Pressure Swing Adsorption Process for Recovering H2 from the

Article pubs.acs.org/IECR

Pressure Swing Adsorption Process for Recovering H2 from the Effluent Gas of a Melting Incinerator Dong-Kyu Moon,† Yo-Han Kim,† Hyungwoong Ahn,‡ and Chang-Ha Lee*,† †

Department of Chemical and Biomolecular Engineering, Yonsei University, Seoul, Korea Institute for Materials and Processes, University of Edinburgh, Edinburgh, United Kingdom

‡

ABSTRACT: A two-bed pressure swing adsorption (PSA) process using activated carbon was studied to recover hydrogen from the effluent gas (H2/CO/CO2, 39.3:35.4:25.3 vol.%) of a melting incineration process. The adsorption dynamics of the activated carbon bed were investigated by breakthrough experiments. Since the product purity needs depend on demands of the incinerator site, various PSA operating conditions, such as purge to feed (P/F) ratio, adsorption (AD) step time, adsorption pressure, and feed flow rate, were investigated to recover H2 with a wide range of purities. The purity varied with P/F ratio, and the recovery varied with AD step time asymptotically because bed purification by H2 product purge approached a limitation. On the other hand, the variations in purity and recovery with P/F ratio, AD step time, and adsorption pressure were almost linear or only slightly curved. The variation in purity with feed flow rate was similar to the variation with P/F ratio while the recovery trend was more similar to the variation with the AD step time. Because the propagation velocity was significantly different for CO and CO2, the PSA performance was mainly affected by CO propagation, but the contribution of the CO2 heat of adsorption to the bed should be considered. The PSA process in this study produced hydrogen with a purity of 75.43−99.99% and a recovery of 90.99−49.29%. When a syngas (H2 and CO) is needed, the PSA process can result in high recovery and productivity because the major impurity in the product is CO.

1. INTRODUCTION High economic growth in many countries has led to changing lifestyles. In particular, waste emissions have increased steeply because improved standards of living have changed consumption patterns and city concentrations. A huge amount of waste is brought into landfills daily because most waste treatment relies on open dumping. In many countries, including Korea, landfills are reaching their limits.1 Therefore, incineration of solid wastes has become an attractive way to treat waste, and its use is increasing in countries with limited landfill sites. Of all of the permanent treatment technologies, properly designed incineration systems can achieve the highest overall degree of destruction and control for the broadest range of hazardous waste streams. Substantial design and operational experience exists in this area, and a wide variety of commercial systems are available.2 Some wastes, such as medical, industrial, and military wastes, are prohibited from landfills, and special pretreatment is needed before incineration. Therefore, a focus for the incineration field is alternative treatments due to ever tightening environmental regulations. Melting incineration using coke or natural gas, which is similar to the blast furnace process using coke in the iron and steel industries, is a candidate for alternative treatment. Off-gases from melting incineration contain hydrogen. Due to global warming, another focus in the incineration field is energy efficiency and carbon dioxide reduction. Therefore, there is interest in recovering hydrogen from effluent gas generated by melting incineration. Melting incinerators are generally installed near wasteemitting sites. Three scenarios to recover a product gas from off-gases are available for melting incinerators: (1) producing H2 and recycling a certain level of H2/CO mixture as fuel to the © XXXX American Chemical Society

process, (2) producing H2 and CO by using a two-stage PSA, and (3) producing syn-gas (H2/CO mixture). The economic evaluation totally depends on the site and the size of the melting incinerator and market needs. Therefore, the quality of recovered hydrogen depends on the needs of industries which are typically located near the sites. Although effluent gas from melting incineration contains some carbon monoxide, a watergas shift reaction to generate more hydrogen may not be required, and the steam-energy balance should be considered. As an alternative, it is considered that the recovered syn-gas (H2 and CO) can be supplied to the incinerator to reduce carbon dioxide emission. Like the COG (coke over gas) process used in the iron and steel industries,3,4 the effluent gas is filtered to remove particles, scrubbed to remove sour components, and cooled to ambient temperature. Separation and purification of hydrogen mixtures, such as refinery fuel gas, coke oven gas, and reformer-off gas, by PSA technology have become an important unit in various industries. In order to develop the well-designed PSA process for hydrogen recovery, the dynamic behaviors of adsorption and temperature in a bed are important criteria. In addition, the PSA simulation is required to design PSA processes and optimize their operating conditions because it is not easy to cover the comprehensive operating results by experiments only. Even though the simulated results are somewhat different from the real operating results, it can still give some directions to Special Issue: Alirio Rodrigues Festschrift Received: January 8, 2014 Revised: March 21, 2014 Accepted: March 24, 2014

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dx.doi.org/10.1021/ie500091r | Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

Industrial & Engineering Chemistry Research

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realize the process if the simulated performance results show the same trend as that of the experimental results within an acceptable deviation range. Many studies have reported that a large variety of binary and multicomponent gas mixtures can be successfully separated by using this technology.5−14 In this study, two-bed PSA packed by activated carbon was investigated to recover hydrogen from the effluent gas of melting incineration for wastes. A ternary gas mixture consisting of H2/CO/CO2 (39.3:35.4:25.3 vol.%), which was the composition after off-gas pretreatment for melting incineration, was used as the feed. Since the CO and CO2 needs as a product strongly depend on the installed site, we only focused on H2 recovery even though the feed contains a large amount of CO (scenario 1). In addition, since the hydrogen purity needs also depend on the site, two-bed PSA was used to produce a wide variety of hydrogen purities. The dynamic adsorption characteristics in an activated carbon bed were studied by breakthrough experiments. The effects of operating variables, such as P/F ratio, AD step time, adsorption pressure, and feed flow rate, on purity and recovery were evaluated experimentally and theoretically. A nonisothermal dynamic model incorporating mass, energy, and momentum balance was used to simulate the breakthrough and PSA results.

Table 1. Characteristics of the Bed and Adsorbent adsorption bed length 100 cm inside diameter 3.50 cm outside diameter 4.04 cm materials of wall stainless steel heat capacity of wall 0.504 J/g·K density of wall 7.83 g/cm3 adsorbents (activated carbon) type

granular

pellet size pellet density bulk density bed porosity heat capacity

10−12 mesh 0.85 g/cm3 0.48 g/cm3 0.43 1.05 J/g · K

and a wall thickness of 0.27 cm. Microwire mesh was placed at the ends of the bed, and compressed glass wool was added to the bed connections. Since significant temperature variation was expected under the nonisothermal conditions because of the heats of adsorption,15 measuring the inside temperature was important. Therefore, four RTDs (residence temperature detectors, Pt 100Ω) were installed at 10, 30, 50, and 75 cm from the feed end of the bed to measure temperature variations during experiments. A pressure gauge was placed at the top of the bed and two pressure transducers at the top and bottom of the bed. A buffer tank and product tank were placed before and after the two-bed PSA system, respectively. A MFC and MFM (mass flow controller and mass flow meter, Bronkhorst, High-tech) were placed on the back of the buffer tank to maintain a constant feed flow rate and measure the amount of supplied gas. Another MFC was installed in front of the product tank to control the flow rate of the purge gas. An electrical BPR (back pressure regulator) was placed in front of the product tank to maintain constant pressure inside the adsorption bed. Check valves and microfilters were installed in front of electrical devices (such as MFC, MFM, and BPR) to prevent reverse flow and contamination. In addition, a wet gas meter (Sinagawa Co.) was used to measure the total gas flow. All solenoid valves, temperature monitors, pressure monitors, and controllers were connected to a computer by a data acquisition system (NI LabVIEW).3 The composition of the gas mixture was analyzed by an online mass spectrometer (Balzers Quadstar 421). Activated carbon (Calgon Co, 2GA-H2J) was used as an adsorbent because it shows better selectivity of CO2 than CO, and the desorption of CO2 is generally easier than that of zeolites. It was activated at 420 K for longer than 6 h in a vacuum oven. After the activated carbon was packed, the system was purged with H2, and a vacuum pump (ULVAC DA-60D, ULVAC KIKO. Inc.) was applied to the system for 1 h after vacuuming under a very low H2 flow for 1 h. Then, the system was pressurized to 10 atm with H2 before experiments. The experiments were done at room temperature (298 ± 1 K) under nonisothermal conditions. 2.2. PSA Process Description. The PSA cycle consisted of the following steps: (i) feed pressurization (PR) of a bed that was partially pressurized from a previous cycle, (ii) highpressure adsorption (AD), (iii) depressurizing pressure equalization (DPE), (iv) countercurrent depressurization (BD), (v) purge with part of a light product (PG), and (vi) pressurizing pressure equalization (PPE). The cyclic sequence for the process and a simple flow diagram are illustrated in Figure 2. Because the operating

2. EXPERIMENTAL SECTION 2.1. Apparatus and Procedures. A schematic of the PSA system and the characteristics of the adsorption bed and adsorbent are shown in Figure 1 and Table 1. A ternary gas

Figure 1. Schematic of the PSA apparatus.

mixture (H2/CO/CO2; 39.3:35.4:25.3 vol.%) as a feed gas was purchased from a gas company. Two adsorption beds with a flange were made from stainless steel with a length of 100 cm, an inside diameter of 3.50 cm, B



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(5) No variation exists in the radial direction for both temperature and concentration. The mass balance for the bulk and fluid phases between particles were as follows: ∂yi ∂t

+

∂ 2y ∂(uy)i RT ⎛⎜ 1 − ε ⎞⎟ ∂qi ρi + − DL 2i = 0 ∂z P ⎝ ε ⎠ ∂t ∂z

(1)

where i represents the component, DL is the axial dispersion coefficient, ε is porosity, u is interstitial velocity, and ρi is particle density. Applying an ideal gas law and overall mass balance for the bulk gas and fluid phases between particles can be written as follows: i=1

∂q ∂u RT ⎛⎜ 1 − ε ⎞⎟ =0 ρs + i + ∑ ∂t ∂z P ⎝ ε ⎠ n

(2)

The energy balance for the gas and solid phases in an adsorption bed is given by (ερg Cpg + (1 − ε)ρs Cps)

∂(uT ) ∂ 2T − Kz 2 ∂z ∂z n ⎛ ∂q ⎞ 2h + (1 − ε)ρs ∑ ⎜ΔHads, i i ⎟ − ε in (T − Tw ) ∂t ⎠ R in i=1 ⎝

Figure 2. Flow diagram and step sequence for six-step, two-bed PSA.

= − ερg Cpg

conditions for melting incineration depend on the type of waste, the effluent gas pressure can differ. In this study, PSA was performed from 5 to 9 atm. In addition, because the H2 quality needs depend on the location, the performance changes resulting from the P/F ratio, AD step time, and feed flow rate were also evaluated. The operating conditions for the PSA process are listed in Table 2.

(3)

where KZ is the effective axial thermal conductivity and hin is internal heat transfer coefficient.19,20 Since the diameter of the adsorption bed used in this study was rather small, the effect of heat transfer through the metallic bed wall could not be neglected.21 Accordingly, another energy balance for the wall of the adsorption bed was constructed by neglecting axial conduction in the wall:

Table 2. Operating Conditions for Two-Bed PSA run number

adsorption pressure (atm)

adsorption time (sec)

purge rate (LPM)

feed rate (LPM)

P/F ratio (purge/ feed)

Run01 Run02 Run03 Run04 Run05 Run06 Run07 Run08 Run09

9 9 9 9 9 9 9 7 5

120 120 120 90 180 120 120 120 120

0.393 0.393 0.393 0.393 0.393 0.773 1.60 0.393 0.393

6.8 4.5 9.1 6.8 6.8 6.8 6.8 6.8 6.8

0.058 0.09 0.04 0.058 0.058 0.114 0.235 0.058 0.058

∂T ∂t

ρw CpwA w

∂Tw = 2πRBihi(T − Tw) − 2πRB0h0(Tw − Tatm) ∂T (4)

where Aw = π(R2B0 − R2Bi). Ergun’s equation was used to calculate the pressure drop, where v is the superficial velocity.22 −

dP = aμυ + bρυ|υ| dz

a=

3. MATHEMATICAL MODEL To understand the dynamic behavior of the adsorption bed and PSA, a set of mathematical models was used.15−17 Equations were used to model each step of the PSA process with a suitable boundary condition. A mathematical model that included mass, energy, and momentum balance was constructed to develop a complete nonisothermal PSA model with the following assumptions:15,18 (1) Ideal gas mixture behavior exists throughout the bed. (2) The flow pattern is described by the axially dispersed plug flow model, comprising mass, energy, and momentum balances. (3) The mass transfer rate is represented by a linear driving force (LDF) model. (4) Thermal equilibrium exists between fluid and particles.

(1 − ε) 150 (1 − ε)2 , b = 1.75 2 3 4R p 2R pε 3 ε

(5)

(6)

The adsorption rate was presented by the LDF (Linear Driving Force) model.3,15 In this study, the LDF parameter was assumed as a constant and was obtained by fitting the simulated breakthrough curves to the experimental results using the reference values as initial values.23 However, the parameter values were changed in the same order of magnitude from the reference values because the equilibrium separation was dominated. ∂qi ∂t

= ωi(qi* − qi)

(7)

where qi* is the equilibrium amount adsorbed at the bulk gas phase and ωi is a single lumped mass transfer parameter. C



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where “in” is the temporal value of the effluent stream during the DPE step. A centered finite difference method (CFDM) of the second order was applied to the spatial partial derivatives. A simulation package, gPROMS Modelbuilder, was used to solve the above mathematical models.

The Langmuir−Freundlich isotherm was used to predict the adsorption equilibrium, and this isotherm was extended to predict the multicomponent adsorption equilibrium by the loading ratio correlation (LRC) equation.23 Since adsorption is exothermic, it is accompanied by changes in temperature. The temperature dependence of the LRC parameter is expressed as follows: qi =

qmiBi Pini n

4. RESULTS AND DISCUSSION 4.1. Adsorption Breakthrough Dynamics in Activated Carbon Bed. The adsorption dynamics of the activated carbon bed were obtained from breakthrough experiments at a 5−9 atm adsorption pressure and 3.0−6.8 LPM feed flow rate. Figure 3 shows the breakthrough curves of the ternary mixture on an activated carbon bed at 6.8 LPM and 5−9 atm.

n

1 + ∑ j = 1 Bi P j j

(8)

qm = k1 + k 2T , B = k 3 ek4 / T , n = k5 + k6/T

(9)

Since a few experimental data points in the proper pressure range from the previous study23 were available for this study, the LRC parameters were obtained by fitting the simulated result to the experimental breakthrough results. And the parameters in the previous study23 were used as initial values. The LRC parameters and LDF coefficients for the PSA process simulation are listed in Table 3. Table 3. LRC (Langmuir−Freundlich) Parameters and LDF Coefficients for Each Component activated carbon k1 (mol/kg) k2 (mol/kg K) k3 (Pa) k4 (K) k5 (−) k6 (K) ω (1/s)

H2

CO

CO2

63.59 −0.1417 2.934 × 10−10 1490 1.192 −90.75 0.70

7.916 −0.0131 3.429 × 10−09 1854 1.312 −100.0 0.075

28.26 −0.0766 2.372 × 10−09 2741 1.592 −189.6 0.028

Solutions of mathematical models of PSA systems often contain sharp profiles with different velocities. When the Danckwerts boundary conditions3,9,24 are applied to a cyclic system, discontinuity in the connected beds is often observed. In this study, the following modified boundary conditions were applied to the connected beds (z = 0 is the feed end and z = L is the product end) to overcome the discontinuity.11,25 The boundary conditions for the PR, AD, and PG steps are ∂yi (α) ∂z

= 0,

∂T (α) = 0, u(β) = −u in , P(β) = Pin ∂z

Figure 3. Breakthrough curves at various adsorption pressures in an activated carbon bed at a 6.8 LPM feed flow rate.

The breakthrough time increased with the adsorption pressure. The increase in breakthrough times for both CO and CO2 when adsorption pressure increased from 7 to 9 atm was lower than those when adsorption pressure increased from 5 to 7 atm. However, the difference in breakthrough times for CO and CO2 increased with adsorption pressure. The variation of CO breakthrough times with adsorption pressure was smaller than that for CO2 because the CO isotherm variation with pressure was smaller than the CO 2 isotherm variation in the experimental pressure range. The CO breakthrough curve was sharper than the CO2 curve because of the competitive adsorption and strong adsorption affinity of CO2. In addition, the roll-up phenomenon was observed in each CO breakthrough curve. The roll-up phenomenon is caused by the weak adsorbate losing adsorption sites during competitive adsorption with the stronger adsorbate. The roll-up of CO stemmed from the strong CO2 adsorption. In addition, the height of CO roll-up and the length of roll-up plateau increased with adsorption pressure because the increase in the amount of adsorbed CO2 with increasing adsorption pressure was higher than that of CO. The simulated breakthrough times agreed with

(10)

where α is z = L and β is z = 0 for the PR and AD steps, and “in” = “feed.” For the PG step, α is z = 0, β is z = L, and “in” = “purge.” The boundary conditions for the DPE and BD steps are ∂yi (α) ∂z ∂yi (β) ∂z

= 0,

∂T (α) ∂P(α) = 0, u(α) = 0, =0 ∂z ∂z

(11)

= 0,

∂T (β) =0 ∂z

(12)

where α is z = L and β = 0 for the BD step and α is z = 0 and β is z = L for the DPE step. The boundary conditions for the PPE step are yi (L) = yin, i , T (L) = Tin , u(L) = −u in , P(L) = Pin ∂yi (0) ∂z

= 0,

∂T (0) =0 ∂z

(13)

(14) D



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the experimental results while the simulated CO roll-up height deviated from the experimental result. Figure 4 shows the variation in breakthrough dynamics due to changes in the feed flow rate at a fixed adsorption pressure of

Figure 5. Temperature breakthrough curves in an activated carbon bed at 9 atm adsorption pressure and a 6.8 LPM feed flow rate.

Figure 4. Breakthrough curves at various feed flow rates in an activated carbon bed at 9 atm adsorption pressure.

9 atm. The reduction in breakthrough time with an increase in feed flow rate was smaller as the flow rate increased from 4.5 to 6.8 LPM than from 6.8 to 9.1 LPM. The difference in MTZ (mass transfer zone) of CO and CO2 narrowed as flow rate increased. The decrease in breakthrough time for both components was almost proportional to the increase in flow rate. The shape of the breakthrough curve was almost constant with changing flow rate. These results indicated that the mass transfer limitation had a relatively minor effect on the breakthrough dynamics in the range of experimental flow rates. Like the adsorption pressure results in Figure 3, the simulated breakthrough time agreed with the experimental results, but the simulated CO roll-up was higher than the experimental result. According to the adsorption pressure and flow rate results, the deviation in CO roll-up may have stemmed from overestimating the CO isotherm or deviation of heat of adsorption on activated carbon. Figure 5 shows the temperature variation in a bed during the breakthrough experiment at 9 atm adsorption and a 6.8 LPM feed flow rate. Each temperature curve was obtained at 10 cm, 30 cm, 50 cm, and 75 cm from the feed end. Significant temperature excursions (about 30 K) occurred at the initial part of the adsorption bed. After the excursion, two temperature excursion peaks were observed, and they were clearly separated along the axial position of the bed due to the difference in propagation of CO and CO2. The higher adsorption affinity and amount of CO2 led to slower temperature propagation and higher temperature excursion (second temperature peak in Figures 6b−d) even though there was less CO2 than CO in the feed. Over time, the temperature profile in the bed decreased due to heat exchange with the surrounding environment and

Figure 6. Temperature variation with time at a cyclic steady-state of an activated carbon PSA process at 9 atm adsorption pressure, 6.8 LPM feed flow rate, and 0.393 LPM purge rate (Run01 in Table 2).

the continuously supplied feed gas. Competitive adsorption is accompanied by the heat of desorption for the less strongly adsorbed component and the heat of adsorption for the more strongly adsorbed component.23,26 Accordingly, the temperature decrease took a long time, and the bed temperature did not reach the inlet temperature until 2500 s. As a result, the CO2 breakthrough curve had a more gentle slope with a smoother increase than did the CO breakthrough curve due to the temperature profile, even though the adsorption affinity of CO2 was much stronger than that of CO. E



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stemmed from not only from the isotherm, but also from bedinside heat effects. 4.2. Two-Bed H2 PSA Performance. In this study, six-step two-bed PSA experiments (described in Figure 2) were carried out under various operating conditions, such as adsorption time, feed flow rate, adsorption pressure, and P/F ratio. The P/F ratio and recovery of the PSA process was evaluated as follows:

As shown in Figure 5, the front wave of the temperature profile could be predicted by the mathematical model while the amount of temperature excursion and profile was overestimated by the simulation. The difference in temperature excursion between experimental and simulated results was over 6 K. Because the average of heat of adsorption of each component was used, it led to a deviation in temperature profile. In addition, because a flange type of bed was used, heat transfer at the flange resulted in some deviation, especially at the product end, as shown in Figure 5d. This temperature deviation had a more significant effect on the breakthrough curve of the weak adsorbate, CO. Therefore, the deviation in the CO roll-up recovery =

purge/feed (P/F) ratio =

(product eluted from AD step − product used in PG step) × product purity feed from PR and AD steps × composition of H2 in feed

from 90 to 120 s, the change in recovery and productivity was slightly higher than that as the AD step time changed from 120 to 180 s. As shown in Figure 6d, the main temperature excursion at 75 cm was observed between 120 and 180 s. A certain amount of impurities (CO and CO2) is expected to propagate into the end of the bed during the AD step time, especially at an AD step time of 180 s. Therefore, as the adsorption step time increased, the product purity decreased drastically with a relatively small increase in recovery. Reducing the adsorption step time from 120 to 90 s resulted in higher H2 purity and productivity at the expense of recovery. As shown in Figure 8, in general, a higher P/F ratio resulted in higher product purity because more of the product was used to regenerate the bed.4,27 The recovery changed linearly with the change in P/F ratio while the variation in purity above 99% was small as the P/F ratio increased. An increase in P/F ratio from 0.05 to 0.1 can significantly push the CO MTZ back, which propagated faster than the CO2 MTZ, to the feed end (Figures 4 and 5). Therefore, the adsorption bed could separate H2 from the mixture more efficiently. However, the effect of a further increase in P/F ratio on removing MTZs of impurities was significantly reduced. Therefore, as shown in Table 4, H2 product of more than 99.99% purity was accompanied by significant sacrifices in recovery and productivity.

productivity =

amount of H 2 used in PG step amount of H 2 inlet in AD step

H 2(L) from AD step − H 2(L) used in PG step total cycle time (s) × adsorbent used (g)

Figure 6 shows the temperature profile at each position inside the bed over the time when the PSA process (Run01 in Table 2) reached a cyclic steady state. In Figure 6b, the temperature rise caused by the heat of adsorption was observed at the PR (10 s) and AD steps (10−130 s). As the CO and CO2 desorption occurred in the bed during the DPE step (for 130−135 s), the temperature decreased slightly. Then the desorption in the BD step (135−145 s) led to a significant temperature decrease. Further desorption in the PG step (145−265 s) was indicated by a continuous decrease in temperature, but the temperature variation was small after a certain period of time. A temperature rise in the bed occurred due to adsorption of the strong adsorbate during the PPE step (265−270 s). The temperature at the feed end, in Figure 6a, decreased with time during the AD step because of heat transfer with the surrounding environment and the supplied feed. The temperature increase during the PPE step was clear at the feed end. In the middle of the bed (Figure 6c), a temperature profile was not developed because the MTZs of CO and CO2 were not fully propagated. At 75 cm (Figure 6d), a small increase in temperature was observed at the end of the BD step and the beginning of the PG step because of readsorption of the desorbed adsorbate. However, because the temperature variation at this position was small, this position worked as a purifier instead of a bulk separator, which differed from other parts of the bed. As mentioned above, H2 purity needs depend on the application. Therefore, understanding a wide range of purities, recoveries, and productivities is important depending on operating variables such as step time, adsorption pressure, feed flow rate, and purge rate. The operating conditions to obtain high hydrogen purity (99%+) from effluent gas should be investigated because such purity is widely applied for various fields. Figure 7 shows the effect of AD step time on PSA performance of Run02, 04, and 05 in Table 2. H2 purity of higher than 99% could be obtained at an AD step time of 90 s. The purity and recovery decreased almost linearly with an increase in AD step time. However, as AD step time changed

Figure 7. Effect of adsorption step time on H2 purity and recovery at 9 atm adsorption pressure, 6.8 LPM feed flow rate, and 0.393 LPM purge rate (Runs 02, 04, and 05 in Table 2). F



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depressurization. Therefore, an adsorption pressure of 7 atm produced the highest productivity at the same feed flow rate and AD step time. Figure 10 shows the effect of feed flow rate on purity and recovery. The purity was 99.86% at a feed flow rate of 4.5 LPM

Figure 8. Effect of P/F ratio on H2 purity and recovery at 9 atm adsorption pressure, a 6.8 LPM feed flow rate, and 120 s adsorption time (Run02, 06, and 07 in Table 2).

Table 4. Experimental Results of the Six-Step PSA Process run number

purity (%)

recovery (%)

Run01 Run02 Run03 Run04 Run05 Run06 Run07 Run08 Run09

99.86 96.48 87.86 99.54 83.87 99.51 99.99 88.04 75.43

68.27 83.91 87.02 76.27 91.68 69.49 49.29 89.63 90.99

productivity (L/g·s) 8.563 1.269 1.608 1.298 1.307 1.035 7.896 1.338 1.217

× × × × × × × × ×

Figure 10. Effect of feed flow rate on H2 purity and recovery at 9 atm adsorption pressure, 120 s adsorption time, and a 0.393 LPM purge rate (Run 01, 02, and 03 in Table 2).

10−5 10−4 10−4 10−4 10−4 10−4 10−5 10−4 10−4

and decreased to less than 87.0% at a feed flow rate of 9.1 LPM. The recovery and productivity increased with feed flow rate, as shown in Table 4. The increase in feed flow rate could advance CO and CO2 MTZs to the product end. Simultaneously, it decreased the P/F ratio to the PSA. Therefore, fewer impurities were removed from the bed as feed flow rate increased. The change in purity with increased feed flow rate was similar to that with a decrease in P/F ratio, as shown in Figure 8. However, the change in recovery variation was more similar to the change in AD step time, as shown in Figure 7, than to the P/F ratio.

5. CONCLUSION A two-bed PSA process using activated carbon was studied to recover hydrogen from effluent gas (H2/CO/CO2, 39.3:35.4:25.3 vol.%) of melting incineration for waste. Before designing the PSA process, the adsorption dynamic behaviors at various pressures and feed flow rates in the activated carbon bed were investigated by breakthrough experiments. The purity, recovery, and productivity of the PSA process were evaluated under various operating conditions. The differences in CO and CO2 breakthrough time increased as adsorption pressure increased. The height of the CO roll-up and length of the roll-up plateau increased with adsorption pressure because the amount of CO2 adsorbed increased more than that of CO. In addition, the difference in CO and CO2 MTZs narrowed as flow rate increased. However, the reduction in the breakthrough time of both components was almost proportional to the increase in flow rate. The breakthrough results indicate that the PSA performance was mainly controlled by CO propagation, but the heat of CO2 adsorption should be considered in the PSA design. In the PSA results, the end part of the bed played the role of adsorbing CO because CO MTZ propagated much faster than CO2 MTZ along the bed during the adsorption step. Therefore, potentially, a mixture of CO and H2 without CO2 can be

Figure 9. Effect of adsorption pressure on H2 purity and recovery at a 6.8 LPM feed flow rate, 120 s adsorption time, and 0.393 LPM purge rate (Run02, 08, and 09 in Table 2).

Similarly, as adsorption pressure increased, the purity increased and recovery decreased, as shown in Figure 9. However, the variation in purity decreased slightly with an increase in adsorption pressure, and the variation in recovery decreased with a decrease in adsorption pressure (Table 4) because the CO isotherm has only slight curvature in the range of adsorption pressures. The decreased recovery was mostly due to hydrogen lost in the feed end during countercurrent G



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hi, ho = heat transfer coefficient of inside and outside walls (J/cm2 s K) k = parameter for LRC model Kz = effective axial thermal conductivity (J/m s K) L = bed length (cm) P = pressure (Pa) qi, qi*, qi̅ = amount adsorbed, equilibrium amount adsorbed, and average amount adsorbed of species i, respectively (mol/g) qm = equilibrium parameter for Langmuir−Freundlich model (mol/g) R = gas constant (Pa·m3/mol K) Rp = radius of pellet (cm) RBi, RBo = inside and outside radius of the bed, respectively (cm) t = time (s) T, Tw, Tatm = solid and gas phase temperature and wall temperature of surroundings, respectively (K) u = interstitial velocity (cm/s) yi = mole fraction of species i (−) z = axial distance in bed from the inlet (cm)

obtained as a second product following the recovery of a high purity H2. The purity and recovery varied asymptotically with the P/F ratio and AD step time, respectively, because the bed purification by H2 purge gas approached a limitation. Other variations in purity and recovery with P/F ratio, AD step time, and adsorption pressure were almost linear or slightly curved. Therefore, the change in productivity was different from that of purity and recovery. Because the variation in feed flow rate led to the change of the P/F ratio and changed the propagation in impurities during PSA operation, the change in purity was similar to that in P/F ratio while the variation trend of recovery was similar to that in the AD step time. The productivity increased with increased recovery and decreased with increased purity. The study showed the feasibility of the PSA process to produce a wide purity range of H2 from the effluent gas in the melting incineration process. The PSA process in this study could produce H2 that was 75.43−99.99% purity with 49.29− 90.99% recovery. PSA to produce more than 99.99% purity H2 was accompanied with a significant loss of recovery and productivity. However, if a syngas (H2 and CO) is needed to supply to the users, the PSA using activated carbon can be performed with high recovery because the main impurity in the product is CO. Alternatively, a certain level of H2/CO mixture in the effluent gas from the PSA process can be recycled to the melting incinerator as a fuel. And the quality of the product gas and the performance improvement of a PSA process using advanced adsorbents can be decided by the economic analysis of the total melting incineration process.

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AUTHOR INFORMATION

REFERENCES

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Corresponding Author

*Tel.: +82-2-2123-2762. Fax: +82-2-312-6401. E-mail: leech@ yonsei.ac.kr. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS This work was supported by a grant from the “International Collaborative Energy Technology R&D Program” of the Korea Institute of Energy Technology Evaluation and Planning (KETEP) funded by the Korea government Ministry of Knowledge Economy (No. 20118510020030)

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DEDICATION The authors submit this original research paper to I&EC ́ E. Research as a collection of papers honoring Professor Alirio Rodrigues of the University of Porto in recognition of his 70th birthday.

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ε = voidage of adsorbent bed (−) ρg, ρs = gas density and pellet density, respectively (g/cm3) μ = viscosity (Pa·s) ν = superficial velocity (cm/s) ω = LDE coefficient (s−1)

NOMENCLATURE

Alphabets

Aw = cross-sectional area of the wall (cm2) Bi = equilibrium parameter of species i for Langmuir− Freundlich model (atm−1) Ci = component concentration of species i in the bulk phase (mol/cm3) Cpg, Cps, Cpw = gas, pellet, and wall heat capacity, respectively (J/kg K) DL = axial dispersion coefficient (cm2/s) −ΔHads = average heat of adsorption (J/mol) H



Article

(12) Rajesh, J.; Gupta, S.; Rangaiah, G.; Ray, A. Multi-objective optimization of industrial hydrogen plants. Chem. Eng. Sci. 2001, 56 (3), 999−1010. (13) Hufton, J.; Mayorga, S.; Sircar, S. Sorption-enhanced reaction process for hydrogen production. AIChE J. 1999, 45 (2), 248−256. (14) Kim, Y. H.; Lee, D. G.; Moon, D. K.; Byeon, S.-H.; Ahn, H. W.; Lee, C. H. Effect of bed void volume on pressure vacuum swing adsorption for air separation. Korean J. Chem. Eng. 2014, 31 (1), 132− 141. (15) Lee, J. J.; Kim, M. K.; Lee, D. G.; Ahn, H.; Kim, M. J.; Lee, C. H. Heat-exchange pressure swing adsorption process for hydrogen separation. AIChE J. 2008, 54 (8), 2054−2064. (16) Da Silva, F. A.; Rodrigues, A. E. Vacuum swing adsorption for propylene/propane separation with 4A zeolite. Ind. Eng. Chem. Res. 2001, 40 (24), 5758−5774. (17) Santos, M. n. P.; Grande, C. A.; Rodrigues, A. E. Dynamic Study of the Pressure Swing Adsorption Process for Biogas Upgrading and Its Responses to Feed Disturbances. Ind. Eng. Chem. Res. 2013, 52 (15), 5445−5454. (18) Silva, F. A. D.; Rodrigues, A. E. Propylene/propane separation by vacuum swing adsorption using 13X zeolite. AIChE J. 2001, 47 (2), 341−357. (19) Sundaram, N.; Wankat, P. Pressure drop effects in the pressurization and blowdown steps of pressure swing adsorption. Chem. Eng. Sci. 1988, 43 (1), 123−129. (20) Malek, A.; Farooq, S. Kinetics of hydrocarbon adsorption on activated carbon and silica gel. AICHE J. 1997, 43 (3), 761−776. (21) Ahn, H.; Kim, M. B.; Lee, C. H. Effects of Heat-Transfer Coefficients on Thermal Dynamics in a Near-Adiabatic Fixed Bed. Sep. Sci. Technol. 2004, 39 (11), 2627−2654. (22) Sereno, C.; Rodrigues, A. Can steady-state momentum equations be used in modelling pressurization of adsorption beds? Gas Sep. Purif. 1993, 7 (3), 167−174. (23) Jee, J.-G.; Kim, M.-B.; Lee, C.-H. Adsorption characteristics of hydrogen mixtures in a layered bed: binary, ternary, and fivecomponent mixtures. Ind. Eng. Chem. Res. 2001, 40 (3), 868−878. (24) Jee, J. G.; Lee, S. J.; Kim, M. B.; Lee, C. H. Three-bed PVSA process for high-purity O2 generation from ambient air. AIChE J. 2005, 51 (11), 2988−2999. (25) Kim, M.-B.; Bae, Y.-S.; Choi, D.-K.; Lee, C.-H. Kinetic separation of landfill gas by a two-bed pressure swing adsorption process packed with carbon molecular sieve: nonisothermal operation. Ind. Eng. Chem. Res. 2006, 45 (14), 5050−5058. (26) Yang, J.; Han, S.; Cho, C.; Lee, C.-H.; Lee, H. Bulk separation of hydrogen mixtures by a one-column PSA process. Sep. Technol. 1995, 5 (4), 239−249. (27) Rege, S. U.; Yang, R. T.; Qian, K.; Buzanowski, M. A. Airprepurification by pressure swing adsorption using single/layered beds. Chem. Eng. Sci. 2001, 56 (8), 2745−2759.

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Pressure Swing Adsorption Process for Recovering H2 from the

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