Biological Oxidation of Air-Borne Volatile Organic Compounds by Pilot Sieve-Plate Absorption Tower

Abstract

An activated sludge aeration tank and a sieve-plate column with six sieve plates were utilized to remove gas-borne volatile organic compounds (VOCs) in air streams. The tank was used for the biodegradation of absorbed VOCs from the column that utilized the activated mixed liquor drawn from the tank as scrubbing liquor. This research proposed a model for VOC absorption to a down-flow activated sludge liquor in a sieve-plate column. Experiments were conducted and the model was verified based on results of tests on removal efficiencies of isopropyl alcohol, toluene, and p-xylene in the system operated at a range of influent VOC concentrations, air application rates, and liquid/gas flow ratios (L/G). Analysis of the model for effects of affecting parameters on VOC removal effectiveness indicates that L/G, plate number N, biodegradation rate constant k, Henry's law constant of VOC H are important.

Introduction

Biological treatment of volatile organic compounds (VOCs) or odors in air streams provides an inexpensive alternative to conventional technologies such as catalytic and thermal oxidation, wet scrubbing, ozonation, and activated carbon adsorption (Don and Feenstra, 1984; Ottengraf, 1986; Neal and Loehr, 2000).

Biological treatment methods include biofiltration, biotrickling, and bioscrubbing. Biofiltration needs solid carrier for immobilization of microorganisms and nutrients, VOCs in the introduced gas to the reactor diffuse to the carrier and are then degraded in biofilms attached to the carriers (Ottengraf, 1986; Devinny and Deshusses, 1999; Yang et al., 2010, 2018; Cheng et al., 2016a, 2016b). Biotrickling also needs solid carrier for immobilization of microbial films and an aqueous solution trickles over the films to provide the necessary nutrients and remove the metabolites (Moussavi and Mohseni, 2008; for example: Tu et al., 2015; Leili et al., 2017).

One of the bioscrubbing systems is the activated sludge aeration (ASA). Waste gas stream to be treated is introduced to an ASA tank for wastewater treatment and water-soluble contaminants in the gas stream are dissolved in the activated sludge mixed liquor and biodegraded therein. Bielefeldt and Stensel (1998, 1999) reported that based on a proposed mass-transfer model, the system with an aeration submerged liquid depth of >2.0 m provides >80% removal for biodegradable VOCs with dimensionless Henry's law coefficients (H) of <0.35. VOCs with the Henry's law coefficients include common solvents such as methanol, ethanol, methyl isobutyl ketone, methyl ethyl ketone, acetone, methylene chloride, and BTEX (benzene, toluene, ethyl benzene, and xylenes).

A related study demonstrated that the treatment efficiency for gas-borne BTEX removal in a laboratory-scale activated sludge tank of 40 cm liquid depth was not significantly impacted by different BTEX mixtures, and ∼99% removal was achieved for volumetric loadings of 11–18 g BTXE/(m³ liquid.h) (Bielefeldt and Stensel, 1998; Burgess et al., 2001). Burgess et al. (2001) extensively reviewed this process and pointed out that by using existing activated sludge plants as bioscrubbers, the foul air generated by other process units of the wastewater treatment system could be treated on site, with no requirement for additional units or interruption of the wastewater treatment (Burgess et al., 2001). The process has been shown to remove hydrogen sulfide in biogas from 2,000 to <20 ppm (Neal and Loehr, 2000).

Although the ASA process has been shown to be effective and economical, the ASA process can only be applied to the treatment of foul air streams of low to medium flow rates such as a few to several tens of cubic meter per minute. One of the main reasons may be the limitation of the permissible aeration intensity of <0.5 cubic meter of gas per cubic meter of aeration tank volume per hour, as specified by ASA system design (Metcalf and Eddy, 1991).

Another similar process is the “tower scrubbing-activated sludge oxidation (SAO)” (Overcamp et al., 1993; Croonenberghs et al., 1994; DeHollande et al., 1998; Hammervold et al., 2000). This SAO process generally consisted of a wet scrubber and a biological oxidation unit. The scrubber absorbs gas-borne pollutants into a biomass slurry such as activated sludge liquor from biological wastewater treatment, and the microorganisms in the oxidation unit oxidize the absorbed pollutants for regeneration of the slurry to retain its absorption ability. The biomass slurry is continuously circulated between the two units.

Overcamp et al. (1993) developed an integrated theory to describe the steady-state operation of a suspended-growth bioscrubber for the control of biodegradable, volatile organic gases. The bioscrubber consists of an N-stage absorber and an oxidation reactor. The biomass slurry is circulated between the absorber and the oxidation reactor, and the pollutant is absorbed and partially oxidized in the absorber. Oxidation is completed in the oxidation reactor. Predictions of the theory show that the removal efficiency is a function of Henry's law constant for the pollutant, the ratio of the liquid flow rate to the gas flow rate, and the number of stages. However, the theory has not been experimentally verified.

Croonenberghs et al. (1994) studied the removal of ethanol in a 12.6 Nm³/min vent gas from a brewery wastewater treatment plant by a single-stage packing tower scrubbed with an activated sludge mixed liquor containing a suspended biomass solid concentration of 500 mg/L. An average VOC removal of 85% was observed with an average influent ethanol concentration of ∼70 mg/Nm³ and a liquid to gas flow ratio (L/G) of 0.0075 m³ liquid/(m³ gas).

Hammervold et al. (2000) used a sorptive slurry biological scrubber to study the control of air-borne acetone. The scrubber had an inner diameter of 15 cm and was equipped with three sieve plates with 5-mm diameter holes of 50% opening area. Using scrubbing liquid with mixed liquor suspended sludge (MLSS) concentrations of 1,440–1,590 mg/L, a gas superficial velocity of U = 0.71 m/s, and an L/G of 0.00267 m³/m³, an average acetone removal efficiency of ∼90% was obtained.

In general, like a wet scrubber, a packed or sieve-plate tower can be served as a biological scrubber and a superficial gas velocity of 1.0–1.5 m³/(m²·s) [equivalent to an air mass loading rate of ∼90–1,300 lb/(ft²·h) for air at 25°C] can be used as a design basis for the cross-sectional area of the biological scrubber (Perry et al., 1994). Hence, a scrubber with a cross-sectional area of 1 m² can be used to treat a foul gas with a flow rate of 60–90 m³/min. The scrubber can thus be used to clean a medium to high flow rate of waste gas if the attached biological oxidation unit has enough capacity to degrade the absorbed pollutants.

As stated previously, the SAO process has the advantage of treating much more gas flow than the ASA with the same activated sludge tank volume. However, there are limited studies focused on the influencing parameters to the VOC removal efficiency of the SAO process.

This research attempted to propose a model for absorbing gas-borne VOCs into a down-flow activated sludge liquor in a sieve-plate column. Experiments were also conducted to verify the model based on the results of tests on the removal efficiencies of water-soluble isopropyl alcohol (IPA) and relatively water-insoluble toluene and p-xylene in a pilot-scale system operated at a range of influent VOC concentrations, air application rates, and L/G ratios. The three VOCs are frequently used in manufacturing paint, printing ink, and synthetic chemicals. Results of this study allow a confirmation of the model and provide a design basis for treating gases contaminated with a range of VOCs.

Model Development

A model was developed in this study to simulate the steady-state performance of a suspended-growth bioscrubber composed of an absorber and a separate oxidation reactor as given in Fig. 1. The biomass slurry is circulated between the absorber and the activated sludge column for oxidation of the absorbed VOCs. The pollutants are both absorbed and partially oxidized in the absorption unit. The extent of pollutant oxidation in the absorber depends on the concentration of biomass, the pollutant degradation rate, and the total hold-up liquid volume in the plates of the absorber. The oxidation reactor allows sufficient liquid residence for the oxidation of the remaining pollutant and sufficient biomass residence time to allow for their growth.

FIG. 1.

Nomenclature for material balances and gas absorption in a N-stage sieve-plate column.

The mechanistic model developed by Bielefeldt and Stensel (1998, 1999) describes the removal of VOCs from a contaminated gas stream sparged into a completely mixed activated sludge reactor as a function of the gas–liquid mass-transfer rate and liquid VOC concentrations. By referring to Fig. 1, considering the hold-up aerated liquid in the stage n as a reactor, application of the model gives: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} \left( { { y_n } - y_n^* } \right) = \left( { { y_ { n - 1 } } - y_n^* } \right) \exp \left( { { \frac { - { \rm { \alpha } } { K_ { \rm { L } } } { a_ { { \rm { VOC } } } } Z } { H ^\prime G / A } } } \right) \tag { 1 } \end{align*} \end{document}

In addition, a material balance over plate n for the VOC removal from the gas phase and the VOC transferred to the liquid and biodegraded therein gives: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} G{ \rm{ }} \left( {{y_{n - { \rm{1}}}} - {y_n}} \right) = L ( {x_n} - {x_{n + { \rm{1}}}} ) + { \rm{ }}{r_n}{V_n} \tag{2} \end{align*} \end{document}

In Equations (1) and (2), A is the cross-sectional area of the sieve plate available for gas flow (m²), G is the volumetric gas flow rate (m³/h), H′ is the dimensionless apparent Henry's law coefficient, L is the volumetric recirculating liquid flow rate (m³/h), V_n is the hold-up aerated liquid volume in plate n (m³), Z is the hold-up aerated liquid depth (m), K_La_VOC is the overall mass-transfer coefficient of the VOC in clean water (1/h), α is a coefficient for correcting K_La_VOC to the case in the activated sludge mixed liquor, r_n is the volumetric rate of VOC biodegradation (mg/m³ h), y is the gas-phase VOC concentration (mg/m³), x is the liquid-phase VOC concentration (mg/m³), and y* is the gas-phase VOC concentration in equilibrium with the liquid phase with a VOC concentration of x. Equation (2) states that the gas-phase VOC decrease rate G(y_n _− 1 − y_n) in stage n equals the net VOC L(x_n − x_n _+ 1) flowing down to the downstream stage n − 1, and the remaining is biodegraded with a rate of r_nV_n. α is ∼1 for mixed liquors with suspended solids concentration of ∼2,000–3,000 mg/L generally used in activated sludge systems (Bielefeldt and Stensel, 1999).

\documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $$y_n^*$$ \end{document} in Equation (1) is related to x_n by a modified Henry's law for partitioning of a specific VOC between gas and an activated sludge mixed liquor at an equilibrium condition: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} y_n^* = H ^\prime {x_n} \tag{3} \end{align*} \end{document}

A correlation model between H′ and H, the Henry's law constant for partitioning of the VOC between gas and clean water has been derived and verified in this laboratory (Lin and Chou, 2006): \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} H ^\prime = H \frac { 1 } { { 1 + { K_ { OC } } { X_d } + { K_ { SS } } { X_S } } } \tag { 4 } \end{align*} \end{document}

where K_OC and K_SS are the partition coefficients of VOC between the dissolved organic carbons and the suspended solids (solid activated sludge) in the liquid and the aqueous phase, respectively; and X_d and X_S are the dissolved organic carbon and the suspended solid concentration in the liquid phase, respectively. According to the model, in general, H′ is a little smaller than the Henry's law constant H for partitioning of the VOC between gas and clean water at the same condition for relatively water-insoluble VOCs such as toluene and xylene, but somewhat higher for water-soluble VOCs such as IPA in this study. It has been shown that H′ for toluene and p-xylene decreases up to 77% and 93%, respectively, in the mixed liquor with activated sludge concentrations of ∼40,000 mg/L. A higher X_S thus helps in the absorption of water-insoluble VOCs and upgrades the performance of the column for VOC removal. In general, H′≈H for X_S < 3,000 mg/L in an activated sludge mixed liquor. The r_n can approximately be related to x_n by a first-order kinetics at a fixed activated sludge or microbial concentration: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} {r_n} = k{x_n} \tag{5} \end{align*} \end{document}

where k is the rate constant (1/h).

For highly volatile compounds such that gas film resistance to mass transfer is relatively smaller than the liquid transfer, K_La_VOC can generally be related to \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} , the overall volumetric mass-transfer coefficient of oxygen in clean water, and the ratio of the molecular diffusivities of the VOC compounds to the dissolved oxygen in water: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} { K_ { \rm { L } } } { a_ { { \rm { VOC } } } } = { K_ { \rm { L } } } { a_ { { { \rm { O } } _ { \rm { 2 } } } } } { \left( { { \frac { { D_ { { \rm { VOC } } } } } { { D_ { { { \rm { O } } _ { \rm { 2 } } } } } } } } \right) ^ { m } } \tag { 6 } \end{align*} \end{document}

where D_VOC is the diffusion coefficient of the VOC in water (m²/h), \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${D_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} is the diffusion coefficient of oxygen in water (m²/h), and m is a constant. Depending on the system power intensity, m ranges from 0.5 to 1.0; 1.0 provides a conservative design (Bielefeldt and Stensel, 1999).

For semivolatile compounds (SVOCs) such that the gas film resistance to the overall mass transfer is evident, the correlation presented in Equation (6) should be corrected as follows (Hsieh, 2000): \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} { K_ { \rm { L } } } { a_ { { \rm { VOC } } } } = { K_ { \rm { L } } } { a_ { { { \rm { O } } _ { \rm { 2 } } } } } { \left( { { \frac { { D_ { { \rm { VOC } } } } } { { D_ { { { \rm { O } } _ { \rm { 2 } } } } } } } } \right) ^ { m } } \left( { { \frac { { R_L } } { { R_T } } } } \right) \tag { 7 } \end{align*} \end{document}

where R_L and R_T are liquid phase and total resistance for VOC transfer, respectively. By the two-film theory, \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} { \frac { { R_L } } { { R_T } } } = \frac { 1 } { { 1 + { k_ { \rm { L } } } { a_ { { \rm { VOC } } } } / H { k_ { \rm { G } } } { a_ { { \rm { VOC } } } } } } \tag { 8 } \end{align*} \end{document}

where k_La_VOC and k_Ga_VOC are the individual liquid- and gas-phase mass-transfer coefficient for VOC transfer, respectively (Perry et al., 1994). As an approximation, from penetration or surface-renewal theory: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} { \frac { { k_ { \rm { L } } } { a_ { { \rm { VOC } } } } } { H { k_ { \rm { G } } } { a_ { { \rm { VOC } } } } } } \approx \frac { { { { \left( { { D_ { { \rm { VOC } } } } / { D_ { { \rm { VOC , G } } } } } \right) } ^ { 1 / 2 } } } } { H } \tag { 9 } \end{align*} \end{document}

where D_VOC,G is the diffusion coefficient of VOC in air (m²/h). In general, (D_VOC/D_VOC,G)≈(10⁻⁵ cm²/s/10⁻¹ cm²/s) = 10⁻⁴, k_La_VOC/Hk_Ga_VOC≈1/(100H) and R_L/R_T≈1/(1 + 0.01/H) from Equation (8) (Perry et al., 1994). For VOCs with H > 0.2, R_L/R_T > 0.95, Equation (6) can be used instead of Equation (7) for estimating K_La_VOC from \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} . However, for SVOCs with H < 0.14, R_L/R_T < 0.93, and gas film resistance is greater than ∼7% of the total resistance and should be considered. For IPA at 25°C, R_L/R_T is ∼0.93 by the data D_VOC = 1.04 × 10⁻⁵ cm²/s, D_VOC,G = 0.098 cm²/s, and H = 0.00615. R_L/R_T is ∼0.96 for toluene (D_VOC = 8.5 × 10⁻⁶ cm²/s, D_VOC,G = 0.087 cm²/s, H = 0.25) and m-xylene (D_VOC = 7.8 × 10⁻⁶ cm²/s, D_VOC,G = 0.07 cm²/s, H = 0.22) at the same temperature. For oxygen at 25°C, D_VOC = 2.11 × 10⁻⁵ cm²/s (Metcalf and Eddy, 1991; Eastem Research Group, 1997). Equation (6) may be safely used for K_La_VOC estimation.

For a sieve plate with an available cross-sectional area A, a hold-up aerated liquid depth Z, and operated at volumetric gas flow rate G, V_n≈Z × A and G≈A × U, the superficial gas velocity over the cross-section. By substituting Equations (3) and (5), H′ = H, α = 1, V_n = Z × A, and G = U × A into Equations (1) and (2), it can be shown that: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} ( { y_n } - H { x_n } ) = ( { y_ { n - 1 } } - H { x_n } ) \exp \left( { \frac { { - { \rm { } } { K_ { \rm { L } } } { a_ { { \rm { VOC } } } } } } { U } \frac { Z } { H } } \right) \tag { 10 } \end{align*} \end{document} \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} \left( { { y_ { n - 1 } } - { y_n } } \right) \; = \; \left( { \frac { L } { G } } \right) \left( { { x_n } - { x_ { n + 1 } } } \right) { \rm { } } + k { x_n } \left( { \frac { Z } { U } } \right) \tag { 11 } \end{align*} \end{document}

By referring to Fig. 1, with given K_La_VOC/U, Z/U, H, L/G, k, influent liquid VOC concentration x_N_+ 1, and specified effluent gas VOC concentration y_N, x_N, and y_N_− 1 could be calculated from Equations (10) and (11). The calculated x_N and y_N _− 1 are then again substituted into the two equations to obtain x_N _− 1 and y_N_− 2. By repeated substitution of x_N and y_N_− 1 into Equations (10) and (11), effluent liquid VOC concentration x₁ and influent gas VOC concentration y₀ could finally be obtained. For a column of fixed specifications and influent gas flow rate, if the calculated y₀ is smaller than that of the influent gas to be treated, an increase of L/G could give a higher value of calculated y₀, as a higher scrubbing liquid for a gas flow increases the VOC absorption capacity of the plate column. For the same case, if the calculated y₀ is higher, a smaller L/G should be used.

By Equations (10) and (11), parameters affecting column performance are K_La_VOC (Z/U), H, x_N_+ 1, L/G, Z/U, and k. For activated sludge mixed liquor with a higher MLSS, H in Equation (10) should be replaced by H′ to take the effect of MLSS on the VOC solubility into consideration. In addition, k is also affected by the MLSS concentration of the absorbing liquor. According to the model, an increase of any one of K_La_VOC (Z/U), L/G, Z/U, and k, or a decrease of H (or H′) or x_N_+ 1 increases the column absorption capacity for the gaseous VOC.

Materials and Methods

Experimental apparatus

The experimental setup consisted of a pilot-scale activated sludge tank and a sieve-plate tower, as given in Fig. 2. The activated sludge tank was constructed from a 40 × 40 × 300 cm (W × L × H) acrylic column and equipped with an air sparger with 20 orifices of 2 mm diameter. The sieve-plate tower was constructed from a 25 × 25 × 162 cm (W × L × H) acrylic column with six custom-made sieve plates. Each plate has 382 holes, 3 mm in diameter arranged on a square pitch. The holes give an open area of 3.82% of the whole plate area for gas flow. Two 25 mm i.d. down-comer pipes were also equipped to allow for the down-flow of the activated sludge liquor. Ports were provided at the column inlet, outlet, and each plate for gas and liquid sampling.

FIG. 2.

(a) Schematic experimental system. 1, Ambient air; 2, air blower; 3, water bath; 4, liquid VOC; 5, gas flowmeter; 6, circulation pump; 7, nutrient solution; 8, metering pump; 9, sieve-plate column; 10, activated sludge tank; 11, vent gas; 12, gas sampling port; 13, liquid sampling port; 14, sieve plate. (b) Plan view of the sieve plate and a part of cut-away view of the sieve-plate column. VOC, volatile organic compound.

Air drawn by a blower was bubbled through a liquid VOC (of toluene, p-xylene, and IPA) in a water-bathed Erlenmeyer flask to generate a VOC-rich air stream. VOC concentrations in the stream were varied by adjusting the temperature of the water bath or the air flow rate. The VOC-rich air was then mixed with a main air stream, which subsequently flowed into the bottom space of the sieve-plate tower. Air flow rates were measured by the rotameter and regulated by the valve. Constant flow rates of activated sludge liquor for gas scrubbing was provided by the activated sludge tank and pumped into the top plate of the column. An average aerated liquid depth (Z) of 5.0 or 8.5 cm could be kept in each sieve plate by adjusting the weir height of the down-comer pipe. VOC-rich effluent liquor from the column bottom was circulated back to the tank for VOC oxidation.

Materials

Reagent-grade toluene, p-xylene and IPA were used as target VOCs. Activated sludge for seeding and acclimation was taken from the wastewater plant of Ta-Lin Refinery, Chinese Petroleum Co. (Kaohsiung City, Taiwan). The sludge was adopted because it exhibited the potential to oxidize the target VOCs. Table 1 lists the VOC influent rates and the solution composition for the activated sludge nutrition. The nutrient solution was daily prepared and fed continuously into the activated sludge tank. All nutrients and salts were reagent grade.

Table 1.

Nutrition and Volatile Organic Compound Influent Rates to Experimental Activated Sludge Tank

Target VOC	MLSS (mg/L)	Nutrient	Loading rate (kg/m³ day)
Target VOC	MLSS (mg/L)	Nutrient	Acclimation	VOC removal test
Toluene	1,800	Toluene (as COD equivalent)	0.009–0.018	0.025–0.131
		Glucose (as COD equivalent)	0.431	0.431
		K₂HPO₄-P	0.0063	0.00625
		NH₂CONH₂-N	0.0123	0.01232
p-Xylene	1,700	p-Xylene (as COD equivalent)	0.010–0.021	0.002–0.035
		Glucose (as COD equivalent)	0.416	0.416
		K₂HPO₄-P	0.0060	0.00604
		NH₂CONH₂-N	0.0118	0.01179
IPA	2,200	IPA (as COD equivalent)	0.004–0.009	0.020–0.098
		Glucose (as COD equivalent)	0.531	0.531
		K₂HPO₄-P	0.0077	0.0077
		NH₂CONH₂-N	0.0152	0.0152

COD equivalents for glucose and the three VOCs are based on the reaction stoichiometries: (1) C₆H₁₂O₆ (glucose) +6 O₂ = 6 CO₂ + 6 H₂O, (2) C₇H₈ (toluene) +9 O₂ = 7 CO₂ + 4 H₂O, (3) C₈H₁₀ (p-xylene) +10.5 O₂ = 8 CO₂ + 5 H₂O, (4) C₃H₈O (IPA) +4.5 O₂ = 3 CO₂ + 4 H₂O. The calculated equivalents are 1.07 kg COD/(kg glucose), 3.13 kg COD/(kg toluene), 3.17 kg COD/(kg p-xylene), and 2.4 kg COD/(kg IPA). VOC influent rates are based on the volumetric mass rates (volumetric influent VOC concentration × volumetric gas flow rate × 8 h/day ÷ liquid volume) for 8 h per day and the others (glucose, N, and P) are for 24 h per day during acclimation.

COD, chemical oxygen demand; IPA, isopropyl alcohol; MLSS, mixed liquor suspended sludge; VOC, volatile organic compound.

Operation

Laboratory experiments were performed to measure (1) the clean water \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} for the aerated hold-up liquid on the sieve-plate as a function of the superficial gas velocity U, and (2) the removal efficiencies of each of the three VOCs in the sieve-plate column operated at the operations listed in Table 2.

Table 2.

Experimental Conditions

Target VOC	H (Note)	L/G (m³/m³)	Influent VOC concentration C₀ (mg/m³ at 300K)	Average aerated liquid depth for each sieve-plate Z (cm)
Toluene	0.273	0.107	225	5.0
		0.080	450	8.5
		0.064
		0.053
p-Xylene	0.22	0.620	150	8.5
		0.310	450
		0.155
		0.103
IPA	0.00615	0.001	150	5.0
		0.002	450
		0.003
		0.005

Literature data of Henry's law constants (H) at 25°C in clean water (Eastem Research Group, 1997).

L/G, liquid/gas flow ratio.

The \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} of clean water was determined by the ASCE Standard Method (ASCE Standards, 1984; Hammervold et al., 2000). It was calculated from a DO versus time data, using linear regression (Perry et al., 1994): \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} \ln \left( { { \frac { D { O^ * } - D { O_0 } } { D { O^ * } - DO } } } \right) = { K_ { \rm { L } } } { a_ { { { \rm { O } } _ { \rm { 2 } } } } } t \tag { 12 } \end{align*} \end{document}

where DO* is the saturated dissolved oxygen concentration in the liquid at the liquid temperature, and DO₀, and DO are, respectively, those values at the beginning of aeration and after an aeration time of t. For simplicity, only \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} of the bottom plate were measured at aerated liquid depths of 5.0 and 8.5 cm for superficial gas velocities of 0.0267–0.160 m/s (96–576 m/h) that corresponded to gas flow rates of 100–600 L/min. During the measurements, clean water was circulated between the 44-L bottom circulating tank and over the bottom plate, and a DO probe with a meter (SensoDirect OXI 200; Lovibond, Germany) was used for measuring the time variations of the dissolved oxygen in the bottom water. Values of \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} were then estimated for the bottom plate. It was assumed that the measured \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} of the bottom plate could be good approximations to the value for the rest five plates.

In the acclimation experiment for VOC removal, the activated sludge tank was initially filled with 320 L of the seed sludge with an MLSS concentration of ∼1,940 mg/L and continuously fed with nutrients (Table 1). A constant air flow rate of 10 L/min was maintained for each full acclimation period of 30 days. Glucose was used as a sole carbon source when there was no VOC in the influent air. The carbon nutrition rate was based on an MLSS concentration of 2,000 mg/L and a food-to-microorganism loading of F/M = 0.30 g COD (chemical oxygen demand)/(g MLSS.day). N and P were added according to a mass ratio of COD:N:P = 150:5:1. Daily doses of glucose, N, and P were dissolved in 15 L tap water and fed into the sludge liquor over 24 h. During this phase, for producing more acclimated sludge for VOC degradation, from around 9 a.m. to 5 p.m., the concentrations of influent VOC in the introduced air were in the range of 188–376, 216–433, and 122–245 mg/m³, for toluene, p-xylene, and IPA, respectively. From 5 p.m. to 9 a.m. the next day, there was no VOC in the introduced air. On a daily basis, 50 mL of the mixed liquor was sampled for MLSS determination. Thereafter, the air flow was stopped for 30 min to settle the sludge down and the supernatant was drained to remove half of the total original liquid volume. The removed liquid was replaced with an equal volume of tap water and the flows of air and the nutrient solution were restarted. The rate of liquid replacement gave an equivalent hydraulic retention time of 2 days for the liquid in the activated sludge tank. The operating conditions were pH = 7–8, MLSS = 1,860–2,380 mg/L, and T = 26°C ± 2°C for the activated sludge tank. A period of 30 days has been shown to be enough for acclimation of the sludge to the target VOCs (Chou and Chang, 2005).

After acclimation, experiments were performed to test the performance of the sieve-plate column for adsorption and biodegradation the influent VOC. Experimental conditions are given in Table 2. Influent gas flow rates were controlled in the range of 100–600 L/min (U = 0.0267–0.160 m/s or 96–576 m/h) at an increment of 100 L/min. For each gas flow rate, circulating liquid flow rates were controlled and varied to give L/G values as given in Table 2. For IPA, L/G was in the range of 0.001–0.005 m³/m³; whereas for toluene and p-xylene it was 0.053–0.620 m³/m³. Each condition was kept at least 3 h to let the system getting a quasi-steady state. After that time, a gas sample from each of the influent gas line and the overhead spaces of all the six plates was taken by an 1-mL gas-tight syringe for VOC analysis. In addition, liquid samples were taken from the inlet and outlet of the sieve-plate column and the activated sludge tank to determine the VOC concentration. After liquid sampling, the mixed liquid was immediately filtered through a 4.7-cm diameter glass fiber paper to separate the sludge and the filtrate was put in screw-thread glass vials with Teflon-lined caps and stored at 4°C until analysis.

Analytical methods

VOC concentrations in the gas samples were measured by a gas chromatograph (GC-14B, Serial No. 9663; Shimadzu) with a flame ionization detector. Gas samples with known VOC contents were used for preparation of the calibration curve. The dissolved VOC concentration in the liquid was determined by sealing a 3-mL liquid sample in a 43-mL sample vial, and the vial was then immersed in a water bath at 80°C. As was verified experimentally in another study, a heating time of 10–20 min was required to get a thermal equilibrium between the liquid sample and the water in the bath (Chou and Chang, 2005). A gas sample taken from the tube was then analyzed to determine the VOC concentration. The concentration of the absorbed VOC in the liquid sample could thus be determined. Liquid samples with known VOC contents were used for calibration. Instrument integrity was verified by injecting VOC standards at the beginning and end of each analytical session and comparing the results with the standard calibration curve. The detection limit was 1 mg/m for an injected gas sample of 1 mL. pH of the mixed liquor was measured by a pH meter (Model 761; Knick Co., Italy).

Results and Discussion

Volumetric oxygen- and VOC-transfer coefficients

Figure 3 provides \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} for clean water with an aerated liquid height of 8.5 cm in the bottom plate (Fig. 2a) of the experimental column. Table 3 lists H, D, and K_La_VOC calculated from Equation (6) at different superficial gas velocities U. Figure 3 also shows K_La_VOC versus U for the three test compounds. An m of 1 was used to give more conservative K_La_VOC values.

FIG. 3.

Variation of \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} and K_La_VOC for the bottom plate with superficial air velocity U across the experimental column. Symbol for \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} : experimental data; regressed line for \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} : \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $${K_{ \rm{L}}}{a_{{{ \rm{O}}_{ \rm{2}}}}}$$ \end{document} = 0.084 U + 175 with R² = 0.9977. K_La_VOC for IPA, p-xylene, and toluene: calculated from Equation (6). IPA, isopropyl alcohol.

Table 3.

H, D and K_La_VOC for Test Volatile Organic Compounds

Element or VOC	D (10⁻⁶ m²/h)at 25°C (Note 1)	H at 25°C (Note 1)	K_La (1/h) for U = 50–500 m/h (Note 2)
Oxygen	7.60	—	179–217
Toluene	2.23	0.273	53–64
p-Xylene	3.17	0.22	75–90
IPA	3.74	0.0062	88–107

Note 1, D for oxygen obtained from ASCE Standards (1984) and for the VOCs from (Eastem Research Group, 1997); H from (Eastem Research Group, 1997).

K_La_VOC calculated from Equation (6) \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} $$ { K_ { \rm { L } } } { a_ { { \rm { VOC } } } } = { K_ { \rm { L } } } { a_ { { { \rm { O } } _ { \rm { 2 } } } } } { \left( { { \frac { { D_ { { \rm { VOC } } } } } { { D_ { { { \rm { O } } _ { \rm { 2 } } } } } } } } \right) ^m } $$ \end{document} with m = 1.

As given in Fig. 3, K_La_VOC increases slightly with increasing U in the range of 50–500 m/h. This phenomenon might arise from the observed fact that many bubbles were formed so quickly and located so closely in the shallow liquid flowing over the plate that they tended to coalesce and the coalescing bubbles did not help in a significant increase of the specific gas–liquid interfacial area for VOC transfer. For design practice, it can reasonably be assumed that the K_La_VOC for a specific VOC to be a constant in the operating U of 50–500 m/h. In this study, however, K_La_VOC for each operating U was substituted into Equation (10) when verifying the model.

VOC treatment

Data on the variations in the effectiveness of VOC removal at a liquid depth of Z for a fixed U of 48–720 m/h and y₀ of 150–450 mg/m³ were obtained for the tested VOCs.

Figure 4 plots variations of toluene removal efficiency with plate number at fixed L/G ratios of 0.053 and 0.107 m³ mixed liquor/(m³·gas) for U = 576 and 288 m/h (influent gas flow rate = 600 and 300 L/min), respectively. Calculated values of the removal efficiency by the model Equations (10) and (11) are also shown in the figure for verification of the model. Table 4 lists the input parameters to the model equations. During the experiments, it was observed that there was only a trace of toluene in the influent to the column and accordingly x₇ = 0 was substituted into the model equations to obtain the calculated data. The kinetic constant k value for toluene degradation in the mixed liquor was estimated to be 702 1/h by the method of least squares from the experimental and calculated data.

FIG. 4.

Variations of toluene removal efficiency with plate number at fixed L/G and Z (symbols: experimental data; lines: calculated from the model). L/G, liquid/gas flow ratio. (a) for y₀ = 220 mg/m³; (b) for y₀ = 450 mg/m³.

Table 4.

Parameters Input to Model Equations (10) and (11)

Tested VOC	H at 25°C	Z (m)	k (1/h)	x₇ (mg/m³)	y₀ (mg/m³)	U (m/h)	L/G	K_La_VOC (1/h)
Toluene	0.273	0.085	700	0	225	288	0.107	58
					225	576	0.053	66
					450	288	0.107	58
					450	576	0.053	66
p-Xylene	0.22	0.085	490	598	450	48	0.620	75
p-Xylene	0.22	0.085	490	102	450	192	0.155	80
IPA	0.0062	0.050	1,000	829	150	480	0.005	106
IPA	0.0062	0.050	1,000	1,021	150	720	0.001	116

Figure 4a shows that at ∼25–27°C and an influent gas toluene concentration of y₀ = 225 mg/m³, the toluene removal efficiency increased approximately linearly with the increasing plate number n. The higher L/G value of 0.107 (m³ liquid)/(m³ gas) gave a higher toluene removal efficiency than the lower one of 0.053 for the same operating conditions other than L/G. This can easily be understood that more mixed liquor absorbed more toluene from the influent gas although the absorption rate was not linearly proportional to the L/G because the driving force (y − Hx) for mass transfer was not proportional to that. Figure 4b with y₀ = 450 mg/m³ shows similar results as Fig. 4a. The limiting removal efficiencies could be attributed to the limited solubility of toluene in the circulating mixed liquor, the low L/G values, and the limited total plate number of the column. Material balance over the whole column if there were no toluene biodegradation in the column liquid gives the following: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} G \left( {{y_0} - {y_N}} \right) \; = \;L \left( {{x_1} - {x_{N + 1}}} \right) \tag{13} \end{align*} \end{document}

The down-flowing circulating mixed liquor is unable to further absorb any VOC if y₀ is in equilibrium with x₁, that is, \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} G \left( {{y_0} - {y_N}} \right) \; = \;L \left( {{y_0} / H - {x_{N + 1}}} \right) \tag{14} \end{align*} \end{document}

The maximum VOC removal efficiency η_max of the column in this case with N = 6 and x_N _+ 1 = x₇ = 0 is thus: \documentclass{aastex}\usepackage{amsbsy}\usepackage{amsfonts}\usepackage{amssymb}\usepackage{bm}\usepackage{mathrsfs}\usepackage{pifont}\usepackage{stmaryrd}\usepackage{textcomp}\usepackage{portland, xspace}\usepackage{amsmath, amsxtra}\usepackage{upgreek}\pagestyle{empty}\DeclareMathSizes{10}{9}{7}{6}\begin{document} \begin{align*} \eta & = \;\left[ {\left( {{y_0} - {y_N}} \right)/{y_0}} \right]{\rm{ }} \times \;100\% \;\; = \;\;{\eta _{{\rm{max}}}} \\& \quad = \;\left[ {\left( {L/GH} \right) - \left( {L/G} \right){\rm{ }}\left( {{x_{N{\rm{ + 1}}}}/{y_0}} \right)} \right] \times 100\% \tag{15} \end{align*} \end{document}

Figure 4 shows that at an L/G of 0.107, the experimental η_max was 42% and 48% for y₀ = 225 and 450 mg/m³, respectively, and the calculated η_max according to Equation (15) is 39.2% (L/GH = 0.107/0.273 = 0.392) for both y₀. Biodegradation of toluene in the column liquid might result in the slightly higher experimental η_max than the calculated ones. Similar results were observed at an L/G of 0.053. The phenomenon of higher experimental η_max than the calculated ones implies that biodegradation in the absorbing column cannot be neglected in the model.

Figure 5 shows variations of p-xylene removal efficiency with plate number at fixed L/G ratios of 0.155 and 0.620 for U = 192 and 48 m/h (influent gas flow rate = 200 and 50 L/min), respectively. Calculated removal efficiencies by the model equations using parameters shown in Table 4 are given in the figure for comparison. The kinetic constant k value for biodegradation of the absorbed VOC transferred to the mixed liquor was estimated to be 490 1/h by the method of least squares. By Equation (15) and the parameters shown in Table 4, η_max were calculated to be 67% and 100% for L/G of 0.155 and 0.620, respectively. Compared η_max with the experimental η_N = 6 of 34% (L/G = 0.155) and 58% (L/G = 0.620), it is apparent that additional plates can be used to absorb more p-xylene from the influent gas to improve the removal efficiency.

FIG. 5.

Variations of p-xylene removal efficiency with plate number at fixed L/G and Z (symbols: experimental data; lines: calculated from the model).

Figure 6 shows variations of experimental and calculated IPA removal efficiencies with plate number at fixed L/G ratios of 0.001 and 0.005 for U = 720 and 480 m/h (influent gas flow rate = 750 and 500 L/min), respectively. Parameters for the calculations are given in Table 4. The kinetic constant k value for biodegradation of IPA in the column was estimated to be 1,000 1/h by the method of least squares. By Equation (15) and the parameters given in Table 4, η_max was calculated to be 15.4% and 77.8% for L/G of 0.001 and 0.005, respectively. Experimental data show that a single plate of aerated liquid depth Z = 0.05 m was required to achieve an IPA removal of >96% for both L/G at the high superficial velocities of 480 and 750 m/h. The Henry's law constant of IPA is 0.00615 at 25°C. Substitution of parameters for IPA listed in Table 4 into Equation (10) gives (y_n − Hx_n)/(y_n − 1 − Hx_n) = 0.136 with U = 480 m/h, L/G = K_La_VOC = 106 1/h, y_n/y_n − 1 = 0.168. The removal could achieve 85% with only one sieve plate, and 97.2% with two plates connected in series. The high efficiency could attribute to the far low Henry's law constant of IPA as compared with toluene and p-xylene. Biodegradation of IPA in the column and the relatively lower Henry's law constant of the VOC could be among the major reasons for the high removal efficiency.

FIG. 6.

Variations of IPA removal efficiency with plate number at fixed L/G and Z (symbols: experimental data; lines: calculated from the model).

The novelty and advantage of the present process is that it could achieve a target removal of a certain VOC with the calculated plate number by a certain operation conditions. The limitations may be it needs a higher construction cost and maintenance labor to clean the plates.

Conclusions

In this study, experiments and theoretical analysis were conducted to determine and predict the removal efficiency of toluene, p-xylene, and IPA from air streams by a perforated-plate column scrubbed with circulating activated sludge mixed liquor. From the results of experiments and model application, the following conclusions can be drawn.

First, a stage-wise perforated-plate column is applicable for absorption of VOCs from waste gas streams by introducing an activated sludge mixed liquor as a scrubbing liquor and partial oxidation of the absorbed VOCs occurs in the column.

Second, a model developed by a material balance for the gaseous and liquid VOC over each plate of the column was developed and experimentally verified in this study. Superficial gas velocity over the column plate (U), plate number (N), volumetric liquid-phase VOC-transfer coefficient (K_La_VOC), aerated liquid depth over the plate (Z), volumetric liquid/gas flow rate ratio (L/G), dimensionless Henry's law constant of the VOC to be absorbed (H), VOC content of the influent scrubbing liquor (x_N _{+ 1}), and the biodegradation rate constant of the VOC in the activated sludge mixed liquor (k) are among the affecting parameters to the effectiveness of the VOC removal.

Nomenclature

α = correction factor

A = cross-sectional area of the sieve plate available for gas flow (m²)

D_VOC = diffusion coefficient of VOC in water (m²/h)

D_VOC, _G = diffusion coefficient of VOC in air (m²/h)

G = volumetric gas flow rate (m³/h)

H′ = dimensionless apparent Henry's Law constant

H = Henry's Law constant

k_La_VOC = individual volumetric VOC-transfer coefficient in liquid phase (1/h)

k_Ga_VOC = individual volumetric VOC-transfer coefficient in gas phase (1/h)

K_La_VOC = overall volumetric VOC-transfer coefficient in clean water (1/h)

K_OC = partition coefficients of VOC between the dissolved organic carbons in the liquid and the aqueous phase

K_SS = partition coefficients of VOC between the suspended solids (solid activated sludge) in the liquid and the aqueous phase

k = biodegradation rate constant (1/h)

L = volumetric recirculating liquid flow rate (m³/h)

m = constant (0.5 – 1.0)

MLSS = mixed liquor suspended solids concentration (mg/L)

R_L = liquid-phase resistance for VOC transfer

R_T = total resistance for VOC transfer

r_n = volumetric rate of VOC biodegradation (mg/m³ hr)

V_n = hold-up aerated liquid volume over plate n (m³)

U = superficial gas velocity over plate cross section (m/h)

X_d = dissolved organic carbon concentration in the liquid phase (mg/m³)

X_s = dissolved suspended solid concentration in the liquid phase (mg/m³)

x = liquid-phase VOC concentration (mg/m³)

x_n = VOC concentration in effluent liquid from plate n (mg/m³)

y = gas-phase VOC concentration (mg/m³)

y_n = VOC concentration in effluent gas from plate n (mg/m³)

y_N = VOC concentration in effluent gas form column (mg/m³)

y₀ = VOC concentration in influent gas (mg/m³)

y* = gas-phase VOC concentration in equilibrium with the liquid phase with a VOC concentration x (mg/m³)

Z = hold-up aerated liquid depth (m)

Footnotes

Author Disclosure Statement

No competing financial interests exist.

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