Due to the possible failures they can present and risk of collateral impacts, bushings can reduce the life expectancy and reliability of power transformers and shunt reactors. It is therefore important to establish an evaluation strategy to monitor and replace these important components in a proactive manner.

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Global expansion of power grids is driving a substantial increase in deployment of new transmission and distribution components, particularly high and ultra-high voltage cables. To ensure reliable design and manufacturing of these cables as well as their accessories, international standards require rigorous type tests.

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This edited contribution to INMR by Dr. Aderibigbe Adekitan at Tridelta Meidensha in Germany examines the issue of uneven voltage distribution across metal-oxide varistor blocks, as commonly associated with tall arresters, and which can lead to thermal stress on the metal-oxide varistor blocks in the upper section of the surge arrester.

In his investigation, finite element analysis was applied to evaluate the voltage distribution across two- and three-unit 435 kV rated surge arresters. Application of a grading ring reduced the U/Uc ratio from 1.22 to 1.18 for the two-unit arrester, indicating an improved voltage distribution. Results reveal that proper arrangement of multi-unit arresters and uniform distribution of the metal-oxide blocks within the arrester can lead to a more uniform voltage distribution. These findings also emphasize the importance of ensuring adequate design considerations in development and manufacture of surge arresters.

Guangzhou MPC Power International Co. Ltd.

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Electrical power systems must be adequately protected to ensure consistent reliability and service longevity. Lightning strikes, switching operations, which are an integral part of power station processes, and other related disturbances can trigger overvoltage transients in electrical systems. Surge arresters are installed to protect equipment from the negative impact of these undesired but sometimes unavoidable system overvoltages and are crucial for maintaining the integrity of electrical systems.

Surge arresters protect insulators, such as bushings, in power systems by clamping surges that can exceed the basic insulation level (BIL) of the protected device, thereby reducing insulation stress and preventing flashover. By implication, risk of thermal or mechanical damage is minimised, thereby enhancing the overall reliability of the power system component. The advent of ZnO blocks led to the development of the first non-gapped metal-oxide surge arrester (NGMOSA) in 1976. Over the years, this concept gained popularity and acceptance across various utilities and industries worldwide. ZnO arresters can handle high non-linear energy and are more reliable than silicon carbide (SiC).

Surge arresters are designed to comply with relevant standards for optimal performance. Design, construction and testing of the NGMOSA are regulated by standards such as IEEE Std C62.11-2020 and IEC 60099-4. The parameters of a surge arrester determine the suitable system voltage and areas of application of the arrester.

As illustrated in Fig. 1, a surge arrester designed in compliance with current industry practices must demonstrate certain capabilities in terms of continuous voltage withstand, energy and charge withstand rating, short-circuit rating, ageing and environmental durability, mechanical strength, etc. Each of these performance evaluation criteria is like a continuous link in a chain that must not be broken for prolonged service reliability. Compliance with these requirements is established during type testing of an arrester design series and routine testing of a newly produced arrester.

Metal oxide varistor blocks (MOVBs) are the primary components in surge arresters. An MOVB is a semiconductor device primarily composed of zinc oxide (ZnO) grains embedded in a polycrystalline matrix. Due to its nonlinear voltage-current characteristics, it is highly effective for surge protection. A stack of MOVBs achieves the protective function of a surge arrester.

Number of blocks ultimately determines the total block height, defines the continuous operating voltage of the arrester, residual voltage under impulse currents and its temporary overvoltage (TOV) withstand rating. Width or diameter of the block defines its current handling capacity, and ultimately, these parameters determine the energy absorption rating of the blocks. In some cases, achieving a desired residual voltage or energy rating may require installation of MOVBs in parallel if a suitable single, larger-diameter block is unavailable or impractical.

To achieve higher voltages, multiple surge arrester units are coupled together vertically, and MOVBs at the top of the uppermost unit are often subjected to more voltage stress than the rest of the blocks in the arrester column due to stray capacitances to the earth acting on the arrester column, resulting in a non-uniform voltage distribution and faster ageing. A conductive ring, referred to as a grading ring, which is placed near the high-voltage end of surge arresters, becomes necessary in such cases to improve the voltage distribution across the blocks and to prevent excessive voltage stress on MOVBs stacked at the top of the column, which could lead to thermal runaway and arrester failure.

Height of the MOVBs determines rated voltage of a surge arrester. MOVBs have a fixed ratio between their rated and residual voltage, and this often requires a trade-off to achieve the desired protection level. A surge arrester with a specific rated voltage consists of a certain number of blocks based on the established manufacturer’s dimensioning factors. It is important to evaluate the effects of the position and arrangement of MOVBs on the stress distribution within the arrester.

Stress distribution can vary significantly for the same number of blocks if placed at different physical positions in the stack. Different tools, such as finite element method (FEM), Maxwell, COMSOL Multiphysics, boundary element method, EMTP-ATP, have all been used in various studies on surge arresters. FEM has been applied to study the voltage distribution across a surge arrester and for other purposes, including the investigation of the electrothermal attributes of surge arresters.

FEM simulations require significant computational resources, particularly for 3D models with fine meshes. Models with finer mesh elements improve accuracy, but increase simulation time and memory requirements, as observed in studies on grading ring optimization.

A comparative surge arrester simulation study in IEC 60099-4 using both 2D and 3D computations established the sufficiency and accuracy of the 2D model with a significant simulation time advantage. This study assessed the practicality of incorporating a ‘virtual’ grading ring in axisymmetric simulations to replicate the effect of grading ring supports, and the application of other model simplifications proposed by IEC 60099-4. FEM is a vital tool for modelling and optimizing surge arrester designs and provides detailed insight into voltage distribution, material performance, and impact of environmental factors.

Ongoing research is being conducted to further improve the design, performance, and reliability of ZnO-based arresters. A key area is the application of high-gradient ZnO materials in gas-insulated surge arresters for gas-insulated switchgear (GIS) applications. High-gradient MOVBs have higher electric field stress withstand and can therefore operate at higher voltage gradients and higher voltage ratings per block. The associated high electric field is not a challenge because of the insulation withstand of the sulphur hexafluoride (SF6) gas in GIS arresters. This innovation extensively minimizes height of GIS surge arresters, improving performance against fast transients and enhancing voltage distribution.

Research studies are being conducted to develop new high-gradient ZnO materials doped with rare earth oxides. These materials also reduce grain size of ZnO varistors, leading to higher voltage gradients and compact arrester designs. Air-insulated surge (AIS) arresters only have low-dielectric air as insulation and can only use standard gradient MOVBs. Optimal arrester design and ZnO block arrangement within the arrester, coupled with height reduction where necessary, are required to ensure good voltage distribution for AIS arresters.

Recent studies have further evaluated the ability of arrester materials to handle very fast transients, such as those caused by high-altitude electromagnetic pulses (HEMP). Continuous research in this field will help shape the design of future surge arresters.

This study presented below investigated voltage distribution across a 435 kV-rated polymer tube design surge arrester and highlights the impact of different MOVB arrangements as well as improved voltage distribution from using grading rings.

Methodology

This study applies a FEM-based approach using FEMM 4.2 to simulate a 435 kV-rated, polymer-tube-design surge arrester with a continuous operating voltage (Uc) of 348 kV rms or 492.15 kV peak. The maximum system voltage is 550 kV. The capacitive arrester model simulation was performed using FEM, while the capacitive model with nonlinear resistive elements, which provides more realistic results, was implemented using ATP EMTP, and the final data review and analysis were performed in MATLAB.

Modelling and analysis of the arrester, including the MOVBs and metal spacers, is based on guidelines and the resistance-capacitance arrester model circuit provided in Annex F of IEC 60099-4:2014.

A surge arrester installed between the phase and the earth operates normally at the nominal phase-to-earth voltage level. The surge arrester can also safely continuously operate at the Uc as illustrated in the sample VI curve in Fig. 2. This study evaluates the variation in the field distribution across the arrester at these two voltage levels. For a particular arrester design series, the number of MOVBs to achieve a specific voltage is established based on the type test results.

Two versions of the arrester are considered in this study: a two-unit arrester with a total height of 4240 mm, where each unit has the same height and number of blocks; and a three-unit version with a total height of 4560 mm, consisting of two identical units with 15 blocks each and a third, longer unit with 34 blocks. Effect of altering position of each unit in the three-unit arrester is investigated.

This study investigates the effect of distributing the same number of ZnO blocks differently in the arrester. It considers MOVBs concentration; uniformly, at the top, at the bottom, at the upper and lower ends (top and bottom) and at the centre as illustrated in the simplified drawing in Fig. 3, from (a) to (e). IEC recommends including the line conductor in the model. This study examines the effect and the level of improvement achieved by incorporating a grading ring with a diameter of 1500 mm in the arrester. The arrester has 64 ZnO blocks stacked vertically in the housing, and the relative permittivity of the MOV block is 800. The pedestal is 2000 mm high and 150 mm wide.

Results

A. Effect of Grading Ring
The first FEM simulation highlights the beneficial effect of a grading ring on a surge arrester operating at the continuous operating voltage. The 435 kV rated arrester with 2 similar units is analysed first without a grading ring and then with a 1500 mm diameter grading ring. The normalized potential distribution across the surge arrester with and without a grading ring is shown in Figs. 4(a) and 4(b). The ratio of the voltage (U) across each block from the FEM simulation divided by the average Uc per block for both cases is plotted in Figs. 5(a) and 5(b).

Without applying a grading ring, maximum U/Uc ratio is 2.18 for the capacitive model and 1.22 for the capacitive-resistive model. With grading ring, values were reduced to 1.15 for the capacitive model and 1.18 for the capacitive-resistive model. These values and the voltage equivalent can be compared with the unbalance factor established during the long-term stability test under continuous operating voltage. This helps confirm if any of the blocks are over-stressed beyond a reasonable limit.

Voltage stress across the MOVBs from the base upward is plotted in Fig. 6. The analysis without a grading ring resulted in a maximum stress of 77.4% per meter with quite a non-uniform voltage stress distribution with peaks around the upper blocks in the second arrester units. Application of a grading ring improved the voltage stress pattern and reduced the maximum value to 40.9% per m.

A similar pattern is observed in Fig. 7, with the bulk of the voltage distributed along the MOVBs in the upper unit for the case without a grading ring. These results show the significance of the stray capacitance on the voltage distribution along the arrester without a grading ring and the beneficial improvements obtained by the introduction of the grading ring. This study does not include further grading ring optimization analysis.

The initial simulation applied a Uc of 492.15 kV peak. In this section, the peak phase-to-earth voltage of 449.1 kV is applied for a comparative review as illustrated in Fig. 2. This represents the peak nominal phase-to-earth system voltage (Uph) that the arrester should be subjected to.

The result for this case is presented in Fig. 8(a) for the capacitive model and Fig. 8(b) for the capacitive-resistive model. There is no difference in the U/Uc distribution for the capacitive model when Uph and Uc were applied, both with and without a grading ring. The result from the capacitive-resistive model is quite different, with a unique data trend for each of these cases.

Coefficient of variation, as shown in Table 1, reveals interesting trends in dispersion. U/Uc distribution from the application of Uc and a grading ring has data points closest to its mean value, while the U/Uc distribution from the application of Uph without a grading ring has data points farthest from its mean value. For this surge arrester under study, the result shows that a lower system voltage, such as Uph, does not necessarily imply a better U/Uc distribution than a higher system voltage, such as Uc. However, actual voltage U per block can be higher for each of the MOVBs for the higher system voltage.

B. Effect of MOVB Arrangement Within Housing
This section examines the effect of concentrating the MOVBs in different sections of the arrester housing. Five different block arrangements are analysed: uniform block distribution, concentration at the top, concentration at the bottom, concentration at both the top and bottom, and concentration at the centre, as illustrated in Fig. 3. A grading ring is included in the simulation, and the applied voltage is Uc.

Peak Uc is applied for the 5 MOVB configurations. U/Uc distribution for the capacitive model and the resistive-capacitive model is presented in Figs. 9 and 10, respectively. Coefficient of variation, which is a measure of dispersion, is presented in Table 2 for the capacitive model. It shows that the concentration of blocks at the top ranks first, the uniform distribution ranks third, and the concentration of blocks at the bottom ranks fifth and last in the dispersion analysis.

Table 3 presents the results for the capacitive-resistive model, which represents the more realistic solution. It shows that the uniform distribution of blocks ranks first, while the concentration of blocks at the top, which previously ranked first in the capacitive analysis, now ranks fifth and last in the dispersion analysis. The concentration of blocks at the extreme ends (i.e. at both the top and bottom) ranks second both in the capacitive and capacitive-resistive models.

The uniform distribution of blocks has the lowest maximum U/Uc ratio of 1.15 and 1.18, respectively, for both models. Voltage stress across the MOVBs from the base upward is plotted in Fig. 11. Peak voltage stress is 40.9% for the uniform configuration, 42.7% for the top, 45.4% for the bottom, 41.7% for the top and bottom and 42.2% for the centre configuration. The uniform MOV block arrangement has the best voltage stress distribution. Fig. 12 shows the voltage distribution for the 5 MOVB configurations.

C. Effect of Interchanging Arrester Units
This section presents results from the simulation of a three-unit version of the 435 kV rated arrester with a height of 4560 mm. The voltage applied is Uc. The longer unit is referred to as A, while the second and third units, B and C, are identical in height and contain the same number of MOVBs. The normalised potential distribution across the 3-unit surge arrester in the ABC configuration is shown in Fig. 13. The ratio of the voltage (U) across each block from the FEM simulation divided by the average Uc per block is plotted in Fig. 14(a) for the capacitive model and 14(b) for the capacitive-resistive model for the three configurations ABC, BAC and BCA considered.

The simulation reveals a consistent order of results for both the capacitive model and the capacitive-resistive model. The ABC configuration with the tallest unit at the top has a maximum U/Uc ratio of 1.10, while the BAC configuration with the tallest unit at the centre has a maximum U/Uc ratio of 1.29, and BCA configuration with the tallest unit at the bottom has a maximum U/Uc ratio of 1.39, representing the worst voltage distribution among the three cases. A similar result pattern is observed for the capacitive-resistive model with U/Uc ratios of 1.16, 1.19 and 1.20, respectively. This indicates that placing the unit with the highest number of blocks at the top resulted in the best voltage distribution for this case study.

Voltage stress along the surge arrester is shown in Fig. 15 for the three configurations. The peak voltage stress is 39.1% per meter for the ABC arrangement, which exhibits a more uniform voltage stress distribution compared to the other two configurations. For the BAC configuration, the peak voltage stress is 45.8% per m, while for the BCA configuration, it is 49.2% per m.

Voltage distribution based on the position of the arrester units is plotted in Fig. 16. The voltage distribution for the ABC configuration has a steeper voltage distribution across the lower blocks, implying a better voltage distribution across the blocks than the BAC and BCA configurations.

The key findings include:
a. The voltage distribution across tall surge arresters can be significantly improved by using grading rings. The grading ring reduces the voltage stress across the ZnO blocks in the upper region of the surge arrester;

b. A measure of dispersion using the coefficient of variation for the capacitive-resistive model revealed that a lower system voltage, such as Uph, may not necessarily imply a better U/Uc distribution than a higher system voltage, for example, Uc. However, the actual voltage U per block may be higher for each ZnO block for a higher system voltage;

c. A uniform distribution of blocks within the arrester housing helps achieve an optimal U/Uc distribution;

d. For multi-unit arresters, the longest unit with the highest number of ZnO blocks should be positioned at the top of the arrester to achieve the best voltage distribution.

Conclusions

Surge arresters are vital for protecting electrical equipment and insulators from switching and lightning impulses in electrical systems. Tall arresters can experience uneven voltage distribution due to stray capacitances to earth.
This FEM study confirmed that application of a grading ring helps in redistributing electric field more uniformly across the arrester and minimizing voltage stress at the top end of the arrester. Applying a grading ring to a 2-unit arrester reduced maximum U/Uc ratio from 1.22 to 1.18 for the capacitive-resistive model.
The study also evaluated and confirmed the impact of 5 different MOVB arrangements within the arrester units. During production, it is important to ensure that MOVBs are uniformly distributed across the length of the arrester to achieve optimal voltage and electric field distribution.

References
1. Huang, S.-J. and C.-H. Hsieh, A method to enhance the predictive maintenance of ZnO arresters in energy systems. International Journal of Electrical Power & Energy Systems, 2014. 62: p. 183-188.
2. Raju, K., et al., Development of high gradient ZnO arrester material for high voltage applications. IEEE Access, 2020. 8: p. 115685-115693.
3. Latiff, N.A.A., et al., Measurement and Modelling of Leakage Current Behaviour in ZnO Surge Arresters under Various Applied Voltage Amplitudes and Pollution Conditions. Energies, 2018. 11(4): p. 875.
4. IEEE C62.11: IEEE standard for metal-oxide surge arresters for AC power circuits (>1 kV). 2020.
5. IEC 60099-4: Surge arresters – Part 4: Metal-oxide surge arresters without gaps for a.c. systems. 2014, Ed. 3.0.
6. Meshkatoddini, M.R., Metal oxide ZnO-based varistor ceramics, in Advances in ceramics – electric and magnetic ceramics, bioceramics, ceramics and environment, C. Sikalidis, Editor. 2011, IntechOpen: Rijeka.
7. Prasad, V., A Review of Voltage Distribution on Metal Oxide Surge Arrester and Suggestions for Improvement in High Voltage Applications. IEEE Latin America Transactions, 2025. 23(6): p. 479-486.
8. Alti, N., A. Bayadi, and K. Belhouchet, Grading ring parameters optimisation for 220 kV metal-oxide arrester using 3D-FEM method and bat algorithm. IET Science, Measurement & Technology, 2021. 15(1): p. 14-24.
9. Jinliang, H., et al., Potential distribution analysis of suspended-type metal-oxide surge arresters. IEEE Transactions on Power Delivery, 2003. 18(4): p. 1214-1220.
10. Meng, P., et al., Breakdown phenomenon of ZnO varistors caused by non-uniform distribution of internal pores. Journal of the European Ceramic Society, 2019. 39(15): p. 4824-4830.
11. Nurul, A.A.L., et al., Parametric Evaluation of 11kV Zinc Oxide Surge Arrester using Finite Element Analysis Model. IOP Conference Series: Materials Science and Engineering, 2021. 1127(1): p. 012038.
12. Waghmare, V.V., V.K. Yadav, and I.M. Desai. Optimization of Grading Ring of Surge arrester by using FEM method, PSO & BAT Algorithm. in 2022 2nd International Conference on Advance Computing and Innovative Technologies in Engineering (ICACITE). 2022.
13. Zhang, C., et al. Electric field analysis of high voltage apparatus using finite element method. in 2010 Annual Report Conference on Electrical Insulation and Dielectic Phenomena. 2010.
14. Tighilt, F., A. Bayadi, and A.M. Haddad. Voltage distribution on ZnO polymeric arrester under pollution conditions. in 45th International Universities Power Engineering Conference UPEC2010. 2010.
15. Seyyedbarzegar, S.M. and M. Mirzaie, Application of finite element method for electro-thermal modeling of metal oxide surge arrester. Computer Applications in Engineering Education, 2015. 23(6): p. 910-920.
16. Kannadasan, R., P. Valsalal, and R. Jayavel, Performance improvement of metal–oxide arrester for VFTs. IET Science, Measurement & Technology, 2017. 11(4): p. 438-444.
17. Bowman, T.C., T. Kmieciak, and L.B. Biedermann, Nanosecond Transient Validation of Surge Arrester Models to Predict Electromagnetic Pulse Response. IEEE Transactions on Electromagnetic Compatibility, 2025. 67(1): p. 286-294.

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But all this changed by the mid to late 1990s and today these types of insulators account for over half the world market. Competition and production volumes have soared and, with this, acquisition costs have typically become less than for porcelain and glass counterparts.

One of earliest applications for composite insulators was as insulating cross-arms, which are indispensable for design of compact lines and so-called aesthetic towers. The former, in particular, are rapidly gaining ground as an alternative to building traditional lines due to higher public acceptance. Moreover, composite insulators play a growing role in cases of line uprating to increase power transfer capacity of existing lines.

This edited past contribution to INMR by Dr. Konstantin O. Papailiou, current CIGRE President, explained the necessary properties of composite insulators as well as examples of such applications.

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Options for Line Compaction

The basic idea behind line compaction is to suppress horizontal movement of the classical suspension string. This way, line supports can become more slender and, at the same time, the right-of-way dimensions needed are reduced. Over time, four different insulator arrangements have come to be used for line compaction: V-strings; horizontal posts; suspended posts; and insulated cross-arms. These four arrangements are shown in Figs. 1 to 4.

Mechanical Design

Fig. 5 shows the loads that act on an insulated cross-arm. These are:

• vertical loads, V, from the conductor and from ice, if present;
• horizontal loads, H, from wind and, in the case of light-angle supports, from angular pull; and
• longitudinal loads, T, possibly from non-uniform conductor tension in adjacent spans or from a conductor failure – a rare exceptional load.

The vertical loads are taken up largely by the brace, depending on the angle, between the brace and post. By contrast, horizontal loads acting in compression, load the post in buckling. The insulator forces, i.e. the compression force, P, on the post and tensile force, B, on the brace are calculated, assuming T = 0, using the following formulas:

Rigid Connection

For voltages up to 245 kV, the post is often rigidly connected to the support (as in Fig. 6). CIGRE WG 22-03 used commercial finite element software to calculate the loading diagram, also called the application curve, for a 63 mm post of 2000 mm length with inclination angle to horizontal of 15° (see Fig. 8). Coupling angle of the 16 mm brace to tower was 45°, this brace being assumed to pivot at either end.

The load on the brace should not be negative (compression) so as to prevent buckling of the brace that would lead to contact between the metal fittings of the two insulators. With a horizontal angle of the post insulator of 15°, as used here, this condition leads to the inequality: V > H tan 15°. In this diagram, the lower straight line corresponds to the equality V = H tan 15°, i.e. the brace is not loaded along this line or, in other words, the insulated cross-arm should not ‘work’ below this line. The upper straight line in the diagram extends parallel to the lower straight line and corresponds to maximum allowable tensile load of the brace. It is good practice to use a so-called fail-safe base for the post, which will be plastically deform in case of overload thus protecting the more sensitive – and more expensive – post.

Pivoted Connection

This is the most commonly used arrangement, particularly for higher voltages, since it has a high level of mechanical strength and is also fault-tolerant in case of transverse loads (see Fig. 8). The composite long rod in the brace is loaded purely in tension and can be easily dimensioned. By contrast, the post is loaded in compression and, since it is articulated at both ends, can be calculated as a Euler beam with maximum tolerable compression load, i.e. the Euler load:

where E is the modulus of elasticity, I the moment of inertia and L the length of the post insulator’s fiberglass core rod.

It must be noted, however, that, due to an often-unavoidable eccentric application of compression load, the post insulator is additionally subjected to bending. The negative influence of this eccentricity can be seen in Fig. 9, where results of buckling tests performed on 63 mm diameter rods are plotted versus maximum buckling load given, as described above from the conventional buckling formula. Considerable reduction can be seen in measured failing load compared to theoretical buckling load and this must be considered in design. Since this and other effects, e.g. partial rigidity of the pivots, cannot reliably be modeled, it is recommended that the insulating cross-arm be type tested before putting it into service (see Fig. 10).

Stability Issues

One particular advantage of pivoting insulated cross-arms is their ability to stabilize in the event of sudden conductor movements. Such movements could take place should temporary differential line tension occur at the tip of a cross-arm on an overhead line section consisting of a number of spans. These movements could be caused by gusts of wind, irregular icing, spans of considerably different length (e.g. as in mountainous terrain) and short circuit forces. Particularly in the case of long cross-arms for higher voltages, it may be that cross-arms become de-stabilized, leading to considerable deflections with associated reduction in safety distance between conductor and tower. In extreme cases, the cross-arm could fail mechanically.

If there are differences in horizontal line tensions in two adjacent spans, the tip of the cross-arm will move toward the higher tensile load. In this case, if angle of rotation of the insulated cross-arm is inclined to the vertical, the tip of the cross-arm is physically raised and vertical line loads generate restoring torque. This causes the equilibrium of forces in the line direction to be re-established. Therefore, the most important design parameter to avoid stability problems is an adequate inclination angle, of the cross-arm’s rotation axis (see Fig. 6). Values of about 20° have proven to offer a good compromise. These types of stability issues have been examined extensively. As mentioned, stability depends greatly on wind speed perpendicular to the line and improves with decreasing number of spans as well as with an increase in the inclination angle, of the cross-arm (see Fig. 11).

In general, wind stability of compact lines with insulated cross-arms can be improved using the following measures:

• Increasing inclination angle, or the angle, between brace and post;
• Increasing vertical loads on the cross-arm, e.g. by adding weights;
• Increasing conductor tension;
• Reducing individual span lengths and/or number of spans within a line section;
• Reducing line angle in angle towers;
• Using ‘stabilizing cross-arms’ for long line sections.

Line Uprating

Because of increasing difficulty to secure approval for new transmission corridors, technologies for line uprating have been more and more developed. The basic goal is to use existing line corridors for transferring more power. One option is to use high temperature low sag (HTLS) conductors to increase current. Another is to modify tower top geometry in to increase voltage. In both cases, composite insulators have offered interesting new opportunities.

Application Examples

First 400 kV Compact Line

The first 400 kV compact line was installed in Switzerland in 1998. The line became necessary since the existing 125 kV line (see Fig. 12a) was not sufficient to carry increasing power demand in the region of Lake Geneva and had to be replaced with a 400 kV line. Conventional lattice towers normally used for this type of line would require considerable right-of-way, which was not available in a certain location close to buildings. The solution was to design slender 2D towers and apply insulated cross-arms with composite insulators for the 400 kV circuits. These could also support two single-phase 132 kV circuits for feeder lines of the Swiss railway (see Fig. 12b). It is worth noting that FRP core rods available at the time were restricted to 76 mm diameter, which could not take the compression loads. So, hollow core composite insulators had to be used.

Hybrid AC/DC Line

Conversion from AC to DC offers advantages for long distance transmission. In particular, conductors can be utilized up to their thermal limit in DC whereas in AC surge impedance loading (SIL) often becomes the limiting factor. In a pioneer project in Germany, one of the two circuits of a 380 kV AC line has been converted to a ±400 kV DC bi-pole with the third conductor used for the metallic earth return (see Fig. 13). A major prerequisite for this project was to use the same conductors and same towers. This would only be possible by replacing existing porcelain strings with composite insulators since these have significantly better contamination performance. As such, their length can be accommodated within existing tower top geometry.

Conversion of 245 kV to 420 kV AC

An interesting project in Austria saw two-fold line uprating. On one side, the original ACSR conductors were replaced by newly designed expanded conductors, i.e. conductors with the same amount of aluminum but with larger diameter to increase current carrying capacity. On the other side, the metallic cross-arms on the lattice steel towers were replaced by insulating cross-arms (see Fig. 14), which enabled increasing line voltage from 245 kV to 420 kV.

CompactLine

This new concept, developed in Germany, utilized existing 220 kV corridors but with a new design of 380 kV low visibility profile pylon. In order to achieve this, conductor sag had to be limited and this was made possible by suspending the quad bundle conductors to a tightly strung high strength steel rope, as used in cable cars. In order to keep the support top geometry narrow, V-string insulator assemblies had to be used and this is only feasible with high strength composite insulators (see Fig. 15).

Conclusions

Composite insulators are probably the most important technical innovation in overhead lines over the past decades. Apart from their increasing application in standard line situations to replace porcelain and glass, they have contributed as well to designing compact lines with more aesthetic support structures. Both have promoted greater community acceptance of new line projects.

In addition, due mainly to their excellent mechanical properties, these insulators have also helped line uprating projects, such as conversion of AC to DC as well as stepping up system voltages.

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This edited past contribution to INMR by expert consultant Dr. Frank Schmuck advised what should be done if discovered.

Surface silicification is a type of deterioration that leads to layers of silica oxide forming on the housing, with thicknesses ranging from 100 to 300 µm. Unlike the virgin polymer surface, elasticity in the cross-section of the silica oxide layer is significantly reduced and shows brittle behavior when subjected to mechanical stress such as bending of sheds. If such deteriorated material is measured with respect to mechanical characteristics, there is reduction in tensile strength – even though tear strength is not necessarily affected. This can be explained by the fact that, when bent, the brittle silica oxide layer will break and the resulting crack leads to a ‘notch effect’ in the polymeric bulk material.

In the past, for example, different severities of damage were identified on silicone housed 420 kV instrument transformers installed in Central Europe and made with an LSR housing. The resulting deterioration was classified as either Type A or B, where in the first case there are only minor cracks when the sheds are bent (see Fig. 1). Type B, by contrast, is characterized by disintegration of the deteriorated material layer (as shown in Fig. 2). The years of manufacture of these insulators was from 1991 to 2007 and there was no direct correlation to time in service for cases of Type A, Type B or no deterioration.

Type B damage was considered the more critical since this mode was found all over the housing, including in areas protected against direct UV exposure. The level of severity was confirmed by scanning electron microscope analysis (see Fig. 3) that showed loss of material and crack formation into the bulk polymer. The LSR surface in this example was later sealed by an RTV coating to prevent pollution or moisture ingress into the cracks, which might risk further damage. This remedial measure proved successful and the affected instrument transformers remained in service.

FTIR analysis suggested hydrolytic processes as the root cause behind such deterioration. Unlike HTV silicone rubber, LSR material is not enriched with aluminum trihydrate (ATH) for improved erosion resistance. Instead, silica is used as filler to provide the mechanical properties required. One hypothesis was that the type of damage observed is possible if this filler has not had appropriate surface treatment.

Deep surface cracking or splitting is another form of deterioration that results in serious cracks through the entire housing. Such damage carries great risk since it exposes the rod, which is not outdoor-resistant. In case of pollution and moisture ingress, the interface between housing and rod is further attacked from the exposed area and a flashunder becomes likely.

This problem was initially investigated using a simple acid storage test with nitric acid having a pH of 0 chosen because this can be generated by corona discharges. Results showed that filler treatment is decisive in determining susceptibility of a material to being attacked by the acid. Without proper filler treatment, for example, one sample was shown to suffer substantially such that it showed deep cracking when bent after the test.

When considering what should be done if such problems are detected in service, there are two key issues:

1. Evaluating deterioration and estimating risk for continued use. IEEE Guidelines for Establishing Diagnostic Procedures for Live-Line Working of Non-Ceramic Insulators can be used as reference. As a ‘rule of thumb’, no level of housing damage should be accepted that allows the core rod to become exposed.

2. Since there is no principle to predict risk of further damage, it is recommended to annually inspect insulators from the supplier and vintage affected that are operating in the service area where this type of deterioration has been observed. In the case of silicification, but where the housing is still hydrophobic, affected insulators can remain in service because the silica oxide layer serves to shield the intact bulk material against further deterioration.

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Lightning has been reported as the main cause of unscheduled outages on overhead sub-transmission and transmission lines (e.g. U.S.: 57%; Brazil: 50-70%; Japan: 70-80%; Denmark: 57%; Colombia: 47-69%). Reducing outages due to lightning has a major impact on overall reliability of both distribution and transmission lines. The main aim of installing transmission line arresters (TLAs) is to reduce tripping/outages on shielded or unshielded lines. Tripping generally arises from insulator flashovers – commonly known as back-flashovers, since the tower is no longer at earth potential but due to the lightning surge at higher voltage than the conductor:

• Unshielded Lines: Lightning strikes to the structures or the phase conductors will, in almost all cases, produce flashovers along the insulator strings.

• Shielded Lines: Strikes to the structure or earthwire have the possibility of backflashover occurring across the insulator strings depending, amongst other parameters, on the level of strike current and the transient grounding system behaviour.

Transmission line lightning performance also depends on correct choice of arrester type and positioning on the structure and along the line. It is possible to significantly reduce lightning caused outages by having TLAs on every phase on every structure but this will rarely be economical and the overall failure rate of the TLAs may then reduce transmission line performance. It is therefore necessary for a power supply utility to specify an acceptable performance level and work from there.

This edited past contribution to INMR by Dr. Brian Wareing, Overhead Lines & Lightning Protection Consultant based in the U.K., provided an example of choice and positioning of TLAs on a double circuit 400 kV line.

Wish Power (Thailand) Co. Ltd

Thailand

Wenzhou Yikun Electric

China

See more suppliers of Arresters

An evaluation was required to increase the reliability of the 400 kV Beauly-Denny line that crosses the North Scottish Highlands from near Inverness to just west of Edinburgh. The major problems affecting reliability are severe weather conditions (snow, ice and winds up to 65 m/s) and lightning triggered outages. Investigations into line design were introduced to reduce the impact of the effect of the severe weather expected, especially in the Corrieyairack area where the line rises to around 800 m. The risk of the line being struck by lightning with consequential back-flashover and tripping out of circuits was then investigated. The back-flashover situation commonly caused by lightning strikes is usually mitigated by a low value of tower earthing. However, in many parts of the line the ground is granite with soil resistivities >20,000Ωm. Obtaining low value tower earths would therefore be expensive and, in places, virtually impossible. The alternative to reducing back-flashovers is the transmission line arrester (TLA) and a full analysis of how this could used to produce a reliable line within financial constraints. This analysis was then used to determine the necessity or otherwise of providing possible costly earthing solutions. So the TLA method was compared with the cost of introducing earthing mitigation measures to produce the most reliable and economic solution. The initial target was to prevent 95% of strikes causing a back-flashover.

Back-flashover Scenario

Back-flashovers

Back-flashover is the situation when the difference between the voltage on the cross-arm (created by strike current to the earthwire travelling to earth down the tower) and that on the phase conductor (both with respect to the same earth) exceeds the BIL of the phase insulator arc gap. This can put a sharp fronted overvoltage waveform onto the phase conductors – a particular problem on substation approach. A backflashover is generally agreed to be the situation where the tower is at high voltage compared with the phase wire instead of approximately at earth level, since it would be in the case of flashover from a direct strike to a phase conductor. The back-flashover-generated surge has a very sharp wavefront as the arc causes the phase wire to jump in <1µs from an induced voltage level (from the surge along the earthwire) to virtually the full lightning surge voltage as present on the tower cross-arm (the arc itself will drop a few hundred volts only).

Back-flashover Probability

The probability of a flashover is dependent on the voltage across the insulator i.e. between the tower crossarm voltage and the voltage on the phase conductor (induced voltage plus normal 50 Hz voltage). To calculate this it is necessary to calculate the surge impedance of all the phase and earthwires, which requires knowledge of local ground resistivity, and determine how much surge voltage is induced by the earthwire onto the phase conductors. This requires a calculation of the coupling factor between the earthwire and the phase conductors and requires knowledge of the ground resistivity. This is also dependent on the tower type (distance of conductors above ground) and the tower footing resistance. The crossarm voltage is also dependent on the strike current, earthwire impedance and the surge impedance of the tower and ground. In dealing with the latter it is always necessary to look at the system when subject to a MHz phenomenon and not a 50 Hz scenario. Hence the process is dealing with surge impedance and not resistance.

Once the surge voltage at the cross-arm is known and compared with the phase wire voltage, the probability of flashover depends on the position on the 50 Hz cycle and the ±15% of the mean impulse (50%) breakdown level of the arc gap. The probability of flashover of the insulator will also depend on the waveform of the lightning surge as it appears at the crossarm level. This is currently the subject of work by SSE at Heriot Watt University in Edinburgh. The frequency of the flashover event can then be determined from the local lightning activity. The Heriot Watt PhD project takes a fresh approach at adding to the understanding of the effect of lightning strikes on an overhead transmission line in terms of its electrical behaviour such as charge and voltage propagation around towers and lines. It is reviewing and developing models to simulate this behaviour. This is to enable the answering of questions such as: where should transmission line arresters (TLAs) be placed, on the top, middle, bottom conductor or all and how efficient they are at reducing back-flashovers.

Evaluating Lightning Risk

General
A simplified method of determining the lightning risk and eventually the back-flashover risk per tower uses the average lightning strike density and strike current as well as a simplified calculation for the tower surge impedance. A full investigation would cover the range of lightning strike currents (not the average) and the surge steepness plus corona losses and calculation of the surge impedance for each section of the tower.

It is important to note that lightning strike parameters vary regionally and it is recommended that local data be used if available to increase the accuracy of the overall evaluation of overvoltages and back-flashover risk on transmission lines. In evaluating the strike risk along the Beauly-Denny route, historical data on strikes were obtained within a 1 km radius of each tower. This radius was chosen since this is the approximate distance that a lightning voltage surge on the line will travel before corona losses, etc cause the strike to dissipate.

Evaluating Strike Rate & Current

The strike rate is obtained from a calculation of the collection width for each line section on the Cigré basis where the horizontal attraction, ld, distance for lightning strikes to an OHL conductor is determined from the positions in space of the various phase conductors and earthwires and is obtained from the expression:

l_d=C •K₀  •I^0.74•h^0.6

where:

l_d  = horizontal attraction distance (m)
C = line factor (dependent upon type of line),
for bare wire C = 0.84
I = lightning current (kA)
h = height above ground of the overhead line conductor (m)
K₀ = topography factor (for flat open terrain, K₀ = 1.0 but can vary from 0.7 for valleys to >2.0 for hill tops)

The above calculation gives the collection area of lightning strikes for each tower on the line as 2 x ld x 2/1000 km² as the tower could, in theory, be reached by surges from 1 km away. Once the collection area is established, it is multiplied by lightning strike density to obtain the lightning risk per annum. The average strike current is then determined from the data. In practice, due to the lower surge impedance of the towers (~200Ω) compared with the conductors (350 to 500Ω) most of the surge current will go down the first tower a surge along the earthwire meets and so this tower will be the most susceptible to back-flashovers. The surge impedance of a tower can be calculated from the size of the individual vertical steelwork sections and the distance they are apart plus an allowance for the closeness of the ground. This is a complex process and a simple, though less accurate, method is given in IEC 60071-2:

Z_t = t.((W.ln(4h/W)/(32.π.AW))^0.5 + 6.5

where:

t is the transit time (s)
W is the tower base width (m)
H is the tower height (m)
AW is the average tower width (m)

Average self and surge impedances of have been calculated for the earth wire of the Beauly-Denny line for the range of tower types and locations. The earth wire surge impedance varies between 498Ω – 548Ω according to the local ground resistivity values. The tower impedance varies from 170Ω for a D55 M3 tower (high angle of deviation and low height) to 244Ω for a DL E15 (in-line suspension extended height tower). The surge impedance of a conductor, Z, is determined from Z = √ (L/C) where L is the inductance and C the capacitance of the conductor. In a simplified format, the earthwire surge impedance, Z_ew, is calculated from:

Z_ew = 60 x ln ((h +S)/r)

where: S = 659 x √(ρ/f)

where:

h – average height of EW above ground (m)
r – Nominal earthwire radius (m)
S – calculated depth of the image of the earthwire in the ground (m)
ρ – ground resistivity (Ωm)
f – frequency (Hz)

The surge impedance of a twin bundled conductor is calculated in a similar fashion but takes into account the lowered inductance of the two phases.

Evaluating Flashover Risk

Incident Surge Impedance

Once the frequency and magnitude of lightning current surges to a particular tower has been established, it is necessary to establish the surge voltage levels. This is obtained by determining the surge impedance of the conductor at the struck point and then the surge travel along the line. Theoretical considerations and practical measurements show that the lightning channel surge impedance is likely to be up to 3000Ω and is thus assumed to be substantially larger than the surge impedance of the struck object (the conductor) that will be ≤600Ω.

At the strike point, for instance on an overhead line earth wire as illustrated in Fig. 1, the injected current is divided equally between the earth wire ends connected to the towers. Therefore the impedance Z seen from the lightning strike is a parallel circuit of earth wires Zew and tower impedances Zt plus the ground impedance Ze as the tower footing is not remote earth – the reference point for the phase conductor voltage. Assuming that the lightning channel impedance is 3000Ω, the earth wire impedance 500Ω, tower impedances 200 Ω and ground impedance 50Ω, the equivalent impedance seen by the lightning strike calculates to 334Ω resulting in the injection of approximately 90% of the strike current at the strike point (Fig. 2).

In round terms, therefore, a 30 kA strike to the earthwire will result in a 14 kA surge going in each direction towards the towers. The surge will travel to a point of impedance change (the pole top) at which there will be a reflected and transmitted component. The pole top junction has several routes available for the current and voltage waves and the surge current will split according to the inverse ratio of the surge impedances of the routes available – say ~200Ω down the tower route and ~500Ω for the continuing earthwire. The 14 kA surge will therefore split approximately 5/7 down the tower i.e. 10 kA. This goes on to the ground and generates a ROEP at the ground impedance.

Coupling Factor Between Earthwire & Phase Conductors

A lightning strike to the earth wire will induce a voltage on the phase conductors. This is beneficial since it reduces voltage stress on the tower insulators. The coupling factor CF can be derived from the following general relationship:

Electrostatic and electromagnetic couplings are calculated using the following simplified formula

where:
CF       Coupling factor.
a =       distance earth wire to phase wire (m)
b =       distance from phase B to image of earth wire in ground (m). Alternatively, for consistency with previous calculations the depth S_e may be used.
h =       height of earth wire above ground (m).
r =        actual radius of earth wire (m).

The electrostatic calculations use an imaging technique for calculation where the image of a conductor is the same distance below the ground surface that the earthwire is above the ground surface. The electromagnetic calculations use conductor above ground height and a ground depth equal to the calculated depth of the image in the ground. The coupling factor of all phases of all tower types may thus be calculated. For a typical standard 400 kV tower with a single earthwire, the CF values would typically be around 0.2-0.3 for the upper phase, 0.11-0.18 for the middle phase and 0.06-0.12 for the lowest phase. These values are affected by ground resistivity and tower type.

Calculation of Cross-arm Voltage

The cross-arm voltage can be determined from the tower surge impedance, dimensions, crossarm positions and the tower footing voltage, which is obtained from the strike current and earthwire and tower impedance values. It is not intended here to go into the determination of the tower footing resistance/impedance as these are likely to be different.

Calculation of Insulator Stress

The stress across the insulator, V_ins, is then determined from knowledge of the phase conductor voltage. The phase conductor voltage relative to remote earth will be the power frequency voltage plus the induced voltage, which comes from the earthwire surge. It is possible for back-flashovers to occur on any of the phases as although the cross-arm voltage will be lower for the lower phases, so will the coupling factor from the earthwire strike.

Effect of Tower Footing Resistance

Looking at a standard 400 kV tower and the various parameters discussed in this paper, the stress across the phase insulators can be calculated. For the lowest phase if the tower footing resistance (TFR) is ~5Ω then with C~0.08, V_ins will vary from 1.09 MV to 1.32 MV. As the 50% probability of failure is 1.4MV, there is a probability of around 10% that this insulator will flashover if the strike occurred at an appropriate point on the power frequency cycle. The utility may accept this rather than use TLAs. However, a substantial part of the Beauly Denny line is on granite with a ground resistivity of over 20,000Ωm and for several towers the TFR cannot be reduced below 100Ω unless several thousand pounds is spent. If R=100Ω, then the lowest phase will have a V_ins of between 1.37 to 1.83 MV and thus an almost 100% probability of flashover. The utility choice is then between using TLAs on the lowest phases or spending a considerable amount of money on reducing earthing levels. It is thus possible to evaluate a backflashover risk per tower. Figs. 4.3 to 4.5 show the backflashover risk per tower against lightning strike density, strike current and ground resistivity.

The lack of correlation with strike density is because other factors such as lightning strike current and ground resistivity can have a significant effect. There is mostly good correlation between high lightning surge currents and flashovers but in one case there appears to be little effect. This is because of the low lightning strike density. Fig. 5 shows a good correlation between high ground resistivity and backflashover rate.

Field Experience In South America

Placement of TLAs

As the Beauly-Denny line is still under construction, performance data is not available. However, data from Cigré concerning the Cambuci – Sto Antonio Padua 69kV and the Antamina 220kV lines in Brazil provides such data. Lightning has been reported as the major cause of non-scheduled outages in Brazilian’s power system, creating many issues and damages for power supply utilities and their consumers. Losses and damages in the Brazilian power supply utilities caused by lightning exceed an annual value of $350 million.

Installing one TLA on an individual tower reduces the probability of flashover (Fig. 6). This shows the effect on individual towers. While the lightning performance is improved, this improvement is restricted since surges will cause problems at neighbouring towers that are not protected. As can be seen, the improvement for one TLA per circuit is only around 30% and in some cases much less. If a second TLA is installed the improvement is only marginal.

Fig. 7 shows the effect of applying TLAs on adjacent towers. The improvement now is commonly over 50% (on the same towers as Fig. 5.1) and the gain from the installation of a second TLA is significant – in some cases reducing the back-flashover rate to less than 10% of the original unprotected level.

The Antamina Mine is located in the Antamina valley in the Andes Mountains in the Ancash region of north-central Peru and is fed by five 220 kV transmission lines located in regions with isoceraunic levels from 15 to 90 thunderstorm days per year. In the period from 2002 to 2006, 80 non-scheduled outages occurred due to lightning which affected the production processes on these 220 kV lines and other outages due to lightning also occurred on an associated 23 kV overhead shielded distribution ring network. From January 2006 till June 2007, approximately 450 units of class 2 line arresters were installed along the distribution network and 265 gapless transmission line arresters rated at 192 kV were installed along the sections of the two 220 kV transmission lines with the worst lightning performance. From October 2006 to February 2008, five outages due to lightning were recorded on sections of the lines which were not protected but no outages on the protected lines (Fig. 8).

Brazil’s transmission system has more than 3000 units of gapless line arresters installed along overhead lines from 34.5 kV up to 230 kV. Analysis and evaluation of lightning performance before and after the installation of line arresters have shown high effectiveness, with average indexes for the improvement greater than 70%.
Lightning Performance with TLAs
Lightning is the most frequent cause of transmission outage and service interruption in the United States, accounting for about 30% of all power outages, and resulting in economic losses approaching $1 billion annually. Reduction in earthing levels can achieve significant reduction in back-flashover rates but this can be very difficult and expensive in areas of high ground resistivity. In general:

• Typically not needed in all phases to get a significant reduction of backflashovers

• Areas with high tower footing resistance are first targets.

• The use of TLAs on adjacent rather than individual towers generally produces better results.

The best performance in reducing flashovers is to put TLAs on both circuits and all phases but this can be uneconomic. Two TLAs in one circuit and one TLA in the most exposed phase of the second circuit may give an even lower outage rate for the two circuits together and still get very close to zero risk for a double circuit tripping. Calculations are required to determine which are the most exposed phases.

Costs: Earthing Versus TLAs

It is possible to calculate the minimum tower earthing level to reduce the back-flashover rate to a level acceptable to the utility. The cost of reducing the tower footing resistance to this target value can then be determined. Other factors to bear in mind are the possibility of copper theft, thereby increasing the resistance and back-flashover rate, and the failure rate of the arresters, thereby reducing the overall reliability of the line. Normally, this failure rate is well known in terms of substation arresters, but with TLAs there is the additional problem of conductor movement to deal with, whether this is due to vibration, galloping or simply excessive movement under high winds. Without these extra factors, it is a relatively easy calculation to determine whether earthing improvement or arrester use is more economic. However, lack of field-testing and necessity or otherwise of increased damping means that the mechanical reliability of TLAs cannot yet be established. If the TLA presence increases vibration levels, for example, then a shortened conductor life may result. Another aspect is the disconnection time related to the circuit breaker operation time. There is therefore a race between the line protection scheme and the TLA disconnector operation and this may vary from incident to incident as well as in different weather conditions. However, this is an electrical/mechanical balance situation that is not within the scope of this article.

Conclusions

There is a wealth of experience of TLA use in many countries, particularly in South-East Asia and South America. There is no doubt that significant improvements in the lightning performance of OHLs can be achieved with TLA use and that in areas of high lightning activity and/or high ground resistivities it can prove to be the most economic option. However, the effect of a significant weight on OHL conductor is not fully known, especially with regard to vibration and galloping scenarios. The disconnector time is also another area that needs to be considered in relation to the protection system used. It is possible to calculate the appropriate phases and towers that would benefit from TLA use and thereby focus TLA use economically to achieve the required performance level. However, the mechanical considerations of vibration and galloping and TLA effects on the OHL conductor are yet to be evaluated and testing in the field is recommended.

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This edited past contribution to INMR by experts at TE Connectivity dealt with applications where wildlife impact with electrical networks.

EB Rebosio SRL, A Gruppo Bonomi Company

Italy

Hivolt Power System

China

See more suppliers of Wildlife Protection Products

Every power provider has suffered periodic disruptions from unplanned outages of one type or other. These have traditionally been classified as Weather, Wildlife or Unknown, with limited knowledge of the proportions due to each. Now, with greater knowledge, such unplanned outages are better categorized as due to:

1. Wildlife (birds & animals)
2. Weather (storms, wind, lightning)
3. Vegetation (tree branches)
4. Human intervention (accidental or deliberate)
5. Unknown or not yet categorized

Problems created by wildlife at substations and on overhead lines typically fall into two main categories: bridging and wildlife guano pollution flashover. While the result, i.e. a system trip and possible arc flashover, may be the same, the way this occurs, and the best solution can differ case-by-case. So do the remedial products that can be applied to compromised equipment.

Bridging is where a bird or animal makes contact between phases or between phase and ground, leading to a short circuit. For example, large birds can easily cause problems across all distribution voltages at substations and on lines. Phase bridging usually results in the creature being electrocuted and how it falls will determine whether the auto-reclosing system operates successfully or not. Often the bird or animal will fall away and allow the auto-recloser to re-energize the circuit. Although this may not normally require investigation by emergency crews, it usually leaves unnoticed but tell-tale evidence of damage.

Although the circuit might operate normally, a series of such events can lead to progressive damage that becomes worse with each successive trip and that finally causes hardware failure, e.g. where each trip burns out a single strand of conductor eventually dropping the line. Moreover, if the dead creature falls between its points of contact and remains there, this will result in the circuit not being able to be re-energized until an emergency crew removes the fault source. At least, this type of event can be logged with some certainty of what happened. By contrast, when the bird or animal falls to the ground or on top of a transformer, it is likely to be removed by a predator and no evidence will remain.

Most bridging problems occur on MV systems (typically <36 kV) designed such that bare conductors, bushings, busbar systems and other equipment have ‘air-spaced’ clearances that operate without issue under normal conditions. Clearances are typically up to 40 cm at substations and about 1m for overhead lines to allow for swinging conductors. Unfortunately, these clearances mean that a wide range of birds and animals can cause problems at substations and on lines. It should be noted that, although line conductors are generally 1 m apart, the most common failure mode is between the cross-arm (usually grounded) and the conductor – a typical clearance of only 30 to 40 cm.

Medium and large birds perch on a cross-arm and bridging usually occurs during landing or take-off. In more unusual cases, depending on location, wildlife such as bears, possums or even snakes climb poles and cause similar problems.

Pollution flashover from bird guano is a less frequent occurrence but can be every bit as damaging as bridging-induced outages. It is mostly associated with line insulators but can also occur at substations and differs from bridging since it can happen across all voltages – even up to EHV. Such flashovers of bushings or insulators result from build-up of guano for weeks or months during dry conditions when there is no rain to wash it away. Birds often favour certain places to perch, and this is often the extremity of a lattice tower arm, directly above the insulator string. While the build-up of guano can be considerable, so long as conditions are hot and dry the relatively small individual amounts of liquid guano dry quickly and surface resistance is not greatly altered.

Flashovers take place when mist, fog or rain returns, changing the dried guano into a semi-liquid state and greatly reducing surface resistance of the porcelain or glass insulator. If the build-up is sufficient to alter the conductivity of the insulator string, flashover can result. It should also be noted that another bird related failure mechanism – guano streamers –can cause flashover when a bird perched above an energized conductor emits a liquid streamer up to 2 m that effectively shorts out the otherwise safe air gap.

While all countries see some level of wildlife induced disruptions, the diverse nature of the birds or animals causing the problem as well as the different types of equipment and service conditions mean that the optimal approach to prevention also differs on a case-by-case basis. Once a failure mechanism is recognized, the latest generation of materials and designs make it largely preventable, often for the entire remaining life of the equipment.

Diverse Problems Across the Globe

The range of wildlife problems experienced by substation and overhead line engineers is too vast to cover in one article so common examples discussed below give a flavour of the diverse issues that can be encountered. Wildlife problems affecting power networks are categorized as ‘migratory related’, ‘local specific’, ‘food-related’ (predator or prey), ‘shelter/security/nesting related’, ‘seasonal’ and ‘other’.

Migratory issues relate only to birds passing through one area en route to some destination to feed and rest. These short time periods can cause local havoc by overcrowding mast tops where there is limited perch space available. In such cases, there is high risk of bridging phase-to-ground or phase-to-phase clearances. Other incidents include birds flying into lines near feeding sites, usually at distribution voltages where conductor clearances are such that large birds colliding with them can bring two conductors together, resulting in flashover. Once birds have completed their migration, they tend to return to the same location year after year. This means that, even if old nests are removed once the birds have left, they will be rebuilt the following year on that precise pole or tower.

The exact nature of any failure will depend on the design of affected equipment, type of bird involved, system voltage etc. but can be as simple as bridging while the bird is perching on a MV cross-arm. Guano induced failures most often impact HV suspension towers. Similar problems can arise from materials falling out of nests or when groups of birds fly in and around jumpers where there is a switch or pole top transformer.

Absolute data regarding the number of birds killed worldwide from interaction with power networks is usually under-reported since many countries do not record such statistics or fear their publication due to possible regulatory consequences. However, a figure measured in tens of millions of birds killed each year seems beyond question.

Seen from the viewpoint of electricity suppliers, this massive loss of wildlife equates to annual losses of over US$10 bln from lost revenues, damaged equipment, emergency call-outs by line crews and fines imposed by regulatory authorities. Some countries use legislation to protect endangered wildlife with severe penalties of up to US$ 500K.

Problems at substations are almost exclusively of a bridging nature when birds or animals bridge phases or phase-to-ground. Issues are generally not migratory related but can be seasonal, e.g. during cold weather when cats and similar size creatures are attracted to warmth close to a transformer. This seasonal problem can also be due to small birds building nests that attract a range of predators such as cats, snakes, larger birds or other species.

At MV substations, the ‘rule of thumb’ is that anything up to a maximum of 40 cm between phases or phase-to- ground is considered vulnerable to bridging by wildlife. This critical distance can be a greater (even up to 1 m) if a substation is plagued by wildlife such as snakes, monkeys and large birds of prey.

Solutions to Prevent Wildlife Outages

Given the huge financial costs, loss of wildlife and inconvenience to the public, it is a surprising that almost all the problems due to direct bird or animal interaction with power networks were preventable. The best solution to such problems is generally one that allows whatever wildlife is involved to safely access their chosen sites without causing electrical breakdown of any part of the system. It is equally important that any solution be cost effective such that the costs of installation are considerably less than the cost of the possible outages to be avoided.

It is still not possible to put a cost to avoiding every potential wildlife problem because of the diversity of issues and creatures. Still, it is estimated that it can vary from as little as US$1K to more than US$ 1mln. An average of circa US$10K is used in the U.S., where most data originates.

The importance of the protective material used cannot be overestimated. This is because any solution to a wildlife problem must not cause issues downstream because of material failure. On ‘day one’ almost any solution will perform to some degree; but the most effective solution will continue to work maintenance-free for the life of the equipment on which it is installed.

This will often mean at least 20 years or longer in challenging environments such as extremes of hot and cold, severe weather, long-term exposure to UV as well as a variety of pollutants, especially uric acid (the main chemical in bird guano). With these environments in mind, materials need to exhibit a minimum of technical criteria according to the relevant international standards, as shown below:

In addition to optimized geometry, materials used must deliver consistent performance, without electrical or mechanical breakdown, for over 30 years. As such, these materials must be robust mechanically, maintain excellent electrical properties while remaining tolerant to wildlife exposure. The images below illustrate what can happen in a short time when minimum levels of such performance are not met.

A variety of alternatives have been tried over the past years, including sonic deterrents, barriers, covers, electrified fences, smooth climbing barriers, chemical barriers and spined perching devices designed to drive the bird or animal away from a specific location. However, in most cases, the problems only re-surface elsewhere. There is also the consideration that migratory birds and even some resident animals always return to the same nest or feeding sites. This reinforces the importance of finding a solution to protect the reliability of electricity supply while also allowing wildlife unhindered access.

Many of the wildlife protective solutions adopted over the years have brought some limited, short-term or localized benefits. But experience has shown that the most successful overall solution is based on a range of electrically high performing polymeric materials, custom-designed for specific applications. Critical material performance factors include excellent long-term resistance to premature ageing, whereby the material is expected to maintain its electrical and physical properties, thermal endurance (continuous operating at up to 105°C) and high UV stability over some 30 years of service.

In parallel with outstanding material performance is good design. The key to successfully preventing bridging or guano flashovers lies in understanding the problem. This means identifying the most vulnerable bare metal places that need insulating. The solution will vary depending on the precise nature of any problem (e.g. a cat bridging phase-to-ground on a MV breaker or a white stork perching atop a cross-arm and making contact with one or more phases). As such, the local substation or line engineer is key to collecting this information along with looking for tell-tale signs of past wildlife interactions, as discussed earlier.

Vulnerable bare metal components can easily be insulated with either heat shrink or cold applied materials (or some combination of both) and typically up to 30 cm further along the busbar or overhead line than the farthest location of the known fault. This, however, is only general guidance and local knowledge should always be taken into account when available. In situations where there is bridging between phase and ground it is most common to protect both the equipment and wildlife by insulating the live side, e.g. busbar or conductor. It is also perfectly acceptable to insulate the ground side instead.

Long-Term Solutions

The following are examples of insulating vulnerable bare lines and substation equipment metalwork that had previously suffered multiple flashovers or were deemed at high risk of such events. In all cases, applying wildlife protective insulating materials effectively eliminated the problem while also allowing the wildlife to continue access.

Case 1: A white stork has a +2 m long wingspan. Thousands of these birds migrate each year from Africa to nest in Spain as well as central and south-eastern Europe. The nest is so big (up to 400 kg in weight and 2m in diameter) its presence will eventually cause flashover from accidental bridging between nest debris and conductor when conditions are wet or when birds bring conductive material (e.g. wire) to the nest Moreover, young birds that stray around the nest risk contacting the conductor. Insulating the bare metal fittings with a wraparound cover and the conductor with wraparound sleeving approx. 1.5 m from the nest in both directions offers both the bird and the power system long-term security.

Case 2: This crow’s nest in South Africa is precariously positioned on a 66 kV line between insulator and pole. The bird’s wingspan is sufficient to cause accidental bridging during landing or take off. Without the combination of conductor sleeve and wrap around connection cover, this bird/nest combination would at some point lead to flashover. The nest itself, while dry conditions last, might not cause problem. However once conditions become wet (rain, mist, fog, etc.) it is likely to short out enough of the insulator to cause flashover. Even in dry conditions, once the young birds are grown enough to move about, it is inevitable they will trigger a flashover.

Case 3: Laws in Germany require power utilities to implement countermeasures to prevent electrocution of birds and to protect both migratory and domestic species. The solution has been to use protective devices on 10/20 kV distribution mast tops. Insulation covers with an overall length of approximately 1.4 m prevent contact between birds and live conductors. There are many designs that can be attached using cable ties or live with insulated poles (hot-sticks). It is also possible to use transparent materials that allow aerial line inspection in terrain difficult to access. This solution may become more widely used as utilities increasingly introduce drones for maintenance.

Case 4: This insulator string demonstrates how the entire length can become heavily polluted with bird guano. The other photo shows that shielding the string makes it possible to keep the surfaces clean enough to never suffer pollution flashover due to wildlife. Here, it is critical that the material and design of the shield give many years of maintenance free service. This requires it be made of a high-performance UV stable polymer and robustly secured at the top of the string. After seasonal rains, accumulated guano will wash away and fall harmlessly away from the string below. In those countries with extreme dry conditions and where it might rain heavily only once in several years, it is even more important to use a material that is also resistant to attack by the uric acid in the guano.

Case 5: A substation in Croatia had all its bare live metalwork insulated, and heat shrink tubing on busbars. Connection points have cold applied wrap-around covers that can be removed for maintenance or inspection and then re-applied. The problem behind this installation was the phase-to-ground clearance on the MV side of the transformer and bus network where the typical 40 cm air gaps were being bridged by cats, crows & pigeons.

Case 6: An ageing MV substation in central Athens suffered from entry by cats and pigeons. With numerous phase-to-ground clearances well below the nominal 40 cm, it therefore experienced multiple flashovers. The solution was to cover all vulnerable bare metalwork with a range of high-performance polymeric covers, wrap-around sleeves and circular barriers that prevent cats climbing over the insulator and simultaneously touching the MV and ground sides.

Case 7:A MV substation in South Africa was built in an area of cleared scrub and trees inhabited by a troop of large monkeys. In its first year of operation it suffered 12 flashovers caused by monkeys that considered this their territory. The solution was to insulate all live bare metalwork across the site with a combination of heat shrink tubes and wraparound covers. Fully understanding the situation was critical to success on this site since phase clearances in some parts were approximately 1m which, if not for the monkeys, would not require additional insulation.

Case 8: A 20 kV substation in rural Germany was insulated against climbing animals such as martens and cats. Wraparound covers were cut on site to accommodate any unusual geometry, including the centre cover where a cut-out was added to allow access for an earth clamp in a relatively safe position. Covers that need such cutting in the field benefit from being made of a polymeric cross-linked material that will not propagate splits if cut roughly.

ABB and PEHLA Laboratories

Germany

GCC Electrical Testing Laboratory

Kingdom of Saudi Arabia

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Conclusions

Awareness of the cost of wildlife-induced outages is growing and, with this, has come an increasing demand for reliable and robust solutions to protect both utility assets and the wildlife causing problems. Every power utility therefore has to contend with wildlife in the design and protection of its network. In finding the best long-term solutions, engineers must balance the need for exclusion/prevention and some tolerance of the wildlife in question.

The main causes of wildlife problems vary by region but the implications and basic design principles deployed are often the same. In addition to reliability enhancement and asset protection, an unacceptably high number of birds and other creatures are killed yearly due to electrocution. Legislation to protect many species is increasing, with countries such as the U.S. and Germany leading the way. Wider EU legislation regarding protection of birds is now being planned.

As with any network component, the material chosen as a solution must not break down and cause a reliability problem of its own during the life of the system. It is critical to specify high performing materials that are similar to those used on other overhead line components, including insulators and surge arresters. Most wildlife induced outages are now largely preventable.

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Pingxiang Huaci Insulators Group Co. Ltd.

China

EB Rebosio SRL, A Gruppo Bonomi Company

Italy

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Fig. 1, which derives from laboratory testing of ceramic insulators plots unified specific creepage distance (USCD) required as a function of pollution severity, as measured by salt deposit density (SDD). This relationship has since been confirmed by field experience.

HVDC lines equipped with ceramic insulators designed according to the red curve will offer good service performance while insulators designed with a lower USCD are likely to encounter problems of flashover. Also illustrated is the fact that, depending on pollution, DC lines may require insulation having a much larger USCD than for AC. This can result in unrealistic design parameters – especially when there is the combination of UHV and heavy contamination.

In the extreme case of very high pollution (e.g. SDD of 1 mg/cm2), a total creepage distance of more than 90 m would be required for 800 kV insulator strings. This would mean an insulator length of some 27 m for insulators with a creepage factor (CF) of 3.4, as typical limit for DC applications.

DC presents a far more severe design scenario for composite insulators as well. However such insulators benefit from the hydrophobic transfer material (HTM) property of their polymeric weathersheds. In this regard, Fig. 2 depicts results of a past analysis within CIGRE WG C4.03.03. Basically, this showed that HTM insulators require a lower USCD at the same SDD than do insulators made from non-HTM materials.

This is demonstrated not only by laboratory investigation (see dotted curve in Fig. 2) but also by service experience (see points in Fig. 2 that refer to real HVDC lines).

Based on Fig. 2, in the extreme case of very heavy contamination (e.g. SDD of 1 mg/cm²), 52 m of creepage would be sufficient for an 800 kV composite insulator set. The corresponding insulator length could therefore be “only about 14 m”, mainly because composite insulators work more efficiently at the higher CF (assumed to be 3.8 in this case against a 3.4 value for cap & pin ceramic).

To be fair, this is but an example that makes reference to extreme pollution conditions that are rare. The goal was only to qualitatively emphasize the criticality of insulation design parameters and to demonstrate the comparative advantage offered by composite types.

Due to feasibility aspects and the cost impact of pollution in DC, insulation design should be highly detailed. This will require accurately assessing site contamination severity by means of measurements on energized insulators while also carrying out laboratory tests to assess performance of the insulators selected. Such performance will depend on profile and characteristics (i.e. the curves shown in Figs. 1 and 2 only represent average performance). A statistical design approach can then be applied by assigning an acceptable risk level for flashover.

Two of the most important aspects of insulators for DC applications are housing material and profile. As far as the first is concerned, a tracking and erosion test was standardized for AC (IEC 60587) but there has been debate whether a similar test is necessary for DC (or if the same AC material ranking can be applied).

Several tracking and erosion investigations have been conducted in different parts of the world applying a DC stress equal to that of ACrms. These showed that DC is indeed more severe in this respect than AC. However, such tests are inherently biased, and their conclusions should not be considered valid from a practical perspective. This view is confirmed by the satisfactory service experience of insulators presently installed on HVDC lines as well as by laboratory experience obtained by long-term ageing tests on different makes of insulators that were properly scaled according to the AC/DC stress ratio.

As can be derived from Fig. 1 (which can be extended to composite insulators by taking into consideration their HTM benefit), insulators will not encounter the same service stress in AC as in DC. Depending on pollution severity, the DC stress can be as little as half that of AC.

Therefore, if a tracking test is to be standardized and made realistic for DC, it should take this into account. Moreover, optimization of insulator geometry in DC will have to consider that creepage distance loses its efficiency in the case of too narrow a profile and too high a CF factor, with more stringent limits than for AC.

PSW – Siemens Energy Testing Laboratories Berlin

Germany

Powertech Labs Inc.

Canada

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