Ind. Eng. Chem. Res. 2007, 46, 591-599
591
High-Purity Oxygen Production by Pressure Swing Adsorption
J. C. Santos, P. Cruz, T. Regala, F. D. Magalhães, and A. Mendes*
LEPAEsDepartamento de Engenharia Quı́mica, 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
592
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
594
Ind. Eng. Chem. Res., Vol. 46, No. 2, 2007
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 +ỹF(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, ỹF, is calculated using an
appropriate nonlinear limiter. As an example, we present the
SMART58 limiter (third-order convergence in smooth regions),
[ (
ỹF ) max ỹP, min
x̃F(1 - 3x̃P + 2x̃F)
x̃P(1 - x̃P)
x̃F(1 - x̃F)
ỹ +
x̃P(1 - x̃P) P
x̃F(x̃F - x̃P)
, 1 (17)
1 - x̃P
ỹP,
)]
and the MINMOD limiter54 (second-order convergence in
smooth regions),
[ (
ỹF ) max ỹP, min
)]
1 - x̃F
x̃F - x̃P
x̃F
ỹP,
ỹP +
x̃P
1 - x̃P
1 -x̃P
(18)
where the normalized variables ỹP, x̃P, and x̃F are calculated using
ỹP )
yP - yU
x P - xU
x F - xU
, x̃P )
, x̃F )
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
596
Ind. Eng. Chem. Res., Vol. 46, No. 2, 2007
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
HR
ỹHR
F ) ỹU + χ( ỹF - ỹU)
{
(20)
The NVF form of the SMART scheme is written as57
1
6
1
5
< ỹU <
6
6
5
< ỹU < 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 < ỹU <
3ỹU
3 3
+ ỹ
ỹF ) 8 4 U
1
ỹU
{
(21)
The equivalent χ formulation is57
16ỹU
3 - 2ỹU
1
6
5
1
< ỹU <
1
6
6
χ)
8(1 - ỹU) 5
< ỹU < 1
3 - 2ỹU 6
0
elsewhere
0 < ỹU <
(22)
Consistency is ensured by forcing all the related equations to
have the same values of χ at any control volume face.57
Adsorption Equilibrium
The single-component adsorption equilibrium isotherms for
AgLiLSX and for Oxysiv 5 were obtained experimentally by
the gravimetric method using a Rubotherm magnetic suspension
balance,60 at three different temperatures (11.7, 23.0, and 29.9
°C for AgLiLSX and 11.7, 19.9, and 36.4 °C for Oxysiv 5).
The adsorption equilibrium was described using the Sips
equation,39
q ) qmax
(bp)1/n
(23)
1 + (bp)1/n
Figure 6. Simulated and experimental pressure history, inside the columns,
achieved on the unit studied for run 9 ((- - -) simulation, (0) experimental).
For Oxysiv 5, the multicomponent adsorption equilibrium was
predicted using the Sips multicomponent equation since the
saturation capacity and the parameter n are the same for the
three species, which indicates consistency from a thermodynamic point of view.39 For AgLiLSX, these parameters are not
the same, and so the multicomponent adsorption equilibrium
was predicted using the ideal adsorption solution theory
(IAST).39
The experimental adsorption data for oxygen, nitrogen, and
argon at 11.7 °C for Oxysiv 5 and AgLiLSX are presented in
Figure 4.
The pore-size distribution was determined using a mercury
porosimeter from Quantachrome (model Poremaster 60). This
porosimeter allows one to determine the pore-size distribution
ranging from 200 µm to 35 Å, by analyzing both intrusion and
extrusion of mercury in the pores. The average diameter of the
pores obtained for both adsorbents was 0.5 µm. The porosity
obtained was 0.336 for Oxysiv 5 and 0.471 for AgLiLSX.
where
[ (
)]
(
)
Results and Discussion
Tref
1
Q Tref
1
+R 1,
-1 ,
)
b ) bref exp
RTref T
n nref
T
[(
and qmax ) qmax,ref exp β 1 -
T
Tref
)]
bref is the affinity constant at the reference temperature, Tref (in
this work, it was considered to be 20 °C), Q is the heat of
adsorption, T is the temperature, qmax,ref is the maximum
adsorption capacity at the reference temperature, and nref, R,
and β are empirical parameters. The experimental values at the
three different temperatures were used for obtaining the Langmuir-Freundlich equation parameters, presented in Table 3.
For avoiding any contamination of the adsorbent, mixtures
of pure gases were prepared in the feed tank with the following
concentrations: 78.12% of nitrogen, 20.95% of oxygen, and
0.93% of argon (similar to air) and 95% of oxygen and 5% of
argon (such as a product from a PSA unit).
Table 4 presents the operating conditions of the experiments
with AgLiLSX at 30 °C (runs 1-3 were conducted with pure
reconstituted air).
The experimental and simulation results (purity, Pur, and
recovery, Rec) of the runs presented in Table 4 are presented
in Table 5. As can be seen in Table 5, this adsorbent allows the
production of high-purity oxygen.
Ind. Eng. Chem. Res., Vol. 46, No. 2, 2007 597
Table 6 presents the valve coefficients determined experimentally (obtained measuring the flow rate passing through the
valves at different pressures) for the runs presented before.
The feed, CFv , vent, CVv , and storage tank, CTv , valve parameters were determined in the same way and are, respectively,
0.025, 0.025, and 0.02.
Figures 5 and 6 present the simulated and experimental
pressure histories for runs 5 and 9, respectively.
By analyzing Table 5 and Figure 5, it is possible to conclude
that the simulator was able to accurately represent the experimental results. It is worth mentioning that the other runs also
show a good match between the simulation and experimental
results.
PSA Optimization
This PSA unit was optimized by simulation, minimizing the
recovery, for producing 50 mLSTP/min of 99.5% of oxygen from
air, operating between 1 and 3 bar at 30 °C. The optimization
procedure described by Cruz et al.36 was used, and the
optimization variables were the pressurization time, tpress, the
production time, tprod, the equalization time, teq, the purge valve
coefficient, CPv , and the equalization valve coefficient, CEv . An
optimum recovery of 5.03% was obtained for tpress ) 6.6 s, tprod
) 8 s, teq ) 3.1 s, CPv ) 0.0015, and CEv ) 0.0020.
As mentioned in the Introduction, Jee et al. presented a PSA
unit with two adsorption columns with ∼380 mL packed with
CMS adsorbent from Takeda.29 The feed was 95% of oxygen
and 5% of argon, and its flow rate was 4 LSTP/min. The high
pressure was 5 bar. This unit produces ∼1.6 LSTP/min of 99.87%
oxygen with a recovery of 56.90%. This unit coupled with a
PSA unit packed with Oxysiv 5 would have a global recovery
∼17%.2 It would be composed by two columns with ∼2 L
(packed with Oxysiv 5) and two columns with 380 mL (packed
with CMS adsorbent).2 If Oxysiv MDX was used instead, it
would have a global recovery ∼45%.2 It would be composed
by two columns with ∼1 L only (packed with Oxysiv MDX)
and two columns with 380 mL (packed with CMS adsorbent).2 The drawback of this technology is that the product is
obtained at a pressure close to the atmospheric pressure, and
so, a third compressor might be needed to pressurize it. With
AgLiLSX and with the cycle presented to produce 1.6 LSTP/
min of 99.5% oxygen, only two columns with ∼3.2 L would
be needed.
Although the recovery given by this combination of two units
is high, the unit is large and complex. In some applications
where the size is important, it might be advisable to use
AgLiLSX to directly produce high-purity oxygen or even a
combination of a PSA with Oxysiv MDX, for example, with
another PSA with AgLiLSX (although this combination is also
complex, the product is obtained at high pressure).
A PSA unit using the Skarstrom cycle with top-to-top
equalization allowed the production of high-purity oxygen
directly from air.
The simulator was able to represent accurately a real unit
using a feed without carbon dioxide and water vapor. With the
optimization strategy, it was possible to achieve better operating
conditions and design characteristics that improved the performance of the units. An optimal recovery of 5.03% was obtained
for producing 99.5% of oxygen from air with a PSA operating
between 1 and 3 bar at 30 °C.
Acknowledgment
The work of João Carlos Santos and Paulo Cruz was
supported by FCT, Grants SFRH/BD/6817/2001 and POSI
SFRH/BPD/13539/2003, respectively. The research was also
supported by funds from Growth GRD1-2001-40257 and FCT
project POCTI/EQU/38067/2001.
Acknowledgment
The authors would like to acknowledge Air Products and
Chemicals, Inc., and Dr. Roger D. Whitley for offering the
adsorbent.
Nomenclature
A ) area, m2
b ) Langmuir affinity constant, bar-1
c ) fluid-phase molar concentration, mol m-3
Cv ) valve parameter
Dax ) effective axial dispersion coefficient, m2 s-1
DeM ) effective homogeneous diffusion coefficient, m s-1
d ) diameter, m
k ) iteration or LDF coefficient, ki ) 15DM,ie/rp2, s-1
Kv ) valve parameter
L ) column length, m
M ) molecular weight, kg mol-1
n ) empirical parameter from Langmuir-Freundlich equation
N ) molar flow rate, mol s-1
nc ) number of mixture components
p ) pressure, bar
Pur ) product purity
Q ) heat of adsorption, J mol-1
q ) molar concentration in the adsorbed phase, mol kg-1
qmax ) isotherm parameter, maximum adsorption capacity, mol
kg-1
rp ) particle radius, m
R ) universal gas constant, J mol-1 K-1
Rec ) product recovery
t ) time variable, s
T ) temperature, K
u ) average (interstitial molar) velocity, m s-1
z ) spatial coordinate, m
Conclusions
Greek Symbols
A study about the high-purity oxygen production from air in
a single stage by pressure swing adsorption (PSA) using an
adsorbent patented by Air Products and Chemicals was presented. The adsorption equilibrium measurements indicate that
this adsorbent has large nitrogen, oxygen, and argon capacities
and selectivities.
The adsorbents studied (Oxysiv 5 and AgLiLSX) showed to
have similar average pore diameters, 0.5 µm; however, they
presented different porosities: 0.336 for Oxysiv 5 and 0.471
for AgLiLSX.
R ) empirical parameter from Langmuir-Freundlich equation
β ) empirical parameter from Langmuir-Freundlich equation
b ) bed void fraction (ratio between the free volume and the
total volume)
p ) particle porosity
Fs ) apparent particle density, kg m-3
τ ) tortuosity factor
Subscripts
d ) downstream
598
Ind. Eng. Chem. Res., Vol. 46, No. 2, 2007
eq ) equalization
i ) component
in ) feed stream (inlet)
out ) outlet stream
p ) particle
pres ) pressurization
prod ) production
purg ) purge
ref ) reference
T ) total
u ) upstream
Superscripts
E ) equalization
F ) feed
H ) high
L ) low
P ) purge
ST ) storage tank
STP ) standard temperature and pressure conditions
V ) vent
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ReceiVed for reView March 30, 2006
ReVised manuscript receiVed November 2, 2006
Accepted November 13, 2006
IE060400G