Ind. Eng. Chem. Res. 2007, 46, 591-599
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High-Purity Oxygen Production by Pressure Swing Adsorption J. C. Santos, P. Cruz, T. Regala, F. D. Magalha˜ es, and A. Mendes* LEPAEsDepartamento de Engenharia Quı´mica, Faculdade de Engenharia, UniVersidade do Porto, Rua Dr. Roberto Frias, 4200-465 Porto, Portugal
The oxygen purity produced by pressure swing adsorption (PSA) processes is limited to 95%, with the rest being essentially argon. This oxygen grade is suitable for many industrial applications. However, medical applications, cylinder filling, oxyfuel cutting in metal fabrications, and fuel cells technology with recirculation loop, among others, require oxygen with a higher purity (99% or above). In this paper, a study of high-purity oxygen production by a PSA unit using a silver exchanged zeolite from Air Products and Chemicals, Inc., with oxygen/argon adsorption selectivity is presented. This study comprehends the determination of adsorption equilibrium isotherms of oxygen, nitrogen, and argon as well as the simulation and optimization of a PSA experimental unit and the corresponding experimental validation. Introduction Oxygen is a gas with many applications. It may be used in chemical processing, fishing farms, medical applications, combustion enhancement, oxyfuel cutting operations in metal fabrication, bleaching in the paper industry, wastewater treatment, fuel cells, etc. Oxygen production (purity below 95%) from air, using nitrogen selective zeolites of type A (5A) or X (13X-NaX, LiX, or LiLSX), by means of pressure swing adsorption (PSA) processes has noticeably increased in the past decade. However, the concentration of the product is limited to 95% oxygen, because of the presence of argon in air, since these adsorbents present similar adsorption capacities for oxygen and argon. There are some applications that require higher oxygen purities. In the case of fuel cell technology, when 95% oxygen is used, the average concentration of oxygen in the cathode recirculation loop is ∼55%.1 To minimize the amount of oxygen that is lost in the purge, which can be seen as an energy loss, a higher concentration of oxygen may be used, which will result in a reduction of the argon buildup and a reduction of the purge flow rate and, thus, in a smaller requirement of this gas.2 Medical applications, such as surgeries, in the United States require oxygen with 99% purity and in Europe require oxygen with 99.5% purity. Thus, it is important to develop a simple and portable technology for producing high-purity oxygen (for example, for use in campaign hospitals). Some research has been carried out in the past years for the development of new adsorbents with selectivity for argon relative to oxygen, allowing the production of oxygen with a purity above 95% by PSA.2 In 1999, Hutson et al. suggested the addition of silver to LiX zeolites for improving the air separation performance, by increasing the nitrogen adsorption capacity, and presented simulation results of a PSA unit producing oxygen with a purity of 96.42% and a recovery of 62.74% from a feed containing 22% of oxygen and 78% of nitrogen (no argon was used).3 Although their adsorbent does not actually present selectivity of argon relative to oxygen, the argon adsorption capacity was slightly increased. In 2002, Air Products and Chemicals, Inc., patented an argon/oxygen selective X-zeolite (U.S. 6432170 B1) designated as AgLiLSX (low silica X).4 In 2003, a vacuum and pressure swing adsorption * To whom correspondence should be addressed. E-mail: mendes@ fe.up.pt. Phone: +351 22 5081695. Fax: +351 22 5081449.
(VPSA) unit for the production of high-purity oxygen from air, using AgLiLSX, was described in a patent by the same company.5 The unit described may use one layer of AgLiLSX or two layers of adsorbents: LiX and AgLiLSX. According to the simulation results presented in this patent, changing the percentage of AgLiLSX in the bed from 50 to 100% allows the production of 99% oxygen with a recovery between 6 and 15%, respectively. The unit is operated between 0.34 and 1.4 bar at 38 °C. According to the Air Products and Chemicals’ patent, the recovery of a VPSA unit with unspecified volume and flow rates and producing 99% of oxygen is ∼4% when fed with a current of 95% oxygen and 5% argon (such as the product of a PSA unit using a traditional adsorbent). Two other technologies may be used for producing highpurity oxygen from air but are more complex and make use of two steps. Both units have a pressure swing adsorption (PSA) unit packed with a zeolite for producing 95% of oxygen and 5% of argon. The product of this unit is fed to either a carbon molecular sieve (CMS) membrane module or another PSA unit packed with CMS adsorbent. In both systems, the product is obtained at low pressure and a compressor must be used to bring the product pressure to the operating requirements.2 The low separation factor of polymeric membranes, which hinders the production of high-purity oxygen, has led to the development of different configurations such as the continuous membrane column (CMC)6-8 and the CMC in two-strippersin-series (TSS) mode6,8 or membrane cascades.9-11 These multistage configurations make use of several membrane modules, and the streams are recycled between the modules. Many arrangements of the modules are already patented.12-21 The combination of polymeric membrane modules with PSA units for high-purity oxygen production can also be found in the literature.8,22 The ability of pore tailoring of the carbon molecular sieve membranes together with its high permeability and selectivity makes the use of this material very appealing since it allows the production of high-purity oxygen with only one module.2 According to simulation results, a CMS membrane module with a selectivity of oxygen relative to argon of 14, using a feed pressure of 6 bar and a permeate pressure of 0.1 bar, is able to produce 99.5% of oxygen with a recovery of 46%.2 This technology associated with a PSA unit packed with Oxysiv MDX is expected to yield a global recovery ∼36%.2
10.1021/ie060400g CCC: $37.00 © 2007 American Chemical Society Published on Web 12/20/2006
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Ind. Eng. Chem. Res., Vol. 46, No. 2, 2007 Table 2. Position of the Valves during the Cycle valve
Figure 1. Schematic representation of the experimental PSA apparatus: FM ) flow meter; FC ) flow controller; PR ) pressure reducer; and ST ) storage tank.
Table 1. Characteristics of the Adsorbent, Adsorption Bed, and Storage Tank bed length bed internal diameter column thickness bed porosity storage tank length storage tank diameter adsorbent porosity adsorbent diameter Peclet number for mass transfer
L d b LST dST p dp Pe
27 cm 2 cm 0.175 cm 0.36 27 cm 2 cm 0.471 1.01 mm 600
The third technology referred to before for producing highpurity oxygen, i.e., a PSA unit packed with zeolite followed by another PSA unit packed with carbon molecular sieve adsorbent, can also be readily found in the literature.23-34 In 2005, Jee et al.29 presented a PSA unit packed with a CMS from Takeda, producing 99.5% of oxygen from a feed mixture of 95% of oxygen and 5% of argon. This unit had a recovery ∼57% if the product was obtained at a pressure close to the atmospheric pressure or ∼80% if the product was obtained at low pressure.29 However, from the three technologies proposed before, this one should result in the largest unit. The choice of the most suitable technology depends on the application characteristics. For small systems, the PSA unit with AgLiLSX should be the most adequate. For systems where size is not so important, an economical balance should be made to find out which of these technologies is the best solution. In this paper, high-purity oxygen production using a zeolite patented by Air Products and Chemicals, Inc., AgLiLSX, is presented. This adsorbent allows the production of high-purity oxygen (above 95%) using only a single separation step, directly from air, which makes this technology very appealing for many applications.
B
B
T
step
VC1
VC2
VC1
VTC2
VF
VProd
1 2 3 4 5 6
2f1 2f1 1f2 1f3 1f3 1f2
1f3 1f3 1f2 2f1 2f1 1f2
1f3 1f2 1f3 1f2 1f2 1f3
1f2 1f2 1f3 1f3 1f2 1f3
on on off on on off
off on off off on off
the bed end were used to measure the pressure variations in the bed. The feed flow rate was measured by a mass flow meter (Bronkhorst High-Tech, F-112AC, 0-20 dm3STP/min). The purge and equalization flow rates were measured using two Bronkhorst High-Tech, F-112C, 0-10 dm3STP/min mass flow meters. The production flow was controlled using a Bronkhorst High-Tech, F-201C, 0-0.1 dm3STP/min mass flow controller. In order to keep the pressure in the adsorption bed constant during the production step, a pressure regulator (Joucomatic) was installed between the feed tank and the adsorption columns. The interchange between adsorption columns is made using two three-way electrovalves (ASCO: EV1, EV2) installed at the columns bottom. At the columns top, there are also two threeway electrovalves (ASCO: EV1, EV2) in order to change between equalization and production steps. Two oxygen analyzers were used: one from M&C, model PMA22, from 0 to 100% with an accuracy of 0.1% FS, and another from Sable Systems, model PA-1B, also from 0 to 100% with an accuracy of 0.01% FS. The analyzers were calibrated each run using pure oxygen as span and a calibrated mixture of 90% oxygen, 5% nitrogen, and 5% argon as zero. Table 2 presents the position of the valves during the cycle. The left column was packed with 101.86 g of AgLiLSX from Air Products and Chemicals and the right column was packed with 101.37 g of AgLiLSX. Mathematical Model The mathematical model used for describing the operation of the pressure swing adsorption unit was based on the following assumptions: perfect gas behavior, axially dispersed plug flow, uniform bed properties along the axial coordinate, negligible pressure drop, negligible radial gradients, interparticle mass transport described by the linear driving force (LDF) approximation,35 and adsorption equilibrium between gaseous and adsorbed phases described using the Langmuir-Freundlich equation. Cruz et al.36 showed that isothermal operation is a reasonable hypothesis for oxygen separation from air, so this was also considered here. According to these assumptions, the model equations can be written as follows,
Total mass balance ∂cT
)-
∂(ucT)
∂t
∂z
nc
-
Ni ∑ i)1
(1)
PSA Experimental Apparatus
Partial mass balance A schematic representation of the experimental apparatus is presented in Figure 1. The adsorption beds are made of stainless steel and are packed with AgLiLSX supplied by Air Products and Chemicals, Inc. The characteristics of the adsorbent and sorption bed are listed in Table 1. The feed tank has an internal volume of ∼60 dm3. A pressure transducer located in the feed stream and two pressure transducers (Druck, PMP 4010, 0-7 bar) located at
(
)
∂ci ∂ ∂(uci) ∂(ci/cT) ) DaxcT - Ni, i ) 1, nc (2) ∂t ∂z ∂z ∂z where cT is the total molar concentration, u is the average (interstitial) molar velocity, z is the spatial coordinate, Dax is the effective axial dispersion coefficient, Ni is the ith component molar flow rate, ci is the ith component molar concentration in
Ind. Eng. Chem. Res., Vol. 46, No. 2, 2007 593
Figure 2. Schematic form of finite volume method discretization.
parabolic profile inside the particle.41 This equation is a firstorder delay in mass transfer from the bulk gas phase to the adsorbed phase and is easily coupled to the conservation equation in the bulk gas phase. The molar velocities across the valve orifices are described by42,43
Figure 3. Definition of local variables.
Table 3. Parameters of the Monocomponent Sips Equation for Oxysiv 5 and AgLiLSX at the Reference Temperature T0 ) 20 °C Oxysiv 5 qmax,ref (mol kg-1) bref (bar-1) Q/R (K) 1/nref R β
AgLiLSX
N2
O2
Ar
N2
O2
Ar
3.091
3.091
3.091
2.636
5.481
7.270
0.1006 0.03670 0.03365 0.2581 0.0301 0.0230 2501.07 1767.86 1791.15 2740.70 1873.82 1771.33 1 1 1 0.4856 0.9484 0.9216 0 0.3007 0 0.0112
the fluid phase, t is the time variable, and nc is the number of components in the mixture. The mass exchange rate between the particle and its surroundings (ith component molar flow rate), assuming the accumulation in the intraparticle (nonadsorbed) gas-phase is negligible, is given by the following equation,
Ni ) Fs
p(1 - b) ∂qi b ∂t
(3)
where p is the particle porosity, qji is the average molar concentration in the adsorbed phase, and cji is the molar average concentration in the fluid phase. The intraparticle mass transfer is described using the linear driving force (LDF) approximation,35
Linear driving force model ∂qi ) ki(qi,s - qi) ∂t
up ) Kvf(pu, pd, T, M)
(5)
where Kv is proportional to the valve parameter, Cv, and is given by
1 pSTP Kv ) (2.035 × 10-2) ‚ STPCv bA T and
f(pu, pd, T, M) )
{
1.179
x
pu
(6)
x
pu2 - pd2 p > 0.53p u T d p dM
1 T pdM
pd e 0.53pu (7)
where pu and pd are the upstream and downstream pressures, respectively, A is the area, T is the temperature, and M is the molecular weight of the gas passing through the orifice. The superscript “STP” stands for standard temperature and pressure conditions. The molar velocity, u, across the orifice in the feed, in the vent, before the storage tank, and across the purge and equalization orifices was considered to be given by
Feed: uinp ) KvFf(pH, p, T, M) ST ST ) KST Storage tank: uST in p v f(p, p , T, M)
(4)
e where ki ) 15DM,i /rp2 is the ith component LDF coefficient,37 which is directly proportional to the effective macropore e is the effective homogeneous diffusivity coefficient, DM,i diffusion coefficient, rp is the particle radius, and qi,s is the molar concentration in the particle surface (adsorbed phase), which is related to the molar concentration in the interparticle gas phase through the adsorption equilibrium isotherm, qi,s ) f(pT, ci). The value of DM,i was estimated using the following equation:38 DM,i ) pDep,i/(p + Fs dq/dc), where Dep,i is the effective diffusivity in the pores. The effective diffusivity in the pores was calculated e e using the Bosanquet equation:38 1/Dep,i ) 1/DK,i + 1/Di,j , where e e DK,i is the effective Knudsen diffusivity and Di,j is the effective molecular diffusivity. The effective Knudsen diffusivity is given e by39 DK,i ) 3.068rpore/τ xT/Mi and the effective molecular diffusivity was obtained using the Chapman-Enskog equation.40 Equation 4 can be deduced from the intraparticle mass balance equation, which is a partial differential equation, considering
Purge: upurgep ) upurge,other columnpother column ) KPv f (p, pother column, T, M) Vent: uout,ventp ) CVv f(p, pL, T, M) Equalization: uequalizationp ) KEv f(p, pother column, T, M) The mass balance equation in each storage tank is given by
V
ST
∂cST T ST ST ST ) -(uST outcT - uin cT ) ∂t
(8)
where VST is the volume of the storage tank (ST). The boundary conditions of eqs 1 and 2 change with the process steps. The unit operates with a standard cycle with six steps and with a top-to-top equalization (TE). In step 1, column 1 is being pressurized while column 2 is depressurizing. In step 2, column 1 produces and column 2 is being purged. In step 3, column 1 provides equalization to column 2. In step 4, column
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1 depressurizes and column 2 pressurizes. In step 5, column 1 is being purged while column 1 produces. Finally, in step 6, column 1 receives equalization from column 2. When column 1 is pressurizing and column 2 is depressurizing (step 1), the boundary conditions are as follows:
Column 1 F
Kv 1 ∂ci ) u(ci - ci,in); uin ) f(pH, p1, T, M) Dax ∂z p1
z ) 0:
z ) L:
∂ci ) 0; u ) 0 ∂z
Column 2 ∂ci KVv z ) 0: ) 0; u ) - f(p2, pL, T, M) ∂z p2 z ) L:
Column 1 ST uST KPv in p f(p1, p2, T, M) + p1 bp1
Column 2 z ) L:
∂k ckT
)-
k k-1 (ukFcT,F - uk-1 F cT,F ) k
∂t
∆x
KPv u ) - f(p1, p2, T, M) p2
n
c ) t T
∑ k)1
( ) ∂c k kT ∂t
n
∆x ) k
Column 1 dci ) 0; u ) 0 dz
Column 2 dci ) 0; u ) 0 dz
ukF ) uk-1 F
∆xk cT
nc
(10)
nc
(cTt +
Nki ), ∑ i)1
k ) 1, n - 1
(11)
k k-1 k-1 k-1 ′ cT,F (cki /ckT)′F - cT,F (ci /cT )F ∂c kki ) Dax k ∂t ∆x k k-1 - uk-1 (ukFci,F F ci,F )
∆x
k cki ) cki +
k
-N k ki (12)
(∆xk)2 k (∆xk)4 k (iv) (ci )′′ + (c ) + ... 24 1920 i
KEv f(p , p , T, M) p2 1 2
Numerical Methods The discretization of eqs 1 and 2 was performed in two stages. First, the space derivatives appearing in the right-hand side were computed using the finite volume method. Then, the resulting initial value problem was integrated explicitly in order to obtain the grid point values at the next time step using the package
(13)
(cki /ckT)′F is the derivative of (cki /ckT) in the face k and is a function of the neighboring cells. Using a second-order approximation (central difference schemesCDS2), we obtain
(cki /ckT)′F )
1 ∂ci ) u(ci - ci|z ) L,column1); Dax ∂z u)
(9)
k c ki is the cell average concentration that is a function of cki as follows:45
KEv dci ) 0; u ) f(p1, p2, T, M) z ) L: dz p1
z ) L:
k ) 1, n
(N kki ∆xk) - cT(uout - uin) ∑ ∑ k)1 i)1
In the top-to-top equalization (column 1 provides equalization to column 2, step 3), the boundary conditions are as follows:
z ) 0:
N kki , ∑ i)1
where uout and uin are the outlet and inlet velocities (uout ) uFn and uin ) u0F), given by eq 5. The partial mass balance is given by
1 ∂ci ) u(ci - ci|z ) L,column1); Dax ∂z
z ) 0:
nc
-
where ∆xk is the volume in k stage. From eq 9, assuming negligible pressure drop, we obtain the velocity profile, i.e.,
∂ci ) 0; u ) 0 ∂z
When column 1 is producing and column 2 is purging (step 2), the following boundary conditions change:
z ) L: u )
LSODA.44 The routine LSODA solves initial boundary problems for stiff or nonstiff systems of first-order ordinary differential equations (ODEs). For nonstiff systems, it makes use of the Adams method with variable order (up to 12th order) and step size, while for stiff systems it uses the Gear (or BDF) method with variable order (up to 5th order) and step size. In the finite volume method, the values of the conserved variables (for example, molar concentration) are averaged across the volume and the conservation principle is always assured. Figure 2 presents the finite volume discretization method in a k schematic form. ukF is the velocity in the face k, and cF,i is the concentration of i species in the face k. The equivalent total mass balance equation using the finite volume method is given by the following equation,
k+1 k k (ck+1 i /cT ) - (ci /cT)
/2(∆xk+1 + ∆xk)
1
(14)
k ci,F is the concentration of species i in the face k and is a function of the neighboring cells, as follows:
k k k+1 k+2 ) f(ck-1 ci,F i , ci , ci , ci )
(15)
Several methods have been proposed in the literature for the k (concentration of species i in the face k), calculation of ci,F such as the first-order upwind differencing scheme (UDS) of k Courant el al.,46 ci,F ) cki , i ) 1, ..., nc; the second-order linear k upwind scheme (LUDS) of Shyy,47 ci,F ) 3/2cki - 1/2ck-1 i , i ) 1, k ..., nc; or the third-order QUICK scheme of Leonard,48 ci,F )
Ind. Eng. Chem. Res., Vol. 46, No. 2, 2007 595
Figure 4. Adsorption equilibrium isotherms for nitrogen (4), oxygen (0), and argon (O) at 11.7 °C for (a) Oxysiv 5 and (b) AgLiLSX. Table 4. Operating Conditions of the Experiments with AgLiLSX at 30 °C
Table 5. Experimental and Simulation Results of the PSA Unit with AgLiLSX
feed composition (%)
experimental results (%)
run
O2
Ar
N2
tpress (s)
1 2 3 4 5 6 7 8 9
21 21 21 21 21 95 95 95 95
0 0 0 1 1 5 5 5 5
79 79 79 78 78 0 0 0 0
6.6 6.6 6.6 6.6 6.6 6.6 6.6 6.6 6.6
tprod (s)
teq (s)
product flow rate (mLSTP/min)
8.28 48.28 8.28 8.28 8.28 18.28 18.28 19.28 12.28
2.1 32.1 2.1 2.1 2.1 12.1 12.1 16.1 16.1
100 30 30 30 50 80 30 30 50
+ cki ) - 1/8(ck+1 + 2cki + ck-1 i i ), i ) 1, ..., nc, which are all upwind biased. Central schemes are often used, such as the k ) 1/2(ck+1 second-order central differences (CDS2), ci,F + cki ), i i ) 1, ..., nc, or the fourth-order central differences (CDS4). All these methods, with the exception of the first-order UDS, suffer from lack of boundedness, and for highly convective flows, the occurrence of unphysical oscillations is usual. The UDS scheme is referred to in the literature using different names: successive stages method, mixed cells in series model, or cascade of perfectly mixed tanks. It is well-known in the context of separation processes;49-52 however, it has only firstorder accuracy and is not normally recommended.53 In order to overcome the occurrence of nonphysical oscillations, an extensive amount of research has been directed toward the development of accurate and bounded nonlinear convective schemes. Several discretization schemes were proposed on the total variation-diminishing framework (TVD)47,54 and, more recently, on the normalized variable formulation (NVF)55 and its extension, the normalized variable and space formulation (NVSF) of Darwish and Moukalled.56 Two different bounded approaches are nowadays commonly used: high-resolution schemes (HRS) and weighted essentially nonoscillatory (WENO) schemes. In this work, we use bounded higher-order schemes following the χ scheme formulation proposed by Darwish and Moukalled,57 which are briefly described below. Considering a general grid, as illustrated in Figure 3, the labeling of the nodes depends on the local velocity, uF, calculated at face F. For a given face F, the U and D nodes refer to the upstream and downstream points, relative to node P, which is itself upstream to the face F under consideration, as shown in Figure 3. According to the NVSF, the face values are interpolated as56 1/ (ck+1 2 i
yF ) yU +y˜ F(yD - yU) where y is the convected variable (for example ci).
(16)
simulation results (%)
run
Pur
Rec
Pur
Rec
1 2 3 4 5 6 7 8 9
100 100 100 98.73 98.64 98.73 99.80 99.65 98.98
19.80 20.41 5.65 5.64 7.60 7.42 2.94 2.93 4.29
100 100 100 98.71 98.55 98.75 99.92 99.80 99.93
19.76 20.35 5.50 5.60 7.45 7.45 3.03 3.10 4.33
The normalized face value, y˜ F, is calculated using an appropriate nonlinear limiter. As an example, we present the SMART58 limiter (third-order convergence in smooth regions),
[ (
y˜ F ) max y˜ P, min
x˜ F(1 - 3x˜ P + 2x˜ F) x˜ P(1 - x˜ P)
x˜ F(1 - x˜ F)
y˜ + x˜ P(1 - x˜ P) P x˜ F(x˜ F - x˜ P) , 1 (17) 1 - x˜ P
y˜ P,
)]
and the MINMOD limiter54 (second-order convergence in smooth regions),
[ (
y˜ F ) max y˜ P, min
)]
1 - x˜ F x˜ F - x˜ P x˜ F y˜ P, y˜ P + x˜ P 1 - x˜ P 1 -x˜ P
(18)
where the normalized variables y˜ P, x˜ P, and x˜ F are calculated using
y˜ P )
yP - yU x P - xU x F - xU , x˜ P ) , x˜ F ) yD - yU xD - xU xD - xU
(19)
More details on this issue, and other high-resolution schemes, can be found in the works of Darwish and Moukalled56 and Alves et al.59 Although these high-resolution schemes are very efficient and bounded, for systems with more than two components, their use may result in inconsistencies in the mass balances and may lead to unphysical solutions. For these cases, the χ-schemes should be used. The χ-schemes are a new class of highresolution schemes that combine consistency, accuracy, and boundedness across systems of equations.57 This new formulation is based on the observation that the upwind scheme and all high-order schemes are consistent. So, if the high-resolution (HR) scheme at a control volume face is forced to share across the system of equations the same linear combination of highorder (HO) schemes, then it will be consistent.57 Otherwise, the mass balance may not be consistent. In this formulation, the
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Table 6. Valve Coefficients of the PSA Unit with AgLiLSX run
CPv
CEv
run
CPv
CEv
1 2 3 4 5
0.0004 0.0004 0.0006 0.0005 0.0010
0.0020 0.0020 0.0050 0.0030 0.0020
6 7 8 9
0.0010 0.0010 0.0010 0.0016
0.0048 0.0048 0.0015 0.0017
value at a control volume face using a high-resolution scheme is written as
y˜ HR ˜ U + χ( y˜ HR ˜ U) F )y F -y
{
(20)
The NVF form of the SMART scheme is written as57
1 6 1 5 < y˜ U < 6 6 5 < y˜ U < 1 6 elsewhere
Figure 5. Simulated and experimental pressure history, inside the columns, achieved on the unit studied for run 5 ((- - -) simulation, (0) experimental).
0 < y˜ U
