Control of an active rectifier with an based algorithm

Control of an active rectifier with an based algorithm
Control of an active rectifier with an
inductive-capacitive-inductive filter using a Twisting
based algorithm
Arnau Dòria-Cerezo
Paul F. Puleston
Cristian Kunusch
Dept. Electrical Eng. and
Inst. Industrial and Control Eng.
Polytechnical University of Catalonia
08028 Barcelona, Spain
Email: arnau.doria@upc.edu
CONICET and LEICI
National University of La Plata
CC91, 1900, La Plata, Argentina
Marie Curie Fellow, IRI, CSIC-UPC
08028 Barcelona, Spain
Email: puleston@ing.unlp.edu.ar
Inst. de Robòtica i Informàtica Industrial
Consejo Superior de Inv. Cientı́ficas
Polytechnical University of Catalonia
08028 Barcelona, Spain
Email: ckunusch@iri.upc.edu
Abstract—This paper presents a novel controller for an
active rectifier with an inductive-capacitive-inductive filter. The
proposed control scheme comprises two levels. The internal
level, is a current controller based on the second order sliding
mode Twisting algorithm, which robustly ensures an unity power
factor at the connection point. The slower external level is a PI
controller in charge of regulating the output DC voltage to a
desired reference. The controller setup also includes a finite-time
observer of the current derivative that can be used to avoid direct
(and sometimes problematic) numerical differentiation. Finally,
simulation results are presented to validate the proposed control
scheme.
I.
I NTRODUCTION
The use of active rectifiers has beed widely increased in the
last decades thanks to its ability to transform the AC electrical
power to a DC power. The advantages of using these power
converters are the power factor correction and its robustness to
guarantee the DC bus voltage for a large range of load values.
However, the commutation of the set of switches implies an
undesired current ripple of the inductor filter. This effect can be
considerably reduced using an inductive-capacitive-inductive
(LCL) filter instead of the traditional first order inductance
filter, and low total harmonic distortion (THD) of the current
can be attained.
The key goal of a single-phase active rectifier is to ensure
a given DC voltage with an unity power factor on the grid
side. These objectives are usually fulfilled by using control
algorithms that measure some of the variables (normally,
voltage and current on the grid side and the DC voltage)
and process that information to provide a suitable switching
policy to reach the desired set points. Several control techniques, from classical proportional-integral-derivative (PID)
controllers to more advanced nonlinear methodologies, have
been applied to different power converter topologies. However,
in the literature, mostly linear based controllers are used for
active rectifiers with an LCL filter, see examples in [1][2]
(classical PI controllers), [3][4] (deadbeat control), [5] (spacestate based PI controller), [6] (resonant PI control). The main
disadvantages of using linear techniques are that the inherent
nonlinear behavior of the full-bridge rectifier is disregarded
and the results are valid only locally.
The use of control methodologies that rely on non-linear
approaches allow a systematic design procedure of robust
controllers, guaranteeing stability and performance in extended
operation ranges. One non-linear control technique specially
appropriate to switched power converters is the Sliding Mode
Control (SMC) approach [7]. Due to the fact that the control
action is commonly discontinuous, it is particularly suitable
to deal with variable structure systems such as power converters [8], resulting in controlled dynamical systems with
enhanced performance. However, one of the main problems
when implementing sliding mode based controllers is that the
switching frequency is variable. A possible solution to obtain
a fixed switching frequency is to apply a control algorithm
which synthesizes a continuous (or even discontinuous) control
law and then process it by means of pulse-width modulation
(PWM) devices.
In this framework, considering the system relative degree
(RD) and specific design issues (variables differentiation, computational burden, etc.) a Twisting Algorithm (TA) is proposed
to solve the rectifier control problem. This algorithm belongs to
the so-called family of Second Order Sliding Mode (SOSM)
controllers and is specially intended for relative degree two
systems, providing finite time convergence and insensibility to
coupled Lipschitz disturbances.
The main contribution of this paper is the development of a
SOSM based control strategy for a single-phase active rectifier
with an LCL filter. The control setup comprises two levels,
namely a robust inner current control, designed using the
Twisting Algorithm methodology, and an outer control level,
in charge of regulating the DC voltage in the presence of load
variations. Additionally, a uniform robust exact differentiator
for the current error is included in order to avoid numerical
problems during the implementation.
The paper is organized as follows: in the next section, the
single-phase active rectifier is introduced, and its dynamical
model is given. In Section III the whole control scheme
is proposed. Then, Section IV presents the design of the
current controller based on the TA, and also includes, the zero
Fig. 3.
III.
Fig. 1.
M ODEL OF
A SINGLE - PHASE ACTIVE RECTIFIER WITH
AN INDUCTIVE - CAPACITIVE - INDUCTIVE FILTER
Figure 1 shows a simplified scheme of a single-phase
rectifier with a LCL filter. The dynamics of this system is
described by
dig
dt
dvC
C
dt
dif
Lf
dt
dvCo
Co
dt
Lg
=
−rg ig − vC + vg ,
(1)
=
ig − if ,
(2)
=
−rf if + vC − uvCo ,
(3)
=
vCo
uif −
,
R
(4)
where ig , if are the input and output currents of the LCL
filter, respectively, vC is the voltage in the filter capacitor,
C, and vCo is the voltage in the DC link. The load has been
considered a pure resistance, R. The power grid voltage is
vg = Vg sin(ωt) (where Vg and ω are the voltage amplitude
and frequency, respectively) and u ∈ [−1, 1] is the control
signal and represents the duty cycle of the set of switches.
Lg and Lf are the LCL filter inductances, C and Co are the
capacitances in the LCL filter and DC link, respectively, and
rg , rf represent the inductor loses. Equations (1) and (4) can
be rewritten in an affine in the control form as
ẋ = f (x, t) + g(x, t)u(t),
where xT = (ig , vC , if , vCo ),
 1

− Lg (rg ig + vC )
 1 (ig − if ) 
C

f (x, t) = 
 L1 (vC − rf if )  ,
f
1
− RC
vCo
o
and
C ONTROL
Simplified scheme of a single-phase rectifier with a LCL filter.
dynamics analysis and the observer for the time derivative
of the current error. In Section V the use of a PI controller
for the regulation of the DC voltage is justified. Section VI
shows the numerical simulations performed to validate and
assess the performance of the proposed controller. Finally, the
conclusions are stated in Section VII.
II.
Sliding variable block diagram.
0


0

g(x, t) = 
− L1f vCo  .
1
Co if


(5)
OBJECTIVES AND OVERALL CONTROL
SCHEME
The control objective is twofold. To regulate the DC link
d
voltage, vCo , to a desired value vCo
and to have unity power
factor at the connection point. These two requirements can be
simultaneously attained by tracking an input current reference
of the form idg = Ig sin(ωt). The amplitude Ig is then
computed by the outer loop to maintain the DC voltage at its
desired value, while the sin(ωt) is obtained from the power
grid voltage vg .
Figure 2 shows a schematic representation of the proposed
control setup, with the SOSM inner-loop controlling the gridside current ig . For the outer control, a PI controller is used to
d
regulate the DC voltage, vCo , to its reference, vC
. The control
o
action generated by this external loop is, precisely, the required
amplitude of the grid current, Ig , while the grid voltage carrier
is obtained by means of a phase-locked loop (PLL). Then, the
current reference idg is constructed by direct multiplication.
IV.
S LIDING
MODE CURRENT CONTROLLER DESIGN
A. Control objective in terms of SMC
The first step for the design of the sliding mode controller is
to define an sliding variable s in terms of the current error e =
id −idg , taking into account that the ultimate current objective is
to track the grid current reference, ensuring robust convergence
to e = 0. Given that the error e is RD 3 with respect to u,
it cannot be used as sliding variable by its own (note that
classical SMC requires s to be RD 1). However, advantage
can be taken of SOSM techniques, which are capable to deal
with relative degree 2. Then, a simple RD 2 sliding variable
can be defined just adding a derivative term
s = e + k ė,
(6)
where k > 0, is a design parameter. Notice that this SOSM
approach needs only one time differentiator in s (see Figure
3), while classical SMC would also required to include full
information of the second time derivative into s, in order to
attain relative degree 1.
Moreover, the time derivative of the current reference
required in (6) can be reconstructed using the PLL outputs
cos(ωt) and ω. An schematic representation of the proposed
computation of s is depicted in Figure 4.
B. Second order SMC design
At this point, with the current control objective adequately
embedded into s, a robust controller can be straightforwardly
designed in the framework of second order sliding mode.
Considering the system characteristics and the proposed RD
2 sliding variable, a suitable SOSM control algorithm that
Fig. 2.
Overall control scheme.
The controller parameters r1 and r2 are tuned such that
they simultaneously satisfy the following conditions
r1 > r2 > 0
Γ′m (r1 + r2 ) − Φ′ > Γ′M (r1 − r2 ) + Φ′
Γ′m (r1 − r2 ) > Φ′ .
Then, in the bounded region of operation, it is proved that the
Twisting controller (7) generates a second order sliding mode
that robustly attracts the trajectories of the system to s = ṡ = 0
in finite time [10].
C. Analysis of the zero dynamics
Fig. 4.
The zero dynamics (or ideal sliding dynamics) can be
interpreted as the dynamics that remains once the controller has
forced the system to reach and stay on the manifold s = ṡ = 0.
Its behavior completes the analysis of the closed loop system
defined by (5) with (7).
Modified scheme to compute the sliding variable, s.
The constraint s = 0 establishes the following algebraic
dependence between ig and vC ,
kvC + (Lg Ig − kVg ) sin(ωt) + kωLg Ig cos(ωt)
. (8)
Lg − krg
Similarly, the second constraint ṡ = 0, relates the dynamics of
if with ig and vC , which together with (8) yields
ig =
Fig. 5.
Twisting Algorithm control scheme.
guarantees robust convergence to s = 0 in finite time (and consequent fulfillment of the aforementioned current objective), is
the so called Twisting Algorithm [9][10],
u = −r1 sign(s) − r2 sign(ṡ).
(7)
Note that its implementation is considerably simple, requiring
low on-line computational burden (see Fig. 5). It has only
two constant parameters, r1 and r2 , tuned in accordance with
certain system bounds as stated in [10], which are obtained
off-line during the design phase.
Specifically, given the sliding variable (6) and the affine
in the control form (5) of the active rectifier, the second time
derivative of s can be written
s̈ = ϕ′ (x, t) + γ ′ (x, t)u(t).
In a bounded region of operation ϕ′ (x, t) and γ ′ (x, t) are
uncertain (disturbed) but uniformly bounded functions, as
follows
0 < Γ′m ≤ γ ′ (x, t) ≤ Γ′M ,
and
|ϕ′ (x, t)| ≤ Φ′ .
vC = c1 if + c2 sin(ωt) + c3 cos(ωt),
(9)
where c1 , c2 and c3 are (10), (11) and (12), respectively.
Finally, it is substituted the discontinuous control action u
with an equivalent control (i.e., a fictitious continuous control
that would make the sliding surface flow invariant), which can
be obtained from s̈ = 0 and takes the form
1
ueq =
(d2 if + d3 sin(ωt) + d4 cos(ωt)),
(13)
d1 vCo
where d1 , d2 , d3 and d4 are given by (14), (15), (16) and (17),
respectively.
Substituting vC and u with (9) and (13) in the system (1)(4), the order of the closed loop dynamics is reduced in two.
It results
d3
dif
d2
i f + c2 −
sin(ωt)
Lf
= − rf − c1 +
dt
d
d1
1
d4
+ c3 −
cos(ωt),
(18)
d1
if
1
dvCo
=
d2 if + d3 sin(ωt) + d4 cos(ωt)
Co
dt
d1
vCo
vCo
.
(19)
−
R
d1
d2
d3
d4
c1
=
c2
=
c3
=
k(Lg − krg )
k 2 + C(Lg − krg )
(kLg ω 2 − rg )C(Lg − krg ) − kLg
Ig + Vg
k 2 + C(Lg − krg )
ω(C(k 2 rg2 − L2g ) − k 2 Lg )
kCω(Lg − krg )
Ig + 2
Vg .
2
k + C(Lg − krg )
k + C(Lg − krg )
(10)
(11)
(12)
= k 3 + C(Lg − krg )k
(14)
= −(rf + rg )k 3 + (rf rg C + Lf + Lg )k 2 − C(Lf rg + Lg rf )k + Lg Lf C
(15)
= rg Cω 2 (Lf − Lg )k 3 + (rg2 C + Cω 2 (Lg (Lg + Lf ) − rg2 CLf ) − Lf − Lg )k 2 + rg C(Lf − Lg )k Ig
− CLg Lf (1 − Lg Cω 2 ) Ig + (1 − CLf ω 2 )(k 3 − rg Ck 2 + Lg Ck)Vg
(16)
= (ω(C(Lg Lf ω 2 + rg2 ) − Lf − Lg )k 3 − Lg ω 3 C 2 Lf rg k 2 + Cω(CLf (L2g ω 2 + rg2 ) − Lg (Lg + Lf ))k − Lg rg C 2 Lf ω)Ig
+Cω(−rg k 3 + (Lf + Lg )k 2 − rg CLf k + Lg CLf )Vg .
(17)
The zero dynamics in (18) only depends on if . If rf − c1 +
> 0, if is stable with oscillations at frequency ω. Then,
the remaining dynamics of (19) can be seen as
d2
d1
vCo Co
v2
dvCo
= − Co + h(t),
dt
R
where h(t), due to the product of if (t) with sinusoidal
functions, is a periodic function with frequency 2ω. Assuming
vCo (t) > 0, it is transformed into the following linear ODE
2
(see [11] for further
with the change of coordinates ξ = 12 vC
o
details)
dξ
2
Co
= − ξ + h(t),
dt
R
with stable solutions for Co , R > 0. This analysis reveals that
a perfect current tracking yields oscillations at frequency 2ω
in the vCo voltage, similar to the results presented in [11].
Fig. 6.
Block diagram of the uniform robust exact differentiator.
Defining the error signal σ = z0 − η, the following
dynamical system can be constructed
ż0
ż1
D. Uniform robust exact differentiator
The RD 2 expression of s designed in (6) depends on
the time derivative of the current ig . However, in some
actual cases, it is preferred to avoid the implementation of
direct differentiators, among other reasons due to practical
noise issues. Therefore, in this control proposal, the variable
dig
dt can be computed by means of an uniform robust exact
differentiator (URED) [12], a time derivate observer based on
the generalized super twisting algorithm reported in [13].
Let η(t) be a Lebesgue-measurable function defined on
[0, ∞) and take it as the input signal. Suppose that η(t) can
be decomposed as
η(t) = η0 (t) + v(t),
where η0 (t) is the unknown base signal to be differentiated
and v(t) corresponds to a uniformly bounded noise signal.
Assuming that η0 (t) is twice continuously differentiable and
that the first derivative is Lipschitz with known Lipschitz
constant L > 0.
=
=
−k1 φ1 (σ) + z1
−k2 φ2 (σ) ,
where
φ1 (σ)
=
φ2 (σ)
=
1
3
|σ| 2 sign(σ) + |σ| 2 sign(σ)
3
1
sign(σ) + 2σ + µ2 |σ|2 sign(σ),
2
2
and k1 and k2 are constant positive gains dependent on L.
The variable z0 and z1 are the estimations of η0 and η̇0 ,
respectively. Indeed, z1 (t) converges exactly to η̇0 in finitetime, with the convergence time independent of the initial differentiation error (see [12] for details). See the block diagram
implementation in Figure 6.
di
Note that an observer for dtg based on the URED converges to its true value in finite-time. This is an important
advantage, given that the current SOSM Twisting controller
can be connected once the convergence time of the URED
has elapsed and, consequently, undesirable errors due to the
observer transient are avoided.
PI
OUTER - LOOP VOLTAGE CONTROL
480
The slower DC bus voltage control loop is designed
assuming an ideal current tracking (i.e. ig = idg ), which is
in fact guaranteed in finite time. In Section IV-C, the zero
dynamics was studied where idg = Ig sin(ωt) with Ig constant.
This analysis concluded that, in steady state, oscillations at 2ω
frequency appeared in the DC bus voltage, vCo . Consequently,
the control objective should be turned into the regulation of
the averaged value1 of the DC bus voltage, v̄Co , rather than of
the instantaneous voltage vCo .
Assuming a perfect current tracking, the control law designed in the previous subsection guarantees
vco
d
vco
470
460
450
DC bus voltage [V]
V.
440
430
420
410
400
390
ig = Ig (t) sin(ωt),
(20)
where Ig (t) is the required amplitude to regulate the DC bus
voltage at the desired value. Replacing (20) in (1), (2), (3)
and (4) one gets a complex dynamics with the oscillations
predicted in Section IV-C and the (first, second and third) time
derivative of Ig (t). Then, averaging the resulting dynamics,
with the following coordinates change
ψ=
dψ
2
= − ψ − a2 Ig2 + a1 Ig − a0 ,
dt
R
(21)
where Ig can be used as a control action to regulate ψ at a
d 2
desired value ψ d = 12 (vCo
) , and
a2
a1
a0
π(rf (ω 2 Lg C − 1)2 + ω 2 rg2 rf C 2 + rg )
,
ω
πVg (1 + 2ω 2 rg rf C 2 )
,
=
ω
= πVg2 ωrf C 2 .
=
Notice that since all the parameters are positive, a2 , a1 , a0 > 0.
Replacing the parameters of the rectifier used in the simulation stage (see Section VI),
a2
a1
a0
= 9.9 · 10−5
= 3.11
= 3.8 · 10−4
reveals that a1 is clearly dominating a2 and a0 , and (21) can
be approximated by the linear system
Co
dψ
2
= − ψ + a1 Ig ,
dt
R
which can be regulated by means of a PI controller
Z
d
Ig = kp (ψ − ψ) + (ψ d − ψ)dt,
d 2
with kp , ki > 0 and ψ d = 12 (v̄C
) .
o
1 The
0
0.01
0.02
0.03
0.04
0.05
time [s]
0.06
0.07
0.08
0.09
0.1
Fig. 7. Simulation results: DC bus voltage after a reference step change,
from 400V to 450V.
VI.
S IMULATIONS
In order to validate the control algorithm presented in
this paper numerical simulations have been performed using
a fixed time step of5 · 10−8 s. The sliding variable has been
implemented using the URED to observe the time derivative
of ig and minimize possible numerical differentiation issues.
1 2
v̄ ,
2 Co
and assuming slow variations of Ig (t), one gets
Co
380
averaged
R t function of a periodic signal f (t) of period T is calculated
as f¯(t) = T1 t−T f (τ )dτ .
In this section two different simulation tests are shown.
The first one consists in a change of the DC voltage reference,
from 400V to 450V, with a load of R = 50Ω. The second
test consists in a sudden change of the load form R = 50Ω to
R = 20Ω. The parameters of the power converter are: Lg =
300µH, rg = 50mΩ, C = 20µF, Lf = 100µH, rf = 50mΩ
and Co = 2mF. The control gains are: k = 10−4 , r1 = 0.7
and r2 = 0.3 (for the Twisting Algorithm), and kp = 0.75 and
ki = 50 (for the external PI controller).
d
In Figure 7, the DC voltage vCo and its reference value vC
o
are depicted. The voltage is regulated after a short transient
time, even when the oscillations predicted in Section IV-C
appear. Meanwhile, the grid current ig is in phase with the
voltage grid vg as Figure 8 shows, implying an unity power
factor. The results of the second test are shown in figures 9 and
10. In this case, the change of the load value implies a transient
before the controller fulfills the control goals. Similarly to the
previous case, the average of the DC voltage is regulated with
an unity power factor.
VII.
C ONCLUSIONS
AND FINAL REMARKS
A general controller for an active rectifier with an
inductive-capacitive-inductive filter was developed. Taking into
account several features, such as the system relative degree,
robustness and implementation simplicity, a non-linear control
approach was proposed using a second order sliding mode
Twisting algorithm in conjunction with an exact differentiator.
This results in a robust and relatively simple control structure
that particularly suits the current tracking problem. In addition,
satisfactory zero dynamics analysis is provided for the current
loop. The controller setup is completed with an outer loop
based on a PI controller for the regulation of the average output
DC voltage.
The feasibility of the combined control strategy was assessed by means of comprehensive numerical simulations
conducted on a benchmark model of an existing rectifier,
simulating realistic and exacting operating conditions. They
demonstrate that the proposed SOSM non-linear approach is a
highly efficient solution for the rectifier control problem under
study, establishing its viability for implementation in actual
single-phase active rectifiers with LCL filter. In accordance
with these encouraging results, the following stage of this
project will be the refinement and implementation of this
SOSM based controller in the laboratory experimental test
bench.
100
i
g
vg
80
60
40
g
i [A],v /4[V]
20
g
0
−20
−40
−60
ACKNOWLEDGMENT
−80
−100
0
0.01
0.02
0.03
0.04
0.05
time [s]
0.06
0.07
0.08
0.09
0.1
Fig. 8. Simulation results: grid current, ig after a change of the DC bus
voltage reference, from 400V to 450V.
480
vco
d
vco
DC bus voltage [V]
470
460
R EFERENCES
450
[1] J. Dannehl, C. Wessels, and F. Fuchs, “Limitations of voltage-oriented
PI current control of grid-connected PWM rectifiers with LCL filters,”
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2009.
[2] M. Liserre, F. Blaabjerg, and S. Hansen, “Design and control of an
LCL-filter-based three-phase active rectifier,” IEEE Trans. on Industry
Applications, vol. 41, no. 5, pp. 1281–1291, 2005.
[3] Y.-R. Mohamed, M. Rahman, and R. Seethapathy, “Robust line-voltage
sensorless control and synchronization of LCL-filtered distributed generation inverters for high power quality grid connection,” IEEE Trans.
on Power Electronics, vol. 27, no. 1, pp. 87–98, 2012.
[4] E. Wu and P. Lehn, “Digital current control of a voltage source
converter with active damping of LCL resonance,” IEEE Trans. on
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[5] J. Dannehl, F. Fuchs, and P. Thogersen, “PI state space current control
of grid-connected PWM converters with LCL filters,” IEEE Trans. on
Power Electronics, vol. 25, no. 9, pp. 2320–2330, 2010.
[6] Y. Tang, P. Loh, P. Wang, F. Choo, and K. Tan, “Improved one-cyclecontrol scheme for three-phase active rectifiers with input inductorcapacitor-inductor filters,” IET Power Electronics, vol. 4, no. 5, pp.
603–614, 2011.
[7] V. Utkin, J. Guldner, and J. Shi, Sliding mode control in electromechanical systems. Taylor & Francis, 1999.
[8] R. Venkataramanan, Sliding mode control of power converters. Dissertation (Ph.D.), California Institute of Technology, 1986.
[9] A. Levant, “Sliding order and sliding accuracy in sliding mode control,”
International Journal of Control, vol. 58, no. 6, pp. 1247–1263, 1993.
[10] L. Fridman and A. Levant, Sliding Mode Control in Engineering,
W. Perruquetti and J. Barbot, Eds. Marcel Dekker Inc., 2002, Chapter
3: Higher Order Sliding Modes, pp. 53-101.
[11] R. Griñó, E. Fossas, and D. Biel, “Sliding mode control of a full-bridge
unity power factor rectifier,” Lecture Notes in Control and Information
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[12] E. Cruz-Zavala, J. A. Moreno, and L. Fridman, “Uniform robust exact
differentiator,” IEEE Trans. on Automatic Control, vol. 56, no. 11, pp.
2727–2733, 2011.
[13] J. Moreno, Lyapunov approach for analysis and design of second order
sliding mode algorithms. Springer, 2011.
440
430
420
410
400
390
380
0.1
0.11
0.12
0.13
0.14
0.15
time [s]
0.16
0.17
0.18
0.19
0.2
Fig. 9. Simulation results: DC bus voltage after a change of the load value,
from R = 50Ω to R = 10Ω.
100
ig
vg
80
60
40
20
0
g
g
i [A],v /100[V]
This work was supported by the Spanish Government
Research Project DPI2010-15110, Project Marie Curie FP72011-IIF-299767/911767 ACRES (EU), CONICET and UNLP
(Argentina). The research of C. Kunusch has been supported
by the Seventh Framework Programme of the European Community through the Marie Curie actions (GA: PCIG09-GA2011-293876), the Puma-Mind project (GA: FCH-JU-20111-303419), the CICYT project DPI2011-25649 (MINECOSpain) and the CSIC JAE-DOC Research Programme.
−20
−40
−60
−80
−100
0.1
0.11
0.12
0.13
0.14
0.15
time [s]
0.16
0.17
0.18
0.19
0.2
Fig. 10. Simulation results: grid current, ig after a change of the load value,
from R = 50Ω to R = 10Ω.
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