HIGH electric field strength ( ) may cause corona on nonceramic

1070 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 22, NO. 2, APRIL 2007 Practical Cases of Electric Field Distribution Along Dry and Clean Nonceramic Insulators of High-Voltage Power Lines Weiguo Que, Stephen A. Sebo, Life Fellow, IEEE, and Robert J. Hill, Member, IEEE Abstract High electric field strength may cause problems, such as corona and deterioration, on nonceramic insulators. Therefore, control of the electric field strength ( ) in the vicinity of such insulators is an important aspect for their design. This paper discusses insulator computation models and their applications to 34.5 kv and 765 kv nonceramic insulators. Possible model simplifications are reviewed and discussed. Various 765 kv insulator designs and tower types are covered; the magnitude and distribution of are reviewed. Verification tests are also described. Index Terms Electric field and voltage distributions, insulator model setup, nonceramic insulators, verification tests. I. INTRODUCTION HIGH electric field strength ( ) may cause corona on nonceramic insulators, which may result in corona cutting, deterioration, and aging of their polymer materials. Therefore, control of the electric field strength around nonceramic insulators is an important aspect for their design and the design of their associated grading devices. Several studies recommend the value of 2.28 kv(rms)/mm (or 3.22 kv /mm) as the dry electric field strength at any point on the surface of a nonceramic insulator in order to prevent dry corona on the insulator [1]. Significant discharge activity on the surface material of nonceramic insulators may develop in the case of high under both dry and wet conditions. That may cause material aging and erosion. Field experience has shown that erosion of the surface material of nonceramic insulators typically starts near the end fittings, primarily near the line-end fittings. In some very severe cases, erosion can progress through the sheath material and reach the fiberglass rod, where electrical tracking may occur. Therefore, one of the most important details of nonceramic insulator design is the design of the triple junction area (i.e., the junction of housing, air, and metal-end fitting). The electric Manuscript received July 18, 2005; revised January 9, 2006. Paper no. TPWRD-00416-2005. Weiguo Que was with The Ohio State University, Columbus, OH 43210 USA. He is now with Axcelis Technologies, Beverly, MA 01915 USA (e-mail: weiguo.que@axcelis.com). S. A. Sebo is with The Ohio State University, Columbus, OH 43210 USA (e-mail: sebo.1@osu.edu). R. J. Hill is with MacLean Power Systems, Franklin Park, IL 60131 USA (e-mail: RJH@MacLeanPower.com). Digital Object Identifier 10.1109/TPWRD.2007.893190 field strength near this junction area should be limited to a value that is less than the corona onset electric field strength. When nonceramic insulators are installed on a three-phase power line, the tower configuration, the conductors, the hardware, and the presence of the other two phases of the three-phase system can influence the electric field and voltage distribution (EFVD) around insulators. Consequently, a three-dimensional model must be set up in order to evaluate the EFVD near and along the nonceramic insulator. Depending on the voltage level, the magnitude of on the surface of the insulator may exceed the corona onset values. Corona rings can be used to modify the electric field distribution and reduce the maximum value of. The presence of water droplets intensifies on the surface of nonceramic insulators and may cause water droplet-triggered corona at the insulator surface. The typical electric field strength threshold value for water droplet corona is 0.45 kv(rms)/mm (or 0.64 kv /mm) [2]. Water droplet corona is a concern primarily on the sheath of the insulator rather than on the sheds. A standard or conventional insulator design is shown in Fig. 1(a). To prevent the sheath damage adjacent to the line-end fitting, special sheds have been designed to stack together. A modified or stacked shed design is shown in Fig. 1(b). Both designs are available with sheds of equal diameters or with alternating diameters, as shown in Fig. 5(a) and (b). The stacked shed design is intended to eliminate exposed sheath material in the high-stress area adjacent to the line-end fitting, minimizing the potential for material aging or degradation due to water droplet corona. This design significantly changes the insulator profile in the high electric field region, resulting in much greater surface area, and lower current density (compared to a standard shed design) when leakage currents flow as a result of contamination and wetting. II. IMPORTANCE Utility companies are utilizing nonceramic insulators for power lines through 765 kv [3]. In this paper, simplified calculation models for nonceramic insulators are introduced first. The voltage distributions along 765 kv nonceramic insulators with two different shed designs are analyzed, assuming the suspension tower and dead-end tower environment. The electric field strength distributions along the insulation distance and along the leakage distance path are described. To verify the calculation results, a series of experiments was conducted in the High Voltage Laboratory of The Ohio State University by using a Positron insulator tester. 0885-8977/$25.00 2007 IEEE

QUE et al.: ELECTRIC-FIELD DISTRIBUTION ALONG DRY AND CLEAN NONCERAMIC INSULATORS 1071 Fig. 1. (a) Nonstacked standard shed design. (b) Stacked shed design. For the studies described in this paper, the commercially available program, Coulomb, based on the boundary element method developed by Integrated Engineering Software, has been employed. Fig. 2. Simplified dimensions of a typical 34.5 kv nonceramic insulator used in the calculations. III. MODEL SETUP A. Simplification of the Nonceramic Insulator Model A typical 765 kv nonceramic insulator has more than 100 weather sheds, and its length is approximately 4.7 m. To obtain accurate results, a large number of elements have to be used for the electric field analysis of a 765 kv nonceramic insulator. When using the boundary element method to calculate the EFVD along nonceramic insulators, the more elements are used, the more computation time is needed. In order to reduce the calculation time when analyzing long insulators, some simplifications of the nonceramic insulator model are necessary. A nonceramic insulator has four main components. They are the fiberglass rod, polymer sheath on the rod, polymer weather sheds, and two metal-end fittings. In order to analyze and decide which component of the insulator can be simplified with the least influence on the accuracy of the calculation results on the EFVD along the nonceramic insulators, a 34.5 kv nonceramic power-line insulator is employed as a test case. It has 12 weather sheds, and its length is about 0.8 m. The detailed geometric dimensions of the 34.5 kv insulator are shown in Fig. 2. Assume that the insulator, equipped with metal fittings at both line and ground ends, is made of silicone rubber with a relative permittivity of 4.3 and a fiber-reinforced polymer (FRP) rod with a relative permittivity of 7.2. These data were obtained from Manufacturer A. The insulator is surrounded by air with a relative permittivity of 1.0. The top metal end fitting is taken as the ground electrode. The bottom electrode is connected to a steady voltage of 1 kv for the purpose of calculations. The insulator is positioned vertically, but it is shown horizontally in Fig. 2 for convenience. Three simplified computation models are used for the step-by-step comparison process. In addition, a full insulator model is set up as reference to study the effects of three simplified computation models of the nonceramic insulator on the EFVD along the insulator. Components considered in these three calculation models are (a) two electrodes and the fiberglass rod, (b) two electrodes, rod, and sheath on the rod without weather sheds, (c) two electrodes, rod, sheath, two weather sheds at the each end of the insulator, and (d) the full 34.5 kv insulator. The equipotential contours along these four computation models are shown in Fig. 3. The units of the and axes of Fig. 3 are centimeters. Case (a) shows that about 22% of the insulation distance sustains about 70% of the applied voltage. The distribution of the equipotential contours for Case (b), with the sheath on the rod, is Fig. 3. Equipotential contours around the four computation models of a 34.5 kv insulator. (a) Electrode and rod. (b) Electrode, rod, and sheath. (c) Electrode, rod, sheath, and two sheds. (d) Full insulator. very close to Case (d), the full insulator model. The presence of the weather sheds changes the equipotential contours somewhat. If more accurate results of the voltage distribution are needed near the line- and ground-end areas, the simplified insulator model with two weather sheds at each end of the insulator Case (c) can be used. Comparing Cases (c) and d), the voltage distributions in the vicinity of the two weather sheds at the line end are very similar. Moreover, the positions of the equipotential lines for Cases (c) and (d) are very close to each other along the surface of the sheath of the insulator. This indicates that the simplification by Case (c) is acceptable for the calculation of the voltage distribution of the full insulator and Case (d), along the sheath surface. The magnitudes of for Cases (c) and (d) along the paths defined close to the surface of the sheath (similar to the path shown in Fig. 5) are also calculated for comparison, which is shown in Fig. 4; the differences between the two plots are very small. The dips (notches) in the electric field strength plot of the insulator modeled with weather sheds are due to the calculation path passing through the weather shed material, which has a relative permittivity of 4.3. in the vicinity of the two weather sheds at each end of the insulator is the same for Cases (c) and (d). There is only a slight change in the electric field strength distribution near the other eight weather sheds. However, outside the weather shed region still has a good correspondence to Cases

1072 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 22, NO. 2, APRIL 2007 Fig. 4. Electric field strength magnitude along the insulation distance at the sheath surface for the full insulator Case (d), and the simplified insulator model Case (c). Fig. 5. Two insulator designs used for the EFVD calculation. The dashed lines show the calculation path. (a) Design I (standard) insulator computation model. (b) Design II (stacked) insulator computation model. (c) and (d). The maximum electric field strength for Case (d) is 0.0256 kv /mm, and for Case (c) is also 0.0256 kv /mm (peak values). This means that the electric field distribution of the insulator with the full number of weather sheds can be estimated through the simplified insulator model with a small number of weather sheds at each insulator end. B. Computation Models of Two 765 kv Nonceramic Insulator Designs It is of practical interest to know the electric field distribution for a full-scale insulator under field conditions. In this study, two typical nonceramic insulator designs are used for 765 kv suspension towers with four-subconductor bundles. Assume that for this example, the insulator is made of silicone rubber with a relative permittivity of 4.0 and an FRP rod with a relative permittivity of 5.5. These data were obtained from Manufacturer B. There are 51 large weather sheds and 52 small weather sheds on the 765 kv insulator. Two insulator designs are considered here. Design I insulator has all standard sheds on the sheath. Design II insulator has 12 stacked sheds at the line end, three stacked sheds at the ground end, and standard sheds for the rest of the insulator. Fig. 5 shows the two insulator designs (see Fig. 1 for details). In order to prevent corona on the end fittings and to reduce in the high stress regions of the nonceramic insulator, a 17-in corona ring is applied on the line end of the insulator, and a 12-in corona ring is applied at the ground end. The dimensions and positions of the two corona rings are shown in Fig. 6. A corona ring keeper is used to attach the corona ring to the end electrode. For simplification, it is modeled as a small corona ring with an internal diameter of 2.4 cm and an outer diameter of 3.8 cm. The first tower is a typical 765 kv suspension tower with foursubconductor bundles. The simplified geometry and major dimensions are shown in Fig. 7(a). The second tower is a dead-end tower for a major conductor direction change of the power line. The simplified geometry and major dimensions are shown in Fig. 7(b). The conductors have been modeled as smooth conductors, positioned parallel to the ground. The subconductor diameter Fig. 6. Dimensions and positions of the (a) line-end corona ring and (b) ground-end corona ring. Fig. 7. Geometry and dimensions of a 765 kv (a) suspension tower and (b) dead-end tower. for the four-subconductor bundles is 2.96 cm. The distance between adjacent subconductors is 45.7 cm. The length of each conductor considered is 60 m. The two ground wires have been ignored in the calculations. The ground plane is modeled as a 50 50 m large plane with zero potential.

QUE et al.: ELECTRIC-FIELD DISTRIBUTION ALONG DRY AND CLEAN NONCERAMIC INSULATORS 1073 IV. ELECTRIC FIELD AND VOLTAGE DISTRIBUTIONS ALONG 765 KV NONCERAMIC INSULATORS TABLE I THREE CASES FOR EFVD CALCULATION ALONG INSULATORS The electric field and voltage distribution along 765 kv nonceramic insulators have been investigated, assuming four subconductor bundles. The center phase bundle is inside the tower window. The instantaneous voltages applied to the three-phase conductor system for the worst case are kv (i.e., peak value of the line-to-ground voltage, at rated voltage); kv; kv. The basic rules for showing the calculation results are as follows. In the following sections, the voltages are expressed either in kilovolts or in percent values, referring to 624.62 kv, which is the applied voltage of the center-phase insulator for the calculations. is always expressed in kv /mm. The insulation distance used in the figures is expressed in centimeter units or in percent values, referring to the length of the insulator. The calculation path is close to the surface of the insulator sheath along a straight dashed line (Fig. 5). Three cases have been modeled and calculated for this paper with various combinations of insulators and towers. Table I describes the insulator and the tower configuration for these three cases. A. Case I Calculation Results The resulting percent equipotential contours inside the 765 kv suspension tower window for a Design I standard nonceramic insulator are shown in Fig. 8. It can be seen that the line-end equipotential contours are greatly influenced by the line-end hardware and the line-end corona ring. They are nearly parallel to the shed surface. The 15 weather sheds near the line end sustain about 48% of the applied voltage. The 15 weather sheds near the ground end sustain about 13% of the applied voltage. The electric field strength magnitude along the path defined close to the surface of the insulator sheath is shown in Fig. 9. The maximum value of near the triple junction point is 1.48 kv /mm. For a detailed view, the electric field strength magnitudes along the insulation distance near the line-end fittings are also shown in Fig. 10. As in Fig. 4, the dips in the electric field strength plot are due to the calculation path passing through the weather shed material. The other sections of the plot represent the along the sheath regions. It can be seen that is much higher at the junction region of the sheath and shed than that at the middle part of the sheath region. All values are below the dry corona threshold value, 3.22 kv /mm. However, water droplet-induced corona can occur on the insulator sheath at values of approximately 0.64 kv /mm [2]. As seen in Fig. 10, values on the sheath between the end fittings and the first shed, and between the first several sheds are above this threshold value for water droplet corona. Fig. 8. Percent equipotential contours for a 765 kv suspension tower with Type I standard insulator under three-phase energization (Case I). (a) Full view of one of the Type I insulators. (b) Enlarged area around the line end. B. Case II Calculation Results Case II considers a 765 kv suspension tower with Design II stacked shed insulators. The resulting percent equipotential contours inside the 765 kv suspension tower window for a nonceramic insulator are shown in Fig. 11. The magnitude of along the path defined close to the surface of the insulator sheath (Case II) is shown in Fig. 12. The maximum value of near the triple junction point is

1074 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 22, NO. 2, APRIL 2007 Fig. 9. Electric field strength magnitude along the percent insulation distance at the surface of the insulator sheath (Case I), standard insulator design. Fig. 10. Electric field strength magnitude along the percent insulation distance at the surface of the insulator sheath (Case I) near the line end, standard insulator design. 1.51 kv /mm. For a more detailed view, the calculated magnitudes along the insulation distance near the line end are also shown in Fig. 13. All values are below the dry corona threshold value, 3.22 kv /mm. As before, the dips in the electric field strength plot are related to the sheds, the other sections of the plot represent the regions between the sheds. There are 12 stacked sheds at the line end, for about 9% of the insulation distance. The sheath is not exposed where the sheds are stacked. The first region where the sheath is exposed is between 9 and 10% of the insulation distance (encircled in Fig. 13). In general, the stacked shed design is intended to eliminate the exposed sheath in the areas of high stress, minimizing the possibility of water droplet corona and its potential aging effects. As can be seen in Fig. 13, the electric field strength in this case is below the 0.45 kv /mm (0.64 kv /mm) threshold value in the first (and subsequent) areas of exposed sheath. C. Case III Calculation Results Case III considers a 765 kv dead-end tower with two Design I standard insulators. The resulting percent equipotential contours along the nonceramic insulator with four subconductor bundles are shown in Fig. 14. It can be seen that the ten weather sheds near the line end sustain about 32% of the applied voltage. The line-end equipo- Fig. 11. Percent equipotential contours for a 765 kv suspension tower with Type II stacked insulator under three-phase energization (Case II). (a) Full view of one of the Type II insulators. (b) Enlarged area around the line end. tential contours are greatly influenced by the line-end hardware, the corona rings, and the jumpers. The magnitude of along the path defined close to the surface of the top insulator sheath (Case III) is shown in Fig. 15. The maximum value of near the triple junction point is 0.89 kv /mm. The calculated magnitudes along the insulation distance near the line end are also shown in Fig. 16. V. ELECTRIC FIELD STRENGTH ALONG LEAKAGE PATH It is also of practical interest to know the electric field strength along the leakage path on the surface of a nonceramic insulator, which is the path along a b c d e f g h i in Fig. 17.

QUE et al.: ELECTRIC-FIELD DISTRIBUTION ALONG DRY AND CLEAN NONCERAMIC INSULATORS 1075 Fig. 12. Electric field strength magnitude along the percent insulation distance at the surface of the insulator sheath (Case II), stacked insulator design. Fig. 15. Electric field strength magnitude along the percent insulation distance at the surface of the insulator sheath (Case III), standard insulator design. Fig. 13. Electric field strength magnitude along the percent insulation distance at the surface of the insulator sheath (Case II) near the line end, stacked insulator design. Fig. 16. Electric field strength magnitude along the percent insulation distance at the surface of the insulator sheath (Case III) near the line end, standard insulator design. Fig. 14. Percent equipotential contours for a 765 kv dead-end tower with Type I standard insulator under three-phase energization (Case III). The heavy lines represent the corona ring. The insulator of Case I in Section IV with standard shed design is calculated first. The electric field distribution along the leakage path on the surface of the insulator with standard sheds near the line-end fitting is shown in Fig. 18. The calculation results show that the electric field strength on the sheath region (b c and f g) is higher than that on the shed region (c d e f). Points with much Fig. 17. Leakage path at the surface of the insulator (a b c d e f g h i) with equipotential lines (standard shed design). (a) Leakage path with equipotential contours. (b) Equipotential contours around the shed edge (d e). higher are at the junction regions between the sheath and the shed (c, f, g). The electric field strength magnitudes on the top of the shed (e f) are very similar to the electric field strength magnitudes on the bottom side of the shed (c d). The standard shed design has two high stressed regions (at f and g) between two adjacent sheds. It is interesting that the electric field strength near the shed edge (along path d e) suddenly drops. The equipotential lines

1076 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 22, NO. 2, APRIL 2007 Fig. 18. Electric field strength magnitude along the leakage path at the surface of the insulator. shown in Fig. 17(b) are close to a direction parallel to the surface of the weather shed. The rounded shed edge presses out the equipotential lines toward the side of the shed edge. The electric field strength is slightly enhanced at the top and bottom sides of the shed edge, but it is much lower at the center of the shed edge. The electric field distribution along the leakage path of the insulator of Case II in Section IV with stacked shed design is also calculated. The leakage path along a b c d e f g h i j and equipotential lines are shown in Fig. 19. The heavy lines represent the corona ring. The electric field distribution along the leakage path on the surface of the insulator with stacked sheds near the line-end fitting is shown in Fig. 20. The calculation results show that the electric field strength on the junction regions (a, e, h) between two stacked sheds is higher than that on the shed region (e f g h). The stacked shed design has only one high stressed region (e.g., at e) between two adjacent sheds. The electric field strength at point c suddenly drops and is lower than at point d. The equipotential lines shown in Fig. 19(b) have to follow the contour of the ring keeper. Since point c is closer to the ring keeper than point d, the ring keeper shows more shielding effect at point c. VI. EFFECTS OF THE POSITION OF THE CORONA RING The calculations assumed fixed dimensions of the corona ring, designed by the manufacturer. One variable that can be adjusted is the position of the corona ring. To investigate the effects of the corona ring position on the electric field distribution in the vicinity of the line-end fitting of a insulator, the insulator with standard shed design and a 765-kV suspension tower is used (Case I). The dimensions and position of the line-end corona ring are shown in Fig. 21. The height of the corona ring above the line-end fitting is defined as. The range of is from 1.8 to 11.8 cm. The maximum electric field strength at the triple junction point as a function of is shown in Fig. 22. The position of the corona ring is very important in order to control the electric field strength distribution in the vicinity of the line-end fittings. As the corona ring is moved from the line end toward the ground end in the range of investigated, the maximum electric field strength at the triple junction point is reduced. However, the dry arcing distance between the line end Fig. 19. Leakage path at the surface of the insulator (a b c d e f g h i j) with equipotential lines (stacked shed design). (a) Leakage path with equipotential contours. (b) Equipotential contours around the line-end electrode. Fig. 20. Electric field strength magnitude along the leakage path at the surface of the insulator. Fig. 21. Dimensions and positions of the line-end corona ring. Fig. 22. Maximum electric field magnitude at the triple junction point as a function of the corona ring position. and the ground end is also reduced as the corona ring is moved toward the ground end.

QUE et al.: ELECTRIC-FIELD DISTRIBUTION ALONG DRY AND CLEAN NONCERAMIC INSULATORS 1077 VII. VERIFICATION TESTS In order to verify the calculation results by experimental results, a series of experiments was conducted in the High Voltage Laboratory of The Ohio State University. The electric field strength distribution along a dry nonceramic insulator was measured by a Positron insulator tester. The 34.5 kv insulator, which is shown in Fig. 2, was used for the verification tests. The insulator was tested in the vertical position. The tower window was simulated by a grounded supporting structure. A 1 m long, 2.3 cm diameter conductor was connected to the line end of the insulator. The experimental setup and dimensions of the grounded supporting structure are shown in Fig. 23. The insulator tester was attached to a horizontal hot stick, which was supported by an insulating stand. The height of the hot stick above the ground plane could be adjusted easily. The insulator tester was moved along a vertical line, which was about 18.5 cm away from the center line of the insulator tested. The applied voltage on the conductor was 30 kv(rms) or 42.43 kv. The insulator tester was moved from 141 to 93 cm above ground by 2.54 cm steps. The vertical component of the electric field strength along the insulator was measured and calculated as well. The measurements and calculation results are shown in Fig. 24. Good agreement between the measurements and calculation results is demonstrated. It shows that the calculation results using the Coulomb software package are good and reliable. The reasons for the small discrepancies are: spacial resolution limitations of the insulator tester; slight movement of the tester during data taking; distortion of the electric field distribution due to the presence of various objects in the laboratory during measurements as well as the mounting (hotstick) of the insulator tester itself. VIII. CONCLUSION 1) Full and simplified models of a dry and clean 34.5 kv nonceramic insulator with 12 weather sheds have been developed for calculations. The electric field and voltage distribution (EFVD) of the insulator with the full number of weather sheds can be estimated through a simplified insulator model with a smaller number of weather sheds at each end of the insulator. 2) Three practical cases have been modeled and their EFVD has been calculated along the insulators used. The cases represented various combinations of 765 kv nonceramic insulators (i.e., standard and stacked shed nonceramic insulator designs), and towers (i.e., suspension and dead-end types). 3) The electric field strength ( ) magnitude was calculated and analyzed i) along a straight path between the two metal fittings of the insulator, generating plots of as a function of the insulation distance in percent, and ii) along the surface of the insulator, generating plots of as a function of the position of the test point of the leakage path. Fig. 23. Verification test setup and dimensions of the grounded supporting structure. Fig. 24. Electric field distribution along a dry insulator measured by the insulator tester (*) and calculated by the simulation model (0). 4) Effects of the position of the corona ring at the line-end fitting of a 765 kv nonceramic insulator on the electric field strength magnitude at the triple junction point were also calculated. 5) Verification tests were conducted to compare the calculation results with measurements. The calculation results and measured values of were in good agreement. APPENDIX SOFTWARE USED For the studies described in this paper, a commercially available program Version 5 of Coulomb (Integrated Engineering Software), based on the boundary element method, has been employed. It is a three-dimensional electrostatic and quasistatic simulation software, which is especially suited for applications where the design requires a large open-field analysis and exact modeling of the boundaries. An AMD Athlon 1.3 GHz computer was used for the calculations with 700 MB random-access memory (RAM). The number of elements used in the calculation model for Case I is 21126 and the calculation time is about 9 hours. As an example, the element configurations on the surface of the insulator and the corona ring at the line end are shown in Fig. 25.

1078 IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 22, NO. 2, APRIL 2007 Weiguo Que received the B.S. and M.S. degrees in electrical engineering from Tsinghua University, Beijing, China, in 1994 and 1997, respectively, and the Ph.D. degree in electrical engineering from The Ohio State University, Columbus, in 2002. In 2002, he joined Axcelis Technologies Inc., Beverly, MA, an ion implanter company. He is a Senior Electrical Engineer, responsible for the high-voltage system and insulation design for the ion implanters. His main research interests are electric field and magnetic field analysis, high-voltage power-supply designs, and ion-beam acceleration. Fig. 25. Element configuration on the surface of the insulator and the corona ring at the line end. ACKNOWLEDGMENT The authors would like to thank C. Armstrong, General Manager of Integrated Engineering Software, for his support, which was invaluable to this study. REFERENCES Stephen A. Sebo (LF 02) received the M.S.E.E. degree from the Budapest Polytechnical University, Budapest, Hungary, in 1957, and the Ph.D. degree from the Hungarian Academy of Sciences, Budapest, in 1966. Between 1957 and 1961, he was a Laboratory and Test Engineer of the Budapest Electric Company, Budapest. He was a Faculty Member of the Budapest Polytechnical University between 1961 and 1967. In 1968, he joined The Ohio State University (OSU), where he was a Full Professor from 1974 to 2003. He has been a Professor Emeritus since 2003 still active in teaching, research, and professional service in the electric power and high-voltage areas. His research focuses on high-voltage engineering, electromagnetic field (EMF), and electromagnetic compatibility (EMC) topics. Prof. Sebo was the 1981 recipient of Edison Electric Institute s Power Educator Award, and received the 1982 Best Paper Award from the IEEE Power Engineering Society with Ross Caldecott. In 1982, he was appointed the American Electric Power Professor at OSU. From 1995 to 2003, he was the Neal A. Smith Professor at OSU. He was named Technical Person of the Year by the Columbus Technical Council in 1994. [1] Electric field modeling of NCI and grading ring design and application, Palo Alto, CA, Tech. Rep. 113977, Dec. 1999, EPRI. [2] J. Philips, D. J. Childs, and H. M. Schneider, Aging of non-ceramic insulators due to corona from water drops, IEEE Trans. Power Del., vol. 14, no. 3, pp. 1081 1089, Jul. 1999. [3] T. Zhao and M. G. Comber, Calculation of electric field and potential distribution along nonceramic insulators considering the effects of conductors and transmission towers, IEEE Trans. Power Del., vol. 15, no. 1, pp. 313 318, Jan. 2000. Robert J. Hill (M 84) is Materials and Product Manager of Nonceramic Insulators with MacLean Power Systems, Franklin Park, IL. He is active in numerous IEEE Insulator Working Groups, as well as NEMA and ANSI activities related to insulator standardization. He serves as the U.S. representative on CIGRE Working Group B2.03 (Insulators) and is active in IEC insulator standardization activities, serving as the U.S. member on IEC TC36WG11 and TC36WG12.

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