LIGHTNING PERFORMANCE OF COMPACT LINES

X International Symposium on Lightning Protection 9th-13th November, 2009 – Curitiba, Brazil

LIGHTNING PERFORMANCE OF COMPACT LINES William A. Chisholm1, John G. Anderson2, Andrew Phillips3, John Chan3 1

Kinectrics/UQAC, Canada – [email protected] 2 3

Consultant, USA – [email protected]

EPRI, USA – [email protected], [email protected] E-mail: [email protected]

1 INTRODUCTION The period of transmission line development starting in the mid 1950s led to worldwide applications of Extra-High Voltage (EHV) networks at system voltages at or above 345 kV. One of the engineering challenges faced and resolved in this development was the problem of insulation coordination for switching-surge overvoltages. Test results in this period suggested that generous dry-arc distances and tower clearances would be needed to provide adequate margin for EHV lines. For example, overvoltage level of 3.2 per-unit on a 345-kV system leads to line-to-ground voltage of 900 kV, corresponding to the critical switching impulse flashover level of a rod-plane gap of 2.9 m [1]. The per-unit (pu) system normalizes switching surges to the peak of the ac line-to-ground voltage. It proved to be better to limit switching overvoltages through the application of pre-insertion resistors in circuit breakers, as shown in Fig. 1, than to increase tower dimensions to accommodate the anticipated 3.2-pu switching surges without them.

Fig. 1 – Effect of Pre-Insertion Resistor Value on Maximum Switching-Surge Overvoltage Level for 500-kV Systems [2]

In cases where pre-insertion resistors could not be used in circuit breakers, the alternative of gapless metal oxide surge arrester protection was also effective for limiting the peaks of switching overvoltages. These technology developments led to the use of reduced phase-to-phase and phase-to-tower clearances on most EHV transmission lines, compared to the practice that had been common in electrical layout of HV transmission lines of 115 to 240 kV system voltage. Utilities obtained satisfactory switching-surge flashover operating experience on their EHV lines with closing resistors and surge arresters. As a consequence, these utilities then started to consider the possibility of using the same “compact” ratios of phase spacing to electrical strength on their HV lines. Many of the issues were summarized in 1978 [2][3]. Since that time, some utilities did make use of the extra margin in HV electrical dimensions to increase system voltage (uprating). Others have used compact dimensions on their new HV lines as a way to reduce visual impact and thus help gain public acceptance for new rights-of-way.

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A recent re-evaluation of the compact line design process [2] found it useful to refine the definition to consider more than just the ratio of phase-to-phase spacing to system voltage. Another measure, which is the restraint of phaseconductor motion, has proved to be fundamental to the reliable operation of compact lines. Phase spacing can be maintained better in adverse weather on compact lines with shorter spans, post insulators rather than strings, and interphase spacers. Either criteria, reduced phase spacing or conductor restraint, is now considered in the classification. Several aspects of compact line design directly related to lightning protection have changed in the thirty years between reviews [2,3]. These include the widespread availability of polymer insulators with suitable mechanical ratings for use as horizontal line posts and as tension braces, along with experience leading to mechanical test standards for inter-phase spacers. Transmission line surge arresters (TLSA) based on polymer insulator technology now offer an alternative for lighting overvoltage protection to overhead groundwires (OHGW). Utilities that carried out voltage upgrading projects, typically doubling the system voltage and at the same time improving ground resistance, reported fairly successful operating experience in general. However, these changes never improved the lightning backflashover performance, and in one case [3] an acceptable lightning outage rate on the old circuit was unacceptable when the uprated line took its new role in the power system. Direct lightning stroke protection with OGHW is still a feasible choice for compact lines, although the additional wires and tower height add to visual impact and provide additional risk of phase-to-OGHW contact during conductor motion. Effectiveness of OGHW depends largely on the lightning impulse impedance achieved at the base of each tower. In lines with reduced insulation, combined with narrow-base poles with reduced soil contact area, it can be difficult to achieve adequate backflashover performance without an extensive network of buried ground electrodes that is vulnerable to theft and damage and raises transferred potentials under fault conditions. Since 1978 [3], utilities have investigated the possibility of substituting transmission line surge arresters (TLSA) for OHGW as the primary protection against direct lightning flashes. This elimination of OHGW supports the main reason for selecting a compact design – improved visual appearance - by trading off fewer wires in the air with more complex towers. This alternative relaxes grounding requirements for backflashover performance, but instead calls for equipotential grading or other countermeasures at each tower to mitigate touch potentials from power system faults. The use of TLSA as a supplement to OHGW, rather than an alternative, is less desirable for compact lines but preserves the important role of OHGW in managing fault currents and touch potentials. The use of an underbuilt groundwire (UBGW) in combination with TLSA is discussed. The concerns with phase conductor sag and motion that normally argue against the use of UBGW are relaxed with compact designs. Also, an UBGW is more effective than a second OHGW for mitigating backflashovers on lines with vertical configuration.. Examples of UBGW applications on three continents are given in the case studies. 2 DEFINITIONS OF A COMPACT LINE There are several different definitions of a “compact transmission line”. One of the best is “A transmission line that looks like the common and acceptable distribution lines now running through urban areas”. The compact transmission line may feature: o Single-pole structures o Relatively short spans (30 to 150 m) o Armless or single-arm construction o Reduced separation between phases o Suitable for narrow right-of-way or roadside installation These features make the compact line designs suitable for use in congested urban areas. There are other benefits compared to multiple distribution lines that could provide similar power transfer. Rebuilding an existing circuit at a higher voltage level on the same right-of-way may be the only feasible way to increase capacity. The compact designs may have lower capital and maintenance costs, lower line losses, and may achieve important reductions in electric and magnetic fields. Compact design also tends to simplify some mechanical issues. When armless or single-arm construction with post insulators is selected, conductors are fixed at each pole. This simplifies tension stringing, thermal rating, wind loading and other design and construction issues. The short spans of a compact design reduce the magnitude of high-temperature sag, galloping and large-amplitude conductor motion. In many cases, line tension can be reduced, relaxing pole strength requirements and concerns about aeolian vibration damage.

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2.1 Phase-to-phase and phase-to-tower spacing AC flashover of rod gaps is a linear function of gap spacing in the range of 20 to 250 cm. Figure 2 shows the critical flashover at peak of ac wave is a high fraction (84%) of the positive critical lightning impulse flashover level (+LI). The ac strength for conductor-to-tower and other geometries is adjusted with the use of gap factors, as described in [1].

Critical Flashover: Positive Impulse +LI compared to AC Peak of Wave

Critical Flashover: Negative Impulse –LI compared to AC Peak of Wave

Fig. 2 – Rod-to-Rod Gap Critical Flashover Levels, adapted from [4]

Additional data from [4] are presented for the critical impulse flashover with nonstandard 1.2/5 μs test waves as well as for standard 1.2/50 μs waves. These are used below to exercise models of flashover using the Destructive Effect (DE) approach, and to optimize the DE parameters for the compact line analysis. For insulators [2], the Basic Impulse Level (BIL, kV) that causes a 10% probability of flashover at standard conditions for insulators of length S (m) is given by:

BIL+ = 125 + 539 ⋅ S

BIL− = 164 + 469 ⋅ S

(1)

+

The BIL values for insulators are rather similar to those for rod-to-rod +CFO values in Fig. 2, when the 1.28-σ statistical margin is considered. These values are relevant for backflashover calculations. There is a considerable reduction in negative impulse strength when insulators are present, compared to rod-rod gap values. This has a negligible effect, because nearly every shielding failure causes a flashover, either from first or subsequent strokes [5]. Typical ratios of phase spacing to ac flashover spacing declined from 10:1 for 138-kV lines to 9:1 for 345-kV systems, 7.8:1 at 500 kV and 6.2:1 for 765-kV lines in the USA [3]. These compactions of electrical layout took advantage of switching surge control measures such as closing resistors for the EHV systems. Figure 3 shows a variety of shielded and unshielded compact configurations, with vertical configurations that have significantly reduced horizontal extent and visual impact compared to the conventional 115-kV horizontal design. Polymer line posts such as those shown in the compact designs of Fig. 3 typically have a 6.5:1 ratio of system line-toground voltage to ac dry flashover level.

Compact 115-kV Configurations: Horizontal Unshielded; Horizontal Shielded; Vertical; Delta; Vertical Delta.

Standard 115-kV Transmission Line Configuration with 3.7-m Phase Spacing and 1.17-m Dry-Arc Distance

Fig. 3 – Typical Compact and Standard Transmission Line Configurations [2]

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Tight spacing among phases of compact lines may lead to challenges, including higher voltage gradients on conductors and insulators. These in turn may lead to unacceptable levels of audible noise, electromagnetic interference, and corona degradation of hardware and polymer insulator materials. If the layout of the line is too tight, it may become impossible to service the line or wash the insulators on potential, using live-line work methods. As conceived originally, compact lines also used reduced insulation levels that lead to decreased lightning performance. In some cases, compact lines and voltage uprating projects led to the use of unshielded designs that, like distribution lines, had no OHGW at all. These will have fewer direct flashes (from reduced height) but almost all will cause lightning faults, leading to elevated fault current at each tower base. The insulation length plays a relatively minor role in the visual impact, expressed by the line cross-section area at the tower. Instead, the visual and physical cross-section of a transmission line is a strong function of the span length. 2.2 Phase Conductor Restraint The restraint of phase conductor motion is the hallmark of a modern compact line design. This restraint is achieved with some combination of inter-phase spacers, rigid insulators at each pole and shorter span length compared to conventional transmission lines. These factors tend to make certain aspects of OHGW lightning protection more effective, while at the same time reduced foundation size of compact poles can make local grounding more difficult. Inter-phase spacers are usually considered to manage relative motion of conductors in design of compact lines. This is especially important in cold climates, where accretion of ice is nearly identical on large-diameter phase conductors and smaller-diameter OHGW. Since the change in weight per unit length is relatively larger for the small OHGW, they may sag and make contact with phases. Inter-phase spacers and spacers from phases to OHGW can hold the conductors apart under these static loads. On de-icing, a second benefit may be derived. Ice tends to fall off phase conductors first, and the cascade leads to upward “sleet jump” that can also cause conductor clashes. In addition, properly placed spacers can control the amplitude of galloping, a large-amplitude skipping motion that occurs in icing conditions and steady winds. Suspension insulators typically use 16-mm fiberglass rods to achieve maximum design tension ratings of 67 kN (15 kip). In contrast, the fiberglass rods are relatively weak in cantilever. This is a problem, because horizontal post insulators in the designs of Fig. 3 transfer conductor and radial ice dead-weight to the towers mainly as cantilever loads. There is thus a significant reduction in load-bearing capability as the length of the post and its impulse flashover level increase. One option is to increase the diameter of the fiberglass rod, and Fig. 4 shows that increases from 64 to 76 mm can double the cantilever capability. Another option is to improve stress transfer. Ongoing progress is seen in the improvement in cantilever rating from 2006 to 2008, using the same 64-mm (2.5”) fiberglass core diameter.

Fig. 4 – Maximum Design Cantilever Load of Typical Polymer Line Post Insulatosr Wtihout Brace

In rough terms, the ratio of typical suspension insulator tension rating to cantilever load of line posts will indicate the ratio of maximum span length. A line with suspension insulators can often achieve 400-m maximum span length using 67-kN suspension insulators of any BIL rating. If this line was constructed with line posts, using the premium products from Manufacturers 2 and 3 in Fig. 4, it would require (20 kN/67 kN)⋅400 m = 120 m spans with 500-kV CFO, or

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(10 kN/67 kN)⋅400 = 60 m spans with 1000 kV CFO. Both lines would qualify as compact designs, with about 2 m of sag for the 120-m span on the “115-kV class” line and less than 0.5 m of sag for the 60-m spans on the “230-kV class” line strung at typical transmission line tensions. However, the number of poles could be prohibitively costly. With limitations in cantilever load rating for line posts with high impulse CFO, it is common to make use of a tension brace to increase vertical load capability. Fig. 15, below, shows Urban Transmission Lines at the Brazilian utility, COPEL. Line posts of suitable cantilever rating were available at 138 kV, but a braced-post design was selected for similar span length at 230 kV. 2.3 Typical Span Lengths of Compact Lines The sag of conductors between attachment points is approximately related to the span length and tension [5] by: ACSR : Sag m = 0.0014 ⋅

Span m2 % RTS

Steel : Sag m = 0.0009 ⋅

Span m2 % RTS

(2)

Typical spans on conventional 115-kV transmission lines like those in Fig. 3 are 200-300 m long, with conductor tension approximately 18% of rated tensile strength (RTS) at 20°C. With high-temperature operation, this tension will fall to about 13% at 100°C or perhaps 11% at 150°C. Even with 200-m spans, the increase in sag is considerable at high temperature. Phase-to-ground clearance established by the height of the insulator at the pole in Fig. 3 provides the sum of required phase-to-ground electrical clearance (typically 6-8 m), the sag at 100°C or150°C (4.3 or 5.1 m), and the length of the insulator (1.2 m). Lightning protection, given by a single OHGW in Fig. 3, adds another 2.5 m to the line height for a total of 14-17 m. The minimum cross-section is (7.4 m wide x 14 m tall) = 104 m2, compared to roughly 20 m2 for the vertical configuration and 32 m2 for the delta configuration in Fig. 3. A compact 230-kV configuration such as the portal arrangement in Fig. 5, using a 105-m span length, was considered at one utility for a 5.6-km line section. However, 26 tapered compact steel pole structures, shown in Figs 5 and 6, were selected instead. Insulation distance was reduced from 14x146 mm standard discs (2.04 m) to 6x159 mm semiconductive glaze bell insulators that provided the same pollution performance.

Conventional Lattice Structure for 240-m Span

Tapered Compact Steel Pole Structure for 240-m Span

Concrete Portal Structure for 105-m Span

Fig. 5 – Alternatives for 230-kV Transmission Lines [2]

Compact lines that made use of reduced dry-arc distance for insulation, such as the one shown in Figs. 5 and 6, generally proved to have poor long-term reliability. The poor lightning performance caused by reduced insulation strength was well understood in the conception phase but was poorly accepted by line operators, who expected all 230 kV-class lines to meet the same security goals, of less than one outage per 100 km per year. It was thought in conception that improved contamination performance would be an acceptable trade-off, and for this reason semiconductive glaze bell insulators were developed and applied. In the long term, current density proved difficult to manage, leading to problems with integrity of metal-to-glaze contacts. Consequently, the 230-kV compact line in Fig. 5 and Fig. 6 was eventually rebuilt with conventional 14x146 mm glass suspension discs.

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Compact and Conventional 230-kV Lines Original Compact Design with 6x159 mm Semiconductive Glaze Bell Insulators (June 2006)

Compact and Conventional 230-kV Lines Rebuilt Subcompact Design with 14x146 mm Glass Discs (Aug 2009)

Fig. 6 – Compact and Conventional 230-kV Steel-Pole Lines in Ontario

It was necessary to extend the center-phase insulator arm in the rebuilt compact steel pole design at the right in Fig. 6, and to add a metal brace to achieve adequate ground and phase clearance for the 1843-kcmil (40.6-mm diameter) phase conductors. 2.3 Summary: The Modern Compact Line The updated definition of a “compact” transmission line has mechanically restrained phases, but no longer requires a reduced insulation length to achieve reductions in phase spacing. Some of the reasons for retaining a conventional dryarc distance include: o o o o

Retention of a lightning outage rate similar to that of a conventional line of the same voltage class Improving cantilever ratings of polymer line post and braced post insulators in the 1-2 m range Reduced electrical stress per mm of leakage distance for long polymer insulator life Retention of adequate electrical performance in icing and contamination conditions 3 LIGHTNING PERFORMANCE ISSUES FOR COMPACT LINES

If the dry-arc distance and critical impulse flashover level is held constant in the design, then the following features of compact transmission lines will lead to improved lightning performance: o o o

Reduction in the average spacing from OHGW to phase conductor Decrease in span length Decrease in average line height

Rigid insulators also solve many of the reliability problems related to flexible connections to transmission line surge arresters (TLSA), and enable the use of externally gapped (EGLA) configurations that tend to be more robust. This makes the alternative of TLSA protection, instead of OHGW, more feasible, offering additional reduction in line height and improved appearance, except at the pole where extra components are apparent. Compact transmission lines should be insulated sufficiently that the rate of flashover from induced overvoltages is negligible. Direct stroke protection can be achieved through the use of overhead groundwires (OHGW), diverting most of the flashes safely to ground, or through the use of transmission line surge arresters (TLSA) on the uppermost phase.

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The TLSA protection will divert stroke currents to ground and limit overvoltage, but will not mitigate the effects of stroke charge on metallurgical damage to the phase conductors. 3.1 Induced overvoltages Lightning flashes that terminate near, but not directly on, a compact transmission line may induce overvoltages of significant magnitude. The peak of the overvoltage is a function of the lateral distance to the flash, the peak stroke current, the return stroke velocity and the local soil resistivity. The overvoltages induced from nearby lightning [6][7] tend to have short times to half value, approximately 5 μs compared to 50 μs for a standard lightning wave [4]. Above, Fig. 2 showed that the CFO for this nonstandard 1.2/5 wave was some 30-37% higher than the 1.2/50 level, depending on rod-to-rod gap distance. There are similar increases in the flashover strength of insulators, which will be discussed in detail for backflashover modelling. The use of a minimum positive lightning impulse (+LI) CFO of 300 kV for a standard 1.2/50 wave is generally sufficient to mitigate induced overvoltage flashovers in areas with low soil resistivity, ρ420 kV. For example [9], the wood pole path in Fig. 7 adds about 210 kV CFO per meter of path length, and the fiberglass standoff path adds another 200 kV/m. If fiberglass poles are used instead of wood, they add 410 kV/m between insulator bases. 3.2 Shielding failures The IEEE [5] recommends that transmission lines with OHGW be designed for a nonzero failure rate on the order of 0.05 shielding failure flashovers per 100 km of line length per year. This design level calls for different shield angles in areas of different ground flash density and for lines of different height. The design curve for compact lines in Fig. 8 [2] relies on a model of stroke incidence, N=Ng/10⋅(28h0.6+b) [10] where N is the number of flashes to the line (per 100 km per year), Ng is the ground flash density (flashes /km2/year), h is the height of the topmost wire at the structure (m), and b is the spacing between topmost wires. For the vertical configuration in Fig. 7, b=0. There are a number of ways to calculate the lateral attractive radius of a conductor, with [11] providing a preferred model from switching-surge physics that agrees closely with predictions of stroke incidence from [10].

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Fig. 8 – Fraction of Shielding Failures versus OGHW Shield Angle [2]

In areas with no reliable measurements of ground flash density, Ng can be estimated from optical transient density OTD using Ng ≈ 0.3 OTD where both are in flashes /km2/year. For a region with Ng=6 flashes / km2 / year, a compact line of vertical configuration with tower height of 20 m will receive about 100 flashes per 100 km per year. To achieve a shielding failure flashover rate of 0.05 per 100 km per year, the ordinate in Fig. 8 must be less than 0.0005%, calling for a shielding angle of θ< 30°. The attractive radius models of the form suggests that freestanding structures can provide direct stroke protection of lines for a circular zone centered at the top of the structure. If the span lengths are sufficiently short, several researchers have suggested it may be possible to achieve direct-stroke protection using tower masts without overhead groundwires. Using [11], the approximate ratio of attractive radius Ra from a mast to the lateral attractive distance of a conductor Da can be read out from Fig. 9.

Attractive Radius of Freestanding Mast, height h

Lateral Attractive Distance of Horizontal Conductor, height h Da=1.57⋅ I0.69⋅ h0.45

Fig. 9 – Attractive Radius of Masts and Conductors [11]

For the low currents of interest for shielding calculations, the ratio Da/Ra is about 0.75, and this leads to a simple result that the length of a span totally protected by the towers, without OHGW, is approximately 1.33 Ra.

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0.66⋅Ra

Ra

Da= 0.75⋅Ra

Fig. 10 – Overhead View of Line Protection from Freestanding Structures

To achive a 1% proportion of shielding failures, a prospective current of 5 kA is appropriate. For a line height of 20 m, the relevant attractive radius is also 20 m, and this means that span lengths less than 27 m would be needed. Wider spacing of up to 40 m could be considered if each pole had a considerable vertical extension. 3.3 Midspan flashovers A transmission line in Champion, NY used line surge arresters to improve outage rates. As part of the research, this line was monitored closely with lightning cameras to find fault locations. Follow-up studies showed several places where burn marks in covered conductor indicated midspan phase-to-phase phase lightning flashovers. It may be prudent to evaluate the corona radius around a stricken conductor, using for example a radial gradient of 1500 kV/m at the edge of the corona envelope [5]. Corona normally increases common-mode voltage coupling to nearby phases and reduces stress on insulation. However, if the corona envelope expands past the nearby conductors, they may be increasingly prone to midspan flashover faults. 3.4 Backflashovers Even if lightning is successfully attracted to one of the poles in Fig. 10, there remains an important risk of flashover. In this simplified case, the stroke current will flow down the pole and into the ground. The thin wire connecting pole top to base will introduce a series inductance of about L=1.5 μH/m. The ground electrode, usually consisting of a driven vertical rod, will have a resistance to remote earth, Rf. The potential rise on the ground system will rise as a function of the impressed current and its rate of rise to V=RfI + L dI/dt at the top of the pole and its connection to the base of the uppermost insulator. Simplified models of lightning [5][12][14][18] assume that it is a current source, and its rate of current rise dI/dt can be represented by a 2-μs ramp, rising from 0 to I in 2 μs. This simplifies the circuit model. During the rising phase of the ramp, the voltage at pole top is given by V=I⋅ (Rf+L/2μs). For Rf of 20 Ω and a 30-μH inductance of a 20-m pole bond, the total is V=I⋅ (20 + 15), giving a wave impedance Zwave= 35 kV/kA at crest of wave. After the peak, the dI/dt is zero and V=I⋅Rf. Generally, the insulation strength of post or suspension insulators – vertical, horizontal, or braced – is a linear function of the smallest dry-arc distance between energized and grounded points, including corona rings and arcing horns. When negative lightning flashes strike the pole, the positive impulse strength of the insulator is tested. Eqn. 1 gives a 90% withstand strength (BIL) of 664 kV for S=1 m, corresponding to a +LI CFO of 639 kV. Backflashover will occur in this simple case whenever the flash current I exceeds a critical current Icrit = BIL / Zwave. For the flash to the pole, the critical current will be Icrit=639 kV / 35 kV/kA = 18 kA. So, while the 20-m poles spaced at 27 m can capture 99% of the lightning (all of the currents exceeding 5 kA), any current above 18 kA will cause a backflashover. The probability that a first negative return stroke will exceed Icrit and cause a flashover is P(I>Icrit)=1/(1+(Icrit/31 kA)2.6) [5]. In the example, this probability is high - 80%. Of every 100 flashes to the line, on average. one will be less than 5 kA and cause a shielding failure, and 80 will exceed Icrit=18 kA and cause a backflashover.

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In general, the backflashover process for transmission lines is very sensitive to the insulation level. The benefits of including fiberglass stand-offs (200 kV/m) and wood pole path (210 kV/m) [9] were shown in Fig. 7. For a wood path of 1 m and 1-m standoffs, the +LI CFO will increase from 639 kV to about (639 + 200 + 210) or 1049 kV. This will raise the critical current in the simple calculation to Icrit=1049 kV / 35 kV/kA = 30 kA. The probability of exceeding this peak current is only 52%. The simple modification will cut the outage rate by approximately one-third. 4 DIRECT STROKE PROTECTION OF COMPACT LINES WITH OHGW Overhead groundwires (OHGW) located above the phase conductors provide four important benefits. They attract flashes, especially away from the towers, that would normally terminate on phase conductors and cause flashovers. They divert a fraction of the lightning surge current to adjacent towers. The voltage rise from the product of this small fraction of current and the OHGW surge impedance illuminates the phase conductors, and the coupled voltage on the phases subtracts from the tower-to-insulator impulse stress. The current fraction reaches adjacent towers rapidly (at the speed of light) and flows into adjacent pole grounds, sending back a cancelling reflected voltage wave that reduces the insulation stress at some delay in time. Finally, if there is a flashover, the OHGW provide a well-controlled 60-Hz impedance that improves the speed of fault identification, allowing faster tripout of the circuit breakers and reducing the hazards of touch potentials near towers and equipment damage from sustained short-circuit duty. 4.1 Statistics of first-stroke peak current, rise time and flash incidence First, negative, downward lightning flashes are the most frequent and tend to have similar characteristics from region to region. The distribution of peak stroke current magnitude is so wide, with values ranging from 2 to 200 kA, that it is common to assume that the logarithm of peak current is normally distributed. Where a standard deviation σ=1 for a distribution with mean of i=31 suggests that 68% of the values are within ±1 (30