Ind. Eng. Chem. Res. 1997, 36, 903-914
903
Bubble Column Reactors for Wastewater Treatment. 3. Pilot-Scale Solvent Sublation of Pyrene and Pentachlorophenol from Simulated Wastewater Jeffrey S. Smith† and Kalliat T. Valsaraj* Department of Chemical Engineering, Louisiana State University, Baton Rouge, Louisiana 70803
Removal efficiencies from a three-phase continuous, countercurrent solvent sublation of pyrene and pentachlorophenol from water are presented and compare well with Series CSTR Model predictions. The effects of the air-to-water flow ratio, the solvent flow rate, and the gas sparger design are discussed. Steady-state efficiencies as high as 96% for pyrene and 94% for pentachlorophenol are reported for an annular shear gas sparger. These results demonstrate the superiority of solvent sublation over bubble fractionation which is shown to be limited to a removal efficiency of 68% or less for these compounds. Moreover, analyses of the solvent (mineral oil) phase show that the separation factor (ratio of the solvent effluent concentration to the water effluent concentration) in solvent sublation can be as high as 300, which is distinctly greater than that observed in solvent extraction. Introduction In wastewater treatment plants, aeration devices, such as diffused aerators and mechanical surface aerators, are often used to facilitate the biological oxidation of organic compounds. In the case of diffused aeration, the process closely resembles that of a short, squat bubble column reactor with multiple air injection points. Recently (Chern and Yu, 1995), it has been reported that the air stripping which takes place in these devices is a serious and growing concern as waste treatment operators are being charged to further control their volatile organic compound (VOC) emissions. The authors stressed the need for improved mass-transfer models which describe VOC emission from these devices and introduced a new oxygen mass-transfer model. Although such activities are indeed very good first steps, the problem of preventing VOC emission from these devices still remains. Moreover, the current system fails to remove organic compounds which are resistant to biological oxidation such as polynuclear aromatic hydrocarbons and chlorinated pesticides and insecticides. Though regulations may not require action now, it is anticipated that new or improved aeration technology that reduces VOC emission as well as targets compounds which resist biodegradation will soon be required. Liquid-liquid extraction is an alternative for the removal of nonvolatile compounds if an appropriate solvent is available (Cusack, 1996). Solvent extraction, however, leaves significant residual solvent in the water which requires a subsequent separation step. One potential alternative to conventional solvent extraction is solvent sublation. The solvent sublation process, which was introduced in parts 1 and 2 of this series (Smith et al., 1996a,b), is a wastewater treatment technology that has the potential to remove volatile and nonvolatile compounds without the problem of residual solvent. Like the diffused aeration device, solvent sublation closely resembles a squat bubble column reactor and is basically operated in the same way; however, the only difference * To whom correspondence should be addressed. E-mail:
[email protected]. Telephone: (504) 388-6522. † Separations Expertise Center, Dow Corning Corp. No. 128, Midland, MI 48686. E-mail:
[email protected]. S0888-5885(96)00524-6 CCC: $14.00
is that a nonvolatile organic solvent is floated upon the aqueous phase. The advantage of the solvent is that it serves as a sink for VOCs as well as nonvolatile organic compounds (NVOCs) which are resistant to biodegradation. Moreover, in the sublation process, the surfaceactive nature of organic compounds is exploited. This phenomenon has a profound favorable effect on the removal efficiencies of these compounds, in particular strongly hydrophobic compounds. The reasons why efficiencies are higher in sublation than in other separation processes are because (i) the adsorbed phase and the bubble wake, which is entrained into the overlying solvent, experience solvent extraction, (ii) the overall driving force for mass transfer in the air-water dispersion is increased, and (iii) the mechanism of adsorption to the bubble-water interface circumvents the air-side mass-transfer resistance which is the dominant resistance for hydrophobic compounds (of low vapor pressures) in conventional air-water contactors. In this paper, results obtained from a pilot-scale sublation process used to treat aqueous streams containing pyrene and pentachlorophenol are reported. The effects of the air-to-water flow ratio, the gas sparger design, and the Henry enhancement factor on fractional removal are discussed. The Henry enhancement factor, which was defined in part 1, is the ratio of the effective Henry law constant, which arises from the combination of bulk phase partitioning and surface adsorption, to the conventional Henry law constant. The results show that solvent sublation is distinctly superior to bubble fractionation (diffused aeration). Furthermore, the Series CSTR Model (SCM) developed in Part 1 is tested for predicting sublation performance. Experimental Section Materials and Methods. The equipment used in this investigation included a 4 in. i.d. × 60 in. long (0.1016 × 1.524 m) sublation column, equipped with both a fine porous glass frit and an annular shear sparger (0.1 ft2, 6.3 ft/s water jet) described in part 2 of this series. By using the two different spargers, we were able to investigate the effect of the different column hydrodynamics. Pyrene and pentachlorophenol (PCP) were selected as the target organics to be removed from the aqueous phase. Pyrene is an example of a poly© 1997 American Chemical Society
904 Ind. Eng. Chem. Res., Vol. 36, No. 3, 1997 Table 1. Operating Conditions and Steady-State Fractional Removal Results for the Solvent Sublation and Bubble Fractionation of Pyrene and Pentachlorophenol run I.D.
compound
sparger
R-73-9-21-94 R-79-9-27-94 R-80-9-29-94 R-87-10-12-94 R-88-10-13-94 R-89-10-15-94 R-92-11-1-94 R-94-11-15-94 R-97-11-17-94 R-99-11-30-94 R-100-12-1-94 R-103-12-7-94 R-106-1-18-95 R-113-2-9-95 R-117-3-7-95 R-120-3-21-95 R-124-4-6-95 R-16-6-22-95 R-24-9-20-95 R-30-9-26-95 R-33-9-29-95 R-37-10-3-95 R-40-10-9-95 R-43-10-10-95
pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene pyrene PCP (pH 9) PCP (pH 3) PCP (pH 3) PCP (pH 3) PCP (pH 3) PCP (pH 3) PCP (pH 3)
porous frit porous frit porous frit porous frit porous frit porous frit porous frit porous frit porous frit porous frit porous frit porous frit annular shear annular shear annular shear annular shear annular shear annular shear annular shear annular shear annular shear annular shear annular shear annular shear
R-131-6-13-95 R-58-11-27-95
pyrene PCP (pH 3)
annular shear annular shear
a
QA, SCCM
Qw, cm3/min
959 959 959 959 2225 4413 443 443 1592 2225 8307 959 2225 4093 5977 5355 2225 4732 4732 7156 7156 2840 2840 3454
Bubble Fractionation ultrasonic 5977 conductivity 7156
instr./control Sublation ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic ultrasonic conductivity conductivity conductivity conductivity conductivity conductivity
Qo, cm3/min
FRss
41.8 27.9 41.8 27.9 27.9 27.9 27.9 55.8 13.9 41.8 41.8 27.9 27.9 27.9 27.9 13.9 55.8 27.9 27.9 27.9 13.9 55.8 41.8 27.9
4 4 4 2 2 2 2 2 2 2 2 2a 2 2 2 2 2 2 2 2 2 2 2 2
0.809 0.863 0.796 0.862 0.892 0.896 0.771 0.684 0.893 0.833 0.901 0.832 0.887 0.945 0.962 0.956 0.882 0.127 0.853 0.907 0.942 0.744 0.783 0.858
27.9 27.9
no solvent no solvent
0.673 0.682
R-103 was conducted with “spent solvent”; that is, the solvent used in R-100 was used again.
nuclear aromatic hydrocarbon (PAH) which is a byproduct from coal gasification, and PCP is an example of a commercially manufactured insecticide. Both compounds pose a significant threat to the environment and are classified as priority pollutants by the U.S. Environmental Protection Agency under the Clean Water Act of 1977. Light-white mineral oil was chosen as the solvent phase as it has several advantages over other organic liquids. For instance, the vapor pressure is low; the density is 0.86 g/cm3; and there is no foul odor. The compounds used in study were obtained from SigmaAldrich, while distilled water and air were supplied by LSU plant utilities. Twenty-six experiments were conducted which investigated the effects of the air, water, and solvent flow rates, the gas sparger design, and the Henry enhancement factor. A summary of the experiments and operating conditions appears in Table 1. The water fed to the process was prepared by spiking containers of distilled water (≈19 L each) with the target organic from gravimetrically prepared acetonitrile (MeCN) solutions. For the pyrene experiments, the concentrations of the feed stocks ranged between 86 and 130 µg/L, while those for the PCP experiments ranged between 3 and 9 mg/ L, with most being equal to 6 mg/L. For those experiments which investigated the removal of neutral PCP, sulfuric acid was added to the feed stocks to adjust the pH between 2.5 and 3.0. Since the pKA of PCP is 4.74 (Montgomery and Welkom, 1990), this ensured that the PCP remained neutral throughout the experiments. In a typical experiment, aqueous influent and effluent samples were taken intermittently and analyzed for the organic compound being sublated. Each analysis included two replicates. Fractional removal was calculated as FR ) 1 - Aeff/Ainf, where Aeff/Ainf is the ratio of the average aqueous, effluent concentration to the average aqueous, influent concentration. A Hewlett-Packard 1046A fluorescent detector was used to determine the concentration of pyrene in water.
Table 2. Settings for Chromatographic Equipment Schimadzu Autoinjector injection volume (µL) 25 no. of replicates 3 HP 1090L HPLC (Reverse Phase) mobile phase 70-85% MeCN (variable) stationary phase Phenomenex Envirosep-PP (3.2 × 125 mm) flow rate (mL/min) 0.5 pressure (bar) 100 retention time (min) 4.5 HP 1046A Fluorescence Detector excitation wavelength (nm) 237 emission wavelength (nm) 385 response time (ms) 5 (2000) scan speed (nm/s) 6 (50) photomultiplier gain 12
The detector was connected in series with a HewlettPackard 3932A integrator, a Hewlett-Packard 1090L high-performance liquid chromatograph, and a Shimazdu autoinjector. The settings for the analytical equipment are listed in Table 2. A three-step procedure was used for preparing aqueous standards for calibration. First, a gravimetric solution of pyrene in MeCN (3192 mg/L) was prepared; second, 1000 µL of the gravimetric solution was diluted with 100 mL of MeCN to make a 30 mg/L solution; finally, the samples were further diluted with distilled water to make solutions ranging from 1 to 135 µg/L. The relative standard deviation (RSD) of the fluorescent detector was 13%, which after error propagation corresponds to a 2.5% RSD in fractional removal. For the analysis of pyrene in mineral oil, a HewlettPackard 8452 UV/vis spectrophotometer was used. A standard solution was prepared gravimetrically by dissolving 0.5 g of solid pyrene in 1 L of mineral oil. To calibrate the detector, aliquots of the standard were diluted with fresh oil to make solutions ranging from 0.5 to 7.40 ppm (wt). The analytical wavelength used
Ind. Eng. Chem. Res., Vol. 36, No. 3, 1997 905
Figure 1. Examples of fractional removal profiles obtained with (a and b) ultrasonic Level Control and (c) conductivity-based level control.
for measuring absorbence was 340 nm. The uncertainty of the measurements (95% C.I.) was 6.48%. For the analysis of PCP in water, the UV spectrophotometer was used to measure absorbence at 214 nm. A 520 mg/L solution of PCP in MeCN was prepared gravimetrically, and aliquots of this standard were diluted with distilled water to make solutions ranging from 0.5 to 10 mg/L. Fifty microliters of sulfuric acid was added to each standard so that the pH would be similar to the actual aqueous samples from the sublation column. The RSD for the fractional removal of PCP was 1.5%. Control Aspects. As described in part 2 of the series, the control of the solvent-water interface position at the top of the column required the most attention as the flows to the process were all well controlled with metering pumps and flow tubes (rotameter). An ultrasonic level switch, which measured the time for sound to traverse the two phases, and a solenoid valve were used in most of the experiments for controlling the interface position. This arrangement served reasonably well. Figure 1a shows a typical fractional removal profile obtained from the sublation process with the ultrasonic-based controller. On occasion, however, process upsets did occur and were the result of the ultrasonic switch failing to detect deviations from setpoint. For example, two upsets occurred during run R-88 as shown in Figure 1b. In both occurrences, the solvent was displaced with water because the switch failed to detect that the interface position had increased. This type of failure was more common than solvent infiltration because of the large Qw/Qo ratios used in the experiments. The process upsets were avoided with simple manual adjustments that included resetting the
ultrasonic switch and adjusting the height of the weir in the overflow reservoir. To eliminate the need for human intervention, a second control scheme, named DAC (data acquisition and control), was constructed. In the DAC system, the control action was based upon the electrical-conductivity profile in the overflow reservoir. As illustrated in Figure 2, six pairs of nichrome electrodes, encased in glass tubes, were immersed in the overflow reservoir to different depths. The probes were arranged to maximize the sensitivity near the interface. Each pair of probes was subjected to a 5 kHz ac potential. (High frequency was necessary as it prevented polarization of the probes.) A multimeter measured the conductances across the probes and transmitted the information to a PC where the interface position was calculated. A model-based control algorithm then determined the output to the solenoid valve. One of the unique features of the controller was that it simulated the valve action as if it had full dynamic range. (Details of the DAC system are given in the Appendix, which is available upon request.) With the DAC system, the control of the process was improved significantly as deviations about steady state were within (0.012 fractional removal points. Figure 1c shows a representative run where conductivity-based control was used. Deoxygenation Experiments for the Determination of KLav. Oxygen-based mass-transfer coefficients were determined with the sublation column operating in a semibatch mode (Qo ) 0, Qw ) 0). A YSI Model 55 oxygen meter was used to measure the dissolved oxygen in the water. In a typical experiment, air, at a known flow rate, was sparged into the column until the meter showed that the dissolved oxygen was near 8.6 mg/L, the solubility at 22 °C, and elevation at 0. (Distilled water was used.) Once steady state had been maintained for several minutes, the feed gas was switched from instrument air to nitrogen. The switch to nitrogen occurred quickly and smoothly without any fluctuation in the gas flow rate. This was possible because the two regulators controlling the source of air and nitrogen were fed into one common regulator which fed the column. During the experiment, the drop in dissolved oxygen was recorded with time. Plotting the data as ln Co/C versus time, as eq 1 suggests, resulted in a slope equal to KLav
ln
Co 3k t ∼ KLavt ) C a(1 - )
(1)
Results and Discussion Effect of Operational Variables (QA, Qw, Qo) on Fractional Removal. The most important variable affecting the sublation of organic compounds from water is the air-to-water flow ratio, QA/Qw. This ratio not only influences the extent of stripping but also contributes greatly to column hydrodynamic properties. In part 2, it was argued that the homogeneous flow regime is the appropriate operating regime for the sublation process. Thus, the values of QA and Qw shown in Table 1 were chosen to ensure homogeneous flow. In Figure 3, fractional removal results obtained from the sublation process are plotted versus QA/Qw. Each point represents a 15-30 h steady-state average. The general trend is that, as the flow ratio is increased, the fractional removal rapidly approaches an upper limit. This behavior is consistent with SCM (Series CSTR Model) predictions presented in part 1, which showed a very
906 Ind. Eng. Chem. Res., Vol. 36, No. 3, 1997
Figure 2. Diagram of conductivity probes and data acquisition and control (DAC) system used to control the solvent-water interface.
Figure 3. Fractional removal of pyrene and neutral pentachlorophenol (PCP) versus the air-to-water flow ratio, QA/Qw.
steep gradient between 0 e QA/Qw e 50 and a plateau at QA/Qw g 150. In fact, the data in the figure suggest the optimum operating condition is in the range of 200 e QA/Qw e 300, with the exact value depending on the compound and sparger type. Defining the optimum range is based on the observation that the gains made in fractional removal are very small when QA/Qw is increased beyond 300. Another reason is that ultimately, as the gas velocity is increased, an unfavorable transition occurs from homogeneous flow to heterogeneous flow. The turbulence associated with heterogeneous flow is undesirable, as it tends to disperse the solvent throughout the system. Sublation experiments were performed to examine the relative effect of the solvent flow rate versus the water flow rate. First, runs R-73 and R-79, which appear at the top of Table 1, were conducted to demonstrate the effect of a -33% step change in water flow rate while holding the solvent flow rate constant at 4 cm3/min. As one can see from the table, the fractional removal of pyrene increased from approximately 80% to 86% in response to decreasing Qw from 41.8 to 27.9 cm3/min.
Given the strong influence of the air-to-water flow ratio observed in Figure 3, this was not surprising. When the water flow rate was returned to its original value (run R-80), the fractional removal returned to 80%. A second experiment, run R-87, was performed which investigated the effect of reducing the solvent flow rate from 4 to 2 cm3/min while holding the water flow rate constant at 27.9 cm3/min. As shown in Table 1, for runs R-79 and R-87, there was no effect of reducing the solvent flow rate on the fractional removal. The reason for this is that the mineral-oil-partition coefficient (Kow) for pyrene is very large. In fact, since most hydrophobic compounds possess very large partition coefficients, the amount of solvent needed for sublation is only that which ensures a steady state. This was shown in part 1. Thus, the solvent flow rate used in most of the experiments was maintained at the lowest pump setting, 2 cm3/min. Because of the large capacity of the solvent, one variation of the process is to recycle the solvent. This is referred to as using “spent” solvent and is a way of minimizing the overall solvent usage. Therefore, to further show the weak dependence of the solvent, run R-103 was conducted using a feed solvent (oil) having a pyrene concentration of 1.57 mg/L; however, the operating conditions of R-103 were identical to those used in R-87. The effect of the spent solvent was to reduce the fractional removal by only 0.03. Since this difference was outside the statistical error limits ((0.02), the effect was statistically significant but, nonetheless, very weak. Effect of Gas Sparger on Mass Transfer. As illustrated in Figure 3, sublation experiments involving pyrene were conducted which investigated the effect of gas sparger type. For QA/Qw < 150, there was no significant difference between the fractional removal data collected from the shear sparger and those collected from the porous frit. However, for QA/Qw g 150, the fractional removal of pyrene was improved by as much as 5%, suggesting that the benefit of the shear sparger is only realized at large QA/Qw ratios. The reason for this behavior may be explained by considering the
Ind. Eng. Chem. Res., Vol. 36, No. 3, 1997 907
unique hydrodynamic properties of the shear sparger. It was observed in part 2, that, for gas velocities greater than about 1 cm/s, the processes of bubble coalescence and breakage were buffered by the combined actions of the water jet and the milky-white plume of microbubbles exiting the shear sparger. It was speculated that this so-called “buffering effect” was responsible for the weak bubble size dependence on gas velocity. As it turns out, in the cases where the shear sparger performed better than the frit, the gas velocities varied from 0.84 to 1.23 cm/s. For the low QA/Qw experiments where the removal efficiencies were about the same for both spargers, the gas velocities were less than or equal to 0.46 cm/s. To investigate whether or not the improved performance was just the result of operating at the higher gas velocities, run R-100 was conducted using the porous frit sparger and a gas velocity of 1.71 cm/s (QA/Qw ) 199). The result of this experiment was no different than those observed at 0.33 (QA/Qw ) 115) and 0.91 cm/s (QA/Qw ) 158) for the porous frit. Thus, one may reasonably conclude that the shear sparger does have an advantage over the frit for superficial gas velocities g 1 cm/s and that the effect is not simply the result of gas velocity. The advantage is attributed to improved hydrodynamic properties of the shear sparger and how they affect the air-water mass-transfer coefficient. The following discussion elucidates this relationship. Consider eq 2, which is the classical two-resistance model for the overall liquid-side mass-transfer coefficient. For many organic compounds such as VOCs,
1 1 1 ) + k κl Hcκg
(2)
air-side resistance is unimportant and the term 1/Hcκg is neglected. However, for strongly hydrophobic compounds such as pyrene and pentachlorophenol, air-side resistance is important, if not the dominant resistance, because of the very low Henry law constants, Hc, that they possess. Thus, the common practice of scaling oxygen mass-transfer coefficients (which are liquid-sidecontrolled) by the Schmidt number is not correct for these compounds. Therefore, other means must be used to determine k. One common approach for predicting the masstransfer coefficient for gas-liquid dispersions is to use a penetration theory such as the Higbie penetration model (Taylor and Krishna, 1993). In this model the average mass-transfer coefficient is defined as
κl )
x
4DAB ) πte
x
4DAB λ π ω
(3)
where DAB is the binary diffusivity and te is the exposure time. The exposure time may be considered as the time a small fluid element of uniform concentration comes in contact with a phase boundary through which mass transport occurs. Like film theory, where the film thickness is difficult to determine, the exposure time is also difficult to determine; however, in some cases, it can be calculated. As suggested in eq 3, the exposure time may be taken as the ratio of some characteristic length, λ, which the fluid element must traverse to some characteristic velocity, ω. For example, in the case of noninteracting bubbles rising up through a liquid (string bubbling flow regime), the exposure time is the mean
bubble diameter, ds, divided by the rise velocity, U (Taylor and Krishna, 1993). This is because the liquid in immediate contact with a gas bubble is replaced in a time equal to that for the mean bubble to rise one diameter. However, it is implied that the resistance to mass transfer resides totally in the liquid phase, that is, the phase which experiences the penetration. Thus, eq 3 is valid for single-phase resistance only. Additionally, it is implied that the bubbles rise in pure plug flow since the effects of backmixing or circulation are ignored. In order to generalize, the key questions which must be addressed are as follows: How does one combine gas-phase resistance into eq 3? How does one determine the exposure time when the flow regime is different from plug flow? Sherwood et al. (1975) derived that the total moles transferred during an exposure period for two resistances in series may be expressed as
∫0t N dt ) e
{x
( ( ) ( x ) )}(
4DABte DAB κi2te + exp × π κi DAB erfc κi
te -1 DAB
Cl -
)
Cg (4) Hc
where κi is the resistance at the interface due to a surfactant. In the derivation of eq 4, the authors assumed that the resistance 1/κi may be added to the liquid-side resistance in a fashion similar to the way gas-side resistance was added to eq 2. Thus, eq 4 may be used as a starting point to generate a general twophase penetration model describing the overall liquidside mass-transfer coefficient. In the spirit of the conventional film theory, we assume that the interface itself offers no resistance but that the gas phase offers significant resistance. Substituting 1/Hcκg for 1/κi and after some rearrangement, one can show
k ) Hcκg
x
4DAB πte exp(z2) erfc(z) - 1 + Hcκg z2
(5)
where z is defined as Hcκg(te/DAB)0.5, and Hc is the bulk phase Henry law constant. When z is approximately greater than or equal to 1, the “correction function”, f(z), becomes effectively zero and eq 5 reduces to eq 3; however, if z is
