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Comparison of the rate of CO2 absorption into aqueous ammonia and monoethanolamine
ARTICLE IN PRESS Chemical Engineering Science 65 (2010) 915–922
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Comparison of the rate of CO2 absorption into aqueous ammonia and monoethanolamine Graeme Puxty, Robert Rowland, Moetaz Attalla CSIRO Energy Technology, PO Box 330, Newcastle NSW 2300, Australia
a r t i c l e in fo
abstract
Article history: Received 13 May 2009 Received in revised form 14 September 2009 Accepted 16 September 2009 Available online 20 September 2009
Aqueous ammonia has been proposed as an absorbent for use in CO2 post combustion capture applications. It has a number of advantages over MEA such as high absorption capacity, low energy requirements for CO2 regeneration and resistance to oxidative and thermal degradation. However, due to its small molecular weight and large vapour pressure absorption must be carried at low temperature to minimise ammonia loss. In this work the rate of CO2 absorption into a falling thin film has been measured using a wetted-wall column for aqueous ammonia between 0.6 and 6 mol L 1, 278–293 K and 0–0.8 liquid CO2 loading. The results were compared to 5 mol L 1 MEA at 303 and 313 K. It was found that the overall mass transfer coefficient for aqueous ammonia was at least 1.5–2 times smaller than MEA at the measured temperatures. From determination of the second-order reaction rate constant k2 (915 L mol 1 s 1 at 283 K) and activation energy Ea (61 kJ mol 1) it was shown that the difference in mass transfer rate is likely due to both the reduced temperature and differences in reactivity between ammonia and MEA with CO2. & 2009 Elsevier Ltd. All rights reserved.
Keywords: Carbon dioxide capture Aqueous ammonia Carbon dioxide absorption rate Wetted-wall
1. Introduction The capture, reversible release and storage of carbon dioxide (CO2) from combustion flue gases (post combustion capture, PCC) is now recognised by government and industry as a viable nearterm option for greenhouse gas abatement (IPCC, 2005; IEA, 2007). It is relevant to electricity generation from fossil fuels (coal, oil and gas) which accounts for approximately 25% of global CO2 emissions (WRI, 2007). This figure has been forecast to increase drastically in the next 25 years (IEA, 2008). PCC has two distinct advantages over other power station CO2 mitigation options such as oxy-firing and integrated gasification combined cycle (IGCC) with pre-combustion capture (Davison, 2007). The first advantage is that being an end-of-pipe technology means it can be retrofitted to existing power stations with minimal modification to the power station itself, or easily integrated into new ones. The second advantage is the ability to dynamically control the energy demand of the PCC plant, allowing additional electricity output to the grid at times of peak load or optimal electricity pricing. PCC technology is also suitable for CO2 capture from other point sources such as steel and cement manufacturing.
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E-mail address: [email protected] (M. Attalla). 0009-2509/$ - see front matter & 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.ces.2009.09.042
The most mature PCC technology is reactive chemical absorption/desorption of CO2 into/from an aqueous alkanolamine absorbent (Davison, 2007). It is a temperature swing process where CO2 is absorbed at low temperature ( 313 K) and released at high temperature ( 393 K), with regenerated absorbent returned to the absorption process. Gas–liquid contacting takes place via packed columns with counter-current gas and liquid flows. The application of PCC to combustion flue gases from electricity generation or other point sources poses a number of technical challenges. The two main issues are the energy requirements of the process and capital cost. The main energy requirements are for heating of the absorbent to release CO2 and the electricity to pump the absorbent around the system. The largest contribution to the capital cost is the cost of materials for and construction of the absorption columns. The size of the absorption columns are defined by the rate of CO2 absorption. The faster the absorption rate, the smaller the gas–liquid contact area required, and thus the smaller the absorption column dimensions required. In an attempt to address these issues, aqueous ammonia (NH3) solutions are now being proposed as an alternative to aqueous alkanolamine based liquid absorbents for PCC. Aqueous ammonia has been shown to achieve higher CO2 loadings (on a molar and mass basis) than sterically free primary alkanolamines such as monoethanolamine (MEA) (Yeh and Bai, 1999). This is mainly due to the CO2–NH3–H2O system favouring bicarbonate over carbamate formation, particularly as CO2 loading increases (Mani et al.,
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2006). Aqueous ammonia has also been shown to require less heat input for desorption than MEA (Yeh et al., 2005). This is due to the smaller reaction enthalpy for CO2 absorption and higher CO2 partial pressure at elevated temperature compared to MEA. Also, ammonia is resistant to oxidative degradation, which is a major benefit when treating oxygen containing gas streams such as those from coal fired power stations. The other main attractive feature is that in the presence of sulphur and nitrogen oxides in the gas stream, the ammonium salts that form may have commercial value as fertilisers. A major drawback in the use of ammonia is its vapour pressure. Due to its small molecular weight ammonia vapour pressure is high (Edwards et al., 1978; Goppert and Maurer, 1998) compared to alkanolamines. To address this it has been proposed that the absorption process take place at lower temperatures to reduce losses via volatilisation. Currently there is no material in the public domain in which the kinetics of CO2 absorption into an aqueous ammonia thin film, which mimics the gas–liquid contacting environment of a packed column, has been presented. More importantly, there is no material in the public domain addressing the impact of lower temperatures on the absorption kinetics. This means it is not possible to accurately size absorber columns and thus predict the capital cost for an aqueous ammonia based CO2 capture process. In this work CO2 absorption rate into aqueous ammonia as a thin film has been measured using a wetted-wall column and compared to MEA. The experiments covered the temperature range 278–293 K and ammonia concentrations of 0.6–6 mol L 1 (1–10% w/w).
Thin liquid film
Thermocou ple
Liquid out
Gas in
Liquid in Fig. 1. Diagram of the wetted-wall column showing the thin liquid film and the thermocouple position.
Gas Out
To Water Bath
2. Background and experimental 2.1. Wetted-wall column A wetted-wall column was used to study the absorption rate of CO2 into aqueous solutions of ammonia and MEA. This type of apparatus allows a flowing gas stream to be contacted with a falling thin liquid film in a counter-current fashion. Such apparatus mimic the gas–liquid contacting that occurs in packed columns where a gas stream enters at the bottom and a liquid stream at the top. The wetted-wall column consisted of a stainless steel column (effective height 8.21 cm, diameter 1.27 cm) over which a thin film of liquid flowed under gravity. Liquid located in a reservoir was pumped up the inside of the column and flowed out the top and down the sides. The liquid flow then returned to the reservoir to form a closed loop. The column was contained within a jacketed glass cover with an internal diameter of 2.54 cm. A mixture of CO2 and N2 was introduced at the bottom of the column and moved upwards, passing the outer column surface before being exhausted from the top. The column gave a known surface area of contact between the liquid and the gas flow as illustrated in Fig. 1. The jacketed cover allowed water from a water bath to control the temperature of the apparatus, as shown in Fig. 2. A T-type thermocouple ( 71 K) was used to measure the temperature of the liquid inside the column. The amount of CO2 absorbed from the gas flow into the liquid film was determined by measuring the CO2 content of the gas entering and exiting the column. This, in conjunction with knowledge of the contact surface area of the liquid film with the gas, allowed determination of the CO2 absorption flux NCO2 . The flux is the number of moles of CO2 absorbed per second per unit area of contact between the liquid and gas. A plot of flux versus applied CO2 partial pressure (logarithmic mean of the CO2
From Water Bath
Fig. 2. A diagram of the wetted-wall column with the jacketed glass cover.
inlet and outlet partial pressures) yields a straight line with a slope equal to the overall mass transfer coefficient KG.
2.1.1. Overall mass transfer The reactive chemical absorption of CO2 into a thin film can be described as a combination of diffusion and chemical reaction processes. The CO2 diffuses from the gas phase, across the gas– liquid interface, into the liquid phase where it undergoes chemical reaction. It is assumed that the film is uniform and the continuous replenishment of the film means there are no long range diffusion processes taking place in the liquid phase between the interface and the bulk liquid (Danckwerts, 1970). The mass transfer processes taking place can be described as a combination of those taking place on the gas side of the interface and on the liquid side, as illustrated in Fig. 3. The concentration of CO2 in the gas phase falls from its bulk gas partial pressure, PCO2 , to its partial pressure at the gas–liquid interface, PCO2;i , according to the gas side mass transfer coefficient kg. [CO2]i represents the dissolved CO2 concentration at the interface. This concentration then falls by diffusion and chemical reaction to the bulk dissolved CO2
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PCO2
rate constant.
Liquid film
Gas film
Bulk gas
rCO2 ¼ k2 ½CO2 ½NH3
Column [CO2]i
[CO2] kl
Assuming significant enhancement of mass transfer by chemical reaction is occurring, absorption occurs according to a pseudo first-order regime if OM0 {Ei where M0 and the infinite enhancement factor Ei are defined according to the following two equations (Danckwerts, 1970).
Gas-liquid interface KG Fig. 3. Illustration showing the gas and liquid mass transfer processes occurring in a segment of the wetted-wall column.
concentration, [CO2], according to the liquid side mass transfer coefficient kl. The liquid side mass transfer coefficient is a function of mass transfer of CO2 without reaction, k01 , and enhancement by chemical reactions occurring in the liquid film that act to consume CO2. These processes are all linked, and the relationships between the absorption flux NCO2 and the overall mass transfer coefficient KG is via the driving force as NCO2 ¼ KG ðPCO2
PCO Þ 2
ð1Þ
is the equilibrium CO2 partial pressure. PCO 2
2.1.2. CO2-Ammonia chemistry The significant liquid phase reactions that are known to occur during reactive chemical absorption of CO2 into an aqueous ammonia solution are summarised in (Mani et al., 2006). þ CO2 þ H2 O"HCO 3 þH
CO2 þ NH3 "NH2 COO þ H þ NH2 COO þ H2 O"NH3 þ HCO 3 NH3 þH þ "NH4þ þ CO2 3 þ H "HCO3
ð3Þ
If CO2 absorption into a thin film is occurring according to a pseudo first-order regime in which the concentration of ammonia is not depleted across the film NCO2 can be described by Eq. (4) (Danckwerts, 1970; Bishnoi and Rochelle, 2000) where DCO2 is the diffusivity of CO2 in the liquid phase. PCO2 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi NCO2 ¼ ð4Þ DCO2 k2 ½NH3 HCO2
PCO2,i kg
Jacket
917
ð2Þ
An equivalent set of reactions also occur for MEA. CO2 reacts with either water to form bicarbonate or ammonia to form a carbamate. The carbamate species may undergo hydrolysis to free ammonia and bicarbonate. At low CO2 loading the carbamate is known to be the dominant species with the equilibria shifting to favour bicarbonate formation at higher loading and lower free ammonia concentration (Mani et al., 2006). This shift allows higher CO2 loadings to be achieved than MEA like systems where CO2 remains predominantly as a carbamate. 2.1.3. Mass transfer with chemical reaction Surface renewal theories of mass transfer with chemical reaction also provide other mathematical relationships that describe NCO2 as a function of the rate of chemical reaction in the liquid phase. It can be assumed that the direct reaction between CO2 and ammonia given in Eq. (2) is much faster than the reaction of CO2 with water, as is the case for primary, secondary and sterically hindered amines (Versteeg et al., 1996). If this is the case the rate limiting step for CO2 absorption is a second-order reaction (subsequent proton exchange can be assumed to be instantaneous). The rate of reaction in the liquid phase rCO2 can then be described by Eq. (3), where k2 is the second-order reaction
M0 ¼
p 4
k2 ½NH3 t
sffiffiffiffiffiffiffiffiffiffiffi sffiffiffiffiffiffiffiffiffiffiffi DCO2 ½NH3 DNH3 þ Ei ¼ DNH3 2½CO2 DCO2
ð5Þ
ð6Þ
In these equations t is the exposure time of the liquid inside the column and DNH3 the diffusivity of ammonia in the liquid phase. 2.2. Experimental setup A gas flow was mixed using two Bronkhurst mass flow controllers. The amount of CO2 and N2 in the gas flow was varied to achieve CO2 partial pressures in the range 0–20 kPa. The gas was passed through a 1/800 stainless steel coil and saturator1 which were both immersed in a water bath (Techne). From the coil and saturator the gas then flowed passed the outside of the wetted-wall column, entering at the bottom and being exhausted at the top. The exhaust gas was then passed through a condenser at 5 1C containing stainless steel wool to increase the surface area followed by a Peltier driven cooler to remove moisture from the gas stream prior to entering a Horiba VA3000 CO2 analyser. The gases used were liquid nitrogen boil off ( 499.99% purity) and 99.5% carbon dioxide, both supplied by BOC Gases Australia. The system was maintained at atmospheric pressure and determination of CO2 partial pressure in the column took into account the water and ammonia content of the gas phase (Fig. 4). 600 mL of the test solution was placed in a Pyrex bottle in the water bath. A large solution volume relative to the gas–liquid contact area was chosen to ensure that during experiments the total CO2 loading of the solution did not change significantly. The solution was pumped in a loop through the wetted-wall column and back to the bottle using a gear pump (Micropump GA-V21) at a rate of 220 mL min 1. This liquid flow rate was chosen to ensure continuous and ripple free film formation. A needle valve and a syringe filled with extra solution were used to provide liquid level control in the system. A full schematic diagram of the complete apparatus is given in Fig. 3. 2.3. Experimental procedure Experiments were carried out by measuring the CO2 content of the inlet and exhaust gas of the wetted-wall column as a function 1 The gas stream was passed through a saturator only when operating above 30 1C. The purpose of this was to minimise evaporative losses from contacting a hot dry gas stream with the thin film. Below 30 1C this was deemed unnecessary due to the small vapour pressure of water below this temperature.
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Saturator
Gas Conditioner
Condenser
N2 CO2 Mass Flow Controllers
Water Bath
CO2 Analyser
Wettedwall Column
Coil Syringe & Needle Valve Sample Reservoir
Pump
Fig. 4. A schematic of the wetted-wall and associated apparatus. Table 1 Overall mass transfer coefficients KG (mmol s 1 m 2 kPa 1) and the y-intercept b (mmol s 1 m 2) from regression of NCO2 versus PCO2 ðNCO2 ¼ KG PCO2 þ bÞ for ammonia (NH3). CO2 loading (mol CO2/mol NH3)
0 0.2 0.4 0.6 0.8
0.6 mol L 1 NH3, 278 K
0.6 mol L 1 NH3, 293 K
3 mol L 1 NH3, 283 K
6 mol L 1 NH3, 278 K
6 mol L 1 NH3, 293 K
KG
b
KG
b
KG
b
KG
b
KG
b
0.30 70.02 0.23 70.01 0.17 70.01 0.13 70.01 0.10 70.01
0.23 0.097 0.039 0.049 0.16
0.27 7 0.02 0.22 7 0.02 0.15 7 0.03 0.14 7 0.01 0.11 7 0.01
0.10 0.014 0.12 0.069 0.35
0.83 70.06 0.59 70.02 0.27 70.03 0.29 70.02 0.19 70.02
0.035 0.091 0.12 0.45 0.57
1.13 70.04 0.72 70.06 0.39 70.02 0.21 70.01 0.21 70.02
0.29 0.87 0.075 0.026 0.24
1.667 0.14 1.097 0.07 0.517 0.07 0.357 0.03 0.22 7 0.01
2.0 0.092 0.060 0.067 0.55
The errors in KG are expressed as one standard deviation.
of solution CO2 loading and CO2 partial pressure. The experimental procedure used was as follows:
1. For ammonia the test solution was made using ammonia solution and preloaded by adding the appropriate amount of ammonium bicarbonate. For MEA the test solution was preloaded to the desired CO2 loading by bubbling pure CO2 through the solution and gravimetrically measuring the amount of CO2 absorbed. 2. The solution was placed in the water bath, the system flushed with nitrogen and allowed to reach thermal equilibrium. 3. The CO2 partial pressure of the inlet gas flow was determined with the gas flow bypassing the wetted-wall column and passing directly to the CO2 analyser. 4. With the solution being pumped through the column, the gas flow was switched from the bypass to the column. 5. Once a steady state was reached (constant CO2 content in the column exhaust gas), the CO2 content of the column exhaust was recorded. 6. The gas flow was changed to the bypass and the flow rates of CO2 and N2 were varied to deliver a different CO2 partial pressure. Steps 3–6 were then repeated until measurements at the required number of CO2 partial pressures had been completed. The CO2 absorption rate of ammonia was studied at the concentrations and temperatures shown in Table 1. This data was also compared to the absorption rate of 5 mol L 1 (30% w/w) MEA at 313 and 333 K. Ammonia solutions were prepared from a 28% w/w (Ajax Univar) aqueous stock solution and ammonium bicarbonate (Sigma Aldrich Z99.0%). MEA solutions were prepared from pure MEA (Sigma Aldrich 99+%). No further purification was carried out and solutions were made with distilled water. For the ammonia concentrations and temperatures listed in Table 1 absorption flux experiments were conducted at liquid CO2
loadings of 0, 0.2, 0.4, 0.6, and 0.8 moles of CO2 per mole of ammonia. For MEA the experiments carried out were at CO2 loadings of 0, 0.1, 0.25, 0.4 and 0.5. At each of these liquid CO2 loadings inlet CO2 partial pressures of 20, 16, 12, 8, 4 and 0 kPa were applied. At the highest CO2 loadings of 0.8 CO2 removal from the gas phase was only 1%. The CO2 partial pressure in the reactor ðPCO2 Þ was calculated as the logarithmic mean of the inlet and outlet CO2 partial pressure. Gas flow rates of 3 L min 1 were used for 0.6 mol L 1 ammonia concentrations and 5 L min 1 for the 3, 6 mol L 1 ammonia and 5 mol L 1 MEA concentrations. The gas path annulus had an inner diameter of 1.27 cm and an outer diameter of 2.54 cm. These gas flow rates were chosen to ensure that gas side mass transfer, kg, did not significantly contribute to overall mass transfer (kg was determined at the conditions used via the method described in Pacheco, 1998) while maintaining measurable CO2 removal. During some of the experiments with ammonia a white solid was formed in the condenser down stream from the wetted-wall column. The solid decomposed at just above room temperature, indicating it was ammonium bicarbonate (ammonium carbonate and carbamate do not decompose at this temperature). It is assumed to have formed from gas phase ammonia and CO2 as water vapour condensed. After a measurement at each CO2 partial pressure the condenser was rinsed with distilled water. The collected solution was titrated with standardised sodium hydroxide to determine the ammonium bicarbonate content. This data was then used to correct the measured CO2 concentration exiting the wetted-wall column by up to 25%. This correction was applied to all CO2 partial pressures for the following test solutions:
0.6 mol L 1 ammonia at 293 K (loadings of 0, 0.2, and 0.4) 3 mol L 1 ammonia at 283 K (loadings of 0, 0.2, 0.4 and 0.6) 6 mol L 1 ammonia at 278 and 293 K (loadings of 0, 0.2 and 0.4) At all other operating conditions no ammonium bicarbonate formation could be detected.
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3. Results and discussion 3.1. Overall mass transfer The CO2 absorption fluxes ðNCO2 Þ and overall mass transfer coefficients (KG) were determined for ammonia at molar CO2 loadings between 0 and 0.8, concentrations between 0.6 and 6 mol L 1 and temperatures between 278 and 293 K. For comparison at industrial conditions, flux values and overall mass transfer coefficients were determined for 5 mol L 1 MEA at molar CO2 loadings between 0.1 and 0.5, and temperatures of 313 and 333 K. is constant for a particular loading so according to Eq. (1), PCO 2 by plotting NCO2 versus applied CO2 partial pressure PCO2 (log mean of the inlet and outlet CO2 partial pressure) KG can be determined by linear regression according to Eq. (7). NCO2 ¼ KG PCO2 þb
ð7Þ
5
919
An example is shown for 0.6 mol L 1 ammonia at 278 K and liquid CO2 loading of 0.2 in Fig. 5. The overall mass transfer coefficients KG and y-intercepts of Eq. (1), determined for ammonia are listed in Table 1 and plotted in Fig. 6. From Fig. 6 it can be seen that KG is dependent on ammonia concentration, CO2 loading, and temperature. KG is most sensitive to increasing ammonia concentration. For unloaded ammonia solutions at 278 K, KG increased from 0.30 to 1.13 mmols 1 m 2 kPa 1 as the ammonia concentration was increased from 0.6 to 6 mol L 1. KG decreases with increasing CO2 loading for the measurements at all ammonia concentrations and temperatures. In Fig. 6 KG is plotted on a log10 scale and the decrease is close to linear with increasing loading. At the highest loading (0.8) values for KG are in the range 0.1–0.2 mmols 1 m 2 kPa 1. This is still an order of magnitude larger than would be expected if only diffusion processes were dominant suggesting enhancement by chemical reaction is still occurring. An increase in KG with temperature was observed for ammonia solutions at 6 mol L 1 when increasing the temperature from 278 to 293 K. Alternatively, the KG values for 0.6 mol L 1 ammonia at
CO2 Absorption Flux, NCO2 (mmol.s-1m-2)
y = 0.2305x - 0.0973
4 Table 2 Overall mass transfer coefficients KG (mmol s 1 m 2 kPa 1) and and the yintercept b (mmol s 1 m 2) from regression of NCO2 versus PCO2 ðNCO2 ¼ KG PCO2 þ bÞ for MEA.
3 2
CO2 loading (mol CO2/mol MEA)
1 0 0
2
4
6
8
10
12
14
16
18
20
-1 PCO2 (kPa) Fig. 5. CO2 absorption flux NCO2 versus applied CO2 partial pressure PCO2 for 0.6 mol L 1 ammonia at 278 K and a liquid CO2 loading of 0.2.
0 0.1 0.25 0.4 0.5
5 mol L 1 MEA, 313 K
5 mol L 1 MEA, 333 K
KG
b
KG
b
2.80 7 0.09 2.34 7 0.03 1.737 0.02 1.207 0.02 0.32 7 0.02
0.48 0.16 0.20 1.2 1.5
3.69 7 0.13 2.89 7 0.05 2.247 0.06 1.257 0.02 0.377 0.05
1.8 0.081 0.92 3.0 5.5
The errors in KG are expressed as one standard deviation.
Fig. 6. Plot of the overall mass transfer coefficient KG on a log10 scale as a function of molar liquid CO2 loading. Data at the same ammonia concentration have the same symbol (square—0.6 mol L 1, triangle—3 mol L 1 and diamond—6 mol L 1) and data at the same temperature have the same colour (light blue—278 K, dark blue—283 K and red—293 K). (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Fig. 7. Plot of the overall mass transfer coefficient KG on a log10 scale as a function of molar CO2 loading. Data at the same ammonia concentration have the same symbol (square—0.6 mol L 1, triangle—3 mol L 1 and diamond—6 mol L 1) and data at the same temperature have the same colour (light blue—278 K, dark blue—283 K and red—293 K). MEA at 5 mol L 1 and 313 and 333 K are shown as dots with thick black dashed and solid lines, respectively. The data in grey are also for 5 mol L 1 MEA at the same conditions from Dugas and Rochelle, 2009. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
both 278 and 293 K are essentially identical. A possible explanation for this is that at this low ammonia concentration there is depletion of ammonia in the liquid film. The consequence of this is that ammonia diffusion starts to play a role in limiting the absorption rate reducing the temperature sensitivity. An important comparison is between the absorption rate of CO2 into ammonia and MEA at representative conditions for the operation of a capture process. This gives an indication of the absorber sizing that would be required for an ammonia process relative to what is known for MEA. For example, a KG half that of MEA indicates twice the gas–liquid contact area would be required to achieve the same amount of CO2 removal from a gas stream at identical driving force. KG determined for 5 mol L 1 MEA at 313 and 333 K are represented in Table 2 and Fig. 7. Comparison with MEA data from Dugas and Rochelle (2009) is also given in the figure. The values of KG for MEA over the CO2 loading range of 0.1– 0.4 are 1.5–2 times greater than those of 6 mol L 1 ammonia at 293 K over the same loading range. This means that if a 6 mol L 1 ammonia solution at 293 K was used, the gas–liquid contact area would need to be 1.5–2 times larger than that of an MEA based process. The loading regime in which a capture plant is operated will also have an impact. MEA plants generally operate with lean/rich loadings2 around 0.2/0.5 (Kohl and Nielsen, 1997). If an ammonia plant is operated with higher lean/rich loadings (e.g. 0.3/0.6), to maintain a predominantly bicarbonate based system to minimise desorption energy requirements and ammonia volatilisation, the absorber sizing will need to be significantly larger than for an MEA process. Also, use of a temperature below 293 K or ammonia concentration below 6 mol L 1 will also require a larger packing
2
The lean/rich loading regime refers to the ‘‘lean’’ CO2 loading at the base of the desorption column and the ‘‘rich’’ CO2 loading at the base of the absorption column.
Table 3 The second-order reaction rate constant kref,2 and the activation energy Ea,2 for the reaction between CO2 and ammonia. kref,2 (L mol 1 s 1) at Tref,2 = 283 K
Ea,2 (kJ mol 1)
915
61
area and thus a larger absorber relative to MEA to achieve the same amount of CO2 removal. 3.2. Chemical reaction rate CO2 absorption flux measurements into unloaded aqueous ammonia were used to estimate the second-order reaction rate constant k2 using Eq. (4). The diffusivity and physical solubility (Henry’s constant) of CO2 in aqueous ammonia, DCO2 ðm2 s1 Þ and 1 HCO2 ðkPa m3 kmol Þ respectively, were assumed equivalent to that in water and were calculated using the following equations (Versteeg et al., 1996). DCO2 ¼ 2:35 106 e2199=T m2 s1
ð8Þ
HCO2 ¼ 2:82 106 e2044=T Pa m3 mol
1
ð9Þ
where T= temperature (K) Eqs. (5) and (6) were used to verify that the CO2 absorption measurements were carried out under a pseudo first-order regime. The diffusivity of ammonia, DNH3 ðm2 s1 Þ, was calculated according to Eq. (10) (Frank et al., 1996) and the exposure time of the liquid in the column was 0.2 s. DNH3 ¼ ð1:65 þ2:47XNH3 Þ 106 e1996:6=T m2 s1
ð10Þ
The condition that OM0 {Ei was only satisfied for the ammonia concentrations of 3 and 6 mol L 1, with Ei being at least 10 times
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CO2 Absorption Flux, NCO2 (mmol.s-1m-2)
G. Puxty et al. / Chemical Engineering Science 65 (2010) 915–922
921
35 6 mol.L-1 NH3 at 293 K
30 25
6 mol.L-1 NH3 at 278 K
20 15
3 mol.L-1 NH3 at 283 K
10 5 0 0
2
4
6
8
10 12 PCO2 (kPa)
14
16
18
20
Fig. 8. Plot of the measured flux (symbols) and calculated flux (lines) using the optimised values of kref,2 and Ea given in Table 3, Eqs. (4) and (11).
larger than OM0 . So only these concentrations were using for estimation of k2. This also supports the supposition that for the 0.6 mol L 1 measurements depletion of ammonia in the film is occurring and the absorption rate is being limited by ammonia diffusion. The value of k2 was determined using non-linear regression (Excels Solver) to fit Eq. (4) to the measured flux data. The temperature dependence of k2 was described using a reparameterised form of the Arrhenius equation (Furusjo¨ et al., 2003). 1 1
k2 ¼ kref;2 eEa =Rðð1=TÞð1=Tref;2 ÞÞ L mol
s
ð11Þ
where Ea is the energy of activation (J mol 1), R the gas constant (8.314 J K 1 mol 1) and T/Tref,2 the temperature (K). In this form of the Arrhenius equation kref,2 represents the value of k2 at the reference temperature Tref. For the purposes of non-linear regression this form of the equation reduces the correlation between the pre-exponential factor and Ea resulting in more robust optimisation. The values of kref,2 and Ea were optimised to minimise the error in the least-squares sense between the measured and calculated absorption flux with the final values given in Table 3. The measured data and the results from the regression are in excellent agreement, as shown in Fig. 8. Inclusion of reaction with hydroxide had a negligible effect on the determined parameter values. These results should be considered indicative only as the CO2 solubility and diffusivity in aqueous ammonia were estimated from those in water and only a small number of concentrations and temperatures were considered. Table 4 shows a comparison between k2 and literature data at three temperatures. The values are larger than those determined by Pinsent et al. (1956) probably due to the higher ammonia concentrations used in this work. Comparison with the results of Derks and Versteeg (2009) is difficult as their mechanistic interpretation means the reaction rate constants are concentration dependent. However, at an ammonia concentration of 1 mol L 1 direct comparison can be made and the agreement is good. To allow comparison with MEA at the same temperature wettedwall measurements were also made using CO2 free 5 mol L 1 MEA at 283 K. The value of k2 determined from this data, again using CO2 solubility and diffusivity in water, was 2634 L mol 1 s 1. This is in good agreement with the literature value 2273 L mol 1 s 1 extrapolated using Eq. (12) (Versteeg et al., 1996) or 2042 L mol 1 s 1
Table 4 The second-order reaction rate constant k2 from this work and literature sources. Temperature (K)
k2 (L mol 1 s 1) This work Pinsent et al. (1956) Derks and Versteeg (2009)a
278 283 293
574 915 2217
107 155 313
300 700 1400
a The reaction rate constants from this reference are concentration dependent. The values given are for 1 mol L 1 ammonia as at this concentration the equations used collapse to those in this work.
calculated from Eq. (13) (Alper, 1990) which covers the temperature range 278–298 K. If the results from Aboudheir et al. (2003) are extrapolated to 283 K, and values for kNH2 and kH2 O determined assuming a termolecular mechanism are converted to the appropriate second-order reaction rate constant, a value of 5507 mol L 1 s 1 is obtained. This deviation could be due to the temperature extrapolation or the different models used to interpret the data. k2 ¼ 4:4 1011 e5440=T L mol
1 1
s
k2 ¼ 8:51 1011 e5617=T L mol
1 1
s
ð12Þ ð13Þ
At a comparable temperature of 283 K the value of k2 for aqueous ammonia is approximately 2.8 times smaller than that for MEA. This indicates that the CO2 absorption rate into aqueous ammonia is smaller than that into MEA due to both lower operating temperature and lower overall reactivity towards CO2.
4. Conclusions In this work the CO2 absorption flux NCO2 into aqueous ammonia and MEA solutions has been compared. The CO2 absorption flux of both systems has been measured as a function of CO2 loading, temperature and CO2 partial pressure. A wettedwall column was used to measure the absorption into a falling thin film and results were expressed as overall mass transfer coefficient KG values. Measurements into unloaded aqueous ammonia were also used to estimate the second-order reaction rate constant between CO2 and ammonia. This is the first such
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study to provide relevant rate information for CO2 absorption into an aqueous ammonia absorbent. As expected for a chemical absorbent combining diffusion and fast chemical reaction, the CO2 absorption flux fell as the liquid CO2 loading increased. KG increased with increasing temperature over the range 278–293 K for the 6 mol L 1 ammonia solution when operating under a regime of mass transfer enhanced by chemical reaction. The same temperature dependence was not seen for the 0.6 mol L 1 ammonia solution probably due to depletion of ammonia in the liquid film. KG was most sensitive to ammonia concentration, with a 3–6 fold increase in when moving from a 0.6 to a 6 mol L 1 ammonia solution. The results for ammonia were also compared to those of a 5 mol L 1 MEA solution at 313 and 333 K. In general the values of KG for the MEA solution were 1.5–2 times greater than the 6 mol L 1 ammonia at 293 K. At temperatures and concentrations below this the value of KG for ammonia was up to 7 times smaller than MEA. Absorption into unloaded 3 and 6 mol L 1 aqueous ammonia at 278, 283 and 293 K were used to estimate the secondorder reaction rate k2 between CO2 and ammonia. Comparison with MEA data at 283 K suggests that the smaller values of KG for aqueous ammonia are driven by both lower temperature and lower reactivity towards CO2. In general, an aqueous ammonia based process will require a larger gas–liquid contact area and thus a larger absorber column than a traditional MEA based process to achieve the same amount of CO2 removal from a gas stream. Whether the size increase is marginal or significant depends on a combination of the temperature, concentration and lean/rich loading regime used for the process.
Notation b DCO2 DNH3 Ea,2 Ei HCO2 k2 kref,2 kg kl1 kl KG M0 NCO2 PCO2 PCO2;i PCO 2 rCO2 R t T Tref,2 [CO2]
y-intercept of flux versus loading linear regression, mmol s 1 m 2 diffusivity of CO2 in water, m2 s 1 diffusivity of NH3 in water, m2 s 1 activation energy for second-order reaction, kJ mol 1 infinite enhancement factor Henry’s constant for CO2 in water, Pa m3 mol 1 second-order reaction rate constant, L mol 1 s 1 second-order reaction rate constant at reference temperature Tref,2, L mol 1 s 1 gas-side mass transfer coefficient, mmol s 1 m 2 kPa 1 liquid-side mass transfer coefficient without reaction, mmol s 1 m 2 liquid-side mass transfer coefficient with reaction, mmol s 1 m 2 overall mass transfer coefficient, mmol s 1 m 2 kPa 1 dimensionless group CO2 absorption flux, mmol s 1 m 2 applied CO2 partial pressure, kPa interfacial CO2 partial pressure, kPa equilibrium CO2 partial pressure, kPa CO2 reaction rate, mol L 1 s 1 gas constant, 8.314 J K 1 mol 1 exposure time of liquid in wetted-wall column, s temperature, K reference temperature for kref,2, K liquid phase CO2 concentration, mol L 1
[CO2]i interfacial liquid phase CO2 concentration, mol L 1 [NH3] liquid phase NH3 concentration, mol L 1
Acknowledgements This work was carried out within CSIRO’s Energy Transformed Flagship Research stream on Post Combustion Capture and was supported by the Australian Government through the Asia Pacific Partnership on Clean Development and Climate. References Aboudheir, A., Tontiwachwuthikul, P., Chakma, A., Idem, R., 2003. Kinetics of the reactive absorption of carbon dioxide in high CO2-loaded, concentrated aqueous monoethanolamine solutions. Chemical Engineering Science 58, 5196–5210. Alper, E., 1990. Reaction mechanism and kinetics of aqueous solutions of 2-Amino2-methyl-1-propanol and carbon dioxide. Industrial & Engineering Chemistry Research 29, 1725–1728. Bishnoi, S., Rochelle, G.T., 2000. Absorption of carbon dioxide into aqueous piperazine: reaction kinetics, mass transfer and solubility. Chemical Engineering Science 55, 5531–5543. Danckwerts, P.V., 1970. Gas–Liquid Reactions. McGraw-Hill, New York. Davison, J., 2007. Performance and costs of power plants with capture and storage of CO2. Energy 32, 1163–1176. Derks, P.W.J., Versteeg, G.F., 2009. Kinetics of absorption of carbon dioxide in aqueous ammonia solutions. Energy Procedia 1, 1139–1146. Dugas, R., Rochelle, G., 2009. Absorption and desorption rates of carbon dioxide with monoethanolamine and piperazine. Energy Procedia 1, 1163–1169. Edwards, T.J., Maurer, G., Newman, J., Prausnitz, J.M., 1978. Vapor–liquid equilibria in multicomponent aqueous solutions of volatile weak electrolytes. A.I.Ch.E. Journal 24 (6), 966–976. Frank, M.J.W., Kuipers, J.A.M., Swaaij, W.P.M.v., 1996. Diffusion coefficients and viscosities of CO2 + H2O, CO2 + CH3OH, NH3 + H2O, and NH3 + CH3OH liquid mixtures. Journal of Chemical & Engineering Data 41, 297–302. ¨ E., Svenssen, O., Danielsson, L.-G., 2003. Estimation of kinetic paraFurusjo, meters from non-isothermal batch experiments monitored by in situ vibrational spectroscopy. Chemometrics and Intelligent Laboratory Systems 66, 1–14. Goppert, U., Maurer, G., 1998. Vapour–liquid equilibria in aqueous solutions of ammonia and carbon dioxide at temperatures between 333 and 393 K and pressures up to 7 MPa. Fluid Phase Equilibria 41, 153–185. IEA, 2007. Capturing CO2. In: Adams, D., Davison, J. (Eds.), Greenhouse Gas R&D Program: 23, IEA, Cheltenham. IEA, 2008. World Energy Outlook 2007. Paris, France. IPCC, 2005. IPCC special report on carbon dioxide capture and storage. In: Metz, B., Davidson, O., Coninck, H.C.d., Loos, M., Meyer, L.A. (Eds.), Working Group III of the Intergovernmental Panel on Climate Change, Cambridge University Press, Cambridge. Kohl, A., Nielsen, R., 1997. Gas Purification. Gulf Publishing Company, Houston, Texas. Mani, F., Peruzzini, M., Stoppioni, P., 2006. CO2 absorption by aqueous NH3 solutions: speciation of ammonium carbamate, bicarbonate and carbonate by a 13C NMR study. Green Chemistry 8, 995–1000. Pacheco, M.A., 1998. Mass transfer, kinetics and rate based modelling of reactive absorption. Ph.D. Thesis, Chemical Engineering, Austin, The University of Texas. Pinsent, B.R.W., Pearson, L., Roughton, F.J.W., 1956. The kinetics of combination of carbon dioxide with ammonia. Transactions of the Faraday Society 52, 1594–1598. Versteeg, G.F., Dijck, L.A.J.V., Swaaij, W.P.M.V., 1996. On the kinetics between CO2 and alkanolamines both in aqueous and non-aqueous solutions. An overview. Chemical Engineering Communications 144, 113–158. WRI, 2007. Climate analysis indicators tool. 2007. /http://www.wri.org/S. Yeh, A.C., Bai, H., 1999. Comparison of ammonia and monoethanolamine solvents to reduce CO2 greenhouse gas emissions. Science of the Total Environment 228, 121–133. Yeh, J.T., Resnik, K.P., Rygle, K., Pennline, H.W., 2005. Semi-batch absorption and regeneration studies for CO2 capture by aqueous ammonia. Fuel Process Technology 86, 1533–1546.
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Comparison of the rate of CO2 absorption into aqueous ammonia and monoethanolamine Graeme Puxty, Robert Rowland, Moetaz Attalla CSIRO Energy Technology, PO Box 330, Newcastle NSW 2300, Australia
a r t i c l e in fo
abstract
Article history: Received 13 May 2009 Received in revised form 14 September 2009 Accepted 16 September 2009 Available online 20 September 2009
Aqueous ammonia has been proposed as an absorbent for use in CO2 post combustion capture applications. It has a number of advantages over MEA such as high absorption capacity, low energy requirements for CO2 regeneration and resistance to oxidative and thermal degradation. However, due to its small molecular weight and large vapour pressure absorption must be carried at low temperature to minimise ammonia loss. In this work the rate of CO2 absorption into a falling thin film has been measured using a wetted-wall column for aqueous ammonia between 0.6 and 6 mol L 1, 278–293 K and 0–0.8 liquid CO2 loading. The results were compared to 5 mol L 1 MEA at 303 and 313 K. It was found that the overall mass transfer coefficient for aqueous ammonia was at least 1.5–2 times smaller than MEA at the measured temperatures. From determination of the second-order reaction rate constant k2 (915 L mol 1 s 1 at 283 K) and activation energy Ea (61 kJ mol 1) it was shown that the difference in mass transfer rate is likely due to both the reduced temperature and differences in reactivity between ammonia and MEA with CO2. & 2009 Elsevier Ltd. All rights reserved.
Keywords: Carbon dioxide capture Aqueous ammonia Carbon dioxide absorption rate Wetted-wall
1. Introduction The capture, reversible release and storage of carbon dioxide (CO2) from combustion flue gases (post combustion capture, PCC) is now recognised by government and industry as a viable nearterm option for greenhouse gas abatement (IPCC, 2005; IEA, 2007). It is relevant to electricity generation from fossil fuels (coal, oil and gas) which accounts for approximately 25% of global CO2 emissions (WRI, 2007). This figure has been forecast to increase drastically in the next 25 years (IEA, 2008). PCC has two distinct advantages over other power station CO2 mitigation options such as oxy-firing and integrated gasification combined cycle (IGCC) with pre-combustion capture (Davison, 2007). The first advantage is that being an end-of-pipe technology means it can be retrofitted to existing power stations with minimal modification to the power station itself, or easily integrated into new ones. The second advantage is the ability to dynamically control the energy demand of the PCC plant, allowing additional electricity output to the grid at times of peak load or optimal electricity pricing. PCC technology is also suitable for CO2 capture from other point sources such as steel and cement manufacturing.
Corresponding author. Tel.: + 61 2 4960 6083; fax: + 61 2 4960 6054.
E-mail address: [email protected] (M. Attalla). 0009-2509/$ - see front matter & 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.ces.2009.09.042
The most mature PCC technology is reactive chemical absorption/desorption of CO2 into/from an aqueous alkanolamine absorbent (Davison, 2007). It is a temperature swing process where CO2 is absorbed at low temperature ( 313 K) and released at high temperature ( 393 K), with regenerated absorbent returned to the absorption process. Gas–liquid contacting takes place via packed columns with counter-current gas and liquid flows. The application of PCC to combustion flue gases from electricity generation or other point sources poses a number of technical challenges. The two main issues are the energy requirements of the process and capital cost. The main energy requirements are for heating of the absorbent to release CO2 and the electricity to pump the absorbent around the system. The largest contribution to the capital cost is the cost of materials for and construction of the absorption columns. The size of the absorption columns are defined by the rate of CO2 absorption. The faster the absorption rate, the smaller the gas–liquid contact area required, and thus the smaller the absorption column dimensions required. In an attempt to address these issues, aqueous ammonia (NH3) solutions are now being proposed as an alternative to aqueous alkanolamine based liquid absorbents for PCC. Aqueous ammonia has been shown to achieve higher CO2 loadings (on a molar and mass basis) than sterically free primary alkanolamines such as monoethanolamine (MEA) (Yeh and Bai, 1999). This is mainly due to the CO2–NH3–H2O system favouring bicarbonate over carbamate formation, particularly as CO2 loading increases (Mani et al.,
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2006). Aqueous ammonia has also been shown to require less heat input for desorption than MEA (Yeh et al., 2005). This is due to the smaller reaction enthalpy for CO2 absorption and higher CO2 partial pressure at elevated temperature compared to MEA. Also, ammonia is resistant to oxidative degradation, which is a major benefit when treating oxygen containing gas streams such as those from coal fired power stations. The other main attractive feature is that in the presence of sulphur and nitrogen oxides in the gas stream, the ammonium salts that form may have commercial value as fertilisers. A major drawback in the use of ammonia is its vapour pressure. Due to its small molecular weight ammonia vapour pressure is high (Edwards et al., 1978; Goppert and Maurer, 1998) compared to alkanolamines. To address this it has been proposed that the absorption process take place at lower temperatures to reduce losses via volatilisation. Currently there is no material in the public domain in which the kinetics of CO2 absorption into an aqueous ammonia thin film, which mimics the gas–liquid contacting environment of a packed column, has been presented. More importantly, there is no material in the public domain addressing the impact of lower temperatures on the absorption kinetics. This means it is not possible to accurately size absorber columns and thus predict the capital cost for an aqueous ammonia based CO2 capture process. In this work CO2 absorption rate into aqueous ammonia as a thin film has been measured using a wetted-wall column and compared to MEA. The experiments covered the temperature range 278–293 K and ammonia concentrations of 0.6–6 mol L 1 (1–10% w/w).
Thin liquid film
Thermocou ple
Liquid out
Gas in
Liquid in Fig. 1. Diagram of the wetted-wall column showing the thin liquid film and the thermocouple position.
Gas Out
To Water Bath
2. Background and experimental 2.1. Wetted-wall column A wetted-wall column was used to study the absorption rate of CO2 into aqueous solutions of ammonia and MEA. This type of apparatus allows a flowing gas stream to be contacted with a falling thin liquid film in a counter-current fashion. Such apparatus mimic the gas–liquid contacting that occurs in packed columns where a gas stream enters at the bottom and a liquid stream at the top. The wetted-wall column consisted of a stainless steel column (effective height 8.21 cm, diameter 1.27 cm) over which a thin film of liquid flowed under gravity. Liquid located in a reservoir was pumped up the inside of the column and flowed out the top and down the sides. The liquid flow then returned to the reservoir to form a closed loop. The column was contained within a jacketed glass cover with an internal diameter of 2.54 cm. A mixture of CO2 and N2 was introduced at the bottom of the column and moved upwards, passing the outer column surface before being exhausted from the top. The column gave a known surface area of contact between the liquid and the gas flow as illustrated in Fig. 1. The jacketed cover allowed water from a water bath to control the temperature of the apparatus, as shown in Fig. 2. A T-type thermocouple ( 71 K) was used to measure the temperature of the liquid inside the column. The amount of CO2 absorbed from the gas flow into the liquid film was determined by measuring the CO2 content of the gas entering and exiting the column. This, in conjunction with knowledge of the contact surface area of the liquid film with the gas, allowed determination of the CO2 absorption flux NCO2 . The flux is the number of moles of CO2 absorbed per second per unit area of contact between the liquid and gas. A plot of flux versus applied CO2 partial pressure (logarithmic mean of the CO2
From Water Bath
Fig. 2. A diagram of the wetted-wall column with the jacketed glass cover.
inlet and outlet partial pressures) yields a straight line with a slope equal to the overall mass transfer coefficient KG.
2.1.1. Overall mass transfer The reactive chemical absorption of CO2 into a thin film can be described as a combination of diffusion and chemical reaction processes. The CO2 diffuses from the gas phase, across the gas– liquid interface, into the liquid phase where it undergoes chemical reaction. It is assumed that the film is uniform and the continuous replenishment of the film means there are no long range diffusion processes taking place in the liquid phase between the interface and the bulk liquid (Danckwerts, 1970). The mass transfer processes taking place can be described as a combination of those taking place on the gas side of the interface and on the liquid side, as illustrated in Fig. 3. The concentration of CO2 in the gas phase falls from its bulk gas partial pressure, PCO2 , to its partial pressure at the gas–liquid interface, PCO2;i , according to the gas side mass transfer coefficient kg. [CO2]i represents the dissolved CO2 concentration at the interface. This concentration then falls by diffusion and chemical reaction to the bulk dissolved CO2
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PCO2
rate constant.
Liquid film
Gas film
Bulk gas
rCO2 ¼ k2 ½CO2 ½NH3
Column [CO2]i
[CO2] kl
Assuming significant enhancement of mass transfer by chemical reaction is occurring, absorption occurs according to a pseudo first-order regime if OM0 {Ei where M0 and the infinite enhancement factor Ei are defined according to the following two equations (Danckwerts, 1970).
Gas-liquid interface KG Fig. 3. Illustration showing the gas and liquid mass transfer processes occurring in a segment of the wetted-wall column.
concentration, [CO2], according to the liquid side mass transfer coefficient kl. The liquid side mass transfer coefficient is a function of mass transfer of CO2 without reaction, k01 , and enhancement by chemical reactions occurring in the liquid film that act to consume CO2. These processes are all linked, and the relationships between the absorption flux NCO2 and the overall mass transfer coefficient KG is via the driving force as NCO2 ¼ KG ðPCO2
PCO Þ 2
ð1Þ
is the equilibrium CO2 partial pressure. PCO 2
2.1.2. CO2-Ammonia chemistry The significant liquid phase reactions that are known to occur during reactive chemical absorption of CO2 into an aqueous ammonia solution are summarised in (Mani et al., 2006). þ CO2 þ H2 O"HCO 3 þH
CO2 þ NH3 "NH2 COO þ H þ NH2 COO þ H2 O"NH3 þ HCO 3 NH3 þH þ "NH4þ þ CO2 3 þ H "HCO3
ð3Þ
If CO2 absorption into a thin film is occurring according to a pseudo first-order regime in which the concentration of ammonia is not depleted across the film NCO2 can be described by Eq. (4) (Danckwerts, 1970; Bishnoi and Rochelle, 2000) where DCO2 is the diffusivity of CO2 in the liquid phase. PCO2 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi NCO2 ¼ ð4Þ DCO2 k2 ½NH3 HCO2
PCO2,i kg
Jacket
917
ð2Þ
An equivalent set of reactions also occur for MEA. CO2 reacts with either water to form bicarbonate or ammonia to form a carbamate. The carbamate species may undergo hydrolysis to free ammonia and bicarbonate. At low CO2 loading the carbamate is known to be the dominant species with the equilibria shifting to favour bicarbonate formation at higher loading and lower free ammonia concentration (Mani et al., 2006). This shift allows higher CO2 loadings to be achieved than MEA like systems where CO2 remains predominantly as a carbamate. 2.1.3. Mass transfer with chemical reaction Surface renewal theories of mass transfer with chemical reaction also provide other mathematical relationships that describe NCO2 as a function of the rate of chemical reaction in the liquid phase. It can be assumed that the direct reaction between CO2 and ammonia given in Eq. (2) is much faster than the reaction of CO2 with water, as is the case for primary, secondary and sterically hindered amines (Versteeg et al., 1996). If this is the case the rate limiting step for CO2 absorption is a second-order reaction (subsequent proton exchange can be assumed to be instantaneous). The rate of reaction in the liquid phase rCO2 can then be described by Eq. (3), where k2 is the second-order reaction
M0 ¼
p 4
k2 ½NH3 t
sffiffiffiffiffiffiffiffiffiffiffi sffiffiffiffiffiffiffiffiffiffiffi DCO2 ½NH3 DNH3 þ Ei ¼ DNH3 2½CO2 DCO2
ð5Þ
ð6Þ
In these equations t is the exposure time of the liquid inside the column and DNH3 the diffusivity of ammonia in the liquid phase. 2.2. Experimental setup A gas flow was mixed using two Bronkhurst mass flow controllers. The amount of CO2 and N2 in the gas flow was varied to achieve CO2 partial pressures in the range 0–20 kPa. The gas was passed through a 1/800 stainless steel coil and saturator1 which were both immersed in a water bath (Techne). From the coil and saturator the gas then flowed passed the outside of the wetted-wall column, entering at the bottom and being exhausted at the top. The exhaust gas was then passed through a condenser at 5 1C containing stainless steel wool to increase the surface area followed by a Peltier driven cooler to remove moisture from the gas stream prior to entering a Horiba VA3000 CO2 analyser. The gases used were liquid nitrogen boil off ( 499.99% purity) and 99.5% carbon dioxide, both supplied by BOC Gases Australia. The system was maintained at atmospheric pressure and determination of CO2 partial pressure in the column took into account the water and ammonia content of the gas phase (Fig. 4). 600 mL of the test solution was placed in a Pyrex bottle in the water bath. A large solution volume relative to the gas–liquid contact area was chosen to ensure that during experiments the total CO2 loading of the solution did not change significantly. The solution was pumped in a loop through the wetted-wall column and back to the bottle using a gear pump (Micropump GA-V21) at a rate of 220 mL min 1. This liquid flow rate was chosen to ensure continuous and ripple free film formation. A needle valve and a syringe filled with extra solution were used to provide liquid level control in the system. A full schematic diagram of the complete apparatus is given in Fig. 3. 2.3. Experimental procedure Experiments were carried out by measuring the CO2 content of the inlet and exhaust gas of the wetted-wall column as a function 1 The gas stream was passed through a saturator only when operating above 30 1C. The purpose of this was to minimise evaporative losses from contacting a hot dry gas stream with the thin film. Below 30 1C this was deemed unnecessary due to the small vapour pressure of water below this temperature.
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Saturator
Gas Conditioner
Condenser
N2 CO2 Mass Flow Controllers
Water Bath
CO2 Analyser
Wettedwall Column
Coil Syringe & Needle Valve Sample Reservoir
Pump
Fig. 4. A schematic of the wetted-wall and associated apparatus. Table 1 Overall mass transfer coefficients KG (mmol s 1 m 2 kPa 1) and the y-intercept b (mmol s 1 m 2) from regression of NCO2 versus PCO2 ðNCO2 ¼ KG PCO2 þ bÞ for ammonia (NH3). CO2 loading (mol CO2/mol NH3)
0 0.2 0.4 0.6 0.8
0.6 mol L 1 NH3, 278 K
0.6 mol L 1 NH3, 293 K
3 mol L 1 NH3, 283 K
6 mol L 1 NH3, 278 K
6 mol L 1 NH3, 293 K
KG
b
KG
b
KG
b
KG
b
KG
b
0.30 70.02 0.23 70.01 0.17 70.01 0.13 70.01 0.10 70.01
0.23 0.097 0.039 0.049 0.16
0.27 7 0.02 0.22 7 0.02 0.15 7 0.03 0.14 7 0.01 0.11 7 0.01
0.10 0.014 0.12 0.069 0.35
0.83 70.06 0.59 70.02 0.27 70.03 0.29 70.02 0.19 70.02
0.035 0.091 0.12 0.45 0.57
1.13 70.04 0.72 70.06 0.39 70.02 0.21 70.01 0.21 70.02
0.29 0.87 0.075 0.026 0.24
1.667 0.14 1.097 0.07 0.517 0.07 0.357 0.03 0.22 7 0.01
2.0 0.092 0.060 0.067 0.55
The errors in KG are expressed as one standard deviation.
of solution CO2 loading and CO2 partial pressure. The experimental procedure used was as follows:
1. For ammonia the test solution was made using ammonia solution and preloaded by adding the appropriate amount of ammonium bicarbonate. For MEA the test solution was preloaded to the desired CO2 loading by bubbling pure CO2 through the solution and gravimetrically measuring the amount of CO2 absorbed. 2. The solution was placed in the water bath, the system flushed with nitrogen and allowed to reach thermal equilibrium. 3. The CO2 partial pressure of the inlet gas flow was determined with the gas flow bypassing the wetted-wall column and passing directly to the CO2 analyser. 4. With the solution being pumped through the column, the gas flow was switched from the bypass to the column. 5. Once a steady state was reached (constant CO2 content in the column exhaust gas), the CO2 content of the column exhaust was recorded. 6. The gas flow was changed to the bypass and the flow rates of CO2 and N2 were varied to deliver a different CO2 partial pressure. Steps 3–6 were then repeated until measurements at the required number of CO2 partial pressures had been completed. The CO2 absorption rate of ammonia was studied at the concentrations and temperatures shown in Table 1. This data was also compared to the absorption rate of 5 mol L 1 (30% w/w) MEA at 313 and 333 K. Ammonia solutions were prepared from a 28% w/w (Ajax Univar) aqueous stock solution and ammonium bicarbonate (Sigma Aldrich Z99.0%). MEA solutions were prepared from pure MEA (Sigma Aldrich 99+%). No further purification was carried out and solutions were made with distilled water. For the ammonia concentrations and temperatures listed in Table 1 absorption flux experiments were conducted at liquid CO2
loadings of 0, 0.2, 0.4, 0.6, and 0.8 moles of CO2 per mole of ammonia. For MEA the experiments carried out were at CO2 loadings of 0, 0.1, 0.25, 0.4 and 0.5. At each of these liquid CO2 loadings inlet CO2 partial pressures of 20, 16, 12, 8, 4 and 0 kPa were applied. At the highest CO2 loadings of 0.8 CO2 removal from the gas phase was only 1%. The CO2 partial pressure in the reactor ðPCO2 Þ was calculated as the logarithmic mean of the inlet and outlet CO2 partial pressure. Gas flow rates of 3 L min 1 were used for 0.6 mol L 1 ammonia concentrations and 5 L min 1 for the 3, 6 mol L 1 ammonia and 5 mol L 1 MEA concentrations. The gas path annulus had an inner diameter of 1.27 cm and an outer diameter of 2.54 cm. These gas flow rates were chosen to ensure that gas side mass transfer, kg, did not significantly contribute to overall mass transfer (kg was determined at the conditions used via the method described in Pacheco, 1998) while maintaining measurable CO2 removal. During some of the experiments with ammonia a white solid was formed in the condenser down stream from the wetted-wall column. The solid decomposed at just above room temperature, indicating it was ammonium bicarbonate (ammonium carbonate and carbamate do not decompose at this temperature). It is assumed to have formed from gas phase ammonia and CO2 as water vapour condensed. After a measurement at each CO2 partial pressure the condenser was rinsed with distilled water. The collected solution was titrated with standardised sodium hydroxide to determine the ammonium bicarbonate content. This data was then used to correct the measured CO2 concentration exiting the wetted-wall column by up to 25%. This correction was applied to all CO2 partial pressures for the following test solutions:
0.6 mol L 1 ammonia at 293 K (loadings of 0, 0.2, and 0.4) 3 mol L 1 ammonia at 283 K (loadings of 0, 0.2, 0.4 and 0.6) 6 mol L 1 ammonia at 278 and 293 K (loadings of 0, 0.2 and 0.4) At all other operating conditions no ammonium bicarbonate formation could be detected.
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3. Results and discussion 3.1. Overall mass transfer The CO2 absorption fluxes ðNCO2 Þ and overall mass transfer coefficients (KG) were determined for ammonia at molar CO2 loadings between 0 and 0.8, concentrations between 0.6 and 6 mol L 1 and temperatures between 278 and 293 K. For comparison at industrial conditions, flux values and overall mass transfer coefficients were determined for 5 mol L 1 MEA at molar CO2 loadings between 0.1 and 0.5, and temperatures of 313 and 333 K. is constant for a particular loading so according to Eq. (1), PCO 2 by plotting NCO2 versus applied CO2 partial pressure PCO2 (log mean of the inlet and outlet CO2 partial pressure) KG can be determined by linear regression according to Eq. (7). NCO2 ¼ KG PCO2 þb
ð7Þ
5
919
An example is shown for 0.6 mol L 1 ammonia at 278 K and liquid CO2 loading of 0.2 in Fig. 5. The overall mass transfer coefficients KG and y-intercepts of Eq. (1), determined for ammonia are listed in Table 1 and plotted in Fig. 6. From Fig. 6 it can be seen that KG is dependent on ammonia concentration, CO2 loading, and temperature. KG is most sensitive to increasing ammonia concentration. For unloaded ammonia solutions at 278 K, KG increased from 0.30 to 1.13 mmols 1 m 2 kPa 1 as the ammonia concentration was increased from 0.6 to 6 mol L 1. KG decreases with increasing CO2 loading for the measurements at all ammonia concentrations and temperatures. In Fig. 6 KG is plotted on a log10 scale and the decrease is close to linear with increasing loading. At the highest loading (0.8) values for KG are in the range 0.1–0.2 mmols 1 m 2 kPa 1. This is still an order of magnitude larger than would be expected if only diffusion processes were dominant suggesting enhancement by chemical reaction is still occurring. An increase in KG with temperature was observed for ammonia solutions at 6 mol L 1 when increasing the temperature from 278 to 293 K. Alternatively, the KG values for 0.6 mol L 1 ammonia at
CO2 Absorption Flux, NCO2 (mmol.s-1m-2)
y = 0.2305x - 0.0973
4 Table 2 Overall mass transfer coefficients KG (mmol s 1 m 2 kPa 1) and and the yintercept b (mmol s 1 m 2) from regression of NCO2 versus PCO2 ðNCO2 ¼ KG PCO2 þ bÞ for MEA.
3 2
CO2 loading (mol CO2/mol MEA)
1 0 0
2
4
6
8
10
12
14
16
18
20
-1 PCO2 (kPa) Fig. 5. CO2 absorption flux NCO2 versus applied CO2 partial pressure PCO2 for 0.6 mol L 1 ammonia at 278 K and a liquid CO2 loading of 0.2.
0 0.1 0.25 0.4 0.5
5 mol L 1 MEA, 313 K
5 mol L 1 MEA, 333 K
KG
b
KG
b
2.80 7 0.09 2.34 7 0.03 1.737 0.02 1.207 0.02 0.32 7 0.02
0.48 0.16 0.20 1.2 1.5
3.69 7 0.13 2.89 7 0.05 2.247 0.06 1.257 0.02 0.377 0.05
1.8 0.081 0.92 3.0 5.5
The errors in KG are expressed as one standard deviation.
Fig. 6. Plot of the overall mass transfer coefficient KG on a log10 scale as a function of molar liquid CO2 loading. Data at the same ammonia concentration have the same symbol (square—0.6 mol L 1, triangle—3 mol L 1 and diamond—6 mol L 1) and data at the same temperature have the same colour (light blue—278 K, dark blue—283 K and red—293 K). (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Fig. 7. Plot of the overall mass transfer coefficient KG on a log10 scale as a function of molar CO2 loading. Data at the same ammonia concentration have the same symbol (square—0.6 mol L 1, triangle—3 mol L 1 and diamond—6 mol L 1) and data at the same temperature have the same colour (light blue—278 K, dark blue—283 K and red—293 K). MEA at 5 mol L 1 and 313 and 333 K are shown as dots with thick black dashed and solid lines, respectively. The data in grey are also for 5 mol L 1 MEA at the same conditions from Dugas and Rochelle, 2009. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
both 278 and 293 K are essentially identical. A possible explanation for this is that at this low ammonia concentration there is depletion of ammonia in the liquid film. The consequence of this is that ammonia diffusion starts to play a role in limiting the absorption rate reducing the temperature sensitivity. An important comparison is between the absorption rate of CO2 into ammonia and MEA at representative conditions for the operation of a capture process. This gives an indication of the absorber sizing that would be required for an ammonia process relative to what is known for MEA. For example, a KG half that of MEA indicates twice the gas–liquid contact area would be required to achieve the same amount of CO2 removal from a gas stream at identical driving force. KG determined for 5 mol L 1 MEA at 313 and 333 K are represented in Table 2 and Fig. 7. Comparison with MEA data from Dugas and Rochelle (2009) is also given in the figure. The values of KG for MEA over the CO2 loading range of 0.1– 0.4 are 1.5–2 times greater than those of 6 mol L 1 ammonia at 293 K over the same loading range. This means that if a 6 mol L 1 ammonia solution at 293 K was used, the gas–liquid contact area would need to be 1.5–2 times larger than that of an MEA based process. The loading regime in which a capture plant is operated will also have an impact. MEA plants generally operate with lean/rich loadings2 around 0.2/0.5 (Kohl and Nielsen, 1997). If an ammonia plant is operated with higher lean/rich loadings (e.g. 0.3/0.6), to maintain a predominantly bicarbonate based system to minimise desorption energy requirements and ammonia volatilisation, the absorber sizing will need to be significantly larger than for an MEA process. Also, use of a temperature below 293 K or ammonia concentration below 6 mol L 1 will also require a larger packing
2
The lean/rich loading regime refers to the ‘‘lean’’ CO2 loading at the base of the desorption column and the ‘‘rich’’ CO2 loading at the base of the absorption column.
Table 3 The second-order reaction rate constant kref,2 and the activation energy Ea,2 for the reaction between CO2 and ammonia. kref,2 (L mol 1 s 1) at Tref,2 = 283 K
Ea,2 (kJ mol 1)
915
61
area and thus a larger absorber relative to MEA to achieve the same amount of CO2 removal. 3.2. Chemical reaction rate CO2 absorption flux measurements into unloaded aqueous ammonia were used to estimate the second-order reaction rate constant k2 using Eq. (4). The diffusivity and physical solubility (Henry’s constant) of CO2 in aqueous ammonia, DCO2 ðm2 s1 Þ and 1 HCO2 ðkPa m3 kmol Þ respectively, were assumed equivalent to that in water and were calculated using the following equations (Versteeg et al., 1996). DCO2 ¼ 2:35 106 e2199=T m2 s1
ð8Þ
HCO2 ¼ 2:82 106 e2044=T Pa m3 mol
1
ð9Þ
where T= temperature (K) Eqs. (5) and (6) were used to verify that the CO2 absorption measurements were carried out under a pseudo first-order regime. The diffusivity of ammonia, DNH3 ðm2 s1 Þ, was calculated according to Eq. (10) (Frank et al., 1996) and the exposure time of the liquid in the column was 0.2 s. DNH3 ¼ ð1:65 þ2:47XNH3 Þ 106 e1996:6=T m2 s1
ð10Þ
The condition that OM0 {Ei was only satisfied for the ammonia concentrations of 3 and 6 mol L 1, with Ei being at least 10 times
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921
35 6 mol.L-1 NH3 at 293 K
30 25
6 mol.L-1 NH3 at 278 K
20 15
3 mol.L-1 NH3 at 283 K
10 5 0 0
2
4
6
8
10 12 PCO2 (kPa)
14
16
18
20
Fig. 8. Plot of the measured flux (symbols) and calculated flux (lines) using the optimised values of kref,2 and Ea given in Table 3, Eqs. (4) and (11).
larger than OM0 . So only these concentrations were using for estimation of k2. This also supports the supposition that for the 0.6 mol L 1 measurements depletion of ammonia in the film is occurring and the absorption rate is being limited by ammonia diffusion. The value of k2 was determined using non-linear regression (Excels Solver) to fit Eq. (4) to the measured flux data. The temperature dependence of k2 was described using a reparameterised form of the Arrhenius equation (Furusjo¨ et al., 2003). 1 1
k2 ¼ kref;2 eEa =Rðð1=TÞð1=Tref;2 ÞÞ L mol
s
ð11Þ
where Ea is the energy of activation (J mol 1), R the gas constant (8.314 J K 1 mol 1) and T/Tref,2 the temperature (K). In this form of the Arrhenius equation kref,2 represents the value of k2 at the reference temperature Tref. For the purposes of non-linear regression this form of the equation reduces the correlation between the pre-exponential factor and Ea resulting in more robust optimisation. The values of kref,2 and Ea were optimised to minimise the error in the least-squares sense between the measured and calculated absorption flux with the final values given in Table 3. The measured data and the results from the regression are in excellent agreement, as shown in Fig. 8. Inclusion of reaction with hydroxide had a negligible effect on the determined parameter values. These results should be considered indicative only as the CO2 solubility and diffusivity in aqueous ammonia were estimated from those in water and only a small number of concentrations and temperatures were considered. Table 4 shows a comparison between k2 and literature data at three temperatures. The values are larger than those determined by Pinsent et al. (1956) probably due to the higher ammonia concentrations used in this work. Comparison with the results of Derks and Versteeg (2009) is difficult as their mechanistic interpretation means the reaction rate constants are concentration dependent. However, at an ammonia concentration of 1 mol L 1 direct comparison can be made and the agreement is good. To allow comparison with MEA at the same temperature wettedwall measurements were also made using CO2 free 5 mol L 1 MEA at 283 K. The value of k2 determined from this data, again using CO2 solubility and diffusivity in water, was 2634 L mol 1 s 1. This is in good agreement with the literature value 2273 L mol 1 s 1 extrapolated using Eq. (12) (Versteeg et al., 1996) or 2042 L mol 1 s 1
Table 4 The second-order reaction rate constant k2 from this work and literature sources. Temperature (K)
k2 (L mol 1 s 1) This work Pinsent et al. (1956) Derks and Versteeg (2009)a
278 283 293
574 915 2217
107 155 313
300 700 1400
a The reaction rate constants from this reference are concentration dependent. The values given are for 1 mol L 1 ammonia as at this concentration the equations used collapse to those in this work.
calculated from Eq. (13) (Alper, 1990) which covers the temperature range 278–298 K. If the results from Aboudheir et al. (2003) are extrapolated to 283 K, and values for kNH2 and kH2 O determined assuming a termolecular mechanism are converted to the appropriate second-order reaction rate constant, a value of 5507 mol L 1 s 1 is obtained. This deviation could be due to the temperature extrapolation or the different models used to interpret the data. k2 ¼ 4:4 1011 e5440=T L mol
1 1
s
k2 ¼ 8:51 1011 e5617=T L mol
1 1
s
ð12Þ ð13Þ
At a comparable temperature of 283 K the value of k2 for aqueous ammonia is approximately 2.8 times smaller than that for MEA. This indicates that the CO2 absorption rate into aqueous ammonia is smaller than that into MEA due to both lower operating temperature and lower overall reactivity towards CO2.
4. Conclusions In this work the CO2 absorption flux NCO2 into aqueous ammonia and MEA solutions has been compared. The CO2 absorption flux of both systems has been measured as a function of CO2 loading, temperature and CO2 partial pressure. A wettedwall column was used to measure the absorption into a falling thin film and results were expressed as overall mass transfer coefficient KG values. Measurements into unloaded aqueous ammonia were also used to estimate the second-order reaction rate constant between CO2 and ammonia. This is the first such
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study to provide relevant rate information for CO2 absorption into an aqueous ammonia absorbent. As expected for a chemical absorbent combining diffusion and fast chemical reaction, the CO2 absorption flux fell as the liquid CO2 loading increased. KG increased with increasing temperature over the range 278–293 K for the 6 mol L 1 ammonia solution when operating under a regime of mass transfer enhanced by chemical reaction. The same temperature dependence was not seen for the 0.6 mol L 1 ammonia solution probably due to depletion of ammonia in the liquid film. KG was most sensitive to ammonia concentration, with a 3–6 fold increase in when moving from a 0.6 to a 6 mol L 1 ammonia solution. The results for ammonia were also compared to those of a 5 mol L 1 MEA solution at 313 and 333 K. In general the values of KG for the MEA solution were 1.5–2 times greater than the 6 mol L 1 ammonia at 293 K. At temperatures and concentrations below this the value of KG for ammonia was up to 7 times smaller than MEA. Absorption into unloaded 3 and 6 mol L 1 aqueous ammonia at 278, 283 and 293 K were used to estimate the secondorder reaction rate k2 between CO2 and ammonia. Comparison with MEA data at 283 K suggests that the smaller values of KG for aqueous ammonia are driven by both lower temperature and lower reactivity towards CO2. In general, an aqueous ammonia based process will require a larger gas–liquid contact area and thus a larger absorber column than a traditional MEA based process to achieve the same amount of CO2 removal from a gas stream. Whether the size increase is marginal or significant depends on a combination of the temperature, concentration and lean/rich loading regime used for the process.
Notation b DCO2 DNH3 Ea,2 Ei HCO2 k2 kref,2 kg kl1 kl KG M0 NCO2 PCO2 PCO2;i PCO 2 rCO2 R t T Tref,2 [CO2]
y-intercept of flux versus loading linear regression, mmol s 1 m 2 diffusivity of CO2 in water, m2 s 1 diffusivity of NH3 in water, m2 s 1 activation energy for second-order reaction, kJ mol 1 infinite enhancement factor Henry’s constant for CO2 in water, Pa m3 mol 1 second-order reaction rate constant, L mol 1 s 1 second-order reaction rate constant at reference temperature Tref,2, L mol 1 s 1 gas-side mass transfer coefficient, mmol s 1 m 2 kPa 1 liquid-side mass transfer coefficient without reaction, mmol s 1 m 2 liquid-side mass transfer coefficient with reaction, mmol s 1 m 2 overall mass transfer coefficient, mmol s 1 m 2 kPa 1 dimensionless group CO2 absorption flux, mmol s 1 m 2 applied CO2 partial pressure, kPa interfacial CO2 partial pressure, kPa equilibrium CO2 partial pressure, kPa CO2 reaction rate, mol L 1 s 1 gas constant, 8.314 J K 1 mol 1 exposure time of liquid in wetted-wall column, s temperature, K reference temperature for kref,2, K liquid phase CO2 concentration, mol L 1
[CO2]i interfacial liquid phase CO2 concentration, mol L 1 [NH3] liquid phase NH3 concentration, mol L 1
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