202
IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 23, NO. 1, MARCH 2008
Wind–Diesel Generation Using Doubly Fed
Induction Machines
Rubén Peña, Member, IEEE, Roberto Cárdenas, Senior Member, IEEE, José Proboste,
Jon Clare, Senior Member, IEEE, and Greg Asher, Fellow, IEEE
Abstract—In this paper, the modeling and control strategy of
a wind–diesel generation system are discussed. In the proposed
topology, the diesel engine and the wind turbine are both variablespeed machines, allowing maximum fuel efficiency and optimal
energy capture from the wind. A vector-controlled doubly fed induction generator is used in each generation system to provide
fixed voltage and frequency to the load. The diesel unit balances
the system power and changes the speed according to the power
demand in order to minimize the fuel consumption. The electrical
torque of the wind system generator is regulated to maximize the
energy capture of the wind turbine. The advantages of operating
a diesel engine at variable speed are discussed. The dynamic and
steady-state operation of the wind–diesel system, including voltage and frequency control, active power balancing, and control of
the reactive power supplied to the grid/load are analyzed in this
paper. Experimental results, from a 3-kW experimental prototype
are presented in this paper.
ρ
Air density.
Electrical angle.
θe
Rotor position angle.
θr
Slip angle.
θslip
Induction machine rotational speed.
ωr
Stator electrical frequency.
ωsG
Slip frequency.
ωslip
Superscripts
∗
Demanded value.
Subscripts
(d, q)
Synchronous rotating coordinates.
G, W
Diesel, wind generation quantity.
r, s
Rotor or stator quantities.
(α, β)
Stator fixed coordinates.
Index Terms—Diesel-driven generators, induction generator, induction motor drives, wind energy.
ARIABLE-SPEED operation of wind turbines has many
advantages that are well documented in the literature [1],
[2]. The torque peaks in the gearbox and shafts are reduced,
the wind turbine can operate with maximum aerodynamic efficiency, and the power fluctuations can be absorbed as an inertial
energy in the blades. In some applications, the wind turbine may
be augmented by an additional source, usually a diesel generator.
These systems are called wind–diesel systems [3], [4] and they
may be used to supply electrical energy to stand-alone loads,
e.g., small villages that are not connected to the main utility.
Most diesel generation systems operate at a constant rotational
speed due to the restriction of constant frequency at the generator
terminals. However, diesel engines have high fuel consumption
when operating at light load and constant speed [5], [6]. Moreover, for light loads at rated speed operation, not all the fuel
is burned by the engine and wetstacking is produced [7], [8].
This increases maintenance costs [8]. In order to improve the
efficiency and avoid wetstacking, a minimum load of about 30%
to 40% is usually recommended by the manufacturers [8].
In recent publications [5]–[9], the operation of variable-speed
diesel engines is proposed. The main advantage of variablespeed operation is increased efficiency, because the fuel consumption can be reduced by up to 40%, especially when the
diesel generator supplies energy to a light electrical load [7], [8].
Moreover, the engine life is increased because the diesel engine
is run at a low speed for a light load. In this way, not only is
the wetstacking avoided, but also the engine is operated with a
lower thermal signature [6].
To explain further the motivation for this paper, Fig. 1(a)
illustrates the electrical load characteristics measured during
October (middle spring) at “Villa Tehuelche,” a small village
100 km from the city of Punta Arenas, Chile. A fixed-speed
V
NOMENCLATURE
General
i
Lm , Ls , Lr
R r , Rs
Te
v
λ
σ
σs
DFIG
DGS
GR
ims
p
PWM
Rb
Tsr
τ
V
WECS
I. INTRODUCTION
Stator or rotor current.
Magnetising, rotor, stator inductance.
Rotor, stator resistance.
Electrical torque.
Stator or rotor voltage.
Stator or rotor flux.
Total leakage coefficient.
Stator leakage coefficient.
Doubly fed induction generator.
Diesel generator system.
Gear ratio.
Magnetizing current.
Number of poles.
Pulsewidth modulator.
Turbine blade radius.
Tip speed ratio.
Time constant.
Wind velocity.
Wind energy conversion system.
Manuscript received April 24, 2006; revised January 29, 2007. This work
was supported in part by Fondecyt under Grant 1010942, in part by the British
Council, and in part by the University of Magallanes. Paper no. TEC-001152006.
R. Peña, R. Cárdenas, and J. Proboste are with the Electrical Engineering
Department, University of Magallanes, Punta Arenas 113-D, Chile (e-mail:
ruben.pena@umag.cl; rcd@ieee.org; jprobost@umag.cl).
J. Clare and G. Asher are with the School of Electrical and Electronic Engineering, University of Nottingham, Nottingham NG7 2RD, U.K. (e-mail:
jon.clare@nottingham.ac.uk; greg.asher@nottingham.ac.uk).
Digital Object Identifier 10.1109/TEC.2007.914681
0885-8969/$25.00 © 2008 IEEE
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PEÑA et al.: WIND–DIESEL GENERATION USING DOUBLY FED INDUCTION MACHINES
203
Fig. 1. (a) Frequency distribution for the load at “Villa Tehuelches.” (b) Fuel
consumption for fixed speed and variable speed operation.
Fig. 2.
diesel system is used to supply electricity to the village, daily
from 6:30 A.M. to 12:00 P.M. During winter, the load increases
because electrical heating is used. Most of the time, the load is
40% to 60% of the nominal value. The maximum load may occur few times during the year, typically, during winter. Fig. 1(b)
shows the fuel consumption characteristic of the generator illustrating the saving that could be made by operating at variable
speed. The overall fuel saving obtained for this month would
be 22%. This saving is in addition to the other benefits outlined
earlier.
In general, for variable-speed operation, power electronic interfaces must be provided in order to have constant frequency
and regulated voltage in the ac load. In [5], the simulation results
for a variable-speed wind–diesel system, with an additional energy store, are presented illustrating the advantages of variablespeed operation. Permanent magnet machines are considered
and a diode rectifier, a chopper, and an inverter are used in each
generator. However, in [5], there is little discussion about load
voltage regulation, performance of the system when load impacts are considered, and grid frequency control. Moreover, the
control of the nonlinear diesel system is not addressed.
DFIGs have long been considered as a good choice for variable speed generation systems [10]–[12]. If a DFIG is operated
in a restricted speed range, the power converters are rated to
only a fraction of the total system power, typically, at 30% of
the machine-rated power. The wind–diesel system proposed in
this paper is based on DFIGs (see Fig. 2). DFIGs are used in
the variable-speed DGS and in the WECS. The machine stators
are connected together to form an ac bus with fixed frequency
and voltage. Three voltage source PWM converters are required
for the proposed wind–diesel system. A single dc link is used to
connect the DGS and WECS rotor converters. These converters
are vector controlled to regulate the rotor currents in both the
machines. A single vector-controlled front-end converter is used
to connect the dc link to the ac bus. Some preliminary results, of
the wind–diesel system, proposed in Fig. 2, are presented by the
authors in [13]. However, in this publication, only simulation
results are presented; the effects of inductive load disturbances
are not considered. Small-signal models and control system design are not discussed. For the wind–diesel system in Fig. 2,
there are at least three modes of operation. The first mode is
when the wind turbine is operating at high wind speed, and the
power captured by the WECS is sufficient to source the load.
In this case, the diesel generator is disconnected and the WECS
DFIG is controlled to operate in a stand-alone mode [14]. The
second mode of operation is when the wind speed is very low.
In this case, the WECS is disconnected and the DGS supplies
the required energy to the grid/load. Finally, the third mode
of operation for the proposed wind–diesel system is when the
WECS is connected to the system, but the power captured by
the wind turbine is not sufficient to feed the load. Therefore, the
load has to be sourced with generation from both the WECS
and the DGS. The first mode of operation is not considered here
since it has been discussed extensively in [14]. The second and
the third modes of operation are addressed in this paper.
The DGS (when connected) is always controlled to regulate the voltage and electrical frequency of the load [14]. The
WECS DFIG is controlled for grid-connected operation [10],
since the DGS forms a virtual grid for the WECS. Therefore,
the generator is synchronized to the grid, and the vector control system is orientated in the stator flux. If the control is run
on a processor other than that used for the DGS system, the
stator flux is estimated from the grid voltage and stator current measurements. Otherwise, the same flux vector position
used to control the DGS generator can be used to control the
WECS generator [13]. The WECS DFIG electrical torque is
controlled to drive the wind turbine to the point of maximum
aerodynamic efficiency, optimizing the energy capture from the
wind [10], [14]. When both the DGS and the WECS are sourcing the load, the power consumption has to be balanced with
the total power generated by both the generators. For instance,
if the energy captured from the WECS increases, the energy
supplied from the DGS decreases, and the rotational speed of
the diesel generator is reduced in order to save fuel and improve
efficiency. On the contrary, if the energy captured from the wind
turbine is reduced, then the DGS has to supply more energy
into the grid, and the rotational speed of the diesel generator
increases. The reactive power required by the system can be
supplied from the DGS generator, the WECS generator, or from
the front-end converter according to some control law, e.g., to
reduce the losses and increase the efficiency of the whole generation system [11], [15]. However, the optimal control of the
reactive power sourcing is outside the scope of this paper. The
rest of this paper is organized as follows. In Section II, the control systems for stand-alone operation of the DGS are discussed.
Fuel consumption curves obtained from a 3-kW, 220-V, 50-Hz
Proposed wind–diesel system.
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IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 23, NO. 1, MARCH 2008
Fig. 3.
Proposed control system for the diesel doubly fed induction machine and front-end converter.
diesel generator system are presented. In Section III, the control
systems for simultaneous operation of the DGS and WECS are
introduced and small-signal models are analyzed. Section IV
presents experimental results for the stand-alone operation of a
variable speed DGS. Also, experimental results related to simultaneous operation of the WECS and DGS, operating at variable
speed, are presented and discussed verifying the validity of the
proposed methodology. Finally, an appraisal of the proposed
wind–diesel topology is presented in the conclusion.
II. MODELING AND CONTROL OF THE DIESEL SYSTEM
A. Modeling and Control of the DFIG
The machine equations of a DFIG in a synchronously rotating
d–q reference frame, with the q-axis aligned along the stator flux
vector position are given by [14]
idsG
LsG
0
λdsG = Lm G im sG
=
0
iqsG
0
LsG
Lm G
idrG
0
+
(1)
0
Lm G
iqrG
vdsG
vqsG
vdrG
vqrG
d λdsG
RsG
idsG
0
=
+
iqsG
0
RsG
dt λqsG
0
−ωsG
λdsG
+
ωsG
0
λqsG
d λdrG
RrG
idrG
0
=
+
0
RrG
iqrG
dt λqrG
0
−ωslipG
λdrG
+
ωslipG
0
λqrG
(2)
of the reference frame on the stator flux vector gives
iqrG = −
LsG
iqsG .
Lm G
(4)
Considering (1) and (2), the following expression is obtained
for the dynamics of the magnetizing current
τsG
dim sG
1 + σsG
+ im sG = idrG +
vdsG
dt
RsG
(5)
with τsG = LsG /RsG and σsG = (LsG − Lm G )/Lm G . The
magnetizing current is controlled via the rotor excitation current
idrG . The rotor current iqrG is controlled in order to follow a
reference current given by (4) to force the orientation of the
reference frame along the stator flux vector position. If iqrG follows the reference under the action of a fast current control loop,
then the orientation of the reference frame along the stator flux
vector will be correct. The vector control schematic is shown in
the left-hand side of Fig. 3. The idrG and iqrG currents are regulated using the PI controllers. Compensation terms are added
to the controller outputs to provide linear transfer functions in
order to simplify the controller design and ensure good tracking
of these currents. The slip angle is given by
(6)
θslipG = θsG − θrG = ωsG dt − θrG
where θrG is the DFIG rotor position. The stator flux angle θsG
is obtained by the integration of the reference stator frequency
ωsG = 2π50 rads−1 . A diesel engine operating at variable speed
regulates the speed of the generator.
B. Control Strategy of the Front-End Converter
(3)
where im sG is the equivalent stator magnetizing current and
ωslipG = ωsG − ωrG is the slip frequency. Aligning the d–q axis
The aim of the front-end or stator-side converter is to regulate the common dc link voltage E, regardless of the direction
of the power flow. The converter currents are controlled with
the conventional vector control approach [10] with a d–q reference frame oriented along the stator voltage vector position
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PEÑA et al.: WIND–DIESEL GENERATION USING DOUBLY FED INDUCTION MACHINES
Fig. 4.
205
Diesel engine model.
θv . The reference frame orientation angle can be derived from
the stator flux vector position of the diesel-driven DFIG as
θv = θsG + π/2. The control schematic is shown in the righthand side of Fig. 3. A PI controller is used to process the error
in the dc link voltage and generate the active power component
reference current. The reactive power component reference current could be set to zero, implying close-to-unity displacement
factor operation, or could be the output of a reactive power
controller.
Fig. 5.
Variation of K p as a function of power and rotational speed.
Fig. 6.
Fuel consumption versus power at various rotational speed.
C. Modeling and Control of the Diesel Engine
The model of the diesel engine used (see Fig. 4) is based on the
previous work on fixed- and variable-speed diesel engines [6]–
[9], [16]–[18], and experimental data obtained from a 3-kW
diesel generator set. The dynamic of the actuator is represented
by a first-order model with a time constant τ2 [16], [17] and a
gain K2 . The combustion system is represented as a variable
gain K1 , which depends on the speed and output power [16],
and a dead time τ1 . This dead time can be calculated as
60
60st
+
(7)
τ1 =
2N n 4N
where st = 2 or four for two- or four-stroke engines, N is the
speed in r/min, and n is the number of cylinders. In Fig. 4, J is
the total system inertia and B is the friction coefficient. The load
torque TeG is the electrical torque calculated according to [14]
TeG = 3
p L2m G
im sG iqrG .
2 LsG
(8)
As shown in Fig. 5, the total gain Kp = K1 K2 of the diesel
engine is dependent on the rotational speed of the engine and
the power supplied by the DGS. Further discussion about the
variation of K1 and K2 for a typical diesel generator set is
presented in [16]. The fuel consumption in a diesel engine depends on the speed and torque of the machine. Fig. 6 shows
the fuel consumption curves obtained experimentally from a
3-kW, 220-V, 50-Hz diesel engine, for five rotational speeds.
According to Fig. 6, at 20% of rated power, there is 50% of additional fuel consumption when the system is operated at rated
speed instead of 0.6 pu. Fuel efficiency also decreases when
the power supplied by the DGS is increased without adjusting
the rotational speed. For instance, if the power supplied by the
DGS is changed from 35% to 45% of the rated value, the en-
gine speed has to be varied from ∼0.7 to ∼0.8 pu in order to
improve the system efficiency. Therefore any increase/decrease
in load power should be accompanied by an increase/decrease
in the rotational speed to improve the system efficiency. From
the fuel consumption characteristic of Fig. 6, a continuous function for the optimal curve in the power-rotational speed plane
for minimum fuel consumption can be obtained. For the diesel
system tested in this paper, the relationship between the optimal
rotational speed and the power supplied by the DGS is almost
linear, as shown in Fig. 7. This is in broad agreement with the
previous research [5]–[7]. The control scheme for the diesel
engine is shown in Fig. 4. If losses in the DFIG are neglected,
the power supplied by the diesel engine to the shaft is given by
PeG = ωrG TeG . In order to minimize the fuel consumption, the
speed demand (optimum speed) for the diesel engine is calculated by using a look-up table where the optimal power–speed
curve (see Fig. 7) is implemented. The input to the look-up table
is PeG , and the output is the demanded speed. The diesel speed
is regulated by using a PI controller. In order to compensate the
variations of Kp , a gain scheduling control system is used [16].
The controller gain is a function of the speed and power and is
adjusted using an additional 2-D look-up table.
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IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 23, NO. 1, MARCH 2008
using (12) and (10), for optimal power capture, the torque component current (iqrW ) is regulated as
i∗qrW =
Fig. 7.
Optimal rotational speed versus power curve.
III. MODELING AND CONTROL OF THE WIND TURBINE SYSTEM
A. Modeling of the Wind Turbine
The power captured and the mechanical torque produced by
a wind turbine are given by [19], [20]
(9)
where Ct (Tsr , β) and Cp (Tsr , β) are the torque and power coefficients, respectively, β is the blade pitch angle, Tsr is the
tip-speed ratio (= ωT Rb /V ), and ωT is the rotational speed of
the blades. In this paper, the blade characteristic Ct (Tsr , β) reported in [21] is used. For each wind velocity, there is a point
of maximum power capture when the turbine is operating at the
maximum power coefficient (Cpm ax ) [19]. If the losses are neglected, it can be shown that, in a steady state, the wind turbine
operates at an optimum power coefficient when the generator
electrical torque is regulated as [20], [22]
TeW =
2
kopt ωrW
(10)
where kopt is a constant that depends on the blade aerodynamic,
gear box ratio, and wind turbine parameters. More information related to the control of variable-speed generators for wind
energy systems is presented in [2], [10]–[12], [14], [15], [19]–
[22].
B. Modeling and Control of the WECS DFIG
The proposed control strategy considers the generator of the
WECS as a grid-connected DFIG. The grid voltage and frequency is established by the DGS system. The WECS DFIG
is vector controlled with a reference frame orientated along the
stator flux. The stator flux position θe is obtained as [10], [14]
λβ sW
θe = tan−1
λα sW
λα sW = (vα sG − RsW iα sW )dt
λβ sW =
(vβ sG − RsW iβ sW )
(11)
where λα sW and λβ sW are the α − β stator flux components.
The electrical torque is obtained as [10]
TeW = 3
p L2m W
(im sW iqrW )
2 LsW
(13)
With some minor modifications, the control system shown
in the left-hand side of Fig. 3 can be applied to grid-connected
DFIGs. In this case, the position of the synchronous reference
frame is given by (11), the rotor q-axis reference current is given
by (13), and the rotor d-axis reference current is set to zero, unless otherwise specified. To avoid the integrator drift produced
by the dc component in the signals, the integrators of (11) are
replaced by bandpass filters [10], [14]. Further information regarding vector control of grid-connected DFIGs is presented
in [10].
C. Control of the Proposed Wind–Diesel System
Tm = 0.5πρCt (Tsr , β)Rb3 V 2
Pm = 0.5πρCp (Tsr , β)Rb2 V 3
2
2LsW kopt ωrW
.
2
3pLm W im sW
(12)
This section discusses the system integration of the WECS
and DGS, and the dynamic and steady-state operation of the
entire system. For each DFIG, and by neglecting the losses, the
relationship between the power supplied from the stator and the
power supplied by the machine rotor is [23]
Pr = −sPs .
(14)
In (14), Ps and Pr have positive values when the power is
supplied from the grid to the machine stator or rotor, respectively, with s = (ωsG − ωre )/ωsG and ωr e = (p/2)ωr . The net
power supplied from each DFIG is
PW = −(PsW + PrW ) = −PsW (1 − sW )
PG = −(PsG + PrG ) = −PsG (1 − sG ).
(15)
The front-end converter is used to supply electrical energy to
both machine rotors. The total power supplied by the front-end
converter to the machine rotors Pfe is
Pfe = −(sW PsW + sG PsG ).
(16)
From (16), it is concluded that not all the power required
by the machine rotors is supplied from the grid through the
front-end converter. If sW and sG have opposite signs, a part of
the energy is directly supplied from the WECS DFIG rotor to
the DGS DFIG rotor (or vice versa), via the common dc link.
The worst case, from the viewpoint of the maximum current
in the front-end converter, may occur when both generators
are operating at a supersynchronous speed (i.e., sW , sG < 0).
However, this case is unlikely to occur unless a high load above
the nominal power of the DGS is connected to the grid. For most
applications, it is considered that, by adopting similar ratings for
the front-end converter and the two rotor converters will provide
a satisfactory system. The total power supplied to the grid/load
is
PL = vdL idL + vqL iqL
(17)
where vdL , vqL , idL , and iqL are the d–q components of the
load voltage and current. Because the stator resistance in each
machine is relatively small, vdsW ≈ vdsG ≈ vdL ≈ 0 and the
load power can be obtained as PL ≈ vqL iqL . For simultaneous
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PEÑA et al.: WIND–DIESEL GENERATION USING DOUBLY FED INDUCTION MACHINES
Fig. 9.
Fig. 8.
207
Small-signal model for stator flux regulation of the DGS generator.
Small-signal model for power component of voltages and currents.
operation of the DGS and WECS, each generator supplies a
fraction of the load power. From (15), the load quadrature current
iqL is obtained as
iqL = −(iqsW (1 − sW ) + iqsG (1 − sG )).
(18)
Fig. 8 shows the small-signal model for the power control
system of the proposed wind–diesel topology. It is assumed that
the wind turbine is driven at maximum aerodynamic efficiency
by regulating the electrical torque of the WECS generator via
(10). The block labeled “nonlinear gain” represents the transfer
function obtained by linearising (13) as
∆i∗qrW
4LsW kopt ωrW 0
=
.
∆ωr
3pL2m W im sW 0
(19)
Fig. 10.
The cross-coupling terms [10] between the d and q axes are
neglected in Fig. 8, because they are compensated at the output of the current controller (see Fig. 3). In Fig. 8, the WECS
generator is similar to a current source supplying the energy
captured from the wind to the system. In order to balance the
power and considering constant load operation (i.e., ∆idL ≈ 0),
the relationship between the quadrature currents supplied from
the WECS and DGS is obtained from (18) as
∆iqsW (1 − sW ) = −∆iqsG (1 − sG ).
(20)
Therefore, a change in the power captured from the wind
turbine is compensated by a change in the opposite direction
for the power generated from the diesel system. Variations in
the power generated by the DGS also produce changes in the
rotational speed of the diesel engine that is regulated to a new
operating point, minimizing the fuel consumption. If the load is
relatively small compared to the energy captured by the WECS,
then the pitch control of the blades [20] or the power dissipation
in a dump load has to be used in order to balance the power in the
system. In the proposed system, the grid voltage and frequency
are regulated by the DGS generator.
The grid voltage control is achieved through stator flux regulation, because the stator resistance voltage drop is negligible.
Fig. 9 shows the closed loop control system used to regulate
the stator flux of the DGS generator. The flux is regulated by
controlling the rotor current idrG . In Fig. 9 idL , iq , and idsW are
the reactive components of the load current, front-end converter
current, and the DFIG stator current, respectively. From Fig. 9, it
is concluded that the load reactive power can also be supplied by
Experimental system.
the DGS generator. However, this is relatively slow, considering
the low bandwidth of the stator flux control loop. Alternatively,
the reactive power can be supplied from the front-end converter
or the WECS generator.
IV. EXPERIMENTAL WORK
The proposed wind–diesel topology has been tested by using
the experimental system shown in Fig. 10. Two DSP boards,
based on the TMS320C31 processor, are used to control the
whole system. The control algorithms for the emulation of the
diesel engine, optimum speed tracking, vector control of the
diesel-driven DFIG, and the control of the front-end converter
are implemented in one of the DSP boards. A speed-controlled
cage induction machine is used to emulate a 3.0-kW diesel engine with 35% transient torque overload and rated friction losses
of 0.2 kW. The algorithms for the emulation of the variablespeed wind turbine, tracking of the optimal rotational speed,
and vector control of the WECS DFIG are implemented in a
second DSP board. The DSP boards are installed in separated
host computers.
Emulation of wind turbines has already been presented in [22]
and [24], and is briefly discussed here. The wind turbine is emulated using a speed-controlled dc machine. To implement the
emulation, wind profiles are sent from the host PC to a secondorder model of the WECS implemented in the DSP. The power
coefficient curve is stored, in the DSP memory, using a look-up
table. Linear interpolation is used to obtain the power coefficient Cp (Tsr , β) from the look-up table. By using a wind turbine
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IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 23, NO. 1, MARCH 2008
model and the methodology discussed in [22] and [24], the
∗
) is calculated in
reference speed of the WECS generator (ωrW
each sampling time (see [22]). The speed-controlled dc machine
forces the DFIG speed to this value. With this emulation technique, the DFIG rotates at the same speed as that of a generator
driven by a real wind turbine. A complete discussion of the
emulation technique used in this paper can be found in [24]
and [25].
Two identical DFIGs are used for the implementation of the
experimental rig. The speed range for both the machines is
from 700 to 1300 r/min (rated speed 1000 r/min), ±30% of
the synchronous speed, while the magnetizing current of the
DGS is controlled at 7-A resulting in a 120-V stator voltage.
The dc-link voltage is regulated at 530 V and a 10 µF/phase
capacitor is connected to the stator to filter the high-order PWM
harmonics. External DSP interfaces are used to measure the
rotor position (using 10000 pulses per revolution encoders), for
signal conditioning and to provide the PWM signals to the power
converters. The converters switching frequency is 1 kHz and the
sampling time is 0.5 ms.
Fig. 11. Control system performance for load steps. (a) Speed tracking performance. (b) Estimated fuel consumption.
A. Emulation of the Diesel Engine
The diesel engine emulator models both the steady state and
dynamic characteristics by controlling the rotational speed of
an induction machine drive (different machines were used for
the emulations because of the availability of equipment in the
laboratory). The rotational speed is a function of the actuator
input u(s) (see Fig. 4) and the DFIG electrical torque. From
Fig. 4, the rotational speed is obtained as
ω(s) =
1
Kp e−sτ 1
u(s) −
TeG (s)
(1 + sτ2 )(sJ + B)
(sJ + B)
(21)
using (8) and the bilinear transform, (21) can be discretized to
Kp [(z + 1)Ts ]2 z −N t
u(z)
ω(z) =
(2τ2 + Ts )(2J + BTs )(z − A)(z − C)
−
A=
(z + 1)Ts p L2m G ∗
3
i
i∗ (z)
J(z − A) 2 LsG m sG qrG
(2J − BTs )
(2J + BTs )
C=
(2τ2 − Ts )
(2τ2 + Ts )
(22)
the variable delay Nt is calculated as τ1 /Ts , where τ1 is obtained from (7) and Ts is the sampling time. Using (21) and
(22), the emulation of the diesel engine is implemented. For
each sampling time, ω(k) is calculated from (22), and is used
as the demand velocity for the speed control system of the cage
induction machine (see Fig. 10) i.e. the induction machine rotates at the same speed as the DGS modeled under dynamic and
steady-state conditions.
B. Experimental Results for the DGS
Experimental results for the DGS supplying energy to a standalone load are presented in this section. The system is tested
for step changes of resistive and inductive loads. Fig. 11(a)
shows the speed tracking performance when step resistive loads
are applied to the stator. The reference and the actual speeds
Fig. 12. Current response corresponding to Fig. 11. (a) Magnetizing and rotor
currents. (b) Stator currents. (c) Front-end converter currents.
are shown. Initially, the system is running with a 0.9 kW load
at the optimum speed of 783 r/min. At t = 50 s, the load is
increased to 1.3 kW, and correspondingly, the control strategy
drives the generator to the new optimum speed of 910 r/min.
A new step load is applied at t = 150 s resulting in a total
load of 2.1 kW. The speed of the system increases and a new
optimal speed of 1190 r/min is reached. At these loads the diesel
is running, considering losses at 45%, 56%, and 85% of rated
power. The load is decreased at t = 275 and 360 s. Fig. 11(b)
shows the estimated fuel consumption for these load conditions,
considering the experimental results of Fig. 6. Extrapolation has
been used to estimate the fuel consumption outside the registered
experimental data. In steady state, for 45% and 56% of rated
power, the fuel saving operating at reduced speed would be 18%
and 10%, respectively.
Fig. 12(a) shows the magnetizing and the d–q axis rotor currents for the conditions corresponding to Fig. 11. The q-axis
rotor current reflects the increase in power demand due to the
load impacts. After the transient, the rotational speed and torque
current settle to a new operating point. The variation in the daxis rotor current is due to the slight variation in the second
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PEÑA et al.: WIND–DIESEL GENERATION USING DOUBLY FED INDUCTION MACHINES
Fig. 13. Voltages and flux corresponding to Fig. 11. (a) DC link and stator
voltages. (b) d–q stator flux components.
Fig. 14. Control system response for an inductive load step. (a) Speed tracking
performance. (b) DC link voltage and magnetizing current.
term of the right-hand side in (5). Fig. 12(b) shows the d–q axis
DFIG stator current. Because the load is resistive the d-axis stator current is nearly zero and the power factor is close to unity.
The q-axis stator current reflects the power changes due to the
load impacts. Finally, Fig. 12(c) shows the front-end converter
currents. The d–q axis front-end converter currents illustrate the
operation close to unity power factor (q-axis current ≈ 0). The
d-axis front-end converter current reverses when the speed is
above synchronous, because power is supplied to the load from
the stator and the rotor of the DFIG.
Fig. 13(a) shows the dc link voltage and the rms stator voltage
corresponding to the test of Fig. 11. The dc link voltage excursion is within ±10 V during these load and speed transients. The
stator voltage is practically constant, and the regulation is very
good because the effect of the stator resistance voltage drop is
negligible. Fig. 13(b) shows the d–q stator flux components for
these load and speed transients. The q-axis stator flux is close
to zero, reflecting the correct orientation of the vector control
system.
Fig. 14 shows the operation of the system when an inductive load step is applied to the stator. Initially, the system is
supplying a 1.9 kW, 0.25 kvar load, with the optimum speed
control strategy enabled, and a speed of about 1128 r/min. At t
≈ 1 s, a load step of about 1.65 kvar is applied. Fig. 14(a) shows
209
Fig. 15. Current response for the test of Fig. 14. (a) Reactive power currents.
(b) Active power currents.
the reference and real speeds whereas Fig. 14(b) shows the dc
link voltage and magnetizing current. Initially, at t ≈ 1 s, the
reference speed reduces because the load impact causes a magnetizing current (and stator voltage) dip of about 30%; hence,
the power supplied by the machine reduces. The settling time
for the magnetizing current is about 0.8 s. After the transient,
the final value for the speed (≈1214 r/min) is slightly higher
than the initial speed because of the additional losses produced
by the increase in the stator current.
In Fig. 14, the inductive load step is disconnected at t ≈
13 s. There is an overshoot in the magnetizing current (and
stator voltage) of about 35%; hence, at t ≈ 13 s, the reference
speed increases due to the increase in power. As the magnetizing
current error goes to zero, the speed of the system settles down to
about 1128 r/min. The variation on the dc link voltage is mainly
produced by changes in the machine stator voltage. The dip
and the overshoot of E are below 20 V. These results illustrate
well-managed interactions between the controllers.
Fig. 15 shows the d–q axis active and reactive power current
components corresponding to the load transients of Fig. 14.
The front-end converter q-axis current (iq ) is regulated at zero
[see Fig. 15(a)], hence, the d-axis rotor current compensates the
increase in the load reactive power. As shown in Fig. 15(b),
the increase in the system losses, produced by the load step,
is compensated by a small increase in the DFIG q-axis rotor
current.
During the load steps, the additional reactive power is supplied by the DFIG with a dynamic that is dependent on the low
bandwidth of the magnetizing current control loop. The dynamic
of the response can be improved if the required reactive power
is supplied by the front-end converter. This is shown in Figs. 16
and 17. The front-end converter reactive power is regulated in
order to supply all the reactive power required by the load (i.e.,
idsG ≈ 0). Fig. 16(b) shows that the dip in the magnetizing current for the reactive load step is reduced to 10% [compare with
Fig. 14(b)]. However, the system losses are increased when the
reactive power is supplied from the front-end converter and the
speed settles to a higher value ≈1300 r/min. A dc link voltage
dip of about 20 V occurs, because during transients, the angle
between the stator flux and the voltage is not π/2 as assumed
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210
IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 23, NO. 1, MARCH 2008
Fig. 16. System performance with reactive power compensation. (a) Speed
tracking. (b) Magnetizing current and dc link voltage.
Fig. 18.
Steady state operation at sub (a, b) and supersynchronous (c, d) speed.
Fig. 17. Reactive power current components. (a) DFIG currents. (b) Front-end
converter currents.
by the front-end converter control system (see Fig. 3). Fig. 17
shows the magnetizing currents supplied by the DFIG and frontend converter corresponding to the test of Fig. 16. The d-axis
stator current is regulated at zero [see Fig. 17(a)]; hence, no
change in the reactive current is seen by the machine during the
inductive load impact.
Fig. 18 shows steady-state results for the voltages and currents of the DFIG stator and front-end converter. These results
are obtained at sub- [Fig. 18(a) and (b)] and supersynchronous
speed [Fig. 18(c) and (d)]. The rotational speeds are 830 and
1300 r/min, respectively. The equivalent per phase stator voltage
va is also shown. The load reactive power is supplied entirely
by the machine; hence; the front-end converter operates at close
to unity power factor.
C. Experimental Results for the DGS and WECS
Assuming that the DGS is already active, connecting the
WECS generator to the grid is equivalent to a step demand
in the reactive power supplied by the DGS generator. As discussed in the previous section, a sudden change in the reactive
power provided by the DGS DFIG may produce a relatively
large variation in the magnetizing current (and stator voltage)
due to the low bandwidth of the magnetizing current loop. This
excursion is reduced if the front-end converter provides the reactive power. However, this reduces the maximum power current
Fig. 19. WECS connection to the system. (a) DGS speed. (b) Magnetizing
current and dc link voltage.
available in the converter. Also, if the current supplied by the
front-end converter increases, it may produce higher switching
and conduction losses in this power converter. To overcome
both of these difficulties, a control strategy is used in which the
required reactive power is initially provided by the front-end
converter. After the transient, the reference reactive power for
the front-end converter is gradually reduced in order to reduce
the losses and avoid variations in the stator voltage.
Figs. 19 and 20 show results for connection of the WECS
generator to the system. Before the connection, the rotor of the
WECS DFIG is opened, and the stator is connected to the grid
by using a manually operated circuit breaker. In this test, the
system is sourcing a load of 1.9 kW with a 0.97 lagging power
factor. Fig. 19(a) shows the speed during the transient. When
the WECS is connected, the DGS speed, initially 1130 r/min,
tends to increase due to the higher losses, but as the reactive
power supplied by the front-end converter reduces, the losses
decrease. The speed finally settles at 1215 r/min. Fig. 19(b)
shows the magnetizing current and the dc link voltage. The dip
and the overshoot in the magnetizing current are below 12%.
The dc link voltage has a maximum overshoot of 4.5%.
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PEÑA et al.: WIND–DIESEL GENERATION USING DOUBLY FED INDUCTION MACHINES
Fig. 20. Rotor and stator current corresponding to the test of Fig. 19. (a)
Reactive power component. (b) Active power current components.
Fig. 22. Control system response for a WECS DFIG q-axis current step. (a)
Currents in the WECS. (b) Currents in the DGS.
Fig. 23.
Fig. 21. Control system response for a WECS DFIG d-axis current step. (a)
Currents in the WECS. (b) Currents in the DGS.
The active and reactive power current components, corresponding to the test of Fig. 19, are shown in Fig. 20. Initially,
the DGS reactive stator current is zero and the total reactive current is supplied from the front-end converter [iq in Fig. 20(a)].
The rotor current idrG sets the flux in the DGS machine. At
t ≈ 3 s, the front-end reactive current is decreased using a slow
ramp. Finally, when iq ≈ 0, the reactive power of the load is
supplied by the DGS. Fig. 20(b) shows the DGS torque currents
and the front-end converter active power current component
(i∗qrW = 0 for this test). The variation in the DGS torque current
is negligible. Therefore, considering the increase in the generator speed (see Fig. 19), the power supplied from the DGS is
slightly higher than that supplied before the connection of the
WECS generator.
The next experimental test shows the power balance produced
when a change in the active and/or reactive power generated by
the WECS is balanced by an opposite variation in the active
and/or reactive power generated from the DGS DFIG [see (20)
and Figs. 8 and 9]. The load connected to the system is 2.0
kW, and before the transient, the power factor seen from the
DGS is 0.8 lagging. The speed of the WECS is controlled at
800 r/min. Fig. 21(a) shows the response of the WECS DFIG
rotor current controller. The q-axis rotor current is regulated at
zero, and a step demand of 3-A reactive current is produced.
Fig. 21(b) shows the reactive and magnetizing currents of the
211
DGS speed for active and reactive power changes from the WECS.
DGS DFIG when the step increase in idrW takes place. Because
a part of the magnetizing current, required by the DGS generator,
is provided from the WECS, the current idrG in the DGS DFIG
reduces. Therefore, the power factor seen from the DGS DFIG
stator increases to 0.95.
The performance of the system for a step change in the torque
current, iqrW , is shown in Fig. 22. The WECS d–q rotor currents, with the reactive current component set to zero, are shown
in Fig. 22(a). For this test, the DGS DFIG q-axis rotor current
does not change significantly, because the power balancing is
produced mainly by changes in the rotational speed of the DGS
and a small variation in the torque current of the DGS generator [see Figs. 22(b) and 23]. For this test, the variation in the
magnetizing current is also low. The DGS speed corresponding
to the tests of Figs. 21 and 22 is shown in Fig. 23. For the idrW
step change, there is an increase in the DGS losses and the speed
of the DGS system rises. Therefore, for this operating point the
DGS DFIG is less efficient when the magnetizing current is supplied from the stator instead of the rotor. When a step increase
in iqrW is applied, a part of the power required by the load is
supplied from the WECS and the speed of the DGS reduces
accordingly.
The performance of the wind–diesel system for a step increase
in wind velocity is shown in Fig. 24. A step in the wind velocity
is not realistic but is a very drastic change, appropriate to verify
the performance of the proposed control system. The optimum
speed for the WECS corresponds to continuous operation at
maximum power coefficient, and for each wind velocity, is given
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212
Fig. 24.
IEEE TRANSACTIONS ON ENERGY CONVERSION, VOL. 23, NO. 1, MARCH 2008
Speed system performance for a step increase in wind velocity.
Fig. 26. Active current components corresponding to the test of Fig. 24. (a)
DGS DFIG current. (b) Front-end converter current. (c) WECS DFIG current.
Fig. 25. Currents and voltages corresponding to the test of Fig. 24. (a) DC
link voltage. (b) DGS DFIG Magnetising current. (c) DGS DFIG stator voltage.
by [19]
ωrW ,opt =
Tsr,opt V
GR
Rb
(23)
where Tsr,opt is the tip speed ratio corresponding to Cpm ax .
Initially, a 2 kW load is connected to the ac bus and the wind
velocity is 5 m/s, i.e. little power is supplied from the WECS to
the load. The DGS optimum speed ≈ 1164 r/min and WECS
optimum speed ≈ 650 r/min. At t ≈ 3 s, the wind velocity is
increased to 10 m/s. The WECS speed increases until it reaches
the new optimum speed (1300 r/min), and because the load is
constant, the power supplied by the DGS reduces to balance the
generated power with the power of the load. This is accompanied
by a reduction of the DGS speed reaching the new optimum
value of ≈ 835 r/min.
The dc link voltage, DGS magnetizing current, and stator voltage, for the condition corresponding to Fig. 24, are
well regulated as shown in Fig. 25(a)–(c). Even for the
relatively large variation in the power generated from the
WECS, the variation in the stator (grid) voltage is very small
[see Fig. 25(c)].
The system active power current components for the condition corresponding to Fig. 24 are shown in Fig. 26. As shown
in this graphic, the DGS q-axis rotor current decreases slightly
because the power reduction is mainly due to the speed variation. After the transient (t ≈ 25 s), the DGS is operating
below the synchronous speed and the WECS above the synchronous speed. The net power supplied from the front-end
Fig. 27. System performance for a typical wind profile. (a) Wind profile. (b)
Optimal speed and WEC real speed. (c) DGS optimal speed and real speed.
converter to the rotor of both DFIGs is increased after the wind
step.
Therefore, as shown in Fig. 26(b), the power current id also
increases.
Experimental results using a wind profile are shown in Fig. 27.
The wind profile used in the emulation is shown in Fig. 27(a).
For this test, the WECS is initially operating with a wind velocity of 5 m/s, for t < 1 s. The WECS optimum speed and the real
speed are shown in Fig. 27(b). Imposing a q-axis rotor current
given by (13), the WECS tracks well the low-frequency components of the optimum speed. In order to balance the power
variations in the energy supplied from the WECS, the rotational
speed and power generated from the DGS are also varied. The
DGS speed is shown in Fig. 27(c). Because of the relatively
slow variation in the power generated from the WECS, the corresponding variation of the DGS speed is also slow. Therefore,
the tracking of the DGS optimum speed is good. As shown in
Fig. 28(a)–(c), the regulation of the dc link voltage, magnetizing
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PEÑA et al.: WIND–DIESEL GENERATION USING DOUBLY FED INDUCTION MACHINES
213
APPENDIX
SYSTEM RATING
Diesel machine (emulated): 3.0 kW, 4 cylinder, 4 stroke,
τ2 = 0.1 s, J = 0.8 kg · m2 . Wind turbine (emulated): 1300
r/min, J = 0.9 kg · m2 , Rb = 1.25 m, GR = 1.70, vrated =
10 m/s. Doubly fed induction machines: 2.5 kW, 6 poles, Stator 220 V delta, rotor 250 V star, Rr = 0.525 Ω, Rs = 0.398,
Ls = 0.0835 H, Lm = 0.0796, Lr = 0.0825. Front-end converter: C = 2600 µF , Lf = 12 mH. To improve current filtering 30 mH is added to the rotor.
REFERENCES
Fig. 28. Control system response for the test corresponding to Fig. 27. (a) DC
link voltage. (b) DGS DFIG magnetizing current. (c) Stator voltage.
current, and stator voltage is very good during the whole wind
profile.
V. CONCLUSION
In this paper, the control strategy of a wind energy system embedded in a hybrid wind–diesel variable-speed energy system
has been proposed. The generation system uses two doubly fed
induction machines with corresponding PWM rotor inverters
connected to a common dc bus. An additional front-end converter connected to the same dc bus is employed to allow the
system to operate below and above the synchronous speed. The
control strategy for the diesel-driven generator allows indirect
control of the stator voltage (the ac system load voltage) by
regulating the stator flux magnitude via the control of the rotor
current. The load frequency is also regulated by the diesel generator by imposing the rotor currents with the slip frequency. The
wind energy system control strategy considers the generator as
connected to a grid. The electrical torque of the WECS generator
is controlled to drive the system to the rotational speed, where
maximum energy capture is obtained. Depending on the load
size and the power supplied by the WECS generator, the control system regulates the DGS rotational speed to minimize fuel
consumption.
An experimental prototype has been set up, emulating the
diesel engine and the wind turbine, in order to experimentally verify the proposed control strategy. Several tests including the connection of the WECS generator to the system,
load impacts, step changes in wind velocity, and operation
of the WECS with a realistic wind profile, have been carried
out.
The experimental results have verified that the stator voltage
regulation is very good. It has been experimentally demonstrated
that the front-end converter can be used to supply reactive power
to improve the transient performance of the magnetizing current
control when a sudden reactive power load is connected to the
stator. The performance obtained from the experimental tests is
excellent, showing the feasibility of the proposed wind–diesel
system.
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Rubén Peña (S’95–M’97) was born in Coronel,
Chile. He received the Electrical Engineering degree
from the University of Concepcion, Concepcion,
Chile, in 1984, and the M.Sc. and Ph.D. degrees
in electrical engineering from the University of
Nottingham, Nottingham, U.K., in 1992 and 1996,
respectively.
He is currently with the Electrical Engineering Department, University of Magallanes, Punta Arenas,
Chile. His current research interests include control
of power electronics converters, ac drives, and renew-
José Proboste was born in Puerto Natales, Chile, on
March 21, 1976. He received the Electrical Engineering degree from the University of Magallanes, Punta
Arenas, Chile, 2004.
He is currently a Research Assistant in the Electrical Engineering Department, University of Magallanes. His current research interests include control
of power electronics converters and ac drives.
Jon Clare (M’90–SM’04) was born in Bristol, England. He received the B.Sc. and Ph.D. degrees in
electrical engineering from the University of Bristol,
Bristol, U.K., in 1979 and 1990, respectively.
From 1984 to 1990, he was a Research Assistant
and a Lecturer at the University of Bristol, where
he was engaged in teaching and research in power
electronic systems. Since 1990, he has been with
the Power Electronics, Machines and Control Group,
University of Nottingham, Nottingham, U.K., where
he is currently a Professor in the Power Electronics
and also the Head of the Research Group. His current research interests include
power electronic converters and modulation strategies, variable-speed drive systems and electromagnetic compatibility.
Prof. Clare is a member of the Institution of Electrical Engineers. He is also
an Associate Editor for the IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS.
able energy systems.
Dr. Peña is a member of the Institute of Electrical and Electronic Engineers.
Roberto Cárdenas (S’95–M’97–SM’07) was born
in Punta Arenas, Chile. He received the Electrical Engineering degree from the University of Magallanes,
Punta Arenas, Chile, in 1988, and the M.Sc. degree in
electronic engineering and the Ph.D. degree in electrical engineering from the University of Nottingham,
Nottingham, U.K., in 1992 and 1996, respectively.
In 1989, he was a Lecturer at the University of
Magallanes, where he is currently a Professor in the
Electrical and Electronics Department. His current
research interests include control of electrical machines, variable speed drives, and renewable energy systems.
Prof. Cardenas is a member of the Institute of Electrical and Electronic
Engineers. He is the author of the paper that received the Best Paper Award
from the Industrial Electronics Society, for the best paper published in the IEEE
TRANSACTIONS ON INDUSTRIAL ELECTRONICS during 2004.
Greg Asher (M’98–SM’04–F’07) received the Graduate degree in electrical and electronic engineering
from Bath University, Bath, U.K., in 1976, and the
Ph.D. degree in bond graph structures and general
dynamic systems from University of Bath, in 1979.
In 1984, he was a Lecturer in Control in the School
of Electrical and Electronic Engineering, University
of Nottingham, Nottingham, U.K., where in 2000, he
was appointed the Professor of Electrical Drives, and
where he is currently the Head of the School of Electrical and Electronic Engineering. He is the author or
coauthor of more than 180 research papers published in various international
journals. His past research interests include motor drive systems, particularly
the control of ac machines.
Prof. Asher was a member of the Executive Committee of European Power
Electronics (EPE) Association until 2003. He is an Associate Editor of the IEEE
Industrial Electronics Society and is currently the Chair of the Power Electronics Technical Committee for the Industrial Electronics Society.
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