Bhel Journal

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BHEL JOURNAL

Volume 27 No. 2 September 2006

Editorial Advisory Committee

Ramji RaiK. RavikumarD. IndranS.K. Goyal

Editor : R.K. Bhattacharya

Associate Editor : D. Roy

BHEL JOURNAL is published quarterly.All correspondence and enquiries are to beaddressed to :

Mr. R.K. BhattacharyaEditor, BHEL Journal

Bharat Heavy Electricals LimitedBHEL House, Siri Fort,New Delhi-110 049

The statements and views expressed in thisJournal are entirely those of the authors, andnot necessarily that of the Organisation.

Contents may be referred to or reproducedpartially with due acknowledgements.

Copyright reserved.

CONTENTS Page

ADVANCES IN MATERIALS FORADVANCED STEAM CYCLE POWER PLANTS 1

SELF-EXCITATION IN 3-PHASE SQUIRRELCAGE INDUCTION GENERATOR FORWINDMILL APPLICATION 20

TURBO-GENERATOR INDUCED VOLTAGEWAVEFORM COMPUTATION ANDTELEPHONE HARMONIC CAPABILITYPREDICTION 26

EFFECT OF PRELOAD FACTOR ANDWORN DEPTH ON THE DYNAMICCOEFFICIENTS AND STABILITY OF A

LOADING ARC (WORN) TWO-LOBEBEARING USED IN TURBO-GENERATOR 35

COLLECTION, HANDLING ANDTREATMENT OF LIQUID EFFLUENTSIN THERMAL POWER PLANT 45

INNOVATIONS FROM BHEL 58

RECENT MAJOR ACHIEVEMENTS OF BHEL(during March'06-August'06) 62

1

2

4

3 Cover Photographs

1. 600MW Western Mountain Gas Turbine Power Plant,

Libya.

2. Sri sailam Hydro-electri c Plant (7x110 MW).

3. NALCO Captive Power Plant (960 MW).

4. Advanced Control Room at 2x500 MW (Stage-I I ) RihandSTPS.


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1BHEL JOURNAL,September 2006

ADVANCES IN MATERIALS FOR ADVANCED STEAM

CYCLE POWER PLANTS

Kulvir Singh

SYNOPSIS

The efficiency of conventional boiler or steam turbine

fossil fuel fired power plants is strongly based on steam

temperature and pressure. Since the energy cri sis of the

1970s, there have been efforts worldwide to increase

both : extensive research has been pursued worldwide.The need to reduce carbon dioxide emissions has provided

further impetus to improve efficiency. Development of

stronger high-temperature materi als is the prime

requi rement. EPRI and many other organizations have

extensively reviewed the materials technology for ultra

supercritical power plants. This article reviews the potential

benefi ts, operational experiences, the present trend and

the advances in materials that requi re special attention,

in respect of power plants wi th supercritical steam

conditions. This wi ll serve as a basis for defining material

issues for both the boilers and the turbines in next-

generation ultra supercri ti cal power plants.

Key Words:

Power Plants; Creep-Resistant Steels; Rotors; Casings;

Boiler; Superheater.

1. INTRODUCTION

An enhanced ecological awareness in the industrialisedcountries prompted increased initiatives world overto reduce CO

2 emission levels in the power plants.

This is essentially achieved by improving the efficiencyof the plants. Figure-1 shows some possibilities ofincreasing power plant efficiency [1]. In conventional

FIG. 1 : EFFICIENCY IMPROVING MEASURES FOR STEAM POWER PLANTS [1]


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FIG. 2 : DEVELOPMENT OF UNIT SIZES AND STEAM PARAMETERS IN JAPAN [1]

FIG. 3 : POWER PLANT EFFICIENCIES IN JAPAN [1]

power plants, a marked improvement of efficiencycan be achieved by advancing steam parameters. Theresulting developments of unit sizes and steamparameters are illustrated in Fig.2 [1]. The powerplant efficiencies achieved and planned for new

plants in Japan are shown in Fig.3. Steam conditionswere raised very rapidly during the 1950s and anumber of sets with supercritical steam conditionswere installed in 1950s and 60s. Subsequently, thetrend was reversed in respect of the steam conditions


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but the capacity of individual sets, however, continuedto increase till a limit with the existing fabricationand handling technologies has been reached. Thereversal in steam conditions is primarily the result ofexperience in 1960s and 70s when several newlycommissioned plants with advanced steam conditionsdid not live up to expectation in respect of availabilityof sets caused by operational problems.

Initially, the performance of the supercritical plantswas so poor that many utilities experiencedconsiderable downtime and significant financial loss[2]. Consequently, it created misconception thatimproved efficiency sacrifices reliability and therewas rapid retrenchment to subcritical units on the

assumption that they would be more reliable.Therefore, plants with operating temperature of538C received a wide favour in late 1960s and 70s.However, the concerted efforts of designers in liaisonwith metallurgists and material scientists inunderstanding of initial problems and takingcorrective steps led to a great deal of improvementin plant performance. Analysis of the historical

records and the stock of accomplishments ofsupercritical plants show that their reliability iscomparable to the conventional units [3-5]. The twooil crises in 1973 and 1978 which caused a drasticincrease in fuel cost and the encouraging operationalresults now available from earlier supercritical unitsprompted a renewed interest in supercriticalconditions to make the best use of the heat rateadvantage provided by these advancements [6-9].

1.1 Potential Benefits

Material development work over the past twodecades has paved the way for large thermal power

plants to be built today with live steam temperaturesof 610C, reheat temperatures of 625C andsupercritical steam pressures. The likely potential forreducing the heat rate by increasing the pressure andtemperature of the steam admitted to the turbine onthe basis of single and double reheating is shown inFig.4 [10]. At live steam conditions of 600C and300 bar with double reheating, for example, the heat

FIG. 4 : NET HEAT RATE IMPROVEMENT FOR SINGLE AND DOUBLE REHEAT CYCLES [10]


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rate can be reduced by 8% compared with the heatrate of today's standard power stations featuringsteam parameters of 540C/180 bar and singlereheat.

This improvement in thermal efficiency helpsconsiderably to conserve fuel resources and reduceCO

2 emission by 20%. This is a substantial

contribution on the part of power generatingindustries towards achieving the Germany's target oflowering CO

2 emission by 25-30% by 2005 [10].

This objective requires an ambitious developmentprogramme for advanced materials, which canwithstand such steam conditions. The researchprogramme has been undertaken simultaneously by

USA, Japan and European nations. It has focused ondeveloping further the existing high-temperature-resistant ferritic martensitic 12% CrMoV Steels forthe production of rotors, casings and chests, pipesand headers capable of operating at 593C, as wellas further development of the existing high-temperature austenitic steels suitable for inlet steamtemperatures up to 649C. For smaller highlystressed components such as first stage moving

blades and bolts, the objective was to employ andfurther develop existing high-temperature-resistantsuperalloys eg. Nim 80A and 90 etc.

Further development of ferritic steels geared to inletsteam temperatures up to approximately 625C in thecontext of COST 501 European research programmewas spurred by research activities in USA and Japan.Figure-5 gives an overview of the international researchprogrammes aimed at developing power plantmaterials[10]. Today, 65 partners from 13 countriesare involved in the European programme backed bya research budget of some DM31 million.

In the case of pulverized-coal-fired boilers, even a

marginal improvement in plant efficiency, say from34 to 37%, is reported to bring a savings of at least$5 million a year for each 1000 MW of capacity [2].However, in estimating the actual gains, the plant netheat rate gain should be weighed against the increasein total plant cost. Increase in steam parametersrequires more expensive materials of construction asthese advancements increase the severity of the serviceconditions the components must undergo.

FIG. 5 : INTERNATIONAL RESEARCH PROGRAMMES FOR DEVELOPING ADVANCED STEAM CYCLE PLANTS [10]


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A study funded by Electric Power Research Institute(EPRI), USA, suggested 750 MW unit as optimalsize for the advanced steam conditions offering a10% improvement in thermal efficiency, comparedwith present commercial designs [2]. The technologypromises a drop in heat rate of as much as 865 Btu/kWh. Assuming a first year fuel cost of $1.38/865BTU/kWh, the study concluded that cost savings inoperation could go as high as $160 million (1978value) over the plant's life time. Capital costs, on theother hand, are estimated to be 3 to 5% more thanthose for conventional capacity ($800/kW versus$775/kW). Keeping in view the current high andescalating level of fuel and capital costs, the potentialfuel savings, increasing environmental consciousness,

the supercritical steam conditions with advancedmaterials offer significant benefits that require seriousattention in selecting future capacity additions.However, it is necessary to consider local conditionssuch as grid size, expected annual utilisation period,cycling duty requirements, fuel cost etc. to work outoptimal unit capacity and operational conditionsincluding the number of reheats.

1.2 Service ExperienceBased on the operating data from supercriticaldouble reheat units in the range from 600-825 MWplant sizes, Westinghouse and GEC reported [2] thatthey achieved average availability of 80%. This ishigher than the average availability of 600-825 MWunits as a whole and is comparable to that of smallunits. A VGB evaluation also shows that theoperational availability of their supercritical plants isapproximately the same as with sub-critical units [5].

Average forced outage rates for the period 1970-83for ABB turbine operating with supercritical mainsteam pressures in several countries in Europe wasonly about 1% [8]. A 100 MW steam turbine hasbeen operating for several ten thousand hours inUSSR, as a test unit under steam conditions of 29.4MPa and 630-650C and reheat steam conditions of3 MPa and 565C. This unit was designed withcooling of many elements of HP housing and theoperating experience shows that it is highly reliable

[9].

Environmental conditions in several cases haveproved to be important in the availability of theplant. Operational errors in the water treatment ledto stress corrosion cracking of austenitic stainlesssteel tubes in some power stations [5]. Due to widespread deterioration of the quality of coals in severalcountries, a number of power stations experiencedfire side corrosion of tubes. In such places, newsubcritical power units installed in the same periodalso showed similar reduced availability [2].Supercritical units under those conditions requireda quite uneconomic purification of basic fuels [13].In case of oil-fired or dual-fired boilers, austeniticstainless steel superheater tubes suffered from excessivecorrosion problem.

1.3 Present Trend

Encouraging operational history of the earliersupercritical units, availability of more versatilematerials at a reasonable cost, progress in designtools such as computer programmes, experiencegained in the designing and manufacture of largesteam turbines and the continuously increasing trendof fuel cost, all together prompted several leadingpower plant suppliers, in the recent past, to revivetheir interest in units with advanced steam conditions.To avoid any technical risk, the developmentprogrammes have been planned in a phased manner.Two such programmes were launched independentlyon similar lines by the EPRI [9] and ToshibaCorporation, Japan [7].

Since most of the conventional materials readilyavailable are expected to meet at least for the first

two phases, the developmental programmes mainlyconcentrated in rationalizing the designs throughcomputer-aided programmed and in perfecting themanufacturing technologies especially in the case oflarge-sized components. Having completed suchexercises for the units of 700 MW capacity withsteam conditions of 566/566/566C and 32.2 MPa,both EPRI and Toshiba are expected to take up theirproduction, while the developmental activities tomeet the remaining two phases will continue to

progress. Mitsubishi Heavy Industries Ltd., Japan,


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on the other hand, since they have already suppliedfive out of the ten 450 MW steam turbines now inoperation in Japan with steam temperatures of566C, is already making proof tests of material andturbine elements, using an actual power plant, forsteam temperatures of 590-650C. A 50 MWprototype unit manufactured by Mitsubishi withsteam temperature of 590-650C was put intooperation in the year 1987 [6]. At present, many ofthe units up to the sizes of 1000MW are operatingin Europe, USA and Japan with the steam parametersof 600/610/610C temperature and 300 bar pressure.

2. REQUIREMENTS OF MATERIALS

FOR HIGH-TEMPERATURE APPLI-CATIONS

When a unit has to be designed and built to a highintegrity, one of starting points is concerned with thechoice of materials. As shown by the experiencegained, this assumes additional emphasis with thesupercritical sets as the equipment reliability andavailability otherwise may nullify the performancebenefits expected to accrue from the advancements

in technology. Generally speaking, the proper materialfor use at elevated temperatures is the one that bestmeets the following requirements at the lowest cost:

1. Adequate strength to resist deformation andrupture when exposed at the design conditionsfor the designed life, to the operatingenvironment.

2. Adequate fatigue strength at the designconditions and damping capacity when

vibratory stresses are involved.3. Suffi cient ducti li ty to accommodate

cumulative plastic strain and notch strengthto resist stress concentrations during theservice life.

4. Good resistance to service environment towithstand oxidation, corrosion and erosion.

5. Structural stability to resist damagingmetallurgical changes at operating conditions.

6. Abili ty to be fabricated with ease, as bymachining, forging, casting and welding.

7. Low coefficient of thermal expansion toresist the thermal stresses imposed by

differential temperatures and thermal cyclingor shocks during heat treatment, weldingand operation.

8. Good thermal conductivity for efficient heattransfer and to minimise thermal gradientalong the wall thickness of the thick walledcomponents so as to reduce thermal stressesduring start-up or quenching due to carryover.

9. Low density to provide a high strength-to-weight ratio for applications such as the laststage blading of the large capacity steamturbines.

10. Availability in the desired size and shape.

11. Enough long-term test data to allow sufficienthigh-temperature analysis to validate thedesign to the satisfaction of the safetyregulations and licensing authorities.

The power plants have depended mainly on the low-alloy steels for metal temperatures up to about580C. At temperatures above this level, theirresistance to creep is such that the resulting wallthickness becomes uneconomical. Also, theiroxidation resistance at higher temperatures is notsufficient. With increased metal temperatures above580C, austenitic stainless steels were employed formost of the power plant components designed inearly 1950s. However, austenitic stainless steels,though they are superior in high-temperature strength,have problems such as steam oxidation, and high-temperature corrosion in oil-fired boilers. Further,they are very expensive, susceptible to stress corrosioncracking and give rise to weld problems. Ferriticsteels, on the other hand, offer several technologicaladvantages such as better workability, high thermalconductivity and lower coefficient of thermalexpansion as compared to austenitic stainless steels.As a result, to bridge the gap between the low-alloy


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ferritic steels and austenitic stainless steels, severalhigh-alloy ferritic steels were developed during thelast three to four decades. These include 5CrMo,6CrMoVWTi, 7CrMoTi, 8CrMoTi, 9CrMo,9CrMoVNb, 9CrMoWVNb, 12CrMoV and12CrMoWVNb steels. Among these, 9Cr and 12Crclass of steels were extensively studied and successfullyemployed in several power stations. In fact, there areover two hundred grades of 12Cr steels withdifferent trade names cited in the literature, of whicha number of steels are generally used for differentapplications in gas turbines, and a few in thermal aswell as nuclear power plants [14,15]. A brief list ofmaterials is given in Tables-1 and 2.

With the operating temperature around and above

600C, austenitic stainless steels are required to beused for high-temperature strength combined withresistance to environmental attack. Simple austeniticsteels of type AISI 304, 316 and 347 have beenextensively used for power plant components. Butseveral complex austenitics of the type Essehete 1250,Alloy 800 H, 17-14CuMo, A286 and NF709 havebeen developed to give improved service performancearound 650C. Based on the actual service conditionsof a component, several higher alloys including nickelbase alloys are also being considered as candidatematerials for meeting exacting requirement to improvereliability. In view of the above, the candidate materialsfor advanced steam cycles are suggested, and theapproaches to meet the higher steam cycles are

discussed, in the following sections.

TABLE-1 : CANDIDATE MATERIALS FOR BOILER TUBES, PIPES AND HEADERS

Sl. No. MATERIALS FOR BOILER TUBES AND PIPES

Sub Critical Super Critical

1. C-Mn Steel HCM2S (T23)

2. Mo (T1) 7 CrMoV TiB 10 10 (T24)3. 1CrMoSi (T11) X20 CrMoV 12 1

4. 2Cr1Mo (T22) X10CrMoVNb 91 (T 91)

5. X20 CrMoV 12 1 X10CrMoWVNb 911 (E911)

6. X10CrMoVNb 91 (T 91) X10CrMoWVNb 92 (T 92-NF616)

7. AISI 304 X10CrMoVNb 12 1 (T 122)

8. AISI 310 X10CrNiMoTiB 15 15

9. AISI 316 X8CrNiMoVNb 16 13

10. AISI 321 X3CrNiMoNb 16 16

11. AISI 347 NF 709

12. E1250 Alloy 800

13. 17-14CuMo HR 3C

14. X10CrNiMoTiB 15 15 HR6W

15. X8CrNiMoVnb 16 13 AC66


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3. BOILER MATERIALS

3.1 Superheater and Reheater Tubing

Superheater and reheater tubes operate in the creeprange since their main function is to provide heattransfer between the hot flue gas and the Pressurisedsteam carried within them. They are, therefore,designed primarily based on the maximum allowable

stress to rupture in 100,000 hours, as specified bythe mandatory codes or standards. The other propertybases for their selection are a combination ofadequate corrosion resistance to both steam and theflue gas and ease of fabrication, particularly in regardto bending and welding. Reheaters receive partiallyexpanded steam from the turbine and serve to raiseits operating temperature to the required inlet levelof the Intermediate pressure turbine. Consequently,they operate at lower pressure and are made of larger

diameter but thinner walled tubes, as compared to

superheaters. Table-1 lists candidate materials forsuperheater and reheater tubing for power plantapplications. Figures-6 and 7 show their maximum

TABLE-2 : MATERIALS USED FOR THE ADVANCED STEAM TURBINES AT HIGH TEMPERATURES [47]

Component 566 C 620 C 700 C 760 C

Casings/shells Cr MoV (cast) 9-10%Cr (W) CF8C+ CCA617

(Valves; steam 10Cr MoVNb 12CrW (Co) CCA617 Inconel 740chests; nozzle Inconel 625box; cylinders) IN 718

Nimonic 263

Bolting 422 9-12%CrMo V Nimonic105 U7009-12% CrMo V A286 Nimonic115 U710Nimonic80A IN 718 Waspaloy U720

IN 718 Nimonic 105Nimonic 115

Rotors/Discs 1 Cr MoV 9-12% CrWCo CCA617 CCA61712 CrMoVNbN 12CrMoWVNbN CCA617 Inconel 74026NiCrMoV11 5 Haynes 230

Inconel 740

Vanes/Blades 422 9-12% CrWCo Wrought Ni Base Wrought Ni Base10 CrMoVNbN

Piping P22, P91 P91, P92 CCA617 Inconel 740

FIG. 6 : MAXIMUM ALLOWABLE STRESSES FOR VARIOU S

BOILER GRADE STEELS [20]


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allowable stresses and 100,000h stress rupturestrength, respectively [14-20].

Carbon steels are suitable and economical up toabout 400 to 450C metal temperature. Low-alloysteels Mo (SA209 T1), 1...Cr MoSi (SA213T11) and 2...Cr1Mo (SA213 T22) are used widelyfor metal temperatures up to 480, 550 and 580Crespectively [21]. T22 steel which has been extensivelyused for the final superheater for conventional unitsoperating at a main steam temperature of 540C hastoo low a creep strength to be accepted as a final

stage superheater material for use with a steamtemperature of 565C. Though these steels can stillbe used for the tube banks operating up to theirexisting allowable temperatures in the boilers ofsupercritical units, with the increase in steam pressuretheir required wall thickness increases. There is astrong incentive in using improved carbon steels nowavailable with higher allowable yield strength atlower temperatures up to 450C and in bringingdown the temperature range of application of thelow-alloy steels such that strong high-alloy ferritic

steels can be used to some extent even at steamtemperatures slightly below 540C. This solution

permits thickness of tubes less than those needed forthe common low-alloy steels for the same operatingconditions as shown in Fig. 8. This results in savingof base material and welding filler material, reducesthermal stresses and welding problems and alsoimproves the heat transfer efficiency.

FIG. 7 : 1,00,000H STRESS RUPTURE STRENGTH FOR BOILER MATERIALS [10]

FIG. 8 : COMPARISON OF T HE SIZE OF THE WALL THICKNESS

OF P22, X20, P91 AND NF616 STEEL PIPES

10CrMo9.10

NF616

X20 T91/P91


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As the boiler steam advances, the metal temperatureof the last banks of the superheater tubes of the565C unit exceeds the maximum allowable limit ofthe low-alloy steels. In these regions, high-alloyferritic steels are required to be used. 9Cr1MoVNb(SA213T91) steel has improved oxidation resistance.The allowable stresses for 9Cr1MoVNb (SA213T91)are about 100% higher than T22 steel at temperaturesabove 540C and are even higher than X20 CrMoV12 1 (DIN17175 - 1.4922) steel in the range of 500to 650C. The other two high-alloy ferritic steels9Cr1MoWVNb (SA213 T92) andX12CrMoWVNbN 10 11 (E911) developed in USAand Japan are much superior in their stress rupturestrength as compared to all other ferritic steels

including the 12Cr steel (X20 CrMoV 12 1) uptoabout 650C. Their allowable stresses are also higherthan TP304H and TP321H at temperatures below600C. 9Cr1MoVNb steel, due to its low carboncontent exhibits good weldability and workabilityand has been giving satisfactory service as superheatersand reheaters, for over two decades in Japan [22].X20 CrMoV 12 1 is also supplied with 0.4 to 0.6%tungsten, which is then designated as X20 CrMoWV12 1 (DIN 17 175-1.4935) [15]. Addition of

tungsten is reported to give greater creep strengththan the steel without tungsten but, based on somelong-time investigations no effect of tungsten hasbeen found for additions up to 1% [14]. It is,however, reported to be beneficial for thick sections.Both X20 CrMoV 12 1 and X20 CrMoWV 12 1were developed and extensively used in Europe forsuperheater and reheater tubes. These steels beingmartensitic grade have a strong self-hardeningproperty. Due to the formation of martensite, the

weld metal and the heat affected zone (HAZ)become very brittle on cooling. For better results,both preheat and post weld heat treatments aremandatory. 9Cr1MoVNb steel has been extensivelystudied for over two decades in USA and its tubesamples are currently in service in the USA & UK[18]. Of the two V and Nb bearing 9Cr steels, thisoffers better rupture strength. Among the varioushigh-alloy ferritic steels, X20 CrMoV 12 1,9Cr1MoVNb (T/P91) and 9Cr1MoWVNb (T/P92)are the three most promising candidates for tubing.

P92 has an edge over the other two due to its highrupture strength.

Austenitic steels would be required for finalsuperheaters of 593C units and for most of the

superheater and reheater of the 650C units. Thedesign stress values derived by various boiler codesdiffer based on the approaches adopted by them.Similarly, ranking of austenitic steels 304, 316, 347,based on their allowable stresses varies depending onthe code. However, 304, 316 and 347 type of tubesare widely used for operation at higher steamtemperatures. TP347H type of tubes are extensivelyused in USA, Japan and Germany [5, 22] due to itssuperior properties such as resistance to fire-side

corrosion, steam-side oxidation and higher thermalfatigue strength, as compared to 316 type of steel.Also, ASME Boilers code allows higher design stressfor TP 347H as compared to TP316H type [15],whereas in the case of BSI, the reverse is true [26].Type 347 and 321 steels are prone to strain-inducedembrittlement due to the formation of strongcarbides like NbC and TiC [26]. As a result, AISI316 steel is preferred in U.K. For superheatersoperating at the highest temperatures of the high-pressure units, stronger austenitic steels like Incoloy800H, 17-14 CuMo, Esshete 1250, NF616 and 15-15N, X8CrNiMoNb 16 13 and X3CrNiMoN 17 13would be required for reliable operation. For mostexacting conditions, materials such as Inconel 617which contains 12.5% Co, though very costly, mayalso have to be considered.

3.2 Steam Piping and Headers

Ferritic steels are preferred because of their higherthermal conductivities and lower coefficient of thermalexpansion coupled with their good fabricability.High-strength high-alloy ferritic steels such asX20CrMoV 12 1, X20CrMoWV 12 1, 9Cr1MoVNband 9Cr1MoWVNb are, therefore, employed fortemperatures up to 625C. These steels are, however,subject to temper embrittlement in the temperaturerange of 540 to 595C, where advanced supercriticalsteam power plants operate. Though the temperembrittlement behaviour is not likely to have any


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effect on their use for piping, it can be controlledby maintaining low levels of manganese and siliconcontents [9]. Also low levels of sulphur should bemaintained, as the toughness of these steels is verysensitive to sulphur content.

At 650C steam condition, it is necessary to usehigher-strength austenitic steels. Niobium stabilizedaustenitic steels of the type 347 are preferred inGermany because of their high fatigue strength andsuperior steam-side oxidation as compared to niobium-free steels [27]. Due to temperature-constrainedexpansion during and after welding, niobium steelsshould be used in the case of wall thickness up to30mm, whereas niobium-free austenitic steels

X8CrNiMoNb 16 13 and X3 CrNiMoN 17 13 aredesirable for thick walled pipes. In case of 316stainless steel, main steam pipe failure due to sigmaphase formation was reported [9]. Extensive databaseup to about 100,000 hours available on Esshete 1250,NF709, HR6W steels confirm their reliability forsteam pipe and header application [19, 46].

3.2.1 FIRE-SIDE CORROSION

In coal-fired boilers, the ash corrosion results fromthe formation of complex alkali iron sulphates in ashdeposits, which become aggressive in molten state.The severity of liquid ash corrosion varies withtemperature and follows a bell-shaped curve [29].Corrosion increases sharply from about 595C to700C. Below 595C, the corrodents in the ashdeposits will be in a fairly dry state and therefore donot aggressively attack the tubes. As the temperatureis increased up to about 700C, the corrodentsbecome molten and the corrosion rate increases.

With further increase in temperature, the moltedsulphates begin to vaporize and become unstable,decreasing the corrosion rate.

Superheaters and reheaters of the supercritical plantsat 566C operate close to the beginning of the bell-shaped curve. 9Cr and 12Cr tubes should be goodenough to serve without experiencing any significantfireside corrosion. But the final superheaters andreheaters of the plants at 593C and 649C wouldoperate at the apex of the curve, where corrosion is

most severe. Alkali sulphate corrosion rate decreaseswith the increase in chromium content and asuperior resistance can be obtained with chromiumcontents of over 25% (Fig.9) [30]. Simulated liquidash corrosion tests carried out on different superheatertubing alloys show that their resistance vary widelyas shown in Fig.10 [31]. Amongst the alloys tested,

FIG. 9 : COMPARISON OF FIRE-SIDE CORROSIONRESISTANCE OF VARIOUS ALLOYS [31]

FIG. 10 : HOT-CORROSION WEIGHT-LOSS wrt Cr CONTENT

FOR VARIOUS ALLOYS [30]


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the 17-14CuMo austenitic steel used for superheatertubing in Eddystone unit [9] experienced the highestcorrosion rate, while the Inconel 671 (50Cr-50Ni)alloy was practically immune to liquid ash corrosion.It can also be seen that the 17-14 Cu Mo alloy, onchromizing, exhibits liquid ash corrosion resistancealmost similar to Inconel 671. There are variouspotential means of overcoming the coal ash corrosionproblem of superheater tubing such as bandageshields of more corrosion-resistant materials, surfacecoatings, grain refinement and composite tubes.Bandage shielding decreases heat transfer efficiencybecause of the insulating air gap, and it may alsoresult in significant increase in corrosion rate of non-bandaged tubing [32]. Surface coatings such as

chromizing [33], chromating [22], chromium plating[33] and calorizing [32] have been attempted. Thesemethods require further detailed study for theireffect on fabrication, ductility and high-temperaturecreep strength. However, studies carried out on achromized austenitic steel showed encouraging results[33]. Since both outside surfaces can be chromizedat the same time, it appears to be a promisingapproach to prevent corrosion of both the surfaces.Grain refinement promotes the grain boundary

diffusion of chromium to the surface [22]. Thisresults in improved corrosion resistance [20] but ithas an adverse effect on the rupture strength of thematerial. For most advanced supercritical steamconditions and under highly aggressive serviceconditions generated for combustion of coal withhigh chlorine contents where a single materialcannot provide an economically viable solution,technical advantage provided by two differentmaterials can be utilised by employing composite

tubes. Composite tubes are produced by co-extrusionof two different materials comprising a high-strengthinner material such as Essehete 1250, alloy 800H,17-14 Cu Mo to provide the stress-bearing capacityand an outer casing of material of high chromiumcontent like TP310 or Inconel 671 for protectionagainst corrosion. The problems expected from suchtubing are thermal fatigue, welding and sigma phaseformation in outer casing [28]. The serviceperformance for several years in a number of powerstations, however, confirms the integrity and

economics of using composite tubes [19]. Fordemanding environmental applications, it maysometimes be necessary to select even nickel orcobalt base alloys, provided economics permit.

3.2.2 STEAM-SIDE OXIDATION

General experience indicates that the oxidationresistance of high-temperature steels in drysuperheated steam is almost the same as theiroxidation resistance to air at the correspondingtemperature [26]. The oxide scales formed on theinternal surface of superheater tubes, reheater tubes,headers and piping spall off, or exfoliate, when thethermal stresses due to differential thermal expansion

between the oxide scale and the base metal exceedthe bond strength of oxide scale. 9Cr and 12Crsteels, due to their low coefficient of thermalexpansion, offer better resistance to oxide exfoliationas compared to austenitic steels, in the temperaturerange of their application. The spalled oxides can besevere enough to clog the bends in superheater andreheater tubing, eventually leading to overheat failures.At high steam pressures, these oxides from headersand steam pipes can be carried to the turbine at high

velocities and cause turbine erosion. Turbine erosiondamage not only causes a loss in efficiency but alsois expensive to repair and increases the time ofturbine overhaul outages [34]. It was reported thatthe erosion damage had led to the destruction of aturbine. Since the rate of steam oxidation variesexponentially with temperature [35], exfoliation canbe much more of a threat to advanced supercriticalsteam plants. Amongst the conventional austeniticstainless steels, TP347H provides better resistance toboth fire-side corrosion and steam oxidation,presumably due to its high chromium content.Steam oxidation, as in the case of fire-side corrosion,is a function of chromium content in the innersurface of the tube. To overcome steam oxidationproblem, several methods such as chromizing,chromating, grain size refinement and cold workingof the inner surface through shot peening/blasting,all of which increase the chromium content at thesurface, are applicable. Chromizing being a veryhigh-temperature process, is applicable to only new


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and replacement components. Chromating, on theother hand, can be performed at lower temperaturesand even on assembled parts. Both grain refinementand surface shot peening methods aim at enhancingthe chromium diffusion rate to the surface, but theireffect is lost during the process of welding or whenstress relief annealing has to be carried out afterbending. Combination of chromizing and chromatingtechniques seems to be the best choice to overcomethe environmental problems.

4. TURBINE MATERIALS

In the case of turbine, the advancement in steam

conditions mainly affects its high pressure (HP) andintermediate pressure (IP) sections. As a result, theassociated rotations as well as stationary parts ofthese sections experience more severe serviceconditions than that of conventional sets. Since theyoperate well within the creep range, their design isbased primarily on the long-term creep strength ofthe material, but the stress levels during steady andnon-steady operating conditions, particularly during

start-up and shut-down periods, must also be takeninto account. Figure-11 shows the temperature rangefor application of different grades of steels [4]. It isclear that the low-alloy ferritic steels are limited touse at temperatures up to about 550C, and theirrange of application further decreases for rotatingcomponents. Table-2 gives a list of candidate materialsof interest, and Fig. 11 depicts their 100,000hrupture strength as a function of temperature [4, 5,36, 37, 38]. For the sake of comparison andcompleteness, some of the low-alloy ferritic steels arealso included.

4.1 HP/IP Rotors

As can be seen from Fig.11, 12Cr steel can beemployed for HP/IP turbine rotors at 566C steamtemperatures. X22 CrMoV 121 has been successfullyused for rotors of the supercritical units for manyyears. With rotor cooling, it can also be used up to595C. Both EPRI [9] and Toshiba [9] have chosen12Cr steel for HP and IP rotors of turbines foroperation at 566C. Their advanced designs at

FIG. 11 : 1,00,000H CREEP STRENGTH FOR STEAM TURBINE APPLICATIONS [10]


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593C contemplate the use of 12Cr steel rotor withsteam cooling to bring the rotor temperature downto 566C, where its creep strength is adequate tomeet the design pressure. However, presently severalsuper 12Cr steels with much superior creep resistanceare available and they should also be consideredbefore a final decision is taken. Above 593C steamtemperature, X12CrMoWVNbN 10 11 and austeniticstainless steel must be considered. Amongst theaustenitic steels, A286 and X8CrNiMoBNb 16 16offer better creep strength for an HP rotor ofadvanced sets operating at 649C.

One of the rotor-related problems is the maximumsize that can be produced from the 12Cr and

austenitic steels. Due to severe segregation inconventional ingots, the size of the austenitic steelrotors used in earlier supercritical units was limitedto small size, as a result of which, it becamenecessary to divide the HP turbine into two stages[9]. It is estimated that a large advanced plant wouldrequire a one-piece super-alloy HP rotor forgingweighing 11,300 kg with a barrel diameter of890mm. Similarly, a double-flow reheat rotor madeof 12Cr steel is expected to be about 1150mm in

diameter and 31,750 kg in weight, which wouldrequire to start with an ingot size of 63,500 kg [9].Significant progress has been made, in recent years,in increasing the size as well as the quality of theforging by employing modern steel making techniquessuch as low sulfur silicon deoxidation (low S),vacuum oxygen decarburization (VOD), vacuumcarbon deoxidation (VCD), central zone refining(CZR), electro slag hot topping (ESHT) and electroslag remelting (ESR). By employing these techniques,either individually or in combination, productionexperience with low-alloy ferritic [39], 12Cr as wellas austenitic steels [4, 40] demonstrate that therotors of the candidates materials can be made to therequired size and quality without experiencing muchproblems.

4.2 Blading

Conventional 12CrMoV steel blades are adequate to

meet the steam temperature at 566C. But a wide

variety of high-temperature blade materials withproven service performance in large gas turbines areavailable, and they should be considered for moreadvanced steam conditions. These include super12Cr steels, austenitic steels, Nimonic 80A, 90, 105,115, In718 and precision casting alloys such asUdimet 500 and IN 738LC.

4.3 LP Rotor

The principal requirements of material for low-pressure (LP) rotor are high yield strength towithstand the high stress imposed on it by longblades and high fracture toughness to minimize sub-

critical flaw growth so as to avoid the possibility offast fracture. 3.5NiCrMoV steel is widely used forLP rotor throughout the power industry. To avoidtemper embrittlement, the maximum operatingtemperature of the LP rotor made of this steel isgenerally limited to about 350C [9]. The inletsteam temperature to LP turbine of the supercriticalunits, on the other hand, is dictated by the exhauststeam from the second IP section. The IP-LPcrossover temperature from advanced supercriticalunits at steam temperatures of 593C and abovewould be 400-455C [9]. To maintain the inletsteam temperature of LP turbine at its presentmaximum allowable limit, it would be necessary tocool the steam either through cooling of the rotoror by adding an additional stage of expansion to theIP turbine. The latter approach would be a difficultdesign task, as it requires usage of long blades at hightemperature, whereas the former approach has tosacrifice a part of the thermal efficiency.

Another approach to the problem would be torender LP rotor material more resistant to temperembrittlement [41]. Efforts are, therefore, beingmade to improve the fracture toughness of the IProtor steel by improved steel making technology andcloser control of chemical composition. Theinteraction between Mn, Si, P and Sn was shown tohave promoted the degree of temper embrittlement[41]. Resistance to temper embrittlement of3.5NiCrMoV rotor steel with low Mn and low Si

contents was found to have greatly improved as


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compared to conventional steel [9]. By utilising themodern steel making technologies, it is now possibleto decrease both Mn and Si contents to levels of0.001 - 0.002%.

4.4 Heavy Section Stationary Parts

High-pressure turbine requires a number of heavysection static components such as the inner and outercasings, static components such as the inner and outercasings, the steam valve, the nozzle box, and the inletpipe. Besides high temperature and pressure, theseparts are subject to thermal cycling. If the sectionsizes are very high, there is a danger of experiencing

thermal cracking as a result of the heavy thermalstresses that might develop during start-up or carry-over. It is, therefore, desirable to minimize the sectionsize by using high-strength steel so as to reducethermal stresses. Depending on the stresses and therequired wall thicknesses, it might be advisable toemploy 12Cr steel at temperatures lower than 566Cand austenitic steels at temperatures as low as 566C.This could be advantageous especially in the case ofhigh-pressure units designed for 31 MPa and above.Given the choice, forgings are preferred as they allowthinner sections but it would be economical to usecastings. Toshiba [7] will be using a 12Cr cast steel(10CrMoVNbN) for these parts of the units at566C, whereas EPRI [9] intends to give preferenceto forgings for the initial advanced units.

In order to minimize differential thermal expansion,it is desirable to make the rotor and the stationaryparts of the turbine of the same material. However,for 593C units, it is likely that the rotor would be

made of 12Cr steel, while the inner casing would bemade of austenitic steel. Under such circumstances,shaft seals must be used to accommodate the greaterthermal expansion of the casing. Larger clearancesare required to be given, when austenitic steels areemployed for rotating and stationary parts, for mostadvanced steam conditions. Both the designrequirements, 12Cr rotor cooling at 593C andlarger clearance to be provided with austenitic steelsat higher temperatures, adversely affect the cycle

efficiency. This, in turn, partly reduces the net heat

rate gain achieved by elevating the steam parameters.In addition, rotor cooling complicates the design ofthe turbine and its external piping, and calls for anoverall economic justification in final selection of thecandidate materials.

Since the outer casing is subjected to cooler and low-pressure steam as compared to the other casing, thiscould be still made of the conventional low-alloyferritic steel.

4.5 Transition Weld Joints

In cases where main steam piping and the outercasing are made of austenitic and low-alloy ferritic

steels, respectively, the inlet piping to the turbinewill have to utilize transition joints. An approachsuggested to this problem is to make the joint inthree sections, utilising a material of intermediatethermal expansion coefficient such as Alloy 800H inbetween the pipe and the casing with nickel-basedfiller metal for welding. Nickel-based filler metalimproves the rupture life as much as five times morethan the austenitic filler metals. Due care must alsobe taken in design to minimize bending stresses, as

life of transition joints is greatly reduced if bendingstresses are superimposed upon stresses from thermalexpansion.

4.6 Bolting

Bolts and studs are used in many joints of theturbine, which need to be separated for maintenanceor repairs, as for example, castings and valves. Thebolts and studs differ from all other turbine

components in that they are notched, subjected tocold as well as hot stressing and a varying patternto stressing due to practice of tightening andretightening. The strain to which bolts are tightenedis based on both the properties of the materials andthe design practice. The usual strain applied in UKis 0.15% [44], whilst it is 0.2% in Germany [36].The elastic strain produced by the initial tighteningof the bolts is progressively converted to creep strain,thereby reducing the effective load on the joint.

Bolts for turbine parts are required to possess


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sufficient resistance to stress relaxation to maintainflange toughness against internal steam pressure atleast for the period between overhauls and should bere-usable throughout the life of the plant. Inaddition, bolting materials must have coefficient ofthermal expansion close to that of flange material,high proof strength, good notch rupture ductilityand resistance to embrittlement.

The stress relaxation behaviour of bolting materialsthat have already demonstrated their successfulservice performance in power stations is shown inFig.12 [36, 45]. 1Cr1MoVTiB steel, which isextensively used in the conventional unitsmanufactured in the country, can still be used for

outer casing joints of the supercritical units, but itis desirable to use 12Cr steel for inner casing flangebolts. At steam temperature of 566C, the differencein the stress relaxation behaviour of these two steels,as shown in Fig.12, is as a result of the differencein their initial strain. But, at the same initial strainlevel of 0.15%, 12Cr bolting steels possess betterresistance to stress relaxation. In some of the earlierunits, 12Cr bolts were also used at highertemperatures through steam cooling [13]. Since the12Cr bolts undergo extensive stress relaxation at593C, consideration is being given by EPRI to anumber of high-strength nickel-based bolting

materials. For units at 649C steam temperature,nickel-based bolting materials are required to beused, as the austenitic steel bolts are not strongenough to meet the requirement. Based on a criticalsurvey of worldwide experience and stress relaxationtests, Incoloy X750, Nimonic 80A, Nimonic 90,Refractaloy 26 and PER 2B have been identified asthe best candidates for use up to 650C.

5. SUMMARY AND CONCLUSIONS

i) The quest for lower-cost power generationled to a rapid progress in the steam conditions,reheat cycles and output capacity of the

power plants during 1950s and 60s.Supercritical power plants offer considerablegain in heat rate.

ii) During the last two decades, ferri tic-martensitic 9 to 12% Cr steels have beendeveloped under international researchprogrammes, which permit (live) inlet steamtemperatures for thermal power stations upto approximately 610C, pressures of up toabout 300 bar and reheat temperatures up to

about 625C. The results have beenimprovements in efficiency of around 8%versus conventional steam parameters.

iii) The newly developed 9-12% Cr steels arealready being used in 12 European and 34Japanese power stations with inlet steamtemperatures of up to about 610C. Theexperience with the components made ofthese steels has been decidedly positive.

iv) Advancements in steam parameters increasethe severity of the service conditions thatmaterials must undergo. The presentlyavailable low and high alloy ferritic steelswith proven service experience can meet therequirements of both boiler and steam turbinecomponents of units designed for steamtemperature at 600C.

v) For advanced steam plants of temperaturesFIG. 12 : COMPARISON OF 30,000H RELAXED STRESS FOR

BOLTING MATERIALS [20]


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above 600C, austenitic steels are required tobe employed for final stages of super heatersand static components of the steam turbine,such as inner casing, steam valve and nozzleblock. 12Cr steel can be used for HP/IProtors subject to reliable design for coolingthem so as to bring the material temperaturedown to 566C.

vi) At steam temperatures of 649C/621C,high-strength martensitic steels are requiredfor HP/IP steam turbine rotors and stationarycomponents. Preventive measures such asapplication of composite or chromized tubesto withstand fireside corrosion of superheater

and reheaters as well as steps to minimizesteam-side oxidation should be taken.

vii) For steam turbine blading and bolting attemperatures above 593C, materials such asNimonics and Refractaloy 26 etc. should beconsidered.

viii) To meet the IP-LP crossover temperature forunits at steam temperatures of 593C and

above, it is necessary to improve the temperembrittlement of the existing 3.5NiCrMoVLP rotors steel, by employing modern steelmaking techniques to eliminate elementssuch as Mn, Si, P and Sn. Alternatively, thesteam should be cooled in the IP turbine tolimit the LP inlet steam temperature to theexisting maximum allowable level of about350C.

References

1. Husemann, R.U., et.al., 'Processing and

Practi cal Application of New Materi als in

Power Plant Constructions', VGB

Kraftwerkstechnik, 75(3), 1995, 241-255

2. EPRI Journal, September 1981, 22

3. Spencer, R.C., Proc. Amer. Power Conf. 42,

1980, 225

4. Haas, H., et al., Proc. Amer. Power Conf., 44,

1982, 330

5. Schneider, A., VGB Kraftwerkstechnilk, 58,

1978, 168

6. Kawai, T., Turbomachinery International, 25,

March 1984, 34

7. Akiba, M. and Aizawa, K., Turbomachinery

International, 25, March 1984, 37

8. Muhlhauser, H., Brown Boveri Review, 71,

1984, 120

9. Gold, M. and Jaffee, R.I., J. Materials for

Energy Systems, 6, 1984, 138

10. Mayer, K.H., et.al., 'New Materi als for

Improving the Efficiency of Fossil Fired Thermal

Power Stations', VGB Power Tech, Jan 1998,

22-27

11. Trojanowskij, B.M., VGB Kraftwerkstechnik,

60 (1980) 525

12. Plastow, B., et al., Int. Conf. on Creep and

Fati gue in Elevated Temperature Applicati ons,

Instn. Mech. Engrs. Sheffield, 1974, Paper

C49/ 74

13. Gemmil l, M .G., ASTM Jl. Testing and

Evaluation, 2, 1974, 3

14. Briggs, J.Z. and Parker, T.D., 'Super 12% Cr

Steels' Climax Molybdenum company, New

York, NY, 1965

15. Section I , Power Boilers, ASME Boiler and

Pressure Vessel Code, ASME, New York, 1983

SI Edition, 176-227

16. Characteri sti cs of HCM9M Steel Tubings for

Boiler application, M itsubishi Heavy Industr ies

Ltd. (MHI) and Sumitomo Metal Industries

Ltd. (SMI). July 1979.

17. Sumitomo, High Strength Boiler Tubes, SMI,

80-F-No.1081

18. Canonico, D.A., The factors that Influence the

Selection of H igh Temperature Materials,

presented at the National Symposium on Creep


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Resistant Steels for Power Plants, BHEL R&D,

Hyderabad (India), January 1983

19. Orr, J. and Nileswar, V.B., Stainless Steels-84,

The Insti tute of Metals, London 1985, 533

20. Viswanathan, R., 'Damage Mechanisms and

Li fe Assessment of H igh Temperature

Components,' ASM Intl., 1989

21. Ranganathan, S., et al., National Symposium

on Creep Resistant Steels for Power Plants,

BHEL R&D, Hyderabad (India), January

1983, Paper No. 1.03

22. Inoue, M ., et al., The Sumitomo Search no. 29

Nov 1984, 64

23. Fricker, H. and Walser, B., Ferri ti c Steels for

Fast Reactor Steam Generators, BNES, London,

1978,35

24. Properti es of 9Cr steel tubes and pipes, SIM,

804-f-No. 1194, February 1984

25. Caubo, M., Improved ferr iti c steels for super

heater tubing, ASME paper No. 63-WA-246,

1963

26. Gemmil l, M .G., The technology and properties

and ferrous alloys for high temperature use,George Newnes Ltd., London, 1966

27. Wyatt, L.M ., Materi als of construction for

steam power plants, appl ied science publisher

Ltd., London, 1976

28. Wyatt, L.M., Mat. Sci. and Engg., 1, 1971,

273

29. Koopman, J.G., et al., Proc. Ameri can Power

Conf., 21, 1959, 236

30. Sumitomo, High alloy composite tubes for

pulveri zed coal fired boiler application, SMI,

803-F-No. 1006

31. Ohtomo, A., et al., 'H igh temperature corrosion

characteristicsof superheater tubes', IHI Engg.

Rev., 16(4), October 1983

32. Flatly, T., and Lathom, E., Materi als in power

plants, spring resident ial course Instn. of

Metallurgists, Chamelon press, London, 1975,

63

33. Sumi tomo, Chromized stainless steel tubes,

SMI 803-F-No. 1079, Jun 1983

34. Haberman, J.A., and Keyton, H., Proc. Amer.

Power conf., 44, 1982, 1970

35. Rehn, I.M., et al., NACE corrosion 80, Chicago,

IL, March 1980, Paper No. 192

36. Warmfeste Hochwarmfeste Werkstoffe Fur

Schrauben and Mattern, Gutevorschri ften, DIN

17240, July 1976

37. Wegst, G.W., Stahlschlussel, Verlag Stahlschlussel

Wegst GmbH , 1983

38. Warmfester Ferritischer Stahlgu?, Tecchnische

Li eferbedingungen, DIN17245, October, 1977

39. Swaminathan, V.P. and Jaffee, R.I., M etal

Progress, 128, December 1985, 52

40. Manufacturing of Trial A 286 Rotor Forging,

Kobe Steel, TKE 83-57, January 1984

41. Todeu, H., et al., M itsubishi Power Systems

Bulletin MBB-82113E, November 1982

42. Watanabe, J. and Murakami, Y., Proc. Amer.

Petroleum Inst., 1981, 216

43. Viswanathan, R., et al., 'Dissimilar Metal

Welds in Power Plants', Presented at AWS-

EPRI Conf. on Joining Dissimilar Metals,

Pittsburgh, PA, August 1982

44. Branch, G.D., et al., Int. Conf. on Creep and

fracture in Elevated Temperature Applicati ons,

Sheffield, 1974, Paper C192/73

45. Bri ti sh Steel Corporati on's Data Sheets on

Durehete 900 and Durehete 1055 Steels

46. Oakey, J.E., Pinder, L.W., Vanstone, R.,

Henderson, M . and Osgerby S, 'Review of

status of advanced materials for power

generati on', Report no. COAL R224, DTI/

URN 02/1509, 2003

47. Wright, I .G., Maziasz, P.J., Ellis, F.V., Gibbons,

T.B. and Woodford, D.A., 'Materials issues for

turbines for operati on i n ultra supercri ti cal

steam', ORNL report, USA.


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Mr. Kulvir Singh graduated in MetallurgicalEngineering from University of Roorkee, (nowIIT, Roorkee) in the year 1981. He completedM.Tech. (Metallurgy) from IIT, Kanpur, in1983.

Thereafter, Mr. Singh joined MetallurgyDepart ment of Corporate R& D, BH EL,Hyderabad. Since the beginning, he was involvedin indigenization of creep-resistant steels for steamturbine and boiler applications. Subsequently, healso studied structure property correlation and

creep crack growth behaviour of power plantsteels. He has also carried out extensive studies onthe creep-rupture behaviour of P91 and X20

steels, their weldments and simulated heat-affectedzones. He is actively involved in residual lifeassessment of steam and gas turbine components.He is also working in the area of indigenousdevelopment of gas turbine buckets and heattreatment of steels by microwaves. His otheractivities include many important failureinvestigations of boiler, steam and gas turbinecomponents and process industry equipment. Heis currently working as Dy. General Manager andheading Creep Lab of the Metallurgy Department.

Mr. Singh has published/presented over 50technical papers in various national andinternational journals and conferences. He hasalso received BHEL Excel Award in the bestTechnical Paper category for the year 2003.


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SELF-EXCITATION IN 3-PHASE SQUIRREL CAGE

INDUCTION GENERATOR FOR WINDMILL APPLICATION

P.K. Khanna

SYNOPSIS

In India, there are certain areas where plenty of energy

is available but it remains untapped. Wind energy is

one such area. Also, i t i s well known that Induction

Generators are best sui ted for power generation through

windmi ll due to their simple construction. In this paper,

an attempt has been made to describe how an Induction

Generator can be made to get self-excitation and thus

be used in any stand-alone situation, especially for

windmill application.

Key Words:

Induction Generator; Self-Excitation; Stand-Alone.

1. INTRODUCTION

With the advancement of technology, a need isalways felt to have as much simplification as ispossible but with high reliability of operation.Induction Generator is one such category of machine,which is most simple in its construction, as well asin operation. As the name implies, the InductionGenerator has a squirrel cage rotor. This obviatesbrush gear assembly, brushless excitation system orpermanent magnet etc. as is necessary for other kind

of generators. For excitation of the inductiongenerator, a capacitor bank is used across the statorterminals. The main advantage of such generators issimple construction, low cost and high reliability.

2. T YPES O F SQUIRREL CAGEINDUCTION GENERATOR

These are of two types:

(i) In the first type, when a squirrel cage motor

is run above its synchronous speed, it starts

functioning as a generator. In this type, it isalways necessary that the machine runs aboveits synchronous speed and it should remainconnected to power supply for excitation.

(ii) In the second type, when a squirrel cagemotor running near to its synchronous speedis switched off and simultaneously a capacitorbank is connected across the motor terminals,it starts functioning as a generator. This typeis characterised by self-excitation.

In this article, the second type of generator only(self-excitation type) has been discussed in detail.

3. PRINCIPLE OF WORKING OF THE

SELF-EXCITATION TYPE

When an Induction motor is running under steadystate condition, an e.m.f. (E) exists across its statorwinding. Now, if the speed of motor is maintainedthrough a prime mover and supply to inductionmotor is switched off, simultaneously connecting itacross a capacitor bank, there is a flow of excitationcurrent in the stator winding as per the load line ofcapacitor (Fig. 1). This current produces rotatingflux, which, in turn, induces e.m.f. in the statorwinding. Under steady state conditions, the e.m.f.induced in the winding has a magnitude anddirection the same as that of the original e.m.f.(E).Thus, the voltage E continues to be sustained.

4. PH ENOMENON OF SELF-

EXCITATION

The most important condition for self-excitation to

take place is that there should be presence of residual


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flux in the rotor. When a capacitor bank is connectedacross the stator winding after switching off itssupply, capacitive current flows in the stator winding.Then the flux produced by this current aids theresidual flux. The increase in flux increases theinduced e.m.f. in the winding and this cycle continuestill there is saturation. Under this condition, thesteady state e.m.f. is given by the intersection of themagnetizing curve of motor with the capacitive loadline. If the residual flux in the rotor is not sufficient,the self-excitation fails to occur and therefore, the

voltage at the terminals does not build up.

The slope of the motor magnetizing curve is calledthe critical slope. The size of the capacitors in thecapacitor bank is chosen in such a way that itscapacitive reactance is less than the critical slope ofthe magnetizing curve, otherwise self-excitation willnot take place. Figure-2 shows the case in which the

load line of capacitance is tangent to the criticalslope and, thus, self-excitation does not take place.

This phenomenon can be observed in a laboratoryalso by connecting three-phase supply to an Inductionmachine driven by a prime mover, At the ratedspeed, when the supply to the motor is switched offand simultaneously a capacitor bank is connectedacross it, a steady state voltage is generated at themotor terminals.

The value of capacitance in the capacitor bankshould neither be too low nor too high. If the valueof capacitance is too low, self-excitation will not takeplace. If the value of capacitance is too high, there

will be inadequate build-up of steady state voltagedue to saturation of flux paths. Hence, there is aneed to use optimum size of capacitor for properbuild-up of voltage and also for keeping the cost ofcapacitor bank low.

5. CO NFIGU RAT IO NS O F STATO R

WINDING AND CAPACITOR BANK

There are various configurations of stator winding

and capacitor bank, which are possible for thepurpose of self-excitation, e.g.:

(i) Star-connected stator winding and Star-connected capacitor bank (Fig. 3)

(ii) Star-connected stator winding and Delta-connected capacitor bank.(Fig. 4)

FIG. 1 : LOAD LINE OF CAPACITANCE CUTT INGTHE MAGNETISING CURVE

FIG. 2 : LOAD LINE OF CAPACITANCE TANGENT

TO THE MAGNETISING CURVE

STAR-CONNECTED STAR-CONNECTEDSTATOR WINDING CAPACITOR BANK

FIG. 3


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(iii) Delta-connected stator winding and Star-

connected capacitor bank (Fig. 5)

(iv) Delta-connected stator winding and Delta-connected capacitor bank (Fig. 6)

In all the above four cases, whenever the capacitorsare connected in Delta, the value of capacitancerequired in each phase reduces to one third of itscorresponding value in Star, but the Peak InverseVoltage (PIV) of the capacitors required becomes3times its corresponding value in Star connectedcapacitor bank. As far as performance of the InductionGenerator is concerned, both Star and Deltaconnected capacitor banks are equivalent for a givenconnection of stator winding, and give the sameperformance.

6. INDUCT ION GENERATOR ONLOAD (Fig. 7)

Under no-load condition, since only the magnetizingcurrent is flowing through stator winding, voltagedrop in stator winding is very small, and thereforethe voltage at the generator terminals is almost equalto the induced e.m.f. in stator winding.

When the Induction Generator is connected to apure resistive load and current flows through statorwinding, there is a voltage drop in the statorwinding, which is usually less than 5% of theinduced e.m.f. Now, if the load current is increasedfurther, then at a certain point, where the criticalslope of magnetizing curve coincides with the loadline of capacitor, the Induction Generator stops

STAR-CONNECTED DELTA-CONNECTEDSTATOR WINDING CAPACITOR BANK

FIG. 4

DELTA-CONNECTED STAR-CONNECTEDSTATOR WINDING CAPACITOR BANK

FIG. 5

DELTA-CONNECTED DELTA-CONNECTEDSTATOR WINDING CAPACITOR BANK

FIG. 6 FIG. 7 : GENERATOR WITH LOAD


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generating any voltage. Thus, in an InductionGenerator with a resistive load, there is no possibilityof overheating of the Induction Generator due toover-load / over-current.

When there is an inductive load put across thegenerator, a part of the capacitive current due tocapacitor bank will get neutralized by the inductiveload current. Thus, the magnetizing current availableto the generator will get reduced. Therefore, morecapacitance will be required to be added in thecapacitor bank for maintaining the terminal voltageof the Induction Generator.

7. D ET ER MIN AT IO N O F CAPACI-TANCE VALUE

The following procedure can be adopted fordetermining the value of capacitance required in thecapacitor bank under no-load condition -

Draw the no-load characteristic of Induction Motor.Calculate the critical slope of the magnetizingcurrent (Fig. 8). Draw a load line of capacitorshaving a slope less than the critical slope of

magnetizing current. From this, back calculate thevalue of capacitance required under no-loadcondition.

To calculate the capacitance under full-load condition,the following procedure may be adopted:

There will be a voltage drop of approx. 5% interminal voltage due to resistive load. Inductive load

current will have to be provided by capacitance.Hence, load line of capacitance with inductive loadis drawn accordingly, as shown in Fig. 8. From this,back calculate the value of capacitance requiredunder full-load condition.

Generators of this type are best suited for resistiveloads.

8. CIRCUIT TO EN SURE SELF-

EXCITATION

Due to any reason, if there is no residual flux in themotor, the process of self-excitation can be initiatedby connecting a battery momentarily across one ofthe capacitors of the capacitor bank (Fig. 9). Withthis, the capacitor bank gets charged and when itdischarges through the stator winding, a flux isproduced in the rotor of Induction Machine.

FIG. 8 : DETERMINATION OF CAPACITANCE VALUE

FIG. 9 : SELF-EXCITATION WITH THE AID OF BATTERY

9. VOLTAGE REGULATION

Though in the above paragraphs, it has beenpresumed that at rated voltage the rotor core getsfully saturated, but practically the core is neversaturated fully and thus, it would always result inlarge variation of the terminal voltage with respect

to load current. This problem of drop in voltage


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with load can be overcome in the following twoways:

(i) Manual Voltage Regulator

In this case, the variable capacitors are usedin the capacitor bank in place of fixedcapacitors. By manually varying thecapacitance with respect to load, the terminalvoltage can be maintained. While selectingthe range for the variable capacitor, it shouldbe ensured that the lower value of capacitancecorresponds to load line of generator underno-load condition, and that the higher valueof capacitance in the range corresponds to

load line of generator under fully-loadedcondition. In this case, while starting, thegenerator operation must be started withlower value of capacitance across the terminals.

(ii) Automatic Voltage Regulator

In this case, there is a main capacitor bank,which is used for no-load operation of thegenerator. Then, a provision is kept foradding another capacitor bank in parallel tothe first one when there is a dip in voltageat the terminals. Connection of anothercapacitor bank in parallel is achieved throughan electronic circuit (Fig.10).

FIG. 10 : AUTOMATIC VOLTAGE REGULATION OF INDUCTION GENERATOR


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10. CONCLUSION

The Induction Machine when operated as a Generatoralways requires its excitation from outside. Either itremains connected to the supply if it is to feed powerto grid or its excitation can be provided through acapacitor bank if it is feeding power to a stand-alonesystem. In stand-alone system like wind mill installedin a remote area having no grid power supply, thereare various options available for providing thenecessary excitation to Induction Generator and alsofor maintaining a constant terminal voltage atdifferent loads.

Acknowledgement

The author wishes to express his sincere gratitude toSh S.K. Goyal, GGM, Corp. R & D, BHEL,Hyderabad, for his continuous encouragement andguidance in writing this paper. Thanks are also dueto Sh M.S. Dhami, AGM (EME) & Sh S.C. Goel,SDGM (MM) for their support and help incompleting this paper.

Bibliography

1. S.S. Murthy, C.S. Jha and P.S. Nagendrarao,

"Analysis of grid connected induction generators

dri ven by hydro/wind turbine under realisti c

system constraints," in IEEE Trans. Energy

Conversion, vol. 5, pp. 1-7, Mar. 1990.

2. L. Shridhar, B. Singh, C.S. Jha and B.P.

Singh, "Analysis of self-exci ted induction

generator feeding induction motor," in IEEE

Power Eng. Soc., Summer Meetings, 1994, pp.

1-7.

3. L. Shridhar, B.Singh, C.S. Jha, B.P. Singh and

S.S. Murthy, "Selection of capacitors for the self

regulated short shunt self-excited inductiongenerator," in IEEE Trans. Energy Conversion,

vol. 10, pp. 10-17, M ar. 1995.

4. S.P. Singh, Sanjay K. Jain and J. Sharma,

"Voltage regulati on optimizati on of compensated

self-excited induction generator wi th dynamic

load," in I EEE Trans. Energy Conversion, vol.

19, pp. 724-732, Dec. 2004.

Mr. P.K. Khanna graduated in ElectricalEngineering from Indian Institute of Technology,New Delhi, in the year 1979.

Mr. Khanna joined BHEL, Haridwar, as an EngineerTrainee in 1979 and was posted in AC MachinesEngineering Department. For more than 24 years,

he has been involved in the electrical & mechanicaldesign of various capacities of tailor-made ACmotors for Thermal Power Station, Irrigation,Cement, Petrochemical and other Industries. Hehas also undertaken retrofit jobs to develop complete

design of AC motors to replace non-BHEL-makeAC motors for various customers. At present, he isworking as Deputy General Manager in ElectricalMachines Engineering at Haridwar and is involvedin design and development of 300 MVA TARI AirCooled Generator.

Prior to this paper, Mr. Khanna has contributedone technical paper in an International Conferenceheld at IIT, Roorkee, recently.


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TURBOGENERATOR INDUCED VOLTAGE WAVEFORM

COMPUTATION AND TELEPHONE HARMONIC

CAPABILITY PREDICTION

C. Prem Kumar

SYNOPSIS

The Fini te Element method has enabled accurate

estimati on of the magnetic field in electr ical machines

and devices. The approach has therefore made possible

the accurate estimati on of the vari ous flux-related

parameters the induced voltage magni tude being one

among them. Though the machine characteristic curve

or the open-circui t characteristi c is readily deducible,

the induced voltage waveform, however, i s not. This

paper presents details of a method for the computation

of turbogenerator induced voltage waveform, i ts harmonic

content and machine Telephone Harmonic Factor or the

Telephone Interference Factor all at design stage.

Key Words:

Voltage Waveform; Harmonics; SynchronousMachines; Telephone Harmonic Capability.

1. INTRODUCTION

Accurate evaluation of the magnetic field distributionsin electrical machines has been made possible by the

Finite Element method. Magnetic field-relatedmachine parameters such as inductances, inducedvoltage magnitudes & waveforms, useful & strayfluxes, leakage co-efficients, induction-related losses,saturation effects etc, in turn, stand accuratelyevaluated. Though the open-circuit characteristic isreadily deducible from the magnetic field mapping,the induced voltage waveform, however, is not. Thetime variation of the induced e.m.f in a conductorof the stator in a synchronous machine has the same

form as the space distribution of the flux density in

the air gap. Therefore, only a sinusoidal wave of theairgap flux density can result in a sine-wave inducedvoltage. Several factors such as rotor saturation,shape of the rotor core and the style of field coildisposition render realisation of a sinusoidal airgapflux wave impossible.

This article details an approach to the evaluation ofinduced voltage waveform in a turbogenerator,quantification of harmonic voltage magnitudes andcomputation of Telephone Harmonic capability ofthe machine all at design stage.

2. FE ANALYSIS

Of the several approaches propagated, the FE methodhas found increasing acceptance from industry. Theformulation of the FEM and its application tomagnetic field analysis has been adequately detailedelsewhere[1 to 8]. The FE approach is the cheapest,fastest and certainly the surest way to the predictionof machine parameters at design stage. The fieldsolution is only the first step in the analysis process.More important and relevant in industry are themachine parameters derivable from the field solutions.

The availability of affordable desk-top computingpower and the arrival of powerful menu-drivensoftware have largely contributed to the acceptanceof such methods in design offices[1,2].

2.1 Turbogenerator Magnetic Field at No-

load

Prediction of the induced voltage waveform in aturbogenerator necessitates estimation of the no-load

magnetic field distribution in the machine. At no-


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load, the singly excited magnetic circuit of theturbogenerator presents a picture of symmetry bothalong the direct and quadrature axis. A symmetricquarter region of the machine cross-section extendingfrom the direct axis to the adjacent quadrature axis issufficient to evaluate the no-load parameters of a two-pole turbogenerator. However, in the case ofhydrogenerators, the region for analysis must alsoensure symmetry of the stator slots. Figure-1 shows theflux distribution in a symmetric quarter section of atwo-pole turbogenerator. The corresponding fluxdensity distribution is shown in Fig. 2.

3. AIR-GAP INDUCTION PROFILE

The airgap flux density profile can be extracted fromthe flux density plot by mapping the flux density onto an arc at the mean radius of the machine airgap.Figure-3 shows the radial component and the fluxdensity magnitude mapped along a mean airgap lineof the machine. As can be seen from the graph, thenormal component of the flux density is equal to themagnitude of the flux density for a large portion ofthe curve except at the quadrature axis where thetangential component contributes significantly tothe flux density magnitude. Of the two componentsof the airgap flux density, only the radial componentcontributes to the induced stator voltage while the

peripheral component does not. In reality, the radialcomponent of the airgap induction is very nearlyequal to the total induction at every point in theairgap of the machine.

FIG. 1 : FLUX DISTRIBUTION AT NO-LOAD

FIG. 2 : FLUX DENSITY DISTRIBUTION AT NO-LOAD

FIG. 3 : AIRGAP FLUX DENSITY PROFILE AT NO-LOAD

3.1 Evaluating the Voltage Inducing Flux

The useful flux is defined as the flux linking thestator winding and causing the induced voltage. Thisflux is lesser than the total flux by the amount ofleakage flux. The useful flux in a turbogenerator canbe arrived at by integrating the radial component ofthe airgap induction along the afore-mentionedairgap line.


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The expression for the useful flux per pole is givenby -

u = {.dl}*L

I*S

f*M

f(wb) (1)

where

u

is the useful flux per pole in webers

B is the radial component of the airgapinduction in Tesla

LI

is the nett length of iron in meters

Sf

is the stacking factor(typically around 0.94)

Mf

is the model factor

4. INDUCED E.M.F COMPUTATION

The induced phase voltage in a 3-phase synchronousmachine can readily be arrived at from the usefulflux per pole computed earlier, together with certainstator winding details.

The induced e.m.f per phase is given by [9,13,14]-

Eph

= 4.44*kd*k

p*f*T

ph*[

u] (Volts) (2)

where

kd

is the stator winding distribution factor

kp

is the stator winding pitch factor

f is the frequency in Hertz

Tph

is the number of series turns per phase of thestator winding

u

is the useful flux per pole in webers.

5. HARMONIC ANALYSIS

The time variation of the induced e.m.f in aconductor of the stator winding in a turbogeneratorhas the same form as the space distribution of theflux density in the airgap. Therefore, only a sinusoidalwave of the airgap flux density can result in a sine-wave induced voltage. Several factors such as rotorsaturation, shape of the rotor core and the style offield coil disposition render realisation of a sinusoidalairgap flux wave impossible.

Non-sinusoidal airgap inductions such as the oneshown in Fig.3 can be resolved into a fundamentaland higher-order components using Fourier Analysis.The symmetric airgap flux density wave results inthe cancellation of even harmonics, leaving a spacedistribution comprising a fundamental and harmonicswhich are odd multiples of the fundamental[9,10,11]-

B = B1sin() + B

3sin(3) + B

5sin(5) + . . +

Bnsin(n) (3)

where B1 is the fundamental component of the

airgap flux density and B3, B

5 etc. are the third

harmonic and fifth harmonic components respectively.

Typical harmonic spectrum of the airgap inductionfor a turbogenerator is shown in Fig. 4.

The decomposed representation of the airgap fluxdensity distribution enables consideration of themachine as having 2p pairs of fundamentalpoles(fictitious), 6p pairs of poles contributing tothird harmonic, 10p pairs of poles contributing tofifth harmonic component and, in general, 2np polescontributing to the nthharmonic component of thefield form. The fundamental as well as the harmonic

pole fluxes generate e.m.fs of corresponding frequencyin the conductors, but the proportion of harmonicsin the phase and line e.m.f waveform is reduced dueto grouping and factors related to the stator windingdisposition.

FIG. 4 : AIRGAP INDUCTION HARMONIC SPECTRUM


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6. INDUCED VOLTAGE WAVEFORM

The r.m.s induced voltage per phase due to the nth

harmonic component of the flux density is given by[9,11,13,14,15] -

Eph n

= 4.44 kdn

kpn

fn

nT

ph (Volts) (4)

where

Eph n

= r.m.s induced phase voltage due to the nth

harmonic

kdn

= sin(n/2) / (n/2) is the distributionfactor for nth harmonic

kpn

= sin(n/2) is the pitch factor for nth

harmonic

fn

= nth harmonic frequency

n

= (Bn r.m.s

*Li* /n) is the nth harmonic flux

n = harmonic numberT

ph= number of series turns per phase

and

= phase-belt angular width in elec. radian

= coil-pitch in elec. radian

Bn r.m.s

= r.m.s value of the nth harmonic flux

densityL

i= active length of iron

= pole-pitch in air-gap

The magnitude of the induced r.m.s voltage due toeach of the harmonics can be evaluated using theabove expression.

The variation in time of the fundamental, harmonicand cumulative voltage computed using the procedure

described above is shown in Fig. 5 for a synchronousgenerator whose line-to-line voltage is 11kV.

And finally, the r.m.s value of the resultant phasevoltage is given by [11,12,13] -

FIG. 5 : TYPICAL COMPUTED HARMONIC VOLTAGE PROFILES FOR A 11kV GENERATOR


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Eph

=[(Eph 1

)2 + (Eph 3

)2 + (Eph 5

)2 + +(Eph n

)2]

(Volts) (5)

Eph

= Eph 1

*[1 + (Eph 3

/ Eph 1

)2+ +(Eph n

/Eph 1

)2]

(Volts) (5.a)

The value in the radical is very nearly unity, leadingto the phase voltage being equal to fundamentalalone since the harmonic magnitudes are small incomparison to the fundamental.

7. TELEPHONE HARMONIC FACTOR

Telephone/Communication lines running parallel topower grid lines can experience severe interferenceby induction, resulting in hum and high pitch noisebecause of the presence of harmful frequencies in thegrid. It is, therefore, necessary to limit the harmoniccontent in the output voltage waveform of everygenerator likely to be connected to the grid.

The IEC test procedure and recommendations onthe tolerable limits of telephone harmonic factor(THF) for synchronous machines is reproduced

below -

8.9.2 Limits : When tested on open circui t and at rated

speed and voltage, the telephone harmonic factor(THF)

of the line-to-line terminal voltage as measured according

to the methods laid down in 8.9.3 shall not exceed the

following values:

Rated output of the machine % THF

300kW(orKVA)< PN < 1000kW(orkVA) 5.0%

1000kW(orKVA)< PN < 5000kW(orkVA) 3.0%

5000kW(orKVA)< PN

1.5%

The section 8.9.3 of IEC details the tests and theapproach to be adopted for the measurement ofsynchronous machine THF. The THF capability ofa machine is given by -

THF(%) = 100.(E12.2+E

22.2+E

32.2+ E

n2.2)

U

where

En

is the r.m.s value of the nthharmonic of the

li ne-to-l ine terminal voltage

U is the r.m.s value of the line-to-line terminal

voltage of the machine

i s the wei ght i ng factor for fr equency

corresponding to nthharmonic

8. " WAVE" T H E SPREAD -SH EET

CODE

The procedures detailed in the article have beencoded into a design office utility package by name"WAVE". Exclusively developed for synchronous

machines, the Microsoft Excel spread-sheet utilitycode computes and displays graphically theinformation on harmonic voltage magnitudes/waveforms, cumulative voltage waveform and machineTelephone Harmonic Factor (THF) for each harmonicand cumulative value up to the 100th harmonic.The program has built-in logic to account for triplenharmonics and even harmonics from the computedvalues for line and phase quantities.

The code is structured in two levels and spread over12 sheets and can be tailored to suit individualdesign office requirements. An overview of the codefollows :

The Machine ID Section of the utility is anidentifier section for design office documentation/records and contains information on customer name,order number and machine nominal rating particulars.The program uses colour coding to distinguishunlocked input cells from locked coded cells.

The Inputs Section seeks machine dimensionalinformation such as gross length, number of radialventilating ducts & their width, rotor diameter,stator inner diameter, coil-throw, bars per slot etc.These input details are made use of in thecomputation through formulae embedded in thecells (Fig. 6).

The Computation Section uses the input data tocompute harmonic winding factors, pole-pitch value

for each harmonic, harmonic flux magnitudes,


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FIG. 6 : THE IN PUTS SH EET IN " WAVE" SPREAD-SHEET UTILIT Y


34/68


induced phase e.m.f due to each harmonic etc forthe fundamental and harmonics up to the 100th. Theprogram can detect and display the type of windingemployed. The program can tackle both integral-slotand fractional-slot chorded winding which are verycommon with practical synchronous machines withtwo-pole cylindrical rotor constructions and multiplesalient-pole low-speed hydrogenerators.

The THF Section uses the embedded harmonicweighting factors to compute the individual andcumulative THF due to each harmonic up topredefined significant harmonic. This section alsogenerates a plot of the %THF versus harmonicnumber, as shown in Fig. 7.

Other Features : Graphical comparison of therelative magnitudes of fundamental and significantharmonics (up to13th) and its variation in time are

provided together with a zoom of significantharmonics. The IEC:1996 recommendations on theweighting factor for various frequencies to be usedfor the computation of Telephone Harmonic Factor(THF) have been built into the program. The airgapflux density harmonic magnitudes are used toreconstruct and compare with the original airgapinduction curve.

9. CONCLUSION

The Finite Element approach has been used for theaccurate estimation of the magnetic field in aturbogenerator at no-load. Harmonic components of

the airgap induction computed from the airgapinduction profile have been used to arrive at theharmonic voltage magnitudes and the cumulativeinduced voltage waveform using a code specifically

FIG. 7 : THF CONTRIBUTION FROM INDIVIDUAL HARMONICS AND CUMULATIVE VARIATION WITH INCREASING HARMONIC NUMBER


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developed. The analysis indicates that for a "n" slotmachine, the (n-1)th harmonic and the (n+1)th

harmonic are predominant[11]. The TelephoneHarmonic Factor or the Telephone InterferenceFactor have been computed for the machine analysed.The harmonic levels of the airgap induction aredictated by the airgap induction profile which, inturn, reflects the radial airgap permeance presentedby the magnetic boundaries constituted by the statorand rotor iron. A separate study investigating thedependance of THF on saturation in the machineis under way.

References

(1) M .V.K. Chari , P.Si lvester, "Analysi s of

Turboalternator M agneti c Fields by Fini te

Elements", IEEE Trans. on PAS, Vol-PAS-90,

No:2, March-Apri l 1971, pp.454 to 464.

(2) M.V.K. Chari "Finite Element Analysis of

Electrical Machinery and Devices", IEEE Trans.

on Magneti cs, Vol.MAG-16, No.5, Sept.1980,

pp. 1014 to 1019.

(3) M.V.K. Chari , "Nonlinear Finite ElementSolution of Electr ical Machines Under No-load

and Full-Load Conditions", pp.686 to 689.

(4) P.Silvester and M.V.K. Chari, "Fini te Element

solution of Saturable Magnetic Field Problems",

IEEE. Trans. on PAS, Vol.PAS-89, No.7, Sept-

Oct 1970, pp.1642 to 1651.

(5) P. Silvester, H.S. Cabayan and B.T. Browne,

"Effi cient Techniques for Fi ni te Element analysis

of Electric Machines", IEEE Trans. on PAS,

Vol.PAS-92 1971,pp.1274 to 1281.

(6) Parviz Rafinejad et al,"Fini te Element

Computer Programs i n D esi gn of

Electromagnetic Devices", IEEE Trans. on

Magneti cs, Vol. MAG-12, No.5, Sept1976,

pp.575 to 578.

(7) Mulukutla S.Sarma, "Magnetostatic Field

Computation by Fini te Element Formulation",

IEEE. Trans. on Magneti cs, Vol.MAG-12, No.6,

Nov1976, pp1050 to 1052.

(8) J.R. Brauer, E.A. Aronson et al,"Three

Dimensional Fi ni te Element Calculation of

Saturable Magnetic Fluxes and Torques of an

Actuator" , I EEE. Trans. on M agneti cs,

Vol.M AG-24, No.1, January 1988, pp455 to

458.

(9) M.G. Say, "The Performance and Design of

Alternati ng Current Machines", Sir I saac

Pitman & Sons, London.

(10) Ralph R.Lawrence & Henry E. Richards,

"Principles of Alternating Current Machinery",

McGraw-Hill Book Company, USA.

(11) Alexander S. Langsdorf, "Theory of Alternating-

Current Machinery", M cGraw-Hi ll Book

Company, USA.

(12) Robert L.Ames, "A.C. Generators: Design and

Application",John Wiley & Sons, USA.

(13) Mulukut la S. Sarma, "Synchronous Machines

- Their Theory, Stability and Excitation

Systems", Gordon & Breach Science Publ ishers,

New York.

(14) Essam S. Hamdi, "Design of Small Electr ical

Machines", John Wiley & Sons, USA.

(15) Brian Chalmers & Alan Will iamson, "A.C.

Machines - Electromagnetics and Design" , John

Wi ley & Sons Inc, USA.


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Mr. C. Prem Kumar obtained his Engineeringdegree from Bangalore University in the year1975.

Mr. Prem Kumar joined BHEL at the Corporate

R&D Division, Hyderabad, after a brief stint withM/s Oblum Electrical Industries, Hyderabad.Currently, he is working at the Electrical Machines

Lab in the R&D Complex, as Senior DeputyGeneral Manager. He specializes in the modellingand analysis of large rotating electrical machinesusing the Finite Element approach.

Mr. Prem Kumar has several papers to hi

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