28.12.1 M.m.f. phasor diagram (Figure 28.20)

The e.m.f. E behind the leakage reactance X₁ requires a field m.m.f. F_c, read from the calculated or tested o.c.c. F_ar is the armature-reaction m.m.f. in rotor terms, obtained by using the equations in Section 28.7 or (if X₁ is known) by calculation from the tested s.c.c. as follows. To circulate stator current I on short circuit, excitation I_n is needed; to generate an e.m.f. to balance IX₁ drop, I_f2 is needed; hence the armature-reaction m.m.f. is F_af = I_f1 − I_f2. E_f is the excitation m.m.f. required for the load current I at terminal voltage V and p.f. angle φ. Figure 28.20(c) shows the diagram for a salient-pole machine with a leading p.f. load. As in Figure 28.14, ab is divided at c such that ac/ab = X_q/X_d to find point f.

28.12.2 The ANSI Potier reactance method (Figure 28.21)

ANSI Potier reactance method (ANSI/IEEE Publication 115, Section 4) requires the tested o.c. and z.p.f. characteristics, and is limited to machines that can be loaded for a z.p.f. test. In Figure 28.21(a), A is the rated-current short-circuit point and D is the rated-current rated-voltage point: to reach D the field current exceeds the rated-load excitation level. Draw DC = A0, and draw CF parallel to the airgap line 0BH. Drop FL perpendicular to DC, and draw triangle 0BAM similar to CFDL. Then FL is the Potier voltage drop IX_P, and DL is the armature-reaction m.m.f.

The argument is that for a given stator current the armature-reaction and leakage-reactance voltage drops are constant, but the latter requires more excitation at the higher levels of saturation.

0G, 0A and FH in Figure 28.20(b) are excitation currents as in (a), with 0G for rated voltage on the airgap line. GH is the total excitation required.

The z.p.f. test may have to be performed at less than rated stator current; the z.p.f.c. is then closer to the o.c.c., and DL and FL are smaller. Nevertheless, the Potier reactance is still considered to be X_p = FL/I up to rated value.

Given only the o.c.c. and s.c.c., and no facility for adequately loading the machine, one procedure is to measure the d axis subtransient reactance from a sudden-short-circuit test (or by the IEEE method below) and to use it as X_p.

The test method for X″_d in ANSI/IEEE 115, Clause 7.30.25, is to apply a voltage E of normal frequency to each pair of stator terminals in turn and to observe the current I with the rotor stationary. Let the three quotients of E/I be A, B and C. Then with E and I in per-unit values of rated phase voltage and current, X″_d = (A + B + C)/6 to an approximation. To avoid rotor overheating, the duration of the test should not exceed the maker’s recommendations (e.g. 0.2 p.u. for a time sufficient to read the meters).

28.12.3 Use of design calculation (Figure 28.22)

Methods can be refined by allowing for the rotor pole-to-pole leakage. From a knowledge of the airgap flux, the m.m.f.s required for the gap, stator teeth and stator core are calculated. The m.m.f. for the rotor pole and body are calculated from the gap flux and pole leakage. The component phasors are added as in the diagram, where the total rotor m.m.f. per pole bc is obtained from 0a (armature-reaction), Ob (gap, teeth and core) and ac (pole and body). Saliency can be allowed for by dividing 0a at d such that 0d/0a = X_q/X_d. Then the total rotor m.m.f. is bc′. This differs very little from bc, but the load angle δ is more accurate.

(1) Draw the envelopes abc and a′b′c′ of one phase-current oscillogram. Then aa′ is the double-amplitude of the prospective current at the instant t = 0 of short circuit. (The first current peak is slightly less than 1/2 aa′ because of the rapid subtransient decrement.) Taking aa′ as scaled in per-unit terms, the r.m.s. current is and

(2) Project the envelopes in the transient region, cb and c′b′, back respectively to d and d′, ignoring the initial rapid subtransient decrement (this cannot be done with great accuracy). Then and .

(3) Repeat steps (1) and (2) for the other two phases and derive the mean values of and .

(4) For a closer estimate, the equation relating the r.m.s. value of the a.c. short-circuit current I₁ to time may be used:

(28.46)

    where I_d = ee′/ is the sustained steady-state short-circuit current, and and are the transient and subtransient r.m s. currents respectively, corresponding to dd′ and aa′, respectively, at t = 0.

(5) Measure (e.g. in centimetres) the double-amplitude between the envelopes at each current peak and subtract ee′ from each. Plot the results as ordinates on a logarithmic scale, to a linear base of time. This gives the curve abc in Figure 28.24(b).

(6) Project cb by a straight line to d at t = 0. Then

    [d (cm)/] × current scale of oscillogram whence

    [If bc is not linear, the decay is not exponential and is not a constant. I′_d and T′_d can be estimated from a straight line drawn through chosen points on the curve bc where the transient current components are b and 0.368b (see IEC 34–4: 1985).]

(7) The subtransient r.m.s. current at t = 0 is I′′_d = I′_d + the rapidly decaying component represented by da. Thus

    and

    The intercepts between ab and db are drawn to extended current and time scales in the lower part of Figure 28.24(b). Point f on the ordinate scale corresponds to da.

(8) Time constants are obtained from the slopes of the current—time plots. At c on line dc the transient component at time t is represented by h on the logarithmic ordinate (centimetre) scale and by hc on the time scale. It is related to I′_d at t = 0 by

(28.47)

(28.48)

    where hc is in seconds and d and h are in centimetres. T′′_d is obtained similarly from the extended-scale plot in Figure 28.24(b). The .

(9) The procedure (5)–(8) is repeated for the other two phases and the mean values are obtained.

(10) The reactance values decrease with increasing short-circuit current (and therefore with increasing open-circuit voltage E_o). Rated-current values are obtained when at t = 0 the transient current I′_d is equal to rated current. A range of short-circuit tests spanning the expected X′_d will give a plot of reactance to a base of transient current I′_d from which the appropriate value can be found. A rated-voltage value, if required, is obtained by testing at 1 p.u. voltage for a small machine without a transformer, but the electromechanical forces on the stator endwindings would be excessive in a large generator. Tests up to 0.7 p.u. voltage simulate a fault on the h.v. side of the transformer in a generator—transformer unit and are more relevant to the service conditions of a large machine.

(11) The armature short-circuit time-constant T_a is determined from the decay of the d.c. components of stator current or from the decay of the a.c. component of induced field current. The latter method is simple: it requires only a log-linear plot of the a.c. component to a base of time, similar to the plot in Figure 28.24(b). The stator d.c. component is represented by the median line kl of the current envelopes in Figure 28.24(a). However, if there are significant even-order harmonics then the median is displaced and a waveform analysis is necessary to find the harmonic effect. This will occur in a salient pole machine that has no effective damper winding, and so is not common.

(1) to control the generator voltage accurately as slow changes of power and reactive loading occur;

(2) to limit the fluctuations of voltage when loads are suddenly imposed or removed;

(3) to maintain steady-state stability; and

(4) to maintain transient stability.

(1) To hold the generator voltage within specified limits, which may be between ±3%, and ±0.5% of the set voltage, the excitation system needs a high d.c. gain, but a moderate ceiling voltage is enough to supply the small and slow changes of excitation needed.

(2) To limit the fluctuation of voltage when load is suddenly applied or removed, a large, rapid and well damped response is needed. However a high initial response system, as defined above is often not necessary. For example, starting a large motor demands first a large reactive output from the generator, then increasing real power as the motor runs up. On a small generator or generator group, to avoid the voltage dip being too great and too prolonged, the generator needs a low X′_d, and the excitation must be kept high for perhaps several seconds. A high ceiling voltage and a response time of say 0.2 s is more useful than a high initial response system with a lower ceiling.

(3) Following a fault and its clearance on a high voltage power system, or a serious sudden loss of generation, a large and rapid response is needed to maintain transient stability. i.e. to restore the voltage and synchronising power flow to hold the several generators in step. Whether or not a high initial response system is essential depends on the particular circumstances, including the inertias of the generator sets and the post-fault reactances of the system. Very low frequency swinging, at 0.5 Hz or less, can occur between generating areas, requiring higher than rated excitation currents to be maintained for several seconds.

(4) Steady-state stability can be improved, in the sense that the generator can run safely close to, or even beyond, the fixed excitation stability limit, if the controller has no dead band, and is designed to have the desired speed of response without introducing a phase shift that causes positive feedback. Such feedback encourages oscillations of rotor angle, etc., which can become intolerable. This is more likely with high-reactance power lines. Long lines with series capacitance inserted to compensate for the line inductance introduce the possible hazard of subsynchronous resonance.^128–133 A minor disturbance, for example normal switching of lines, can cause transient current to flow at the line natural frequency f_n (hertz), often in the range 20–40 Hz. Torque is developed on a generator rotor at system frequency f_s−f_n (hertz). If this is close to a natural frequency of torsional oscillation of some part of a turbogenerator shaft system, the oscillation may increase, sustaining the subsynchronous current. Fatigue damage has occurred on a few turbogenerator sets in this way.

28.14.1 D.c. exciters

These have been superseded by brushless or static thyristor a.c. systems. The exciter was a d.c. generator coupled to the shaft of the synchronous machine, feeding its output to the main field through slip-rings. For high-speed generators of more than about 50 MW, the exciter had to be driven at a lower speed (typically 1000 or 750 rev/min) through gears, or separately by a motor, in order to avoid difficulties of construction and of commutation. On very-low-speed hydrogenerators a directly coupled exciter would be excessively large, so a higher-speed exciter, driven by a motor or perhaps a small water turbine, was used.

For small ratings, the exciter was shunt excited; however, most were separately excited from a directly coupled shunt-excited pilot exciter. Control of the generator excitation was provided by controlling the field current of the main exciter.

28.14.2 A.c. exciters with static rectifiers

Satisfactory service experience with brushless and static thyristor systems has made these early (diode) systems obsolescent. Many remain in service, but many have been replaced by modern equipment. The advent of solid-state rectifiers made it possible to avoid commutators by using an a.c. exciter, directly coupled to the generator and feeding its output via floor-mounted rectifiers to the generator held winding through slip-rings. The exciter can operate at any economically convenient frequency, usually between 50 and 250 Hz, and the system is suitable for generators up to the largest ratings.

The diode cubicles may be cooled by natural convection or forced air flow. Alternatively, especially for large ratings. the diodes may be mounted on water-cooled bus-bars; this greatly reduces the size of the cubicles so that they can, if desired, be mounted on the sides of the main exciter frame, thus avoiding the need for long runs of a.c. and d.c. busbars or cables.

The main exciter field is supplied by an a.c. pilot exciter, often a permanent-magnet generator. The excitation of the main generator is controlled by controlling the main exciter field current via the automatic voltage regulator.

The main exciter is usually three-phase, and the diodes are connected in the six-arm bridge circuit, usually with a fuse in series with each diode to interrupt the fault current should a diode break down (which almost always causes it to conduct in both directions, i.e. to act as a short circuit). Diodes are available with current ratings up to 1000 A mean d.c, and peak inverse voltage up to 5kV (but not both together in one diode). For other than small ratings, each bridge arm has several diodes in parallel; all the diodes are fused and have sufficient current margin to enable the bridge to carry full-load excitation continuously with one or more of the diodes failed and isolated by their fuses. Hence the generator can remain in service until maintenance can be carried out conveniently. Some installations on large generators had a.c. and d.c. isolators to allow parts of the bridge to be worked on without taking the generator out of service.

The diodes must be able to withstand induced transient currents and voltages resulting from system short circuits, asynchronous running, pole-slip and faulty synchronising, as well as from faults in the excitation system itself. Their continuous duty rating must leave some margin for imperfect sharing between parallel paths and for the possible loss of one or more paths, as noted above.

The fuse characteristic is co-ordinated with that of the diode so that fuses should not blow unless a diode fails or there is a short circuit on the d.c. output. The fuses must clear the fault current under the most onerous condition, which is usually that of a failure with the exciter at ceiling voltage. Fuse blowing is easily indicated by a microswitch operated by a striker pin that is ejected from an indicator fuse in parallel with the main fuse.

28.14.3 Brushless excitation

The a.c. exciter has a rotating armature with three or more phases and a stationary field system. It usually is designed for a frequency of 0–5 times the power system frequency. The pilot exciter, when there is one, is usually a permanent magnet generator operating at around 6–8 times system frequency. The diodes and fuses are mounted on the rotor, and the rectified output is led directly to the generator field winding without need of brushes and slip-rings. The diodes are mounted on well-ventilated heat sinks, and special designs of fuse are used to withstand the centrifugal force on the fusible link.

On units small enough to require only one diode and fuse per arm, failure of one diode or fuse leaves the exciter with one phase unloaded; exciters are usually designed to supply full-load excitation in this condition without damage, so that the generator can remain in service until the fault can be repaired conveniently. However, experience shows that the failure rate of diodes is extremely low and that more often fuse links fail mechanically. Hence some makers supply salient-pole generators, up to say 25 MW, with no fuses at all, but use generously rated diodes to provide a large margin. These generators would use up to three diodes in parallel per bridge arm. For turbogenerators up to about 70 MW, some designs use two diodes in series—each of full duty, with one, two or more series pairs in parallel per bridge arm—and no fuses. On large units, redundant parallel paths, individually fused, are provided as in static equipments.

For units that use fuses, the striker-pin indicator type can still be used, the pin being observed by causing it to interrupt a light beam falling on to a photoelectric cell, or it can be observed visually with a stroboscope. Alternatively a neon lamp is connected across the fuse and glows when the fuse blows.

When diodes are in parallel, whether fused or not, if one becomes open circuit the system will continue to function apparently normally unless the remaining diodes are overloaded and eventually fail also. If an unfused diode fails by short circuiting, the short-circuit current in the exciter armature induces fundamental (exciter) frequency current in the exciter field winding. This can be detected and used to trip the set before serious damage is done. Another method of detection is to use stationary pick-up coils to see whether the diode connections are carrying current as they should as they pass the coils.

More elaborate indication, perhaps coupled with measurements of current and voltage and indication of earth fault, can be arranged by telemetry, but the telemetry may be less reliable than the diodes. Frequently instrument slip-rings are used, with solenoid-operated brushes that make contact only when readings are required.

The diodes, and fuses too if they are used, must be rated for the normal duty, including field forcing, and to withstand the abnormal conditions noted in Section 28.14.2.

28.14.4 Thyristor excitation

Direct control of the field current of the synchronous machine by thyristors gives quicker response than can be obtained by controlling the exciter field current, because the time delay in the exciter is eliminated and the machine field current can be forced down by using the thyristors to reverse the machine field voltage. (By contrast, with a diode bridge, the machine field voltage can only be reduced to zero by reversing the exciter field voltage.) This is valuable for generators and synchronous compensators in certain power-system situations: for example, to minimise the voltage dip caused by large and possibly frequent load changes; to maintain transient stability of a generator under short-circuit conditions on the power system; to enable a synchronous compensator to maintain close control of the system voltage by rapid change in its reactive load, to minimise the voltage rise following sudden load rejection; to reduce more quickly the current resulting from a fault between the generator and its nearest protective circuit-breaker when field suppression is the only means available.

The synchronous machine requires slip-rings and brushes, and this is a disadvantage, especially for large machines for which brushgear maintenance may become a significant inconvenience.

The excitation power may be supplied by direct coupled main and pilot exciters, the main exciter working continuously at ceiling voltage. This makes the power supply independent of voltage fluctuations on the power system. Usually though the excitation is supplied from the generator terminals through a step down transformer. This is usually designed to provide the required ceiling voltage when its primary voltage is reduced to about 60% of normal. This ensures that some field forcing can be done even when the power system voltage is depressed by a fault. It does subject the generator field winding to a rather high peak voltage when the system voltage is normal. A lower ceiling is adequate if power-rated current transformers are added in order to derive some excitation from the machine output current, so boosting excitation during the fault. The set is shortened by the absence of the exciter, and this may save costs on foundations and building. For very-low-speed generators the scheme may well be cheaper than a direct-coupled exciter and diodes.

Some excitation systems use diodes and thyristors in combination, e.g. in a full-wave half-controlled bridge circuit. One patented scheme uses a full-wave diode bridge with thyristor ‘trimmer’ control fed from a special excitation winding on the generator stator and from compounding current transformers.

Rotating thyristor systems have not yet been developed commercially, mainly because of technical difficulties in transferring control signals from the stationary equipment and problems concerning the reliability of rotating control circuitry.

28.14.5 Excitation systems circuits

Typical systems are shown in Figure 28.26.

Scheme (a) is capable of the most rapid response. In (b) and (c) some delay is introduced by the exciter time-constant; consequently a high exciter ceiling voltage and a large output from the pilot exciter are needed to obtain a more rapid response.

28.14.6 Excitation control

When a generator operates alone the excitation is controlled to maintain the steady-state voltage within the necessary limits, and to prevent unacceptable variations of voltage when large and sudden changes of load occur. Generators running in parallel may need additional control signals to share the total reactive load correctly between them. In an interconnected system, control of steady state and transient stability is a vital duty. Manual control of the excitation is inadequate, and automatic control is provided.

Electromechanical voltage regulators, in use only in old installations, may be of the carbon-pile, vibratory-contact (Tirrill) or rolling-sector (Brown Boveri) type. These have been superseded, initially by magnetic amplifiers, with or without amplidynes, and now by solid-state control systems using transistor amplifiers at low power levels, with thyristors for field-circuit power. The new systems are continuously acting (i.e. have no dead band) and can be arranged to respond to many control signals besides that from the terminal voltage, so the term ‘automatic excitation controller’ is more logical than ‘automatic voltage regulator’ (a.v.r.). Power supply for the control circuits is derived from the machine terminals or the pilot exciter.

28.14.7 Basic principles of voltage control

A direct voltage proportional to the generator average terminal voltage is derived via voltage transformers and a diode rectifier circuit. This voltage is compared with a stable reference voltage generated within the regulator. Any difference (the ‘error voltage’) is amplified and used to control the firing of a thyristor circuit which supplies the excitation, either to the field of the synchronous machine or to its main exciter-field winding. Thus the excitation is raised or lowered to restore the machine voltage to the desired level and the error voltage returns to near zero. The set level is obtained by adjusting the proportion of the machine voltage that is compared with the reference voltage or by adjusting the reference voltage itself. The basic circuit (Figure 28.27) is incorporated in Figures 28.28 and 28.29.

28.14.8 Additional control features

28.14.9 Overall voltage response

The overall voltage response is defined in terms of steady-state and transient behaviour with the generator on open circuit and on no load, and under the control of the excitation system. For a generator operating alone, the following conditions may be relevant, their importance depending on the duty required of the generator:

Conditions of importance under transient conditions are the voltage rise and recovery time of the generator voltage when load is suddenly removed, and the voltage dip and recovery time when a large motor is switched on to the generator terminals. For a generator running alone, BS 4999-140:1987, ‘Specification for voltage regulation and parallel operation of a.c. synchronous generators’, specifies various grades of voltage regulation, for steady state and transient conditions. Accuracy of voltage control may conform to ±1% or ±2.5% or ±5%. Voltage dip must not exceed 0.15 p.u. when a current of 0.35 or 0.6 or 1.0 p.u. of rated generator current is suddenly demanded. The voltage is required to recover to 0.94 or 0.97 p.u. within 1.5 or 1.0 or 0.5 s. Values of the temporary voltage rise that occur when rated load at p.f. 0.8 is thrown off are specified with a range 0.35–0.15 p.u. The more severe the conditions the more powerful the a.v.r. control must be. Also, the lower must be the generator X″_d and X′_d in order to reduce the immediate fall or rise in voltage that the a.v.r. cannot affect. There is a consequential increase in the short-circuit current and in the generator frame size; both these increases raise the cost of the generator and, perhaps, of its switchgear. The response under transient conditions is a convenient way of expressing the overall performance and of testing it during commissioning. The terms used are

Normally, either V₁ or V₂ is the rated voltage V_r, depending on whether a ‘step-up’ or a ‘step-down’ change is being tested. The voltage-time curve is a well-damped transient, settling at V₂ after a few oscillations. Typical values of the quantities defined above are

For a given step change (V₂ − V₁), t₁ is reduced by increasing the ceiling voltage V_c, by increasing the excitation system gain, and/or by reducing the system time constants. The parameters of the generator and exciter cannot be changed once the machines are made, but the a.v.r. parameters are designed to be adjustable. Changes that reduce t₁ will increase the overshoot v and the settling time t₂; hence settings of a.v.r. gain, time constants and feedback signals (if any, e.g. exciter voltage or exciter field current) are calculated to achieve the desired compromise, and performance is checked over a range of values of step change during commissioning (say steps of 1%, 5%, 10% and 20% of V_r).

When generators operate in parallel, as most do, the excitation systems must be designed and adjusted to achieve the best compromise between highly accurate voltage control, steady-state stability, and transient stability following a system disturbance. Thus some step-change tests, or tests by injecting low-frequency sinusoidal voltage into the reference circuit, are desirable, to confirm the calculated performance.

See IEEE Standard 421A, ‘Guide for identification, testing and evaluation of the dynamic performance of excitation control systems’.

28.14.10 Digital control¹⁷⁰

Increasing use is being made of a dedicated microprocessor to replace the analogue control system. The input signals (voltage, current, MW, MVAR, etc.) are converted to digital form, and the processor is programmed to respond to these to provide an analogue signal to the firing circuits. The characteristics needed in the excitation controller are reproduced digitally in the processor, and limits of relevant quantities are introduced. The whole can be set up and tested in the works, so that site commissioning is simpler and quicker. In service, settings do not drift or suffer from poor rheostat contacts, but can be readily changed, even while the set is in service, to suit changed operating conditions. If a component does fail, a new card can be inserted without repeating the commissioning tests.

There is a delay of up to 10 ms while the microprocessor scans the input signals and readjusts its output signal, if necessary, but this is very small compared with the machine time constants.

The thyristor firing pulses may also be generated digitally. Development is proceeding of adaptive controllers that will automatically tune their characteristics to suit the operating conditions.¹⁶⁹

28.15.1 Main dimensions

The output coefficient C ranges typically from about 0.5 MVA s/m³ for a rating of 20 MVA to about 2.0 MVA s/m³ for a 1000 MVA unit. For 3000 rev/min machines these figures correspond to D²L of 0.8 and 10 m³, respectively. Economic diameters for these outputs range from approximately 0.75 m to 1.3 m, the latter being a limit set by centrifugal stresses in the endrings and in the rotor teeth. Hence outputs range from approximately 17 to about 170 MVA per metre of core length. The higher values of output coefficient are made possible by enhanced cooling techniques. Typical dimensions for a 660 MW two-pole hydrogen-cooled machine may be: rotor diameter, 1.15 m; core length, 6.8 m; overall shaft length, 13.5 m; core outside diameter, 2.7 m; outer casing, 4.8 m in diameter and 10.3 m long; total weight, 480 t.

28.15.2 Rotor body

The output available from a turbogenerator is largely determined by the excitation m.m.f. that can be carried on the rotor with acceptable winding temperatures. The high centrifugal stresses make cylindrical (i.e. non-salient-pole) construction essential.^92,93 Within the chosen diameter the number, shape, size and spacing of the winding slots have to be optimised to obtain the maximum m.m.f. capability with acceptable stresses in the teeth and slot wedges, with adequate insulation, with acceptable magnetic flux densities and with ducts for ventilation that enable temperature guarantees to be met. For air-cooled machines of medium output the manufacturing simplicity afforded by parallel-sided slots and solid copper conductors of rectangular cross-section may outweigh the loss of optimum performance and provide the cheapest design. For larger ratings tapered slots are used to accommodate more copper, while giving approximately constant mechanical stress and magnetic flux density along the radial length of the teeth.

A rotor is forged from a single steel ingot, the largest of which approach 500 t in weight; this would produce a rotor weighing 250 t, enough for a four-pole machine of about 1250 MW at 1500 rev/min. The forgings contain the alloying elements nickel, chromium, molybdenum and vanadium; according to size and speed, ultimate tensile strengths of the forgings range from 650 to 800 MN/m², while their 0.2% proof stresses range from 550 to 700 MN/m². The forgings are inspected with ultrasonics and magnetic-particle ink before use. Many generator makers now rely on these examinations and do not bore the forging axially along its centre line, except for large forgings or if ultrasonics reveals defects that can be removed by boring.

The endwindings (the parts projecting beyond the ends of the slots) must be supported against centrifugal forces by endrings (retaining rings) from which they are insulated by, for example, resin bonded fibreglass or aramid paper (Nomex) or combinations of synthetic insulating sheets. The endrings are steel forgings, usually shrunk on to the ends of the rotor body. In some older designs they were shrunk on to discs that were shrunk on to the shaft outboard of the windings, and they were not tight on the rotor body.

From the 1950s until 1982, the endrings¹⁰⁶,¹⁰⁷ were of austenitic steel (18% Mn, 4–5% Cr, 0.3% C) warm worked to give high strength, up to 1100 MN/m² proof stress and 1220 MN/m² ultimate in the highest grade. This alloy is very susceptible to stress corrosion if it gets wet, e.g. by condensation from moist air or leakage of cooler water. In 1982 an austenitic alloy became available with 18% Mn, 18% Cr;¹⁰⁸,¹⁰⁹ this is not subject to stress corrosion under any likely operating conditions, and has mechanical properties up to 1200 MN/m² proof stress (0.2% strain) and 1300 MN/m² ultimate strength. Endrings large enough for 1500 MVA two-pole or four-pole generators can be obtained, and higher strengths have been developed. Many endrings of the older alloy have been replaced, after some years in service, using the 18–18 material.¹⁰⁶,¹¹⁶

Rotor vibration at running speed must be low—typically about 50 μm peak-to-peak measured on the shaft near the bearings, though up to twice this is commercially acceptable. Hence balance weights must be carefully positioned, axially as well as circumferentially, and the design of the rotor, its bearings and its supports must ensure that its critical speeds are sufficiently far from rated speed. Small 3000 rev/min rotors will have one critical speed below 3000 rev/min, say about 1700–2000 rev/min, but as ratings (and therefore the bearing span) increase, two or even three criticals will occur below 3000 rev/min (typically around 650, 1750 and 2500 rev/min). When the rotor is coupled to the turbine, the critical speeds are usually raised slightly, so that behaviour in both the coupled and the uncoupled condition must be acceptable, for site running and works testing (without the turbine) respectively.

Electrical faults on the power system or on the machine itself produce abnormally high oscillatory torques on the rotor, at system frequency and often at twice this frequency also. Where series capacitance is used in long lines to compensate for inductive reactance drop, electrical oscillation at the natural frequency f_n may cause torques on generator rotors at system frequency minus f_n. These torques may resonate with torsional natural frequencies of the shafts and produce unacceptable fatigue stresses. From all these causes, complex torsional oscillations develop in the shaft system, with components determined by the inertias and stiffnesses of the several shafts of the turbine, generator and exciter. The shaft dimensions must be chosen to avoid a serious loss of fatigue life during such incidents as well as to satisfy the critical speed criteria mentioned above. There are many articles on the subject; references 128–135 are a few of these.

Bearings are of the white-metalled cylindrical type, with forced oil lubrication and, except on small sets, high-pressure oil jacking also. Jacking allows the set to run slowly (typically 3–20 rev/min) on the turning gear to cool off the turbine and generator rotors before the unit is finally stopped. Without this, the rotors would bend because temperature gradients would occur across the diameter of each rotor, and vibration would occur on the next run-up.

All turbogenerator rotors bend under their own weight, and with two-pole rotors for outputs more than about 30 MW the amount of bend would be significantly more when the pole axis is horizontal than when it is vertical. Hence a vibration would occur at a frequency corresponding to twice the running speed; it is caused by the changing stiffness of the rotor in the vertical plane, and therefore cannot be removed by mechanical balancing. The stiffnesses about the polar (direct) axis and the slot (quadrature) axis must be made as nearly equal as possible under running conditions (when centrifugal force on the windings increases the stiffness in the plane of the quadrature axis). This is done either by cutting axial slots along the pole areas (and filling them with magnetic steel if necessary to avoid magnetic saturation) or by cutting narrow arcuate grooves circumferentially across the poles, sufficient grooves being spaced down the length of the rotor to reduce the stiffness to match that in the quadrature plane.^29,93

28.15.3 Rotor winding

Excitation currents range from say 400 A at 200 V for a 30 MW generator to 5.7 kA at 640 V for a 1000 MW two-pole machine and up to around 7 kA at 650 V for a 1500 MVA machine. For rotors using indirect cooling each coil is wound with a continuous length of copper strap, bent on edge at the four corners. The copper contains about 0.1% silver and is hard drawn to increase its strength and so avoid the coil-shortening effect that occurred with plain soft copper as a result of heating while part of the copper was prevented (by centrifugal force) from expanding axially.

Directly cooled coils are usually made of larger section conductors of silver-bearing copper containing grooves and holes to provide gas passages. Half-turns are brazed together in the end regions after they have been positioned in the rotor slots. The slots may be parallel-sided, or tapered to contain more copper without increased tooth stress.

Insulation between turns is usually provided by interleaves of resin-bonded glass fabric or some other synthetic material. The coils are insulated from the rotor body by U-or L-shaped troughs of resin-bonded fibreglass, Nomex or melamine, or combinations of such materials. Similar insulating strips insulate the top conductor from the slot wedge; in direct-cooled rotors the strips must have through-holes to allow the cooling gas to escape from the rotor; they must be thick enough to provide adequate creepage distance to withstand the specified h.v. tests.

The end-windings are packed, partly or wholly, with blocks of insulating material to avoid distortion and the consequent risk of short circuits between turns.

28.15.4 Stator core

The stator core is built up of segments of electrical sheet steel, usually 2–3% silicon, 0.35 mm thick, cold rolled and non-oriented. To minimise weight, the core is worked at the highest flux density consistent with reasonable losses. In a two-pole machine the magnetic force across the airgap subjects the core to an elliptical distortion that rotates with the rotor, so producing a double-frequency (2f) vibration. The core depth must be chosen so that its natural frequency of vibration in this elliptical mode is well away from 2f; usually 3f or more is practicable without excessive depth of core. Grain-oriented steel has better permeability and lower losses than non-oriented steel, but as it has a lower modulus of elasticity its other advantages cannot be realised without accepting higher vibration. This, and its higher cost, severely limit its use in turbogenerators.

28.15.5 Stator casing

Air-cooled machines may have bearing pedestals on a bedplate; the stator frame then merely supports the core and forms the ventilation enclosure. Alternatively, it may be a more rigid box frame with the rotor bearings carried in end-brackets.

A hydrogen-cooled machine must have a totally enclosed and gas-tight construction; the end-bracket bearing arrangement is adopted to minimise the bearing span and to raise the critical speeds. Hence the frame must be rigid enough to provide proper support for the bearings and to contain the gas pressure that might occur in the unlikely event of a hydrogen—air explosion inside the frame. This could produce pressures up to about 1400 kN/m²; therefore this pressure, rather than the continuous working pressure of hydrogen, becomes the design criterion.

In large two-pole machines (which are invariably hydrogen cooled) some form of flexible mounting is needed between the core and the casing, and the casing should not have any natural frequency near to 2f. This is to avoid excessive magnetic noise and the risk of unacceptable vibration on the casing, coolers or pipework.

Where transport facilities are inadequate for handling the complete stator, the core and windings must be made separately from the outer casing, separately transported, and assembled on site before the rotor is inserted. By contrast, smaller machines can be transported complete to some sites, with the rotor clamped in temporary supports; this arrangement facilitates erection.

28.15.6 Stator winding

For small machines the voltage is usually fixed at a standard network voltage (e.g. 6.6 or 11 kV), but for large machines, where a generator-transformer is used, the designer has a free choice. A high voltage avoids difficulties due to high currents, but valuable space in the slots has to be sacrificed to insulation; a compromise is thus about 15 kV for 100 MW and 200 MW machines, and up to 22 or 25 kV for the larger sets. Even so, generators of more than about 50 MW rating will have two circuits in parallel per phase; for more than 1000 MW it may be necessary to use special winding arrangements to have four parallel paths in a two-pole machine. In four-pole generators, four circuits occur naturally and can be in parallel or in series-parallel.

The winding is of the two-layer basket type, almost always with integral slots per phase per pole, as described in Sections 28.3 and 28.4. In the slots²⁷ and in the endwindings¹⁰¹ the coils must be supported to resist electromagnetic forces.⁹³ These are, on normal load, continuous though fairly low vibratory forces at twice supply frequency. When a short circuit occurs close to the generator, transient oscillatory forces occur, 50–100 or more times greater than those on load.

In the slots the forces are radial, directed towards the bottom of the slot, except where different phases occupy the same slot; there the force on the top conductor is towards the wedge for part of each cycle. To support the coils along the whole core length, conformable packing strips of, for example, resin impregnated polyester fleece are placed beneath and between the coil sides. Slot wedges are fitted that apply a known radial load to the coils, greater than the electromagnetic forces. Packing may be fitted down the sides of the coils, and there may be a fibreglass ripple spring between the wedge and the top coil side to take up any small shrinkage in service.^27,29,93

The endwindings are secured to a strong structure of insulating materials,^27,29,93,100 fibreglass rings carried on brackets of resin-bonded wood laminate have been used very widely. For larger ratings a solid cone of filament-wound resin bonded fibreglass is used for greater strength and long-term rigidity. The coils are bedded to the structure with conformable packing material. Usually the slot and endwinding packings are cured while the coils are held by temporary wedges and clamps, which are then replaced by the permanent ones. The complete structure is bolted to the end of the core; some axial movement may be allowed, to accommodate expansion of the coils relative to the core.

28.15.7 Cooling

Efficiencies are between 96.5% and 99%, increasing with the rated output. However, the losses will be 0.5–15 MW, appearing as heat that must be removed by circulation of an appropriate cooling medium: oil for removing bearing-friction losses, and air, hydrogen or water for other losses. Details of the cooling media used for stator and rotor windings are given in Table 28.6. The heat transfer coefficients are typical, but depend considerably on velocity and duct size.

Traditionally, indirect air cooling was adopted for outputs up to 70 MVA, and indirect hydrogen cooling for outputs above about 50 MVA. Direct hydrogen cooling of rotors developed rapidly for ratings from about 100 MVA up to the largest, about 1500 MVA, that have been built. Some manufacturers used hydrogen cooled stator coils up to about 650 MVA: others adopted water cooling above about 150 MVA where indirect cooling was becoming difficult. In recent years cheap and simple designs have been developed up to 200 MVA rating using air cooling, direct in the rotor winding and indirect for the stator.

28.17.1 Introduction^{26–28, 30, 53, 138–156}

The design of hydrogenerators is determined mainly by mechanical considerations. Outputs range from less than 1 MVA to over 800 MVA, and speeds from 50 to 1000 rev/min, depending on the water head available and the type and size of the turbine. The low speeds require the generators to be physically large, and it is often necessary to transport them to site in sections. The inertia required in the set is determined by turbine governing or speed regulation requirements, or by the transient stability of the associated power system. The turbine contributes little flywheel effect, so the generator inertia must often be more than that of a design that satisfies the electrical specification in the least expensive way. The diameter must be increased, or in extreme cases a separate flywheel coupled to the shaft. This is often the best arrangement in fairly small horizontal shaft units.

A water turbine runs up to a high overspeed when load is suddenly removed, because the flow of water cannot be suddenly stopped without causing a high and probably damaging rise of pressure at the turbine gates or valves. The ratio of overspeed to normal speed is, approximately, for impulse turbines (Pelton) 1.7 to 1.9, for reaction (Francis) 1.8 to 2.1, and for propeller (Kaplan) 2 to 2.2. If the governor should fail during load rejectien, a Kaplan turbine could run up to 3 times normal speed.

The rotors must be designed to be safe at overspeed, where the factor of safety on the proof stress of the material used is normally not less than 1.5. A figure closer to 1.1 is acceptable for the very rare runaway condition of a Kaplan unit.

The first critical speed is required to be above the overspeed.

28.17.2 Construction

28.17.3 Cooling

Except for small machines, closed-circuit air cooling is used. Most hydrogenerators rotate in one direction only, and radial-flow or axial-flow fans mounted on the shaft are commonly used. The heat is removed by air-to-water heat exchangers mounted on the back of the stator frame. About 4 m³/s of air is required per kilowatt of loss removed. In cold climates, some of the hot air may be taken from the machine and used for heating the station.

Reversing sets for pumped storage would have to use radially bladed fans, which have poor performance and efficiency. For these machines, and for highly rated machines needing carefully controlled ventilation, several motor driven fans are used.

Hydrogen cooling has not been applied to hydrogenerators, mainly because of the cost and practical difficulties of making an explosion-proof casing. However, when inertia is not a controlling factor, a useful reduction in physical size can be achieved by water cooling the stator or rotor windings, or both.²⁶ Water cooling of only the stator winding introduces a multiplicity of joints which must not leak, but greatly increases the current capacity, or conversely allows the slot size to be reduced for a given rating; this slightly reduces the outside diameter of the core and significantly reduces the leakage reactance.

Water cooling of the rotor winding introduces more constructional difficulties and affects the design more profoundly. The excitation capability is increased without risk of exceeding the specified temperature limits, so a higher short-circuit ratio (longer airgap) is possible, improving the underexcited (line charging) capability. Alternatively a smaller machine with higher electric loading is possible with the stator design suitably adjusted. However, it may be necessary to adopt a size larger than the smallest determined from purely thermal considerations. This larger size may be needed to attain the desired inertia constant, or the lower capitalised cost of its lower losses may be enough to offset the lower first cost of the smaller frame. In some circumstances, water-cooled stator and rotor windings may be economically justified at ratings as low as 150 MVA; conversely, water-cooled stator windings with some form of improved air cooling for the rotor may be preferred on grounds of adequacy and simplicity for ratings as high as 700 MW.

28.17.4 Excitation

Vertical-shaft generators may use main and pilot a.c. exciters mounted above the generator, or separate motor-driven exciters. These often have a flywheel to maintain the exciter speed and output during momentary interruptions or reduced voltage of the motor supply.

Static thyristor equipment is now more often used, because it simplifies the mechanical arrangement of the unit, reduces the height, and avoids the possibility of an unacceptable run-out at the top of a tall shaft assembly. Thyristors also have the advantage of inherently high response, which is valuable when the generator feeds along transmission lines.

Slip-rings and brushgear are rarely troublesome, as the peripheral speed can be fairly low; 40 to 50 m/s is a usual limit.

Horizontal-shaft generators may use brushless exciters, overhung or two bearing depending on size and speed. Again static thyristors simplify the mechanical layout, especially if the generator has a turbine at each end.

28.17.5 Pumped storage units

Pumped storage units were originally installed as peak-levelling units, running as generators at times of high system load, and as motors pumping water up to the top reservoir during light-load periods. A unidirectional set has separate pump and turbine, whereas a reversing set uses the same hydraulic machine either as a pump or as a turbine, depending on the direction of rotation.

The original purpose has been extended to provide spinning reserve and to deliver power into the system at a ‘few seconds’ notice to assist in maintaining stability if other generation, or a system interconnection, is suddenly lost. This introduces particular problems of thermal cycling and mechanical fatigue, especially with reversing sets, which may be required to go from full-speed pumping to reversed full-speed generating within seconds, and to do this perhaps several times a day. For reversing units, separately driven fans are usually provided because rotor-mounted fans designed for both directions of rotation have low efficiency. Water cooling may be applied, as for generators, and may be particularly valuable for the damper cage if this is used for ‘induction-motor’ starting.

Figure 28.33 shows a cross-section of one half of the motor/generator of a reversible pumped-storage unit rated at 330 MVA, 18 kV, 0.95 p.f., 500 rev/min. The design at this output and speed approaches the limit achievable with present materials and air cooling. The outside diameter of the stator core is 6.2 m, the rotor diameter 4.5 m and the active core length 3.6 m. The mechanical design of the rotor is dominated by centrifugal stresses and the fatigue effects of reversals. The rotating parts weigh about 440 t, and the hydraulic thrust raises the load on the thrust bearing to almost 600 t.

28.18.1 Applications

Salient-pole generators are used both for stand-by applications and to continuously supply power. They can therefore be connected to public power systems, incorporated in marine installations (ships and oil rigs) or in a great variety of industrial plants. Where a public system requires unit ratings up to about 40 MVA, high speed, medium speed or low speed diesel sets can be installed more cheaply and more quickly than a gas or steam turbine set with its boiler and auxiliaries. The internal combustion engine set can deliver full load within a few minutes of starting, which can be done remotely if necessary; it will respond rapidly to changing load demand, and will have better efficiency at part load than the turbine. The slow two-stroke diesel is particularly suitable for the larger outputs (15–60 MVA), running economically on low grade fuel and with low maintenance costs. The recent focus of attention on emissions has had a marked effect on the choice of type of i.c. engine to be used. Engines burning gas are able to achieve lower level emissions of nitrous and sulphur oxides and are therefore proving very popular for new installations. In contrast, achieving acceptable levels of emissions with 2 stroke diesel engines burning low grade fuel is more difficult and costly.

Combined cycle installations are being increasingly used to attain higher thermal efficiency. Exhaust heat from a gas turbine that drives one generator is used to supply steam to a turbine that drives another. The turbines are high speed, geared down to, usually, the four-pole speed. For outputs greater than the gearbox limit of about 60 MW, two-pole cylindrical rotor generators are directly coupled to the turbines.

For marine applications, high-speed or medium-speed generators are used. These can either be driven by dedicated diesel engines, shaft driven from the engine used for propulsion or, in larger installations by gas turbines.

For ship propulsion, medium-speed diesel or gas turbine driven generators supply d.c. through rectifiers, or variable frequency a.c. to motors coupled to the propellers. The domestic load may be supplied from the same generators, but often separate generators are used—this has the advantage of isolating the domestic supply from the converter harmonics.

Oil rigs can either use internal combustion engine-driven generator sets or gas-turbine-driven generators for their power depending on their size. The generators are required to supply the large drilling motors and mud-pump motors, and the domestic supply. The number, rating and reactances of the generators are chosen to avoid excessive voltage dip when a large motor drive is started, so that other loads are not badly affected. Rated outputs can be around 15 MW where gear-driven generators running at 1800 rev/min would be used or up to 25 MW directly coupled to the power turbine at 3600 rev/min.

Industrial power stations commonly use steam turbines if there is need for process steam, which is delivered through a back-pressure or pass-out turbine; or if waste heat is available to raise steam for generating electricity.

Gas and diesel engines are often used, alone or in combined cycle installations. Gas turbines are expensive to run, and are used only if rapid response to load is essential, or the fuel is locally not so expensive, e.g. in an oil refinery.

28.18.2 Construction

The rotor construction for these salient pole machines is determined by the rating of the unit and its peripheral speed. High-speed rotors with four or six poles and ratings up to about 30 MW can have integral poles which are made up of a series of laminations. These can be punched from sheet material using a die or a nibbling type press; alternatively, they can be manufactured using a laser cutting machine. The thickness of punchings is normally limited to less than 3 mm if one of these processes is used, which has the advantage of limiting the pole-shoe losses. Thicker plates up to 100 mm thick can also be used, in which case the pole profile is cut using a numerically controlled flame cutting machine. With this type of construction, due consideration must then be given to the extra surface losses which will occur in the pole-shoes.

For larger machines, i.e. above 30 MW, with bigger diameter rotors, the peripheral speeds will result in rotational stresses in excess of the safe limits for the laminated type of rotor materials. In this case, forged solid rotors with pole pieces integral with the shaft are used. Solid forged pole-shoes are secured by high tensile strength bolts after the coils have been put on.

For the lower peripheral speeds associated with multipole rotors, separate laminated pole pieces may be secured by dovetails or T heads in slots in the square or hexagonal middle portion of the forged shaft. Alternatively, the laminated poles may be bolted or dovetailed to the wide cylindrical rim of a flywheel on the shaft; or they can be fixed by dovetails or T-head slots in a laminated rim.

The laminated-pole construction is cheaper than a forged rotor, and reduces the cost of the stator winding too, because fewer, wider, stator slots can be used with acceptably low eddy current losses in the pole faces.

Generators may have one or two bearings. Many diesel driven generators have only an outboard bearing, the driving end of the shaft being solidly coupled to the engine crankshaft flange. This shortens the set and saves the cost of a drive end bearing and the flexible coupling. It also allows the generator to be partly supported on a Society of Automobile Engineers (SAE) flange on the engine crankcase. The crankshaft bearing must carry half the weight of the generator rotor. A gearbox bearing can rarely do this, so gear-driven generators normally have two bearings.

The generator may be supported on sole plates grouted into the foundations, or may be bolted directly to the engine base plate, where this has been designed for the purpose. If transport and site lifting facilities permit, the generator can be delivered complete on its own bed plate; this is to be preferred as it saves erection time on site.

Often it is desirable to limit the vibration transmitted to the foundations, especially when the driver is a diesel engine. A common method is to put flexible mountings beneath the combined bed plate. These need to be chosen, or tuned, to isolate the vibration and not to permit resonance at any of the disturbing frequencies.

Class F insulation is always used on high voltage stator and some rotor windings, although class B temperature limits are often specified to extend the life of the machine and give some margin for an overload capability. Class H systems are common on low voltage generators and also on some medium voltage machines. Often class F temperature rises will still be specified even for machines with class H insulation systems in order to achieve an extended life.

Stator coils are usually of the pulled diamond type (see Section 28.4.1). Rotor coils are generally similar to those of hydrogenerators, but as they are smaller they are more often of the wound on edge construction using rectangular strip copper rather than fabricated. Where the rotor has integral poles, the field copper is wound directly onto the pole pieces by rotating the complete rotor or the laminated rotor core-pack about its longitudinal axis.

28.18.3 Ventilation and cooling

These generators are always air-cooled. If an adequate supply of clean air is available, open ventilation with an appropriate class of protection is used. If the air is only slightly dirty, it may be ducted to and from the machine, with filters on the inlet side. In a short machine the air is drawn through from one end to the other by a single centrifugal shaft-mounted fan. Longer machines have radial ducts in the stator core; the air enters the stator bore at both ends and is expelled radially through the ducts to the back of the core. In this case, shaft mounted axial-flow fans are used to drive the air through the machine.

In dirtier surroundings a totally enclosed machine is used. The primary (internal) air is circulated by shaft fans through an air to air heat exchanger (c.a.c.a. arrangement). The secondary (external, ambient air) is driven through the heat exchanger by one or more separate motor driven fans. Alternatively a water-cooled heat exchanger is used (c.a.c.w. arrangement). The materials used in the heat exchanger must be chosen to suit the quality of the water, e.g. tubes of aluminium brass or cupro-nickel if the water is salty. The limits of temperature rise may be adjusted in accordance with the maximum temperature of the secondary coolant. (see Section 28.5.)

28.18.4 Particular design requirements

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121. Emery, F.T., Harrold, R.T. ‘On line incipient arc detection in large t g stator windings’. T-PAS. 1980;99-6:2232–2238. November/December

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133. Dunlop, R.D., et al. ‘Torsional oscillations and fatigue of steam t g shafts caused by system disturbances and switching events’. Proc Cigre.. 11-06, 1980.

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(1) Electrical and Electronic Abstracts, published monthly by the IEE as part of its Inspection Service

(2) Cumulative Index of IEEE Transactions on Power Apparatus and Systems, 1975-1984, and for 1985

(3) Combined Index for IEEE Transactions on Power Delivery, Power Systems, Energy Conversion, published annually from 1986 onwards