Carbon Dioxide Absorption and Desorption in Aqueous

Ind. Eng. Chem. Res. 2007, 46, 2823-2833

2823

SEPARATIONS Carbon Dioxide Absorption and Desorption in Aqueous Monoethanolamine Solutions in a Rotating Packed Bed Majeed S. Jassim,*,† Gary Rochelle,‡ Dag Eimer,§ and Colin Ramshaw| Department of Chemical Engineering, UniVersity of Bahrain, P.O. Box 32038, Bahrain, Department of Chemical Engineering, The UniVersity of Texas at Austin, Austin, Texas 78712, Hydro Oil & Energy, Research Centre, Porsgrunn, Norway, and Process Intensification and InnoVation Center, School of Chemical Engineering and AdVanced Materials, UniVersity of Newcastle upon Tyne, Newcastle upon Tyne, NE1 7RU, United Kingdom

The absorption and desorption of carbon dioxide in aqueous monoethanolamine (MEA) was measured in a rotating packed bed of size 398 mm outside diameter, 156 mm inside diameter, and axial depth 25 mm. The effect of lean amine temperature (20 and 40 °C), peripheral rotor gravity (31 and 87 g), and various MEA concentrations were investigated. Using MEA concentrations above 30 wt % achieved lower CO2 penetration levels. This is particularly pronounced for the 100% MEA solution. Comparison with conventional columns showed the advantages of using rotating packed beds in terms of saving size and space and efficient operation. 1. Introduction Imperial Chemical Industries (ICI) patented the HIGEE (high gravity) or the rotating packed bed (RPB) concept in late 1970s.1,2 The HIGEE is one of the process intensification (PI) technologies that promotes size and weight reduction, enhances inherent safety with lower inventories, improves energy consumption, lowers capital cost, and addresses environmental concerns.3 This technology takes advantage of centrifugal fields as stimulants for process intensification.4 The dynamic behavior of multiphase fluids is dictated by the interphase buoyancy factor ∆Fg. Therefore, increasing the centrifugal acceleration improves the slip velocity, which in turn improves the flooding characteristics and interfacial shear stress, and consequently boosts the mass transfer coefficient. CO2 capture and sequestration is one approach for reducing the emissions that cause global climate change and the burden of the carbon tax levy. Absorption/stripping is considered the state-of-the-art technology for CO2 removal. Aqueous alkanolamine solution absorbs acid gas components from flue gas streams in a countercurrent operation. The alkanolamine is then regenerated in a reboiled stripper. An extensive literature for the reactive mass transfer system of CO2-MEA-H2O is available.5-12 Chambers and Wall13 designed a mild steel centrifugal absorber with intermeshing concentric rings and no packing to remove 10-15% CO2 from air using pure monoethanolamine (MEA) solution. The corrosive nature of pure MEA caused deterioration of the rich solution, and their results were not expressed in terms of overall gas mass transfer coefficient (KGa) * To whom correspondence should be addressed. Tel.: (+973) 17 876189. E-mail address: [email protected]. † University of Bahrain. ‡ The University of Texas at Austin. § Hydro Oil & Energy. | University of Newcastle upon Tyne.

Figure 1. Flowsheet of experimental facility.

because it was claimed that Henry’s Law is not applicable for very short contact time in the centrifugal absorber. Bucklin et al.14 investigated the application of a rotating packed bed in selective H2S removal with MDEA. The loading of acid gas in rich solution was unexpectedly high but there was 25% error in the dry chemical analysis and that influenced the calculations of mass balances. The capacity of the rotating packed bed was fully exploited as runs with higher circulation rates and packing thickness were not possible due to flooding of the rotor. Lin et al.15 investigated the effects of operating parameters on the overall mass transfer coefficient (KGa) using low MEA concentrations of 6.1 and 12.2 wt % in a rotating packed bed with stainless steel wire mesh packing, 0.96 porosity, and 803 m2/ m3 specific surface area. The MEA solvent achieved the highest KGa values in comparison to those for the sterically hindered amine 2-amino-2 methyl-1-proanol (AMP) and NaOH. A comprehensive review and appraisal of HIGEE technology has recently been published by Rao et al.16 The objective of our work was to test the performance of a pilot-plant scale rotating packed bed in absorption and desorption of carbon dioxide using MEA solutions. The resistances to mass transfer in both modes were identified. The effect of rotor speed, lean amine temperature, and amine strength were investigated.

10.1021/ie051104r CCC: $37.00 © 2007 American Chemical Society Published on Web 04/03/2007

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Ind. Eng. Chem. Res., Vol. 46, No. 9, 2007

Figure 2. Cross-sectional view of the Higee rig.

The rotating packed bed was compared to a conventional absorber and a simulated stripper. 2. Experimental Section The flowsheet of the experimental facility is shown in Figure 1. The MEA solution was stored under a nitrogen blanket in a polypropylene feed tank (420 L). A valve was used to control both the temperature and the homogeneity of the aqueous solution prior to being routed to the experimental rig. The irrigation rate was manually controlled by two polypropylene diaphragm valves (George Fischer, Type 315) and the solution was pumped via a flameproof Polyvinylidene Fluoride (PVDF) magnetic drive centrifugal pump. The rig was located inside a flameproof enclosure, which was equipped with a ventilation system. The gas phase could either be a CO2/air mixture or steam depending on the mode of the operation. The rotor was entirely made from stainless steel whereas the piping and the storage tanks were made from polypropylene. Figure 2 shows the cross-sectional view of the rotor. The orientation of the rotor axis is horizontal. The MEA solution was released via a four-arm distributor onto the inner surface of the packing in an outward direction. A turbulent CO2/air mixture entered via a rotary union (Deublin, part number 450191-029) positioned at the end of the horizontal shaft and then passed into the gas plenum chamber before entering 72 holes drilled in the stainless steel disc located on the outer side of the packed bed. It then flowed inwardly within the packing in a countercurrent direction to the flow of the MEA solution. In the desorption runs, all the components of the absorption experiments remained the same with the exception of the rotary

union (Deublin, part number 525-086-026) which was replaced in order to safely handle the saturated steam. The packed bed was sandwiched between the stainless steel disc and a perspex disc. The dimensions of the packed bed were 398 mm outside diameter, 156 mm inside diameter, and an axial depth of 25 mm. The overall gas flow area passed into the packing was 0.031 m2. The perspex was used as a transparent front cover in order to aid the visual observation of the packed bed during operation. There were eight equispaced cut-outs of 3 mm deep in the perspex disc to allow liquid an exit path. The packing rotor was made from expanded stainless steel small mesh (707S), or expamet packing, and was cut into a doughnut shape. The characteristics of the packed bed were the following: density 1877.8 kg/m3, porosity 0.76, total surface area per unit volume 2132 m2/m3, and the volume of the packing 2.84‚10-3 m3. Detailed calculations of these characteristics can be found in the work of Jassim.17 The aqueous MEA solution irrigated the rotating packed bed using a stationary stainless steel four-arm distributor. There were eight holes drilled in each arm with a diameter of 2 mm parallel. The calculated pressure drop across the distributor was 68.24 kPa based upon an amine flowrate of 50 L/min and taking into account vena contracta of flow (detailed calculations are shown in the Appendix). Two types of vortices were created inside the rotating packed bed. A forced vortex was formed because the gas was forced to rotate like a solid body due to its interaction with the packing. The gas then attempts to maintain its angular momentum as it reaches the “eye” of the rotor, and consequently, it speeds up in a free vortex. Therefore, a low-pressure region was created in the eye of the rotor and the four-arm liquid distributor broke off both the free and the forced vortices and subsequently reduced any effects of windage in the liquid distribution area. The distributor irrigated the packed bed in a more uniform way as the liquid jets traveled at an average injection velocity of 8.24 m/s (Appendix 1) using 32 injection points and the spacing between the 4 prongs and the inner packing surface was kept adequately close. These factors ensured minimum deflection of liquid trajectories by central vortices and lower channeling of liquid solution over the rotated packed bed in comparison to the fixed bed. Hassan-beck18 demonstrated that the four-arm distributor achieves lower height of transfer unit (HTU) values compared to the single pipe distributor. A dry steam system ensured a maximum delivery of 300 kg/h dry saturated steam to the rig. A CO2/air system delivered a steady, turbulent, and homogeneous sour gas mixture for the duration of the experiment (15 min). A sequence of absorption experiments with a typical industrial concentration of 30 wt % MEA solution was initially completed.

Table 1. Comparison of GCs Used by Jou et al.19 and the Current Investigation Jou et al.11

current investigation

column

1.63 m × 3.175 mm stainless steel packed column of chromosorb 104

detector

TCD

column 1: 25 m × 0.53 mm fused silica packed column of CP wax for amines. Column 2: 1 m × 0.32 mm silcosteel packed column of Carbosieve S-II. catalytic converter prior to FID

injection port column detector carrier gas, flowrate injection vol (µL)

300 250 200 He, 22 cm3/min 5

CO2/H2O/MEA

0.44/1.04/8.20

Oven Temperatures (°C) 375 100-150 250 He, 10.8 cm3/min 0.2 Retention Time (min) 4.28/-/7.02

Ind. Eng. Chem. Res., Vol. 46, No. 9, 2007 2825

Figure 3. Flooding data on a Sherwood diagram.

organic gas could be detected by a flame ionization detector (FID). After 5.5 min, the divert valve opened and hence the midpoint gas pressure directed the flow of the ethanolaminerich gas from the first column directly to the FID. The rationale of this action was to avoid bringing the ethanolamine into contact with the catalyst converter because it could poison the catalyst. The total retention time for a complete gas/liquid sample analysis was 10 min in which the CO2 peak appeared at 4.28 min and the MEA peak at 7.02 min. The GC in the current investigation is compared with that of Jou et al.19 in Table 1. Multiple analyses of identical samples showed the reproducibility of CO2 measurement in gas samples was (0.6%, and it was ( 1.6% and ( 1.4% for CO2 and MEA measurements in the liquid sample, respectively. 4. Results and Discussion

Figure 4. Variation of gas superficial mass velocity with rotor speed at a constant liquid superficial mass velocity during flooding experiments.

The mass transfer calculations showed unexpected low CO2 recovery. Therefore, a new experimental campaign was carried out with higher MEA concentrations: 100, 75, and 55 wt %. Liquid analysis showed that the difference between the lean and the rich loading measurements was trivial compared to the loading level. Hence, the system can be analyzed considering liquid loading as constant. The average lean/rich loading measurements were 0.023, 0.048, 0.094, and 0.329 mol CO2/ mol MEA for 100, 75, 55, and 30 wt % MEA. The aqueous MEA solution was then upgraded to 64 wt % by pure MEA and then loaded to 0.43 mol CO2/mol amine before desorption runs were carried out at MEA strengths of 64, 54, and 34 wt %. 3. Analysis Method. A gas chromatograph developed by Unicam Chromatography (UK) was used to detect the concentration of CO2 and MEA in the vapor and liquid phases. The GC has two columns and a catalytic converter. The dimensions, column packing details, and the operational parameters are shown in Table 1. A total volume of 0.2 µL liquid sample was injected manually through a rubber septum into the sample injector port with the aid of a 1 µL Hamilton syringe. The barrel of the syringe was made from glass, and the needle was stainless steel. The entire sample was contained in the needle, and it was vaporized in the injector port as the temperature of the injector oven reached 375 °C. The gaseous sample was then mobilized by helium at a volumetric flowrate of 10.8 cm3/min through the 25 m × 0.53 mm fused silica packed column of CP Wax. Both CO2 and air would diffuse much faster than MEA. The second 1.0 m × 0.32 mm Silcosteel packed column of Carbosieve S-II separated CO2 from air, and then, it was catalytically converted to methane using hydrogen; hence, the

4.1. Flooding. The flooding experiments were carried out by varying the superficial mass velocity of gas, superficial mass velocity of liquid, and rotor speed. The flooding points were determined by fixing two of the operating variables and manipulating the third one. Excessive splash of MEA solution in the eye of the rotor was the selected flooding criterion. Two procedures were carried out to confirm the results prior to the calculation of the abscissa and ordinate in the Sherwood plot (Figure 3). For expamet packing with 0.76 porosity () and 2132 m2/m3 specific surface area (ap), the experimental results are represented by the following:

ln

[ ( )] uG2ap FG rω23 FL

) -3.01 - 1.40 ln

(x ) [ ( x )] L G

FG FL

0.15 ln

L G

FG FL

2

(1)

The flooding velocities in the Higee were higher than those in dumped rings and were closer to the the values in stacked rings. This confirms the higher hydraulic capacity of expamet in comparison to wire gauze packing that showed similar behavior to Sherwood correlation for dumped rings.20 The effect of rotor speed on superficial mass velocity of gas (G) at the constant superficial mass velocity of liquid (L) is shown in Figure 4. Lockett21 correlated a relationship between Sherwood and Wallis flooding representations. Equation 2 shows the transformation of a Wallis-type correlation to a Sherwood-type flooding correlation. For expamet packing, the coefficients in eq 2 are the following: β ) 2.03 and γ ) 2.058. Figure 5 shows the experimental flooding data using a Sherwood-Wallis plot.

CG )

[

]

βa -0.25Ng0.22µ -0.03 L FG 0.5 0.5 1+γ G FL

(( ))

2

(2)

4.2. Absorption Runs. 4.2.1. Mass Transfer. A series of absorption experiments were performed under atmospheric pressure and at four different MEA concentrations: 30, 55, 75, and 100 wt %. Figure 6 shows the performance of rotating packed bed in the absorption mode that was operated at several rotor speeds and lean amine temperatures. The experimental

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equation could be given in terms of the area of transfer unit (ATU).20,23

QG NTUOG KGaeZ

π(ro2 - ri2) ) ATUOGNTUOG )

(5)

Rearranging eq 5, the overall gas-phase mass transfer coefficient is

K Ga e )

QG

Tables 3-6 show the operating conditions, KGa, and the CO2 penetration for the absorption runs at 30, 55, 75, and 100 wt % MEA. 4.2.3. Mass Transfer Control in a Conventional Absorber. The primary resistance to CO2 mass transfer in a conventional MEA absorber for this system is not in the gas film as can be shown by tray efficiency. Kohl24 observed that the chemical absorption of CO2 by alkanolamines gave 8-25% plate efficiency. In comparison, the gas film controlling processes (such as the absorption of ammonia) gave >80% plate efficiency. Therefore, it could be concluded that the absorption of CO2 into alkanolamines solutions is a liquid film controlled process. 4.2.4. Mass Transfer Control in a Rotating Packed Bed. 4.2.4.1. Liquid Film Resistance, Equilibrium Reaction. If instantaneous reversible reactions dominate the rate of absorption, the mass transfer flux will be independent of reaction rates but dependent on the diffusion of reactants and products. An expression was introduced by Tung and Mah25 using the penetration model to describe the liquid mass transfer behavior in the rotating packed beds. In their procedure, the kL value is calculated using eq 7 and the total gas-liquid interfacial area is calculated with the Onda et al.26 correlation (eq 8).

Figure 5. Experimental data on a Sherwood-Wallis flooding plot.

() [ ()

Figure 6. Effect of MEA concentration on CO2 penetration at a liquid flowrate of 2.4 m3/h, gas flowrate of 64.4 m3/h, 4.4 vol % CO2 in sour gas, and average lean/rich loading (mol CO2/mol MEA).

at kL,calcdp ) 0.919 DL a

data were presented in terms of CO2 penetration that is defined as follows:

σc a ) 1 - exp -1.45 at σ

(

CO2 penetration (%) ) 1 -

)

yCO2,in - yCO2,out yCO2,in

× 100 (3)

Figure 6 shows that the rotating packed bed was most efficient at higher MEA concentrations. However, it surprisingly showed high penetration at a typical conventional absorber concentration of 30 wt % MEA. Section 4.2.7 explains the reasons for such behavior. 4.2.2. Calculation of KGa. The difficulty of separation is expressed in terms of the number of transfer units (NTU). The number of overall transfer units based upon the change in gas concentration is defined by Colburn:22

NTUOG )

∫yy

2

1

( )

yCO2,in dy ) ln y - y* yCO2,out

(6)

NTUOG

π(ro - ri2)Z 2

(4)

Equation 4 assumes that the equilibrium partial pressure of CO2 is negligible (y* ) 0) because the CO2 loading (mol/mol MEA) is low relative to yCO2 and because of the fast chemical reaction between CO2 and the concentrated amine solutions. The equilibrium pressure at 40 °C and a loading of 0.33 is approximately 0.04 kPa.19 The rotating packed bed design

1/3

ScL1/2ReL2/3GrL1/6

0.75

ReL0.1WeL0.2FrL-0.05

(7)

]

(8)

The liquid diffusion coefficient (DL) is predicted using eqs 1720. The effective pore diameter of packing (dp) is 675 µm. It should be noted that the Onda et al.26 correlation was derived from data other than the rotating packed beds. The present application of this model is thus an extrapolation. The experimental liquid film mass transfer coefficient is calculated using eqs 9-12.27

kL,exp )

L

NTUL

aeFπ(ro2 - ri2)Z

(9)

The number of transfer units in the liquid phase is defined in terms of CO2 loading (R, mol/mol MEA), and this is analogous to eq 4:

NTUL )

X 2 - X 1 R2 - R1 ) ∆XLM ∆RLM

(10)

From the experimental gas-phase mole balance, the loading difference could be calculated:

(R2 - R1) )

QG (y - yCO2,out) QLCT CO2,in

(11)


4.2.4.2. Fast Reaction in the Boundary Layer. Freguia and Rochelle28 found this mechanism to be important for CO2 absorption by MEA solution, and it falls into the interface pseudofirst-order (IPFO) reaction regime. Dankwerts30 used the surface renewal model to give the rate of CO2 absorption for this case.

NCO2 ) kL

x

1+

k1DCO2 ∆P k 2 HCO2

(13)

L

where k1 is the rate coefficient for the pseudo-first-order reaction defined by

k1 ) k2[MEA] Figure 7. Distribution of kl,calc/kl,exp ratio at different MEA concentrations.

(14)

where k2 is the kinetic rate constant and given by Hikita et al.31 for the temperature range of 5-80 °C:

log10 k2 ) 10.99 - (2152/T)

(15)

Equation 13 can be simplified because the second term under the square root is much greater than unity. Thus, the overall gas-phase mass transfer coefficient for the fast reaction of CO2 with amine is given by

K Ga )

Figure 8. Effect of rotor speed on kl,calc/kl,exp for 30 wt % MEA solution.

The log mean loading driving force is a function of the measured loading and the equilibrium loading (R*) at the operating CO2 partial pressures.

∆RLM )

(R2 - R/2) - (R1 - R/1) ln

( ) R2 - R/2

(12)

R1 - R/1

The R* values were generated using the FLASH module based on the Freguia-Rochelle model28 in the Aspen Plus environment. Figure 7 shows the distribution of kl,calc/kl,exp ratio at different MEA concentrations. The calculated kl is greater than the experimental kl by an average factor of 6.2. The deviation might be attributed to the fact that the calculated kl uses the viscosity and CO2 diffusion coefficient in MEA solution with zero loading. In practice, in our data, even if the bulk solution has zero loading, the interface will have a loading of 0.5. According to Weiland et al.,29 the viscosity and therefore the diffusion coefficients were extrapolated to increase by factors of 0.5, 2.2, and 7.4 for 30, 55, and 75 wt % solutions, respectively. Also, the CO2 diffusion coefficient of the lean solutions is predicted from data measured at 30 wt % (eq 20). Furthermore, the estimation of the overall kl neglects any resistance because of the fast reaction. Although the fast reaction does not dominate the rate, it may offer a significant resistance in the more concentrated amine solutions. Another reason for the apparent low values of the liquid film mass transfer coefficient is the use of Onda’s equation to predict wetted area as the occurrence of channeling is possible and hence the reduction of the available gas-liquid interfacial area. Figure 8 shows that model correctly predicts the effects of variation of rotor speed, although this variable has only a power of 0.05 in the Froude number in the Onda expression.

xkappDCO a 2

(16)

HCO2

where “a” refers to the total geometric area as it is assumed that the packing was completely wet. Aboudheir et al.32 provided the apparent reaction-rate constant (kapp) at operational MEA concentrations and temperatures, and these are given in Table 2. The diffusivity and physical solubility of CO2 in MEA solution was determined using the N2O analogy (eq 17). The diffusivities of CO2 and N2O in water as a function of temperature are given by Versteeg et al.33 in eqs 18 and 19. The diffusivity of N2O in aqueous MEA is correlated by Ko et al.34 in eq 20.

( ) ( ) DN2O DCO2

)

amine

DN2O DCO2

(17)

water

DCO2_H2O ) 2.35 × 10-6 exp(-2119/T)

(18)

DN2O_H2O ) 5.07 × 10-6 exp(-2371/T)

(19)

DN2O_MEA ) {5.07 × 10-6 + 8.65 × 10-7camine + 2.78 × 10-7camine2} exp

(

)

-2371 - 93.4camine (20) T

The physical solubility correlations for CO2 and N2O in water are given in eqs 21 and 22.32 The solubility of N2O in pure MEA is given by Wang et al.35 at operating temperatures of 20.4 and 40.2 °C to be 2425 and 3262 kPa m3/kmol. A summary of results is shown in Table 2.

HCO2_H2O ) 2.82 × 106 exp(-2044/T)

(21)

HN2O_H2O ) 8.55 × 106 exp(-2284/T)

(22)

This model (eqs 16-22) predicts higher KGa values of 15.855.6 1/s, which are much greater than the measured values (1-7 1/s). As we are seeking limiting mechanisms, we expect that

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Table 2. Calculated KGa Values for Fast Reaction Mechanism with HCO2_MEA ) 0.745 and 0.638 at 293 and 313 K DCO2_MEA (109 m2/s)b

kapp (1/s)a

MEA conc

KGa (1/s)

wt %

mol/dm3

293 K

313 K

293 K

313 K

293 K

313 K

30 55 75 100

4.9 9.0 12.3 16.4

30 750 80 000 160 000 180 000

90 000 216 666 350 000 450 000

1.12 0.67 0.38 0.17

1.97 1.28 0.78 0.37

16.8 21.0 22.4 15.8

44.5 55.6 55.3 43.4

a

Aboudheir et al.32

b

Calculated using eq 16.

the calculated KGa from each the limiting mechanisms should always be greater that the measured KGa. If the calculated value is much greater than the measured value, then its respective mechanism probably plays little role in the mass transfer process. The fast reaction mechanism is thus not solely limiting, but may contribute a significant resistance to the mass transfer. In contrast to measured values, the predicted values are sensitive to temperature but not to the variation in MEA concentration, indicating that the mass transfer of CO2 is not limited by the mechanism of fast reaction in the boundary layer. Therefore, this mechanism does not predict the magnitude or trends of the experimental results. 4.2.4.3. Reaction in Liquid Bulk. An extreme case is when the reaction takes place in the liquid bulk rather than the boundary layer. There is also no significant loss of CO2 concentration across the liquid boundary layer. In this case, the resistance to mass transfer depends on the bulk chemical reaction and is hence controlled by liquid holdup. The KGa for this case could be derived from flux equations:36

NA ) KG(PAG - P/A) NA )

(23)

kapp (C - CAL)L a Ai

(24)

Rearranging eq 23 by 24 yields

K Ga )

kapp HCO2_MEA

(25)

L

where HCO2_MEA is the Henry’s law constant and it is defined as the ratio of delta concentrations in the case when Henry’s law does not hold. The liquid holdup (L) for high voidage structured packing in a rotating packed bed is correlated by Burns et al.:37

L ) 0.039

() ( )() g g0

-0.5

U U0

0.6

V V0

0.22

(26)

where the gravitational acceleration is defined for rotating equipment as

g ) rω2

(27)

Figure 9. Effect of MEA concentration on experimental KGa at a liquid flowrate of 2.4 m3/h and gas flowrate of 64.4 m3/h.

The following are characteristic values for acceleration (g0) 100 m/s2, superficial flow velocity (U0) 1 cm/s, and kinematics viscosity (V0) 10-6 m2/s. The kapp and HCO2 values were determined as described in the previous section. The range of the predicted KGa values was 0.46-73.80 1/s. This range overpredicts the measured KGa values by a 0.8-5.2 order of magnitude indicating that the resistance to mass transfer is not dictated by liquid bulk reactions. 4.2.5. Effect of Rotor Speed. The rotating packed bed was operated between 600 and 1000 RPM giving a gravitational acceleration of 31 and 87 g, respectively. Tables 3-6 show the variation of rotor speed on the kl,exp for the investigated MEA concentrations at different operating conditions. It is believed that the higher rotor acceleration causes a subsequent rise to the interphase velocity throughout the packed bed. This fact in conjunction with a large specific area causes the creation of thin films and small droplets, and hence, the overall result is a better mass transfer performance as indicated by the higher mass transfer coefficients achieved at higher rotor speeds. According to Burns and Ramshaw,38 the rotating packed bed is more efficient at higher rotor speeds due to reduction of maldistribution as the type of flow in the rotating packed bed is primarily dictated by the rotor speed. The pore flow or rivulets are dominant at low rotor speeds (25 g). Therefore, the range of the experimental rotor speeds indicates the region of flow type to be droplet and/or film flow, and this ensures minimization of maldistribution levels. 4.2.6. Effect of MEA Concentration. Lower CO2 penetration was achieved with concentrated MEA as shown by Figure 6. The sharp performance enhancement might be attributed to accelerated absorption kinetics especially when coupled with higher temperature. The total solubility of CO2 and the driving forces for the forward instantaneous reaction are proportional to MEA concentration. Figure 6 and Tables 3-6 show that the CO2 loadings in the average lean/rich solution are lower in the highly concentrated solution, and that might be the reason for higher CO2 penetration using 30 wt % MEA.

Table 3. Pilot Plant Results for CO2 Absorption at 30 wt % MEA and PT ) 1 atm MEA (wt %)

rotor speed (RPM)

lean T (˚C)

superficial liquid velocity (m/s)

superficial gas velocity (m/s)

average lean/rich loading (mol CO2/mol MEA)

CO2 in (vol%)

KGa (1/s)

CO2 penetration (%)

30

600

40.5 22.5 40.6 22.8 39.9 23.1 39.3 23.8

0.00633 0.00633 0.00633 0.00633 0.00317 0.00317 0.00317 0.00317

0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988

0.336 0.323 0.339 0.328 0.330 0.325 0.333 0.320

4.5 4.4 4.3 4.3 4.4 4.4 3.6 4.2

1.05 0.71 1.22 0.79 0.80 0.76 0.64 0.72

47.4 60.1 41.8 57.1 56.6 58.1 63.4 59.9

1000 600 1000

Ind. Eng. Chem. Res., Vol. 46, No. 9, 2007 2829 Table 4. Pilot Plant Results for CO2 Absorption Runs at 55 wt % MEA and PT ) 1 atm MEA (wt %)

rotor speed (RPM)

lean T (˚C)




CO2 in (vol%)

KGa (1/s)

CO2 penetration (%)

55

600

39.6 20.7 40.1 20.9 39.5 22.3 39.6 22.6

0.00633 0.00633 0.00633 0.00633 0.00317 0.00317 0.00317 0.00317

0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988

0.080 0.092 0.080 0.094 0.105 0.100 0.103 0.098

4.7 4.6 4.5 4.5 4.4 4.5 4.4 4.1

4.18 2.48 4.32 2.86 2.87 2.58 3.22 2.78

5.1 17.0 4.6 13.0 13.0 15.9 10.1 13.8

1000 600 1000


rotor speed (RPM)

lean T (˚C)




CO2 in (vol%)

KGa (1/s)

CO2 penetration (%)

75

600

41.0 21.4 40.2 21.0 40.8 22.1 39.4 20.6

0.00633 0.00633 0.00633 0.00633 0.00317 0.00317 0.00317 0.00317

0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988

0.051 0.040 0.049 0.038 0.061 0.047 0.055 0.044

4.4 4.4 4.3 4.3 3.6 4.4 4.4 4.5

5.63 2.59 5.19 3.42 5.47 2.60 5.54 3.38

1.8 15.8 2.5 8.8 2.0 15.7 1.9 9.0

1000 600 1000


rotor speed (RPM)

lean T (˚C)




CO2 in (vol%)

KGa (1/s)

CO2 penetration (%)

100

600

40.7 20.6 40.5 24.3 41.1 20.4 40.9 20.3

0.00633 0.00633 0.00633 0.00633 0.00317 0.00317 0.00317 0.00317

0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988 0.16988

0.010 0.041 0.007 0.026 0.014 0.043 0.004 0.039

3.4 4.5 4.5 4.2 4.5 4.2 4.3 4.4

6.09 5.59 7.11 6.48 6.21 4.85 6.88 6.25

1.3 1.9 0.6 1.0 1.2 3.1 0.7 1.2

1000 600 1000

Even though the liquid side mass transfer coefficient decreases with highly viscous MEA solutions, the larger driving forces contribute to lower CO2 penetration. A significant increase in the overall gas phase mass transfer coefficient was observed when the MEA concentration was increased as shown in Figure 9. 4.2.7. Comparison with Conventional Absorber. According to operating data by Kohl and Nielsen,36 for an aqueous MEA gas treating plant, a typical absorber consists of two 7-m (23-ft) beds of polypropylene saddles with 14.5 m diameters. The average inlet CO2 composition was 12.5 vol % and the average outlet was 0.2 vol %; hence, the number of transfer units for the absorber is 4.1, and thus, the HTUOG is computed to be 3.4 m. In comparison, the average HTUOG for the Higee machine using 30 wt % MEA solution could reach a maximum of 21 cm as shown in Table 7. 4.3. Desorption Runs. The desorption runs were carried with 30, 54, and 60 wt % MEA solutions, preheated to 57-70 °C, and at atmospheric pressure as shown in Table 8. The flow of steam inside the rotating shaft, the gas plenum, and the holes was subsonic as the Mach number (Ma) was approximately 0.43. There were no operational problems in dealing with high temperatures and high MEA concentrations as the rig was made from stainless steel. 4.3.1. Calculation of KGa. The expression for the number of the overall gas transfer units in the rotating packed bed at constant Henry’s law values is given by McCabe et al.:39

NTUOG )

y1 - y2 ∆ylm

(28)

Assuming linear equilibrium and operating lines, the log

Table 7. Comparison between Conventional Absorber36 and RPB conventional absorber36 diameter (m)

4.42

packing height (m) internals

14.02 2 beds of polypropylene saddlesa 1 (40˚C)b 10-18 10-15 0.1-0.3 0.062 0.415 4.45 40c 3.4 0.068b 180a 144 0.198 0.0163

pressure (atm) temperature (°C) MEA conc (wt %) feed CO2 (vol %) outlet CO2 (vol %) lean loading (mol/mol) rich loading (mol/mol) L/G (kg/kg) flooding (%) HTUOG (m) KGa (1/s) at (m2/m3) aw (m2/m3)26 kGa (1/s)41 kLa (1/s)25

RPB 0.398 (OD); 0.156 (ID) 0.025 axial depth stainless Perspex sheets 1 20/40 30 3.5-4.5 0.03-0.13 refer to Table 3 16.30-30.75 6.6-15.5 0.14-0.27 0.64-1.22 2132 992 0.233 2.336-4.916

a Assume 1.5 in saddles. b Calculated using the Kohl and Nielsen36 equation at 40 °C. c Eckert flooding diagram.40

mean driving force is given by

∆ylm )

(y1 - y/1) - (y2 - y/2) ln

( ) y1 - y/1

(29)

y2 - y/2

where y1 and y2 are the vapor mole fractions and y/1 and y/2 are the equilibrium mole fractions. The Freguia and Rochelle28 model was used to generate equilibrium values using the FLASH module in the Aspen Plus environment as no experimental

2830


Table 8. Pilot Plant Results for Desorption Runs with PT ) 1 atm MEA (wt %)

rotor speed (RPM)

lean temp (°C)

34

800 800 800 800 600 1000 600 1000 600 1000 600 1000

68.0 69.0 67.1 70.0 58.2 59.7 56.9 57.2 58.8 58.4 58.9 59.3

54

60

loading (mol CO2/mol MEA) rich lean 0.405 0.346 0.379 0.399 0.414 0.407 0.431 0.422 0.402 0.403 0.437 0.432

solubility data were available in the literature for high MEA concentrations at high temperatures. The FLASH module temperature was set to 100 °C because the saturated steam at atmospheric pressure was the stripping medium. The KGa was calculated using eq 6. The effect of the specific steam rate on the KGa at three amine concentrations is shown in Figure 10. The stripping cycle is believed to be gas film controlled because the mass transfer is observed to be independent of the liquid rate and rotor speed. Also, the operation of the rotating packed bed causes the creation of thin films, and the stripping at high-temperature causes lower solution viscosities; thus, the liquid film resistance is considered to be negligible. Table 9 shows the dimensions of the rotating packed bed and the operating conditions for a typical desorption run. On the basis of the overall energy balance calculations, the average molar H2O/CO2 ratio was 45.2 and hence the estimated outlet CO2 partial pressure was 2.64 kPa. The equilibrium partial pressure of CO2 at an average loading of 0.4 mol/mol and 100 °C in 30 wt % solution was 20 kPa.35 Thus, there is no rich end pinch in these mass transfer experiments. 4.3.2. Comparison with Conventional Desorber. Table 9 shows the necessary dimensions for a simulated conventional stripper to achieve similar performance to the rotating packed bed at an MEA concentration of 34 wt %. The stripper was modeled using the Aspen Plus environment, and the reactions were in equilibrium. The simulated stripper was randomly filled with metallic cascade mini rings (CMR; type 2) that has a specific surface area of 144 m2/m3 and a voidage of 0.971. The

Figure 10. Variation of KGa with G/L at a constant steam rate (250 kg/h).

specific steam rate (kg steam/L soln)

KGa (1/s)

H2O/CO2 (mol/mol)

0.09 0.20 0.34 0.37 0.12 0.12 0.20 0.20 0.43 0.43 0.14 0.14

38.3 57.5 73.1 76.7 49.2 48.8 50.3 52.0 85.9 85.9 54.4 54.4

39.5 79.3 45.2 38.7 34.5 46.9 46.8 56.0 43.8 44.3 39.7 41.8

0.398 0.329 0.321 0.329 0.404 0.399 0.408 0.403 0.332 0.334 0.424 0.419

Table 9. Comparison between Simulated Stripper and RPB at an MEA Concentration of 34 wt % simulated stripper rich solution temp (°C) lean loading (mol CO2/mol MEA) rich loading (mol CO2/mol MEA) Gm/Lm dimensions (m) mol H2O/ mol CO2 PCO2 (kPa), at top of column

RPB

67.1 0.3221

67.1 0.3211

0.3792

0.3788

0.49 packing: metallic, random, CMR (#2) 4.5 ID × 0.21 height 55.6 1.79

0.49 0.398 OD 0.156 ID 0.025 thickness 45.2 2.64

height of the column was varied while the rich solution loading, the rich solution temperature, and the molar L/G were kept constant in order to achieve the lean loading specification. The results show that the conventional stripper height necessary to achieve similar performance to the rotating packed bed is greater by a factor of 8.4 and its diameter is greater by a factor of 11.3. 5. Conclusions Because of the complex nature of the mass transfer accompanied by chemical reaction for absorption of CO2 by concentrated MEA solutions in a rotating packed bed, limiting mechanisms were used to calculate the expected performance of the mass transfer machine. The mechanisms of mass transfer with equilibrium reactions matches the data best, but the predicted rates were higher than the measured rate. A possible explanation for this is that the reactions are not instantaneous and thus the measured rate is slower than the predicted rate. At the other extreme, the measured rate is also much slower than the predicted rate assuming a fast but a finite rate of reaction with no depletion of MEA. The most important factor that influenced the low CO2 penetration was the MEA concentration as the driving forces are proportional to this variable. The range of rotor speeds giving accelerations of 31-87 g ensured minimum maldistribution and achieved high mass transfer coefficients. It was also noticed that a higher lean amine temperature contributed to lower CO2 penetration and better mass transfer coefficients. In the desorption mode, the experimental results indicate that the mass transfer in the rotating packed bed is gas film controlled. The comparison with a simulated stripper confirmed the size reduction advantage of a rotating packed bed at similar operational conditions. Therefore, the rotating packed bed is an attractive option to use for gas sweetening applications in offshore facilities where size and space are important.


Thus, the pressure drop is

∆Pf ) -

Figure A. Sudden contraction to the liquid flow in a four-arm distributor.

2 1013‚8.242 1 - 1 ) -8.34 kPa 2 0.67

[

Thus, the total pressure drop in each arm due to the sudden contraction is 66.72 kPa. The pressure drop in each arm due to frictional loss can be calculated as follows:

mean velocity ) Acknowledgment The authors would like to pay tribute to the generous financial assistance of Hydro Oil & Energy to carry out this project.

Q)

Appendix 1: Pressure Drop Calculation in the Liquid Distributor Figure A shows the configurations of the four-arm liquid distributor:

number of arms ) 4

]

Q πdi2/4

0.00083 ) 0.0002075 m3/s 4

πdi2 π(0.01)2 ) ) 0.0000785 m2 4 4 Thus, the mean velocity ) 2.64 m/s.

Reynolds number ) Re )

length of each arm ) 12 cm ) 0.12 m

Fud 1013‚2.64‚0.01 ) ) µ 2 × 10-3 1.3 × 104

number of holes per arm ) 8

pipe roughness ) 0.000045 m (commercial steel)

diameter of each hole ) 2 mm ) 0.002 m

Thus, the relative roughness ) e/d ) 0.000045/0.01 ) 0.0045. Accordingly, the fanning friction factor ) 0.009.43 The frictional pressure drop per unit length is given by

The physical properties of 30 wt % ethanolamine solution at 30 °C are the following: density ) 1013 kg/m3 and dynamic viscosity ) 2 × 10-3 Pa‚s. The following calculations are based upon an ethanolamine solution flowrate of 50 L/min ) 0.00083 m3/s.

(hole diameter)2 ) 4 2 π(0.002 /4) ) 3.141 × 10-6m2

cross-sectional area of each hole ) π

flowrate per hole )

0.00083 total flowrate ) ) number of holes 8‚4 25.9 × 10-6 m3/s -6

velocity of solution for each hole )

Re ) Fud/µ )

25.9 × 10 Q ) ) A 3.141 × 10-6 8.24 m/s

1013‚8.24‚0.002 ) 8.347 2 × 10-3

The pressure drop for a sudden contraction to the effective area of the flow could be calculated as follows:42

∆Pf ) -

[

Fu2 1 -1 2 Cc

]

2

where Cc is the contraction coefficient and it varies between 0.6 and 1.0 as the ratio of the pipe diameters varies from 0 to 1 (the ratio is 0.2 in our case), assuming a common value for Cc of 0.67. The velocity of the fluid (u) refers to the smaller pipe.

∆Pf ) 4f

()

( )

L Fu2 1 1013‚2.642 ) 4‚0.009 ) 12.7 kPa/m di 2 0.01 2

The distance between the holes in each arm is approximately 0.5 cm. Thus, the frictional pressure drop is only 0.06 kPa between the two holes. This loss is negligible in comparison to the total pressure drop due to the sudden contraction in the holes. The length of each arm is 0.12 m; thus, the frictional pressure drop per arm is 1.52 kPa.

total pressure drop in each arm ) 66.72 + 1.52 ) 68.24 kPa Nomenclature A ) gas-liquid interfacial area (m2/m3) ac ) centrifugal acceleration (m/s2) a, at, ap ) total specific surface area of packing (m2/m3) ae ) effective specific surface area (m2/m3), eqs 5, 6, and 9 ATU ) area of a transfer unit (m2) CG ) gas capacity factor (m/s) camine ) concentration of amine in solution (kmol/m3) CA ) molar concentration of A (kmol/m3) CAi ) molar concentration of A at the interface (kmol/m3) CAL ) molar concentration of A in the bulk of the liquid (kmol/ m3) CT ) total concentration of MEA solution (kmol/m3) DCO2 ) diffusion coefficient of CO2 (m2/s) DN2O ) diffusion coefficient of N2O (m2/s) DL ) diffusion coefficient of liquid (m2/s) dp ) diameter of packing pore (m) dp ) 6(1 - )/at g ) gravitational acceleration (m2/s) or acceleration due to centrifugal field (eq 27) g0 ) characteristic acceleration value (eq 26) ) 100 m/s2 G ) superficial mass velocity of gas (kg/s‚m2)

2832


HCO2 ) Henry’s law constant (Pa/kmol‚m3) kL,calc ) calculated mass transfer coefficient of liquid (m/s) kL,exp ) experimental mass transfer coefficient of liquid (m/s) k1 ) pseudo-first-order reaction rate constant (1/s) k2 ) second-order reaction rate constant (1/s) kapp ) apparent pseudo-first-order reaction rate constant (1/s) KG ) overall gas phase transfer coefficient (m/s) KGa ) overall volumetric mass transfer coefficient (1/s) L ) superficial mass velocity of liquid (kg/m2‚s) Ng ) dimensionless acceleration ) ω2r/g NCO2 molar rate of absorption of CO2 per unit (kmol/s‚m2) NTUOG ) overall gas phase number of transfer units (-) NTUL ) number of transfer units for the liquid phase (-) ∆P ) pressure drop (N/m2) PAG ) partial pressure of A in bulk of gas phase (N/m2) P/A ) partial pressure of A in equilibrium with a solution having the composition of the main body of liquid (N/m2) QG ) volumetric flowrate of gas (m3/s) QL ) volumetric flowrate of liquid (m3/s) PT ) total pressure (atm) r ) radial position (m) ri ) inner radius of the rotor (m) ro ) outer radius of rotor (m) T ) temperature (K) uG ) superficial gas velocity (m/s) U ) superficial flow velocity (m/s) Uo ) characteristic superficial flow velocity (m/s) VL ) superficial liquid velocity (m/s) X1, X2 ) mole fraction of CO2 in lean/rich solution (-) Xlm ) log mean mole fraction in liquid (-) yCO2,in ) mole fraction of CO2 in inlet stream (-) yCO2,out ) mole fraction of CO2 in outlet stream (-) y/ ) gas-phase mole fraction of CO2 in equilibrium with CO2 concentration in liquid (-) Z ) axial height of the packing (m) Dimensionless Numbers FrL ) liquid Froude number (VL2at/g) GrL ) liquid Grashof number (dp3ac/νL2) ReL ) liquid Reynolds number (VL/atνL) ScL ) liquid Schmidt number (νL/DL) WeL ) liquid Webber number (VL2FL/atσ) Greek Letters ) voidage (-) L ) liquid holdup (-) R1, R2 ) CO2 loading in lean and rich solutions (mol/mol MEA) RLM ) log mean loading defined by eq 11 (mol/mol MEA) µ ) viscosity (Pa‚s) FL ) liquid density (kg/m3) FG ) gas density (kg/m3) σ ) liquid surface tension (N/m) σc ) critical surface tension (N/m) νL ) kinematics liquid viscosity (m2/s) ω ) rotational speed (rad/s) AbbreViations CMR ) cascade mini rings HTU ) height of transfer unit MEA ) monoethanolamine

N2O ) nitrous oxide NTU ) number of transfer unit Literature Cited (1) Mallinson, R. H.; Ramshaw, C. Mass Transfer Apparatus and Process. European Patent 0053881, 1982. (2) Ramshaw, C.; Mallinson, R. H. Mass Transfer Process. U.S. Patent 4,283,255, 1981. (3) Green, A. Process Intensification: the key to survival in global markets? Chem. Ind. 1998, 168-172. (4) Ramshaw, C. Opportunities for exploiting centrifugal fields. Heat RecoVery Syst. CHP 1993, 13 (6), 493-513. (5) Astarita, G.; Savage, D. W.; Bisio, A. Gas Treating with Chemical SolVents; John Wiley & Sons: New York, 1983. (6) Astarita, G. Carbon dioxide absorption in aqueous monoethanolamine solutions. Chem. Eng. Sci. 1961, 16, 202-207. (7) Zarzycki, R.; Chacuk, A. Absorption: fundamentals and applications; Pergamon Press Ltd: Elmsford, NY, 1993. (8) Caplow, M. Kinetics of carbamate formation and breakdown. Am. Chem. Soc. 1968, 90, 6795-6803. (9) Danckwerts, P. V. The reaction of CO2 with ethanolamines. Chem. Eng. Sci. 1979, 34 (4), 443-446. (10) Dankwerts, P. V. Gas-Liquid Reactions; McGraw-Hill: New York, 1970. (11) Versteeg, G. F.; Van Swaaij, W. P. M. Solubility and Diffusivity of Acid Gases (CO2, N2O) in aqueous alkanolamine Solutions. J. Chem. Eng. Data 1988, 33, 29-34. (12) Blauwhoff, P. M. M.; Versteeg, G. F.; VanSwaaij, W. P. M. A study on the reaction between carbon dioxide and alkanolamines in aqueous solutions. Chem. Eng. Sci. 1984, 39 (2), 207-225. (13) Chambers, H. H.; Wall, M. A. Some factors affecting the design of centrifugal gas absorbers. Trans. Inst. Chem. Eng. 1954, 32, S96-S107. (14) Bucklin, R. W.; Buckingham, P. A.; Smelser, S. C. The Higee demonstration test of selective H2S removal with MDEA. Presented at The 38th Laurence Reid gas conditioning conference; Norman, Oklahoma, March, 1988. (15) Lin, C.-C.; Liu, W.-T.; Tan, C.-S. Removal of carbon dioxide by absorption in a rotating packed bed. Ind. Eng. Chem. Res. 2003, 42, 23812386. (16) Rao, D. P.; Bhowal, A.; Goswami, P. S. Process Intensification in Rotating packed beds (HIGEE): An Appraisal. Ind. Eng. Chem. Res. 2004, 43, 1150-1162. (17) Jassim, M. S. Process Intensification: Absorption and desorption of carbon dioxide from monoethanolamine solutions using HIGEE technology. Ph.D. thesis, University of Newcastle upon Tyne, UK, 2002. (18) Hassan-beck, H. M. Process Intensification: mass transfer and pressure drop for counter-current rotating packed bed. Ph.D. thesis, University of Newcastle upon Tyne, UK, 1997. (19) Jou, F.-Y.; Mather, A. E.; Otto, F. D. The solubility of CO2 in a 30 mass percent monoethanolamine solutions. Can. J. Chem. Eng. 1995, 73, 140-147. (20) Singh, S. P.; Wilson, J. H.; Counce, R. M.; Villiers-Fisher, J. F.; Jennings, H. L.; Lucero, A. J.; Reed, G. D.; Ashworth, R. A.; Elliott, M. G. Removal of volatile organic compounds from groundwater using a rotary air stripper. Ind. Eng. Chem. Res. 1992, 31, 574-580. (21) Lockett, M. J. Flooding of rotating structured packing and its application to conventional packed columns Chem. Eng. Res. Des., Part A 1995, 73, 379-384. (22) Colburn, A. The Simplified calculation of diffusional process. General consideration of two-film resistances. Trans. AIChE 1939, 35, 211. (23) Kelleher, T.; Fair, J. R. Distillation studies in a high-gravity contactor. Ind. Eng. Chem. Res. 1996, 35, 4646-4655. (24) Kohl, A. Plate Efficiency with Chemical Reaction - Absorption of Carbon Dioxide in Monoethanolamine Solutions. AIChE 1956, 2, 264270. (25) Tung, H.; Mah, R. S. H. Modeling Liquid Mass Transfer in HIGEE Separation Process. Chem. Eng. Commun. 1985, 39, 147-153. (26) Onda, K.; Takeuchi, H.; Okumoto, Y. Mass Transfer Coefficient between Gas and Liquid Phases in Packed Columns. J. Chem. Eng. Jpn. 1968, 1, 56. (27) Coulson, J. M.; Richardson, J. F.; Backhurst, J. R.; Harker, J. H. Chem. Eng. 2002, 2 (5), 564-570. (28) Freguia, S.; Rochelle, G. T. Modeling of CO2 capture by aqueous monoethanolamine. AIChE J. 2003, 49, 1676-1686.

Ind. Eng. Chem. Res., Vol. 46, No. 9, 2007 2833 (29) Weiland, R. H.; Dingman, J. C.; Cronin, D. B.; Browning, G. J. Density and viscosity of some partially carbonated aqueous alkanolamine solutions and their blends. J. Chem. Eng. Data 1998, 43, 378-382. (30) Dankwerts, P. V. Gas-Liquid Reactions; McGraw-Hill: New York, 1970. (31) Hikita, H.; Asai, S.; Ishikawa, H.; Honda, M. The kinetics of reactions of carbon dioxide with monoethanolamine, diethanolamine and triethanolamine by a rapid mixing method. Chem. Eng. J. 1977, 13, 7-12. (32) Aboudheir, A.; Tontiwachwuthikul, P.; Chakma, A.; Idem, R. Kinetics of the reactive absorption of carbon dioxide in high CO2-loaded, concentrated aqueous monoethanolamine solutions. Chem. Eng. Sci. 2003, 58, 5195-5210. (33) Versteeg, G. F.; Van Swaaij, W. P. M. Solubility and diffusivity of acid gases (carbon dioxide, nitrous oxide) in aqueous alkanolamine solutions. J. Chem. Eng. Data 1988, 33, 29-34. (34) Ko, J.-J.; Tsai, T.-C.; Lin, C.-Y.; Wang, H.-M.; Li, M.-H. Diffusivity of Nitrous Oxide in Aqueous Alkanolamine Solutions. J. Chem. Eng. Data 2001, 46, 160-165. (35) Wang, Y. W.; Xu, S.; Otto, F. D.; Mather, A. E. Solubility of N2O in alkanolamines and in mixed solvents. Chem. Eng. J. 1992, 48, 31-40. (36) Kohl, K.; Nielsen, R. Gas Purification, 4th ed.; Gulf Publishing Company: Houston, TX, 1985.

(37) Burns, J. R.; Jamil, J. N.; Ramshaw, C. Process intensification: Operating characteristics of rotating packed beds - determination of liquid hold-up for a high-voidage structured packing. Chem. Eng. Sci. 2000, 55, 2401-2415. (38) Burns, J. R.; Ramshaw, C. Process intensification: visual study of liquid maldistribution in rotating packed beds. Chem. Eng. Sci. 1996, 51, 1347-1352. (39) McCabe, W.; Smith, J.; Harriott, P. Unit Operations of Chemical Engineering, 5th ed.; McGraw-Hill: New York, 1993. (40) Eckert, J. Selecting the Proper Distillation Column Packing. Chem. Eng. Prog. 1970, 66 (3), 39-44 (41) Onda, K.; Sada, E.; Takeuchi, H. Gas absorption with chemical reaction in packed columns. J. Chem. Eng. Jpn. 1968, 1, 62-6. (42) Coulson, J. M.; Richardson, J. F.; Backhurst, J. R.; Harker, J. H. Chem. Eng. 1999, 1 (6), 72. (43) Holland, F.; Bragg, R. Fluid Flow for Chemical and Process Engineers, 2nd ed.; Butterworth-Heinemann: Woburn, MA, 1995.

ReceiVed for reView October 3, 2005 ReVised manuscript receiVed January 30, 2007 Accepted February 16, 2007 IE051104R

Carbon Dioxide Absorption and Desorption in Aqueous

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