Transformer Design And Application Considerations For Nonsinusoidal Load Currents.pdf

TRANSFORMER DESIGN AND APPLICATION CONSIDERATIONS FOR NON-SINUSOIDAL LOAD CURRENTS Linden W.Pierce, Member, IEEE General Electric Company 1935 Redmond Circle Rome, Georgia 30165-1319

Abszrucf-The use of adjustable speed drives requires transformers capable of withstanding high levels of harmonic currents under normal operating conditions. Experience has been that overheating problems are much more common with dry type transformers than with liquid filled transformers. Transformer insulation life is determined by the hot spot temperature but confirmation of hot spot temperature rise is one performance characteristic which is ignored in industry standards. This is especially important for transformers rated for non-sinusoidal load currents. Hot spot allowances used in IEEE standards for ventilated dry type transformers were developed in 1944 and recent data indicates that revisions are required. The design of transformers for non-sinusoidal load currents should include an analysis of the eddy loss distribution in the windings and calculation of the hot spot temperature rise. Calculations and thermal tests giving only average winding temperature rises are not sufficient. Thermal tests with non-sinusoidal currents and measurements of hot spot temperature rises are extremely difficult on large transformers. The combination of testing and analysis may be the only economically practical approach. Analysis indicates that the dry type transformer hot spot temperature is very sensitive to the eddy loss magnitude and distribution. The UL K-factor rated dry type transformer and the recommended practices given in ANSIlIEEE C57.110 are reviewed. When purchasing transformers subject to non-sinusoidal load currents, considerations should be given to the manufacturer's development program and capability to calculate the eddy loss distribution and hot spot temperatures. I. INTRODUCTION

Common sources of harmonics in industrial electrical systems are rectifiers, dc motor drives, uninterruptible power supplies, and arc furnaces. The use of adjustable speed drives requires transformers capable of withstanding high levels of harmonic currents under normal operating conditions. Heating due to nonsinusoidal load currents has become and important (hot) topic in the last several years. Today almost every conference of the IEEE Power Engineering or Industry Application Societies has a panel session on power quality or harmonics. Many of the speakers at the panel sessions are consultants who continue to present the same information at conference after conference to promote their companies. Many trade press articles with erroneous information have also appeared. Scare headlines and advertisements such as, "Harmonics-cancer in the electrical body", and "Don't let the harmonic demon destroy your transformer", are used to promote

0-7803-2456-0195 $4.00 0 1995 IEEE-IAS

a particular product or expertise. Failures have been reported [ 1,2] in dry type transformers loaded at less than nameplate kVA. The K-factor rated transformer listed by UL has been advertised as a solution for these applications. Manufacturers advertise their Kfactor rated transformers while in their trade association (NEMA) correspondence to UL claim that it is an invalid approach. Problems due to harmonics started to be recognized in the early 1980's and articles such as those by Stratford [3,4,5] started to appear in the IEEE Industry Applications Society Transactions and Conferences. Application considerations for handling the effects of harmonics in cement plants were also reported in the paper by Smith and Stratford [6]. Indoor transformer alternatives for industrial plants include ventilated dry type, cast resin, silicone, and high temperature hydrocarbon transformers. Outdoor units are usually oil immersed unless mounted near or on the roofs of buildings. For each of these transformer options, the cost or application advantages must be evaluated. McCann [q gives a review of these options for cement plants. 11. DEVELOPMENT OF IEEE STANDARD C57.110

A. History In the March 1980 meeting of the Transformers Committee of the IEEE Power Engineering Society there was a discussion of the effect of non-sinusoidal load currents on transformer temperature rise. For industrial systems the non-sinusoidal load was a steadily increasing percentage of the total load of a load center transformer. It was suggested that some guide should be developed to assist in estimating transformer loading capability based on the amount of distortion. In May 1980 a study committee was formed and had its first meeting in October 1980. After the first meeting the study group was elevated to a IEEE Working Group of the Performance Characteristics Subcommittee. 22 representatives from manufacturers and users constituted the Working Group. An IEEE paper by Alexander D. Kline [8] of Southem Transformer Company was presented at the IAS Annual meeting in 1981 and distributed to the Working Group membership. This paper first presented the methodology used in the C57.110 document by considering the eddy loss to vary as the square of the current and harmonic order. After several preliminary drafts the first draft for balloting of C57.1101D1, "Recommended Practice for Establishing Transformer Capability When Supplying Nonsinusoidal Load Currents", was completed October 29, 1982. The methodology was also described in a 1984 paper by Rice [9] using information from a draft copy of the document. The document [101 was completed and issued as IEEE Std. C57.110 in 1986. In 1989 this Working Group received the IEEE Outstanding Working Group Award recognizing contributiom-to the industry, originality of work, and timeliness of publication. Kennedy and Ivey [ll] in their 1990 paper described application and design considerations for transformers containing harmonic currents based on the IEEE (257.110-1986 Recommended Practice.

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The intent of IEEE C57.110-1986was to give procedures to determine the capability of an existing transformer for nonsinusoidal load currents. The methodology determines the derated current magnitude to allow for the increased harmonics. The symbolization used in the 1986 document is overly confusing to those not knowledgeable of transformer design or terminology. For example, Balda [12]erroneously calculated that the more a small single phase distribution transformer was derated the more load it could carry! An IEEE Working Group of the Transformers Committee. has been formed to revise IEEE (37.110. It has produced a draft [13] for discussion but has not been balloted.

The eddy loss is assumed to vary with the square of the rms current and the square of the frequency (harmonic order h), i. e., 'EC

h

L 1, PER

Transformer losses consists of no-load or core loss and load losses. This can be expressed by the equation below. Pr

=

Pc + PU

where PC PLL PT

core or no-load loss, watts load loss, watts total loss, watts.

Core or no-load loss is due to the voltage excitation of the core. For non-sinusoidal load currents the primary voltage waveform is assumed to be sinusoidal. There is no allowance in (37.110 for any increase in core loss for non-sinusoidal load currents. Although the magnetizing current does consists of harmonics, these are extremely small compared with the load current and their effect on the losses is minimal. Load losses consists of f R loss, eddy loss, and stray loss, or in equation form, pU.

=

I ~ R+ peC + pSL

where 12R p, PSL

loss due to load current and d.c. resistance of the windings, watts winding eddy loss, watts stray losses in clamps, tanks,etc., watts.

The IZRloss is due to the currents in and the dc resistance of the windings. Surprisingly this group of terms has no name other than IZR although "ohmic losses" has been suggested by Ravot and Kreuzer [14] and Blume et. a1 [I51 refer to it as "(d.c) ohmic losses." The ohmic loss is affected by the magnitudes of the harmonic currents but not the frequency. The ohmic loss is determined by measuring the d.c. resistances using a d.c. current and voltage then a calculation is performed using the winding currents. There is no test method to determine individual winding eddy loss or to separate transformer stray loss from eddy loss. The total stray and eddy loss is determined by measuring the total load loss during the impedance test. The total stray and eddy loss is determined by subtracting the ohmic loss from the load loss, i.e.,

PeC

+

=

pU. - PR

(3)

= 'EC,R

2

h2

(4)

where

IR

B. Review of Transformer Losses

2 [?]

k=h,

harmonic order, 1,2,3,etc. the greatest harmonic order to be considered current at harmonic order h, amperes rated current, amperes eddy loss at rated current and frequency.

Bewley [16]in discussing eddy loss stated, "Thus at very high frequencies it varies as the square root of the frequency, whereas at low frequencies it varies as the square of the frequency." Stigant and Franklin [lq stated, "Various formulae have been propounded from time to time for calculating this lqs, but there are so many factors which enter into the calculation that it is more usual and practical to add a percentage on to the IzR loss rather than to attempt to calculate it by means of formulae. The percentages which it is customary to add are based upon experience with the particular type of transformer under consideration." Similar statements are made about the stray loss. The eddy loss depends on the square of the strand dimension perpendicular to the leakage flux field. At the ends of the winding the flux field bends and the larger dimension of the rectangular strand is perpendicular to a vector component of the leakage flux field. Equalking the height of the primary and secondary windings reduces the concentrated eddy loss at the winding ends. However the magnitude is still greater than the middle of the winding due to this bending of the leakage flux field. This is illustrated by a figure in the Tutorial of C57.110-1986. Approximately equal heights of primary and secondary windings can be achieved with any winding design. Reducing strand size reduces the per cent eddy loss but increases ohmic loss unless more than one strand per turn is used. If more than one strand per turn is used, eddy losses may increase due to circulating currents unless the strands are transposed throughout the winding. Small transformers inherently have small strand sizes due to low currents. Stray loss occurs due to the stray flux which introduces losses in the core, clamps, tank, and other iron parts. Stray loss may raise the temperature of the structural parts. For dry type transformers increased temperatures in these regions does not contribute to an increase in the winding hot spot temperature. For liquid immersed transformers, the stray loss increases the oil temperature and thus the hot spot temperature of the windings. Kennedy and Ivey [111 state, "Even then manufacturers disagree on the methods to use in order to calculate these stray losses". In a proposed standard [18]for semi-conductor rectifier transformers the stray loss is assumed to vary with the square of the current times the frequency (harmonic order), as shown by (5).

C. Assumptions used in C57.110

(37.110 was intended as a guide and the methodology uses data from the transformer test report. The methodology is to determine the per unit current with the harmonic load which gives the same loss density at the hot s h t location as with rated sinusoidal load

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current. The methodology applies to either liquid immersed or dry type transformers. Since the individual winding eddy loss and stray losses cannot be determined from the test data, the following assumptions are made:

2.

I

"The K-factor equals ~ Z , ( ~ Uh2) ~ b -1

where: 1) Conservative Assumptions: a) All stray loss is assumed to be winding eddy loss, b) Eddy loss varies with the square of the current and harmonic order (eq. 4),and c) The per unit eddy loss at the hot spot location is assumed to be 4 times the average per unit eddy loss. 2) Other Assumptions: a) The division of eddy loss between the primary and secondary winding is assumed to be 60 96 in the inner winding and 40 % in the outer winding for transformers having a turns ratio of 4:l or less and 70 % in the inner winding and 30 96 in the outer winding for transformers having a turns ratio greater than 4:l and having one or more windings with a maximum self-cooled current rating greater than 1 O00 amperes. b) Although not stated, it is also assumed that the transformer hot spot temperature rise does not exceed values given in the transformer standards.

I,(PU)

= the rms current at harmonic "h" (per unit of

h

= the harmonic order."

rated rms load current) and 3. "K-factor rated transformers have not been evaluated for use with harmonic loads where the rms current of any singular harmonic greater that the tenth harmonic is greater than l/h of the fundamental rms current. "

C. Relationship between K-Factor and C57.110 The UL definition of K-factor is based on using the transformer rated current in the calculation of per unit current in (6). Substituting the rated current into the UL equation for the K-factor gives;

Transformers

B. UL Definition of K-Factor

The only d e f ~ t i o nfor the K-factor rating for transfora?efsis in Underwriters Laboratory Standards 1561 and 1662 [21]. Per Paragraph 7B.l added to UL 1562 on May 12, 1992, UL defines K-factor as follows: 1. "K-FACTOR -- A rating optionally applied to a transformer indicating its suitability for use with loads that draw nonsinusoidal currents. "

h2

1~

where; 1,

In a 1988 tutorial paper Kerszenbaum, Majur, Mistry, and Frank [191 presented a tabulation based on C57.110 methodology which produced a summation of per unit current squared times the harmonic order squared which the authors' called the "K coefficient". Winding eddy loss and average winding temperature rises were calculated for transformers rated from 225 through 2 500 kVA. Calculated and tested eddy loss at 60 HZ only were compared. Hot spot temperatures were discussed but not calculated and eddy loss at the hot spot location also was not calculated. A temperature rise test was performed on a 15 kVA transformer connected to a rectified load at 58.3 96 rated. Average temperature rise of the secondary winding was measured but the hot spot temperature rise was not measured. For this small transformer, at reduced load, the average temperature rise was 33.56 'C which was only a 2.31 'C increase above the temperature rise with sinusoidal current. The paper received the Ralph E. Lee award in 1990 from the Commercial Power System Department of the Industry Application Society of the IEEE and was published in the Transactions in 1991. International Transformer approached UL and the term "K coefficient" was changed to "K-factor" and ITC achieved a marketing coup by convincing UL to adopt the K-Factor rathg approach in December 1990 and becoming the first manufacturer to obtain a UL listing for K-factor rated transformers. Presently most manufacturers of dry type transformers below 500 kVA have UL listed K-factor rated transformer designs. Frank [20]reviewed the origin of the k-factor rated transformer in a 1994 pper.

=

b=1

111. K-FACTOR RATED TRANSFORMERS

A. Introduction of K-Factor Rating of Dry

e["] 2

E FACTOR

rated rms load current of transformer.

Equation 7 is similar to (4)from C57.110-1986. Trade press articles and marketing brochures have contributed to much confusion by giving examples of K-factor incorrectly calculated using the fundamental current to calculate per-unit current instead of the transformer rated current. When the harmonic currents are less than the rated current this results in a higher but inaccurately calculated K-factor. If the definition by UL is used then the calculated K-factor for a given set of harmonic currents varies with the transformer kVA. Bishop and Gilker [22,23,24]report on a PC controlled meter which allows input of the transformer rated current to calculate the K-factor. K-factor alone is meaningless unless the rated current used to calculate the K-factor is also given. A suitable statement would be, "A K-factor of -based on a rated current of amperes".

D. UL Testing Requirements For K-factor rated dry type transformers UL requires that the stray and eddy losses for rated sinusoidal current be determined by test and then multiplied by the K-factor. A thermal test is permitted using additional sinusoidal current to generate load loss equivalent to that determined by the calculation and average temperature rises are measured. Sinusoidaltest current used for all but small units. The test method gives a lower hot spot rise than with non-sinusoidal test current. With harmonic currents the added eddy losses are concentrated at bottom and the hot spot location at the top of the windings. Current UL listing requirements require no substantiation that the hot spot temperature is less than the rated insulation system temperature class for K-factor rated designs. The effect of l d i eddy loss on the hottest spot temperature is ignored. The National Electrical Manufacturer's Association (NEMA) (representingdry type manufacturerswith UL K-factor listings) is on record with UL that the UL K-factor listing requirements are inadequate since they ignore increased hot spot temperatures due to increased eddy losses caused by harmonic currents at the hot spot location.

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E. Scope of UL 1562 UL 1562 covers single-phase or three-phase, dry type, distribution transformers provided with either ventilated or nonventilated enclosures and are rated for a primary or secondary voltage from 601 to 35 OOO volts and from 1 to 5 OOO kVA. UL 1562 does not cover arc furnace transformers, rectifier transformers, mining transformers, motor starting transformers, or transformers under the exclusive control of electrical utilities. F. Neutral Conductors

In 1988 Electrical Construction and Maintenance magazine published and article [25] entitled "Double the neutral and derate the transformer--or else", summatizing the Computer and Business Equipment Manufactures Association recommendations to prevent damage to the distribution system. Gruzs 1261 presents a survey of the magnitude of the neutral current in three-phase computer power systems. Excessive neutral currents may also occur in three-phase fluorescent lighting circuits [27]. To address this problem UL 1561 and 1562 contain special requirements for the sizes of the transformer neutral conductor to allow for the possible additive nature of triplen harmonic currents. G. Trade Press and K-Factor

When the first manufacturer obtained its K-factor listing from UL and began promoting its product, the industry did not understand the concept. A review of the only consensus industry standard, C57.110-1986, did not reveal the definition. There is a K in the document but it is used to describe other parameters. After the K-factor rated transformer became well known, advertisements and trade press articles appeared and contributed to the confusion about the K-factor. However, one of the better articles is that by Nallen [28]. It appeared that many consultants and manufacturers portrayed the K-factor concept as a complicated and mysterious subject for which only they had particular expertise. In 1993 Electrical Construction and Maintenance 1291 magazine published a series of articles from different Viewpoints and in the final article by Moravek and Lethert [30] (which was suppose to clear up the mystery) the "H-factor" was introduced without definition! IV. APPLICATION OF K-FACTOR TRANSFORMERS

.

A. K-Factor Rated Transformers in Codes and Standards Statements which have appeared in advertising literature are: "The National Electrical Code (NEC) requires that if harmonic currents are present, then transformers must be K-factor listed per the National Electrical Code sections 90-4 and 110-3", and "Transformers ahead of a variable speed drive must be listed for the purpose to comply with the NEC sections !X-4 and 110-3". A recent IEEE paper by Massey E311 stated, "The K-factor is the only acceptable method for estimating harmonic content and for specdymg distribution transformers for use under non-linear load conditions according to ANSI, IEEE, the 1993 National Electrid Code, and Underwriter's Laboratory". The facts are that K-factor is defined only in UL standards. NEMA has stated to UL that the UL K-factor listing requirements are not adequate. The 1993 National Electrical code [32] and the National Electrical Code Handbook, sixth edition [33] contain no references to K-factor rated transformers or requirements for transformers ahead of a variable speed drive. NEC sections 90-4

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and 110-3 apply to all electrical equipment [34] and are not specific to transformers. The information in these and related sections is condensed (based on the NEC and NEC Handbook) as follows: "All equipment required or permitted by the NEC shall be acceptable only if approved. Approval of equipment is the responsibility of the electncal inspection authority. It is the responsibility of the local authority enforcing the NEC to interpret the specific rules of the NEC. Listing or labeling is considered as a means of establishing suitability." The 1993 NEC does include one fine print note (FPN) after article 222-22, Feeder Neutral Load, that states, "A 3-phase, 4-wire power system used to supply power to computer systems or other similar electronic loads may necessitate that the power system design allow for the possibility of high harmonic neutral currents". A proposal was made for the 1996 NEC to require that all transformers be labeled with a K-factor rating even if it was 1. This was rejected. The 1994 NFPA Ad hoc Committee on NonLiear Loads in its report [35] agreed with the code making panel action and stated, "The loading of a transformer must account for the total load including the increased heating effects of nonlinear load currents. Several methods for dealing with the heating effects of nonlinear loads have been presented including; derating, oversizing, increased insulation ratings, thermal protection systems, and K-factor transformers. The optimum method for dealing with transformer overheating will vary depending upon several technical and economic factors, and should be considered in the design phase for the electrid system". The user must decide (based on the local inspection authority and requirements) if the NEC applies to any electrical equipment in the instabtion. For an excellent discussion of this topic refer to the recent pap& by Goldberg and Castenschiold [36]. The majority of the transformers supplied by manufacturers for variable speed drive applications are not listed by UL. Generally these units are designed to sustain specified unique nonsinusoidal load currents.. There are no ANSI standards which define K-factor. "K-factor" is not defined in current IEEE standards. The letter "K" appears in many iEEE standards to describe many other parameters. The term does appear in the 1994 IEEE Red Book [37] which states "An equation developed in IEEE Std C57.110-1986 produces a value referred to as the K-factor and has helped in rating a transformer's ability to wry harmonic currents"; and, "An equation presented in E E E Std. (37.110-1986 includes a tabulation of the per-unit current squared times the frequency squared that, when summed up for each harmonic, produces a value refmed to as the K-factor. The K-factor is related to the eddy-current loss in the winding conductors." B. Application Considerations for Combining Non-Sinusoidal Load Currents

According to the definition in UL standards, K-factor is a rating applied to dry type transformers. In the UL definition the per unit current to be used to calculate K-factor is obviously the transformer rated secondary current at rated kVA. However, the industry has begun to assign K-factors to the various electricalequipment which produce harmonic load currents; computers, incandescent lights, etc. Because these equipment "K-factors" are based on per unit currents different than the transformer per unit rated current, a method is needed to determine the required transformer rating for new installations. Methods have proposed and some consultants claim particular expertise in this field and have published magazine articles to publicize their names. Tests reportedby Arthur [38] show that when many individual harmonic

producing devices are present the combined load K-factor is less than the individual K-factors of the devices. Methods are needed to design systems with many different types of equipment of differing non-sinusoidal loading in order to specify the correct transformer rating. One method has been proposed by Massey [3 11 however it has not been confirmed by testing. Present methods consider the non-sinusoidal load current content to be constant with time. Load diversity, load cycle variation during the day, week, and year should be considered and these topics are starting to be addressed. The time duration of harmonics was addressed by Kaprielian and Emanuel [39] using statistical methods. The summation of randomly fluctuating loads produced by acldc power converters has also been studied by Wang, Pierrat, and Wang [MI. Simple, practical methods, are needed.

V. TRANSFORMER INSULATION LIFE AND RELIABILITY A. Introduction In 1970 articles appeared in the trade press about the life expected of dry type transformers. The first of these was by Jerome M. Frank [41,42] of Sorgel Electric Corp. Frank claimed a 2.3 year life at 220 ' C and a 4.6 year life at 210 'C. Franks's conclusion was based on a premise (supposedly in industry standards) that half-life was reached in 20 OOO hours at rated temperature. Rebuttal articles appeared by Dutton and Antalis of General Electric [43] quoting industry standards and stating that service experience supported a longer life than that stated by Frank. At the time of the articles, the loading guide [U] for dry type transformers only gave a per unit life curve and the 150 'C average temperature rise class applied to sealed gas transformers only and not ventilated .dry units. These articles must have caused confusion to specifyers of dry type transformers. Lazar [45] (an electrical consultant) stated in a 1977 article, "The weakest point of the class H insulation is its life expectancy at the ultimate temperature of 220 'C. Indeed Class H materials must withstand 150 'C rise for a minimum of 20,000 hr. At full load and assuming a possible 40 'C ambient for 24-hr. continuous operation, the 20,000 hr would disappear in 2.3 years. Based on these assumptions, it would be good practice to design for a Class H with 115 'C rise only." However Reason [&I, assistant editor of Power, wrote in 1980 about ventilated dry type transformers, "Their normal life expectancy of 20 years can be drastically increased if they are operated at less than rated capacity. " Concerns about life of transformers is also evident in modem times, In response to a 1990 readers quiz [47] in Electrical Construction and Maintenance, one reader replied, "As a good engineering practice the transformers should not be loaded more than 80 96 of the nameplate rating". Another response was, "A.A.G may add additional load of up to 450 A, thus bringing the transformer loading up to 1 800A. However, if this total load is continuous, the transformer will probably not last longer than ten more years, as transformers are generally not designed for fullrated continuous load." The quiz statement did not state whether the unit was a liquid or dry type. B. Dry Type Transformers

Test procedures for thermal evaluation of dry type transformer models is documented in IEEE standard C57.12.56 [48]. A life of 40 OOO hours is to be demonstrated for the insulation temperature class at the maximum rated ambient temperature. In dry type transformers standards [55] the kVA rating is based on an average

24 hour ambient of 30 'C with a 40 'C maximum. The rated hot spot temperature based on the 30 'C ambient is 210 'C for 150 'C average winding rise system and the maximum hot spot temperature is 220 "C. The life for the model would then be expected to be approximately 80 OOO hours at 210 'C. The life to be demonstrated by model testing is thus one-half the life expected for full size equipment as shown in the 1989 loading guide [49]. This guide shows a life of 175 680 or 20.05 years at the rated the hot spot temperature of 210 'C.

C. Liquid Immersed Transformers For liquid filled distribution transformers thermal evaluations are conducted per C57.100 [51]. The present document states in one place that a 60 OOO hour life is required to be demonstrated by model testing. This appeats to be a carry over from the first proposed method of thermal evaluation [52]. Initially the principles in the liquid thermal evaluation and the dry type thermal evaluation guides were similar. That is, that model aging tests were overly pessimistic and that the life of actual eqspment would be longer. After manufacturers conducted thermalaging programs on distribution transformers, no failures occurred after aging times several times that shown in the loading guide. Currently the distribution transformer loading guide is based on a 20 year life at rated hot spot temperature of 110 'C. For distribution transformers industry experience [53] has also been to require a safety factor of 5 times the life in the loading guide. The revision of C57.100 [54] will also require that a life of 180 OOO hours at rated hottest spot temperature be demonstrated by using full size units for distributions transformers and models for power transformers with a safety factor of 5 for distribution transformers and 2 for power transformers.

D. Cast Resin Transformers

Dry type transformers with solid cast windings of epoxy resin were developed in Europe, and this transformer design began to be widely accepted in the United States in the 1980's. The solid cast or residencapsulated transformer was incorporated into the first IEEE standard [ S I in 1989. There are many cast-resin transformer designs available with different insulation temperature classes. Current practice by manufacturers has been to rely on tests of individual materials to determine the rated insulation temperature class to assign to their designs. Operating experience indicates this gives acceptable life when cast-resin transformers are operated at nameplate ratings. At the present time, the industry has not established Arrhenius insulation aging curves to give loss of insulation life for cast-resin transformer windings operated above the rated insulation temperature class. A trial use standard NEE (37.12.60 [Sa] for thermal evaluation of insulation systems for solid cast-resin transformers will serve as a standard test method for life test models for d e " h b g the rated insulation temperature class of cast-resin transformer windings. The materials and coil design techniques used in cast-resin transformers necessitated a document to recognize factors such as the effect of glass transition temperature, higher resin to air and metal ratios, filler contents, and conductor identity on aging and performance. An arbitrary extrapolation criteria of 40 OOO hours was selected for the evaluation. In addition to determining the rated insulation temperatwe class, thermal evaluation standards give data for preparing loading guides. At this time no test data in accordance with the trial use standard IEEE CJ7.12.60 bas been reporkd. There have been reports of difficulties in using the test methods of C57.12.60 and an IEEE Working Group of the Transformers Committee has been

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formed to revise the document. The effect of cracking or softening of the epoxy due to thermal cycling to elevated temperatures could be the limiting factor in overloading cast-resin transformers and not insulation aging. Due to the lack of aging data, the draft [57] guide for loading cast-resin transformers uses a concept of loading above rating based on limiting hot spot temperatures to determine loading capability of cast-resin transformers.

E. Transformer Reliability

Dry type transformers are replacing liquid filled units in many industrial plants. Thermal failures have been reported on dry type transformers subjected to non-sinusoidal load currents even when the RMS current was below rated. The aging model thermal evaluation requirements discussed previously indicate a longer life expectancy for liquid immersed transformers. There is a difference in philosophy between liquid and dry type thermal evaluation standards. The dry type thermal evaluation requirement is that model aging demonstrate a life of one-half the 20 year life given in the loading guide. For liquid immersed transformers the next thermal evaluation test requirement will be that a life of 2 or 5 times the 20 year life in the loading guide be demonstrated. There is an argument that model aging testing is generic and that dry type ageing models could also be expected to last 5 times the life given in the loading guide, similar to liquid units. Except for the limited Crouse and Hutchinson data [50] for inorganic systems which are no longer used, there is no published aging model data to support this conclusion for the organic insulation systems used currently in dry type transformers. Although insulation life is based on the hot spot temperature, investigation of hot spot temperatures has been ignored until recently. Hot spot temperature increments in excess of 30 C may be a factor in failures of dry type transformers with non-sinusoidal load currents and this needs further study. The last reliability survey [58] of transformers in industrial plants was conducted in 1979. There is a need for an up to date reliability survey. VI. TRANSFORMER DESIGN CONSIDERATIONS A. Introduction Many transformer manufacturers have developed designs rated for nonsinusoidal load currents while optimizing their product costs. In addition to material costs the optimizationdepends on the manufacturer’s labor rates, production volume, and investments in expensive production equipment required for automated manufacturing. Every manufacturer claims its own specific design has features which reduce stray and eddy loss. Design of transformers for non-sinusoidal load currents should include an analysis of the eddy loss distribution in the windings and calculation of the hot spot temperature rise. Eddy losses due to the leakage flux distribution are concentrated in the ends of the winding. Thermal studies should be conducted using embedded thermocouples installed in test windings and prototype transformers to measure hot spot temperature to refine mathematical models to A three dimensional calculate the hot spot temperature. mathematical model should allow for eddy loss variation in winding. Analysis of the eddy loss distribution may be performed using finite element or other type computer programs. Commercial programs are available. A combination of testing and analysis may be the only economically practical approach for large transformers above about 300 kVA.

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B. Electromagnetic Analysis The subject of harmonics has received much publicity in recent times leading to the belief that the industry is only beginning to understand the effect of harmonics and to calculate the increased eddy losses. It seems that each generation must relearn the engineering principles discovered by its ancestors. The facts are that many companies and individuals have made significant contributions to the design, construction, and testing of transformers supplying non-linear loads. As early as 1906 Field [59] presented a paper giving analytical results for the eddy losses in conductors located in a magnetic field. Although this was for conductors located in an iron slot, this had similarity to transformer windings located within the window region of the iron core studied by B a g [a]. Eiarly investigations were highly mathematical such as that by Roth [61] in 1928 or graphical flux plotting such as by Stevenson and Park [62,63] in 1926 and 1927. The flux plots given in these early papers were every bit as detailed and probably as accurate as those produced by modem computer programs. An English language review and summary of the many French language papers of Roth is given by Hammond [MI. Stephens [65] in 1934 and Kaul [66] in 1957 also developed methods to determine eddy current losses in windings. With the invention of the computer, methods were developed to use this machine to compute electrical fields and eddy losses in transformers. Mamak and Laithwaite [67] used finite difference methods. The paper by Rabins [68] from 1956 is another example. In a 1985 paper Konrad [69] reviewed 171 references about eddy curreats and modeling. Figures 3 and 5 from his paper illustrate flux and current density plots for a prototype transformer manufactured by the author’s company. Many commercial computer programs are currently available and many run on a PC. A list is given in the 1989 ZEEE S p e c ” article by Cendes [70]. These computer programs produce elegant plots however their accuracy cannot be proven. One example is the paper by Pavlik, Johnson, and G ~ g i [71]. s In a comparison of calculations with test results, the total loss including the easily calculable ohmic loss which indicated a higher accuracy in the calculation of eddy and stray losses than actually existed. The ohmic loss should be subtracted from the total load loss to obtain the stray and eddy loss. A comparison of test and calculated stray and eddy loss should then be made. Losses due to harmonic currents in rectifier transformers also appears to be well understood as evidenced by the papers of DeBlieux [72,73] in 1931 and 1937. More current papers are those by Crepaz [74,75] from 1970 and 1975. Crepaz used the letter “k“as a ratio of eddy loss with harmonics divided by eddy loss with the fundamental harmonic only. Recently Ram, Forrest, and Swift [76], and Forrest [77] reported on test methods to determine the variation of losses with frequency. Small currents at different frequencies are input and the losses measured. A curve fit equation was developed to predict the eddy current and stray loss variation with current and frequency. Manufacturers who have attempted to use the methods have reported varied results. C. Thermal Analysis Although hot spot temperature is a performance parameter [55] to be met by the manufacturer, there are currently no test methods or requirements that this parameter be measured on production or prototype transformers. This seems extremely peculiar since it is the fundamental parameter determuun ‘ ’ g the life of the equipment and the thermal evaluation standards give test methods for determining the rated hot spot temperature of an insulation system. Most manufacturers of dry type transformers simply add 30 C to

the average temperature rise (calculated using empirical equations) and claim this is "per the standards", although the IEEE standards do not state that this is permissible practice. IEEE standard (37.12.01-1989 requires that both average winding temperature rise and hot spot temperature are limits to be met at rated kVA. The difference between these two limits is 30 ' C but the 30 " C does not appear as a separate number. Many dry type transformer manufacturers would like users to believe that the subject of dry type hot spot temperatures is a mysterious subject. They would like you to believe that almost nothing is known on this subject with no published papers, etc. They also claim that the hot spot temperature is difficult to measure and that since there is no IEEE test method, they are under no obligation to conduct testing, etc. They also convey the impressionthat temperatures in transformer windings are somewhat uniform and the 30 'C allowance is for localized heating at a spot which cannot be predicted. They also try to give the impression that the 30 ' C allowance was adopted in standards as a maximum number and that it is usually less that this value. Some even claim that most manufacturers take steps to see that the allowance is always less than 30 "C in their designs. ANSIHEEE Standard 1 [78] supports these arguments by stating that the hot spot allowance value is arbitrary, difficult to determine, and depends on many factors, such as size and design of the equipment. The hot spot is simply the highest temperature in the transformer windings. It is a naturally occurring phenomena due to the generation of losses and the dissipation of those losses by the heat transfer mechanisms. All transformers have a hot spot since at some place in the winding the temperature is the highest. Many commercial computer programs are advertised to predict temperatures in equipment. Readers of magazines such as Muchine Design or Design News in which these programs are described would naturally expect manufacturers to have the capability to calculate the maximum temperature in the transformer windings. Similarly readers of magazines such as the IEEE Specfrutn would naturally expect manufacturers to have the ability to calculate the eddy loss distribution within the windings. With modem computer technology it should be possible to predict the hot spot temperature. Dry type transformers have unique heat transfer characteristics which are not well known. For this reason the author developed a proprietary 3-D program for heat transfer calculations in dry type windings. The program was refined by tests on windings and prototype transformers with embedded thermocouples. There have been extensive investigations of liquid filled transformer heat transfer phenomena with many published papers. Some companies have chosen not to publish their test results and computer calculation methods and expertise depends on the manufacturer. Harold Moore, a consultant, recently reported that information submitted by manufacturers of large transformers during design reviews show that some have detailed computer analysis of the thermal design but others calculate the hot spot simply by adding 15 'C to the average temperature rise. An accurate electromagnetic and thermal analysis is essential for transformers designed for non-sinusoidal load currents. An example of this type analysis for liquid immersed transformers is given by the paper of Preiningerova, Kahoon, and Doiezel[79] and in the book by Karsai, Kerenyi, and Kiss [80]. Heating due to harmonic currents may also occur in utility transformers [811.

VII. THERMAL INVESTIGATIONS FOR DRY TYPE TRANSFORMERS A. History of the 30 'C Allowance Based on the 1944 experimental works of Stewart and Whitman [82] and Satterlee [83], NEE standards used a hottest spot allowance of 30 * C for 80 'C average temperature rise. The conclusion of Stewart and Whitman from 1944, that the ratio of hottest spot temperature rise to average temperature rise was constant, appears to have been forgotten in present industry standards for ventilated dry type transformers. The 30 'C hottest spot temperature allowance established in 1944 for 80 'C average temperature rise was approximately correct. At that time average kVA ratings were less than the present. The 150 ' C average winding temperature rise for the 220 'C insulation temperature class was extended to ventilated units initially in NEMA Standards [84] and later in IEEE standards [55] and the 30 'C hottest spot allowance incorrectly retained. The 1989 LEEE standard [55] used a constant 30 'C hottest spot allowance for all insulation temperature classes and all size transformers. The IEC standard [85] uses a variable hottest spot allowance from 5 " Cto 30 'C.

B. Recent Hot Spot Investigations-Dry Type Transformers The author and his company have conducted the following tests on dry type transformer windings and prototypes and the results shared with the industry [86,87,88,89,90]: 1. 2 OOO kVA cast resin transformer, 300 imbedded thermocouples, 22 test runs. 2. Six thermal test coils of different duct sizes, 142-171 thermocouples, 97 test runs for fundamental heat transfer data. 3. 2 500 kVA prototype with disc high voltage windings and embedded thermocouples. 4. 2 500 kVA prototype with sheet conductor low voltage winding containing embedded thermocouples. The above investigations represent the most comprehensive investigation of dry type transformer heat transfer since that of Stewart and Whitman [82] 50 years ago. The author is the only investigator to determine the effect of core loss on the hot spot temperature by conducting comparative loading back and short circuit heat run tests. The author discovered that the ratio of hot spot rise to average rise was essentially constant for a given winding and depended on duct size and winding height.

C. Hot Spot Temperature of Cast Resin Transformers The author's test data indicated a ratio of hot spot rise to average rise of 1.20 to 1.25 for the cast resin transformer design tested. This is lower than the ratio of 1.4 to 1.5 for large ventilated dry type transformers due to the extremely large cooling duct sizes and the large amount of winding surface area available for radiation heat loss. Cast resin designs with smaller internal cooling ducts would be expected to have hot spot ratios similar to ventilated dry type transformers.

D. Hot Spot Temperatures of Ventilated Dry Type Transformers The author discovered that there is a large thermal gradient from the bottom to the top of the windings of large (above 500 kVA) drytype transfor". Thisgradientis dependant onthe

41

height (length) of the winding. IEEE standards utilize a constant 30 'C allowance between the hot spot temperature rise and the average temperature rise. To meet hot spot temperature rise limits requires that the average temperature rise be reduced below guarantees because the difference between hot spot and average temperature rise may be more than 30 'C. The recent data indicated that the hot spot increment may be up to 60 'C on large transformers (generally above 500 kVA). A comparison of the author's test results with the IEEE assumed temperature gradient for a typical winding is shown in Figure 1. Changes in IEEE standards to more accurately reflect the relationship between average temperature rise and hot spot rise in large (above 500 kVA) dry type transformers has been proposed. An IEEE Working Group has been formed to review these proposals for a future revision of (37.12.01 Table 4A. It will be difficult to reconcile the divergent opinions on this subject and most manufacturers are unwilling to change IEEE standards since it increases their transformer cost.

TEST -

The analysis was performed with the calculated average eddy loss of 1.46 per cent of the ohmic loss for K-factors of 1 and 13. A calculation was performed with a slightly higher eddy loss of 2 % to show the effect of the uncertainty on the hot spot rise and average rise. The results are summarized in Table I1 and the temperature distribution for eddy loss distribution case number 2 is shown in Figure 2. The sensitivity analysis indicates that the hot spot temperature rise is influenced considerable more than average winding rise by changes in eddy loss magnitude and location. Test methods which simply increase sinusoidal current to give the same losses will give approximately the same average temperature rise but hot spot temperatures may vary considerable. TABLE I EDDY LOSS DISTRIBUTION DISTRIBUTION

DESCRIPTION

Case 1

Eddy loss was distributed throughout the winding and the eddy loss at the top and bottom five per cent of the winding was four times the average eddy loss distribution.

case 2

All the eddy loss was assumed to be at the top and bottom five per cent of the winding.

IEEE

..........

................

U

TABLE I1 CALCULATED TEMPERATURE FUSE VARIATION WITH EDDY LOSS MAGNITUDE AND DISTRIBUTION [/-.cBOTTOU

0

20

I

TEST 40

40

60

80

100

PER CENT HEIGHT

Fig. 1. Comparison of test data with IEEE standards.

VIII. EXAMPLE CALCULATIONS FOR VENTILATED DRY TYPE TRANSFORMER A dry type transformer was required rated AA-T-60-300@ 13800-950with a harmonic current content equivalent to a k-Factor of 13. Loss data was as follows: PR LV 1% HV Eddy Loss LV Eddy Loss HV stray Loss

7 650 watts 15 446 watts 326 watts 225 watts 1 744 watts

KDIST. EL FACTOR CASE %

AVE. FUSE 'C

HOT SPOT RISE 'C

1 1

1 1 2

1.46 2.00 1.46

80.2 80.5 80.2

128.1 128.7 129.7

13 13 13

1 1 2

18.98' 26.018.98'

89.0 92.3 87.5

151.9 160.7 176.9

1

'1.46 x 13 = 18.98

0

"2.0 x 13 = 26.0

K = l -

K = 13

----

For this transformer the high voltage winding had a higher hot spot temperature even though the eddy loss was lower than the low voltage winding. The low voltage winding used a sheet conductor which provided good heat conduction from the high eddy loss region. The electromagnetic analysis indicated a eddy loss of 1.46 % of the ohmic loss for the high voltage winding. The three dimensional mathematical model [W] developed to predict hot spot temperatures in ventilated dry type transformers allows for eddy loss variation throughout the winding. The computer program was used to perform a sensitivity analysis of the relationship of eddy loss magnitude and distribution on the average and hot spot temperature rises. The distribution of the eddy loss for two cases is described in Table I.

42

PER CENT HEIGHT

Fig. 2. Effect of harmonic losses on temperature rise.

IX. HOT SPOT INVESTIGATIONS FOR LIQUID FILLED TRANSFORMERS In the decade of the 1990's the technology focus in the transformer industry is on thermal performance and loadability. Many transformers undergo planned overloading by electric utilities to maximize the return on the investment in this expensive equipment. As reported recently by W. J. McNutt [91], "...all loadability decisions are based on hottest conductor temperature, but that parameter is never measured during a thermal test, and some manufacturers never actually calculate what it would be for rated loading. ....The thermal equations provided in the loading guide are recognized to be grossly inaccurate for some cooling modes and not precisely correct for others". Since the loading guide equations were first proposed [92] in 1945, many papers and trade press articles have been written on transformer loading using these equations. It is generally stated that the equations are conservative without giving test data to substantiate this statement. The equations have been incorporated into computer programs by consultants and utilities. Until recently little research has been performed on measurements of hot spot temperatures in windings of full size transformers during transient loading. Thermal research has centered on computer modeling of steady state performance with tests on model windings to obtain data to validate the models. The purpose of these studies was to more accurately predict thermal performance at nameplate ratings to minimize the manufacturer's cost. Transformer manufacturers have investigated liquid filled transformer heat transfer phenomena for over a hundred years with many published papers. Many of the design equations are based on proprietary test data obtained by many researchers. The author [93] recently reparted on an investigation on a full size test winding with embedded thermocouples to investigate transient loading on the hot spot temperature. The tests indicated that the industry loading guides were not valid for all modes of cooling and all loading conditions. During overloads there is a time lag between the top oil temperature rise in the tank and the oil rise in the winding cooling ducts. Thermal tests were also performed on a 5 600 kVA prototype transformer at various loads to compare the hot spot temperature with oil and silicone fluids. The test data was used to refine an analytical computer program to calculate hot spot temperature. Modem computer technology now pertnits more refined calculations of loading capability. An improved system of equations was developed by the author [94]for inclusion in the next loadiqg guide [95]. The improved system of loading equations is based on the fluid flow conditions occurring in the transformer during transient conditions. The equations and computer program also incorporated an improved method of calculation for variable load cycles and variable ambient temperatures. The properties of the low flammability fluids polydimethylsiloxane (silicone) and high temperature hydrocarbon (hthc) were also incorporated into the improved loading equations. In the 1980s direct measurement of winding hot spot using fiber optic detectors became feasible. In the 1990's thermal research is moving into studies of overload performance using these fiber optic hot spot detectors. Due to the high cost and poor reliability of the fiber optic devices their use appears to be primarily as a research tool to measure hot spot during factory thermal tests on transformers and for use on selected units to obtain dynamic loading data under actual field conditions. Due to the poor reliability, redundant probes must be used. Experience such as that reported in a recent Doble paper by Troisi [96] indicates that even with extreme care in installation, about 25 to 33 percent of the probescanbe expected to failduetodamage. Reported

data and'the author's experience indicates that the location of the hot spot is difficult to determine. A recent CERE [97] survey reported that two to eight sensors would be adequate if placed in the winding where the higher temperature is expected but for prototype transformers it was estimated that twenty to thirty sensors would be required.

X. EXAMPLE CALCULATION FOR LIQUID FILLED TRANSFORMER A oil immersed transformer rated OA-T-60-3600 kVA-22000Y575 was designed for a specified harmonic current content. After installation actual harmonic currents were measured and the current spectrum supplied to the manufacturer with a request to check the temperature rises. At rated load and 60 Hz.the tested losses were:

4 072 watts 27 821 watts 4 060 watts 35 953 watts

No Load PR Stray and eddy loss Total loss

Temperature rises above ambient were:

48.1 'C 47.6 'C 47.2 'C 55.3 'C

HV Ave. Rise LV Ave. Rise Top Oil Rise Hot Spot rise

A calculation using the harmonic current distribution gave;

Engineering analysis indicated the division of the eddy and stray loss to be; eddy loss 316watts stray loss 3 744 watts With the harmonic load currents the losses are; No Load I* eddy loss 6.28 x 316 = stray loss 1.62 x 3 744 = Total loss

4 072 watts 27 821 watts 1 984 watts 6 065 watts 44 014 watts

As indicated by equations in the IEEE Loading Guide [95], for the OA cooling mode, the top oil rise is proportional to the losses to the .8 exponent and may be estimated for the harmonic lossessfrom the test data at rated load and losses as shown below: =

014 %.2 [44M I

= 55.5

-c

The maximum per unit eddy loss occurred in the high voltage winding and was calculated to be 2.0 per cent of the ohmic loss. Assuming that the maximum eddy loss at the hot spot region to be

43

four times the average eddy loss would give an eddy loss of 8.0 per cent of the ohmic loss density at the hot spot location. The hot spot rise over top oil can be calculated as,

The hot spot rise over ambient = 55.5

+ 10.5 = 66.0 'C.

This is less than the permissible hot spot rise of 80 "C

XI. SUMMARY AND CONCLUSIONS Although sophisticated computer programs are available for electromagnetic analysis and numerous technical papers have been produced, most manufacturer's simple add a percentage to the ohmic (d.c.) losses for winding eddy and stray loss. This percentage is determined from empirical data. There is no test method to distinguish the winding eddy losses from the stray loss which occurs in structural parts. For hot spot temperature calculations many manufacturers simple add 15 ' C to the average winding rise for liquid units and 30 'C for dry type transformers. Although hot spot temperature is a performance parameter to be met by the manufacturer, there are currently no test methods or requirements that this parameter be measured on production or prototype transformers. Thermal failures have been reported on dry type transformers subjected to non-sinusoidal load currents even when the RMS current was below rated. Hottest spot temperature increments in excess of 30 'C may be a factor in these failures. For large dry type transformers the difference between hottest spot rise and average temperature rise may be 50 to 60 ' C instead of the U) ' C . The hot spot temperature in dry type transformers is more sensitive to the eddy loss magnitude and distribution within the winding. Model aging tests suggest a longer life for liquid immersed transformers than for dry type transformers. A current reliability survey of transformers in industrial plants is needed for comparison with data from 1979. Simple, accurate, practical methods are needed to determine the required transformer rating for new installations with nonsinusoidal load currents. The methods should consider the load diversity from multiple source harmonic load currents and load cycle variation during the day, week and year. For existing installations, with transformers not designed for harmonic currents, the IEEE C57.110-1986 procedures give conservative estimates of the load capability provided the transformer met industry standard hot spot temperature performance criteria. When required by the local inspection authority, the K-factor listed and labeled transformer at 300 kVA and below appears to be a viable option for commercial buildings with harmonic currents from sources such as personal computers. For industrial plants with dedicated loads such as drive systems requiring transformers above 300 kVA, the harmomc current content should be specified. For these transformers a more sophisticated analysis than the UL K-factor approach is required. Design of transformers for non-sinusoidal load currents should include an analysis of the eddy loss distribution in the windings and calculation of the hottest spot temperature rise. When purchasiig transformers subject to non-sinusoidal load currents, considerations should be given to the manufacturer's research programs and ability to calculate the eddy loss distribution and hot spot temperatures.

44

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BIOGRAPHY

1901 L. W. Pierce, "Predicting hottest spot temperatures in ventilated drv tvpe transformers," IEEE Tram on Power Delivery. Vol. 9 No. 2, pp. 1160-1172. April 1994. 1911 W. J. McNutt. "Notes from Bill McNutt". me Dobfe Exchange, Vol. 10, No. 3, pp. 9-10, Sept., 1992. 1921 AIEE Transformer Subcommittee, "Guides for operation of transformers, regulators, and reactors," Trans. AIEE, Vol. 64. QQ. 797-805; disc. pp. 957-958, Nov. 1945. 1931 L. W. Pierce, "An investigation of the thermal performance of an oil filled transformer winding," IEEE Trans. on Power Delivery, Vol. 7, No. 3. pp. 1347-1358. July 1992. I941 L. W. Pierce "Predicting liquid tilled transformer loading capability," IEEE Trans. on Industy Applicm'ons, Vol 30. No. 1, pp. 170-179. JanJFeb. 1994. I951 IEEE Guide forbading Mineral Oil Immersed Transfomn. PC57.91, Draft 11.3, Oct. 24. 1994 Annex G. [96] J. F. Troisi. "Experience with fiber optic hot-spot sensors in 450 MVA SMIT autotransformer," Proc. I992 Doble Conference. [971 Working Group 09 of Study Committee 12. "Direct measurement of hot-spot temperature of transformers". CIGRE Electra, No. 129, pp. 48-51. March 1990.

Linden W. Pierce (M'70) received the B. S. degree in Mechanical Engineering from The University of Texas, Austin in 1963. completed the General Electric Advanced Course in Engineering in 1966, and received the M. S. Degree in Mechanical Engineering from the University of Tennessee, Knoxville in 1973. He is a Registered Professional Engineer in the State of Georgia. In 1963 he joined the General Electric Company and since 1965 has worked for the Transformer Department at Rome, Georgia with various positions in transformer design. development, and program management. He is currently Senior Engineer, Product Technology. He holds eight patents and is the author of ten technical papers. His paper, "An Investigation of the Thermal Performance of an Oil Filled Transformer Winding", was selected by the IEEE Transformers Committee for its 1992 Prize Paper Award. Mr. Pierce is a member of the IEEE Industry Applications. Power Engineering, Magnetics, and Dielectrics &d Electrical Insulation Societies, and is also a member of CIGRE. He is a member of the IEEE Transformers Committee. Chairman of the Working Group on Guides for Loading Liquid Filled Transformers, Chairman of the Working Group on Development of the Loading Guide for Cast-Resin Transformers, and Chairman of the Insulation Life Subcommittee.

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