Cigre_544 -- Metal Oxide (MO) Surge Arresters - Stresses and Test Procedures
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544 MO Surge Arresters Stresses and Test Procedures
Working Group A3.17
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)
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ISBN: 978-2-85873-239-5
MO Surge Arresters-Stresses and Test Procedures
MO Surge Arresters STRESSES 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 Page 2
MO Surge Arresters-Stresses and Test Procedures 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
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EXECUTIVE SUMMARY The 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 to maximize 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 < Us 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 biological growth 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 Resistors 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, improved and 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 highcurrent 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.
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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 detailed information 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 electronic applications, 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 many different 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 in HVDC 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. Functional parameters and design of 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 1 gives 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 with a 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 from the 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 GENERAL CIGRE 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 relatively long 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 rejection overvoltages 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.
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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 sufficiently probable, 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 OVERVOLTAGES Slow front overvoltages generated by switching have traditionally been associated with lines, load or fault clearing. With the emergence of flexible AC transmission system (FACTS) devices such as static var compensators(SVC), static compensators (STATCOM), and thyristor controlled series capacitors(TCSC) which require power-electronic types of switching and SA protection for both the electronic and power components.
1.2.3.1 Line Reclosing Random 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
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considered necessary to ensure line switching performance and reliability when controllers are out-of-service or CBs misoperate.
1.2.3.2 Capacitive Loads Capacitor 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-toground 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 loads Switching 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) Devices Power 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 TRVS Interrupting 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
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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.
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1.3 Stresses from HVDC networks Authors in charge: Bengt Johnnerfelt and Reinhard Göhler
1.3.1 INTRODUCTION The 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 operating 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 rd th th th be stressed with sinusoidal voltages but with higher frequencies, like for example 3 , 5 , 11 , 13 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.
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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 IEC are 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 energy impulses 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 can be 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 metaloxide 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
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MO Surge Arresters-Stresses and Test Procedures
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 voltages The 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 with respect 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:
CCOV
3
U dim
(equation 1.1)
2 U v0
Udio
ideal no-load direct voltage (IEC 60633)
Udim
maximum value of Udio
U v0
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.
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Figure 1.1: Typical waveforms of continuous operating voltages at various locations in the converter station. All individual figures show the voltage over time. 1.3.2.2 Sources and types of overvoltages Overvoltages 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 shortcircuits within the converter unit.
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1.3.2.2.2 Fast-front, very fast-front and steep-front overvoltages Travelling 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 AIR The 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 of contaminated 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 HVDC applications. 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 ARRESTERS Metal-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.
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The amplitude of the current for which the protective level is specified, which is referred to as the co-ordination current, 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 rated voltage 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 continuous operating voltage is defined differently because the voltage wave shape which continuously appears across the arresters consists, in many cases, of superimposed direct, fundamental and harmonic components and, in some cases, 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, different from standard tests usually applicable for a.c. arresters. The required energy capability of the arresters shall consider the applicable wave shapes as well as the amplitudes, duration and the number of respective discharges.
Figure 1.2: Operating voltage of a valve arrester, rectifier operation
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MO Surge Arresters-Stresses and Test Procedures
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 converter station
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
Current extinction
X
X
Loss of return path, monopolar operation or commutation failure
X
Earth faults and switching operation, a.c. side
X
Lightning from a.c. system
X
Station shielding failure (if applicable)
X
X
X
X X
X
Table 1.1: Events stressing the different arresters
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X
MO Surge Arresters-Stresses and Test Procedures
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-flashovers and 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 lightning stresses. 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.
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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 filter capacitor discharge transients resulting from earth faults on the d.c. pole.
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1.4 Stresses in traction systems Author in charge: Bernhard Richter
1.4.1 GENERAL The 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].
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1.4.2 VOLTAGES IN TRACTION SYSTEMS The 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
Highest non-permanent voltage
Umax1
Umax2
V
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 of regenerative breaking.
Table 1.2: Voltages in traction systems
Figure 1.4: Highest values of voltage occurring in the system depending on time duration
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MO Surge Arresters-Stresses and Test Procedures
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: Voltage at the current collector of a d.c. traction vehicle for a period of 15 min, urban transportation system 1.4.3 MO SURGE ARRESTERS FOR D.C. TRACTION SYSTEMS The 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 Qt = 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].
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1.5 Stresses from Lightning Authors in charge: Trond Ohnstad and Yoshihiro Ishizaki
1.5.1 INTRODUCTION Lightning 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 lightning 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 lightning 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- (U s up to 1 kV) and MV- systems (1 kV < U s
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 < U s
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 < Us
800kV) and UHV-systems above 800kV
Transmission 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.
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MO Surge Arresters-Stresses and Test Procedures
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 0.23 - 52
52 - 245
245 - 800 and above
Insulation characteristic BIL < induced lightning surges. Determined by lightning overvoltages BIL determined by lightning overvoltages.
BIL > induced lightning surges. BIL > surges caused by direct stroke to the phase. BIL determined by switching surges.
Lightning overvoltage Induced surges can cause flashover. Direct stroke to a line or substation is very rare. Induced surges can cause flashover. Direct stroke to shield wires with back flashover, and direct stroke to phase conductors. Direct stroke to the power line, to a phase conductor or shield wire.
Surge arrester stress No critical stress due to lightning.
Critical stress from lightning can occur. High energy ratings needed.
Critical stress only due to no shielding or shielding failure or high earth resistance combined with a 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 145 kV with shielding wires 420 kV with shielding wires
2,20 kJ / kV 0,02 kJ / kV 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 of Norway gave the following energy levels : -
Direct stroke to the phase wire (10kA) Stroke to the shielding wire (100kA) Stroke to a tower without shielding wires
0.88kJ/kV 0.04kJ/kV 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
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MO Surge Arresters-Stresses and Test Procedures
-
Stroke to a shield wire (50kA)
0.10 kJ/kV
1.5.4 LIGHTNING STATISTICS Since 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
1
23
17
18
0.6
37
22
24
-
-
-
24 (51kA 16% value)
0.1 - 200
-
-
35
Norway
Norway west coast Japan (winter)
Cigré (all world)
Table 1.4: Lightning parameters in different parts of the world 1.5.5 WINTER LIGHTNING.
1.5.5.1 General Norway 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.
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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 with the cycle of a thundercloud (Chisholm and Renick, 1972); in the lower part of this figure the temporal variation of each radar echo reflectivity in a convective cloud is illustrated in correspondence with the echo life 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, lightning in 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.
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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 current For the cumulative probability distribution of lightning stroke peak current, a variety of them have been proposed mainly targeting at summer lightning, 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.
(equation 1.2)
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.
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probability (%)
MO Surge Arresters-Stresses and Test Procedures
Figure 1.7: Various data of cumulative probability distribution 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 currents Cumulative 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.
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Summer lightning
probability
probability (%)
Winter lightning
Wave front duration µs
Data of Fukui & Kashiwazaki Type A
Wave front duration µs
Figure 1.8: Probability of the wave front duration of lightning current [Ike 1981][Asa 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.
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MO Surge Arresters-Stresses and Test Procedures
probability (%)
Wave tail duration 50%: 50.1 µs 16%: 794 µs
Wave tail duration µs
Figure 1.9: Probability of lightning current wave tail duration of winter lightning [CRI 1989]
Wave tail duration µs
As understood from the relationship between wave front duration and wave tail duration in Figure 1.10, the longer the wave front duration, the wave front duration also tends to be longer. According to data of Figure 1.11, which were measured by Berger [Ber 1975], the 10% value of the wave tail duration of negative first stroke currents exceeds 150 s. When focused on the positive polarity lightning, the 10% value exceeds 1 ms.
Wave front duration µs
Figure 1.10: Relationship between the wave front and wave tail duration of the lightning current of winter lightning [CRI 1989]
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probability (%)
MO Surge Arresters-Stresses and Test Procedures
1: negative first lightning stroke 2: negative 3: positive lightning stroke
Wave tail duration µs
Figure 1.11: Probability of wave tail duration of lightning current [Ber 1975]
: Total
probability
30% : 2.9C
Amount of electric charge (C) Figure 1.12: Probability of electric charges of winter lightning current [Miy 1992]
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Current peak value
MO Surge Arresters-Stresses and Test Procedures
Time
Figure 1.13: Example of positive polarity current shape of winter lightning measured at Fukui in Japan in February 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 the cumulative 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.
probability
-
1: negative first lightning stroke 3: positive lightning stroke
Electrical charge C
Figure 1.14: Probability of electrical charges of summer lightning [Ber 1975]
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MO Surge Arresters-Stresses and Test Procedures
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.
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MO Surge Arresters-Stresses and Test Procedures
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 and the 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.
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MO Surge Arresters-Stresses and Test Procedures
Year
1965-
1970-
1975-
1980-
1985-
1990-
st
1
JEAG
Static 0.5 G
-5003
1995-
2000-
2005-
Revision
Dynamic 0.3 G
(1980) st
(1998)
IEEE
1
Revision
Revision
-693
(1984)
(1997)
(2005)
68-3-3 [Note]
68-2-57
(1991)
1463
62271-2
(1989)
1166
(1996)
(2002)
IEC
(1993) Tottoriken Niigata -Seibu (1964)
Off-
Nihonkai
Hokaido
Hyogoken
Off-
Miyagi
-Chuubu
-Nanseioki
-Nanbu
Tokachi
(1978)
(1983)
(1993)
(1995)
Earthquake
(2000)
in Japan
Niigata -Chuuetu (1968) (2004)
Earthquake
Sanfernand
Northridge
in US
(1971)
(1994) Indonesia Kocaeli,
Other earthquake
Off-
Pakistan
Sumatra
(2005)
Shu-shu (1999) (2004)
[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.
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MO Surge Arresters-Stresses and Test Procedures
Arrester standard
IEC60099-4
IEEE C62.11
JEC-2371
(Circuit breaker)
IEEE-693
JEAG-5003
IEC 60068-3-3
(General)
(General)
IEC 62271-300 Referred guide for seismic test (General) 90kV Test Voltage rating
170 kV
Not detailed ( 90kV Analysis) 0.5 G High
Input acceleration
0.5 G High 0.3 G
0.3 G Moderate 0.25 G Moderate 0.2 G Low
Frequency range Seismic test method
0.5 35 Hz
0.3 33 Hz
0.5 10 Hz
Artificial earthquake which complies with RRS
Artificial earthquake which complies with RRS
Resonant 3 cycles (Note 2) sine wave
Waveform
RRS Max. acceleration response at 2% damping
RRS
1.4G (High) 1.62G (High) 2.35G
0.85G (Moderate) (Single-degree-of-freedomsystem)
Acceptance criteria
For structural parts For arrester performance
0.81G (Moderate) 0.56G (Low) Specified
Specified
Specified
in the guide
in the guide
in the guide
Not detailed
Not detailed
Not detailed
(Note 1)
(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
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MO Surge Arresters-Stresses and Test Procedures
(3.1) Test condition of 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 manufacturer and user in Annex M and refers the IEC61166 (IEC62271-300), which is the guide for circuit breakers. As arresters of higher rating have more sensitive and complex responses against the earthquake, the actual seismic test may be necessary. Therefore the guide for the seismic test and evaluation method for arresters is required. (3.1.1) Background of the test condition (1) Acceleration Figure 1.15 shows the map that indicates the value of horizontal acceleration on the ground surface expected in Japan once at the interval of 75 years (return period of 75 years). According to this figure, the horizontal accelerations are less than 0.3 G in almost all regions in Japan.
Figure 1.15: Distribution of horizontal acceleration on the ground surface expected once at the interval of 75 years [JEA 1998] Table 1.7 shows the seismic intensity scale used by the Japanese Meteorological Agency. The acceleration of 0.3 G (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 the past records of the acceleration are below 0.3 G (294 Gal) in almost all regions in Japan. Please note that the mentioned scale was revised in 1997.
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MO Surge Arresters-Stresses and Test Procedures
Intensity scale
Designation
Acceleration (Gal)
0
No feeling
< 0.8
A slight earthquake
0.8
2.5
A weak earthquake
2.5
8
A rather strong earthquake
8
A strong earthquake
25
80
A very strong earthquake
80
250
A disastrous earthquake
250
A very disastrous earthquake
> 400
25
400
Table 1.8: Seismic intensity scale used in 1949 – 1997 by the Japanese Meteorological Agency [JEA 1998]
Hokkaido 107 12
0
Tohoku 116 25
0
Hokuriku 17
3
1
Kanto
13
1
Chubu
93
Chugoku
7
205 3
0
32
Kansai
95 17
4
[25-80] [80-250] [250-400]
1
Sismic intensity scale [Acceleration (Gal)]
Chugoku 58
13
Kyushu 4
1
0
0
Figure 1.16: Seismic records in the past 75 years from 1921 to 1995 in Japan [JEA 1998]
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MO Surge Arresters-Stresses and Test Procedures
(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 the vibration, therefore the quasi-resonant method by resonant N cycle sine wave is adopted as the Japanese seismic test. (2.1) Frequency The predominant frequencies of earthquakes in Japan are 0.5–10 Hz, which overlaps the natural frequencies of porcelain type apparatuses, bushings, arresters, etc. So the severest test condition is realized when the test wave is 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 sine waves 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 (1.1 times) are considered. As the correction factor of 1.3 (1.2x1.1) corresponds to the ratio of the response factor to resonant 3 cycles sine wave to that to resonant 2 cycles sine wave (6.1/4.7 as shown in Figure 1.16), resonant 3 cycles sine wave is adopted as the seismic test wave.
Figure 1.17: Comparison of the response factors between a simplified single-degree model to resonant 1-3 cycles sine waves and the actual 615 earthquake records in Japan [JEA 1998]
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MO Surge Arresters-Stresses and Test Procedures
(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 Item Natural frequency [Hz]
Seismic test JEAG-5003
of
Y
Z
24
17
> 30
Wave shape
3 cycles sine wave of 10 Hz
Input acceleration [G]
0.3
0.31
0.31
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
Response factor
Seismic test with
X
Wave shape
Actual earthquake wave
Input acceleration [G]
0.35
0.35
---
Tank (A4)
1.1
1.0
---
Middle of Internal parts (A3)
1.3
1.0
---
Top of Internal parts (A2)
1.3
1.1
---
the EL Centro Response earthquake wave factor
Table 1.9: Seismic test results of 500 kV GIS arrester in accordance with JEAG-5003 [Shi 1997]
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MO Surge Arresters-Stresses and Test Procedures
Insulating rod (S1) Acceleration
Tank (A4) Middle of internal parts (A3) Top of internal parts (A2) Input acceleration (A1) Time a: Example of wave shapes
Input acceleration (A1)
Tank (A4) X Z Y 1.5 m
Top of internal parts (A2) Middle of internal parts (A3)
Shield ZnO element Shaking table b: 500 kV GIS shaking table, seismic test for horizontal installation
Insulating rod (S1) Y
c: Internal view of 500 kV GIS arrester
X Z
Figure 1.18: Seismic test of 500 kV GIS arrester in accordance with JEAG-5003 [Shi1997]. The axis X,Y and Z give the directions of acceleration, see Table 1.9.
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MO Surge Arresters-Stresses and Test Procedures
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 test procedures 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.
3 Risk of "internal" partial discharges, degradation of the MO resistors and deterioration of the supporting structure
1 Risk of external flashover (see IEC 60507)
2 Risk of partial heating of the active parts (see Annex F of IEC 60099-4)
Figure 1.19: Possible risks due to pollution
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MO Surge Arresters-Stresses and Test Procedures
MO column Gas
Conductive layer
Uaxial, int
Solid
Uradial
Figure 1.20: Possible voltage distribution of an arrester unit under polluted conditions Table 1.10 gives an overview about proposed methods for pollution tests of polymer housed MO surge arresters. The main discussed and open point is the way of treating the polymer surface of the insulating housing to get realistic results in the test, which can be compared to long term stresses in the system. Item
Solid layer
Standard to be referred
IEC 60507
Artificial pollution procedure
After the polluted insulator dried, the pollutant wetted by clean fog
Pollutant slurry sprayed on insulator
Salt fog generated
Soluble component
NaCl
NaCl
NaCl
NaCl
Kieselguhr
Kaolin
Tonoko
SiO2
or Tonoko
or Bentonite
Non-solible component
Note
Sprayed pollutant
Salt fog
JEC 0201 IEC 60507 IEEE C62.11
--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, is described for information in the following, see Figure 1.21.
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MO Surge Arresters-Stresses and Test Procedures
Remark 1: Tonoko Tonoko is fine inorganic powder from particular kinds of natural rock, which has been originally used as fine abrasive 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. Granulometry (cumulative distribution) mm
SiO2
Al2O3
Fe2O3
H2O
16%
50%
84%
Volume conductivity s20 (S/m)
60 - 70
10 – 20
4-8
-
0.8 - 1.5
3–5
8 - 15
0.002-0.01
Weight composition (%)
Table 1.11: Characteristics of Tonoko [IEC 1991]
Figure 1.21: Procedure to remove hydrophobicity 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 of MOR temporarily for the testing, without any damage on the surface or any additional chemical agent in the pollutant. 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.
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MO Surge Arresters-Stresses and Test Procedures
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 (U c) 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.
10 min. contaminant application
UC
3 min.
Figure 1.22: Continuous voltage U c applying method according to IEC standard [IEC 2006]
×4 sequences
E2
E1 1 min
contaminant application
3 min
1 min
1 min
1 min
1 min
E1
continuous operating voltage
E2
temporary over voltage
1 min
1 min
1 min
first sequence
Figure 1.23: Temporary overvoltage (TOV) applying method according to JEC standard [JEC 2003]
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1.6.3 HUMIDITY Author in charge: Kari Lahti
1.6.3.1 Ambient humidity stress The 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 world’s climatic zones. The most humid climates in equatorial and tropical zones are indicated by colors [Mar92]
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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.
Ng
0.04TD1.25 -2
(equation 1.3) -1
Ng = Ground flash density (km yr ) TD= Number of days with thunderstorms
Figure 1.25: Annual number of days with thunderstorms [Mar 1992]
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1.6.3.3 Polymer housed surge arresters under humidity stresses Tightness 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 arresters Figure 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
Leakage Current ( A)
Table 1.12: Tested medium voltage arrester types according to internal structure
100 90 80 70 60 50 40 30 20
Group I Group II Group III
10 0 0
100
200
300
400
500
600
700
800
Testing Time (days)
Figure 1.26: internal leakage currents of MO surge arresters according to their internal structure during the humidity chamber test
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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 STRESSES Continuous 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 partial pressures 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 +30 C to +50 C. 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. The behavior 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.
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10 9
5
D1 D2 D3 D4 D5
4
D6
8 Power losses (W)
7 6
3 2 1 0 0
10
20
30
40
50
60 Test duration (weeks)
70
80
90
100
Figure 1.27: Internal power losses of type D (internal structure type II) at 12 kV during the test
E1 E2 E3 E4 E5 E6 E7 F1
10 9 8 Power losses (W)
7 6 5 4 3 2 1 0 0
10
20
30
40
50
60 Test duration (weeks)
70
80
90
100
Figure 1.28: Internal power losses of type E (internal structure type II) and F (internal structure type III) at 12 kV during the test
5 H1 H2 H3 H4 A1 A2
Power losses (W)
4 3 2 1 0 0
10
20
30
40
50
60 Test duration (weeks)
70
80
90
100
Figure 1.29: Internal power losses of type A and H (both of internal structure type I) at 12 kV during test
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1.6.5 EXPOSURES TO LOW AMBIENT TEMPERATURES Exposures 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 coverings Studies 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 thus decrease 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.
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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 a minor 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 to rocks, 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.
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MO Surge Arresters-Stresses and Test Procedures
Figure 1.30: Example of biological growth.
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1.7 Short circuit currents Author in charge: Bernhard Richter An arrester overload is a very rare event. However, it cannot in principle be ruled out, not even in the case of an over dimensioned arrester. Possible causes are, for instance, direct lightning strikes occurring near the arrester, or power frequency voltage transfer from a higher to a lower voltage system, for example on a transmission line with several voltage levels crossing each other and suffering a conductor failure or galloping. In transmission systems this occurs considerably less frequently than in distribution systems. In some special applications an MO arrester is intentionally 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 will occur internally in the housing between the two flanges. The full short circuit current of the net, which appears where the arrester is actually installed, flows through this arc. As a result, an abrupt increase of pressure develops within 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 is valid for all arrester designs with considerable enclosed gas volume. To avoid dramatical failures (violent shattering) pressure relief devices, integrated in the flanges, are needed, typically on both sides of each arrester unit. The pressure relief devices have to open within a few milliseconds so that the arc can commutate to the outside 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 tested with specific “pressure relief tests”. Figure 1.31 shows the principle steps of a pressure relief of a porcelain or hollow core housed arrester unit.
Figure 1.31: Pressure relief of a porcelain housed arrester unit. left: puncture and flashover of individual MO resistors middle: internal arc along the full length of the unit right: 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 their housing, 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 an arbitrary 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 overloading the housing must either remain intact, or if it breaks, the housing fragments and ejected parts must fall down in a narrow and defined area.
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Figure 1.32: Short circuit behavior of a cage design polymer housed arrester unit (courtesy ABB). 1: the arrester has failed and gas begins to be expelled through the housing 2: the gas streams trigger 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 as symmetric current in r.m.s. value with power frequency. The manufacturer has to state the rated short circuit current the arrester is intended to withstand. Table 1.13 gives the rated short circuit currents for short circuit tests as they are defined in IEC 60099-4.
Rated short circuit current 200 ms / A 80 000 63 000 50 000 40 000 31 500 20 000 16 000 10 000 5 000
Reduced short circuit currents 200 ms / A 25 000 50 000 12 000 25 000 25 000 12 000 25 000 12 000 6 000 12 000 12 000 6 000 6 000 3 000 6 000 3 000 3 000 1 500
Low short circuit current 1s/A
600 ± 200
Table 1.13: Standard short circuit currents for tests purposes of MO arresters Once the rated short circuit current is claimed for an MO arrester the type tests have to be performed with all the short circuit currents given in Table 1.13. The required tests at reduced short circuit currents shall help avoiding that the short circuit performance is optimized for high short circuit currents only while at lower short circuit currents the housing might violently shatter before the pressure relief devices can open.
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To avoid misunderstandings, it has to be pointed out that the short circuit current with power frequency, driven by the system voltage, is flowing through the surge arrester, or more likely through the arc, only after the arrester was overloaded. 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 the results 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 in a well-defined enclosure. The MO arrester is of course destroyed in both cases, but it failed in a well-defined manner.
Figure 1.33: Examples for successful short circuit tests of porcelain housed arrester units. left: intact porcelain body, right: secondary break of porcelain body(courtesy Siemens). Rated short circuit current tested 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 are to be considered positive. In both cases the arrester units are destroyed, but without cracking or ejecting hard parts like MO resistors or parts of them. In case of the cage design, Figure 1.34 right, only soft polymeric material is ripped off.
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Figure 1.34: Example of successful short circuit tests on polymer housed MO arrester units. left: FRP hollow insulator (courtesy Siemens), right: cage design (courtesy ABB). In both cases the short circuit current was 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, surge arresters have to be designed for an overload. Important is that the arrester fails in a controlled and safe way. This is an important safety issue when arresters are installed close to public places, sensitive infrastructure or for instance on the roof of a traction vehicle in railway applications.
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2.
Functional parameters and design of MO Surge Arresters
2.1 Function and relevant parameters Author in charge: Bernhard Richter
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: Voltage-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).
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In Figure 2.2 a more technical diagram is given, indicating the standardised definitions.
Figure 2.2: Voltage-current characteristic of a MO arrester with I n = 10 kA, line discharge class 2. The voltage is normalized to the residual voltage of the arrester at I n . The values are given as peak values for the voltage (linear scale) and the current (logarithmic scale). Shown are typical 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.
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MO Surge Arresters-Stresses and Test Procedures
Uc
Isw
Ur
In
Uref
Ist
Ihc
Figure 2.3: Typical voltage and current waveforms of a MO surge arrester. -
Uc Ur Uref Isw
continuous operating voltage rated voltage reference voltage 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 operating 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.
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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 voltagecurrent 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 with DC 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 discharge 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 lightning 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 1 s 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 voltagecurrent 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
LD
kA
Us
L
kV
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 l d 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 electrical equipment. 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.
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It makes sense to choose the continuous operating voltage Uc a little 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 occurring voltages in the system, the withstand voltage of the electrical 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 nonlinear 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.
Figure 2.5a: Linear presentation of the characteristic of a metal-oxide resistor for the high voltage sector
Figure 2.5b: Log-log plot of the normalized J(E)-characteristics of a typical ZnO varistor [Gre 1989]
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 U v Continuous operating voltage (AC, 50 Hz) Up Residual voltage, 8/20 s Resistivity Non-linearity exponent (U)
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By careful dimensioning of the geometry and controlled manufacture of the metal-oxide ceramic, the breakdown 4 voltage per element can be set to values within a very wide range (UB 3 V to >10 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, which each 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. In the 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 P v(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 U B . 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 can also 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 RESISTORS The 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 10-20 m 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-
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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 microstructure of a ZnO-varistor after fracturing preferentially along the grain boundaries (dark: doped ZnO grains, white: Bi2O3-phase at triple points, grey: Spinel secondary phase Most of the admixed bismuth oxide collects as a separate phase at the triple points at which the adjacent grain edges 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 PROCESS The 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 1. 2. 3. 4. 5. 6. 7.
2
3
4
5
6
7
Figure 2.7: Manufacturing process for metal-oxide resistors Production of homogeneous slurry by wet-mixing of oxide powders Drying and granulation in a spray dryer Compacting the granulate to form resistor blocks Sintering to obtain dense ceramic bodies Addition of electrical contacts and a protective coating Electrical testing MO resistors ready for assembly
The basic material used to manufacture metal-oxide resistors is very fine grain ZnO with particle sizes of about 1 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 the water 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
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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 and applying 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 VARISTORS Before 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 with production 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 BLOCKS Early 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 the “electrical” microstructure of a MOvaristor [Gre 1989]
Figure 2.9: Band diagram of a single grain boundary, showing the Double Schottky Barriers DSB formed by charge trapping in interface states [Gre 1989] (dashdotted/dotted lines: Fermi or quasi Fermi level [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. It 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, E F : 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 of the potential barrier B(U) and eventually even -values 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 2 increase and stabilize B(U) at a high level (note that: B(U) ~ Qi /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 literature 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 state physics 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)·d B/dU, i.e. U and d B/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 17 3 side of the grain boundary. For sufficiently high doping of the bulk of the ZnO ( N0 > ca. 10 /cm ) 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
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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 switchingor 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 breakdown mechanism [Gre 1989]
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Figure 2.11: Electroluminescence observed under the light microscope [Gre1990], [Gre 1998]. A surface contact geometry is used with a contact at the top and bottom of the picture. Each grain boundary lights up by a short line roughly perpendicular to the current filament connecting both electrodes. Some filaments can be seen to disappear under the surface and to reappear closer to the other surface electrode. The material is stressed in the breakdown region U> U B
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 achieved for 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
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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 distribution of current in a random network containing varistor-type grain boundaries for different positions on the U/I-characteristics: a) ohmic region with a homogeneous distribution of the current density (V=1.5), b) in the breakdown with clear filamentary conduction (V=3.5), c) higher up in the breakdown region at the transition to the upturn (V=4.3) and d) in the upturn-region where the current distribution reverts to a more homogeneous situation (V=6) [Bar 1996]. The gray-level represents the relative value of the current through a grain boundary, normalized to the total current.
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Figure 2.13: Infrared thermal images taken with a high speed camera for thin slices of low voltage varistors with 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 ) the current flow is rather homogeneous throughout the structure, whereas in the high nonlinear regime strong 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.
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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 BLOCKS In 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 nonhomogeneities, which can be present in a non-perfect varistor block [Ste 2004].
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Figure 2.14: Left: Cross section along the axis of a station arrester block (Ø 63 mm), viewed by a fast IR-camera upon exposure to a square wave impulse of a few ms duration and a current amplitude of 100-200 A [Ste 2004]. Right: light intensity profile proportional to the temperature, which in turn is proportional to the current flowing through the block. For a complete simulation of the failure behavior, the electrical, thermal and mechanical properties of an inhomogeneous varistor block have to be calculated in 3 dimensions in a coupled mode. This is a very demanding task and has not yet been done with finite element methods. Parts of the problem however have been addressed and are illustrated in Figure 2.15 and Figure 2.16. Figure 2.15 shows the coupled thermal electrical simulation for the 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 high local 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 ceramic locally is heated to above the melting point (> 750-850 C), where the melt can be progressively ejected, starting from the surface of the hot spot. The local heating simultaneously also creates thermo-mechanical stresses and, if these 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, if sufficient energy is deposited in the channel. Both phenomena, puncturing and cracking under long duration impulse, hence have the same origin, but the outcome will depend on such parameters as the local energy input rate, 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 energy testing 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 and tangential stresses were then calculated self consistently on the basis of a theoretical U/I-characteristic for different current densities and two different block sizes. Although the assumed inhomogeneity is rather high compared to today’s high quality blocks (and for symmetry reasons was located in the center of the block), the model predicts quite well the regimes for the most likely failure modes to occur: thermal runaway, puncture, cracking under tension and cracking under compression. The results for a distribution arrester disk are illustrated in Fig. 2.16. Note that depending on the range of current density, different failure modes can show up on a block.
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Figure 2.15: Equithermal contour plot for an edge margin of 2 mm of a varistor block as obtained from coupled thermal electrical simulations [And 2000].
Figure 2.16: a) Time to failure t f and b) energy handling capability of a distribution-class type of MO disk; the different limiting failure modes are shown [Bar 1996]. For the simpler case, where the temperature field across the block is known, the mechanical stresses have been calculated by Nied [Nie 1989]. Despite the simplifications made the work from Bartkowiak et.al. provides a rather
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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 10 7 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 (c 4-5000 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: 1st principle stresses for 1D and 3D model [Len 2000], right: typical “midplane” crack under high current impulse for a distribution 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 st st maximum 1 principle stress is normal to the axis, while for a high aspect ratio the maximum 1 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
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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]. Poor control 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 VARISTORS In 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 a reducing 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 N 2 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.
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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 Bi 2O3 [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.
1.4
1.2
1.2
power loss ratio P / Po
pow er loss ratio P / Po
1
0.8
1
0.8
0.6
0.6
stable
0.4
unstable
stable varistor in air or N2 unstable varistor in air unstable varistor in N2
0.4
0.2
0.2
0
0 0
100
200
300
400
500
600
700
800
900
1000
0
Figure 2.18: Power loss ratio vs. time for stable and unstable varistors during accelerated ageing tests at 115°C and slightly elevated AC operating voltage.
100
200
300
400
500
600
700
800
900
1000
time [h]
time [h]
Figure 2.19: Power loss ratio vs. time depending on the surrounding medium (air, N 2 ) during accelerated ageing test at 115 °C and continuous AC voltage.
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 T/10) factor AF was estimated with AFT = 2.5 , 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).
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Upper limit of ambient temperature
Test duration at 115 °C
°C 40 65 95
h 1000 2000 7000
Equivalent time at upper limit of ambient temperature Years 110 22 5
Table 2.2: Test duration and equivalent time at upper limit of ambient temperature according IEC 60099-4 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 will 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 recently been 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 field 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.
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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 will 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 ACbehavior. 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.
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2.3 Design of surge arresters Authors in charge: Volker Hinrichsen and Bernhard Richter
2.3.1 FOREWORD Latest 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 capabilities, 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 subsolutions 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 composite hollow 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 role so 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 subtransmission 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 STRUCTURE Figure 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
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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.
Porcelain; Polymer Type A
Polymer Type B1b
Polymer Type B1a
Polymer Type B2
MO column
Gas
FRP supporting structure
Solid/semi-solid material
Outer housing
Metal end fittings
Type A: gas volume enclosed, separate sealing system, pressure relief vents Type 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
Figure 2.20: Classification of basic arrester designs; cross sectional side and top views
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Top cover plate Flange with venting outlet Sealing ring Pressure relief diaphragm Compression spring
MO resistor column Composite hollow core insulator (FRP tube/ rubber sheds)
Figure 2.21: 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. Some of the potential benefits of the Type A design are extremely high mechanical strength, the safest possible shortcircuit 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 differences.
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MO column
FRP wrap
main orientation of glass fibers
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 preimpregnated 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.
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The resulting gap between MO and FRP material is then filled by solid or semi-solid material. Again, slots may be provided 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 metal end fittings by FRP loops or rods, applying an enormous axial pre-stress in the range of 100 kN [Ste 2003]. The basic mechanical design is shown in Figure 2.23. The left variant is designed with loops, which are fixed in notches in the end flanges. This design was first introduced for distribution arresters [Sch 1996] and then extended to HV. A sub-variant (Figure 2.23, middle) uses an additional bondage from polymeric material in order to achieve the mechanical and short-circuit characteristics required for application in HV and EHV [Sky 2002]. Also the other variant of Type B2 arresters (Figure 2.23, right) was first realized in distribution arresters and then further developed for HV. Here, FRP rods are applied, which are mechanically fixed in holes in the end flanges by a proprietary clamping system. Figure 2.24 shows photographs of real arrester modules as they are produced by two different manufacturers. Main technical advantages of the cage designs are that they offer comparatively high mechanical strength combined with an inherently good short-circuit performance.
Sealing
Bonded by molding process
Gap filled by solid material
Figure 2.25: Possible implementations of sealing
Figure 2.24: Internal mechanical structure of Type B arrester Left: loop design, Right: rod design
With regard to the commercial aspects it is impossible to give any statement here on production cost, which greatly depends on the total manufactured quantities, the degree of automation and process optimization, the quality of the applied materials, the degree of type diversification and so on. However, it can be noticed that the Type B arrester in general constitutes the most economical way to produce an arrester. At the same time it offers a technical performance, which in most cases ranks between comparable porcelain and polymer housed Type A arresters. The Type B arrester is therefore the typical "low cost" arrester mentioned earlier, one of the reasons for the success of this design on the market.
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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 individually 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. The alternative 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 ARRESTERS For 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 few designs 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.
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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 voltage arresters Left: design group I, middle: design group II and III, right: example for design group I before molding in silicon
2.3.6 CONCLUSION Polymer housed HV arresters actually have a market share of roughly 30%. First installations have been made around 1990, and so far there is no indication that they will not show the same good performance as MO arresters in general or polymer housed distribution arresters, which have successfully been in service for nearly 20 years now. The technology can be considered mature. Thus it can be predicted that the share of polymer housed arresters in HV will continuously increase, because they usually offer economical as well as technical advantages over 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 designs usually 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 "price orientated" 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 the domain 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 A designs. Especially with the latter nearly any required mechanical, electrical or safety feature can be achieved, which has its price, however.
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2.4 Special designs of surge arresters Author in charge: Roger Perkins
2.4.1 SEPARABLE AND DEAD FRONT ARRESTERS
2.4.1.1General Comments These 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 to be 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 Designs Figure 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 arrester. In this example the connector is also equipped with load-breaking capability.
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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 Conditions The characteristics of separable and dead front arrester are fundamentally the same as those for any other metaloxide 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 time. 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 for separable 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.
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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 4/10 s may be considered adequate where an overhead system requires 65kA 4/10 s. 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 Comments This 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 failure detection.
Figure 2.28: Appearance and internal design of a typical under-oil metal-oxide 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 96
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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 Conditions Special 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 19 th 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 will 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.
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2.5 SF6 gas insulated MO surge arresters Author in charge: Bernhard Richter 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 installation of a GIS arrester at the line entrance of a gas insulated substation (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 1 reduced self-inductance . 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 98
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insulating plates MO resistors current path
Figure 2.30: Design of the active part (mechanically three columns, electrically one column) of an EHV GIS 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 flowing, 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].
insulator electrical connection, high voltage pressure relief device grading hood MO column vessel ground plate
Figure 2.31: Design of GIS arresters (principle) for different system voltages. Left: single phase design for system voltage up to 170 kV. Middle: Three phase design up to 170 kV. Right: single phase design for system voltages above 170 kV with electrical one phase, but mechanically three columns of MO resistors, courtesy Siemens.
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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 is reduced drastically, which is an important argument in the today’s discussion about greenhouse gasses.
Figure 2.32: GIS arresters for 550 kV systems with MO resistors with “normal” field strength (left) and high 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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MO resistors of 200 V/mm
MO resistors of 400 V/mm
Figure 2.33: GIS arresters for 550 kV systems with MO resistors with “normal” field strength of 200 V/mm and 400 V/mm in a mechanical meandering design, courtesy Mitsubishi.
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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 in 1991. There were two characteristic developments in this period. The first was the polymer-housed arrester made possible by the use of improved fiberglass-reinforced components that could be reliably used as structural support of the device when a polymeric housing should be used that in itself could not generate this capability whilst still providing safe pressure relief in the event of failure. The second major development was the increasing performance and reliability of the metal-oxide resistor itself. A characteristic consequence of these facts is the emergence during this period of continually more complex devices with a greater degree of integration with other electrical transmission and distribution installations; in other words devices or apparatus with more than one primary function or capability. Documented examples of these that incorporate 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 more important requirements. Improved reliability, improved overvoltage protection, improved environmental protection have all been claimed as well. However, most likely the major benefits have still to be demonstrated, since the increasing 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 effectively test a device with multiple, sometimes interrelated functionality. These and other special considerations have been reviewed 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 as suspension insulator/line arrester in a medium voltage trial line 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 center break disconnector. System voltage 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 Reinhard
2
3.1 Summery This 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 next revisions 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 MOarresters. 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 find that this is obviously not the case. There is still 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 Page 104
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standards were published. Till now, the standards do reflect this situation. In terms of energy handling issues, one still recognizes historical approaches from the gapped SiC arrester era. It is, therefore, the time to think about new definitions of energy handling capability and about revised, appropriate test procedures, based on actual knowledge and most recent findings. Besides others, this was the task given to Cigré WG A3.17: to check the actual literature and standards about energy handling issues, to contribute to some of the open questions by a comprehensive practical research program, and to work on proposals for energy handling issues in future revisions of 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 reflected in 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 approximately uniform temperature distribution in the MO resistors, o "repeated" impulse stress, where the time interval between impulses is sufficiently long to obtain cooling of the MO resistors close to their initial temperature (this includes durability and degradation aspects).
3.3.1 THERMAL ENERGY HANDLING CAPABILITY Thermal energy handling capability can only be considered for complete arresters, as besides the MO material properties it is mainly affected by the heat dissipation capability of the overall arrester design. The situation is schematically depicted in Figure 3.1.
Figure 3.1: About thermal energy handling capability 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 increases non-linearly but moderately with the temperature difference to ambiance. Electrical power losses under normal operating conditions are usually very small, in the range of tens of milliwatts per kilovolt of rated voltage for distribution 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 stable operating 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 MO resistors 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
W V
2,59
J J 0, 0044 cm³ K² cm³ K
(equation 3.1)
²
where W is the contained energy in J, V is the MO volume in cm³ and dependence is shown in Figure 3.2.
is the MO temperature in °C. This
Energy per volume in J/cm³ Energie/Volumen in J/cm³
900 800 700 600 500 400 300 200 100 0 0
10
20
30
40
50
60
70
80
90 100 110 120 130 140 150 160 170 180 190 200 210 220 230
Temperature Temperaturin in °C °C
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
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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. by current impulse stress, must be considered by appropriate conditioning procedures.
3.3.2 IMPULSE ENERGY HANDLING CAPABILITY At first glance, impulse energy handling capability may easily be defined as well: it is just the energy, which is injected into the arrester by one single impulse, and if a limit value is exceeded, one or several MO resistors will mechanically fail as exemplarily shown in Figure 3.3, finally leading to an overall arrester failure.
Figure 3.3: Examples of MO resistors, mechanically failed by single impulse energy overload. Left: failed by thermo-mechanical cracking; Right: failed by flashover of the coating. 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 only some small punctures in or at the edge of the metallization, but furthermore the MO resistor may look intact. In other cases, no damage at all might be seen by a visual examination, but the MO resistor is pre-damaged and will not pass any further energy input. Or its electrical characteristic may be dramatically changed such that if this happened 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 operating voltage, since otherwise the arrester would suffer a thermal runaway even if its MO resistors were able to handle the 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 statistical approach is more appropriate, in the same way as for the dielectric strength definition of external insulation (in which case "withstand" voltage stands for a "10 % flashover probability" voltage). Verification of a withstand energy is 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 be demonstrated by the following example. An arrester for a 420 kV system is made up from approximately n = 65 MO resistors. If each resistor has a failure probability of 0,1 % (p = 0,001) at its "rated" energy, the full arrester, at the n same rating, will have a failure probability of P = 1 – (1 – p) = 0,063 or roughly 6 %, respectively. Higher failure probabilities for the full arrester are hardly acceptable! Vice versa, if the full arrester shall have a failure probability of only 1 % (P = 0,01) at its "rated" energy the individual MO resistors in this case must have a failure probability 1/n -3 p = 1 – (1 – P) = 0,155·10 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 be explained 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 LITERATURE 3
In a comparison between SiC and ZnO 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) mm height, 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. To date, one usually speaks exclusively of "MO" arresters and resistors in this context. Page 108
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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 the metallization, 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 extreme forces 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 approximately 1 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 4 Voronoi network [Bar 1996b] [Bar 1006c]. The network is made up from three components: "good" grain boundaries 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 are applied 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 were multiple 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 investigated recovery 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 difficult 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 111
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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 electroquasistatic 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 A3.17 When 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 currents Commercially 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.
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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 ambient temperature. 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 introduced since 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.
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Initial measurements Uch,1 at Jch = 0,12 mA/cm² (after 5 s) Ures,1 at I = In
Impulse test (energy injection)
Visual inspection: mechanically failed?
yes
no Measurement of characteristic voltage Uch,2 at Jch = 0,12 mA/cm² (after 5 s)
Exit measurements
95% U ch,1
105% Uch,1 ?
Uch,2
no
yes Measurements at lightning current impulse Ures,2 at I = In Imd at J = 1,5 kA/cm²
Visual inspection: mechanically failed?
yes
no no
95% Ures,1
105% Ures,1 ?
Ures,2
yes
OK Figure 3.4: Flowchart of the test and evaluation procedure
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defect
MO Surge Arresters-Stresses and Test Procedures
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-Icharacteristic. 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
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solved in the near future. Actually, MO resistors have not been optimized for the extreme current stress of the lightning current test as specified in Annex N of [IEC 2009].
Figure 3.5: Mean failure energy vs. amplitude of current density for “Size 1” MO resistors For make "T", tests at long-duration current impulse stress were also performed up to mechanical failure (rather than to apply the "complex failure criterion"). It can be seen from Figure 3.6 that in this case the mean failure energy follows the expected dependence of current density. This demonstrates that the differences in Figure 3.5 between the a.c. and the impulse current tests are not related to the different current shapes but only to the different applied failure criteria. It further allows concluding that energy handling test can also be made with alternating current stress if this is considered more convenient. The only difference to impulse current testing will then be the lower failure energies due to the lower achievable current densities, which, however, can easily be taken into account by correction factors.
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1800 Diameter 60 mm Heigth 40..45 mm
1600
mean failure energy in J/cm³
1400 1 ms 1200
2 ms 100 ms 4 ms
1000 8s
T, complex failure criterion
800
Ringler 97, until mechanical failure 600
T, until mechanical failure 90/200 µs
400 AC 200 0 0,1
1
10
100
1000
10000
peak current density in A/cm²
Figure 3.6: Mean failure energy vs. amplitude of current density for “Size 1” MO resistor of make T; comparison of failure criteria “until mechanical failure” and “complex” for the 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.
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Figure 3.7: Mean values of current density amplitude vs. time to failure
Figure 3.8: Mean failure energy vs. amplitude of current density for “Size 2“ MO resistors
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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 with those 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.
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100
T
100
80
80
S
60
60
%
%
40
40
20 0 CR BR
FO ÜB
PU DU
MF Uch Uref Ures Ures
20
4 ms 2 ms 1 ms 90/200 µs
4 ms 2 ms 1 ms 90/200 µs
0 CR
FO
PU
MF Uch Ures
V 100
100 80 %
80
U
60
60
% 40
40 4 ms
20
4 ms 2 ms 1 ms 90/200 µs
20
2 ms
0
0
1 ms
CR BR
FO ÜB
DU PU
CR FO
90/200 µs
MF
Uch Uref Ures Ures
100
100
PU
MF Uch Ures
Z
80
80
60
X
60
%
%
40
40 4 ms 2 ms 1 ms 90/200 µs
20 0 CR BR
FO ÜB
20
DU PU MF Uch Uref Ures
0 CR
FO
PU MF Uch Ures
4 ms 2 ms 1 ms 85/180 µs
Failure mechanisms: CR … Cracking FO … Flashover PU … Puncture
MF … Mechanical failure during exit measurement Uch ... Change of "characteristic" voltage Ures... Change of residual voltage
Figure 3.9: Failure mechanisms of “Size 1” MO resistors
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100
100
80
80
60
S
%
60
%
U
40
40
20 4 ms 1 ms 4/10 µs
Uch Ures Uref Ures
FO
1 ms PU
MF
Uch
U ref
MF
4/10 µs Ures
PU DU
MF
FO ÜB
ÜB
BR CR
CR
BR
0
4 ms
0 DU
20
Ures
100 100
80 80
V
60 %
60
40
W
% 40
20 0 CR BR
ÜB FO
DU PU MF Uref Uch Ures Ures
4 ms 1 ms 4/10 µs
20 0 CR BR
X
ÜB FO
DU PU
MF Uref Uch Ures
4 ms 1 ms 4/10 µs
100
100
80
80
60 60
Y
%
%
40
40 4 ms 2 ms 1 ms 90/200 µs 4/10 µs
20 0 CR
FO
PU
MF Uch Ures
20 2 ms
0 BR CR
4/10 µs ÜB FO
DU PU
MF
Uref Uch Ures Ures
Failure mechanisms: CR … Cracking FO … Flashover PU … Puncture
MF … Mechanical failure during exit measurement Uch ... Change of "characteristic" voltage Change of residual voltage Ures...
Figure 3.10: Failure mechanisms of “Size 2” MO resistors
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Figure 3.11: Failure mechanisms (from left) “cracking” (CR), “flashover” (FO), “puncture” (PU) and “flashover” (FO) in the special case of 90/200 impulse current stress For the "Size 1" MO resistors, main failure mechanism of make "S" is change of the characteristic voltage and cracking; only at the 90/200 µs impulse, it mainly failed by flashovers. For make "X" the dominating failure mechanism is cracking, and even in case of the 90/200 µs impulse it is much more the change of the characteristic voltage 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 important observation is that change of the residual voltage (by more than 5 %) is not a concern at all. This criterion is often used 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 and flashover. This is not only the case for 4/10 µs and 90/200 µs, where this might have been expected, but also for the long duration current impulse stress in one example. Particularly for the "Size 2" resistors one can identify a typical 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 injected energy. Current peak value in A 5 40.000 0
60.000
80.000
100.000
120.000
140.000
160.000
180.000
200.000
220.000
Change of U ch in %
-5 -10 S U V W X Y
-15 -20 -25 -30 -35 -40
Figure 3.12: Change of characteristic voltage vs. 4/10 µs impulse current peak value
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Energy in J/cm³ 5 0
100
200
300
400
500
600
700
800
900
1.000
0
Change of U ch in %
-5 -10 S U V W x Y
-15 -20 -25 -30
"Size 2" Diameter 40 mm Height (30..40)
-35 -40
Figure 3.13: Change of characteristic voltage 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.
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140
350 Temperaturanstieg in K Temperature increase Current in kA Strom Energy in J/cm³ Energie
120
300
100
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80
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50
0
Energie in in J/cm³ J/cm³ Energy
Temperaturanstieg in kAin kA Temperature increaseininKK/ Strom / Current
MO Surge Arresters-Stresses and Test Procedures
0 -2
0
2
4
6
8
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14
Zeit in µs µs Time
7000
140
6000
120
5000
100
4000
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60 Kraft Forcein N Strom Current in kA
2000
Strom in Current inkA kA
Force kN Kraft in N
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
40
1000
20
0
0 -10
0
10
20
30
40
50
60
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Zeit in Time inµs µs
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
c
E
100 GPa 5420 kg/m³
4300
m s
4, 3
mm µs
(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.
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120
Strom in Current in kA kA
100 80 60 40 20 0 -20 -5
0
5
10
15
20
25
30
35
40
35
40
9 17,3 µs
Force in kN kN Kraft in
8
9,1 µs
Höhe 37 Height 37 mm mm Height 27,8 mm mm Höhe 27,8
7
11,6 µs
6 5 4 6,3 µs
3 -5
0
5
10
15
20
25
30
Time µs Zeit ininµs
Figure 3.16: Current and force measured on MO resistors of different 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 alternating current 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.
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99,8 99,5 99 98
95
95
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Wahrscheinlichkeit in %
99,8 99,5 99 98
80 70 60 50 40 30 20
Probability in %
Probability in %
Wahrscheinlichkeit in %
MO Surge Arresters-Stresses and Test Procedures
Meßpunkte Measuring
points
Treppenkurve Step curve Verteilungsfunktion function Distribution Vertrauensbereiche Verteilung Conf. interv. distribution Vertrauensbereiche Quantile Conf. interv. quantile
80 70 60 50 40 30 20 10
10 5
5
Verteilungsfunktion function Distribution
2 1 0,5
Verteilung Vertrauensbereiche Conf. interv. distribution Vertrauensbereiche Quantile Conf. interv. quantile
0,2 1470
1570
1670
1770
2 1 0,5 0,2
1870
810
860
910
Energie in in J/cm³ Energy J/cm³
960
1010
1060
Energiein in J/cm³ Energy J/cm³
Figure 3.17: Examples of statistical evaluations (Normal Distribution) of impulse current tests (left) and alternating current tests (right) with 95 % confidence intervals 1,8
Mean failure energy in p.u.
1,6 1,4 1,2 1 0,8 0,6 0,4 0,2
Ø
6 cm
Ø
8 cm
Ø
10 cm
0 0
2
4
6
8
10
12
Amplitude of current density in A/cm²
Figure 3.18: Mean failure energy (incl. standard deviation) vs. current density amplitude for MO resistors of same make and same height but different 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.
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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 a probability p(W') of failure at the same specific energy is
P W'
1
1 p W'
n
(equation 3.3)
Solving equation 3.3 to p(W') leads to
p W'
1
n
1 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 difficult 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 longduration 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).
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1200
Energie in J/cm³ Failure energy in J/cm³
1000
800
600
400
200
0 1
5
9
13
17
21 25 29 33 37 Test and sample number Versuchsnummer
41
45
49
53
Figure 3.19: Example of failure energies at alternating current tests up to mechanical failures 3.4.2.5 Conclusions and Outlook In 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 MO resistors 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 a pass 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 time in this study. Flashover or significant alterations in the
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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 on topics 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 first 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 longduration 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
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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 lightduty station arresters (Cl. 8.5.4 of [IEC 2009]) represent a considerable energy stress for this class of arresters, but are 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-theart 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 help of 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
Surge impedance of the line Z, in
Virtual duration of peak T, in µs
Charging voltage UL, in kV (d.c.)
1
4.9 · Ur
2000
3.2 · Ur
2
2.4 · Ur
2000
3.2 · Ur
3
1.3 · Ur
2400
2.8 · Ur
4
0.8 · Ur
2800
2.6 · Ur
5
0.5 · Ur
3200
2.4 · Ur
Ur = 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]
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LD class
U s (kV)
1
245
2
300
3
420
4
550
5
800
Table 3.2: Recommended line discharge classes depending on system voltage
MO Surge Arresters-Stresses and Test Procedures
Figure 3.20: Specific energy injection (by two consecutive line discharges) during switching impulse operating duty test in kJ/kV of rated voltage dependent on the ratio of switching impulse residual voltage U res to the r.m.s. value of the rated voltage U r of the arrester (in accordance with [IEC 2009]) As no direct conclusions about the energy stress can be drawn from this table, Figure 3.20 depicts the converted energy in a test object during the switching impulse operating duty test (injected by two line discharges), with reference to its rated voltage. This energy is not a fixed value, but depends on the arrester's switching impulse residual 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 is applied with a given amount of specific thermal energy handling capability, then the arrester can, depending on its actual residual voltage, be assigned to different LD classes, as shown in the following example (red dashed lines in Figure 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 the apparently "better" LD 3 arrester might possibly be worse for the intended application, since its protective level is higher. In order to reach LD 3 while maintaining a ratio of Ures/Ur = 2, a design must be used with a thermal energy handling capability of almost 6 kJ/kV, as indicated by the blue dotted lines in Figure 3.20, that would mean application of MO resistors with greater diameters. Inversely, one can only draw conclusions from the LD class in connection with the residual voltage as to the (thermal) energy handling capability of an arrester, and thus about the used MO resistors. For standard applications, one can simply count on recommendations in the application guide [IEC 2000], based on the 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 the demands of the extra-high-voltage systems with Us > 550 kV. In fact, at this voltage level, and sometimes even at the 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 kV d.c. and 1100/1200 kV a.c. applications, where specific energy handling values of (25…50) kJ/kV of rated voltage will be required at switching impulse protective levels in the range of only 1,85 or even less [Ric 2007]. It can easily be seen from Figure 3.20 that LD 5 is by far not sufficient for these applications: an LD 5 operating duty test will inject 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 discussed if other test parameters, such as the time duration of the impulses, may be changed. Summarizing, at least three work items can be identified for a future revision of the standard: the possible replacement of the LD system by a purely energy based rating system, clear definitions and differentiation of different kinds of energy handling capability, and specification of higher energy handling values than today and the related test procedures.
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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 values specified 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-currentcharacteristics (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.
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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 powerfrequency 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 induced mechanical 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 be specified 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 compromise to 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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impulse current 85/180 (kA)
16 14 12 10 8 6 4 2 0 LD 2 590 A / 2000 µs
LD 3 720 A / 2400 µs
LD 4 880 A / 2800 µs
Figure 3.21: Required 90/200 impulse current amplitudes for injecting the same energy into an MO resistor 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
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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 nonstandardized 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 covers gapless 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]. 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
Station 15 kA
20 kA, LD 4 or 5
Station 20 kA
20 kA, LD 4 or 5
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 of 40 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:
6
-
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;
Voltage that causes the same watt losses in the actual MO resistors as a voltage equal to MCOV would do in aged resistors of the same make and of the highest specified watt losses. Page 136
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-
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 values resulting 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 to only 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.
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3.5.4 ENERGY HANDLING ISSUES IN OTHER NATIONAL STANDARDS When 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 JEC2371 [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 1,5kA 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 991A-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 OUTLOOK It 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 first 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. Summ ary Author in charge: Bernhard Richter MO arresters are nowadays installed in all kind of electrical power systems from low voltage up to UHV. They are intended to protect equipment and installations against overvoltages. Due to the various applications the MO arresters have to withstand severe stresses from the electrical system, from lightning and from ambient. In this new TB the different types of stresses are listed and severe stresses, e.g. winter lightning, seismic stresses and severe pollution 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 ambient stresses, 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 with international participation on energy handling withstand capability of MO resistors and arresters. The results show that 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 modes depending 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 times higher than the design energy of the MO arresters as proved in the relevant operating duty tests. This gives a good safety margin and confidence in the todays designs and materials. However, with increasing system voltage the number of MO resistors easily reaches several hundreds of blocks per arrester, and then statistical evaluations may have to be considered. For a more sensitive evaluation of the MO resistors in this study, a new complex failure criterion was developed and used, including the change of the electrical characteristic of the MO resistors in addition to the simple failure by cracking or flashover. The different types of failure modes depending on the different current wave shapes can lead to further improvements 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 same concept 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 single high voltage equipment against overvoltages. This is especially the case for UHV a.c. and d.c. systems. The development of MO resistors and arresters is ongoing with the goal of size and cost reduction but in the same time keeping the high quality and reliability. This leads to the development and use of MO resistors with increased field strength to reduce the size of the complete design, e.g. in GIS applications. Further on, with the very tall arresters for 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 capability of MO resistors and arresters are dealt with in working group A3.25 of SC A3 and will be published in a separate Technical Brochure.
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APPENDIX 1 Following Technical Brochures of Cigré are dealing with surge arresters and their application: TB 60 Metal Oxide Arresters in AC Systems by WG 06 of SC 33, 1991 TB 287 Protection of MV and LV Networks against Lightning. Part 1: Common Topics by CIGRE-CIRED JWG C4.4.02, 2006 TB 441 Protection of MV and LV Networks against Lightning. Part 2: Lightning protection of Medium Voltage Networks by CIGRE-CIRED JWG C4.4.02, 2010 TB (XXX) Protection of MV and LV Networks against Lightning. Part 3: Lightning protection of Low-Voltage Networks by CIGRE WG C4.408, to be published 2013 TB 440 Use of Surge Arresters for Lightning Protection of Transmission Lines by CIGRE WG C4.301, 2012 TB 455 Aspects for the Application of Composite Insulators to High Voltage ( 72 kV) Apparatus by CIGRE WG A3.21, 2011 This TB addresses the special case of surge arresters with composite insulators.
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