2578 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 24, NO. 11, NOVEMBER 2009 An Autotuning Digital Controller for DC–DC Power Converters Based on Online Frequency-Response Measurement Mariko Shirazi, Student Member, IEEE, Regan Zane, Senior Member, IEEE, and Dragan Maksimovic, Senior Member, IEEE Abstract—This paper describes a hardware-descriptionlanguage-coded autotuning algorithm for digital PID-controlled dc–dc power converters based on online frequency-response measurement. The algorithm determines the PID controller parameters required to maximize the closed-loop bandwidth of the feedback control system while maintaining user-specified stability margins and integral-based no-limit-cycling criteria, as well as ensuring single-crossover-frequency operation and sufficiently high loop gain magnitude at low frequencies. Experimental results are provided for five different pulsewidth-modulated dc–dc converters, including a well-damped synchronous buck, a lightly damped synchronous buck with and without a poorly damped input filter, a boost operating in continuous-conduction mode, and a boost operating in discontinuous-conduction mode. Index Terms—DC–DC power conversion, digital control, frequency response, identification, pulsewidth-modulated (PWM) power converters, switched-mode power supplies (SMPS), tuning. I. INTRODUCTION ONVENTIONAL offline design of controllers for dc–dc power converters is complicated by unknown load characteristics, as well as uncertainties within the converter itself. The converter component values are subject to manufacturing tolerances, and converter parasitics are notoriously difficult to model. In light of these uncertainties, the designer must make some assumptions regarding the expected range of load and core converter dynamics, and design a controller that will maintain acceptable stability margins under worst-case conditions. A robust design capable of handling wide variations in dynamics will be overly conservative by design over most of the expected range, resulting in degraded performance. The benefit of an autotuning controller is the ability to perform online control design in the presence of actual system dynamics, resulting in a more optimal design over the full range of system characteristics. The ability to embed such algorithms into the existing feedback controller represents a significant advantage of digital controllers for switched-mode power supplies (SMPS) over their analog counterparts. The terminology adopted in this paper is consistent with that used in [1] and [2], namely the term automatic tuning, or au- C Manuscript received March 7, 2009; revised May 21, 2009. Current version published December 18, 2009. Recommended for publication by Associate Editor C. K. Kong Tse. The authors are with the Colorado Power Electronics Center, University of Colorado, Boulder, CO 80309 USA (e-mail: [email protected]; [email protected]; [email protected]). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TPEL.2009.2029691 totuning, refers to a tuning process that executes, upon startup, event detection, regularly scheduled interval, or external command. The resulting controller parameters are then held constant until the next time the process is run. In contrast, the term adaptation, or adaptive control, refers to the process of continuously updating the controller parameters in a feedback loop to accommodate changing system dynamics or external disturbances. The same adaptive methods can be used in an autotuning context if enabled only at discrete intervals; however, there also exist dedicated open-loop autotuning methods. The distinction between the two is pointed out here in order to facilitate a discussion of the recent applications of these techniques to SMPS and to motivate the autotuning work described in this paper. Adaptive control techniques based on model reference adaptive control (MRAC) [1] have been successfully applied to PIDcontrolled continuous-conduction mode (CCM) buck converters in [3]–[5], and CCM, as well as discontinuous-conduction mode (DCM) buck converters, in [6]. In these works, small oscillations are injected into the duty cycle command with the converter operating in closed loop, and the PID controller parameters are adjusted in an adaptive feedback loop to achieve crossover frequency and phase margin specifications. Adaptive control using the self-tuning regulator (STR) concept [1], [7], [8] has been used to tune predictive controllers for a CCM buck converter in [9] and a phase-controlled rectifier in [10]. The STR concept has also been applied in an autotuning context to deadbeatcontrolled CCM buck converters in [11]. A challenge associated with any adaptive control scheme is the selection of the parameter update algorithms that ensure stability of the adaptive loop, and therefore, convergence of controller parameters. As the loop dynamics depend on the openloop plant being controlled, it is clear that sufficient a priori knowledge of the expected range of plant dynamics is required to select parameter update rates and dynamics that ensure stability over the entire range. Particular care must be taken with nonminimum phase (NMP) systems. Another question specific to MRAC techniques is how the loop will behave if, for the given plant dynamics and compensator structure, there is no solution in terms of compensator parameters that will cause the closedloop system to have the desired dynamics. In contrast, open-loop autotuning techniques can be applied that require little a priori knowledge of the plant being controlled. Open-loop autotuning methods based on inducing limit-cycle oscillations (LCOs) have been successfully applied to PIDcontrolled CCM buck converters in [3], and [12]–[14], as well 0885-8993/$26.00 © 2009 IEEE Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. SHIRAZI et al.: AUTOTUNING DIGITAL CONTROLLER FOR DC–DC POWER CONVERTERS Fig. 1. Digitally controlled PWM converter with integrated frequencyresponse measurement and autotuning capabilities. as PID-controlled CCM boost converters in [12] and [15]. The amplitude and the frequency of the LCOs provide information regarding the loop frequency response, and compensator parameters can be tuned to achieve specified crossover frequency and phase margin. One limitation common to all LCObased autotuning techniques, as well as the single-frequency injection MRAC schemes of [4]–[6], is that the knowledge of the loop frequency response can be obtained only at the frequencies at which the system is excited. In particular, none of these methods are able to measure or take into account gain margin specifications. In addition, multiple crossover frequencies and undesirable flattening of the loop gain magnitude at low values before the crossover frequency can neither be detected nor avoided. The MRAC approaches are further unable to address no-limit-cycling considerations, while in the LCObased approaches such considerations can be incorporated only as a final check on the controller design. Autotuning methods based on online identification of the converter control-tooutput frequency response can overcome these limitations Such methods have been successfully applied to CCM buck converters in [14], [16], and [17]. In [16], the frequency response is parametrically identified, and then, a compensator is designed offline using model inversion techniques. The nonparametric frequency response is used in [17] to design a PID compensator offline using time- and frequency-domain simulations iterated within an optimization framework to meet multiple design criteria. PID compensator parameters are iterated and loop frequency response data are computed online in [14] to maximize crossover frequency subject to phase margin and integral-based no-limit-cycling constraints. The autotuning approach presented in this paper is an extension of the identification-based autotuning work presented in [14]. An overall block diagram of the system is depicted in Fig. 1. As described in [18] and [19], the system-identification algorithm injects a pseudorandom binary sequence (PRBS) perturbation, stores the resulting output voltage perturbations, and computes and stores the converter open-loop frequency response. In Fig. 1, the input stimulus dstim [n] is injected into 2579 the digital pulsewidth modulator (DPWM) on top of the compensator output dcomp [n], which is frozen at its steady-state value during identification. With the converter open-loop frequency response stored in memory, the autotuning algorithm can construct the entire loop frequency response for any arbitrary compensator structure and parameters. For example, the PID parameters can be iterated computationally to force the loop frequency response to meet specifications at arbitrary frequencies, without the need for further system perturbations. The final PID parameters are then exported to the programmable digital PID controller. Section II of this paper gives an overview of the autotuning procedure. Experimental results, including the autotuning controller hardware implementation, are presented in Section III. In order to demonstrate the versatility of this autotuning approach, results are provided for five different PWM dc–dc converters, including a well-damped synchronous buck, a lightly damped synchronous buck with and without a poorly damped input filter, a boost operating in CCM, and a boost operating in DCM. The conclusions, including a summary of the relative merits of the autotuning method presented here with respect to MRAC-based techniques, are given in Section IV. II. AUTOTUNING ALGORITHM The digital PID control law of Fig. 1 was obtained by application of the backward rectangular version of Euler’s method to the continuous-time PID differential equation. The Z-transform of this control law can be expressed in either parallel or cascade form as follows: KI + KD 1 − z −1 Gc (z) = KP + 1 − z −1 1 − z1 z −1 1 − z2 z −1 . (1) = Kcomp (1 − z −1 ) The three independent PID parameters have different interpretations depending on whether the PID is expressed in parallel or cascade form. In the parallel form, the parameters KP , KI , and KD represent the proportional, integral, and derivative gains, respectively. In the cascade form, the parameters are expressed as the gain Kcomp , and the discrete-time zero locations z1 = e−2π f z 1 T s and z2 = e−2π f z 2 T s , where Fs = 1/Ts is the sampling frequency, which is also equal to the converter switching frequency. Since the autotuning method implemented here is a loop-shaping method, it is most convenient to work with the cascade form and iterate zero locations. For any given converter operating with a PID feedback controller, the locations of the PID zeros determine the phase margin of the closed-loop system. However, the solution is not unique. In this paper, an a priori decision is made to place fz 1 , the frequency of the first PID zero, at the frequency f−90 ◦ , where the phase of the measured open-loop frequency response Hv Gv d (ej ω T s ) first drops below −90◦ . The variables Gv d (ej ω T s ) and Hv are the converter open-loop frequency response and the sensing gain, respectively, as shown in Fig. 1. The frequency of the second PID zero fz 2 then uniquely determines the phase margin. The goal of the autotuning algorithm is to maximize the closed-loop bandwidth of the feedback control system, as indicated by the Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. 2580 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 24, NO. 11, NOVEMBER 2009 Fig. 2. Analytical Bode plots for an exemplary tuning case. (a) Converter control-to-output transfer function G v d (z). (b) PID controller G c (z). (c) Resulting loop transfer function T (z). Square: f−9 0 ◦ ; circle: fc ; upward pointing triangle: fz 1 ; downward pointing triangle: fz 2 ; PM: phase margin; GM: gain margin. loop-transfer-function crossover frequency fc , while maintaining user-specified stability margins and integral-based no-limitcycling criteria, as well as ensuring single-crossover-frequency operation and sufficiently high loop gain magnitude at low frequencies. To this end, the algorithm iteratively reduces the target crossover frequency fc,targ until it reaches one for which PID parameters exist, which ensure that the resulting closed-loop system meets specifications. The target crossover frequency is initialized at Fs /8. Then, the basic autotuning procedure is as follows: the first PID zero is placed at f−90 ◦ , constraints are placed on the second PID zero to ensure that the integral gain remains sufficiently low to satisfy the integral-based no-limitcycling criteria [20], [21], and then, the second PID zero is iterated over its allowable range to attempt to provide the required phase lead at fc,targ to meet the phase margin specification. If the phase margin specification cannot be met, the target crossover frequency is reduced and the procedure repeated. If a suitable zero is found, the compensator gain Kcomp is computed to ensure that the actual crossover frequency occurs at fc,targ . Finally, the complete loop frequency response is constructed. A check is then run to measure the gain margin, and detect multiple crossover frequencies and insufficient loop gain magnitude at low frequencies. If the conditions are not met, the procedure is repeated using a reduced target crossover frequency. Fig. 2 shows relevant tuning variables graphically displayed on analytically derived Bode plots of the converter control-to-output transfer function Gv d (z), PID controller Gc (z), and the resulting loop transfer function T (z) of an exemplary tuning case. Details of the autotuning procedure, which, in many aspects, resembles the approach a designer would follow, are discussed in the following sections. A. Allowance for Integral-Only Control If fc,targ iterates down to a value less than f−90 ◦ , where no phase lead is required to meet the phase margin specification, then the controller reverts to an integral-only controller, as the inclusion of a zero at f−90 ◦ in this case would only reduce the loop gain roll-off. The integral-only test is expressed as Gv d ej 2π f c , targ T s ≥ ϕm − 90◦ (2) Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. SHIRAZI et al.: AUTOTUNING DIGITAL CONTROLLER FOR DC–DC POWER CONVERTERS 2581 TABLE I INTEGRAL-BASED NO-LIMIT-CYCLING CONSTRAINTS TABLE II HARDWARE REQUIREMENTS FOR SYSTEM-IDENTIFICATION AND AUTOTUNING ALGORITHMS where ϕm is the desired phase margin. In this case, both z1 and z2 are set to zero, and the algorithm skips to computation of Kcomp . B. Embedding Integral-Based No-Limit-Cycling Criteria Early in the Design Process Under certain conditions, due to quantization effects, a digitally controlled converter can exhibit generally undesirable limit cycling [20], [21]. To avoid limit cycling, it has been shown in [20] and [21] that the controller integral gain Ki must not be too high, i.e., it should satisfy Fig. 3. CCM buck converter hardware prototype. Fig. 4. CCM/DCM boost converter hardware prototype. Ki Hv Gv d0 < α (3) j 0 where α < 1 is a safety factor and Gv d0 = Gv d e is the dc gain of the control-to-output transfer function. It is possible to design a compensator, and then check to see whether the criterion (3) is satisfied. Instead, the tuning algorithm presented here embeds this criterion much earlier in the design process by using it to constrain the target crossover frequency or the PID zero locations. The relevant constraint depends on whether fz 2 is less than or greater than fc,targ . For cases where fz 2 < fc,targ , these constraints are in the form of a maximum allowed frequency for fz 2 . For cases where fz 2 ≥ fc,targ , these constraints are in the form of a minimum allowed magnitude of Hv Gv d ej 2π f c , targ T s . In both cases, the location of fz 1 relative to fc,targ dictates the expression to use according to the results shown in Table I. A derivation of these expressions is given in the Appendix. In implementing these constraints, since the location of fz 1 has been fixed, the current fc,targ immediately dictates whether case 1 or 2 holds. In order to determine whether the A or B case holds, fz 2 is initially placed at fc,targ , and the loop frequency response at fc,targ is constructed using the measured frequency response, and by computing the real and imaginary parts of a Gc with the specified zeros, and setting Kcomp = 1 for simplicity. The loop data can then be checked to see whether the phase margin is low or high with fz 2 = fc,targ . If the phase margin is high, or within the specified range, then fz 2 must be placed at or above fc,targ , and case B1 or B2 holds. If the phase margin is low, then fz 2 must be placed below fc,targ , and case A1 or A2 holds. Details of how fz 2 is iterated are described next. C. Iterating the Second PID Zero Location If case B1 or B2 holds, thenthe relevant expressionin Table I is evaluated and compared to Hv Gv d ej 2π f c , targ T s in order to determine whether the integral-based no-limit-cycling criteria Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. 2582 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 24, NO. 11, NOVEMBER 2009 Fig. 5. Experimental loop frequency responses. (a) Buck #1. (b) Buck #2. (c) Buck #3. (d) Boost #4. (e) Boost #2. In all plots, the experimental loop frequency response data (gray dots) are plotted against the loop transfer function obtained using the autotuned PID parameters and a discrete model of the converter (black line). Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. SHIRAZI et al.: AUTOTUNING DIGITAL CONTROLLER FOR DC–DC POWER CONVERTERS 2583 TABLE III COMPONENT VALUES TABLE IV DISCRETE-TIME CONTROL-TO-OUTPUT TRANSFER FUNCTIONS D. Computing Kcomp , and Checking Gain Margin and Other Loop Gain Constraints Once the PID zeros have been specified, the gain Kcom p is computed as Kcomp = 1 . Hv Gv d (ej 2π f c , targ T s )Gc,K comp =1 (ej 2π f c , targ T s ) (4) Finally, the entire loop frequency response can be constructed as T ej 2π f T s = Hv Gv d ej 2π f T s Gc ej 2π f T s . can be met at this fc,targ with fz 2 adjusted to meet the phase margin specification. If the no-limit-cycling criteria cannot be met, then fc,targ is reduced and the fc,targ loop reiterated. If the no-limit-cycling criteria can be met, then the maximum allowed fz 2 is initialized to Fs /2, the minimum allowed fz 2 initialized to fc,targ , and fz 2 itself initialized to the maximum value. The phase margin is checked at each value of fz 2 , and fz 2 is adjusted as necessary to increase or decrease the phase margin. If the phase margin matches the specification, or if the phase margin is high, but fz 2 is saturated at its maximum allowed value, then the algorithm proceeds to compute Kcomp . If the phase margin is low or high, but fz 2 has not saturated, then fz 2 is iterated using the bisection method. If the phase margin is high, then the minimum allowed fz 2 is set equal to the current fz 2 , and if the phase margin is low, then the maximum allowed fz 2 is set equal to the current fz 2 . The next fz 2 is computed as fz 2 ,next = (fz 2 ,max + fz 2 ,min )/2. If case A1 or A2 holds, the relevant expression in Table I is evaluated and the maximum allowed fz 2 is initialized to this value. The minimum allowed fz 2 is initialized at (1/64)Fs , and fz 2 is itself initialized to the minimum value. If the phase margin is low with fz 2 placed at this minimum value, then there is no fz 2 location that can meet phase margin specifications. In this case, fc,targ is reduced and the fc,targ loop reiterated. In all other cases, fz 2 is iterated, as described previously for the B cases. (5) Gain margin is checked by finding the −180◦ frequency (where the real part is negative and the imaginary part first becomes positive) and computing the magnitude. The presence of multiple crossover frequencies is detected by checking whether the magnitude of the loop frequency response drops below 1 for frequencies less than fc,targ or rises above 1 for frequencies greater than fc,targ . Insufficient loop gain magnitude at low frequencies is detected by searching for the first −45◦ frequency (identified by positive real part, negative imaginary part, and magnitude of the real part becoming greater than the imaginary part). If this frequency is less than fc,targ , then the loop gain begins to flatten since the PID includes an integrator. Flattening of the loop gain at frequencies below the crossover frequency is undesirable because, for many systems, it results in insufficient loop gain magnitude at frequencies below fc,targ , which results in a lower effective closed-loop bandwidth than expected. If insufficient gain margin, multiple crossover frequencies, or flattening of the low-frequency loop gain is detected, then fc,targ is reduced and the fc,targ loop reiterated. The only exception is the case where only the gain margin specification cannot be met and fz 2 = Fs /2. In this case, the zero z2 is directly set equal to zero and the specifications are rechecked. III. EXPERIMENTAL RESULTS A. Controller Hardware Implementation In general, the system-identification procedure to compute the converter frequency responses and the autotuning algorithm described in Section II could be implemented using a microcontroller or dedicated logic gates. The experimental prototypes described in this paper are based on dedicated logic gate implementation coded in Verilog on a Xilinx Virtex-IV fieldprogrammable gate array (FPGA) clocked at 25 MHz. Logic Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. 2584 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 24, NO. 11, NOVEMBER 2009 TABLE V MEAN VALUES OF AUTOTUNING RESULTS OVER 100 AUTOTUNING REPETITIONS Fig. 6. Experimental reference steps. (a) Buck #1: 1.5–1.75 V; time: 400 µs/division; ch. 1: output voltage, 250 mV/division, dc coupled; ch. 2: reference step enable. (b) Buck #2: 1.5–1.75 V; time: 400 µs/division; ch. 1: output voltage, 250 mV/division, dc coupled; ch. 2: reference step enable. (c) Buck #3: 1.5–1.75 V; time: 400 µs/division; ch. 1: output voltage, 250 mV/division, dc coupled; ch. 2: reference step enable. (d) Boost #1: 30–35 V; time: 400 µs/division; ch. 1: output voltage, 5.0 V/division, dc coupled; ch. 2: reference step enable. (e) Boost #2: 30–35 V; time: 400 µs/division; ch. 1: output voltage, 5.0 V/division, dc coupled; ch. 2: reference step enable. Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. SHIRAZI et al.: AUTOTUNING DIGITAL CONTROLLER FOR DC–DC POWER CONVERTERS 2585 Fig. 7. Experimental load steps. (a) Buck #1: 0–9 A; time: 200 µs/division; ch. 1: output voltage, 20 mV/division, ac coupled; ch. 2: load step enable. (b) Buck #2: 0–9 A; time: 200 µs/division; ch. 1: output voltage, 50 mV/division, ac coupled; ch. 2: load step enable. (c) Buck #3: 0–9 A; time: 400 µs/division; ch. 1: output voltage, 200 mV/division, ac coupled; ch. 2: load step enable. (d) Boost #1: 0.3–0.6 A; time: 200 µs/division; ch. 1: output voltage, 2.0 V/division, ac coupled; ch. 2: load step enable. (e) Boost #2: 0.3–0.6 A; time: 200 µs/division; ch. 1: output voltage, 1.0 V/division, ac coupled; ch. 2: load step enable. gate and memory (RAM and ROM) requirements to compute the frequency response are given in [18]. In particular, for a PRBS based on a 10-bit shift register, the system-identification algorithm requires four 1024 × 18-bit RAM blocks to compute and store the frequency response, a 256 × 16-bit ROM block for the complex exponential lookup table (LUT), and a 512 × 16-bit ROM block for the discrete-zero LUT. The autotuning algorithm shares these memory blocks. The algorithm accesses the frequency response stored in one RAM pair and stores the computed loop frequency response in the other RAM pair. To provide finer tuning, the autotuning algorithm extends the discrete-zero LUT to 2048 × 16 bits to achieve four times the frequency resolution of the discrete-zero LUT used by the system-identification algorithm. In addition, the autotuning algorithm requires a 512 × 24-bit LUT for the imaginary part of the discrete integrator (the real part is exactly equal to −0.5 for all frequencies except dc). All calculations were performed using fixed-point arithmetic, and require one unsigned 16 × 16 multiplier, two signed 16 × 16 bit multipliers, and one unsigned 16-bit divider. The multipliers are shared with those used by the system-identification algorithms, but no effort was made to share the unsigned 16-bit divider. Resource requirements are listed in Table II. B. Autotuning Results Hardware prototypes for the three buck configurations and the two boost configurations are shown in Figs. 3 and 4, respectively. Table III lists the component values. In Table III, and in all the following figures and tables, buck #1, buck #2, and Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. 2586 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 24, NO. 11, NOVEMBER 2009 buck #3 refer to the well-damped buck, lightly damped buck, and lightly damped buck with poorly damped input filter, respectively. Similarly, boost #1 and boost #2 refer to the CCM boost and DCM boost converters, respectively. All converters are operated at Fs = 195 kHz. The buck converter prototype uses Vishay Siliconix Si 4888DY MOSFETs. A Texas Instruments THS1230 A/D converter (ADC) samples the output voltage once-per-switching cycle, 1.08 µs prior to the rising of the trailing-edge DPWM signal. The THS1230 is a 12-bit ADC, but only 9 bits are used. The sensing gain is Hv = 1 and the quantization interval is qAD,buck = 7.8 mV. The boost converter prototype uses a Fairchild Semiconductor NDT3055 JFET and a Fairchild Semiconductor SS16 Schottky diode. An Analog Devices AD7822BR ADC samples the output voltage 720 ns prior to the rising of the trailing-edge DPWM signal. As shown in Fig. 4, the sensing gain Hv = 12/406 for an effective quantization interval qAD,boost = 262 mV. The DPWM and digital PID were implemented on the same FPGA as the system-identification and autotuning algorithms. The DPWM is a 12-bit hybrid DPWM clocked at 25 MHz. With Fs = 195 kHz and the PRBS generated by a 10-bit shift register, the resolution of the measured frequency response is ∆f = 191 Hz. For all autotuning results, the phase margin requirement is 50◦ , the gain margin requirement is 10 dB, and the integral-based no-limit-cycling safety factor α is set equal to 0.7. Fig. 5(a)–(e) shows the experimentally computed loop frequency responses of the well-damped buck, lightly damped buck, lightly damped buck with poorly damped input filter, CCM boost, and DCM boost, respectively. These results are compared to plots of the loop transfer function obtained using the PID coefficients from the FPGA and a discrete-time model of the converter [22], [23]. For the buck with input filter, the extra-element theorem is applied to the converter averaged model to obtain an averaged continuous-time model [24], which is then converted to discrete-time model using the impulse invariant mapping [25]. The discrete-time control-to-output transfer functions for each of the converters are listed in Table IV as a reference. The complete system-identification and autotuning procedure was run 100 times for each converter. Table V gives the mean values of compensator parameters and the mean values of crossover frequency, stability margins, and integral-based no-limit-cycling measures, based both on the measured open-loop frequency response, as well as the converter discrete-time model, over the 100 runs. The standard deviation of the crossover frequencies based on the analytical Gv d is 0.117, 1.022, 0.020, 0.085, and 0.482 kHz for the well-damped buck, lightly damped buck, buck with input filter, CCM boost, and DCM boost, respectively. Table V reveals the bandwidth-limiting specification(s) for each case. The well-damped buck is limited primarily by the integral-based nolimit-cycling criterion, while the lightly damped buck is limited both by the no limit cycling and the gain margin constraint. In both of these cases, although the phase margin itself is near the specification, the second PID zero has not saturated anywhere near its lower limit, indicating that the bandwidth could have been pushed higher if it was not otherwise limited. Further- TABLE VI MAXIMUM DURATIONS OF SYSTEM-IDENTIFICATION AND COMPLETE AUTOTUNING PROCEDURE OVER 100 REPS more, the bandwidth for the lightly damped tuning is at or near the maximum allowed Fs /8. Both the buck with poorly damped input filter and the CCM boost require integral-only control with very low bandwidths to meet all specifications. In particular, the phase margin constraint pushes the bandwidth near, but not necessarily below, the converter resonance(s). The bandwidth is then pushed even lower, to well below the resonances, in order to avoid multiple crossover frequencies and meet gain margin constraints. Finally, the DCM boost bandwidth is limited by the integral-based no-limit-cycling criterion. For the buck with filter, the gain margin computed using the analytical Gv d is much lower than the gain margin predicted by the tuning algorithm. This is due to the finite-frequency resolution of the measured frequency responses coupled with the high Q-factor of the input filter. Fig. 6(a)–(e) shows reference steps of 1.5–1.75 V for the buck converters and 30–35 V for the boost converters. Fig. 7(a)–(e) shows load steps of 0–9 A for the buck converters and 0.3–0.6 A for the boost converters. As expected, Fig. 7(c) and (d) shows that the closed-loop output impedance of the buck with poorly damped input filter, as well as the CCM boost, share the same dynamics as the respective open-loop output impedance due to bandwidth being less than the resonant frequency (or frequencies in the case of the buck with input filter) of the converter. An additional reason for the large output voltage deviations for both boost cases is the very small output capacitance of these prototypes (1.8 µF). The small output capacitance was chosen to present a challenging tuning case for the CCM boost, i.e., a case that requires the crossover frequency be pushed to a very low frequency below the resonant frequency to meet all specifications. Finally, Table VI shows the maximum duration of system identification, as well as complete autotuning duration (including system identification), for each converter over 100 repetitions. IV. CONCLUSION This paper has described a hardware description language (HDL) coded autotuning algorithm for digital PID controlled dc–dc power converters based on online identification of control-to-output frequency response. The algorithm determines the PID controller parameters required to maximize the closed-loop bandwidth of the feedback control system while Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. SHIRAZI et al.: AUTOTUNING DIGITAL CONTROLLER FOR DC–DC POWER CONVERTERS maintaining user-specified stability margins and integral-based no-limit-cycling criteria, as well as ensuring single-crossoverfrequency operation and sufficiently high loop gain magnitude at low frequencies. The combined Verilog-coded implementation of the identification and self-tuning algorithms requires 84000 logic gates and 15 kB of memory. Versatility of the proposed autotuning method was demonstrated by applications to several experimental prototypes. The exact same identification and tuning algorithms are able to successfully tune, in 350 ms or less, five different digitally controlled PWM dc–dc converters, including a well-damped synchronous buck, lightly damped synchronous buck, lightly damped synchronous buck with a poorly damped input filter, CCM boost, and DCM boost, covering a representative range of dynamics commonly encountered in dc–dc power converters: well-damped second-order system, lightly damped second-order system, fourth-order system with a lightly damped plus a nearly undamped resonance, non-minimum-phase lightly-damped second-order system and first order system, respectively. Compared to alternative autotuning or adaptive tuning approaches, advantages of the proposed method include the ability to take into account multiple design criteria: stability margin specifications, no-limit-cycle criteria, single-crossoverfrequency operation, and sufficiently large low-frequency loop gain magnitude, which can be ensured without requiring specific a priori information regarding converter power stages. On the other hand, it is also clear that adequate stability margins must still be maintained to account for the expected variation in measured frequency response due to A/D quantization noise. Furthermore, the disruption of normal operation during frequency-response identification prohibits this technique from continuously running, and thus, from being able to easily track system dynamics that change over time. In contrast, the MRAC techniques of [4]–[6] are not only better suited to tracking system dynamics over time, but they also enjoy greater signal-to-noise ratios, and generally lead to simpler implementations. The autotuning and adaptive control schemes are indeed complementary. If sufficient a priori knowledge were already available, an adaptive only control scheme would suffice. If changes in system dynamics could be detected, allowing retriggering of the autotuning procedure, an autotuning-only control scheme could be implemented. But, in the most general case, they could be used in combination. In this case, the autotuning algorithm executes upon converter start-up to initialize the controller, and provide the necessary a priori information regarding converter dynamics and achievable bandwidth/phase margin constraints to the adaptive algorithm, which can then take over to fine tune the controller and adapt to changes in system dynamics. APPENDIX DERIVATION OF THE INTEGRAL-BASED NO-LIMIT-CYCLING EXPRESSIONS OF TABLE I The results of Table I are derived by first converting (3) into an expression for the maximum allowed 0-dB crossing frequency 2587 TABLE VII G c (ej 2 π f c , targ T s ) BASED ON BODE STRAIGHT-LINE ASYMPTOTES of the integral asymptote of the PID. At very low frequencies, the PID behaves like an integrator whose magnitude response crosses 0 dB at fk Ki Ki Ki (6) 1 − ej 2π f k T s ≈ 2πfk Ts = 1 ⇒ fk = 2π Fs . The fk frequency is still relevant even if the magnitude response of the actual PID never reaches 0 dB due to large integral gain and/or low placement of the first PID zero—in this case, fk is the projected 0-dB crossing of the integral asymptote. However, for such a high-gain PID, it is possible that the approximation made in (6), which holds only for fk Fs , may no longer be valid. Fortunately, the whole point of implementing these constraints is to prevent fk from being too large, i.e., the assumption is made that fk Fs and the resulting constraints on PID zero locations will ensure that the assumption holds. The constraint in (3) can now be expressed as fk < α Fs . 2πHv Gv d0 (7) Knowing that fk and fz 1 both lie on a −20-dB/decade asymptote, with magnitude equal to 0 dB at fk (whether or not the asymptotes of the complete PID pass through this point) allows the construction of approximate expressions, based on Bode straight-line asymptotes, for the magnitude of the PID controller Gc (ej 2π f c , targ T s ) in terms of fk , fc,targ , fz 1 , and fz 2 . It can be shown that the resulting expressions, shown in Table VII, hold no matter what the relationship of fk to fc,targ , fz 1 , and fz 2 . The expressions of Table VII will be used here to illustrate the concept; in practice, however, appropriate scaling factors are applied to account for the deviation of the exact magnitude plots from the asymptotes at and near the locations of the zeros. The expressions of Table VII, along with the required inequality of (7), impose design constraints when inserted into the defining equation for crossover frequency Hv Gv d ej 2π f c , targ T s Gc ej 2π f c , targ T s = 1. (8) For the B1 and B2 cases, where fz 2 ≥ fc,targ , these constraints are in the formof a minimum allowed magnitude of Hv Gv d ej 2π f c , targ T s . For the A1 and A2 cases, where fz 2 < fc,targ , these constraints are in the form of a maximum allowed frequency for fz 2 . Table I lists the results. REFERENCES [1] K. J. Astrom and B. Wittenmark, Adaptive Control, 2nd ed. MA: Addison-Wesley, 1995. Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply. Reading, 2588 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 24, NO. 11, NOVEMBER 2009 [2] K. J. Astrom, T. Hagglund, C. C. Hang, and W. K. Ho, “Automatic tuning and adaptation for PID controllers—A survey,” Control Eng. Practice, vol. 1, no. 4, pp. 699–714, 1993. [3] L. Corradini, P. Mattavelli, and D. Maksimovic, “Robust relay-feedback based autotuning for DC–DC converters,” in Proc. IEEE Power Electron. Spec. Conf., 2007, pp. 2196–2202. [4] L. Corradini, P. Mattavelli, W. Stefanutti, and S. Saggini, “Simplified model reference-based autotuning for digitally controlled SMPS,” IEEE Trans. Power Electron., vol. 23, no. 4, pp. 1956–1963, Jul. 2008. [5] J. Morroni, R. Zane, and D. Maksimovic, “Design and implementation of an adaptive tuning system based on desired phase margin for digitally controlled DC–DC converters,” in IEEE Trans. Power Electron., vol. 24, no. 2, pp. 559–564, Feb. 2009. [6] J. Morroni, L. Corradini, R. Zane, and D. Maksimovic, “Robust adaptive tuning of digitally controlled switched-mode power supplies,” in Proc. IEEE Appl. Power Electron. Conf., 2009, pp. 240–246. [7] K. J. Astrom and B. Wittenmark, “On self-tuning regulators,” Automatica, vol. 9, no. 2, pp. 185–199, 1973. [8] K. J. Astrom, U. Borrison, L. Ljung, and B. Wittenmark, “Theory and application of self-tuning regulators,” Automatica, vol. 13, no. 5, pp. 457– 476, 1977. [9] A. Kelly and K. Rinne, “A self-compensating adaptive digital regulator for switching converters based on linear prediction,” in Proc. IEEE Appl. Power Electron. Conf., 2006, pp. 712–718. [10] S. Jeong and S. Song, “Improvement of predictive current control performance using online parameter estimation in phase controlled rectifier,” IEEE Trans. Power Electron., vol. 22, no. 5, pp. 1820–1825, Sep. 2007. [11] S. Saggini, W. Stefanutti, E. Tedeschi, and P. Mattavelli, “Digital deadbeat control tuning for dc–dc converters using error correlation,” IEEE Trans. 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Power Electron. Conf., 2005, pp. 2729– 2735. [17] A. Davoudi, N. Kong, M. Hagen, M. Muegel, and P. Chapman, “A general framework for automated tuning of digital controllers in multi-phase DC– DC converters,” in Proc. IEEE Appl. Power Electron. Conf., 2009, pp. 626– 630. [18] M. Shirazi, J. Morroni, A. Dolgov, R. Zane, and D. Maksimovic, “Integration of frequency response measurement capabilities in digital controllers for DC–DC converters,” IEEE Trans. Power Electron., vol. 23, no. 5, pp. 2524–2535, Sep. 2008. [19] B. Miao, R. Zane, and D. Maksimovic, “System identification of power converters with digital control through cross-correlation methods,” IEEE Trans. Power Electron., vol. 20, no. 5, pp. 1093–1099, Sep. 2005. [20] A. V. Peterchev and S. R. Sanders, “Quantization resolution and limitcycling in digitally controlled PWM converters,” IEEE Trans. Power Electron., vol. 18, no. 1, pp. 301–308, Jan. 2003. [21] H. Peng, A. Prodic, E. Alarcon, and D. Maksimovic, “Modeling of quantization effects in digitally controlled DC–DC converters,” IEEE Trans. Power Electron., vol. 22, no. 1, pp. 208–215, Jan. 2007. [22] C. C. Fang and E. H. Abed, “Sampled-data modeling and analysis of the power stage of PWM DC–DC converters,” Int. J. Electron., vol. 88, no. 3, pp. 347–369, Mar. 2001. [23] D. Maksimovic and R. Zane, “Small-signal discrete-time modeling of digitally-controlled DC–DC converters,” IEEE Trans. Power Electron., Lett., vol. 22, no. 6, pp. 2552–2556, Nov. 2007. [24] R. W. Erickson and D. Maksimovic, Fundamentals of Power Electronics, 2nd ed. Norwell, MA: Kluwer, 2001. [25] A. V. Oppenheim, R. W. Schafer, and J. R. Buck, Discrete-Time Signal Processing, 2nd ed. Upper Saddle River, NJ: Prentice-Hall, 1999. Mariko Shirazi (S’09) received the B.S. degree in mechanical engineering from the University of Alaska, Fairbanks, in 1996, and the M.S. degree in electrical engineering in 2007 from the University of Colorado, Boulder, where she is currently working toward the Ph.D. degree in electrical engineering. From 1996 to 2004, she was an Engineer at the National Wind Technology Center, National Renewable Energy Laboratory, where she was involved in the design and deployment of hybrid wind-diesel power systems for village power applications. Her current research interests include system identification and autotuning of digitally controlled switched-mode power supplies. Regan Zane (S’98–M’00–SM’07) received the B.S, M.S., and Ph.D. degrees in electrical engineering from the University of Colorado, Boulder, in 1996, 1998, and 1999, respectively. From 1999 to 2001, he was with the GE Global Research Center, Niskayuna, NY, where he developed custom IC controllers for power electronic circuits and systems. From 2001 to 2007, he was an Assistant Professor of electrical and computer engineering at the University of Colorado, where he has been an Associate Professor since 2008. He is engaged in research programs in energy-efficient lighting systems, adaptive algorithms and digital control techniques in power electronics systems, and low power energy harvesting for wireless devices. Dr. Zane received the 2004 National Science Foundation CAREER Award, the 2005 IEEE Microwave Best Paper Prize, the 2008 IEEE POWER ELECTRONICS SOCIETY (PELS) TRANSACTIONS Prize Letter Award, and the 2008 IEEE PELS Richard M. Bass Outstanding Young Power Electronics Engineer Award. He received the 2006 Inventor of the Year award, the 2006 Provost Faculty Achievement Award, and the 2008 John and Mercedes Peebles Innovation in Teaching Award, all from the University of Colorado. He is currently an Associate Editor for the IEEE TRANSACTIONS ON POWER ELECTRONICS LETTERS and a Member-At-Large of the IEEE PELS AdCom. Dragan Maksimovic (M’89–SM’04) received the B.S. and M.S. degrees in electrical engineering from the University of Belgrade, Belgrade, Yugoslavia, in 1984 and 1986, respectively, and the Ph.D. degree from California Institute of Technology, Pasadena, in 1989. From 1989 to 1992, he was with the University of Belgrade. Since 1992, he has been with the Department of Electrical and Computer Engineering, University of Colorado, Boulder, where he is currently a Professor and the Director of the Colorado Power Electronics Center. His research interests include digital control techniques and mixed-signal IC design for power electronics. Prof. Maksimovic received the National Science Foundation CAREER Award in 1997, the IEEE POWER ELECTRONICS SOCIETY TRANSACTIONS Prize Paper Award in 1997, the Bruce Holland Excellence in Teaching Award in 2004, and the University of Colorado Inventor of the Year Award in 2006. Authorized licensed use limited to: UNIVERSITY OF COLORADO. Downloaded on December 25, 2009 at 21:47 from IEEE Xplore. Restrictions apply.

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