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Texas Instruments 1Q 2013Issue Analog Applications Journal Application notes
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High-Performance Analog Products
Analog Applications
Journal
First Quarter, 2013
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Contents
Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4
Data Converters
Grounding in mixed-signal systems demystified, Part 1. . . . . . . . . . . . . . . . . . . . . . . . . . . 5
Maintaining system performance with ADCs and DACs in a noisy digital environment is dependent upon
using good circuit-design techniques like proper signal routing, decoupling, and grounding. This article
is the first of a two-part series that looks closely at the grounding techniques used in mixed-signal
systems. Part 1 explains typical terminologies and ground planes and introduces partitioning methods.
Add a digitally controlled PGA with noise filter to an ADC. . . . . . . . . . . . . . . . . . . . . . . . 9
Some application designs with a highly dynamic signal range can be enhanced by adding an external
programmable gain amplifier (PGA) in front of the ADC. This article describes how to implement an
economical PGA by using a single resettable integrator that allows gain to be digitally controlled and
calibrated. In addition, the PGA provides low-pass filtering and the zero-level voltage reference can be
externally controlled.
Power Management
Design of a 60-A interleaved active-clamp forward converter. . . . . . . . . . . . . . . . . . . . 13
Forward converters are typically avoided for supply outputs greater than 30 A or about 250 W. This
article presents an interleaved forward-converter design capable of up to 60 A or about 500 W. This
design retains the benefits of synchronous rectification while avoiding the problems of load sharing
parallel supplies. A complete schematic, bill or materials, and test results are available.
Power MOSFET failures in mobile PMUs: Causes and design precautions. . . . . . . . . 17
Automotive and mobile systems can expose power MOSFETs to harsh operating environments and
intense transients that result in component electrical overstress (EOS). This article addresses special
design considerations and failure analysis of high-frequency switchers and regulators employing
external feedback components for mobile applications. The discussion includes common device-failure
mechanisms, general precautions, PCB layout techniques, and component-selection tips.
35-V, single-channel gate drivers for IGBT and MOSFET
renewable-energy applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22
Renewable-energy applications present designers with challenges that include high voltages and power
levels, extensive safety and reliability requirements, and overall complexity to interconnect multiple
systems. This article describes how a single-channel gate driver contributes to meeting these challenges.
How to pick a linear regulator for noise-sensitive applications. . . . . . . . . . . . . . . . . . . 25
A low-noise power solution is essential to preserving signal accuracy and integrity. This article addresses
criteria and parameters to consider in designing a low-noise power solution, including important
specifications for picking a linear regulator.
Index of Articles. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
TI Worldwide Technical Support . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34
To view past issues of the
Analog Applications Journal, visit the Web site:
www.ti.com/aaj
Subscribe to the AAJ:
www.ti.com/subscribe-aaj
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Introduction
The Analog Applications Journal is a digest of technical analog articles
published quarterly by Texas Instruments. Written with design engineers,
engineering managers, system designers and technicians in mind, these “howto” articles offer a basic understanding of how TI analog products can be used
to solve various design issues and requirements. Readers will find tutorial
information as well as practical engineering designs and detailed mathematical
solutions as they apply to the following product categories:
• Data Converters
• Power Management
• Interface (Data Transmission)
• Amplifiers: Audio
• Amplifiers: Op Amps
• Low-Power RF
• General Interest
Analog Applications Journal articles include many helpful hints and rules of
thumb to guide readers who are new to engineering, or engineers who are just
new to analog, as well as the advanced analog engineer. Where applicable,
readers will also find software routines and program structures.
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Grounding in mixed-signal systems
demystified, Part 1
By Sanjay Pithadia, Analog Applications Engineer,
and Shridhar More, Senior Analog Applications Engineer
Introduction
Figure 1. AGND and DGND pins in a data converter
Every signal-processing system
requires mixed-signal devices,
Digital Data
such as analog-to-digital converters
(ADCs) and/or digital-to-analog con­
verters (DACs). The need for proc­
V = Li (di/dt)
essing analog signals with a wide
Ri
Li
Ri
Li
Digital
dynamic range imposes the requireDigital
DGND
Supply
Section
ment to use high-performance ADCs
and DACs. Maintaining performance
L i – Internal inductance
in a noisy digital environment is
Cstray
Cstray
Cstray – Stray capacitance
dependent upon using good circuitR i – Internal resistance
design techniques like proper signal
Ri
Li
Ri
Li
Analog
routing, decoupling, and grounding.
Analog
AGND
Supply
Section
Undoubtedly, grounding is one of
the most discussed subjects in system
Inside Data Converter
design. Though the basic concepts are
relatively simple, the implementation
Analog Input Signal
is difficult. For linear systems, the
ground is the reference against which
the signal is based; and, unfortunately,
it also becomes the return path for the power-supply
Interpretation of AGND and DGND pins
­current in unipolar supply systems. An improper applicain mixed-signal devices
tion of grounding strategies can degrade the performance
Digital- and analog-design engineers tend to view mixedin high-accuracy linear systems. There is no “cookbook”
signal devices from different perspectives, but every engithat guarantees good results, but there are a few things
neer who uses a mixed-signal device is aware of analog
that, if not done properly, can cause issues.
ground (AGND) and digital ground (DGND). Many are
This article is the first of a two-part series that looks
confused about how to deal with these grounds; and, yes,
closely at the grounding techniques used in mixed-signal
much of the confusion comes from how the ADC ground
systems. Part 1 explains typical terminologies and ground
pins are labeled. Note that the pin names, AGND and
planes and introduces partitioning methods. Part 2 explores
DGND, refer to what’s going on inside the component and
techniques for splitting the ground planes, including pros
do not necessarily imply what one should do with the
and cons. It also explains grounding in systems with multigrounds externally. Data-converter datasheets usually sugple converters and multiple boards. Part 2 will appear in a
gest tying the analog and digital grounds together at the
future issue of Analog Applications Journal.
device. However, the designer may or may not want the
A term often used in system design is star ground. This
data converter to become the system’s star ground point.
term builds on the theory that all voltages in a circuit are
What should be done?
referred to as a single ground point, or star ground point.
As illustrated in Figure 1, the grounds inside a mixedThe key feature is that all voltages are measured with
signal IC are typically kept separate to avoid coupling digirespect to a particular point in the ground network, not
tal signals into the analog circuits. An IC designer cannot
just to an undefined ground wherever one can clip a probe.
do anything about the internal inductance and resistance
Practically, it is difficult to implement. For example, in a
(negligible compared to the inductance) associated with
star ground system, drawing out all signal paths to miniconnecting the pads on the chip to the package pins. The
mize signal inter­action and the effects of high-impedance
rapidly changing digital currents produce a voltage (di/dt)
signal or ground paths causes implementation problems to
in digital circuits, which inevitably couples into the analog
arise. When power supplies are added to the circuit, either
circuits through the stray capacitance.
they add unwanted ground paths or their supply currents
The IC works well in spite of such coupling. However, in
flowing in the existing ground paths are large enough or
order to prevent further coupling, the AGND and DGND
noisy enough to corrupt the signal transmission.
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pins should be joined together externally to the same lowimpedance ground plane with minimum lead lengths. Any
extra external impedance in the DGND connection can
cause more digital noise and, in turn, can couple more
digital noise into the analog circuit through the stray
capacitance.
plane, the analog input signal is going to have digital noise
summed with it, because it is probably single-ended and
referenced to the analog ground plane. Connecting the
pins to a quiet analog ground plane can inject a small
amount of digital noise into it and degrade the noise margin of the output logic. This is because the output logic is
now referenced to the analog ground plane and all the
other logic is referenced to the digital ground plane.
However, these currents should be quite small and can be
minimized by ensuring that the converter output does not
drive a large fan-out.
It is possible that the devices used in a design have
either low digital currents or high digital currents. The
grounding scheme is different for both cases. Traditionally,
data converters may be thought of as low-current devices
(such as flash ADC). But today’s data converters with onchip analog functions are becoming more and more digitally intensive. Along with the additional digital circuitry
come larger digital currents and noise. For example, a
sigma-delta ADC contains a complex digital filter that adds
considerably to the digital current in the device.
Analog or digital ground plane, or both?
Why is a ground plane needed? If a bus wire is used as a
ground instead of a plane, calculations must be done to
determine the bus wire’s voltage drop because of its imped­
ance at the equivalent frequency of most logic transitions.
This voltage drop creates an error in the final accuracy of
the system. To implement a ground plane, one side of a
double-sided PCB is made of continuous copper and is
used as a ground. The large amount of metal has the lowest possible resistance and lowest possible inductance
because of the large, flattened conductor pattern.
The ground plane acts as a low-impedance return path
for decoupling high-frequency currents caused by fast
digital logic. It also minimizes emissions from electromagnetic interference/radio-frequency interference (EMI/RFI).
Because of the ground plane’s shielding action, the circuit’s
susceptibility to external EMI/RFI is reduced. Ground
planes also permit high-speed digital or analog signals to be
transmitted via transmission-line (microstrip or stripline)
techniques, where controlled impedances are required.
As mentioned earlier, the AGND and DGND pins must
be joined together at the device. If the analog and digital
grounds have to be separated, should both be tied to the
analog ground plane, the digital ground plane, or both?
Remember that a data converter is analog! Thus, the
AGND and DGND pins should be connected to the analog
ground plane. If they are connected to the digital ground
Grounding data converters with
low digital currents
As mentioned, a data converter (or any mixed-signal
device) is analog. In any system, the analog signal plane is
where all the analog circuitry and mixed-signal devices are
placed. Similarly, the digital signal plane has all the digital
data-processing circuits. The analog and digital ground
planes should have the same size and shape as the respective signal planes.
Figure 2 summarizes the approach for grounding a
mixed-signal device with low digital currents. The analog
ground plane is not corrupted because the small digital
Figure 2. Grounding data converters with low internal digital currents
PCB
Low digital current
generates low noise.
Keep the local
decoupling capacitor
(de-cap) loop as
short as possible.
Analog ground plane’s shape and size
same as for analog signal plane
Digital ground plane’s
shape and size same as
for digital signal plane
RC Filter
or Ferrite
Bead
VA
VDig
VA
Analog
Signal
Conditioning
Back-to-back
Schottky diodes
between AGND and
DGND keep maximum
ground-potential
difference at < 0.3 V.
Data
Converter
VD
Local
De-Cap
for
Digital
Digital
Data
Processing
AGND
VA
DGND
Digital
Supply
Analog
Supply
VD
To System Star Ground
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transient currents flow in the small loop between VDig, the
local decoupling capacitor, and DGND (the green line).
Figure 2 also shows a filter between the analog and digital
power supplies. There are two types of ferrite beads: high-Q
resonant beads and low-Q nonresonant beads. Low-Q
beads are commonly used for power-supply filtering in
series with the power connection.
Grounding data converters with
high digital currents
The circuit in Figure 2 depends on the decoupling capacitor between VDig and DGND to keep the digital transient
currents isolated in a small loop. However, if the digital
currents are significant enough and have components at
DC or low frequencies, the decoupling capacitor may have
to be so large that it is impractical. Any digital current
that flows outside the loop between VDig and DGND must
flow through the analog ground plane. This may degrade
performance, especially in high-resolution systems. An
alternative grounding method for a mixed-signal device
with high levels of digital currents is shown in Figure 3.
The AGND pin of the data converter is connected to the
analog ground plane, and the DGND pin is connected to
the digital ground plane. The digital currents are isolated
from the analog ground plane, but the noise between the
two ground planes is applied directly between the
device’s AGND and DGND pins. The analog and digital
circuits must be well isolated. The noise between AGND
and DGND pins must not be large enough to reduce
­internal noise margins or cause corruption of the internal
­analog circuits.
Connecting analog and digital ground planes
Figures 2 and 3 show optional back-to-back Schottky
diodes connecting the analog and digital ground planes.
The Schottky diodes prevent large DC voltages or lowfrequency voltage spikes from developing across the two
planes. These voltages can potentially damage the mixedsignal IC if they exceed 0.3 V, because they appear directly
between the AGND and DGND pins.
As an alternative to the back-to-back Schottky diodes, a
ferrite bead can provide a DC connection between the two
planes but isolate them at frequencies above a few megahertz where the ferrite bead becomes resistive. This protects the IC from DC voltages between AGND and DGND,
but the DC connection provided by the ferrite bead can
introduce unwanted DC ground loops and may not be suitable for high-resolution systems. Whenever AGND and
DGND pins are separated in the special case of ICs with
high digital currents, provisions should be made to connect them together if necessary.
Jumpers and/or strap options allow both methods to
be tried to verify which gives the best overall system
performance.
Isolation or partitioning: Which is important
for ground planes?
A common concern is how to isolate the grounds so that
the analog circuit does not interfere with the digital circuit.
It is a well-known fact that digital circuitry is noisy.
Saturating logic draws large, fast current spikes from its
supply during switching. Conversely, analog circuitry is
quite vulnerable to noise. It is not that the analog circuit
might interfere with the digital logic. Rather, it is possible
that the high-speed digital logic might interfere with the
low-level analog circuits. So the concern should be how to
prevent digital-logic ground currents from contaminating
the low-level analog circuitry on a mixed-signal PCB. The
first thought might be to split the ground planes to isolate
DGND from AGND. Although the split-plane approach can
Figure 3. Grounding data converters with high internal digital currents
PCB
Analog ground plane’s shape
same as for analog signal plane
VA
Digital ground plane’s shape
same as for digital signal plane
VA
Analog
Signal
Conditioning
Back-to-back
Schottky diodes
between AGND and
DGND keep maximum
ground-potential
difference at < 0.3 V.
VD
VD
Digital
Data
Processing
Data
Converter
AGND
VA
DGND
Digital
Supply
Analog
Supply
VD
To System Star Ground
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be made to work, it has many problems--especially in
large, complex systems.
There are two basic principles of electromagnetic compatibility (EMC):
1. Currents should be returned to their sources locally and
as compactly as possible. If not, a loop antenna should
be created.
2. A system should have only one reference plane, as two
references create a dipole antenna.
During EMC tests, most problems are observed when
traces are routed across a slot or a split in a ground or
power plane. Since this routing causes both radiation and
crosstalk issues, it is not recommended.
It is important to understand how and where the ground
currents in a split plane actually flow. Most designers think
only about where the signal current flows and ignore the
path taken by the return current. The high-frequency signals have a characteristic of following the path of least
impedance (inductance). The path’s inductance is determined by the loop area that the path encloses. The larger
the area that the current has to travel to return to the
source, the larger the inductance will be. The smallest
inductance path is directly next to the trace. So, regardless of the plane--power or ground--the return current
flows on the plane adjacent to the trace. The current
spreads out slightly in the plane but otherwise stays under
the trace. The actual distribution is similar to a Gaussian
curve in nature. Figure 4 illustrates that the return-current
flow is directly below the signal trace. This creates the
path of least impedance.
The current-distribution curve for the return path is
defined by
i (A/cm) =
IO
×
πh
1
 D
1+  
 h
2
,
where I O is the total signal current (A), h is the height of
the trace (cm), and D is the distance from the trace (cm).
From this equation it can be concluded that digital ground
currents resist flowing through the analog portion of the
ground plane and so will not corrupt the analog signal.
Figure 4. Distribution of return current
w
εr
t
h
i (A/cm)
D
For reference planes, it is important that the clearance
sections of vias do not interfere with the return current’s
path. In the case of an obstacle, the return current finds a
way around it, as shown in Figure 5. However, this rerouting will most likely cause the current’s electromagnetic
fields to interfere with the fields of other signal traces,
introducing crosstalk. Moreover, this obstacle adversely
affects the impedance of the traces passing over it, leading
to discontinuities and increased EMI.
Part 2 of this two-part article series will discuss the pros
and cons involved in splitting the ground planes and will
also explain grounding in systems with multiple converters
and multiple boards.
References
1. H. W. Ott, “Partitioning and layout of a mixed-signal
PCB,” Printed Circuit Design, pp. 8–11, June 2001.
2. “Analog-to-digital converter grounding practices affect
system performance,” Application Report. Available:
www.ti.com/sbaa052-aaj
Related Web sites
Data Converters:
www.ti.com/dc-aaj
For examples of grounding for precision data converters,
visit: www.ti.com/e2egrounding-aaj
Subscribe to the AAJ:
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Figure 5. Return current with and without slot
Circuit Trace
Load
Ground-Plane
Disruption
Driving
Gate
High-Speed
Return Current
Return Path
Around Obstacle
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Add a digitally controlled PGA with noise
filter to an ADC
By Kai Gossner
Field Application Engineer
Introduction
In some applications, a signal with high dynamic range
needs to be digitized. A common method of digitization is
to add an external programmable gain amplifier (PGA) in
front of the analog-to-digital converter (ADC). Only a few
microcontrollers have internal PGAs. However, nowadays
PGAs are available in a single chip with one or multiple
input channels. Such PGAs add additional costs to the
system and usually consume more power as a fixed-gain
solution.
This article describes how to implement a PGA by using
just a single resettable integrator, with the following
benefits:
• The solution is economical and easy to design.
• Gain can be digitally controlled and calibrated.
Figure 1. Basic block diagram of the PGA
fRES
fSH
RES
VIN
∫
VINT
ADC
Output
• Signal noise is reduced with a low-pass filter, which is
especially useful in noisy microcontroller environments
and for small analog signals. The cutoff frequency automatically adjusts with the chosen sample rate.
• The zero-level voltage reference can be controlled
externally, which makes it handy for single-supply
circuits where the zero level usually is set to VREF /2.
The basic circuit
Figure 1 shows the basic circuit, where an integrator is
added in front of the ADC. The integrator can be reset
with the signal fRES (1 = integrator is reset). The ADC is
controlled with the signal fSH, which connects to the ADC’s
sample-and-hold (SH) unit (1 = sample, 0 = hold). A falling
edge starts the analog-to-digital conversion cycle.
Figure 2 shows a single analog-to-digital (A/D) conversion cycle with the circuit from Figure 1. The cycle is split
into four periods:
1. Integrator reset period: Resets the integrator to “0.”
2. Integration period: The integrator reset signal is
released and the integrator starts to integrate.
3. Sample period: The ADC’s sample-and-hold unit
samples the integrator output, VINT.
4. A/D conversion period: The sample-and-hold unit
holds the voltage, and the ADC starts to convert.
Figure 2. Single A/D cycle with gain = 1
fRES
fSH
VIN
VINT
Undefined (previous value)
VSH
Integrator Integration Sample
Reset Period Period
Period
Start of
A/D Conversion
Period
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Figure 3. Single A/D cycle with PGA gain = 2
fRES
fSH
VIN
VINT
Undefined (previous value)
VSH
Integrator
Reset Period
Integration
Period
The duration of the integration period defines the PGA’s
gain, as the voltage on its input influences the slope linearly: A doubling of the integration time doubles the gain.
Figure 3 demonstrates this influence. The integration
period is doubled and the voltages VSH are increased by a
factor of two.
A nice benefit from this integration scheme is that the
input signal is averaged during the integration period,
which reduces out-of-band noise from the input signal,
VIN. The filter’s impulse response is of finite duration and
is comparable to the behavior of a digital FIR filter rather
than to that of a standard low-pass filter.
Start of
A/D Conversion
Period
Sample
Period
Practical configuration of a PGA
An inverting amplifier can be built with a single operational amplifier (Figure 4). The integrator can be reset by
short-circuiting the capacitor, C, with the switch element, S.
The components R and C influence the integrator’s gain.
The signal VCOM defines the integrator’s zero-level voltage and can be set, for example, to VREF /2, where VREF is
the ADC’s reference voltage. The integrator is reset to this
voltage when the capacitor is discharged. Usually a VCOM
signal is present in the system anyway. Often it is used as
a virtual ground or bias voltage for single-supply analog
signal chains.
Figure 4. Practical configuration of the PGA
MCU/ADC
fRES
S
C
f SH
R
VIN
–
VINT
+
+
VSH
–
A/D
Op Amp
VCOM
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Figure 5. SPICE-simulation results of circuit in Figure 4
1.90
ADC Sample Points
ADC Output
(Interpolated)
1.80
Output (V)
VINT
1.70
VIN
1.60
0
250
500
750
1000
Time (µs)
Figure 5 shows SPICE-simulation results of the circuit
in Figure 4. The blue dots mark the sample moments of
the ADC. As shown, the signal VIN is amplified by a factor
of about –8. The red signal is inverted to the green due to
the integrator’s inverting behavior.
How it works
The sample rate, the maximum desired gain, and the A/D
conversion time influence the selection of the integration
constant defined by the components R and C. As shown in
Figures 2 and 3, the integrator needs enough time to reach
the gain, G, within the duration of the integration period, t.
The dependency of G and t can be calculated as
G=
−t
.
R×C
Calibration
Tolerances of R and C lead to modification of the gain
­factor. The capacitor should have a very small piezo effect
to get a very linear integration. Capacitors can have an
especially large tolerance--for example, 20%. This is just
the initial tolerance, which can be calibrated once.
Tolerances due to aging effects are very small (less than
1% per year).
The gain and offset can be calibrated in the same way as
with a standard ADC by applying known voltages to the
input and calculating correction values for offset and gain
based on expected and actual values. The calibration can
be done for each gain factor used in the application.
Circuit variations
The close time (integrator reset period) of the switch
(S) depends on the impedance of the switch and the value
of the capacitor (C).
Using the PGA as a low-pass filter only (gain = 1)
In case input-signal amplification is not wanted, it is possible to use the PGA circuit only as a noise filter. The integrator constant can be set to a value that leads to a fixed
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gain of 1. In this case the integration phase can start
immediately after the sample, and the hold stage can be
set to hold mode (Figure 6).
Non-inverting integration
The circuit in Figure 4 uses an inverting integrator. When
this inversion is not acceptable, it is possible to use a noninverting integrator by adding a single-supply inverting
buffer in front of the integrator.
Conclusion
This article has presented a cost-effective and simple way
to implement PGA functionality in cost- and power-driven
applications. Its filtering properties also reduce costs by
eliminating the need for an external filter, which is often
present in front of ADCs. Nevertheless, this method cannot replace a PGA in all cases; for example, high sample
rates or very large gain variations make such a solution
difficult to realize.
Related Web sites
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Figure 6. PGA circuit used as a filter only (gain = 1)
Start of A/D Conversion
Period
fRES
fSH
VIN
VINT
VSH
Integration Period
Sample
Integrator
Period Reset Period
Sample
Integrator
Period Reset Period
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Design of a 60-A interleaved active-clamp
forward converter
By Brian King
Applications Engineer, Member Group Technical Staff
Introduction
In 48-V-input telecommunications systems, power supplies
with a capacity of 100 to 250 W are sufficient to cover many
applications. Forward converters are a good choice for
these applications. At lower output voltages, synchronous
rectification in the secondary circuitry improves efficiency
and simplifies system thermal design. Active-clamp forward
converters work well in these applications because of the
ease of implementing synchronous rectification.
In most cases, the output currents of forward converters
are commonly limited to around 30 A. Beyond this current,
the inductor design and conduction losses in the secondary circuitry become difficult to manage. From a power
standpoint, the primary circuitry (number of parallel
FETs) becomes a limiting factor for power ratings above
250 W. In systems with higher power, it is necessary to
move to a different topology like the full bridge, or operate
two or more forward converters in parallel to increase the
output power.
Load-share ICs work great for paralleling supplies that
use diodes to rectify their outputs. Diode-rectified supplies
allow current to be sourced only from the power supply.
Power supplies with synchronous rectifiers, however, can
both source and sink power, which can wreak havoc with
some load-share controllers. This is particularly true at
start-up, where the feedback loop is overridden by the
primary controller’s slow-start circuit, and the two paralleled supplies could attempt to regulate the output to
different voltage levels. These issues can be circumvented
by interleaving two separate power stages. This article
pre­sents the design of a 5-V, 300-W interleaved isolated
supply that is powered from a standard 36- to 72-V telecom input.
Designing the interleaved power stage
In this design example, splitting the power into two interleaved power stages reduces the current in the secondary
of each phase to 30 A. This is much more manageable than
the 60 A that would be required in a single-phase supply.
Both phases actually need to be designed to carry a little
more than 30 A to account for phase imbalances. Design­
ing the power stage begins by selecting the turns ratio and
inductance for the power transformers. A feature of the
active-clamp forward converter is its ability to run at duty
cycles of over 50%. It is best to design for a maximum duty
cycle of no greater than 75% so that the transformer’s
reset voltage does not become excessive. In this example,
a turns ratio of 4.5:1 results in a duty cycle of around 63%
at a 36-V input. Switching each phase at 200 kHz provides
a good balance between size and efficiency. Setting the
primary inductance at 100 µH ensures that sufficient magnetizing current is flowing to drive the commutation of the
power MOSFETs during the switching transitions. The primary inductance and switching frequency determine the
value of the resonant capacitor in the clamp. In this case,
a 0.1-µF capacitor sets the resonant frequency at 50 kHz.
The output inductors are determined just as in any
buck-derived topology. An inductance of 2 µH is used,
resulting in 8.5 A of peak-to-peak ripple current in each
phase with a worst-case input of 72 V. Accounting for a
20% phase imbalance, the inductor must be able to carry
at least 41 A of peak current without saturating.
The output capacitors are selected to meet the requirements for output ripple voltage and for voltage excursions
due to load transients. Interleaving the power stages results
in some cancellation of the ripple current seen by the output capacitors. The amount of ripple-current can­cel­lation
is dependent on the duty cycle and the phase angle
between the two phases. Total cancellation occurs with a
50% duty cycle only when the two phases are synchronized 180° out of phase. This reduction in ripple current
reduces the number of capacitors required based on the
ripple-voltage requirements and the RMS current ratings of
the capacitors. For this design, four 180-µF polymer capac­
itors rated for 4-A RMS each are sufficient to keep the
peak-to-peak ripple voltage below 50 mV. More capacitance
can be added to support large load transients if necessary.
Selecting the primary MOSFETs is also straightforward.
The peak drain voltage is the sum of the input voltage and
the resonant transformer’s reset voltage. The RMS primary
current comprises the reflected load current and the trans­
former magnetizing current. It is important to select a
minimal number of cost-effective transistors and to keep
the power loss in each transistor manageable. For this
design, each phase uses two 150-V, 50-mW MOSFETs in
parallel, with a worst-case loss per FET of approximately
700 mW.
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Figure 1 shows how self-driven synchronous rectifiers
are implemented in each phase of the active-clamp forward converter. One set of synchronous rectifiers (Q4, Q5,
and Q6) sees the input voltage reflected through the
transformer, while the other set (Q1, Q2, and Q3) sees the
transformer’s reset voltage reflected to the secondary side.
With the selected turns ratio, MOSFETs rated at 30 V are
sufficient for this design. Most of the power loss in these
components is due to conduction loss. Paralleling multiple
7-mW MOSFETs for each phase results in a worst-case loss
per FET of around 800 mW. This ensures that the junction
temperatures are reasonable, even with a 20% phase
imbalance. The gate-drive components Q12, Q13, Q15, and
Q16 serve two functions. First, they protect the MOSFET
gates from voltage spikes on the switching waveforms.
Second, they provide a buffer so that the transformer’s
secondary windings are not directly connected to a large
amount of gate capacitance. This is important to ensure
that the power MOSFETs commutate quickly during the
switching transitions.
Figure 1. Gate-drive conditioning circuitry for a self-drive synchronous rectifier
+
+
Ω
Ω
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Figure 2. Interleaved controllers sharing feedback network and soft-start circuit
Master Controller
Ω
Slave Controller
Ω
Figure 3. Variation in offsets can lead to phasecurrent imbalance
60
55
50
48-V Input, Worst-Case Imbalance
45
40
Current (A)
Figure 2 shows how two controllers can be
connected in parallel so that they share a
common feedback signal and soft-start circuit.
With peak-current-mode control, each power
stage behaves as a current source that is controlled by the voltage at the feedback pin. A
single error amplifier regulates the output volt­
age by simultaneously controlling the feedback
pins of the two controllers. Current imbalance
between the two phases is mostly determined
by variations of the offsets inside the controllers and by the tolerances of the current sense
and slope compensation. Figure 3 plots the
current in each phase versus the feedback
voltage for a total tolerance resulting in the
maximum error between phases. This is not of
much concern at high load levels, as one stage
will just carry a heavier burden. At light loads,
however, the error can allow one phase to sink
current, forcing the other phase to source
extra current. This leads to increased losses at
Worst-Case
Maximum Per Phase
35
Total Output Current
30
25
20
15
Worst-Case
Minimum Per Phase
10
5
0
–5
1.65
1.72
1.79
1.86
1.93
2
2.07
Feedback Voltage (V)
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light loads. The phase imbalance must also be considered
when the current limit is programmed.
Synchronization is implemented by designating one controller as the master and the other as the slave. The clock
frequency of the slave controller is set 10% slower than
that of the master clock to ensure synchronization. The
gate-drive signal of the master is used as the clock for the
slave. Some conditioning components are needed to shape
the magnitude and duration of the synchronization pulse.
For proper start-up, timing is critical. Start-up must be
completed before the VDD voltage on either chip falls
below the UVLO OFF level, or neither controller will be
able to start. Tying the two soft-start pins together ensures
that both converters initiate the start-up sequence at the
same time. In case of a fault, this also allows both controllers to be disabled by discharging the soft-start capacitance.
The efficiency of this power supply is shown in Figure 4.
With a nominal 48-V input and a load current of 60 A, the
supply’s efficiency is over 92%. The converter’s ability to
convert to an isolated and regulated 5-V output with no
intermediate bus and minimal power loss simplifies the
system design and reduces the power demand on the
upstream AC/DC rectifier.
Conclusion
In summary, interleaving active-clamp forward power
stages can result in a cost-effective and efficient design.
The design must account for current imbalances between
the phases and ensure proper synchronization and startup. If properly designed, interleaving extends the practical
power range of the active-clamp forward converter to
around 500 W and easily supports load currents of up
to 60 A.
Please visit www.ti.com/tool/PMP2214 for more information on this design, including the complete schematic, bill
of materials, and test results.
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Figure 4. Synchronous rectification enables very
high efficiency
95
94
Efficiency (%)
93
92
91
90
Input Voltage, VIN
–36 V
–48 V
–72 V
89
88
0
6
12
18
24
30
36
42
48
54
60
Load Current (A)
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Power MOSFET failures in mobile PMUs:
Causes and design precautions
By Kern Wong
Principal Applications Engineer
Introduction
It was suspected that a parasitic NPN (formed by n+ (S),
p– (well), and n+ (D) as shown in Figure 1) may have
turned on hard when the p– (well) base biased up the
emitter from n+ (S), a classic EOS scenario in power
devices. Figure 2 shows an equivalent-circuit model of a
MOSFET device with parasitic components.
Power MOSFETs in automotive systems and in mobile
devices being charged or operated in automobiles may be
subjected to harsh operating environments and intense
transients from power equipment and transmitters.
Moreover, caustic contaminants in the atmosphere and on
exposed conductive surfaces of circuit boards
can induce low-impedance paths. Over time,
Figure 1. Cross-section of a typical MOSFET
these low-impedance paths and transient
structure and relevant parasitic elements
events like overloading, electromagnetic coupling, and inductively induced spikes from
B_sub
D
G
S
the operating environment can cause destrucRG
tive electrical overstress (EOS) conditions.
Poly
Such conditions may cause a large current to
n+
n+
flow across a MOSFET power switch in a very
CDB
CGDCGS
e
short time.
Parasitic b
This article addresses special design considRDS(on)
NPN
c Rb
erations and failure analysis of high-frequency
p– (Well)
switchers and regulators employing external
feedback components for mobile and automotive applications. The goal is to help familiarn– (Epitaxial)
ize designers with various mechanisms and
circumstances that may lead to destruction of
on-chip power switches. Techniques for avertn+ (Substrate)
ing and eliminating the effects of EOS conditions are discussed to help improve end-user
products and PCB designs. This article also
presents tips for conducting lab tests and suggests good engineering practices to obviate
Figure 2. Model of a typical MOSFET with
potential problems from occurring in highassociated parasitic elements
density/ultracompact mobile designs.1, 2
Case studies
B
Diode
Drain
In 2011, a designer reported a shorted NMOS
switch in the step-down DC/DC converter of
the Texas Instruments (TI) LM26484 PMU
during in-house testing. This regulator was
designed into a new instrumentation panel.
The banks of LEDs powered by a buck converter were operating in light-load conditions.
TI asked the designers to monitor the voltage
at the supply pins around the clock for transients above 6 V. They confirmed that transient spikes were peaking at over 8 V for
­hundreds of nanoseconds, which occurred
­frequently. The device’s absolute maximum
limit on the supply pin is VIN = 6 V!
CGD
R DS(on)
CDB
Gate R G
Diode
Parasitic
NPN
Rb
CGS
Source
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Examining the PCB layout (Figure 3) revealed that the
Figure 3. PCB with two LM26484s provides four
top traces of the power pins had a single via tapped into
buck converters and two LDOs
the power plane, and their longer tracks made the bypass
capacitors ineffective. To prevent this situation from arising again, TI has suggested improved design guidelines.
For example, adequately large bulk capacitors need to be
added between the VIN and ground planes. Also, local
bypassing needs to be augmented with additional capacitors covering broader frequency bands. These precautions,
shown implemented in Figure 4, will keep large transients
from stressing the PMU’s integrated circuit.
A more involved solution for eliminating EOS is to place
the bypass capacitors closer to the power and ground
pins, as shown in Figure 5. Note that the power-ground
tracks have been widened and include liberal use of larger
vias. This recommendation became a viable solution for
the customer.
In 2012, another customer reported experiencing some
failures with another PMU of the same family that had
dual buck converters and dual LDOs. The
buck-converter switches either shorted out or
Figure 4. Example of a more robust line filter
opened soon after the system left the factory.
and bypass
This PMU was powered from a stepped-down
supply in an automotive application. With
many infotainment and safety systems
7
1
becoming standard equipment in cars starting
in 2014, the PMU production rate is projected
8 VIN2
VIN1 24
20-µF,
to increase by approximately tenfold, creating
47-µF to
PMU with Dual
1-nF, and
Input CLC Pi
a concern for all parties involved. Although no
220-µF
Buck Converters
Line
100-pF
Voltage
Bulk
and LDO
anomalies have been discovered in the cusBypass
Filter
Capacitors
tomer’s rigorous testing for device- and board11
21
Capacitors
PGND2 PGND1
level stress, some infrequent failures have
12
20
occurred. In general, there are many known
mechanisms and opportunities involved in
vehicular applications that potentially could
induce abnormal input-voltage transients,
leading to device damage.
Figure 5. Improved layout with bypass
capacitors closer to power and ground pins
100 pF
220 µF
7
1
8 VIN2
VIN1 24
LM26484
PMU
11
12
PGND2
PGND1
21
100 pF
220 µF
100 pF
10 nF
20 µF
Many EOS conditions on PMUs arise from
inadequate design considerations or overlooking subtle parasitics in some systems. This is
especially true in industrial/automotive applications, wherein unusual ambient conditions
or differences in the electromechanical layout
can manifest reliability issues. EOS can also
be related to the manufacturing process, testing, and component aging.
The following discussion presents some of
the most common EOS culprits. Appropriate
design tips and suggestions are included to
help designers eliminate EOS problems. A
typical means of identifying failure mechanisms is welldocumented. It is strongly suggested that readers seeking
more information also study the physics of failure via
­failure-mode mechanisms and effects analysis (FMMEA).
100 pF
10 nF
20 µF
Common causes of EOS
20
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EOS caused by battery and wiring in automotive
applications
Whenever a vehicle’s 12-V battery voltage falls too low, as
in cold weather and cranking operations, the onboard
PMU’s control, timing, and decision-making circuits may
malfunction before undervoltage lockout (UVLO) comes
to its rescue. As a result, undesirable effects such as
shoot-through and disengaged clamping can stress the
MOS switches and cause permanent damage over time.
High-voltage, fast edge-rate transients are another commonly encountered cause of instantaneous device damage.
Another example, load dumping, is when the 12-V battery
is momentarily removed from the alternator connection.
Due to the inductive effect from the long wirings involved,
the loads can experience a sudden increase in potential at
over 100 V, which may last for hundreds of milliseconds
before it decays to normal levels.
High-voltage spikes in fast transients can propagate
from the MOSFET’s drain terminal to the gate via terminal
capacitance. This can rapidly bias up the gate, potentially
leading to runaway conditions. Normally, slightly exceeding the recommended maximum operating supply voltage
might not be a destructive event. However, when the
suppply voltage exceeds the maximum level and sustains
sufficient energy, it can cause the device to short-circuit in
a few nanoseconds or lead to an avalanche breakdown.
Moreover, loose or poorly secured battery-cable connections can manifest similar high-voltage transients if subjected to strong and abrupt mechanical vibrations.
Inadequate or poor power-supply bypassing
Inadequate supply bypassing can cause abnormal operation that may lead to shoot-through stress from timing
issues. A proper bypass capacitor must have a voltage
­rating that adequately covers peak voltage transients.
Leakage and parasitic inductance from traces are among
the sources that cause the largest, most severe L(di/dt)
overstress pulses created at the pulsing terminal of a
switcher. These high-energy pulses can lead to device
breakdown as previously described. Hence, taking proper
precautions to eliminate unwanted inductive paths is
imperative. For example, bypass capacitors should be
placed as close as possible to the device rail pins. A thick
metal trace should be used as much as is allowable on all
high-transient paths to further cut down parasitic inductance. Finally, transient-suppressing elements or similar
techniques should be used as appropriate to attenuate
potentially destructive high-voltage spikes.
Shorted output from overloading and/or a defective load
capacitor
When a switcher’s output current (IOUT load) exceeds the
rated limit, built-in protection circuits usually prevent any
immediate damage to the device. However, frequent overcurrent events can lead to accumulated EOS conditions,
which over time may cause permanent device damage.
Such damage is associated with the finite delay time, typically in the range of microseconds, required before the
protection circuit kicks into action. Other than true loading
shorts, a defective output capacitor can effect a lowimpedance path that creates a dynamic short-circuit
­current in parallel with the maximum loading—thus producing another continuous EOS condition.
Temporary high-overcurrent operation with synchronous
switches
The MOSFET body diode generally has a long reverse
recovery time compared to that of the MOSFET switch
itself. If the body diode of one MOSFET is still conducting
when the opposing complementary device has switched
on, then a short-circuit condition similar to shoot-through
occurs. This can happen due to timing issues from parasitics or from the circuit or device design (see Figures 1 and
2). Furthermore, internal parasitic inductance and capacitance can store energy that, under certain conditions,
additional current may freewheel through the body diodes
of the FET switches as one turns off and the other turns
on. This is the classic parasitic-capacitance mechanism,
C(dv/dt), with high-speed switching that can lead to continuous high-peak-current transients with no dependence
on load conditions.
This type of EOS increases dramatically when coupled
with power-rail integrity issues as discussed before. The
circumstance can be improved or eliminated with more
accurate design and simulation of the power-train circuitry
and/or by augmenting protective devices, such as a
Schottky diode across the drain and source of the
MOSFET. Using a Schottky diode is a proven technique to
prevent the body diode from being turned on by the freewheeling current. Eliminating excessive undershooting
below ground that could cause noise and turning on parasitic pn junctions also lends another benefit—the Schottky
diodes may moderately increase switcher efficiency.
Device-failure verification and analysis
Failure analysis (FA) utilizes visual inspection, impedance
measurements, X-rays, SAT.Sam, emission hot-spot
OBIRCH analysis, SEM, and SCM tools and techniques,
etc., to identify failure-mode mechanisms and root causes
of device failure. Failure analysis also examines whether
general oversights in a customer’s design or manufacturing
process may be the cause. When the cause is identified, TI
issues relevant advisory and containment actions to internal and external customers to help prevent failure from
reoccurring.
Failure-mode mechanisms
1. Electrostatic-discharge (ESD) destruction or gate surge:
Device-junction or oxide-rupture damage (a short or
leakage) can occur as a result of improper handling during assembly and testing of the device and system.
These mechanisms introduce electrostatic charges onto
the device and/or create external high-voltage surge
events that reach the switch circuit.
For example, an ESD event between a fingertip and
the communication-port connectors of a cell phone or
tablet may cause permanent system damage. As processtechnology nodes continue to shrink, device-level ESD
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protection becomes inadequate on a system level. A
transorb, or a transient-voltage suppressor such as TI’s
TPD1E10B06 protection diode, is a good ­remedy.
Figure 6. High-side pFET shorted to the VIN rails
2. Wear-and-tear mechanisms:
• A die fracture may occur in extreme temperature cycling
• Over time, high-voltage stress may induce dielectric
breakdown that will become a gate-oxide short circuit
• Wire bond and metal routes can open due to EOS from
current overload, etc.
• A voltage transient on the supply lines can cause
­damage to passive and active devices on the die
3. PCB elements and environment:
• A circuit failure may occur due to humidity, presence
of a contaminant, or filaments becoming conductive
• A die fracture may occur due to shock, vibration,
­material fatigue, etc.
• Loss of polymer strength, known as glass transition
­failure, may occur under high-temperature stress
• Bypass and load capacitors may be leaky or shorted
• Inductor windings may short-circuit due to wear and
tear of insulation under high-temperature stress or
mechanical vibration
4. Component aging and inadequacy:
Because aging components may contribute to MOSFET
failures even if they initially meet datasheet specifications,1 manufacturing and product-engineering departments are encouraged to perform testing and burn-in of
parts at ratings slightly above datasheet limits. This
ensures that marginal devices with inherent waferdefect density and random process-related issues are
weeded out. It may be better to lose some yield at production than to be accountable for and spend valuable
resources on field failures later on.
Failure-analysis results
In the 2012 case study mentioned earlier, where the
switch’s drain and source channels were fused together in
an automotive application, the customer could not determine that the PMU IC, the circuit board, or the subsystem
had a reliability problem. Each was rigorously tested and
stressed beyond specification limits, and no failure ever
surfaced. The culprit might have been the layout; the electrical plumbing; the system installation; and/or the operating conditions, such as cold cranking, a weak battery, or
intermittent connection of long/loose power cabling.
Because the customer and its subcontractors were
unable to reproduce the initial failure in their lab, they
needed confirmation and sought assistance from TI.
Examples of in-house failure-analysis results are depicted
in Figures 6 and 7.
Figure 7. Low-side nFET shorted to ground
Failure analysis suggested that the burn marks reflected
in the deprocessed dies were likely the consequence of
EOS conditions. To validate this assumption, it was demonstrated that the failures could be induced in lab setups
for (1) 5-V operation and (2) start-up conditions. By using
a Keithley 2420 3-A source meter—a versatile power supply whose amplitude, frequency, and on/off times can be
programmed—VIN was programmed at 5 V and injected
with a 50-ms pulse that repeated at 100-ms intervals. With
loading at 200 mA and above, the pulse amplitude was
increased at 0.5-V increments at 5-minute intervals until
abnormal current was observed. The part was then
decapped to visually confirm EOS. The results revealed
that when the peak-to-peak pulse voltage reached approximately 7.5 V or more, the switches shorted out. Moreover,
if pulses were to peak further to 9 V, the ESD structure
might also be damaged.
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Reproducing a short circuit from the switches during
start-up was more challenging, however. With a bench
supply cycling the buck converters on and off, VIN issued
relatively slow and smooth start-up transients and settled
in at about 6 ms (Figure 8). Even with the supply set to
slightly over 7 V, the switchers did not fail over days of
stress testing.
In order to make the operation mimic in-vehicular conditions more closely, the cable length between the supply
and the device was increased from about 30 cm to about
1.5 m. These longer wires, typically routed from the 12-V
battery to the device, created more inductance. Further­
more, the soft power cycling from the power supply was
replaced with a mechanical toggle switch such that the
mechanical bounce and chatter behaved more like transients introduced by mechanical relay contacts (Figure 9).
The tests were conducted with the power-supply output
set at 5.0 V, then the toggle switch was flip-flopped 20
times. If no overcurrent failure was detected, the supply
voltage was increased by 0.2 V, the switch was again toggled on and off 20 times, and the process repeated until
the part failed. The result was a stunning success! The
buck converter’s high- or low-side switch became shorted
with the power-supply output at about 7.5 VDC. The VIN
pins monitored with a 10-pF probe exhibited faster turnon transients, which caused an overshoot above 11 V in
20 µs. The actual L(di/dt) could have been a lot higher,
creating a repeatable destructive EOS condition. The
­customer was elated that this bench setup replicated the
same failures as in the field.
Figure 8. Power supply off/on ramping in ~6 ms
Soft Power On
(2 V/div)
1
Time (2 ms/div)
Figure 9. An 80-µs transient induced by switch
and longer wire
Power On
with Toggle Switch
(2 V/div)
1
Conclusion
This article has discussed common device-failure mechanisms related to MOSFET transistors in integrated powermanagement and voltage-regulator circuits. General precautions, specific PCB layout techniques, and componentselection tips have been presented to help mitigate and
eliminate EOS concerns. It is hoped that this article will
help system and PCB designers be aware of the EOS
effects of seemingly benign parasitic elements that can be
subjected to transients in the PMU operating environment.
Product and field support personnel may also find this
article useful for understanding the cause and effect of
EOS to facilitate their interface with customers.
Acknowledgments
The author thanks Ann Cocannon, Steven Jacobson,
ChiYoung Kim, Maxwell Goodman, Andy Strachan, and
Gavin Kirihara for their valuable input and support.
Time (40 µs/div)
2 Kern Wong. (2012, June 21). “High-performance loadand line-transient test jigs for mobile regulators.” EDN
[Online]. Available: www.edn.com
3 “Power MOS FET application note,” Rev.2.00, Renesas
Electronics, REJ05G0001, Aug. 23, 2004 [Online].
Available: www.renesas.com (use Search, Advanced
Search, Document Search to find REJ05G0001)
4 “Analysis of MOSFET failure modes in LLC resonant
converter,” Fairchild Semiconductor, AN-9067, Nov. 5,
2009 [Online]. Available: www.fairchildsemi.com/an/AN/
AN-9067.pdf
Related Web sites
References
1 Kern Wong. (2012, April 27).“PCB and ESR subtleties
in switching regulator and LDO designs.” EDN [Online].
Available: www.edn.com
Power Management:
www.ti.com/power-aaj
www.ti.com/lm26484-aaj
www.ti.com/tpd1e10B06-aaj
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35-V, single-channel gate drivers for IGBT and
MOSFET renewable-energy applications
By John Stevens
Systems Engineer, High-Performance Isolated Power
Introduction
The electronics market segment labeled as renewable
energy is a complex and diverse arena for electric power
conversion. In point-of-load applications, the switching
power converter typically is non-isolated; power levels are
fairly low (<200 W); and the power is usually converted
from one DC voltage to another, such as from 12 V to 3.3 V.
Further, the power-stage switches are integrated or capable
of being driven by low-current controllers or transistors.
Integration between the controller and power stage is
being realized today. Silicon (Si) MOSFETs dominate this
arena, where higher switching frequencies are preferred
and can reach speeds of greater than 1 MHz. These power
switches generally are driven by a 5- or 12-V IC gate driver
or similar solution.
Challenges to efficiently managing renewableenergy systems
in the first half of 2012, according to the Danish Energy
Agency. According to its parent agency, the Danish
Ministry of Climate, Energy and Building, Denmark has
committed to having 50% of its total power supply come
from wind by 2020. When wind energy makes up that
large a portion of a country’s total power, reliability of the
conversion system becomes critical. This—together with
the high-power connection to the grid, isolation safety
requirements, and the cost of large renewable-energy conversion systems—means that system reliability is always
the design priority, followed by efficiency. Therefore, protection features and reliability are preferred at all levels,
from the controller all the way down to the FET/IGBT
driver itself.
Typical power-management configuration
High power levels lead to higher system voltages, and
therefore higher standoff voltages, for the components
used within the converter. For lower power loss at greater
than 400 V, most circuit designers prefer to use insulatedgate bipolar transistors (IGBTs) or the latest silicon carbide
(SiC) FETs. These devices can have standoff voltages of
up to 1200 V, with lower ON resistance than equivalent Si
MOSFETs. These complex power systems often are managed by a digital signal processor, a microcontroller, or a
dedicated digital power controller. Thus, they usually
require both power and signal isolation from the noisy
switching environment of the power stage. Even during
steady-state switching cycles, the circuit can see massive
changes in both voltage and current that can create significant ground bouncing.
In the electronic power train from a wind or photovoltaic
power generator, there are some unique performance
challenges. Typical power levels for renewable energy can
range from 1 to 3 kW for micro-inverters, 3 to 10 kW for
string inverters, and 10 kW to 1 MW for large centralinverter stations. In addition to DC-to-DC conversion,
DC-to-AC and AC-to-DC conversion can also be used, and
sometimes a combination of the two.
Older wind turbines were tied directly to the power grid
but had to run at the power-line frequency. This made
them inefficient across the many operating points they
experienced. Newer wind turbines (Figure 1) often convert AC to DC and then DC back to AC so that the winddriven generator can run at variable speeds for maximum
efficiency.
Conversely, photovoltaic cells produce
Figure 1. Simplified power flow from wind
DC voltage/current. Generally, the voltage is
turbine to grid
boosted higher and then sent through a
DC-to-AC inverter before being tied to
the grid.
Renewable-energy trends
For most countries, generating renewable
energy from sources such as wind and solar
power makes up only a small percentage of
their total power portfolio. In recent history,
growth has been consistent year by year.
There are places where renewable energy
makes up a large share of the available power.
Denmark, for example, generated nearly 34%
of its total electricity from wind power alone
Generator
AC
to
DC
DC
to
AC
Grid
Gearbox
22
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Figure 2. Basic structure of single-phase inverter
Isolator
E/2 +
–
Isolator
Gate
Driver
Gate
Driver
Controller
Isolator
E/2 +
–
Figure 2 shows that even for a single-phase
DC-to-AC inverter, there are many gate
­drivers needed to properly switch the IGBTs
in the power stage. As a single-channel gate
driver, the Texas Instruments UCC27531 can
drive any of the switches in the switch bridge
if it has the necessary signal and bias isolation.
Signal isolation is achieved by using an optocoupler or digital isolator. For bias isolation,
the designer can use a bootstrap circuit with
a diode and a ­capacitor, or an isolated-bias
supply. Another option is to connect the gate
driver on the same side of the isolation as the
controller, then drive the switch through a
gate transformer after the gate driver itself.
This option allows the driver to be biased with
a non-isolated supply on the control side.
Isolator
Gate
Driver
Gate
Driver
Figure 3. Driving a power switch with FET/IGBT
single-gate drivers
UCC27531
EN
–
IN
VDD
+
OUTH
OUTL
GND
+ 18 V
–
13 V
+
–
Gate drivers in renewable energy
As a small, non-isolated gate driver, the singlechannel UCC27531 is a good fit for the environment described. Its input signals to the IC are provided
by an optocoupler or digital isolator. Its high supply/outputdrive voltage range of 10 to 35 V makes it ideal for 12-V Si
MOSFET applications as well as for IGBT/SiC FET applications. Here, a higher positive gate drive is typical, as well
as a negative voltage pull-down on shutoff to ­prevent the
power switch from false turn-on. Typically, SiC FETs are
driven by a +20/–5-V gate driver relative to the source.
Similarly, for IGBTs, system designers may use a +18/–13-V
gate drive, for example (see Figure 3).
Since the UCC27531 is a rail-to-rail driver, OUTH pulls
up the power-switch gate to its VDD of 18 V relative to the
emitter. OUTL pulls down the gate to the driver’s GND of
–13 V relative to the emitter. The driver effectively sees
+18 to –13 V, or 31 V from VDD relative to its own GND.
Further, the 35-V rating provides a margin to prevent
overvoltage failure of the IC due to noise and ringing.
The split output with both OUTH and OUTL permits the
user to control the turn-on (sourcing) current and turn-off
(sinking) current separately. This helps to maximize
­efficiency and maintain control of the switching times to
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comply with requirements for noise and electromagnetic
interference. Further, even with a split output, the singlegate driver maintains a minimum inductance on the output stage, preventing excessive ringing and overshoot. By
­having an asymmetrical drive (2.5-A turn-on and 5-A
turn-off), the UCC27531 is optimized for average switch
timing in high-power renewable-energy applications.
Further, with the low pull-down impedance, this driver
increases reliability by ensuring that the gate does not
experience voltage spikes that could lead to false turn-on
from the parasitic Miller-effect capacitance between the
collector and gate for IGBTs and between the drain and
gate for FETs. This internal capacitance can lead the gate
to exceed the turn-on threshold voltage by pulling up on
the gate when the collector/drain voltage rapidly increases
during turn-off of the switch.
The input stage of the UCC27531 is also designed for
high-reliability systems like renewable energy. It has a socalled TTL/CMOS input that is independent of the supply
voltage, allowing for compatibility with standard TTL-level
signals. It provides a higher hysteresis of about 1 V when
compared to the usual 0.5-V hysteresis seen in classic
TTL. If the input signal is lost and becomes floating for
any reason, the output is pulled low. Also, with the large
changes in voltage on the GND of the driver IC, it is possible for the input signals to appear negative if the GND
bounces high during a switching edge. This driver addresses
this concern by handling up to –5 V continuously on the
input (IN) or enable (EN) during these events.
The UCC27531 comes in a 3 x 3-mm, industry-standard
SOT-23 package, which is very competitive with a discrete
two-transistor solution that has a discrete level shifter
without negative-input capability or added protections.
Beyond the obvious space savings, integrating the
UCC27531’s functions into a single IC package increases
the system’s overall reliability.
This single-channel driver is an attractive option
because it can be located very close to the power-switch
gate. Placement is more flexible than for a combination
high-side/low-side gate driver in a single IC. This flexibility
helps minimize the inductance between the driver and
power switch and gives the designer better control of the
switch’s gate. Figure 2 shows how many high-power
switches are in just a single phase of a DC-to-AC stage.
Over a complete three-phase system with multiple conversions between DC and AC and back, and with boost stages
of DC-to-DC conversion also needed in some applications,
there becomes a need for many gate drivers. Each one
must be strategically placed on the PCB to ensure a
proper design.
Conclusion
In renewable-energy applications, conversion of power
generated from solar arrays and wind turbines presents
unique challenges to the system designer. These challenges include high voltages and power levels, meeting
safety and reliability requirements, and the overall complexity of the completely interconnected system. Although
gate drivers for power switches seem like a small part of
the total system control and power flow, they are actually
very important to the overall design performance.
References
1. F. Blaabjerg et al., “Power electronics and reliability in
renewable energy systems,” presented at 21st IEEE
International Symposium on Industrial Electronics,
Hangzhou, China, May 28–31, 2012.
2. Veda Prakash Galigekere and Marian K. Kazimierczuk.
(year, month). “Role of power electronics in renewable
energy systems” [Online]. Available under “White
Papers” at www.magnelab.com
3. Bob Calanan. (2011, Jan.). “Application considerations
for silicon carbide MOSFETs” [Online]. Available under
“Power Application Notes” at www.cree.com/power/
document-library
Related Web sites
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www.ti.com/ucc27531-aaj
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How to pick a linear regulator for
noise-sensitive applications
By Sureena Gupta
Marketing Engineer
Noise-sensitive applications require a power supply that
generates low internal noise and rejects noise from the
power source. These applications include test and
­measurement applications, medical equipment, communication equipment, base stations, and many others. A lownoise power supply is used to power a signal chain that
includes data converters, amplifiers, clocks, jitter cleaners,
PLLs, analog front ends and many other devices. A lownoise power solution is essential to preserving signal
­accuracy and integrity. This article addresses criteria and
parameters to consider in designing such a power solution,
including important specifications for picking a linear
­regulator.
The terms “power supply ripple rejection” (PSRR) and
“linear regulator” often are used together. The linear regulator’s high ripple rejection makes it an integral part of a
power solution. PSRR is a measure of how well the regulator filters a circuit by rejecting noise or ripple coming from
the power-supply input at various frequencies. In both
low-dropout regulators (LDOs) and linear regulators,
PSRR is a measure of output ripple compared to the input
ripple over a frequency range.
Since PSRR is calculated as ripple rejection, it is
expected to be a negative number. However, it is represented as a positive number in the datasheet so that a
higher number denotes higher noise rejection.
Mathematically, it is expressed in decibels as
 VIN _ ripple 
PSRR = 20 × log 
.
 VOUT _ ripple 
The PSRR of a linear regulator can be divided into three
frequency-range regions. The first region extends from DC
to the roll-off frequency. The ripple rejection in this region
is mostly dominated by open-loop gain and the bandgap
reference. The second region extends from the roll-off frequency to the unity-gain frequency. The PSRR in this
region is usually higher than in the first region and is
mainly dominated by the open-loop gain of the regulator.
The third region’s frequency range is above that of the
unity-gain frequency. The output capacitor, along with the
linear regulator’s parasitics (in the VIN-to-VOUT path),
dominates this region. Therefore, the values of the
selected output capacitor and its equivalent series resistance are quite important. This information can be found
in any datasheet.
In addition to VIN, VOUT, and system load requirements,
an engineer needs to know the frequency range of ripple
and noise in the system or power supply in order to select
linear regulators with a good PSRR in that frequency
range. For example, a switcher that switches at a frequency of 2 MHz may require a linear regulator that has a
high PSRR at around 2 MHz. Figure 1 shows a linear regulator’s high PSRR of about 55 dB at 2 MHz that helps to
remove input noise. Also, when PSRR graphs in the regulator datasheets are evaluated, it is always good to note
the dropout voltage at which the PSRR is measured. High
dropout voltage leads to better PSRR but reduces the
device’s efficiency.
Figure 1. Plot of linear regulator’s widebandwidth, high PSRR
90
80
PSRR Magnitude (dB)
70
60
50
40
30
VIN = 3 V
VOUT = 1.87 V
CIN = None
20
10
0
10
100
1k
10 k
100 k
Frequency (Hz)
1M
10 M
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Figure 2 shows a switching regulator’s spectral noise
that is fed to a linear regulator. It can be seen that the
switcher is operating at 500 kHz. Figure 3 shows the
­output spectrum of the Texas Instruments TPS7A4700
­linear regulator. The spike caused by the switcher at
500 kHz has been attenuated. If the power solution is not
designed for noise attenuation with high-PSRR linear
­regulators, the spike may show up at the output of the RF
voltage-­controlled oscillators, which after mixing affect
the PA performance. The spike may also fold back into
the audio band and create noise in an audio application.
Usually, noise and PSRR parameters are lumped
together in a linear regulator’s datasheet, which causes a
lot of confusion because noise and PSRR are two very
different characteristics. Noise is purely a physical
­phenomenon that occurs with transistors and resistors
Figure 2. Typical noise spectrum from a
switching regulator
Output Spectral Noise Density (µV Hz)
100
fSW = 500 kHz, VOUT = 4.2 V, IOUT = 500 mA
10
1
0.1
0.01
0.001
10
100
1k
10 k
100 k
1M
10 M
1M
10 M
Frequency (Hz)
Figure 3. Output noise spectrum of TPS7A4700
linear regulator with attenuated 500-kHz spike
Output Spectral Noise Density (µV Hz)
100
TPS7A4700 Output: VOUT =3.3 V, I OUT = 500 mA
10
1
0.1
0.01
0.001
10
100
1k
10 k
100 k
Frequency (Hz)
26
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inside the linear regulator on a very fundamental level.
This type of noise may include thermal, flicker, and shot
noise. Noise is usually indicated as a curve showing spectral noise density (in μV/√Hz) versus frequency (Figure 4).
Noise can also be indicated as integrated output noise (in
μVRMS), listed under the electrical characteristics table in
the datasheet (Figure 5). The output noise (in μVRMS) is
the spectral noise density integrated over a certain frequency range and can be seen as the total noise in a specified frequency range.
The next obvious question is whether an engineer
should look at spectral noise density or integrated output
noise, or both. The answer depends purely on the engineer’s application. For example, in RF applications where
the signal does not have any dependency on the frequency,
it makes more sense to look at the linear regulator’s spectral noise density. However, in applications where the
noise will be integrated by the system, such as powering
DACs and ADCs, the engineer should look at the linear
regulator’s integrated output noise instead.
Conclusion
This article has discussed the important specifications that
design engineers need to consider when picking a linear
regulator. It has also covered the criteria and parameters
to consider in designing a power solution for low-noise
applications. Given these guidelines, engineers should be
able to preserve signal accuracy and integrity in their
applications.
Reference
1. Thomas Neu, “Power-supply design for high-speed
ADCs,” Analog Applications Journal (1Q 2010).
Available www.ti.com/slyt366-aaj
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Figure 4. Spectral noise density for the TPS7A4700
Output Noise (µV/ Hz)
100
VOUT
VOUT
VOUT
VOUT
10
= 15 V, VNOISE = 12.28 µVRMS
= 10 V, VNOISE = 7.25 µVRMS
= 5 V, VNOISE = 4.67 µVRMS
= 1.4 V, VNOISE = 4.17 µVRMS
IOUT = 500 mA
COUT = 50 µF
CNR = 1 µF
BWRMSNOISE (10 Hz, 100 kHz)
1
0.1
0.01
10
100
1k
10 k
Frequency (Hz)
100 k
1M
Figure 5. Excerpt from TPS7A4700 datasheet showing integrated output noise voltage
PARAMETER
VNOISE
Output noise voltage
TEST CONDITIONS
TYP
UNIT
VIN = 3 V, VOUT(NOM) = 1.4 V, COUT = 50 µF,
CNR = 1 µF, BW = 10 Hz to 100 kHz
4.17
µVRMS
VIN = 6 V, VOUT(NOM) = 5 V, COUT = 50 µF,
CNR = 1 µF, BW = 10 Hz to 100 kHz
4.67
µVRMS
27
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Index of Articles
Texas Instruments Incorporated
Index of Articles
Title
Issue
Data Converters
Page
Grounding in mixed-signal systems demystified, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q 2013 . . . . . . . . . . 5
Add a digitally controlled PGA with noise filter to an ADC . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q 2013 . . . . . . . . . . 9
WEBENCH® tools and the photodetector’s stability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2012 . . . . . . . . . . 5
How delta-sigma ADCs work, Part 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2011 . . . . . . . . . . 5
How delta-sigma ADCs work, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2011 . . . . . . . . . 13
Clock jitter analyzed in the time domain, Part 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2011 . . . . . . . . . . 5
The IBIS model, Part 3: Using IBIS models to investigate signal-integrity issues . . . . . . . . . . . . . 2Q, 2011 . . . . . . . . . . 5
The IBIS model, Part 2: Determining the total quality of an IBIS model . . . . . . . . . . . . . . . . . . . . 1Q, 2011 . . . . . . . . . . 5
The IBIS model: A conduit into signal-integrity analysis, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2010 . . . . . . . . . 11
Clock jitter analyzed in the time domain, Part 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2010 . . . . . . . . . . 5
Clock jitter analyzed in the time domain, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2010 . . . . . . . . . . 5
How digital filters affect analog audio-signal levels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2010 . . . . . . . . . . 5
How the voltage reference affects ADC performance, Part 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2009 . . . . . . . . . . 5
How the voltage reference affects ADC performance, Part 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2009 . . . . . . . . . 13
Impact of sampling-clock spurs on ADC performance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2009 . . . . . . . . . . 5
How the voltage reference affects ADC performance, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2009 . . . . . . . . . . 5
Stop-band limitations of the Sallen-Key low-pass filter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2008 . . . . . . . . . . 5
A DAC for all precision occasions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2008 . . . . . . . . . . 5
Understanding the pen-interrupt (PENIRQ) operation of touch-screen controllers . . . . . . . . . . . 2Q, 2008 . . . . . . . . . . 5
Using a touch-screen controller’s auxiliary inputs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2007 . . . . . . . . . . 5
Calibration in touch-screen systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2007 . . . . . . . . . . 5
Conversion latency in delta-sigma converters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2007 . . . . . . . . . . 5
Clamp function of high-speed ADC THS1041 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2006 . . . . . . . . . . 5
Using the ADS8361 with the MSP430™ USI port . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2006 . . . . . . . . . . 5
Matching the noise performance of the operational amplifier to the ADC . . . . . . . . . . . . . . . . . . . 2Q, 2006 . . . . . . . . . . 5
Understanding and comparing datasheets for high-speed ADCs . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2006 . . . . . . . . . . 5
Low-power, high-intercept interface to the ADS5424 14-bit, 105-MSPS converter for
undersampling applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2005 . . . . . . . . . 10
Operating multiple oversampling data converters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2005 . . . . . . . . . . 5
Simple DSP interface for ADS784x/834x ADCs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2005 . . . . . . . . . 10
Using resistive touch screens for human/machine interface . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2005 . . . . . . . . . . 5
Implementation of 12-bit delta-sigma DAC with MSC12xx controller . . . . . . . . . . . . . . . . . . . . . . 1Q, 2005 . . . . . . . . . 27
Clocking high-speed data converters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2005 . . . . . . . . . 20
14-bit, 125-MSPS ADS5500 evaluation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2005 . . . . . . . . . 13
Supply voltage measurement and ADC PSRR improvement in MSC12xx devices . . . . . . . . . . . . 1Q, 2005 . . . . . . . . . . 5
Streamlining the mixed-signal path with the signal-chain-on-chip MSP430F169 . . . . . . . . . . . . . 3Q, 2004 . . . . . . . . . . 5
ADS809 analog-to-digital converter with large input pulse signal . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2004 . . . . . . . . . . 8
Two-channel, 500-kSPS operation of the ADS8361 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2004 . . . . . . . . . . 5
Evaluation criteria for ADSL analog front end . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2003 . . . . . . . . . 16
Calculating noise figure and third-order intercept in ADCs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2003 . . . . . . . . . 11
ADS82x ADC with non-uniform sampling clock . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2003 . . . . . . . . . . 5
Interfacing op amps and analog-to-digital converters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2002 . . . . . . . . . . 5
Using direct data transfer to maximize data acquisition throughput . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2002 . . . . . . . . . 14
MSC1210 debugging strategies for high-precision smart sensors . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2002 . . . . . . . . . . 7
Adjusting the A/D voltage reference to provide gain . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2002 . . . . . . . . . . 5
Synchronizing non-FIFO variations of the THS1206 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . 12
SHDSL AFE1230 application . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . . 5
Intelligent sensor system maximizes battery life: Interfacing the MSP430F123 Flash
MCU, ADS7822, and TPS60311 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2002 . . . . . . . . . . 5
A/D and D/A conversion of PC graphics and component video signals, Part 2: Software
and control . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . July 2001 . . . . . . . . . 5
A/D and D/A conversion of PC graphics and component video signals, Part 1: Hardware . . . . . February 2001 . . . . 11
Lit. No.
SLYT499
SLYT500
SLYT487
SLYT438
SLYT423
SLYT422
SLYT413
SLYT400
SLYT390
SLYT389
SLYT379
SLYT375
SLYT355
SLYT339
SLYT338
SLYT331
SLYT306
SLYT300
SLYT292
SLYT283
SLYT277
SLYT264
SLYT253
SLYT244
SLYT237
SLYT231
SLYT223
SLYT222
SLYT210
SLYT209A
SLYT076
SLYT075
SLYT074
SLYT073
SLYT078
SLYT083
SLYT082
SLYT091
SLYT090
SLYT089
SLYT104
SLYT111
SLYT110
SLYT109
SLYT115
SLYT114
SLYT123
SLYT129
SLYT138
28
High-Performance Analog Products
www.ti.com/aaj
1Q 2013
Analog Applications Journal
Index of Articles
Texas Instruments Incorporated
Title
Issue
Page
Lit. No.
Data Converters (Continued)
Using SPI synchronous communication with data converters — interfacing the
MSP430F149 and TLV5616 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . . . 7
Building a simple data acquisition system using the TMS320C31 DSP . . . . . . . . . . . . . . . . . . . . . February 2001 . . . . . 1
Using quad and octal ADCs in SPI mode . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . . 15
Hardware auto-identification and software auto-configuration for the
TLV320AIC10 DSP Codec — a “plug-and-play” algorithm . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . . . 8
Smallest DSP-compatible ADC provides simplest DSP interface . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . . . 1
Efficiently interfacing serial data converters to high-speed DSPs . . . . . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . 10
Higher data throughput for DSP analog-to-digital converters . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . . 5
New DSP development environment includes data converter plug-ins . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . . 1
Introduction to phase-locked loop system modeling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . May 2000 . . . . . . . . . 5
The design and performance of a precision voltage reference circuit for 14-bit and
16-bit A-to-D and D-to-A converters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . May 2000 . . . . . . . . . 1
The operation of the SAR-ADC based on charge redistribution . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2000 . . . . 10
A methodology of interfacing serial A-to-D converters to DSPs . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2000 . . . . . 1
Techniques for sampling high-speed graphics with lower-speed A/D converters . . . . . . . . . . . . . November 1999 . . . . 5
Precision voltage references . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 1999 . . . . 1
Evaluating operational amplifiers as input amplifiers for A-to-D converters . . . . . . . . . . . . . . . . . August 1999 . . . . . . . 7
Low-power data acquisition sub-system using the TI TLV1572 . . . . . . . . . . . . . . . . . . . . . . . . . . . August 1999 . . . . . . . 4
Aspects of data acquisition system design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 1999 . . . . . . . 1
Power Management
Design of a 60-A interleaved active-clamp forward converter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q 2013 . . . . . . . . . 13
Power MOSFET failures in mobile PMUs: Causes and design precautions . . . . . . . . . . . . . . . . . . 1Q 2013 . . . . . . . . . 17
35-V, single-channel gate drivers for IGBT and MOSFET renewable-energy applications . . . . . . 1Q 2013 . . . . . . . . . 22
How to pick a linear regulator for noise-sensitive applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q 2013 . . . . . . . . . 25
Simple open-circuit protection for boost converters in LED driver applications . . . . . . . . . . . . . .4Q, 2012 . . . . . . . . . 21
LDO noise examined in detail . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2012 . . . . . . . . . 14
Harnessing wasted energy in 4- to 20-mA current-loop systems . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2012 . . . . . . . . . 10
Designing a Qi-compliant receiver coil for wireless power systems, Part 1 . . . . . . . . . . . . . . . . . . 3Q, 2012 . . . . . . . . . . 8
Easy solar-panel maximum-power-point tracking for pulsed-load applications . . . . . . . . . . . . . . . 3Q, 2012 . . . . . . . . . . 5
Design considerations for a resistive feedback divider in a DC/DC converter . . . . . . . . . . . . . . . . 2Q, 2012 . . . . . . . . . 18
Charging a three-cell nickel-based battery pack with a Li-Ion charger . . . . . . . . . . . . . . . . . . . . . 2Q, 2012 . . . . . . . . . 14
Remote sensing for power supplies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2012 . . . . . . . . . 12
A solar-powered buck/boost battery charger . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2012 . . . . . . . . . . 8
Controlling switch-node ringing in synchronous buck converters . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2012 . . . . . . . . . . 5
High-efficiency AC adapters for USB charging . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2012 . . . . . . . . . 18
Downslope compensation for buck converters when the duty cycle exceeds 50% . . . . . . . . . . . . 1Q, 2012 . . . . . . . . . 14
Benefits of a multiphase buck converter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2012 . . . . . . . . . . 8
Turbo-boost charger supports CPU turbo mode . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2012 . . . . . . . . . . 5
Solar lantern with dimming achieves 92% efficiency . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2011 . . . . . . . . . 12
Solar charging solution provides narrow-voltage DC/DC system bus for
multicell-battery applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2011 . . . . . . . . . . 8
A boost-topology battery charger powered from a solar panel . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2011 . . . . . . . . . 17
Challenges of designing high-frequency, high-input-voltage DC/DC converters . . . . . . . . . . . . . . 2Q, 2011 . . . . . . . . . 28
Backlighting the tablet PC . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2011 . . . . . . . . . 23
IQ: What it is, what it isn’t, and how to use it . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2011 . . . . . . . . . 18
Benefits of a coupled-inductor SEPIC converter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2011 . . . . . . . . . 14
Implementation of microprocessor-controlled, wide-input-voltage, SMBus smart
battery charger . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2011 . . . . . . . . . 11
Fine-tuning TI’s Impedance Track™ battery fuel gauge with LiFePO4 cells in
shallow-discharge applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2011 . . . . . . . . . 13
An introduction to the Wireless Power Consortium standard and TI’s compliant solutions . . . . . 1Q, 2011 . . . . . . . . . 10
Save power with a soft Zener clamp . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2010 . . . . . . . . . 19
A low-cost, non-isolated AC/DC buck converter with no transformer . . . . . . . . . . . . . . . . . . . . . . 4Q, 2010 . . . . . . . . . 16
Computing power going “Platinum” . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2010 . . . . . . . . . 13
Coupled inductors broaden DC/DC converter usage . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2010 . . . . . . . . . 10
SLYT137
SLYT136
SLYT150
SLYT149
SLYT148
SLYT160
SLYT159
SLYT158
SLYT169
SLYT168
SLYT176
SLYT175
SLYT184
SLYT183
SLYT193
SLYT192
SLYT191
SLYT501
SLYT502
SLYT503
SLYT504
SLYT490
SLYT489
SLYT488
SLYT479
SLYT478
SLYT469
SLYT468
SLYT467
SLYT466
SLYT465
SLYT451
SLYT450
SLYT449
SLYT448
SLYT440
SLYT439
SLYT424
SLYT415
SLYT414
SLYT412
SLYT411
SLYT410
SLYT402
SLYT401
SLYT392
SLYT391
SLYT382
SLYT380
29
Analog Applications Journal
1Q 2013
www.ti.com/aaj
High-Performance Analog Products
Index of Articles
Texas Instruments Incorporated
Title
Issue
Power Management (Continued)
Page
Designing DC/DC converters based on ZETA topology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2010 . . . . . . . . . 16
Discrete design of a low-cost isolated 3.3- to 5-V DC/DC converter . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2010 . . . . . . . . . 12
Power-supply design for high-speed ADCs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2010 . . . . . . . . . 12
Li-Ion battery-charger solutions for JEITA compliance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2010 . . . . . . . . . . 8
Fuel-gauging considerations in battery backup storage systems . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2010 . . . . . . . . . . 5
Efficiency of synchronous versus nonsynchronous buck converters . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2009 . . . . . . . . . 15
Designing a multichemistry battery charger . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2009 . . . . . . . . . 13
Using power solutions to extend battery life in MSP430™ applications . . . . . . . . . . . . . . . . . . . . 4Q, 2009 . . . . . . . . . 10
Reducing radiated EMI in WLED drivers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .3Q, 2009 . . . . . . . . . 17
Selecting the right charge-management solution . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2009 . . . . . . . . . 18
Designing a linear Li-Ion battery charger with power-path control . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2009 . . . . . . . . . 12
Taming linear-regulator inrush currents . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2009 . . . . . . . . . . 9
Using a portable-power boost converter in an isolated flyback application . . . . . . . . . . . . . . . . . . 1Q, 2009 . . . . . . . . . 19
Cell balancing buys extra run time and battery life . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2009 . . . . . . . . . 14
Improving battery safety, charging, and fuel gauging in portable media applications . . . . . . . . . . 1Q, 2009 . . . . . . . . . . 9
Paralleling power modules for high-current applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2009 . . . . . . . . . . 5
Designing DC/DC converters based on SEPIC topology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2008 . . . . . . . . . 18
Compensating and measuring the control loop of a high-power LED driver . . . . . . . . . . . . . . . . . 4Q, 2008 . . . . . . . . . 14
Getting the most battery life from portable systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2008 . . . . . . . . . . 8
New current-mode PWM controllers support boost, flyback, SEPIC, and
LED-driver applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2008 . . . . . . . . . . 9
Battery-charger front-end IC improves charging-system safety . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2008 . . . . . . . . . 14
Understanding output voltage limitations of DC/DC buck converters . . . . . . . . . . . . . . . . . . . . . . 2Q, 2008 . . . . . . . . . 11
Using a buck converter in an inverting buck-boost topology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2007 . . . . . . . . . 16
Host-side gas-gauge-system design considerations for single-cell handheld applications . . . . . . 4Q, 2007 . . . . . . . . . 12
Driving a WLED does not always require 4 V . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2007 . . . . . . . . . . 9
Simultaneous power-down sequencing with the TPS74x01 family of linear regulators . . . . . . . . 3Q, 2007 . . . . . . . . . 20
Get low-noise, low-ripple, high-PSRR power with the TPS717xx . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2007 . . . . . . . . . 17
TPS6108x: A boost converter with extreme versatility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2007 . . . . . . . . . 14
Power-management solutions for telecom systems improve performance, cost, and size . . . . . . 3Q, 2007 . . . . . . . . . 10
Current balancing in four-pair, high-power PoE applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2007 . . . . . . . . . 11
Enhanced-safety, linear Li-Ion battery charger with thermal regulation and
input overvoltage protection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2007 . . . . . . . . . . 8
Power management for processor core voltage requirements . . . . . . . . . . . . . . . . . . . . . . . . . . . . .1Q, 2007 . . . . . . . . . 11
LDO white-LED driver TPS7510x provides incredibly small solution size . . . . . . . . . . . . . . . . . . . 1Q, 2007 . . . . . . . . . . 9
Selecting the correct IC for power-supply applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2007 . . . . . . . . . . 5
Fully integrated TPS6300x buck-boost converter extends Li-Ion battery life . . . . . . . . . . . . . . . . 4Q, 2006 . . . . . . . . . 15
bq25012 single-chip, Li-Ion charger and dc/dc converter for Bluetooth® headsets . . . . . . . . . . . 4Q, 2006 . . . . . . . . . 13
A 3-A, 1.2-VOUT linear regulator with 80% efficiency and PLOST < 1 W . . . . . . . . . . . . . . . . . . . . . 4Q, 2006 . . . . . . . . . 10
Complete battery-pack design for one- or two-cell portable applications . . . . . . . . . . . . . . . . . . . 3Q, 2006 . . . . . . . . . 14
Single-chip bq2403x power-path manager charges battery while powering system . . . . . . . . . . . 3Q, 2006 . . . . . . . . . 12
TPS65552A powers portable photoflash . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2006 . . . . . . . . . 10
TPS61059 powers white-light LED as photoflash or movie light . . . . . . . . . . . . . . . . . . . . . . . . . . .3Q, 2006 . . . . . . . . . . 8
Powering today’s multi-rail FPGAs and DSPs, Part 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2006 . . . . . . . . . 18
Wide-input dc/dc modules offer maximum design flexibility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2006 . . . . . . . . . 13
TLC5940 PWM dimming provides superior color quality in LED video displays . . . . . . . . . . . . . . 2Q, 2006 . . . . . . . . . 10
Practical considerations when designing a power supply with the TPS6211x . . . . . . . . . . . . . . . . 1Q, 2006 . . . . . . . . . 17
TPS79918 RF LDO supports migration to StrataFlash® Embedded Memory (P30) . . . . . . . . . . . 1Q, 2006 . . . . . . . . . 14
Powering today’s multi-rail FPGAs and DSPs, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2006 . . . . . . . . . . 9
TLC5940 dot correction compensates for variations in LED brightness . . . . . . . . . . . . . . . . . . . . 4Q, 2005 . . . . . . . . . 21
Li-Ion switching charger integrates power FETs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2005 . . . . . . . . . 19
New power modules improve surface-mount manufacturability . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2005 . . . . . . . . . 18
Miniature solutions for voltage isolation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2005 . . . . . . . . . 13
Understanding power supply ripple rejection in linear regulators . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2005 . . . . . . . . . . 8
Understanding noise in linear regulators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2005 . . . . . . . . . . 5
A better bootstrap/bias supply circuit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2005 . . . . . . . . . 33
Tips for successful power-up of today’s high-performance FPGAs . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2004 . . . . . . . . . 11
Lit. No.
SLYT372
SLYT371
SLYT366
SLYT365
SLYT364
SLYT358
SLYT357
SLYT356
SLYT340
SLYT334
SLYT333
SLYT332
SLYT323
SLYT322
SLYT321
SLYT320
SLYT309
SLYT308
SLYT307
SLYT302
SLYT294
SLYT293
SLYT286
SLYT285
SLYT284
SLYT281
SLYT280
SLYT279
SLYT278
SLYT270
SLYT269
SLYT261
SLYT260
SLYT259
SLYT256
SLYT255
SLYT254
SLYT248
SLYT247
SLYT246
SLYT245
SLYT240
SLYT239
SLYT238
SLYT234
SLYT233
SLYT232
SLYT225
SLYT224
SLYT212
SLYT211
SLYT202
SLYT201
SLYT077
SLYT079
30
High-Performance Analog Products
www.ti.com/aaj
1Q 2013
Analog Applications Journal
Index of Articles
Texas Instruments Incorporated
Title
Issue
Page
Power Management (Continued)
LED-driver considerations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2004 . . . . . . . . . 14
UCC28517 100-W PFC power converter with 12-V, 8-W bias supply, Part 2 . . . . . . . . . . . . . . . . . 4Q, 2003 . . . . . . . . . 21
UCC28517 100-W PFC power converter with 12-V, 8-W bias supply, Part 1 . . . . . . . . . . . . . . . . . 3Q, 2003 . . . . . . . . . 13
Soft-start circuits for LDO linear regulators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2003 . . . . . . . . . 10
Auto-Track™ voltage sequencing simplifies simultaneous power-up and power-down . . . . . . . . 3Q, 2003 . . . . . . . . . . 5
Using the TPS61042 white-light LED driver as a boost converter . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2003 . . . . . . . . . . 7
Load-sharing techniques: Paralleling power modules with overcurrent protection . . . . . . . . . . . 1Q, 2003 . . . . . . . . . . 5
Understanding piezoelectric transformers in CCFL backlight applications . . . . . . . . . . . . . . . . . . 4Q, 2002 . . . . . . . . . 18
Power conservation options with dynamic voltage scaling in portable DSP designs . . . . . . . . . . . 4Q, 2002 . . . . . . . . . 12
Using the UCC3580-1 controller for highly efficient 3.3-V/100-W isolated supply design . . . . . . . 4Q, 2002 . . . . . . . . . . 8
Powering electronics from the USB port . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . 28
Optimizing the switching frequency of ADSL power supplies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . 23
SWIFT™ Designer power supply design program . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . 15
Why use a wall adapter for ac input power? . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2002 . . . . . . . . . 18
Comparing magnetic and piezoelectric transformer approaches in CCFL applications . . . . . . . . 1Q, 2002 . . . . . . . . . 12
Power control design key to realizing InfiniBandSM benefits . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2002 . . . . . . . . . 10
Runtime power control for DSPs using the TPS62000 buck converter . . . . . . . . . . . . . . . . . . . . . July 2001 . . . . . . . . 15
Power supply solution for DDR bus termination . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . July 2001 . . . . . . . . . 9
–48-V/+48-V hot-swap applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . . 20
Optimal design for an interleaved synchronous buck converter under high-slew-rate,
load-current transient conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . . 15
Comparison of different power supplies for portable DSP solutions working from a
single-cell battery . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . . 24
Understanding the load-transient response of LDOs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . . 19
Optimal output filter design for microprocessor or DSP power supply . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . 22
Advantages of using PMOS-type low-dropout linear regulators in battery applications . . . . . . . August 2000 . . . . . . 16
Low-cost, minimum-size solution for powering future-generation Celeron™-type
processors with peak currents up to 26 A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . May 2000 . . . . . . . . 14
Simple design of an ultra-low-ripple DC/DC boost converter with TPS60100 charge pump . . . . May 2000 . . . . . . . . 11
Powering Celeron-type microprocessors using TI’s TPS5210 and TPS5211 controllers . . . . . . . . February 2000 . . . . 20
Power supply solutions for TI DSPs using synchronous buck converters . . . . . . . . . . . . . . . . . . . February 2000 . . . . 12
Understanding the stable range of equivalent series resistance of an LDO regulator . . . . . . . . . . November 1999 . . . 14
Synchronous buck regulator design using the TI TPS5211 high-frequency
hysteretic controller . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 1999 . . . 10
TI TPS5602 for powering TI’s DSP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 1999 . . . . 8
Migrating from the TI TL770x to the TI TLC770x . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 1999 . . . . . . 14
Extended output voltage adjustment (0 V to 3.5 V) using the TI TPS5210 . . . . . . . . . . . . . . . . . . August 1999 . . . . . . 13
Stability analysis of low-dropout linear regulators with a PMOS pass element . . . . . . . . . . . . . . . August 1999 . . . . . . 10
Interface (Data Transmission)
Design considerations for system-level ESD circuit protection . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2012 . . . . . . . . .
How to design an inexpensive HART transmitter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2012 . . . . . . . . .
Data-rate independent half-duplex repeater design for RS-485 . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2012 . . . . . . . . .
Extending the SPI bus for long-distance communication . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2011 . . . . . . . . .
Industrial data-acquisition interfaces with digital isolators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2011 . . . . . . . . .
Isolated RS-485 transceivers support DMX512 stage lighting and special-effects applications . . 3Q, 2011 . . . . . . . . .
Designing an isolated I2C Bus® interface by using digital isolators . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2011 . . . . . . . . .
Interfacing high-voltage applications to low-power controllers . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2010 . . . . . . . . .
Magnetic-field immunity of digital capacitive isolators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .3Q, 2010 . . . . . . . . .
Designing with digital isolators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2009 . . . . . . . . .
Message priority inversion on a CAN bus . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2009 . . . . . . . . .
RS-485: Passive failsafe for an idle bus . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2009 . . . . . . . . .
Cascading of input serializers boosts channel density for digital inputs . . . . . . . . . . . . . . . . . . . . . 3Q, 2008 . . . . . . . . .
When good grounds turn bad—isolate! . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2008 . . . . . . . . .
Enabling high-speed USB OTG functionality on TI DSPs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2007 . . . . . . . . .
Detection of RS-485 signal loss . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2006 . . . . . . . . .
Improved CAN network security with TI’s SN65HVD1050 transceiver . . . . . . . . . . . . . . . . . . . . . 3Q, 2006 . . . . . . . . .
28
24
15
16
24
21
17
20
19
21
25
22
16
11
18
18
17
Lit. No.
SLYT084
SLYT092
SLYT097
SLYT096
SLYT095
SLYT101
SLYT100
SLYT107
SLYT106
SLYT105
SLYT118
SLYT117
SLYT116
SLYT126
SLYT125
SLYT124
SLYT131
SLYT130
SLYT140
SLYT139
SLYT152
SLYT151
SLYT162
SLYT161
SLYT171
SLYT170
SLYT178
SLYT177
SLYT187
SLYT186
SLYT185
SLYT196
SLYT195
SLYT194
SLYT492
SLYT491
SLYT480
SLYT441
SLYT426
SLYT425
SLYT403
SLYT393
SLYT381
SLYT335
SLYT325
SLYT324
SLYT301
SLYT298
SLYT271
SLYT257
SLYT249
31
Analog Applications Journal
1Q 2013
www.ti.com/aaj
High-Performance Analog Products
Index of Articles
Texas Instruments Incorporated
Title
Issue
Page
Interface (Data Transmission) (Continued)
Device spacing on RS-485 buses . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2006 . . . . . . . . .
Maximizing signal integrity with M-LVDS backplanes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2005 . . . . . . . . .
Failsafe in RS-485 data buses . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2004 . . . . . . . . .
The RS-485 unit load and maximum number of bus connections . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2004 . . . . . . . . .
Estimating available application power for Power-over-Ethernet applications . . . . . . . . . . . . . . . 1Q, 2004 . . . . . . . . .
Power consumption of LVPECL and LVDS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2002 . . . . . . . . .
The SN65LVDS33/34 as an ECL-to-LVTTL converter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . July 2001 . . . . . . . .
The Active Fail-Safe feature of the SN65LVDS32A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . .
A statistical survey of common-mode noise . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . .
Performance of LVDS with different cables . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . .
LVDS: The ribbon cable connection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . May 2000 . . . . . . . .
LVDS receivers solve problems in non-LVDS applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2000 . . . .
Skew definition and jitter analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2000 . . . .
Keep an eye on the LVDS input levels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 1999 . . .
TIA/EIA-568A Category 5 cables in low-voltage differential signaling (LVDS) . . . . . . . . . . . . . . . August 1999 . . . . . .
Amplifiers: Audio
Precautions for connecting APA outputs to other devices . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2010 . . . . . . . . .
Audio power amplifier measurements, Part 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2002 . . . . . . . . .
Audio power amplifier measurements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . July 2001 . . . . . . . .
An audio circuit collection, Part 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . July 2001 . . . . . . . .
An audio circuit collection, Part 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . .
Notebook computer upgrade path for audio power amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . .
1.6- to 3.6-volt BTL speaker driver reference design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . .
An audio circuit collection, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . .
PCB layout for the TPA005D1x and TPA032D0x Class-D APAs . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2000 . . . .
Power supply decoupling and audio signal filtering for the Class-D audio power amplifier . . . . . August 1999 . . . . . .
Reducing the output filter of a Class-D amplifier . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 1999 . . . . . .
Amplifiers: Op Amps
Using a fixed threshold in ultrasonic distance-ranging automotive applications . . . . . . . . . . . . . . 3Q, 2012 . . . . . . . . .
Source resistance and noise considerations in amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2012 . . . . . . . . .
Measuring op amp settling time by using sample-and-hold technique . . . . . . . . . . . . . . . . . . . . . . 1Q, 2012 . . . . . . . . .
Converting single-ended video to differential video in single-supply systems . . . . . . . . . . . . . . . . 3Q, 2011 . . . . . . . . .
Using single-supply fully differential amplifiers with negative input voltages to drive ADCs . . . . 4Q, 2010 . . . . . . . . .
Operational amplifier gain stability, Part 3: AC gain-error analysis . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2010 . . . . . . . . .
Operational amplifier gain stability, Part 2: DC gain-error analysis . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2010 . . . . . . . . .
Interfacing op amps to high-speed DACs, Part 3: Current-sourcing DACs simplified . . . . . . . . . . 1Q, 2010 . . . . . . . . .
Signal conditioning for piezoelectric sensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2010 . . . . . . . . .
Operational amplifier gain stability, Part 1: General system analysis . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2010 . . . . . . . . .
Interfacing op amps to high-speed DACs, Part 2: Current-sourcing DACs . . . . . . . . . . . . . . . . . . 4Q, 2009 . . . . . . . . .
Using fully differential op amps as attenuators, Part 3: Single-ended unipolar input signals . . . . 4Q, 2009 . . . . . . . . .
Using the infinite-gain, MFB filter topology in fully differential active filters . . . . . . . . . . . . . . . . 3Q, 2009 . . . . . . . . .
Interfacing op amps to high-speed DACs, Part 1: Current-sinking DACs . . . . . . . . . . . . . . . . . . . . 3Q, 2009 . . . . . . . . .
Using fully differential op amps as attenuators, Part 2: Single-ended bipolar input signals . . . . . 3Q, 2009 . . . . . . . . .
Using fully differential op amps as attenuators, Part 1: Differential bipolar input signals . . . . . . 2Q, 2009 . . . . . . . . .
Output impedance matching with fully differential operational amplifiers . . . . . . . . . . . . . . . . . . 1Q, 2009 . . . . . . . . .
A dual-polarity, bidirectional current-shunt monitor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2008 . . . . . . . . .
Input impedance matching with fully differential amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2008 . . . . . . . . .
A new filter topology for analog high-pass filters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2008 . . . . . . . . .
New zero-drift amplifier has an IQ of 17 µA . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2007 . . . . . . . . .
Accurately measuring ADC driving-circuit settling time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2007 . . . . . . . . .
Low-cost current-shunt monitor IC revives moving-coil meter design . . . . . . . . . . . . . . . . . . . . . . 2Q, 2006 . . . . . . . . .
High-speed notch filters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2006 . . . . . . . . .
Getting the most out of your instrumentation amplifier design . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2005 . . . . . . . . .
So many amplifiers to choose from: Matching amplifiers to applications . . . . . . . . . . . . . . . . . . . . 3Q, 2005 . . . . . . . . .
Auto-zero amplifiers ease the design of high-precision circuits . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2005 . . . . . . . . .
Lit. No.
25
11
16
21
18
23
19
35
30
30
19
33
29
17
16
SLYT241
SLYT203
SLYT080
SLYT086
SLYT085
SLYT127
SLYT132
SLYT154
SLYT153
SLYT163
SLYT172
SLYT180
SLYT179
SLYT188
SLYT197
22
26
40
34
41
27
23
39
39
24
19
SLYT373
SLYT128
SLYT135
SLYT134
SLYT145
SLYT142
SLYT141
SLYT155
SLYT182
SLYT199
SLYT198
19
23
21
29
26
23
24
32
24
20
23
19
33
24
21
33
29
29
24
18
22
14
27
19
25
24
19
SLYT481
SLYT470
SLYT452
SLYT427
SLYT394
SLYT383
SLYT374
SLYT368
SLYT369
SLYT367
SLYT360
SLYT359
SLYT343
SLYT342
SLYT341
SLYT336
SLYT326
SLYT311
SLYT310
SLYT299
SLYT272
SLYT262
SLYT242
SLYT235
SLYT226
SLYT213
SLYT204
32
High-Performance Analog Products
www.ti.com/aaj
1Q 2013
Analog Applications Journal
Index of Articles
Texas Instruments Incorporated
Title
Issue
Page
Amplifiers: Op Amps (Continued)
Active filters using current-feedback amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2004 . . . . . . . . . 21
Integrated logarithmic amplifiers for industrial applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2004 . . . . . . . . . 28
Op amp stability and input capacitance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2004 . . . . . . . . . 24
Calculating noise figure in op amps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2003 . . . . . . . . . 31
Expanding the usability of current-feedback amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2003 . . . . . . . . . 23
Video switcher using high-speed op amps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2003 . . . . . . . . . 20
Analyzing feedback loops containing secondary amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2003 . . . . . . . . . 14
RF and IF amplifiers with op amps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1Q, 2003 . . . . . . . . . . 9
Active output impedance for ADSL line drivers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2002 . . . . . . . . . 24
FilterPro™ low-pass design tool . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2002 . . . . . . . . . 24
Using high-speed op amps for high-performance RF design, Part 2 . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2002 . . . . . . . . . 21
Using high-speed op amps for high-performance RF design, Part 1 . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . 46
Worst-case design of op amp circuits . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . 42
Fully differential amplifier design in high-speed data acquisition systems . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . . 35
Designing for low distortion with high-speed op amps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . July 2001 . . . . . . . . 25
Frequency response errors in voltage feedback op amps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . . 48
Pressure transducer-to-ADC application . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . . 38
Fully differential amplifiers applications: Line termination, driving high-speed ADCs,
and differential transmission lines . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . . 32
Analysis of fully differential amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . . 48
Thermistor temperature transducer-to-ADC application . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 2000 . . . 44
Reducing PCB design costs: From schematic capture to PCB layout . . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . 48
The PCB is a component of op amp design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . 42
Fully differential amplifiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . 38
Design of op amp sine wave oscillators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . August 2000 . . . . . . 33
Using a decompensated op amp for improved performance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . May 2000 . . . . . . . . 26
Sensor to ADC — analog interface design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . May 2000 . . . . . . . . 22
Matching operational amplifier bandwidth with applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2000 . . . . 36
Reducing crosstalk of an op amp on a PCB . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 1999 . . . 23
Single-supply op amp design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . November 1999 . . . 20
Low-Power RF
Selecting antennas for low-power wireless applications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2008 . . . . . . . . . 20
Using the CC2430 and TIMAC for low-power wireless sensor applications:
A power- consumption study . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2008 . . . . . . . . . 17
General Interest
High-definition haptics: Feel the difference! . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3Q, 2012 . . . . . . . . .
Applying acceleration and deceleration profiles to bipolar stepper motors . . . . . . . . . . . . . . . . . . 3Q, 2012 . . . . . . . . .
Industrial flow meters/flow transmitters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2012 . . . . . . . . .
Analog linearization of resistance temperature detectors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4Q, 2011 . . . . . . . . .
Spreadsheet modeling tool helps analyze power- and ground-plane voltage drops
to keep core voltages within tolerance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2007 . . . . . . . . .
Analog design tools . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2Q, 2002 . . . . . . . . .
Synthesis and characterization of nickel manganite from different carboxylate
precursors for thermistor sensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . February 2001 . . . .
Lit. No.
SLYT081
SLYT088
SLYT087
SLYT094
SLYT099
SLYT098
SLYT103
SLYT102
SLYT108
SLYT113
SLYT112
SLYT121
SLYT120
SLYT119
SLYT133
SLYT146
SLYT144
SLYT143
SLYT157
SLYT156
SLYT167
SLYT166
SLYT165
SLYT164
SLYT174
SLYT173
SLYT181
SLYT190
SLYT189
SLYT296
SLYT295
29
24
29
21
SLYT483
SLYT482
SLYT471
SLYT442
29
50
SLYT273
SLYT122
52
SLYT147
33
Analog Applications Journal
1Q 2013
www.ti.com/aaj
High-Performance Analog Products
Index of Articles
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High-Performance Analog Products
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1Q 2013
Analog Applications Journal
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