ETC AB-075

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With their low-input currents, FET input op amps are universally used in monitoring photodetectors, the most common of which are photodiodes. There are a variety of
amplifier connections for this purpose and the choice is
based on linearity, offset, noise and bandwidth considerations. These same factors influence the selection of the
amplifier with newer devices offering very low-input currents, low noise and high speed.
Photodetectors are the bridge between a basic physical
indicator and electronics resulting in the largest single usage
of FET op amps. As a measure of physical conditions, light
is secondary to temperature and pressure until the measurement is made remotely with no direct contact to the monitored object. Then, the signals of a CAT scanner, startracking instrument or electron microscope depend on light
for the final link to signal processing. Photodiodes have
made that link economical and expanded usage to detector
arrays that employ more than 1000 light sensors. Focus then
turns to accurate conversion of the photodiode output to a
linearly related electrical signal. As always, this is a contest
between speed and resolution with noise as a basic limiting
element. Central to the contest is the seemingly simple
current-to-voltage converter which displays surprising multidimensional constraints and suggests alternative configurations for many optimizations.
from the input. That is the key to the basic current-to-voltage
converter connection of Figure 1b. It provides an input
resistance of R1/A where A is the open-loop gain of the op
amp. Even though R1 is generally very large, the resulting
input resistance remains negligible in comparison to the
output resistance of photodiodes.
The energy transmitted by light to a photodiode can be
measured as either a voltage or current output. For a voltage
response, the diode must be monitored from a high impedance that does not draw significant signal current. That
condition is provided by Figure 1a. Here, the photodiode is
in series with the input of an op amp where ideally zero
current flows. That op amp has feedback set by R1 and R2 to
establish amplification of the voltage diode just as if it was
an offset voltage of the amplifier. While appealing to more
common op amp thinking, this voltage mode is nonlinear.
The response has a logarithmic relationship to the light
energy received since the sensitivity of the diode varies with
its voltage.
Constant voltage for a fixed sensitivity suggests current
output instead and that response is linearly related to the
incident light energy. A monitor of that current must have
zero input impedance to respond with no voltage across the
diode. Zero impedance is the role of an op amp virtual
ground as high-amplifier loop gain removes voltage swing
1994 Burr-Brown Corporation
eO = (1 + R2/R1) (KT/q) In (1 + IP/IS)
A1: OPA128
D1: HP5082-4204
eO = IPR1
A1: OPA111
D1: HP5082-4204
FIGURE 1a. Photodiode Output Can be Monitored as a
Voltage; or, 1b, as a Current.
Diode current is not accepted by the input of the op amp as
its presence stimulates the high amplifier gain to receive that
current through the feedback resistor, R1. To do so, the
amplifier develops an output voltage equal to the diode
current times the feedback resistance, R1. For that current-tovoltage gain to be high, R1 is made as large as other
constraints will permit. At higher resistance levels, that
resistor begins to develop significant thermal DC voltage
Printed in U.S.A. January, 1995
drift due to the temperature coefficient of the amplifier input
current. To compensate this error, an equal resistance R2 is
commonly connected in series with the op amp noninverting
input, as shown, and capacitively bypassed to remove most
of its noise. The remaining DC error is determined by the
mismatches between the amplifier input currents and between the two resistors. A drawback of this error correction
is the voltage drop it creates across the diode and the
resulting diode leakage current. That leakage can override
the correction achieved with R2, as photodiodes typically
have large junction areas for high sensitivity. Leakage current is proportional to that area which can become much
larger than the op amp input currents.
Only zero diode voltage can eliminate this new error source
but that is in conflict with control of a second attribute of
large diode area. Large parasitic capacitance is also present
creating often severe amplification of noise as will be
described. To reduce that capacitance, a large reverse-bias
voltage is sometimes impressed on the diode greatly complicating DC stability and making current noise from the
photodiode an additional error factor. Larger diode area may
actually degrade overall accuracy and higher photo sensitivity should first be sought through optical means such as a
package with an integral molded lens. Monitor-circuit configurations that maintain zero diode voltage are also candidates in this optimization and are described with Figures 6,
7 and 9.
The value of the feedback resistor in a current-to-voltage
converter largely determines noise and bandwidth as well as
gain. Noise contributed directly by the resistor has a spectral
density of √4KTR1 and appears directly at the output of a
current-to-voltage converter without amplification. Increasing the size of the resistor not only raises output noise by a
square root relationship but also increases output signal by
a direct proportionality. Signal-to-noise ratio, then, tends to
increase by the square root of the resistance.
a response zero due to CD and begins a rise that is terminated
only because of a second parasitic capacitance. Stray capacitance, CS, shunts the feedback resistor resulting in a response
pole leveling the gain at 1+ CD/CS. For large area diodes CD
can be hundreds of picofarads causing the noise gain to peak
in the hundreds as well. That gain continues to higher
frequencies until rolled off by the op amp bandwidth limit.
As feedback resistance increases, the pole and zero of this
gain peaking move together to lower frequencies encompassing a greater spectrum with high gain.
|A| (dB)
Open-Loop Gain
eO = IPR1 + InR1 +
A1: OPA111
1 + JωR1CD
1 + JωR1CS
I to V Gain
1/(2πR1 CS)
(CS/CD) fC
Noise Gain
1/(2πR1 CD)
Noise from the op amp also influences the output with a
surprising effect introduced by high feedback resistance and
the diode capacitance. The amplifier noise sources are modeled in Figure 2a as an input noise current, in, and the input
noise voltage, en. The current noise flows through the feedback resistor experiencing the same gain as the signal
current. It is the shot noise of the input bias current, IB, and
has a noise density of √2qIB1. Choice of an op amp having
input currents in the picoamp range makes this noise component negligible for practical levels of feedback resistance.
Input noise voltage of the amplifier would at first seem to be
transferred with low gain to the output. That is true at DC
where its gain 1 + R1/RD is kept small by the large diode
resistance, RD. Capacitance, CD, of the diode alters the
feedback at higher frequencies adding very significant gain
to en. As both the capacitance and the feedback resistance are
commonly large, the effect can begin at fairly low frequencies. Figure 2b illustrates the effect with an op amp gain
magnitude curve plotted with the reciprocal of the feedback
factor or the “noise gain.” That gain curve first experiences
1 + CD/CS
Log f
FIGURE 2a. Due to Diode Capacitance in the Feedback of
the Basic Current-to-Voltage Converter, 2b,
Op Amp Noise Receives Gain and Bandwidth
Not Available to the Signal.
First signs of this gain peaking phenomena are familiar to
anyone who has used high resistance op amp feedback in
more general circuits. High output to input resistance with
an op amp results in overshoot, response peaking, poor
settling or even oscillation all due to the resistance interaction with amplifier input capacitance. Together the resistance and capacitance form another pole in the feedback
loop resulting in the classic differentiator feedback response.
Shown by the dashed line for more general op amp cases, the
associated feedback factor reciprocal intercepts the amplifier open loop magnitude response with a 12dB/octave rate
of closure corresponding to feedback phase shift approaching or equal to 180°. The common cure for this condition is
a capacitor across the feedback resistor, which for the very
high resistances of current-to-voltage converters, automatically results from stray capacitance. Such capacitance degenerates the added feedback pole to control phase shift in
the feedback loop.
In understanding current-to-voltage converter noise performance it is important to note that the signal current and the
noise voltage encounter entirely different frequency responses. The current-to-voltage gain is flat with frequency
until the feedback impedance is rolled off by stray capacitance as shown. Gain received by the amplifier noise voltage, on the same graph, extends well beyond that roll-off and
is high in that extended region. The majority of the op amp’s
bandwidth often serves only to amplify that noise error and
not the signal. This is typically the dominant source of noise
for higher feedback resistances.
Relative effects of the major noise sources of a current-tovoltage converter can be seen with the curves of Figure 3.
Those curves show output noise for the basic current-tovoltage converter of Figure 1b including the effects of the
noise gain represented in Figure 2b. Plotted are total output
noises for three cases as a function of feedback resistance
and each is the rms sum of the components produced by the
feedback resistor and an op amp. Represented are three FET
op amps having different performance specialties that cover
the spectrum of photodiode applications with low noise,
low-input bias current and high speed. While all three types
have low-noise designs and low-input currents, the OPA111
offers the lowest noise in the FET op amp class at 6nV/√Hz,
and the OPA128 has the lowest input current at 0.075pA.
Without neglecting performance in these categories, the
OPA404 design pushes bandwidth to 6.4MHz. Noise due to
the op amp is found by integrating the amplifier noise
density spectral response over the noise gain response2. Also
shown, by a dashed line, is the noise due to the resistor alone
for the OPA111 and OPA2111 case. This resistor noise
curve is different for the other op amps as each amplifier has
a different bandwidth rolling off noise due to the resistor.
Different factors control the noise curves for different ranges
of feedback resistance. At low resistance levels, the noise
curves are largely flat with the op amp voltage noise the
dominant contributor. That domination makes initial resistance increases have little effect except for the case of the
very low-voltage noise of the OPA111/ OPA2111. In this
region noise gain peaking has not yet been encountered so
the output noise remains small. Between 10kΩ and 1MΩ,
resistor noise is dominant and the curves track that error
source as the dashed line shows for the OPA111/OPA2111.
Here, the curves demonstrate the square root relationship
with the resistance and differ only because of amplifier
bandwidths. At still higher resistance, noise gain peaking
takes effect returning the op amp noise to dominance and
boosting the curves higher. That effect is first demonstrated
by the increased slope of the OPA404 curve as that amplifier’s
wide bandwidth first encompasses the peaking. The noise
curves level off when essentially the full amplifier bandwidth is encompassed by the gain peaking. Moving to yet
higher resistance, resistor noise would return the curves to
rising slopes, but resistor bandwidth is by then rolled off by
stray capacitance. In this upper region, any increase in
resistance is accompanied by a matching reduction in noise
bandwidth so that the total resistor noise becomes a constant.
Variables of diode and stray capacitances alter the point of
onset of gain peaking errors, but the characteristic shape of
the output noise curves remains the same for any case. Each
will display ranges dominated by op amp noise, resistor
noise and gain peaking effects.
eno (µVrms)
Gain Peaking Range
Total Output Noise (dB)
(for OPA111 Bandwidth)
I to V Converter
CS = 0.5pF
D1: HP5082-4204
10M 100M 1G
Feedback Resistance (Ω)
FIGURE 3. As the Feedback Resistance of a Current-toVoltage Converter Increases, the Dominant
Noise Source Changes from the Op Amp to the
Resistor and Back to the Op Amp under Gain
Peaking Conditions.
Comparing the curves shows that the OPA111/OPA2111
provide the lowest noise in two of the characteristic ranges.
While the OPA128 shows a lower noise curve in the middle
range, that is due to the amplifier’s lower bandwidth and a
bandwidth reduction technique to be described, removes that
difference for the OPA111. Where the OPA128 excels is in
very low DC error as its input currents are a mere 0.075pA
which is 1/20th that of its low-noise contender. The third op
amp, OPA404, produces higher total output noise overall,
but that again is largely a bandwidth phenomenon. The
6.4MHz response of that amplifier accommodates noise
over a much greater frequency range. While the noise curve
for this amplifier is consistently higher than that of the
OPA128, the OPA404 actually has lower noise density but
it has six times the bandwidth. That 6.4MHz bandwidth is
available to signals for feedback resistances up to 50kΩ and
the amplifier still offers the best bandwidth for resistances
up to 150kΩ. As the OPA404 is a quad op amp, its economy
suggests consideration for use at even higher resistances
along with bandwidth reduction that provides more competitive output noise.
Only a five dimensional graph could display the output
noise, resistance, DC error, diode area and signal bandwidth
considered in current-to-voltage converter design. Each specific application’s requirements are evaluated separately
with respect to these factors. To avoid suboptimizing a given
design for one factor such as gain, the various effects of
increasing feedback resistance are anticipated at each step.
Choices such as large diode area are made considering the
related capacitance and its effect on output noise and overall
circuit sensitivity.
Gain peaking effects are the primary noise limitation with
the commonly preferred high feedback resistances. To limit
this effect, or to eliminate the gain rise entirely, additional
capacitance is commonly added to bypass the feedback
resistor. The capacitance level required can be very small for
some values of R1 and the relative significance of unpredictable stray capacitance make tuning desirable. Combined,
these requirements are a challenge better resolved with a
capacitor tee network as described in Figure 4a. It is capable
of even subpicofarad tunable capacitance with little effect on
stray capacitance in the tuning operation. The tee uses a
capacitive divider formed with C2 and C3 to attenuate the
signal applied to C1 at the circuit input. With only a fraction
of the output signal on C1, it supplies far less shunting
current to the input node as would a much smaller capacitor.
Controlling the attenuation ratio is the tunable C3, which is
the largest of the capacitors, so its capacitance value is more
readily available in tunable form. Since that capacitor is
grounded, it has a shielding advantage to reduce stray
capacitance influence while tuning.
Another option for practical feedback bypass exists with a
resistor tee which is a commonly considered replacement for
the high value feedback resistor. The latter is replaced in
Figure 4b by elements of more reasonable value but introduces greater low frequency noise. Its operation is the dual
of the capacitor tee above with R2 and R3 attenuating the
signal to R1 so that the latter appears as a much larger
resistor to the input node. A similar opportunity for the DC
error compensation resistor R2 does not exist. DC error due
to amplifier input current is no different with the tee so the
large compensation resistor is still needed.
Stray capacitance across the feedback is somewhat reduced
with the resistor tee by the added physical spacing of the
feedback with three elements. Also, stray capacitance across
each individual element has much less effect with their
lower resistances. Sensitivity to other stray capacitance from
the op amp output to its input has the same effect as before.
In the attenuation network of the feedback is the opportunity
for intentional bypass with reasonable capacitor values.
Bypassing the moderate resistance of R2 removes the attenuation at higher frequencies leaving the net feedback resistance at the level of R1. This operation differs from true
feedback bypass in that impedance levels off, rather than
continuing to fall with frequency, but the dramatic drop in
CEQ = C1C2/(C1+ C2 + C3)
= 0.5pF to 5pF
A1: OPA111
D1: HP5082-4204
REQ = R1 + R2 + (R1R2/R3)
= 1000MΩ
A1: OPA128
D1: HP5082-4204
FIGURE 4a. Removal of Amplifier Gain Peaking Through
Small Capacitive Bypass of Large Feedback
Resistance is More Feasible with a Capacitor
Tee; or, 4b, Bypass of One Element of a Feedback Resistor Tee.
equivalent resistance serves the circuit requirement. Another
benefit offered by the resistor tee is more accurate DC error
Reduced high frequency noise with the tee element bypass
is accompanied by an opposing increase at lower frequencies. Below the frequency of the bypass, noise gain is
increased by the feedback attenuation of the tee network.
That amplifies the noise and offset voltages of the op amp as
well as the noise of resistor R1 by a factor of 1 + R2/R3.
Countering the latter is the resistor’s smaller value so that
shown to continue its attenuating amplifier action well
beyond the unity gain crossover of A1. This avoids a second
gain peak that could cause oscillation. Signal bandwidth of
the current-to-voltage conversion is essentially unaffected as
R1 has not been influenced .
Where the Figure 5 technique is most useful is with lower
level signals that have greater sensitivity to noise. In higher
level applications that circuit can encounter a voltage swing
limitation but another use of the second amplifier offers
similar noise improvement. The swing limitation results
from the maximum output voltage limit of A1 and its
attenuation by A2. If the output of A1 has a peak swing of
12V and A2 has the gain of –1/10 illustrated, the final output
is limited to a 1.2V peak swing. For lower-level signals this
will be acceptable as the maximum practical level of feedback resistance already limits output swing.
Higher-level signals are not as sensitive to noise and better
tolerate a more straight forward approach to filtering. An
active filter following the conventional current-to-voltage
converter also removes the high frequency noise. Setting
filter poles at the frequency of the signal bandwidth results
this effect is increased only by the square root of the new
noise gain. Most important, however, is the bypass capacitor
removal of high frequency gain as it eliminates the greatest
portion of previous noise bandwidth. In the absence of other
means to remove the high frequencies, the bypassed resistor
tee provides lower total output noise for the higher ranges of
feedback resistance.
Adding feedback capacitance is an effective means of reducing noise gain but it also decreases signal bandwidth by the
same factor. That bandwidth is already low with high feedback resistance and the end result can be a response of a
kilohertz or less. A more desirable solution to the noise
problem is to limit amplifier bandwidth to a point just above
the unavoidable signal bandwidth limit. Then, the high
frequency gain which only amplifies noise is removed. Op
amps with provision for external phase compensation offer
this option, but those available lack the low-input currents
and low-voltage noise needed for photodiode monitoring.
To achieve this bandwidth limiting with better suited op
amps, a composite amplifier uses two op amps with the
added one for phase compensation control as in Figure 5a.
Note the reversal of the inverting and noninverting inputs of
A1 needed to retain a single phase inversion with two
amplifiers in series. With the composite structure, internal
feedback controls the frequency response of the gain added
by A2. At DC, that feedback is blocked by C1 and overall
open-loop gain is the product of those of the two amplifiers
or 225dB for those shown. That gain is rolled off by the
open-loop pole of A1 and by the integrator response established for A2 by C1 and R3. As this is a two pole roll-off, it
must be reduced before intercepting the noise gain curve to
establish frequency stability. A response zero does this due
to the inclusion of R4. Above the frequency of that zero, R4
also replaces the integrator response with that of an inverting
amplifier having a gain of –R4/R3. Making that gain less than
unity drops the net gain magnitude curve below that of a
single amplifier at high frequencies. Graphically, the noise
gain response of Figure 5b is moved back in frequency much
as if the op amp bandwidth had been reduced.
Eliminated is the shaded area of noise gain, which visually
may not appear dramatic, but that is because of the logarithmic-frequency scale. Actually, the associated noise reduction is large because most of the amplifier’s bandwidth is
represented in this upper end of the logarithmic-response
curve. Moving the unity gain crossover of the noise gain
from 2MHz to 200kHz, as shown, drops the output noise due
to A1 by about a factor of three. To achieve the same result
with feedback bypass, the signal bandwidth would have
been reduced a factor of ten. That bandwidth is unaffected
with the Figure 5a approach. No noise, or offset, is added by
A2 as this amplifier is preceded by the high gain of A1. With
the exceptionally low noise of the OPA111 input amplifier,
this improvement reduces noise to the fundamental limitation imposed by that of the feedback resistor. This condition
is retained for all practical levels of high feedback resistance. For the second amplifier, the wideband OPA404 is
A1: OPA111
A2: 1/4 OPA404
|A| (dB)
Modified O. L. Gain
Initial O. L. Gain
I to V Gain
Noise Gain
Log f
FIGURE 5a. Noise Reduction Results with a Composite
Amplifier that, 5b, Restricts Noise Bandwidth
Without Reducing that of the Signal.
amp unity gain bandwidth, fC. Interrelating the controlling
factors at the optimum bandwidth point is the expression
defining the choice of R1:
in a system bandwidth that does not extend beyond that of
useful information. Such a filter is not enclosed in a feedback loop with the converter so the input noise and offset
voltage of the second amplifier are added to the signal.
R1 = R τ / 2πC S f C
Signal bandwidth requirements are an integral part of the
current-to-voltage converter noise considerations for two
reasons. Total output noise increases in proportion to the
square root of system bandwidth simply because a broader
noise spectrum is encompassed. Added is conflict between
optimum signal-to-noise ratio and signal bandwidth. That
optimum occurs for very high gain but high gain current-tovoltage converters are bandwidth limited far below the rolloff of the op amp. To the signal current, the amplifier
feedback factor is unity which would normally make the full
amplifier unity gain bandwidth available. Yet the very highfeedback resistances that produce the desired gain are shunted
by stray capacitances at much lower frequency. Just 0.5pF
stray capacitance around a 100MΩ feedback resistor pulls
signal bandwidth from megahertz level unity gain crossovers down to 3.2kHz. To minimize the stray shunting, low
capacitance resistors and assembly precautions are used.
Mounting the feedback resistor on standoffs reduces capacitive coupling with printed circuit boards and such standoffs
are normally Teflon™ insulated to reduce leakage currents.
That mounting must be rigid to avoid introduction of noise
through the microphonic effects of mechanical stress from
There is an ultimate limit to the effects of such measures as
capacitive coupling through the air around the resistor body
always remains. Bandwidth beyond that imposed by such
residual limits requires lower feedback resistance and accompanying lower converter gain. To restore gain, several
options are available with a first shown in Figure 6a. A
second amplifier with voltage gain is simply added following the current-to-voltage converter to retain the net input to
output transimpedance for RT = AVR1. Then, the high-value
resistance is reduced by a factor equal to the voltage gain for
a bandwidth increase by as much as the same factor.
While an obvious alternative, its overall effect on bandwidth
and noise are not so immediate. Bounding the upper end of
the bandwidth increase is the response limitation of the
second amplifier. The bandwidth of the two op amp circuit
for a net transimpedance of 100MΩ is plotted in Figure 6b
as a function of the voltage gain involved in the overall
conversion. Bandwidth initially increases linearly with the
voltage gain as the reduction in R1 diminishes the roll-off
effect of stray capacitance. However, the added demands of
the voltage gain on A2 eventually make that amplifier’s
bandwidth the controlling factor. For a given set of conditions there is an optimum gain. AV produces the peak
bandwidths shown for the three example amplifiers. That
peak occurs when the amplifier closed loop bandwidth
equals the stray limited bandwidth of R1. Variables affecting
this peak are the net transimpedance, RT, and the second op
Bandwidth is extended to 100kHz from the original 3kHz
using the wideband OPA404 for the second amplifier. That
wideband op amp offers the best frequency response in
Figure 6 and, although its total output noise result is greater,
that is again largely due to the greater available bandwidth.
If even greater bandwidth is required, either a faster op amp,
with typically poorer noise performance, or lower
transimpedance are the choices. Less bandwidth demands
are encountered by A1 with its unity feedback factor so an
FET amplifier focused on low noise is used there like the
OPA111 shown.
The price paid for improved bandwidth through voltage gain
is increased output noise from that gain as well as from the
presence of the added amplifier. While the lower value of R1
does reduce its noise density, that effect is counteracted by
the increase in bandwidth for a net zero change is resistor
noise. That noise is now amplified by the voltage gain of the
eo = AVR1IP
AV = 1 + R4/R3
A1: OPA111
A2: 1/4 OPA404
D1: HP5082-4204
R T = AV R 1
= 100MΩ
1: OPA404
2: OPA128
3: OPA111
Output Noise (mV)
Bandwidth (Hz)
Voltage Gain (AV)
1: OPA404
2: OPA111
3: OPA128
FIGURE 6. For Greater Bandwidth and the Same Net
Transimpedance, (a) Voltage Gain is Added
Providing (b) Bandwidth That Increases Faster
Than Noise.
second amplifier causing an associated increase in output
noise proportional to the voltage gain. Added to that is the
noise from the op amps with the net result also shown in
Figure 6b. Those noise curves are continuations of the ones
presented in Figure 3 with the transition beginning at the
100MΩ level for the present example. In the lower gain
ranges from one to ten, the noise is first determined largely
by the op amps and their gain peaking but those effects give
way to resistor noise dominance before the end of this range.
Also in that range, the associated signal bandwidth plotted in
Figure 6b is controlled by stray capacitance and shows a
linear increase with increasing gain due to the corresponding
decrease in resistance. Above the gain of ten and before 100,
bandwidth begins to drop due to the encounter of A2 limits.
Simultaneous with this drop is a flattening of the output
noise curve. Roll-off of the amplifier bandwidth and the
simultaneous resistance drop nullify the effect of increasing
voltage gain leaving output noise a constant. In the voltage
gain range from 100 to 1000, these trends continue and
degrade optimum performance since bandwidth is lost while
noise remains constant.
While it is accepted that noise degrades with the voltage gain
replacement of resistance, the overall circuit figure of merit
gains. Including bandwidth in that measure shows that its
improvement more than offsets the drop in signal-to-noise
ratio. Mentioned before was the fact that the simple currentto-voltage converter suffers from greater bandwidth for the
amplifier voltage noise than for the signal current. That
discrepancy is removed with Figure 6 as the voltage gain
increases and A2 begins to filter out the higher frequencies.
Evidence of this is in the noise curves that increase more
gradually than the bandwidth curves—Figure 6b—up to the
optimum bandwidth point. At this optimum point, no bandwidth is afforded to noise that is not also available to signal.
In effect, A2 now also serves as the output active filter
discussed earlier. While each of these curves is drawn for a
specific 100MΩ transimpedance and the amplifiers and
photodiode specified, similar optimums are considered for
any design case.
For some of the more common photodiode applications, a
significant drawback of the above circuit is the need for two
op amps per photodetector. Often hundreds of detectors are
employed in a large arrays. As a compromise, one op amp
can be made to provide the same transimpedance, still
without the very large resistors, if some bandwidth and noise
degradations can be accepted. A single op amp can both
perform the current-to-voltage conversion and provide the
subsequent voltage gain. With traditional techniques the task
would be performed as in Figure 7a using R2 for the
conversion and R3 and R4 to set voltage gain. Current from
D1 flows in R2 resulting in a signal voltage at the input of a
noninverting amplifier. However, that signal voltage is also
across the photodiode and this condition produces a nonlinear response as described before.
Instead, the diode is connected directly between the op amp
inputs where zero diode voltage is maintained. Shown in
Figure 7b, the resistors perform the same functions as in the
last circuit but a linear response results. Current from the
photodiode still flows in R2 developing the same signal
voltage. That current also flows into the feedback network
but has little effect with the low resistances there. For the
resistor values shown, an equivalent transimpedance of
100MΩ results—just as with the two op amp example, but
bandwidth improvement is less. At 20kHz, it is increased a
factor of seven rather than a factor equal to the voltage gain
as in Figure 6a. A new bandwidth limitation accounts for the
difference and occurs due to the new placement of the highvalue resistance. That resistor is now shunted by the common-mode input capacitance of the op amp instead of just
the smaller stray capacitance. To maximize bandwidth, this
new shunting effect is made to coincide with the amplifier
roll-off through choice of R2 and the voltage gain. A second
benefit from this choice is that resistor noise beyond the
signal bandwidth encounters a two pole roll-off.
Final output noise from the resistor has the expected increase over the basic circuit by the square root of the voltage
gain. Added to that would have been a small component due
to the op amp as the normal source of gain peaking is
eO = –IPR2 (1+ R4/R3)
A1: OPA111
D1: HP5082-4204
eO = IPR2 (1+ R4/R3)
A1: OPA111
D1: HP5082-4204
FIGURE 7. Combining Current-to-Voltage Conversion and
Voltage Gain Using One Op Amp, (a) Impresses
Unwanted Voltage on the Diode that (b) is
Removed by Connecting the Diode Between the
Op Amp Inputs.
Once diminishing returns impose a limit on reduction of the
noise due to the circuit itself, consideration must be given to
external noise sources. With its very high resistance, a
current-to-voltage converter is extremely sensitive to noise
coupling from electrostatic, magnetic and radio frequency
sources. Those sources require attention to shielding, grounding, and component physical location(3), or they could otherwise become the dominant noise contributors. In each case,
physical separation of the noise source from the sensitive
circuitry is the most important step, but this becomes a
compromise warranting other measures as well.
Electrostatic coupling, such as from the power line, supplies
noise signals through the mutual capacitances that exist
between any two objects. Voltage differences between the
objects are impressed on those capacitances and any voltage
variation couples a noise current from one to the other. To
avoid that error signal, electrostatic shielding is used to
intercept the coupled current and shunt it to ground. In this
case, ground must be earth ground as that is the common
reference for the separate objects. Such shields, however,
create parasitic capacitances between the components shielded
and the shields must also be returned to the signal common
to avoid that coupling. Then shield carried capacitive currents from the output of a current-to-voltage converter are
also shunted to ground and represent no bandwidth restriction to the feedback resistor. Even still, the shield produces
a capacitance from the converter input to ground, possibly
adding to gain peaking and its effect on total output noise.
As electrostatic coupling is most often of power line frequency and common to all points, it is a natural candidate for
removal through the common-mode rejection of an op amp.
At the line frequency, op amp CMR is very high but it is not
utilized by the conventional current-to-voltage converter.
This is a result of single-ended rather than differential input
configurations, but that can be altered for improved noise
rejection and DC error benefits as well. Op amp CMR is not
a total replacement for shielding as electrostatic coupling
will not perfectly common-mode to amplifier inputs. As a
second defense, that rejection capability is most useful in
removing the residual coupling that passes through shield
removed. However, a new source is included in the circuit of
Figure 7b, again due to the diode capacitance, as modeled in
Figure 8a. Amplifier voltage noise, en, is impressed directly
across that capacitance developing a noise current that is
supplied to R2. That creates a noise voltage at the input of the
noninverting amplifier which is a multiple of en. The capacitive feedback network of CD and CICM produces a noise gain
that peaks at 1 + CD/CICM and which exists in addition to the
normal voltage gain of the noninverting amplifier.
Effects on frequency response are plotted in Figure 8b and
they again produce a high-frequency peak in the noise gain.
Its incidence is at a much higher frequency than with the
basic current-to-voltage converter because of the lower resistance involved and it is truncated earlier by the op amp
roll-off. For the low capacitance diode used in both example
circuits, it now encompasses little area in the response plot
corresponding to less noise effect. Larger diodes do not
escape the effect, however, as represented by the dashed line
for a capacitance around 200pF. Even still, the spectrum
covered by the peaking is not the high end of the op amp
bandwidth as it was for the basic circuit. Hence, op amp
noise does not become the overriding source.
|A| (dB)
Open-Loop Gain
A1: OPA111
AV = 1 + R4/R3
Noise Gain
The differential input capability of an op amp fits exactly
with the signal from a photodiode. Since the diode signal is
a current, it is available at both terminals of that sensor and
can drive both amplifier inputs as in Figure 9a. Here, the
diode current is no longer returned directly to common, but
drives the amplifier noninverting input in that path. That
creates a second signal voltage to double the circuit gain
when R2 = R1 for, compensation. For a given gain level, the
resistor value need be only one-half the normal for a similar
reduction in error sensitivity to amplifier input currents. This
also removes DC voltage from the diode as it is now directly
across the inputs of an op amp. With the voltage between
those inputs being essentially zero, photodiode leakage current is avoided.
I to V Gain
Log f
FIGURE 8. Photodiode Capacitance (a) Adds a Positive
Feedback Path to Figure 7b for (b) a New but
Lesser Source of Gain Peaking.
Aside from these benefits is the added improvement in the
common-mode rejection of coupled noise. Electrostatic coupling to this current-to-voltage converter is modeled in
Figure 9b along with the converter’s parasitic capacitances.
Zero signal is assumed there to illustrate only the electrostatic coupling effects. The electrostatic noise source, ee,
couples error currents, ie, through mutual capacitances, CM,
to the circuit’s two inputs. It might seem that the coupling
effects would be different to the two points because feedback makes the R1 input node a virtual zero impedance and
the other node is high impedance. Yet, the noise coupling is
via currents through capacitances that only depend on voltage signals on the capacitances. Both input nodes have the
same voltage due to amplifier feedback, and thus receive the
same level of noise current ie. Those equal currents develop
canceling ene noise voltage effects on the two circuit resistors
for a zero final output signal.
Accuracy of the error cancellation is determined by three
matching conditions involving the mutual capacitances, the
resistors and the parasitic capacitances shunting them.
Matched mutual capacitances are best assured by locating
the resistors equidistant from any significant noise source
not effectively blocked by a shield. Equal resistance values
assure accurate cancellation of error signals until frequencies are reached where capacitive shunting imbalances net
impedances. Shunting R1 will be only about 0.5pF of stray
capacitance but across R2 is the much larger common-mode
input capacitance of the op amp. For the 3pF of the OPA111
and the 50MΩ resistance shown, a pole occurs at about
1kHz, leaving the impedances of interest unbalanced. This
shunting by CICM also imposes a signal bandwidth limitation
at a lower frequency than normally encountered. The bandwidth of R2 is rolled off earlier than that of R1 creating a
response with two plateaus separated by a factor of two in
swing on common-mode input capacitances for improved
bandwidth in electrostatic suppression and signal gain. Note
that those noninverting inputs are not connected through
high resistances for input current error correction. That is not
necessary, as the input currents of A1 and A2 produce
matching voltages at their amplifier outputs. Those voltages
are a common-mode signal to the input of the INA105, so
they too are rejected.
eO = IP (R1 + R2)
A1: OPA111
D1: HP5082-4204
For the most common electrostatic coupling at power line
frequency, the above capacitive shunting has little effect. To
better reject higher frequencies, capacitance can be added
around R1 to restore impedance matching, or signal swing on
the common-mode input capacitance can be avoided. The
latter option offers a more accurate solution and avoids the
bandwidth limitation of CICM as well by using a second
differential connection. Shown in Figure 10, the photodiode
is connected between the inputs of two current-to-voltage
converters whose outputs drive an INA105 difference amplifier. Again the diode current flows in two equal resistances
that will receive equal electrostatic noise coupling. The
diode current creates a differential output on the resistances,
but the noise coupling generates a common-mode signal.
Supplied to the INA105, those signals are separated with the
diode signal passed to the output and the noise rejected.
Retained with the new differential input circuit are the 2:1
lower individual resistance and a zero diode voltage. The
latter is assured by the grounded noninverting inputs of both
current-to-voltage converters which establishes zero voltage
on both diode terminals. These connections also avoid signal
EOE = 0
FIGURE 9. Exploiting the CMR Capabilities of the Op Amp,
(a) the Differential Inputs are Driven Giving (b)
Rejection of Electrostatic Coupling.
Another function available with the differential structure of
Figure 10, is difference monitoring of two photodiodes.
Instead of D1, the two diodes shown in dashed lines are
connected separately to the two input current-to-voltage
converters. Their currents produce independent voltages at
the outputs of A1 and A2 where they are processed by the
difference amplifier to remove any common-mode portion.
Left is an output proportional to the difference between the
two input photocurrents as a measure of relative light intensity. A relative intensity measure is the type of signal used
in position sensing or optical tracking control to direct
feedback correction.
Magnetic coupling of noise can be more difficult to eliminate than the electrostatic, but its effects are also reduced by
the differential input connections. Coupling is through mutual inductances in this ease, so minimum sensitive loop area
is key to its control, along with shielding and maximum
separation of source and receiver. Its effects are not removed
by the electrostatic shield, so the first step is control of the
source itself.(3) Power transformers that cannot be placed at
a distance are internally shielded to largely terminate their
magnetic fields at the transformer boundaries. Remaining
magnetic coupling is addressed through physical and circuit
configurations. High value resistors used in photodiode
monitoring are sensitive to this coupling and connections
must be kept short between those resistors and high impedance op amp inputs. Coupling effects that remain are made
common-mode to be rejected by the op amp through loop
size and distance matching. In Figure 9 and Figure 10, the
high resistance is divided into two equal elements that are
then physically mounted with the same orientation to and
spacing from magnetic coupling sources. Noise coupled to
the two resistors then causes equal signals that have canceling effects at the circuit output.
may be close to the photodiode monitor because of digital
circuitry that is most likely co-resident in the system. Due to
the high frequencies involved, op amps have little gain or
common-mode rejection remaining for rejection of such
signals. Because of this same amplifier limitation, and the
basic voltage-to-current converter bandwidth restriction,
desired signals will not exist in the radio frequency range.
Filtering can then be used to largely remove the unwanted
signal if applied in front of the op amp. Later filtering is less
effective as the op amp can act like an RF detector separating a lower frequency envelope from a carrier.(4) Further
reduction of that noise is achieved with an RF shield and a
ground plane layer in printed circuit boards.
(1) Tobey, G., Graeme, J., Huelsman, L., Operational Amplifiers—Design
and Applications, McGraw-Hill, 1971.
(2) OPA101 product data sheet, PDS-434A, Burr-Brown Corp., 1980.
(3) Morrison, R., Grounding and Shielding Techniques in Instrumentation,
second edition, John Wiley & Sons, 1977.
(4) Sutu Y., and Whalen, J., Statistics for Demodulation RFI in Operational
Amplifiers, IEEE International Symposium on Electromagnetic Compatibility, August 23, 1983.
TeflonTM E.I. du Pont de Nemours & Co.
With the third class of noise coupling, radio frequency
interference, less can be removed by the amplifiers so
shielding and filtering are the best defenses. Sources of RFI
EO = 2 IPR1
EO = (Ip2 – Ip1)R1
A1, A2: 1/2 OPA2111
FIGURE 10. Differential Inputs with Wider Band CMR and Gain Result with Virtual Grounds Across Amplifier Common-Mode
Input Capacitances.
The information provided herein is believed to be reliable; however, BURR-BROWN assumes no responsibility for inaccuracies or omissions. BURR-BROWN assumes
no responsibility for the use of this information, and all use of such information shall be entirely at the user’s own risk. Prices and specifications are subject to change
without notice. No patent rights or licenses to any of the circuits described herein are implied or granted to any third party. BURR-BROWN does not authorize or warrant
any BURR-BROWN product for use in life support devices and/or systems.