Two-Stage VPSA Process for CO2 Capture from Flue Gas Using

Article pubs.acs.org/IECR

Two-Stage VPSA Process for CO2 Capture from Flue Gas Using Activated Carbon Beads Chunzhi Shen, Zhen Liu, Ping Li,* and Jianguo Yu* State Key Laboratory of Chemical Engineering, College of Chemical Engineering, East China University of Science and Technology, Shanghai 200237, China ABSTRACT: Carbon dioxide removal from flue gas with a two-stage vacuum pressure swing adsorption (VPSA) process, which uses activated carbon (AC) beads as the adsorbent, was investigated both theoretically and experimentally. First, single-column VPSA experiments were studied for CO2/N2 separation with high CO2 feed concentration. Then, a two-stage VPSA process composed two columns for each stage was designed, and the effects of different parameters were investigated. The first-stage VPSA unit operates with a four-step Skarstrom cycle, which includes feed pressurization, adsorption, blowdown, and countercurrent purge with N2. For the second-stage VPSA process, a cycle with feed pressurization, adsorption, pressure equalization, blowdown and pressure equalization was employed. With the proposed two-stage VPSA process, a CO2 purity of 95.3% was obtained with 74.4% recovery. The total specific power consumption of the two-stage VPSA process is 723.6 kJ/kg-CO2, while the unit productivity is 0.85 mol-CO2/kg·h.

1. INTRODUCTION With the increasing focus on global warming mainly caused by a large quantity of CO2 emissions, carbon capture and storage (CCS) is seen as an effective mitigation strategy and has attracted considerable research development and demonstration efforts over the last two decades.1−3 In particular, since the capture cost is the major component of the overall cost of CCS, many capture technology solutions have been proposed and investigated, such as monoethanolamine (MEA) chemical absorption, membranes, cryogenic and adsorption.3 Among the capture technology candidates, adsorption process is a promising option for separating CO2 from flue gas because of the reusable nature of the adsorbents used, its low capital investment cost and easiness to achieve automatic operation. Particularly, pressure/vacuum swing adsorption (PSA/VSA) processes for capturing CO2 from fossil fuel combustion have been intensively studied.4−8 The configuration of the PSA/VSA process for CO2 recovery varies according to the CO2 concentration of the feed. When the feed concentration of CO2 is higher than 25%, high purity of CO2 (over 99%) can be easily produced by zeolite 13X at the recovery of 70% using a one-stage PSA.5 However, for the recovery of CO2 from low concentration flue gases (e.g., 15% CO2), it is difficult to achieve high CO2 purity with high recovery using the one-stage VPSA. To improve the PSA performance, a lot of research effort has been attracted to develop a more efficient PSA cycle. Ishibashi et al.9 investigated the performance of the pressure and temperature swing adsorption (PTSA) process for a 1000 N m3/h scale pilot plant. The pilot plant completed 2000 h of continuous operation without incident and no decrease in the effectiveness of the CO2 adsorbent was observed after 4000 h of operation. Reynolds et al.10 used high temperature PSA with potassium promoted hydrotalcite (HTlc) as adsorbent, and the purge/feed ratio, cycle step time, and pressure ratio were studied and optimized. Ko et al.11 conducted optimizations of PSA and fractionate vacuum pressure swing adsorption (FVPSA) by © 2012 American Chemical Society

simulation under normal temperature and high temperature and concluded that high temperature FVPSA is much better than PSA, but the average power for FVPSA is greater than that for PSA.12 A dual reflux PSA process was studied by Diagne et al.8 where feed is introduced to an intermediate position of the adsorber not to the adsorber end. They showed that the process can enrich both the weakly adsorbed component and strongly adsorbed component simultaneously. Moreover, Park et al.13 introduced a two-stage PSA process using 13X and made numerical analysis on the power consumption of the first stage PSA. In spite of the large number of studies, most of the research was concerned with the adsorption process using zeolites (13X or 5A) as the adsorbent. Compared to zeolites, carbonaceous materials are water tolerant and can be produced with novel morphologies (monolith, bead, fiber, granular, respectively). Additionally, they are less expensive than other adsorbents such as zeolites.14 The spherical nature and hardness of AC beads minimizes dust formation and attrition losses during adsorption and regeneration processes. They also exhibit excellent fluidization properties both in gas and liquid application. These characteristics make AC bead the material of choice for higher performance in carbonaceous materials application. However, carbonaceous materials have relative low CO2 capacity and low selectivity of CO2/N2 in the comparison with zeolites, which result in the difficulty to concentrate CO2 from flue gas up to above 95% (which is required for CO2 storage) directly with a relative high recovery. To use carbonaceous materials for CO2 capture, high feed pressure is usually employed to obtain the separation objective. To compress the flue gas directly from atmospheric pressure to a high pressure, high energy consumption is Received: Revised: Accepted: Published: 5011

September 13, 2011 March 10, 2012 March 10, 2012 March 10, 2012 dx.doi.org/10.1021/ie202097y | Ind. Eng. Chem. Res. 2012, 51, 5011−5021

Industrial & Engineering Chemistry Research

Article

required because it contains about 85% of N2. Therefore, a process comprising two different VPSA units operating in series was designed in this work to concentrate CO2 from flue gas to above 95% with relative high recovery using AC beads. At the first stage of the two-stage VPSA, CO2 is concentrated to about 40−60%, with the flue gas feeding at almost atmospheric pressure. Then, the product of the first stage is compressed and feed to the second stage of the two-stage VPSA, where CO2 is further concentrated to above 95%. By this way, the power consumption for compressing the product is much less than that for compressing the flue gas directly. On the other hand, a product with purity higher than 95% can be obtained. Before the two-stage VPSA numerical simulation, single-column VPSA experiments were performed with high CO2 concentration in the feed stream to study the effect of different operation conditions, also to supply optimal operating conditions for the simulation of two-stage VPSA units. Typical flue gas from coalfired power plants contains N2, CO2, H2O, and minor impurities such as SOx and NOx. It was assumed that H2O, SOx, and NOx were removed in the pretreatment process so that a simulating flue gas composed of CO2 and N2 was employed in the study.

Table 1. Physical Properties of Adsorbent and Column Characteristics Used in the VPSA Experiments for CO2/N2 Separation column length [m] inner column radius [m] column porosity bulk density [kg/m3] weight of adsorbent [g] adsorbent specific area [m2/kg] pellet diameter [m] pellet density, [kg/m3] pellet porosity solid specific heat [J/(kg·K)] specific heat of column wall [J/(kg·K)] thickness of column wall [mm]

0.557 0.0125 0.32 671 183.3 846 1−1.18 × 10−3 984 0.51 850 500 3.5

and micropores. The gas mixture interacts with the solid surface following the mass transfer governing equations between gas− solid interface, gas−crystal inside the pellet and the thermodynamic equilibrium. Temperature inside the pellet is seen as uniform. Column wall is the shell of the fixed bed and only interchanges energy with the gas phase inside the column and with the external environment. The resulting model of the fixed bed is a combination of the mass, momentum and energy conservation equations. The model assumptions adopted are summarized below:15 (1) The gas phase behaves as an ideal gas. (2) No mass, heat, or velocity variations in the radial direction. (3) Macropore and micropore mass transfer resistances are expressed with the linear driving force (LDF) model. (4) The bed porosity is uniform along the bed. The mass balance for each component in the gas phase is given by

2. EXPERIMENTAL SECTION VPSA experiments were performed in our laboratory unit, which mainly contains three sections including gas mixture section, VPSA column section, and analytical section. Desired feed concentration and gas flow rate were controlled by mass flow controllers in the gas mixture section, while the effluents leaving the column were analyzed in the analytical section, which contains a gas chromatograph and a CO2 infrared online analyzer. The concentric column has an inner-tube inner diameter of 25 mm and out-tube inner diameter of 55 mm. The inner tube is the adsorption column filled with AC beads, and the annulus is filled circularly with water to keep the inner packed column at a specific temperature. The temperature of the water is controlled by a thermostat bath. The VPSA equipment is connected to a computer where the individual gas flow rates and pressures at the inlet and outlet of the column are stored together with temperature measured in two different points of the column (0.04 and 0.507 m from the inlet). A full description of this laboratory unit is found elsewhere.15 A four-step cycle sequence including feed pressurization, adsorption, cocurrent depressurization, and blowdown was employed for the second-stage single-column VPSA experiments. Because the main objective of the second-stage VPSA process is to further purify CO2, an N2 purge step, which may lead to the decrease of product purity, was removed, as compared to the first-stage VPSA cycle reported in the previous paper.15 A cocurrent depressurization step was introduced because the column was at high pressure after adsorption step. The AC beads used here are synthesized from coal tar pitch through the emulsion method by the cooperators in our laboratory.16 Some parameters relative to the column and AC beads are summarized in Table 1.

εc

∂y ⎞ ∂(uCi) ∂Ci ∂⎛ − (1 − εc) = ⎜εcDax, iCT i ⎟ − ∂z ⎠ ∂z ∂t ∂z ⎝ a′k f, i Bii + 1

× (Ci − ci̅ )

(1)

where εc is the column porosity, Ci is the gas phase concentration of component i, Dax,i is the axial dispersion coefficient, CT is the total gas concentration, yi is the molar fraction, u is the superficial velocity, a′ is the pellet specific area, kf,i is the film mass transfer resistance, Bii = Rpkf,i/(5εpDp,i) is the Biot number, and ci̅ is the averaged concentration in the macropores. The LDF model used to describe the mass transfer rate from the gas phase to the macropores can be expressed as follows: εp

15Dp, i Bii ∂⟨q ̅ ⟩ ∂ ci̅ + ρp i = ε p (Ci − ci̅ ) ∂t ∂t R p2 Bii + 1

(2)

where Dp,i is the pore diffusivity, Rp is the pellet radius, ρp is the particle density, εp is the pellet porosity, and ⟨qi̅ ⟩ is the pellet averaged adsorbed phase concentration. The LDF equation for the crystals averaged over the entire pellet is

3. MATHEMATICAL MODELS 3.1. Fixed Bed Adsorption Model. The real adsorption column is mainly divided in gas phase, solid phase, and column wall. The gas phase exchanges mass and energy with the solid phase while only energy is exchanged with the column wall. The adsorbent particles are considered as bidisperse with macropores

∂⟨qi̅ ⟩ ∂t 5012

=

15Dc, i rc2

(qi* − ⟨qi̅ ⟩)

(3)

dx.doi.org/10.1021/ie202097y | Ind. Eng. Chem. Res. 2012, 51, 5011−5021


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Figure 1. Adsorption equilibrium of carbon dioxide (a) and nitrogen (b) on activated carbon beads. Solid lines: Virial model.

where Dc,i is the crystal diffusivity, rc is the crystal radius, and qi* is the adsorbed phase concentration in equilibrium with the concentration of component i averaged over the pellet, ci̅ . The energy balance of the gas phase can be written as

εcC TC̃ v

∂Tg ∂t

=

with αw =

∂Tg ∂C ∂ ⎛ ∂Tg ⎞ ⎜λ ⎟ − uC TC̃ p + εcR gTg T ∂z ∂t ∂z ⎝ ∂z ⎠ − (1 − εc)a′h f (Tg − Ts) 2h − w (Tg − Tw ) Rw

α wl =

(1 − εc)[ε p

(4)

i=1 n

∂⟨q ̅ ⟩ ∂Ts ∂⟨c ⟩ = (1 − εc)ε pR gTS i + ρ b ∑ ( −ΔHi) i ∂t ∂t ∂T i=1

(8)

(9)

⎛ N ⎞ N N 3 ⎜2 ⎟ exp⎜ ∑ A ijqj+ 2 ∑ ∑ Bijk qjqk ⎟ Pi = ⎜ ⎟ KH, i S 2S j = 1 j = 1 ⎝ j=1 ⎠ qi

(5)

where ρb is the bulk density of the column, C̃ v,ads,i is the molar constant volumetric specific heat of component i in the adsorbed phase, and (−ΔHi) is the isosteric heat of adsorption of component i. The wall energy balance can be expressed by ρw C̃ pw

⎛D + e⎞ ⎟ (Dw + e) ln⎜ w ⎝ Dw ⎠

where P is the total pressure, μg is the gas viscosity, dp is the pellet diameter, and ρg is the gas density. The pure adsorption equilibrium data was fitted with Virial isotherm model, and its multicomponent extension was used to predict the behavior of the binary CO2−N2 mixture. The multicomponent extension of this model is17−19

n

+ (1 − εc)a′h f (Tg − Ts)

1

150μg(1 − εc)2 1.75(1 − εc)ρg ∂P =− + |u|u u ∂z ε3c d p2 ε3c d p

∑ ⟨ci⟩C̃ v, i + ρp ∑ ⟨qi̅ ⟩C̃ v,ads, i + ρpC̃ ps] i=1

(7)

where ρw is the density of the column wall, C̃ pw is the specific heat of the column wall, αw is the ratio of the internal surface area to the volume of the column wall, αwl is the ratio of the logarithmic mean surface area of the column shell to the volume of the column wall, Dw is the internal diameter of the column, e is the wall thickness, U is the global external heat transfer coefficient, and T∞ is the environmental temperature. The momentum balance only considers the terms of pressure drop and velocity changes and relates them by the Ergun equation:

where C̃ v is the molar constant volumetric specific heat of the gas mixture, Tg is the temperature of the gas phase, λ is the axial heat dispersion, C̃ p is the molar constant pressure specific heat of the gas mixture, Rg is the universal gas constant, hf is the film heat transfer coefficient between the gas and the solid phase, Ts is the solid temperature, hw is the film heat transfer coefficient between the gas phase and the column wall, Rw is the column radius, and Tw is the wall temperature. The solid-phase energy balance is described as follows: n

Dw e(Dw + e)

(10)

with A ij =

∂Tw = α w h w (Tg − Tw ) − α w1U (Tw − T∞) ∂t

Ai + Aj 2

Bijk =

Bi + Bj + Bk 3

where S is the adsorbent specific surface area, A and B are Virial coefficients, and KH is the Henry constant ([mol·g−1·kPa−1]),

(6) 5013



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Table 2. Parameters for Pure-Component Virial Equation and Micropore Diffusion Constants of CO2 and N2 on AC Beads Used For VPSA Simulation gas CO2 N2

K∞ [mol/(kg·kPa)] −6

3.41 × 10 2.38 × 10−6

−ΔH0 [kJ/mol]

A0 [kg/mol]

A1 [(kg·K)/mol]

B0 [(kg/mol)2]

B1 [(kg/mol)2·K]

0 Dc,i /rc2 [s−1]

Ea,i [kJ/mol]

23.17 18.11

6.889 −2.017

−0.0871 688.392

0.0244 0.348

−8.652 −105.164

12.995 9.492

18.050 12.391

different parameters: product purity, product recovery, unit productivity, and the total power consumption of the process. The unit productivity is calculated as the amount of CO2 produced per mass of adsorbent and per hour. The product purity, product recovery and the unit productivity are defined as follows:

which has an exponential dependence with temperature described by Van’t Hoff equation: ⎛ ⎞ ΔH 0 ⎟ KH = K∞ exp⎜⎜ − ⎟ ⎝ R gT ⎠

(11)

where K∞ is the adsorption constant at infinite temperature and −ΔH0 is the heat of adsorption at zero coverage. Adsorption equilibrium of pure carbon dioxide and nitrogen in activated carbon beads at different temperatures are reported in Figure 1. It can be observed that carbon dioxide is the strongly adsorbed species on activated carbon beads toward nitrogen. Adsorption equilibrium and kinetic parameters for pure CO2 and N2 are shown in Table 2.20 The physical properties of CO2 and N2 can be found elsewhere.21 The fixed bed model involves several transport parameters that were calculated using frequently used correlations list in Table 3.22−26 The microspore diffusivities were measured in

⎧ t blow purity = ⎨ CCO2u|Z = 0 dt ⎩ 0

∫

+

⎧ t blow /⎨ (CCO2 + C N2)u|Z = 0 dt ⎩ 0 t purge ⎫ (CCO2 + C N2)u|Z = 0 dt ⎬ + ⎭ 0

∫

∫ t

Table 3. Correlations Used for Estimation of Mass and Heat Transfer Parameters axial mass dispersion coefficient

recovery =

εDax = 20 + 0.5Sci Re Dm, i

film mass transfer coefficient

k f , id p/Dm = 2.0 + 1.1Re0.6Sc1/3

molecular diffusion

n ⎛ y ⎞ Dm, i = (1 − yi )/ ∑ ⎜⎜ i ⎟⎟ j = 1 ⎝ Dij ⎠

D k, i = 9700rp

axial heat dispersion coefficient

λ = 7 + 0.5PrRe kg

film heat-transfer coefficient

hf d p/k g = 2.0 + 1.1Re0.6Pr1/3

t

∫0 feed

+ t press

CCO2u|Z = 0 dt

(14)

The power consumption of the VPSA process is mainly consumed by the blower and the vacuum pump. For the calculations of power consumption, we have assumed that all compression/decompression steps are performed by devices that operate under an adiabatic regime such that the specific power consumption Wt for the VPSA process can be described by the following equation:

Tg

⎛ 1 1 1 ⎞⎟ = τp⎜⎜ + Dp, i D k, i ⎟⎠ ⎝ Dm, i

t

∫0 blow CCO2u|Z = 0 dt + ∫0 purge CCO2u|Z = 0 dt (13)

Mw

macropore diffusion

(12)

t t 3600(∫ blow CCO2u|Z = 0 dt + ∫ purge CCO2u|Z = 0 dt ) 0 0 productivity = tcyclewads

j≠i

Knudsen diffusion

⎫

t

∫0 purge CCO2u|Z = 0 dt ⎬⎭

Wt = Power blower + Powervacuum ⎧ ⎤ ⎡ γ− 1/ γ ⎪ t press+ t feed + trinse γ ⎢⎛ P high ⎞ ⎥d t ⎜ ⎟ − =⎨ n1̇ R gT 1 ⎥ γ − 1 ⎢⎝ P low ⎠ ⎪ 0 ⎦ ⎣ ⎩ ⎫ ⎤ ⎡⎛ γ− 1/ γ t blow + t purge ⎪ γ ⎢ P high ⎞ ⎟ ⎜ − 1⎥⎥dt ⎬ + n2̇ R gT ⎢ γ − 1 ⎝ P low ⎠ 0 ⎦ ⎪ ⎣ ⎭ t blow t purge ⎧ 44 ⎫ ⎬ /⎨( CCO2u|Z = 0 dt + CCO2u|Z = 0 dt )A × 1000 ⎭ ⎩ 0 0

∫

our previous work20 and are listed in Table 2. The heat transfer coefficient at the wall (hw) and external convective film transfer coefficient (U) were also previously fitted by the breakthrough curves.15 The values of hw and U were 50 W/(m2·K−1) and 100 W/(m2·K−1), respectively. To solve the partial differential eqs 1−9, boundary and initial conditions are necessary. The adsorbent bed is filled with N2 before the cyclic process starts, and the final distributions of concentrations, temperature, and pressure along the column for one step are the initial conditions for the next step. In a VPSA process, different cycle steps require different boundary conditions. Thus, boundary conditions for the cycle steps used in this work are listed in Table 4. 3.2. Process Performance. The performance of the VPSA experiments and simulations is evaluated according to four

∫

∫

∫

(15)

where γ = CP/Cv, ṅ1 is the molar flow rate to be compressed in the blower, ṅ2 is the molar flow rate to be decompressed in the vacuum pump, and A is the column area. If the efficiency of the vacuum pump and blower is 1, then the real power consumption of those machines are given by eq 15. However, because of the mechanical and electrical losses, the real power consumption of the machines is higher than that calculated by eq 15. Since the mechanical efficiencies depend on the type of the vacuum pump, system configuration, and manufacturer, it is 5014



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Table 4. Boundary Conditions for Different Steps Used in the VPSA experiments and Simulations Employing AC Beads pressurization with feed inlet, z = 0

z=L

P(0) = Pinlet

−

ui(L) = 0

εcDax, i ∂y(i , 0) ui(0) ∂z

−λ

∂Tg(0)

∂y(i , L) =0 ∂z z−

+ y(i , 0)|z + − y(i , 0)|z − = 0 z+

∂Tg(L)

+ uiC tC̃ pTg(0)|z + − uiC tC̃ pTg(0)|z − = 0

∂z

∂z

z+

=0 z−

feed inlet, z = 0

outlet, z = L

P(L) = Pexit

ui(0) C(i , 0)|z + = ui(0) C(i , 0)|z − −

εcDax, i ∂y(i , 0) ∂z ui(0)

−λ

∂Tg(0)

∂y(i , L) =0 ∂z z−

+ y(i , 0)|z + − y(i , 0)|z − = 0 z+

∂Tg(L)

+ uiCt C̃ pTg(0)|z + − uiCt C̃ pTg(0)|z − = 0

∂z

∂z

z+

=0 z−

pressure equalization (depressurization) z=0

outlet, z = L

ui(0) = 0

P(L) = Pexit

∂y(i , 0) =0 ∂z z+

∂y(i , L) =0 ∂z z−

∂Tg(0)

∂Tg(L)

∂z

=0

∂z

z+

=0 z−

counter-current blowdown outlet, z = 0

inlet, z = L

P(0) = Pexit

ui(L) = 0

∂y(i , 0) =0 ∂z z+

∂y(i , L) =0 ∂z z−

∂Tg(0)

∂Tg(L)

∂z

=0

∂z

z+

=0 z−

counter-current purge (N2) outlet, z = 0

inlet, z = L

P(0) = Pexit

ui(L) C(i , L)|z + = ui(L) C(i , L)|z −

∂y(i , 0) =0 ∂z z+

−

∂Tg(0) ∂z

λ

=0 z+

εcDax, i ∂y(i , L) ∂z ui(L)

∂Tg(L) ∂z

+ y(i , L)|z − − y(i , L)|z + = 0 z−

+ uiC tC̃ pTg(L)|z − − uiC tC̃ pTg(0)|z + = 0 z−

pressure equalization (pressurization) z=0

inlet, z = L

ui(0) = 0

ui(L) C(i , L)|z + = ui(L) C(i , L)|z −

∂y(i , 0) =0 ∂z z+

−

∂Tg(0) ∂z

λ

=0 z+

εcDax, i ∂y(i , L) ui(L) ∂z ∂Tg(L) ∂z

+ y(i , L)|z − − y(i , L)|z + = 0 z−

+ uiC tC̃ pTg(L)|z − − uiC tC̃ pTg(0)|z + = 0 z−

3.3. Numerical Method. The resulting mathematical model consists of a set of partial differential and algebraic equations that is solved in gPROMS (PSE, UK). Orthogonal collocation on

of little use to calculate the power consumption with a specific efficiency. Therefore, the idea power consumption given by eq 15 was used to compare the process performance. 5015



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Table 5. Experimental Conditions and Performance of the Second-Stage VPSA Experiments for the Separation of CO2/N2 Using AC Beads casea VPSA VPSA VPSA VPSA VPSA a

1 2 3 4 5

tfeed [s]

tpres [s]

tblow [s]

yCO2 [%]

purity [%]

recovery [%]

productivity [mol/(kg·h)]

600 600 600 420 300

180 210 240 210 210

430 550 610 460 370

0.40 0.50 0.60 0.50 0.50

84.96 88.75 94.14 87.50 83.98

86.04 87.22 85.08 91.03 94.44

3.16 3.67 4.10 3.74 3.91

For all the VPSA experiments: Tfeed = 303 K; Pfeed = 202.65 kPa; Plow = 10 kPa; Qfeed = 1.0 SLPM; tdes = 20 s.

Figure 2. Experimental and simulated results from the second-stage VPSA process for CO2/N2 separation using AC beads, when cyclic steady state was reached (VPSA 2): (a) pressure history of one cycle; (b) CO2 molar flow rate outlet. Steps are (1) feed; (2) depressurization; (3) blowdown; (4) pressurization. Solid lines are theoretical model predictions, and solid points are experimental values.

feed pressure, operating temperature, and vacuum pressure were experimentally studied and reported in a previous paper.15 This first-stage VPSA processes could concentrate CO2 from 15% to (40−60)% with recovery varied from 40% to 85%, under a vacuum pressure of 10 kPa when the flue gas was fed at almost atmospheric pressure. Usually, two-stage VPSA processes were employed to further improve the purity of CO2. In the previous study, a VPSA process with a CO2 feed concentration of 50% was performed, and a CO2 purity of 93.7% and recovery of 78.2% were obtained, which proved the viability of the two-stage VPSA process. However, an explicit study of the second-stage VPSA was not performed yet. Therefore, mixed gas with CO2 concentrations varying from 40 to 60% was prepared as the simulated product gas of the first-stage VPSA and fed to the second-stage VPSA. Table 5 shows the main operating conditions and the performance of the experimental VPSA runs. An example (VPSA 2 in Table 5) of the results obtained in VPSA experiments is depicted in Figure 2. The pressure history and molar flow rate shown is for cycles 6−11, where no significant variation of pressure, molar flow rate, or temperature was experimentally detected and thus we may assume that the represent the behavior of the cyclic steady state. The experimental and simulation results agree quite well. CO2 loss in the feed step can be observed in Figure 2b, which may result in the decrease of recovery. Figure 3 shows the effect of different CO2 feed concentrations on the second-stage VPSA performance. With the increasing of feed CO2 concentration, the product purity and unit productivity increase significantly, while the recovery decreases. It can be seen that, with the feed concentration of 60%, a purity of 95.5% and recovery of 83.4% can be obtained with the four-step cycle, which means that the two-stage VPSA process is viable. Not only the CO2 feed concentration but also the amount of feed is one of the important operating variables. When the feed flow rate is constant, the amount of feed is directly related to the time of

finite elements method (OCFEM) is used to discrete the spatial domain with 50 finite elements and two interior collocation points in each element of the adsorption bed. The set of ordinary and algebraic equations (ODAE) were integrated with the DASOLV solver, which is based on backward-differentiation formulas (BDF). The solver uses a value of 1 × 10−5 for absolute tolerance. At all collocation points, eqs 2, 3, 6, and 9 are discretized into a set of algebraic equations that are solved numerically by the Gauss method. Equations 1, 4, and 5 are discretized in the axial direction, leading to a set of ordinary differential equations. The sequence for solving numerically the VPSA process at each step is as follows: (a) Calculate adsorption capacity of CO2/N2 at the specific P and T using eq 10. (b) Calculate the adsorptive loading derivative by eqs 2 and 3. (c) Solve eq 9 for the pressure. (d) Solve eq 1 for the component mole fraction and velocity derivative. (e) Solve eqs 4−6 for the temperature derivative. At the end of each step, the gas- and solid-phase component concentrations, temperature, and pressure were stored and used as the initial conditions for the next step in one cycle.

4. RESULTS AND DISCUSSION 4.1. Second-Stage VPSA Runs. Before the multibed numerical simulation, single-column VPSA experiments with different operating conditions were performed to validate the mathematical model and evaluate the viability of the two-stage VPSA. The first-stage VPSA processes employed a four-step Skarstrom-type cycle, including cocurrent pressurization with feed stream, feed, counter-current blowdown, and counter-current purge with N2. Separation performance of the first-stage VPSA processes with varying total feed flow rate, feed composition, 5016



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are that the feed gas amount was less and a N2 purge step was removed in this work. The results show that the purity of CO2 is influenced by the amount of both the feed gas and the purge gas. That is, the more feed gas and the less purge gas, the higher the CO2 purity and lower recovery. Also, the decrease of the purity of the cycles without N2 purge indicated that the amount of the feed gas play the dominant role to the CO2 purity under the operating conditions in this work. 4.2. Design of the Two-Stage VPSA Process. Based on the experimental study of the effect of different operating parameters, a two-stage VPSA process comprising two columns for each stage is designed to obtain a high purity above 95%, with a relative high recovery and low specific power consumption, as shown in Figure 5. This design considers that the flow rate in most of the steps will not be constant, and thus, two tanks were associated to the two-stage VPSA process. The flue gas after the pretreatment process, where the impurities such as H2O, SOx, and NOx were removed, was separated in the first-stage VPSA at almost atmospheric pressure (131.325 kPa) and room temperature (303 K). Then, the product from the first-stage VPSA process was compressed and further separated in the second-stage VPSA process. It should be noted that the real flue gas included extensive amount of water, which is strongly adsorbed in the adsorbent and difficult to be removed. In this work, we assume the water in the flue gas was removed before the two-stage VPSA processes for CO2 capture. A similar strategy was employed by Ishibashi et al.9 previously, and about 3% curtailment of the required power consumption could be obtained. The product purity of the two-stage VPSA process is the product purity of the second-stage, while the product recovery of the two-stage VPSA process equals to the recovery of the first-stage VPSA multiplied by the recovery of the second-stage VPSA. Cycle sequences and step time of the designed two-stage VPSA process for CO2 capture from flue gas using AC beads are shown in Figure 6. Four-step Skarstrom cycle comprising feed pressurization, adsorption, blowdown, and N2 purge was employed in the first-stage VPSA process. Because the main objective of the first-stage VPSA process is to obtain a high recovery (>90%), product rinse steps, which may cause the decrease of CO2 recovery, were not used. Due to the high pressure of adsorption step in the second-stage VPSA, pressure equalization was employed to save the mechanical energy while the N2 purge step, which lowers the product purity, was not used. The dimensions and properties of the columns in the first-stage VPSA process are the same as that used in the VPSA

Figure 3. Purity, recovery, and unit productivity of the second-stage VPSA process with the feed concentration (tfeed = 600 s). Solid lines are theoretical model predictions, and solid points are experimental values.

Figure 4. Purity, recovery, and unit productivity of the second-stage VPSA process with the feed time (feed concentration, 50%). Solid lines are theoretical model predictions, and solid points are experimental values.

adsorption. Figure 4 shows the effect of different adsorption times on the performance of the second-stage VPSA process with the CO2 feed concentration of 50%. With the increase of adsorption time, the concentration wavefront of CO2 at the end of the adsorption step becomes close to the outlet, and the CO2 purity increases. However, more CO2 is lost in the adsorption step, which results in the decrease of recovery. It is noted that the purity of the VPSA cycles for VPSA (2, 4, 5) in this work is slightly lower than the purity obtained in our previous study with the same feed CO2 concentration and pressure. The differences

Figure 5. Simplified scheme of the two-stage VPSA process to capture CO2 from flue gas using AC beads. 5017



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variable to take into account is the feed pressure of secondstage VPSA. As shown in Table 6, simulations carried out at higher feed pressure result in higher recovery and lower purity. The optimal feed pressure should be established from comparison of results at same product purity and overall power consumption. Not only high product purity and recovery but also high productivity and low energy consumption are equally crucial for a VPSA process. By comparison of the simulation results shown in Table 6, it can be observed that, with the vacuum pressure of 10 kPa and feed flow rate of 2.5 SLPM (run 2), the performance of the integrated two-stage VPSA process is as follows: CO2 purity, 95.3%; CO2 recovery, 74.4%; unit productivity, 0.85 mol-CO2/(kg·h); and specific power consumption, 723.6 kJ/kg-CO2. In Figure 7, the mass balance of the twostage VPSA process (run 2) is presented, showing how we have achieved the desired CO2 purity and recovery. It can be observed that the CO2 concentration of the outflow in the second-stage VPSA is quite high (13.2%), resulting in the low recovery. To increase the recovery, the outflow can be recycled to the firststage VPSA, as shown in Figure 4 (the dash line). In run 2 (Table 6), with the recycled flow, the CO2 feed concentration in the first-stage decreases to 14.7%, which is slightly less than 15%. The performance indicators of the two-stage VPSA process for capturing CO2 using AC beads, such as CO2 purity, CO2 recovery, and specific power consumption, were compared with those of other processes for CO2 capture reported previously in the literature and summarized in Table 7. Ishibashi et al.9 studied a two-stage VPSA operation in their pilot plant for CO2 capture from flue gases with zeolite 13X. Under the desorption pressure of 0.5−1.5 bar, a removal efficiency of 90% and 99% purity of CO2 can be achieved, and the power consumption can be reduced to 2.02 MJ/kg-CO2. Cho et al.27 also investigated a two-stage VPSA with two beds each stage with zeolite 13X, the CO2 can be concentrated from 10.5% to 99% with a recovery of 80%, and the experimentally power consumption was 2.31−2.79 MJ/kg-CO2, while the theoretically power consumption was 513 KJ/kg-CO2. It is noted that the CO2 purity and recovery in this work are slightly lower, which is mainly due to the lower CO2 capacity and selectivity of activated beads. The energy consumption varies as a result of the different process configurations and operation conditions employed. The specific power consumption required in this study is in the same magnitude as that in their works. New effort in VPSA cycle design and optimization is still required to improve the power consumption of the whole capture system.

Figure 6. Cycle sequence of the two-stage VPSA process for CO2 capture from flue gas using AC beads.

experiments, as described in Table 1 (0.273 L each column), while the columns of the second-stage VPSA (0.044 L each column) are smaller than that of the first-stage VPSA. Simulations with different vacuum pressures, feed flow rates, and feed pressures of second-stage VPSA were performed to study their effects on the performance of the integrated twostage VPSA process. The process parameters and performance of the two-stage VPSA process for CO2 capture were shown in Table 6. With the increase of total feed flow rate in the firststage VPSA process, both the CO2 purity and unit productivity increase significantly, while the recovery decreased. Decreasing the vacuum pressure (from 10 to 3 kPa) results high product purity, recovery, and unit productivity. However, the total power consumption increases obviously due to the increase of power consumption of the vacuum pump. Another important

Table 6. Simulations of Two-Stage VPSA Processes for CO2/N2 Separation Using AC Beads with Different Operating Parametersa first-stage VPSA

second-stage VPSA

integrated two-stage VPSA

run

Qfeed [SLPM]

Plow [kPa]

purity [%]

recovery [%]

Phigh [kPa]

Plow [kPa]

purity [%]

recovery [%]

purity [%]

recovery [%]

Wtot [kJ/kg-CO2]

productivity [mol/(kg·h)]

1 2 3 4 5 6 7 8 9

1.5 2.5 3.0 4.0 2.5 2.5 2.5 2.5 2.5

10 10 10 10 5 3 10 10 10

32.49 43.86 45.92 47.24 44.51 44.50 43.86 43.86 43.86

99.85 90.94 81.18 63.51 96.21 97.16 90.94 90.94 90.94

350 350 350 350 350 350 200 250 450

10 10 10 10 5 3 10 10 10

85.28 95.29 96.26 96.76 96.34 96.56 95.93 96.00 93.33

88.21 81.78 79.35 77.71 83.90 85.28 57.44 67.43 89.76

85.28 95.29 96.26 96.77 96.34 96.56 95.93 96.00 93.33

88.07 74.36 64.42 49.36 80.72 82.86 52.23 61.32 81.62

840.42 723.56 726.27 759.19 829.28 902.41 670.96 694.28 749.64

0.59 0.85 0.89 0.91 0.93 0.95 0.61 0.71 0.93

a

For all the simulations: yCO2 = 15%; T = 303 K. 5018



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Figure 7. Mass balance of the two-stage VPSA process for CO2 capture from flue gas using AC beads (case 1).

Table 7. Comparison of Performance Employing Different Processes for CO2 Capture from Flue Gas process

adsorbent

two-stage VPSA two-stage VPSA PTSA 323−373 K one-stage PSA two-stage PSA second-stage: 13.34 kPa FVPSA VSA PSA MEA absorption TSA VSA

AC beads AC beads Ca-X type zeolite 99 13X 13X 99 13X 13X 13X 5A 13X

yCO2,feed [%] 15 15 11.5 90 10 10.5 80 15 13 13 13 10 12

CO2 purity [%]

regeneration method 10 kPa 5 kPa 0.05−0.15 atm 2016 kJ/kg-CO2 6.67 kPa first-stage: 6.67 kPa 513.2 KJ/kg-CO2 0.1−0.7 atm 0.05 bar (Pfeed = 1.5 bar) 1 bar (Pfeed = 6 bar) 423 K 5 kPa (Pfeed = 130 kPa)

95.36 96.40 9 50−70 27 88.9 48 48 >99 ≥94 >90

CO2 recovery [%]

power consumptiona

ref

723.19 kJ/kgCO2 831.53 kJ/kgCO2

this study this study

30−90

90−1100 kJ/kgCO2

13

96.9 85 85 90 75−85 >70

150.4 KJ/kgCO2 $51/ton CO2 avoided $57/ton CO2 avoided $49/ton CO2 avoided 6120−6460 kJ/kgCO2 4−10 kW/TPDc

11 28 28 28 29 30

73.62 80.42

a

In the VPSA processes, the power consumption is for compressors and/or vacuum pump; in the TSA process, the power consumption comes from the water vapor heating the adsorbent; for the MEA absorption process, the power consumptions comes from the electricity to heat the boiler.

cost of adsorption could be higher than the MEA absorption process.27 This is caused by the higher energy penalty for compressor due to a lower purity of CO2 in the product stream, also because much lower unit productivity leads to higher adsorbent unit and adsorbent replacement costs.

The unit productivity is directly related the size of the absorber and the amount of the adsorbents. The productivity in this work (0.85 mol-CO2/(kg·h)) is much higher than that of the other VPSA processes using zeolites (such as 0.33 molCO2/(kg·h) from the two-stage VPSA work of Liu et al.32). This is because the lineally isotherms of CO2 on activated carbon and a higher feed pressure of the second-stage VPSA in this work, which lead to much higher CO2 cyclic capacity than that of the VPSA process using zeolites. Also, the parameters of the blowdown process should be noted. Usually, lower vacuum pressure results in higher CO2 purity and recovery, and thus, unit productivity. However, if the vacuum pressure continues to be reduced to some low level, the power consumption could be increase as a result of the significantly increased Ph/Pl ratio. The duration of vacuum time also plays a similar role in determining the process performance. With the same vacuum level reached, longer evacuation time implies smaller vacuum flow rate, which leads to lower pressure drop in the column than short evacuation times. Thus, longer evacuation results in lower average pressure in the column and leads to higher CO2 purity and recovery.33 However, the productivity depends on the dominant effects of the recovery and vacuum duration. It should be noted that the real CO2 capture costs include the capital and operating costs, which include all process equipment and fixed general maintenance costs and labor cost.28 The power consumption of adsorption (2.02− 2.79 MJ/kg-CO2)9,27 is much lower than that of absorption with MEA (4−6 MJ/kg-CO2).31 However, the CO2 capture

5. CONCLUSIONS In this work, a two-stage vacuum pressure swing adsorption (VPSA) process, which uses activated carbon beads as the adsorbent to recover CO2 from flue gas, was investigated through experiments and simulation. At the first-stage VPSA, CO2 is enriched to (40−60) % at almost atmospheric pressure and then further concentrated to above 95% at the second-stage VPSA. Singlecolumn VPSA experiments were performed first to study the effect of different operating conditions and to evaluate the viability of the two-stage VPSA. With the feed concentration of 60%, a purity of 95.5% and recovery of 83.4% can be obtained with the four-step cycle, which indicated that the two-stage VPSA process is viable. On the basis of the experimental results, a two-stage VPSA process composed of two columns for each stage was simulated and the parameters were optimized. The first-stage VPSA unit operates with a four-step Skarstrom cycle, which includes feed pressurization, adsorption, blowdown, and N2 purge. For the second-stage VPSA process, a cycle with feed pressurization, adsorption, pressure equalization, blowdown, and pressure equalization was employed. A CO2 purity of 95.3% was obtained with 5019



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Wt = total specific power consumption (kJ/kg-CO2) P = pressure (Pa) Pr = Prandtl number q = adsorbed phase concentration (mol/kg) qi* = adsorbed phase concentration in equilibrium (mol/kg) ⟨qi̅ ⟩ = average amount adsorbed of component i (mol/kg) Re = Reynolds number rc = crystal radius (m) Rg = universal gas constant (J/(mol·K)) Rp = pellet radius (m) Rw = column radius (m) S = adsorbent specific area (m2/kg) t = time (s) T = temperature (K) Tg = temperature of the gas phase (K) Ts = solid temperature (K) Tw = wall temperature (K) T∝ = environmental temperature (K) u = superficial velocity (m/s) U = global external heat transfer coefficient (J/(mol·K)) yi = molar fraction of component i z = partition of the column length L (m)

74.4% recovery. The total specific power consumption of the two-stage VPSA process is 723.6 kJ/kg-CO2, while the unit productivity is 0.85 mol-CO2/(kg·h). The energy consumption required by the two-stage VPSA process was compared with other processes reported in the previous literature. In future work, the design of a scale-up two-stage VPSA process will be performed with respect to the product purity, recovery, and the energy consumption.

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AUTHOR INFORMATION

Corresponding Author

*Phone: +86 21 64252826. Fax: +86 21 64252826. E-mail: [email protected]; [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors are grateful for the financial support of the China 863 Program (Grant No. 2008AA062302), Shanghai International Cooperation Project (Grant No. 08160704000), and Shanghai Pujiang Talent Program (Grant No. 08PJ14034). The authors also acknowledge Professor Ling Licheng for providing the activated carbon beads.

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Greek Letters

NOMENCLATURE

Latin Letters

a′ = Pellet specific area (m−1) A = first Virial coefficient (m2/mol) B = second Virial coefficient (m4/mol2) Ci = gas phase concentration of component i (mol/m3) ci̅ = averaged concentration in the macropores (mol/m3) C̃ p = molar constant pressure specific heat of the gas mixture (J/(mol·K)) C̃ pw = specific heat of the column wall (J/(mol·K)) C̃ v = molar constant volumetric specfic heat of the gas mixture (J/(mol·K)) C̃ v,ads,i = molar constant volumetric specific heat of component i in the adsorbed phase (J/(mol·K)) CT = total gas concentration (mol/m3) Dax = axial dispersion coefficient (m2/s) Dc = micropore diffusion coefficient (m2/s) Dc0 = limiting diffusion coefficient at infinite temperatures (m2/s) Dij = binary molecular diffusivity (m2/s) Dk = Knudsen diffusivity (m2/s) Dm = molecular diffusion coefficient (m2/s) dp = pellet diameter (m) Dp = pore diffusivity (m2/s) Dw = internal diameter of the column (m) e = wall thickness (m) Ea = activation energy of micropore diffusion (kJ/mol) hf = film heat transfer coefficient between gas and solid phase(W/(m2·K)) hw = film heat transfer coefficient between gas and column wall (W/(m2·K)) kf = film mass transfer coefficient (m/s) KH = Henry constant (mol/(kg· kPa)) K∝ = adsorption constant at infinite temperature (mol/ (kg· kPa)) L = column length (m) wads = mass of adsorbent (kg) MW = molecular weight of the gas (kg/mol)

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αw = ratio of the internal surface area to the volume of the column wall (m−1) αwl = ratio of the logarithmic mean surface area of the column shell to the volume of the column wall (m−1) (−ΔH0) = heat of adsorption at zero coverage (kJ/mol) (−ΔHi) = isosteric heat of adsorption of component i (kJ/mol) λ = axial heat dispersion (W/(m2·K)) εc = column porosity εp = pellet porosity μg = gas viscosity (Pa· s) ρb = bulk density of the column (kg/m3) ρg = density of the gas phase (kg/m3) ρp = density of the adsorbent (kg/m3) ρw = density of the column wall (kg/m3) τp = pore tortuosity

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Two-Stage VPSA Process for CO2 Capture from Flue Gas Using

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