Wind–Diesel Generation Using Doubly Fed Induction Machines

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. 

Index Terms—Diesel-driven generators, induction generator, induction 

motor drives, wind energy. 

NOMENCLATURE 

General 

i Stator or rotor current. 

Lm ,Ls,Lr Magnetising, rotor, stator inductance. 

Rr ,Rs Rotor, stator resistance. 

Te Electrical torque. 

v Stator or rotor voltage. 

λ Stator or rotor flux. 

σ Total leakage coefficient. 

σs Stator leakage coefficient. 

DFIG Doubly fed induction generator. 

DGS Diesel generator system. 

GR Gear ratio. 

ims Magnetizing current. 

p Number of poles. 

PWM Pulsewidth modulator. 

Rb Turbine blade radius. 

Tsr Tip speed ratio. 

τ Time constant. 

V Wind velocity. 

WECS 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-00115- 

2006. 

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 

ρ Air density. 

θe Electrical angle. 

θr Rotor position angle. 

θslip Slip angle. 

ωr Induction machine rotational speed. 

ωsG Stator electrical frequency. 

ωslip Slip frequency. 

Superscripts 

∗ 

Subscripts 

Demanded value. 

(d, q) Synchronous rotating coordinates. 

G, W Diesel, wind generation quantity. 

r, s Rotor or stator quantities. 

(α, β) Stator fixed coordinates. 

I. INTRODUCTION 

VARIABLE-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 

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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. 

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. 

Fig. 2. Proposed wind–diesel system. 

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 



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] 

� � � �� 

λdsG = LmGimsG LsG 0 idsG 

= 

0 

0 LsG iqsG 

� �� 

LmG 0 idrG 

+ 

(1) 

0 LmG iqrG 

� vdsG 

vqsG 

� vdrG 

vqrG 

� 

= 

� 

= 

� 

RsG 0 

�� 

idsG 

0 

� 

0 

RsG iqsG 

�� 

−ωsG λdsG 

+ 

ωsG 

0 

� 

+ d 

dt 

� 

λqsG 

� λdsG 

λqsG 

� �� 

RrG 0 idrG 

+ 

0 RrG iqrG 

d 

� 

λdrG 

dt λqrG 

� 

�� 

0 −ωslipG λdrG 

+ 

ωslipG 0 λqrG 

(3) 

where imsG is the equivalent stator magnetizing current and 

ωslipG = ωsG − ωrG is the slip frequency. Aligning the d–q axis 

� 

� 

(2) 

of the reference frame on the stator flux vector gives 

iqrG = − LsG 

iqsG. (4) 

LmG 

Considering (1) and (2), the following expression is obtained 

for the dynamics of the magnetizing current 

dimsG 

τsG 

dt + imsG = idrG + 1+σsG 

RsG 

vdsG 

with τsG = LsG/RsG and σsG =(LsG − LmG)/LmG. 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 

� 

θslipG = θsG − θrG = ωsG dt − θrG (6) 

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 

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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(5)


Fig. 4. 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. 

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 

τ1 = 60st 60 

+ (7) 

2Nn 4N 

where st =2or 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 L 

2 

2 mG 

imsGiqrG. (8) 

LsG 

As shown in Fig. 5, the total gain Kp = K1K2 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- 

Fig. 5. Variation of Kp as a function of power and rotational speed. 

Fig. 6. Fuel consumption versus power at various rotational speed. 

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 = ωrGTeG. 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. 



using (12) and (10), for optimal power capture, the torque component 

current (iqrW) is regulated as 

i ∗ qrW = 2LsWkoptω 2 rW 

3pL 2 mW imsW 

. (13) 

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] 

Tm =0.5πρCt(Tsr,β)R 3 b V 2 

Pm =0.5πρCp(Tsr,β)R 2 b V 3 

(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 (Cpmax) [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 = koptω 2 rW 

(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] 

θe = tan −1 

� � 

λβ sW 

λαsW 

� 

λαsW = 

� 

λβ sW = 

(vαsG − RsWiαsW)dt 

(vβ sG − RsWiβ sW) (11) 

where λαsW and λβ sW are the α − β stator flux components. 

The electrical torque is obtained as [10] 

TeW =3 p L 

2 

2 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 

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 ωre =(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 + sGPsG). (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 = vdLidL + vqLiqL 

(17) 

where vdL, vqL, idL, and iqL are the d–q components of the 

load voltage and current. Because the stator resistance in each 

mW 

(imsWiqrW) 

LsW 

(12) 

machine is relatively small, vdsW ≈ vdsG ≈ vdL ≈ 0 and the 

load power can be obtained as PL ≈ vqLiqL. For simultaneous 



Fig. 8. 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 

∆ωr 

= 4LsWkoptωrW0 

3pL2 mWimsW0 . (19) 

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.InFig.9idL, 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 

Fig. 9. Small-signal model for stator flux regulation of the DGS generator. 

Fig. 10. 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 



model and the methodology discussed in [22] and [24], the 

reference speed of the WECS generator (ω∗ rW ) is calculated in 

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. 

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 

Kpe 

ω(s) = 

−sτ1 

1 

u(s) − 

(1 + sτ2)(sJ + B) (sJ + B) TeG(s) (21) 

using (8) and the bilinear transform, (21) can be discretized to 

ω(z) = 

A = 

Kp [(z +1)Ts] 2 z −N t 

(2τ2 + Ts)(2J + BTs)(z − A)(z − C) u(z) 

(z +1)Ts 

− 

J(z − A) 3p 

2 

(2J − BTs) 

(2J + BTs) 

L2 mG 

i 

LsG 

∗ msG i ∗ qrG(z) 

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. 11. Control system performance for load steps. (a) Speed tracking performance. 

(b) Estimated fuel consumption. 

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 =50s, 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 



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 

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 



Fig. 16. System performance with reactive power compensation. (a) Speed 

tracking. (b) Magnetizing current and dc link voltage. 

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. 18. Steady state operation at sub (a, b) and supersynchronous (c, d) speed. 

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%. 



Fig. 20. Rotor and stator current corresponding to the test of Fig. 19. (a) 

Reactive power component. (b) Active power current components. 

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 

=0for 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 

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. 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 



Fig. 24. Speed system performance for a step increase in wind velocity. 

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,optV 

Rb 

GR (23) 

where Tsr,opt is the tip speed ratio corresponding to Cpmax. 

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. 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. 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


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. 

APPENDIX 

SYSTEM RATING 

Diesel machine (emulated): 3.0 kW, 4 cylinder, 4 stroke, 

τ2 =0.1 s, J =0.8 kg· m 2 . Wind turbine (emulated): 1300 

r/min, J =0.9 kg· m 2 , 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 =12mH. To improve current filtering 

30 mH is added to the rotor. 

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[24] R. Cárdenas, R. Peña, G. Asher, and J. Clare, “Emulation of wind turbines 

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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 renewable 

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. 

JoséProbostewas 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. 

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.

Wind–Diesel Generation Using Doubly Fed Induction Machines

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