This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
Switching Strategies for Fault Tolerant
Operation of Single DC-link Dual Converters
Jose A. Restrepo, Member, IEEE, Alberto Berzoy, Antonio E. Ginart, Senior Member, IEEE,
Jose M. Aller, Ronald Harley, Fellow, IEEE, and Thomas Habetler, Fellow, IEEE
Abstract
This paper proposes a method that allows the operation with direct torque control under incipient faults of the
power switching devices in single DC-link dual-inverters. Once the fault onset of the switching device is detected
the troubled devices are put under trigger suppression to regain operation. The proposed control algorithm makes
use of additional states, termed opportunistic states to maintain control of the induction machine. Simulations and
experimental results confirm the capabilities of the technique for up to six devices under trigger suppression.
Index Terms
AC motor drives, Fault tolerance, Drives, Inverters
J. A. Restrepo, A. Berzoy and J. M. Aller are with the Universidad Simón Bolı́var, Caracas, Venezuela (e-mail: restrepo@ieee.org; aberzoy@usb.ve; jaller@usb.ve).
A. E. Ginart is with Impact Technologies, LLC, Atlanta, GA 30308 USA (e-mail: aginart@ieee.org).
R. G. Harley is with the School of Electrical and Computer Engineering, Georgia Institute of Technology, Atlanta, GA 30332. He is also
Professor Emeritus and an Honorary Research Associate at the University of KwaZulu-Natal, Durban, South Africa. (e-mail: rharley@gatech.edu).
T. Habetler is with the School of Electrical and Computer Engineering, Georgia Institute of Technology, Atlanta, GA, 30332 USA (email: rharley@gatech.edu, thabetler@ee.gatech.edu).
Manuscript received , ; revised , .
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
1
Switching Strategies for Fault Tolerant
Operation of Single DC-link Dual Converters
I. I NTRODUCTION
The migration from traditional electromechanical and hydraulic drives in critical equipment to power electronic
drives raises concern about the reliability of critical equipment controlled by power electronic components and
drives. Examples of this migration can be observed in the aeronautic industry as in the proposed “fly by wire”
(aircraft control) initiative and in the automotive industry with the electric steering wheel concept. Furthermore, there
is a strong push in the aeronautics industry to produce complete solid-state technology based power converters and
for military applications to implement the main parts of hybrid vehicles using a similar approach. The same trend
discussed above can be observed in the main power drive where reliability is essential. Additionally, Power electronic
converters, drives and related modern electric machines continue to play a more critical role in the performance
and overall operation of ships, aircraft, and ground vehicles. The electric power and propulsion systems developed
for the military are being rapidly modernized [1]. The new trend is to create designs that integrate motor drive
systems (power electronics), into a monolithic structure for propulsion and generation, power transformers, new
storage devices, power cells, etc. These new technologies present new challenges in methods of operation and
automated health management in order to accomplish the goal of creating a more versatile, robust, efficient, and
reliable system.
The problem of increasing the system reliability can be tackled following two strategies. The first approach is
to improve the performance of standard topologies and early/post diagnostics of the power devices [2], [3]. The
second approach is to develop new fault tolerant topologies for power drives. Standard power drives are based on a
controller with the traditional six transistor structure that drives induction or synchronous Permanent Magnet (PM)
machines. Fault tolerant operation of multilevel converters has been investigated in literature [4]–[9]. In this work,
the use of a dual converter is proposed and analyzed as an alternative for a more reliable operation of drives under
faulty conditions. This topology has been previously described in the literature, mainly for its multilevel operation
[10]–[14]. As with any multilevel topology, the power and voltage ratings of the inverter are increased without
increasing the voltage rating of the switching device, and the dv/dt of the converter is reduced [15]. Also, dual
converters have the drawback of doubling the switching devices count when compared with conventional two-level
converters, but present a higher potential for extended operation in critical operation [16], [17]. The analysis and
advantages of different fault tolerant strategies in single DC-link dual converters are presented in this paper.
II. E ARLY DIAGNOSTIC TECHNIQUE
There are several methods for the early identification of the degradation of power electronic devices. These
diagnostic techniques allow and enhance the application of fault tolerant techniques [4], [18], [19]. Early pre-failure
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
2
characterization of switching devices is mainly based on the operating condition of a power drive. The operating
frequency band of a standard power drive can be divided in three ranges. The low frequency range, usually below
200 Hz, provides the main electrical signal for the motor and achieves the control objective of the power drive. The
second range, between 10 kHz and 200 kHz is due to the Pulse Width Modulation (PWM) signal, and is used to
modulate the main or fundamental signal of the power drive. The third band corresponds to frequencies greater than
1-MHz and is the consequence of system’s resonance and high frequency harmonics in the PWM step waveforms.
This high frequency band is of main interest for modeling and diagnosing transistor aging through changes in the
“ringing characteristics” of the power devices [20]. In the ringing characterization method specific frequencies are
identified and tracked for changes such as the devices degradation over time [21]. In the ringing characterization
method the system’s response to high frequency components injected during the inverter switching can be used as
a signature to track the aging in the power devices. These high frequency components interact with the parasitic
elements present in all power devices, giving rise to the “ringing” response that changes with the deterioration of
the device. When the early diagnostics fail, a strategy based on the de-saturation protection used in intelligent drives
is employed. In this case the protection circuitry triggers a shutdown of the troublesome device, and this period
allows the recovery of the device thanks to its self-healing properties [18], [22]–[24].
A supervisory fault detector system monitoring the switching devices is required, and depending on the fault
detected a corresponding corrective action should be taken by the controller for the different type of faults, such
as shorted or open switching device, missing or randomly suppressed gate signals, etc. This paper proposes what
action to take when a fault requires trigger suppression of the relevant switching devices.
Once a satisfactory early diagnostic technique is in place the goal is to continue the operation of the converter,
while the troubled switching device is under trigger suppression.
III. D UAL CONVERTER
A diagram of the converter being analyzed in this paper, for switch suppression operation, is shown in Fig. 1. The
particular topology of this dual converter is comprised of two three-phase two-level inverters sharing the same DC
link. In this topology the dual converter can be considered as a six phase converter, where the number of possible
states are 46 =4096, of which 46 -36 =3367 are forbidden states because they short-circuit the DC link, from the valid
states 36 -26 =665 have at least one “fly-wheeling” floating branch and only 26 =64 are well established and in general
considered as usable states for normal operation, resulting in 19 different voltage space vectors shown in Fig. 2.
For simplicity the power drives are some times analyzed assuming switching with on and off states, and assuming
the loads to be without regeneration capabilities or purely resistive. This generates what has been previously called
usable states. However, when the load has regenerative capabilities or energy storage properties such as inductances
or electrical machines, the energy storage provides additional control alternatives expressed in additional transient
usable states. In the case of fault tolerant strategies, these additional states could provide an important tool for
optimization during trigger suppression. In this paper these states are defined as “opportunistic states” since they
can only be used when they become available, and the control algorithm can favor their occurrence since they are
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
3
determined by current circulating in the winding of the machine under control. When these “opportunistic states”
are available, the application of “regenerative vectors” can be achieved.
Since the two sub-converters making up the dual-converter share a common DC-link, the modulation strategies
used to synthesize the space vector in each sub-converter need to avoid the injection of zero sequence components,
as they will result in corresponding zero sequence currents flowing through the load [25], [26].
In this dual converter topology the voltage levels in each phase of the induction machine are: {+VDC }, obtained
when the associated branch in sub-converter 1 is in high level and its paired branch in sub-converter 2 is in low
level, and when there is enough energy stored in the machine inductances the following voltage levels can be
applied; {0}, when the both branches feeding the phase winding are simultaneously in high or low level in both
sub-converters, and {-VDC }, when the associated branch in sub-converter 1 is in low level and its paired branch in
sub-converter 2 is in high level. In this way the space vectors in the dual converter can be classified in the following
sets.
•
First a set of six “positive” small size voltage space vectors obtained with the combination of voltage levels
{0, +VDC }.
•
A second set of six “negative” small size voltage space vectors obtained with the combination of voltage levels
{-VDC , 0}.
•
A Third set of six medium size voltage space vectors obtained with one phase set to -VDC , the other phase
to 0, and the last phase to +VDC between each winding terminals.
•
A fourth set of six large size voltage space vectors obtained using voltage levels VDC and +VDC between the
winding terminals.
•
Finally a set of three zero magnitude voltage space vectors obtained when the voltage in the three winding
terminals are the same {(-VDC ,-VDC ,-VDC ), (0, 0, 0), (+VDC , +VDC , +VDC )}.
The resulting space vector locations are shown in Fig. 3. The zero sequence voltage, responsible for the zero
sequence current in the induction machine phases, is defined as [25]:
vZS =
1
(va1 a2 + vb1 b2 + vc1 c2 )
3
(1)
The control algorithm used in this paper for healthy as well as switch suppressed operation is the classical direct
torque control (DTC) as proposed by Takahashi and Noguchi [27]. During normal operation DTC can take advantage
of the additional space vectors produced by the dual converter, and produce less ripple in the system’s state variables
[28]–[30]. However, in this paper the focus is on the operation of the machine during switch suppression. Thus,
the classical DTC algorithm using the medium size space vectors is used for healthy switching operation.
IV. DTC ALGORITHM
The classical direct torque control (DTC) was originally proposed by I. Takahashi and T. Noguchi [27]. It is
based on the computation of the instantaneous stator flux and electric torque in the machine. The calculated values
are compared to the flux and torque references to produced error signals that together with the information of the
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
4
angular stator flux location are employed in the switching state selection for the inverter. In the classical DTC
algorithm the hexagonal space vector area is divided in six zones, the requirements of flux and torque for each
zone and the corresponding space vector for two level-converters are summarized in Table I. In this paper the space
vectors are obtained using Clarke’s transformation,
f = fx + jfy =
2
fa α0 + fb α1 + fc α2
3
(2)
where α = ej2π/3 . The stator flux linkage Ψs can be obtained by integrating the stator equation of the induction
machine model [27].
s =
Ψ
vs − Rsis dt
(3)
and the electric torque by multiplying the stator flux linkage and the stator current space vectors.
s × is
τe = Ψ
(4)
In the simplified algorithm, used in this paper for normal operation, one branch is at -VDC , another at 0 V while
the last one is at VDC . Therefore, the voltage supplying the machine does not contain a zero sequence component.
The states that satisfy this combination are termed medium size voltage space vectors, and from Fig. 2 they are
vm1 = {14, 36}, vm2 = {24, 35}, vm3 = {21, 65}, vm4 = {41, 63}, vm5 = {42, 53}, vm6 = {12, 56}, and the
null vector v0 = {00, 11, 22, 33, 44, 55, 66, 77}. Figure 4 shows the medium size space vectors and the resulting
hexagonal region covered by this set of space vectors. From (1) it follows that an advantage of the medium size
vectors for the dual converter is that they do not inject zero sequence voltage, hence there are no zero sequence
a1H
b1H
c1H
a2H
b2H
c2H
IM
VDC
a2
a1
b2
b1
c1
a1L
b1L
c1L
c2
a2L
b2L
c2L
Fig. 1: Diagram of the dual converter with one DC supply feeding an open-end induction machine.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
5
25
21
20
27
65
23
45
61
41
63
01
05
31
35
24
75
64
15
26
34
04
30
14
74
37
36
06
66 00 55
07 44
10 76
33
17 54
7011 22 77
60
32
03
02
12
40 73
13 72
47 62
46 57
56
51
50
71
67
43
42
16
52
53
Fig. 2: Voltage space vectors with their respective switching states in base 8.
vm2
vl3
vs3
vm3
vl2
vs2
vs4
vl4
vm4
vl1
vs1
vs5
vl5
vm1
vm6
vs6
vm5
vl6
Fig. 3: Sets of voltage space vectors in the dual converter.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
6
Z2
vm2
Z3
Z1
vm
3
v m1
Z6
vm
6
v m4 Z 4
vm5
Z5
Fig. 4: Medium size space vectors.
currents injected into the load. Any other selection of space vector will inject a zero sequence component and its
influence on the load needs to be considered before using them.
TABLE I: Switch selection for classical DTC [27].
OΨ
↓
↑
Oτ
Z1
Z2
Z3
Z4
Z5
Z6
↓
v5
v6
v1
v2
v3
v4
=
v0
v0
v0
v0
v0
v0
↑
v3
v4
v5
v6
v1
v2
↓
v6
v1
v2
v3
v4
v5
=
v0
v0
v0
v0
v0
v0
↑
v2
v3
v4
v5
v6
v1
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
7
TABLE II: Common mode voltage for the positive and negative small voltage space vectors.
va
vb
vc
VDC
VDC
VDC
0
0
0
-VDC
-VDC
-VDC
-VDC
1
V
3 DC
VDC
0
0
0
-VDC
-VDC
VDC
VDC
0
0
0
-VDC
0
VDC
0
-VDC
0
-VDC
0
VDC
VDC
-VDC
0
0
0
0
VDC
-VDC
-VDC
0
VDC
0
VDC
0
-VDC
0
voltage space vector
vZS
VDC
vs0
vs1
vs2
0
- 23 VDC
2
V
3 DC
- 13 VDC
1
V
3 DC
vs3
- 23 VDC
2
V
3 DC
vs4
- 13 VDC
1
V
3 DC
vs5
vs6
- 23 VDC
2
V
3 DC
- 13 VDC
V. DTC OPERATION FOR SWITCH TRIGGER SUPPRESSION
The large amount of switching states in the dual converter allow the implementation of different machine control
algorithms. For this work the DTC is chosen for implementation. Depending on the amount of switches under trigger
suppression there is a certain number of voltage space vectors available to perform the DTC on the machine, or to
enforce the conditions to use all the space vectors. Since the DTC uses a computation of the stator flux linkage,
the use of (3) requires the voltage at the machine terminals. The use of the switching state of the dual inverter
for estimation of the stator voltage assumes that each terminal of the machine is connected at +VDC or 0. This
restriction limits the failure to one switch per branch. A fully asymmetric operation of the inverter is considered in
the case of six switches with trigger suppression, i.e. trigger suppression in three upper switches in one sub-converter
and trigger suppression in three lower switches in the other sub-converter. The classical DTC algorithm remains
unchanged for healthy or trigger-suppressed operation, but for the dual converter, the healthy case uses the medium
size space vectors versus the use of the small size space vectors in the trigger suppression mode. This requires a
30 degrees rotation in the region selection of the stator flux (Z1 , Z2 , ..., Z6 ) shown in Fig. 4.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
8
A. Trigger suppression in one and two devices
In this case, trigger suppression is applied to one or two devices. The effect of trigger suppression in one switch is
that the current does not flow in the direction controlled by the suppressed device. However, if regenerative operation
allows for a valid state to be established then all the original voltage space vectors are available for the control of
the load. When the conditions do not allow the establishment of a known state, trigger suppression in one device
affects 32 out of the 64 usable states shown in Fig. 3, but due to the redundancy of the dual converter only five
voltage space vectors out of 19 are affected. An additional trigger suppression in the complementary sub-converter
and complementary level of the former device affects 16 additional states, corresponding to five additional voltage
space vectors. The effects of the trigger suppression for one and two power devices in phase ‘a’ are illustrated in
Fig. 5. The devices under trigger suppression are: sub-converter 1, high side phase ‘a’ (a1H ) and sub-converter 2,
low side phase ‘a’ (a2L ). Trigger suppression and no regeneration in one of the aforementioned devices prevent the
establishment of the outer space vectors in Fig. 5 (vl2 , vm1 , vl1 , vm6 , vl6 ), and the additional trigger suppression in
the second device prevent the establishment of the additional space vectors (vm2 , vs2 , vs1 , vs6 , vm5 ). Some states
associated with voltage space vectors (v0 , vs3 , vs4 , vs5 ) are affected, but the redundancy in the number of states
in the dual converter allows the establishment of the associated voltage space vectors. If the second device with
trigger suppression is (a2H ), the voltage space vectors affected are (vl3 , vm3 , vl4 , vm4 , vl5 ). A proper control of the
load with DTC can be achieved through the injection of a zero sequence component in the load current. Then all
the switching states are available for control thanks to regenerative operation in the suppressed devices. This zero
sequence injection can be controlled with the proper selection of the switching state and its associated common
mode voltage.
Due to the symmetry of the dual converter, the pattern of voltage space vectors affected by trigger suppression
in one or two devices are rotated by the same amount as the associated phase. Some resulting patterns for two
devices with trigger suppression are shown in Fig. 6.
For the case depicted in Fig. 5 the voltage space vectors available for DTC are those with phase ‘a’ at −VDC ,
shown in Table III:
For two devices associated with the same phase and the same sub-converter there is no possible combination of
remaining switching devices that can impress the desired current value in that phase. Therefore, no opportunistic
states are available for control and a different control strategy should be considered.
B. Trigger suppression in six switches (asymmetric operation)
For this kind of operation only 50% of the switching devices are available for controlling the current in the load.
In this case the resulting power stage topology is similar to the asymmetric converter employed in the control of
switched reluctance machines (SRM) [31]. In this work, the operation of the converter under this circumstances is
termed asymmetric operation, and is considered when trigger suppression is done in complementary devices of the
same phase as shown in Fig. 7. Different modes of operation for the induction machine fed by an asymmetrical
converter have been presented in [32]. In this work the DTC algorithm is employed using the positive and negative
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
9
vm2
vl3
vs3
vm3
vl2
vs2
vs4
vl4
vm4
vl1
vs1
vs5
vl5
vm1
vm6
vs6
vm5
vl6
Fig. 5: Voltage space vectors for one and two trigger suppressed devices ‘a1H ’ and ‘a2L ’.
a1H or a2L
a1L or a2H
b1H or b2L
a1H and a2L b1H and b2L
c1H and c2L a1H and a2H b1H and b2H c1H and c2H a1H and b1H
b1H and c1H c1H and a1H a1H and c2H a1H and b2H a1H and c2L
Fig. 6: Example of different space vectors affected in the case of trigger suppression in one and two power devices.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
10
TABLE III: States available for trigger suppression in two switches of phase ‘a’ as presented in Fig. 5.
a
b
c
voltage space vector
vZS
-VDC
-VDC
-VDC
vs0
-VDC
-VDC
-VDC
0
vs5
- 23 VDC
-VDC
-VDC
+VDC
vl5
- 13 VDC
-VDC
0
-VDC
vs3
- 23 VDC
-VDC
0
0
vs4
- 13 VDC
-VDC
0
+VDC
vm4
0
-VDC
+VDC
-VDC
vl3
- 13 VDC
-VDC
+VDC
0
vm3
0
-VDC
+VDC
+VDC
vl4
1
V
3 DC
or regenerative set of space vectors. Under this operating conditions all the switching states become available thanks
to the opportunistic states described above.
a1H
b1H
c1H
a2H
b2H
c2H
a2L
b2L
c2L
IM
VDC
a1L
b1L
c1L
Fig. 7: Diagram of the dual converter under six switches trigger suppression.
For trigger suppression in the devices shown in Fig. 7 the only known state that can be impressed on the
machine at any time is {1, 1, 1}, corresponding to the null vector. All the other states require a current flow in each
particular branch to operate with regenerative states. Once the establishment of regenerative states is assured all the
64 states and 19 voltage space vectors become available for DTC. However, the regenerative states are obtained at
the expense of zero sequence injection into the load. The strategy to control the zero sequence injection is similar
to the one employed in the case of trigger suppression with one and two devices. The opportunistic states (space
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
11
TABLE IV: Induction machine parameters.
1 HP, 220/380V, 3.3/1.92 A, fpnom =0.78, 50 Hz, 1400 rpm
Stator Resistance (Rs )
3.8 Ω
Stator Inductance (Ls )
282.7 mH
Rotor Resistance (Rr )
2.5 Ω
Rotor Inductance (Lr )
282.7 mH
Mutual Inductance (Lm )
270.5 mH
vectors synthesized with 0 and −1 states) are used until a predefined threshold is reached. Then a space vector is
synthesized using states 0 and 1 which increases the energy in the machine leakage reactances and goes back to
the use of opportunistic states.
VI. S IMULATION AND EXPERIMENTAL RESULTS
The proposed fault-tolerant strategies are first simulated in SIMULINK using the Piece-wise Linear Electrical
Circuit Simulation (PLECS) package for the machine and power electronics modules. The circuit diagram in PLECS
is shown in Fig. 8, while the control strategy is implemented in SIMULINK. The parameters of the machine used
for the simulation and experimental tests are presented in Table IV, the sampling rate of the controller is fixed at
10 kHz.
For the experimental test, the proposed algorithm is implemented on a custom build floating-point DSP-(ADSP21369) based test rig, shown in Fig. 9. The power stage in each sub converter uses six 50 A, 1200V insulated
gate bipolar transistors (IGBTs) with a 2200 μF, 450 V capacitor in the DC-link. The interface between the DSP
processing unit and the peripherals (power devices and sensors) is done with a SPARTAN-Xilinx field programmable
gate array (FPGA). This interface is programmed to enable or disable each IGBT present in the power stage, using
a 6-bit mask, one mask per sub-converter. Also, the switching frequency for the controller, and for sampling the
sensors (10 kHz), is programmed in this FPGA. Once the number of suppressed IGBTs is known, the control
algorithm selects the proper common mode injection strategy to ensure the use of all the available space vectors,
thanks to the availability of opportunistic states. In all cases the DTC operates with TABLE I, but using the space
vectors shown in Fig. 3, depending on the operation mode, medium size space vector for healthy operation and the
small size space vectors for operation under trigger suppression.
In the first test the single DC-link dual converter is simulated for healthy operation with the set of medium voltage
space vectors. The advantage of this selection for healthy operation is that there is not zero sequence injection at
any time. Figure 10 shows the simulated stator currents, stator flux, electric torque and rotor speed for a healthy
operation of the dual converter. It can be seen from Fig. 10a that no zero sequence component is present in the stator
current. Also, since the classical DTC algorithm is employed, the torque shows the large ripple that is characteristic
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
12
1
Iabc
gate signals
1
3
Vdc
V
2-level
IGBT
conv. 2
V_dc
A
Ia
A
Ib
2-level
IGBT
conv. 2
A
Ic
Tm m
IM
*
/
-c-
2
wr
Fig. 8: Circuit diagram.
of this kind of controllers. Figure 11 shows the experimental results obtained for healthy operation with DTC with
the medium size set of space vectors, the experimental stator current waveforms confirm that for healthy operation
with the selected set of space vectors there is not zero sequence injection.
Figure 12 shows the corresponding simulated state variables for trigger suppression in one or two switches
associated to one phase winding (a1L and a2H ), in this case the current flowing through phase ‘a’ can take only
positive values, and to insure known switching states in the dual converter the phase current has to be positive
at all time (isa > 0). If the trigger suppression is in the complementary switches (a1H and a2L ) the current will
be required to be negative at all times to ensure operation of the dual converter under fault operation with known
states. For trigger suppression in devices (a1L and a2H ) the control strategy is to use the set of regenerative space
vectors whenever the current flowing through the affected phase is above certain threshold that insures the use of
opportunistic states, otherwise use the non regenerative set of small space vectors to increase the stored energy in
the affected winding for use of opportunistic states in subsequent control cycles.
Although trigger suppression in six switches will affect all the space-vectors synthesized by the dual converter,
there are still known states available to store energy in the machine inductances. This brings the possibility of
achieving proper control of the dual converter if enough energy can be stored in the machine inductances so that
through regeneration the opportunistic states become available. Thereby all the space vectors can be used and the
switching state of the dual converter is known. Figure 14 shows the simulated state variables for the dual converter
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
13
SubConverter 1
SubConverter 2
DSP
Load
Open End
Ind. Machine
Fig. 9: Experimental test rig for the dual converter.
Electric Torque (N.m)
Ψsy (Wb)
Current (A)
1.5
Time (s)
(a) Currents
Ψsx (Wb)
(b) Stator Flux
1.25
1
0.75
0.5
0.25
0
0
0.2
0.4
0.6
Time (s)
0.8
1
(c) Electric Torque
Fig. 10: Simulated stator currents, stator flux, rotor speed and torque for normal operation with DTC using medium
size voltage space vectors.
operating under six switches with trigger suppression. The switches with trigger suppression are (a, b, c)1L and
(a, b, c)2H . The control strategy used in this case is similar to the one used previously and consists of applying
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
14
0.4
1.5
Electric Torque (N.m)
0.3
Ψsy (Wb)
0.2
0.1
0
-0.1
-0.2
-0.3
-0.4
-0.4 -0.3 -0.2 -0.1
0
1.25
1
0.75
0.5
0.25
0
0.1 0.2 0.3 0.4
0
Ψsx (Wb)
(a) Currents
0.2
0.4
0.6
0.8
1
Time (s)
(b) Stator Flux
(c) Electric Torque
Fig. 11: Experimental stator currents, stator flux, rotor speed and torque for normal operation with DTC using
medium size voltage space vectors.
0.4
1.5
0.2
Ψsy (Wb)
4
3
Current (A)
Electric Torque (N.m)
0.3
2
1
0
0.1
0
-0.1
-0.2
-1
-0.3
-2
-3
0
0.04
0.08
0.12
Time (s)
0.16
0.2
(a) Currents
-0.4
-0.4 -0.3 -0.2 -0.1
0
0.1 0.2 0.3 0.4
Ψsx (Wb)
(b) Stator Flux
1.25
1
0.75
0.5
0.25
0
0
0.5
1
Time (s)
(c) Electric Torque
Fig. 12: Simulated stator currents, stator flux, rotor speed and torque for trigger suppression in a1L and a2H .
the opportunistic states as long as there is enough energy stored in the machine leakage inductances, and then
switch to the complementary states to increase the stored energy when the energy is not sufficient to use the
opportunistic states. For the case under test the opportunistic states can be used if the phase currents are above a
predefined minimum value, i.e. each phase circuit is operating in a continuous switching mode. Figure 15 shows the
experimental results, that confirm the results obtained previously in the simulation. Although there is a high zero
sequence voltage and current component injected in the machine terminals, this zero sequence does not produce any
braking torque because there is not zero sequence component in the stator flux. The disadvantage of this operating
mode is that for the same electrical torque the machine is operating with higher rms currents and hence with
increased power losses and reduced system efficiency. This results in higher thermal stress and reduced life span
for the operating power components.
Figure 16 shows the instantaneous transition between healthy operation and switch trigger suppression in two
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
15
0.4
1.5
Electric Torque (N.m)
0.3
Ψsy (Wb)
0.2
0.1
0
-0.1
-0.2
-0.3
-0.4
-0.4 -0.3 -0.2 -0.1
0
1.25
1
0.75
0.5
0.25
0
0.1 0.2 0.3 0.4
0
0.2
Ψsx (Wb)
(a) Currents
0.4
0.6
0.8
1
Time (s)
(b) Stator Flux
(c) Electric Torque
Fig. 13: Experimental stator currents, stator flux, rotor speed and torque for trigger suppression in a1L and a2H .
0.4
1.5
0.2
Ψsy (Wb)
5
Current (A)
4
3
2
0.1
0
-0.1
1
-0.2
0
-0.3
-1
Electric Torque (N.m)
0.3
0
0.05
0.1
Time (s)
0.15
-0.4
-0.4 -0.3 -0.2 -0.1
0.2
0
1.25
1
0.75
0.5
0.25
0
0.1 0.2 0.3 0.4
0
0.2
Ψsx (Wb)
(a) Currents
0.4
0.6
0.8
1
Time (s)
(b) Stator Flux
(c) Electric Torque
Fig. 14: Simulated stator currents, stator flux, rotor speed and torque for trigger suppression in six switches,
{a1L , b1L , c1L } and {a2H , b2H , c2H }.
0.4
1.5
Electric Torque (N.m)
0.3
Ψsy (Wb)
0.2
0.1
0
-0.1
-0.2
-0.3
-0.4
-0.4 -0.3 -0.2 -0.1
0
0.1 0.2 0.3 0.4
Ψsx (Wb)
(a) Currents
(b) Stator Flux
1.25
1
0.75
0.5
0.25
0
0
0.2
0.4
0.6
0.8
1
Time (s)
(c) Electric Torque
Fig. 15: Experimental stator currents, stator flux, rotor speed and torque for trigger suppression in six switches,
{a1L , b1L , c1L } and {a2H , b2H , c2H }.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
16
0.4
1.5
Electric Torque (N.m)
0.3
Ψsy (Wb)
0.2
0.1
0
-0.1
-0.2
-0.3
-0.4
-0.4 -0.3 -0.2 -0.1
0
0.1 0.2 0.3 0.4
Ψsx (Wb)
(a) Currents
1.25
1
0.75
0.5
0.25
0
0
0.2
0.4
0.6
0.8
1
Time (s)
(b) Stator Flux
(c) Electric Torque
Fig. 16: Experimental stator currents, stator flux, rotor speed and torque for the transition between healthy operation
and trigger suppression in a1H and a2L .
power devices. Since the stator flux and torque are maintained constant before and after the change in operating
mode, the transition of the remaining state variables towards steady state is fast.
Table V shows the rms value of the stator currents and the THD for the dual-converter when operating under
healthy condition, as well as with two switches under trigger suppression and also with six switches under trigger
suppression. From these results it is clear that there is an increase in both rms current and the THD in most phases,
which in turn results in an increase in the machine’s temperature for the same output torque.
TABLE V: RMS current and THD for healthy operation, two switches trigger suppression and six switches trigger
suppression.
Healthy
2sts
6sts
RM S (A)
T HD
ia
0.9731
0.0647
ib
0.9397
0.0574
ic
1.0160
0.0608
ia
0.2913
−−−
ib
1.7797
0.2143
ic
1.8834
0.2325
ia
1.8883
1.4729
ib
1.7342
1.2646
ic
1.7967
1.4013
VII. C ONCLUSION
In this paper a fault-tolerant DTC algorithm for a single DC-link dual-converter feeding an induction machine has
been proposed and verified for trigger suppression in up to six power devices. This extended operation compensates
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
17
and improves over the apparent reduction in reliability due to the increase in power devices. The proposed algorithm
shows that operation of the machine is still possible, and although there is a penalty in the thermal stress of the
system. It presents an alternative to maintain operation in critical systems. The experimental test at different operating
conditions verify the proposed control strategy and the simulation results.
R EFERENCES
[1] Peng Xiao, G. Venayagamoorthy, and K. Corzine, “Seven-level shunt active power filter for high-power drive systems,” IEEE Trans. Power
Electron., vol. 24, no. 1, pp. 6 –13, Jan. 2009.
[2] Bin Lu and S. Sharma, “A literature review of igbt fault diagnostic and protection methods for power inverters,” IEEE Trans. Ind. Appl.,
vol. 45, no. 5, pp. 1770 –1777, Sep./Oct. 2009.
[3] Sangshin Kwak, “Fault-tolerant structure and modulation strategies with fault detection method for matrix converters,” IEEE Trans. Power
Electron., vol. 25, no. 5, pp. 1201 –1210, May. 2010.
[4] Q.-T. An, L.-Z. Sun, K. Zhao, and L. Sun, “Switching function model-based fast-diagnostic method of open-switch faults in inverters
without sensors,” IEEE Trans. Power Electron., vol. 26, no. 1, pp. 119 –126, Jan. 2011.
[5] C. Turpin, P. Baudesson, F. Richardeau, F. Forest, and T. Meynard, “Fault management of multicell converters,” IEEE Trans. Ind. Electron.,
vol. 49, no. 5, pp. 988 – 997, Oct. 2002.
[6] F. Richardeau, P. Baudesson, and T. Meynard, “Failures-tolerance and remedial strategies of a pwm multicell inverter,” IEEE Trans. Power
Electron., vol. 17, no. 6, pp. 905 – 912, Nov. 2002.
[7] X. Kou, K. Corzine, and Y. Familiant, “A unique fault-tolerant design for flying capacitor multilevel inverter,” IEEE Trans. Power Electron.,
vol. 19, no. 4, pp. 979 – 987, Jul. 2004.
[8] A. Chen, L. Hu, L. Chen, Y. Deng, and X. He, “A multilevel converter topology with fault-tolerant ability,” IEEE Trans. Power Electron.,
vol. 20, no. 2, pp. 405 –415, Mar. 2005.
[9] M. Ma, L. Hu, A. Chen, and X. He, “Reconfiguration of carrier-based modulation strategy for fault tolerant multilevel inverters,” IEEE
Trans. Power Electron., vol. 22, no. 5, pp. 2050 –2060, Sep. 2007.
[10] D. Casadei, G. Grandi, A. Lega, C. Rossi, and L. Zarri, “Switching technique for dual-two level inverter supplied by two separate
sources,” in Applied Power Electronics Conference, APEC 2007 - Twenty Second Annual IEEE. Anaheim, CA, USA: IEEE, Feb. 2007,
pp. 1522–1528.
[11] E. G. Shivakumar, K. Gopakumar, S. K. Sinha, A. Pittet, and V. T. Ranganathan, “Space vector PWM control of dual inverter fed openend winding induction motor drive,” in Applied Power Electronics Conference and Exposition, 2001. APEC 2001. Sixteenth Annual IEEE,
vol. 1. Anaheim, CA: IEEE, Mar. 2001, pp. 399–405.
[12] V. T. Somasekhar, S. Srinivas, and K. K. Kumar, “Effect of zero-vector placement in a dual-inverter fed open-end winding induction motor
drive with alternate sub-hexagonal center PWM switching scheme,” IEEE Trans. Power Electron., vol. 23, no. 3, pp. 1584–1591, May.
2008.
[13] S. Srinivas and V. T. Somasekhar, “Space-vector-based PWM switching strategies for a three-level dual-inverter-fed open-end winding
induction motor drive and their comparative evaluation,” IET Electric Power Applications, vol. 2, no. 1, pp. 19–31, Jan. 2008.
[14] H. Stemmler and P. Guggenbach, “Configurations of high-power voltage source inverter drives,” in Power Electronics and Applications,
1993., Fifth European Conference on, Brighton, Sep. 1993, pp. 7–14.
[15] S. Mekhilef and M. Abdul Kadir, “Voltage control of three-stage hybrid multilevel inverter using vector transformation,” IEEE Trans.
Power Electron., vol. 25, no. 10, pp. 2599 –2606, Oct. 2010.
[16] W. dong Jiang, S. wu Du, L. chen Chang, Y. Zhang, and Q. Zhao, “Hybrid PWM strategy of SVPWM and VSVPWM for NPC three-level
voltage-source inverter,” IEEE Trans. Power Electron., vol. 25, no. 10, pp. 2607 –2619, Oct. 2010.
[17] L. Maharjan, T. Yamagishi, H. Akagi, and J. Asakura, “Fault-tolerant operation of a battery-energy-storage system based on a multilevel
cascade pwm converter with star configuration,” IEEE Trans. Power Electron., vol. 25, no. 9, pp. 2386 –2396, Sep. 2010.
[18] M. Baybutt, C. Minnella, A. Ginart, P. Kalgren, and M. Roemer, “Improving digital system diagnostics through prognostic and health
management (phm) technology,” IEEE Trans. Instrum. Meas., vol. 58, no. 2, pp. 255 –262, Feb. 2009.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
18
[19] M. Musallam and C. Johnson, “Real-time compact thermal models for health management of power electronics,” IEEE Trans. Power
Electron., vol. 25, no. 6, pp. 1416 –1425, Jun. 2010.
[20] A. Ginart, P. Kalgren, M. Roemer, D. Brown, and M. Abbas, “Transistor diagnostic strategies and extended operation under one-transistor
trigger suppression in inverter power drives,” IEEE Trans. Power Electron., vol. 25, no. 2, pp. 499 –506, Feb. 2010.
[21] A. Ginart, D. Brown, P. Kalgren, and M. Roemer, “Online ringing characterization as a diagnostic technique for igbts in power drives,”
IEEE Trans. Instrum. Meas., vol. 58, no. 7, pp. 2290 –2299, Jul. 2009.
[22] D. C. Katsis, “Thermal characterization of die-attach degradation in the power MOSFET,” Ph.D. dissertation, Virginia Politech. Inst.,
Blacksburg, 2003.
[23] E. Miranda and J. Su, “Electron transport through broken down ultra-thin SiO2 layers in MOS devices,” Microelectron. Reliab., vol. 44,
no. 1, pp. 1 – 23, 2004. [Online]. Available: http://www.sciencedirect.com/science/article/pii/S0026271403003871
[24] A. Ginart, I. Barlas, J. Dorrity, P. Kalgren, and M. Roemer, “Self-healing from a phm perspective.”
IEEE, Sep 2006, pp. 697–703.
[Online]. Available: http://ieeexplore.ieee.org/lpdocs/epic03/wrapper.htm?arnumber=4062464
[25] V. T. Somasekhar, S. Srinivas, B. Prakash Reddy, C. Nagarjuna Reddy, and K. Sivakumar, “Pulse width-modulated switching strategy
for the dynamic balancing of zero-sequence current for a dual-inverter fed open-end winding induction motor drive,” IET Electric Power
Applications, vol. 1, no. 4, pp. 591–600, Jul. 2007.
[26] V. T. Somasekhar, S. Srinivas, and K. K. Kumar, “Effect of zero-vector placement in a dual-inverter fed open-end winding induction-motor
drive with a decoupled space-vector PWM strategy,” IEEE Trans. Ind. Electron., vol. 55, no. 6, pp. 2497–2505, Jun. 2008.
[27] I. Takahashi and T. Noguchi, “A new quick-response and high-efficiency control strategy of an induction motor,” IEEE Trans. Ind. Appl.,
vol. IA-22, no. 5, pp. 820–827, Sep. 1986.
[28] A. Shukla, A. Ghosh, and A. Joshi, “Improved Multilevel Hysteresis Current Regulation and Capacitor Voltage Balancing Schemes for
Flying Capacitor Multilevel Inverter,” IEEE Trans. Power Electron., vol. 23, no. 2, pp. 518–529, Mar. 2008.
[29] M. F. Escalante, J. C. Vannier, and A. Arzande, “Flying capacitor multilevel inverters and DTC motor drive applications,” IEEE Trans.
Ind. Electron., vol. 49, no. 4, pp. 809–815, Aug. 2002.
[30] F. Khoucha, S. M. Lagoun, K. Marouani, A. Kheloui, and M. El Hachemi Benbouzid, “Hybrid cascaded h-bridge multilevel-inverter
induction-motor-drive direct torque control for automotive applications,” IEEE Trans. Ind. Electron., vol. 57, no. 3, pp. 892–899, Mar.
2010.
[31] S. Vukosavic and V. Stefanovic, “Srm inverter topologies: a comparative evaluation,” IEEE Trans. Ind. Appl., vol. 27, no. 6, pp. 1034–1047,
Nov./Dec. 1991.
[32] J. Restrepo, J. Aller, V. Guzman, M. Gimenez, and J. Viola, “Switching strategies for dtc on asymmetric converters,” in Power Electronics
and Applications, 2009. EPE ’09. 13th European Conference on, 8-10 2009, pp. 1 –9.
Jose A. Restrepo (S’89-M’92) received the B.Sc. and M.Sc. degrees in electronics engineering from the Universidad
Simón Bolı́var, Caracas, Venezuela, in 1988 and 1991, respectively, and the Ph.D. degree from the University of
Manchester Institute of Science and Technology, Manchester, U.K., in 1996. He is currently a Full Professor of
electronics engineering in the Departamento de Electrónica y Circuitos, Universidad Simón Bolı́var. He has been a
Visiting Professor at the School of Electrical and Computer Engineering, Georgia Institute of Technology, Atlanta, in
2002 and 2010. His research interests include advanced control of electrical machines, artificial intelligence, power
quality and digital signal processors applied to motion control.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
19
Alberto Berzoy received the B.Sc. and M.Sc. degrees in electronics engineering from the Universidad Simón Bolı́var,
Caracas, Venezuela, in 2003 and 2008. He is currently an Assistant Professor of electronics engineering in the
Departamento de Electrónica y Circuitos, Universidad Simón Bolı́var. His research interests include advanced control
of electrical machines, artificial intelligence, power quality, active power filtering, maximum power point tracking,
control of dc-dc converters, and signal processing.
Antonio Ginart (S’89-M’01-SM’07) Received the B.Sc. and M Sc degrees in electrical engineering from Simon Bolivar
University, Caracas, Venezuela in 1986 and 1990, respectively, and the Ph.D. in electrical engineering from the Georgia
Institute of Technology in 2001. He has over 20 years of experience in motors, electronic drives, and industrial controls.
He was an Instructor, Assistant Professor, and later Associate Professor at Simon Bolivar University from 1989 to 2002.
He was a consultant for Aureal Semiconductors, Inc. in power amplification from 1999 to 2000, where he pioneered the
effort to develop Class AD amplifiers. At Impact Technologies, he is responsible of developing intelligent automated
monitoring systems for electrical and electronics equipment for industrial and military applications. His research has
led to over 50 publications.
Jose M. Aller was born in Caracas, Venezuela, in 1958. He received the Eng. degree in electrical engineering from the
Universidad Simón Bolı́var, Caracas, in 1980, the M.S. degree in electrical engineering from the Universidad Central
de Venezuela, Caracas, in 1982, and the Ph.D. degree in industrial engineering from the Universidad Politécnica de
Madrid, Madrid, Spain, in 1993. He has been a Lecturer at the Universidad Simón Bolı́var for 28 years, where he
is currently a Full Time Professor in the Departamento de Conversión y Transporte de Energı́a teaching electrical
machines and power electronics. He was the General Secretary of the Universidad Simón Bolı́var from 2001 to 2005.
He has been visiting professor at the Georgia Institute of Technology, Atlanta, USA, in 2000 and 2007. His research
interests include space-vector applications, electrical machine control, power electronics and monitoring of electrical machines.
Ronald Harley (M’77SM’86F’92) received the M.Sc.Eng. degree (cum laude) in electrical engineering from the
University of Pretoria, Pretoria, South Africa, in 1965, and the Ph.D. degree from London University, London, U.K.,
in 1969. In 1971, he was appointed as the Chair of Electrical Machines and Power Systems at the University of Natal,
Durban, South Africa. He is currently the Duke Power Company Distinguished Professor in the School of Electrical
and Computer Engineering, Georgia Institute of Technology, Atlanta. He is the coauthor of some 410 papers published
in refereed journals and international conference proceedings. He is the holder of three patents. His research interests
include the dynamic behavior and condition monitoring of electric machines, motor drives, power systems, and their
components and controlling them by the use of power electronics and intelligent control algorithms.
Dr. Harley is a Fellow of the Institution of Engineering and Technology, U.K., a Fellow of the Royal Society in South Africa, and a member
of the Academy of Science in South Africa. He received the Cyril Veinott Electromechanical Energy Conversion Award from the IEEE Power
Engineering Society for “outstanding contributions to the field of electromechanical energy conversion” in 2005 and the IEEE Richard H.
Kaufmann Field Award with the citation “for contributions to monitoring, control and optimization of electrical processes including electrical
machines and power networks” in 2009.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.
This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.
20
Thomas G. Habetler (S’85-M’89-SM’92-F’02) received the B.S. and M.S. degrees in electrical engineering from
Marquette University, Milwaukee, WI, in 1981 and 1984, respectively, and the Ph.D. degree from the University of
Wisconsin, Madison, in 1989.
From 1983 to 1985, he was a Project Engineer with the Electro-Motive Division, General Motors. Since 1989, he
has been with the School of Electrical and Computer Engineering, Georgia Institute of Technology, Atlanta, where he
is currently a Professor of electrical engineering. He is a regular consultant to industry in the field of condition-based
diagnostics for electrical systems. His research interests include electric machine protection and condition monitoring,
switching converter technology, and drives. He has published over 150 papers in the field.
Dr. Habetler currently serves as the IEEE Division II Director-Elect, is the Past-President of the IEEE Power Electronics Society, and is the
Past-Chair of the Industrial Power Converter Committee of the IEEE Industry Applications Society. He was the recipient of four conference
prize paper awards from the IEEE Industry Applications Society.
Copyright (c) 2011 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing pubs-permissions@ieee.org.