3364 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 27, NO. 7, JULY 2012Instead of modifying the voltage loop, adding a current loopwith an inductor current sensing circuit is considered as an al-ternative to implementing variable-duty-cycle control of DCMPFC rectiﬁer. The current loop can effectively suppress currentdistortion by the initial inductor current perturbation . How-ever, digitally sampling the inductor current in DCM operationis not as straightforward as it is in continuous conduction mode(CCM) . In DCM, the sampled inductor current needs to becorrected to properly represent the average inductor current perswitching period , . The digital controller should deviatethe sampling instant within the switching period and calculatea correction factor to determine the average inductor current,which consumes considerable resources of the controller. Sam-pling the inductor current more than once in a switching cyclemay help to achieve precise average inductor current, though itrequires a high-performance A/D converter for high-rate sam-pling .In this paper, an average current-mode control for DCM PFCrectiﬁer is digitally implemented to realize the variable-duty-cycle control. With a direct sensing of the average inductorcurrent, the high-bandwidth current loop dramatically reducesthe input current distortion. It uses the conventional sensingcircuit and ﬁxed sampling instant to sample the average in-ductor current. The sampling occurs once per switching cycle,which does not need high-rate sampling A/D converter. Thecontrol algorithm of the proposed control implementation re-quires the same calculation burden with that of conventionaltwo-loop-controlled converter, and additional computation suchas correction factor calculation is not necessary.This paper consists of the following sections. In Section II,circuit conﬁguration, mode analysis, and operation of the sens-ing circuit as well as the two-loop control strategy are given.With the mathematical expression of properly controlled induc-tor current shape, it is explained how the variable-duty-cyclecontrol can achieve lower current distortion than constant-duty-cycle control. Section III illustrates design procedure of the pro-posed control implementation. It includes the design of boostinductance, sensing capacitance, and digitally realized controlloops. Experimental results with 200-W prototype hardware arepresented in Section IV. Conclusions are presented in Section V.II. CONTROL TECHNIQUE FOR DCM PFC RECTIFIERA. Operation of Average Sensing CircuitThe proposed control technique follows the general structureof the conventional two-loop control to build an average current-mode control, as shown in Fig. 1. The average sensing circuit,the shaded box in Fig. 1, gives the average inductor currentin each switching cycle to establish the current-mode control.The sampling of the average inductor current occurs once in aswitching period like other controller inputs such as rectiﬁedinput voltage and output voltage. It does not need high-ratesampling, which samples twice or more per switching period;thus, it can save the resources of the digital controller.Circuit conﬁguration and operational waveform of the aver-age sensing circuit are shown in Fig. 2. As shown in Fig. 2(a),the circuit looks the same as the conventional current sensingFig. 1. Proposed control technique applied to the PFC rectiﬁer for digitalaverage current-mode control.Fig. 2. Average sensing circuit. (a) Circuit structure. (b) Operationalwaveforms.circuit utilized in charge current-mode control , but its oper-ation differs in the proposed control. In the charge current-modecontrol, the sensing capacitor voltage vC s represents the aver-age switch current per switching cycle by integrating the switchcurrent during ON-time. The voltage is reset to zero right af-ter the main switch turns off by the inversed main gate signal.
SHIN AND CHO: DIGITALLY IMPLEMENTED AVERAGE CURRENT-MODE CONTROL IN DISCONTINUOUS CONDUCTION MODE 3365Fig. 3. Comparison of current waveform shapes. (a) Average current-mode-controlled CCM. (b) Constant-duty-cycle-controlled DCM. (c) Variable-duty-cycle-controlled DCM.In the proposed control, on the other hand, the same circuit isutilized to integrate not the switch current, but both the switchand the diode current. The integrating interval of the sensingcircuit is extended to near the end of the switching cycle by de-laying the timing of the reset signal. Fig. 2(b) shows the typicalwaveform of the proposed average sensing circuit operation inDCM PFC rectiﬁer. vgs, iL , reset, and TS are the gate signal formain switch, the inductor current, the gate signal for auxiliaryswitch Qaux in average sensing circuit, and the switching pe-riod, respectively. d1 and d2 are the duty ratios for the switchconduction interval and diode conduction interval in the powerstage. To take advantage of the analog integrating circuit andavoid the high-rate sampling, the circuit integrates the scaledinductor current in the sensing capacitor CS and samples vC sonce in the switching cycle. In the ﬁrst interval in kth switchingcycle, kTS < t < t1, the main switch is in ON-state. iL increaseslinearly and vC s increases in parabolic way because the linearterm of iL is integrated, as expressed in the following:vCs(t) =1nCStkTSiL (t)dt. (1)In (1), n is the turn ratio of the current transformer. In the secondinterval, t1 < t < t2, the main switch is in OFF-state. iL decreasesand the sign of the parabolic curve of vC s is inverted. In the thirdinterval, t2 < t < (k + 1)TS , of which the time duration is d3TS ,iL becomes zero and the sampled value vC s[k] fully representsthe average inductor current, as follows:vCs[k] =TSnCS1TS(k+1)TSkTSiL (t)dt =TSnCSiavg [k]. (2)The term in the braces in (2) is the average inductor current in thekth switching cycle iavg . Thus, vC s[k] represents the scaled valueof the average inductor current. Qaux in the average sensingcircuit turns on after the sampling to discharge CS before thenext switching period.Sampling vC s[k] in the third interval occurs when the re-maining time of the switching cycle is Tcal. Tcal is the minimumrequired time that includes A/D conversion and reset of vC s, andcalculation and update of the duty cycle for the next switchingcycle. Because the resetting time of vC s[k] becomes negligibleby selecting Qaux with low ON-state resistance, Tcal is effec-tively ﬁxed by the calculation time according to the controlalgorithm programmed in the microcontroller. Therefore, thesampling time can also be ﬁxed within the switching cycle. Thesampling time and the inductance design to secure sufﬁcientTcal will be explained later.B. Current Waveforms in DCM PFC RectiﬁerFig. 3 compares the inductor current waveforms of averagecurrent-mode-controlled CCM, constant-duty-cycle-controlledDCM, and variable-duty-cycle-controlled DCM over a half lineperiod. The continuous-time-domain current waveforms suchas ipk(t) and iavg (t) in Fig. 3 are the approximated ones of thediscrete-time-domain waveforms ipk[k] and iavg [k], assumingthat the switching frequency of the rectiﬁer is much higherthan the line frequency. The waveforms of average current-mode-controlled CCM PFC rectiﬁer in Fig. 3(a) resemble thesinusoid of the line voltage because the switching ripple ofiL is relatively small. However, unlike in CCM rectiﬁer, theenvelope of the inductor current in DCM PFC rectiﬁer shouldnot look like a sinusoid to demonstrate low current distortion.In Fig. 3(b), constant-duty-cycle control constructs the envelopeof the inductor current ipk(t) to be sinusoidal. However, theresultant ﬁltered inductor current or the line current iavg (t) ishighly distorted. Analysis of line current distortion accordingto line voltage in constant-duty-cycle-controlled DCM is givenin . In contrast to Fig. 3(b), Fig. 3(c) shows that in variable-duty-cycle-controlled DCM, iavg (t) follows the sinusoidal curvethough ipk(t) does not.If the average inductor current per switching cycle is con-trolled to be proportional to the sinusoidal input voltage by thevariable-duty-cycle control, ipk(t) can be derived from the dutyratios d1 and d2, as expressed in (3) and (4)d1 =Lipk(t)|vin(t)| TS(3)d2 =Lipk(t)(vO − |vin(t)|) TS(4)where L, vO , and vin(t) are inductance in power stage, outputvoltage, and line voltage, respectively. iavg (t) is derived fromthe geometry of the waveform shown in Fig. 2(b) with unknownconstant X, as shown in the following:iavg (t) =12(d1 + d2)ipk(t) = X sin ωL t (5)where vin(t) is assumed to be the pure sine wave with peakvalue Vpk and line period TL = 2π/ωL , which is expressed inthe following:vin(t) = Vpk sin ωL t. (6)Substituting (3), (4), and (6) into (5) yields the following:ipk(t) =2TS VpkXL1 −Vpk sin ωL tvOsin ωL t. (7)
3366 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 27, NO. 7, JULY 2012Fig. 4. (a) ipk curves of variable-duty-cycle-controlled DCM PFC rectiﬁer forvarious line voltages. (b) Normalized ipk curves compared with the sine wave.The unknown constant X is derived by considering input power.If Pin, PO , and η are input power, output power, and efﬁciencyof the rectiﬁer system, respectively, i.e., input ﬁlter and rectiﬁer,the following can be achieved:Pin =POη= IrmsVrms =12VpkX. (8)Irms and Vrms in (8) represent rms of line current and line voltagein a half line cycle, respectively. Rearranging and substituting(8) into (7) yields the following:ipk(t) = 2TS POηL1 −Vpk sin ωL tvOsin ωL t. (9)Fig. 4(a) compares various ipk(t) curves in case the variable-duty-cycle-controlled DCM rectiﬁer outputs 400 V and 200 W,assuming that the efﬁciency of the rectiﬁer system is unity. Themaximum inductor current occurs at the midpoint of the halfline period when line voltage is lower than 188.6 Vac. When theline voltage is higher than 188.6 Vac, a local minimum forms inthe middle of the curve and the maximum current occurs twicein the half line period. Detailed derivation of maximum ipk(t)is discussed in Section III-A. Fig. 4(b) shows the normalizedipk(t) with the pure sinusoid. Low line voltages such as 90 and115 Vac demonstrate ipk(t) curves, which are slightly changedfrom the sine wave. On the contrary, curves for high line voltagessuch as 230 and 264 Vac are highly deviated and show thelocal minimum in the middle of the half line period. This isbecause the second square-root term in (9) is not a constant,but a function of sin ωL t. The deviation of the term over halfline cycle becomes dominant when the line voltage is high, i.e.,Vpk/vO is high. Considering that the pure sinusoid is the ipkwaveform of the conventional constant-duty-cycle-controlledrectiﬁer, the performance improvement of the proposed controltechnique may be trivial with low line voltages like 90 or 115Vac. However, it can be emphasized if the high line voltage suchas 230 or 264 Vac is applied.III. RECTIFIER DESIGN WITH THE PROPOSEDCONTROL TECHNIQUEPractical implementation of the proposed average current-mode variable-duty-cycle control to boost PFC rectiﬁer is shownin Fig. 5. Loop conﬁguration of the proposed control followsthe general two-loop control with the input voltage feedforward.Output of the voltage compensator GV and the sensed rectiﬁedinput voltage are multiplied together and compared with thesensed average inductor current. Current compensator GC out-put is again compared with the counter value of the pulse widthmodulation (PWM) generator and provides the gate signals tothe power stage.A. Inductance DesignA large inductance of L is effective because it decreases therms current in DCM PFC rectiﬁer. However, the maximum limitof the inductance Lcrit exists to maintain the DCM operationthroughout the line cycle and load variation. Otherwise, the rec-tiﬁer operation may go into CCM and lose control even thoughthe control loop is designed to be stable. Speciﬁcally, the timeduration of the third interval d3TS should be considered in theinductance design because the sampling of vC s and calculationof the next duty cycle is executed in the third interval. Fig. 6illustrates the shrinkage of d3TS according to the increase of L.Assuming the same level of iavg [k], d3TS becomes shorter whenL is larger because the slopes of iL become more gradual.Lcrit is derived in the following to properly reset vC s andobtain sufﬁcient calculation time Tcal: d3min, the minimum dutyratio of the third switching interval, is expressed in (10) byconsidering (3), (4), (6), and (9)d3m in= 1 − (d1 + d2) = 1 −2VpkLcritPOηTS×11 − (Vpk/vO ) sin ωL t. (10)Equation (10) notes that d3min occurs in the middle of the halfline cycle, i.e., sin ωL t = 1. d3min is plotted as the function ofLcrit for various line voltages and PO in Fig. 7, assuming thatTS is 15.385 μs (reciprocal of 65 kHz) and η is a unity. The
SHIN AND CHO: DIGITALLY IMPLEMENTED AVERAGE CURRENT-MODE CONTROL IN DISCONTINUOUS CONDUCTION MODE 3367Fig. 5. Implementation of the proposed control technique to boost PFC rectiﬁer.Fig. 6. Decrease of d3 TS according to the increase of L.worst case for d3min is high line voltage and full load situation,which is represented by the solid curve at the bottom.Selection of d3min affects the design of Lcrit, and conse-quently, the maximum and the rms inductor currents. The max-imum inductor current in a half line period, Ipk max, is derivedby differentiating and equating (9)ddωL tipk(t)=TS POηL(cos ωL t) (2 − (3Vpk/vO ) sin ωL t)1 − (Vpk/vO ) sin ωL t=0.(11)Fig. 7. d3m in as the function of L to determine the critical inductance Lcrit .For 0 < t < (TL /2), one obvious solution for (11) is whencos ωL t = 0 or ωL t = π/2. The other solutions exist onlywhen (2/3)(vO /Vpk) < 1, considering the term in the sec-ond parentheses of the numerator in (11). If vO is 400 V and(2/3)(vO /Vpk) > 1, i.e., Vrms < 188.6 V, (9) has the maxi-mum at ωL t = π/4 and does not have any local minimum. IfVrms > 188.6 V, (9) has the local minimum at ωL t = π/4 andthe maximums occur at ωL t = sin−1(2/3)(vO /Vpk). Substitut-ing ωL t = π/2 and ωL t = sin−1(2/3)(vO /Vpk)in (9) accord-ing to the condition of Vpk, Ipk max is mathematically expressed
3368 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 27, NO. 7, JULY 2012Fig. 8. Ipk m ax versus Vpk for various d3m in .in (12)Ipk max = 2TS POηL1 −VpkvOif23vOVpk> 1 (12a)Ipk max =43√3TS POηLVpkvOif23vOVpk< 1 . (12b)The rms inductor current in a half line period Irms should also beconsidered in d3min design. Irms is expressed, as in the follow-ing, according to the analysis in , assuming that the rectiﬁerefﬁciency is a unityIrms =VpkTSL2TLTL /20VO d1(t)3 sin2ωL t3(VO − Vpk sin ωL t)dt (13)where d1(t) is given in (14) by combining (3), (6), and (9)d1(t) =2VpkLPOηTS1 −VpkVOsin ωL t (14)The critical inductances for several d3min such as 0.1, 0.2, 0.3,and 0.4 are determined as 62, 88, 112, and 144 μH, respectively,from the worst-case curve in Fig. 7, the one for 264 Vac and200 W. Ipk max and Irms for the selected critical inductances andline voltages are plotted in Figs. 8 and 9, respectively, assumingthat the rectiﬁer outputs maximum power 200 W. Figs. 8 and9, respectively, show that Ipk max and Irms are proportional tod3min. Therefore, selecting large d3min increases the peak andthe rms inductor current. However, the effect of d3min on Irms isless considerable than that on Ipk max. For example, for increaseof d3min from 0.1 to 0.4, Irms rises by 0.33 A, while Ipk max isincreased by 2.4 A in 230 Vac operation.In this paper, Tcal requires approximately 4 μs to properlycalculate and update the duty cycle, which spends 26% of theunit switching cycle. With 4% margin, d3min is set to be 0.3and Lcrit is determined as 88 μH, as indicated in Fig. 7. Fig. 10shows the variation of d3 throughout the half line cycle with theresultant L when full load is applied. In correspondence withthe curves in Fig. 7, 264 Vac curve demonstrates the smallestFig. 9. Irm s versus Vpk for various d3m in .Fig. 10. Variation of d3 within a half line cycle when L is 88 μH.d3min on the middle of the line cycle, followed by 90, 115, and230 Vac curves by the ascending order of d3min.B. Sensing Capacitance and Current Sensing GainSensing capacitance in the average sensing circuit CS is di-rectly related to the performance of the proposed control tech-nique. The range of CS is limited by practical reasons. SmallCS may not be appropriate because vC s may exceed the max-imum input voltage of the microcontroller. Large CS is eithernot feasible because vC s may not be reset correctly and the A/Dconversion resolution for vC s may become insufﬁcient. The ca-pacitance is recommended to be within the range from a fewtens to a few hundreds of nanofarads.The minimum boundary of CS is not dependent on the induc-tance L, but on the average inductor current in a unit switchingcycle. One can notice from Fig. 6 that vC s at the sampling in-stant is constant though L changes. The boundary is determinedby considering the maximum average inductor current in a halfline cycle Iavg max and the maximum input voltage of the mi-crocontroller VC s max as expressed in the following:TSnCSIavg max < VC s max. (15)
SHIN AND CHO: DIGITALLY IMPLEMENTED AVERAGE CURRENT-MODE CONTROL IN DISCONTINUOUS CONDUCTION MODE 3369Fig. 11. Small-signal diagram of the proposed control technique.If power factor of the rectiﬁer is a unity and the waveformiavg (t) is sinusoidal, Iavg max and Irms have the relationshipshown in the following:Irms =1√2Iavg max (16)and substituting (16) into (8) achieves the following:Iavg max =√2POηVrms min(17)where Vrms min in (12) is the minimum rms line input voltage.From (15) and (17), the minimum CS can be deﬁned as follows:CS >√2TS POnηVC s maxVrms min. (18)In this paper, the minimum CS is 293 nF by considering TS , n,PO , VCs max, η, and minimum line voltage as 15.385 μs, 50,200 W, 3.3 V, a unity, and 90 Vac, respectively.The current sensing gain of the average sensing circuit isconstant and does not change in the frequency domain. Thoughthe proposed sensing circuit integrates the inductor current, theintegrating operation does not continue over consecutive switch-ing cycles. Thus, in contrast to standard analog integrators, thesensing circuit does not have any frequency-dependent dynam-ics such as analog integrators. The effect of the average sensingcircuit on the small-signal modeling appears as the ﬁxed gain Ki,which is the shaded box in Fig. 11. Ki is deﬁned to be TS /nCSas the coefﬁcient of iavg [k] in (2).C. Control Loop ImplementationControl loops are realized by the algorithm programmed in themicrocontroller Microchip dsPIC33FJ16GS502. Current loopgain Ti is deﬁned as in Fig. 11 and the following:Ti = GidKiHeKadcGC FM (19)where Gid is open-loop control-to-inductor current transferfunction iL / ˆd. He, Kadc, GC , and FM are sampling effect,internal A/D conversion gain, current compensator gain, andPWM modulator gain of the digital controller, respectively. TiFig. 12. Current compensator output and PWM counter to derive the modu-lator gain FM.Fig. 13. Bode plots of the current control loop Ti for 200–264 Vac line voltage.is similar to the characteristic of the conventional CCM averagecurrent-mode-controlled converter , .A/D conversion gain and PWM modulator gain are depen-dent on the characteristic of the microcontroller. By the fea-ture of the internal A/D converter and conversion algorithm ofdsPIC33FJ16GS502, Kadc is deﬁned as follows:Kadc =1VC s max= 0.303. (20)PWM modulator gain is only dependent on the maximum valueof the counter in PWM generator in microcontroller. In the ana-log average current-mode control, the modulator gain is the com-plicated function of switching frequency, current compensator,modulator ramp slope, and sensed inductor current slope .In the digital control, however, FM is simply a reciprocal of vpshown in Fig. 12 because the output of GC does not ﬂuctuatewhen it is compared with the PWM counter. The unit step of thecounter is set by the PWM clock signal as 1.04 ns and 14 793steps are required to realize 65-kHz switching frequency. FM isdetermined by the following based on Q-15 data formatFM =114793/(215 − 1)= 2.215. (21)Gid, which can be derived from the averaged circuit model, , shows ﬂat gain up to half of the switching frequency.
3370 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 27, NO. 7, JULY 2012Fig. 14. Steady-state waveforms of the proposed and conventional control methods when line voltage is 230 Vac (iL 5 A/div). (a) Proposed control with 50%load (iavg 1 A/div). (b) Proposed control with 100% load (iavg 2 A/div). (c) Conventional control with 50% load (iavg 2 A/div). (d) Conventional control with100% load (iavg 2 A/div).Current compensator GC is ﬁrst designed in s domain to have anintegrator with single pole to attenuate the HF noise, as shownin the following:GC (s) =ωis11 + (s/ωp ). (22)GC (s) is interpreted into z domain with zero-order-hold dis-cretization method with the sampling frequency 65 kHz. Resul-tant GC (z) is shown in the following:GC (z) =a0 + a1z−11 + b1z−1 + b2z−2. (23)Integrator gain ωi and single pole ωp in (22) are determinedto set the bandwidth and phase margin of Ti to 1 kHz and 75◦for universal line voltage and entire load range. For example,ωi and ωp are 143 Hz and 20 kHz, respectively, in s domain,and the corresponding a0, a1, b1, and b2 in (23) are 0.007772,0.004123, −1.145, and 0.1447, respectively, for 200–264 Vacline voltage. Resultant Ti curves are shown in Fig. 13.As the voltage loop compensation, GV is designed through thesimilar steps of GC design. The rectiﬁed input voltage and theoutput voltage are sensed through the resistive voltage dividershown in Fig. 5. Their sampling instant and A/D conversion gainof vO in voltage loop are the same as those of vC s in current loop.Capacitors in the resistive dividers are for noise reduction andsufﬁciently small not to affect the dynamic characteristics of thecontrol loops. In Fig. 11, Kv and Kﬀ in the voltage loop representthe divider gain of the output voltage and the feedforward gainof the input voltage, respectively. GV is designed to let thevoltage loop Tv have 10-Hz bandwidth and 90◦phase margin toattenuate the half-of-line-period ripple.The calculation procedure for loop compensation algorithmcontains only two vector inner products for each loop and sin-gle multiplying operation for sensed input voltage and voltagecompensator output. It requires the same calculation burden withconventional digitally controlled CCM PFC rectiﬁer. There isno additional square root, multiplying, or dividing operation,which spends a lot of the resources of the microcontroller.
SHIN AND CHO: DIGITALLY IMPLEMENTED AVERAGE CURRENT-MODE CONTROL IN DISCONTINUOUS CONDUCTION MODE 3371Fig. 15. Steady-state waveforms of the proposed and conventional control methods when line voltage is 115 Vac (iL 10 A/div, iavg 5 A/div). (a) Proposed controlwith 50% load. (b) Proposed control with 100% load. (c) Conventional control with 50% load. (d) Conventional control with 100% load.IV. EXPERIMENTAL RESULTSThe proposed current control technique is veriﬁed by a 200-Wlaboratory prototype, as shown in Fig. 5. vO is regulated to be400 V with the universal line input voltage range and the loadrange of 50–200 W. Circuit components and parameters that arenot mentioned in previous sections are stated in the following.1) Lf : 250 μH.2) Cf : 1 μF.3) Dbr : D25XB60.4) L: 70 μH.5) Q: STW20NM50.6) D: FSF10A60.7) C: 220 μF.8) DZ : 1N4744.9) CS : 660 nF.10) Qaux: FDD8770.11) Resistive input voltage sensing gain: 0.0089.Fig. 16. THD comparison between the proposed and the conventionalconstant-duty-cycle control.
3372 IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 27, NO. 7, JULY 2012Fig. 17. Operation of the proposed control technique in load transition when line voltage is 230 Vac (vO ac coupled, 10 V/div). (a) When the load steps from25% to 100%. (b) When the load steps from 100% to 25%.12) Resistive output voltage sensing gain: 0.0025.13) Switching frequency: 65 kHz.14) Gate driver: TC4420.15) Line frequency: 60 Hz.It should be noted that the smaller sensing capacitor CS ispreferred considering the efﬁciency of the rectiﬁer. In case theline voltage is speciﬁed as high range, such as 200–264 Vac, CSsmaller than 660 nF should be selected.Fig. 14 illustrates the current waveforms in steady state whenthe load is 100 W, and 200 W when line voltage is 230 Vac. Asshown in the Fig. 14(a) and (b), the line current of the proposedcontrol follows the sinusoidal shape, as analyzed in Section II,and the variable-duty-cycle-controlled inductor current demon-strates the local minimum in the middle of the half line period.The reason the inductor current shape in a half line cycle isnot symmetrical is that the output of the voltage compensatorGV is not ﬁxed. However, the effect of the asymmetry on thecurrent distortion is negligible. For comparison, Fig. 14(c) and(d) shows the waveforms of the constant-duty-cycle-controlledDCM PFC. The inductor current looks like a sinusoid, but theline current is highly distorted in the middle of the half lineperiod. The same comparison in case the line voltage is 115 Vacis also shown in Fig. 15. In Fig. 15, the proposed control methodstill shows smaller current distortion than constant-duty-cyclecontrol. The inductor currents of the two control method seemto have similar waveform with each other, but the ﬁltered linecurrent waveforms demonstrate different distortion.Total harmonic distortion (THD) measurement results for var-ious loads and rms line voltages are shown in Fig. 16. All thedata are collected by Yokogawa WT210 harmonics measure-ment functions. The proposed control shows smaller THD thanthe conventional constant-duty-cycle control. According to thepower increase, the proposed control further decreases the THD,while constant-duty-cycle control increases it. The impact onTHD improvement between the control methods is larger in 230Vac than in 115 Vac, as analyzed in Section II.Fig. 17 shows the operation of the proposed control when theload steps from 100% to 25% and vice versa when 230 Vac isapplied. The proposed control shapes the inductor current in thetransient operation as well. From the ac-coupled waveform ofvO , the overshoot and undershoot do not exceed 10 V.V. CONCLUSIONA digital average current control method for DCM PFC rec-tiﬁer to achieve low current distortion has been proposed. Themain feature of the proposed control method is that the aver-age inductor current is directly sensed to eliminate additionalcalculation in the microcontroller. The control method achievessmaller current distortion than the conventional constant-duty-cycle control method by employing a conventional sensing cir-cuit. Circuit operation, inductor current shaping, and rectiﬁerdesign procedure have been analyzed and explained with math-ematical expressions. The proposed control technique has beenveriﬁed and compared with the conventional approach basedon the experimental results achieved with a 200 W prototype.Waveforms and measured data have proven the performanceimprovement of the proposed control method.REFERENCES K. Billings, Switchmode Power Supply Handbook, 2nd ed. New York:McGraw-Hill, 1999, pp. 4.3–4.9. Limits for Harmonic Current Emissions (Equipment Input Current < 16A per Phase), IEC 61000-3-2 International Standard, 2001. R. D. Middlebrook, “Topics in multiple-loop regulators and current-modeprogramming,” IEEE Trans. Power Electron., vol. PE-2, no. 2, pp. 109–124, Apr. 1987. P. Barbosa, F. Canales, J. Crebier, and F. C. Lee, “Interleaved three-phase boost rectiﬁers operated in the discontinuous conduction mode:Analysis, design considerations and experimentation,” IEEE Trans. PowerElectron., vol. 16, no. 5, pp. 724–734, Sep. 2001. K. Liu and Y. Lin, “Current waveform distortion in power factor correctioncircuits employing discontinuous-mode boost converters,” in Proc. IEEEPower Electron. Spec. Conf., 1989, pp. 825–829. J. Lazar and S. ´Cuk, “Open loop control of a unity power factor, discon-tinuous conduction mode boost rectiﬁer,” in Proc. IEEE Int. Telecommun.Energy Conf., 1995, pp. 671–677.
SHIN AND CHO: DIGITALLY IMPLEMENTED AVERAGE CURRENT-MODE CONTROL IN DISCONTINUOUS CONDUCTION MODE 3373 J. Sebasti´an, J. A. Martinez, J. M. Alonso, and J. A. Cobos, “Voltage-follower control in zero-current-switched quasi-resonant power factor pre-regulators,” IEEE Trans. Power Electron., vol. 13, no. 4, pp. 727–738,Jul. 1998. K. De Gussem´e, D. M. Van de Sype, A. P. M. Van den Bossche, andJ. A. Melkebeek, “Input-current distortion of CCM boost PFC convertersoperated in DCM,” IEEE Trans. Ind. Electron., vol. 54, no. 2, pp. 858–865,Apr. 2007. Z. Lai, K. M. Smedley, and Y. Ma, “Time quantity one-cycle control forpower-factor-correctors,” IEEE Trans. Power Electron., vol. 12, no. 2,pp. 369–375, Mar. 1997. K. Yao, X. Ruan, X. Mao, and Z. Ye, “Variable-duty-cycle control toachieve high input power factor for DCM PFC boost converter,” IEEETrans. Ind. Electron., vol. 58, no. 5, pp. 1856–1865, May 2011. J. Lazar and S. ´Cuk, “Feedback loop analysis for ac/dc rectiﬁers operatingin discontinuous conduction mode,” in Proc. IEEE Appl. Power Electron.Conf., 1996, pp. 797–806. Z. Z. Ye and M. M. Jovanovi´c, “Implementation and performanceevaluation of DSP-based control for constant-frequency discontinuous-conduction-mode boost PFC front end,” IEEE Trans. Ind. Electron.,vol. 52, no. 1, pp. 98–107, Feb. 2005. D. Weng and S. Yuvarajian, “Constant switching frequency ac–dc con-verter using second harmonic injected PWM,” IEEE Trans. Power Elec-tron., vol. 11, no. 1, pp. 115–121, Jan. 1996. D. S. Schramm and M. O. Buss, “Mathematical analysis of a new harmoniccancellation technique of the input line current in DICM boost converters,”in Proc. IEEE Power Electron. Spec. Conf., 1998, pp. 1337–1343. L. Huber, B. T. Irving, and M. M. Jovanovi´c, “Effect of valley switch-ing and switching-frequency limitation on line-current distortions ofDCM/CCM boundary boost PFC converter,” IEEE Trans. Power Elec-tron., vol. 24, no. 2, pp. 339–347, Feb. 2009. D. M. Van de Sype, K. De Gussem´e, A. P. M. Van den Bossche, and J.A. Melkebeek, “A sampling algorithm for digitally controlled boost PFCconverters,” IEEE Trans. Power Electron., vol. 19, no. 3, pp. 649–657,May 2004. K. De Gussem´e, D. M. Van de Sype, A. P. M. Van den Bossche, and J.A. Melkebeek, “Digitally controlled boost power-factor-correction con-verters operating in both continuous and discontinuous conduction mode,”IEEE Trans. Ind. Electron., vol. 52, no. 1, pp. 88–97, Feb. 2005. F. Chen and D. Maksimovi´c, “Digital control for improved efﬁciency andreduced harmonic distortion over wide load range in boost PFC rectiﬁers,”IEEE Trans. Power Electron., vol. 25, no. 10, pp. 2683–2692, Oct. 2010. G. Zhou and J. Xu, “Digital average current controlled switching DC-DCconverters with single-edge modulation,” IEEE Trans. Power Electron.,vol. 25, no. 3, pp. 786–793, Mar. 2010. W. Tang, F. C. Lee, R. B. Ridley, and I. Cohen, “Charge control: Modeling,analysis and design,” IEEE Trans. Power Electron., vol. 8, no. 4, pp. 396–403, Oct. 1993. L. Dixon, “Average current mode control of switching power supplies,”presented at the Unitrode Power Supply Design Seminar, 1990. F. A. Huliehel, F. C. Lee, and B. H. Cho, “Small-signal modeling of single-phase boost high power factor converter with constant frequency control,”in Proc. IEEE Power Electron. Spec. Conf., Jun. 1992, pp. 475–482. W. Tang, F. C. Lee, and R. B. Ridley, “Small-signal modeling of aver-age current-mode control,” IEEE Trans. Power Electron., vol. 8, no. 2,pp. 112–119, Apr. 1993. V. Vorp´erian, “Simpliﬁed analysis of PWM converters using the model ofthe PWM switch: Parts I and II,” IEEE Trans. Aerosp. Electron. Syst.,vol. 26, no. 2, pp. 490–496, Mar. 1990. R. W. Erickson and D. Maksimovi´c, Fundamentals of Power Electronics,2nd ed. Boston, MA: Kluwer, 2001, pp. 420–427.Jong-Won Shin (S’08) received the B.S. degree inelectrical engineering from Seoul National Univer-sity, Seoul, Korea, in 2006, where he is currentlyworking toward the Ph.D. degree.His research interests include ac–dc and dc–dcpower conversion and their control.Bo-Hyung Cho (M’89–SM’95–F’11) received theB.S. and M.S. degrees from the California Insti-tute of Technology, Pasadena, and the Ph.D. degreefrom Virginia Polytechnic Institute and State Uni-versity (Virginia Tech), Blacksburg, all in electricalengineering.Prior to his research at Virginia Tech, he was amember of the Technical Staff with the Departmentof Power Conversion Electronics, TRW Defense andSpace System Group. From 1982 to 1995, he was aProfessor in the Department of Electrical Engineer-ing, Virginia Tech. In 1995, he joined the School of Electrical Engineering,Seoul National University, Seoul, Korea, where he is currently a Professor.His current research interests include power electronics, modeling, analysis,and control of spacecraft power processing equipment, and distributed powersystems.Dr. Cho is a recipient of the 1989 Presidential Young Investigator Award fromthe National Science Foundation. He chaired the 2006 IEEE Power ElectronicsSpecialists Conference. He is a member of Tau Beta Pi.