Ind. Eng. Chem. Res. 2001, 40, 3361-3368
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Removal of Ammonia from Aqueous Systems in a Semibatch Reactor Marc H. Heggemann and Hans-Joachim Warnecke Chemie & Chemietechnik, Technische Chemie & Chemische Verfahrenstechnik, Universita¨ t Paderborn, Warburgerstrasse 100, Paderborn, Germany
Hendrik J. Viljoen* Department of Chemical Engineering, University of Nebraska, 224 Avery Laboratory, Lincoln, Nebraska 68588-0126
Ammonia has been stripped from aqueous solutions in a semibatch reactor with steam. This reactor system has proven to be an economical design for smaller users. Data have been collected in a laboratory-scale reactor at different conditions of steam flow rates, liquid volumes, pH, and temperature. A mathematical model of the reactor has been developed, and the experimental results are compared with theoretical values. The liquid phase is described by either a perfectly mixed model or a finite dispersion model. The steam is modeled as a plug-flow system. The best match between experimental results and the model has been obtained for the finite dispersion model, and this model is proposed for design and scale-up applications. 1. Introduction In the past the emissions of larger operations have been closely monitored, but environmental stewardship is now also expected from smaller entities. Groundwater contamination enjoys closer scrutiny because of pressure from animal confinement operations, purification plants, and urban waste storage facilities to expand. Large quantities of nitrogen-bearing salts are released into the ground where partial oxidation by microorganisms into nitrate pollutes drinking water and adversely affects ecosystems. Large volumes of contaminated wastewater accumulate at treatment plants, feed lots, and waste storage facilities. Another concern is uncontrolled effluent due to seepage from wastewater lagoons and storage tanks. One of the major contaminants is ammonia, and its removal is an important unit operation in wastewater treatment. The agricultural sector has lagged other sectors in the implementation of wastewater treatment protocols. There are several reasons for this, but it is an undeniable fact that the agricultural sector is in need of a technology to remove ammonia from wastewater in an economical but technically simple manner. Research at the University of Paderborn under the auspices of the Westfa¨lisches Umweltzentrum (WUZ, Westphalian Environmental Center) has been focused on the development of an attractive technology to fill this need. Several unit operations have been tested and compared. The basis for this comparison includes operating cost, complexity of operation, capital layout, versatility, and ability to serve fluctuations in load and volume.1,2 This paper represents the result of this study, and it is proposed that the semibatch reactor offers the most practical method to remove ammonia from wastewater at agricultural operations. 1.1. State of the Art. Different physicochemical processes for ammonia removal are in use worldwide. A biological process (nitrification/denitrification) is one possibility, but it is limited to low ammonium concentrations.3,4 Microorganisms oxidize ammonia to nitrate * Author to whom correspondence should be addressed.
under aerobic conditions. Nitrate is reduced to elementary nitrogen under anaerobic conditions. The oxidation step is done batchwise, but the anaerobic breakdown is normally done in a continuous way. However, smaller agricultural operations do not prefer continuous processes. Other disadvantages of this process are as follows: The maximum concentration of ammonium in the wastewater is only 0.1 g/L (NH4+-N). Furthermore, the valuable compound ammonia is destroyed during the denitrification step. Moreover, the design and operation of this process is not without complications, and the conversion of ammonia itself is slow compared to competing processes. Precipitation of magnesium ammonium phosphate (MAP) is a further method for the removal of ammonia.3 The main investment cost is a mixer-settler vessel, but the operating costs are high compared to those of other methods. This process will not be able to compete with other methods on an economical basis. The major operating cost is the purchase of magnesium phosphate. Although MAP is a potential fertilizer, it has not been commercially applied as such, and the success of this method hinges on the value of MAP. Membrane techniques such as reverse osmosis generally are dirt-sensitive continuous processes.5-7 Solids in the liquid have to be removed by sand filtration for instance. Wastewater or similar liquids usually have a high total salt concentration. Thus, high operation pressures are needed for this process, which again makes it an unprofitable method. Air stripping is another possibility to remove ammonia.3,4,8-10 For this process columns with a large number of plates are necessary, because the transfer is sluggish. A more exotic reactor type is a plate scrubber system.6 The pH of the liquid is raised by adding a base in order to shift the ammonia/ammonium equilibrium toward the ammonia. Ammonia is blown out by air. The evaporation energy for ammonia from aqueous systems is needed which has to be taken from the liquid itself. A further problem in this process is the contaminated
10.1021/ie010106i CCC: $20.00 © 2001 American Chemical Society Published on Web 06/21/2001
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Figure 2. Semibatch reactor.
2. Objective Figure 1. Influence of pH and temperature on the distribution of NH3 and NH4+ in water.
air at the outlet. This ammonia containing air has to be washed by mineral acids such as sulfuric acid. Inert gas stripping is a comparatively expensive method.11 It is normally used in bubble columns for stripping of volatile organic compounds (VOCs), which are sensitive to oxidation. Thus, this process is not used in connection with the removal of ammonia because of the costs. Steam stripping is similar to the above-mentioned air treatment.3,4,9,10,12 The application of steam stripping by an aerosol reactorsa novel type of reactor for the removal of ammoniaswas investigated.1 Although the system has excellent mass-transfer characteristics, problems have been experienced with fouling and clogging of the steam nozzle. This design lacks the robustness to be used in an agricultural setting. Columns with a large number of theoretical plates have also been investigated. The pH is raised by adding a base, e.g., lime (a cheap commodity compared to magnesium phosphate). The equilibrium is shifted toward the NH3, and removal is more effective.
NH3 + H2O h + NH4+ OHammonia water ammonium hydroxide In this case steam is blown through the liquid phase. The removal of ammonia from aqueous systems involves mass transfer of NH3 from the liquid into the gaseous phase. The pH decreases during the removal, and consequently the equilibrium changes. When the pH drops, the equilibrium changes from ammonia to ammonium over a relatively small pH region. This region marks the limit of NH3 removal. In Figure 1 the variation in the equilibrium NH3 concentration is plotted as a function of pH. The equilibrium is also affected by temperature. On the basis of equilibrium arguments, it is more advantageous to operate the reactor at higher temperatures. Remark: If one wants to assign a single value to the transition region, the inflection point serves as a good indicator of the limiting concentration that will remain after stripping. Usually the starting pH in wastewaters is too low for effective stripping; hence, the pH is raised before steam is fed into the reactor. After leaving the reactor, the ammonia-containing steam is condensed as highly concentrated ammonia water, which can be used for further processes, e.g., for reduction of NOx to N2 and H2O in waste combustion plants and power stations. This process is economical if steam is available, but generation of steam is cheap and condensation heat can be recovered.
The high ammonia concentrations in the wastewater require a powerful removal method. Diverse approaches to contact the liquid and gaseous phases have led to a variety of reactor designs and operating procedures. In this study we will focus on a steam-stripping semibatch operation from wastewater. The advantages of this design are the high efficiency despite the costs to maintain a low pressure and steam generation. The vacuum increases the relative volatility of ammonia. Even heating of the liquid is not really necessary, and the removal efficiency is still high.1 The steam provides the evaporation energy that is necessary for the mass transfer of ammonia. Higher investment costs compared to those of competitive methods (e.g., air stripping) are more than off-set by the higher efficiency of this type of process. Also, the method presented here reduces the processing time compared to air-stripping methods. The authors hold the belief that the stirred semibatch reactor is an attractive system, especially for highly contaminated liquids. The stirrer promotes the dispersion of the steam phase and creates a high masstransfer area. The vessel is not very susceptible to fouling or dirt and is easy to clean because of the absence of internal packings and other fixtures. In addition, the reactor layout is simple and space requirements are modest, it has a high operating reliability, lower investment costs compared with fully continuous plants, and reasonable operating cost, it leaves nearly no residual waste, and the operation is extremely flexible. These features make this process attractive for smaller and medium-sized companies. 3. Mathematical Model of the System A schematic of the reactor is shown in Figure 2. Steam is fed at the bottom at a rate of n˘ S,in (mol/s) where a stirrer disperses the steam. The mixer has flat paddles to support tangential flow and at the same time induces minimal axial motion. Its major role is to distribute steam evenly at the inlet, but backmixing must be kept to a minimum to maximize the driving force for mass transfer. Therefore, only moderate stirring was employed, to minimize the backmixing. Under these conditions, the gas phase behaves as in a bubble column. It is assumed that the steam bubbles have uniform diameter φ (m) and this diameter does not change; i.e., hydrostatic pressure changes and NH3 transfer into the bubble do not affect the bubble dimensions. Furthermore, we do not consider radial or angular variations in concentration. While the bubbles rise to the top, NH3 diffuses across the liquid/gas boundary layers and the NH3 concentration in the gaseous phase is given by
(
)
∂cG cG ∂cG ) -v + Ka cL ∂t ∂x H
(1)
Ind. Eng. Chem. Res., Vol. 40, No. 15, 2001 3363
v denotes the bubble velocity, and it will be discussed later in this section. The last term on the right-hand side describes the mass transfer between the two phases. The rate of mass transfer depends on the structure of the boundary layers and the microscopic transfer processes in these layers. It is difficult to make an experimental assessment; thus, we resort to a theoretical description. The mechanism of mass transfer within these layers is by molecular diffusion as described by Fick’s law.13 The boundary layer thickness on the liquid side is δL, and on the gaseous side, it is δG. The quotient of the molecular diffusion coefficient Di and the boundary layer thickness δi is equal to the mass-transfer coefficient ki (i ) L, G). The concentrations of the transferred compound at the phase boundary are assumed to be in equilibrium; the Nernst law gives
cG* ) HcL*
(2)
The distribution coefficient H is an equilibrium constant. At steady state the mass flow on the liquid side n˘ L is equal to the mass flow on the gaseous side n˘ G and becomes
(
)
(
)
kLkGH cG cG n˘ ) n˘ L ) n˘ G ) A - cL ) KA - cL kL + kGH H H (3) In terms of a specific area density a ) A/VG, the transfer rate is
(
)
(
)
cG n˘ A cG - cL ) Ka - cL )K VG VG H H
(4)
The sum of NH3 and NH4+ concentrations in the liquid phase is denoted as cT:
cT ) cL + cNH4+
(
)
cG ∂ cT ∂cT ) EL 2 - Kaη cL ∂t H ∂x
(6)
The dispersion coefficient EL is much larger than the molecular diffusion coefficient because of the mixing action of the stirrer. Usually the hold-up ratio η is defined as η(x) ) VG(x)/VL(x), because it must be recognized that bubble concentrations vary through the reactor. Because charge neutrality is maintained in the liquid phase, it follows that
cH+ + cNH4+ + cY+ ) cOH- + cX-
(7)
The balance accounts for cations and anions that have been added to the liquid phase to adjust the pH. Using the water equilibrium
KW ) cH+cOH-
(8)
and ammonia equilibrium
KA ) cNH4+/cH+cL
cL )
(9)
cT K Ac H + + 1
(10)
The charge neutrality equation (7) can now be written as
cT KW + cY+ - cX- ) 0 (11) cH+ + KcH+ KcH+ + 1 cH+ This cubic equation in cH+ combines with eqs 1, 6, and 10 to form a coupled differential-algebraic equation system that describes the behavior of the semibatch reactor. Remark: Because the input pH is always greater than 7, the numerical accuracy is better if eq 11 is solved in terms of cOH-. The bubble velocity v and bubble concentration cB still need to be described. A steam balance over the reactor gives
π π ∂ cBv φ3FS,in ∂ cB φ3FS,in ∂cB ∂(cBv) 6 6 + ) + ) 0 (12) ∂t ∂x ∂t ∂x
(
) (
)
because the steam density and bubble diameter remain constant. To minimize steam condensation or evaporation from the liquid phase, the steam is fed at slightly superheated conditions. Two operating pressures have been investigated, 150 and 70 mbar. The feed temperatures of the steam are 60 and 40 °C, respectively, and the corresponding saturation temperatures are 53.97 and 39 °C. Other parameters for these two sets of conditions are listed in Tables 4 and 5. For steady state, one can write
(5)
Dispersion is included in the liquid-phase balance and the total N balance in the liquid phase is 2
one can substitute for cL as follows:
cBv ) cB,invin )
n˘ S,inMH2O π FS,inAT φ3 6
(13)
It is assumed that the steam is immediately dispersed after exiting the feed pipe and the inlet velocity is calculated as follows:
vin )
n˘ S,inMH2O FS,inAT
(14)
The mass of a bubble is mB ) (π/6)φ3cSMH2O, and the equivalent mass of liquid it displaces is mL ) (π/6)φ3FH2O. The dominant terms that determine the bubble dynamics are buoyancy, the force on an accelerating body, and drag.14 Furthermore, bubble-bubble interaction is neglected.
(m
B
1 π + mL v˘ ) (mL - mB)g - FLv|v|φ2Ψ(v) 2 8
)
(15)
Because φ is constant, eq 15 can be integrated a priori to determine the total height, residence time, and v(x). The drag coefficient Ψ is dependent on v and can be described as functions of Re in certain intervals.14 The calculated velocities v(x) as well as the number of bubbles per total volume cB(x) are illustrated versus the normalized total height of the mixture in an example (Figure 3).
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Figure 3. Velocity v and bubble concentration cB versus height x of the mixture.
The bubble concentration follows directly from eq 13. The presence of the gas phase changes the total column height xT as follows:
xT ) xL + xG )
VL + AT
∫0x cBπ6φ3 dx ) T
VL π 3 + φc v AT 6 B,in in
dx ∫0x v(x) T
(16)
Because ∫x0Tdx/v(x) ) ∫τ0Gdt, eq 16 can be written as
xT )
VL n˘ S,inMH2O + τ AT FS,inAT G
(17) Ka )
The residence time of the steam follows directly from eq 17:
(
τG )
)
VL F A AT S,in T n˘ S,inMH2O
xT -
(18)
The hold-up ratio can now be expressed in terms of the bubble diameter as follows:
(19)
V
∫0x η(x) dx ) VGL T
(20)
is the average hold-up ratio that will be used in the perfectly mixed model. The mass-transfer coefficient still needs to be determined. To this end, we use the Fro¨ssling equation (eq 21), which is valid for Re > 100.15 This is fulfilled over the whole velocity range in each considered case.
Sh ) 2 + 0.552Re1/2Sc1/3 ) 365
φ
φ
) 0.05 s-1
(22)
∂cG 1 ≈ (cG - cG,in) ∂x τG
(23)
where τG is given by eq 18. Thus, the perfectly mixed model becomes
( )
)
cG dcG 1 ) - (cG - cG,in) + Ka cL dt τG H
because cB(x) is known.
1 xT
(Sh)DNH3,L 6
Equations 1 and 6 constitute the finite dispersion/plug flow model for the semibatch reactor, and in the rest of this study we will refer to it as the FDPF model. If it is assumed that the liquid phase is perfectly mixed, the ideal mixed state is obtained: limELf∞ EL (∂cT/∂x) ) 0 and the convection term in the gas phase NH3 balance simplifies as follows:
v
π cB φ3 6 η(x) ) π 3 1 - φ cB 6
η j)
Sh ) 2.0 + 0.6Re1/2Sc1/3. This correlation is applicable between a very small Reynolds number and a maximum of 1000. However, mass transfer to drops or bubbles is further complicated by internal circulation currents that lead to enhanced mass transfer:21 Sh ) 1.13Re1/2Sc1/2. In reality the circulation is often not as strong as in the ideal case, and the mass transfer is only slightly larger than that for solid spheres. It must be realized that it is difficult to predict the mass-transfer coefficient accurately. However, eq 21 should provide reasonable, if somewhat conservative, estimates of coefficients for mass transfer for this study. In Table 3 the parameters are listed to calculate the Re and Sc numbers. The bubble diameter that best fits the experimental data is φ ) 0.015 m, and different steam flow rates and liquid volumes did not affect this value. The model is not too sensitive to the value of φ, as is demonstrated by the comparison in Figure 4 between three different values φ ) (0.01, 0.015, 0.02) m. The operating conditions are the same for all three cases: VL ) 2 L, m ) 500 g/h, T ) 60 °C, p ) 150 mbar, pH ) 12. The case of bubble diameter φ ) 0.01 m (solid line) has the largest interfacial transfer area, and ammonia is removed fastest. The dashed line depicts the case φ ) 0.02 m, and it lies above the line for φ ) 0.015 m (dotted line) because removal is slower for larger bubble diameters. The mass-transfer coefficient is calculated as
for {Re} > 100 (21)
Geankoplis20 recommends use of the Fro¨ssling equation for Schmidt numbers between 0.6 and 2.7 and for Reynolds numbers between 1 and 48 000. McCabe et al.21 propose a modification of the Fro¨ssling equation:
(
cG dcT ) -Kaη j cL dt H
(24)
(25)
The system is closed by the algebraic relations (10) and (11), and we will refer to this system as the perfectly mixed model (PM). The hold-up ratio η j is defined by eq 20. In the following section the experimental procedures are described followed by a comparison between the experimental and theoretical results. 4. Experimental Section The removal of ammonia from wastewater is done in a stirred semibatch reactor. The stirrer is necessary to disperse the steam bubbles. The pH, steam flow, liquid volume, and temperature are varied. The liquid is a (NH4)2SO4 solution as a model system which contains approximately 4 g of NH4+-N/L. This concentration is
Ind. Eng. Chem. Res., Vol. 40, No. 15, 2001 3365 Table 1. Ammonia Removal Parameters parameter
definition and value
concentration at the beginning
water equilibrium ammonia equilibrium at room temperature ammonia equilibrium during the removal16 concentration of anions {calculated at pH ) 6 w no addition of NaOH, c(Y+) :) c(Na+) ) 0 mol/L} concentration of sodium for c(X-) ) 0.286 mol/L pH ) 8.7 pH ) 9.7 pH ) 10.7 pH ) 12 for c(X-) ) 0.536 mol/L pH ) 12 pH ) 13.5 distribution coefficient H vapor-liquid-phase equilibrium17
definition and value pR ) 150 and 70 mbar TR ) 60 and 40 °C AT ) 0.02 m2 g ) 9.81 m/s2 EL ) 10-3 m2/s
Table 3. Parameters for Estimation the Mass Transfer Coefficient K‚a
estimated diffusion coefficient of ammonia in the liquid (Wilke-Chang)19 Reynolds number Schmidt number
7.5 g/L ) 0.536 mol/L 14 g/mol
KW ) c(H+) c(OH-) ) 10-14 mol2/L2 KA ) 1/Kacid :) 7 × 109 KA ) 1/Kacid ) 6.4 × 108 at 40 °C KA ) 1/Kacid ) 1.9 × 108 at 60 °C c(X-) ) c(H+) + c(NH4+) + c(Na+) - c(OH-) ) (10-6 + 0.286 + 0-10(14-6)) mol/L ) 0.286 mol/L, and c(X-) ) 0.536 mol/L w c(Na+) ) c(X-) - c(H+) - c(NH4+) + c(OH-) with c(Na+) ) 0.019 11 mol/L w c(Na+) ) 0.1194 mol/L w c(Na+) ) 0.2515 mol/L w c(Na+) ) 0.2940 mol/L
Table 5. Relevant Parameters for Calculations at T ) 40 °C and p ) 70 mbara
parameter
viscosity of liquid18
c0 )
H ) cG/cL ) 2.5 × 10-3 at 40 °C H ) cG/cL ) 2.7 × 10-3 at 60 °C
pressure of the vessel temperature of the vessel cross-sectional area of the vessel gravitational constant dispersion coefficient of the liquid phase
mean bubble velocity bubble diameter density of liquid18
4 g/L ) 0.286 mol/L >14 g/mol
w c(Na+) ) 0.5423 mol/L w c(Na+) ) 0.8521 mol/L
Table 2. Operating, Reactor, and Fluid Dynamics Parameters
parameter
c0 )
m ˘ G (g/h) VL (L) 2.5 2.5 2.5
v0 ) V˙ G/AT (m/s) τG (s) xT (m) η j
50
83
250
0.014 0.25 0.128 0.028
0.024 0.25 0.130 0.048
0.071 0.27 0.144 0.156
a The values of the parameters listed in the second column are read across in the third through fifth columns.
definition and value vj ) 0.6 m/s φ = 0.015 m FL60°C ) 983 kg/m3 FL40°C ) 992 kg/m3 ηL60°C ) 4.67 × 10-4 kg/m‚s ηL40°C ) 6.53 × 10-4 kg/m‚s 60°C DNH ) 5.24 × 10-9 m2/s 3,L 40°C DNH ) 3.52 × 10-9 m2/s 3,L
Re60°C ) vφFL/ηL ) 2.15 × 104 Re40°C ) vφFL/ηL ) 1.41 × 104 Sc60°C ) ηL/DNH3,LFL ) 91 Sc40°C ) ηL/DNH3,LFL ) 187
Table 4. Relevant Parameters for Calculations at T ) 60 °C and p ) 150 mbara m ˘ G (g/h) VL (L) 2 2 2 4 4 4
v0 ) V˙ G/AT (m/s) τG (s) xT (m) η j τG (s) xT (m) η j
250
500
750
0.035
0.071 0.20 0114 0.141 0.38 0.226 0.135
0.106
0.36 0.212 0.064
0.40 0.242 0.214
a The values of the parameters listed in the second column are read across in the third through fifth columns.
consistent with ammonia loading at cattle feedlots. Experiments with wastewater of a dump in Germany showed the same removal behavior at the same conditions.1 Furthermore, washing water of a battery producer containing approximately 7.5 g of NH4+-N/L is used as a real system.
Figure 4. Effect of bubble diameter.
4.1. Structure and Operation of the Process. The semibatch process consists of a liquid batch phase and a continuously fed steam phase. For effective mass transfer, a large phase boundary is needed. This is achieved by injecting steam through a small pipe, and the steam bubbles are further broken down and dispersed by a stirrer. Advantages of the semibatch process are the lower investment costs and the simpler and more adaptable operation compared to continuous processes. The semibatch reactor offers smaller entities the flexibility to do other processing in these reactors at time (several users have indicated to WUZ the desirability of equipment to process different streams).
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Figure 6. Influence of pH on ammonia removal: VL ) 2 L, m ˘ s) 500 g/h, T ) 60 °C, and p ) 150 mbar.
Figure 5. Flowsheet of the semibatch plant.
Furthermore, no baffles are needed which are able to silt up, and the cleaning of the reactor is easy as well. The following schematic (Figure 5) shows the semibatch plant. It consists of a tempered vessel R with a maximum load of the liquid phase of approximately 7 L. The stirrer S promotes and maintains a large interfacial area between the phases as mentioned before. The thermostat TH controls the vessel at the set temperature. The vacuum pump VP generates, with the aid of the controller VR and valve VV, the desired pressure in the plant. The manometer M indicates this pressure. The pump P transports water from the tank VT though the steam generator DE, which produces steam with a temperature of 200 °C and a pressure of 4 bar. The valve V maintains this pressure, which prevents the condensation of the steam already in the vessel. The pump control PR maintains a constant steam flow, which is switched on by valve H. The effective experimental time starts at this moment (t ) 0). After the ammonia has been stripped from the liquid phase in vessel R, the enriched steam is condensed in condenser K. The condensate is collected in container VG. Samples are taken via tap PN. 4.2. Experiments. To characterize the ammonia removal rate, the total concentration cT is determined as a function of time. Over a period of 90 min (180 min, respectively), samples are taken and analyzed at increasing intervals of time, because the concentration changes faster at the beginning. The temperature of the liquid is 60 °C (40 °C, respectively). The pressure in the plant is 150 mbar (70 mbar, respectively). It is important to adapt the pressure to the boiling point of the liquid that is an aqueous system. This prevents steam condensation or water evaporation. The influence of the pH on the removal is examined in the first series of experiments. In the second series the effect of the liquid volume is treated. In the third series the steam mass flow rate is varied. In the fourth series a typical industrial application is considered and different pH values and various steam mass flows are investigated. 5. Results The FDPF model has been discretized using a secondorder central difference scheme. The number of discreti-
zation nodes has been varied, and for the parameters as listed in Tables 1 and 2, it has been found that the difference between 10 and 25 nodes is negligible. For the purpose of the computations, a dispersion coefficient EL value of 10-3 m2/s has been used. By comparison of the results for values between 10-2 and 10-5 m2/s, it has been found that the choice of EL in this interval is not sensitive and differences between the results are negligible. Even EL ) 10-8 m2/s gives realistic results, but the slope of the computed cT vs t curve differs noticeably from the experimental values. This is an indication that there is a lower bound on EL. The initial value problem has been integrated by an explicit method in the FDPF as well as the PM model, varying the time step until changes are negligible (relative error e10-6). 5.3.1. pH Influence. In the first series of experiments, the liquid volume is 2 L and the steam mass flow rate is controlled at 500 g/h. The experimental results are presented in Figure 6 at four different initial pH values by discrete points. They are compared with the FDPF model (dashed lines) and the PM model (solid lines) as functions of time. The removal increases with increased initial pH because of the shift in the equilibrium between ammonia (NH3) and ammonium (NH4+) (cf. Figure 1). The PM model (solid lines) describes the experiments reasonably well. Because perfect mixing is assumed in each phase, the driving forces are smaller than those in reality with the result that the PM curves tend to predict a slower removal. In the FDPF model the varying driving force is accounted for, and this theoretical model is closer to the experimental values. Furthermore, it has been found that the dispersion in the liquid phase plays a negligible role and variations in EL between 10-2 and 10-5 m2/s have no influence on the outlet NH3 concentrations. The FDPF model gives a better description of the bubble column character of the experimental unit than the PM model. 5.3.2. Influence of the Liquid Volume. For evaluation of the influence of the volume of liquid, an additional experiment is performed at a pH of 10.7 and a steam flow rate of 500 g/h (Figure 7). The removal of ammonia from 4 L of liquid takes a little longer because the doubling in VL implies the removal of a larger amount of ammonia. The residence time is longer because the level of liquid is approximately doubled. The remaining concentration in the liquid in both cases is the same in accordance with the equilibrium theory. The FDPF model simulates the experimental data better, and the slower transfer rates in the PM model are consistent with the lower driving force.
Ind. Eng. Chem. Res., Vol. 40, No. 15, 2001 3367
Figure 7. Influence of liquid volume VL on ammonia removal: m ˘ s ) 500 g/h, pH ) 10.7, T ) 60 °C, and p ) 150 mbar.
Figure 9. Influence of pH and the steam mass flow m ˘ s on the ammonia removal: VL ) 2.5 L, T ) 40 °C, and p ) 70 mbar.
result in a substantial reduction in volume. After the steam is condensed, it does not contain solids and other contaminants. In addition, liquid temperatures above 40 °C are impractical, because the heating of the liquid charge would take too much time in an industrial-scale plant. Thus, the pressure has to be adapted to 70 mbar. The differences in the removal rates of ammonia, both in the experiment and in the models, are negligible for pH 12 and 13.5. Despite the temperature and pressure changes (consequently, changes in equilibrium constants, etc.), both the PM and FDPF models fit the experimental data very well. At a temperature of 40 °C and a pressure of 70 mbar, the mass-transfer coefficient changes to Figure 8. Influence of the steam mass flow m ˘ s on the ammonia removal: VL ) 4 L, pH ) 10.7, T ) 60 °C, and p ) 150 mbar.
5.3.3. Influence of the Steam Flow Rate. A pH of 10.7 and a liquid volume of 4 L are selected. At lower steam flow rates, longer processing times are required to remove the same amounts of ammonia. When the steam flow is 0 g/h, it is called flashing. In this case the driving force is maintained just by drawing a vacuum on the reactor. When the steam flow rate is increased, the incremental removal improvement becomes smaller. Consider Figure 8. Because the initial pH values are the same in all cases, the limiting concentrations are the same. The difference in the removal rates is a result of the differences in the driving forces caused by varying numbers of bubbles per total volume cB(x); i.e., the volume of steam and the area for mass transfer are different. In Figure 8 a comparison between the FDPF model, the PM model, and experimental results for different steam flow rates is shown. The FDPF model fits the experimental data very well, and the reader is reminded that the parameters such as the bubble diameter have remained fixed. The results for the PM model lie consistently above the FDPF model and experimental results. 5.3.4. Experiments for an Industrial Application. Various experiments are done to improve the operating efficiency of this process (Figure 9). It follows from the results of section 5.3.1 that complete removal of ammonia can only be achieved if the pH is at least 12. In practical applications, it is necessary to use smaller steam/liquid ratios. Besides from the cost of steam, the purpose of stripping ammonia from wastewater in the first place is to reduce the volume. The steam can contain between 20 and 25% ammonia by weight and
Ka )
(Sh)DNH3,L 6 φ
φ
) 0.035 s-1
(26)
As the steam mass flow decreases, the differences between the PM and FDPF models become larger. Even the more fundamental FDPF model does not fit the data satisfactorily. Thus, the calculations have been repeated for a dispersion coefficient value of 10-8 m2/s. The reduced steam flow rate leads to a decrease in the mixing of the liquid phase. These results fit the experimental results, and a better correspondence is observed between the slopes of the experimental and theoretical results. Remark: To calculate the concentrations cNH4+ and cY+ at the different pH values, the ammonia equilibrium constant KA must be known. KA has an Arrhenius-type temperature dependence, and the preexponent room temperature value is needed. This value is a little smaller in the literature than the chosen parameter in Table 1. The fact that all of the calculations both at 60 °C for the removal from the (NH4)2SO4 solution and at 40 °C for the treatment of a real wastewater by using our choice for KA lead to realistic results justifies the use of the modified value. The (NH4)2SO4 system plus a base like sodium hydroxide is a complex system with several equilibria which is difficult to describe. However, the complexity of a real wastewater system will defy any theoretical attempt to predict KA. 5.4. Conclusions and Further Investigations. A stirred semibatch reactor has been researched for the removal of ammonia from wastewater. It has proved to be a powerful and reliable system that is furthermore simple to operate. This is of particular interest to users who must process small or fluctuating amounts of wastewater with high ammonia concentrations.
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A mathematical model for the presented stirred semibatch reactor has been developed. It has been found that for our steam inlet configuration, the dispersed phase is well described by a bubble diameter of φ ) 0.015 m. The mass-transfer coefficient Ka ) 0.05 s-1 at 60 °C and Ka ) 0.035 s-1 at 40 °C. The FDPF model simulates the experiments very well and is recommended over the PM model for design applications. A scale-up by using the developed mathematical models is a further interesting investigation step for evaluating the model for reactor design. Nomenclature A ) area [m2] a ) specific area [m-1] Bo ) Bodenstein number c ) concentration [mol/m-3] cB ) concentration of steam bubbles [m-3] cG ) concentration of NH3 in the gaseous phase [mol/m-3] cL ) concentration of NH3 in the liquid phase [mol/m-3] cT ) total N concentration in the liquid phase (N ) NH3 + NH4+) [mol/m-3] D ) diffusion coefficient [m2/s] E ) dispersion coefficient [m2/s] g ) gravitational constant [m/s2] H ) distribution coefficient i ) index J ) mole flow density [mol m-2 s-1] K ) total mass-transfer coefficient [m/s] KA ) equilibrium constant of ammonia/ammonium KW ) equilibrium constant of water k ) mass-transfer coefficient [m/s] M ) molecular mass [kg/mol] m ) mass [kg] m ˘ ) mass flow [kg/s] n ) amount of the substance [mol] n˘ ) mole flow [mol/s] p ) pressure [Pa] Sc ) Schmidt number Sh ) Sherwood number T ) temperature [K] t ) time [s] v ) linear velocity [m/s] V ) volume [m3] x ) height [m] y ) length [m] δ ) thin layer thickness [m] φ ) diameter [m] η ) dynamic viscosity [kg m-1 s-1] η ) hold-up ratio F ) density [kg/m3] τ ) residence time [s] Ψ ) drag coefficient Subscripts * ) equilibrium B ) bubble G ) fluid phase G G ) gas in ) in L ) fluid phase L L ) liquid NH4+ ) ammonium S ) steam T ) total T ) tank
X- ) anions Y+ ) cations
Literature Cited (1) Heggemann, M. H. Untersuchungen zur Abreicherung von Ammoniak aus wa¨ssrigen Systemen in verschiedenen Reaktoren. Diplomarbeit, Universita¨t-GH Paderborn, Paderborn, Germany, 1997. (2) Heggemann, M. H. Desodorierung: Mathematisch-mechanistische Modellierung, Experimente. Dissertation in Vorbereitung, Universita¨t Paderborn, Paderborn, Germany, 2001. (3) Rentrop, M.; Gossel, C.; Merrath, S. Ammoniakstrippung in der Wasser- und Abwassertechnik mit Energieru¨ckgewinnungssystem und Erzeugung eines leicht zu vermarktenden Wertstoffes. Forsch. Plan. Betrieb 1993, 18, 29. (4) Gro¨mping, M.; Rautenbach, R.; Kollbach, J.-S. Separate Stickstoffelimination aus dem Prozesswasser der Schlammentwa¨sserung entlastet kommunale Kla¨ranlagen. Wasser, Luft Boden 1997, 41, 34. (5) Rautenbach, R.; Machhammer, O. Aufbereitung Ammoniak enthaltender Abwa¨ssersVerfahrens- und Kostenvergleich. Chem. Ing. Tech. 1988, 60, 23. (6) Beg, S. A.; Obaid-ur-Rehman, S.; Hassan, M. M. Studies of Ammonia Stripping in a Cocurrent Orifice Plate Scrubber. Chem. Ing. Tech. 1989, 61, 338. (7) Jakobs, D. MembranverfahrensStofftransport durch Nanofiltrationsmembranen unter Beru¨cksichtigung von Biofilmen. Dissertation in Vorbereitung, Universita¨t Paderborn, Paderborn, Germany, 2001. (8) Sacarro, G.; Genon, G. High-temperature ammonia stripping and recovery from process liquid wastes. J. Hazard. Mater. 1994, 37, 191. (9) Maier, K.-H. Stickstoffelimination durch thermisches Strippen. Wasser, Luft Boden 1996, 40, 18. (10) Sackewitz, M.; Maier, K.-H. Strippverfahren zur Teilstrombehandlung auf Kla¨ranlagen. Wasser, Luft Boden 1999, 43, 34. (11) Bin, A. K.; Machniewski, P. Removal of trace amounts of volatile pollutants from liquids using stripping to an inert gas. Inz. Chem. Proces. 1998, 19, 371. (12) Fritzmann, J.; Meckl, S.; Wunder, R. Strippen von Ammoniak aus Abwa¨ssern. Verfahrenstech. Mechanisch., Chemischen Biologischen Abwasserbehandlung, Prepr. (GVC-Kongress) 1992, 2, 129. (13) Lewis, W. K.; Whitman, W. G. Principles of Gas Absorption. Ind. Eng. Chem. 1924, 16, 1215. (14) Chow, Ch.-Y. An Introduction to Computational Fluid Mechanics, Wiley: New York, 1979. (15) Brauer, H. Stoffaustausch einschliesslich chemischer Reaktionen; Verlag Sauerla¨nder: Aarau und Frankfurt am Main, Germnay, 1971. (16) Emerson, K.; Russo, R. C.; Lund, R. E.; Thurston, R. V. Aqueous Ammonia Equilibrium Calculations: Effect of pH and Temperature. J. Fish. Res. Board Can. 1975, 32, 2379. (17) Knapp, H.; Do¨ring, R.; Oellrich, L.; Plo¨cker, U.; Prausnitz, J. M. Vapor-Liquid Equilibria for Mixtures of Low Boiling Substances; Chemistry Data Series, Vol. VI; DECHEMA: Frankfurt, 1981. (18) Atkins, P. W. Physikalische Chemie, 2nd ed.; korr. Nachdr. d. 1. Aufl.; VCH Verlagsgesellschaft: Weinheim, Germany, 1990. (19) Jakubith, M. J. F. Grundoperationen und chemische Reaktionstechnik. Einfu¨ hrung in die Technische Chemie; VCH Verlagsgesellschaft: Weinheim, Germany, 1998. (20) Geankoplis, C. J. Transport Processes and Unit Operations; Prentice-Hall: Englewood Cliffs, NJ, 1993. (21) McCabe, W. L.; Smith, J. C.; Harriott, P. Unit Operations of Chemical Engineering; McGraw-Hill: New York, 1993.
Received for review February 2, 2001 Revised manuscript received May 4, 2001 Accepted May 7, 2001 IE010106I
