from Flue Gas - American Chemical Society

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Ind. Eng. Chem. Res. 2002, 41, 4122-4131

Numerical Analysis on the Power Consumption of the PSA Process for Recovering CO2 from Flue Gas Jong-Ho Park, Hee-Tae Beum, Jong-Nam Kim, and Soon-Haeng Cho* Korea Institute of Energy Research, 71-2, Jangdong, Yusungku, Taejon, 305-343, Korea

To recover 99% CO2 from flue gas containing 10-15% CO2, a two-stage pressure swing adsorption (PSA) process is usually employed. At the first stage of the two-stage PSA, CO2 is concentrated to 40-60% and then concentrated to 99% at the second stage. Because two stages are coupled with each other, the overall optimization of the two-stage PSA process is quite a complicated task. In this paper, we only considered the first stage of the two-stage PSA process to simplify the analysis. Effects of the process configuration and operating variables such as the P/F ratio and desorption pressure on the specific power consumption were investigated. The specific power consumption at the blower was reduced with the increase of the P/F ratio. On the contrary, that at the vacuum pump was increased with the increase of the P/F ratio. Because of these two effects, there was an optimum P/F ratio, which minimized the specific power consumption. While the compression ratio is reduced with the increase of the desorption pressure, the CO2 purity is decreased and the amount of the power used to compress nitrogen is increased. As a result, the specific power consumption was insensitive to the desorption pressure within the range studied here. Employing the pressure equalization step, the CO2 purity could be increased without much increase of the specific power consumption. With the rinse step, which is often used to increase the purity of the strongly adsorbed component, the CO2 purity could be increased. However, because more gases should be pumped to produce a given amount of CO2,, the specific power consumption was significantly increased at a given recovery. Introduction The emission of CO2 from power plants that burn fossil fuels is the major cause for the accumulation of CO2 in the atmosphere, which causes long-range environmental problems. One option to alleviate the emission of CO2 is to capture CO2 from the emission sources and store it to the ocean or the depleted oil field. CO2 recovery has been achieved by gas absorption employing solutions of carbonates and alkanolamines. However, this process is energy-intensive for the regeneration of solvent and is also plagued by corrosion problems. Recently, the pressure swing adsorption (PSA) process has been considered as an alternative to the absorption process.1 The PSA processes have been widely applied for the removal of CO2 from various feed mixtures. For example, the CO2 in the steam reformer off gas,2 landfill gas, and natural gas3,4 are successfully removed by the PSA processes. A major concern of these examples is purifying a weakly adsorbed component not enriching the strongly adsorbed CO2. For the disposal of CO2 to the ocean or depleted oil field, it is necessary to concentrate the CO2 up to 99% to reduce the compression and transportation cost. When the concentration of CO2 in the feed is high, above 25%, 99% CO2 is easily obtained with relatively high recovery, over 70% using a one-stage PSA.5 However, when the concentration of CO2 is low like the flue gas of the coal power plant, which ranges 10-15%, it is difficult to recover 99% CO2 with high recovery using the one-stage PSA. Thus, a two-stage PSA is employed to enrich CO2 from low concentration CO2.6 At the first stage of the two-stage * Corresponding author. Phone: 82-42-860-3021. Fax: 8242-860-3102. E-mail: [email protected].

PSA, CO2 is concentrated to 40-60% and then concentrated to 99% at the second stage. Besides the two-stage PSA, several other processes were proposed. A dual reflux PSA process was introduced by Diagne et al.7,8 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. However, it was not verified that the power consumption could be reduced. Pugsley et al.9 theoretically investigated the performance of the pressure-temperature swing adsorption (PTSA) process using the circulating fluidized bed. Their results showed that the process could enrich 15% CO2 to 90% with 70% recovery. A piston-driven rapid PSA process using hydrophobic zeolite was studied by Suzuki et al.10 Though the productivity of the RPSA process was over 10 times larger than those of conventional PSA processes, the enrichment of CO2 achieved was no higher than 1.3. The major obstacle of applying the CO2 capture process to the power plants is its high energy consumption. However, little effort was given to optimize the PSA process with respect to the energy consumption. In the two-stage PSA process, the effluent of the second stage is recycled to the first stage. Thus, the overall process optimization of the two-stage PSA process is quite a complicated task because the first and the second stages are coupled with each other. According to the preliminary study, the first stage consumed more energy than the second stage. Thus, as the first step for the optimization of the two-stage PSA process, the first stage of the two-stage PSA process is dealt with in this paper. The effects of the operating variables and cycle configurations on the process performance, espe-

10.1021/ie010716i CCC: $22.00 © 2002 American Chemical Society Published on Web 07/10/2002

Ind. Eng. Chem. Res., Vol. 41, No. 16, 2002 4123

cially on the power consumption, are investigated theoretically. Process Description. Usually, the two-stage PSA is applied for the recovery of CO2 from the flue gas containing 10-15% CO2. At the first stage, CO2 is concentrated to about 40-60% and then concentrated to over 99% at the second stage. As the first step to optimize the process with respect to the power consumption, the effects of the operating variables and cycle configuration of the first stage PSA are investigated. Three different cycles are compared: (I) Skarstrom cycle, which undergoes an adsorption, an evacuation, a purge, and a feed pressurization step; (II) Skarstrom cycle with a pressure equalization step (adsorption, pressure equalization, evacuation, purge, pressure equalization, and feed pressurization steps); and (III) Skarstrom cycle with the rinse step (adsorption, rinse, evacuation, purge, and feed pressurization steps). Figure 1 shows the cycle sequences studied here. Step times are also shown in Figure 1. The evacuation and feed pressurization step times were set arbitrary as 1/7 of the adsorption time. The pressure equalization step is conducted for 1/14 of the adsorption time. The relative lengths of each step were fixed regardless of the cycle time. Mathematical Model The mathematical model adopted is a nonisothermal, nonequilibrium, bulk separation model with a nonlinear multicomponent equilibrium isotherm (the IAS theory). The assumption used to derive the model included the following: ideal gas behavior, negligible axial pressure gradient, no radial gradient, no axial dispersion and no axial heat conduction, and thermal equilibrium between gas phase and adsorbents. Mass transfer is represented by the linear driving force (LDF) approximation. On the basis of these assumptions, the mass balance for each component of the mixture and the total mass balance are written as follows:

∂yi ∂t

+u

∂yi

+

∂z

ji (1 - ) RgT ∂q Fp P ∂t jj (1 - ) RgT ∂q Fp ) 0 (1) yi P ∂t j

∑

component mass balance total mass balance

∂C ∂t

+

∂(uC)

+

∂z

∑i

ji 1 - ∂q Fp )0 ∂t

(2)

Figure 1. Flow directions and cycle sequences of three different processes.

∂Tw ) 2πhwRi(T - Tw) - 2πUwRo(Tw - TF) cpwFwaw ∂t (4) Mass transfer is expressed by the LDF approximation.

Energy balance for the gas-phase includes the heat transfer to the column wall

∑i

∂T ∂t

+ cpguC

(5)

∂T

∂z ∂q j i 2hw (-∆Ha)i(1 - )Fp (T - Tw) ) 0 (3) + ∂t Ri

(cpgC + (1 - )cpsFp)

∂q ji ) ki(qi* - q j i) ∂t

In eq 3, the last term accounts for heat transfer to the column wall. Energy balance around the column wall is given by

The steady-state momentum balance is given by Ergun’s equation 2

-

1- ∂P 150µu (1 - ) ) + 1.75Fgu2 3 2 ∂z dp dp3

(6)

In the PSA process, most of the power is consumed in the vacuum pump and blower. Work required for the

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Ind. Eng. Chem. Res., Vol. 41, No. 16, 2002

purge step

Table 1. Langmuir Isotherm Parameters and Heats of Adsorption

N2 CO2 O2

ai,1 × 103 (mol/g)

ai,2 (K)

bi,o × 107 (1/mmHg)

bi,1 (K)

-∆Ha (cal/mol)

-1.03 -1.95 -0.37

0.61 2.03 0.94

2.81 87.84 19.4

2520 2200 1031

5952 8333 2976

adiabatic compression of gas from a suction pressure, P1, to a discharge pressure, P2, is given by the following equation:

(( )

P2 k Ws ) n˘ RgT1 k - 1 P1

(k-1)

)

/k - 1

(7)

where n˘ is the molar flow rate. If the efficiency of the vacuum pump and blower is 1, then the power consumption of those machines are given by this equation. However, because of the mechanical and electrical losses, the power consumption of the vacuum pump and blower is higher than that given by this equation. Because the efficiencies of the vacuum pump and blower depend on the type of the vacuum pump, system configuration, and manufacturer, it is of little use to calculate the power consumption with a specific efficiency. Therefore, the ideal power consumption given by eq 7 was used to compare the power consumption with the operating conditions. Though the estimation differs from the actual power consumption, it can be used for the comparison of the process performance. Typical flue gas consists of nitrogen, carbon dioxide, oxygen, water vapor, and minor impurities such as SOx, and NOx. It was assumed that water, SOx, and NOx were removed in the pretreatment process and that the feed of the CO2 PSA process was mainly composed of carbon dioxide, nitrogen, and oxygen. Zeolite 13X was the best adsorbent for the separation of carbon dioxide from the above feed.5 The multicomponent adsorption equilibrium of CO2 and N2 was well-represented with the IAS model incorporating the Langmuir isotherm for the pure component adsorption equilibrium.11

qi )

qsibi pi 1 + bi pi

(8)

where the parameters bi and qsi depend on the temperature in the following way:

bi ) bio exp(b1/T)

qsi ) ai,1 + ai,2/T

(9)

The Langmuir isotherm parameters and the heats of adsorption are listed in Table 1. Parameters for O2 were obtained elsewhere.12 In all simulations, it was assumed that the bed is saturated with pure nitrogen. Thereafter, the final conditions of the previous step become the initial conditions for the next step. Boundary conditions for each step are written as follows: adsorption step

yi(0,t) ) yi,F, T(0,t) ) TF, u(0,t) ) uH (10a)

|

∂P ) -K1uH - K2uH2, P(L,t) ) PH (10b) ∂z z)0 2

1- 150µ (1 - ) , K2 ) 1.75Fg K1 ) 3 2 dp dp 3

(10c)

yi(L,t) )yji,AD, T(L,t) ) TF, u(L,t) ) uPU (11a)

|

∂P ) -K1uPU - K2uPU2, ∂z z)L u(0,t) ) uS(1 - e-(P|z)0-Pd)/a), uS ) Q/(A) (11b) EQ-BD step

u(0,t) ) 0

u(L,t) ) CV(P|z)L - PEQ-PR) (12)

evacuation step

u(L,t) ) 0, u(0,t) ) uS(1 - e-(P|z)0-Pd)/a) uS ) Q/(A) (13) EQ-PR step

yi(L,t) ) yi,EQ-BD, T(L,t) ) TF, u(L,t) ) uEQ-BD

PEQ-BD (14) P|z)L

FP step

yi(0,t) ) yi,F, T(0,t) ) TF, u(0,t) ) CV(PH - P|z)0) (15) During the purge step, the velocities at both ends of the adsorber were specified. The gas velocity at the gas outlet depends on the suction capacity of the vacuum pump, which varies with the suction and discharge pressure as usual. But, a certain type of vacuum pump gives almost constant suction capacity above the ultimate vacuum pressure. In this study, the variation of the suction capacity with the suction pressure was modeled with the exponential function as shown in eq 11b, where Q is the suction capacity of vacuum pump at room temperature and atmospheric pressure. Above the ultimate desorption pressure, Pd, uS is nearly constant because the exponential term is negligibly small. Assigning large value for uS in eq 11b, the ultimate desorption pressure, Pd, one of the important operating variables, could be maintained nearly constant regardless of the purge amount. The parameter a in eq 11b was introduced to obtain smooth function. If the suction capacity varied abruptly around the ultimate desorption pressure, the integrator converged very slowly to the solution. By adjusting parameter a, it was possible to accelerate the convergence. It was assumed that the velocity of gas leaving the adsorber during the EQ-BD step, uEQ-BD, was proportional to the pressure difference between tops of two columns. Therefore, for the calculation of the velocity, the information on the pressure history at the top of the column in the EQ-PR step, PEQ-PR, was required. Assumed pressure history was used in the first cycle simulation. Thereafter, it was obtained from the simulation on the EQ-PR step. The adsorber characteristics with the heat and mass transfer coefficients and other constants used in the simulations are listed in Table 2. As the operating variables of the PSA process, the followings are defined:

P/F ratio ) amount of gas used in the purge step/ amount of feed introduced in the adsorption and feed pressurization steps (16)

Ind. Eng. Chem. Res., Vol. 41, No. 16, 2002 4125 Table 2. Parameters Used in the Simulations Physical Properties of Bed and Adsorbents bed inner diameter cm 8.9 bed outer diameter cm 9.3 bed porosity 0.36 particle density g/cm3 1.17 particle diameter cm 0.25 density of wall g/cm3 8.0 heat capacity of gas cal/(mol K) 7.3 heat capacity of adsorbent cal/(g K) 0.22 heat capacity of wall cal/(g K) 0.11 total bed length cm 150 hw Uw kCO2 kN2 kO2

Mass and Heat Transfer Coefficients cal/(cm2 s K) cal/(cm2 s K) s-1 s-1 s-1

Other Constants CV for EQ-BD step cm/(s mmHg) CV for pressurization step cm/(s mmHg) a defined in eq 11b mmHg

6.0 × 10-4 1.0 × 10-4 0.02 0.1 0.1

Table 3. Cumulative Volumes of CO2 and Gas Mixture Flowing into and out of the Column at Each Step of the Skarstrom Cycle when the Bed Utilization Factor Is 0.284 GCO2 (L) feed 94.5 adsorption effluent 21.88 purge gas 0.688 evacuation effluent 19.97 purge effluent 58.07 feed pressurization 4.24 recovery (%) 79.03 bed utilization factor

Gtot (L)

YCO2 (%)

945 889.51 28 44.49 82.76 42.43 P/F ratio 0.284

10 2.459 2.459 44.88 70.16 10 0.0283

power consumption (kWh/Nm3‚CO2) evacuationa

purgea

blowerb

0.123

0.167

0.0358

a

1.6 0.5 2

Based on the amount of CO2 produced in the corresponding step. b Based on the total amount of CO2 produced during the evacuation and purge steps.

recycle ratio ) amount of gas used in the rinse step/ total amount of gas pumped out of the adsorber in the evacuation and purge steps (17) To solve this system of partial differential equations, the spatial derivatives are divided using a backward difference scheme, and the resulting ordinary differential equations are solved with the GEAR method. Results and Discussion Effects of the P/F Ratio. The composition and cumulative volumes of CO2 flowing into and out of the adsorber at each step are listed in Table 3. The bed utilization factor, P/F ratio, and the desorption pressure were 0.28, 0.0283, and 50 mmHg, respectively. The bed utilization factor is defined as the ratio of the amount of CO2 fed during the adsorption and feed pressurization steps to the equilibrium adsorption capacity for CO2. Overall balance of CO2 in a cyclic steady state was satisfied within 0.5% error. The average CO2 concentration of the evacuation effluent is 44.88% and that of the purge effluent is 70.16%. The specific power consumption at the evacuation and purge steps based on the amount of CO2 produced at the corresponding steps are 0.123 and 0.167 kWh/Nm3‚CO2, respectively. That is, more energy is consumed at the purge step to produce a given amount of CO2. At first glance, it looks like the specific power consumption can be minimized when the process is operated without a purge step. However, this is not true because the power consumption at the blower should be accounted for. Figure 2 shows the specific power consumption, CO2 purity, and recovery with the P/F ratio at the bed utilization factor of 0.28. The specific power consumptions shown in Figure 2 are based on the total amount of CO2 desorbed in the evacuation and purge steps. As the P/F ratio increases from zero, the recovery increases monotonically. On the other hand, once the purity reaches a maximum near the P/F ratio of 0.005, it gradually decreases. It is seen in Figure 2 that the specific power consumption at the blower increases with the decrease of the P/F ratio. As shown in eq 7, the power consumption is the function of the molar flow rate, compression ratio, and the inlet gas temperature. Considering that the amount of CO2 produced through the evacuation and purge steps, at the desorption steps,

Figure 2. Recovery, purity, and power consumption of the Skarstrom cycle with the P/F ratio: WB, specific power consumption at the blower; WV, specific power consumption at the vacuum pump; WTOT, overall specific power consumption (feed flow rate, 135 L/min; Pd, 50 mmHg; PH, 860 mmHg; yN2/yCO2/yO2 ) 81.3/10/ 8.7; bed utilization factor, 0.28).

is FyCO2,FR, the specific power consumption at the blower can be represented by the following equation:

WB ) RgT1 f

()

P2 /(yCO2,FR) P1

(18)

where R is the recovery and P2 is the discharge pressure of the blower. The discharge pressure, P2, depends on the pressure drop in the adsorber, which varies with the velocity profile developing in the adsorber. However, if the feed flow rate is constant, the variation of the pressure drop during the adsorption step is negligibly small. In addition, the feed temperature T1 is also constant in these simulations. Thus, the specific power consumption at the blower is mainly dependent on the process recovery at a given feed composition and adsorption pressure. Because the recovery increases with the P/F ratio, the specific power consumption at the blower decreases with the increase of the P/F ratio, as shown in Figure 2. Contrary to the specific power consumption at the blower, the specific power consumption at the vacuum

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pump increases with the P/F ratio. Considering that the total amount of gas pumped during the desorption steps is FyCO2,FR/yjCO2, the specific power consumption at the vacuum pump can be represented by the following equation:

WV ) RgT1 f

() P2

P h1

/yjCO2

(19)

The discharge pressure P2 is constant, atmospheric pressure. P h 1 is a pressure between the adsorption pressure, PH, and the desorption pressure, PD, and it approaches the desorption pressure as the purge amount increases. yjCO2 in eq 19 is the average CO2 composition of desorption effluents, the evacuation and purge effluents. It is clear from eq 19 that the average CO2 composition of desorption effluents plays a more important role in the specific power consumption than the recovery. During the evacuation and purge steps, not only CO2 but also N2 and O2 are desorbed and compressed by the vacuum pump. As the amounts of N2 and O2 in the desorption effluents increase, more and more energy is used to compress N2/O2. This is why the average composition of CO2 is important. The average CO2 composition of the desorption effluents is reduced above the P/F ratio of 0.005, which plays a role to increase the specific power consumption. Moreover, with the increase of the P/F ratio, the pressure, P h 1, approaches the desorption pressure because more and more gas is obtained at lower pressure. This increases the compression ratio. Because of a higher compression ratio and lower CO2 purity, the specific power consumption at the vacuum pump increases with the increase of the P/F ratio. The situation is somewhat complicated below the P/F ratio of 0.005. Both the CO2 purity and the compression ratio are increased with the increase of the P/F ratio. However, because the increase of the compression ratio is more significant than the increase of the purity, the specific power consumption at the vacuum pump below the P/F ratio of 0.005 is also increased with the increase of the P/F ratio. Because of contradicting effects of the P/F ratio on the specific power consumption at the vacuum pump and blower, there is an optimum P/F ratio that minimizes the specific power consumption. It is seen in Figure 2 that the optimum purge amount appears about the P/F ratio of 0.005. If the pressure drop is increased or feed composition of CO2 is reduced, then the specific power consumption at the blower is increased. In those cases, the optimum P/F ratio will appear at a higher value. Effects of Bed Utilization. Not only the P/F ratio but also the amount of feed introduced at the adsorption and the feed pressurization steps, F, is one of the important operating variables. The bed utilization factor defined previously is directly related to the amount of feed, F. Here, the effects of the bed utilization factor on the specific power consumption are investigated. Two different ways can be used to change the bed utilization factor. One is to increase the feed flow rate at a fixed cycle time and the other is to increase the cycle time at a fixed feed flow rate. The latter method was employed in this study. If the bed utilization is reduced, the concentration wave front of CO2 at the end of the adsorption step becomes close to the feed end and the CO2 purity of the evacuation effluent is decreased. The composition and

Table 4. Cumulative Volumes of CO2 and Gas Mixture Flowing into and out of the Column at Each Step of the Skarstrom Cycle when the Bed Utilization Factor Is 0.194 GCO2 (L) feed 63 adsorption effluent 13.13 purge gas 0.415 evacuation effluent 14.26 purge effluent 40.66 feed pressurization 4.34 recovery (%) 81.57 bed utilization factor

Gtot (L)

YCO2 (%)

630 591.33 18.67 42.34 59.55 43.4 P/F

10 2.22 2.22 33.69 68.29 10 0.0277 0.194


purgea

blowerb

0.127

0.166

0.0339

a


Figure 3. Relations between power consumption, purity, and recovery of Skarstrom cycle at different bed utilization factors. Arrows indicate the direction of the increase of P/F ratio (feed flow rate, 135 L/min; Pd, 50 mmHg; PH, 860 mmHg; yN2/yCO2/yO2 ) 81.3/ 10/8.7).

cumulative volumes of CO2 flowing into and out of the adsorber at the bed utilization factor of 0.19 are listed in Table 4. It is seen in Table 4 that the CO2 purity of the evacuation effluent is decreased about by 10% as compared to that at the bed utilization factor of 0.28 (see Table 3 for the comparison). The total amount of gas desorbed during the desorption steps, evacuation and purge steps, at the bed utilization factor of 0.288 is 127 L and 2/3 of which is obtained at the purge step. However, the total amount of gas desorbed during the desorption steps (evacuation and purge steps) at the bed utilization factor of 0.19 is 101 L and 1/2 of which is obtained at the purge step. This means that the average compression ratio of the bed utilization factor of 0.28 is higher than that at the bed utilization factor of 0.19. Figure 3 shows the purity and specific power consumption with the recovery at different bed utilization factors. All of the curves were obtained by varying the P/F ratio maintaining the bed utilization factor nearly constant. In fact, the bed utilization factor changed slightly with the P/F ratio because the amount of feed required to pressurize the adsorber changed slightly


with the P/F ratio. For example, the bed utilization factors of the curve denoted as 0.28 varied from 0.286 to 0.280 with the P/F ratio. When the bed utilization factors are 0.28 and 0.19, there are optimum P/F ratios which maximize the CO2 purity. However, if the bed utilization factor is further reduced to 0.1, then the CO2 purity is highest when the cycle is operated without the purge step. Similar behavior is observed on the specific power consumption. That is, there are optimum P/F ratios, which minimize the specific power consumption when the bed utilization factor are 0.28 and 0.19. However, when the bed utilization factor is 0.1, the specific power consumption is lowest when the cycle is operated without the purge step. It is seen in Figure 3 that the CO2 purity increases with the increase of the bed utilization factor at all recoveries. The enhancement is more significant when the bed utilization factor is increased from 0.1 to 0.19 than when the bed utilization factor is increased from 0.19 to 0.28. The specific power consumption at a given recovery does not increase so much when the bed utilization factor is increased from 0.1 to 0.19. As mentioned previously, if the recovery is fixed, then the specific power consumption at the blower is almost the same, regardless of the operating conditions. On the other hand, in the vacuum pump, the CO2 purity and compression are more important than the recovery. The CO2 purity increases with the increase of the bed utilization, which plays a role to reduce the specific power consumption. However, another factor, the compression ratio, increases with the increase of the bed utilization factor because the fraction of the CO2 obtained at the purge step increases, as mentioned previously. This plays a role to increase the specific power consumption. When the bed utilization factor is increased from 0.1 to 0.19, the relative increase of the CO2 purity is enough to compensate for the increase of the compression ratio. Therefore, the specific power consumption remains almost constant despite of the increase of the compression ratio. On the other hand, when the bed utilization factor is increased from 0.19 to 0.28, the relative increase of the CO2 purity is small as compared to the case that the bed utilization factor is increase from 0.1 to 0.19. So, the increase of compression ratio outweighs the enrichment achieved with high bed utilization. As a result, the specific power consumption increases slightly when the bed utilization factor is increased from 0.19 to 0.28. Effects of the Desorption Pressure. The specific power consumption at the vacuum pump is influenced by the compression ratio and the CO2 purity of desorption effluents. That is, the lower the compression ratio and the higher the CO2 purity, the lower the specific power consumption. While the compression ratio is reduced with the increase of the desorption pressure, the CO2 purity of the desorption effluents is reduced. Therefore, the specific power consumption at the vacuum pump can be increased or decreased depending on the dominant effect. Figure 4 shows the purity and specific power consumption with the desorption pressure. The curves were obtained varying the P/F ratio at fixed bed utilization factor of 0.28. It is clearly seen in Figure 4 that the purity increases with the increase of the compression ratio. The specific power consumptions at two desorption pressures, 50 and 80 mmHg, are almost the same. This

Figure 4. Relations between power consumption, purity, and recovery of Skarstrom cycle at different desorption pressure. Arrows indicate the direction of the increase of P/F ratio (feed flow rate, 135 L/min; PH, 860 mmHg; yN2/yCO2/yO2 ) 81.3/10/8.7; bed utilization factor, 0.28; filled symbols, CO2 purity; empty symbols, power consumption).

is because the increase of compression ratio is compensated by the increase of the CO2 purity. However, if the desorption pressure is further decreased to 30 mmHg, the relative increase of the CO2 purity is not enough to compensate the increase of the compression ratio. Thus, the specific power consumption at the desorption pressure of 30 mmHg is higher than those at other desorption pressures, especially at lower recovery where the relative increase of the CO2 purity is smaller as compared to that at higher recovery. Effects of Pressure Equalization Step. During the pressure equalization step, the weakly adsorbed component is preferentially desorbed and the strongly adsorbed component is concentrated in the adsorber. Therefore, employing the pressure equalization step, one can increase the purity of the strongly adsorbed component of the desorption effluents. On the other hand, because the desorption starts at a lower pressure as compared to the Skarstrom cycle, a higher compression ratio is required for the desorption of given amount of the gas mixture. Because of these two contradicting effects, the specific power consumption could be increased or reduced depending on dominant effect. A simulation result for the cycle employing pressure equalization step is represented in Table 5, where CO2 composition and cumulative volume of carbon dioxide flowing into or out of the adsorber are listed. The bed utilization factor and P/F ratio of this simulation were 0.28 and 0.029, respectively. The pressure at the end of the pressure equalization step was 300 mmHg in this simulation. The cumulative volume of gases vented during the EQ-BD step is 16.5 L, which is exactly balanced with the volume of the gases introduced to the adsorber during the EQ-PR step. During the EQ-BD step, the weakly adsorbed component is preferentially vented. As a result, the average concentration of the CO2 produced at the evacuation step is increased by 24% (see Table 3 for the comparison). The total amount of gas pumped at the evacuation step is about 60% of that pumped at the Skarstrom cycle. However, the specific power consumption of the evacuation step based on the amount of CO2 produced at the evacuation step are

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Table 5. Cumulative Volumes of CO2 and Gas Mixture Flowing into and out of the Column at Each Step of the Skarstrom Cycle with the Pressure Equalization Step GCO2 (L) feed 94.5 EQ-BD 1.09 purge 0.68 evacuation effluent 18.47 purge effluent 58.16 EQ-PR 1.09 feed pressurization 2.59 recovery 78.93 bed utilization factor

Gtot (L)

YCO2 (%)

945 16.5 28 26.77 82.56 16.53 25.89 P/F

10 6.61 2.42 68.97 70.44 6.61 10 0.0288 0.280


purgea

blowerb

0.123

0.163

0.036 464 4

a


Figure 5. Effects of the pressure equalization step on the process performance. Arrows indicate the direction of the increase of P/F ratio (feed flow rate, 135 L/min; Pd, 50 mmHg; PH, 860 mmHg; yN2/yCO2/yO2 ) 81.3/10/8.7; bed utilization factor, 0.28; filled symbols, CO2 purity; empty symbols, power consumption).

similar in both cycles. This is because higher compression ratio is required in the cycle, including the pressure equalization step where the desorption starts at 300 mmHg. The carbon dioxide vented in the EQ-BD step is transferred to the other adsorber in the corresponding EQ-PR step, and then the carbon dioxide is vented to atmosphere at the following adsorption step. If the amount of carbon dioxide vented during the pressure equalization step is small, then the recovery is not seriously affected by employing the pressure equalization step. According to the simulation result, the amount of carbon dioxide vented during the pressure equalization step is about 1 L, and the recovery is lower by 2.5% as compared to the Skarstrom cycle. For the rather complete comparison, the performance of the cycle including the pressure equalization step was investigated varying the P/F ratio at the fixed bed utilization factor of 0.28. The results are compared to the Skarstrom cycle in Figure 5. It is clear that the CO2 purity of the cycle including the pressure equalization step at a given recovery is always higher than that of the Skarstrom cycle. While an optimum P/F ratio, which

maximize the CO2 purity, is clearly seen in the Skarstrom cycle, that is unclear in the cycle with pressure equalization step. The specific power consumption of the cycle including the pressure equalization step is slightly larger than that of the Skarstrom cycle, especially at the recovery of 60%. As shown in Tables 3 and 5, the relative increase of the CO2 purity obtained at the evacuation step is about 53%. Thus, at the evacuation step, a higher compression ratio of the cycle including the pressure equalization step is compensated by the relative increase of the CO2 purity. As a result, the specific power consumptions at the evacuation and purge steps of both cycles are almost the same, as shown in Tables 3 and 5. However, the fraction of CO2 produced at the purge step of the cycle, including the pressure equalization step, is higher than that of the Skarstrom cycle. That is, the fraction of CO2 produced with high specific power consumption is increased including the pressure equalization. Thus, the specific power consumption increases slightly if the pressure equalization step is included. Effects of Rinse Step. For the enrichment of a strongly adsorbed component, the product rinse step is often employed.13,14 That is, some of the product is recycled to the adsorber, which just finished the adsorption step. In this section, the effect of the product rinse step on the power consumption is investigated. From the analysis on the rinse step, the effects of the gas recycled from the second stage to the first stage can be understood. In the two-stage PSA process, the effluent of the second stage, which contains more CO2 than the feed, is recycled to the first stage. The situation is similar to the product rinse in the point that gas containing more CO2 than feed is introduced to the first stage adsorber. During the rinse step, the weakly adsorbed component in the bed void and the adsorbent is replaced with the strongly adsorbed component. Thus, employing the rinse step, the CO2 purity in the subsequent desorption steps is enhanced substantially. However, because some portion of the desorption effluents is recycled, much more amount of gas should be pumped for the production of a given amount of CO2. Therefore, the specific power consumption could be increased or decreased according to the dominant effect between the enhancement of the CO2 purity and the increase of pumping volume. The specific power consumption of the cycle including the rinse step could be represented by the following equation:

WV ) RgT1 f

() P2

P h1

/(yjCO2(1 - rr))

(20)

where rr is the recycle ratio, ratio of the amount of gas used in the rinse step to the total amount of gas pumped out during the evacuation and purge steps. Roughly speaking, the specific power consumption of the cycle including the rinse step can be reduced if yjCO2(1 - rr) is increased. That is, if the relative increase of the CO2 purity is large enough to overcome the increase of the volume to be pumped, then the specific power consumption could be reduced. The specific power consumption and purity of the cycles employing the rinse step are compared to the Skarstrom cycle in Figure 6. The bed utilization factors of both simulations were 0.28. The recycle ratio was 0.5.


Figure 6. Effects of the rinse step on the process performance: WB, specific power consumption at the blower; WV, specific power consumption at the vacuum pump. Arrows indicate the direction of the increase of P/F ratio (feed flow rate, 135 L/min; Pd, 50 mmHg; PH, 860 mmHg; yN2/yCO2/yO2 ) 81.3/10/8.7; bed utilization factor, 0.28; recycle ratio, 0.5).

Figure 7. Effects of the rinse step on the process performance when the feed contains 5% CO2: WV, specific power consumption at the vacuum pump; WTOT, overall specific power consumption. Arrows indicate the direction of the increase of P/F ratio (feed flow rate, 135 L/min; Pd, 50 mmHg; PH, 860 mmHg; yN2/yCO2/yO2 ) 86.3/ 5/8.7; bed utilization factor, 0.20; recycle ratio, 0.5).

The definition of the bed utilization factor in the cycle including the rinse step is the same as in the Skarstrom cycle. It is seen in Figure 6 that the purity at a given recovery is increased employing the rinse step, as expected. However, the relative increase of the purity is about only 20% at most. The quantity is insufficient to compensate for the increase of the pumping volume. As a result, the specific power consumption of the cycle with rinse step is higher than that of the Skarstrom cycle. The relative increase of the specific power consumption at the vacuum pump is about 45% at most. When the values, yj CO2(1 - rr), of two cycles are compared, about 65% increase of the specific power consumption is expected. The difference is attributed to the different compression ratios of two cycles. Employing the rinse step, the amount of CO2 desorbed at the evacuation step is increased. Therefore, the average compression ratio becomes low, which plays a role to decrease the specific power consumption. These results imply that the specific power consumption of the first stage increases as the amount of gas recycled from the second stage to the first stage increases. If the feed composition of CO2 becomes low, then the CO2 purity of the desorption effluents of the Skarstrom cycle also becomes low. In this case, the relative increase of CO2 purity caused by the rinse step might be significant so that the rinse step may play a role to increase the product purity with a relatively small increase of the specific power consumption at the vacuum pump. Figure 7 compares the specific power consumption and purity of the Skarstrom cycle and the Skarstrom cycle with the rinse step at the feed CO2 concentration of 5%. The relative increase of the CO2 purity and the specific power consumption at the vacuum pump are about 50% and 36%, respectively. Compared to the previous results obtained at the feed CO2 concentration of 10%, the relative increase of the specific power consumption at the vacuum pump is small and the relative increase of the purity is large. It should be noted that the optimum P/F ratio appears at a higher recovery. That is attributed to the increase of

the power consumption at the blower. As shown in eq 18, the specific power consumption at the blower increases with the decrease of the feed composition of CO2. Thus, to reduce the power consumption at the blower, the recovery should be increased. It is noted that the specific power consumption is about doubled as the feed CO2 concentration is reduced by 50%. Conclusions To recover 99% CO2 from flue gas containing 10-15% CO2, a two-stage PSA process is usually employed. Because two stages are coupled with each other, the overall optimization of the two-stage PSA process is quite a complicated task. In this paper, the specific power consumption of the first stage with the process configuration and operating variables was investigated theoretically. In the PSA processes, the blower and vacuum pump consume most of the energy. The specific power consumption at the blower was proportional to the inverse of the recovery. That is, the higher the recovery, the lower the specific power consumption at the blower. On the other hand, the specific power consumption at the vacuum pump was not dependent on the recovery but on the compression ratio and the CO2 purity. The lower the compression ratio, and the higher the CO2 purity, the lower the specific power consumption at the vacuum pump. With the increase of the P/F ratio, the recovery was increased so that the specific power consumption at the blower was reduced. However, the specific power consumption at the vacuum pump was increased with the increase of the P/F ratio due to the increase of the compression ratio and the decrease of the CO2 purity. As a result, there was an optimum P/F ratio minimizing the specific power consumption of the CO2 PSA process. According to the simulation results, the optimum P/F ratio appeared at a value giving recovery of 60%. If the power consumption at the blower becomes high due to the increase of the pressure drop, then the specific power consumption will be minimized at a P/F ratio giving higher recovery.

4130


With the increase of the bed utilization, the CO2 purity was increased. However, the specific power consumption at a given recovery was insensitive to the bed utilization within the range studied here. A bed utilization factor of 0.28 gave the highest product purity. If the desorption pressure is increased, the compression ratio is reduced. However, because the CO2 purity is reduced with the increase of the desorption pressure, the specific power consumptions of the desorption pressure 80 and 50 mmHg at a given recovery were almost the same, regardless of the desorption pressure. According to the simulation results, the desorption presure of 50 mmHg consumed less energy than the desorption pressure of 30 mmHg, especially at a lower recovery. On the other hand, the purity of CO2 obtained at 30 mmHg was higher than that obtained at 50 mmHg. If high purity of CO2 is introduced to the second stage, the power consumption at the second stage can be reduced. Therefore, the desorption pressure of the first stage should be determined after a complete analysis on the two-stage PSA process is performed. Another way to increase the CO2 purity without much increase of the specific power consumption was to employ the pressure equalization step. During the pressure equalization step, the weakly adsorbed component is preferentially desorbed so that the purity of the strongly adsorbed component of the desorption effluents is increased. However, because the desorption starts at a lower pressure, the compression ratio for the production of the given amount of CO2 is increased employing the pressure equalization step. As a result, despite the increase of the CO2 purity, the specific power consumption of the Skarstrom cycle with the pressure equalization step at a given recovery was almost the same as that of the Skarstrom cycle. Therefore, it is clear that the Skarstrom cycle with the presssure equalization step is better than the Skarstrom cycle. The rinse step, which is often used to increase the purity of the strongly adsorbed component, was another way to increase the CO2 purity. However, it increased the specific power consumption at a given recovery. The increase of the specific power consumption at the first stage can be compensated by the reduction of the specific power consumption at the second stage if high purity of CO2 is fed to the second stage. However, it is clear that the power consumption at the first stage is increased by the rinse step. This means that the amount of CO2 recycled from the second stage to the first stage should be minimized to reduce the power consumption at the first stage. Acknowledgment The research was financially supported by the Ministry of Science and Technology (MOST), Korea. Nomenclature A ) cross-sectional area of column, cm2 a ) parameter defined in eq 11b, mmHg aw ) cross-sectional area of column wall, cm2 bi ) Langmuir isotherm parameter, 1/mmHg C ) gas-phase concentration, mol/cm3 cpg ) heat capacity of gas phase, cal/(mol K) cps ) heat capacity of particle, cal/(g K) cpw ) heat capacity of column wall, cal/(g K) CV ) valve coefficient

dp ) diameter of particles, cm hw ) heat transfer coefficient at inner column wall, cal/ (cm2 s K) (-∆Ha)i ) heat of adsorption of ith component, cal/mol k ) ratio of specific heats, Cp/Cv ki ) mass transfer coefficient of ith component, 1/s K1 ) parameter defined in eq 10 K2 ) parameter defined in eq 10 P ) pressure, mmHg Pd ) the lowest desorption pressure, mmHg qi* ) equilibrium amount adsorbed, mol/g q j i ) amount adsorbed of ith component, mol/g qsi ) saturation amount adsorbed of ith component, mol/g Ri,Ro ) inner and outer radius of column, cm Rg ) gas constant t ) time, s T ) gas-phase temperature, K TF ) feed temperature, K Tw ) wall temperature, K u ) interstitial velocity, cm/s uH ) interstitial velocity of feed, cm/s uPU ) purge gas velocity, cm/s Uw ) wall heat transfer coefficient at outer surface, cal/ (cm2 s K) yi ) mole fraction of ith component in the gas phase yi,AD ) composition of ith component in the adsorption effluent yi,EQ-BD ) composition of ith component in the effluent of the EQ-BD step yi,F ) composition of ith component in the feed z ) the distance along the length of the column, cm Greek Symbols ) bed void fraction µ ) gas viscosity, cP Fg ) gas density, g/cm3 Fp ) density of particles, g/cm3 Fw ) density of wall, g/cm3

Literature Cited (1) Kikkinides, E. S.; Yang, R. T.; Cho, S. H. Concentration and recovery of CO2 from flue gas by pressure swing adsorption. Ind. Eng. Chem. Res. 1993, 32, 2714-2720. (2) Sircar, S.; Kratz, W. C. Simultaneous production of hydrogen and carbon dioxide from steam reformer off-gas by pressure swing adsorption. Sep. Sci. Technol. 1988, 23, 2397-2415. (3) Sircar, S. Separation of methane and carbon dioxide gas mixtures by pressure swing adsorption. Sep. Sci. Technol. 1988, 23, 519-529. (4) Kapoor, A.; Yang, R. T. Kinetic separation of methanecarbon dioxide mixture by adsorption on molecular sieve carbon. Chem. Eng. Sci. 1989, 44, 1723-1733. (5) Chue, K. T.; Kim, J. N.; Cho, S. H.; Yang, R. T. Comparison of activated carbon and zeolite 13X for CO2 recovery from flue gas by pressure swing adsorption. Ind. Eng. Chem. Res. 1995, 34, 591598. (6) Ishibashi, M.; Ota, H.; Akutsu, N.; Umeda, S.; Tajika, M.; Izumi, J.; Yasutake, A.; Kabata, T.; Kageyama, Y. Technology for removing carbon dioxide from power plant flue gas by physical adsorption method. Energy Convers. Manage. 1996, 37, 929-933. (7) Diagne, D.; Goto, M.; Hirose, T. New PSA process with intermediate feed inlet position operated with dual refluxes: Application to carbon dioxide removal and concentration. J. Chem. Eng. Jpn. 1994, 27 (1), 85-89. (8) Diagne, D.; Goto, M.; Hirose, T. Parametric studies on CO2 separation and recovery by dual reflux PSA process consisting of both rectifying and stripping sections. Ind. Eng. Chem. Res. 1995, 34, 3083-3089.

Ind. Eng. Chem. Res., Vol. 41, No. 16, 2002 4131 (9) Pugsley, T. S.; Berruti, F.; Chakma, A. Computer simulation of a novel circulating fluidized bed pressure-temperature swing adsorber for recovering carbon dioxide from flue gases. Chem. Eng. Sci. 1994, 49, 4465-4481. (10) Suzuki, T.; Sakoda, A.; Suzuki, M.; Izumi, J. Recovery of carbon dioxide from stack gas by piston-driven ultra-rapid PSA. J. Chem. Eng. Jpn. 1997, 30, 1026-1033. (11) Kim, J. N.; Chue, K. T.; Kim, K. I.; Cho, S.-H.; Kim, J. D. Nonisothermal adsorption of nitrogen-carbon dioxide mixture in a fixed bed of zeolite-X. J. Chem. Eng. Jpn. 1994, 27, 45-51. (12) Kim, J. N. A study on pressure swing adsorption process for the production of more strongly adsorbed component. Ph.D. Dissertation, KAIST, Korea, 1994.

(13) Cen, P. L.; Yang, R. T. Separation of binary gas mixture into two high-purity products by a new pressure swing adsorption cycle. Sep. Sci. Technol. 1986, 21, 845-864. (14) Baksh, M. S. A.; Kapoor, A.; Yang, R. T. A new composite sorbent for methane-nitrogen separation by adsorption. Sep. Sci. Technol. 1990, 25, 845-868.

Received for review August 30, 2001 Revised manuscript received April 24, 2002 Accepted May 1, 2002 IE010716I

from Flue Gas - American Chemical Society

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