution arrester! This temperature rise occurs on a time scale much faster than

the material can adapt in its thermal expansion, which is limited by the sound velocity c of the ceramic (c4-5000

m/s; < h/c, h=block dimension). Even in a perfectly homogenous varistor block this can lead to very high tensile

and compressive stresses. Depending on the block geometry and block properties, these stresses may reach the

material strength and can cause fracture of the varistor block. Hence this thermo-elastic stress is an intrinsic

limitation for all varistor blocks exposed to such short and energetic impulses. This adds the mechanical strength

as an additional parameter for a good material. Existing non-homogeneities in the block not only limit the

mechanical strength of the material, but they can also create additional thermo-mechanical stresses if they lead to

large current non-homogeneities and temperature gradients inside the material [Per 2003], [Bal 2004], [Hag 1997],[Voj 1997], [Len 2000]. Analytical calculations in simple 1D [Hag 1997], [Voj 1997] or full 3D-analysis [Mah 2001],

as well as 3D finite element simulations [Len 2000], predict these oscillating mechanical stresses in the varistor

body under high current impulse, see Figure 2.17. With the full three dimensional models, the reflections of the

mechanical stress waves from the different surfaces of the varistor body are included, showing that the stress fields

vary between tensile and compressive stresses in time and space in a rather complex manner. Figure 2.17 also

compares the simple 1D-model [Hag 1997], [Voj 1997] with a 3D-simulation, showing that the reflections cause

even higher stresses than predicted from the 1D model only. If the stresses exceed the mechanical strength of the

ceramics, the blocks will crack. Very characteristic cracking patterns can be observed, like the one shown in Figure

2.17, where due to the short duration of the impulse quite often no signs of any discharges are present.

Figure 2 .17 :left: 1 st principle stresses for 1D and 3D model [Len 2000] ,right: typical “midplane” crack under high current impulse for a distr ibut ion block with high aspect ratio(height/ diameter) [Hag 1997] , [Voj 1997] , [Len 2000] , [Per 2003 ].

From the simulations it is evident that the size and aspect ratio of the varistor block have a clear influence on the

thermo-elastic stresses generated. For varistors with an aspect ratio (height/diameter) less than about 0.5, the

maximum 1st principle stress is normal to the axis, while for a high aspect ratio the maximum 1

st principle stress has

axial direction, in agreement with the fracture modes observed (see Figure 2.17). Minimal high current stresses are

predicted for an aspect ratio of ca. 0.9 [Len 2000], [Mah 2001].

Today’s simulations do not yet consider possible influences of electrical non-homogeneities, electrodes, metal

spacers, contact forces, encapsulation etc. as present in the real arrester design. Other minor effects also not




Page 83

considered yet in the simulations of the block performance are the slight, beneficial influence of the small positive

temperature coefficient (PTC) in the grain resistivity in the high current range [Cab 2004] or the pressure

dependence of the U/I-characteristics [Dor 1985].

Very little information is available in the literature on flashover phenomena on single MO-blocks [Rei 2008]. Poorcontrol of the metallization edges of course is one of the possible origins, besides such parameters like contacting

electrodes, ambient conditions, surface passivation, surface contaminations etc. Also care must be taken by

separating a true dielectric surface flashover from discharges being triggered by near edge puncture (as e.g.

expected at high energies for rim situations like in Figure 2.15).

2.2.7 LONG-TERM STABILITY OF ZNO VARISTORSIn the early period of development, ZnO varistors showed significant degradation in the U/I-characteristics in the

accelerated ageing test under continuous AC or DC operating voltage: the leakage current and power loss

increased with time and applied voltage, either right from the beginning or after passing through a minimum (see

Figure 2.18). As such changes of the characteristics under long term stress affect the thermal stability of the

arrester; the ageing of the varistor elements has to be assessed before performing thermal stability tests.

Instabilities can be the result of an intrinsic behavior of the bulk of the material or can be related to the near surface

areas, where the atmosphere and the passivation are additional points of concern, as illustrated in Figure 2.19.From the manufacturer’s point of view, the main variables regarding the long term stability are the recipe and the

thermal processes, which both can have a decisive influence and are not well described in the literature. From the

application side, AC vs. DC operating voltage and the surrounding medium may have a significant effect on the

long term stability of a varistor block. From steady material improvements, most of today’s available varistors show

no degradation under AC operating voltage, i.e. a stable or decreasing power loss vs. time in the accelerated

ageing test. Generally, the long term stability is assessed by measuring the power loss at elevated voltage and

temperature (115°C) for duration of 1000 h and stability is defined as a stable or monotonically decreasing power

loss versus time curve. Long term stability under DC voltage however is more difficult to achieve than under AC.

The efforts of realizing a varistors with good DC long term stability are higher and for this application often special

DC material formulations and thermal processing means are used.

As mentioned above, the medium surrounding the varistor may also cause additional degradation of the U/I

characteristic. In general, oxygen in the surrounding atmosphere helps for the long term stability, whereas areducing atmosphere may have negative effects on the stability of the rim area of a block. Internal partial

discharges in a faulty surge arrester may change the gas composition surrounding the varistor blocks to become of

reducing nature and can trigger degradation. The established approach today to verify performance of varistors

under critical atmospheric conditions is to perform accelerated aging tests under N2 or SF6 (for GIS-arresters) with

low oxygen concentration (< 0.1%). This ensures that even in the total absence of oxygen, the long term stability of

the varistor can be granted. Similar, application specific tests are needed e.g. for under oil operation in

transformers, which in addition to the surrounding medium, also have to consider possible higher service

temperatures.

A variety of empirical long term test results with different time-evolutions of Pv(t,T,U) are reported in the literature,

but a physics based, microscopic model for the degradation mechanism is still missing [Lev 1989], [Bue 2008], [Gre

1989], ]Per 2003], [Stu 1990], [Gre 1995], [Gup 1990]. The intrinsic ageing behavior of the bulk material is known to

be a grain boundary phenomenon. Migration of charged defects (e.g. Zinc interstitials Zni or Oxygen vacancies VO)within the space charge regions [Gre 1995], [Gup 1990] and the reduction of the excess oxygen at the grain

boundaries [Stu 1990], [Gre 1995] are expected to take place or have been observed during accelerated ageing

tests. From an electrochemistry point of view, this is not surprising, given the high electric fields present at the grain

boundaries. The charge rearrangements from migrating defects will lead to a time dependent distortion of the

electrostatic potential barriers and can qualitatively explain the observed changes in the U/I-characteristics during

ageing tests [Lev 1989], [Stu 1990], [Gre 1995], [Gup 1990]. For ageing under DC or unipolar impulse stresses the

U/I-curve becomes not only displaced but also asymmetric, with different shapes depending on the polarity [Lev

1989], [Gre 1995], [Gup 1990]. This implies that the space charge regions at the grain boundaries are no longer

symmetric.




Page 84

Note that most of the degradation or polarization phenomena are reversible and can be healed out by heating to

200-300°C without applied voltage. Thereby small thermally stimulated “depolarization” currents are observed [Lev

1989], either due to the migration of the charged defects or the de-trapping of electrons back to their original

equilibrium configurations. Empirically, some dopants are known for their positive or negative influence on the long

term stability [Lev 1989], [Gup 1990], [Fan 1993], [Bin 1993]. It is assumed that they act either directly via forming

migrating or blocking defects or indirectly via their influence on the defect equilibrium in ZnO (e.g. the density of

Zni) and other phases like Bi2O3 [Gre 1995], [Gup 1990], [Fan 1993], [Bin 1993]. For a more detailed understanding

of the ageing phenomena, certainly more research work is still needed on such electrochemical processes

occurring near the grain boundaries and the adjacent triple point phases.

As mentioned above, in the accelerated ageing tests in the early days of ZnO varistors, the materials showed

increasing leakage currents or power losses vs. time. These changes were found to follow, to a good

approximation, an Arrhenius-type law when varying the test temperature [Lev 1989], [Gup 1990]. Based on these

early observations, accelerated ageing procedures were established in the IEC standard 60099-4. An acceleration

factor AF was estimated with AFT = 2.5T/10)

, a test temperature of 115°C was chosen as well as a test time of

1000h. As an example, the lifetime prediction derived from this test would be equivalent to 110 years at 40°C

ambient temperature. In the past, this accelerated ageing procedure provided good confidence on life expectancy

of metal-oxide blocks. For applications with higher ambient temperatures (e.g. 65°C) in general the test durations

must be extended to demonstrate acceptable equivalent life times under such condition (see Table 1). A further

increase of the test temperature to shorten the test time seems not applicable, as a change of the ageing

mechanisms at higher temperatures cannot be excluded, making the extrapolation from this high temperature

range to near room temperatures questionable. Also an increase of the test voltage, as another acceleration factor,

is quite limited. With test voltages above the reference voltage, self-heating and thermal runaway may occur and

new transport mechanisms come into play (e.g. hole generation, Figure 2.10).

0

0.2

0.4

0.6

0.8

1

1.2

0 100 200 300 400 500 600 700 800 900 1000

time[h]

p o w e r l o s s r a t i o

P

/ P o

stable

unstable

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 100 200 300 400 500 600 700 800 900 1000

time[h]

p o w e r l o s s r a t i o

P

/ P o

stable varistor in air or N2

unstable varistor in air

unstable varistor in N2

Figure 2 .18 : Power loss rat io vs. time forstable and unstable varistors during

accelerated ageing tests at 115°C and

slightly elevated AC operating voltage.

Figure 2 .19 : Power loss rat io v s. timedepending on the surrounding medium (air ,

N 2 ) dur ing accelerated ageing test at

115 °C and cont inuous AC volt age.




Page 85

Upper limit of ambient temperature

°C

Test duration at 115 °C

h

Equivalent time at upper limit ofambient temperature

Years

40 1000 110

65 2000 2295 7000 5

Table 2.2: Test duration and equivalent time at upper limit of ambient temperature according IEC60099-4

The problem with the presently used accelerated ageing concept is that most of today’s established materials show

a decreasing power loss with time and the decrease normally is higher at higher test temperatures. Hence these

stable materials show behavior just opposite to an Arrhenius-type law, making the standardized test rather

questionable and without a sound physical basis. Presently we do not understand the microscopic origin of the

initially fast and then slower continuous decrease of the power loss observed in most established varistor materials

nor do we have a deeper insight in such cases as shown in Figures 2.18 and 2.19, where the losses go through a

minimum, followed by a steady increase. Such behavior suggests that several mechanisms might be at work on

different time scales. It is well possible that the increase after the minimum corresponds to a mechanism of

instability and perhaps is thermally activated as in the case of an Arrhenius-type mechanism. If this holds then the

standardized test procedure from the early days might still be basically appropriate and only would need some

reformulation. Certainly, further research efforts are also needed here and the existing test procedure should be

kept until an improved understanding puts this sensitive issue on a better founded basis.

2.2.8 TRENDS AND OPEN ISSUES

Overvoltage protection on all voltage levels from electronics to extra high voltages with passive components will

remain a vital technology and further developments will be realized as manufacturing technologies and basic

understanding make further progress. Overall the last decade has seen a decline in fundamental material research,

if we specifically look at the varistor technology. But much progress has been made regarding the process- and

manufacturing-technology of the ceramic as well as in numerical simulation techniques. A variety of new

microscopic characterization methods have been developed, directly probing materials on the atomic scale and we

can expect that these tools wi ll be helpful to further understand the fascinating mechanisms behind the varistor

materials as well. Also varistor research has moved away from solely looking at the electrical function and has

addressed also additional aspects like their mechanical properties [Bal 2004], [Hag 1997], [Len 2000], which are of

concern for new compact designs and for high stress situations. Various new developments have appeared which

are still in the development phase or are currently making their way into new products. As an illustration we add a

few examples:

i) High field varistor materials and devices, which allow to build more compact components, like e.g. for

GIS-surge arresters, light weight (gapped) line arresters etc.

ii) Alternative varistor materials have and will be a topic of ongoing research. Examples, where materials

with reasonably high, controlled and reversible nonlinearities have been demonstrated are varistors based

on SrTiO3, CaCu3TiO12, grain boundary doped SiC, SnO2 etc. Much progress has been made in the area

of SnO2-varistors and electrical characteristics approaching the one of good ZnO-materials have recentlybeen demonstrated [Bue 2008], [Met 2007]. However, ZnO remains an unique material when it comes to

such practical aspects like availability, costs, purity, particle size, electrical transport properties, economic

sintering temperatures etc and it will be hard to beat it for the big volumes of varistor applications.

iii) Microvaristors as a by-product of the traditional varistor manufacturing have been developed and are

used as a functional filler in polymers, e.g. for f ield grading in cable accessories [Str 1995], [Str 2001].

iv) Multilayer varistors: Impressive and hardly published progress has been made for applications in the

electronics area, where these protection devices of (sub)mm-dimensions are produced on highly

automated production lines in huge numbers or were thick-film varistors with substantially improved

characteristics are evolving.




Page 86

v) Varistor integration: New concepts and ideas have been evaluated and partly demonstrated for

integrating the surge arrester function into other components [Per 2005]. This becomes possible with the

improvements in the design- and simulation-methods as well as the further progress in arrester

manufacturing. In oil filled distribution transformers integrated solutions show very positive field records in

Japan and USA and proof that integrated solutions can be attractive. Further work, in particular on the test

philosophy, however, is needed to gain a broad acceptance of such new approaches.

Regarding the basic understanding of ZnO varistors, several topics still lack a deeper and quantitative

understanding and hopefully wi ll be addressed in future research activities. A very challenging area certainly is the

ageing mechanism, which seems to be closely linked to the poorly understood defect chemistry in ZnO. The latter

also is of interest for understanding the role of individual dopants and the electrical activation of the grain

boundaries. With the recent discovery of p-type ZnO thin film materials, a revival of ZnO-research has started and

this will also add to a deeper understanding of the atomistic phenomena in varistor materials. A puzzling issue,

which is not conclusively treated in the literature, is the true quality and spread in the junction properties within a

varistor ceramic, where a certain fraction of inactive or bad varistor grain boundaries is reported, however without

quantifying the possible artifacts from sample preparation.

For the transport properties, there is quite a good understanding available today for the DC- and small signal AC-behavior. For the large signal AC-response at 50 or 60 Hz however no attempts have been made so far to apply

the existing junction models to describe the capacitive and nonlinear resistive current components in a quantitative

way. The same holds for the different impulse shapes used in the upturn region, where only a preliminary

simulation study on the transient response has been made so far [Tua 1988]. A related field is the more frequent

use of varistor elements in power electronic circuits, where very steep transients with modest energy stresses, but

very high repetition rates, are typical and only limited knowledge is available today from experiments and from

modeling.




Page 87

2.3 Design of surge arresters

Authors in charge: Volker Hinrichsen and Bernhard Richter

2.3.1 FOREWORDLatest by end of the 1980s MO arresters had definitely been established as state of the art, since their technical

and commercial benefits are quite evident. MO arresters offer low protection levels, high energy absorption capa-

bilities, and stable operation even under severe pollution conditions and lifetimes which easily may exceed thirty

years. Knowledge about principal design of conventional porcelain housed MO arresters can be presumed;

therefore this technology will not be addressed here. However, the very simple structure of an MO arrester – the

active part basically consists of a stack of cylindrical MO resistor elements – supported the development of polymer

housed arresters at a very early stage. They were introduced for the first time in the mid-1980s in the distribution

voltage level. After about 15 years of development some few basic design principles of polymer housed arresters

can be distinguished with a variety of individual sub-solutions. While there are less basic design variants for

distribution than for HV arresters (only cost effective, no "high tech" solutions are applied), the variety of sub-

solutions is much larger, one reason for this given by the fact that there are far more manufacturers of distribution

arresters on the market than in HV. Accordingly, more design variants had to be created in order to avoid patent

conflicts. As mentioned, for HV arresters at least one more basic design principle is applied – the use of compositehollow core insulators – but there are not so many sub-solutions, as there are not too many manufacturers of HV or

EHV polymer housed arresters worldwide. Therefore a classification of today's polymer housed HV arrester

designs is comparatively easy. It must be mentioned, though, that there still does not exist any official

nomenclature. Designations like "Type A" or "Type B", as they are used in this contribution, must not be mixed up

with other emerging classifications. For instance, the new IEC document 37/317/CDV on arrester short-circuit

testing [IEC 37] has introduced "Design A" and "Design B" arresters. These designations have different meanings,

as they serve for classification with regard to short-circuit performance only.

This contribution exclusively focuses on constructional design principles. Other important design aspects such as

the different performance characteristics can be looked up in [Hin 2003].

2.3.2 DESIGN PRINCIPLES OF POLYMER HOUSED HV ARRESTERS

"HV" actually ranges from Us = 72.5 kV up to Us = 800 kV (higher levels do exist but do not play an important roleso far). This wide range may be further sub-divided into two parts where different philosophies govern the decision

process for an arrester purchase. In the voltage levels up to Us = 300 kV, i.e. the lower transmission and the sub-

transmission levels, in most cases just technical standard requirements apply. There is only little need for special

features like extra-high mechanical strength or safety considerations. These are the voltage levels of standard

applications, where more and more the same criteria as in distribution systems are applied and not too much time

or money is spent to optimize the arrester layout for a particular location. It is the domain of "low cost" (in its

positive meaning) arresters. For the EHV levels, Us = 360 kV and more, requirements especially on mechanical

characteristics play an increasingly important role, which cannot easily be fulfilled by the "low cost" designs.

Further, users are less willing to take any risk of possible arrester failures. The electrical and mechanical

requirements on the arresters are often evaluated by system studies, and in many cases the user has detailed

knowledge and information about the system configuration and clear ideas about the optimal arrester for his

particular application. This is the domain of "special feature" arresters. Both types of arresters are available today in

polymer housed design.

2.3.3 THE MECHANICAL SUPPORTING STRUCTUREFigure 2.20 gives a classification of the mechanical design principles of arresters, which is not only limited to

polymer housed arresters. According to this suggestion a differentiation is made between designs using a hollow

core insulator with an intentionally enclosed gas volume and such designs, where the housing is put onto the MO

column without any intentional internal gas volume. With respect to the polymer housed arresters, the mechanical

designs can be characterized as follows:

Type A: This design – the "tube design" – is a more conventional approach, looking quite similar to that of a

porcelain housed arrester as shown in Figure 2.21. The stack of MO resistor elements is mechanically supported

by an internal cage structure, for example made from FRP rods. This insert is clamped between the end flanges by




Page 88

help of compression springs. Additional supporting elements (not shown in the figure) may be necessary to fix the

insert in radial direction. What is important and sometimes criticized is the fact that this arrester due to its enclosed

gas volume needs a sealing and pressure relief system. This has in fact to be designed and manufactured with the

same care as it is the case with porcelain housed arresters. But just for this reason this does not constitute a real

problem for an experienced manufacturer of HV arresters. There are numerous makes of HV arresters on the

market, which have an excellent service record (over twenty years or even more) also with respect to their sealing

systems. Therefore, whether this arrester design has problems with moisture ingress or not is a matter of the

manufacturer's know-how and production quality, as it has always been with porcelain housed arresters and – by

the way – as it is also with the other designs of polymer housed arresters, which by far are not all inherently leak

age sealed.

Figure 2.20 : Classifi cation of basic arrester designs; cross sectional side and top v iew s

Porcelain;

Polymer Type A

Polymer Type B1a Polymer Type B2Polymer Type B1b

Type A: gas volume enclosed, separate sealing system, pressure relief ventsType B: no intentional gas volume included

...1: wrapped mechanical structure or tube... a: FRP material directly wrapped onto the MO blocks... b: FRP tube with distance to the MO blocks, gap filled by other material

... 2: cage design

MO column FRP supporting structureGas

Solid/semi-solid material Outer housing Metal end fittings

Porcelain;

Polymer Type A

Polymer Type B1a Polymer Type B2Polymer Type B1b

Type A: gas volume enclosed, separate sealing system, pressure relief ventsType B: no intentional gas volume included

...1: wrapped mechanical structure or tube... a: FRP material directly wrapped onto the MO blocks... b: FRP tube with distance to the MO blocks, gap filled by other material

... 2: cage design

MO column FRP supporting structureGas

Solid/semi-solid material Outer housing Metal end fittings




Page 89

Figure 2.2 1: Polymer housed arrester Type A

Another question in this context is if vapor could permeate directly through the sheds and walls of the housing or

through the bonding area between flanges and FRP tube [Hin 1994]. Both investigations and service experience

(first arresters of this design have been installed in 1990) have shown that this is not the case. The amount of

moisture ingress due to these mechanisms is below the quantities which can pass through a good sealing system.

Thus this is no issue, since these quantities can easily be controlled by internal desiccants, as it is done in nearly

every HV device in the electric power system. Actually, some research is being done in order to better understand

these mechanisms and to derive minimum design requirements on composite hollow core insulators used for

arrester applications.

Quite evidently, the Type A arrester cannot be cheaper in production than a comparable porcelain housed arrester,

since composite hollow core insulators are still far more expensive (on the market) than porcelain housings. The

Type A arrester is therefore the typical "special feature" arrester, which in most cases offers technical advantages

both over the "low cost" polymer housed and the porcelain housed arresters – and which have to be paid for. Someof the potential benefits of the Type A design are extremely high mechanical strength, the safest possible short-

circuit performance, or the possibility of making tall units which can serve as single-unit arrester up to 300 kV

system voltage. It will greatly depend on the market price development of composite hollow core insulators if this

arrester design will be mainly limited to EHV or not in the future.

Type B1a: this type, which is often called the "wrapped design", was basically the very first design principle of

polymer housed arresters, when they were introduced in distribution in the mid-1980s. It has then be extended to

HV arresters and can be found in HV also for another reason: one possible (and occasionally implemented) way of

building an HV arrester is to connect a large number of distribution arresters in parallel and in series. Common to

all Type B1a arresters is that the FRP mechanical structure is directly wrapped onto the MO resistor elements (in

some cases applying a thin intermediate foil between the MO stack and the wrap). It can be imagined that this may

be done in nearly infinite ways, and a large variety of sub-solutions has been brought to the market for technical,

commercial and patent reasons. Without being complete, Figure 2.22 gives an idea about some of the main differ-ences.

Flange with ventingoutlet

Sealing ring

Pressurerelief diaphragm

Compressionspring

MO resistor column

Composite

hollow core insulator

Top cover plate

(FRP tube/ rubber sheds)

Flange with ventingoutlet

Sealing ring

Pressurerelief diaphragm

Compressionspring

MO resistor column

Composite

hollow core insulator

Top cover plate

(FRP tube/ rubber sheds)




Page 90

Figure 2.22 : Internal designs of Type B1a arresters

Figure 2.23: Internal design of Type B2 arresters

Possibly the most economical variant is shown on the left. Fiber glass rovings soaked in uncured epoxy resin or

pre-impregnated ribbons are wound crosswise around the MO stack, and the module is then cured in an oven. The

resulting rigid ribbons provide the required mechanical strength. They do not fully overlap and thus form rhombic

"windows". These are important, technically for short-circuit performance and commercially for minimizing the

amount of material (which is also a technical concern in order to minimize the amount of inflammable material). If

the windows are too large, however, the mechanical strength of the module may become insufficient.

In the middle of Figure 2.22 a variant is shown, where no windows remained open, implemented by full overlapping

of the ribbons or by using pre-impregnated FRP mats with appropriate orientation of the glass fibers. This gives

high mechanical strength but forms a closed tube, in which internal pressure can be built up in case of overloading,

possibly leading to violent breaking of the housing. In order to improve the short-circuit performance slots can be

provided on the surface, which function as predetermined breaking points.

The variant on the right hand of Figure 2.22 also shows a design, which is completely closed, realized by a pre-

impregnated mat wound around the MO stack. But in this case the glass fibers are nearly exclusively arranged in

axial direction. This is also a possible means to improve the short-circuit performance: if carefully designed the tube

will easily tear open in case of an internal pressure build-up.

Arrester Type B1b is quite similar to the closed tube variants of Figure 2.22. The difference is in the manufacturing

process. The tube is not produced by wrapping FRP material onto the MO column. Instead, a pre-fabricated FRP

tube is used, which must have a diameter larger than that of the MO column, of course, in order to push it over.

MO column FRP wrap

main orientation of glass fibers

MO column FRP wrap

main orientation of glass fibers




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The resulting gap between MO and FRP material is then filled by solid or semi-solid material. Again, slots may beprovided on the surface in order to improve short-circuit performance.

Type B2: this is a completely different design concept and usually called the "cage design". While the mechanical

strength for Type A and Type B1 arresters is exclusively provided by the FRP structure (closed or partly open tube)this is done by the MO resistors themselves with Type B2. For this purpose, they are clamped between the metalend fittings by FRP loops or rods, applying an enormous axial pre-stress in the range of 100 kN [Ste 2003]. Thebasic mechanical design is shown in Figure 2.23. The left variant is designed with loops, which are fixed in notchesin the end flanges. This design was first introduced for distribution arresters [Sch 1996] and then extended to HV. Asub-variant (Figure 2.23, middle) uses an additional bondage from polymeric material in order to achieve themechanical and short-circuit characteristics required for application in HV and EHV [Sky 2002]. Also the othervariant of Type B2 arresters (Figure 2.23, right) was first realized in distribution arresters and then furtherdeveloped for HV. Here, FRP rods are applied, which are mechanically fixed in holes in the end flanges by aproprietary clamping system. Figure 2.24 shows photographs of real arrester modules as they are produced by twodifferent manufacturers. Main technical advantages of the cage designs are that they offer comparatively highmechanical strength combined with an inherently good short-circuit performance.

With regard to the commercial aspects it is impossible to give any statement here on production cost, which greatlydepends on the total manufactured quantities, the degree of automation and process optimization, the quality of theapplied materials, the degree of type diversification and so on. However, it can be noticed that the Type B arresterin general constitutes the most economical way to produce an arrester. At the same time it offers a technicalperformance, which in most cases ranks between comparable porcelain and polymer housed Type A arresters. TheType B arrester is therefore the typical "low cost" arrester mentioned earlier, one of the reasons for the success ofthis design on the market.

Bondedby

moldingprocess

Sealing

Gap filled

by solidmaterial

Bondedby

moldingprocess

Sealing

Gap filled

by solidmaterial

Figure 2.24: Inter nal mechanicalstructure of Type B arrester

Left : loop design, Right: rod design

Figure 2 .25 : Possible implementat ionsof seal ing




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2.3.4 OUTER HOUSING AND SHEDS As for the mechanical design, there are numerous possibilities to implement the outer housing and sheds. With

regard to material, however, with only few exceptions there has been a clear tendency towards SR (silicone

rubber). All other materials, such as EPDM (ethylene propylene diene copolymer), EPDM/SR blends, EVA

(ethylene vinyl acetate), which are widely used in distribution and may perform well there, are usually not being

accepted in HV or even EHV. The reason is quite obvious: only SR offers hydrophobicity (i.e. the ability to repel

water from its surface), which lasts for decades, and from its chemical structure it is inherently least sensitive to

solar radiation, because its basic component – poly-dimethyl-siloxane – has bonding energies above the intrinsic

energy of UV light, the main aging factor for polymeric materials. Again this is a benefit, which has to be paid for.

Market prices for SR may be in the range of two times the prices for an EPDM.

For Type A arresters there has never been an alternative to SR, for production as well as for performance reasons,

since these arresters mostly belong to the family of "special feature" arresters, where traditionally "high-end"

materials have been used. From a production point of view, the only way to cover a hollow core tube is by using

SR. One of the very first designs used an insulator with sheds from HTV (high temperature vulcanizing) SR indi-

vidually slipped over the FRP tube [Hin 1994], but today's composite hollow core insulators are in the majority

covered by a direct molding process, using RTV (room temperature) SR or LSR (liquid silicone rubber). The latter

obviously will be the material of the future, offering some important benefits in production, i.e. a reasonable

compromise between ease of handling, process temperatures and pressures and vulcanization time (not to forget

the market price!).

The same applies for the Type B2 arresters, which only can be covered by a direct molding process. All common

types of SR can be found with the actual arresters of this design on the market.

Most alternatives exist for the Type B1 design, where the housing basically can be produced either by direct

molding or by pre-fabricated housings slipped over the modules. The latter concept offers highest flexibility in

production [INMR 2002], but special care is required for its implementation. Since most of the Type B1 designs do

not have a smooth surface, a sealing material (e.g. a silicone compound) must be put between the internal parts

and the outer housing. This has to be done in a way to ensure absolute freedom from internal voids, which would

affect the long term performance (potential locations of partial discharges and moisture). Further, an appropriate

sealing system at the end fittings must be provided. Figure 2.25 shows possible ways of implementation. Thealternative shown on the right, which inherently offers the best reliability due to the chemical bonding of the housing

to the end fittings, can only be achieved by direct molding and is therefore reserved to the SR insulated designs.

It should finally be recalled that it is impossible to create a good housing from poor materials, but that it is easily

possible to make a poor housing even when applying excellent materials. In other words: besides the discussion of

materials it should never be forgotten that the final performance of the product is also, if not mainly, influenced by

the design and – in the long term – by the production quality. Only the first aspect might be evaluated by

accelerated aging tests (e.g. the "weather aging test" under salt fog). The latter, however, bears a remaining

permanent risk and should carefully be considered.

2.3.5 DESIGN PRINCIPLES OF POLYMER HOUSED MV ARRESTERSFor distribution arresters the designs can be grouped according to the manufacturing technique and the internal

structure, similar to the differentiation of the design principles of HV arresters. In the distribution field only very fewdesigns have a type A (“tube design”), and if, for special applications only. The designs of distribution arresters are

differentiated as follows:

Group I: The polymeric material (e.g. silicon) is directly molded onto the internal parts: the MO-resistors and the

mechanical structure. A primer is used to ensure chemical bonding of the different materials with the silicon. End

caps are not needed.

Group II: The insulating housing is pressed or slipped over the separately manufactured active part. The materials

are only attached mechanical one to the other. A sealing system with end caps is needed.

Group III: As Group II, but with considerable internal gas space. This gas space may be intended, due to the

design, or not intended due to an uncontrolled manufacturing process.




Page 93

The differences in the design differentiation of HV and MV arresters are due to the fact that in HV the mechanical

strength of the design is of more importance, while in the MV field the production process is of higher importance

(for cost reasons). The design principle B2 for HV arresters is almost identical to Group I for MV arresters.

Figure 2 .26 : Principle designs of medium vol tage arresters

Left: design group I, middle: design group II and III, right: example for design group I before molding insilicon

2.3.6 CONCLUSION

Polymer housed HV arresters actually have a market share of roughly 30%. First installations have been madearound 1990, and so far there is no indication that they will not show the same good performance as MO arrestersin general or polymer housed distribution arresters, which have successfully been in service for nearly 20 yearsnow. The technology can be considered mature. Thus it can be predicted that the share of polymer housedarresters in HV will continuously increase, because they usually offer economical as well as technical advantagesover porcelain housed designs.

Different design principles have emerged in the meantime, which basically can be divided in Type A arresters("tube design"), Type B1 arresters ("wrapped design") and Type B2 arresters ("cage design"). The Type B designsusually have lower and the Type A designs higher market prices than comparable porcelain housed arresters.These different designs normally serve different market segments, which can be classified as "low cost" or "priceorientated" on one hand and "high performance" on the other. Primarily this is a matter of the system voltage level.

In systems of 72.5 kV Us 300 kV ("HV" systems) mostly the "low cost" variant is preferred, since this is thedomain of standard applications where no exceptional requirements on electrical or mechanical performance exist.The majority of polymer housed arresters have been optimized for this market. The arrester designs coming into

question here are the Type B designs, where in general the B2 variant offers higher mechanical strength.

In systems of Us > 300 kV ("EHV" systems) mechanical requirements usually favor the Type B2 or even the Type Adesigns. Especially with the latter nearly any required mechanical, electrical or safety feature can be achieved,which has its price, however.




Page 94

2.4 Special designs of surge arresters

Author in charge: Roger Perkins

2.4.1 SEPARABLE AND DEAD FRONT ARRESTERS2.4.1.1General CommentsThese two distribution-type arrester are similar in that each is typically used to protect connections between

underground cables and ground-mounted (pad-mounted) or underground transformers. As such they are usually

directly attached to the cable connectors on each phase and between high potential and earth. In this sense they

offer optimal protection of the cables and the transformers since the protection distance (between arrester and

equipment) is very small. It is worth noting that the installation enclosures can easily become submersed during

flooding or other events.

Appropriate standards for separable connectors, of which these arresters form a sub-group, are IEC 60502-4, IEC

61442 and IEE 386. These contain related test procedures. Appropriate arrester standards are IEC 60099-4 and

IEEE 62.11. They contain definitions for separable and/or dead front arresters as follows:

- IEC – an arrester assembled in an insulated or screened housing providing system insulation, intended tobe installed in an enclosure for the protection of distribution equipment and systems. Electrical connection

may be made by sliding contact or by bolted devices; however, all separable arresters are dead-break

arresters.

- IEEE – an arrester assembled in a shielded housing providing system insulation and conductive ground

shield, intended to be installed in an enclosure for the protection of underground and pad-mounted

distribution equipment and circuits.

Note: these arrester types are often referred to as “elbow” arresters, with reference to their characteristic form or

shape.

2.4.1.2 Differences between Arrester DesignsFigure 2.27 shows an example of a dead front arrester that illustrates the major elements of the design. Critical are

the insulation material, which is typically either EPDM or silicone rubber, the internal arrangement of MOV disks

and the methods by which they are contained, the separable high-potential contact system and its termination, the

earth connection, and the internal and external shields that are made of typically graphite-containing and therefore

semi conducting EPDM or silicone rubber.

Figure 2 .27: Appearance and internal design of a typical dead front, separable metal- oxide ar rester. Inthis example the connector is also equipped w ith load-breaking capabilit y.




Page 95

Various methods are used to contain the MOV disks within the housing. They may be contained by composite

elements similar to those used in overhead distribution arresters. Alternatively they may be directly bonded to each

other with appropriate conducting adhesives. Less favorably they may only be contained within and by the housing

itself and its ground termination elements.

The major difference between these two arrester types lies in the presence or not of a screened or earthed shield

that provides solid protection whilst deliberately or accidentally making contact with the device whilst under

potential. The external screen is either a discrete element separately molded onto the insulated housing or simply

but less favorably painted onto this housing.

2.4.1.3 Special Test ConditionsThe characteristics of separable and dead front arrester are fundamentally the same as those for any other metal-

oxide surge arrester. Correspondingly they are generally subjected to the same conditions of routine and type

tests. However there are differences in the operating conditions of these arrester types that require consideration in

their test methods and these will be described in the following.

Environmental and Accelerated Aging Tests

As stated above these arrester types are usually contained within enclosures that contain other types of electrical

equipment, usually transformers. For this reason the maximum ambient temperature of the arrester is usually

higher than that in other distribution arresters, there is little air-flow and there is little direct exposure to atmospheric

conditions. The upper limit of the ambient temperature is usually elevated to +65ºC. During the accelerated aging

test the ambient temperature is retained at 115ºC but the condition of elevated ambient is given special

consideration in that the test duration is extended from 1’000h to 2’000h. Of course, the test samples for this test

must be contained within a housing equivalent to that in the actual device; this should include the screen, if used.

The applied voltage during the aging test should represent the maximum that the MO resistors will experience in

the application. Especially for screened arresters this may require special determination e.g. field calculations, to

ensure representative aging conditions.

The elevated ambient temperature should be given consideration during other tests such as both the high-current

short-duration and low-current long-duration impulse withstand test, as well as the operating duty and temporary

overvoltage withstand tests.

An environmental test of the type necessary for outdoor equipment is not required for separable and dead front

arresters. Neither salt-fog tests nor exposure to UV radiation tests are necessary. Whilst the familiar multi-stress

tests applied to outdoor arresters are also not suitable for those considered here, it is certainly necessary to

consider equivalent tests that achieve the same purpose of combining a variety of severe application conditions

that place extraordinary stress on the device. Such tests can include immersion in a high-conductivity liquid

medium at elevated temperatures and voltages for extended periods of t ime. Whilst these tests are not yet

established in the relevant standards, they are used by qualified manufacturers who wish to ensure good field

performance of their products.

Insulation withstand tests

This testing is carried out on the arrester housings without active internal components. It is particularly important forseparable and dead front arresters because of the particular electrical stresses existing in the relevant enclosures,

the short inter-phase separation of the devices and, in the case of the dead front design, the high internal stresses

generated by the proximity of the ground plane to the conducting parts. High potential withstand testing must be

extended to include partial discharge testing.

Short circuit tests

This test is very important for the type of installations and arresters considered here. This is because of the

importance of the equipment contained within the enclosures and the necessity to prevent both collateral damage

to it and major electrical breakdown and associated system interruptions.




Page 96

For this reason, in cases where the fault currents can be high i.e. in low-impedance earthed systems, the failure

mode during short circuit testing is required to be benign with respect to the immediate surroundings. This usually

means that the housing of the device may either not vent any debris at all or it may only do so in a downward

direction away from the neighboring phases.

HCSD test

The conditions of this test are very similar to those for other arrester designs (apart from the ambient temperature).

However it is worth noting that generally high current impulses do to penetrate underground systems to the same

extent as overhead systems. The rated high-current short-duration impulses should therefore be lower. For

example, 40kA 4/10s may be considered adequate where an overhead system requires 65kA 4/10s.

For dead front-installations this test is more severe than for live-front. The reason is that the proximity of the shield

or ground plane increases the electrical stresses on the outer edges of the MOV disk to much higher values that

can cause it to fail, where it would otherwise withstand the same impulse current conditions in a live-front

assembly. This means that the dead front arrester must be HCSD tested with a sample (model) representing the

same electrical situation as in the final

2.4.2 UNDER-OIL ARRESTERS

2.4.2.1 General CommentsThis type of arrester, also referred to as “liquid-immersed” arrester in the standards, is typically mounted inside the

tank of a transformer (most typically MV but also HV) and therefore exposed to the higher temperature and

potentially corrosive nature of its fluid Kno 1985, Kno 1986, Hen 1989. An example is shown in Figure 2.28. Its

particular benefit is its proximity to the transformer core windings and therefore the optimal protection to

overvoltage, and in particular to steep transients, that it offers them. Experience shows that failure rates of both

arresters and transformers are significantly lower. Further advantages are correction of capacitive effects, a space

saving assembly, factory testing of complete system, reduction of on-site assembly cost and increased personal

safety. However, there are other benefits worth mentioning. In particular the influence of environmental conditions

typically afflicting outdoor arresters is absent. There is also usually no outdoor-type housing, which removes a

performance-influencing variable. Potential problems are cover retention during fault currents, testing and failuredetection.

Figure 2 .28: Appearance and internal design of a t ypical under-oil meta l-ox ide arrester .

The under-oil arrester has been employed since 1980 in the USA, with more than 500’000 units installed, and since

1987 in Japan with an even larger quantity. They have in each case demonstrated excellent reliability as devices




Page 97

and have significantly reduced the lighting-related damage to transformers. In the case of Japan, this has been

impressively documented [Ish 2004].

2.4.2.2 Special Test ConditionsSpecial Note on Transformer Test Precautions

Since the under-oil arrester is factory-mounted inside the transformer tank it must be fitted with a bypass or

disconnect switch in order to isolate it from the overvoltages used during testing of the transformer . If this is not

done the arrester will certainly be damaged!

Elevated Ambient Temperature

As mentioned above the typical ambient temperatures experienced by the under-oil arrester are significantly

higher. A maximum ambient temperature of +95ºC and a maximum temperature of the device +120ºC are usually

assumed in the relevant standards. Both IEC 60099-4 and IEEE C62.11 specifically address these arresters.

The elevated ambient temperature and the immersion liquid should be given consideration during other tests such

as both the high-current short-duration and low-current long-duration impulse withstand test, as well as the

operating duty and temporary overvoltage withstand tests. It should be noted that no special test models are

defined in the standards but they should meet the general requirements regarding similarity to the final device with

respect to the surrounding medium and to the thermal equivalence. For instance the HCSD test is done in the liquid

at 75ºC ± 5ºC during and after test at MCOV, whilst the LCLD and duty cycle conditioning are carried out in the

liquid at 20ºC ± 5ºC during the shots and 120ºC ± 5ºC prior to 19th shot. Similarly the temporary overvoltage, TOV,

test also requires to be done at 120ºC ± 5ºC

Environmental and Accelerated Aging Tests

Accelerated aging test are carried out at 115ºC ± 2ºC, which is the same as for other arrester types. However the

duration of energization at MCOV is 7000 h whilst heated in mineral insulating oil meeting ASTM D3487-00

requirements. This is considerably longer than is usually the case for other arrester types but is necessary not only

because of the higher ambient temperature but also because of the exposure to the transformer or other fluid. So

this test takes the place of environmental tests.

Short circuit tests

A particular requirement of the under-oil arrester is that it should be specifically designed and tested to provide

either a failure mode of either open-circuit or closed-circuit during short circuit conditions. This must be specified by

manufacturer. The reason for this is that the conditions of application must allow for this difference. For example, a

fail-open design will mean that should the arrester fail, it wi ll no longer provide overvoltage protection and will in

addition not easily be identified as failed (since it is not visible from outside). In contrast a fail-short design will both

short-out the installation and subject it to full fault currents. The installation must allow this to occur without

collateral damage.

The short circuit test must be carried out in the actual application condition, which means inside the transformer

tank. This is necessary to ensure that the installation fails safely in the event of a short-circuit caused by the

arrester.

The conditions of the test are similar to those for other arrester types in that a thermal failure is generated by

application of a suitably sustained power-frequency overvoltage. It should be noted that no interruption capacity is

expected from the arrester.




Page 98

2.5 SF6 gas insulated MO surge arresters


SF6 gas insulated MO surge arresters (GIS arresters) are nowadays used in sub transmission and transmission

systems up to the UHV range of Us = 1100 kV.

Main advantage of GIS arresters over air insulated designs are their favorable performance under seismic stress

and the excellent behavior under pollution conditions, the very high availability and the possibility to be integrated

into the SF6 substation for optimized protection of the equipment [Pry 1998].

SF6 gas insulated substations (GIS) are normally well protected by arresters installed at the line entrance only,

whereas large stations must be protected by additional arresters installed at suitable locations inside the GIS.

Traditionally, the arresters that are installed inside the GIS have a similar rating as those arresters installed at the

line entrances.

Figure 2.29 : Typical instal lat ion of a GIS arrester at the line entrance of a gas insulat ed substati on (Us =420 kV) , example Siemens.

For GIS arresters a main engineering target is a compact, space saving design [Sch 1992]. GIS arresters

principally consist, as all other MO arrester designs, of one or more parallel columns of MO resistors installed within

a housing, in this case an earthed metal vessel filled up with SF6 gas. The individual columns are built up by

connecting MO resistor elements in series. In order to achieve an economical and space saving design and to

minimize the impact of stray capacitances to the earthed vessel often a meandering mechanical design of the

active part has to be used. This reduces the overall physical length of the active part and additionally contributes to

reduced self-inductance1. Insulating plates of extreme high electric withstand have to be applied in order to insulate

the layers of MO resistors from each other. Figure 2.30 shows the principle design of the active part of an EHV GIS

arrester. Each of the insulating plates is electrically stressed by the voltage drop across eight MO resistors.

1 The self-inductance per unit length of GIS arresters is typically assumed as 0.3 µH/m, while air insulated designs have a

typical value of 1 µH/m.




Page 99

Figure 2.30 : Design of the active part (mechanically three columns, electr ically one column) of an EHVGIS arrester using conventional MO resistors.

Due to the very short radial distance between the active part on high voltage potential and the earthed vessel a

relatively high capacitive stray current is f lowing, which leads to an unfavorable axial voltage distribution along the

active column of MO resistors. For this reason countermeasures such as metallic grading elements (hoods or rings)

or capacitive grading elements have to be taken. Figure 2.31 shows the principle design of GIS arresters for

different system voltages if standard MO resistors with 200 V/mm are used. For systems up to 170 kV system

voltage GIS exist with only one phase or with all three phases in one metallic enclosure. Accordingly, the GIS

arresters are designed in the same way. Up to a system voltage of Us = 170 kV the MO column consists typically of

a linear column. For higher system voltages generally a mechanically three or four column design is used [Göh

2006].

Figure 2.31 : Design of GIS arresters (principle) for di fferent system vol tages. Left: single phase design forsystem volt age up to 170 kV . Middle: Three phase design up to 170 kV . Right: single phase design for

system volt ages above 170 kV w ith electrical one phase, but mechanically t hree columns of MOresistors, courtesy Siemens.

insulator

electrical connection,high voltage

pressure relief device

grading hood

MO column

vessel

ground plate

insulating plates

MO resistors

current path




Page 100

Actually, there are two major directions of MO resistor development: increasing the energy handling capability (in

terms of kilojoules per cubic centimeter of volume) and increasing the field strength (in terms of volts per millimeter

of height). Typically, a MO resistor for high-voltage arrester applications has a field strength of 200 V/mm at a direct

current of 1 mA. But MO resistors of 400 V/mm have successfully been developed in the mid-1990s [Ima 1984] [Shi

1997] and are commercially available, and MO resistors with 600 V/mm are reported [Fuk 2012].

The benefit of this progress can less be utilized for air insulated arresters as the high field strength causes severe

dielectric stress across the external surface and the heat after energy injection cannot be dissipated to the ambient.

But GIS arresters, where dielectric problems along the active part do not occur due to the high electric strength of

the surrounding SF6, and where the heat transfer is much better than in air, take advantage from the fact that the

overall length of the MO column can be drastically reduced when using high field MO resistors. With the new high

field MO resistors GIS arresters for application in voltage systems up to 550 kV can be built using a simple linear

stack of MO resistors instead of the meandering mechanical design.

Figure 2.32 shows the difference in dimensions of a GIS arrester containing an active part built up with MO

resistors with “normal” field strength (200 V/mm) and three column meandering design (left), and designs using

high field MO resistors of 400 V/mm (middle) 600 V/mm (left). All of these designs are for application in 550 kV

systems and have the same energy rating. Besides the simpler and space saving design the SF6 volume isreduced drastically, which is an important argument in the today’s discussion about greenhouse gasses.

Figure 2 .32 : GIS arresters for 5 50 kV systems w ith MO resistors w ith “normal” f ield strength (left ) andhigh field MO resistors of 400 V/mm (middle) and 600 V /mm (right) , courtesy Toshiba.

GIS arresters with meandering mechanical design using high field MO resistors are designed to reduce the arrester

height and increase the mechanical strength against seismic stresses even in horizontal installations, which gives

more flexibility in positioning of the arresters in an optimized GIS layout, see Figure 2.33.




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Figure 2 .33 : GIS arresters for 550 kV systems w ith MO resistor s w ith “normal” field strength of 200V/mm and 400 V/mm in a mechanical meandering design, court esy Mit subishi.

MO resistors of 200 V/mm MO resistors of 400 V/mm




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2.6 Integrated Arrester Systems

Author in charge: Roger Perkins

Surge arresters have developed considerably in the period since the first CIGRÉ Technical Brochure TB 60 in1991. There were two characteristic developments in this period. The first was the polymer-housed arrester madepossible by the use of improved fiberglass-reinforced components that could be reliably used as structural supportof the device when a polymeric housing should be used that in itself could not generate this capability whilst stillproviding safe pressure relief in the event of failure. The second major development was the increasingperformance and reliability of the metal-oxide resistor itself.

A characteristic consequence of these facts is the emergence during this period of continually more complexdevices with a greater degree of integration with other electrical transmission and distribution installations; in otherwords devices or apparatus with more than one primary function or capability. Documented examples of these thatincorporate surge arresters are:

- Supports, Post Insulators- Bushings, Cable Termination, Connectors

- Disconnectors- Transformers and Reactors- Cutouts- Fuses

There are various potential attractions of these devices; for example reduced cost or size, both increasingly moreimportant requirements. Improved reliability, improved overvoltage protection, improved environmental protectionhave all been claimed as well. However, most likely the major benefits have still to be demonstrated, since theincreasing freedom of the above developments afford the engineer more opportunity for innovative design.

However, integrated arrester systems bring special concerns with them; not the least of which is how to effectivelytest a device with multiple, sometimes interrelated functionality. These and other special considerations have beenreviewed in a recent publication of this Working Group [Per 2005].

Perhaps the most obvious of dual applications for MO surge arresters is the use as post or suspension insulators,given their high cantilever strength and dielectric withstand. In Figure 2.34 and Figure 2.35 examples are given.

Figure 2.34 : MO arresters used as post insulators in a 420 kV substation (left , example Siemens) and assuspension insulator/ line arrester in a medium voltage tr ial l ine in Norway ( right, example ABB).




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Figure 2 .35 : MO surge arresters with porcelain housing and modified grading ring integrated in a centerbreak disconnector. System volt age 420 kV , example Siemens.

Arresters have a limited spatial protection range due to travelling wave effects. A lightning impulse voltage may

reach twice the value of the related arrester’s lightning impulse protection level at the terminals of the device to be

protected, depending only on the steepness of the incoming overvoltage and on the distance between arrester and

the device. This “protective zone” or “separation distance” is typically in the range of several ten meters in high

voltage applications down to only a few meters in distribution systems. Therefore, the integration of MO surge

arresters directly into other equipment of the substation improves naturally the protection of the substation. This is

of course an additional very important benefit besides the space saving for the substation.




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3. Energy handling capability of MO surge arresters

Authors in charge: Volker Hinrichsen and Max Reinhard2

3.1 SummeryThis part of the Brochure covers energy handling capability of MO resistors and arresters. It starts with an

introduction and a short subsumption of the different aspects of energy handling, basically divided into "thermal"

and "impulse" energy stress. It then reviews the state of knowledge by evaluating some of the most important

published literature on this subject. Cigré WG A3.17 has initiated an experimental research program on energy

handling, which is being performed at Technische Universität Darmstadt. It is the most comprehensive investigation

on this subject performed so far. Though this project is still going on, some important results can just be

summarized. Several thousand MO resistors from different manufacturers worldwide were tested. The test

specimens were of approximately 60 mm diameter and 45 mm height, as typically applied in HV arresters, and of

approximately 40 mm diameter and 45 m height, as used in high duty distribution arresters. Energy was injected by

long duration and double exponential current impulses, among them also the new "lightning discharge" impulse

(sine half-wave of approximately 230 µs base time, resulting in an impulse current 90/200 µs) that was recently

introduced to the standard IEC 60099-4 (Annex N). Furthermore, alternating current stress was imposed as well, in

order to investigate the impact of this kind of stress, but also to check if this test approach can be favorably applied

in the future. A very important aspect was the introduction of a "complex failure criterion", which means that the MO

resistors were not only stressed up to visible mechanical damage, but that also deterioration of the electrical

characteristics was considered. The investigations have basically confirmed the typical dependence of energy

handling capability from current density, as published before. But there are exceptions at extremely high current

densities, where in many cases the coating of the resistors would fail. A general increase in energy handling

capability, expressed in terms of 50 % failure energy, by 20 % in average and up to 70 % in some cases, compared

with values published in the late 1990s, can be observed. The "complex failure criterion", however, leads to more

pessimistic statements, as it turns out that the MO resistors for distribution applications typically (but with

exceptions) show remarkable degradation of their electrical characteristics before they fail mechanically. Finally,

this chapter ends with a critical review of the existing arrester standards with regard to energy handling definitions

and test procedures. Some lacks are identified and suggestions for improvement are given. Based on the actual

and updated knowledge of energy handling capability it should be possible to improve the standards in their nextrevisions accordingly in order to better fulfill the requirements and expectations from manufacturers and users of

MO surge arresters.

3.2 Introduction

Since simple spark gaps for overvoltage protection were replaced by surge arresters, the arresters' energy

handling capability has become an important issue. While in series gapped SiC-arresters, especially in the EHV

systems, the energy during charge transfer to ground is shared among the arcs burning in the gaps and the series

connected SiC-resistors, this energy has to be dissipated exclusively by the MO-resistors in case of gapless MO-

arresters. On one hand, this results in high requirements on the non-linear resistors, which have to act as nearly

perfect "insulators" under normal operating conditions and as high-performance overvoltage limiting "energy sinks"

under overvoltage stress. On the other hand, one should expect that energy handling definitions, specifications and

test procedures would have become simpler, as only one element – the MO-resistor – has to be considered. When

looking to the published literature and to the actual surge arrester standards, one will f ind that this is obviously notthe case. There is stil l a certain lack of general knowledge and theoretical understanding about some energy

handling capability aspects, for instance the impact of the way the energy is injected or degradation effects caused

by multiple or repeated energy stress. Of course, the theoretical background has been improved, simulation tools

have been developed which allow many effects to be modeled and simulated, and finally thirty years of experience

with MO arrester application have given a high degree of confidence in their reliable performance. However, the

fact should not be underestimated that MO resistor manufacturing requires a rather complex technology, and

therefore the final products' performance will always strongly depend on production technology and quality. With

respect to the MO arrester standards: they had to be developed in a time when MO technology was quite new and

still emerging, and it took about ten years after introducing the first MO arresters to the systems that first related

2 With assistance of Maximilian Tuczek, TU Darmstadt




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standards were published. Till now, the standards do reflect this situation. In terms of energy handling issues, onestill recognizes historical approaches from the gapped SiC arrester era. It is, therefore, the time to think about newdefinitions of energy handling capability and about revised, appropriate test procedures, based on actualknowledge and most recent findings. Besides others, this was the task given to Cigré WG A3.17: to check theactual literature and standards about energy handling issues, to contribute to some of the open questions by acomprehensive practical research program, and to work on proposals for energy handling issues in future revisionsof the international arrester standards. The basic results of this work are reported below.

3.3 The different aspects of “energy handling capability”

Energy handling capability of MO arresters has many different aspects, which are only partly or not at all reflectedin the actual standards. At least, though this list may not be complete, they have to be divided into

- "thermal" energy handling capability,- "impulse" energy handling capability,

o "single" impulse stress, withstand values (deterministic approach), values related to a certain failure probability (statistical approach),

o "multiple" impulse stress, i.e. impulses in time intervals too short to obtain an approximatelyuniform temperature distribution in the MO resistors,

o "repeated" impulse stress, where the time interval between impulses is sufficiently long to obtaincooling of the MO resistors close to their initial temperature (this includes durability anddegradation aspects).

3.3.1 THERMAL ENERGY HANDLING CAPABILITYThermal energy handling capability can only be considered for complete arresters, as besides the MO materialproperties it is mainly affected by the heat dissipation capability of the overall arrester design. The situation isschematically depicted in Figure 3.1.

Figure 3 .1: About thermal energy handling capabilit y of an MO arrester

An arrester's heat dissipation capability (heat flow; measured in Watts) is determined by thermal conduction,convection and radiation. In the interesting temperature range (operating temperature below 250 °C) it increasesnon-linearly but moderately with the temperature difference to ambiance. Electrical power losses under normaloperating conditions are usually very small, in the range of tens of milliwatts per kilovolt of rated voltage fordistribution arresters up to several hundreds of milliwatts per kilovolt for line discharge class five (LD 5) arresters.However, due to their temperature dependence, the power losses are much higher at higher temperatures, e.g. by




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a factor of ten to twenty at 150 °C compared with 20 °C. This power loss characteristic is specific to a particular MO

material and make.

Under continuous operating conditions, an arrester will adopt an operating temperature slightly above ambient

temperature, and the generated heat can easily be dissipated to the ambiance: the arrester adjusts itself to a stableoperating point (left intersection of the two curves in Figure 3.1). However, once a high amount of energy is injected

into the arrester under overvoltage conditions, the arrester temperature will be increased in form of a step function,

with typical values of temperature increase under nominal energy stress up to 100 K or even more. The operating

point will instantaneously jump to the right on the electrical power loss curve. As long as it remains left of the

second intersection point of the two curves, the generated heat can still be dissipated to the ambiance, and the

arrester will cool back to its normal operating temperature within five of its thermal time constants. But if the right

intersection point – the limit of thermal stability – is reached or even exceeded, the arrester will generate more heat

than can be dissipated and electrical power losses will further increase and finally destroy the MO material by

excessive heat (puncture at several hundred Degrees Celsius).

It is evident that, on one hand, the thermal stability limit depends on the overall arrester design. Arresters with MO

resistors directly covered by a polymeric housing, for instance, will have a thermal stability limit at higher

temperatures than conventional porcelain housed arresters, since they can better transfer heat from the MOresistors to ambiance. On the other hand, also the MO material properties (electrical power losses and their

temperature dependence) have an effect, because the more pronounced the increase of power losses with

temperature is, the more will the right intersection point of the two curves be shifted to the left, i.e. to lower

temperatures. As well, the curve of electrical power losses versus temperature is affected by possible impulse

degradation, i.e. it will be shifted upwards [Hei 2001], which again changes the limit of thermal stability to lower

temperatures.

However, definition and verification of the thermal energy handling capability is a comparatively easy task. Injected

energy per volume and temperature increase are simply linked by the heat capacitance, which has a non-linear

dependence of temperature, and can, acc. to [Lat 1983], be calculated as

J J2,59 0,0044 ²

cm³ K cm³ K²

W

V

(equation 3.1)

where W is the contained energy in J, V is the MO volume in cm³ and is the MO temperature in °C. This

dependence is shown in Figure 3.2.

Figure 3 .2: MO resistor energy per volume vs. temperature, acc. to [Lat 1983]

In order to verify thermal energy handling capability, energy may thus be injected into the arrester by any suited

method that will rise its temperature to a value related to the specified energy, because the only purpose of this

0

100

200

300

400

500

600

700

800

900

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190 200 210 220 230

Temperatur in °C

E n e r g i e

/ V o l u m e n i n J / c m ³

E n e r g y p

e r v o l u m e i n J / c m ³

Temperature in °C




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verification is to demonstrate that the arrester is able to cool back afterwards. Of course, the operating conditions(applied power-frequency voltage) must be specified, and possible electrical aging of the MO material, e.g. bycurrent impulse stress, must be considered by appropriate conditioning procedures.

3.3.2 IMPULSE ENERGY HANDLING CAPABILITYAt first glance, impulse energy handling capability may easily be defined as well: it is just the energy, which isinjected into the arrester by one single impulse, and if a limit value is exceeded, one or several MO resistors willmechanically fail as exemplarily shown in Figure 3.3, finally leading to an overall arrester failure.

Figure 3.3 : Examples of MO resistors, mechanicall y fai led by single impulse energy over load. Left: fa iledby thermo-mechanical cracking; Right: failed by flashover of t he coat ing.

When looking deeper into the details, however, the matter is more complex. One issue is the definition of a"failure". Not in all cases will the MO resistors fail so obviously as to be seen in Figure 3.3. There may be onlysome small punctures in or at the edge of the metallization, but furthermore the MO resistor may look intact. Inother cases, no damage at all might be seen by a visual examination, but the MO resistor is pre-damaged and willnot pass any further energy input. Or its electrical characteristic may be dramatically changed such that if thishappened in a complete arrester the arrester would become thermally instable.

At this point it shall be noted that in general use of any impulse energy handling capability higher than the "thermal"energy limit can be made only if the arrester is not applied to an operating voltage close to its continuous operatingvoltage, since otherwise the arrester would suffer a thermal runaway even if its MO resistors were able to handlethe excessive impulse energy input.

Another point that has to be addressed is if "withstand" capability shall be specified in a deterministic way –meaning that no single failure is allowed when the MO resistor is stressed by its withstand energy – or if a statisticalapproach is more appropriate, in the same way as for the dielectric strength definition of external insulation (inwhich case "withstand" voltage stands for a "10 % flashover probability" voltage). Verification of a withstand energyis a difficult task anyway, as even for the statistical approach acceptable failure rates of individual MO resistors in a

complete arrester at "rated" energy handling capability are in the range of only 0,1 % or less. This shall bedemonstrated by the following example. An arrester for a 420 kV system is made up from approximately n = 65 MOresistors. If each resistor has a failure probability of 0,1 % (p = 0,001) at its "rated" energy, the full arrester, at thesame rating, will have a failure probability of P = 1 – (1 – p)n = 0,063 or roughly 6 %, respectively. Higher failureprobabilities for the full arrester are hardly acceptable! Vice versa, if the full arrester shall have a failure probabilityof only 1 % (P = 0,01) at its "rated" energy the individual MO resistors in this case must have a failure probabilityp = 1 – (1 – P)1/n = 0,155·10-3 or approximately 0,015 % only.

But so far, there is no effective test procedure to reliably verify failure probabilities of only 0,1 % or even less.

The observation that the actual failure rate of high-voltage arresters in service is obviously close to zero can beexplained by one or more of the following reasons. In general, the energy stress in real service may be far below




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the actual impulse energy handling limits that are verified during type tests. Energy of the "withstand" level may be

injected only few times during the arrester's life time. The actual failure probability at "rated" energy may be much

lower than can be verified by the type tests. The type tests as actually specified therefore obviously ensure good

operational performance in service, but they may cause overdesign of arresters, and they will not give any

information about the real limits of energy handling capability.

Furthermore, the question comes up how helpful to the user the information is that an arrester passes one single

energy input, but not a second one a certain time later on. This might be interesting in only some few special

cases, but in general, the expectation will be that an arrester can be stressed by its "withstand" energy several

times during its total service life. But what is the meaning of "several": three times, ten times, eighteen times (the

actual number of energy stresses in the long duration current impulse withstand test according to [IEC 2009]), or

even, e.g., one hundred times? It is assumed to date that energy handling capability decreases with the number of

stresses, but in an actually unknown dependence (this is being investigated in a follow up research program of

Cigré WG A3.25 and will be published later). It is also questionable how meaningful a type test on three MO

resistors by eighteen energy injections each is (this is again the long duration current impulse withstand test

according to [IEC 2009]), as the impulse withstand capability is not only a material issue but depends at least to the

same degree on production quality.

Finally, multiple stress (i.e. energy injections in time intervals of only a few milliseconds, as it may happen by

multiple lightning strikes) has turned out to result in interesting effects, which are not only related to the MO

material properties but to the overall system of the MO resistor and the arrester [Dar 1998].

Thus, definition of impulse energy handling capability is by far not trivial, and the same is true for appropriate test

procedures. Users have become familiar, for instance, with the line discharge classes of [IEC 2009]. These are

very helpful and easy to apply in standard applications. But more and more users have very special system

configurations, and in many cases, by means of system analyses they are in a position to give detailed information

on the energy duties from the system. However, none of the to date’s arrester standards gives satisfying answers

and information on the various different aspects of energy handling, and it must be the objective of any research on

MO arrester energy handling capability to give better guidance on this matter in the near future.

3.4 State of knowledge about energy handling of MO arresters

In the following, a short overview about the most relevant literature on energy handling capability of MO arresters

and resistors, respectively, will be given, and the chapter will be finalized with a report about recent findings of an

energy handling research program that was initiated and scientifically accompanied by Cigré WG A3.17.

3.4.1 A BRIEF REVIEW OF THE RELEVANT LITERATUREIn a comparison between SiC and ZnO

3 arresters, Sakshaug [Sak 1989] concluded that ZnO resistors in general

have a higher energy handling capability than SiC resistors. He gave a value of (170…200) J/cm³ for the thermal

capability and mentioned at the same time that under alternating current stress up to mechanical failure values of

(450…700) J/cm³ were observed.

Eda [Eda 1984] performed one of the very first methodical experimental investigations about energy handling

capability of MO resistors. For this purpose, he produced resistors of (10…110) mm in diameter and of (1…20) mmheight, i.e. he worked with non-commercial test specimens. He reported about two different failure modes under

impulse energy stress, that is to say cracking and puncture. Flashover as a possible failure mechanism was not

observed on these specimens. Most of his published results are related to small discs of only 1,3 mm height. He

found an impulse energy limit value of 750 J/cm³ for discs of 1,1 cm diameter. For discs of 2,76 cm diameter, he

found an energy limit of 520 J/cm³ when stressed by a 2 ms impulse and of 615 J/cm³ for a 20 µs impulse. These

findings indicate two important tendencies: energy handling capability decreases with increasing diameter and

volume, respectively, which can be explained by a worse homogeneity of the ceramic material with larger diameter

and volume, and it increases with shorter impulse durations or, in other words, with higher current densities. Any

3 All MO resistors are basically made from ZnO. In the beginning, "ZnO" arrester and resistor was a common terminology. Todate, one usually speaks exclusively of "MO" arresters and resistors in this context.




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further interpretation of the results and comparison with modern MO resistors seems problematic as Eda used very

thin discs where the contact system assumedly had strong influence on the experimental results (e.g. due to a

comparatively high amount of heat transfer). Furthermore, it is not for sure that self-made resistors of 114 mm

diameter and 10 mm height can really be compared with to dates commercial products of the same dimensions.

Studies of Martinez and Zanetta jr. [Mar 1996] basically confirmed Eda's findings.

One of the best known systematic investigations was performed by Ringler et al. [Rin 1997]. Ringler's group

investigated 350 commercially available MO resistors from three different manufacturers. Their diameters were in

the range of (62…64) mm, and their height was (23…24) mm. The rated voltage was approximately 3 kV. Energy

stress tests until mechanical failure were performed with very long duration current impulses and with 60 Hz

alternating current on a batch of (25...50) specimens each at current amplitudes of 0,84 A up to 35 kA,

corresponding to current densities of (0,03…1130) A/cm².

Failure energies varied from approximately 460 J/cm³ at a current density of 0,03 A/cm² up to nearly 1700 J/cm³ at

a current density of 1130 A/cm². There was a distinct increase of energy handling capability with increasing current

density and thus decreasing stress time duration. These findings basically correlate with those reported by Eda in

1984.

At this point it must be recalled that failure energies (e.g. related to 50 % failure probability) are of course far

beyond those energies that are the basis for an arrester specification. The latter are in the range of only 200 J/cm³,

for reasons that were explained above. One finding of Ringler et al. by statistical evaluation was that varistors with

a 50 % failure probability of more than 400 J/cm³ may have a failure probability of still 1 % at 200 J/cm³.

When the average current amplitude during energy stress is plotted as a function of the average time to failure in a

double logarithmic scale, this gives a linear dependence in form of a sloping straight line over five orders of

magnitude. This relationship (log I = const. – log t) was published by Ringler et al. for the very first time and since

then belongs to the basic knowledge about energy handling capability of MO resistors. The failure charge for

Ringler's investigations was (11…17) As, corresponding to failure charge densities of (0,35…0,55) As/cm².

Ringler et al. reported furthermore that under alternating current stress the resistors would mainly fail by puncture

close to the edges. Under impulse current stress up to 35 kA they reported about little holes on the surface of themetallization, including the edges. This may have been caused by the quality of the metallization edge. It is well

known today that energy handling capability can be increased by optimization of the metallization. In most cases

the investigated resistors failed by puncture or by some kind of "tracking" along the outer coating. In only few cases

mechanical cracking was reported to be the failure mechanism.

Another important observation of Ringler et al. was that none of the considered distribution functions – “Normal”,

“Weibull” and “Gumbel” – could be given a preference. Failure probabilities of MO resistors cannot consistently be

described by any of these distributions. All these distribution functions covered the observations with deviations in

different details, but the final outcome (e.g. in terms of 50 % failure energy) was comparable for all of them.

Therefore, it was suggested to apply the Normal distribution as a generally known function and acceptable

approximation when results about failure energies shall be compared.

Boggs et al. [Bog 2000] investigated energy handling capability depending on how the metallization is

implemented. It was the objective of these investigations to find out if the metallization should preferably be applied

exactly up to the edge of the MO resistor or if a certain clearance to the edge should be kept, and how this would

affect energy handling capability. Performed simulations indicated a distinct temperature increase directly at the

edge of the metallization. It was finally proposed to keep a distance of (0,3…0,6) mm to the edge of the MO resistor

in order to get an optimized energy handling capability. For smaller distances, it was stated that dielectric strength

would be affected (risk of external flashovers). This conclusion has not generally been accepted. Many MO resistor

manufacturers have successfully implemented a metallization exactly up to the edges. This has to be seen not only

from the point of optimized energy handling capability; it is also a concern of manufacturing technology, since

implementation of a metallization up to the edge is difficult. However, Boggs' investigations have shown that the

way of metallization does have an influence on energy handling capability. If the metallization ends too far away

from the MO resistor's edge local current densities may reach values that increase the risk of puncture of the MO




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material. Also the quality (smoothness) of the edge on a microscopic scale has effect on the overall energy

handling capability.

Bartkowiak et al. [Bar 1996a] worked intensely on the simulation of energy stress and energy handling limits. Two

different kinds of MO resistors were considered: one with a diameter of 32 mm and a height of 45 mm,representative for medium-voltage (distribution) arresters, and the other with a diameter of 63 mm and a height of

23 mm, typically applied in high-voltage (station) arresters. Apart from the different typical applications, these MO

resistors thus differ distinctly in their aspect ratios, i.e. their ratios of height over diameter, which was considered in

particular and which has influence on the failure mode as could be demonstrated by Bartkowiak at al.

With regard to failure mechanisms, it was distinguished between puncture, cracking and thermal instability (where it

has to be noted that thermal instability is not an impulse energy failure mechanism as such; it just starts a process

that finally leads to puncture of the MO resistor by overheating). Only heat dissipation in radial direction was

considered, as it is the case in a real arrester (apart from its ends). Material inhomogeneity was modeled by a

straight small channel in the center of the MO resistor, having a varistor voltage reduced by 5 % compared with the

overall characteristic. Energy injection was simulated by a direct current, heating the material. A further boundary

condition was the possibility of free movement of the material in any direction. It must be critically noted that in real

arresters this is not always the case. In many designs the MO columns are mechanically clamped by extremeforces in the range of 100 kN or even more. Therefore, in a real arrester any thermal expansion of the MO material

may be limited or even totally suppressed, which may result in different distributions of pressure and tensile forces

in the material. This must be kept in mind when interpreting the simulation results.

Some more details about the simulation are explained in [Bar 1999]. Following preconditions were assumed to

result in failure of the MO resistor:

- a tensile force larger than 480·106 N/m² in axial direction will cause cracking,

- a tensile force larger than 140·106 N/m² in radial direction will cause cracking,

- an average overall temperature above 190 °C will cause thermal instability,

- a temperature of the center channel above 800 °C will cause puncture.

The performed simulations showed that the minima of energy handling capability depend on the failure mode. For

the distribution MO resistors, the minimum was found to be 310 J/cm³ at a current density of approximately1 A/cm². The related failure mode is "cracking". Energy handling capability then increases with current density. For

"puncture", the minimum failure energy is found at a current density of 1 A/cm² as well, but at a higher level of

about 600 J/cm³. For the failure mode "thermal instability" the failure energy is about 980 J/cm³ at extremely low

current densities of less than 0,0001 A/cm², and it then reaches a nearly constant value of 580 J/cm³ over the full

range of current density from 0,001 A/cm² up to 50 kA/cm². For distribution MO resistors the minimum of energy

handling capability thus will be found at current densities of about 1 A/cm², for puncture as well as for cracking.

This is not the case for the high-voltage resistors. For the failure mode "puncture" the minimum was calculated to

be 420 J/cm³ at a current density of about 0,1 A/cm², while for "cracking" the minimum energy is slightly higher –

500 J/cm³ – but at much higher current densities of about 20 A/cm². These findings, however, could not all be

verified by recent experimental investigations, as will be reported later in section 3.4.2.

Bartkowiak also modeled the behavior of an MO resistor on the basis of a two-dimensional, randomly generated

Voronoi network4 [Bar 1996b] [Bar 1006c]. The network is made up from three components: "good" grainboundaries with extremely non-linear voltage-current-characteristic, "bad" grain boundaries with poor non-linearity

and "ohmic" grain boundaries.

An interesting recent publication is from China [He 2007]. He and Hu report about tests on two different makes of

commercial varistors: type A with a height of 10 mm and 32 mm in diameter; type B with a height of 10 mm but a

diameter of 52 mm. For energy tests with long duration current impulses of 2 ms and 8 ms time duration they quote

cracking and puncture as dominating failure mechanisms. For the MO resistors of 32 mm diameter they give

surprisingly low failure energy values of only (216…575) J/cm³. Such low values for these comparatively small

4 Voronoi polygons acc. to the Russian mathematician Georgi Feodosjewitsch Woronoi (1868-1908). Voronoi polygons areapplied in material sciences to simulate a random crystal constellation in polycrystalline materials.




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resistor elements are interesting, as they show that though MO resistor technology is considered to be mature, one

has always to be aware that products will come to the market that do not fulfill the general expectations on energy

handling capability. It is thus once more important to have clear definitions and related test procedures in future

arrester standards that allow an easy and simple evaluation of energy handling capability of a MO resistor.

Another publication is surprising as well [Ver 1992]. Energy handling tests on commercially available MO resistors

are reported there. The resistors were of 22 mm height and 53 mm in diameter, they had a continuous operating

voltage of 2,5 kV and a nominal discharge current of 10 kA. Unfortunately, no absolute values of energy handling

capability is given, the information is limited to general tendencies which, however, are remarkable. While all

reports published so far indicate an increase of energy handling capability with current density, the contrary is the

case here: energy handling capability decreases with increasing current density. Compared with all other findings

and to date's knowledge, this has to be judged as an error, for which reason ever.

Darveniza et al. [Dar 1998] investigated the performance of distribution arresters under multiple impulse stress

which is motivated by the nature of lightning flashes5. They tested 21 porcelain housed arresters from six different

manufacturers. Additionally, they investigated MO resistors with different coating systems of one manufacturer, and

further MO resistors of other manufacturers in different surrounding media. Impulse currents 8/20 µs at amplitudes

from 5 kA to 11 kA were applied, as well as 4/10 µs impulses from 40 kA to 100 kA. The 8/20 µs impulses weremultiple impulses at time intervals of (15…150) ms. No information of the energy handling capability is published,

but an interesting observation is reported. The MO resistor stacks flashed over. Tests with surrounding gases

modified for different dielectric strengths (air at normal density, air at lowered pressures, SF6) resulted in the same

behavior. Thus, obviously, the flashover under this kind of multiple impulse is initiated not outside the resistor but

directly underneath or within the coating, a phenomenon that has also been observed for the new lightning current

impulse 90/200 µs (see later in section 3.4.2).

Dengler [Den 1998] intensely investigated electrical degradation of MO resistors

(in terms of watt loss and leakage current increase) under an extremely high number (up to 400 impulses) of

lightning impulse current stresses at amplitudes around nominal discharge current and finally derived online

monitoring procedures from his findings. He investigated two different kinds of MO resistors (material A: height

46 mm, diameter 38 mm, nominal discharge current 5 kA; material B: height 40 mm, diameter 74 mm, nominal

discharge current 20 kA). Besides the impact of current amplitude and front steepness, he also investigatedrecovery effects at different temperatures or time intervals between the individual impulses. As cause of electrical

degradation, he suggests migration of negative oxygen ions towards the inner region of the ZnO grains. This

changes the oxygen ion concentration at the grain boundaries and has effect on the barrier voltage. Recovery may

take place under certain conditions by negative oxygen ions travelling back to the boundaries.

Klein [Kle 2004] investigated changes of material properties by impulse currents, expressed by changes of leakage

current, reference voltage and power loss. He also looked very closely to fine cracks on the resistor surface. As a

good approach for generalization, he introduced a common reference current density of Jref = 0,12 mA/cm²

(according to the standards, manufacturers are free to specify their reference current in any suited way).

Unfortunately, in case of asymmetries in polarity (which is typical after unipolar impulse current stress), he used the

average of positive and negative current amplitudes for the determination of the reference voltage, which makes

comparison of similar results diff icult that are found according to the definitions in the standards.

The contributions cited so far dealt with the particular aspects of impulse energy handling capability. Only few

publications can be found on thermal stability issues. First studies were performed and published by [Lat 1983] [Lat

1985], where the thermal behavior and thermal stability limit of individual MO resistors and MO distribution

arresters was investigated and simulated by means of a transient network analysis. St.-Jean et al. [StJ 1990]

reported about a similar approach for high-voltage arresters up to 120 kV rated voltage, were the MO temperature

is approximately evenly distributed along the arrester axis as well. The problem becomes more complex, however,

if high-voltage arresters of several meters in height are considered, since they represent structures of distributed

parameters with a spatial distribution of all electrical and thermal quantities. In [Hin 1987], [Hin 1989] and [Hin

1990] approaches and performed simulations on such arresters are reported. They were also based on a

5 In 55% of all cases lightning flashes are composed of two or more individual strikes [And 1980].




Page 112

comparatively coarsely structured distributed parameter network. Though further progress has been made in this

field – for instance [Hin 2008] reports about successful coupling of a three-dimensional non-linear electro-

quasistatic and a thermal field problem for UHV arresters – till now really satisfying approaches to simulate thermal

stability limits under special consideration of the axial temperature distribution in HV-, EHV- and UHV arresters

have not been published.

3.4.2 RESULTS OF AN EXPERIMENTAL INVESTIGATION INITIATED BY CIGRÉ WG A3.17When Cigré WG A3.17 started their work on surge arresters, the investigations on energy handling capability of

Ringler et al. [Rin 1997] dated back more than ten years. With this background, the following goals for the first part

of a research program on energy handling capability were formulated:

- Confirm and update the test results of [Rin 1997];

- extend the investigation to MO varistors of dimensions that are commonly applied today in station and

distribution arresters (different diameters, larger height);

- extend the investigation to samples of several manufacturers worldwide;

- make use of standard impulse currents as usually applied for arrester testing;

- also investigate the impact of 4/10 µs high current impulses and of the new lightning impulse discharge

current (approx. 90/200 µs) according to IEC 60099-4, Annex N [IEC 2009];- extend the failure criterion to general significant changes of the material properties, instead of mechanical

failure only.

In a second part of the program (in the work frame of Cigré WG A3.25) issues such as durability (impact of number

of impulses) and the problem of single versus multiple impulse stress will be addressed as well as statements on

failure risk (failure probability versus absorbed energy), and finally the impact of uneven axial temperature

distribution in high-voltage arresters on the thermal stability limits. This will allow deriving rated energies for specific

applications of surge arresters. Finally, better energy definitions and simpler test procedures for future revisions of

the arrester standards shall be derived and suggested.

Results were published so far in [Rei 2008a] [Rei 2008b] [Tuc 2009] [Hin 2009] and are – including some new and

most recent results – summarized below.

3.4.2.1 Test specimens and test currentsCommercially available MO resistors from eight well established American, European and Japanese manufacturers

(named as S, T, U, V, W, X, Y, Z) were tested. Two basically different sizes of MO resistors were considered. The

first size, denominated as "Size 1", is typically applied in 10 kA station class arresters of line discharge class 3.

Their height is 40 mm to 45 mm (except for one make of only 26 mm) and their diameter around 60 mm. The

second size – "Size 2" – is typically applied in 10 kA distribution class arresters. Their height varies from roughly 30

mm to roughly 40 mm, their diameter is around 40 mm. Eight different types of current stress were applied for

testing: alternating current (50 Hz) at three levels of current amplitudes î 10 A, 100 A, 300 A, long-duration

current impulses of about 1 ms, 2 ms and 4 ms time duration, high current impulse 4/10 µs and lightning impulse

discharge current 90/200 µs (time parameters ± 10%). For each test series 40...50 samples were tested. For some

tests the sample number was even increased up to 80. From the total number of MO resistors that have been

announced and delivered, respectively, for the test program, more than 3000 pieces have been tested. This has

thus been the most extensive investigation on MO energy handling capability so far.

Different from other test programs standard impulse current shapes were used, except for the alternating current

stress, which is not specified in any arrester standard. Reason for the latter is that when testing up to mechanical

failure each alternating current stress will contribute directly to a failure energy distribution, whereas impulse tests

(each impulse results in "passed" or "not passed" for the test sample) requires higher statistical efforts in

determining a mean failure energy (i.e. an energy that would lead to 50 % failure probability), and it is more difficult

to give statements about very low failure probabilities. It was one of the objectives of this test program to show if

alternating current stress can be favorably applied for this purpose. Another reason was that applications do exist

where alternating current stress is imposed (e.g. overvoltage protection of series capacitor banks) and it is

interesting to know if energy handling capability is affected in any direction by this kind of energy input.




Page 113

It was for the first time, as well, that a systematical investigation with the new 90/200 µs impulse ("lightning impulse

discharge" acc. to [IEC 2009]) was performed.

3.4.2.2 Test setup A pneumatic test fixture allowing a rapid change of the test specimens was especially developed. It consists of a

pneumatic actuator, a fixed and a moving contact electrode consisting of copper foil and two heat insulating blocks.

The air pressure of the pneumatic actuator was adjusted for each type of varistor to ensure a pressure of p = 3,0

N/mm² on the electrodes. This value results in a comparatively high contact force and was chosen to prevent any

bouncing of the electrodes during the tests with high-current impulses (4/10 µs). The heat insulating blocks prevent

axial heat flow during the tests. They are made from fiber silicate with high compression strength and low thermal

conductivity. To ensure identical contact conditions for each test and to avoid flashover problems as a result of

damaged electrodes, two new aluminum electrodes (discs of 5 mm thickness with rounded edges) were used for

each test. The diameter of the electrode discs was adopted for each type of varistors such that the electrode

diameter was (1...2) mm was smaller than the diameter of the varistor, in order to avoid any dielectric problems.

3.4.2.3 Test procedure A further difference to former investigations on energy handling capability, where test were performed up to

mechanical failure of the samples, was the introduction of a "complex" failure criterion. This was to take into

account the fact that not only visible mechanical damage but also non-visible pre-damage or deterioration of the

electrical characteristic would constitute an arrester failure in a real system (because the arrester would not pass

any further energy input or degradation effects would cause thermal instability).

At the beginning of a test series an initial measurement procedure was performed to acquire the electrical

characteristics of the MO resistors (see Figure 3.4). During these initial measurements a "characteristic" voltage

Uch (per definition similar to the reference voltage) in the leakage current range of the U-I-characteristics was

measured at a current Ich corresponding to a peak current density of 0,12 mA/cm², five seconds after voltage

application. Then the residual voltage was measured at nominal discharge current of 10 kA, 8/20 µs.

After the initial measurement, the energy stress test was carried out. Thereafter, the MO resistors were visually

inspected to determine mechanical failure such as cracking, puncture or flashover. If there was no obvious

mechanical failure the MO resistors were again tested for their electrical characteristics after cooling to ambienttemperature. These measurements were performed exactly in the same way as the initial measurement, with one

exception: one additional impulse current Imd for check of mechanical pre-damage, shape 8/20 µs at an increased

discharge current corresponding to a current density of 1,5 kA/cm², was applied after the residual voltage test with

10 kA.

The following set of failure criteria was specified for the impulse current tests (the alternating current tests were

performed until mechanical failure, see below, and thus did not require these special considerations). First criterion

was mechanical integrity, determined by visual inspection. If the MO resistor passed mechanically the exit

measurements were performed. If the characteristic voltage Uch had changed by more than 5 % the MO resistor

was considered as failed. This criterion was introduced since such change in the voltage would clearly constitute a

change of the material characteristics. If a mechanical failure such as cracking, puncture or flashover occurred

during the exit impulse current tests (two impulses 8/20 µs, first at In = 10 kA and then at Imd corresponding to 1,5

kA/cm²) the MO resistor was considered as failed as well. Both of these additional criteria had to be introducedsince in many cases apparently sound but actually severely pre-damaged MO resistors could be identified only this

way. As an example, eventually the metallization of the MO resistors was punctured at the edges or within the

electrode surface. Only the Uch and the Imd criteria allowed to decide if these MO resistors still performed

satisfactory or not. Finally, changes of the residual voltage were evaluated. If the residual voltage had changed by

more than 5 % the MO resistor had failed.

A drawback of this extended and very sensitive evaluation procedure is that it is rather time consuming.




Page 114

Figure 3 .4: Flowchart of the test and evaluat ion procedure

Impulse test

(energy injection)

Visual inspection:mechanically failed?

OK defect

Initial measurements

Uch,1 at Jch = 0,12 mA/cm² (after 5 s)Ures,1 at I = In

ch,1 ch,2 ch,195% 105% ?U U U

Measurement of characteristic voltage

Uch,2 at Jch = 0,12 mA/cm² (after 5 s)

Measurements at lightning current impulse

Ures,2 at I = InImd at J = 1,5 kA/cm²

E x i t m e a s u r e m e n t s

yes

no

no

yes

yes

res,1 res,2 res,195% 105% ?U U U no

yes

no


Impulse test

(energy injection)


OK defect

Initial measurements

Uch,1 at Jch = 0,12 mA/cm² (after 5 s)Ures,1 at I = In

ch,1 ch,2 ch,195% 105% ?U U U

Measurement of characteristic voltage

Uch,2 at Jch = 0,12 mA/cm² (after 5 s)

Measurements at lightning current impulse

Ures,2 at I = InImd at J = 1,5 kA/cm²

E x i t m e a s u r e m e n t s

yes

no

no

yes

yes

res,1 res,2 res,195% 105% ?U U U no

yes

no





Page 115

Some of the energy test series for station class MO resistors were performed at alternating current stress. In this

case, due to restrictions of the test setup, voltage was applied until mechanical failure of the MO resistor.

Therefore, no initial or exit measurements had to be performed for this test series. After failure of the MO resistor

the short circuit current of the transformer was interrupted by a vacuum circuit breaker. Since the transformer short

current was flowing for a time of up to 40 ms, it was not in all cases possible to exactly determine the failure mode

of the MO resistor.

3.4.2.4 Test results and discussion

The station class MO resistors "Size 1" were tested with the following impulses: lightning impulse discharge current

90/200 µs, long duration current impulse (1, 2, 4 ms) and alternating current of 10 A, 100 A and 300 A (peak).

The impulse tests were carried out at energies related to approximately 50 % failure probability. The alternating

current tests were carried out until MO resistor breakdown.

The distribution class MO resistors "Size 2" were tested with high current impulses 4/10 µs, lightning impulse

discharge currents 90/200 µs and long duration current impulses (1, 2, 4 ms). As for the "Size 1" samples, the

impulse tests were carried out at energies that would lead to approximately 50 % failure probability.

Figure 3.5 shows, for the station class MO resistors "Size 1" of six different manufacturers, the mean failure energy

of the failed samples as a function of current density amplitudes. For comparison, the failure energies published in

[Rin 1997] are also included. It has to be noted that the alternating current and the impulse current measurements

cannot directly be compared, as different failure criteria were applied for both of these test series. If the alternating

current tests had been interrupted before mechanical failure (which was not possible for practical reasons) and the

"complex" criterion of the impulse tests had been applied, the failure energies would have been lower. Vice versa,

the mean impulse current failure energies would be higher if the tests had been performed up to mechanical failure.

This is shown for one make of MO resistors in Figure 3.6. In Figure 3.5, a direct comparison (in terms of absolute

values) with the results of [Rin 1997] is, therefore, only possible for the alternating current tests.

The following can be concluded from Figure 3.5:

1. Mean failure energies approximately vary from 400 J/cm³ up to 1200 J/cm³, for very fast impulses even up to

1700 J/cm³.

2. Energy handling capability increases with current (density) amplitude basically in the same way as reported in

[Rin 1997]. From the alternating current tests, where the same failure criteria were applied, one can see that

the mean failure energy is increased by up to 70 % (in average by at least 20 %) compared with the

investigations of [Rin 1997]. This probably reflects the continuous improvements industry has made over the

last decade in processing, material formulation and MO resistor design. And evidently, cost pressure on the

market has not resulted in lower qualities. However, the wide spread by a factor of 1,7 among the different

makes is remarkable.

3. Application of the "complex" failure criterion (in this case applied for the impulse current tests) results in

approximately 50 % lower mean failure energy values than if tests were carried out up to visible, mechanical

failure. For further investigations it is therefore important to discuss if the "complex" failure criterion shall

generally be applied, and if yes, what may be considered as acceptable limits of changes in the U-I-

characteristic.

4. Not all investigated makes of MO resistors exhibit the expected increase of energy handling capability for

extreme values of current densities. Two makes show an unexpected decrease of the failure energy down to

values of only 500 J/cm³. The reason is a different dominant failure mechanism: resistors "S" and "U" would fail

by a dielectric failure of their coatings, finally resulting in a flashover. It is thus not a material problem of the

bulk ZnO but a characteristic of the coating system. The situation will probably not improve when the MO

resistors are directly covered by a polymeric housing, as the flashovers develop from a breakdown of the

coating material and/or the interface between ZnO and coating, respectively. Assumedly, this problem will be




Page 116

solved in the near future. Actually, MO resistors have not been optimized for the extreme current stress of thelightning current test as specified in Annex N of [IEC 2009].

Figure 3 .5: Mean fai lure energy vs. amplitude of cur rent density for “Size 1” MO resistors

For make "T", tests at long-duration current impulse stress were also performed up to mechanical failure (ratherthan to apply the "complex failure criterion"). It can be seen from Figure 3.6 that in this case the mean failureenergy follows the expected dependence of current density. This demonstrates that the differences in Figure 3.5between the a.c. and the impulse current tests are not related to the different current shapes but only to thedifferent applied failure criteria. It further allows concluding that energy handling test can also be made withalternating current stress if this is considered more convenient. The only difference to impulse current testing willthen be the lower failure energies due to the lower achievable current densities, which, however, can easily betaken into account by correction factors.




Page 117

Figure 3.6: Mean failure energy vs. amplitude of current density for “Size 1” MO resistor of make T;comparison of fai lure cri ter ia “until mechanical fai lure” and “complex” for t he long-duration current

impulse stress

The linear dependence between logarithm of current (density) and logarithm of time to failure could basically be

verified, see Figure 3.7 (for the impulse tests the prospective impulse time is used for the time scale; this is due to

the test conditions, as the test was performed with standard current impulses and not with an impulse current

lasting up to mechanical failure of the sample). It is interesting with this kind of depiction that for the short impulse

times and high current densities, respectively, the different failure mechanism of resistors "S" and "U" (flashover

instead of breaking) can clearly be identified by a change of the rate of rise of the curve – but only if the curve is

carefully interpreted. One may also (erroneously!) conclude that the same linear log-log dependence is valid over

the full covered range, as the dramatically decrease of the energies at the left end of the curve looks quite

"harmless" in the logarithmic scale and can easily be ignored. In general, this way of depiction points out general

dependencies, but is too coarse for quantitative evaluations.

0

200

400

600

800

1000

1200

1400

1600

1800

0,1 1 10 100 1000 10000

peak current density in A/cm²

m e a n f a i l u r e e n e r g y i n J / c m ³

T, complex failure criterion

Ringler 97, until mechanical failure

T, until mechanical failure

AC

8 s

100 ms 4 ms

90/200 µs

Diameter 60 mm

Heigth

40..45 mm

1 ms

2 ms




Page 118

Figure 3 .7: Mean values of current density amplit ude vs. t ime to failure

Figure 3 .8: Mean failure energy vs. amplitude of curr ent density for “Size 2“ MO resistors




Page 119

The results on the distribution MO resistors "Size 2" from six different manufacturers are shown in Figure 3.8. The

following can be derived from this picture:

1. For the long duration current impulses, mean failure energies are in the range of (600…1000) J/cm³, with

increasing values for increasing amplitudes of current density. These values can be directly compared withthose of Figure 3.5 which are typically in the range of (800…1200) J/cm³. The distribution MO resistors

thus have (15…25) % lower energy handling capability. This is explained by the different dominating failure

mechanisms, as will be shown later in more detail. In general, the "Size 2" resistors would much more often

fail due to a change of the U-I-characteristic, which can only be found with the help of the "complex" failure

criterion.

2. Except for one make ("X"), beginning with a current density of several hundred A/cm², the mean failure

energy does not increase any more but even decreases, down to values of only (150…650) J/cm³ at high

current impulse stress. This is, of course, again due to the application of the complex failure criterion. For

tests carried out up to mechanical failure an increase would have been expected in this range.

At this point some details about the typical failure mechanisms shall be given. Figure 3.9 shows the failure

mechanisms depending on the impulse shape for the station class MO resistors "Size 1", makes "S", "U" and "X".

Figure 3.10 gives the same information for the distribution MO resistors "Size 2", makes "S", "U", "V", "W" and "Y".Figure 3.11 gives an idea about the meaning of failure mechanisms "cracking" (CR), external "flashover" (FO),

"puncture" (PU) and the special characteristic of a "flashover" in case of the 90/200 impulse current stress, which

originates from a puncture of the coating.




Page 120

Figure 3.9: Failure mechanisms of “Size 1” MO resistor s

Failure mechanisms:

CR … Cracking MF … Mechanical failure during exit measurement

FO … Flashover Uch ... Change of "characteristic" voltage

PU … Puncture Ures... Change of residual voltage

BR ÜB DU MF Uref Ures

90/200 µs1 m s

2 ms4 m s

0

20

40

60

80

10 0

%

S

CR FO PU Uc h UresMFBR ÜB DU MF Uref Ures

90/200 µs1 m s

2 ms4 m s

0

20

40

60

80

10 0

%

S

CR FO PUBR ÜB DU MF Uref Ures

90/200 µs1 m s

2 ms4 m s

0

20

40

60

80

10 0

%

S

CR FO PU Uc h UresMF

BR ÜB D U MF Uref Ures

90/200 µs

1 m s

2 m s

4 ms

0

20

40

60

80

10 0

%

U

CR FO PU MF Uc h Ures

BR ÜB D U MF Uref Ures

90/200 µs

1 m s

2 m s

4 ms

0

20

40

60

80

10 0

%

U

CR FO PU MF Uc h Ures

BRÜB DU MF Uref Ures

90/200 µs1 m s2 m s

4 m s

0

20

40

60

80

100

%

X

CRFO PU Uch UresMF


90/200 µs1 m s2 m s

4 m s

0

20

40

60

80

100

%

X

CRFO PU


90/200 µs1 m s2 m s

4 m s

0

20

40

60

80

100

%

X

CRFO PU Uch UresMF

C R FO PU MF Uch Ures

90/200 µs1 m s

2 m s4 m s

0

20

40

60

80

100

%

V

CR FO P UMF Uch Ures

85/180 µs1 m s

2 ms4 m s

0

20

40

60

80

100

%

Z

CR FO PU MF Uch Ures

90/200 µs1 m s

2 m s4 m s

0

20

40

60

80

100

%

T




Page 121

Figure 3.10: Failure mechanisms of “Size 2” MO resistors

BR ÜB DU MF Uref Ures

4/10 µs

1 m s

4 m s

0

20

40

60

80

10 0

% S

C R FO PU MF Uc h UresBR ÜB DU MF Uref Ures

4/10 µs

1 m s

4 m s

0

20

40

60

80

10 0

% S

C R FO PU MF Uc h Ures

B R

Ü B

D

U

M

F

U

r e f

U r e s

4/10 µs

1 m s

4 m s0

20

40

60

80

10 0

% U

CRFO

PUMF

Uc hUres

B R

Ü B

D

U

M

F

U

r e f

U r e s

4/10 µs

1 m s

4 m s0

20

40

60

80

10 0

% U

CRFO

PUMF

Uc hUres

Failure mechanisms:

CR … Cracking MF … Mechanical failure during exit measurement

FO … Flashover Uch ... Change of "characteristic" voltage

PU … Puncture Ures... Change of residual voltage

MFB R ÜB D U MF Uref Ures

4/10 µs

1 m s

4 m s

0

2 0

4 0

6 0

8 0

10 0

%

V

CR FO PU Uc h UresMFB R ÜB D U MF Uref Ures

4/10 µs

1 m s

4 m s

0

2 0

4 0

6 0

8 0

10 0

%

V

CR FO PU Uc h Ures BR ÜB DU MF Uref Ures

4/10 µs

1 m s

4 m s

0

20

40

60

80

10 0

% W

CR F O PU M F Uc h UresBR ÜB DU MF Uref Ures

4/10 µs

1 m s

4 m s

0

20

40

60

80

10 0

% W

CR F O PU M F Uc h Ures

B R ÜB D UMF Uref Ures

4/10 µs

2 m s

0

20

40

60

80

100

% Y

CR FOPU M F Uc h Ures

B R ÜB D UMF Uref Ures

4/10 µs

2 m s

0

20

40

60

80

100

% Y

CR FOPU M F Uc h Ures

C R FO PU M F Uc h Ures

4/10 µs90/200 µs

1 m s2 ms

4 m s

0

20

40

60

80

100

%

X




Page 122

Figure 3.11: Failure mechanisms (from lef t) “cracking” (CR), “flashover” (FO), “puncture” (PU) and

“flashover” (FO) in the special case of 90 / 200 impulse current stressFor the "Size 1" MO resistors, main failure mechanism of make "S" is change of the characteristic voltage andcracking; only at the 90/200 µs impulse, it mainly failed by flashovers. For make "X" the dominating failuremechanism is cracking, and even in case of the 90/200 µs impulse it is much more the change of the characteristicvoltage than flashover. It is also interesting that in general puncture obviously is not a common failure mechanism,a finding that is in contradiction to what was published in [Rin 1997] (see section 3.4.1). Another importantobservation is that change of the residual voltage (by more than 5 %) is not a concern at all. This criterion is oftenused in the arrester standards and has to be questioned for further revisions.

For "Size 2" MO resistors, the dominating failure mechanisms are change of the characteristic voltage andflashover. This is not only the case for 4/10 µs and 90/200 µs, where this might have been expected, but also forthe long duration current impulse stress in one example. Particularly for the "Size 2" resistors one can identify atypical failure mechanism pattern of a certain make. Here again, change of the residual voltage does not take

place, and puncture is not a relevant failure mechanism, either.

Figure 3.12 and Figure 3.13 give detailed information about the change of the characteristic voltage for the "Size 2"distribution voltage MO resistors, depending on the 4/10 µs impulse current peak value and the related injectedenergy.

Figure 3 .12 : Change of characteri sti c vol tage vs. 4/ 10 µs impulse current peak value

-40

-35

-30

-25

-20

-15

-10

-5

0

5

40.000 60.000 80.000 100.000 120.000 140.000 160.000 180.000 200.000 220.000

Current peak value in A

C h a n g e o f U

c h i n %

S

UV

W

X

Y




Page 123

Figure 3.13: Change of character isti c vol tage vs. energy injection by 4/ 10 µs impulse current

The differences among the different makes of MO resistors are impressing. Only two of them – "V" and "W" –

exhibit a change of less than –5 % in the characteristic voltage after one 100 kA current impulse 4/10 µs, which is

the standard test impulse during the operating duty test on distribution arresters according to [IEC 2009]. The other

three makes are at –5 % and at –15 % for this current amplitude. Interesting as well is the fact that some makes

are obviously optimized for minimum impulse current degradation. "V" reaches the –5 % limit at a current of 150

kA, and "W" has a decrease in the characteristic voltage of only 7 % at a current amplitude of 190 kA, the slope of

the curve being extremely flat. From Figure 3.12 it can be seen that all investigated resistors easily reach energy

handling values (without mechanical failure) of (700…900) J/cm³ – assumedly even higher, but the impulse current

generator was at its limits. This example demonstrates how important the introduction of a "complex" failure

criterion is. It is acknowledged that all of the investigated MO resistors will perform well in a complete distribution

arrester because the decrease of the characteristic voltage (for practical applications this might be the reference

voltage) of –(5….15) % at 100 kA current amplitude is taken into consideration for the dimensioning of the arresters

– all of them are designed to pass the operating duty test. However, deterioration of the material has definitely

taken place (showing a linear dependence from the current amplitude), and it may just be discussed if, for instance,

a change of –10 % in the characteristic voltage should be used in the complex failure criterion rather than –5 %.

This would result, e.g. for make "Y", in an increase of the failure energy from actually 200 J/cm³ to 300 J/cm³, see

Figure 3.13.

For the first time, mechanical shock waves in MO resistors under high current impulse stress, so far only

theoretically predicted by simulations (see section 2.2), could be measured in this investigation. Figure 3.14 shows

the calculated temperature increase acc. to Equ. 3.1 in an MO resistor of 40 mm diameter and 45 mm height under

100 kA high current impulse stress and under assumption of a homogeneous temperature distribution in the

material. Temperature increases by nearly 110 K within a time of 12 µs. This adiabatic step increase causes

extreme thermo-mechanical stress, and mechanical shock waves will travel through the material.

-40

-35

-30

-25

-20

-15

-10

-5

0

5

0 100 200 300 400 500 600 700 800 900 1.000

Energy in J/cm³

C h a n g e o f U

c h i n %

S

U

V

W

x

Y"Size 2"

Diameter 40 mm

Height (30..40)




Page 124

Figure 3 .14: Calculated energy injection and temperature increase under 100 kA high current impulse

stress on an MO resistor of 40 mm diameter and 45 mm height

Figure 3.15: Current and force measured on an MO resistor 40 mm diameter and 45 mm height

The result of an actual force measurement on such MO resistor under this kind of stress is shown in Figure 3.15. It

must be noted, however, that the diagram shows the uncompensated output signal of the force sensor, which may

be affected by the test setup, especially the supporting structure and the force sensor itself. Therefore, the result

has to be carefully interpreted and the general validity to be further verified.

The propagation speed c of an acoustic shock wave in MO ceramics can be calculated as

100 GPa m mm

4300 4,35420 kg/m³ s µs

E

c (equation. 3.2)

where E is the module of elasticity and is the density (values taken from [Len 2000]. Thus for MO resistors of 27,8

mm height and of 37 mm height, respectively, especially cut to these heights for this investigation, the required time

for travelling two times along the height would be 12,9 µs and 17,2 µs, respectively. This is quite well correlated

with the comparative measurements shown in Figure 3.16. These investigations will be continued.

0

1000

2000

3000

4000

5000

6000

7000

-10 0 10 20 30 40 50 60 70 80 90

Zeit in µs

K r a f t i n N

0

20

40

60

80

100

120

140

S t r o m i n k A

Kraft in N

Strom in kA

C u r r e n t i n k A

F o r c e i n k N

Time in µs

ForceCurrent

0

1000

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7000

-10 0 10 20 30 40 50 60 70 80 90

Zeit in µs

K r a f t i n N

0

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120

140

S t r o m i n k A

Kraft in N

Strom in kA

C u r r e n t i n k A

F o r c e i n k N

Time in µs

ForceCurrent

0

20

40

60

80

100

120

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-2 0 2 4 6 8 10 12 14

Zeit in µs

T e m p e r a t u r a n s t i e g i n K / S t r o m i n k A

0

50

100

150

200

250

300

350

E n e r g i e i n J / c m ³

Temperaturanstieg in K

Strom in kA

Energie in J/cm³

T e m p e r a t u r e i n c r e a s e i n K / C

u r r e n t i n k A

E n e r g y i n J / c m ³

Time in µs

Temperature increaseCurrent

Energy

0

20

40

60

80

100

120

140

-2 0 2 4 6 8 10 12 14

Zeit in µs

T e m p e r a t u r a n s t i e g i n K / S t r o m i n k A

0

50

100

150

200

250

300

350


Temperaturanstieg in K

Strom in kA

Energie in J/cm³

T e m p e r a t u r e i n c r e a s e i n K / C

u r r e n t i n k A

E n e r g y i n J / c m ³

Time in µs

Temperature increaseCurrent

Energy




Page 125

Figure 3.16: Current and force measured on MO resistors of di fferent heights

Finally, an important result of the tests with alternating current stress shall be discussed here. It was one of the

goals of this research project to find a simple test procedure allowing statements on very low failure probabilities –

below 1% – even if only a comparatively small batch of MO resistors is tested. This was one reason for introducing

the alternating current test to the research program. Figure 3.17 shows two examples of statistical evaluations (in

form of a Normal Distribution), for a current impulse test series (left) and for a test series at alternating current

(right). For the impulse test case, each particular point in the depiction represents at least 10 individual impulse

tests. Therefore, in sum the result of about at least 40 individual impulse tests is evaluated. For the alternatingcurrent tests, each measuring point represents one energy stress up to mechanical failure; thus in sum the

evaluation is made for about 50 individual tests. The benefit of the test at alternating current up to mechanical

failure is obvious: the 95 % confidence intervals are smaller, and a reasonable statement about a 1 % failure

probability is possible from a test on only 50 samples. This is not the case for the impulse tests on approximately

the same number of samples, which can only give reliable information on the 50 % failure energy. It is thus

worthwhile thinking about tests at alternating current stress. If these tests, however, shall be performed with

application of the "complex" failure criterion, which would require a test setup that is able to interrupt the test

current at a specified injected energy level before mechanical failure, there is no difference to the impulse test any

more. On the other hand, a conversion factor from "test with complex failure criterion" to "test up to mechanical

failure" could be taken from Figure 3.5. This is open for discussion in the future. However, failure probabilities of

0,1 % or less will remain difficult to determine.

3

4

5

6

7

8

9

-5 0 5 10 15 20 25 30 35 40

Zeit in µs

K r a f t i n k N Höhe 37 mm

Höhe 27,8 mm

-20

0

20

40

60

80

100

120

-5 0 5 10 15 20 25 30 35 40

S t r o m

i n k A

6,3 µs

9,1 µs

11,6 µs

17,3 µs

C u r r e n t i n

k A

F o r c e i n k N

Time in µs

Height 37 mm

Height 27,8 mm

3

4

5

6

7

8

9

-5 0 5 10 15 20 25 30 35 40

Zeit in µs

K r a f t i n k N Höhe 37 mm

Höhe 27,8 mm

-20

0

20

40

60

80

100

120

-5 0 5 10 15 20 25 30 35 40

S t r o m

i n k A

6,3 µs

9,1 µs

11,6 µs

17,3 µs

C u r r e n t i n

k A

F o r c e i n k N

Time in µs

Height 37 mm

Height 27,8 mm




Page 126

Figure 3.17 : Examples of statistical evaluati ons (Normal Distr ibution) of impulse current tests (left ) andalternating current tests (right) w ith 95 % confidence intervals

Figure 3 .18 : Mean failure energy ( incl. standard deviation) vs. current density amplitude for MOresistors of same make and same height but dif ferent diameters

Since tests at alternating current up to mechanical failure can be made at comparatively low effort, this procedure

might be given preference in certain cases. Figure 3.18 shows an example. Here, the objective was to compare the

influence of the MO resistor diameter on energy handling capability. Each measuring point represents energy tests

at the given current density amplitude on about 50 MO resistors of the given diameter. Such test program can be

performed in comparatively short time.

0

0,2

0,4

0,6

0,8

1

1,2

1,4

1,6

1,8

0 2 4 6 8 10 12

Amplitude of current density in A/cm²

M e a n f a i l u r e e n e r g

y i n p . u

Ø 6 cm

Ø 8 cm

Ø 10 cm

0,2

0,5

1

2

5

10

20

30

40

50

60

70

80

90

95

98

99

99,5

99,8

1470 1570 1670 1770 1870

Energie in J/cm³

a r s c e n c

e

n

Verteilungsfunktion

Vertrauensbereiche Verteilung

Vertrauensbereiche Quantile

P r o b a b i l i t y i n %

Energy in J/cm³

Distribution function

Conf. interv. distribution

Conf. interv. quantile 0,2

0,5

1

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1470 1570 1670 1770 1870

Energie in J/cm³

a r s c e n c

e

n

Verteilungsfunktion

Vertrauensbereiche Verteilung

Vertrauensbereiche Quantile


Energy in J/cm³


Conf. interv. distribution

Conf. interv. quantile

99,8

99,5

99

98

95

90

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10

5

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1

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810 860 910 960 1010 1060

Energie in J/cm³

Meßpunkte

Treppenkurve

Verteilungsfunktion

Vertrauensbereiche VerteilungVertrauensbereiche Quantile


Energy in J/cm³


Conf. interv. distributionConf. interv. quantile

Measuring points

Step curve

99,8

99,5

99

98

95

90

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60

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30

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5

2

1

0,5

0,2

810 860 910 960 1010 1060

Energie in J/cm³

Meßpunkte

Treppenkurve

Verteilungsfunktion

Vertrauensbereiche VerteilungVertrauensbereiche Quantile


Energy in J/cm³


Conf. interv. distributionConf. interv. quantile

Measuring points

Step curve




Page 127

The interesting finding here is not that failure energy increases with current density (which has been expected) but

that it decreases with diameter. This can only partly be explained by statistical (volume) effects as it is

demonstrated in the following example:

The probability P (W') of failure at a given specific energy W' for n MO resistors in parallel, each having aprobability p(W') of failure at the same specific energy is

1 1 n

P W ' p W ' (equation 3.3)

Solving equation 3.3 to p(W') leads to

1 1n p W ' P W ' (equation 3.4)

where p(W') is the required failure probability of an individual MO resistor at the specific energy W' in order to

achieve the overall failure probability P(W') when a number of n resistors are connected in parallel. The number n

of parallel resistors can also be a volume factor of a larger diameter resistor.

When looking at the mean (or 50%) specific failure energy (where p = 0,5) of a resistor of 60 mm diameter, the

same specific energy injected into a resistor of the same make, but of 100 mm diameter (i.e. 2,8 times larger

volume at the same height; or n = 2,8), the failure probability acc. to equation 3.3 would be P = 0,856. If, vice versa,

the failure probability of the larger resistor shall be P = 0,5, the related failure probability of a 60 mm resistor

(volume factor = 1/2,8) would be, acc. to equation 3.4, p = 0,22.

From Fig. 3.16 (right) the 22% specific failure energy is approximately 4% lower than the 50% specific failure

energy. One might thus expect that a 100 mm diameter resistor has an approximately 4% lower mean specific

failure energy than a 60 mm resistor of the same make. Fig. 3.17, however, shows that the difference is much

bigger, i.e. in the range of 10%.

Therefore, the decrease of specific failure energy with diameter cannot be explained by statistical (volume) effects

alone. An additional influence may be the (in)homogeneity of the material. The larger the diameter the more diff icult

it is to achieve homogeneity. However, the effect is not too much pronounced and particularly covered by the wide,

overlapping deviations from the average values. It may, anyway, be concluded that a diameter of 60 mm evidently

represents a kind of optimum where change of the U-I-characteristic under high current densities (compare Figures

3.5 and 3.8) and the effects of material in homogeneities at low current densities both have minimum impact on

energy handling capability.

Another outcome of this investigation is shown in Figure 3.19. Typically, in a test series of several hundred

specimens with energy injection up to mechanical failure there will be one or more "outliers", i.e. MO resistors that

fail at extremely low energy levels. Such performance of a batch was found for all investigated makes of MO

resistors, and it shows that there will always remain a certain unavoidable risk when going to the limits of specified

energy handling capability. This is less a concern for standard applications, where an arrester is made up from

comparatively few MO resistors and is very likely never stressed to its limits, but it is definitely an issue for the large

arrester banks for overvoltage protection of series capacitors. These outliers, by the way, can better be found with

this way of testing, i.e. test at energy stress up to mechanical failure. In an impulse test with a standard long-duration current impulse, where the outcome of each individual energy stress would be only be "passed" or "failed",

the information about the low failure energy of sample number 32 in Figure 3.19 is usually not available (unless the

actual failure energy during each impulse is measured and evaluated).




Page 128

Figure 3.19: Example of failure energies at alternating current tests up to mechanical failures

3.4.2.5 Conclusions and OutlookIn this experimental investigation a statistically significant number of station and distribution class MO varistors from

eight well established American, European and Japanese manufacturers was tested to probe the energy limits for

different impulse and alternating current stresses. A major part of the test program is concluded and some general

observations and conclusions can right now be formulated:

- Compared to an earlier comparative study [Rin 1997] an increase of up to 70 % and in average of at least

20 % in energy handling capability can be seen for the materials studied here (only the alternating current

tests are directly comparable, where the same failure criteria were applied), despite the fact that MOresistors of larger (about two times) volume were tested, which in doubt would result in lower energy values

due to volume effects. This probably reflects the continuous improvements industry has made over the last

decade in processing, material formulation and block design. And evidently, cost pressure on the market

has not resulted in lower qualities.

- So far no new, emerging suppliers of MO varistors could be included in the study. However, recent

references (e.g. [He 2007]) indicate that substantial differences might exist in the performance of such

materials, which should be studied further.

- For the first time a "complex" failure criterion was introduced, which allows very sensitive evaluation of

different failure mechanisms. Since also changes of the U-I-characteristic are evaluated it allows

considering the impact of impulse energy stress on thermal stability issues. Some aspects of this failure

criterion might be further discussed, e.g. which alterations of the U-I-characteristic may be accepted as apass criterion, but basically this procedure has proven to be effective. Its application is, however, time

consuming.

- For the first time such a comparative study included MO resistors designed for distribution and station

arresters. They differ mostly in their geometrical dimensions (cross-section, aspect ratio) and also show

distinctive differences in their performance, presumably due to the fact that different features are optimized

for the two different applications.

- Energy stresses caused by fast impulses 90/200 µs as they are specified for line arrester applications in

[IEC 2009] were as well evaluated for the first t ime in this study. Flashover or significant alterations in the

0

200

400

600

800

1000

1200

1 5 9 13 17 21 25 29 33 37 41 45 49 53Versuchsnummer


F a i l u r e e n e r g y i n J / c m ³

Test and sample number

0

200

400

600

800

1000

1200

1 5 9 13 17 21 25 29 33 37 41 45 49 53Versuchsnummer


F a i l u r e e n e r g y i n J / c m ³

Test and sample number




Page 129

U-I-characteristics show up as the major failure modes when reaching the energy limits under these fast

impulses. It has thus been found that the rule "increase of failure energy with higher current density"

(published by [Rin 1997] for the first time and basically confirmed by this investigation) must be confined,

as at a certain current density stress level the performance of the coating system may become dominant,

or in other words: it has not been optimized for this kind of stress. Practical implications of these findings

for specific applications in medium or high voltage arresters have to await further discussion.

- Tests at alternating current stress have turned out to be suited for fast comparison of material properties

(benchmark tests), when the "complex" failure criterion shall not be applied. They allow statements on

failure probabilities down to approximately 1 %, which is not possible by impulse current tests. But it is still

an open problem how to reliably specify energies that would result in 0,1% failure probability or less.

- MO resistors of approximately 60 mm diameter may constitute an optimum with regard to impulse energy

handling capability.

- The change in residual voltage after energy absorption is not an issue and should be removed as failure

criterion in the standards. Change of power loss would be a sensitive failure criterion but is not sufficiently

reproducible. Change of the reference voltage could be considered for this purpose, instead.

One final general remark is important: all comparisons among different makes presented in this study are based on

mean failure energies (50 % failure probability). For these figures the statistical confidence is high, and they can

thus preferably be used for benchmark purposes. The mean failure energies, however, should not be mixed up with

the design energies, which are in the range of only 200 J/cm³ and thus far below the values reported here. No

reliable and serious statement can be given on, e.g., a 0,1 % failure energy, for reasons explained before, and

there is definitely no basis for any comparison of the investigated makes at such low probability failure energy

values. One should therefore be very careful with rash conclusions that a certain make might be "better" or "worse"

than another one, only based on this study. All investigated samples came from well-established manufacturers of

excellent reputation, and it is well known that failure rates of real arresters in the systems are only a few percent in

distribution and close to zero in transmission applications.

As the next steps the ongoing measurements will be completed. Thereafter, further studies might be necessary ontopics such as multiple impulse stress, durability, and impact of aspect ratio, probabilistic aspects in energy

handling, high gradient MO resistors or low performance makes.

3.5 Energy handling capability in international arrester standards

3.5.1 GENERAL

This section critically reviews the many different aspects of MO surge arresters' energy handling capability in

international standards, with main focus on IEC 60099-4 [IEC 2009] and IEEE C62.11 [IEE 2005]. Some national

standards have additionally been checked for differences to these standards. Requirements as well as test

procedures have been evaluated, and suggestions for future improvements are made. This was fi rst published in

[Hin 2007] and is summarized and extended here.

The basic definitions of energy handling capability – the different aspects of impulse energy and the thermal energy

– are not distinctly specified in [IEC 2009], the most important international arrester standard. Instead, energy

handling capability is only indirectly described by means of the line discharge class. But the switching surge

operating duty tests (Cl. 8.5.5 of [IEC 2009]), based on this classification, is only a thermal stability test. The long-

duration current impulse withstand test (Cl. 8.4 of [IEC 2009]) could give valuable information on the durability of

the MO arresters if it were to be performed at higher energy levels. Thus, though not mentioned in any standard,

virtually all arrester manufacturers specify a long-duration current withstand capability, usually (but not always)

based on the test procedure of the long-duration current impulse withstand test (i.e. stress by 18 consecutive

impulses in a given test sequence), but at a fixed time duration of the current impulse of e.g. 2 ms and at the

maximum permitted current amplitude at this time duration. Problem with this approach is the definition via a

current amplitude, which does not take into account differences in the injected energy caused by the wide




Page 130

tolerances of the impulse current parameters. An improvement therefore would be to also state charge transfer (in

As) or specific energy in kJ per kV of rated voltage.

The high current impulses during the high current impulse operating duty test on distribution arresters and light-

duty station arresters (Cl. 8.5.4 of [IEC 2009]) represent a considerable energy stress for this class of arresters, butare not a good basis for an energy handling capability specification either as the injected energy may vary strongly

when utilizing all allowed tolerances. Historically, this test has been introduced primarily as a dielectric withstand

test for the gapped SiC arresters, rather than an energy handling test.

This situation in the standards may have been sufficient for "classical" arrester applications. But with state-of-the-

art MO arresters many new applications have become possible and usual. Examples are line arresters, arrester

banks for protection of FACTS (particularly series compensation capacitors), shunt capacitor and reactor

protection, arresters in HVDC applications and others. Their application require a more sophisticated consideration

of energy handling capability, and an increasing number of users has profound knowledge about the arising energy

stresses and specify requirements on energy handling capability. But today's definitions and tests do not give

sufficient guidance in this respect. Therefore, activities have been started in IEC as well as in IEEE to improve the

arrester standards in this respect.

3.5.2 ENERGY HANDLING ISSUES IN STANDARD IEC 60099-4

3.5.2.1 Line discharge class of an arrester

The line discharge (LD) class is – besides the nominal discharge current – the actual determining characteristic of

a high-voltage arrester. Presently it is the only possible way of specifying the energy handling capability of an

arrester in accordance with the IEC standard. It focuses only on the aspect of thermal stability, and the meaning of

the LD class may be difficult to understand for an uninformed user. Its definition is based on the assumption that a

long transmission line, charged to a certain overvoltage during de-energization, will discharge into a connected

arrester in the form of a travelling wave. The current will flow at a value that is determined by the overvoltage value

and the surge impedance of the line, for a duration given by twice the length of the line and the propagation speed

of an electro-magnetic wave. Ideally, the resulting current is a rectangular (long-duration) current impulse. This

process is simulated in the laboratory in a line discharge test, where the current impulse is generated with the helpof a distributed constant impulse generator. Five different LD classes are defined with increasing demands from

LD 1 to LD 5, in which line discharge parameters are established, and the resulting energy content has to be met in

the test (Table 3.1). These parameters are derived from typical characteristic values of high-voltage transmission

lines (see also IEC 60099-1, Table C.1 or [IEC 2000] or Table 3.2).

LD class U s (kV)

1 245

2 300

3 420

4 550

5 800

Table 3.2:Recommended linedischarge classes

depending on system

voltage

LD class

Surge impe-

dance of the

line Z , in

Virtual

duration of

peak T , in µs

Charging voltage

U L, in kV (d.c.)

1 4.9 · U r 2000 3.2 · U r

2 2.4 · U r 2000 3.2 · U r

3 1.3 · U r 2400 2.8 · U r

4 0.8 · U r 2800 2.6 · U r

5 0.5 · U r 3200 2.4 · U r

U r = rated voltage of the test sample as an r.m.s. value in kV

Table 3.1: Test parameter for the line

discharge test [IEC 2006]




Page 131

Figure 3.20: Specific energy injecti on (by two consecutive line discharges) during sw itching impulseoperating duty test in kJ/ kV of rated volt age dependent on the rat io of sw it ching impulse residual

voltage U res to the r.m.s. value of the rated volt age U r of the arrester (in accordance w ith [ IEC 2009] )

As no direct conclusions about the energy stress can be drawn from this table, Figure 3.20 depicts the convertedenergy in a test object during the switching impulse operating duty test (injected by two line discharges), withreference to its rated voltage. This energy is not a fixed value, but depends on the arrester's switching impulseresidual voltage. The higher the residual voltage, the less energy the arrester absorbs during the line discharge.With the help of Figure 3.20 a typical problem related to the LD class definition shall be explained. If a design isapplied with a given amount of specific thermal energy handling capability, then the arrester can, depending on itsactual residual voltage, be assigned to different LD classes, as shown in the following example (red dashed lines inFigure 3.20): when using a design that can absorb an energy of 4 kJ/kV during the operating duty test, the arrester

is of LD 2 at a ratio of Ures/Ur = 2. However, it can also be assigned to LD 3 at the ratio of Ures/Ur = 2,35. But theapparently "better" LD 3 arrester might possibly be worse for the intended application, since its protective level ishigher. In order to reach LD 3 while maintaining a ratio of Ures/Ur = 2, a design must be used with a thermal energyhandling capability of almost 6 kJ/kV, as indicated by the blue dotted lines in Figure 3.20, that would meanapplication of MO resistors with greater diameters. Inversely, one can only draw conclusions from the LD class inconnection with the residual voltage as to the (thermal) energy handling capability of an arrester, and thus aboutthe used MO resistors.

For standard applications, one can simply count on recommendations in the application guide [IEC 2000], based onthe system voltage level (Table 3.2). In practice, however, users often tend to select the next higher LD class,respectively, in the table. That leads to the problem that the current highest LD 5 can frequently not meet thedemands of the extra-high-voltage systems with Us > 550 kV. In fact, at this voltage level, and sometimes even atthe 550-kV-level itself, MO resistor diameters and/or parallel connections of resistors are used, which have much

higher energy handling capability than specified by LD 5. This is also a particular problem of the emerging 800 kVd.c. and 1100/1200 kV a.c. applications, where specific energy handling values of (25…50) kJ/kV of rated voltagewill be required at switching impulse protective levels in the range of only 1,85 or even less [Ric 2007]. It can easilybe seen from Figure 3.20 that LD 5 is by far not sufficient for these applications: an LD 5 operating duty test willinject only (25…50) % of the required energy. For UHV arresters, the test procedure has therefore to be modified.For instance, energy could be injected by more than two long duration current impulses. It could also be discussedif other test parameters, such as the time duration of the impulses, may be changed. Summarizing, at least threework items can be identified for a future revision of the standard: the possible replacement of the LD system by apurely energy based rating system, clear definitions and differentiation of different kinds of energy handlingcapability, and specification of higher energy handling values than today and the related test procedures.




Page 132

3.5.2.2 Long-duration current impulse withstand test

The long-duration current impulse withstand test as per clause 8.4 of [IEC 2009] is a durability test – at least it

covers one possible aspect of durability. The test consists of eighteen discharge operations divided into six groups

of three discharges each. Time interval between individual discharges shall be fifty to sixty seconds and between

the groups such that the sample cools to ambient temperature. The following critical remarks seem adequate:

- The relevance of a number of eighteen current impulse applications is questionable: eighteen impulses

each on three samples results in 54 energy stresses. This, even if identical characteristics of each of the

three samples are assumed in the best case, cannot validate a failure probability of less than

p = 1/(3·18) = 0,0185 or 1,85 %, respectively, a value totally insufficient for real arrester applications, see

also section 3.3.2. Furthermore, it does not seem sufficient to specify such test as type test only, as energy

handling capability is also or distinctly a matter of production quality.

- As the test sample temperature has effect on the test result, the thermal conditions of the test setup and

the test procedure should be better specified.

- The injected energy per line discharge is usually less than the long-duration current withstand valuesspecified by the manufacturers; higher values should possibly be specified. It must be noted, however, that

not all manufacturers clearly specify how the long duration current withstand values are defined and

determined; therefore, comparisons with the line discharge currents have to be made with reservations.

- The maximum allowed change of 5 % in residual voltage after the test should be subject to discussion, as

such extreme changes are only seldom observed, while other relevant regions of the voltage-current-

characteristics (e.g. reference voltage) react more sensitively to impulse degradation; see also section

3.2.5.

- While in case of 10-kA- and 20-kA arresters the long-duration current impulse shall be a line discharge as

defined by the parameters shown in Table 3.1, for 2,5-kA- and 5-kA arresters (i.e. mainly distribution

arresters), the relevant current impulses shall be only (50 A / 500 µs) and (75 A / 1000 µs), respectively.

These values are in fact too low to inject any notable energy even into a light-duty distribution arrester. The

applied MO resistors for these arresters are usually specified for much higher values to date, e.g. in the

range of (200 A / 2 ms). Therefore, this requirement should be discussed as well.

3.5.2.3 Operating duty tests

The operating duty tests cover the aspect of thermal energy handling capability. They shall demonstrate the ability

of an arrester to cool back to normal operating temperature after a specified energy injection and under various

worst case assumptions, in other words: that it is not subject to thermal runaway under any circumstances. Though

the final thermal stability test itself looks quite simple, the whole test procedure is rather complex. The following has

to be done in the given order:

a) calculate (simulate) the arrester's axial voltage distribution and derive an appropriate power-frequency test

voltage for the accelerated aging test;

b) perform an accelerated test for thousand hours and derive corrected power-frequency test voltages for the

thermal stability test;

c) pre-condition the test samples by different kinds of current impulse stress in order to provoke worst case

degradation of the voltage-current-characteristic;

d) perform the thermal stability test, which consists of pre-heating, energy injection and a following application

of power-frequency voltages: U Ur for ten seconds, then U Uc for thirty minutes.




Page 133

Procedures a) and b) are covered quite well in the actual version of the standard, latest since the accelerated aging

test has been extended to the case that part of the arrester is stressed by voltages higher than the reference

voltage of the MO resistors. Item c), however, is questionable: the application of twenty lightning current impulses,

superimposed to an applied power-frequency voltage, originates from the era of gapped SiC arresters and served

for preconditioning of the series gaps, which had to interrupt the power-frequency follow current each time. For

today's MO arresters the test procedure in its actual version seems meaningless, and as this test is particularly

difficult to perform it should be simplified in the future by removing the requirement for a superimposed power-

frequency voltage.

The operating duty tests shall be performed on prorated arrester sections that represent the electrical and thermal

characteristics of the full arrester (Cl. 8.5.3.2 of [IEC 2009]). But the requirements on these sections are partly

contradictory, in that thermal equivalence is required on one hand and use of the same material and dimensions for

the housing as for the real arrester on the other. A revision of this part of the standard has therefore to make the

requirements more consistent. Furthermore, since the high-current impulse applications also impose mechanical

and dielectric stress to the sample, presence of the mechanically supporting structure should be required.

3.5.2.3.1 High current impulse operating duty test

This test according to clause 8.5.4 of [IEC 2009] applies to all 1,5-kA-, 2,5-kA, 5-kA arresters, to 10-kA LD 1

arresters and to "high lightning duty arresters" (specified in Annex C of the standard). In the majority, these are

actually the arresters in distribution systems. As a first critical comment, it does not seem appropriate to test a

10-kA LD 1 arrester according to this procedure. An arrester that has been designated an LD class should be

tested by a switching surge operating duty test (see next section). In the high current impulse operating duty test,

which has to be performed at a current amplitude of 65 kA on 5-kA arresters and at 100 kA on 10-kA

LD 1 arresters, the two high current impulses inject energy at the limits to the test samples. Furthermore, they may

cause a temporary impulse degradation (to a great extent reversible under applied continuous operating voltage

stress) with an increase in power loss by a factor of around two, and the residual voltage across the terminals is in

the range of 1,7 to 1,8 times the lightning impulse protection level. Thus additional dielectric stress is imposed,

which, historically, has been the main intention of this test. Particularly for distribution arresters that make use of

MO resistors of a high aspect ratio (ratio of height over diameter), also the stress due to thermally inducedmechanical shock waves will be extreme. The test therefore results in a very meaningful test of the overall design.

It should nevertheless be noted that an arrester, even when tested at high current impulses of 100 kA, cannot

survive a direct lightning strike of 100 kA, because this will have a much longer time duration than the standard

high current impulse 4/10 µs.

There is a particular problem when specifying energy handling capability by the high current amplitude. This is due

to the permitted tolerances in the time parameters T1 = (3,5...4,5) µs, T2 = (9...11) µs and the amplitude

î = (90...110) % of its nominal value, as well as the usual tolerances of the arrester's U-I-characteristic (the

protection level may vary by 10 % for distribution arresters of the same rated voltage). These tolerances highly

affect the amount of energy injected during high current impulse application. There is a factor of about 1,7 in

energy injection when the upper or the lower limits of all allowed tolerances are accordingly combined. This

problem can easily be overcome in a future revision of the standard by specifying a minimum charge requirement;

the current amplitude has then to be adjusted accordingly. In general, energy related tests should never bespecified by current amplitudes alone but (also) by transferred charge or injected energy.

3.5.2.3.2 Switching surge operating duty test

This test according to clause 8.5.5 of [IEC 2009] applies to 10-kA arresters of LD classes 2 and 3 and to 20-kA

arresters of LD classes 4 and 5, i.e. to virtually all high-voltage station arresters (and should generally apply to all

arresters that have an LD class, as discussed earlier). The test procedure differs from that of the high current

impulse operation duty test in that the specified energy is not injected by the high current impulse applications

(which in this case only serve for dielectric validation and impulse degradation of the electrical characteristic) but by

two line discharges. These line discharges shall exclusively heat up the MO resistors, without any additional

extensive dielectric stress or impact on the voltage-current-characteristics. It may therefore be asked if the test




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procedure could not be simplified by allowing to inject the energy by any kind of long duration current impulses

without requirements on virtual duration of the peak and amplitude, as long as the energy requirements are fulfilled.

For instance, the test could in general be performed with two or even more (in case of extra-high energy injection

requirements) rectangular impulses of variable amplitudes. The main uncertainty to answer this question – namely

the different reactions of the test sample expected for different virtual durations of the peak and different current

densities, respectively – has been clarified by the energy handling research program reported in section 3.4.2.

However, if more than two impulses shall be allowed for energy injection, possible cooling of the sample between

the individual impulses should be taken into account by correction factors for the required energy input.

3.5.2.4 High current impulse operating duty test on high lightning duty arresters

This test according to Annex C of [IEC 2009] is intended to be performed only on 20-kA high lightning duty

arresters, especially applicable for high lightning density areas with highest system voltages in the range

1 kV Us 52 KV. The energy stress is imposed by three impulse currents 30/80 µs of 40 kA peak value, one

minute apart without cooling. After the third impulse thermal stability has to be verified as in the other operating

duty tests. No published information is available about the severity of this test compared with the switching surge

operating duty test. Obviously, there is only little need for this test as this kind of arrester seems to be quite

uncommon. And in fact, Annex C is unknown to many of the users. It may thus be asked if it can be totally removed

in a future revision of the standard.

3.5.2.5 Test procedure to determine the lightning impulse discharge capability

The "test procedure to determine the lightning impulse discharge capability" (Annex N of [IEC 2009]) was

introduced to the standard in 2006 and is intended to close the gap between tests at lightning impulse currents

8/20 µs and line discharges in the range of some milliseconds. It is applicable to high-voltage (Us > 52 kV)

transmission line arresters (TLA) only.

Background of this test is that TLAs are expected to divert currents having a duration of several tens of

microseconds for arresters applied on shielded lines and several hundreds of microseconds for arresters on

unshielded lines, which considerably differs from waveforms specified in the operating duty test and in the long

duration current impulse test. An impulse duration of (200…230) µs has been considered as a suitable compromiseto cover both the typical TLA applications and also the effect of multiple strikes. It has sometimes been criticized

why not the existing 10/350 µs current impulse has been adopted. However, this would require very special test

generators, quite common for low voltage surge protective devices, but unrealistic for high-voltage MO resistors.

The 200 µs current, in contrast, can be generated with the equipment available in a standard high-voltage arrester

test laboratory.

The test is performed on single MO resistors (three samples) in still air and is thus an impulse energy handling test,

not considering any thermal stability issues. In order to not exceed the thermal stability limit, the specified impulse

energy must not be higher than the total energy injected during the operating duty test. If this is not the case, the

operating duty test has to be repeated with increased energy to cover the claimed energy rating.

The test procedure is the same as for the long-duration current impulse withstand test. The rated lightning impulse

discharge capability of the arrester is then the combination of the following:

- the lowest average peak current,

- an energy value lower than or equal to the lowest specific energy and

- a charge value lower than or equal to the lowest average charge

for any of the three test samples. Energy and charge values are taken from tables with standard values. These

tables give steps of rated energy values up to 20 kJ/kV(Ur ) and charge values up to 10 J (for comparison: a typical

LD 5 arrester has a thermal energy handling capability of about 13 kJ/kV(Ur ) and a single impulse (2 ms) energy

handling capability of about 8 kJ/kV(Ur )). For transmission line arresters, which are commonly of LD classes 2 to 4,

Figure 3.21 shows the current amplitudes which inject the same energy to the sample as the respective line

discharge, resulting in peak values up to about 15 kA (and two times this value if the same energy shall be injected

as by two line discharges in the operating duty test).




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Figure 3 .21 : Required 90 / 200 impulse current amplitudes for injecting the same energy into an MOresistor as by one line discharge equivalent to LD classes 2 to 4

It is a new approach of the IEC arrester standard to specify an arrester's energy handling capability by three

different parameters. Background is that, depending on application and system voltage level, charge transfer

capability, energy handling capability or current carrying capability may be of higher importance. As gapless line

arresters represent a comparatively new application it is actually difficult to make a commitment exclusively for one

of these parameters. Thus this set of three parameters has found international consensus. Considering the fact that

energy handling capability of MO resistors depends on current density and impulse time duration, respectively, this

new test procedure is a reasonable approach to verify TLA performance under short-duration, high-amplitude

current stress. The results of the energy handling research program clearly indicate that there are distinct

deviations from the expected energy handling capability for this current impulse (see, e.g., Figure 3.5). Together

with the existing long-duration current test two important extremes of current parameters are thus covered by the

standard.

3.5.2.6 Power-frequency voltage-versus-time characteristic of an arrester

Resistance of an arrester against temporary overvoltages (TOV) is an often raised question. The standard [IEC

2009] specifies a "procedure to verify the power-frequency voltage-versus-time characteristics of an arrester" in

Annex D, which is basically formed by the last part of the operating duty test. The test procedure starts with the

step "preheat to 60 °C", i.e. no pre-conditioning is required. Application of a voltage equal to the rated voltage for

ten seconds is replaced by application of the claimed overvoltage for the claimed time duration. This procedure

follows the general approach of the standard to specify worst case conditions: the start temperature is 60 °C, and

the injected energy is the same as in the operating duty test. However, in most cases users are interested in further

information, e.g. for lower start temperatures and for lower (or no) injected energy. Some manufacturers therefore

offer this information in addition, for instance by publishing different U-t-curves for different start temperatures or

injected energies.

3.5.2.7 Additional energy handling information from the manufacturers

Users require arrester parameters that are easier to compare than it is actually the case. Future standards have to

serve this demand. It has been mentioned before that manufacturers usually give additional information on energy

handling capability, exactly for this reason. One widely-used parameter is the long-duration current withstand

capability. Its advantage over all other kind of energy handling information is that it allows a direct conclusion to the

applied MO resistors' diameter and quality. While two to three different LD classes can be covered by one single

type of MO resistor, depending on the actual protective level (see section 3.5.2.1), a long-duration current value is

strictly related to a certain MO resistor diameter and quality. It can therefore much better be used for comparison of

different arrester makes. Most, though not all, manufacturers apply the same test procedure as in the long-duration

current impulse withstand test (i.e. stress by 18 current impulses, thus making this test a kind of durability test), with

one basic difference: the parameters of the current impulses are not derived from the line discharge class

0

2

4

6

810

12

14

16

LD 2590 A / 2000 µs

LD 3720 A / 2400 µs

LD 4880 A / 2800 µs

i m p u

l s e c u r r e n

t 8 5 / 1 8 0 ( k A )




Page 136

requirements. Instead, a fixed current amplitude and a fixed time duration are chosen, which in most cases lead to

higher stress than the standard long-duration current impulse withstand test. It has just been addressed and

motivated in section 3.5.1 that a charge transfer capability or a specific energy handling capability would be a better

definition for this purpose.

Anyway, it cannot be excluded that other manufacturers specify, for instance, a single impulse energy handling

capability with the same term (probably leading to higher current values). Thus the problem with this kind of non-

standardized energy handling definition becomes evident: the test procedure is not specified and in many cases

not even explained in the manufacturers' catalogues, the time duration of the current impulse may be chosen to

different values, and any current value may be specified, leading to a "battle of catalogue values". The actual

situation is thus not satisfying, and such kind of a well-defined energy impulse handling specification should

officially be adopted in the standard and become mandatory.

3.5.3 ENERGY HANDLING ISSUES IN STANDARD IEEE C62.11

Though IEEE standards are national standards, they serve a huge (also non-American) market and are therefore

also of international relevance. The actual version of the IEEE arrester standard C62.11 dates from 2005 [IEE

2005]. In comparison to the IEC standard, it follows a different approach in at least two aspects: firstly, it coversgapless as well as gapped MO arresters. Therefore, when comparing energy handling tests with those of the IEC

standard, one has always to recall that also the gap performance shall be validated. Secondly, the IEEE arrester

classification is application oriented, whereas the IEC classification is based on the nominal discharge current and

the line discharge class. The different arrester classes can only roughly be compared, as it is done in Table 3.3.

Further comparisons of the standards can be found in [Ham 1992] [Ost 1992].

Table no. 3.3: Comparison of arrester classes acc. to IEEE and IEC standards

3.5.3.1 High current short duration withstand test

This test (clause 8.12 of [IEE 2005]) has to be performed on a complete arrester or an arrester section, i.e. the

housing is included. Two high current impulses shall be injected, followed by application of power-frequency

recovery voltage6 for at least thirty minutes. The intention is to demonstrate the design's dielectric strength as well

as thermal stability. This may be well achieved for distribution arresters, which are tested at current amplitudes of40 kA, 65 kA or even 100 kA. But it imposes neither a notable dielectric nor energetic stress on intermediate and

station arresters that have to be tested at amplitudes of 65 kA only. For these arresters, the test seems

meaningless. Also the following is doubtful:

- the current wave shape shall be 4/10 µs (–0/+50%), which causes the same problem of an undefined

energy injection level as in the high current impulse operating duty test of [IEC 2009];

- no tolerance at all is given for the current amplitude, which is unrealistic for practical testing;

6 Voltage that causes the same watt losses in the actual MO resistors as a voltage equal to MCOV would do in aged resistors ofthe same make and of the highest specified watt losses.

ANSI/IEEEC62.11 IEC 60099-4

Light duty distribution 2,5 kA

Normal duty distribution 5 kA

Heavy duty distribution 10 kA, LD 1

Intermediate 5 kA, LD 1 or 2

Station 10 kA 10 kA, LD 3 or 4






Page 137

- the allowed time interval of five minutes between impulse application and energization at power-frequency

voltage is too far away from real service conditions; thermal stability cannot really be demonstrated by this

test procedure.

3.5.3.2 Low-current long-duration withstand test

This test is a combination of a durability and a thermal stability test. Eighteen long-duration current impulses are

applied to the test sample – a section including housing – arranged in three groups of six applications. After heating

to 60 °C two further impulses are applied, and thermal stability is verified by application of power-frequency

recovery voltage for at least 30 minutes. For distribution arresters, the long-duration current is specified at

amplitudes between 75 A and 250 A and at a time duration of two milliseconds (clause 8.13.2 of [IEE 2005]), which

is more meaningful than the IEC long-duration current impulse withstand test requirements (see section 3.5.2.2,

last item). For intermediate and station arresters, the long-duration current shall be a transmission line discharge

(clause 8.13.1 of [IEE 2005]). Different from the IEC approach, the requirements are based on the system voltage

level rather than on energy handling demands. The following items are worth a discussion:

- the test generator must be very carefully adjusted for each transmission line discharge level, making

testing rather complicated (the IEC standard, instead, only requires that the energy and time valuesresulting from the test parameters given in Table 3.1 are met, which allows less precise adjustment of the

test generator);

- placing the impulses in three groups of six applications each may lead to unnecessarily severe (and

unrealistic) thermal stress to the MO resistors;

- the requirements for system voltage levels above 400 kV are to weak, leading to the strange situation that

an arrester in a 362 kV system is specifically less stressed than an 800 kV arrester; reason is that for EHV

systems the charging voltage of the line has been set to only 2,0 p.u. of the system voltage, whereas for

the lower system voltage levels 2,6 p.u. are assumed;

- a time interval of five minutes between injection of the last impulse and power-frequency application is too

long

- possible impulse degradation by high-current impulses is not considered.

3.5.3.3 Duty-cycle test

This test (clause 8.14 of [IEE 2005]) is again a combination of durability and a thermal stability test. Twenty

lightning impulse currents having a peak value of the classifying current are applied to the test sample, which is a

section including housing, while at the same time the duty cycle voltage is applied. All twenty impulses are injected

one minute apart, leading to notable thermal stress during the conditioning phase. The test is then continued at a

given start temperature of 60 °C. Two further lightning current impulses are injected, and thermal stability is verified

by application of power-frequency recovery voltage (within five minutes after the last impulse application) for at

least 30 minutes. While the conditioning phase may be meaningful for gapped arresters – but not for gapless types

– the thermal stability part is irrelevant in any case, as the two lightning impulse current applications will lead toonly negligible temperature increase, and the power-frequency voltage is applied too late, as previously discussed.

Keeping this test for gapless MO arresters is questionable at all, and for gapped type arresters at least the thermal

stability part needs consideration.

3.5.3.4 Temporary overvoltage (TOV) test

Different to the IEC approach, the IEEE standard requires a temporary overvoltage test as a mandatory design test

(clause 8.15 of [IEE 2005]). It is also more stringent in that it requires a test for five time ranges (IEC: three) on five

samples each (!), and for intermediate and station arresters a "prior duty" test besides the "no prior duty" test. The

test procedures are adequate, but the test effort (number of tests and samples) seems extremely high for the

intended purpose.




Page 138

3.5.4 ENERGY HANDLING ISSUES IN OTHER NATIONAL STANDARDSWhen looking for further international arrester standards, one will find the European standards (EN), published by

CENELEC. However, they are identical with the IEC standards, with some national exceptions to be found in

special Annexes. From the arrester standard series, only IEC 60099-6 [IEC 2002] was not adopted by CENELEC.

Only very basic information on some national arrester standards can be given here. The Japanese standard JEC-

2371 [JEC 2003] is, with regard to energy handling issues, very similar to the IEC standard. The test procedures

are the same, but the test parameters are partly different.

The Australian standard AS 1307.2 [AS 1996] is also very close to the IEC standard, with some additional

requirements. It contains e.g. an optional "Multipulse lightning impulse current operating duty test" (Appendix O;

informative) and thus takes into consideration the observations reported in [Dar 1998].

The Chinese standards on MO arresters are GB 11032 [GB 2000] for gapless MO arresters in general and

equivalent to IEC 60099-4:1991, JB/T 8952 [JBT 2005] for polymer housed gapless MO arresters, based on IEC

60099-4:2001, and DL/T 815 [DLT 2002] for polymer housed gapped and gapless transmission line MO arresters.

They are all basically identical to the IEC standard 60099-4. But, to mention a difference with respect to energy

handling requirements, [GB 2000] addresses the long-duration current impulse withstand of 5-kA, 2,5-kA and 1,5-kA arresters in a more appropriate way, i.e. the requirements are higher than in the IEC standard (see section

3.5.2.2, last item).

The Canadian standard CAN/CSA-C233.1-87 [CAN 2004] represents a mix of the IEC and the IEEE standards.

Basically the material for this standard originates from IEEE/ANSI Standard C62.1-1967 and IEC Publication 99-

1A-1965 as well as IEEE/ANSI Standard C62.11 and IEC draft documents on MO arresters prior to 1987. It was

thus prepared in the transition period from gapped SiC to gapless MO arresters, with a latest reaffirmation in 2004,

however.

3.5.5 CONCLUSION AND OUTLOOKIt took a long time to publish the first standards on MO arresters, more than ten years after the first MO arresters

appeared on the market. One has to keep in mind that writing a standard on power system equipment based on a

totally novel technology is challenging. Admittedly, the outcome has been more than satisfying. MO arresters

tested to these standards belong to the most reliable devices in electrical power systems today. However, the time

has come to reconsider requirements and tests that were considered necessary and appropriate twenty years ago.

Knowledge on MO arresters has increased, allowing for improving the standards today. The need for changing

some of the requirements and test procedures is obvious. The question is less if these issues have to be

reconsidered than how they should be addressed. A verification of energy handling capability should take into

account aspects of durability, impulse degradation and thermal stability. The complicated approach of transmission

line discharges should be replaced by a simple requirement on energy handling capability under long duration

current impulse stress. A fi rst new proposal in this direction has been presented by an IEEE/SPDC Working Group

and will be further developed in cooperation with the responsible group within IEC TC37. In this context, it is seen

as an important progress that standardization is being more and better harmonized at least among IEC and IEEE.

As most of the national standards refer to either the IEC or the IEEE standards it should be possible to introduce

new, internationally agreed MO arrester energy handling requirements and test procedures in reasonably short

time.




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4. Summary


MO arresters are nowadays installed in all kind of electrical power systems from low voltage up to UHV. They areintended to protect equipment and installations against overvoltages. Due to the various applications the MOarresters have to withstand severe stresses from the electrical system, from lightning and from ambient. In this newTB the different types of stresses are listed and severe stresses, e.g. winter lightning, seismic stresses and severepollution of polymeric housings, are shown and examples for test procedures are given.

Main focus is given on the progress in arrester technology and application in the past 20 years.For understanding the interaction of the modern MO surge arresters with the system conditions and the ambientstresses, the basics of the MO material and the various designs of MO arresters on the market are given in detail.The working group engaged in a critical review of the applicable standards and initiated a research program withinternational participation on energy handling withstand capability of MO resistors and arresters. The results showthat the impulse withstand capability has increased by app. 20% for the manufacturers participating in this study,compared to previous investigations. However, differences exist in the mean failure energy and the failure modesdepending on the type of impulse stress. Unfortunately, materials from emerging countries were not available for

this study when it was initiated.For clarity it should be noted, that the observed mean failure energy of the MO resistors is three to four timeshigher than the design energy of the MO arresters as proved in the relevant operating duty tests. This gives a goodsafety margin and confidence in the todays designs and materials. However, with increasing system voltage thenumber of MO resistors easily reaches several hundreds of blocks per arrester, and then statistical evaluationsmay have to be considered.For a more sensitive evaluation of the MO resistors in this study, a new complex failure criterion was developedand used, including the change of the electrical characteristic of the MO resistors in addition to the simple failure bycracking or flashover.

The different types of failure modes depending on the different current wave shapes can lead to furtherimprovements of the MO resistors.

The results from the research project and the review of the existing standards, together with new applications,

leaded to a new classification of the energy capability of MO surge arresters. The new classification concept,charge transfer classes instead of line discharge classes, is introduced in Rev. 3.0 of IEC 60099-4. The sameconcept is adapted for instance in EN 50 526-1, which is a test standard for MO surge arresters to be used in d.c.traction systems.

MO arresters are applied more and more for insulation coordination reasons and not only for protection of a singlehigh voltage equipment against overvoltages. This is especially the case for UHV a.c. and d.c. systems. Thedevelopment of MO resistors and arresters is ongoing with the goal of size and cost reduction but in the same timekeeping the high quality and reliability. This leads to the development and use of MO resistors with increased fieldstrength to reduce the size of the complete design, e.g. in GIS applications. Further on, with the very tall arrestersfor UHV systems, which are easily taller than 10 m, the question of how to test complete arresters comes up.Simulations may help to reduce testing. These questions and ongoing research on the energy withstand capabilityof MO resistors and arresters are dealt with in working group A3.25 of SC A3 and will be published in a separateTechnical Brochure.




Page 140

APPENDIX 1

Following Technical Brochures of Cigré are dealing with surge arresters and their application:

TB 60 Metal Oxide Arresters in AC Systemsby WG 06 of SC 33, 1991

TB 287 Protection of MV and LV Networks against Lightn ing. Part 1: Common Topicsby CIGRE-CIRED JWG C4.4.02, 2006

TB 441 Protection of MV and LV Networks against Lightning. Part 2: Lightnin g protection of MediumVoltage Networksby CIGRE-CIRED JWG C4.4.02, 2010

TB (XXX) Protection o f MV and LV Networks against Lightn ing. Part 3: Lightn ing protection o f Low-VoltageNetworksby CIGRE WG C4.408, to be published 2013

TB 440 Use of Surge Arresters fo r Lightning Protection of Transmission Linesby CIGRE WG C4.301, 2012

TB 455 Aspects for the Application of Composi te Insulators to High Voltage ( 72 kV) Apparatusby CIGRE WG A3.21, 2011This TB addresses the special case of surge arresters with composite insulators.




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Cigre_544 -- Metal Oxide (MO) Surge Arresters - Stresses and Test Procedures