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  1. 1. CHAPTER 6 LIGHTNING PERFORMANCE OF COMPACT LINES John G. Anderson William A. Chisholm Andrew Phillips _____________________________________________________________________ Abstract: This chapter reviews issues related to the occurrence of lightning on compact transmission lines, including the types of lightning flashover that can occur, the factors that govern flashover, and the performance of shielded and unshielded compact lines. John G. Anderson is one of the original authors of the EPRI Transmission Line Reference Book, including the chapter on line lightning performance. He has had more than 50 years of high-voltage engineering experience, and is a Life Fellow of the IEEE and an elected member of the National Academy of Engineering. He is a former manager of the General Electric High Voltage Laboratory and served as a consulting engineer for General Electric and also as a senior consultant for Power Technologies, Inc. He was one of the original researchers at the Lenox, Massachusetts EPRI Project EHV and later Project UHV, and also carried out lightning research at the Empire State Building in New York City. He is the author/coauthor of more than 40 technical papers and coauthor of three books concerned with high-voltage transmission, lightning, and insulation performance. Dr. William A. Chisholm is an internationally acknowledged expert in lightning protection, insulation, and thermal rating of power systems. He is a Senior Research Project Manager in the Transmission and Distribution Technologies group of Kinectrics, the former Research Division of Ontario Hydro, now a division of AEA Technologies PLC. In this capacity he has completed research contract and project work for more than 40 electric utilities, manufacturers and research organizations. Dr. Chisholm was recognized as the editor of the “Best Standard” award in 1999 for the IEEE Standard 1243, Guide to Improving the Lightning Performance of Overhead Transmission Lines. He is a corresponding member of CIGRÉ Study Committee 33 working groups on lightning and insulator icing test methods. He is the chairman of the IEEE Power Engineering Society Lightning and Insulator Subcommittee and a member of the PES Editorial Board. Dr. Andrew Phillips is a Senior Program Manager in the Transmission Lines and Increased Power Flow Research program area of the Power Delivery and Markets Sector. His current research activities focus on the Overhead Transmission, Underground Transmission, Increased Power Flow, and HVDC (high-voltage direct current) programs. Dr. Phillips’ special areas of interest are nonceramic insulators (NCI), lightning and grounding, inspection and assessment of components, sensor development, and daytime corona inspection. Before joining EPRI, Dr. Phillips worked at J. A. Jones Power Delivery, where he was a Project Manager and lead researcher in the fields of insulation, aging equipment, and lightning. Prior to that, Dr. Phillips worked at the University of the Witwatersrand performing research for the South African electric power industry. 1
  2. 2. Dr. Phillips received BSc, MSc, and PhD degrees in electrical engineering from the University of the Witwatersrand in Johannesburg, South Africa. Dr. Phillips holds three U.S. Patents and is the author of over sixty journal and conference publications. 2
  3. 3. 6.1 INTRODUCTION Compaction of transmission lines can have significant benefits to lightning performance in many cases. If one or more overhead shield wires exist, a reduced spacing to nearby phases creates increased coupling, resulting in reduced insulator voltages. Also reduced structure height makes any compact line a smaller lightning target. Finally, the shorter spans of most compact lines—when accompanied by shield wires—can also reduce the voltage challenges to insulators. Conversely, possible reductions in insulator lengths and clearances increase the risk of flashover unless appropriate countermeasures are taken, and elimination of shield wires in the compaction process exposes the phases to more direct lightning hits. The technology involved in voltage upgrading of lines is similar to that of compaction. However, the risks that might be acceptable in voltage upgrading are driven by economic issues that may not exist in new line designs. Thus reductions in design tolerances in an upgrading would not be considered acceptable for a new line. For example, eliminating shield wires in a voltage upgrade of an existing line might be beneficial in order to minimize the outage time and reuse of the structures and make or break the upgrading. There may also be times when construction of a new compact line without shield wires is desirable, perhaps to make it look like a distribution line or to fit in an overall pole height. Lightning phenomena are no different for compact lines, and the same computer design programs are used. Compaction may improve or degrade performance, and this chapter concentrates on the beneficial aspects. Lightning performance is part of the design process with its own set of tradeoffs. The principal strategies reviewed in this chapter for ensuring good lightning performance of compact lines include: • Minimizing support structure footing resistance. • Increasing line insulation. • Installing shield wires to prevent lightning hits to the phases. • Installing transmission line surge arresters (TLSA) to prevent excessive phase overvoltages during lightning hits. • Installing underbuilt shield wire(s). Estimating improvements in transmission line lightning performance gained by any of these strategies is an evolving art, limited in accuracy by annual variations of local weather patterns, as well as by changes in lightning flash magnitudes and ground flash densities from year to year. If detailed and reliable data are available, median flashover rates can be calculated fairly well by line lightning performance prediction software such as EPRI TFlash (EPRI 2005b). These software programs generally use median lightning ground flash density (GFD) data that provide estimated median flashover rates for the line designer. However, extreme value flashover probability estimates based on year-toyear variations are not yet reported by most programs. This chapter does not discuss the fine details of line backflashover voltages, leader progression, grounding nonlinearities, counterpoise propagation, volt-time curves, and 3
  4. 4. other arcane subjects related to the technology. These topics are explained in many references (EPRI 1982; IEEE 1997b; EPRI 2005a), and will not be repeated here. Section 6.2 of this chapter starts by defining and reviewing the various types of lightning flashovers that occur on transmission lines and the importance of each. By their smaller size, a compact line will offer a smaller target to lightning than will a conventional line, but the act of compaction minimizes wire spacings wherever possible, and smaller spacings can then increase risk of flashover for the fewer lightning incidents predicted to occur. Section 6.3 details the major factors that govern when flashovers will occur, starting with the mean number of lightning hits per year that can be expected, the expected characteristics of the lightning flashes that will challenge the line’s insulation, the mechanisms of shielding failures, midspan flashovers and backflashovers, and the influence of ground resistances on the latter. Alternative methods of estimating line lightning performance are briefly covered. Section 6.4 reviews performance of compact transmission lines having shield wires, stressing the benefits of proper grounding and providing numeric examples of the effects of various footing resistance magnitudes and the number of overhead or underbuilt shield wires. Significant economic and compaction benefits can result from omitting shield wires. Section 6.5 examines the pros and cons of this option. Unless transmission line surge arresters (TLSA) are properly selected and connected to vulnerable phases, each lightning hit to these unshielded lines will cause one or more flashovers. Thus care must be taken in selection of arrester locations and selection of arrester energy capabilities, otherwise excessive maintenance and reduced line reliability may result. For line designers who do not have access to line lightning simulation programs, such as EPRI TFlash, Section 6.6 suggests some alternative methods of estimating arrester energy challenges on unshielded compact lines. The electrical effects can be calculated using one of the Electromagnetic Transients Programs (EMTP, ATP or equivalent), but all the associated failure probabilities must be computed separately. Appendix 6.1 expands these procedures. Appendix 6.2 covers possible acceptance tests on qualitycontrol samples of proposed TLSA. Section 6.7 summarizes important points in the chapter. 6.2 LIGHTNING FLASHOVER ISSUES This section briefly describes the characteristics of lightning flashovers in a simple form, the factors that influence each of the mechanisms and their relevance to compact line design. Induced Flashovers occur when a ground flash terminates near, but not directly striking, an energized phase conductor. The intense electromagnetic field from the return stroke current causes large, short-duration overvoltages that can be calculated using (Rusck 1958; Agrawal et al. 1980; Chowdhuri 1996; Rachidi et al.1997). A simplified model has recently been described by (Darveniza 2007). Influencing factors in induced flashovers relate to the distance between the stroke and the line, the effective line height of the phases, and the insulation strength. Like all lightning flashovers, the number of induced flashovers will scale linearly with local ground flash density. Experience has shown that compact lines with insulation strengths of 400 kV BIL or higher and conductor-to-tower clearances of 0.5 to 0.8 m will be largely immune to induced flashovers. 4
  5. 5. Shielding Failures occur when lightning terminates directly on an energized phase conductor. Lightning current is injected directly into the phase conductor. Typically, at least one stroke of a random flash has sufficient current and a fast rate of current rise, together with a high conductor surge impedance, to cause a local overvoltage that may exceed the lightning impulse strength of the struck phase insulator or the dielectric strength of air gap between phases and results in a flashover to ground or mid-span to a nearby phase. Influencing factors in shielding failures relate mainly to the parameters that influence the number, rather than the nature, of the surges. These include the relative height of the line above local terrain and surrounding objects and the line width. Overhead shield wire protection (located above the phase conductors) eliminates more than 90% of the shielding failures on typical lines. Compact lines have reduced height and width, and these parameters tend to reduce the number of lightning flashes. Compact lines that do not use overhead groundwires must absorb shielding-failure surges in some other way, such as line surge arresters. Mid-span Flashovers occur when lightning terminates directly on an overhead shield wire or a phase conductor at mid-span. The lightning stroke current is injected into the shield wire or phase conductor. For long span lengths, the stroke current magnitude and rate of rise, together with the surge impedance, cause a local voltage rise that may exceed the lightning impulse strength of the conductor-to-conductor (or shield wire) gap and result in a mid-span flashover. Influencing factors in midspan flashovers relate mainly to the span length and the line geometry (wire spacing). The insulation strength between two parallel conductors for conventional lines is usually much greater than the impulse strength of the insulators, so midspan flashovers are not common. Hileman 1999 explains the mid-span flashover mechanism in detail. Backflashovers occur when lightning terminates directly on a transmission structure or overhead shield wire. The surge current, together with the impedance of the structure and the structure footing resistance, causes a transient potential rise on the structure and at the structure support end of every insulator (Figure 6.3-6). For high surge currents or highresistivity soil, these local overvoltages may exceed the lightning impulse strength of the structure-to-conductor or ground wire-to-conductor gaps. Influencing factors in backflashover are the local structure footing resistance, the surge current magnitude and rate of rise, and conductor-to-structure (or ground wire) clearances and impulse strength of the insulators. Overhead shield wires couple a fraction of shield wire voltage onto the local (upper) phase conductors, reducing stress and improving performance. Adding an underbuilt shield wire(s) couples the lower phase conductor(s), reducing stress in these conductors and further improving performance. The number of flashes to the line and the tower surge response are affected by the line height. Compact lines typically have reduced conductor-to-tower clearances and small towerbase dimensions. Both of these factors can degrade backflashover performance to such an extent that alternatives, including unshielded operation and use of line surge arresters, may be more practical. Lightning Flash vs. Lightning Stroke A lightning flash consists of one or more lightning strokes traveling down the same channel in less that a second. The median 5
  6. 6. number of strokes in a flash is three, but as many as 10 have been observed. Figure 6.2-1 shows that each stroke consists of a rapidly-rising high current peak, which can, in some cases, exceed 100 kA in a few microseconds. After peak stroke current is reached, the current decays in roughly 100 microseconds or more to 50-400 amperes, which can flow for hundreds of milliseconds before the next stroke current begins (EPRI 1982; EPRI 2005a). Most of the charge delivered by a stroke does not exist near the high current peak but is delivered during the continuing current (Component C in Figure 6.2-1). The peak and rate of rise are important in insulator or air gap design, while total charge delivered is of great importance in arrester design. Figure 6.2-1 Components of lightning flashes (MIL-STD 1757). 6.3 FACTORS INFLUENCING LIGHTNING PERFORMANCE 6.3.1 Flashes to a Line A useful equation developed by Eriksson (Eriksson 1987) can be used to approximate the average expected number of lightning flashes to a line per 100 km per year:  28h 0.6 + b   6.3-1 N = GFD    10   Where: N = number of line hits per 100 km per year. GFD = Ground flash density at the line location, flashes to earth per km2/year.h = height of the topmost wire at the structure, meters. b = spacing between topmost shield wires, meters (if no shield wires are present, b = maximum horizontal spacing between outermost phases). In North America, regional GFD is displayed on GFD maps published by various entities (EPRI 2005a). If GFD data are not available, thunder-day (TD) data can be used for estimating purposed. The relationship between TD and GFD is approximately (EPRI 1982). GFD = 0.04TD1.25 6.3-2 Where: TD = average number of days that thunder is heard per year in the line vicinity. TD maps are published by the World Meteorological Organization (WMO 1953), and a United States map is published in McGorman et al. 1984. Also, NASA (Boccippio 2001) published a climatology of optical transient density (OTD), consisting of both cloud and ground lightning flashes observed from space with optical sensors. Cross-checks suggest that these OTD observations can be used to predict ground flash density with a calibration factor in Equation 6.3-3. GFD = 0.3 ⋅ OTD 6.3-3 Where: OTD = average number of optical flashes per km2 per year (Boccippio 2001). Figure 6.3-1 shows the variables in Equation 6.3-1. The number of hits to each configuration is directly proportional to the lightning activity (which can vary by as much as 3:1 from year to year). The expected number of hits varies roughly by the 0.6 power of the line height, and, for a given height, linearly as the line width. Wires with lower heights will receive fewer strokes and have better lightning performance, if all other variables remain unchanged. 6
  7. 7. Figure 6.3-1 Line profile for line hit and shielding failure calculations. In Figure 6.3-1, the compact line with no shield wires will receive fewer lightning flashes than the other configurations, but they will all be shielding failures. For unshielded lines at any voltage level, most of the line hits will result in line flashovers if the line is unprotected by arresters (TLSAs). Adding a shield wire adds a meter or two of height above the phase conductors and thereby collects a few more lightning strikes. However, if the shield wire takes the strikes as it is intended to, shielding failures or strikes to the phase conductor are avoided, which greatly enhances the overall line performance. If a second shield wire is added, the structure height (h) may be slightly reduced, but b is larger in Equation 6.3-1. With one or two shield wires, there is not much difference in the susceptibility to strikes, but two wires increase the coupling effect and reduce backflashovers, resulting in an improvement in overall performance. Smaller lines receive fewer lightning hits because of their relatively smaller size. Trees and other nearby objects along the right-of-way boundary and having heights comparable to the line can intercept lightning that would otherwise be attracted to the line. Lightning to some trees with deep internal sap ducts (such as pines) tend to detonate from the pressures of interior arcing and scatter debris as a result. Lightning also has a propensity to start fires in trees during dry seasons, and tree trimming along rights-of-way can be difficult to access, so again there are tradeoffs. Usually, however, the benefits in allowing flashes to strike to nearby trees or other tall objects rather than to a compact line can be substantial, particularly in reducing backflashovers. Note, however, that lightning hitting a tree creates essentially the same electromagnetic radiation as mentioned in Section 6.2. This electromagnetic radiation couples into nearby line conductors, causing flashovers for lines with BIL ratings lower than approximately 400 kV. 6.3.2 Stroke Characteristics The essential lightning stroke characteristics that initiate line flashovers are usually represented by cumulative probability curves (IEEE 1985). Figure 6.3-2 shows a stroke peak current magnitude probability. Conservatively, 50% of any first strokes to a line exceed 31 kA crest current, and about 0.8% exceed 200 kA. Three or more strokes can be expected in any flash over one second in duration. Figure 6.3-2 Cumulative distribution of peak of negative first stroke current (Anderson and Eriksson 1980). Figure 6.3-3 shows the relationship between stroke peak amplitudes and maximum rate of rise in the current surge for flashes to transmission towers (Takami and Okabe 2007). Figure 6.3-3 Relationship between peak of negative first stroke current (peak amplitude (kA)) and maximum rate of rise (kA/μS). 6.3.3 Insulator Flashover Strengths The dielectric strength of a line insulator varies with the waveshape of the applied voltage, atmospheric pressure, and humidity, but two rough rules of thumb for lightning transients derived from (EPRI 1982; Hileman 1999) are: BIL+ = 125 + 539 S 6.3-4 BIL- = 164 + 469 S 6.3-5 Where: BIL+ = basic impulse insulation level (crest kV) for a positive polarity wave. 7
  8. 8. (backflashover calculations). BIL- = basic impulse insulation level (crest kV) for a negative polarity wave (shielding failure calculations). S is the striking distance (insulator dry-arc distance) in meters. Basic impulse insulation level is that voltage that will cause flashover 10% of the time for standard conditions of atmospheric pressure and humidity. The flashover path around a line insulator exhibits a “volt-time” effect that is important in fixing the line flashover performance (IEEE 1985). Normally, insulation strength is evaluated just before the return of traveling waves from adjacent towers. At 2 µs, only the towers within 300 m of the flash termination share any fraction of the surge current. For shielded compact lines with minimum-height towers and short span lengths, this time comes sooner, and the corresponding insulation strength is higher than for conventional lines. Also, the overvoltage waveshape across insulators has a fast-falling time to half value of 3-10 µs, compared to the 50-µs value in the standard test waveshape (IEEE 1998). The tail times for overvoltages induced from nearby lightning will be even shorter, at 2-5 µs. The use of a volt-time curve approach for these cases will usually be too conservative for accurate estimates. Advanced methods for these cases are found in EPRI 2005a. 6.3.4 Shielding Failures A shielding failure occurs when a lightning flash gets by any overhead shield wire protection and hits a phase directly. The outboard phases are invariably involved, and the stroke current is usually under 20 kA. Simulation programs such as EPRI TFLASH (EPRI 2005b) or the L-1 and L-2 applets accompanying EPRI 2005a can be used to evaluate expected shielding failures for any proposed shield wire – phase geometry. Figure 6.3-4 Shield angles for line with overhead groundwires. The shield angles Ө shown in Figure 6.3-4 between groundwires GW and phases A and C are particularly important in mitigating shielding failures. Figure 6.3-5 demonstrates how the ratio of phase hits to total line hits can vary with shield angle Ө for different line heights for a single shield wire and two outboard phases. The lower the line height, the greater the permitted shield angle should be for the same performance. The greater the line height, the less the shield angle should be for good performance. Compact lines that are double circuit or have distribution underbuild will generally be taller and wider than a single circuit line and may require two shield wires and a shield angle of 20° or less to attain infrequent shielding failures. Compact, single-circuit lines at lower voltage may perform adequately with a 30° shield angle. Figure 6.3-5 Hits to outboard phases for various shield angles and line heights. For lines with no overhead groundwires, nearly every hit terminates on a phase conductor rather than a pole, and is thus a shielding failure. For practical purposes, every shielding failure results in a flashover on at least one unprotected insulator. Some shielding failures can cause severe insulator damage. First return-stroke currents of 100 kA have corresponding steepness of 50 kA/µs in Figure 6.3-3. These currents are injected into the high surge impedance of the phase conductors Zphase of about 300 Ω in corona, in each of two directions. The steepness of the resulting transient voltages, given by dI/dt times 8
  9. 9. Zphase/2, gives voltage rates of change of more than 7.5 MV/µs. This puts a current of more than 20 A through a typical 3-pF insulator string capacitance, increasing risk of insulator damage by puncture or shed fragmentation in weak regions of the insulation, whether porcelain or polymer. The same extreme dV/dt voltage steepness also stresses any parallel line surge arresters, installed to limit the overvoltage level across insulators to prevent shielding failure flashovers. 6.3.5 Midspan Flashovers For most lines in areas of high lightning activity, at least 40% of lightning hits occur on the spans between the structures, and particularly for compact lines with closer spacings between shield wires and phase conductors, adequate shield wire-spacing must be maintained to prevent flashover from shield wires and a phase conductor. EPRI 2005a and Hileman 1999 explain this flashover process in detail. If an overhead groundwire is struck, the shorter the span length and the lower the tower footing resistances, the less likely a midspan flashover is to occur. This is because current reflections from the nearby towers reduce the voltage stress at midspan, and can interrupt a midspan flashover that would otherwise occur (Hileman 1999). 6.3.6 Backflashovers and Structure Ground Resistance A low tower ground resistance is of extreme importance in regions of high lightning activity if lightning backflashovers are to be minimized on shielded lines or lines with line surge arresters. A backflashover occurs when a lightning flash strikes a structure top or overhead shield wire, and the resulting potential differences across one or more of the insulators exceed the insulator dielectric strength. Figure 6.3-5 shows a simplified equivalent single phase circuit for a hit to a tower top. The stroke current splits, part IS going out into the shield wire impedances ZS, and part IT flowing into the structure (or ground wire) inductance LT and the ground resistance RG. It should be clear that—for a given stroke current—the higher the ground resistance RG, the higher the voltage at the tower top and the greater the voltage VP across the insulator. There is also a voltage induced on the phase conductor by “coupling” from all overhead shield wires and from any underbuilt distribution circuits. This coupling is shown in Figure 6.3-6 as capacitances CSP and CPG, although in reality, magnetic coupling also exists. Figure 6.3-6 Simplified single-phase circuit for a hit to a tower. Reflections from adjacent towers are not shown in Figure 6.3-6. This situation thus persists for two span travel times, given by twice the span length divided by 90% of the speed of light (EPRI 1981; IEEE 1985). 6.3.7. Alternate Methods of Estimating Flashover Performance For those who do not have access to the EPRI TFlash program (EPRI 2005b), other alternatives—though not as comprehensive—can be used for rough estimates of line lightning performance. Reference IEEE 1997 is an IEEE Standard and contains a small computer program diskette called FLASH. This program is essentially a digital version of an estimating method published in EPRI 1982. It does not handle arresters. However, some success has been attained by regarding each phase protected by arresters at every tower as if they were underbuilt shield wires, then using FLASH or other programs to evaluate the voltages on the unprotected phases. 9
  10. 10. 6.4 LIGHTNING PERFORMANCE OF SHIELDED COMPACT LINES 6.4.1 General Insulation and Spacing Considerations A compact line is defined as a line that takes advantage of reduced phase-to-phase spacing, and controls overvoltages through other means such as surge arresters or closing resistors in circuit breakers, controls conductor motion through rigid insulators and short spans, and controls audible noise and EMI as described in other chapters of this book. It can be argued that any extra-high-voltage line is “compact,” since most designs at 345 kV and above have higher phase-to-phase voltage stress per meter of phase separation than lines built for system voltages below 300 kV. One direct consequence of reduced phase spacing is that the insulator dry-arc distance in compact lines may also be reduced. In contaminated areas, additional sheds or deeper skirts can add the necessary leakage distance, but large-diameter sheds do not increase the metal-to-metal dry arc distance significantly. This means that compact lines built with 1-m dry arc distances have lightning performance that is significantly worse than conventional lines built with 2-m electrical clearances in the same lightning and grounding environment. The experience of a utility in Minnesota, Otter Tail, described in Chapter 2—although not a typical case—is an instructive example. It is a particularly good example of the sorts of problems that can arise when a utility is required to be extremely aggressive in modifying lightning performance of a compact line that does not meet initial expectations. A 41.6-kV subtransmission line was initially upgraded to a 115-kV line with compact dimensions and no overhead groundwires. As Otter Tail gained experience with this compact 115 kV design, they began to make many modifications, since their lightning outage rates proved unacceptable to their customers. Based upon breaker operation data, Otter Tail determined that 290 km of these compact115-kV lines had a nine-year average of 18.3 trip-outs per 100 km/yr during lightning storms. To reduce these outages, Otter Tail added overhead groundwires and line arresters in some sections to improve the outage performance. Photos of their two alternatives are shown in Figure 6.4-1. Figure 6.4-1 Retrofit of overhead groundwire (left) and center-phase transmission line surge arrester (right) to Otter Tail 115-kV compact line. The top phase conductor with transmission line surge arrester (TLSA) protection can be closer to the unprotected phases than an overhead groundwire. This means that more of the lightning surge voltage appears on the unprotected phases—and the stress on the insulators is the difference between these voltages. Additional details of TLSA protection are found in Section 6.5. While the reduced insulation level degrades lightning performance, closer phase-tooverhead groundwire spacing of compact lines generally improves the electromagnetic coupling, indicated by Csp in Figure 6.3-6. This means that a higher fraction of the towertop impulse voltage appears on the floating phase conductors. Since the insulator impulse strength must withstand the difference in voltage between the tower voltage and the phase conductor, any increase in the coupled voltage on the parallel phase conductors reduces impulse voltage appearing across the phase conductor insulators. An example of the increase in coupled voltage with compact geometry is given by Table 6.4-1. 10
  11. 11. Table 6.4-1 Effects of Compact Line Insulation and Phase Spacing on Lightning Performance (Outages per 100 km per year) of Line with Horizontal Configuration OHGW and ±2 m at 14-m height shielding 11-m phase conductor height at tower Insulation 2 m (14 discs) dry-arc distance 1 m (7 discs) dry-arc distance Phase Spacing 4m 2m 4m 2m 1.1 0.9 10 9 Rfooting = 20 Ω 2.1 1.8 17 15 Rfooting = 30 Ω 4.0 3.3 23 20 Rfooting = 50 Ω 9.6 8.2 35 33 Rfooting = 100 Ω 2 Ng = 6 flashes / km -year. Both phase configurations have good shielding from direct flashes. The improved coupling with close phase-to-phase spacing offers a significant benefit of between 10 and 20% better lightning backflashover performance. However, this alone cannot make up the deficit in backflashover rate related to the reduced insulation. A number of options can be considered to improve the dismal lightning performance of the compact transmission line with 2-m phase spacing and 1-m dry arc distance. Some possible options would be: • Improve the grounding resistance • Improve the coupling by bundled pairs of overhead groundwires together to increase the geometric mean radius • Increase the insulation strength • Add line surge arresters • Install an underbuilt shield wire • Install twin bundled ground lead (unless steel pole or tower) Figure 6.4-2 illustrates the structure configurations assumed for these examples. Performance improvement for arranging the conductors in a more compact delta configuration is given in Table 6.4-2. The relative merits of each option listed above are ranked in Table 6.4-2. Table 6.4-2 Options for Improving Compact Line Lightning Performance (Outages per 100 km per year) OGHW -2 and 2 m at 14-m height 11-m phase conductor height at tower 1 m (7 disc) dry-arc distance Horizontal, 2-m phase spacing Rfooting 20 Ω Rfooting 30 Ω Rfooting 50 Ω Rfooting 100 Ω 9 6.5 2.6 15 11 5.7 20 16 12 33 28 20 2.4 1.7 4.4 3.2 7.5 5.8 17 13 Basic performance Twin bundled shield wires Instead, cut Rfooting by 50% Arrange conductors in a delta shape Instead, increase to 1.5-m insulation Instead, convert to delta, single shield 11
  12. 12. wire on pole, 1.5-m braced posts, with bottom phase TLSA for coupling Improving coupling by adding more OHGW gives a modest improvement in lightning performance compared to the reference compact line. However, it seems to be more practical to convert the horizontal circuit to a delta arrangement using braced posts, and then to convert the bottom phase to a shield wire by fitting it with a suitable transmission line surge arrester. Figure 6.4-2 Options for improving compact line lightning performance. Hoffmann (Hoffmann et al. 2003) offers a design alternative, also discussed in Section 2.3.7, that takes further advantage of coupling to ensure acceptable lightning performance for urban 230-kV lines in Brazil. Figure 6.4-3 shows the vertical orientation of twoconductor bundles with a phase separation of 3 m. With a maximum 200-m span and a pole height of 22 m, and using the same ground flash density Ng of 6 flashes per km2 per year as in Tables 6.4-1 and 6.4-2, the lightning tripout rate of this line is computed to be: • 4.6 outages per 100 km per year with a single overhead groundwire • 2.4 outages per 100 km per year if two overhead groundwires were located conventionally, above the phase conductors • 1.9 outages per 100 km per year with the above- and below- arrangement of shield wires shown in Figure 6.4-3. This example illustrates that it is feasible to obtain a good lightning performance from a compact transmission line using conventional overhead shield wire protection, and also that the use of an underbuilt horizontal (messenger) ground wire makes a further improvement. The underbuilt ground wire can contribute a performance similar to adding a surge arrester to the bottom phase conductor. Figure 6.4-3 Example of 230-kV compact line with underbuilt shield wire protection. 6.4.2 Wood and Fiberglass as Insulating Materials for Compact Lines Wet wood or wet fiberglass (without sheds) exhibits very little power frequency dielectric strength and can be subject to fires from leakage currents at voltages at 230 kV or above if lengths are inadequate. Prudence requires that no power frequency dielectric strength be attributed to these materials when used as crossarms or poles. If unbonded, they should always have polymer, porcelain, or glass insulators as leakage current suppression resources. However, these crossarms do exhibit significant lightning impulse dielectric strengths, and wet wood crossarms are also generally regarded as having some power frequency follow-current suppression capabilities. A seasoned wet wood crossarm alone can have an impulse flashover strength of over 300 kV per meter. Darveniza (1980) tested the lightning impulse strengths of many combinations of wet wood and porcelain insulators in series, and Figure 6.4-4 shows interpolated values from his report. Some deterioration in lightning impulse strength has been reported when porcelain or glass insulators have been used with unbonded fiberglass crossarms or structural components because of the much different dielectric characteristics of the two materials. 12
  13. 13. Figure 6.4-4 Impulse flashover strengths of wet wood crossarms and porcelain insulators in series. Additional dielectric strength benefits can sometimes be achieved by using wood or fiberglass standoff struts to locate pole ground wires away from the surface of the pole in the vicinity of the pole top to add additional air gap spacing to the crossarm-insulator geometry. 6.5 LIGHTNING PERFORMANCE OF UNSHIELDED COMPACT LINES 6.5.1 General Engineering Considerations Any unshielded compact line—unprotected by TLSAs—will essentially flash over each time lightning strikes the line, and these flashovers can involve several towers and phases for the same event. The only benefits of reducing footing resistances for unshielded lines without arrester protection are increased fault currents for relaying purposes and possibly reduced step-and-touch potentials at a faulted structure. Increasing insulation on an unshielded line may also have little effect. Even a strike directly to the low impedance of an EHV phase conductor bundle usually creates such enormous voltages that something is almost certain to flash over regardless of insulation level. Thus the number of lightning flashovers per year for any line with an insulator BIL of 400 kV or higher and without TLSA protection is essentially equal to the number of lightning hits to the line per year, and an estimate of the median value per year can be found by Equation 6.3-1. For unshielded lines—unprotected by TLSA—with insulator BILs under 400 kV, additional flashovers may be created by lightning hits to the earth near the line. The powerful electromagnetic fields created by these hits in the line vicinity induce high-voltage transients onto the phases that can exceed the phase insulation strength, as discussed in Section 6.2. Modern compact line technology is undergoing a fundamental change, specifically in application of TLSAs to replace or augment shield wires for lightning protection. This application affects almost every other aspect of compact line engineering, including phase spacing, clearances, foundations, grounding, safety, maintenance, and general line performance. A compact transmission line does not automatically require a TLSA. However, a properly applied TLSA application can: • Absorb an appropriate amount of surge energy without failure. • Limit overvoltages, providing a major improvement in line compaction. • Provide better lightning performance than shielded lines with conventional dimensions. However, direct exposure of arrester-protected phases with no overhead shield wires for lightning diversion has—in some cases—required line rebuilds with addition of shield wires to avoid excessive arrester failures, breaker operations, and maintenance problems. Energy capabilities and mechanical capabilities of TLSAs can vary widely, as can ground flash densities, so care must be taken. • Supplant shield wires completely in some applications. On the other hand, improperly applied TLSAs can: • Result in an excessive number of arrester failures by poor arrester selection. • Waste money with little or no improvement in lightning performance by improper TLSA location strategies. 13
  14. 14. • • Transfer overvoltages to insulation that would otherwise be un-stressed. Degrade lightning performance by arrester mechanical failures that reduce striking distances. • Degrade general line reliability under other adverse-weather conditions such as fog or galloping, in some cases to levels worse than if no TLSA were applied. • Create excessive radio noise by poor connections or blown disconnectors or improperly configured gaps. • If disconnectors do not function, the failed arrester becomes a momentary short that is difficult to locate unless fault-locating relays are available to assist with pinpointing the failed arrester. Modern polymer-housed TLSAs are designed and tested to have nonexplosive failure modes that allow application in urban environments, where they have the greatest benefit to customers. The goals of improved lightning performance are usually reached with arresters selected for a Transient Overvoltage (TOV) rating that is well above the system voltage, and most manufacturers of TLSAs specify coordination that forces station-class arresters to absorb this energy. The use of TLSAs rather than shield wires also exposes aluminum phase conductors to lightning arc damage. Test standards such as IEC 60794, based on component C in Figure 6.3-1, have been developed to reproduce lightning damage to aluminum-strand optical fiber groundwires (Chisholm et al. 2001). These test standards are also appropriate for selection of phase conductors on unshielded compact transmission lines, with or without TLSA. TLSAs have two principal applications: • Reducing insulator backflashovers in local areas of high footing resistance • Reducing shielding failure flashovers on lines with no overhead groundwires In many cases, installation of shield wires adds little additional initial cost to a transmission line, but additional structure mechanical loads from these wires require more costly structures and foundations to cope with increased overturning moments and increased ice loading. Continuous and peak overhead groundwire losses and line maintenance issues can sometimes be of concern. However, if shield wires are removed and arresters substituted to manage lightning overvoltages, care must be taken to install and maintain them properly and to be certain that each arrester has adequate energy capacity to survive severe lightning discharges. This is not to say that TLSA applications are necessarily a better alternative than shield wires—each has its place according to lightning exposure and grounding conditions. Selection of the proper operating voltage of any gapless arrester continuously connected to a phase must recognize that the arrester must withstand the maximum continuous phase to ground operating voltage (MCOV) without overheating. The manufacturer will provide this voltage rating for each production arrester, but the designer should also specify to the manufacturer the maximum expected voltage and duration on any unfaulted phase during a phase-ground fault. A conventional MCOV rating can be supplemented with an arrester maximum coulombs rating or maximum energy capability of each block in Joules per cm3, assuming adiabatic heating close to the melting point of the zinc oxide material. For arresters in series with a gap, there is no MCOV, except the maximum arrester resistance drop during the conduction period. 14
  15. 15. A number of developments have completely shifted the paradigm for lightning protection of compact lines. In the previous edition (EPRI 1978), expensive reductions in footing resistance were needed to obtain the marginal improvements described in Section 6.4. The alternative of improving coupling and limiting overvoltage levels using TLSAs was introduced there, and is expanded in detail in this section. 6.5.2 Recent TLSA Applications on Compact Lines With energy absorption capability at destruction in excess of 1500 Joule per cm3, and corresponding charge levels of 10-20 C for IEC Class C (62-63 mm diameter) zinc oxide blocks, there is now a possibility that overhead groundwires can be completely eliminated. This reduces the height and visual impact of lines, giving a new axis of line compaction. A partial list of recent TLSA developments includes: • Adoption of metal oxide varistors as the nonlinear element • Production experience with a wide range of MOV formulations • Control of the V-I characteristics that affect energy sharing • Adoption of sealed polymer housings to improve reliability • Introduction of external series gaps as an alternative to explosive disconnects and hinged installations to reduce cost and improve reliability • Proliferation of suppliers, leading to a competitive market TLSAs have proved to be quite reliable in both shielded and unshielded applications if they are properly applied. Local area reduction of lightning backflashovers by TLSA is described in detail in References (EPRI 1997; Koch et al. 1985; Shih et al. 1985; EPRI 2005a), and this application has had considerable success. Arresters have also been applied on unshielded 69- to 230-kV transmission lines, with a 400-kV application considered for a compact design in Norway (Loudon et al. 1998) and a 500-kV application in Japan (Ishida et al. 1992). Williamson (2007) reports application of gapless TLSAs on 138-kV lines of New Brunswick Power Transmission Corporation. They were quite successful in improving lightning performance in high ground resistance areas, but did encounter mechanical and hardware problems that constituted valuable experience. 6.5.3 Precautions for TLSA Use on Compact Lines Calculations based on the V-I characteristics of line surge arresters suggest that the sharing of energy among parallel arresters on protected unshielded circuits could be a problem. The EPRI TFLASH program, for example, contains algorithms to calculate the effectiveness of all reasonable TLSA applications and estimates the total energy absorption as heat for a statistical distribution of stroke currents. TFLASH compares calculated energy values with manufacturers’ ratings, and “red flags” excessive events. However, to date, there have been few failures from excess energy. Experience has instead shown that manufacturers’ ratings tend to be conservative. Arrester energy testing performed by EPRI on a specific design of TLSA showed that less than 4% of the varistor blocks failed at energies lower than twice their rated values (EPRI 2000). Each ungapped arrester has a MCOV (maximum continuous operating voltage) and a manufacturers’ designated set of discharge voltages for various arrester surge currents. This discharge voltage should be at least 10% below the insulator BIL for the highest expected discharge current, usually negative for most lightning hits to phases on unshielded lines. A gapped arrester has no MCOV since no voltage exists across the arrester until a lightning hit occurs. 15
  16. 16. Most problems with arrester applications have been related to misapplication beyond manufacturers’ mechanical ratings—for example, when arresters are used to restrain conductor motion. In the modern compact line, restraint of conductor motion is important for maintaining minimum phase-to-phase and phase-to-ground clearances under highwind and galloping conditions. These clearances affect the choice of tower head geometry and insulator configurations, as described in other chapters. The fact that conductor motion is restrained at the tower also makes application of line surge arresters more practical. Arresters can be installed in parallel with one of the rigid elements, and the use of braced post designs with two-section posts can automatically provide the necessary series gap for consistent operation of gapped TLSAs. Figure 6.5-1 shows a gapless TLSA installation, with seven discs on a 115-kV line. The action of wind on the conductor puts tension or compression loads onto the body and fittings of the arrester, which is a polymer-encapsulated fiberglass tube with a limited mechanical rating. Mechanical failure of the arrester or fittings by tension or compression is likely in high winds. Figure 6.5-1 Polymeric TLSA restraining conductor motion on a 115-kV Line. Arresters tend to fail shorted, and some means must be applied to disconnect a failed gapless arrester from the line to prevent a permanent fault. One method utilizes a disconnector at the line end of the arrester, shown by the small knob in Figure 6.5-1 (EPRI 1997). The power frequency fault current through a failed arrester heats a small cartridge that blows open the arrester connection to the line, allowing the line breakers to clear the fault in the newly formed air gap. The hanging connection or arrester supplies a visible signal that the arrester has failed, and any subsequent lightning flashes to this newly unprotected location will cause shielding failure flashovers. When arresters are added without an existing shield wire, arrester disconnectors become another failure mode. Disconnector operation from direct lightning strikes can sometimes occur even if the arrester survives. Figure 6.5-2 shows a gapless arrester application on a compact, unshielded 138-kV line (Williamson 2007). With five years of operation, mechanical defect rates of 28% and 0.6% were found with two different TLSA suppliers, and no arresters were reported to have failed from excessive energy dissipation. Figure 6.5-2 138-kV arrester and insulator post assembly. The second means for preventing permanent faults employs an air gap in series with the arrester. The simple external air gap sparks over for lightning overvoltages, connecting the arrester to the protected phase. The air gap has enough dielectric strength to interrupt the power-frequency follow current at the first zero crossing for normal operating voltage. This method is popular in Asia and Europe. Older engineers in North America are familiar with the series gaps needed for successful application of silicon carbide arresters, capable of interrupting about 100 A of ac power-follow current. The new metaloxide technology performs much better than silicon carbide, with an ac follow current of less than 1 A. This means that the complex magnetic coils and grading circuits are no longer necessary. If the arrester fails, breakers clear the faulted arrester, and the gap recovers its dielectric strength and disconnects the arrester. The use of an external series gap has the advantage that the arrester varistor blocks and housing are under stress only when the arrester is operating. This allows the length of the arrester column to be reduced while still meeting 16
  17. 17. all interruption requirements. The short time lag for the gap sparkover does not significantly interfere with the TLSA protective characteristics. Since gap sparkover for an operating arrester limits power follow current, electrode ablation is not usually a problem, and sparkover directly to the phase conductor is often possible, or to an armor rod around the conductor in some cases. Care should be taken that either gapped or gapless TLSAs do not at any time inject significant radio noise into the protected phase. For a gapped arrester that is not an integral part of the arrester housing, the manufacturer will specify the minimum gap dimensions to clear an arc to a phase. For unrestrained conductor motion and a gap that is not an integral part of the arrester housing, it is usually not necessary to be concerned about lightning performance under maximum wind swing conditions since experience has shown that most lightning hits occur with little or no wind swing. However, the gap must always be sufficient to withstand power frequency voltages, and—for that reason—gapped installations are usually applied to restrained conductor designs. TLSA housings are invariably polymeric, making them lighter to install and less likely to fragment in case of a fault than porcelain housings. The housings can be designed to radially vent high-pressure gases during an internal fault, limiting fragmentation. Handling issues are particularly important for long TLSAs, and care must be taken that they are not bent during installation. 6.5.4 Case Study: TLSA Application on Existing 115-kV Compact Line One applicable and instructive compact phase arrangement uses a three-phase delta wire configuration on single poles, as shown in Figure 6.5-3. Figure 6.5-3 Typical 115-kV compact line geometry from 1980, using polymer post and semiconductive glaze bell insulators (courtesy Ontario Hydro 1980). The line in Figure 6.5-3 was designed to be a 115-kV upgrade of a 44-kV subtransmission line, with reliability achieved by meshing of multiple supply lines rather than through good lightning and contamination performance of any individual line. With a height h at the tower of 11.3 m, bottom-phase conductor separation of b = 2 m and a ground flash density of 2 flashes per km2 per year, Equation 6.3-1 suggests that there will be 24 shielding failures per 100 km per year. The rate of 24 outages per 100 km per year would not be suitable for adequate power quality in a dual-supply system, given the number of customers affected by the voltage dips associated with each 115-kV momentary outage. An unshielded compact transmission line designed today could treat the center phase, supported on a polymer post, with a line surge arrester. This would make it function as a shield wire under lightning conditions. The arrester voltage will add to the tower-top voltage, and this will actually improve the lightning performance of the unprotected phases by increased coupled voltage to the phases. It is also possible to stack a pair of post insulators in series to provide the recommended air gap from central flange to phase conductor, given by those manufacturers that supply externally-gapped TLSA rather than, or as well as, TLSA with explosive disconnects. This configuration has been used extensively, with more than a million “Current Limiting Arcing Horns” used for distribution systems in Japan. With a top-phase arrester in parallel with the central post insulator at every pole, there will be no more shielding failures. Instead, there will be some backflashovers on the unprotected 0.914-m three-bell insulator strings in Figure 6.5-3. Figure 6.5-4 shows how 17
  18. 18. the backflashover rate of this line is predicted by the FLASH program (IEEE 1997) to vary with footing resistance, treating the arrester-protected phase as a groundwire. Figure 6.5-4 Calculated lightning performance of compact 115-kV line with top-phase TLSA. A target performance of four outages per 100 km per year can be achieved in this area of moderate ground flash density if it is possible to maintain the footing resistance of each tower at less than 20 Ω. 6.5.5 Case Study: TLSA Application on Single-Circuit 138-kV Unshielded Line Figure 6.5-5 displays one example of a single-circuit 138-kV compact line without shield wires. The phase conductors are shown as “A,” “B,” and “C”. The following dimensional data and voltage data apply: Ground flash density: 10 flashes/km2/year Insulator BIL: -760 kV Average footing resistance (concrete): 25 Ω Conductors: Drake 795 kcmil Line length: 100 km Span distance: 200 m Conductor sags: 3 m Arrester MCOV: 106 kV Arresters on Phase B Figure 6.5-5 Single-circuit delta post-insulator line without shield wires. Table 6.5-1 shows TFLASH lightning flashover performance calculations for the cases of no arrester on the top phase, arresters on the top phase at every pole, and arresters on the top phase at every other pole. Table 6.5-1 Single Circuit Lightning Performance with and ithout Arresters per 100 Kilometers/Year Phas e A Flashes to Phase 2.8 0.3 3.1 B 0 111.4 111.4 C 2.8 0.3 3.1 A 5.1 0.3 5.4 B 0 0 0 C Case Backflashovers Total Phase Flashovers 5.1 0.3 5.4 A 4.8 0.3 5.1 B 0 105.0 105.0 No Arresters Arresters on Every Pole On Top Phase B Only Arresters on Every Other Pole on Top Phase B Only 18
  19. 19. C 4.8 0.3 5.1 With arresters on the top phase at every pole, there is an approximate 20:1 improvement over the performance with no arresters. With arresters on the top phase at every other pole there is little improvement. This is because the flashovers are simply pushed over to the unprotected poles. Again, caution should be observed that adequate energy capability exists for the top phase arresters for high charge lightning flashes to the line. For arresters on the top phase, as shown in Figure 6.3-3, either gapped or ungapped arresters can be positioned. 6.5.6 Case Study: TLSA Application on Double-Circuit 138-kV Unshielded Line Figure 6.5-6 displays a double-circuit 138-kV line with no shield wires. The following dimensional data and voltage data apply: Ground flash density: 10 flashes/km2/year Insulator BIL: -760 kV Footing resistance (concrete): 25 ohms Conductors: Drake 795 mcm Line length: 100 km. Span distance: 200 m. Conductor sags: 3 m. Arrester MCOV: 106 kV No wood bonding, i.e. no electrical conductors along the wood surface. Figure 6.5-6 Double-circuit compact line without shield wires. Table 6.5-2 shows TFLASH lightning performance calculations for Figure 6.5-6 for the following cases • no arresters, • arresters on the top and bottom phases of Circuit #1, • arresters on the top phases of Circuits #1 and #2, and • arresters on the top phase of Circuit #1 only. In this case study. the arresters are placed on every tower. The results are presented in Table 6.5-2 19
  20. 20. Table 6.5-2 Double-circuit Flashovers with and without Arresters per 100 Kilometers per Year Case Circuit Phase 0 4.32 63.3 0 0 B 3.36 0 4.01 0 4.01 0 B 3.36 0 0 0 0 0 B 0 0 0 0 A 3.04 0 B 0 0 0 63.32 A 0 0 B 0 0 C 1.46 0 A 1.46 0 B 0 0 C 2 1.58 C 1 B C Arresters – Top Phase of Ckt 1 Only 0 A 2 0.39 C 1 0 A Arresters on Top & Bottom Phases Ckt 1 Only 0.39 C 2 64 A 1 0 C Arresters on Top Phases Ckts 1 & 2 1.58 A 2 B C No Arresters Flashes to Phase 63.3 Total Flashovers A 1 Backflashovers 4.32 0 63.32 64 7.4 7.4 0 64 1.5 64 In Table 6.5-2, some flashes to the phases also caused backflashover on other phases, leading to multiphase-ground faults. For this insulation of 760-kV BIL, every flash causes a flashover. Even if the first stroke is less than the critical current of about 3 kA, it 20
  21. 21. is likely that one of the subsequent strokes will exceed 12 kA. When a lightning flash with high peak amplitude causes a flashover of the stricken phase, the lightning current then enters the pole ground, and this can—in some cases—increase the ground potential sufficiently to flash over another phase. The “Backflashovers” column shows the occurrence of simultaneous multiphase flashovers due to this ground potential rise. The following conclusions can be drawn from Table 6.5-2: 1. With arresters on both top phases at every pole, only a few backflashovers on the lower phases will occur per 100 kilometers per year, and the top phases essentially act as shield wires. The coupling effect is usually higher because of the tight phase spacing on a compact line, and hence the backflashover rate is better than it would be with larger phase spacing—i.e., not a compact line. Provided arrester energy capacities are adequate and arrester long-term environmental durability is satisfactory, this should be an adequate replacement for shield wires. 2. With arresters on the top and bottom phases of only one circuit at every pole, backflashovers of the protected circuit rarely occur. The unprotected circuit will frequently flash, primarily from every strike to that side of the line. 3. Arresters connected only to the top phase of one circuit at every pole will perform almost as well as arresters on both top and bottom phases of that circuit. This result is dependent on the footing resistance; for high values of footing resistance, flashovers to the bottom phase will increase. Roughly speaking, any phase with an arrester connected to it acts somewhat like a shield wire. 6.6 CALCULATION OF ARRESTER ENERGIES The principal problem in planning TLSA protection on unshielded lines is determination of the minimum size arrester block to use. The greater the diameter of the block, the more expensive the arrester, but an inadequate block diameter (or quality) can lead to an unacceptable failure rate. Blocks can puncture from excessive surge voltage or fail from excessive heat energy injection. Calculation of maximum expected TLSA energies can involve substantial errors because of a paucity of field data. Failures can be expected to occur primarily when very severe flashes strike the line, but very little information about the probability distributions of very severe flashes exists. Severe flashes constitute only a few percent of all available data, and will vary with latitude and other factors. Where transmission lines have shield wires, experience has shown that TLSAs can function very well in suppressing insulator voltages, and even without shield wires, TLSAs should function satisfactorily if the quality and diameter of the arrester metal-oxide blocks are sufficient. Figure 6.6-1 shows energy test results on some earlier model ungapped TLSAs having 41 mm diameter blocks. Figure 6.6-1 Example TLSA failure probabilities vs injected energy in kilojoules per MCOV for an ungapped arrester design with 41-mm diameter blocks. The arresters began to fail at slightly above the manufacturer’s rating of 2.2 kilojoules per MCOV, but the 50% failure rate did not occur until 9.8 kilojoules per MCOV was reached. Appendix 6.1 discusses how these failure probability curves can be used to calculate TLSA energies during lightning hits on unshielded lines. On lines with shield wires, computer programs—such as EPRI TFlash—signal conditions wherein TLSA energies are exceeding their manufacturers ratings. The frequency distribution is not normal, and resembles a Weibull distribution skewed toward the low end. 21
  22. 22. 6.7 HIGHLIGHTS 1. Lightning Performance. From the standpoint of lightning performance, compact transmission lines—if designed properly—can have adequate lightning performance. Compact transmission lines usually collect a smaller number of lightning hits than conventional lines because of their smaller footprint. However, compaction may require some sacrifice of air gap clearances and insulator lengths, so that critical stroke currents for flashover can be less. Design comparisons of lightning performance of proposed compact lines and conventional lines can only be made with a comprehensive lightning simulation program such as EPRI TFlash. 2. TLSAs. Since the ultimate in compaction requires removal of shield wires, interest is increasing in utilization of TLSAs to limit compact line lightning flashovers, and designs up to 500 kV have been developed. TLSAs can help maximize compaction of lines. 3. Gapped versus Gapless Designs. TLSAs are widely available in both gapless and externally gapped designs. An early decision must be made as to which is more applicable for a specific line design. Gapless arresters have an advantage for some tower geometries, particularly where conductors are allowed to swing. The blown disconectors clearly signal either arrester failure or disconnector failure. However, some disconnectors have also proved to have poor mechanical reliability. Externally gapped arresters can sometimes be positioned on the pole horizontally and in parallel with the phase conductors, so that sparkover can be made to occur directly from a restrained conductor to the live end of the arrester. This gives a simple and convenient installation. Gapped arresters have a further advantage that voltage exists on the arrester only during the sparkover process, permitting shorter arresters with reduced leakage distance and no concern over heating from operating voltages. Factoring in the improved reliability with rigid connections, externally gapped arresters would seem to be a better overall choice for compact line designs that use line post insulators, and gapless designs rated for the necessary tension, cantilever, and compression loads would seem more appropriate for restraining motion with suspension insulators. 4. Product Test Data. Since there can be a significant variation in mechanical and electrical quality and energy capability of TLSAs now being marketed, it becomes important to have access to good test data of failure probabilities and modes for any proposed product. 5. Inspection and Maintenance Strategies with TLSAs. If TLSAs are used to eliminate shield wires, it should be expected that occasional failures due to severe lightning hits will occur, and that there will be mechanical problems as well. Inspection, replacement, and maintenance strategies should be factored into the lifecycle cost analysis in comparison to 25- to 50-year, maintenance-free experience with overhead groundwires. APPENDIX 6.1 CALCULATION OF ARRESTER FAILURE PROBABILITIES OF UNSHIELDED LINES Arrester failure probabilities on an unshielded line depend on a variety of parameters: • Gapped or direct-connected arrester designs and their ratings and quality • Distribution of arresters on a struck phase • Span distances 22
  23. 23. • Tower footing resistances • Regional ground flash densities • Shielding effects of nearby objects (trees, structures, regional topology, etc.) • Statistical variation of storms and flash intensities from year to year The following software procedures are suggested to make the calculations in the absence of dedicated software facilities: Estimate number of hits to the line per year. The mean number of hits per year N to a compact line can be estimated using Equation 6.3-1, recognizing that the hit rate can double in some years. Using this number N—if regional information of the percentage of positive flashes is not known—assume that 10% of N flashes to the line will be positive, and that any positive flash to a line will be severe enough to fail an arrester. This is an overly simple, but conservative, assumption. For each line section of nine contiguous structures, select a phase protected by TLSA, and determine an equivalent circuit for a lightning hit to the phase at center structure 5. Figure A6.1-1 shows the conceptual circuit for a hit to a protected phase. Figure A6.1-1 Conceptual circuit for a TLSA energy analysis. The volt-ampere characteristics of the arresters “A” (all arresters assumed to have the same MCOV) can be approximated by: I A = αV Aβ A6.1-1 Where: IA = arrester current, kA. VA = instantaneous voltage across the arrester, kV. α,β = constants to be determines from manufacturer’s specifications. Test values of VA at IA = 5 kA and 40 kA can be used to calculate α and β. The resistors RG are tower dynamic footing resistances, and can be assumed to be the low-frequency values if better data are not available. Note that the tower inductance is not shown, since at maximum arrester current, the dI/dt and inductive voltage drops are zero. For gapped arresters, the gap can be represented by an open switch in series with the arrester. The switch closes when the voltage across it reaches a preset value determined by the manufacturer’s acceptance test data. Once closed, the switch remains closed for the duration of the event. The impedance ZO is the surge impedance of the struck phase, degraded somewhat by corona. A value of 500 ohms for a single conductor should be sufficient for these calculations, and somewhere around 300 ohms for a bundle. Select a flash severity S to be used in the EMTP calculation. Assigning a “severity” of 8 to a flash to be used in an EMTP calculation means that 8 out of 10 random flashes will have a charge magnitude less that this flash. A severity of 5 assigned to a flash means that 5 out of 10 random flashes will have a charge magnitude less that this flash. It is unrealistic to expect TLSA on a compact unshielded line to withstand all possible lightning flashes, since some flashes will be extremely severe, and consequently a few TLSA failures should be expected. Figure A6.1-2 is taken from a flash coulomb probability curve, and provides this “Severity Rating” on a scale of 1-10 in terms of the corresponding coulombs delivered for a negative flash. For example, the severity of 8 mentioned above corresponds to a total flash charge of approximately 16 23
  24. 24. coulombs, whereas a severity rating of 10 corresponds to approximately 100 coulombs and should be extremely rare. Using the selected flash severity, find the corresponding flash coulombs delivered. Figure A6.1-2 is used to find the total charge delivered Figure A6.1-2 Total negative flash coulombs vs. severity rating. Calculate the stroke current waveshape and stroke magnitude to be used in the EMTP analysis. Metal-oxide arresters are very nonlinear devices, and their heat absorption for a total flash coulomb content can vary widely for different current waveshapes, primarily because an arrester is always in parallel with the constant surge impedance of a conductor and is influenced by current-sharing with other arresters on the same phase. The conventional 8 x 20 current waveshape used to test station arresters does not represent a flash current injected into a phase. This flash current will usually consist of a sequence of several strokes over a time period of many milliseconds, and the low currents between the high current peaks can be carried off by the phase surge impedance with very little flowing into the arresters. Since inductance is absent in Figure A6.1-1, dI/dt becomes of little importance in adiabatic heating of the arresters, and a simplified stroke current waveshape similar to that shown in Figure A.6.1-3 is suggested, wherein four of these strokes in rapid succession can be assumed to constitute a flash. Figure A6.1-3 Suggested simplified flash current waveshape for TLSA energy calculations. If four identical strokes with Figure A6.1-3 waveshapes constitute a negative flash, then the peak current Imax of each stroke is given by: I max = 2.5QF A6.1-2 Where: QF = total charge (coulombs) delivered by the flash. Imax = crest current, kA. Since the strokes are identical, only the arrester energy delivered by one stroke has to be calculated by EMTP and the result multiplied by four to get the total energy injected into the arrester at the flash location. Use EMTP to calculate the total kilojoules injected into each of the arresters of Figure A6.1-1. This calculation is made using one of the stroke current waveshapes of Figure A6.1-3. Divide the total kilojoules in each arrester by the arrester MCOV and multiply by four since the lightning flash consists of four of the Figure A6.1-3 transients. Compare the maximum calculated kJ/MCOV by test data exemplified by Figure 6.6-1. As a conservative approximation, if the calculated kJ/MCOV equals or exceeds the 50% failure rate, assume that a failure will occur. This assumption is acceptable since—as shown by Figure 6.6-1—a substantial number of failures will occur below the 50% value. A more accurate approach would utilize a distribution of charge magnitudes in a flash (see Figure A6.1-4), calculate the failure rate for each charge magnitude, and sum all the failures. However, this will require dedicated software, and the distribution of negative charge at high values in Figure A6.1-4 is dubious due to lack of field data. Sum positive and negative failures to find total expected failures for this selected arrester application. 24
  25. 25. If the number of failures per 100 km per year is unacceptable, investigate the use of larger diameter blocks in the arresters. This requires test data similar to Figure 6.6-1 and volt-ampere test data for the proposed arrester to determine α and β in Equation A6.1-1. Figure A6.1-4 Distribution of expected charge in a negative flash (EPRI 2005a). APPENDIX 6.2 TESTING TRANSMISSION LINE SURGE ARRESTERS PRIOR TO INSTALLATION If TLSAs are to replace shield wires on a compact transmission line, it is vital that they survive existing regional electrical and environmental challenges over a period of many years. Challenges will include: • Weather (icing, solar radiation, internal water accumulation, condensation, etc) • Mechanical motion • Continuous and momentary electrical stresses and energy challenges • Installation abuse • Pollution • Corrosion of metallic parts • Vandalism Arrester blocks can fail for a variety of reasons: • Thermal runaway from excessive heating when exposed to severe lightning currents or excessive power frequency voltages • Insulation breakdown at the insulating collar surrounding one or more arrester blocks • Insulation puncture through a block leading to a cascading failure • Block cracking Most of these failures occur near the maximum energy levels of the blocks, usually substantially beyond the manufacturers’ ratings, and in the field may not be detectable in some cases until some sort of visible mechanical disruption occurs. Adequate quality control should be ensured by electrical and environmental tests prior to line construction. Test Requirements Environmental TLSAs may be subject to environmental challenges not anticipated by the manufacturer or overlooked in the electrical or mechanical design. Accelerated life tests in environmental chambers—including voltage, temperature, humidity, UV light, and mechanical cycling—are recommended. Figure A6.2-1 shows a 500-kV accelerated aging environmental chamber at the EPRI Lenox, Massachusetts Research Center. Although mostly utilized for insulator life testing, the facility can evaluate a variety of 25
  26. 26. arresters. A few weeks can simulate a year or more of normal exposure. The aging acceleration varies from 7 to 14 depending on the type of test. Figure A6.2-1 EPRI 500-kV Aging Test Chamber. High Current Electrical Stresses Severe lightning transient currents passing through a TLSA create severe electrical stresses in the arrester blocks, and poor quality or poor design can result in catastrophic failures. This kind of electrical stress can be applied to a proposed arrester by an impulse generator. The impulse test can consist of a single current surge test of 40 kA or to the maximum available current from the generator. Arrester current oscillograms are then inspected and compared with short-circuit oscillograms and lower arrester current oscillograms to look for signs of block failure by abrupt increases in current. Temporary Overvoltage Acceptance Tests Every gapless TLSA must withstand expected temporary overvoltages (TOV) that will occasionally be present during faults and abnormal system conditions. The manufacturer will often supply TOV ratings for various durations of overvoltage. For gapped arresters, the gap should not flash over during TOV periods accompanied by simulated severe weather conditions. If a flashover does occur, the gap must clear when the voltage drops to normal. A gapless test arrester is first energized at normal voltage via a hipot transformer and the arrester current and normal watts loss observed. Then a temporary overvoltage from an autotransformer and current limiting resistor is applied to the arrester via a single-pole double-throw switch for a few seconds—arrester current and watts measured in the interval—and normal operating voltage then reapplied by the switch. The decay in arrester current and watts loss during the following normal voltage is then monitored as the arrester recovers. If desired, this test can be repeated at higher and higher overvoltages until failure to establish failure probability curves. TLSA Energy Failure Tests For line arresters, failure energies can exceed 1000 kilojoules for larger arresters, and this exceeds even the energy capability of large impulse generators. Under these conditions, Ringler et al. 1998 showed that TLSA energy failure tests can be carried out by injecting a 60 Hz overcurrent into a test arrester and oscillographically measuring the total kilojoules into the arrester until failure occurs. This test takes advantage of the adiabatic accumulation of heat per cycle. The product of test duration and current amplitude (charge) was found to be constant over five orders of magnitude, from 0.8 A to 35 kA, because failures occurred before significant amounts of heat could escape from the blocks. This test is essentially a continuation of the TOV and transmission line discharge tests mentioned above but at a higher voltage and current until failure. REFERENCES Agrawal, A. K., H. J. Price, and S. H. Gurbaxani. 1980. “Transient Response of Multiconductor Transmission Lines Excited by a Nonuniform Electromagnetic Field.” IEEE Transactions on Electromagnetic Compatibility. Vol. EMV-22. May. pp.119-129. Anderson, R. J. and A. J. Eriksson. 1980. “Lightning Parameters for Engineering Application.” Electra. No. 69. March. pp. 65-102. Boccippio, D. J., K. L. Cummins, H. J. Christian, and S. J. Goodman. 2001. “Combined Satellite and Surface Based Estimation of the Intracloud-Cloud-to-Ground Lightning 26
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