CO2 Capture by Temperature Swing Adsorption: Use of Hot CO2-Rich

Article pubs.acs.org/IECR

CO2 Capture by Temperature Swing Adsorption: Use of Hot CO2‑Rich Gas for Regeneration Augustine Ntiamoah,†,‡ Jianghua Ling,†,‡,§ Penny Xiao,†,‡ Paul A. Webley,*,†,‡ and Yuchun Zhai§ †

CO2CRC Limited, The University of Melbourne, Victoria 3010, Australia Department of Chemical & Biomolecular Engineering, The University of Melbourne, Parkville, VIC 3010, Victoria, Australia § School of Material and Metallurgy, Northeastern University, Shenyang, Liaoning 110004, People’s Republic of China ‡

S Supporting Information *

ABSTRACT: Temperature swing adsorption (TSA) is an attractive technology for CO2 removal from gas streams. CO2 capture by a TSA process in which the recovered CO2 product is heated and used as regeneration purge gas has been examined. Our study is based on cyclic experiments performed on a single adsorption column packed with the commercially available zeolite NaUSY adsorbent. The commercial Aspen adsorption simulator was used to simulate the experimental system, where the model predictions agreed quite well with experimental results in terms of breakthrough and results for cycle designs based on indirect heating followed by hot product gas purge. The validated model was used to simulate the case of regeneration using only hot product gas purge, which was difficult to examine experimentally due to constraints of the experimental system used. With a three-step cycle of (1) adsorption, (2) hot gas purge, and (3) cooling, this case yielded product purities of >91% CO2 and maximum recoveries of 55.5, 76.2, and 83.6% at specific (thermal) energy consumptions of 3.4, 3.8, and 4.5 MJ/kg of CO2 for regeneration temperatures of 150, 200, and 250 °C, respectively. Calculated productivities also varied from 0.024, 0.037, and 0.047 kgCO2/kgads·h for the various regeneration temperatures. Incorporation of a product CO2 purge prior to desorption with hot CO2 purge gas increased the purity to 96% at a recovery of 90.8%these conditions are suitable for CO2 sequestration.

1. INTRODUCTION Significant reduction in the energy consumption of CO2 capture systems remains a challenge. Clearly, the source and type of the energy used by the capture plant will play an important role in the overall operating cost. Many capture researchers are currently exploring integration options with the power plant for the potential use of cheaper, low-grade thermal energy available at the power plant or heat recovered from flue gas cooling in order to offset part of the energy requirements for the capture plant as a means of reducing operating cost.1 The temperature swing adsorption (TSA) process is of interest due to its ability to directly utilize these low grade thermal energy resources for regeneration.2 So far, only a few of the adsorbent-based studies for CO2 capture reported in the literature are based on TSA compared to capture by pressure/vacuum swing adsorption (P/VSA). Table 1 shows the regeneration methods being employed in some of these studies at the laboratory scale. In conventional TSA applications such as air and natural gas drying, the adsorbent is often regenerated by direct purge with a hot nonadsorbing gas or steam.3 For CO2 capture application, because the adsorbate is the desired product, the large volume of purge gas required to heat the bed (because of the low heat capacity of gases) could result in significant dilution of the extracted CO2 product when hot nonadsorbing gas is used for regeneration. Hence, as shown in Table 1, the bed is often first heated indirectly to the desired regeneration temperature using several means including heating jackets, electric heating tapes, or coils wrapped around the adsorber, and hot/cold fluid carrying tubes. The purge gas (usually hot air or N2 which is © 2015 American Chemical Society

Table 1. Some Studies on CO2 Capture by TSA, Showing the Regeneration Methods Used ref 26 4 27 17 21, 28 18

regeneration method combined PTSA/PSA process; purge gas heated indirectly by steam in a heat exchanger bed is heated indirectly by heating ribbon/tape wrapped around the column + hot N2 purge indirect heating of bed by copper coils filled with hot ethylene glycol/ water mixture + hot N2 purge bed is heated by a heating tape coiled around the column + hot N2 purge bed is heated indirectly by means of steam condensation; column consists of two concentric tubesadsorbent is packed in the annulus created by the tubes while steam and cold water flow through the inner tube to provide heating and cooling, respectively indirect heating using two concentric tubes with adsorbent in the inner tube and heat-conducting oil in the outer tube, with additional hot N2 purge

used at the laboratory scale) is then only applied for a shorter time as “sweep gas” to recover the desorbed components. These arrangements appear to work well in the small laboratory-scale systems considered. However, the indirect heating modes are likely to have scale-up issues and are not practical for large scale beds (>2 m diameter). A study by Tlili et al.4 reveals that the use of N2 even as only a “sweep gas” can Received: Revised: Accepted: Published: 703

May 10, 2015 December 2, 2015 December 2, 2015 December 2, 2015 DOI: 10.1021/acs.iecr.5b01384 Ind. Eng. Chem. Res. 2016, 55, 703−713

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Industrial & Engineering Chemistry Research still lead to significant dilution of the final product with attempts to increase recovery. Regeneration by direct steam stripping has been largely avoided probably because of the reduction in capacity of currently well-developed CO2 adsorbents in the presence of water.5,6 Direct steam purge has, however, been successfully used with regenerable (weak) chemisorbents for CO2 capture in fluidized bed TSA systems.7−9 The main issues of concern here are the high sorbent attrition rates associated with these adsorber configurations and the narrow working temperature range of the amine chemisorbents increasingly used in these systems, requiring careful heat treatment of the feed gas and strict monitoring of process temperature. Electrical swing adsorption (ESA), a type of TSA where the regeneration heat is provided by running an electric current through a conducting monolithic adsorbent, is also being tested for CO2 capture.10−13 Availability of the special adsorbents and construction materials for adsorbers are the key development challenges.13 Thus, the search for efficient TSA regeneration methods for CO2 capture applications remains an important research objective. Given this background, the current study considered using the CO2 product itself as the regeneration purge gas in fixedbed TSA systems. This ensures that at least the actual obtainable CO2 purity under the given adsorption−desorption conditions is maintained. Grande et al.13 proposed this idea for the electric swing adsorption process, and showed analytically that, while enhancing product purity, about 50% savings in electrical energy consumption can be made when the CO2 regeneration gas stream is heated by available heat. In this study we have carried out experiments and simulations (using the Aspen adsorption simulator) to provide an initial assessment of this method.

Figure 1. Adsorption equilibrium isotherms of (a) CO2 and (b) N2 on zeolite NaUSY adsorbent. Symbols represent experimental data, while lines represent fits to the dual-site Langmuir model.

2. ADSORBENT 2.1. Adsorbent. The adsorbent is at the heart of the adsorption process, and it strongly dictates the overall performance and/or the operating conditions required. Chemisorbents such as supported ammines which are known to perform better with the TSA technique2,14−16 are not widely available on the market. They are usually synthesized and used by individual research groups. Other researchers have considered other commercially available physisorbents to test the TSA principle for CO2 capture including activated carbon,17 zeolite 13X,18−20 and zeolite 5A.4,21 Many of these studies yielded limited product purities and recoveries in comparison with performance of the materials in the VSA technique. The less commonly used zeolite NaUSY was chosen to study the proposed TSA process concept. The study focuses on demonstrating the new process cycle and not on presenting NaUSY as a superior adsorbent for CO2 capture by the TSA principle. A powder sample of NaUSY with SiO2/Al2O3 ratio of 5:1 was obtained from Tosoh, Japan, and then mixed with 30% binder and cast into extrudates in our laboratory. Isotherms of CO2 and N2 on the adsorbent were measured at selected temperatures of 30, 100, 150, and 200 °C and pressures up to 113 kPa, using the ASAP2010 (gas adsorption analyzer) apparatus. The experimental isotherm data were fitted to the dual-site Langmuir model for use in the adsorption simulator. Figure 1 illustrates the data and the fit to the data. 2.2. Theoretical Working Capacity and Selectivity. For a feed gas mixture of 85% N2 and 15% CO2, and adsorption conditions of 30 °C and 1 bar pressure (representing CO2

partial pressure of 0.15 bar) and desorption conditions of elevated temperature (T), 1 bar, and 100% CO2 (representing CO2 partial pressure of 1 bar), theoretical working capacities of the adsorbent for CO2 can be estimated as 1.3, 1.7, and 2.0 mol of CO2/kg of adsorbent, at proposed elevated regeneration temperatures, T, of 150, 200, and 250 °C, respectively. For example, the amount of CO2 adsorbed at 30 °C, 0.15 bar (q1), is 2.08 mmol/g, and by heating to 200 °C, 1 bar (q2), the amount adsorbed is 0.38 mmol/g; hence the working capacity (q1 − q2) is 1.70 mmol/g at 200 °C. The calculation is shown in Supporting Information, Table S1. Since there are no experimental isotherm data points at 250 °C, the analytical relation obtained from the Langmuir fitting parameters was used. It must be noted that this analysis overestimates the adsorbent performance in a process equipment as it does not consider several performance limiting factors such kinetic and thermal effects, and process equipment and instrument inefficiencies. Observation of the N2 isotherm shows that very little N2 is adsorbed onto NaUSY at both adsorption and desorption conditions, which also indicates the material is suitable for separating CO2 from dry flue gas. The adsorption capacities estimated from the isotherms generally compare well with those from other physical sorbents in the studies cited above. The heat of adsorption values obtained from the isotherm fits are however lower (average of 25.8 MJ/kmol of CO2), indicating less strong dependence of the adsorption 704

DOI: 10.1021/acs.iecr.5b01384 Ind. Eng. Chem. Res. 2016, 55, 703−713

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Industrial & Engineering Chemistry Research

Figure 2. Simplified schematic diagram of the single column TSA experimental unit used in the study. Legend: MFC = mass flow controller; PT = pressure transmitter; TT = temperature transmitter; AT = CO2 analyzer.

equilibrium on changes in temperature2 in comparison with zeolites 13X and 5A.

without the solenoid valves on the regeneration purge gas line (upstream). Cold N2 purge (cooling step) is carried out in a similar fashion as the feed, but with the feed manual valve closed and that of N2 opened. 3.2. Description of Process Cycles and Operating Conditions. Three different TSA cycles based on different combinations of the cyclic steps I−IV shown in Figure 3 were tested. These basic process steps can be grouped into adsorption, heating, and cooling steps. 3.2.1. Cycle 1 (Regeneration by Indirect Heating). This was carried out as a preliminary cycle to determine the product purity that can be obtained at the given adsorption conditions and regeneration temperatures. The observed purity then

3. EXPERIMENTAL SECTION 3.1. Experimental Setup. A schematic diagram of the system used is presented in Figure 2. The column is made of stainless steel, and has an original height of 1000 mm and inside diameter of 10 mm. An electric-powered ceramic furnace is mounted around the column for indirect/external heating of the packed adsorbent. A gas heating chamber is also connected upstream of the column for heating the purge gas. The unit is equipped with mass flow controllers, pressure transducers, and flow meters to monitor and measure stream flow rates and pressures. A PLC/SCADA system (Allen Bradley and Factory Talk) is used for process control, monitoring, and data acquisition. In the current study, the column was packed to a height of 800 mm (from the top) with the main NaUSY adsorbent, while the bottom 200 mm section was packed with nonadsorbing adsorbent. Due to the small diameter of the column used, inbed temperature sampling was not feasible. The three thermocouples attached to the column wall (Figure 2) measure the temperatures between the furnace and the column wall. The simulation tool was relied upon to understand the bed dynamic profiles during operations. A standard mixed gas of 15% CO2 and 85% N2 in a gas cylinder was used as the source of feed gas. CO2 concentrations in the waste and product streams were determined with the help of a Servomex CO2 analyzer. The cycles were operated in a semiautomated manner. During adsorption, the manual valve on the N2 line is closed, while that on the feed line is opened together with the solenoid valves on the bed inlet and outlet (vent line) to allow the feed gas to flow through the bed from the bottom. Depending on the cycle under consideration, desorption is carried out by switching on the solenoid valves on the product line with or

Figure 3. Cycle designs (cycle steps and their sequence) examined in the study. Legend: I = feed; II = indirect bed heating (with or without CO2 product recovery); III = direct countercurrent hot gas purge with CO2 product recovery; IV = cooling by cocurrent N2 purge. F = feed gas; W = waste gas; P = product gas (CO2); H = hot purging gas; C = cooling N2 gas. 705


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column with a one-phase jacketed heat exchanger was chosen to represent the experimental system. The software simulates the process using coupled equations for mass, momentum, and energy balances between the gas and the adsorbent particles. General assumptions made include ideal gas behavior and a plug flow without axial mixing. Other specific modeling choices are described below. Mass and heat transfer coefficients were not experimentally available but were estimated from standard correlations within the software. The input parameters used in the simulations including column and adsorbent characteristics, and gas physical properties are presented in Table 2.

served as a guide to prepare the gas mixture used for hot regeneration in subsequent cycles. The step sequence (Figure 3) is as follows: feed (I); regeneration (indirect heating + product) (II); cooling (IV). Feed: The feed gas is fed to the column from the bottom at ambient temperature and pressure, and a flow rate of 2 SLPM for CO2 adsorption. The step is terminated when CO2 concentration in the effluent reaches 3%. Regeneration: Following adsorption, the bed is heated indirectly to the desired regeneration temperature (varied from 150, 200, and 250 °C) by the electric furnace/jacket mounted over the column. The final temperature was set on the control panel of the experimental rig. The column valves are closed during heating; then the product valve is opened for 2 min to collect the desorbed components into a product tank. Closing all the column valves during bed heating could cause significant increase in bed pressure, but this will reduce when the bottom valve is suddenly opened to recover the product. This was done to ensure appreciable flow of product to enable more accurate measurement by the flow meter. To compare performance on the same basis, a heating duration of 30 min was used for this step. Simulations revealed the bed fully attains all the set desorption temperatures by this duration. Cooling: After desorption, the bed is cocurrently purged with N2 gas at room temperature and a flow rate of 2 SLPM. N2 both cools and cleans the bed for the next adsorption step. The duration of this step equals the heating time used in the preceding step. Although nitrogen was used in our laboratory work, a more practical implementation in the field would be to use air. This has the added advantage that the air heated by the bed can be directed to the combustion chamber of the power plant to improve the combustion efficiency. The air exiting the bed in the first few seconds may contain high concentrated CO2 and this could be recovered by recycling to the product. 3.2.2. Cycle 2 (Regeneration by Indirect Heating Followed by Hot Product Gas Purge). The step sequence (Figure 3) is the following: feed (I); indirect heating (II); hot gas purge (III); cooling (IV). Both adsorption and cooling are the same as in cycle 1. However, after indirect heating with column valves closed as in cycle 1, the bottom valve is opened and a hot product gas is simultaneously introduced from the top of the bed as purge gas to desorb the adsorbate. As the steps were carried out in batch manner, and also because of the small size of the system, recycling a fraction of the collected product through a heat exchanger to use as a hot purge gas could not be done. A CO2/ N2 mixture was prepared to a composition equivalent to the product purity obtained in cycle 1 to mimic the product gas from the current cycle. The flow of hot purge gas was fixed at 1.3 SLPM for a duration of 2 min. 3.2.3. Cycle 3 (Regeneration by Hot Product Gas Purge Only). The step sequence (Figure 3) is as follows: feed (I); hot gas purge (III); cooling (IV). Adsorption and cooling conditions are the same as in cycles 1 and 2. However, bed heating/regeneration is achieved by a direct contact with a hot product gas only. Thus, indirect heating of the bed is avoided. The hot purge gas composition is the same as the purity realized in cycle 1.

Table 2. Column and Bed Parameters Used in the Simulations parameter

value

Hb1

0.2

Hb2

0.8

Db

0.01

Wt Ei

0.0025 0.35

Ep

0.6

RHOs Rp SFac Cps

700.28 2.0 0.83 1.0

ΔH(“CO2”)

−25.8

ΔH(“N2”)

−15.5

MTC(CO2)

0.5

MTC(N2)

0.3

units

description

m

height of bed (layer 1, inert material) m height of bed (layer 2, main adsorbent) m internal diameter of adsorbent layer m wall thickness used of bed m3 of void/m3 of interparticle voidage bed m3 of void/m3 of intraparticle voidage bed kg/m3 bulk solid density of adsorbent mm adsorbent particle radius n/a adsorbent shape factor kJ/kg·K adsorbent specific heat capacity MJ/kmol constant for heat of adsorption MJ/kmol constant for heat of adsorption 1/s constant mass transfer coefficient 1/s constant mass transfer coefficient

4.1. Equilibrium Model. Gas equilibrium loading onto the adsorbent was described by the dual-site Langmuir model, which was found to provide a good fit to the experimentally measured isotherm data (see Figure 1). Equation 1 shows the version of the model supported in Aspen adsorption, while values of the model parameters are given in Table 3. IP P ( T ) i qi* = IP 1 + IP3i exp( T )Pi

IP1i exp

2i

4i

IP P ( T ) i + IP 1 + IP7i exp( T )Pi

IP5i exp

6i

8i

(1)

where qi is the equilibrium loading of component i (kmol/kg of adsorbent), Pi is the equilibrium pressure of component i (bar), and T is temperature (K). Units of the model parameters are derived as IP1 = kmol·kg−1·bar−1, IP2 = K, IP3 = bar−1, IP4 = K, IP5 = kmol·kg−1·bar−1, IP6 = K, IP7 = bar−1, and IP8 = K. 4.2. Rate of Mass Transfer. Mass transfer is described by the solid phase lumped parameter linear driving force model (eq 2). Values of the mass transfer coefficient (MTC) were estimated by considering resistances in the external fluid film and in the macropore of the particle, thus accounting for molecular and Knudsen diffusion. The different values used for CO2 and N2 were based on their relative molecular weights. Preliminary experiments carried out with different MTC values showed that mass transfer resistance was not a major factor at

4. SIMULATION IN ASPEN ADSORPTION The TSA process was simulated in Aspen adsorption. The simulations were conducted with the view to gaining better insights into the dynamics of the system studied. An adsorber 706


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Table 3. Isotherm Fitting Parameters of the Dual-Site Langmuir Model for CO2 and N2 on Zeolite NaUSY Measured at Temperatures 30, 100, 150, and 200 °C and Pressures 0−1.13 bar isotherm parameter component

IP1

IP2

IP3

IP4

IP5

IP6

IP7

IP8

CO2 N2

0 4.20 × 10−7

30001 1765.77

0 0

30184 2.81489

1.99 × 10−6 6.23 × 10−15

2609.32 3.25178

0.03954 1.47395

932.058 1.19157

4.5. Simulation Flow Sheet. A picture of the simulation flow sheet used is shown in Figure 4. The feed gas enters

the high temperature operation of our experiments. Indeed, our simulations and data suggest that the system operated close to equilibrium. ∂qi ∂t

= MTCi(qi* − qi)

(2)

where q*i is the equilibrium adsorbent loading of component i (mol/kg) and MTCi is the mass transfer coefficient of component i (1/s). 4.3. Momentum Balance. The pressure drop through the adsorbent bed is estimated by the Ergun equation (eq 3). ⎛ 150 × 10−5μ (1 − ε )2 i ∂P g = −⎜⎜ υg 2 3 ∂z (2rpψ ) εi ⎝ +

1.75 × 10−5M w ρg (1 − εi) (2rpψ )εi 3

⎞ υg 2⎟⎟ ⎠

(3)

where ψ is the particle sphericity or shape factor, rp is the particle radius, and εi is interparticle voidage. 4.4. Energy Balance. The energy balance model selected is that for nonisothermal adsorption with no solid conduction. The model considers independent balances in the fluid, solid, and wall phases. The general wall balance equation includes a term for the heat content of the wall, which is based on the specified values of the wall density and the specific heat capacity of the wall material. Gas−solid heat transfer is expressed in terms of a film resistance, where the heat transfer area is proportional to the area of the adsorbent particles, as given in eq 4.

Figure 4. Aspen adsorption flow sheet for simulating TSA cycles.

through F1 and exits at W1 for adsorption. The hot purge gas is admitted through F2, and the product is collected through P1. The composition, temperature, and pressure of the inlet feed and purge gas streams were specified in F1 and F2, respectively. The N2 cooling step was carried out similarly to the feed step, but with the composition specified as 100% N2 in the cycle organizer. For the indirect heating, the adsorber column was specified as a one-phase jacketed heat exchanger, and the heating requirements and/or conditions were specified in the cycle organizer. 4.6. Model Solution. Aspen adsorption uses the numerical method of lines to solve the time-dependent partial differential equations (PDEs) describing the mass, momentum, and energy transport between the gas and the adsorbent particles at each step of the process cycle. We chose the first-order upwind differencing scheme (UDS1) as the method of spatial discretization of the PDEs because of its relatively higher stability, while the default implicit Euler integrator was chosen for the integration of the resulting ordinary differential equations (ODEs), also for reasons of stability. A large number of nodes (150 nodes) were used in the simulations. This was found to provide consistent results after trying several different nodal values. Although it leads to more computation time, using higher discretization nodes avoids the “smearing” of mass transfer fronts and minimizes the “numerical dispersion” in results often given by the UDS1 method.22 The simulations were continued until cyclic steady state (CSS) was reached at the specified CSS tolerance value of 1 × 10−5, which is the difference allowed between the thermal and concentration profiles of the current and previous cycles.

rate of heat transferred per unit volume = (HTC)a p(Tg − Ts)

(4)

where ap is the particle external surface area per unit volume of bed, m−1, and HTC is the gas−solid heat transfer coefficient (MJ/m2/K). HTC is also estimated by the software based on the Colburn j-factor correlation22,23 (eq 5). HTC = jCp,gvgρg Pr −2/3

(5)

where Cp,g is the fluid phase heat capacity, J/kg/K; ρg is the fluid-phase density, kg/m3; vg is the fluid phase velocity, m/s; and Pr is the dimensionless Prandtl number. j = 1.66Re−0.51 for Re < 190; otherwise j = 0.983Re−0.41. Heats of adsorption (ΔHads) of CO2 and N2 as a function of loading were calculated from the measured isotherm data by using the Clausius−Clapeyron relation (eq 6). ⎡ ∂ ln p ⎤ ΔHads = −R ⎢ ⎥ ⎣ ∂(1/T ) ⎦n

(6)

where p and n respectively are the partial pressure and component loading. 707


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5. RESULTS AND DISCUSSION 5.1. Breakthrough Data and Simulation. The ability of the simulation model to represent the experimental system for a single adsorption step is shown in Figure 5this is the

Figure 5. Breakthrough curves for CO2 from experiment and simulation. Open circles = experimental data; line = simulation. Inlet feed gas conditions: CO2/N2 = 0.15/0.85; pressure = 1 bar; feed gas velocity = 0.42 m/s.

breakthrough curve for the feed mixture on a clean bed. The movement of the temperature front during the breakthrough run (simulated results) is also shown in Figure 5. As can be seen, the bed has not yet been saturated with CO2 (i.e., Co/C < 1.0) and hence the temperature has not cooled to the feed value. Both breakthrough curves form similar pattern fronts, with the center of the waves matching very well until the feed composition is approached where the experimental curve shows a more dispersed front. CO2 breakthrough occurred earlier in the simulation (about 4 min), while it took about 4.5 min in the experiment. This may be due to a capacity issue where the capacity for the bed estimated by the model is less than the actual capacity in the experiments. The dead volume in the experimental system between the end of the column and the CO2 analyzer could also contribute to the later breakthrough in the experimental system. 5.2. General Performance of Experimental and Simulation Systems. Cycles 1 and 2 were studied experimentally. Experimental runs were repeated for each regeneration temperature, and consistent results in terms of product purity and recovery were obtained (see Supporting Information, Figures S1 and S2). The simulations also predicted product purities and recoveries quite well as discussed below. Flows of the product during regeneration in cycles 1 and 2 are shown in Figure 6 for both experiments and simulations. The flow of CO2 product stopped after about 40 s after the bed inlet was opened following the indirect heating (cycle 1). In cycle 2, desorption was considered to be complete when the flow rate of the product stream was reduced to that of the hot purge gas, meaning no more CO2 was being recovered from the bed. This also occurred under 1 min of hot gas purge at the given purge flow rate. Product purities and recoveries for cycle 2 (regeneration by indirect bed heating followed by hot “product gas” purge) are compared in Figure 7 for both experiments and simulations.

Figure 6. Flow of product during desorption. (a) Cycle 1: regeneration by indirect heating of bed only. (b) Cycle 2: regeneration by indirect heating of bed followed by hot product gas purge. Line = simulation; symbol = experiment.

Figure 7. CO2 recovery as a function of regeneration temperature (cycle 2): comparison of experimental and simulated results. Line = simulation, symbol = experiment.

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This can be expected as more CO2 is recovered at a higher temperature. The effluent during the first few seconds could potentially be recycled to the product tank (because of the high CO2 concentration at this stage), to reduce product loss and increase recovery. The analysis in Table S2 shows that by recycling the initial N 2 purge effluent up to a CO 2 concentration of 50% directly to the product stream, the product recovery can increase from 52 to 77%, from 77 to 88%, and from 86 to 94% CO2 for regeneration temperatures of 150, 200, and 250 °C, respectively. This, however, lowers product purity from 92% CO 2 to 84.7, 87.9, and 89.2% CO 2 respectively. Figure 9 shows the CSS profiles for the adsorption and desorption steps. Gas-phase concentration profiles for adsorption and desorption steps for a reference regeneration temperature of 200 °C are shown in Figure 9a. There are significant changes in loading at the end of desorption. However, the profile did not change for the various regeneration temperatures, due to the longer purge with N2 resulting in an almost clean bed. The amount of CO2 allowed to break through the column during adsorption (3%) was significant (Figure 9b); while this favors CO2 product purity, it contributes to loss in recovery. Figure 9c shows that a higher regeneration temperature leads to higher change in loading amounts at the end of regeneration steps, which explains the trend in product recovery with the regeneration temperature. Obtaining a high purity product is very important in the carbon capture and storage (CCS) application as this reduces the downstream processing requirements before pipeline transportation. A purity of at least 95% CO2 is usually specified. Purity can be improved by employing a CO2 rinse step prior to desorption and recycling the effluent to avoid loss of CO2 through the rinse outlet.24 Moreover, with two or more adsorption columns, pressure equalization can also be employed to enhance the purity. Such a cycle design has been demonstrated by Jain25 in a combined temperature/ vacuum swing adsorption process. After adsorption, the gas at the top of the column which is predominant in N2 (the lighter component) is transferred to another bed at a lower pressure until their pressures equalize. The pressure equalization step enriches the bed providing the equalization gas in CO2 leading to higher product purity in the subsequent desorption step. 5.3. Purity, Recovery, and Productivity (Regeneration by Hot Product Gas Purge Only). Regeneration by hot product gas only was also studied. However, given the limitation of our experimental system as mentioned earlier, this cycle was investigated using the simulation tool only once it had been validated against experimental data for cycles 1 and 2. The goal here was to determine the maximum recoveries obtainable and the corresponding purge gas requirements at the same regeneration temperatures considered in the previous cycles. Feed and cooling conditions and hot purge gas flow rates and other operating conditions used were the same as in previous cases studied. The amount of purge gas used plays an important role in the time required to heat the bed to reach the regeneration temperature. A higher gas flow can be used which allows a more rapid increase in temperature. Alternatively, a smaller gas flow can be used with a longer heating time. The latter approach was used to minimize pressure drop across the bed. Figure 10 shows the transient bed temperature profiles during heating. With the hot gas flow rate fixed, the time required for the bottom of the bed to reach the desired regeneration temperature decreases

Each experimental data point in Figure 7 represents the average result from six runs (see Figure S1). Purity and recovery are defined in eqs 7 and 8, respectively. Because the purge gas was not recycled from the product tank, the amount of CO2 used in the hot purge gas is subtracted from the total product in order to show the net recovery based on the feed CO2. As expected, recovery increased with regeneration temperature, while purity remained almost constant at about 92% CO2 for both experiments and simulations. It can be seen from Figure 7 that there is a good match between experimental and simulated performances. purity =

∑ (CO2 concn·prod. flow rate·Δt ) ∑ (prod. flow rate ·Δt )

(7)

recovery = (amt CO2 in prod.) − (amt CO2 in reg purge gas) amt CO2 in feed

(8)

There were no significant differences in terms of CO2 product purity and recovery between case 1 (indirect heating of bed for regeneration) and case 2 (indirect heating of bed followed by countercurrent purge with hot product CO2 gas). This means the use of hot product gas as sweep gas may have little or no effect on product recovery. Tlili et al.4 observed in their study that there is a limit to the extent of recovery with bed heating only, without a means to lower the adsorbate partial pressure below atmospheric. In their study, additional gain in recovery was made by purging with hot N2 gas (which lowered the CO2 partial pressure in the bed) after the flow of product had stopped with bed heating. As mentioned earlier, this has the potential to lower product purity by means of dilution depending on the amount of N2 gas used. Analysis of the effluent of the cold N2 purge step shows that an appreciable amount of CO2 is purged out of the bed initially during this step, with the amount decreasing with increasing regeneration temperature as shown in Figure 8, with quantitative data in the Supporting Information, Table S2.

Figure 8. Flow of CO2 in effluent of N2 purge step (for cooling) following desorption at various regeneration temperatures in cycle 2 (regeneration by indirect heating of bed + hot product gas purge): experimental results. Flow velocity = 0.42 m/s. 709


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Figure 9. Simulated cyclic steady state profiles for cycle 2. (a) Outlet CO2 mole fraction during adsorption and regeneration at 200 °C. (b) Axial gasphase concentration profiles during adsorption and desorption at regeneration at 200 °C. (c) Axial solid-phase concentration during adsorption and regeneration at different temperatures.

from 150 to 250 °C. For a given regeneration temperature, recovery increased with purge gas amount (increasing purge time at fixed gas flow rate) until the recovery leveled off and even began to decrease with further increase in purge gas amount (Figure 11). At this point the specified regeneration temperature was reached and hence further purge gas did not

result in a net increase in recovery. Thus, it is no longer beneficial to input more energy. As shown in Figure 11, maximum recoveries observed were 56, 76, and 82% for regeneration temperatures of 150, 200, and 250 °C, respectively. These values correspond to a ratio of CO2 in the purge gas to CO2 in the feed gas of 24.5 for 150 °C, 20.8 for 710


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exchanger: the product CO2 removed then represents a slip stream from this recycle loop. To keep the system in balance, the final product drawn at the end of each cycle must be less than the amount of CO2 in the feed stream. Very high product purity (>95 mol % CO2) and recovery (>90% CO2), which are commonly required for CCS applications, could not be achieved with the system studied. The lower recovery, in particular, agrees with the performance of simple TSA cycles on physical adsorbents4,18,21 which is probably due to a larger thermal mass of the materials. Recovery can further be enhanced by applying a moderate vacuum pressure after the hot gas purge (when the bed is still hot)4,25 or introducing a product rinse step before the hot regeneration step (as is often done in vacuum swing adsorption systems where a heavy reflux is used). We have evaluated this concept as explained below. 5.4. Achieving Higher CO2 Product Purity and Recovery with Product Rinse. Achieving higher product purity (>95 mol % CO2) and recovery (>90% CO2) are major design objectives in CO2 capture programs targeted at sequestration. The previous runs were conducted by allowing CO2 to break through the adsorbent to about 3% of the entire effluent stream mole concentration. This loss of CO2 is significant and affected the product recovery (based on the amount of gas fed to the bed). By varying the adsorption step time (by means of simulations), it was found that higher CO2 product recovery (>90%) can be achieved by stopping the feed step short of breakthrough; however, there is a reduced purity of 88% CO2. Appreciable increase in the adsorption time did not yield the desired >95% CO2 product purity, which is due to inadequate selectivity of the adsorbent material. To determine under what conditions a purity greater than 95% CO2 could be achieved, the cycle run yielding >90% CO2 was selected and rinsed with part of the product CO2 after adsorption in order to increase the amount of CO2 adsorbed on the bed prior to desorption. A feed pressure of 1.2 bar was used with a shorter step time of 40 s which ensured the (cold) CO2 purge front stayed within the column. This run yielded 96% CO2 product purity with 90.8% recovery using a regeneration temperature of 250 °C. The Supporting Information, Figure S3, shows the CO2 loading profiles (moles of CO2 per kilogram of adsorbent) as functions of normalized distance (x/L) from the feed end at the ends of adsorption + product rinse and desorption steps. The ratio of difference in loadings of CO2 and N2 between adsorption and desorption (calculated by integrating the areas under the loading profiles) gave the actual adsorbent working CO2 selectivity of 43.3 (Supporting Information, eq S1). The amount adsorbed at the end of the product rinse step is 3.1 mol/kg, which reduces to 0.2185 mol/kg after desorption, resulting in a CO2 working capacity of 2.88 mol/kg at 250 °C. This means an adsorbent material capable of thermal regeneration and showing actual CO2 working selectivity of 43.3 can provide >95% product purity as well as higher recovery with the proposed cycle design without necessarily including the product rinse step which comes with a small extra compression cost. The residual amount of CO2 after desorption was the same for the cycles with and without the cold product rinse step, which indicates the regeneration temperature and/or regeneration energy consumption is not affected by the amount of gas adsorbed on the adsorbent.

Figure 10. Effects of purge gas amount on CO2 recovery for cycle 3 (regeneration by hot gas purge only).

Figure 11. Simulated cyclic steady state transient temperature profiles during regeneration by hot gas purge (cycle 3) at different temperatures. Direction of gas flow is from the top (L = 80 cm) to bottom (L = 0 cm) of bed.

200 °C, and 17.2 for 250 °C. As in the previous cases, product purities were seen to be less sensitive to regeneration temperature and purge gas amount. Recoveries also compare very well with those from cycle 2. Summarized results for this cycle are presented in Table 3. Because the purge gas is the sole source of heat, larger amounts were required to effectively heat the bed. This means that enough product gas must be available initially to recycle through the heat exchanger and the bed to effect the regeneration. The practical implementation of the hot purge gas system will require a recycle loop with a heat 711


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amount of hot gas recycled through the bed in order to heat the bed to the required regeneration temperature. It can be inferred from eq 9 that a much lower specific regeneration energy can be achieved with an adsorbent with a small ΔT and higher CO2 thermal working capacity (CO2 recovered per mass of adsorbent).

Productivity (defined in eq 9) is also an important variable in the overall evaluation of adsorption systems. It shows how frequent the adsorbent bed can be used in a cyclic operation. productivity =

(amt CO2 in prod.) − (amt CO2 in reg purge gas) (mass of adsorbent)(cycle time)

6. CONCLUSION The key objective of this work has been to assess the potential of a TSA cycle based on using the recovered product as regeneration purge gas. Both experiments and simulations (using Aspen adsorption) have been carried out. Based on a simple one-bed/three-step cycle and zeolite NaUSY adsorbent, it has been shown that product purities of >91% can be obtained with recoveries of 55, 76, and 84% for regeneration temperatures of 150, 200, and 250 °C, respectively, using desorption by a purge with a hot gas stream with composition equivalent to that of the product. The results of the study, along with those from other studies (discussed above), indicate that regeneration temperatures of >150 °C are necessary to achieve reasonably high separation efficiency using current commercial physisorbents. However, such high temperatures may fall short of the expectation of using cheaper waste heat as the source of energy for the regeneration step. In addition, the large quantities of purge gas required would definitely place significant demands on the size of the system required when capture from a power plant flue gas with higher flow rate is envisaged. Notwithstanding this, the concept has been well demonstrated and may work cost-effectively under different conditions and for different applications. The system performance depends strongly on the operating conditions chosen; hence, it is possible that purities and recoveries could be further enhanced when optimal operating variables are determined and used in future studies. Because the performance is also tied to the properties of the adsorbent, improvements in the results may also be realized by using adsorbents with better properties than the one used here.

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The need to use longer times to heat and cool the adsorbent bed is a major drawback of this NaUSY TSA system, which results in relatively lower productivities. Productivities of 0.024, 0.029, and 0.041 kg of CO2/kg of ads/h (and related bed sizing factors (BSFs) of 27, 22, 20 kg of ads/kg of CO2 per cycle) were obtained at the different regeneration temperatures in the current study (cycle 3). Cycle times are 1.40, 1.23, and 1.06 h for regeneration at 150, 200, and 250 °C, respectively. As noted above, there is an option to reduce the heating time by using larger hot gas flow rate, but this may have a kinetic effect on the process. 5.5. Energy Consumption. A critical factor in the regeneration process is the energy required to heat the adsorbent bed. This has been estimated for the cycle employing only hot purge gas for regeneration by using the general heat balance equation used by Berger and Bhown,14 as shown in eq 10. This is the sum of the sensible heat required to heat the bed to the specified regeneration temperature and the desorption heat (equivalent to the heat of adsorption) needed to desorb the carbon dioxide and nitrogen during the regeneration step.14 It is noted that the source of energy and the potential energy losses associated with the method of heat exchange with the regeneration gas may impact the calculated energy consumption. Higher regeneration temperature leads to higher energy consumption because of the higher sensible heat requirement, but it also recovers more of the adsorbed CO2 (higher recovery). Therefore, to allow for a better comparison, the total energy is divided by the mass of product recovered to obtain the specific energy consumption (see Table 4). Q reg = madsCp ,adsΔT + ΔHads

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The Supporting Information is available free of charge on the ACS Publications website at DOI: 10.1021/acs.iecr.5b01384. Theoretical working capacities from isotherm data at different regeneration temperatures (Table S1); summary of molar flows of CO2 into and out of the bed: cycle 2, regeneration by indirect heating followed by hot gas purge (Table S2); CO2 recoveries and purities at the same operating conditions under repeated experimental runs (cycle 2) (Figures S1 and S2); cyclic steady state CO2 and N2 loading profiles on the adsorbent at the ends of product rinse and desorption steps (Figure S3) (PDF)

Table 4. Summarized Performance Obtained by Simulating Cycle 3 (Regeneration Using Hot Product Gas) regen temp (°C)

recovery (% CO2)

purity (% CO2)

sp energy (MJ/kg of CO2)

productivity (kg/kg of ads/h)

150 200 250

55.5 76.2 83.6

91.1 91.3 91.4

3.37 3.82 4.50

0.024 0.029 0.041

ASSOCIATED CONTENT

S Supporting Information *

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where Q is the regeneration heat input (kJ), mads is the mass of sorbent (kg), ΔT is the temperature difference between adsorption and regeneration (K), Cp,ads is the sorbent specific heat (kJ/kg K), and ΔHads is the heat of adsorption (kJ/mol). (We used ΔH = nCO2ΔHCO2 + nN2ΔHN2.) A reference temperature of 30 °C is used. nCO2 and nN2 are the respective moles of CO2 and N2 recovered from the bed. The use of hot purge gas for regeneration may be considered advantageous since the hot gas contacts and heats the adsorbent bed directly, without heating from outside of the column wall in which the supplied energy goes to heat both the wall and the adsorbent. The relatively higher specific energy consumption values realized in this case result from the large

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AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors acknowledge the funding provided by the Australian Government through its CRC Program to support this CO2CRC research project. We thank Dr. Ranjeet Singh 712


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(20) Mérel, J.; Clausse, M.; Meunier, F. Carbon Dioxide Capture by Indirect Thermal Swing Adsorption Using 13X Zeolite. Environ. Prog. 2006, 25, 327. (21) Clausse, M.; Merel, J.; Meunier, F. Numerical Parametric Study on CO2 Capture by Indirect Thermal Swing Adsorption. Int. J. Greenhouse Gas Control 2011, 5, 1206. (22) Aspen Adsorption v8.0 user help file. Aspen; Aspen Technologies, Inc. (23) Bird, R. B.; Stewart, W. E.; Lightfoot, E. N. Transport Phenomena; John Wiley and Sons: New York, 1960. (24) Kikkinides, E. S.; Yang, R. T.; Cho, S. H. Concentration and Recovery of CO2 from Flue Gas by Pressure Swing Adsorption. Ind. Eng. Chem. Res. 1993, 32, 2714. (25) Jain, R. Carbon Dioxide Recovery. U.S. Patent US 2010/ 0251887 A1, 2013. (26) Ishibashi, M.; Ota, H.; Akutsu, N.; Umeda, S.; Tajika, M.; Izumi, J.; Yasutake, A.; Kabata, T.; Kageyama, Y. Technology for Removing Carbon Dioxide from Power Plant Flue Gas by the Physical Adsorption Method. Energy Convers. Manage. 1996, 37, 929. (27) Mulgundmath, V.; Tezel, F. H. Optimisation of Carbon Dioxide Recovery from FLue Gas in a TPSA System. Adsorption 2010, 16, 587. (28) Merel, J.; Clausse, M.; Meunier, F. Experimental Investigation on CO2 Post−Combustion Capture by Indirect Thermal Swing Adsorption Using 13X and 5A Zeolites. Ind. Eng. Chem. Res. 2008, 47, 209.

and Mr. David Danaci in our research group for their help in the measurement and analysis of the isotherm data. We also acknowledge Messrs. Barry Hooper and Trent Harkin of CO2CRC for their contributions to this work.

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REFERENCES

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CO2 Capture by Temperature Swing Adsorption: Use of Hot CO2-Rich

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