# Digital Control of a Continuous Stirred Tank Reactor

Authors: M. P. Di Ciccio; M. Bottini; P. Pepe
Journal: Mathematical Problems in Engineering
 (2011)

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## Abstract

We project a novel digital control law for continuous stirred tank reactors, based on sampled measures of temperatures and reactant concentration, as it happens in practice. The methodology of relative degree preservation under sampling is used. It is proved that a suitably approximated sampled system, obtained by Taylor series expansion and truncation, in closed loop with the projected control law, is asymptotically
stable, provided that a condition on the sampling period is verified. Such condition allows for values of the sampling period larger than necessary in practical implementation with usual technology. Many simulations show the high performance of the proposed digital control law.


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## Body

## 1. Introduction

The control of the operation of chemical reactors has attracted the attention of researchers for a long time. The underlying motivation relies on the fact that industrial chemical reactors are frequently operated at unstable operating conditions, which often corresponds to optimal process performance [1–19]. Polymerization processes [5] and bioreactors fermentors [18] are important examples of large-scale chemical reactors operated at unstable conditions. As well known, the measures of the reactant concentration cannot be achieved on continuous time. In practice, these measures are available at certain time intervals. It is therefore very important to project control laws which are piecewise constant and are updated at each measures update. That is, for practical use, a digital control law has to be projected. Applications of nonlinear digital control theory to some chemical reactors can be found in [20, 21]. In [11] a discrete-time controller is obtained by linear approximations successively executed at each sampling time, and an application to a continuous stirred tank reactor is shown.In this paper we deal with a novel digital controller design for a continuous stirred tank reactor, where an irreversible, exothermic, liquid-phase reactionA→B evolves. The controlled output is the reactor temperature. The jacket temperature dynamics is considered. The control law is actuated by means of a pneumatic valve which regulates the cooling water flow rate. This actuation process is easier to implement than the ones based on the use of the inlet reactor concentration or of the jacket temperature as manipulated inputs. The model consists of three nonlinear ordinary differential equations. We develop a digital control law by means of a suitably approximated sampled model, obtained by Taylor series expansions (see [21, 22]). The preservation of relative degree methodology shown in [22] is here used. For the underlying continuous time model, the relative degree is not full. The relative degree is preserved for the approximated sampled model by introducing a dummy output. Then a feedback control law is projected, by which the approximated sampled system becomes asymptotically stable. Many performed computer simulations show the high performance of the proposed digital controller. Saturation effects have been taken into account in simulations. The convergence of the state variables to the arbitrarily chosen operating point is always obtained with the many considered system initial conditions (even start-up initial conditions).


## 2. The Model of the CSTR

Here we study a CSTR with jacket cooling in which a first-order irreversible exothermic reaction takes place [23]: A→B (Figure 1).Figure 1
Schematic reactor type CSTR.The liquids in the reactor and in the jacket are assumed perfectly mixed, that is, with no radial, axial, or angular gradients in properties (temperature, concentrations). This allows to consider reactor and jacket temperatures and reactant concentrations as space invariant variables. If physical properties are assumed constant (densities and heat capacities), the reactor volume is constant and the jacket volume is constant, the mathematical model of the CSTR is given by the following set of nonlinear functional differential equations:(2.1)ẋ1(t)=F(Ca0-x1(t))Vr-x1(t)K0e-E/Rx2(t),ẋ2(t)=F(T0-x2(t))Vr-λx1(t)K0e-E/Rx2(t)ρ-1Cp-1-UAj(x2(t)-x3(t))VrρCp,ẋ3(t)=u(t)(Tcin-x3(t))Vj+UAj(x2(t)-x3(t))VjρjCj,
where x1=Ca = reactant concentration (kmol/m3), x2=TR = reactor temperature (K), x3=TJ = jacket temperature (K), u = flow rate of coolant (control input) (m3/s), K0 = pre-exponential factor (s−1), E = activation energy (J/kmol), R = universal gas constant, 8314 Jkmol−1K−1, F = flow rate of feed and product (m3/s), ρ = density of product stream (kg/m3), Ca0 = concentration of reactant A in feed (kmol/m3), VR = volumetric holdup of liquid in reactor (m3), T0 = temperature of feed (K), Cp = heat capacity of product (Jkg−1K−1), λ = heat of reaction (J/kmol), U = overall heat transfer coefficient (WK−1m−2), Aj  =  jacket heat transfer area (m2), ρj = density of coolant (kg/m3), Cj = heat capacity of coolant (Jkg−1K−1), and Tcin = supply temperature of cooling medium (K).Letϰ∈[0,1] be the reaction conversion(2.2)ϰ=Ca0-CaCa0.
Let ϰeq be the chosen operating point reaction conversion. By (2.2) it follows that the operating point concentration of reactant in the reactor Caeq is given by(2.3)Caeq=Ca0(1-ϰeq).
From equations in (2.1), setting zero the left-hand side, it follows that the operating point reactor temperature (TReq), the operating point jacket temperature (Tjeq), the necessary coolant water flow rate (ueq) for the chosen operating point are given by
(2.4)TReq=-E(ln⁡(-F(-Ca0+Caeq)CaeqK0Vr))-1R-1,TJeq=λCa0FϰeqUAj-FρCpT0UAj-E(ln⁡(Fϰeq(1-ϰeq)K0Vr))-1R-1-FρCpE(ln⁡(Fϰeq(1-ϰeq)K0Vr))-1R-1U-1Aj-1,ueq=UAjEρj-1Cj-1(Tcin-TJeq)-1(ln⁡(Fϰeq(1-ϰeq)K0Vr))-1R-1+UAjTJeqρjCj(Tcin-TJeq).


## 3. Sampled System and Digital Control Law

The system (2.1) is in the form(3.1)ẋ(t)=f(x(t))+g(x(t))⋅u(t),
where the functions f, g can be easily defined by (2.1) as(3.2)f(x(t))=[f1(x(t))f2(x(t))f3(x(t))]=[F(Ca0-x1(t))Vr-x1(t)K0e-E/Rx2(t)F(T0-x2(t))Vr-λx1(t)K0e-E/Rx2(t)ρ-1Cp-1-UAj(x2(t)-x3(t))VrρCpUAj(x2(t)-x3(t))VjρjCj],g(x(t))=[g1(x(t))g2(x(t))g3(x(t))]=[00Tcin-x3(t)Vj].
In order to rewrite (3.1) in normal form, as required in [22], let us now consider the following variables:(3.3)ξ(t)=[ξ1(t)ξ2(t)ξ3(t)]=[x2(t)-TReqf2(x(t))x1(t)-Caeq]=ϕ(x(t)),
from which the easy inverse relation holds(3.4)x(t)=Φ-1(ξ(t))=[ξ3(t)+Caeqξ1(t)+TReq-FρCpUAj(T0-ξ1(t)-TReq)+λVr(ξ3(t)+Caeq)K0e-E/R(ξ1(t)+TReq)UAj+ξ2(t)VrρCpUAj+ξ1(t)+TReq].
Let the control input be equal to(3.5)u(t)=ueq+v(t).
Let the controlled output be(3.6)y(t)=h(ξ(t))=ξ1(t).
We want to control to zero this output, which corresponds to drive the reactor temperature to the reactor temperature operating point TReq.We can write the system equations by using the new variablesξ(t) as follows:(3.7)ξ̇(t)=p1(ξ(t))+p2(ξ(t))v(t),
where p1, p2 are given by
(3.8)p1(ξ(t))=[p11(ξ(t))p21(ξ(t))p31(ξ(t))],p11(ξ(t))=ξ2(t),p21(ξ(t))=-λK0FCa0Vr-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1-ueqUAjξ1(t)VjVrρCp+ueqFT0VjVr+λξ3(t)K0FVr-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1-ueqUAjTReqVjVrρCp-ueqFξ1(t)VjVr+λK0CaeqFVr-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1-ueqξ2(t)Vj+ueqUAjTcinVjVrρCp+λK02ξ3(t)(eE/R(ξ1(t)+TReq))-2ρ-1Cp-1-ueqFTReqVjVr-UAjξ2(t)VjρjCj+λK02Caeq(eE/R(ξ1(t)+TReq))-2ρ-1Cp-1-ξ2(t)FVr-ξ2(t)λK0Eξ3(t)ρ-1Cp-1R-1(ξ1(t)+TReq)-2(eE/R(ξ1(t)+TReq))-1-ξ2(t)λK0ECaeqρ-1Cp-1R-1(ξ1(t)+TReq)-2(eE/R(ξ1(t)+TReq))-1-UAjξ2(t)VrρCp+UAjFT0VjρjCjVr-UAjFξ1(t)VjρjCjVr-UAjFTReqVjρjCjVr-UAjλK0ξ3(t)Vj-1ρj-1Cj-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1-UAjλK0CaeqVj-1ρj-1Cj-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1-ueqλK0ξ3(t)Vj-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1-ueqλK0CaeqVj-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1,p31(ξ(t))=F(Ca0-ξ3(t)-Caeq)Vr-(ξ3(t)+Caeq)K0e-E/R(ξ1(t)+TReq),p2(ξ(t))=[p12(ξ(t))p22(ξ(t))p32(ξ(t))],p12(ξ(t))=0,p22(ξ(t))=TcinUAjVjVrρCp+FT0VjVr-Fξ1(t)VjVr-FTReqVjVr-UAjξ1(t)VjVrρCp-TReqUAjVjVrρCp-ξ2(t)Vj-λK0ξ3(t)Vj-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1-λK0CaeqVj-1ρ-1Cp-1(eE/R(ξ1(t)+TReq))-1,p32(ξ(t))=0.Sinceξ=0 corresponds to x=[CaeqTReqTJeq], it follows that p1(0)=0. The introduction of the new variables allows to obtain the CSTR equations in normal form [22], and the origin is an equilibrium point. Therefore, the stability of the origin for the system described by (3.7) is equivalent to the stability of the chosen operating point for the system described by (2.1). Let us now sample and approximate, by Taylor series expansion and truncation, the system (3.7) as follows: setting v(t)=v(kδ)=vk, kδ≤t<(k+1)δ, k=0,1,2,…,(3.9)[ξ̃1ξ̃2]((k+1)δ)=[1δ01][ξ̃1ξ̃2](kδ)+[δ22δ]R0(kδ),ξ̃3((k+1)δ)=ξ̃3(kδ)+δp31(ξ̃(kδ))+δ22q(ξ̃(kδ)),
where ξ̃(kδ) is the approximation of ξ(kδ); R0 is given by (Lie derivatives are used), (3.10)R0(kδ)=lp12h(ξ̃(kδ))+lp2lp1h(ξ̃(kδ))v(kδ);lp12h(ξ̃(kδ))=∂lp11h(ξ̃(kδ))∂(ξ̃(kδ))p1(ξ̃(kδ));lp11h(ξ̃(kδ))=∂h(ξ̃(kδ))∂(ξ̃(kδ))p1(ξ̃(kδ));lp2lp1h(ξ̃(kδ))=∂lp1h(ξ̃(kδ))∂(ξ̃(kδ))p2(ξ̃(kδ));q(ξ̃(kδ)) is given  by (3.11)q(ξ̃(kδ))=∂p31(ξ̃(kδ))∂ξ̃(kδ)p1(ξ̃(kδ)),q(ξ̃(kδ))=-K0Eξ̃2(kδ)ξ̃3R-1(ξ̃1(kδ)+TReq)-2(eE/R(ξ̃1(kδ)+TReq))-1-K0Eξ̃2(kδ)CaeqR-1(ξ̃1(kδ)+TReq)-2(eE/R(ξ̃1(kδ)+TReq))-1-F2Ca0Vr2+F2ξ̃3Vr2+F2CaeqVr2+2FK0ξ̃3Vr-1(eE/R(ξ̃1(kδ)+TReq))-1+2FK0CaeqVr-1(eE/R(ξ̃1(kδ)+TReq))-1+K02Caeq(eE/R(ξ̃1(kδ)+TReq))-2-FK0Ca0Vr-1(eE/R(ξ̃1(kδ)+TReq))-1+ξ̃3(kδ)K02(eE/R(ξ̃1(kδ)+TReq))-2.The approximated sampled system (3.9) is obtained by exploiting the normal form of (3.7)  (see [22]) and by neglecting, in the Taylor series expansion, the terms which are 0(δ3) in the expression of ξ1((k+1)δ) and in the expression of ξ3((k+1)δ), and the terms which are 0(δ2) in the expression of ξ2((k+1)δ) (see terms Nr,δ2 in [22, Proposition 3.2]). Notice that, since (∂p31(ξ̃(kδ))/∂ξ̃(kδ))p2(ξ̃(kδ))=0, the inclusion of a term which is 0(δ2) in the approximation of ξ3((k+1)δ) does not cause a new presence (besides the one in R0) of the control input vk. As well, neglecting terms which are 0(δ2) in the expression of ξ2((k+1)δ) avoids to have further terms (besides R0) which involve the control input. This fact is instrumental for building the digital control law here proposed, as will be shown below.The approximated sampled system (3.9) admits relative degree equal to 1, thus not preserving the relative degree equal to 2 of the original continuous time system (2.1) with respect to the output x2(t) (reactor temperature). Thus, a digital input/output linearizing feedback control law built up on the basis of the given output (3.6) yields a further internal dynamics. In order to preserve the relative degree, let us introduce the dummy output as proposed in [22],(3.12)yδ(kδ)=[1-δ2][ξ̃1(kδ)ξ̃2(kδ)],
and let us define the variables(3.13)χ(k)=[χ1(k)χ2(k)χ3(k)]=[yδ(kδ)yδ((k+1)δ)ξ̃3(kδ)]=Mξ̃(kδ),
where the nonsingular matrix M is given by(3.14)M=[1-δ201δ20001].
Let M1 be the upleft submatrix of M(3.15)M1=[1-δ21δ2].
We obtain, from (3.9),(3.16)[χ1(k+1)χ2(k+1)]=M1[1δ01]M1-1[χ1(k)χ2(k)]+M1[δ22δ]R0(kδ),χ3(k+1)=χ3(k)+δp31(M-1χ(k))+δ22q(M-1χ(k)),
where R0, rewritten with the variables χ, becomes(3.17)R0(kδ)=lp12h(M-1χ(k))+lp2lp1h(M-1χ(k))vk.Theorem 3.1.
Let(3.18)Γ=[Γ1Γ2]∈R1x2
be chosen such that the matrix [01Γ1Γ2] has eigenvalues inside the open complex unitary circle. Let the sampling period δ>0 satisfy the following inequality:
(3.19)δ<δmax⁡=2(F/Vr)+K0e-E/RTReq.
Then, the following discrete time feedback control law
(3.20)vk=-lp12h(M-1χ(k))lp2lp1h(M-1χ(k))+[1-2][χ1(k)χ2(k)]+Γ[χ1(k)χ2(k)]δ2lp2lp1h(M-1χ(k))
is such that the trivial solution of the closed-loop system (3.16)–(3.20) is asymptotically stable. Moreover, the eigenvalues of the Jacobian (evaluated at 0) of the right-hand function describing the closed-loop system (3.16)–(3.20) are given by the eigenvalues of the matrix [01Γ1Γ2] and by the eigenvalue
(3.21)1-δ(FVr+K0e-E/RTReq)+δ22(FVr+K0e-E/RTReq)2,
which takes its minimum, equal to 1/2, at δ=δ(1/2)=1/((F/Vr)+K0e-E/RTReq).Proof.
By the form of the matrixM1, the following equation holds for χ1 and for χ2, in (3.16)
(3.22)[χ1(k+1)χ2(k+1)]=[01-12][χ1(k)χ2(k)]+[0δ2]R0(kδ).
Notice that, by construction, such subsystem (and thus the overall system (3.16)) admits relative degree equal to 2, in a neighborhood of the origin, with respect to the dummy output χ1(k). Indeed, taking into account of the form of R0(kδ), we have
(3.23)χ1(k+1)=χ2(k+1),χ1(k+2)=χ2(k+1)=[-12]χ(k)+δ2(lp12h(M-1χ(k))+lp2lp1h(M-1χ(k))v(kδ)),
and the term multiplying v(kδ) is nonzero in a neighborhood of the origin, since lp2lp1h(M-1χ(k)) is smooth in a neighborhood of the origin and lp2lp1h(0)≠0. Therefore, taking again into account the form of R0, the proposed discrete-time feedback control law yields the following equations for the closed-loop-system (3.16), (3.20)
(3.24)[χ1(k+1)χ2(k+1)]=[01Γ1Γ2][χ1(k)χ2(k)],χ3(k+1)=χ3(k)+δp31(M-1χ(k))+δ22q(M-1χ(k)).
In order to show the asymptotic stability of the nonlinear system (3.24), it is sufficient to show that the linear system obtained by linearizing (3.24) is asymptotically stable. The Jacobian, at zero, of the function in the right-hand side of (3.24) is given, taking into account of p31 and of q, by
(3.25)J=[010Γ1Γ20αβγ],
where
(3.26)γ=1-δ(FVr+K0e-E/RTReq)+δ22(FVr+K0e-E/RTReq)2
and α,β are suitable reals. Given the triangular structure of the Jacobian J, the eigenvalues are given by the eigenvalues of the matrix [01Γ1Γ2] and by γ. The eigenvalues of the matrix [01Γ1Γ2] are inside the open unitary circle. The eigenvalue γ is always positive (for positive sampling period), and is inside the open unitary circle if the inequality (3.19) is satisfied. Moreover, by using the derivative of γ with respect to δ, we obtain that γ takes the minimum value, 0.5, at δ(1/2).Now, going back to initial variablesx, by means of the (approximated) relation χ(k)≊Mϕ-1(x(kδ)), the digital feedback control law here proposed is given, for kδ≤t<(k+1)δ,  k=0,1,2,…, by(3.27)u(t)=4VjλK0FCa0(Tcin-x3(kδ))-1U-1Aj-1(eE/Rx2(kδ))-1+2VjVrρCpTReqΓ1δ2(Tcin-x3(kδ))UAj-2VjFx2(kδ)(Tcin-x3(kδ))Vr+VjρCpF2T0(Tcin-x3(kδ))UAjVr-2VjVrρCpx2(kδ)δ2(Tcin-x3(kδ))UAj+2Vjx2(kδ)δ(Tcin-x3(kδ))-2VjλK0Fx1(kδ)(Tcin-x3(kδ))-1U-1Aj-1(eE/Rx2(kδ))-1-VrVjλK02x1(kδ)(Tcin-x3(kδ))-1U-1Aj-1(eE/Rx2(kδ))-2+VjFT0λx1(kδ)K0E(Tcin-x3(kδ))-1U-1Aj-1R-1x2(kδ)-2(eE/Rx2(kδ))-1-VjλK0Fx1(kδ)E(Tcin-x3(kδ))-1U-1Aj-1x2(kδ)-1R-1(eE/Rx2(kδ))-1-VrVjλ2x1(kδ)2K02E(Tcin-x3(kδ))-1U-1Aj-1ρ-1Cp-1(eE/Rx2(kδ))-21Rx2(kδ)-Vjλx1(kδ)K0(Tcin-x3(kδ))-1ρ-1Cp-1(eE/Rx2(kδ))-1-2VjVrρCpΓ1x2(kδ)δ2(Tcin-x3(kδ))UAj-Vjλx1(kδ)K0E(Tcin-x3(kδ))-1ρ-1Cp-1x2(kδ)-1R-1(eE/Rx2(kδ))-1-VjUAjx2(kδ)(Tcin-x3(kδ))VrρCp+Vjx3(kδ)F(Tcin-x3(kδ))Vr+VjFT0(Tcin-x3(kδ))Vr-VjρCpF2x2(kδ)(Tcin-x3(kδ))UAjVr+Vjx3(kδ)λx1(kδ)K0E(Tcin-x3(kδ))-1ρ-1Cp-1R-1x2(kδ)-2(eE/Rx2(kδ))-1+VjUAjx3(kδ)(Tcin-x3(kδ))VrρCp-UAjx2(kδ)(Tcin-x3(kδ))ρjCj+UAjx3(kδ)(Tcin-x3(kδ))ρjCj-2Vjx3(kδ)δ(Tcin-x3(kδ))+2VjVrρCpTReqδ2(Tcin-x3(kδ))UAj-2VjρCpFT0δ(Tcin-x3(kδ))UAj+2VjρCpFx2(kδ)δ(Tcin-x3(kδ))UAj+2VrVjλx1(kδ)K0δ-1(Tcin-x3(kδ))-1U-1Aj-1(eE/Rx2(kδ))-1.The digital control law (3.27) is applied to the CSTR system, described by the differential equations (2.1), by means of a zero-order hold device. Such control law can be easily implemented on a microprocessor. It is a static feedback control law, that is, no dynamics are involved. The control law has to be computed at every sampling period which is very much larger than the computation time (many minutes versus some seconds at maximum). The great advantage of this digital control law is also given by the fact that only sampled-data measures, provided by a suitable device, are necessary. Take into account that, in practice, continuous time measures are difficult to obtain, or even impossible. The feedback control law here presented is therefore easily realizable. The closed-loop system (2.1), (3.27) (with the zero-order hold) is simulated in the following section.


## 4. Simulation Results

Simulations, using MatLab (The MathWorks, Inc.) Software Package, have been carried out to verify the effectiveness of the proposed method. The values of the model parameters used in simulation are given in Table1 (taken from [23]). In all the carried out simulations, a minimum and a maximum value for the input u, specifically the volumetric flow rate of the refrigerant, is imposed, to take into account of the actuator saturation. It is considered: (4.1)FJmin⁡≤u(t)≤FJmax⁡
with the following acceptable physical values for a control valve (4.2)FJmin⁡=0.002m3/s=7.2m3/h,FJmax⁡=0.08m3/s=288m3/h.
With these parameters, δmax in condition (3.19) is equal to 0.57 h, which corresponds to about 34 min. The value of δ(1/2) is equal to about 0.28 h which corresponds to about 16 min. We performed several sets of simulation runs and all showed the high performance of the digital control law here proposed. We choose the reactor conversion ϰeq=0.85, to which the following unstable operating point corresponds: Caeq=1.2015 kmol/m3, TReq=350 K, TJeq=312 K, ueq=69.86 m3/h. In the simulations here presented, the process is initially assumed to be at the start up, that is x1(0)=Ca0 (concentration of reactant in feed) = 8.01 kmol/m3, x2(0)=T0 (temperature of feed) = 294 K, x3(0)=Tcin (supply temperature of cooling medium) = 294 K. This means that the reactor is very far from the operating point.Table 1
Reactor and controller parameters values.
K0=7.47×1010 h−1Ca0=8.01 kmol·m−3E=69.71×106 J·kmol−1F=15.76 m3·h−1ρ=801 kg·m−3Tcin=294 KR=8314 J·kmol−1·K−1T0=294 Kρj=1000 kg·m−3Aj=45.2 m2Cp=3137 J·kg−1·K−1Vr=30.3 m3Cj=4183 J·kg−1·K−1Vj=9 m3λ=-69.71×106 J·kmol−1Γ1=-6·10-6U=3.0636×106 J·h−1·m−2·K−1Γ2=-5·10-3In Simulation 1 (see Table2, Figure 2) we consider the reactor at nominal parameters values (see Table 1), without any uncertainty. Figure 2 shows the evolution of the reactant concentration, of the reactor temperature, of the jacket temperature, and of the piecewise constant control signal, respectively, with sampling period equal to 10−1 h (6 min).Table 2
Process parameters of Simulation 1.
Reactor operating pointInitial stateχeq=85%{Ca(t=0)=Ca0TR(t=0)=T0TJ(t=0)=Tcinδ=0.1 h = 6 minCaeq=1.2015 kmol·m−3TReq=350 KTJeq=312 Kueq=69.86 m3·h−1Parameters mismatch: absentResults of Simulation 1.
(a)(b)(c)(d)As can be seen, all the state variables converge to the operating point ones in a very good fashion, without dangerous oscillations. The reactor temperature settling time is estimated from plots to be about 6 h, the steady-state error is equal to 0.In Simulation 2 (see Table3, Figure 3) we consider again the reactor at nominal parameters values (see Table 1), without any parameters mismatch, and a sampling period equal to 0.25 h (15 min) is considered. The same desirable convergence to the operating point is achieved, with a reactor temperature settling time of about 8 h. Again no dangerous oscillations appear and the reactor temperature is driven in a very good fashion to the operating point, with steady-state error equal to 0.Table 3
Process parameters of Simulation 2.
Reactor operating pointInitial stateχeq=85%{Ca(t=0)=Ca0TR(t=0)=T0TJ(t=0)=Tcinδ=0.25 h = 15 minCaeq=1.2015 kmol·m−3TReq=350 KTJeq=312 Kueq=69.86 m3·h−1Parameters mismatch: absentResults of Simulation 2.
(a)(b)(c)(d)In Simulations 3, 4 (see Table4, Figure 4, and Table 5, Figure 5, resp.) we consider the reactor whose parameters are uncertain, with sampling time again equal to 0.1 h and 0.25 h, respectively. We apply the controller obtained using the nominal values of the parameters (see Table 1), to a reactor whose parameter values are different with respect to the nominal ones as reported in Table 5. In this case we observe a reactor temperature steady-state error of about 5 K. Take into account that the parameters mismatch has been deliberately chosen very critical, in order to show the robustness of the proposed controller.Table 4
Process parameters of Simulation 3.
Reactor operating pointInitial stateχeq=85%{Ca(t=0)=Ca0TR(t=0)=T0TJ(t=0)=Tcinδ=0.1 h = 6 minCaeq=1.2015 kmol·m−3TReq=350 KTJeq=312 Kueq=69.86 m3·h−1Parameters mismatch:F=F/1.3, T0=T0·1.05;Ca0=Ca0/1.1, Tcin=Tcin·1.03U=0.8·UTable 5
Process parameters of Simulation 4.
Reactor operating pointInitial stateχeq=85%{Ca(t=0)=Ca0TR(t=0)=T0TJ(t=0)=Tcinδ=0.25 h = 15 minCaeq=1.2015 kmol·m−3TReq=350 KTJeq=312 Kueq=69.86 m3·h−1Parameters mismatch:F=F/1.3, T0=T0·1.05,Ca0=Ca0/1.1, Tcin=Tcin·1.03,U=0.8·UResults of Simulation 3.
(a)(b)(c)(d)Results of Simulation 4.
(a)(b)(c)(d)The performed simulations show the high performance of the proposed controller, as evidenced, in the case the reactor parameters are exactly known, by zero steady state error, no dangerous oscillations, good settling time. When parameters mismatch occurs, the proposed controller yields a very good behavior for the reactor, evidenced by small steady-state errors, no dangerous oscillations, and good settling time. We stress the fact that the proposed controller allows for sampled-data measures of the state variables, which is very useful in practical applications.


## 5. Conclusion

In this paper we have presented a digital control law for a continuous stirred tank reactor, constructed on the basis of an approximated sampled model and of the preservation of the relative degree methodology. The proposed feedback control law is easy to implement and simulations show high performance. A condition on the sampling period is found, by which the asymptotic stability of the closed-loop approximated sampled system is proved. Such condition is verified by sampling periods belonging to feasible sets, allowing for values larger than necessary in practical implementation. Future work will concern the construction of a discrete-time observer, in order to avoid reactant concentration measures, which are difficult to do and expensive (see [24–27]). Another topic which will be investigated is the problem arised by the recycle time-delay (see [21, 28–31]), on the basis of the results in [32].



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*Source: 439785-2011-04-17.xml*