Computer and Information Science » Computer Science and Engineering » "Applications of MATLAB in Science and Engineering", book edited by Tadeusz Michałowski, ISBN 978-953-307-708-6, Published: September 9, 2011 under CC BY-NC-SA 3.0 license

Chapter 5

Modeling & Simulation of Hysteresis Current Controlled Inverters Using MATLAB

By Ahmad Albanna
DOI: 10.5772/21143

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Overview

Single-phase half-bridge inverter with fixed-band hysteresis control.
Figure 1. Single-phase half-bridge inverter with fixed-band hysteresis control.
Reference voltage calculation and the instantaneous outputs.
Figure 2. Reference voltage calculation and the instantaneous outputs.
Detail ofea(t).
Figure 3. Detail ofea(t).
Effect of va* on the error current duty cycle.
Figure 4. Effect of va* on the error current duty cycle.
Effect of changing β on the harmonic spectrum.
Figure 5. Effect of changing β on the harmonic spectrum.
AC harmonics transfer to the inverter dc side.
Figure 6. AC harmonics transfer to the inverter dc side.
Hysteresis inverter operating with distorted system voltages.
Figure 7. Hysteresis inverter operating with distorted system voltages.
Inverter ac outputs under fixed-band hysteresis control.
Figure 8. Inverter ac outputs under fixed-band hysteresis control.
Simulation results obtained from the developed model and Simulink.
Figure 9. Simulation results obtained from the developed model and Simulink.
Ea(f)| when Vc is decreased to 350V.
Figure 10. Ea(f)| when Vc is decreased to 350V.
|Ea(f)| when the system inductance is decreased by 25%.
Figure 11. |Ea(f)| when the system inductance is decreased by 25%.
Results from reducing ε by 50%.
Figure 12. Results from reducing ε by 50%.
Effect of injecting the 11th ac harmonic voltage on the inverter ac outputs.
Figure 13. Effect of injecting the 11th ac harmonic voltage on the inverter ac outputs.
Instantaneous frequency of ea(t) when vs is distorted.
Figure 14. Instantaneous frequency of ea(t) when vs is distorted.
Error spectrum when vsa contains the 11th oscillator voltage.
Figure 15. Error spectrum when vsa contains the 11th oscillator voltage.
Effect of injecting the 8th dc harmonic voltage on the inverter ac outputs.
Figure 16. Effect of injecting the 8th dc harmonic voltage on the inverter ac outputs.
Frequency of ea(t) when the input dc is distorted.
Figure 17. Frequency of ea(t) when the input dc is distorted.
Error spectrum when the 8th dc background distortion exists.
Figure 18. Error spectrum when the 8th dc background distortion exists.
Instantaneous outputs of the variable-band hysteresis control.
Figure 19. Instantaneous outputs of the variable-band hysteresis control.
Comparing model results to Simulink.
Figure 20. Comparing model results to Simulink.
Error spectra when Vc=350 V.
Figure 21. Error spectra when Vc=350 V.
Error spectra when relay bandwidth is halved.
Figure 22. Error spectra when relay bandwidth is halved.
Inverter harmonic response to 25% reduction in L.
Figure 23. Inverter harmonic response to 25% reduction in L.
DC current harmonics under variable-band control.
Figure 24. DC current harmonics under variable-band control.
Error current under distorted dc and ac system voltages.
Figure 25. Error current under distorted dc and ac system voltages.
A1. Approximation of f(x)=sin(πx)x(1−x) by f¯(x)=(4−π)⋅sin(πx)+π.
Figure 26. A1. Approximation of f(x)=sin(πx)x(1−x) by f¯(x)=(4−π)⋅sin(πx)+π.

Modeling & Simulation of Hysteresis Current Controlled Inverters Using MATLAB

Ahmad Albanna1

1. Introduction

Hysteresis inverters are used in many low and medium voltage utility applications when the inverter line current is required to track a sinusoidal reference within a specified error margin. Line harmonic generation from those inverters depends principally on the particular switching pattern applied to the valves. The switching pattern of hysteresis inverters is produced through line current feedback and it is not pre-determined unlike the case, for instance, of Sinusoidal Pulse-Width Modulation (SPWM) where the inverter switching function is independent of the instantaneous line current and the inverter harmonics can be obtained from the switching function harmonics.

This chapter derives closed-form analytical approximations of the harmonic output of single-phase half-bridge inverter employing fixed or variable band hysteresis current control. The chapter is organized as follows: the harmonic output of the fixed-band hysteresis current control is derived in Section 2, followed by similar derivations of the harmonic output of the variable-band hysteresis controller in Section 3. The developed models are validated in Section 4 through performing different simulations studies and comparing results obtained from the models to those computed from MATLAB/Simulink. The chapter is summarized and concluded in section 5.

2. Fixed-band hysteresis control

2.1. System description

Fig.1 shows a single-phase neutral-point inverter. For simplicity, we assume that the dc voltage supplied by the DG source is divided into two constant and balanced dc sources, as in the figure, each of valueVc. The RL element on the ac side represents the combined line and transformer inductance and losses. The ac source vsa represents the system voltage seen at the inverter terminals. The inverter line currentia, in Fig.1, tracks a sinusoidal reference ia*=2Ia*sin(ω1t+ϕ) through the action of the relay band and the error currentea(t)=ia*ia.

In Fig.2, the fundamental frequency voltage at the inverter ac terminals when the line current equals the reference current is the reference voltage,va*=2Va*sin(ω1t+θ). Fig.2 compares the reference voltage to the instantaneous inverter voltage resulting from the action of the hysteresis loop.

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Figure 1.

Single-phase half-bridge inverter with fixed-band hysteresis control.

Referring to Fig.2, when valve Q is turned on, the inverter voltage isva=Vc>va*; this forces the line current ia to slope upward until the lower limit of the relay band is reached atea(t)=ε. At that moment, the relay switches on Q and the inverter voltage becomesva=Vc<va*, forcing the line current to reverse downward until the upper limit of the relay band is reached atea(t)=ε.

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Figure 2.

Reference voltage calculation and the instantaneous outputs.

The bang-bang action delivered by the hysteresis-controlled inverter, therefore, drives the instantaneous line current to track the reference within the relay band(ε,ε). With reference to Fig. 3 and Fig. 4, the action of the hysteresis inverter described above produces an error current waveform ea(t) close to a triangular pulse-train with modulating duty cycle and frequency.

2.2. Error current mathematical description

The approach described in this section closely approximates the error current produced by the fixed-band hysteresis action, by a frequency-modulated triangular signal whose time-varying characteristics are computed from the system and controller parameters. Subsequently, the harmonic spectrum of the error current is derived by calculating the Fourier transform of the complex envelope of frequency modulated signal.

Results in the literature derived the instantaneous frequency of the triangular error current fia(t) in terms of the system parameters (R=0). Using these results and referring to Fig. 3 (Albanna & Hatziadoniu, 2009, 2010):

t1=2εLVc[1+Msin(ω1t+θ)],t2=2εLVc[1Msin(ω1t+θ)]
(1)

and therefore:

fia(t)=1T=fc+VcM28εLcos(2ω1t+2θ)
(2)

where the average switching (carrier) frequency fc is given by

fc=Vc4εL[1M22]
(3)

and M is the amplitude modulation index of the inverter expressed in terms of the peak reference voltage and the dc voltage as:

M=2Va*Vc
(4)
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Figure 3.

Detail ofea(t).

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Figure 4.

Effect of va* on the error current duty cycle.

Examining (Eq. 2), the instantaneous frequency fia(t) of the error current ea(t) consists of the carrier frequency fc and a modulating part that explicitly determines the bandwidth of the error current spectrum, as it will be shown later in this chapter. Notice that the modulating frequency is twice the fundamental frequency, that is,2f1.

Now, with the help of Fig.3, we define the instantaneous duty cycle of the error current D(t) as the ratio of the rising edge time t1 to the instantaneous periodT. Noting thatD(t)=t1fia(t), we obtain after using (Eq. 1), (Eq. 2) and manipulating,

D(t)=0.50.5Msin(ω1t+θ)
(5)

Implicit into (Eq. 3) is the reference voltageva*. The relation between the instantaneous duty cycle and the reference voltage can be demonstrated in Fig.4: the duty cycle reaches its maximum value at the minimum ofva*; it becomes 0.5 (symmetric form) at the zero ofva*; and it reaches its minimum value (tilt in the opposite direction) at the crest ofva*. Next, we will express ea(t) by the Fourier series of a triangular pulse-train having an instantaneous duty cycle D(t) and an instantaneous frequencyfia(t):

ea(t)=n=12επ2(1)nn2sin[nπ(1D(t))]D(t)[1D(t)]sin(2πn0tfia(τ)dτ)
(6)

As the Fourier series of the triangular signal converges rapidly, the error current spectrum is approximated using the first term of the series in (Eq. 6). Therefore truncating (Eq. 6) to n=1 and using (Eq. 2) yields

ea(t)2επ2sin[πD(t)]D(t)[1D(t)]sin(ωct+βsin(2ω1t+2θ)φ)
(7)

whereφ=βsin(2θ). The frequency modulation index

β=VcM28εL12f1
(8)

determines the frequency bandwidth

BW=4(β+1)f1
(9)

that contains 98% of the spectral energy of the modulated sinusoid in (Eq. 7). To simplify (Eq. 7) further, we use the following convenient approximation (see Appendix-A for the derivation): Given that, 0D(t)1, then

sin[πD(t)]D(t)[1D(t)](4π)sin[πD(t)]+π
(10)

Therefore (Eq. 7) becomes,

ea(t)2επ2[π+(4π)sin(πD(t))]sin(ωct+βsin(2ω1t+2θ)φ)
(11)

Substituting D(t) from (Eq. 5) into (Eq. 11) and manipulating, we obtain

ea(t)2επ2(π+(4π)cos(πM2sin(ω1t+θ)))sin(ωct+βsin(2ω1t+2θ)φ)
(12)

Next, the cosine term in (Eq. 12) is simplified by using the infinite product identity and truncating to the first term. That is,

cos(x)=n=1(1x2π2(n0.5)2)14x2π2
(13)

Substituting (Eq. 13) into (Eq. 12) and manipulating, the error current approximation becomes:

ea(t)[(8k)+kcos(2ω1t+2θ)]e1(t)[επ2sin(ωct+βsin(2ω1t+2θ)φ)]e2(t)
(14)

wherek=(4π)M2. The harmonic spectrum Ea(f) of the error current is the convolution of the spectra of the product terms e1(t) and e2(t) in (Eq. 14). Therefore,

Ea(f)=[(8k)δ(f)+k2ej2θδ(f2f1)+k2ej2θδ(f+2f1)]E2(f)
(15)

where denotes convolution. In order to calculateE2(f), we rewrite e2(t) as

e2(t)=εj2π2(ejωctejφejβsin(2ω1t+2θ)ejωctejφejβsin(2ω1t+2θ))
(16)

The positive frequency half of the spectrum E2(f) is therefore given by

E2(f)=εj2π2n=Jn(β)ej(2nθφ)δ(fc+2nf1)
(17)

where δx=δ(fx) is the Dirac function, and Jn is the Bessel function of the first kind and ordern. Substituting (Eq. 17) into (Eq. 15), and convoluting, we obtain:

Ea(f)=εj2π2n=(k2[Jn1(β)+Jn+1(β)]+(8k)Jn(β))ej(2nθφ)δ(fc+2nf1)
(18)

Using the recurrence relation of the Bessel functions,

Jn+1(β)+Jn1(β)=2nβJn(β),
(19)

the positive half of the error current spectrum takes the final form:

Ea(f)=n=n=Enej(2nθφ)δ(fc2nf1)
(20)

where,

En=εj2π2(knβk+8)Jn(β)
(21)
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Figure 5.

Effect of changing β on the harmonic spectrum.

The calculation of the non-characteristic harmonic currents using (Eq. 20) is easily executed numerically as it only manipulates a single array of Bessel functions. The spectral energy is distributed symmetrically around the carrier frequency fc with spectrum bands stepped apart by2f1. Fig.5 shows the harmonic spectrum of the error current as a function of the frequency modulation indexβ. If the operating conditions of the inverter forces β to increase toβ, then the spectral energy shifts to higher carrier frequencyfc. Additionally, as the average spectral energy is independent of β and depends on the error bandwidthε, the spectral energy spreads over wider range of frequencies, 4(β+1)f1, with an overall decrease in the band magnitudes to attain the average spectral energy at a constant level as shown in Fig.5. The Total Harmonic Distortion (THD) of the line current is independent of β and is directly proportional to the relay bandwidthε.

2.3. Model approximation

The harmonic model derived in the previous section describes the exact spectral characteristics of the error current by including the duty cycle D(t) to facilitate the effect of the reference voltage va* on the error current amplitude and tilting. Moreover, the consideration of D(t) in (Eq. 6) predicts the amplitude of the error current precisely, which in turn, would result in accurate computation of the spectrum bands magnitudes according to (Eq. 20). The model can be further simplified to serve the same functionality in without significant loss of numerical accuracy. As the instantaneous frequency of the error current, given by (Eq. 2), is independent ofD(t), the spectral characteristics such as fc and BW are also independent of D and therefore, setting D(t) to its average value 0.5 will slightly affect the magnitude of the spectrum bands according to (7). Subsequently, the error current harmonic spectrum simplifies to

Ea(f)=4εjπ2n=n=Jn(β)ej(2nθφ)δ(fc2nf1)
(22)

where the carrier (average) frequency fc is given by (Eq. 3), the frequency modulation index β is given by (Eq. 8). The 3 dB frequency bandwidth BW that contains 98% of the spectral energy is given by (Eq. 9).

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Figure 6.

AC harmonics transfer to the inverter dc side.

2.4. Dc current harmonics

The hysteresis switching action transfers the ac harmonic currents into the inverter dc side through the demodulation process of the inverter. As the switching function is not defined for hysteresis inverters, the harmonic currents transfer can be modeled through balancing the instantaneous input dc and output ac power equations.

With reference to Fig. 1, and assuming a small relay bandwidth (i.e.ia*ia), the application of Kirchhoff Current Law (KCL) at node a gives:

id1=ia*+id2
(23)

The power balance equation over the switching period when Q is on is given by:

id2=1Vcva(t)ia(t)
(24)

Using the instantaneous output voltage

va=va*L(ddtea)
(25)

in (Eq. 24), the dc current id1 will have the form:

id1(t)=va*ia*Vc+ia*LVcia*(ddtea)
(26)

where x˙ is the derivative of x with respect to time. Using the product-to-sum trigonometric identity and simplifying yields:

id1(t)=2MIa*2cos(θϕ)+ia*2MIa*2cos(2ω1t+θ+ϕ)LVce˙aia*
(27)

The positive half of the dc current spectrum is thus computed from the application of the Fourier transform and convolution properties on (Eq. 27), resulting in

Id1(f)=I0δ0+I1δf1+I2δ2f1+fIh(Ea(f+f1)Ea(ff1))
(28)

where Ea(f) is the error current spectrum given by (Eq. 22). The average, fundamental, and harmonic components of the dc current spectrum are respectively given by

I0=22MIa*cos(θϕ),I1=2j2Ia*ejϕ,I2=24MIa*ej(θϕ),andIh=2πVcLIa*.
(29)

Each spectrum band of the ac harmonic current creates two spectrum bands in the dc side due to the convolution process implicitly applied in (Eq. 28). For instance, the magnitude of the ac spectrum band at fc is first scaled by fc according to (Eq. 28) then it is shifted by ±f1 to create the two dc bands pinned at fc±f1 as shown in Fig.6. Consequently, every two successive bands in the ac spectrum create one corresponding dc spectrum band that is located half the frequency distance between the two ac bands.

2.5. Harmonic generation under distorted system voltages

The harmonic performance of the hysteresis inverter in Fig. 7 under distorted dc and ac system voltages is analyzed. The presence of background harmonics in the ac and dc voltages will affect the instantaneous frequency of the inverter according to (Eq. 30) as

fia(t)=Vc4εL[1(va*Vc+vhVcvkVc)2]
(30)

where the dc distortionvk, and the distortion of the ac system voltage, vh, are given as:

vk=2Vksin(kω1t+γk)vh=2Vhsin(hω1t+γh)
(31)
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Figure 7.

Hysteresis inverter operating with distorted system voltages.

Notice that in (Eq. 31), kand h need not be integers. Substituting (Eq. 31) in (Eq. 30) and assuming small distortion magnitudes, the instantaneous frequency of the error current ea simplifies to:

fia(t)=fia(t)Δfac(t)+Δfdc(t)
(32)

where fia(t) is given by (Eq. 2) and

Δfac(t)=Vc4εLMMh[cos([h1]ω1t+γhθ)cos([h+1]ω1t+γh+θ)],Δfdc(t)=Vc4εLMMk[cos([k1]ω1t+γkθ)cos([k+1]ω1t+γk+θ)],
(33)

are the frequency noise terms due to the system background distortions. The amplitude modulation indices of the ac and dc harmonic distortions are given by :

Mh=2VhVc,andMk=2VkVc
(34)

Integrating (Eq. 32), the error current ea(t) is thus approximated by the frequency-modulated sinusoid:

ea(t)=8επ2sin[ωct+βsin(2ω1t+2θ)φ+ψac+ψdc]
(35)

In (Eq. 35): the carrier frequency fc is given by (Eq. 3); the frequency modulation index β is given by (3);φ=βsin(2θ); and

ψac=βh+sin([h+1]ω1t+γh+θ)βhsin([h1]ω1t+γhθ)+φac,ψdc=βksin([k1]ω1t+γkθ)βk+sin([k+1]ω1t+γk+θ)+φdc.
(36)

whereφac=βhsin(γhθ)βh+sin(γh+θ),andφdc=βk+sin(γk+θ)βhsin(γkθ). The corresponding ac and dc frequency modulation indices are given by

βh+=Vc4εLMMh1(h+1)f1;βh=Vc4εLMMh1(h1)f1;βk+=Vc4εLMMk1(k+1)f1;βk=Vc4εLMMk1(k1)f1.
(37)

Applying the Fourier transform and convolution properties on (Eq. 35), the positive half of the frequency spectrum Ea(f) simplifies to:

Ea(f)=Ea(f)ach(f)dck(f)
(38)

Where Ea(f) is given by (Eq. 22) and

ach=ejφac(n=Jn(βh+)ejn(γh+θ)δ(h+1)nf1)(n=Jn(βh)ejn(γhθ)δ(h1)nf1),dck=ejφac(n=Jn(βk+)ejn(γk+θ)δ(k+1)nf1)(n=Jn(βk)ejn(γkθ)δ(k1)nf1),
(39)

are the ac and dc modulating spectra. Generally, for any H number of ac voltage distortions and K number of dc distortions, (Eq. 40) is applied first to calculate the total ac and dc modulating spectra, then (Eq. 38) is used to compute the error current harmonic spectrum.

acH(f)=Hach,  dcK(f)=Kdck.
(40)

3. Variable-band hysteresis control

3.1. Error current mathematical description

The harmonic line generation of the half-bridge inverter of Fig.1 under the variable-band hysteresis current control is derived. The constant switching frequency of the error current in (Eq. 2), i.e.fia(t)=fo, is achieved by limiting the amplitude of the error current to stay within the variable band [54, 55]:

εa(t)=εo[10.5M2+0.5M2cos(2ω1t+2θ)]
(41)

where the maximum value of the modulating relay bandwidth is

εo=Vc4Lfo
(42)

and fo is the target switching frequency. Subsequently, the error current is approximated by the amplitude-modulated sinusoid of frequency fo as:

ea(t)=8π2εa(t)sin(2πfot)
(43)

Substituting (Eq. 41) in (Eq. 43) and then applying the Fourier transform, the positive half of the frequency spectrum of Ea(f) is:

Ea(f)=4εojπ2[(1M22)δfo+M24(ej2θδ(fo2f1)+ej2θδ(fo+2f1))]
(44)

The error current spectrum in (Eq. 44) consists of a center band at the switching frequency fo and two side bands located atfo±2f1. The frequency bandwidth that contains the spectral energy of (Eq. 44) is simply4f1.

3.2. Dc current harmonics

The approach developed in 2.2.4 also applies to compute the dc current harmonic spectrum when the variable-band hysteresis control. The positive half of the dc current harmonic spectrum is computed by substituting (Eq. 44) in (Eq. 28).

3.3. Harmonic generation under distorted system voltages

The presence of background harmonics in the ac and dc voltages, given in (Eq. 31) will affect the instantaneous frequency of the inverter according to (Eq. 30). Subsequently, to achieve the constant switching frequencyfo, the modulating error band in (Eq. 41) will also contain the corresponding distortions terms as

εa(t)=εao(t)εaac(t)+εadc(t)
(45)

where εao(t) is the error under zero background distortion given by (Eq. 41), and

εaac(t)=2εoMMhsin(hω1t+γh)sin(ω1t+θ),εadc(t)=2εoMMksin(kω1t+γk)sin(ω1t+θ),
(46)

where MhandMk define the modulation index of the ac and dc background distortion terms respectively as (Eq. 34).

The new terms introduced by the background distortion appear as amplitude modulations in (Eq. 45). The error current ea(t) is then expressed as:

ea(t)=8π2[εao(t)εaac(t)+εadc(t)]sin(2πfot)
(47)

The harmonic spectrum of the error current Ea(f) simplifies to

Ea(t)=Ea(f)Eaac(f)+Eadc(f)
(48)

where Ea(f) is the zero-background-distortion error as in (Eq. 44), and the new terms due to background distortion:

Eaac=2MMhεojπ2[ej(γhθ)δfo+(h1)f1+ej(γhθ)δfo(h1)f1 ej(γh+θ)δfo+(h+1)f1ej(γh+θ)δfo(h+1)f1],Eadc=2MMkεojπ2[ej(γkθ)δfo+(k1)f1+ej(γkθ)δfo(k1)f1 ej(γk+θ)δfo+(k+1)f1ej(γk+θ)δfo(k+1)f1].
(49)

Examining (Eq. 49), the presence of the harmonic distortions in the system tends to scatter the spectrum over lower frequencies, more specifically, tofo(h+1)f1, for h>kor to fo(k+1)f1 fork>h.

4. Simulation

The harmonic performance of the half-bridge inverter under the fixed- and variable-band hysteresis control is analyzed. Results computed from the developed models are compared to those obtained from time-domain simulations using MATLAB/Simulink. Multiple simulation studies are conducted to study the harmonic response of the inverter under line and control parameter variations. The grid-connected inverter of Fig.1 is simulated in Simulink using:Vc=400V, Vsa=120Vrms, f1=60Hz, R=1.88Ω, andL=20mH. In order to limit the THD of the line current to 10%, the line current tracks the sinusoidal reference ia*=215sin(ω1t)A within the maximum relay bandwidth ofε=εo=2.82A.

4.1. Fixed-band hysteresis current control

The ac outputs of the half-bridge inverter under the fixed-band hysteresis current control are shown in Fig.8. the fundamental component va* of the bipolar output voltage va has a peak value of 263.7 V. the inverter line current ia tracks the sinusoidal reference within an absolute error marginε. The error current resulting from the fixed-band hysteresis action resembles a frequency-modulate triangular signal of constant amplitude. The implicit relation between the error current duty cycle and the reference voltage va* is clearly seen in Fig.8. The symmetric duty cycle, i.e.D=0.5, happens whenever the reference voltage approaches a zero crossing.

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Figure 8.

Inverter ac outputs under fixed-band hysteresis control.

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Figure 9.

Simulation results obtained from the developed model and Simulink.

The harmonic parameters of the model are computed the system and controller parameters as follows: substituting the reference voltage in (Eq. 4) results in an amplitude modulation index ofM=0.659; from (Eq. 3), the carrier frequency isfc=23.05f1=1383Hz; and from (Eq. 8), the frequency modulation index is β=3.2 3.2. Fig. 9 compares the harmonic spectrum of the error current Ea(f) computed from (20) to that obtained from the Fourier analysis of the time-domain simulation results using Simulink. The figure shows a good agreement between the two spectra in terms of frequency order, magnitude and angle.

The spectrum bands are concentrated around the order of the carrier frequency and are stepped apart by two fundamental frequency orders 2f1 as shown in Fig. 9. With reference to (Eq. 9) and Fig. 9, it is shown that 98% of the spectrum power is laying in the bandwidthBW=4(β+1)f116f1. Therefore, the spectrum bands outside this range contribute insignificantly to the total spectrum power and thus can be truncated from the spectrum for easier numerical applications.

To study the effect of line parameter variations on the harmonic performance of the inverter, the DG source voltage is decreased to have the dc voltageVc=350V, then the harmonic spectrum is recomputed using the model and compared to the results obtained from Simulink. Decreasing Vc will increase M and β according to (Eq. 4) and (Eq. 8) respectively, but will decrease fc according to (Eq. 3).

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Figure 10.

Ea(f)| when Vc is decreased to 350V.

With reference to the results shown in Fig.10, the harmonic spectrum Ea(f) will shift to the lower frequency order of, approximately, 18, and will span a wider range, as β is greater. The frequency bandwidth has slightly increased to 18f1 from the previous value of 16f1 due to the slight increase in β=3.2 toβ=3.66.

The total spectral energy of the error current depends on the relay bandwidth ε and it is independent ofβ. As β increases the spectrum energy redistributes such that the bands closer to fc decrease in magnitude and those that are farther from fc increase as shown in Fig. 10. The Total Harmonic Distortion (THD) of the line current thus will not be affected by changingVc.

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Figure 11.

|Ea(f)| when the system inductance is decreased by 25%.

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Figure 12.

Results from reducing ε by 50%.

Next, the system and control parameters are set to their original values and the inductance is decreased by 25% toL=15mH. The results are shown in Fig. 11. Lower inductance results in higher switching frequency according to (Eq. 3) and higher β according to (Eq. 8). The harmonic spectrum Ea(f) shifts to higher frequencies as fc is increasing, and the spectrum spans a wider range as β is increasing. The amplitude modulation index M and D are affected by the system inductance variation since the inverter reference voltage va* depends on system inductanceL.

The width of the relay band is reduced by half while maintaining the rest of the parameters at their base values. As (Eq. 4) indicates, Mis independent of ε and thus it remains unchanged from its value of 0.659. Referring to Fig. 12, as the error band is reduced by half, the carrier frequency doubles and the harmonic spectrum Ea(f) will be concentrated around, approximately, the order of 46. The frequency modulation index β doubles and thus the spectrum spreads over a wider frequency range overall decreasing in magnitude, as seen in Fig. 12. Under these conditions, the THD of the line current will decrease to approximately 5% as the spectral energy of the spectrum is proportional to the relay bandwidthε.

To study the harmonic performance of the inverter under distorted system voltages, the system and control parameters are set to the original values and the 11th order voltage oscillator v11(t)=15sin(11ω1t)V is included in the source voltage vs to simulate a distorted ac network voltage. The simulation is run for 30 fundamental periods to ensure solution transients are vanishing, and the last fundamental period of the inverter ac outputs are shown in Fig.13.

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Figure 13.

Effect of injecting the 11th ac harmonic voltage on the inverter ac outputs.

Comparing Fig. 8 and Fig. 13, the reference voltage is distorted due to the presence of the 11th voltage oscillator in the source. The output voltage of the inverter is still bipolar, i.e.va=±400V. Fig. 14 compares the instantaneous frequency of the error current under sinusoidal ac voltage fia to that under the distorted ac system voltagefia.

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Figure 14.

Instantaneous frequency of ea(t) when vs is distorted.

According to (Eq. 32), the carrier frequency fc=23.05f1 is constant and independent of the distortion terms. The amplitude modulation index M11=0.038 is computed from (Eq. 34), subsequently, the harmonic parameters β11+=0.062 and β11=0.074 are computed from (Eq. 37).

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Figure 15.

Error spectrum when vsa contains the 11th oscillator voltage.

Fig. 15 compares the harmonic spectrum Ea(f) obtained from (Eq. 38) to that computed from the Fourier analysis of Simulink outputs with very good agreement in terms of frequency order and magnitude. The spectral energy is centered on the carrier frequency fc=23.05f1 with spectrum bands are stepped apart by2f1. The frequency bandwidth increases due to the distortion terms, and as Fig. 15 shows, the spectrum bands leaks to as low of a frequency order as 5. Notice that the THD of the line current did not change as the controller bandwidth did not change.

Similar analysis is performed to study the harmonic performance of the inverter when the dc voltage contains the distortionv8(t)=28.2sin(8ω1t)V. The inverter instantaneous outputs obtained from Simulink are shown in Fig.16. Notice that the voltage va is still bipolar but distorted.

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Figure 16.

Effect of injecting the 8th dc harmonic voltage on the inverter ac outputs.

The dc distortions impose additional noise component on the instantaneous frequency, see Fig.17, and subsequently, according to (Eq. 38) the harmonic spectrum is drifting to lower order harmonics as shown in Fig.18.

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Figure 17.

Frequency of ea(t) when the input dc is distorted.

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Figure 18.

Error spectrum when the 8th dc background distortion exists.

4.2. Variable-band hysteresis control

The harmonic performance of the same half-bridge inverter used in section 2.4.1 is analyzed when the variable-band hysteresis current control is employed. Similar harmonic studies to those in the previous section are performed to compute the spectral characteristics of the inverter harmonic outputs using the developed models in section 2.3 and compare them with results obtained from time-domain simulations using Simulink.

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Figure 19.

Instantaneous outputs of the variable-band hysteresis control.

The instantaneous line outputs of the single-phase inverter operating under variable hysteresis control are shown in Fig.19. With the maximum relay band εo is set to 2.82, the error current ea(t) resulting from the variable-band control is an amplitude-modulated triangular signal of carrier frequencyfo. Regardless of the adopted switching pattern, the reference voltage is va*=263.7sin(ω1t+37°)Vand hence,M=0.659. From (Eq. 42), the average frequency isfo=29.4f1. Fig.20 compares the spectrum Ea(f) computed from (Eq. 44) to that computed from the harmonic analysis of time-domain simulation of the inverter using Simulink. The figure shows a good agreement between the two spectra in terms of frequency order and magnitude. The center band is located at fo=29.4f1 and the side bands are stepped by 2f1 as shown in Fig.20. The spectral energy of Ea(f) is distributed over the frequency range 27.4f1 to 31.4f1 (i.e.BW=4f1).

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Figure 20.

Comparing model results to Simulink.

The dc voltage Vc was decreased to 350V while all other parameters remain unchanged from Study 1. Decreasing Vc will decrease fo according to (Eq. 42).

The new values are shown in Fig.21. Consequently, the spectrum Ea(f) will shift to the lower frequency order of, approximately, 25.7, while spanning over the constant bandwidth of4f1. The spectral magnitudes of Ea(f) depend on the relay bandwidth εo andM; therefore, with fixing εo and decreasingVc, according to (Eq. 44), the center band magnitude decreases as M is increasing. While the magnitudes of the side bands are directly proportional toM, their magnitudes will increase. This is clear from comparing the harmonic in Fig.21 to that of Fig.20. Similar to the fixed-band control, the Total Harmonic Distortion (THD) of the line current is independent ofVc.

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Figure 21.

Error spectra when Vc=350V.

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Figure 22.

Error spectra when relay bandwidth is halved.

when εois halved, the carrier frequency fo doubles and the harmonic spectrum Ea(f) will be concentrated around, approximately, the order of 59. The THD of the line current will decrease to as low as 5% since εo decreases. This is demonstrated when comparing the harmonic spectra of Fig.22 and Fig.20.

The value of the inductance is decreased toL=15mH. The results are shown in Fig.23.

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Figure 23.

Inverter harmonic response to 25% reduction in L.

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Figure 24.

DC current harmonics under variable-band control.

Lower inductance results in higher switching frequency. The harmonic spectrum Ea(f) shifts to higher frequencies as fo is increasing to39.2f1. As M is directly proportional to the system inductance, Mdecreases and therefore, the magnitude of the center band slightly increases while the side bands decrease in magnitude as shown in Fig.23. The dc current harmonics are computed from substituting (Eq. 44) in (Eq. 28). The resulting spectra are shown in Fig.24 with good agreement in terms of frequency orders and magnitudes.

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Figure 25.

Error current under distorted dc and ac system voltages.

The harmonic performance of the inverter under distorted system voltages is studied by simulating the system with the distorted 8th order dc voltage v8(t)=28.2sin(8ω1t)V and the 11th order ac voltagev11(t)=15sin(11ω1t)V. Results obtained from model using (Eq. 48) and (Eq. 49) are compared to those computed from Simulink in Fig.25, the model predicts the frequency distribution of the dc current harmonics and accurately predicts their magnitudes.

4.3. Comparison and discussion

The spectral characteristics of the line current under the fixed- and variable-band hysteresis control are compared in this section. For identical system configurations and controller settings, i.e.ε=εo, the analytical relation between fc and fo is stated in terms of the amplitude modulation index M as:fc=(10.5M2)fo. The inverter operates at higher switching frequency when it employs the variable-band hysteresis control. In addition, from a harmonic perspective, the frequency bandwidth of Ea(f) in the variable-band control mode is constant (4f1) and independent of the system and controller parameters; unlike the fixed-band controller where the bandwidth BW depends implicitly on the system and controller parameters through the frequency modulation indexβ.

The THD of the line current is directly proportional to relay bandwidth. For similar controllers settingε=εo, the THD is constant as the average spectral energy of the line current is constant. In fixed- and variable-band modes, the variation of system parameters shifts the spectral energy of Ea(f) to higher or lower frequency orders (depending on the carrier frequency), while simultaneously redistributing the spectral energy over the frequency bandwidthBW. The spectral energy of the error current is independent of system parameters; and hence, the THD of the line current is constant for different system settings.

5. Conclusion

A closed-form numerically efficient approximation for the error current harmonic spectrum of single-phase two-level inverters employing either fixed- or variable-band hysteresis current control is derived. The models are based on the amplitude and frequency modulation theorems.

The instantaneous frequency of the inverter is first derived. Then it is used to closely approximate the error current by a modulated sinusoid. The error current harmonic spectrum is basically the Fourier transform of error current complex envelop. In the case of the fixed-band control, the spectrum reduces to a series of Bessel functions of the first kind whose argument is implicitly expressed in terms of the system and controller parameters, where as in the variable-band mode, the spectrum reduces to a 3-element array.

The spectral characteristics such as the carrier frequency and frequency bandwidth are derived analytically and related to line parameters; it is a development useful in inverter-network harmonic interactions. Unlike time-domain simulators, the developed models provide fast numerical solution of the harmonic spectrum as they only involve numerical computation of single arrays. Simulation results agree closely with the developed frequency-domain models in terms of frequency order, magnitude and angle.

In addition to the single-phase two-level inverter, the proposed approximations apply also to the harmonic output of certain three-phase two-level inverters where independent phase control is applicable, such as the neutral point inverter, and the full-bridge inverter in bipolar operation.

6. Future directions of research

The models detailed in this chapter can be extended in a number of ways, both in terms of improving the proposed models as well as in the application of the models in other PWM applications.

The developed models neglected the dynamics of the Phase-Locked Loop (PLL) and assumed that the inverter line current tracks a pure sinusoidal reference current. Possible extensions of the models include the effect of the harmonic current propagation through the ac network and the deterioration of the terminal voltage at the interface level and its effect on the reference current generation. As the PLL synchronizes the reference current with the terminal voltage, the propagation of harmonic currents might affect the detection of the zeros-crossings of the terminal voltage resulting in generating a distorted reference current. The hysteresis controller consequently will force the line current to track a non-sinusoidal reference which, in turn, modifies the harmonic output of the inverter.

The implementation of an LC filter at the inverter ac terminals could trigger a parallel-resonance which tends to amplify the harmonic voltages and currents in the ac network leading, in some cases, to potential harmonic instabilities. The improvement of the developed models to include the effect the filter capacitance on the harmonic performance of the inverter is an interesting improvement.

Reviews of the developed models show that hysteresis current controlled inverters can have a ‘switching function’ notation similar to those inherit with the Sinusoidal PWM inverters. The switching function is based on the error current characteristics which implicitly depend on the system and controller parameters. Such development will enable the various time- and frequency-domain algorithms developed for the harmonic assessment of linear PWM inverters to be applied to hysteresis controlled inverters.

Harmonic load flow studies of systems incorporating inverters with hysteresis current control can be formulated based on the developed models. The iterative solution of the harmonic load flow shall incorporate the harmonic magnitudes and angles obtained from the developed models for a faster convergence to the steady state solution.

7. Appendix - A

Functionf(x)=sin(πx)x(1x), x[0,1]satisfieslimx0f(x)=limx1f(x)=π, is continuous, convex and has even symmetry in[0,1]. The approximation of f(x) in[0,1],f¯(x)=Asin(πx)+π ,

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Figure 26.

A1. Approximation of f(x)=sin(πx)x(1x) by f¯(x)=(4π)sin(πx)+π.

satisfies the same properties in[0,1]. Subsequently, constant A is calculated such that the square error over the interval [0,1] Δf2=01(f(x)f¯(x))2dx is minimized. Substituting the expressions of f(x) and f¯(x) into the mean-squared-error Δf2 and evaluating the integrals numerically yields a function of the parameterA, asΔf2(A)=0.5A20.858A+0.3833. Therefore, the value of A that minimizes Δf2 is A=0.858 orA=4π. Functions f(x) and f¯(x) are shown in Fig. 1.

References

1 - A. Albanna, C. J. Hatziadoniu, 2009Harmonic Analysis of Hysteresis Controlled Grid-Connected Inverters, proceedings of the Power Systems Conference and Exposition, Seattle, WA, March, 2009.
2 - A. Albanna, C. J. Hatziadoniu, 2009Harmonic Modeling of Single-Phase Three-Level Hysteresis Inverters, proceedings of the North American Power Symposium, Starkville, Mississippi, October, 2009
3 - A. Albanna, C. J. Hatziadoniu, 2009Harmonic Modeling of Three-Phase Neutral-Point Inverters, proceedings of the North American Power Symposium, Starkville, Mississippi, October, 2009
4 - A. Albanna, C. J. Hatziadoniu, 2010Harmonic Modeling and Analysis of Multiple Residential Photo-Voltaic Generators, proceedings of the Power and Energy Conference at Illinois, Urbana-Champaign, Illinois, February 2010
5 - A. Albanna, C. J. Hatziadoniu, 2010Harmonic Modeling of Hysteresis Inverters in Frequency Domain”, IEEE Transactions on Power Electronics, May 2010