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A New Method for Start-up of Isolated Boost Converters Using Magnetic- and Winding-
Integration
Lindberg-Poulsen, Kristian; Ouyang, Ziwei; Sen, Gökhan; Andersen, Michael A. E.
Published in:
Annual IEEE Applied Power Electronics Conference and Exposition
Link to article, DOI:
10.1109/APEC.2012.6165841
Publication date:
2012
Link back to DTU Orbit
Citation (APA):
Lindberg-Poulsen, K., Ouyang, Z., Sen, G., & Andersen, M. A. E. (2012). A New Method for Start-up of Isolated
Boost Converters Using Magnetic- and Winding-Integration. In Annual IEEE Applied Power Electronics
Conference and Exposition: Proceedings (pp. 340-345). IEEE. https://doi.org/10.1109/APEC.2012.6165841
A New Method for Start-up of Isolated Boost
Converters Using Magnetic- and
Winding-Integration
Kristian Lindberg-Poulsen
Technical University of
Denmark, Dept. of
Electrical Engineering
k.lindberg.p@gmail.com
Ziwei Ouyang
Technical University of
Denmark, Dept. of
Electrical Engineering
zo@elektro.dtu.dk
Gökhan Sen
Technical University of
Denmark, Dept. of
Electrical Engineering
gs@elektro.dtu.dk
Michael A.E. Andersen
Technical University of
Denmark, Dept. of
Electrical Engineering
ma@elektro.dtu.dk
Abstract—A new solution to the start-up and low output
voltage operation of isolated boost family converters is presented.
By the use of integrated magnetics and winding integration, the
transformer secondary winding is re-used during start-up as a
flyback winding coupled to the boost inductor. The traditional
added flyback winding coupled to the boost inductor is thus
eliminated from the circuit, bringing substantial cost savings,
increased efficiency and simplified design. Each subinterval of
the converter operation is described through electrical and
magnetic circuit diagrams, and the concept is extended to other
isolated boost family topologies. The principle of operation is
demonstrated with a 800W isolated boost prototype, and a 1600W
primary parallel series secondary isolated boost converter. Effi-
ciency measurements of both prototypes are presented, including
measurements during both start-up and normal boost operation.
I. INTRODUCTION
Isolated boost converters have been shown to be the most
efficient topology for high power, low input voltage, high out-
put voltage applications [1]–[5]. Suitable applications include
distributed generation systems, backup systems, fuel cell con-
verters, electric vehicle applications and avionic applications.
However, a disadvantage of the topology is that the boost
characteristic sets a lower limit for the output voltage, which
introduces in-rush current during start-up from zero output
voltage, as well as during fault situations such as output short
circuit.
Fig.1 shows a common solution to the start-up problem [6],
[7]. A flyback winding is arranged on the boost inductor, such
Fig. 1. Isolated full-bridge boost converter with traditional start-up method
using added external flyback winding and diode.
that the converter may operate as a flyback converter during
start-up with the associated buck-boost voltage modulation
factor allowing control of the output voltage all the way down
to zero volts. During normal boost operation, the flyback wind-
ing is completely inactive, but occupies part of the winding
window of the boost inductor, leaving less copper area for the
boost inductor, which in turn causes a drop in efficiency. Being
a power transferring magnetic component, the flyback winding
is a relatively expensive circuit element and complicates the
assembly of the boost inductor. Fig.2 and Fig.3 respectively
show the simplified and complete circuit diagrams of the new
circuit topology, where the magnetic integration of the boost
inductor with the transformer allows the boost inductor to
couple to the secondary winding during start-up, such that
the secondary winding acts as flyback winding.
II. PRINCIPLE OF OPERATION
Figures 4 to 7 explain the principle of operation as
applied to the basic full-bridge isolated boost topology. Each
Fig. 2. Electrical circuit diagram of new start-up method.
Fig. 3. Electrical and magnetic circuit diagram of new start-up method using
EI core with center leg air gap.
978-1-4577-1216-6/12/$26.00 ©2012 IEEE 340
subinterval of operation is shown with an electric circuit
diagram followed by the corresponding and a magnetic core
diagram. Relevant voltage polarities are marked with +/-
signs, and current directions are marked with arrows while
inactive elements are dimmed. The electric circuit diagrams
correspond to their respective core diagrams by the ports
marked a, b, c, d. The core diagrams include the flux rate,
dφ
dt
= φ
0
, induced by the inductor, as well as the flux rate
induced by the transformer. The circuit operation can be
understood by analysing voltage polarities and flux rates
together with the right hand rule, while taking into account
the switch states, current directions and diode bias.
Boost mode operation (D
sw
≥ 0.5)
In the familiar isolated boost topology, the full-bridge
switch duty cycle must be above 0.5, such that there is always
a current path available for the boost inductor. This is defined
as the normal boost mode operation. The gate signal is
phase-shifted 180
◦
between each diagonal high-side/low-side
switch pair. Since D
sw
≥ 0.5, the phase shifted gate signals
overlap such that either all switches are on, or two diagonal
switches are on.
When all switches are on, V
in
is applied to the boost inductor,
and the input current is increasing. Known as the charging
or boosting subinterval, the circuit operation during this is
shown in fig. 4. As seen in fig. 4(b), the flux rate induced by
the inductor is uncoupled from the transformer windings due
to the fact that the voltage drops induced on the transformer
windings on each side leg are of opposite polarity.
The second sub-interval, referred to as the boost mode
discharge subinterval, is shown in fig. 5. When two diagonal
switches are on and the other two are off, the inductor
current passes through the primary winding, allowing the
corresponding diagonal diode pair of the output rectifier to
become forward biased. The output voltage is reflected to the
primary winding, such that a negative voltage drop is applied
over the inductor, decreasing the current. The corresponding
core diagram in fig. 5(b) shows that the inductor flux rate
is decreasing, and is still uncoupled from the transformer.
Additionally, it is noted that the flyback diode D
f
is reverse
biased.
Start-up mode operation (D
sw
< 0.5)
When the switch duty cycle is reduced below 0.5, the
diagonal switch pairs are no longer overlapping in on-state,
meaning that either two diagonal switches are on or all
switches are off. When two switches are on, the circuit
operation is identical to the boost mode discharge subinterval,
except that the inductor current is now decreasing due to the
fact that the primary voltage is lower than the input voltage,
causing a positive voltage drop over L. Because of this,
the subinterval is referred to as the start-up mode charging
subinterval, and the corresponding circuit and core diagrams
are seen in fig. 6.
(a) Boost mode, charging subinterval, circuit diagram.
(b) Boost mode, charging subinterval, core diagram.
Fig. 4. Boost mode, charging subinterval. All switches are on, inductor
current is increasing. Core diagram shows that inductor flux rate is decoupled
from transformer.
(a) Boost mode, discharging subinterval, circuit diagram.
(b) Boost mode, discharging subinterval, core diagram.
Fig. 5. Boost mode, discharge subinterval. Two diagonal switches are on,
inductor current is decreased by transferring energy through the transformer.
It is noted that D
f
is reverse biased..
When all switches are off, there needs to be an alternative
current path such that the inductor flux remains continuous
in order to avoid MOSFET avalanche mode clamping of the
stored energy. This current path is provided by the secondary
transformer winding, which thereby acts as a flyback winding
during the flyback discharge subinterval seen in fig. 7. During
switch turn off, the inductor current is decreasing, creating a
341
(a) Startup mode, charging subinterval, circuit diagram.
(b) Startup mode, charging subinterval, core diagram.
Fig. 6. Start-up mode, charging subinterval. Two diagonal switches are on,
charging the inductor while also transferring energy through the transformer.
reverse in the flux rate, as seen by comparing φ
0
L
on fig. 6(b)
and fig. 7(b). Following the right hand rule, it is evident that
this induces respective voltage drops over each half of the
secondary winding, with polarities such as to allow D
f
to
become forward biased. The two halves of the secondary
winding will then be working in parallel to discharge the
energy stored in the air gap, with the full output voltage
applied to each winding half.
It should be noted that it would also be possible to run the
converter in "pure flyback mode", wherein either all switches
are on or all switches are off. However, this means that
zero energy is transferred to the output during the boosting
subinterval, leading to higher current ripple at same output
power. The "hybrid" startup mode shown in figs. 6 and 7 has
been described in-depth as applied to a conventional external
flyback winding [7], and the specific timing analysis of the
previous work can be directly applied to the new integrated
magnetic approach.
III. EXTENSION TO OTHER ISOLATED BOOST FAMILY
TOPOLOGIES
The concept can readily be applied to numerous isolated
boost derived topologies, such as flyback- current-fed push-
pull [8], dual inductor [9], and parallel primary isolated boost
[3]. It can also be applied to various rectification circuits,
including voltage doubler and center tap rectifier. Figure 8
shows the principle applied to an isolated boost converter with
a center tap rectifier. In this case, the flyback diode D
f
is no
longer required, and the start-up functionality is gained "or
free" using only the specified integrated magnetic structure,
which may be beneficial in itself by reducing magnetic com-
(a) Startup mode, discharging subinterval, circuit diagram.
(b) Startup mode, discharging subinterval, core diagram.
Fig. 7. Start-up mode, flyback discharge subinterval. All switches are turned
off. The drop in inductor current causes a reverse in the associated flux rate,
which couples to the secondary transformer windings. From the polarity of
the induced voltages, it is evident that this allows D
f
to be forward biased,
such that the energy stored in the air gap can be discharged to the converter
output.
Fig. 8. New start-up method applied to isolated boost with center tap
rectification circuit, where a flyback diode is not required.
ponent count and increasing efficiency [10]. Figure 9 shows
the principle applied to the parallel primary topology, which
has been shown to be an efficient way of scaling isolated boost
converter design for higher power [3], [4]. Additionally, the
shown implementation features a shared center leg for flux
cancellation, resulting in increased efficiency [11]. It is noted
that only a single flyback diode is required, regardless of the
degree of paralleling.
Additional applications include dual inductor isolated boost
and Weinberg/push-pull isolated boost topologies.
IV. EXPERIMENTAL VERIFICATION
An 800W isolated boost prototype as well as a 1600W
parallel primary isolated boost prototype have been built
in order to verify the start-up functionality, as well as to
demonstrate the possibility of achieving high efficiency and
high power density by application of the integration method.
Both converters are hard-switched, and rely on extensive
interleaving to achieve a low transformer leakage inductance
342
Fig. 9. New start-up method applied to parallel primary isolated boost
topology.
of 91nH and 124nH respectively. The converters are designed
for fuel-cell application, where a fuel cell output voltage of
25-35V is expected.
Figure 11 shows current measurements during start-up mode
of the isolated boost prototype. The converter is running at
V
in
= 25V , D
sw
= 40%, V
out
= 60V and 100Ω load. C3
(blue) shows the AC component of the input current, C4(green)
shows the AC component of the current through a high side
output rectifier D
1
, while C1(yellow) and C2(red) show the
two gate signals.
When the gate signal C1(yellow) is high, D
1
and D
3
are
forward biased, conducting a constant current through the
secondary winding to the load. During this period, the boost
inductor current C3(blue) is rising. After C1(yellow) goes to
zero, all MOSFETs are turned off, and the boost inductor
current quickly drops. At this moment, current continues
flowing through D
1
, safely discharging the energy stored in
the boost inductor to the output. In fig. 11, C4(green) shows
the current of the flyback diode D
f
, clearly demonstrating the
commutation of the boost inductor current.
Fig. 10. Currrent waveforms of isolated boost prototype during startup mode.
C3 (blue) shows the AC component of the input current, C4(green) shows the
current through a high side output rectifier D
1
, while C1(yellow) and C2(red)
show the two gate signals.
Fig. 11. Currrent waveforms of isolated boost prototype during startup mode.
C3 (blue) shows the AC component of the input current, C4(green) shows the
current through the flyback diode D
f
, while C1(yellow) and C2(red) show
the two gate signals.
A. Efficiency Measurements
In order to measure the efficiency during start-up of the par-
allel primary prototype, the duty-cycle was gradually increased
from zero to 74%. A fixed load of 100Ω is used across the
range, corresponding to the rated 1600W at V
out
= 400V .
Figure 12 shows the resulting efficiency measurements, as
well as the measured output voltage at each duty cycle. The
local maximum of efficiency around D
sw
= 20% is caused
by discontinuous conduction mode. Close to D
sw
= 50%, the
efficiency rises dramatically, as an increasingly greater propor-
tion of the total power is transferred through the transformer
rather than the flyback operation. The converter is capable of
operating continuously in start-up mode without overheating,
and the voltage transitions smoothly across D
sw
= 50%.
The efficiency during normal operation was measured for
both the 800W isolated boost prototype and the 1600W
parallel primary. Figure 13 shows the isolated boost prototype
measurements, while fig. 12 shows the results for the parallel
primary prototype. Both prototypes have a peak efficiency
Fig. 12. Efficiency measurements of parallel primary prototype, showing
efficiency and output voltage as a function of duty cycle, with a fixed load
of 100Ω.
343