Open access peer-reviewed chapter

Minimum Dissipation Conditions of the Mass Transfer and Optimal Separation Sequence Selection for Multicomponent Mixtures I

By A.M. Tsirlin and I.N. Grigorevsky

Submitted: February 20th 2012Reviewed: October 19th 2012Published: July 24th 2013

DOI: 10.5772/54546

Downloaded: 782

1. Introduction

1.1. The mass transfer process with minimum irreversibility

In many processes, heat and mass transfer are distributed in time or space. The problem of thermodynamically perfect organization lies in the choice of such concentration and temperature change, in space or time, laws to minimize the entropy production σ. Below we consider stationary processes and a spatial distribution, for definiteness.

1.1.1. Optimal organization of an irreversible mass transfer process

Consider the irreversible process of mass transfer, in which from one flow to another one substance is transmitted. The problem of minimal irreversibility of this process at a given average intensity of mass transfer takes the form:

σ=0Lg(c1,c2)T(μ1(c1)μ2(c2))dlminE1

under conditions

0Lg(c1,c2)dl=g¯;E2
d(G1c1)dl=dN1dl=g(c1,c2);N1(0)N10.E3

The minimum is searched by selecting the concentration c2(l). change law. (In the equation (1) μi(ci)denotes the chemical potential of the i-th flow dependence on the concentration of the redistributed substance in it). Here G1and c1is an amount and molarity of the redistributed substance respectively, N1-number of moles of that component.

Minimal irreversibility conditions of mass transfer arise from the solution of (1) — (3). They can be described as follows [24]: In the mass transfer process with minimum irreversibility the ratio of flow g and chemical potential μ2 derivatives with respect to the concentration c2 is proportional to the ratio of relative flow g square to the temperature, in any cross-section of l.

g/c2μ2/c2=ξg2TE4

Indeed, the entropy production after transition from dtto dNis:

σ=N10N1L1T(μ1(c1)μ2(c2))dN1minE5

under the condition

N10N1L1g(c1,c2)dN1=L.E6

The Lagrange function of problem (5), (6) takes the form

R=1T(μ1(c1)μ2(c2))+λ1g(c1,c2),E7

stationarity conditions with respect to c2:

Rc2=01Tμ2c2+λ1g2(c1,c2)gc2=0E8

lead to the equation (4). The proportionality coefficient ξin (4) is defined from the initial data of the current problem.

For a specific task g(c1,c2)condition (4) allows us to find a relation between the c1and c2for the optimal mass transfer organization. For example, if

g=k(μ1(c1)Tμ2(c2)T),E9

we'll get the following equation from (4) :

μ1(c1)μ2(c2)=const.E10

At the constant temperature and pressure, this condition leads to the equation

c1(l)/c2(l)=constE11

and the constancy of flow g(c1,c2)for any l.

For the mass transfer law of the form

g=k(c1(l)c2(l))E12

Derivatives are

g/c2=K;μ2c2=RTc2.E13

After their substitution into (4), we obtain

(c1(l)c2(l))2/c2(l)=const.E14

During the mass transfer between phases the driving force of the process is expressed as the difference between the concentration of a redistributed component in one phase c1and the equilibrium concentration c1p(c2)linearly independent of c2(concentration of the same component in another phase). In this case, c1pis substituted in (11) or (14) instead of concentration c2.

1.1.2. Example

Let optimality conditions of irreversible mass transfer have the form (11). From the view of flow g=g¯/Lconstancy from (3) follows:

c1*(l)=c10g¯rL;c1p*(l)=c10c1*(l)M,E15

where M denotes the right side of (11). Substituting c1*and c1p*in the expression for the mass flux and taking the constancy of this flux into account, we obtain

kT(μ1(c1*)μ1(c1p*))=g¯L,E16

Or kRlnM=g¯/L,from which M=eg¯/kLR.

Assuming a linear dependence

c1p(c2)=ac2+b,E17

where a, b — are some constants determined by processing experimental data of the equilibrium. Then, find the optimum profile

c2*(l)=c10rLg¯larLeg¯/kRLba.E18

2. Irreversible work of separation and heat-driven separation

2.1. Introduction

The minimal amount of energy needed for separation a mixture with a given composition can be estimated using reversible thermodynamics. These estimates turn out to be very loose and unrealistic. They also do not take into account kinetic factors (laws and coefficients of heat and mass transfer, productivity of the system, etc.). In this paper we derive irreversible estimates of the work of separation that take into account all these factors.

The majority of separation systems are open systems that exchange mass and energy with the environment. If mass and heat transfer coefficients (determined by the size and construction of the apparatus) are finite and if the productivity of the system is finite then the processes in such systems are reversible. The energy flows, the compositions of the mass flows, and the productivity of the system are linked via the balance equations of energy, mass, and entropy. The latter also includes entropy production in the system. Minimal energy used for separation corresponds to minimal entropy production in the system subject to various constraints. This allows us to estimate this minimal energy.

There is a qualitative as well as a quantitative difference between the reversible and irreversible estimates obtained in this paper. For example, the irreversible estimate of the work of separation for poor mixtures (where the concentration of one of the components is close to one) tends to a finite nonzero limit, which depends on the kinetics factors. The reversible work of separation for such mixtures tends to zero. The reversible estimate differs from the amount of energy needed in practice for separation of poor mixtures by a factor of 105.

For heat-driven separation processes the novel results obtained in this paper include the estimate of the minimal heat consumption as a function of kinetic factors and the thermodynamic limit on the productivity of a heat-driven separation.

2.2. Thermodynamic balances of Separation Processes and the Link between Energy Consumption and Entropy Production

Consider the system, shown in Figure 1, where the flow of mixture with rateg0, composition x0, temperature T0, and pressure P0is separated into two flows with the corresponding parameters gi, xi, Ti,Pi(i=1,2). The flow of heat q+with the temperature T+can be supplied, and the flow of heat qwith the temperature Tcan be removed. The mechanical work with the rate (power) pcan be supplied.

In centrifuging, membrane separation, and adsorption–desorption cycles that are driven by pressure variations, no heat is supplied/removed and only mechanical work is spent. In absorption–desorption cycles, distillation, and so forth, no mechanical work is spent, only heat is consumed (heat-driven separation). In some cases the number of input and output flows can be larger. As a rule one can still represent the system as an assembly of separate blocks, whose structure is shown in Figure 1.

Figure 1.

Simplified schematic of thermodynamic balances for separation processes.

2.2.1. Heat-driven separation

Consider a heat-driven separation (p=0)and assume that each of the vectors xi=(xi1,....,xij,....,xik)i=0,1,2consists of kcomponents which denote the molar fraction of the j-th substance in the i-th flow. The thermodynamic balance equations of mass, energy, and entropy here take the following form

g0x0jg1x1jg2x2j=0,j=1,...,kE19
j=1kxij=1,i=0,1,2E20
q+q+g0h0g1h1g2h2=0E21

where hiis the enthalpy of the i-th flow;

q+T+qT+g0s0g1s1g2s2+σ=0E22

σdenotes entropy production. From eq (19), eq (20) follows that g0=g1+g2. After elimination of g0from eqs (21) and (22) and introduction of enthalpy increments Δhand entropy increment Δswe get

q+q+g1Δh01+g2Δh02=0E23
g2Δs02+g1Δs01+q+T+qT+σ=0E24

Here, Δh0i=h0hi,Δs0i=s0si(i=1,2).

Elimination of qusing eq (23) and its substitution into eq (24) yields

i=12gi(Δs0iΔh0iT)+q+(1T+1T)+σ=0

and the flow of used heat for heat-driven separation is

q+=T+T+T[i=12gi(ΔsoiTΔhoi)+σT]E25

The first term in the square brackets depends only on the parameters of the input and output flows and represents the reversible work of separation per unit of time (reversible power of separation). The second term there represents the process kinetics and corresponding energy dissipation.

For mixtures that are close to ideal gases and ideal solutions, molar enthalpies and entropies hiand siin the eqs (21) and (22) can be expressed in terms of compositions and specific enthalpies and entropies of the pure substances. We obtain for each of the flows

Δh0i=j=1k[x0jhj(T0,P0)xijhj(Ti,Pi)]Δs0i=j=1k[x0jsj0(T0,P0)xijsj0(Ti,Pi)R(x0jlnx0jxijlnxij)],i=1,2E26

where R is the universal gas constant. The reversible energy consumption here is

q+0=1ηki=12gij=1k[[x0jsj0(T0,P0)xijsj0(Ti,Pi)R(x0jlnx0jxijlnxij)]T+xijhj(Ti,Pi)x0jhj(T0,P0)]E27

We denote here the Carnot efficiency of the ideal cycle of the heat engine as

ηC=T+TT+

Condition (25) can be rewritten as

q+=1ηC(p0+σT)E28

Here, p0is the reversible power of separation that is equal to the reversible flow of heat given by eq (27) multiplied by the Carnot efficiency. When eq (28) was derived we took into account only the irreversibility σof the separation process (the irreversibility of the heat transfer was not taken into account). In reality heat can be supplied/removed with a finite rate only irreversibly. Any transformation of heat into work with finite heat transfer coefficients and finite power is irreversible. This leads to a lower efficiency than the Carnot efficiency. The closed form expression for this efficiency was obtained in ref [16]. It depends on the power p and on heat transfer coefficients for heat supply and heat removal α+and α. For the Newton (linear) law of heat transfer it has the form

ηp=maxpq+=112T+(T++T4pα(T+T)2+(4pα)28pα(T++T))E29

where it is assumed that there is constant contact of the working body with the heat reservoirs and

α=4α+αα++αE30

It is easy to show that if p0then ηptends to the Carnot efficiency.

Substitution of ηpinstead of ηCin eq (28) allows us to derive a tighter estimate for the heat consumption in heat-driven separation processes by finding the minimal possible entropy production σsubject to various constraints

q+q+min=pminηp(pmin,α,T+,T)E31

where

pmin=p0+σminTE32

Conditions (29-31) single out the area of thermodynamically feasible heat-driven separation systems.

Expressions (27) and (28) and eq (25) can be further specified by assuming the constancy of heat capacities, that the mixture is binary, and so forth.

2.2.2. Mechanical separation

Consider a separation system that uses mechanical work with rate p. Assume that no heat is supplied/removed (q+=q=0)and that input and output flows have the same temperature T and the same pressure. Multiplication of eq (24) by T and subtraction of the result from the energy balance eq (23), where (q+q)is replaced with the supplied power p, yields

p=Tσ+g0i=12γi(TΔs0iΔh0i)E33

here γi=gi/g0

After taking into account eq (27) that the enthalpy increment Δh0iin a mechanical separation is zero, we get

p=g0RT[i=12γij=1kxijlnxijj=1kx0jlnx0j]+Tσ=p0+TσE34

The first term in this expression represents the minimal power for separation that corresponds to the reversible process (σ=0). This power p0is equal to the difference between the reversible power for complete separation of the input flow p00=g0RTjx0jlnx0jand the combined reversible power of separation of the output flows p10and p20.

Here

pi0(xi)=RTg0γij=1kxijlnxij,i=0,1,2E35

is the reversible power of separation of the i-th flow into pure substances.

2.3. Minimal work of separation in irreversible processes

2.3.1. Assumptions and problem formulation

Assume that the components of the input mixture are close to ideal gases or ideal solutions. The chemical potential of the i-th component can then be written in the following form

μi(T,P)=μ0(T,P)+RTlnxi,i=1,...,kE36

where xiis the concentration of the i-th component.

First we consider a system that includes three elements, a reservoir with the time independent temperature T, pressure P, and vector of concentrations x0={x01,...,x0k}(therefore its chemical potential μ0is also time independent), the finite capacity output subsystem with chemical potential μ1that depends on the current compositions of the mixture and of the working body that has controllable values of chemical potential μowand μ1w, at the points of contact with reservoir and output subsystem. At the time the intensive variables of the output subsystem coincide with the values of the reservoir’s intensive variables, and the number of moles in it is given and equal to N0. At time τthe number of moles N(τ)and the composition x(τ)in the output subsystem are given. The mass transfer coefficients between the reservoir and the working body and the working body and the output subsystem are finite and fixed. The minimal necessary work required for the separation is sought.

We do not consider here how to implement the derived optimal dependence of the chemical potential of the working body because of two reasons. First, our main objective is to derive a lower bound on the work of separation. However, imposing constraints on feasible variations of chemical potential would lead to an increase in energy consumption. Second, we will demonstrate that for the majority of mass transfer laws the optimal mass transfer flow is time independent, and its implementation is straightforward.

The work of separation in an isothermal process for an adiabatically insulated system can be found using the Stodola formula in terms of the reversible work A0and the entropy increment ΔS

A=A0+TΔSE37

The reversible work is equal to the increment of the system’s internal energy. Since as a result of the process (N(τ)N(0))moles of mixture with the composition x0is removed from the reservoir, and the energy of the output subsystem rises because of the increase of the amount of moles in it from N(0)to N(τ)and its composition from x0to xτ, the total change of the system’s internal energy is

A0=N(τ)i=1kΔμi=N(τ)RTi=1k[xi(τ)lnxi(τ)xi0lnxi0]E38

and it is independent of N(0). Because A0is determined by N, x(τ), x(0), the minimum of A corresponds to the minimum of the entropy increment

ΔS=1T0τi=1k[g0i(μ0iμiw)+g1i(μiwμ1i)]dt=1T0τi=1k(g0iΔμ0i+g1iΔμ1i)dtE39

Because the working body’s parameters have the same values at the beginning and at the end of a cycle

0τgi0dt=0τgi1dtN(τ)xi(τ)N(0)xi(0)=Δ(Nxi),i=1,2,...,kE40

2.3.2. Optimal solution

The problem of minimization of ΔSsubject to constraints eq (40) on g0i0, g1i0becomes simpler in a common case where the chemical potentials’ increments Δμ0i, Δμ1iare unique functions of flows g0iand g1i, correspondingly. If processes are close to equilibrium then this dependence is linear.

Assume

Δμ0i=ϕ0i(g0i),Δμ1i=ϕ1i(g1i)

then the problems (39) and (40) can be decomposed into 2k problems

ΔSji=0τσji(gji)dtmin/0τgjidt=Δ(Nxi)j=0,1,i=1,2,...,kE41

where σji=gjiϕji(gji)is the function that determines dissipation.

Problems eq (41) are averaged nonlinear programming problems. Their optimal solutions gji*are either constants and equal to

gji*=g1i*=Δ(Nxi)τE42

or switches between two so-called basic values on the interval (0,τ), the solution eq (42) corresponding to the case where the convex envelope of the function σji(gji)is lower than the value of this function at gji*. Characteristic forms of the function σji(gji)for the constant and switching regimes are shown in Figure 2.

Figure 2.

Dependence of the entropy production on the rate for the constant (a) and switching (b) solutions (g1i* and g2i* are the basic values of the rate).

If the function σjiis concave then the optimal rate gjiis always constant. Let us calculate the second derivative of σon g (we omit subscripts for simplicity). If it is positive then the constancy of the rate in the optimal process is guaranteed.

σ(g)=2ϕ(g)+gϕ(g)0E43

The first term in this expression is always positive because the chemical potentials’ difference is the driving force of mass transfer and monotonically depends on the flow. For the majority of laws of mass transfer the inequality eq (43) holds. In particular, it holds if the flow of mass transfer is proportional to the difference of chemical potentials in any positive degree.

Consider mass transfer flow that depends linearly on the chemical potential difference for all i, j. Then

gji=αjiΔμjiϕji=gjiαjiE44

It is clear that the conditions eq (43) hold and the optimal rates of flows obey equalities (42).

Equalities (42) hold for any nonswitching solution. The minimal increment of the entropy production for such solution is

ΔSmin=ijΔSjimin=τijσji(Δ(Nxi)τ)E45

and the minimal work of separation is

Amin=A0+τTijσji(Δ(Nxi)τ)E46

The optimal rates are determined by the initial and final states which allows us to specify the estimate eq (46).

Near equilibrium the flows obey Onsanger’s kinetics eq (44), and from eq (46) it follows that

Amin=A0+τi=1kgi2(1α0i+1α1i)=A0+1τi=1kΔ2(Nxi)α¯iE47
α¯i=α0iα1iα0i+α1iE48

is the equivalent mass transfer coefficient on the i-th component and the minimal entropy production is

σmin=1Tτ2i=1kΔ2(Nxi)α¯iE49

The lower bound for the average power of separation is

pmin=Aminτ=A0τ+1τ2i=1kΔ2(Nxi)α¯iE50

p0=A0/τis the reversible power of separation.

If

N(0)=0,Δ(Nxi)=Nxi(τ)

then expressions (47) and (50) take the form

Amin=A0+N2τi=1kxi2(τ)α¯iE51
pmin=p0+g2i=1kxi2(τ)α¯iE52

Where

A0=NRTi=1k[xi(τ)lnxi(τ)xilnxi]E53

Figure 3.

Reversible (A0) and irreversible (Ar) estimates of the minimal work of separation of binary mixture as functions of key component’s concentrations.

Figure 4.

Separation of the system with finite capacity on m subsystems.

Note that the irreversible estimate of the work of separation eq (51) does not tend to zero for poor mixtures when the concentration of one of the components tends to one (Figure 3).

If system includes not one but a number of output subsystems then it is clear that the estimate for the minimal work of separation is equal to the sum of the estimates for each subsystem.

Amin=j=1Aminj,pmin=j=1pminjE54

The superscript j here denotes the subsystems.

2.3.3. Separation of a System with finite capacity into m subsystems

Consider a system that is shown in Figure 4. Its initial state is described by the vector of concentrations x0, the number of moles of the mixture N0, and its final state by the number of moles Nj, j=1,...,min each of the subsystems and their concentrations, xj. The mass balances yields

j=1mN¯j=N0j=1mN¯jx¯ji=N0x0i,i=1,2,...,kE55

The work in the reversible separation process here is

Ar0(x0,x¯)=RT[j=1mN¯jix¯jilnx¯jiN0ix0ilnx0i]=Ar00(x0,N0)j=1mArj0(x¯j,N¯j)E56

The reversible work of separation is equal to the difference of the reversible work of separation of the initial mixture into pure components and the reversible work of separation for mixtures in each of the subsystems.

We again assume that flows gjhave components gjiproportional to the difference of the chemical potential of the subsystem and the working body with the coefficient αji. Here, the condition of minimal work of separation corresponds to the condition of flow constancy

gji=N¯jx¯jiτ,i=1,2,...,k,j=1,...,mE57
Δμji=gjiα¯ji,j=0,1,...,mE58

Here, α¯jiis the equivalent mass transfer coefficient calculated using eq (48) for the flow into the j-th output subsystem of the i-th component. Similarly as was done above for the system with the reservoir and one finite capacity output subsystem and flows proportional to the final concentrations eq (57), these concentrations in the output subsystems are time independent and equal to x¯j, correspondingly, and the number of moles N¯j(t)depends linearly on time. The power p here is constant

p=RTτj=1mN¯jix¯jilnx¯jix0i+1τ2j=1mNj2ix¯ji2/α¯jiE59

The minimal work of separation for the mixture with concentrations x0into m subsystems with concentrations x¯iover the time τis

Ar=RTN0j=1mγjix¯jilnx¯jix0i+N02τj=1mγj2ix¯ji2/α¯jiE60

Here, γj=Nj/N0, α¯ji=αjiα0i/(α0i+αji)

The first term here coincides with the reversible work of separation Ar0of the mixture of N0moles with concentration x0into subsystems with number of moles N¯jand concentrations x¯j. The second term takes into account irreversibility of the process. Ardecreases monotonically and tends to Ar0when process duration τand mass transfer coefficient α¯jiincreases.

2.3.4. Example

Consider separation of the binary mixture into pure components in time τ. In this case N1=x0N0, N2=(1x0)N0, where x0is the concentration of the key component, x¯11=x¯22=1. From the formula (60) we get

Ar=RTN0(x0lnx0+(1x0)ln(1x0))+N02τ(x02α¯11+(1x0)2α¯22)=Ar0(x0)+N02τ(x02α¯11+(1x0)2α¯22)E61

The estimate eq (61) was derived in ref [1] by solving the problem of optimal separation of the binary mixture in the given time τin Van’t Hoff’s thought experiment with movable pistons and semitransparent membrane where α¯11and α¯22are the permeability coefficients on the first and second component. If flows do not depend explicitly on the chemical potentials’ differentials, for example, are proportional to the concentrations’ differential, then an estimate similar to the one obtained above can be constructed by solving the following auxiliary nonlinear programming problem

Δμi(P0i,Pi)minP0i,Pi/gi(P0i,Pi)=gi,i=1,2,...E62

Here, (P0i,Pi)are partial pressures of the components in contacting subsystems that depend on the chemical potentials’ differentials Δμi. The flow gidepends on the same differentials. Minimums in these problems are sought for different values of constant gi>0and nonpositive P0iand PiWe denote the minimal values of the objective in each of these problems Δμimin(gi)as Δμi*(gi). This dependence can be used in the estimate eq (41) of the irreversible work of separation.

2.3.5. Example

Assume Δμ=RTln(P0/P), g(P0,P)=(P0P)/α, and 0<P<Pmax. Let us express P0in terms of g and P:

P0i=αigi+Pi,i=1,2

Δμ=RTln(αg/P+1)attains its minimum at P=Pmaxg.

Therefore, Δμi*(gi)=RTln(αigi/Pmax+1).

2.4. Potential application of obtained estimates

We will illustrate the possibilities of the application of the derived estimates.

2.4.1. Estimate of the power of separation in a continuous separation system

Consider a continuous separation system with the input flow g0with concentration x0and m output flows gj(j=1,...,m)with concentrations xj={xj0,xj1,...,xjk}. Here, the temperatures on the input and output flows are close to each other.

Equation (59) allows us to estimate the minimal power required for continuous separation in such system

pmin=j=1mp0j+g02j=1mγj2i=1kxji2αjiE63

Where

γj=gjg00,j=1mγj=1E64
p0j=g0γjRTi=1k[xjilnxjix0ilnx0i]p0j=γjMj(g0,xj)E65

Mass balance equations yield

j=1mγjxji=x0i,i=1,...,k1,i=1kxji=1,j=0,...,mE66

The number of conditions eq (66) is k1, because the concentration of one of the components is determined by the conditions eq (64).

If the number of flows m>k, and their compositions are given, then the removal fractions can be chosen in such a way that the power of separation is minimal subject to constraints eqs (64) and (66). The Lagrange function of this problem is

L=j=1m{γjMj+γj2rjλ0γji=1kλiγjxji}E67

here

rj(g0,xj)=g02i=1kxjiαji

L is the concave function on γj, and its conditions of stationarity determine the flows that minimize the power for separation for a given flow’s compositions

γj*=λ0Mj+i=1k1λixji2rj,j=1,...,mE68

We have k linear equations for λ0and λi

12[j=1mλ0Mjrj+i=1k1λij=1mxjirj]=1E69
12[j=1mxji(λ0Mjrj+1rji=1k1λixji)]=x0i,i=1,...,k1E70

2.4.2. Example

Assume m=3, k=2, g0=1mol/s, T=300K, and the compositions and transfer coefficients are

x01=x02=0.5x11=0.9;x12=0.1;α¯11=α¯12=0.004mol2/(J s)x21=0.3;x22=0.7;α¯21=α¯22=0.01mol2/(J s)x31=0.1;x32=0.9;α¯31=α¯32=0.06mol2/(J s)

From eq (65) we obtain M1=910, M2=197, M3=910, and r1=205, r2=580, r3=137.

Equations (69) and (70) for λ-multipliers take the form

12[λ0M1r1+λ0M2r2+λ0M3r3+λ(x11r1+x21r2+x31r3)]=112[x11(λ0M1r1+λ1x11r1)+x21(λ0M2r2+λ1x21r2)+x31(λ0M3r3+λ1x31r3)]=x01

We obtain λ0=894,λ1=183. Their substitution in eq (68) yields γ1*=0.36, γ2*=0.64, γ3*=0and the corresponding estimate for the minimal irreversible power of separation eq (63) is

pmin=718wt

2.4.3. The selection of the separation sequence for a multicomponent mixture

In practice, separation of multicomponent mixtures is often realized via a sequence of binary separations. So, a three-component mixture is first separated into two flows, one of which does not contain one of the components. The second flow is then separated into two unicomponent flows. The reversible work of separation (that corresponds to the power p0) does not depend on the sequence of separation, because p0is determined by the rates and compositions of the input and output flows of the system as a whole. The irreversible component of the power Δpin eq (63) depends on the sequence of separation and can be used to find the optimal one.

Consider a three-component mixture with concentration x0=(x01,x02,x03), and rate g0we set to one. We denote the mass transfer coefficients at the first and second stages of separation as α1and α2. They depend on the construction of the apparatus. First, we assume for simplicity that these coefficients do not depend on the mixture’s composition (in the general case they do depend on it). We consider irreversible power consumption for two cases:

  1. The first component is first separated, then the second and the third are separated.

  2. The second component is separated, and then the first and the third are separated.

We assume that the separation at each stage is complete. We get up to the constant multiplier

Δpa=Δpa1+Δpa2=x012/α1+(x02+x03)2α1+(x02+x03)2+(x022/α2+x032/α3)E71

The first two terms in this sum represent the loss of irreversibility during the first stage of separation. For g0=1and complete separation the output rates of this stage g1and g2are x01and (x02+x03), correspondingly.

Consider the first stage of case a for g0=1and complete separation and view the second and third component as the same substance with the output rate x02+x03=1x01. The irreversible expenses eq (63) are

Δpa1=x012α1+(1x01)2α1=2x012+12x01α1E72

When the second flow is separated into two flows their rates are

g22=x02(1x01),g23=x03(1x01)

and the irreversible power is

Δpa2=1α2(1x01)2(x022+x032)

The combined irreversible power is

Δpa(x01,x02)=2x0122x01+1α1+x022+(1x01x02)2α2(1x01)2

Similarly in case b we get

Δpb(x01,x02)=2x0222x02+1α1+x012+(1x01x02)2α2(1x02)2

The differential between these two values is

Δpab=Δpa+Δpb=2α1[(x012x022)(x01x02)]+1α2(1x01)(1x02)[(1x02)2(x022+x032)(1x01)2(x012+x032)]E73

If Δpab>0, then sequence b is preferable.

Note that it is not possible to formulate the general rule to choose the optimal separation sequence for a multicomponent mixture, in particular, on the basis of the reversible work of separation. It is necessary here to compare irreversible losses for each sequence.

2.4.4. Example

Assume that the composition of the input three component mixture is x01=0.6, x02=0.3, x03=1x01x02; the mass transfer coefficients are α1=0.01mol2/(J s), α2=0.02mol2/(J s). From (eq 73) we find that the difference in power between sequences a and b is

Δpab=ΔpaΔpb=7.82J

The comparison of the combined minimal irreversible power for the same initial data shows that the power for separation of a mixture using sequence b is higher than the power used for sequence a, that is, Δpab<0.

Thus, sequence a is preferable, and it is better to perform the complete separation by separating the first component.

2.5. Limiting productivity and minimal heat consumption for a heat-driven separation

In many separation processes a heat engine is used to create the differential of the chemical potential between the working body and the reservoirs (the driving force of mass transfer). Here, the working body is heated during contact with one reservoir and is cooled during contact with the other reservoir. One can represent the heat-driven separation system as a transformer of heat into the work of separation that generates power p, consumes heat flow from hot reservoir g+, and rejects flow gto the cold reservoir. Heat transfer coefficients for contacts with the hot and cold reservoir α+and αare fixed.

It was shown in refs [12] and [6] that the potential of the direct transformation of heat to work is limited and the maximal generated power for the working body with the distributed parameters is

pmax=α¯(T+T)2E74

In this expression α¯=(α+α)/(α++α)is the equivalent heat transfer coefficient for continuous contact with the reservoirs; α¯=(α+α)/(α++α)2is the equivalent heat transfer coefficient for sequential contact.

The maximal power determines the heat flow consumed from the hot reservoir. Further increase of heat consumption for given values of heat transfer coefficients requires an increase of the temperature differential between the reservoirs and the working body and reduces the power.

The dependence of the used power on the productivity of irreversible separation processes is monotonic eq (63). Therefore, the limiting productivity of heat-driven separation processes corresponds to the maximal possible power produced by transformation of heat into work. Further increase of heat consumption q+reduces power and therefore reduces the productivity of separation process.

For the Newton (linear) law of mass transfer and heat–work transformer the dependence of the power on the heat used is

q+(p)=pηp=2p(pα¯T++ηC)+(pα¯T++ηC)24pα¯T+E75

Here, ηC=(T+T)/T+is the Carnot efficiency, T+and Tare the hot and cold reservoir’s temperatures, and α¯=(α+α)/(α++α)is the equivalent heat transfer coefficient.

The minimal heat consumption q+as a function of productivity g0for a heat-driven separation can be obtained by substituting expression (75) instead of p in the right-hand side of eq (63). The result holds for ppmaxand therefore for g0g0max. The duration here must not exceed the maximal possible duration.

Substitution of the right-hand side of eq (74) instead of p in eq (63) yields the maximal possible productivity of the system (where α¯is chosen according to the type of contact between the transformer and reservoir). We denote

B=RTjγj2ixjilnxjix0i,D=Tjγj2ilnxji2αjiE76

We obtain

pmax=α¯(T+T)2=Bgmax+Dg0max2

and the limiting productivity is

g0max=B+B2+4α¯D(T+T)22DE77

Formulas (76) and (77) allow us to estimate the limiting productivity of a heat-driven separation process for Newton’s laws of heat transfer between the working body and reservoirs and mass transfer proportional to the differentials in chemical potentials (mass transfer is close to isothermal with the temperature T).

2.5.1. Example

Consider heat-driven monoethanamide gas cleansing. One of the components is absorbed by the cold solution from the input gas mixture. This solution is then heated and this component is vaporized. The input mixture’s parameters are T¯=350K, the key component’s molar concentration x=0.5, the rate of mixture g0=5mol/s. The temperatures of heat supplied/removed are correspondingly Th=400K, Tc=300K, and the heat transfer coefficients are α+=8.368kJ/(s K) and α=16.736kJ/(s K). The concentrations of the key components in the output flows are x1=0.9, x1=0.1; the mass transfer coefficients for each of the components (integral values over the whole contact surface) for the hot and cold reservoir’s contacts are α1=0.07mol2/(kg s), α2=0.03mol2/(kg s).

Because the solution circulates and is heated and cooled in turns, the limiting power for transformation of heat into work is given by the expression (74) with the corresponding α¯

pmax=20.711kJ/s

The power for separation is given by eq (63).

We have

p0=RTg0j=1mγjixjilnxjix0i=5.397kJ/s

The minimal work required for a system with Onsanger’s equations are (see eq (63))

Δp=g02j=1mγj2ixji2αji=7.238kJ/s

Thus, p=p0+Δp=12.636kJ/s <pmax. The work needed for separation does not exceed the maximal possible value for given heat transfer coefficients.

Let us estimate the minimal heat consumption. From eq (75) we get

q+=32.426kJ/s

If the temperatures of the input and output flows are not the same then the minimal energy required for separation can be estimated using the thermodynamic balance equations (31) and (32) and the expression for σmineq (49).

2.6. Conclusion

New irreversible estimates of the in-principle limiting possibilities of separation processes are derived in this paper. They take into account the unavoidable irreversibility caused by the finite rate of flows and heat and mass transfer coefficients. They also allow us to estimate the limiting productivity of a heat-driven separation and to find the most energy efficient separation sequence/regime of separation for a multicomponent mixture.

3. Optimization of membrane separations

3.1. Introduction

As the properties of membranes improve, the membrane separation of liquids and gases is more widely used in chemical engineering [8,10,11,20]. Since the mathematical modeling of membrane separations is simpler than that for most of the other separation processes, they could be controlled by varying the pressure, contact surface area, and the like during the separation process.

The minimal work needed to separate mixtures into pure components or into mixtures of given compositions can be minorized using well-known relationships of reversible thermodynamics [15]. However, this estimate is not accurate because it ignores the mass transfer laws and the properties of membranes, process productivity, possible intermediate processes of mixing, and so on. The estimates based on reversible thermodynamics are not suitable for determining the optimal sequence of operations in the separation of multicomponent systems, because they depend only on the compositions of feeds and end products and do not reflect the sequence of operations in which the end product was obtained. The work needed for separation consists of its reversible work and irreversible energy losses. The losses are equal to ΔST, where ΔSis the increment of the system entropy due to the irreversibility of the process. Below, the minimum possible production of entropy (that is, the minimal additional separation work) will be found for the separation of one component at a specified production rate and transport coefficients. Also, we will determine the dependence of this minimum on the input data for one or another process flowsheet at a fixed production rate.

3.2. Batch membrane separation

We will first consider a batch separation of a mixture in a system consisting of two chambers separated by a membrane permeable to only one active (to be separated) component of the mixture (Fig. 5). Let and Gi(t), Ci(t), μi(Ci,and Pi)denote the amount, the concentration of the active component, and its chemical potential in chamber I, respectively. These parameters can be varied during the process. At the initial moment of time (t=0), the parameters that are specified include the mixture amount G1(0)=G10in chamber 1 and the concentration C1(0)=C10of the active component passing through membrane 3 at a mass transfer rate g, which depends on its chemical potentials on both membrane sides, μ1and μ2. In turn, the potentials depend on the variation of the pressure and mixture composition in the first and second chambers. The pressure in the first chamber can be varied using piston 4. The process is isothermal, and the temperature T is specified and remains unchanged.

The intensive variables in the second chamber are the pressure P2(t)and the chemical potential μ2(t), which varies with time due to the accumulation of the active component in the chamber and the variation of the external conditions. Assume that the laws of this variation are known. The specification of the initial composition of the mixture C10, the number of moles G of the component that passed through the membrane in time τ, and the initial number of moles G10determines the final composition in the first chamber,

C1(τ)=G10C10GG10G,

and, hence, the reversible work of separation, which is equal to the increment of the free energy of the system:

A0=G10RT[C10lnC10+(1C10)ln(1C10)]+(G10G)RT{C1(τ)lnC1(τ)+[1C1(τ)]ln[1C1(τ)]}.E78

Figure 5.

Batch separation of a mixture: 1, chamber with the mixture to be separated; 2, chamber to which the active component passes; 3, membrane; 4, piston.

Consequently, the minimum of the produced work corresponds to the minimum of the irreversible losses of energy, which is proportional to ΔS.

The increment of entropy in the system, the minimum of which should be determined for a separation process of duration τ, is equal to the product of the flux and driving force:

ΔS=1T0τg(μ1,μ2)(μ1μ2)dtmin.E79

The amount of the active component that passed through the membrane is written as

G=0τg(μ1,μ2)dt.E80

The process duration τwill be fixed.

The variation of G1and concentration C1are determined by the equation:

d(G1C1)dt=dG1dt=g(μ1,μ2).E81

It follows from Eq. (81) that d[G1(1C1)]dt=0, implying that G1(t)(1C1(t))=constfor any moment of time. The latter is equal to the amount of the “inert” component of the mixture in the first chamber. It will be denoted as G˜=G1(0)[1C1(0)].

The solution of Eq. (81) determines the dependence of the mixture amount in the first chamber on the active component concentration G1(C1):

G1(t)=G˜1C1(t).E82

After expression (82) is substituted into Eq. (81), the latter takes the form

dC1dt=1G˜1(1C1)2g(μ1,μ2),C1(10)=C10.E83

First, we will find such time variation of, chemical potential μ1(t)that the increment of entropy takes the minimum value at a specified value of G. Then, for a specific form of chemical potential, we will find the time variation of pressure P1(t)corresponding to the found optimal variation of the chemical potential.

We will write the Lagrangian function F for the problem given by Eqs. (79) and (80) in view of the fact that the constant factor 1/T does not affect the optimality condition:

F=g(μ1,μ2)(μ1μ2λ).

The mass transfer rate g is equal to zero when μ1=μ2and increases monotonically with increasing μ1. As a result, the function F is, as a rule, convex with respect to μ1. Consequently, this dictates the stationary of F in the solution of the problem and this solution is unique:

Fμ1=0gμ1(μ1μ2λ)=g(μ1,μ2).

To cancel out λ, we integrate the both sides of this equality from zero to τin view of Eq. (80) to obtain

λ=G+0τgμ1(μ1μ2)dt0τgμ1dt.

Consequently, to determine μ1(t)with a convex function F, we have the equation determining the optimal variation of μ2(t)in the function g(μ1,μ2)for any μ1(t)and mass transfer law μ2(t):

g(μ1,μ2)=gμ1[G+0τgμ1(μ1μ2)dt0τgμ1dt(μ1μ2)].E84

If the flux is proportional to the difference of chemical potentials,

g=α(μ1μ2),E85

it follows from optimality condition (84) that

μ1*(t)=μ2(t)+Gτα,g*=Gτ=const.E86

The variation of P1(t)corresponding to μ1*(t)depends on the form of the chemical potential.

For mixtures close in properties to ideal gases, the chemical potential (molar Gibbs energy) of the active component of the mixture is written as

μ1=μ11(T,P1)+RTlnC1=μ01(T)+RTlnP1+RTlnC1,E87

where μ01is the standard chemical potential for P1=C1=1.

The variation of C1*(t)(t) is determined by Eq. (83) with known mass transfer rate g. After C1*(t)and μ1*(t)are substituted into Eq. (87), we obtain an expression for the pressure in the first chamber:

P1*(t)=1C1*(t)exp[μ1*(t)μ01(T)RT].E88

For the flux defined by Eq. (85) and defined by Eq. (86), Eq. (83) takes the form:

dC1dt=G(1C1)2τG˜1=G(1C1)2G10(1C10)τ,C1(0)=C10.

The solution to this equation is written as

C1*(t)=G10C10GτtG10Gτt.E89

Substituting the latter into Eq. (88) gives the time variation of the pressure:

P1*(t)=G10GτtG10C1(0)Gτtexp(μ2+Gατμ10(T)RT).

After the optimal variation of μ1*(t), or optimal value of this chemical potential, is found, we can determine ΔSminby substituting μ1*and μ2into Eq. (79). Using the flux defined by Eq. (85) and relationship (79), we obtain

ΔSmin=G2Tατ.

Figure 6.

Optimal time variation of (a) pressure and (b) mole fraction of the active component in the first chamber for a gas mixture.

The optimal variation of the pressure and mole fraction of oxygen in the first chamber is shown in Fig. 6. It corresponds to the separation of a gas mixture composed of carbon dioxide, 120 moles of CO2, and oxygen, 180 moles of O2 (active component), when G10=300moles, C10=0.6, G=150moles, τ=90s, α=2.13×103mol2/(s J), P2=101330Pa, C2=1.0, and T=283K. At the moment when the process is terminated, C(τ)=0.2. The production of entropy is ΔS=ΔSmin=415J/K.

The produced work is A=A0+TΔS=415730J, where according to Eq. (78) A0=298300J.

Although the chemical potential for ideal solutions is written like Eq. (87), the function μ11(T,P1)for them takes a different form. This is caused by the fact that the chemical potential μ1(T,P1,x1)is the molar Gibbs energy of the active component and the derivative of the chemical potential with respect to pressure is the molar volume of this component v1[15]. In contrast to gases, the molar volume of liquids is virtually independent of pressure and varies vary little with temperature. As

μ1P1=μ11P1=v,

we obtain

μ1(T,P1,x1)=μ01(T)+vP1+RTlnC1.E90

Figure 7.

Optimal time variation of (a) pressure and (b) the mole fraction of the active component in the first chamber for a near-ideal solution.

For the flux defined by Eq. (85) and μ1*defined by Eq. (86), the variation of C1*(t)for liquids can be written in the same way as for gases in Eq. (89). After μ1*(t)and C1*(t)are substituted into Eq. (90), we obtain an equation for the optimal variation of pressure in the first chamber:

P1*(t)=1v1[μ2+Gταμ01(T)RTln(G10C1(0)GτtG10Gτt)].

For illustration, we considered the separation of water with a high salt concentration. Like ocean water, it contained 36 g/l of salt (inert component). The other process parameters were G10=552.3moles, C10=0.989, G=250moles, τ=3600s, α=9.92×104mol2/(s J), P2=101330Pa, C2=1.0, and T=283K. The time variation of the optimal pressure of the liquid and the mole fraction of water in the first chamber are illustrated in Fig. 7. At the moment when the process is terminated, C(τ)=0.979. The production of entropy is ΔS=ΔSmin=61.8J/K. The produced work is A=A0+TΔS=26470J, where according to Eq. (78) A0=8973J.

Figure 8.

Continuous separation of a mixture.

3.3. Membrane separation process distributed along the filter

The parameters of the system can vary with length rather than with time, as in the previous system. The flow diagram of this system is shown in Fig. 8. The mixture to be separated, which is characterized by a molar flux g1(0)=g10and concentration C1(0)=C10, is continuously supplied to the first chamber, the overall length of which is L. As the mixture travels over the length l, the active component passes across the membrane into the second chamber. The concentration of the active component in the mixture to be separated at the outlet of the first chamber is C1(L)=C1L. The chemical potential of this component in the second chamber, μ2(l), should be chosen so that in the isothermal process the increment of entropy in the system should be a minimum for the specified values of production rate g¯and total membrane surface area s(L). In irreversible continuous separation, the power p expended for separation is the sum of the reversible component

p0=g1(0)RT[C10lnC10+(1C10)ln(1C10)]+(g1(0)g¯)RT[C1(L)lnC1(L)+(1C1(L))ln(1C1(L))],E91

which is determined at the given conditions, and the irreversible losses pH=Tσ. Consequently, the minimal production of entropy σcorresponds to the minimal separation work p.

The flux of the component to be distributed at section l is equal to g[μ1(l),μ2(l)]. The production rate is specified as

0Lg(μ1,μ2)dl=g¯.E92

The production of entropy is determined by the expression

σ=1T0Lg(μ1,μ2)(μ1μ2)dlminμ2(l).E93

Assume that μ2(l)is the control parameter.

If the operating regime in the first chamber is close to plug flow, the material balance equations for section l give equations analogous to Eqs. (81).

ddl(C1g1)=ddlg1=g(μ1,μ2).E94

The above equation can be used to obtain a relationship analogous to Eq. (83):

dC1dl=(1C1)2g˜1g(μ1,μ2),C1(0)=C10,g1(l)=g˜11C1(l),E95

where g˜1=g1(0)(1C10)is the molar flux of the inert component through the first chamber.

Equations (92), (93), and (95) represent an optimal control problem in which C1is the state coordinate and the potential μ2is the control action. This problem can be simplified using the fact that for optimal processes the right-hand side of Eq. (95) never change the sign and C1monotonically varies with time. The independent variable l can be replaced by C1. It follows from Eq. (95) that

dl=g˜1dC1(1C1)2g(μ1,μ2).

In view of this replacement, the problem given by Eqs. (92), (93), and (95) can be written as

σ=g˜1TC1LC10(μ1μ2)dC1(1C1)2minμ2E96

with the constraints

C1LC10dC1(1C1)2=g¯g˜1,E97
C1LC10dC1(1C1)2g(μ1,μ2)=Lg˜1.E98

The concentration C1Lis determined by the initial concentration C10and production rate g¯in constraint (97) or (94). Using constraint (94), we obtain

C1Lg1(L)=C10g1(0)g¯,g1(L)=g1(0)g¯

and, hence,

C1L=C10g1(0)g¯g1(0)g¯.E99

The same follows from constraint (97) with g˜1[C10,g1(0)]. Consequently, after C1Lis found using constraint (99), constraint (97) can be ignored.

In distinction to batch membrane processes, the control action in a continuous membrane separation can be additionally represented by the coefficient of heat transfer α(l), because the membrane surface area can be varied from section to section, which corresponds to the variation of heat transfer coefficient α. Let αbe a function of C1. After α(C1)and C1(l)are found, we can pass to α(l). The mass transfer equation can be written as

g(μ1,μ2)=αg0(μ1,μ2)E100

where g0is called the specific mass transfer rate. In this case, the total surface area of the membrane and, hence, the overall value of the heat transfer coefficient will be bounded:

C1LC10α(C1)dC1=α¯.E101

In constraint (98), the mass transfer rate can be written as Eq. (100), and equality (101) can be added to the constraints of the problem. The resulting problem, given by Eqs. (96), (98), and (101), is an isoperimetric variation problem. The necessary condition for the optimality of its solution is the requirement that the Lagrangian function should be stationary with respect to μ2and α:

F=1(1C1)2[μ1(C1)μ2λ1αg0(μ1,μ2)]λ2α,

where the multipliers λ1and λ2correspond to constraints (98) and (101). The conditions for the stationary of F with respect to the desired variables are written as

Fμ2=0λ1g0/μ2αg02(μ1,μ2)=1,Fα=01(1C1)2λ1α2g0(μ1,μ2)=λ2.

The above equations give the process optimality conditions:

λ1g0/μ2α(C1)g02(μ1,μ2)=const=1λ1,E102
α2(C1)(1C1)2g0(μ1,μ2)=const=λ1λ2.E103

From constraints (98) and (103) we obtain

λ1λ2=g˜1C1LC10α(C1)dC1L=g˜1α¯L.E104

It follows from (101) and (102) that

1λ1=1α¯C1LC10g0/μ2g02(μ1,μ2)dC1.E105

After expressions (105) and (104) are substituted into conditions (102) and (103), respectively, we can use the known function μ1(C1)to find the functions μ2*(C1)and α*(C1)that are optimal in terms of minimal irreversibility, which with the help of Eq. (95) determine C1*(l)and, hence, μ2*(l)and α*(l).

Let us write the above relationships specifically for the function g written as a linear function of the difference of chemical potentials, Eq. (85), and chosen functions μi(Ci). Assume that the specific mass transfer rate takes the form:

g0(μ1,μ2)=μ1μ2.

Constraints (102)–(105) lead to the equations

1α(C1)(μ1(C1)μ2(C1))2=1α¯C1LC10dC1(μ1(C1)μ2(C1))2,E106
α(C1)(1C1)2[μ1(C1)μ2(C1)]=g˜1α¯L.E107

For brevity, we will introduce the notation μ1μ2=Δμand the right-hand sides in constraints (106) and (107) will be denoted as R1and R2. In this case, the above equations can be written as

1αΔμ2=R1,α2(1C1)Δμ=R2,

and we obtain

Δμ*(C1)=(1C1)2/3R21/3R12/3,α*(C1)=R11/3R22/3(1C1)4/3.E108

The concentration of the active component in the first chamber declines with increasing l. Therefore, under optimal operating conditions, Δμ*(l)=Δμ*[C1(l)]increases while the surface area of the membrane, which is proportional to α*(l)=α*[C1(l)], decreases.

To find R1, we will substitute Eq. (108) into Eq. (106) to obtain the equation

R14/3R22/3α¯C1LC10dC1(1C1)4/3=R1.

The evaluation of the integral gives us the desired formula:

R1=α¯L2B3g˜12,E109

where

B=(31C1(0)331C1(L)3).

Equation (108) in view of Eq. (109) yields the optimal dependence of the difference of chemical potentials on the concentration C1:

Δμ*(C1)=(1C1)2/3B2g˜1α¯L,E110
α*(C1)=α¯B(1C13)4.E111

Consequently, Eq. (95) takes the form:

dC1dl=(1C1)2g˜1α*(C1)Δμ*(C1)=(1C1)4/3BL.

Integrating this equation with specified initial conditions, we can find the variation of the concentration of the active component over the length of the first chamber under optimal operating conditions:

C1*(l)=127L3(3L1C1(0)3Bl)3.E112

Substituting this expression into Eqs. (110) and (111) yields the variation of the desired variables over the length:

Δμ*(l)=9LB2g˜1α¯(3L1C1(0)3Bl)2,E113
α*(l)=α¯(3L1C1(0)3Bl)434BL4.E114

The minimal value of the production of entropy corresponding to the above solution is written as

σ¯min=g˜TC1LC10Δμ*(C1)dC1(1C1)2=g˜12LTα¯B3.E115

We will introduce α, the mass transfer coefficient per unit area of the membrane surface, and ds(l), the elementary membrane surface area. If α=constant, then

α¯=α0Lds(l)=αs(L),

where s(L) is the total contact surface area.

If the specific mass transfer coefficient of the membrane material αand the total contact surface area s(L) are known, we can find the optimal distribution of the membrane surface area over the length of the filter:

ds*(l)=α*(l)α¯.

For near-ideal gas mixtures, we can write

μ1(C1)=μ10(P1,T)+RTlnC1,

where P1and T are assumed to be specified, and

μ2(P2,T,C2)=μ20(T)+RTlnP2+RTlnC2.E116

When Δμ*(C1)is known, expression (116) can be used to find the pressure function in the second chamber for which σminis achieved:

P2*(C1,C2)=1C2exp{1RT[μ1(C1)μ20(R)Δμ*(C1)]}.E117

The optimal curves for the pressure and mass transfer coefficient are plotted in Fig. 9, in which the data refer to the separation of a gas mixture composed of carbon dioxide CO2 and oxygen O2 (active component) whenC10=0.6, g1(0)=3.33mol/s, g¯=1.66mol/s, P1=2×106Pa, α¯=3.18×104mol2/(s J), α=0.013mol2/(s J), L=2.5m, and T=283K.

At the filter outlet, C1(L)=0.2. The production of entropy is σ=σmin=4.6J/(s K).

The consumed power is p=p0+Tσ=4600J/s, where according to Eq. (91) p0=3292J/s.

For ideal solutions, the calculation is almost the same except for the form in which the chemical potentials are written. For the first chamber,

μ1(C1)=μ10(T)=vP1+RTlnC1,

where v is the molar volume of the active component.

For the second chamber,

μ2(P2,T,C2)=μ20(T)+vP2+RTlnC2.

The dependence of the solution pressure in the second chamber on the concentration is written as

P2*(C1,C2)=1v[μ1(C1)μ20(T)Δμ*(C1)RTlnC2].

Figure 9.

Optimal variation of (a) pressure and (b) the membrane surface area over the filter length for a near-ideal gas mixture.

For illustration, we considered the separation of water with a high salt concentration. Like ocean water, it contained 36 g/l of salt (inert component). The other process parameters were g1(0)=0.153mol/s, C10=0.989, g¯=1.66mol/s, α¯=2.92×106mol2/(s J), α=0.011mol2/(s J), P1=7.7×106Pa, L=3m, and T=283K. The profile of optimal pressure in the first chamber and the variation of the mass transfer coefficient over the filter length are illustrated in Fig. 10. At the filter outlet, C1(L)=0.979. The production of entropy is σ=σmin=0.017J/(s K).

The consumed power is p=p0+Tσ=7.35J/s, where according to Eq. (91) p0=2.47J/s.

Figure 10.

Optimal variation of (a) pressure and (b) membrane surface area over the filter length for a near-ideal solution.

3.4. Conclusion

The minimal losses of energy for irreversible membrane separations with specified production rates are estimated. The variation of the driving force (difference of chemical potentials) and the distribution of the membrane surface area over the filter length corresponding to the process with minimal energy losses are found.

The obtained estimates can be used for assessing the deviation of the actual membrane separation from the optimal process and for comparing the thermodynamic efficiency of membrane separation processes with different flow diagrams, as well as for formulating and solving problems regarding the optimal sequence of operations in the separation of multicomponent mixtures.

4. Optimization of diffusion systems

4.1. Introduction

The problem of deriving work from a irreversible thermodynamic system and the inverse problem of maintaining its irreversible state by consuming energy are central in thermodynamics. For systems that are not in equilibrium with respect to temperature, the first (direct) of the above problems is solved using heat engines and the second one (inverse) is solved using heat pumps. For systems that are not in equilibrium with respect to composition, the second problem is solved using separation systems and the first one is solved using diffusion engines. As a rule, separation systems and diffusion engines are based on membranes.

There is a lot of studies of membrane separation systems and diffusion engines in the literature [5,7]. In the present paper, these systems will be considered using the theory of finite-time thermodynamics. The finite-time thermodynamics, which evolved in the past years, studies the limiting performance of irreversible thermodynamic systems when the duration of the processes is finite and the average rate of the streams is specified [14, 17]. For example, some problems for heat engines, such as maximizing the power at given heat transfer coefficients and maximizing the efficiency at given power for different conditions of contact between the working body and surroundings, are already solved. In this case, the irreversible processes of the interaction of subsystems each of which is in internal equilibrium are considered.

For systems that are not uniform in concentration, it is most important to study the limiting performance of separation systems. In this case, however, the inverse problem of studying the performance of diffusion engines is of definite interest as well. The simplest variant of this problem was first formulated by Rozonoer [17]. The review of the literature shows that this problem was discussed rather superficially.

In the present paper, we will study the limiting performance of membrane systems in the separation processes with fixed rates, focusing on the following problems:

  1. Minimizing the amount of energy necessary for the separation of a feed mixture with a given composition into separation products with given compositions at a given average production rate.

  2. Maximizing the power and efficiency of diffusion engines.

The solution of these problems depends strongly on whether the feed mixture used by the engine is gaseous or liquid because this determines the form of the chemical potentials of components and, hence, the driving forces of the process. For near-ideal gas mixtures, the chemical potential of component I of the mixture takes the form [15]:

μi(T,Pi)=μ0(T)+RTlnPi,i=1,2,...,

where Piis the partial pressure of component I and μ0(T)is the chemical potential of the pure component. Assuming that the ratio of the partial pressure to the total pressure is equal to xi,

Pi=Pxi=PNiN,i=1,2,...,

we can rewrite the expression for the chemical potential in the form:

μi(T,P,xi)=μ1(T,P)+RTlnxi,E118

where μ1(T,P)=μ0(T)+RTlnP.

Although the chemical potential for liquids has the same form as Eq. (118), the form of the function μ1(T,P)is different. This is caused by the fact that the chemical potential μi(T,P,xi)represents the molar Gibbs energy of component I and its derivative with respect to pressure is equal to the molar volume of this component νi[15]. In contrast to gases, the molar volume of liquids is virtually independent of pressure and weakly dependent on temperature. As

μiP=μ1P=νi,

we obtain

μi(T,P,xi)=μ1(T)+RTlnxi.E119

It is assumed that the processes are isothermal and the temperatures of all subsystems are equal to T. The problems listed above will be considered for gaseous mixtures and then for liquid solutions.

4.2. Limiting performance of diffusion systems for gaseous mixtures

4.2.1. Maximum work in a membrane process

Consider a system consisting of a thermodynamic reservoir, the intensive variables of which are fixed and are independent of mass transfer fluxes, and a working body, the intensive variables of which can be varied with time by one or another way. The system can consume external energy or generate work. In the first case, the work will be negative; in the second, positive.

The reservoir and the working body interact through a membrane that is permeable only to one (active) component of the mixture. The mass transfer rate g depends on the chemical potentials of the active component in the reservoir μ0and in the working body μ(t). When these chemical potentials are equal to each other, the flux is equal to zero. In the particular case under consideration,

g(μ0,μ)=α(μ0μ),E120

where αis the mass transfer coefficient. The working-body temperature T is maintained constant and equal to the reservoir temperature.

When the process duration τand the total amount of the component G0transferred from the reservoir to the working body and in the reverse direction are fixed in the process characterized by a finite mass transfer coefficient, the chemical potentials μ0and μ(t)should differ from each other at every moment of time and the mass transfer process should be irreversible. For definiteness, we assume that μ0>μ(0)and that the component is transferred from the reservoir to the working body.

The variation of the system entropy will be caused by the decrease in the reservoir entropy, the increase in the entropy of the working body, and the production of entropy due to the irreversible mass transfer σ. For a given initial state of the system (that is, the compositions of mixtures at the initial moment of time, the total amount of the substance in the working body) and a given constant value of the quantity

G0=0τg[μ0,μ(t)]dtE121

the variation of the entropies of the reservoir and working body with time τare completely determined and the minimal increase in the system entropy corresponds to the minimum of the entropy production:

σ¯=1T0τ[μ0μ(t)]g[μ0,μ(t)]dtmin.E122

In this case, the function μ(t)should be chosen.

Let us find the quantitative relationship between the work A, which can be extracted (consumed) in this process, and the value of σ¯. For simplicity, we assume that the mixture in the reservoir and working body consists of two components (a more general case can similarly be considered by introducing an equivalent component). If the concentrations of the active component in the reservoir and working body are x0and x(t), the concentrations of the second component will be equal to 1x0and 1x(t), respectively. The variation of the substance amount G and the concentration x(t)of the active component in the working body are determined by the differential equations:

ddt(Gx)=dGdt=g(μ0,μ),G(0)fix,x(0)fix.E123

As the amount of the second component is maintained constant, we obtain

G(0)[1x(0)]=[G(0)+G0][1x(τ)].E124

It follows from (123) and (124) that

dxdt=1G(0)[1x(0)](1x)2g(μ0,μ),x(0)fix.E125

The equations for the material, energy, and entropy balances around the system take the form:

G0x0=G(τ)x(τ)G(0)x(0),E126
G0h0[G(τ)h(τ)G(0)h(0)]=A,E127
G0s0=[G(τ)s(τ)G(0)s(0)]+σ¯=0,E128

where h0and h, s0and s are the molar enthalpies and entropies of the mixture in the working body and reservoir, respectively. They are related by the equation [15]:

s=1T(hi=12μixi),E129
s0=1T(h0i=12μi0xi0).E130

The pressure in the working body can vary with time, provided that P(0)=P(τ). For the chemical potentials defined by Eq. (118), the equation of entropy balance (128) in view of (127), (129), and (130) can be rewritten as

A/T=σ¯+R{G0[x0lnx0+(1x0)ln(1x0)]+G(0)[x(0)lnx(0)+(1x(0))ln(1x(0))]G(τ)[x(τ)lnx(τ)+(1x(τ))ln(1x(τ))]}.E131

The second term in the right-hand side of this equality can be calculated using G0, x0, G(0), and x(0). The latter ones are related through (124) and (126) to the values of G(τ)and x(τ). Let us denote the second term as B[G0,x0,G(0),x(0)]. It can be either positive or negative. It follows from equality (131) that

Amax=T(Bσ¯min).E132

The maximum of the produced (minimum of the spent) work corresponds to the minimum of entropy production in the mass transfer process.

The problem of finding the minimum of σ¯when constraint (121) is valid (or the equivalent problem for the maximum of G0at a given constant value of σ¯) is an averaged nonlinear programming problem [22]. Unlike the problem for the constrained maximum of a function, its optimal solution can vary with time. This solution is a piecewise constant function that can take not more than two values. We will not calculate these values and the fraction of the whole process time during which μ*(t)takes each of these values because in the most common case, where the Lagrangian function for the unaveraged problem

L=g[μ0,μ(t)][μ0μ(t)λ]

is convex with respect to μ(second derivative of L with respect to μis positive), the solution to the formulated problem is constant. Consequently, the constancy condition depends on the validity of the inequality:

2gμ2(μ0μ(t)λ)2gμ0.E133

The multiplier λ, which is equal to the derivative of the minimum value of σ¯with respect to G0, should be positive due to the physical nature of the problem. The second derivative of L with respect to μfor the mass transfer rate in the form of (120) is equal to 2λαand is known to be positive. In all cases where inequality (133) holds, the optimal value of the chemical potential of the active component for the working body is constant and determined by the equation:

g(μ0,μ*)=G0τ.E134

Consequently, the chemical potential of the active component of the working body for any rate satisfying (133) should be controlled so that the mass transfer rate should be constant.

Figure 11.

Schematic diagram of a diffusion engine with a constant contact between the working body and sources.

The law of variation of the control variable, such as the working-body pressure, corresponding to this solution will not be constant in time because the mixture composition is varied during the process according to Eq. (125), in which the flux is determined by Eq. (134).

For mass transfer law (120), the minimal entropy produced is σ¯min=G02/ατ. It follows from equality (132) that positive work can be extracted from the system under study only when τ>τmin=G02/αB. It is easy to see that the process duration τ*, for which the average extraction rate of work A*(τ)/τis maximal, is twice larger than τmin.

In the case where the system contains a source of a finite capacity at constant temperature and pressure instead of the reservoir (source of an infinite capacity), the fraction of the active component varies according to an equation similar to (125). As a result, the chemical potential μ0is changed. However, here also, the minimum of the entropy production for mass transfer law (120) corresponds to such variation of μ(t)that the mass transfer rate is maintained constant.

Instead of the calendar time, the problem can be studied using the time of contact, when the working body moves and its parameters at every point of the loop remain constant. This can be used to determine the optimal laws of pressure variation for the zones of contact between the working body and source.

4.2.2. Diffusion-mechanical cycle for maximum power

Let us consider the direct cycle of work extraction in a system consisting of a working body and two reservoirs with different chemical potentials. In the first reservoir, the chemical potential of the key element is equal to μ+; in the second,μ; for definiteness, μ+>μ(Fig. 11). The process is cyclic: the increase in entropy, internal energy, and mass of the key component of the working body around the cycle is equal to zero. The temperatures are the same for all subsystems.

Alternating contact with reservoirs. Consider the case where the working body alternately contacts the first and second reservoirs and its parameters are cyclically varied with time. Let τstand for the cycle duration and μ0(t)stand for the source chemical potential, which can take the values of μ+and μ. The formulation of the problem dealing with the production of maximum work A in a given time τtakes the form:

A=0τμg(μ0,μ)dtmaxμ0,μE135

with the constraints placed on the increment in the amount of the working-body:

ΔG=0τg(μ0,μ)dt=0.E136

Figure 12.

Dependence of the maximum of the Lagrangian function with respect to μ onλ.

To calculate the basic values of μand μ0in the problem given by (135) and (136), we can write the Lagrangian function and find its maximum with respect to μ0and μand its minimum with respect to λ:

L={g(μ0,μ)(μλ)}maxμ0,μmin.λ

The number of basic values of μ0is equal to two: one of them corresponds to μ0=μ+and the other to μ0=μ. For the Lagrangian function L that is strictly convex with respect to μ, the basic values of μsatisfy the conditions:

Lμ=gμ(μλ)+g(μ0,μ)=0

or

g(μ0,μ)(μλ)=g(μ0,μ)μ.

The roots for this equation for μ0=μ+and μ0=μwill be denoted by μ1and μ2, respectively. As L is maximal at the basic points, we can write

L(μ+,μ1,λ)=L(μ,μ1,λ),E137

which determines the value of λ.

Let us specify the obtained relations for

g(μ0,μ)=α(μ0)(μ0μ).

It follows from (137) that

μ1=μ++λ2,μ2=μ+λ2.E138

Substituting μ1and μ2into the function L for each basic value gives its dependence on λ:

L+=L(μ+,μ1)=α+4(μ+λ)2,L=L(μ,μ2)=α4(μλ)2.

The maximum of L with respect to μ0and μreaches its minimal value with respect to λ(Fig. 12) when

L+(λ)=L(λ)λ*=α+μ++αμα++α.E139

The fractions of time τof contact with reservoirs are determined by Eq. (136) and can be written as

γ+=αα+αα++α+α,γ=α+ααα++α+α.

The maximal work in time τtakes the form:

A*(τ)=τ[γ+μ1α+(μ+μ1)+γμ2α(μ2μ)],

where μ1and μ2can be determined from (138) after the value of λfrom (139) is substituted into this expression. The maximal power is equal to

A*(τ)τ=[γ+μ1α+(μ+μ1)+γμ2α(μ2μ)].

Constant contact with sources. In heat engines, there can be either alternate or constant contact between the working body and sources. In the latter case, the parameters of the working body are distributed and the process in it can be regarded close to reversible if the distribution of the parameters is caused by the conductive flux. Likewise, a constant contact with sources is possible in systems that are not homogeneous in concentration, such as separation systems and diffusion engines.

In this case, the maximal power takes the form of a nonlinear programming problem:

p=[g1(μ+,μ1)μ1g2(μ2,μ)μ2]maxμ1,μ2

with the constraint

g1(μ+,μ1)g2(μ2,μ)=0.E140

The optimality constraint for this problem leads to the relation:

μ1μ2=g2(μ2,μ)g2/μ2g1(μ+,μ1)g1/μ1,E141

which together with equality (140) determines the desired variables.

Let g1and g2are proportional to the difference between the chemical potentials:

g1=α1(μ+μ1),g2=α2(μ2μ).

Equality (141) can be written in the form:

μ1μ2=μ+μ2.E142

The constraint g1=g2results in

α1μ1+α2μ2=α1μ++α2μ.E143

The solution to Eqs. (142) and (143) can be written as

μ2*=12(α1+α2)[μ+α1+μ(α1+2α2)],μ1*=12(α1+α2)[μ+(α2+2α1)+μα2].

The value of maximal power corresponding to this choice is

pmax=α¯4(μ+μ)2,

where the equivalent mass transfer coefficient is defined as

α¯=α1α2α1+α2.

4.3. Limiting performance of diffusion systems for liquid mixtures

The result obtained above for the membrane systems consisting of a working body and a source of finite or infinite capacity using gaseous mixtures can be translated in the same form to liquid solutions with allowance for the different form of the chemical potential. Diffusion engines are most often designed for the treatment of saline water. Let us consider two flow-sheets of liquid diffusion engines.

4.3.1. Diffusion engine with a constant contact between the working body and the sources

Let the system consist of two liquids with the same temperature separated by a semipermeable membrane. One of the liquids is a pure solvent and the other is a solution in which some substance of concentration C is dissolved. The membrane is permeable only to the solvent. The equilibrium in the system is reached as soon as the chemical potentials calculated by formula (119) become equal to each other:

ν0P0νrPr=RTlnxr.

Let the difference of pressure across the membrane be denoted as π. Also, we will keep in mind that the molar volumes ν0and νrfor low concentrations are equal to each other. The mole fraction of the dissolved component will be denoted as x1. If its value is low, then lnxr=ln(1x1)x1. In this case,

π=RTx1ν0=RTC.E144

Equation (144) is called the Van’t Hoff equation for osmotic pressure.

Consider the system shown in Fig. 13. The chamber to the left of the membrane contains a pure solvent at an environmental pressure equal to P0. The chamber of volume V to the right of the membrane contains a continuously replenished solution in which the concentration of the dissolved component is C. The pressure in the right chamber is P2and the solution is assumed to be ideal. When an equilibrium is reached in the right chamber (that is, the flux through it is equal to zero), the pressure established in it will exceed P0by the value of osmotic pressure π. The osmotic pressure value is related to the concentration and temperature in the chamber by the Van’t Hoff equation. When the solution in the chamber is replenished, the pressure P2<P0+π, giving rise to a solvent flux g across the semipermeable membrane. Conventionally, the diffusion flux is taken to be equal to the difference between the actual and equilibrium pressures:

g=α(P0+πP2)=α(πΔP),E145

where ΔP=P2P0

Figure 13.

Schematic diagram of a diffusion engine with a constant contact between the working body and sources.

Let p1stand for the power of the pump supplying the concentrated solution, g1stand for the flow rate of this solution, and C1stand for the solution concentration. Assuming that the pump efficiency is 100%, we obtain

p1=ΔPg1.

The additional flux across the membrane increases the volume of the solution, which drives a turbine and generates power p2:

p2=(g1+g)ΔP.

Consequently, the power r and efficiency ηof the saline diffusion engine can be written as

p=p2p1=gΔP=α(πΔP)ΔP,η=Pg1=α(πΔP)ΔPg1.

where the diffusion engine efficiency is the work extracted from 1 m3 of the concentrated solution. From here on, according to the accepted system of units, the units of power and efficiency referred to a unit membrane surface area are J/(m2 s) and J/m3, respectively. If the relationship between πand ΔPis ignored, the power reaches a maximum when ΔP=π/2and its upper limit is written as

p¯=απ2/4=α/4(CRT)2.

AsC<C1, the value of the power is always less than

p¯*=α/4(C1RT)2.E146

which is the upper bound for the maximal power.

The estimate produced by Eq. (146) can be refined if we take into consideration that g, ΔP, and C are related to each other by Eq. (145) and the equation of material balance on the dissolved component

(g1+g)C=g1C1.E147

Expressing C and ΔPin terms of g from Eqs. (145) and (147) and substituting them into p and η, we obtain

C=g1C1g1+g,ΔP=CRTgα,E148
p=gΔP=RTC1g1gg1+gg2α,E149
η=α(πΔP)ΔPg1=RTC1gg1+gg2αg1.E150

The points of maximum with respect to g for two concave functions (149) and (150) coincide. Consequently, to find the optimal value of g*, we will use one of the functions, specifically the expression for p. The condition for the maximum with respect to g leads to the inequality:

g(g1+g)2=αRTg12C12.E151

Equation (151) can be rewritten as

g3g12+2g2g1+g=αRTC12E152

and its right-hand side can be denoted for brevity as M. Its solution will be denoted as g*. It is obvious that it satisfies the inequality:

0<g<M.

Numerical solution of Eq. (152) makes it possible to refine the value of the limiting power of the diffusion engine and find the corresponding operating conditions. Equation (151) determines g*for the chosen values of g1and C1; Eq. (148), for C*and ΔP*.

It should be noted that the ideal solution bounds the value of the concentration of the working solution:

C=C1g1g1+g.

The concentration should not be very high: otherwise, the molecules of the dissolved component will interact with each other and relation (144) is upset.

Diffusion Engine with an Alternate Contact between the Working Body and Sources. In the schematic diagram of the diffusion engine discussed in the preceding section, the working body was an open system working in constant contact with two sources under steady-state conditions. One of them supplied a concentrated solution and the other supplied a solvent.

Figure 14 shows the schematic diagram for a diffusion engine in which the working body alternately contacts each of the sources, receiving a solvent through one membrane and giving it up to a concentrated solution through another membrane. In this case, the pressure and flow rate of the working body are periodically varied: pressure increases for a lower flow rate (power p1is consumed) and decreases for a higher flow rate (power p2is generated).

We will write the balance equations for this diagram and study its limiting performance, ignoring the energy losses for driving the flow of the concentrated solution through the bottom chamber and assuming that the concentration of the dissolved component in the g2flow is equal to unity and that the pressure of the surrounding medium is equal to P0. For simplicity, flow rates will be used instead of mole fluxes

Figure 14.

Schematic diagram of a diffusion engine with an alternate contact between the working body and sources.

Figure 15.

Variation cycle for the parameters of the working body in a diffusion engine.

The engine power is

p=p2p1=(g1+g)ΔP21g1ΔP21=gΔP21,

where ΔP21=P2+P1

The efficiency will be defined as the ratio of power p to the flow rate g2of the dissolved component:

h=pg2=gg2ΔP21.

The rate of mass transfer is determined by the relations:

g=α1(P0+πP2)=α1(πΔP20)=α2[(P1+π2)(P0+π1)]=α2(Δπ21+ΔP10),E153

where ΔP20=P2P0, Δπ21=π2π1, ΔP10=P1P0. Equation (153) corresponds to the condition that the mass of the working body averaged over the cycle is constant.

Figure 15 demonstrates the cycle of the working body of this diffusion engine. The power p1is equal to the area of the rectangular P2dcP1, and the power p2to the area of P2abP1. The engine power p is equal to the area of the hatched rectangular abcd.

The power of the diffusion engine will be determined when the relationship between the osmotic pressures in the chambers and the flow rates is ignored. To do it, we will solve the problem of constrained optimization:

p=(P2P1)gmaxP1,P2

with the constraints:

α1(P0+πP2)=α2(P1P0+π2π1)=g.E154

It follows from Eq. (154) that

P1=gα2+P0+π1π2,P2=P0+πgα1.

Let us introduce the equivalent permeability:

α¯=α1α2α1+α2

and write the equation:

P2P1=ππ1+π2gα.

Then

p=g(ππ1+π2gα)=g(π+Δπ21gα)maxg.E155

The maximum of this expression, which is equal to

p*=α¯(ππ1+π2)24=α(π+Δπ21)24,

is reached at

g*=α¯(ππ1+π2)2=α(π+Δπ21)2.

Keeping in mind that the osmotic pressures in the chambers are related to the concentrations by Van’t Hoff equation (144) and the concentrations are related to the flow rates g1, g2, and g, we obtain

π=CRT=C1g1RTg1+g,Δπ21=(C2C1)RT=(g2C20+gC1g2+gC1)RT.

In view of these relations, expression (155) for the engine power takes the form:

p=g[RT(C1g1g1+g+g2C20+gC1g2+gC1)gα]=g[RT(g2C20+gC1g2+gC1gg1+g)gα]maxg.E156

The expression for the efficiency is written as

η=gg2[RT(g2C20+gC1g2+gC1gg1+g)gα]maxg.E157

The points of maximum with respect to g for the criteria (156) and (157) coincide. Therefore, we can use either of them in the conditions of optimality to find g*. The stationarity condition of p with respect to g leads to an equation for the optimal flow rate:

g=α¯RT2[(g22C20+2gg2C1+g2C1(g2+g)2)C1g(g+2g1)(g1+g)2].E158

The solution to Eq. (158) will be g*: it is the optimal value of flow rate g at which the efficiency ηand power p take their maximal values. The values of flow rate g*depends on the values of g1, g2, and C1. Its substitution into the equations for p and ηdetermines the maximal power p*(g1,g2,C1)and efficiency η*(g1,g2,C1). The nonnegative nature of p*and η*imposes constraints on the possible values of g1, g2, and C1. For example, increasing g1and g2or decreasing C1causes an increase in p*.

4.4. Conclusion

The estimates obtained in the present paper for the limiting performance of diffusion engines can be used to make their reversible-thermodynamics analysis more accurate and consider the influence of the kinetic factors (mass transfer relations, membrane permeabilities) and production flow rate. These estimates can also be used for the optimization of more complex membrane systems. The capacity of membrane systems increases in proportion to the membrane permeability. In this case, the performance of membranes is decreased by the nonuniformity of concentrations in the solution, polarization phenomena, and the other factors ignored in obtaining the above estimates.

© 2013 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution 3.0 License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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A.M. Tsirlin and I.N. Grigorevsky (July 24th 2013). Minimum Dissipation Conditions of the Mass Transfer and Optimal Separation Sequence Selection for Multicomponent Mixtures I, Mass Transfer - Advances in Sustainable Energy and Environment Oriented Numerical Modeling, Hironori Nakajima, IntechOpen, DOI: 10.5772/54546. Available from:

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