DESIGN AND TESTING OF THE 210 MW AIR-COOLED SERIAL

TURBOGENERATOR FOR THE JYVASKYLA CHP IN FINLAND

N.Pinchuk, O.Antonuk, M.Roytgarts

INTRODUCTION

Air-cooled turbogenerators are distinguished by high reliability, explosion safety, easy

and convenient in operation. In view of increasing demand, a new series of air-cooled

turbogenerators for steam and gas turbines, having the common design solutions and similar-

type ventilation system [1], was developed at the Branch Electrosila of the OJSC Power

Machines in St. Petersburg. By the present moment about 60 turbogenerators of this series

have been manufactured and under successful operation at sites both in Russia and abroad.

The air-cooled turbogenerators of 100 MVA and 200 MVA output have been supplied for

steam and gas turbines in Hungary, Croatia, Iraq, Kazakhstan and other countries of the

world.

Type Т3ФП-220-2У3 turbogenerator with the rated output of 233 MVA for the

Jyvaskyla Project (Finland) was manufactured on the bases of the series 265 MVA

turbogenerator to suit the requirements of the Contract.

Main contractual requirements

Rated output 233.3 MVA (guaranteed)

Rated power factor 0.90

Rated voltage 15.75 kV

Rated frequency 50 Hz

Rated secondary coolant temperature 25°C

Efficiency, no minus tolerance 98.7% (guaranteed)

Brushless exciter

Indirect air-cooling of stator winding

Direct air-cooling of rotor winding

The winding insulation thermal classification 155(F)

The insulation temperature under the rated conditions 130(B)

The bearing vibrations measured in three mutually perpendicular directions shall

comply with zone A (3.8 mm/s) of ISO 10816-2 (guaranteed value).

The shaft vibrations shall comply with the zone A (80 micron) of ISO 7919-2

(guaranteed value).

The average noise pressure level measured at a distance of 1 m from the generator

surface shall not exceed 82 dB(A) according to ISO 3746 (guaranty).

Generator shall comply with IEC Standards, EC Directives and official regulations and

requirements of authorities valid in Finland at the time of the take-over.

The guarantee excluding electronic components of exciter shall be valid four years

from the take-over.

These rather strict requirements were taken into account when manufacturing the

above-mentioned turbogenerator (Fig. 1).

Fig. 1. Type Т3ФП-220-2У3 turbogenerator at the manufacturing plant test rig

TURBOGENERATOR DESIGN FEATURES

STATOR WINDING AND CORE

In the series air-cooled type Т3ФП-220-2У3 turbogenerator the use was made of the

conventional double-layer bar-type winding with transposition of strands both in the slot part

and in the end-winding part in order to reduce the stray-load losses. Upper and lower bars

have different number of strands in order to equalize the specific losses along the bar height.

The stator winding high-voltage insulation quality is of fundamental importance. The

use of dry thin insulation for bars with subsequent vacuum-pressure impregnation of the stator

complete with the winding as per the Global VPI technology (Fig. 2) reduced the labour

input, permitted to reduce the temperature drop in the insulation thickness and to increase the

specific current load with the assurance of heating the class F insulation as per class B. During

the vacuum-pressure impregnation of the stator complete with the winding, the clearances

between the winding and the stator core steel are filled with compound, the insulation thermal

conductivity is improved and the stator rigidity relative to bending vibrations under the action

of radial electromagnetic forces is increased.

Fig. 2. Stator before the vacuum-pressure impregnation

In order to decrease the losses and heating as well as to improve the reliability of

turbogenerator stator end zone, the direction of winding the stator (arrangement of the leading

and lagging layers of the double-layer winding) was chosen based on the numerical modelling

with due account of the field rotation direction [2].

The cold-rolled electric steel with decreased magnetic losses was used in order to

reduce the stator core losses. When fastening the core onto the dovetail bars, the end packs

were insulated against the fastening dovetail bars in order to limit the ways of eddy current

flowing into the end packs. The flexible copper links were used to close the eddy currents

induced by the magnetic stray field of the stator core in the fastening dovetail bars. The

copper links reduce the stray-load losses and the heating of the end packs and pressure plates

of the turbogenerator stator core.

ROTOR WINDING AND SHAFT

To manufacture the rotor the use was made of high-strength steel forging with alloy

additions. The rotor winding was laid in the rectangular slots. The direct cooled rotor design

is shown in Fig. 3.

Fig. 3. Rotor of the air-cooled turbogenerator

When laying the winding into the rotor slots, the technology of high-frequency

induction brazing of half-turns was used. The rotor multi turn coil winding distribution in

slots was chosen from the condition of achieving the best shape of the field, with which the

excitation magnetic field is of the maximum use to create the first voltage harmonic at the

turbogenerator terminals. In the process of designing both the basic losses in the field winding

and the excitation-field harmonic stray-load losses on the stator surface were minimized.

In order to eliminate the unequal rotor rigidity in the longitudinal axis and the

transverse axis, which causes the rotor vibration at the double speed of rotation, longitudinal

slots on the big tooth of each pole were made in the rotor body. The dimensions and the

number of these slots were chosen so that the main moments of the rotor inertia would have

been equalized. The equalizing slots were filled with a magnetic material with cloth-based

laminate gaskets. When the rotor is rotating, the longitudinal forces acting on the filler

material are reduced due to a shearing strain, and the filler material does not make the rotor

body more rigid. The filler material is held in the equalizing slots by means of wedges, which

are a part of a complete rotor damping system. The system of equalizing the unequal rigidity

provides the ultimate lower vibration level of the rotor and bearings with the double speed of

rotation.

Fig. 4. Assembly of the stator with the rotor inserted

VENTILATION SYSTEM AND THERMAL CONDITION

To increase the efficiency of active parts cooling, new engineering solutions (Fig. 5)

allowing rational distribution of airflows in the cooling system were applied.

In the exhaust circuit of ventilation chosen, the air circulation circuits in the stator core

and rotor winding are separated. The centrifugal fans, which provide the cooling air

circulation in the stator core ducts, were used. The air motion in the rotor cooling ducts is

initiated exclusively through a pressure created by these ducts.

Fig.5. Turbogenerator cooling system

Fig.6. Turbogenerator rotor and stator ventilation

As in many world large companies, for cooling of the rotor winding the radial ducts

are used for the slot portion and the axial ducts are used for the end windings. The radial ducts

are provided with cooling air by means of under-slot ducts.

The main advantage of this ventilation system is the uniform gas supply to the stator

core ducts and, as a consequence, the practically uniform distribution of the winding and

active steel temperatures along the circumference and along the stator length.

As shown by the tests performed under operational conditions of various

turbogenerators of the series considered, the difference of the maximum measured

temperature from the average value for the active steel does not exceed a value of 5°-7°C.

Practically uniform temperature distribution in the stator was achieved due to the fact

that the losses generated in the slot portion were removed by the system of U-shaped ducts,

which do not communicate with the air gap between the stator and rotor. The system of U-

shaped ducts is shown in Fig. 6. In the design with U-shaped ducts, cold air from a discharge

chamber located at the stator periphery reaches the radial ducts between the active steel packs

and successively passes first towards the bore, then comes back in the opposite direction

towards the rarefaction chambers, also located at the stator periphery.

The type Т3ФП-220-2У3 turbogenerator design was updated with due account of the

contract required rated power decrease with respect to the series turbogenerator output as well

as of the requirements to the efficiency value. The ventilation and mechanical losses for the

air-cooled turbogenerators amount to about 40% of the total losses [1]. To ensure the

efficiency guarantee value the mechanical losses are decreased through the reduction of the

cooling airflow through the rotor and stator. By using the centrifugal fan of lower capacity the

airflow through the stator is decreased, by changing the cross-section of the rotor under-slot

cooling ducts the airflow through the rotor is also decreased. In doing so, the ventilation

system efficiency was proved by calculations.

NOISE-PROTECTION ENCLOSURE

The special feature of the turbogenerator noise-protection enclosure is that it is also

the housing, along the inner surface of which the cooling air flows at a high speed. The results

of the experimental studies on the turbogenerator noise level and spectral composition showed

that the generator internal noise is of the broadband nature and the noise level amounts to

about 116 dBA. To meet the contract requirements, the noise had to be decreased by 34 dB,

that is the value had to be 50 times less. For this purpose the high-effective noise- protection

composite materials were chosen and the acoustically closed enclosure was manufactured.

Taken into account was that the enclosure design efficiency was less that the efficiency of the

materials used. The through metal elements of the structure and slits are the sound conductors

that lead to the generation of secondary sound sources on the enclosure surface, which

substantially decrease the noise protection efficiency. The noise- protection enclosure

structure covers the turbogenerator itself, the brushless exciter and bearings as well as the

neutral terminals box, the housing of which is made of aluminium in order to meet the

electromagnetic compatibility requirements (Fig. 1).

RESULTS OF FINAL FACTORY RIG TESTS

The generator with exciter and synchronous motor were assembled and balanced for

the carrying out the final factory tests.

When analyzing the test results, let us limit ourselves with the check of meeting the

most important operational guarantees.

LOSSES AND EFFICIENCY

According to the contract the turbogenerator guaranteed efficiency at the rated

parameters without negative tolerance shall amount to at least 98.7%.

When testing the generator at the manufacturing plant test rig, the efficiency η is

determined by using the indirect method of separated losses by the following formula:

( )[ ] %,PP/P1100 2∑ ∑+−⋅=η (1)

where 2P and ∑P are the active power and the sum of losses in the turbogenerator

respectively, kW

The sum of losses is determined by the following equation:

∑ +++= fscfemech PPPPP (2)

where:

mechP are the mechanical losses, kW;

feP are the stator active steel losses, kW;

scP are the short-circuit losses, kW;

fP are the excitation losses, kW.

The losses Pmech, Pfe and Psc are determined simultaneously with the determination of

the no-load and short-circuit characteristics. When taking the readings of each point of

characteristics of the generator under test, the measurements of power P1 consumed by the

drive motor as well as of the current and voltage at the motor terminals are taken. The above-

mentioned measurements permit to determine the motor losses ΣРmot based on the earlier

performed calibration results.

The losses of the generator under test are determined as a difference between the

consumable power by the drive motor and the losses of the motor itself.

∑−= mot1gen PPP (3)

The generator mechanical losses Pmech are determined as the averaged losses obtained

in the no-load and short-circuit tests in points U = 0 and I = 0. In this case the generator and

motor total mechanical losses are taken to calculate the efficiency.

The excitation losses are determined by the following formula:

ex75f2

ff PRIP +⋅= (4)

where: If is the turbogenerator excitation current at the rated load conditions (design value) А;

Rf is the rotor winding resistance (measured value) reduced to the design operating

temperature, Ohm;

Рex are the exciter losses (without the mechanical ones) when loaded with rotor rated

current If, reduced to the design operating temperature, kW.

Results of the losses determination are shown in Table 1.

Table 1. Turbogenerator losses and efficiency

Description Designation Value

Mechanical losses, kW Pmech 1216

Active steel losses, kW Pfe 362

Short-circuit losses, kW Psc 634

Excitation losses, kW Pf 407

Total losses, kW ΣP 2619

Efficiency, % η 98.77

The values included in formula (1) are determined with some uncertainty, which shall

be taken into consideration in the final result. For the quantitative evaluation of the

measurements accuracy it is expedient to point out the dependence of the efficiency indirect

measurement results upon the uncertainty of the separated losses determination (truncation

error). Absolute error of the efficiency values measured is expressed by the following

formula:

( ) 100/P)100/1( ∑⋅−⋅= δηηη∆ , % (5)

Therein ∑Pδ is the relative error of the separated losses sum determining, in percent.

Shown in Fig. 7 are the curves of the efficiency determination absolute error versus both

the efficiency value measured and the losses sum determination error.

From formula (5) and Fig. 7 it follows that, for example, in case the losses

determination error is 1%, the generator efficiency value measured being up to 99% has no

more than three true significant digits.

The separated losses determination error in its turn depends upon the procedure and the

uncertainty of these losses measurements. The separated losses determination errors have the

practically normal law of distribution; the error confidence limits, with the specified

probability within which is the error, are used for their calculation.

Taking into account the truncation error for the method of separated losses the generator

efficiency value is determined by the following formula:

( )[ ]100/P)100/1(1t ∑⋅−±⋅=±= δηηη∆ηη (6)

With due account of the measurements uncertainty the efficiency of the tested

turbogenerator was equal to 98.77±0.05%, which value for sure corresponded to the contract

requirements. The record of the efficiency measured value without showing the uncertainty

would have meant that the efficiency determination uncertainty did not exceed the one half of

unit of the value the last digit. As per formula (5) the uncertainty of the separate losses sum

determination in this case shall be no less than 0.4%, which is practically unachievable.

Fig. 7. Truncation error of the efficiency determination using the indirect method

HEAT RUN TEST

Temperatures of the structural parts were measured by means of the standard

temperature monitoring system, i.e. RTDs.

The stator winding and active steel temperature rise was determined by the following

formula:

noexnlshc ∆Θ∆Θ∆Θ∆Θ −+= , (7)

where noexnlshc ,, ∆Θ∆Θ∆Θ are the temperature rises under the following conditions: short-

circuit conditions with the rated current in the stator winding, no-load conditions with the

stator rated voltage and no-load conditions without excitation, respectively.

The rotor winding temperature was determined by the resistance value, the use was

made of the exciter load characteristics and the results of the exciter excitation voltage and

current measurements under the no-load and short-circuit conditions. Under the no-load

conditions without excitation the measuring current was supplied to the slip rings from the

external power source. The rotor winding temperature rises were obtained by extrapolating

the temperature rise values in relation to the losses in the winding.

Table 2. Temperature rises and generator heating

Stator winding Stator active steel Rotor winding Operating

conditions Temperature rise, ºC

Heating, ºC

Temperature rise, ºC

Heating, ºC

Temperature rise, ºC

Heating, ºC

Without excitation 6 38 4 36 14 46

No-load 26 58 17 49 17 49 Short-circuit 42 74 20 52 37 69

Rated load 62 94 33 65 70 102

As shown in the table, the main generator active assemblies have the heating margin

corresponding to 10% of increase in power. The heating margin improves the generator

operational reliability.

To meet the contract requirement, the neutral terminals box is of closed design, thus

excluding the heat exchange with the surrounding space. In this connection the maximum

temperature of neutral terminals plate (six points) and that of the current transformer surface

(six points) was measured under the steady-state thermal conditions by using the temperature-

indicating tags. These measurements showed that under the sustained short-circuit conditions

with the stator rated current and at a cold air temperature downstream the air coolers of 32°C,

the maximum temperature of the clamping plate is 75ºC, while the temperature of the current

transformer surface is 59ºC. Thus, during the generator operation no overheating occurs in the

neutral terminals box.

VIBRATION DETERMINATION

In the course of the factory tests the vibration velocity spectrum components were

measured in three mutually perpendicular directions in the frequency band of 10…500 Hz.

The root-mean-square value of the vibration velocity shall be no more than 3.8 mm/s. The

shaft vibration double (peak-to-peak) amplitude shall be no more than 80 µm.

The results of the vibration measurements taken at the test rig are shown in Table 3.

Table 3. Bearing vibration, mm/s, and shaft vibration, micron

Bearing, turbine side Bearing, exciter side Operating conditions Vertical Transverse Axial Vertical Transverse Axial

Shaft vibration

Without excitation 1.3 1.7 3.5 1.8 0.7 2.6 53

No-load U=Urated

1.4 1.8 3.5 2.1 0.9 3.2 53

Short-circuit I=Irated

1.3 2.1 3.7 2.5 1.0 3.4 53

The low levels of vibration are provided by:

- removing the rotor shaft and bearing natural frequencies in excess of 15% of the

rotational speed (the first critical frequency is at 940 rpm, while the second one is at 2430

rpm);

- highly accurate rotor balancing, the generating unit assembling and vibration

adjustment at the test rig.

TURBOGENERATOR NOISE

In accordance with the contract the average level of sound pressure created by the

generator and exciter at the power plant shall not exceed 82 bB(A) when the measurements

are taken at a distance of 1 m from the generator housing as per ISO 3746.

The noise measurements were conducted at the manufacturing plant test rig with the

generator operating under the motor mode without load at the rated speed of rotation and the

rated stator voltage. The measurement points were located around the generator at a distance

of 1 m from the generator surface. In each point the sound level and the noise pressure levels

were determined in 1/3 of octave-frequency band being within 50 to 10000 Hz. 49 points

were used to measure the air noise. The measurements were conducted using the precision

instrumentation of Bruel&Kjaer Company made.

Fig. 8 shows the root-mean-square values of the sound pressure level measurement

results. The spectral analysis showed that the generator arranged in the complete set of noise-

protection enclosure is the source of the broadband noise with the relatively flat spectrum as

well as that the slightly defined discrete tones at the frequencies of 50 Hz and 100 Hz and at

the speed of the fan blades rotation are available. Based on the measurement results the

average sound level around the generator was 78 dBA. The noise protection efficiency was 38

dB relative to a sound level of 116 dBA measured under the housing, that is the noise was 80

times reduced. The neutral terminal box have the aluminium housing and above it the

additional noise-protection enclosure. With the additional neutral terminal-box noise-

protection enclosure removed the average sound level was 82 dBA. Based on the spectral

analysis, the noise protection efficiency did not change at lower frequencies and worsened at

medium and high frequencies of the sound range.

Fig. 8. Values of the generator sound pressure levels in 1/3 of octave-frequency band

The study of the test rig acoustical conditions showed that the results reproducibility has

the root-mean-square deviation less than 3 dB. Hence it follows that at a confidence

probability of 90% the noise level measurement uncertainty is within ± 4 dB.

Thus the measured noise of the turbogenerator arranged in the noise-protection

enclosure with a probability of 90% does not exceed 82 dB(A).

TESTS ON INSULATION OF THE STATOR WINDING BARS

The tests on insulation of two stator winding bars were conducted in accordance with

European Standard EN 50209 at a specially manufactured slot test mock-up prior to the

winding laying into the stator slots. The Standard requirements cover the insulation made of

preliminarily impregnated tapes as well as the insulation obtained by applying the vacuum-

pressure impregnation method.

The first step of the test is the measurement of the dielectric loss tangent.

The voltage varied from 0.2 to 1 Urated with an increment of 0.2 Urated. The

measurements were conducted in the initial conditions and after heating up to 90°C with

subsequent cooling to an ambient temperature.

As the evaluation criteria the use was made of the maximum value of the ∆tgδ change

for the measured interval of 0.2Urated before and after the sample heating. The results of

measurements before and after heating shall not differ by more than 0.2% and shall not be

more than 0.7%. Results of measurements are shown in Table 4.

Table 4. Loss tangent change during the tests

,%tgδ∆

Test sample 1 Test sample 2 Range of change

In the initial conditions

After heating up to 90ºC and subsequent

cooling

In the initial conditions

After heating up to 90ºC and subsequent

cooling 2.04.0 tgtg δδ − 0.025 0.010 0.004 0.012

4.06.0 tgtg δδ − 0.040 0.010 0.009 0.004

6.08.0 tgtg δδ − 0.046 0.030 0.006 0.006

8.00.1 tgtg δδ − 0.024 0.011 0.004 0.003

0.5( )tgtg 2.06.0 δδ − 0.033 0.010 0.007 0.008

Then the bars withstood the overvoltage tests with a voltage of 32.5 kV, 50 Hz, applied

for 1 min. With the voltage being raised up to 70 kV, no breakdown or flashover occurred.

Thereby the insulation quality and compliance with Standard EN 50209 requirements were

proved.

After completing the rig tests the final high-voltage dielectric strength tests on the stator

winding were conducted with a voltage of 32.5 kV, 50 Hz, applied for 1 min. The tests were

conducted for each phase with other phases being shorted to the ground housing. The

generator has successfully withstood these tests.

CONCLUSION

The factory rig tests of the air-cooled type Т3ФП-220-2 turbogenerator manufactured

for Jyvaskyla TPP showed the complete compliance with European standards, norms and

contract requirements imposed on the guaranteed technical parameters, including the rated

output, efficiency and heating levels of the main assemblies, winding insulation durability,

noise and vibration levels, etc. The factory tests analysing was made taking into account not

only of the directly measured values but of the methodological and statistical uncertainty of

measurements.

At the present time the Т3ФП-220-2 turbogenerator is erected at the power plant in

Finland and is under the complex of the commissioning tests being a part of the power-

generating unit.

The demand for high-reliable and easy-to-operate turbogenerators directs the design

and research works for the further increase in the unit power of air-cooled turbogenerators,

which corresponds to the tendencies of the world electric machine industry development.

REFERENCES

1. O.Antonyuk, T.Kartashova, M.Roytgarts. Series of air-cooled turbogenerators of

power up to 410 MVA. Intetnational Conference POWER-GEN Europe 2009, 26-29 May

2009, Cologne, Germany.

2. M.Roytgarts, Yu.Varlamov, A.Smirnov. Electromagnetic computation in the end

zone of loaded turbogenerator. ADVANCED COMPUTER TECHNIQUES IN APPLIED

ELECTROMAGNETICS/ Springer, 2008, IOS PRESS. Chapter C. Applications/

C1.Electrical Machines and Transformers.

3. Y.Samorodov. Defects of generators, M, 2005, 350 p. (In Russian)

APPLICATION

FOREIGN OBJECTS PENETRATION INSIDE THE TURBOGENERATOR

It is well known that the penetration of foreign objects inside the turbogenerator is

impermissible and may result in severe accidents and emergency [3].

The procedure of calculating the electromagnetic fields in the end zone, which was

developed at Branch Electrosila, permitted to perform the numerical simulation of possible

emergency with stator winding damage due to the foreign objects getting inside the

turbogenerator. Figs 9 and 10 show the model for the numerical analysis and the results of

calculating the electromagnetic field in the turbogenerator end zone at the rated stator winding

current [2].

Fig.9. Numerical simulation of the turbogenerator end zone

Fig.10. Results of calculating the electromagnetic field

a) b)

Fig. 11. Distribution of eddy currents (a) and losses (b) in a ferromagnetic object

Fig.12. Ferromagnetic object heating on the stator-winding surface

When the foreign object gets on the stator winding end portions, the eddy currents are

induced in this object under the action of the turbogenerator three-dimensional electromagnetic

field (Fig. 11a) and the losses are generated (Fig. 11b). In case the object is of magnetic steel, the

eddy currents have the drastically non-uniform distribution along the object. These eddy currents

are displaced to the surface and flow in the layer with a thickness equal to the depth of the field

penetration. For magnetic steel this value is 1.3 mm. The value of the losses generated in the

object are the quadratic function of the magnetic flux density, therefore these losses are

generated in even much thinner layer of the steel object (Fig. 11b).

The heating is determined based on the obtained distribution of losses in the object

volume, taking into account the cooling conditions (Fig. 12). According to the calculation results

the steel object temperature on the bar insulation surface reaches 700ºC, that is more than enough

for the bar ground and turn-to-turn insulation destruction. Thus, the ferromagnetic object getting

inside the turbogenerator inevitably leads to the emergency situation in the turbogenerator.