Special_Field_Test.pdf

10/12/10

NATIONAL ELECTRIC COIL

GENERATOR REWIND

SECTION 16210

PARAGRAPH 3.26

GENERATOR SPECIAL FIELD TEST REPORT

Submittal #20.1

U.S. ARMY CORPS OF ENGINEERS

BIG BEND GENERATORS

UNITS 4&6

CONTRACT NO. W9128F-07-C-0016

October, 2010 Rev.1

GENERATOR SPECIAL FIELD TESTS

TEST REPORT

Units 4 and 6 Big Bend Dam

WH Vertical Waterwheel Generator Uprated to 70816 KVA at 90oC stator rise, 2963 A

0.95 power factor

13.8 kV, 81.8 rpm, 60 Hz for

U.S. Army Corps of Engineers

Contract No. W9128F-07-C-0016 by

National Electric Coil Columbus, Ohio

EXECUTIVE SUMMARY

1. All generator special field tests were performed on unit 4 in accordance with the requirements of Contract No. W9128F-07-C-0016 and of IEEE Std 115-1995

– Guide: Test Procedures for Synchronous Machines.

2. The stator core loss was measured as 340.8 kW. The guaranteed value of 369 kW was not exceeded.

3. The stator winding DC I2R loss was determined as 172.73 kW at 61,579 kVA, 0.95 power factor and the stator winding resistance corrected to 75oC. The expected I2R loss was 189.06 kW.

4. At 61,579 kVA, 0.95 power factor The measured stray load loss is 144.0 kW.

The calculated stray load loss value was 134.36 kW.

5. At 61,579 kVA, 0.95 power factor:

The measured combined stator dc I2R (copper) loss and stray load loss, including resistance of the circuit rings, is 316.73 kW.

The guaranteed value of the combined stator dc I2R (copper) loss and stray load loss, including resistance of the circuit rings, is 323.42 kW.

6. At 61,579 kVA, 0.95 power factor The measured combined losses of the new winding and core are 657.53 kW, The guaranteed combined losses of the new winding and core are 692.42 kW.

7. During temperature tests full loading of the generator was not possible.

The stator winding temperature rise was measured at three partial loads and the results were extrapolated to the new generator rating of 70,816 kVA at 0.95 power factor. The temperature rise of 56.6oC has been determined by extrapolation. (In 1967 this temperature rise was measured directly and found to be 63.5oC). A temperature rise of 90oC was guaranteed for the new rating.

8. The ozone concentration was measured as 0.016 ppm. The allowable limit of 0.1 ppm was not exceeded.

9. The stator voltage wave form deviation does not exceed the allowable limit of 10%. The measured telephone influence factors are within allowable limits.

10. All determination tests specified in Section 16210, paragraph 3.26.3 have been performed in compliance with IEEE Std 115-1995. The required machine constants have been determined by the determination tests.

1. PURPOSE OF TESTS

The Big Bend Dam generating units 4 and 6 are equipped with Westinghouse AC generators. They are 3-phase, 60 Hz, 81.8 rpm, 13.8 kV AC generators originally rated at 61,579 kVA at 0.95 power factor.

National Electric Coil designed, manufactured and installed new stator windings and the generators were uprated to 70,816 kVA at 90oC stator winding temperature rise at

0.95 power factor.

The generator special field tests were performed in March-April 2010 to determine baseline data for further performance monitoring and to verify compliance with guaranteed I2R losses, stray load losses, and temperature limits.

The following tests were performed on unit 4:

1. Losses tests using the retardation method.

2. Temperature test (heat run) at 52402 kVA, 55153 kVA and 58120 kVA,

0.95 power factor, 13.8 kV.

3. Ozone level measurement in the generator air housing during the temperature test

4. Open-circuit saturation test

5. Short-circuit saturation test

6. Zero power factor saturation test

7. Stator voltage waveform deviation test

8. Test to determine the Telephone Influence Factor

9. Sustained short circuit tests to determine the negative sequence reactance and the zero sequence reactance.

10. Sudden short circuit tests to determine transient and subtransient reactances and time constants

2. INSTRUMENTATION SETUP

2.1. Terminal Voltage. NEC’s 14,400/120V PT’s secondary voltages were measured by a Yokogawa WT-1600 power analyzer. During retardation tests the secondary voltages were converted to a 0-10VDC analog signal using certified voltage transducers and were recorded by a Yokogawa DX208 automatic data acquisition station.

2.2. Stator Current. The secondary circuits of the NEC CT’s were measured by a Yokogawa WT-1600 power analyzer. During retardation tests NEC used 6000A/5A CT’s. One of them failed and was replaced by a 3000A/5A CT. The secondary voltages were converted to a 0 – 10VDC analog signal. The signal was recorded by the Yokogawa DX208 data acquisition station. NEC used another set of 2500A/5A CT’s for the rest of the tests.

2.3. Field Current. A certified DC Ammeter shunt device was used to monitor the field current. The rating of this shunt is 100 mV for 1000 amps. This millivolt signal was transmitted to the data acquisition station for recording field current during the tests.

2.4. Shaft Speed. The WT_1600 power analyzer measured the generator frequency. For retardation data logging a black background tape was placed on the turbine shaft. A transverse reflecting tape, about 1/2” wide, was placed on the shaft circumference. A DC light source illuminated the tape at the reading area. A photocell pickup reported passing of the tape to a Monarch ACT Signal Conditioner. It read the pulse signal and converted the pulse rate to an analog signal which was recorded by the data acquisition station.

2.5. Stator Winding Temperature

In accordance with IEEE 115 the stator winding temperature was measured by the embedded detector method. The DX230 data acquisition station recorded the output signal of the embedded RTD. The highest RTD reading was used as the stator winding temperature.

2.6. Stator Core and Ambient Temperatures

Per IEEE 115, the ambient temperature is the average of the temperatures of the coolant leaving the heat exchangers and entering the machine. Type T thermocouples were installed at the air coolers discharge and at the stator core back iron. The data acquisition station recorded their output signals. The average of the air temperature readings was used as the ambient temperature.

2.7. Stator Winding Resistance. The resistance of each of the phases was measured at the machine terminals using a certified digital low resistance micro-ohmmeter after winding installation was complete. The temperature of the stator winding at the time of the resistance measurement was recorded.

For use in adjusting the losses measured during the retardation runs, the resistance of the short circuit path (the leads and the breaker) was measured, and then adjusted as necessary for the temperatures found during the tests.

2.8 Ozone measurement. An ozone sensor was installed in the generator barrel.

The ozone concentration in the generator barrel was measured by a S-200L type Ozone Solutions monitor with a low range (0.001-0.5 ppm) display.

3. DC I2R LOSS, CORE LOSS AND STRAY LOAD LOSS TESTS

(RETARDATION TESTS)

The retardation method is based on the relationship between the rate of deceleration of a rotating mass, its weight and radius of gyration, and the power loss tending to decelerate it.

Machine losses are obtained from retardation tests made under conditions such that the power tending to decelerate the machine is the loss (combination of losses) to be determined.

Knowing the rate of deceleration, the loss is determined by the following equation:

6, 10 dnLosses kW k J n dt

Where n = rotational speed = 81.8 rpm dn/dt = rate of deceleration as determined from the slope of the speed-time curve at “n” rpm k = 0.4621 J 115.0 x 106 lb*ft2 = moment of inertia of the rotating parts.

The value of J was obtained from the Westinghouse commissioning test report data, pages 19 and 25 of the 1967 report. The input as synchronous motor test data, p.19, was extrapolated to zero volts per IEEE Std 115-1995, 4.3.11 and Fig. 4.8. The Excel straight line fitting algorithm was used for test points in the low saturation region (50% rated volts and below) where the curve is a straight line:

Input as a syncronous motor test WH test report, page 19 y = 0.1821x - 58.296

200 300 400 500 600 700 800 900

Net F & W + core loss, kW

Sq ua re o f p ha se k

V

Square of volts, kV^2 Low saturation Linear (Low saturation)

The resulting F & W loss is 320 kW. The moment of inertia was then determined from 1967 retardation data, p. 25. The obtained value of 115.0 x 106 lb*ft2 is in perfect agreement with the original Westinghouse calculated value of 115.4 x 106 lb*ft2

The resulting equation for the loss calculation is:

Loss, kW = 4347 x dn/dt

The testing was performed in accordance with the requirements of IEEE 115, Section 4.4.

When testing hydraulic-turbine-driven generators, the machine under test is driven electrically from another unit.

For synchronous starting, the armatures of the driving and driven machines are connected together electrically while the machines are at rest. Approximately normal no-load full-voltage field current is applied to the driving machine and approximately 80% of normal no-load full-voltage field current is applied to the driven machine. The turbine of the driving machine is then started slowly and the two electrically connected machines are brought up to the desired speed of approximately 115% (~94 rpm).

Then the stator of the machine under test is separated from the driving machine stator. The unit under test is allowed to decelerate at a desired field setting.

The speed, armature and field currents, and terminal voltage of the machine under test were recorded by a Yokogawa DX208 data acquisition station. The recorded data file was then converted to a Microsoft Excel spreadsheet. The speed - time curves were made by the Microsoft Excel 3rd order polynomial fitting algorithm. The resulting equations are included in the report.

Unit 3 was used as the driving unit.

Turbine 4 was dewatered, with the runner at a minimum of four feet above the water line. Unit 3 was used to supply variable voltage and variable frequency power to unit 4.

The two units were interconnected by installing temporary jumpers and a NEC supplied breaker between them.

During the stray loss tests a three-phase short circuit was applied to the generator terminals. NEC supplied and temporarily connected a shorting breaker to the U4 generator terminals with cables.

The excitation of units U3 and U4 was obtained from the unit exciters powered from the 13.8 kV bus. Temporary jumpers were installed by NEC between excitation PPT’s and the 13.8 kV bus. A NEC exciter specialist controlled U3 and U4 exciters during all tests.

The table below gives a summary of the retardation runs performed on March 31 and April 1:

Big Bend U4 Retardation test summary table

Run # Time Field Amps

Line Volts

Line Amps dn/dt, rpm/sec

Loss kW

31-Mar at 81.8 rpm

FW3 0.06301 273.9

OC1 17:07:57 224.6 6536 0.07759 337.3

OC2 17:14:15 402.4 11473 0.11167 485.4

OC3 17:19:05 440.1 12346 0.1212 526.9

OC4 17:23:58 482.1 12969 0.1306 567.7

OC5 17:27:53 458.2 12720 0.1261 548.2

OC6 17:32:00 477 13125 0.1318 572.9

OC7 17:39:11 577.5 14.653 0.1552 674.7

FW4 0.0628 273.0

1-Apr max

RTD, C

FW5 0.06338 275.5

SC1 11:35:14 185.5 42.6 1184 0.0802 348.6

SC2 11:40:47 306.3 42.9 2005 0.1076 467.7

SC3 11:45:25 378 43.3 2497 0.1322 574.7

SC4 11:50:30 439.1 44.4 2909 0.1547 672.5

SC5 11:55:06 457.5 46.6 3055 0.1636 711.2

FW6 0.06268 272.5

Table 1

3.1. Friction and Windage Loss

The speed-time curves fitted as the 3rd order polynomials and their equations are shown in the Appendix (see Fig. 1). The results of the slope calculations at 81.8 rpm are given above in Table 1.

The first two of the test runs performed on 3/31/10 were done before actual beginning of the retardation tests to make sure the losses stabilized, the 3rd and the 4th were performed before and after the open circuit retardation tests. The fifth one and the sixth one were done before and after the short circuit retardation tests.

The average value of the friction and windage loss of 273.5 kW was used for calculating the core loss, and the average of 274 kW was used for calculating the short-circuit losses.

3.2. Core Loss

The open circuit retardation tests were done at seven different terminal voltages.

The speed-time curves fitted as the 3rd order polynomials and their equations are shown in the Appendix (see Fig. 2). The results of the total open circuit loss are given above in Table 1. The core loss interpolated from the fitted curve equation to 13800 Volts is 340.8 kW. The core loss was guaranteed not to exceed 369 kW.

Big Bend Unit #4 open circuit retardation tests of 3/31/10 Core loss at 13.8 kV is 340.8 kW. (13.8 kV)^2 = 190.44.

y = 0.00357x2 + 1.04408x + 12.50291

0 50 100 150 200 250

Squared stator voltage, kV^2

C or e lo ss , k

W

Core loss test Rated Volts Poly. (Core loss test)

Fig. 1

3.3. DC I2R Loss

DC I2R Loss at the original rating of 61.579 kVA, 2576 A and 75oC is calculated from the following winding phase resistance measurements taken on 3/10/10 after winding installation:

STATOR WINDING

RESISTANCE, Ohms, measured at 75 oF

STATOR WINDING

RESISTANCE, Ohms, corrected to 75 oC T1 – T4 0.00724 0.00867 T2 – T5 0.00725 0.00868 T3 – T6 0.00725 0.00868

Table 2

(I2R Loss at 75oC) = 25762 x (0.00867+0.00868+0.00868) / 1000 = 172.73 kW.

The guaranteed value of DCI2R Loss at 75oC is 189.06 kW.

At the new rated current of 2963 A the loss value is (I2R Loss at 75oC) = 29632 x (0.00867+0.00868+0.00868) / 1000 = 228.53 kW.

3.4. I2R and Stray Load Loss during retardation tests

These following phase resistances of the winding including breaker leads and the shorting breaker were measured before the retardation tests.

Winding phase-to-phase resistances, Ohms, including breaker and breaker leads at t0 = 37.2 oC

T1 – T2 0.016750 T2 – T3 0.016753 T3 – T1 0.016753 Average 0.016752

Table 3 The short circuit retardation tests were done at five different stator currents. The speed-time curves fitted as the 3rd order polynomials and their equations are shown in the Appendix (see Fig. 3). The results of the short circuit loss and the stray load loss calculations are given in Table 4.

The stray load loss was calculated as follows.

Line-to-neutral resistance including shorting cables and breaker = 0.008375 Ohms at 37.2 deg. C

F+W loss = 274.0 kW

Stator Total loss t, per max

RTD

Copper loss I2R SLL

Amps kW deg.C Ohms kA2 kW kW kW 1184 348.6 40 0.008461 1.402 74.6294 35.6 39.0 2005 467.7 40.1 0.008464 4.020 193.7372 102.1 91.7 2497 574.7 41.2 0.008498 6.235 300.6734 159.0 141.7 2909 672.5 42.5 0.008538 8.462 398.4809 216.8 181.7 3055 711.2 43.8 0.008578 9.333 437.1692 240.2 197.0

Table 4 The total short circuit phase (line to neutral) resistance is half of the phase-to-phase resistance of 0.016752 Ohm measured at 37.2 oC (See Table 3 above), i.e. Rph = .008375 Ohms. At a line current of 2497A during run #3 (See Table 4 above), the I2R loss due to the currents in the winding and the short circuit links is

)5.234( )5.234(3 t tRIRI ph 3 x 24972 x 0.008375(41.2 + 234.5)/(37.2 +234.5)

= 159.0 kW at 41.2 oC.

The stray load loss is determined by subtracting the friction & windage loss and

I2R loss from the total loss measured during short circuit retardation. For run #3 (see Table 4 above) we obtain 574.7 kW – 274.0 kW – 159.0 kW = 141.7 kW.

To find the stray load loss at rated conditions, a linear function of kW = 21.709 x kA2 was fit to the data of measured stray load loss vs. squared stator current

Big Bend U4 Losses during short-circuit retardation test of 4/1/2010.

At 61,579 kVA the stray load losses are 144 kW. (2576 A/1000)^2=6.6358

At 70,816 kVA the stray load losses are 190.6 kW.

y = 21.709x

0 1 2 3 4 5 6 7 8 9 10

Stator current, kA^2

Lo ss es , k

W

COMBINED "GROSS" I2R

AND STRAY LOAD LOSS I2R LOSS

(Includes shorting breaker and leads)

STRAY LOAD LOSS

Rated 2963 Amps

Fig. 2

This function was then evaluated at the rated current to find the stray loss. The stray load loss calculated from linear fit is 21.709 x (2576/1000)2 =190.6 kW at the new rating of 70,816 kVA, 2963 stator Amps. The stray load loss is 144.0 kW at the previous rating of 61.579 kVA, 2576 A. The guaranteed value for the stray load loss was 134.36 kW.

4. SHORT CIRCUIT SATURATION CURVE

The short-circuit saturation curve (see the saturation curves chart below) was obtained from the readings taken during the short-circuit retardation test.

5. OPEN CIRCUIT SATURATION TEST.

The test was done in accordance with IEEE 115-1995, Clause 4.2.5.

Under water turbine power the unit was brought to rated speed no load without field. The terminal voltage readings were taken for a few values of increasing field current. Stator voltage and field current were recorded at each field level

The terminal voltage (line-to-line) readings of all three phases were taken under constant conditions of excitation and speed.

Field Amps Average Line Volts 0 0 72 1577 99 2490 151 4200 196 5700 245 7200 299 8700 320 9300 344 10000 364 10600 394 11400 425 12100 463 12900 492 13500 535 14200 588 14900 649 15500 734 16400

Table5

6. ZERO POWER FACTOR SATURATION TEST

The test was done in accordance with IEEE 115-1995, Clause 4.2.10.

In order to obtain readings as close as possible to the rated current, unit under test was connected to the grid and operated at rated speed and voltage. The field was adjusted for maximum possible vars overexcited. The reactive kilovoltampere loading was redistributed among other machines. Under this condition the stator voltage and current and the field current were measured.

Due to system limitations it was not possible to achieve the rated stator current.

The closest test point had the stator current of 2370 A, line voltage of 13.766 kV and a field current of 290 ADC.

Field Amps Stator Amps Terminal Volts PF 527 300 13487 0.25 576 577 13458 0.09 618 780 13434 0.17 663 1005 13507 0.07 719 1270 13507 0.03 819 1700 13500 0.04 858 1880 13555 0.037 900 2050 13607 0.028 944 2209 13662 0.024 993 2370 13766 0.044

Table 6. Zero power factor test readings

COE Big Bend Unit 4 Saturation Curves and Potier Reactance determination from tests of April 2010

10000

11000

12000

13000

14000

15000

16000

17000

18000

0 100 200 300 400 500 600 700 800 900 1000 1100 1200

Field Amps

Li ne

V ol ts

Li ne

A m ps a d b c d'

Air-gap line

Zero PF rated current saturation

Open-circuit saturation

Short-circuit saturation

Test point at zero PF, 13766 V 2370 A

IFS

Fig. 3

The field current at uprated base armature current IFSI = 457 A from the short-circuit test. The field current at base voltage on the air-gap line IFG = 473 A from the open-circuit test.

The Potier reactance was determined from the open circuit saturation curve and from the above test point (point d in Fig. 3) per IEEE 115-1995, Section 5.2.2. To the left of d the length ad was laid off equal to the field current 457.5 ADC required for the armature current of 2370 A under sustained short circuit conditions (determined from the actual short circuit saturation curve). Trough a the line ab was drawn parallel to the air-gap line. The intersection of this line with the actual open circuit saturation curve locates point b.

The vertical distance (2500 V) from point b to line ad, expressed in per unit (2500/13800 = 0.181 p.u.), is equal to the product of the per unit Potier reactance, Xp , and per unit armature current (2370/2963 = 0.800 p.u.). The Potier reactance value is

0.226 p.u.).

The Potier reactance voltage was calculated for the rated current of 2963 A and point d’ was located. Using the rated Potier triangle, the zero power factor saturation curve was drawn up from 0 to 13800 V.

7. DETERMINATION OF DIRECT-AXIS SYNCHRONOUS

REACTANCE

The direct-axis synchronous reactance was calculated from the results of the open-circuit test and the short-circuit saturation test per IEEE-115, Section 10.3 and 10.4.4. It is equal to the ratio of the field current IFSI = 457 A at base armature current, from the short-circuit test, to the field current IFG = 473 A at base voltage on the air-gap line:

Xdu = IFSI/IFG = 457/473 = 0.966.

8. TESTS TO DETERMINE QUADRATURE-AXIS SYNCHRONOUS

REACTANCE

The quad-axis synchronous reactance was determined by the slip test method.

Upon completion of the retardation tests, the unit 4 turbine was filled with water. The U4 rotor was driven by its turbine at a speed slightly different from synchronous with the field open-circuited. Unit 3 was used as a source providing U4 stator winding with three-phase, rated frequency, positive sequence power at a voltage below the point on the open-circuit saturation curve where the curve deviates from the air-gap line.

The oscillograms of the armature current and the armature voltage were recorded, and the voltage across the open-circuit field winding was monitored.

The Unit 4 speed was 80.4 rpm and the Unit 3 speed was equal to the rated speed of 81.8 rpm. The minimum and maximum ratios of the armature voltage (red wave form) to the armature current (blue wave form) were obtained from oscillograms (see a sample below).

Fig. 4

The quadrature-axis synchronous reactance Xqu was determined per IEEE 115- 1995, Section 10.4.2.2 as:

max min max min

I I

U UXX duqu

The slip test was run at 2 levels of U3 field current and the average Xqu was calculated.

U3 Field Current (A) Min Max Xqu in p.u.

I 8.4 9.8 130 U 10 12.8 0.646 I 6.4 7.8 100 U 8.2 9.8 0.663

Average 0.654

Table 7

9. SUSTAINED SHORT CIRCUIT TESTS TO DETERMINE THE

NEGATIVE SEQUENCE REACTANCE

The negative sequence reactance has been determined by Method 3 per Section 10.5.4 of IEEE-115. Two phases were shorted together with a shorting link. Unit 4 generator was brought to rated speed by its turbine. A series of readings was taken at reduced excitation.

Voltage, E, current, I, and Watts, P, were recorded by Yokogawa WT-1600.

Fig. 5

The following base values were used to get the p.u. values: the base line current = 2963A, and the base single-phase power = 23605 W. The per unit negative sequence reactance was obtained from the following equation:

322 I

PX

Volts rms recorded

Amps rms recorded

Watts recorded

P.F.

recorded

I p.u. P p.u. X2

613 324 197772 0.994 0.1095 0.0084 0.4034 612 324 196949 0.994 0.1093 0.0083 0.4032

1213 639 769837 0.994 0.2156 0.0326 0.4051 1209 638 766651 0.994 0.2154 0.0325 0.4041 1793 953 1698647 0.994 0.3217 0.0720 0.4015 1793 952 1696405 0.994 0.3213 0.0719 0.4018 2375 1286 3035675 0.994 0.4341 0.1286 0.3939 2381 1286 3042795 0.994 0.4340 0.1289 0.3952 2956 1635 4803777 0.994 0.5519 0.2035 0.3858 2953 1634 4796646 0.994 0.5516 0.2032 0.3856 4274 2482 10542991 0.994 0.8378 0.4466 0.3674 4263 2486 10533088 0.994 0.8391 0.4462 0.3659

Table 8

The values of reactance were plotted as a function of the negative sequence current and extrapolated to the rated current. The value of negative sequence reactance at rated current is 0.35.

Big Bend Unit 4 Negative Sequence Reactance y = -0.0537x2 - 0.0028x + 0.4061

0.300

0.320

0.340

0.360

0.380

0.400

0.420

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

Line current, p.u.

Li ne

-to -li ne v ol ta ge , p

.u

Fig. 6

10. SUSTAINED SHORT CIRCUIT TESTS TO DETERMINE THE ZERO

SEQUENCE REACTANCE

The zero sequence reactance was determined by Method 3 per Section10.6.5 of IEEE-115. Two phases were shorted to ground. The generator was brought to rated speed, and field current adjusted to give a series of readings of stator current and voltage.

Fig. 7

Ea per unit value based on 7968 rated phase Volts In per unit value based on 2963 rated phase Amps

Field Amps

Ea, Volts read

In, Amps read

Ea, p.u. In, p.u. Xo

40 380 563 0.0477 0.1900 0.2510 100 741 1088 0.0930 0.3672 0.2533 150 1108 1623 0.1391 0.5478 0.2539 200 1423 2150 0.1786 0.7256 0.2461 250 1766 2725 0.2216 0.9197 0.2410

Table 9 The zero sequence reactance was obtained from the following equation:

n a

I E

X 0

Xo = 0.24 at the rated current of 2963 Amps.

11. SUDDEN SHORT CIRCUIT TESTS TO DETERMINE THE

TRANSIENT AND SUBTRANSIENT REACTANCES AND TIME

CONSTANTS

The machine was operated at rated speed and reduced voltage (30%, 35% and 40% rated) at the open terminals. The shorting breaker was closed, simultaneously short circuiting all three phases. Excitation remained constant. The machine operated in this condition until reaching steady state stator current. Oscillograms of the stator current were recorded throughout the transient and steady state operation. The transient and subtransient reactances and time constants were determined from the recorded oscillograms (see a sample below).

Big Bend Unit 4 3 Phase Sudden Short Circuit Test

0.0

0.1

1.0

10.0

0 4 8 12 16 20 24 28 32 Time in Cycles

A rm at ur e

SC

C ur re nt in p

.u . T''d

Terminal Volts in p.u 0.39 Line current in pu 0.42 Unsaturated/Test Saturated/Calc.

Transient reactance X'd in p.u. 0.43 0.381 Subtransient reactance X''d in p.u. 0.32 0.244 Transient time constant T'd, sec 1.27 Subtransient time constant T''d, sec 0.032

Table 10. Calculated reactances and time constants

The tested unsaturated values are in satisfactory agreement with those obtained in the course of the 1967 test. The saturated values were obtained by multiplying the testes unsaturated values by the saturated/unsaturated ratios determined from the 1967 tests.

12. DETERMINATION OF THE UNSATURATED QUADRATURE-

AXIS SUBTRANSIENT REACTANCE (X”q)

The quadrature-axis subtransient reactance is determined from the data obtained in the determination of the direct-axis subtransient reactance (X”d) and the negative sequence reactance (X2).

From "" 2 qd XXX the following is obtained: 47.0

32.0

39.0 2

2" d q X

XX

13. DETERMINATION OF THE OPEN-CIRCUIT TRANSIENT TIME

CONSTANT (t’do)

The open-circuit transient time constant is determined from the data obtained in the determination of the transient time constant (t’d) and transient and subtransient direct axis reactances (Xd and X’d).

From d d dd X

Xtt

' the following is obtained: sec85.2 43.0

966.027.1' d d dd X

Xtt

14. WAVEFORM DEVIATION FACTOR

The waveform of the voltage of each phase and line-to-line voltages was recorded with the machine operating at rated voltage and speed and open circuit.

Big Bend unit 4 Phase A Voltage Wave Form

0.000 0.002 0.004 0.006 0.008 0.010

Time, sec

PT

O ut pu t V ol ts

Test wave Sine wave

Fig 8

Big Bend unit 4 Phase A Voltage Wave Form Deviation

-10%

-8%

-6%

-4%

-2%

0%

2%

4%

6%

8%

10%

1 8 15 22 29 36 43 50 57 64 71 78 85 92 99

No. of point within half-cycle

W av e Fo rm D ev ia tio n, Fig 9

Big Bend Phase B Voltage Wave Form

0.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0 80.0 90.0

100.0

0.000 0.002 0.004 0.006 0.008 0.010 Time, sec

PT

O ut pu t V ol ts

Fig 10

Big Bend Phase B Voltage Wave Form Deviation

-10%

-8%

-6%

-4%

-2%

0%

2%

4%

6%

1 8 15 22 29 36 43 50 57 64 71 78 85 92 99 106

No. of point within half-cycle

W av ef or m d ev ia tio n

Fig 11

Big Bend Phase C Voltage Wave Form

0.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0 80.0 90.0

100.0

0.000 0.002 0.004 0.006 0.008 0.010 Time sec

PT

O ut pu t V ol ts

Fig 12

Big Bend Phase C Voltage Wave Form Deviation

-10%

-8%

-6%

-4%

-2%

0%

2%

4%

6%

1 8 15 22 29 36 43 50 57 64 71 78 85 92 99 106

No. of point within half-cycle

W av ef or m d ev ia tio n

Fig 13

15. TELEPHONE INFLUENCE FACTOR

The TIF was measured with Yokogawa WT-1600 (see a screen below). The generator was operated at no-load, rated voltage, rated speed.

Fig. 14

The following results were obtained:

Phase Line-to-neutral TIF

Phase Volts

A 13.20 7931 B 13.10 7938 C 12.40 7953

Average 12.90

Table 11

Phase Line-to-line TIF

Phase Volts

A-B 7.33 13734 B-C 7.41 13737 A-C 7.64 13733

Average balanced TIF

7.46

Table 12 Residual TIF is:

52.1046.790.12)()( 2222 FbalancedTIneutralTIFtolineFresidualTI

16. TEMPERATURE TEST (HEAT RUN)

The generator was tested per paragraph 6.2.1 Conventional Loading of IEEE 115.

at loads of 52,402 MVA, 55,153 MVA and 58,120 MVA. Higher test loads could not be achieved because of the grid limitations.

The above operating conditions were held constant until the machine reached constant temperature. Readings of the stator volts, amps, winding temperature (by RTD) and the coolant temperatures were taken by thermocouples installed as shown below in Fig. 16. The test continued until winding temperatures and the ambient temperature became constant within +/- 2% of the temperature rise value for three consecutive half-hourly readings (see Tables 1 & 2 in the Appendix for test data).

The measured results were extrapolated to the rated load of 70816 kVA:

Big Bend Unit 4 Temperature tests of 4/8/10 Stator Winding Temperature Rise

2000 2500 3000 3500 4000 4500 5000 5500

MVA^2

D eg re es

C

Rated 70,816 kVA

Fig. 15. Extrapolation of the test results to the new rating (MVA^2 = 5014).

The extrapolated value is 56.6 oC. The generator loads during heat run tests were relatively low and their range was relatively narrow. In 1967 Westinghouse tested unit 7 at 70,563 kVA, i.e. essentially at the new rating of 70,816 kVA. In 1967 the measured value of the stator winding temperature rise was 63.5oC.

Fig. 16

The reference (cold) field resistance Rb = 0.3172 Ohms at tb = 34oC was measured before the tests by the Kelvin bridge. At the end of the last heat run, the test resistance, Rt, = 0.3602 Ohms, was determined by measuring the voltage across slip rings (290 V) and dividing it by the field current (If = 805 A).

The field temperature, tt oC at the end of the last heat run was calculated from the hot field resistance value and the reference resistance by the following equation:

Ckt R

RR

tt o b b bt bt 4.70)5.23434(

3172.0 3172.03602.034)( where k = 234.5 for copper.

The rotor winding temperature rise is tt = 70.4oC – 26.1oC = 44.3oC at 58,120 kVA load.

The temperature rise of field winding is linearly proportional to field I2R loss. The Potier reactance obtained from the zero power factor test and the IEEE Std 115-1995, Eq.

5-11 were used to calculate the field current at the new generator rating:

Amps

IIIII

oo

FSIFSIFGFSFL

992)19.18cos457()19.18sin457472(240

)cos()sin(

The measured open-circuit saturation curve had to be extrapolated to obtain IFS.

The air temperature rise through the fan at the new rating was estimated as:

C lossTotal losswindageFrictiontt o statorfan 2

)203191352282( 2825.06.56

)_(5.0

The following equation calculates the field temperature rise above the temperature of the cooling air leaving the fan. It neglects the effects of the change in resistance (see IEEE Std 115-1995, Section 6.2.3.1, Equation 6-5).

The estimated field winding temperature rise at the new rating is:

Ctt I Itt o fant f

FL

fans 2.6623.44

This estimated value is an overstatement. The field temperature rise was directly measured by Westinghouse at 70,563 kVA, 0.96 PF. Actual field current was 891 Amps.

And the actual temperature rise was 52.8oC. Both the actual current and the actual rise differ substantially from the above calculated values. Extrapolated values are always less reliable than the measurements. The measured values of the field current (891 A) and temperature rise (52.8oC) should be used when they are needed.

17. OZONE CONCENTRATION TEST

During the temperature test the ozone concentration in the generator barrel was measured by the S-200L type Ozone Solutions monitor with a low range (0.001-0.5 ppm) display. At the end of the 3rd heat run the ozone concentration was 0.016 ppm. The allowable limit of 0.1 ppm was not exceeded.

APPENDIX

Fig.1. Friction and windage retardation test data.

Fig.2. Open-circuit retardation test data.

Fig.3. Short-circuit retardation test data.

Table 1. Heat run test data.

Table 2. Stator core temperatures.

Fig. 4. Big Bend generator ventilation system.

Bid Clarification Questions for NEC

Calibration certificates of the measuring instruments

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Fig. 4. Big Bend generator ventilation system