. Pipelined ADC Design Tutorial

By: Imran Ahmed, Copyright 2004-2008



1.1: Overview

1.2: Multiplying Digital to Analog Converter (MDAC)

1.3: MDAC design considerations (matching, thermal noise, switch sizes)

1.4: Opamp design - gain requirement

1.5: Stage-ADC/Sub-ADC comparator design

1.6: Summary




1.1: Overview


his tutorial discusses circuit implementations and related design issues for 1.5 bit/stage pipeline ADCs.  The key sub-blocks discussed are: the stage MDAC, the stage ADC, and the stage amplifier. 


1.2: Multiplying Digital-to-Analog Converter (MDAC)

As pipeline stages operate on discrete time signals (since each stage has a sample and hold), switched capacitor circuits are used for pipeline ADCs.  With switch capacitor circuits it is possible to perform highly accurate mathematical operations such as addition, subtraction, and multiplication (by a constant), due to the availability of capacitors with a high degree of relative matching.  Switch capacitor circuits also facilitate multiple, simultaneous signal manipulations with relatively simple architectures.  It is possible to combine the functions of sample and hold, subtraction, DAC, and gain into a single switched capacitor circuit, referred to as the Multiplying Digital-to-Analog Converter (MDAC) as shown in Fig. 1.


Fig. 1: MDAC functionality in dashes


Fig. 2 shows a single ended circuit implementation of the MDAC of Fig. 1, using a switched capacitor approach.

Fig. 2: stage MDAC


The MDAC of Fig. 2 is shown single ended for simplicity, although in practice fully differential circuitry is commonly used to suppress common-mode noise [3].  A 1.5 bits/stage architecture has one of three digital outputs, thus the DAC has three operating modes:

ADC output = 01: No over range error (stage input is between –Vref/4 and Vref/4. 

During : QC1=C1Vin, QC2=C2Vin

During : C1 is discharged, thus by charge conservation: C1Vin + C2Vin = C2Vout (noting negative feedback forces node Vp to a virtual ground).  Thus

 è if C1=C2, then: Vout=2Vin                                                            (0.1)


ADC output = 10: Over range error – Input exceeds Vref/4, thus subtract Vref/2 from input

During : QC1=C1Vin, QC2=C2Vin

            During : C1 is charged to Vref, thus by charge conservation

C1Vin + C2Vin = C1Vref +C2Vout

 è if C1=C2, then: Vout=2Vin-Vref =2(Vin-Vref/2)          (0.2)


ADC output = 00: Under range error – Input below -Vref/4, thus add Vref/2 to input

During : QC1=C1Vin, QC2=C2Vin

            During : C1 is charged to -Vref, thus by charge conservation

C1Vin + C2Vin = C1(-Vref )+C2Vout

             è if C1=C2, then: Vout=2Vin+Vref =2(Vin+Vref/2) (0.3)


Thus the switched capacitor circuit implements the stage sample-and-hold, stage gain, DAC, and subtraction blocks.


Signal dependent charge injection is minimized by using bottom plate sampling, where the use of an advanced clock , makes charge injection signal independent [4].  A non-overlapping clock generator is thus required for the MDAC.




1.3: MDAC design considerations - Capacitor matching/linearity

From equations (3.1)-(3.3) it is clear stage gain is determined by the ratio of capacitors C1 and C2.  Thus to ensure a gain which is at least 10-bit accurate, C1 and C2 must match to at least 10-bit accuracy or within 0.1% for the first stage in the pipeline.  To obtain at least 0.1% matching a high quality capacitor such as a Metal-Insulator-Metal (MIM) capacitor must be used.  If properly designed in layout, MIM capacitors can achieve matching between 0.01-0.1% [5]. MIM capacitors however are often unavailable in purely digital processes, necessitating alternative capacitor structures.  Alternatively metal-finger capacitors, which derive their capacitance from the combination of area and fringe capacitance between overlapping metal layers can be used in digital processes to achieve sub 0.1% matching.  Metal-finger capacitors however can have large absolute variation (>20%), thus require a conservative design approach.  Alternatively a digital calibration algorithm can be employed to significantly minimize mismatch-induced gain errors (and finite opamp gain errors) [6], [7], [8], [9].  Due to additional design complexity, calibration schemes are beyond the focus of this dissertation.  We note however that calibration techniques are emerging as essential approaches for high-resolution pipeline ADCs due to the relaxed accuracy constraints afforded.


 In addition to capacitor matching, it is essential the ratio of capacitors C1 and C2 be linear for the desired input range to minimize harmonic distortion.  Thus non-linear parasitic gate capacitance (MOS-caps), or other active capacitors should be avoided for C1 and C2 in high precision pipeline ADCs.  Passive MIM, and metal-finger capacitors are linear well beyond the 10-bit level, thus are typically used.


The MDAC shown in Fig. 2 is a popular MDAC architecture, as the capacitor sizes of C1 and C2 are equal.  Since C1=C2, identical layouts can be used for C1 and C2 - maximizing layout symmetry and hence maximizing accuracy.  As MIM capacitors only have a marginal matching for 10-bit accuracy, a high degree of capacitor matching is essential to minimize INL/DNL errors.  Another advantage of the architecture of Fig. 2 is a high beta value (feedback factor), which maximizes the bandwidth of the closed loop system [10].

1.3.2: MDAC design considerations - Thermal noise

Although capacitors are ideally noiseless elements, in a sampled system, sample and hold capacitors capture noise generated by noisy elements such as switch resistors, opamps, etc.  Consider the following noise analysis of a capacitor sampling resistor noise as shown in Fig. 3:

Fig. 3: RC noise model


from [1] it is shown equivalent noise bandwidth is ,


                                                è                                        (0.4)

From the above example it is clear increasing the size of the sampling capacitor reduces the power of thermal noise.  As thermal noise represents a dynamic noise source that reduces ADC SNR, a minimum capacitance (i.e. C1, C2) must be driven to ensure a sufficient accuracy – thus thermal noise imposes a tradeoff between power and accuracy. For the MDAC of Fig. 2, the effective input referred thermal noise, which includes switch, and opamp noise is derived in [11] and found to be


where is the equivalent output load capacitance, and Copamp the input capacitance to the opamp.  The relationship between SNR and minimum capacitor size for a full scale signal swing of 0.8V, and C1=C2=Copamp=0.5pF is shown in Fig. 4.


Fig. 4: Variation of SNR due to thermal noise (ignoring quantization error, full scale=0.8V, C1=C2=Copamp=0.5pF)


From Fig. 4 it is clear thermal noise can alone limit accuracy to less than 10-bits (SNR=62dB) if capacitors are not sufficiently sized.  As thermal noise represents only one of several precision limiting factors (others include: quantization noise, power supply noise, capacitor mismatch, etc.), it is desirable to place the noise floor beyond the 10-bit level (e.g.) for thermal noise less than 1/4 LSB è thermal noise floor should be at least -72dB.  The stage accuracy requirements are relaxed for subsequent pipeline stages.  Thus it is possible to increase the noise floor for subsequent stages by using smaller capacitors - maximizing opamp bandwidth and minimizing overall power.


1.3.3: MDAC design considerations - Switch sizing

When sizing a MOS switch two key issues should be considered: 1.) The desired RC time constant, and 2.) The maximum distortion tolerable through the switch. 


As switched-capacitor circuits have a finite time to settle, it is essential the switches be sized large enough such that the sampled signal settle to the desired accuracy in the allotted time.  Since , switch resistance can be minimized by increasing the MOS switch W/L ratio. However an increased W/L ratio implies a larger area, which imparts a larger parasitic capacitance to the circuit.  As described in [1], a sufficiently large parasitic capacitance can alter charge-sharing equations, and introduce harmonic distortion through charge injection.  Thus switch transistors must be carefully sized, where switches should be large enough to ensure a sufficient RC time constant, but small enough to minimize parasitic induced errors.


A consequence of the switch’s resistance dependency on Veff is an RC time constant that is signal dependent, hence non-linear.  A non-linear RC time constant can lead to significant distortion if the switch passes a continuous time signal, as is the case in front-end sample and hold inputs.  Signal–dependent RC time constants also affect discrete time signals, as the MOS switch must be sized sufficiently such that the worst-case RC time constant (i.e. when Veff is smallest) is sufficient for the desired sampling speed.  Non-linear RC time constants can be significantly minimized however using a bootstrapping approach [4], which maintains a constant and maximal Veff, thereby minimizing signal dependent variations.


1.4: Opamp design - Gain requirement

The charge transfer relations derived in equations (3.1)–(3.3) were based on the assumption of a perfect virtual ground at node Vp in Fig. 2, which only occurs when the opamp gain is infinite.  In practice opamp gain is finite - introducing an error into the charge balance equations. As such opamp gain must be made sufficiently large to minimize finite gain error.


Consider the closed loop gain of a negative feedback system H(s), as shown in Fig. 5:


Fig. 5: basic linear feedback structure

Ideally as A(s) tends to infinity, H(s) è 1/b. Thus the relative error () is


As switch capacitor circuits settle to DC values, DC gain affects charge transfer equations:


Hence for an error due to finite opamp gain to be less than ¼ LSB, i.e. 1/(4x1024)=1/(4096), with b=0.5 implies A > 8192, or A >78dB.  Fig. 6 illustrates the variation of relative error with opamp gain.


Fig. 6: gain error variation with opamp gain

Attaining 78dB of DC gain while maintaining a reasonable bandwidth is near impossible with a simple single stage configuration (e.g. differential pair) for sub-micron technologies.  Thus two-stage or gain-boosted configurations are necessitated for 10-bit pipeline ADCs (a detailed description of high gain opamps is given in [1], [12]).  It is noted that stage accuracy requirements decrease along the pipeline, thus latter stages may have less gain, allowing for simpler opamps (single stage, or no gain-boosting), thus reducing power.

It should be noted that alternative MDAC architectures exist which employ gain-error cancellation methods, facilitating much lower opamp gains [6], [7], [8], [9] than those required by (3.8).  Such approaches however introduce a design overhead, and increase design time, thus are not considered in this dissertation.


1.4.2: Opamp design - Bandwidth requirement

Switched capacitor circuits have a finite time in which to settle, thus to ensure a minimum settling accuracy, opamp bandwidth must be optimized.  If the opamp is modeled as a first order system, the opamp transfer function near the unity gain frequency is given by: [1]. Thus the MDAC step response, during  is given by


where , and slew rate is ignored.  Since, where x is the settling accuracy in bits, the available time to settle is


As the available time t to settle is half the clock period,

 , (0.11)



where for settling within ¼ LSB,  for a 10-bit ADC.  Figure Fig. 7 graphically illustrates the required opamp unity gain bandwidth to achieve a desired sampling rate and settling accuracy.


Fig. 7: required opamp unity gain frequency versus sampling frequency and settling accuracy


From Fig. 7 and equations (3.11)-(3.12), a unity gain frequency much larger than sampling frequency is required to obtain high accuracy settling. Since the MDAC opamps must drive large capacitive loads (to minimize thermal noise), much power is consumed by the opamps.  As such, the power consumption of opamps in a pipeline ADC often consumes 60-80% of the total ADC power.  However, the accuracy requirements decrease along the pipeline, thus the unity gain frequency of subsequent stages along the pipeline can be reduced, minimizing total power [2].



1.5: Stage ADC design - Comparator

A flash architecture is commonly used for the stage ADCs, due to low accuracy required by the stage ADCs.  Flash ADCs consist of comparators at the various thresholds of the ADC.  For a 1.5-bit/stage pipeline architecture stage flash ADCs require comparators at thresholds of +/-Vref/4 and 0.  Digital error correction could be used to relax the tolerable offset on stage-ADC comparators (up to +/-Vref/4).  For Vref=0.8V, the comparator offset can be as high as 200mV, which allows for minimum size devices in the comparator (hence minimizing parasitic capacitance, thus minimizing power). The relaxed offset constrains also afford simpler dynamic comparator architectures, which do not require pre-amp gain stages, or static comparators (e.g.: as used in. 6-bit flash ADCs [13], [14]).  Like digital logic, dynamic comparators only consume power on clock edges according to fCV2 thus have a power that scales linearly with sampling frequency. For pipeline ADCs one of two dynamic comparators are typically used [15]: the Lewis and Gray comparator [16] (Fig. 8), or the charge-distribution comparator (Fig. 9).


Fig. 8: Lewis and Grey comparator


Fig. 9: switched capacitor/charge distribution comparator


The Lewis and Gray comparator compares two fully differential signals , and  (Fully differential comparators are highly desirable to reduce common-mode noise which can be large in digital environments).  Comparators at Vref/4 and –Vref/4 are required to implement the 1.5bit/stage architecture, and comparators at Vref/2, and –Vref/2 for the 2-bit flash at the end of the pipeline.  Rather than supply multiple reference voltages for each unique threshold, it is possible using the architecture of Fig. 8 to derive an arbitrary threshold by appropriate device sizing.  Transistors M1-M4 operate in triode while the remaining transistors implement positive feedback to resolve the differential input [11].  The equivalent triode conductance of M1 and M2 from Fig. 8 are:



The comparator threshold occurs when the circuit is perfectly symmetric, i.e. when G1=G2, thus if W1=W4, and W2=W3


where Vin = Vin+  - Vin-, and Vref = Vref+  - Vref-


Thus it is possible to achieve thresholds at ±Vref/4, and ±Vref/2 by providing a common differential reference voltage to each comparator in the pipeline, but sizing each comparator to yield the desired threshold (e.g.: W2 = 4W1 for a threshold of Vref/4, W2 = 2W1 for a threshold of Vref/2, etc.).  As the comparator is fully differential, thresholds at –Vref/4 and –Vref/2 can be realized by reversing the polarity to the reference voltage.  Thus all required thresholds for a 1.5 bit/stage pipeline can be realized by only supplying only one fully differential reference potential to the chip.


A drawback of the Lewis and Gray comparator is the threshold is a significant function of device symmetry.  As the value resolved by the comparator operates by comparing the integral of the ratio of current to node capacitance at nodes V1 and V2, circuit symmetry is crucial to reduce offset.  Thus the layout of the Lewis and Gray comparator requires great care, and parasitic extraction for full characterization of input-referred offset.  In [15] the Lewis and Gray comparator is shown to have an offset of >200mV for a 0.35mm CMOS process,


Alternatively a charge distribution approach can be used to achieve a lower offset at the cost of increased power.  As shown in Fig. 9, the charge distribution approach uses charge conservation to derive a comparator threshold, which depends on the ratio of capacitors rather than the ratio of device widths and parasitic capacitances.  Using a two-phase clock (,), capacitors Cin and Cref are charged to  and  respectively (in a differential sense) on the first clock phase.  The charge is forced to redistribute between both capacitors during the second clock phase, where according to charge conservation the effective threshold of the comparator is found to be [15]


As the threshold is primarily a function of passive components and largely independent of parasitic capacitance, a lower offset can be achieved using the charge-distribution comparator.  An analysis in [15] compares fabricated implementations (in 0.35mm CMOS) of the Lewis and Gray, and charge distribution comparators, where the following silicon measured results were obtained:


Table 0‑1: Comparison of comparator area, offset, and power



Power @ 100Msps


Lewis and Grey




Charge distribution





As other offsets besides device mismatch (e.g. noise) affect the stage transfer function, it is desirable to keep comparator offsets below Vref/4.  It should be noted the reduced offset of the charge distribution comparator comes at the cost of increased power (due to the dynamic charging of the sampling capacitors, and switches) and area.  Thus the choice of which comparator architecture to use requires a tradeoff between tolerable offset, desired power consumption and area.


1.6: Summary

In this chapter circuit level implementation and design related issued were discussed for key components in a 1.5 bit/stage pipeline ADC: the stage MDAC and stage ADC comparators.  It was shown for a desired settling accuracy, MDAC opamps require a minimum gain and unity gain bandwidth.  Noise limitations due to thermal and opamp noise were shown limit minimum MDAC sampling and feedback capacitor sizes.  Two popular dynamic comparators were examined: the Lewis and Gray comparator, and the charge distribution comparator, where it was shown the optimal comparator was a tradeoff between power and input referred offset. 

. References


[1]               Johns, David and Martin, Ken. Analog Integrated Circuit Design. John Wiley & Sons, Inc: New York, 1997.

[2]               P.T.F. Kwok et al, “Power Optimization for Pipeline Analog-to-Digital Converters”, IEEE Transactions on Circuits and Systems--II: Analog and Digital Signal Processing, vol 36, May 1999, pp. 549-553

[3]               Y. Park et al, “A low power 10 bit, 80MS/s CMOS pipelined ADC at 1.8V power supply”, 2001 IEEEE International Symposium on Circuits and Systems (ISCAS), vol 1, pp. 580-583

[4]               A. Abo, “Design for Reliability of Low-voltage, Switched-capacitor Circuits”, Doctor of Philosophy in Electrical Engineering, University of California Berkeley, 1999

[5]               C. Diaz et al, “CMOS Technology for MS/RF SoC”, IEEE Transactions on Electron Devices, vol 50, March 2003, pp. 557-566

[6]               J. Li et al, “Background Calibration Techniques for Multistage Pipelined ADCs With Digital Redundancy”, IEEE Transactions on Circuits and Systems – II: Analog and Digital Signal Processing, vol 50, September 2003, pp. 531-538

[7]               Y. Chiu et al, “Least Mean Square Adaptive Digital Background Calibration of Pipelined Analog-to-Digital Converters”, IEEE Transactions on Circuits and Systems – I: Regular Papers, vol 51, Janurary 2004, pp. 38-46

[8]               S. Chuang et al, “A Digitally Self-Calibrating 14-bit 10-MHz CMOS Pipelined A/D Converter”, IEEE Journal of Solid-State Circuits, vol 37, June 2002, pp. 674-683

[9]               B. Murmann et al, “A 12-bit 75-MS/s Pipelined ADC Using Open-Loop Residue Amplification”, IEEE Journal of Solid-State Circuits, vol 38, December 2003, pp. 2040-2050

[10]           W. Yang et al, “A 3-V 340-mW 14-b 75 Msample/s CMOS ADC with 85dB SFDR at Nyquist Input”, IEEE Journal of Solid State Circuits, Brief Paper, vol 36, December 2001, pp. 1931-1936

[11]           T. Cho, “Low power Low voltage A/D conversion techniques using pipelined architecture”, Doctor of Philosophy in Engineering, University of California Berkeley, 1995

[12]           Razavi, Behzad.  Design of Analog CMOS Integrated CircuitsMcGraw-Hill, New York,  2000

[13]           Uyttenhove et al, “A 1.8-V 6-bit 1.3-GHz flash ADC in 0.25mm CMOS”, IEEE Journal of Solid-State Circuits, vol 28, July 2003, pp. 1115-1122

[14]           M. Choi et al, “A 6-b 1.3-Gsample/s A/D converter in 0.35-mm CMOS”, IEEE Journal of Solid-State Circuits, vol 36, December 2001, pp. 1847-1858

[15]           L. Sumanen et al, “CMOS dynamic comparators for pipeline A/D converters”, 2002 IEEE International Symposium on Circuits and Systems (ISCAS), vol 5, 2002, pp. 157-160

[16]           L. Sumanen et al, “A mismatch insensitive CMOS dynamic comparator for pipeline A/D converters”, 2000 International Conference on Electronics, Circuits and Systems (ICECS), pp. 32-35